ORNL-4676.txt

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ORNL-4676
UC-80 — Reactor Technology

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MOLTEN-SALT REACTOR PROGRAM
SEMIANNUAL PROGRESS REPORT
FOR PERIOD ENDING FEBRUARY 28, 1971

OAK RIDGE NATIONAL LABORATORY
operated by

UNION CARBIDE CORPORATION
for the
U.S. ATOMIC ENERGY COMMISSION

Printed in the United States of America. Available from
National Technical Information Service
US. Department of Commerce
5285 Port Royal Road, Springfield, Virginia 22151
Price: Printed Copy $3.00; Microfiche $0.95

This report was prepared as an account of work sponsored by the United
States Government. Neither the United States nor the United States Atomic
Energy Commission, nor any of their employees, nor any of their contractors,
subcontractors, or their employees, makes any warranty, express or implied, or
assumes any legal liability or responsibility for the accuracy, completeness or
usefulness of any information, apparatus, product or process disclosed, or
represents that its use would not infringe privately owned rights.

Contract No. W-7405-eng-26

MOLTEN-SALT REACTOR PROGRAM
SEMIANNUAL PROGRESS REPORT
For Period Ending February 28, 1971

- M. W. Rosenthal, Program Director
R. B. Briggs, Associate Director
P. N. Haubenreich, Associate Director

AUGUST 1971

OAK RIDGE NATIONAL LABORATORY
Qak Ridge, Tennessece
operated by
UNION CARBIDE CORPORATION
for the
U.S. ATOMIC ENERGY COMMISSION

ORNL-4676

MARIETTA ENERGY SYSTEMS LIBRARI

AN

3 445k 0428254 O

ORNL-2474
ORNL-2626
ORNL-2684
ORNL-2723
ORNL-2799
ORNL-2890
ORNL-2973
ORNL-3014
ORNL-3122
ORNL-3215
ORNL-3282
ORNL-3369
ORNL-3419
ORNL-3529
ORNL-3626
ORNL-3708
ORNL-3812
ORNL-3872
ORNL-3936
ORNL4037
ORNL4119
ORNL4191
ORNL-4254
ORNL4344
ORNL4396
ORNL-4449
ORNL4548
ORNL-4622

This report is one of a series of periodic reports in which we describe the progress of the program. Other reports
issued in this series are listed below.

Period Ending January 31, 1958
Period Ending October 31, 1958
Period Ending January 31, 1959
Period Ending April 30, 1959
Period Ending July 31, 1959
Period Ending October 31, 1959

Periods Ending January 31 and April 30, 1960

Period Ending July 31, 1960
Period Ending February 28, 1961
Period Ending August 31, 1961
Period Ending February 28, 1962
Period Ending August 31, 1962
Period Ending January 31, 1963
Period Ending July 31, 1963
Period Ending January 31, 1964
Period Ending July 31, 1964
Period Ending February 28, 1965
Period Ending August 31, 1965
Period Ending February 28, 1966
Period Ending August 31, 1966
Period Ending February 28, 1967
Period Ending August 31, 1967
Period Ending February 29, 1968
Period Ending August 31, 1968
Period Ending February 28, 1969
Period Ending August 31, 1969
Period Ending February 28, 1970
Period Ending August 31, 1970
Contents

INT RODUCTION .ottt et et et ettt e ettt ittt ie e ee s ix

QUMM AR Y . .o e e e e e e X1

PART 1. MOLTEN-SALT REACTOR EXPERIMENT

1. POSTOPERATION EXAMINATIONS . ... ...\ uttttneteeneneet et e ieaeaaes I
1.1 Outline of Program . .. .. ... i e 1
1.2 Reactor Vessel and Core .......... ST e 2
1.3 Fuel PUmp ..o et 6
1.4 Heat EXChanger .. ... ... . e e e ettt e 11
1.5 Leak at Freeze Valve FV-105 . .. . i e e e i et e 12
1.6 Other EXaminations ... ... .. ... ittt itiae iy e 12
1.7 Evaluation of Tools and Procedures .......... . ... i 13
2. FURTHER INVESTIGATIONS .. ... .. i P 17
2.1 Test of Coolant Salt Flowmeter and Conclusions . ............ .. . ... 17
2.2 Inventories of Residual Uranium and Plutonium .......... ... . .. . i 18
2.3 Search for Unrecovered 235U ..o en ettt e 20

PART 2. MSBR DESIGN AND DEVELOPMENT

3. DESIGN .. e e e e e e e 21
3.1 Single-Fluid 1000-MW(e) MSBR Design Study Report ... ... ]
3.2 Molten-Salt Demonstration Reactor Design Study ........ ... .. .. .. ... i 22

320 Introduction . ... . . s e 22
3.2.2 Addition of Third Salt-Circulation Loop ........ ... it 22
3.2.3 Salt Overflow and Gas Stripping Systems . . .. .. ... i 24
3.2.4 Primary Drain Tank . ... ... o e e 24
3.2.5 Drain Valves for Salt Service . .......... .. ... .. il 25
3.2.6 Heat Exchangers . . .. ... ..o i e e 25
3.2.7 Building and Containment . ........ ... . it e 28
3.3 Initial Temperature Transients in Empty MSBR “Reference Design” MSBR Heat Exchangers . .. . .. 32
3.4 The Consequences of Tubing Failure in the MSBR Heat Exchanger ......................... 34
3.5 Tritium Distribution and Control inthe MSBR ....... ... ... .. ... .. . i i 35
3.6 Industrial Study of 1000-MW(e) Molten-Salt Breeder Reactor ............. ... ............. 36

111
iv

37 MSBE Design . ... et 36
370 General . ... e e e e e e 36

372 MSBE Core Design . ...t e e e e 36

3.7.3 MSBE Primary Heat Exchanger Design . ............ ... . ... ... .. . .. 39

4. REACTOR PHY SICS ..ot ettt e et e e et it 41
4.1 Physics Analysis of MSBR . ... ... e 41
4.1.1 Single-Fluid MSBR Reference Design .. ..... ... ... ..o 41

4.1.2 Fixed-Moderator Molten-Salt Reactor .......... ... ... .. . . i 43

4.2 MSR Experimental Physics . . .. A 45
4.2.1 HTLTR Lattice Experiments . ......... ...ttt iieiiieaieieeenens 45

5. SYSTEMS AND COMPONENTS DEVELOPMENT ........... PN 49
5.1 Gaseous Fission Product Removal ................ e e e e e e e e 49
5.1.1 Gas Separator and Bubble Generator ....... ... ... ... ... . ... . . i 49

5.2 Gas System Test Facility ......... ... e 51
5.3 Molten-Salt Steam Generator . .... PR B 51
5.3.1 Steam Generator Industrial Program . ........ .. ... .. ool 52

5.3.2 Steam Generator Tube Test Stand (STTS) ... ... ... i i 52

5.3.3 Molten-Salt Steam Generator Technology Facility (SGTF) .......... .. ... ... ... ... 52

5.3.4 Development Bases for Steam Generators Using Molten Salt as the Heat Source .......... 52

5.4 Sodium Fluoroborate Test Loop .......... ... .. ... e 52
54.1 Pump Bowl and Rotary Element . ... ... . ... ... . .. . . 53

5.4.2 BF; Feed and Salt Level Bubbler Tube ... .. ... . . . ... . ... . i i, 55

5.5 -Coolant Salt Technology Facility ............ .. i i 57
56 MSBRPumps ............... .. .. ...... e e e e e e e 58
5.6.1 MSREMark2 Fuel Pump .. ... ... ... . . i ... 58

5.6.2 MSRE Salt Pump Inspection .......... .. .. . . e 58

5.6.3 ALPHA Pump ... ..o e e e 59

5.7 Remote Welding ... .... P 59
5.7.1 Automatic Controls . ... ... ... .t e e e 59

5.7.2 Pipe Cleaning Tests . . . .. ...ttt e 60

6. MSBR INSTRUMENTATION AND CONTROLS . ... .. e it ciee e ... 6l
6.1 Development of a Hybrid Computer Simulation Model of the MSBR System .................. 61
6.1.1 Introduction . ...... ... . . i e e 61

6.1.2 Steam Generator Model .. ... ... . . e 61

7. HEAT AND MASS TRANSFER AND THERMOPHYSICAL PROPERTIES ........................ 64
7.1 Heat Transfer . .. ... o 64
7.2 Thermophysical Properties .. ... ... ... 67
7.2.1 Wetting Studies ............ ... ... . ... ..., e S 67

7.2.2 Thermal Conductivity . ... ... ... .. e 69

7.3 Mass Transfer to Circulating Bubbles .. ... ... . .. ... . . .. .. . . 69
7.3.1 EXPeriment . .. ...ttt e e 69

7.32 Theory . ... e 70
8. FISSION PRODUCT BEHAVIOR .. ................ S 73
8.1 Determination of Tritium and Hydrogen Concentrations in MSRE ‘Pump Bowl! Gas ............. 73
8.1.1 Calibration Apparatus ..............uremimmneinnnannnnn. PP e 74
8.1.2 Analysis for Hydrogen .......................... e e e e 75
8.1.3 Tritium Diffusion-Studies . . .. ... .. e 76
8.2 Examination of Deposits from the Mist Shield in the MSRE Fuel Pump Bow! ......... e 76
B8.2.1 TritiUIN . vttt it et e e e e e e e e e e e 82
8.3 Synthesis of Niobium Fluorides . ............ ... ... .. ... ... e 85
84 Reaction Kinetics of Molybdenum and Niobium Fluoride in Molten Li, BeF, Solutions . ......... 85
8.5 Mass Spectroscopy of Niobium Fluorides . ... ... ... . i 86
9. COOLANT SALT CHEMISTRY AND TRITIUM CONTROL .. ... .. ... i 88
9.1 Studies of Hydrogen Evolution and Tritium Exchange in Fluoroborate Coolant . .. .............. 88
9.2 Reaction of Sodium Fluoride—Sodium Tetrafluoroborate with Water ........................ 90
9.3 Identification of Corrosion Products in the Bubbler Tube of the Fluoroborate Test Loop ... .. A
9.4 Mass Spectroscopy of Fluoroborate MSR Coolants . ...... ... .. .. ... ... ... . o, 93
9.5 Spectroscopic Investigations of Hydrogen- and Deuterium-Containing Impurities :
in NaBF, and NaF-NaBF, Eutectics ... ... ... . i i i 94
9.6 Raman Spectra of the High-Temperature Phase of Polycrystalline NaBF, ..................... 96 |
9.7 A New Method for Synthesis of NaBF;0H ... ... . ... 98
9.8 Solubility of BF; Gasin Fluoride Melts . . .. ... ... ... . i i 98
9.9 Equilibrium Phase Relationships in the System RbF-RbBF, .. .. U 100
9.10 Activities in Alkali Fluoride — Fluoroborate Mixtures . . . ... ... ... .. ..o i ... 100
9.11 High-Temperature Crystal Structure and Volume of Sodium Tetrafluoroborate
and Related Compounds . ...t e 101
9.12 Hydrogen Permeation through Oxide-Coated MELalS .« v vee e e e e e 103
10. PHYSICAL CHEMISTRY OF MOLTEN SALTS .. ... .. .. i PR 107
10.1 Thermodynamics of LiF-BeF, Mixtures from EMF Measurements of Concentration Cells .. ... ... 107
10.2 Equilibrium Phase Relationships in the System LiF-BeF,-CeF; ........... ... ... ... . 109
10.3 Electrical Conductivity of Molten and Supercooled NaF-BeF, (40-60!'M01e_ V) P 109
10.4 Glass Transition Temperatures in the NaF-BeF, System .............. N e 110
10.5 Raman Spectra of BeF, 2~ in Molten LiF and NaF to 686°C .. ......... ... . ... ........... 112
10.6 Bubble Formation by Impingement of a Jet Stream on a Fluid Surface .................... ... 115
107 The Solubility of Hydrogen in Molten Salt . . . . .. ... .. i, . 115
108 Enthalpy of UF, from 298 to 1400°K ... ... .. .....oiieeaianies, e 117
109 Absorption Spectroscopy of Molten Fluorides: The Disproportionation
Equilibrium of UF3 Solutions .. ... i e 118
10.10 The Oxide Chemistry of Pa** in the MSBR Fuel Solvent Salt .. ... ........vuerinrenennnn... 119
10.11 The Redox Potential of Protactinium in MSBR Fuel Solvent Salt ................. e 120

PART 3. CHEMISTRY
11.

12.

13.

14.

vi

10.12 The Crystal Structures of Complex Fluorides . . ... ... ... .. i e 122

10.12.1 The Crystal Structure of CsUgFa5 ... ..o 122
10.12.2 The Crystal Structure of a-KThgF,s . ... ..o ol 122
10.12.3 The Crystal Structure of Li,MoFg . .. ... ..o e 122
10.12.4 The Crystal Structure of RbTh3F,3 ... ... i I 124
10.12.5 The Crystal Structure of RbgZraFay ..o 125
10.12.6 DiSCUSSION ..t tv vt ee i ee et e e ee et e e e e, 125
10.13 Noncrystalline BeF, at 25°C: Structure and Vibrational Motion . ......................... .. 125
10.14 Relationship between Entropy and Sonic Velocity in Molten Salts .......................... 126
10.15 A Reference Electrode System for Use in Fluoride Melts . ............. ... ... .. ... .. .. .. 127
CHEMISTRY OF MOLTEN-SALT REACTOR FUEL TECHNOLOGY ...........iiiiiiiiiinannn, 129
11.1 Extraction of Rubidium and Cesium from MSBR Fuel Solvent into Bismuth by :
Reduction with Lithiumat 650°C ... ... ... i, S 129
112 Distribution of Thorium between MSBR Fuel Solvent and Bismuth '
Saturated with Nickel and Thorium at 650°C ... ... ... ... . ... e, 130
11.3 Bismuth-Manganese Alloys as Extractants for Rare Earths from
MSBR Fuel SOIVENt . . .ottt et e et e e et e e 131
11.4 Removal of Solutes from Bismuth by Fractional Crystallization ............................ 131

DEVELOPMENT AND EVALUATION OF ANALYTICAL METHODS FOR

MOLTEN-SALT REACTORS ....... e e e e L. 134
12.1 Electroanalytical Studies of Titanium(IV) in Molten
LiF-NaF-KF (46.5-11.542.0Mole 20) . ... oot e ettt eeeeans 134
12.2 Reference Electrode Studiesin Molten NaBF, . . ... ... . . i 135
12.3  Electrochemical-Spectral Studies of Molten Fluoride Salt Solutions ......................... 136
12.4 Spectral Studies of Molten Fluoride Salts . .. ......... ... ... . . . i 136
12.5 Analytical Studies of the NaBF, Coolant Salt . ....... ... ... ... . . . . . . . ... ... . ..., 137
12.6 In-Line Chemical Analysis .. ... ... ...ttt ittt ettt eaannns 138

PART 4. MATERIALS DEVELOPMENT

EXAMINATION OF MSRE COMPONENT S . .. ..ot e et et e e .. 139

13.1 Examination of a Graphite Moderator Element ........ ... .. .. ... .. ... . ... . ... 139
13.2 Auger Analysis of the Surface Layer on Graphite Removed from the

Core of the MSRE .. .. ... e e e 143
13.2.1 Auger Electron SpectrosCopY . ..ottt ittt ettt et 144
13.2.2 Results and Discussion .. ... ... ...ttt ittt et 145
" 13.3 Examination of Hastelloy N Control Rod Thimble .................cc.uuun... e 147
13.4 Examination of the Sampler Assembly . .. .. ... .. . 150
13.5 Examination of a Copper Sample Capsule ....... ... .. . . i 154
13.6 Examination of the Primary Heat Exchanger .......... ... ... .. . . .. 156
13.7 Examination of Freeze Valve 105 .. ... .. .. .. . i e 160
GRAPHITESTUDIES .................... e e PP 167

14.1 Graphite Irradiations in HFIR .. .. .. .. .. 167
15.

16.

17.

vii

14.2  Graphite Fabrication ............ e e e e e 169

14.3 Graphite Development — Chemistry .. ... ot i 170
144 Graphitization Study of a Lampblack-Pitch Carbon .. ............. e 171
14.5 Reduction of Graphite Permeability by Pyrolytic Carbon Sealing ........... [ .. 173
14.6 Fundamental Studies of Radiation Damage Mechanisms in Graphite ......................... 174
14.7 Lattice Dynamics of Graphite ...........c.. i i it 176
HASTELLOY N o et e et ettt e 179
15.1 Status of Laboratory Heat Postirradiation Evaluation .............. ... .. ... ... i ... 179
15.2 Postirradiation Creep Testing of Hastelloy N ............ e e 180
15.3 The Unirradiated Mechanical Properties of Several Modified Commercial Alloys . ............... 181
154 The Weldability of Several Modified Commercial Alloys ........... ... i 185
15.5 Postirradiation Properties of Several Commercial Alloys ......... ... .. . oo 188
15.6 Status of Development of a Titanium-Modified Hastelloy N ......... e 192
157 COIToSion STUGIES . . o v ot et e e e ettt et ettt e e e e 192

15.7.0 Fuel Salts ..o oii ittt ettt e et sttt e e 194

15.7.2 Fertile-Fissile Salt . . . .. P T R RETRES 196

15.7.3 BlanKet Salt . .. v oottt it i e 196

1574 Coolant Salt .. ...t e EERTTER. 196

15.7.5 Analysis of H, O Impurities in Fluoroborate Salts ............ ... oo 197
15.8 Forced-Convection Loop Corrosion Studies ............... e 202

15.8.1 Operation of Forced-Convection Loop MSR-FCL-1 ....... ... ...t 202

15.8.2 Metallurgical Analysis of MSR-FCL-1 .. ... ... .. i 204

15.8.3 Forced-Convection Loop MSR-FCL-2 ........ e e 206
159 Retention of Tritium by Sodium Fluoroborate ....... ... ... i 210
15.10 Support for Components Development Program ........ ... ... oot 211

15.10.1 Metallurgical Examination of Inconel Bubbler Tube from PKP-1 Pump Loop .. ......... 211
15.11 Corrosion of Hastelloy Nin Steam . .. .. .ottt ittt et am e 216
SUPPORT FOR CHEMICAL PROCESSING . ... ... e 218
16.1 Construction of a Molybdenum Reductive-Extractive Test Stand . .......... ... .. .. ... ... 212§
16.2 Fabrication Development of Molybdenum Components ............ ..ot 219
16.3 Welding Molybdenum . . ......... ..o .. e R 220
16.4 Development of Bismuth-Resistant Filler Metals for Brazing Molybdenum .................... 221
16.5 Compatibility of Materials with Bismuth ......... ... ... ... i 225
16.6 Chemically Vapor Deposited Coatings ........... oo, e 231
16.7 Molybdenum Deposition fromMoFg ... .. .o i 232

PART 5. MOLTEN-SALT PROCESSING AND PREPARATION

FLOWSHEET AN ALY SIS ..ttt et ittt e it et 235
17.1 Protactinium Isolation Using Fluorination and Reductive Extraction ........................ 235

17.2 Combination of Discard Streams from the Protactinium Isolation System
and the Metal Transfer System ........... ... i 237
18.

19.

20.

viii

17.3 Protactinium Isolation Using Oxide Precipitation .. . ... e e e 237
17.4 Stripping of Rare-Earth Fission Products from LiCl in the

Metal Transfer SyStem . . ..ottt et e e e e e e 240
17.5 Importance of Uranium Inventory in an MSBR Processing Plant . ........................... 240
PROCESSING CHEMISTRY . ... ..ot e e i e e 242
18.1 Measurement of Distribution Coefficients in Molten-Salt—Metal Systems ..................... 242
18.2 Solubilities of Thorium and Neodymiurh in Lithium-Bismuth Solutions ............. e L. 244
18.3  Oxide Precipitation Studies ............ ... ..o e 245
ENGINEERING DEVELOPMENT OF PROCESSING OPERATIONS ........ ... ... ... .. ... ..... . 249
19.1 Engineering Studies of the Metal Transfer Process for Rare-Earth Removal .................... 249
19.2 Design of the Third Metal Transfer Experiment .. .......... ... .. .. ... .., 254
19.3 Development of Mechanically Agitated Salt-Metal Contactors ............. ... ... . ... ... 255
19.4 - Reductive Extraction Engineering Studies .. ... e 256
19.5 Contactor Development: Pressure Drop, Holdup, and Flooding in

Packed Columns ........... ... ... .. ... ... ... .. e e h s e e 260
19.6 Development of a Frozen-Wall Fluorinator ...... ... ... ... ... . . i i, 262
19.7 Estimated Corrosion Rates in Continuous Fluorinators . ......... ... ... ... ... ... ... .. ... 264
19.8 Axial Dispersion in Simulated Continuous Fluorinators . ....... PP 265
19.9 Engineering Studies of Uranium Removal by Oxide Precipitation ........................... 267
19.10 Design of a Processing Materials Test Stand and the Molybdenum

Reductive Extraction Equipment ... .. ... ... . .. . . e 267
CONTINUOUS SALT PURIFICATION SYSTEM ... ................ JE 269

ORGANIZATIONAL CHART .. .. e i ettt et aaan 271
Introduction

The objective of the Molten-Salt Reactor Program is
the development of nuclear reactors which use fluid
fuels that are solutions of fissile and fertile materials in
suitable carrier salts. The program is an outgrowth of
the effort begun over 20 years ago in the Aircraft
Nuclear Propulsion program to make a molten-salt
reactor power plant for aircraft. A molten-salt reactor —
the Aircraft Reactor Experiment — was operated at
ORNL in 1954 as part of the ANP program.

Our major goal now is to achieve a thermal breeder
reactor that will produce power at low cost while
simultaneously conserving and extending the nation’s
fuel resources. Fuel for this type of reactor would be
2331JF, dissolved in a salt that is a mixture of LiF and
BeF,, but it could be started up with 23°U or
plutonium. The fertile material would be ThF, dis-
solved in the same salt or in a separate blanket salt of
similar composition. The technology being developed
for the breeder is also applicable to high-performance
converter reactors.

A major program activity through 1969 was the
operation of the Molten-Salt Reactor Experiment. This
reactor was built to test the types of fuels and materials
that would be used in thermal breeder and converter
reactors and to provide experience with operation and
maintenance. The MSRE operated - at 1200°F and
produced 7.3 MW of heat. The initial fuel contained 0.9
mole % UF,, 5% ZrF,, 29% BeF,, and 65% " LiF; the
uranium was about 33% 233U. The fuel circulated
through a reactor vessel and an external pump and heat
exchange system. Heat produced in the reactor was
transferred to a coolant salt, and the coolant salt was
pumped through a radiator to dissipate the heat to the
atmosphere. All this equipment was constructed of
Hastelloy N, a nickel-molybdenum-iron-chromium
alloy. The reactor core ‘contained an assembly of
graphite moderator bars that were in direct contact
with the fuel.

Design of the MSRE started in 1960, fabrication of
equipment began in 1962, and the reactor was taken
critical on June 1, 1965. Operation at low power began
in January 1966, and sustained power operation was
begun in December. One run continued for six months,
until terminated on schedule in March 1968.

ix

Completion of this six-month run brought to a close
the first phase of MSRE operation, in which the
objective was to demonstrate on a small scale the
attractive features and technical feasibility of these
svstems for civilian power reactors.

We concluded that this objective had been achieved
and that the MSRE had shown that molten-fluoride
reactors can be operated at 1200°F without corrosive
attack on either the metal or graphite parts of the
system, that the fuel is stable, that reactor equipment
can operate satisfactorily at these conditions, that
xenon can be removed rapidly from molten salts, and
that, when necessary, the radioactive equipment can be
repaired or replaced.

The second phase of MSRE operation began in
August 1968, when a small facility in the MSRE
building was used to remove the original uranium
charge from the fuel salt by treatment with gaseous F,.
In six days of fluorination, 221 kg of uranium was
removed from the molten salt and loaded onto ab-
sorbers filled with sodium fluoride pellets. The decon-
tamination and recovery of the uranium were very
good.

After the fuel was processed, a charge of 23U was
added to the original carrier salt, and in October 1968
the MSRE became the world’s first reactor to operate
on 233U. The nuclear characteristics of the MSRE with
the 233U were close to the predictions, and the reactor
was quite stable.

In September 1969, small amounts of PuF; were
added to the fuel to obtain some experience with
plutonium in a molten-salt reactor. The MSRE was shut
down permanently December 12, 1969, so that the
funds supporting its operation could be used elsewhere
in the research and development program.

Most of the Molten-Salt Reactor Program is now
devoted to the technology needed for future molten-
salt reactors. The program includes conceptual design
studies and work on materials, on the chemistry of fuel
and coolant salts, on fission product behavior, on
processing methods, and on the development of com-
ponents and systems.

Because of limitations on the chemical processing
methods available at the time, until three years ago
most of our work on breeder reactors was aimed at
two-fluid systems in which graphite tubes would be
used to separate uranium-bearing fuel salts from tho-
rium-bearing fertile salts. In late 1967, however, a
one-fluid breeder became feasible because of the de-
velopment of processes that use liquid bismuth to
isolate protactinium and remove rare earths from a salt
that also contains thorium. Our studies showed that a
one-fluid breeder based on these processes can have fuel

utilization characteristics approaching those of our
two-fluid designs. Since the graphite serves only as
moderator, the one-fluid reactor is more nearly a
scaleup of the MSRE. These advantages caused us to
change the emphasis of our program from the two-fluid
to the one-fluid breeder; most of our design and
development effort is now directed to the one-fluid
system.
Summary -

PART 1. MOLTEN-SALT REACTOR EXPERIMENT

1. Postoperation Examinations

A limited program of postoperation examinations was
completed during this period. The work on the radio-
active systems was done through the maintenance shield
using tools specially developed and tested in mockups.
The control rods, rod thimbles, and one graphite
moderator element were removed, and the interior of
the reactor vessel was viewed. The interior of the fuel
pump bowl was also viewed through a hole left by
excision of the sampler cage. A section of the primary
heat exchanger shell was cut out, and portions of six
tubes were removed. The salt leak that occurred during
the final shutdown was located at a freeze valve and was
cut out for inspection. The tools and procedures
worked well, and conditions in the reactor were found
to be generally very good.

2. Further Investigations

The differential-pressure system on the coolant salt
flowmeter was tested and found to be in error by 6.7%.
Correction of all known errors brings the heat-balance
value for full power down to 7.65 MW.

Less uranium was recovered from the UF, absorbers
than had been expected on the basis of salt inventory
measurements at the time of the fuel processing in
1968. Careful review of all evidence indicates that
about 2.6 kg of uranium (33% 23°U) was left some-
where in the MSRE processing plant. A uranium search
procedure based on neutron interrogation of closed
vessels was tested but proved insufficiently sensitive.

PART 2. MSBR DESIGN AND DEVELOPMENT

3. Design

The comprehensive report on the conceptual design
of a 1000-MW(e) single-fluid MSBR power station has
received final editing and is scheduled for distribution
in June 1971. | |

Exploratory design and evaluation studies of a
300-MW(e) molten-salt demonstration reactor (MSDR)

xi

were continued. In these studies, the design conditions
are made less stringent than those in our MSBR
reference design. For example, the reactor would
operate as a converter so as not to await demonstration

‘of the advanced processing system being designed for

the breeder, and the power density would be reduced so
that the graphite core would have a life of 30 years. The
graphite would not have to be sealed to reduce’the
permeation by xenon, and rapid fuel reprocessing
would not be necessary. By substituting periodic salt
replacement for continuous fuel processing the reactor
could operate with a conversion ratio of about 0.8 until
the chemical plant was fully developed.

The MSDR reactor vessel design has not been revised
since last reported, but the general flowsheet, the drain
tank, the primary heat exchangers, steam generators
and reheaters, and the cells and building have all been
changed in an investigation of a salt-circulation loop
interposed between the secondary system and the steam
system or to otherwise modify the system parameters.
The heat transport fluid used in the third loop would be
a nitrate-nitrite mixture which would form water of any
tritium diffusing into it from the coolant salt and would
thus block escape of tritium into the steam system. Use
of the nitrate-nitrite salt also makes it possible to
construct ‘the steam generators and reheaters of less
expensive materials and to deliver feedwater and reheat
steam to the boilers and reheaters at conventional
temperatures rather than at the abnormally high values
specified for the MSBR. The fluid used to transport
heat from the primary heat exchangers to the secondary
exchangers would be changed from the previously
proposed sodium fluoroborate to ’LiF-BeF,, a salt
used successfully in the MSRE and one which poses few
problems if it were to leak into the fuel salt. ,

In the MSDR the drain tank is not used as an
overflow volume for the pump bowls, and only one

‘small pump is used to transfer salt from the tank when

filling the primary system. The sump tanks of the main
circulation pumps now provide the surge volume. A

wvalve in which a thin film of salt is frozen between 2

movable poppet and the seat to effect the final
leak-tight closure is proposed for use in the reactor
drain line. The valve in the drain line from the cell catch
basin would be sealed with a membrane that would be
ruptured in the unlikely event that the line is needed.

The primary and secondary heat exchangers were
redesigned to use the new secondary and tertiary salts
and to account for lowering the fuel salt temperature
from the reactor outlet to 1250°F. As could be
expected, the exchangers have more surface than those
in previous concepts.

The design of the heated equipment cells was changed
to incorporate water cooling of the cell walls. The
method of heating the cells was changed from use of
radiant electrical heaters to circulation of hot nitrogen
gas. The gas can also be cooled and used to cool some
equipment.

The temperature transients following shutdown of the
reactor and draining of the primary and secondary salts
were calculated for heat exchangers of the design
proposed for the MSBR reference plant. Decay of
fission products on the metal surfaces provides the heat
source, and the heat is radiated to the surroundings,
which are at 1000°F. It is conservatively estimated that
the maximum temperatures would reach 2150 and
1850°F, respectively, in heat exchangers of 563- and
141-MW capacity at about 2.7 hr after shutdown.

An analysis was started to determine the conse-
quences of the mixing of primary and secondary salts
that would result from the rupture of a tube in a
primary heat exchanger of an MSBR. :

Additional studies were made of methods for keeping
small the amount of tritium that reaches the steam
system in an MSBR plant; 0.2% or less would reach the
steamn if essentially all the tritium and tritium fluoride
were stripped from a side steam of 10% of the
circulating fuel salt flow in a countercurrent contactor.
Use of helium containing a small volume percentage of
water vapor as the coolant in the secondary system in
place of the sodium fluoroborate salt would inhibit the
transport to the steam. Continuous addition and removal
of hydrogen fluoride in the sodium fluoroborate in the
secondary system would be effective in reducing the
transport of tritium to the steam if the rate were more
than 507 times the tritium production rate and the
hydrogen fluoride did not react rapidly with the metal
walls.

The Ebasco Services group, consisting of Ebasco
Services, Continental QOil, Babcock and Wilcox, Cabot,
Union Carbide, and Byron-Jackson companies, was
selected to perform the industrial design study of a
1000-MW(e) MSBR plant.

A report was issued that outlined the objectives and
design bases of the MSBE and provided a brief
description of a reference reactor. Our reference reactor
has a graphite core 45 in. in diameter and 57 in. high

X11

containing 15 vol % salt. The core is centered in a
7.5-ft-ID spherical vessel, with the space between the
graphite and the vessel wall filled with fuel salt. The
design power is 150 MW(t), and the start of life
breeding ratio is 0.96. ,

We continued design studies to determine the prob-
lem areas and to evaluate possible solutions. Major
emphasis was on maintenance of the core graphite and
the primary heat exchanger. We also prepared new
layouts of the cell and primary system to help indicate
how the problems would be handled in the different
configurations.

Alternate core moderator element configurations
were investigated. A cylindrical element design appears
attractive except that it requires a 20.5% salt fraction,
as compared with 15% for the prismatic element. Since
the elements do not interlock, this concept lends itself
to removal of individual elements by a handling
machine. -

We sized the primary heat exchanger, holding the
tube length constant at the 28 ft proposed for the
MSBR. This design, with salt on the tube side, utilizes
1340 tubes *% in. in diameter in a 31.5-in.-ID shell.

4. Reactor Physics

Calculations of the neutronic performance of the
reference single-fluid MSBR have been brought up to
date by modifying fission product removal rates to
conform to the recently adopted metal transfer process.
In addition, a few minor data corrections and cross
section revisions were included in these calculations.
The results indicate a slightly higher breeding ratio for
the new processing scheme (1.071 as compared with
1.063) and a very slightly lower fissile inventory (1487
kg as compared with 1504 kg). The most important

differences in the neutron balance are the absence of -

absorptions in plutonium, with the new process, re-
duced absorptions in fission products, and a higher
value of 7, because a higher proportion of the fissile
material is 233 U,

Studies of the possible performance of a 1000-MW(e)
molten-salt converter reactor have been. continued.
Some recent calculations were based on the core design
of a “permanent-core” MSBR, that is, one whose peak"
power density is low enough (i.e., 9 W/cm?) to permit
the graphite to have a design life of 24 full-power years.
In place of the breeder’s continuous, rapid chemical
processing, however, we assumed the occasional discard
of the carrier salt, with recovery and recycle only of the
uranium in the salt. Batch cycles of six and eight
full-power years and variations in salt composition were
studied. Results indicate that the batch-cycle converter
reactor should be operated with about the same salt
composition as the breeder, for a given core design
optimized for breeding. Reoptimization for different
salt compositions might well reduce the apparent
sensitivity of the reactor performance to salt composi-
tion. The calculations indicate that a conversion ratio of
09 or higher, with a fuelcycle cost of 0.55 to 0.65
mill/kWhr(e), can be achieved with plutonium feed
(60% 23°Pu, 24% 2*°Pu, 12% 2% 'Pu, 4% ***Pu) and a
batch fuel replacement cycle of six full-power years.
Reactor physics experiments with an MSBR lattice
configuration have been initiated in the High-Tempera-
ture Lattice Test Reactor at the Pacific Northwest
Laboratory. These experiments include measurements
of the neutron multiplication factor at temperatures
from 300 to 1000°C, along with the reactivity effects
of varying fuel density, of changing the lattice configu-
ration, and of inserting various materials in the lattice,

xiii

including simulated control rods. Results of these

measurements will be used to check the accuracy of
nuclear data and computational models used in MSBR
design studies.

5. Systems and Components Development

The construction of the water test loop for testing
MSBE-scale gas separators and bubble generators was
completed, and operation was begun. The loop was
operated primarily on demineralized water at liquid
flow rates of 200 to 550 gpm and at gas flow rates of O
to 2.2 scfm. The loop was also operated with water
containing small amounts of n-butyl alcohol and so-
dium oleate and with a 41.5% glycerin-water mixture
which is hydraulically similar to fuel salt. The bubble
generator operated satisfactorily, and the bubble sepa-
rator . operated satisfactorily with water. An unex-
plained reduction in bubble size was observed when the
test fluid was changed from demineralized water to the
other fluids, and the separator was unable to remove
the small bubbles at the required rate. Tests were
started to determine if the production of small bubbles
was influenced by the pump efficiency or only the
pump head. It is believed that the small bubbles are a
characteristic of the test fluid and that they will not be
produced in salt.

The conceptual design of a molten-salt loop for
testing gas systems was completed, and the conceptual
system design description was written. Work is now
beginning on the preliminary design. The facility will be
used for developing the technology of the fuel salt for
MSRs and in particular for tests of the bubble generator
and separator. The facility is scheduled for initial
operation in early FY 1973. |

A program plan for obtaining reliable steam genera-
tors for the MSRP was outlined, and activities in the
first of three phases were started. We received affirma-
tive responses to an inquiry of interest in participating
in a conceptual design study from eight of nine
industrial firms contacted, and we are proceeding to
obtain proposals and to contract with one firm for the
studies.

The conceptual systems design description of the
3-MW test facility to be built as part of the molten-salt
steam generator program was completed, and further
work was suspended until late in FY 1972.

Work was begun on preparation of a development
basis report for molten-salt steam generators. In this
report we expect to evaluate the elements of the
LMFBR and other programs which have a bearing on
the moltensalt technology and then to point out
problem areas which need further study and outline a
program for such studies. This report should be finished
in the second quarter of FY 1972.

Following completion of the fluoroborate test pro-
gram in the PKP test loop, the salt pump rotary
element, the bubbler tube for BF; feed and salt level
indication, and other items of hardware were removed
for examination. The appearance of the pump rotary
element indicated that the fluoroborate service did not
cause excessive corrosion damage to the Inconel system.
Deposits of Na;CrFg and NaNiF; found in the bubbler
tube were attributed to reaction of the fluoroborate salt
with moisture introduced in the gas feed. A deposit of
mefallic nickel which blocked the mouth of the bubbler
tube was -probably formed by transfer of corrosion
product nickel from the bulk salt. The condition of the
gas pressure control valve was found to be like new.

The conceptual system design description for the
coolant salt technology facility was completed, and the
detailed design of the facility and components was
begun. We make maximum use of the drawings from
the PKP-1 test stand and of the components and
materials to be salvaged from the MSRE. The expected
completion date of the facility is late December 1971.

The salt was drained satisfactorily from the main loop
of the MSRE Mark 2 pump test stand after replacing
the plugged drain line which connects the loop piping
to the storage tank. :

The rotary element and pump tank of the MSRE
coolant salt pump were inspected visually as the pump
was removed from the coolant salt system. Except for
evidence of leakage oil from the lower shaft seal on
shield plug and tank surfaces, the pump appeared to be
in very satisfactory condition.

The water test program to qualify the ALPHA pump
for the hydraulic conditions required in the MSR-FCL-2
test facility was concluded satisfactorily. The design
and fabrication of the remaining parts needed for the
pump for the facility were then completed.

Most of the work previously supported by the MSRP
remote welding program has been transferred into an
automated welding program sponsored by the LMFBR
program to meet their needs in reactor pipe construc-
tion and maintenance. We are, however, completing a
small program to develop and test weld-torch position-
ing mechanisms and control circuitry; to define remote
maintenance inspection, viewing, and alignment criteria;
and to investigate pipe cleanliness requirements for

maintenance welding in salt systems. Preliminary tests.

have shown that small pipe filled with solid molten salt
can be welded after a rather simple cleaning procedure
and that these welds satisfy nuclear code x-ray inspec-
tion standards. |

6. MSBR Instrumentation and Controls

A hybrid computer simulation model of the reference
1000-MW(e) MSBR is being developed. The steam
generator is modeled mathematically on the hybrid
machine in continuous space and discrete time using
sets of differential equations derived from the conserva-
tion of momentum, energy, and mass. The integrations
are performed by the analog computer, while the digital
computer calculates the terms of the derivatives of the
differential equations and provides storage and control
for the calculations. The thermodynamic properties of
water are stored in the digital computer as two-
dimensional tables.

The model of the reactor, primary heat exchanger,
piping, etc., is a continuous-time model similar to those
traditionally used on analog computers and is time
scaled to 0.01 of real time. The discrete-time steam
generator calculations are stored and sampled at 1-sec
intervals, representing 0.01 sec in simulation time, then
smoothed and applied to the continuous-time analog
model.

The hybrid program for the steam generator has been
written and nearly debugged. The analog model has
been developed, but has not yet been patched. Integra-
tion of the two models will require some additional
time, and the total simulation is expected to be in
operation during the next reporting period.

7. Heat and Mass Transfer and Physical Properties

Heat transfer studies using the inert-gas-pressurized
flow system have shown that the average heat transfer
coefficient for a proposed MSBR fuel salt at 1070°F
and a Reynolds modulus of 3300 is 15% higher with a

Xiv

hydrodynamic entrance length than without. These
results are difficult to explain in terms of commonly
accepted theories of the combined development of
hydrodynamic and thermal boundary layers, but can be
explained by a recent theory which suggests that the
flow of a fluid whose viscosity has a large negative
temperature dependence will be stabilized by heating.
The present results tend to substantiate this theory.

A new technique for determining the wetting charac-
teristics of liquids was used to study wetting behavior
of the molten salt LiF-BeF,-ZrF,-ThF,-UF; (70-23-
5-1-1 mole %) on a Hastelloy N surface at 700°C
(1292°F). It was found that the typical nonwetting
condition of this salt could be changed to a wetting
condition within several minutes by the introduction of
a zirconium rod. Several hours were required to change
from the wetting back to the nonwetting condition by
the addition of 1 wt % nickel fluoride.

Water calibration of an improved variable-gap thermal
conductivity apparatus designed for use with molten
salts at 830°C gave results in excellent agreement with
the specialized room-temperature measurements pub-.

lished in the literature. Thermal conductivity measure- ..

ments are being made for the molten fluoride salt
system LiF-BeF,. g , ,
MSBR-related mass transfer experiments involving
diffusion of oxygen dissolved in glycerin-water solu-
tions into helium bubbles have been extended to
include the case of horizontal flow. The volume
fraction of bubbles, which is needed to determine the
interfacial area per unit volume, was found to correlate
with the ratio of axial to thermal velocity of bubble
rise. By making use of the interfacial area derived from
the bubble volume fraction correlation, overall mass
transfer coefficients (including the separator) have been
extracted from the measured concentration decay rate
of dissolved oxygen for a Reynolds modulus range from
26,000 to 66,000 at one value of the Schmidt modulus,
1228. When these recent results are compared with

" earlier results for vertical flow, it is found that, at a

sufficiently high Reynolds modulus for gravitational
forces.to be negligible compared with inertial forces,
the mass transfer coefficients for vertical and horizontal
flow become identical. Mass transfer coefficients for the
horizontal test section, corrected by subtracting the
measured mass transfer which occurs in the separator,
are correlated with Reynolds modulus for three values
of mean bubble diameter.

A theoretical description of mass transfer from a
turbulent liquid to bubbles moving at the local liquid
velocity has been developed which includes variation in
eddy diffusivity with concentration gradient near a
bubble interface, velocity, and frequency components
of the turbulence. A computer program has been
written to solve the pertinent equations.

PART 3. CHEMISTRY

8. Fission Product Behavior

Laboratory apparatus for determination of the con-
centration of tritium and hydrogen concentrations in
the MSRE fuel pump bowl gas was constructed and
calibrated.

Examination of the sampler cage and mist shield
excised in January 1971 from the MSRE pump bowl
revealed that all surfaces were covered with a deposited
film, generally gray-black, over unattacked metal. These
films contained carbon, lithium and beryllium fluorides
(and, doubtless, uranium and zirconium fluorides),
structural metals, fission products dominated by noble
metal isotopes and '37Cs, and tritium. Much, but
probably not all, of the structural metals could be
debris from cutting operations. Tritium/carbon atom
ratios of ~1 X 107* are consistent with nontrivial
exchange with hydrogen in tars resulting from the
cracking of lubricating oil which leaked into the pump
bowl. A very high proportion, relative to inventory, was
observed for all noble metal fission isotopes: °°Nb,
99Tc, '93Ru, '%%Ru, 12°8b, and 27" Te. The relative
proportions of '®*Ru and '®°Ru isotopes and the
quantities of deposit suggest that these deposits were
accumulated over a long period and that several percent
of the reactor inventory of noble metal isotopes is
deposited in the pump bowl. Thickest deposits were
noted on the sample cage rods and inner mist shield
walls below the surface, suggesting the steady agglomer-
ation and deposition of suspended material in the
less-agitated regions.

Attempts to synthesize niobium fluorides of lower
oxidation numbers provided additional quantitative
information concerning their disproportionation and
fractional sublimation. Studies of the stability of
molybdenum and niobium fluorides in molten Li, BeF,
were extended to temperatures as high as 700 and
900°C respectively. Evidence is presented which indi-
cates that the mechanism of removal of Mo from
molten Li,BeF; changes as temperatures exceed
500°C. Both increased nobility of niobium metal and
instability of NbF, solutions were observed in the
700-t0-900°C range as contrasted with the range 500 to
700°C. Mass spectroscopy of the niobium fluorides
yielded new information on the pentafluoride poly-
mers, the fluorination of niobium metal, the dispropor-
tionation of NbF,, and the behavior of the associated
oxy fluorides.

XV

9. Coolant Salt Chemistry and Tritium Control

The results of experiments performed with nickel as
the only metal in contact with molten sodium fluoro-
borate indicated that adequate but small concentrations
of chemically bound hydrogen can be retained in the
salt as a potential means for controlling the distribution
of tritium in molten-salt reactors.

Preliminary values for the solubility of HF in the
NaF-NaBF, eutectic mixture were measured in the
temperature ra‘nge 400 to 600°C.

Analysis of materials removed from an engineering
test loop showed that the corrosion experienced during
operation originated with the introduction of water
vapor as a contaminant in the cover gas.

Mass spectroscopic studies of the vapors over molten
fluoroborate melts were initiated, examining first the
effects of NaOH and moisture as impurities in fluoro-
borate systems.

Evidence for the existence of hydrogen-containing
impurities in NaBF,; was obtained from near-infrared
spectra of the molten salt and from mid-infrared spectra
of pressed pellets of the pure material.

Raman spectra were measured with polycrystalline
NaBF, at just above and below the dimorphic crystal
transition temperature. The results rationalize anom-
alies in x-ray diffraction data for NaBF, and related
structures.

Laboratory studies were continued in attempts to
synthesize hydrogen-containing species for the reten-
tion of tritium in molten fluoroborates.

The solubility of BF; was determined in five molten-
salt solvents composed of LiF and BeF,. For concentra-
tions of BF; in the melt below 1 mole %, Henry’s law is
obeyed.

An investigation of the equilibrium phase diagram of
the system RbF-RbBF,; was completed.

Component activities at the liquidus in alkali fluoride—
fluoroborate systems (Na, K, Rb) were computed and
compared with the ideal liquidus. In each case the
liquidus shows positive deviations at low concentration
of fluoroborates but exhibits negative deviation as it
approaches the eutectic.

The relative volume increase of the alkali fluorobo-
rates and perchlorates and alkaline-earth sulfates were
correlated for a range of temperatures up to and
through that of the dimorphic transition as a function
of cation size.

An experimental program was initiated to study
hydrogen permeation of metals under conditions
closely analogous to those which might be expected in a
molten-salt reactor steam generator heat exchanger.
10. Physical Chemistry of Molten Salts

Excess chemical potentials and partial molar enthal-
pies of mixing in molten LiF-BeF, were derived from
emf measurements using concentration cells.

An investigation of the equilibrium phase relation-
ships in the system LiF-BeF,-CeF; was continued.

Preliminary electrical conductance measuréments of
supercooled NaF-BeF, melts were obtained with an
all-metal conductance cell. Glass transition tempera-
tures were determined in the glass-forming region of the
NaF-BeF, system to permit comparison with theo-
retical glass transition temperatures obtained from the
temperature dependence of activation energies of elec-
trical conductance in BeF, -containing systems.

The Raman spectrum of BeF,?” was measured in
melts containing excess F~ ion; the results indicated
that in the molten mixtures the BeF4?™ anion retained
tetrahedral symmetry under a wide variety of composi-
tions and temperatures as in crystalline Li; BeF,.

Parameters affecting bubble formation were studied

experimentally by impingement of a liquid jet on a
fluid surface.

Apparatus was constructed and and used in prelimi-
nary measurements of the solubilities of helium and
hydrogen in molten salt.

The enthalpy of UF, was measured from room
temperature to temperatures above the melting point.
In contrast with recent reports in the Soviet literature,
the compound was found not to be dimorphic.

Factors affecting the stability of uranium trifluoride
as a dilute species contained in graphite were examined.
The stability, as measured using absorption spectros-
copy, was found to be less than predicted; stability was
also found to exhibit a marked solvent composition
dependence.

Because of the promising simplicity and efficiency of
a protactinium removal process based on its precipita-
tion as Pa,Og, work has continued on the oxide
chemistry of protactinium in molten fluorides. The
solubility of a Pa**-oxide phase was found to be
approximately four times higher than that of Pa;Os.
Together with measurements of the redox potential of
the couple Pa**/Pa®*, this leads to the prediction that
Pa, O; will not precipitate in the core of an MSCR or
MSBR in the event that accidental contamination with
oxide occurs.

The crystal structure of several complex compounds
of the alkali fluorides and heavy metal fluorides were
determined employing automated collection of X-ray
diffraction data from single crystals. Structural charac-
teristics of CsUgF,5, «KThgF,5, Li;MoFg,
RbTh;F,;,and RbsZr, F,, were established.

Xvi

A program of investigation of molten salts by x-ray
and neutron diffraction was resumed; initial studies
were devoted to fluoride glasses and SiO, .

The characteristic velocity at which a disturbance
(e.g., a thermal spike) is propagated in a fluid (molten
salts included) is approximately equal to the sonic
velocity in that fluid. A new method, based on the
Debye theory of specific heats, was developed which
allows sonic velocity to be estimated reliably from the
entropy (either measured or estimated) of molten salts.

In continuing efforts to develop an Ni-NiF, reference
electrode for use in fluoride melts, it was shown that
beryllium is very slightly soluble in LiF-BeF, melts and
that conductance in LaF; crystals in such melts has a
small electronic component.

11. Chemistry of Molten-Salt Reprocessing Technology

Distributions of rubidium and cesium in the fluo-
ride/bismuth extraction at 650°C were measured; the
order of extractability of the alkali metals from the salt
was shown to be Cs, Na, Rb, K.

The distribution of thorium between MSBR fuel
solvent and bismuth saturated with nickel and thorium
at 650°C was measured; the results indicated that the
distribution of soluble thorium was unaffected by the
presence of ThNiBi, in the bismuth phase.

The potential application of bismuth-manganese al-
loys as extractants for rare earths from MSBR fuel
solvent was assessed experimentally. The results were
sufficiently encouraging to warrant further investi-
gation.

The removal of solutes from bismuth by fractional
crystallization was examined in attempts to rationalize
the behavior occasionally observed in metal samples
removed directly from liquid bismuth mixtures.

12. Development and Evaluation of Analytical
Methods for Molten-Salt Reactors

An investigation of the electroanalytical behavior of
titanium (a constituent of Hastelloy N) in several
molten fluoride salts was initiated. Initial observations
in LiF-NaF-KF show that Ti(IV) is reduced reversibly
to Ti(lll) and, at a more cathodic potential, to the
metal. Evidence also suggests that Ti(IV) is reduced to
Ti(Ill) by graphite or nickel. The behavior of titanium
in Li, BeF, and NaBF, is presently being studied. The
Ni/NiF, (LaF3) reference electrode, which performs
satisfactorily in MSRE-type salts, has been found
unsuitable for molten NaBF, . This appears to be due to
the low solubility of Ni(II) in NaBF,. Other electro-
chemical couples are being investigated for use .as a
reference system for molten NaBF, .
In an investigation of the electrochemical generation
and spectral characterization of solute species in molten
fluorides, Mn(IIT), Cu(Il), Co(III), U(III), and CrO,;*"
were generated electrolytically. Also, superoxide ion
was spectrally characterized in fluoride melts. A con-
tinuing study has shown that several other oxygenated
species, including VO,%, NO,~, and —OH can be
observed in coolant and/or fuel melts. These observa-
tions have led to a cooperative spectral study of —OH
and —OD in NaBF,. The generation of CrO,* is of
particular interest because it offers a means for the
spectrophotometric determination of traces of chro-
mium and possibly oxide in molten fluorides. In
LiF-BeF, and NaBF,, hexavalent chromium was found
to exist as Cr,0,2". The generation of Cr,0,% by
chemical oxidation of CrF, has been demonstrated in
LiF-BeF, melts but has not yet been achieved in
NaBF, .

By careful purification of NaBF,, melts transparent
to wavelengths as short as 200 nm were obtained.
Absorbance measurements below 300 nm were found
to provide a sensitive method for the detection of iron
as well as other impurities.

An evaluation of the azeotropic distillation method
for the separation of water from NaBF, showed that
significant positive errors can be introduced by high
background (blank) titrations and possible contamina-
tion during sample additions. With an improved proce-
dure, previously undetectable concentrations (<50
ppm) of distillable water were found in samples from
coolant salt test loops. The azeotropic distillation
method is under development for measurement of
hydrolysis products in coolant cover gas as well as for
salt analysis. ‘

Components for the first in-line applications of
electroanalytical methods to fuel salt are now being
tested. Voltammetry will be used to establish the
U(IIT)/U(IV) ratio and to determine the concentrations
of certain corrosion products in the flowing fuel in a
thermal-convection loop. A reference electrode system
will be used for continuous measurement of the redox
potential. A PDP-8/I computer coupled with a voltam-
meter of improved design will be used for the unat-
tended measurement of U(III)/U(IV) ratios. Before
installation, this system will be tested on a melt in a
grounded cell.

PART 4. MATERIALS DEVELOPMENT

13. Examination of MSRE Components

Several components from the MSRE primary circuit
were examined. The graphite moderator element was in

xvii

excellent condition, with no detectable dimensional
changes. There was no metallographic evidence of
corrosion. There was a shallow surface layer about 2
mils deep having a modified structure that we attribute
to surface working. All metal surfaces exposed to the
fuel salt were embrittled to a depth of 5 to 10 mils as
evidenced by grain boundary cracking to a depth of 5
to 10 mils. The failure was located in freeze valve 105
and was attributed to thermal fatigue. A copper sample
capsule was retrieved and found to be extremely brittle;
all surfaces were coated inhomogeneously with Ni, Fe,
Cr, and Mo.

14. Graphite Studies

The routine evaluation of vendor-furnished graphites
has been essentially completed, and the irradiation
program is rapidly shifting to the investigation of
experimental graphites exhibiting the desired structural
variations. In support of this, the fabrication effort on
graphite has expanded. In the study of precursor
chemistry, the effect of hetero atoms (N, S, O) is being
investigated. These materials lead to a deterioration in
crystal perfection, but also lead to the desired iso-
tropicity of the cokes and graphites. Raw cokes and
blacks are currently being fabricated into isotropic
graphites with the required monolithic structure. This
work is still in its early stages, but the fabricated bodies
look most promising.

The black-based graphites discussed in the previous
semiannual report continue to be irradiated. Heat-
treatment series have been prepared and characterized.

At the level of fundamental understanding, the
distortions around an interstitial cluster have been
calculated from elastic continuum theory, and the
phonon dispersion curves have been measured by
neutron scattering.

15. Hastelloy N

Postirradiation tests on laboratory melts of modified
Hastelloy N have shown that acceptable properties can
be obtained with additions of 1.5 to 2.5% Ti, at least
0.5% Hf, and multiple additions of Ti, Hf, and Nb.
Small commercial melts have confirmed the beneficial
effects of Ti and Nb but have not reproduced the
beneficial effects of Hf. The creep strength of all of the
modified alloys is higher than that of standard Hastel-
loy N. Weld metal cracking was encountered with high
Hf and Zr concentrations, but suitable welds were made
using standard Hastelloy N as a filler metal. The
evidence is quite encouraging that an alloy with 1.5 to
2.5% Ti will have adequate postirradiation properties.
Corrosion tests of Hastelloy N in LiF, BeF,, UF,,
ThF, salts continue to give very low corrosion rates,
<0.1 mil/year. Tests in sodium fluoroborate now
involve four thermal convection loops and two pump
loops. The corrosion rates vary from 0.1 to several
mils/year, depending on salt purity. Evidence indicates
that the impurities most affecting the corrosion are
water and oxygen. A test to study tritium retention in
sodium fluoroborate by exchange with H showed that
the tritium was released quickly by the salt, indicating a
lack of stable hydrogen in the salt. Hastelloy N
specimens exposed to steam at 538°C continue to show
a metal loss rate of <0.25 mil/year.

16. Support for Chemical Processing

We have begun fabrication of components for a
molybdenum test stand. A 5-ftlong, 1'%-in.-diam by
0.080-in.-wall-thickness section of molybdenum piping
was fabricated by forward extrusion. Several 37%-in.-
diam by 0.125-in.-wall closed-end, cylinders were back
extruded and machined for welding studies. Procedures
have been developed for electron beam welding a girth
joint between two of the 3%-in.-OD molybdenum
headers and for electron beam welding tube-to-header
joints between Y%-, Y%-, %-, %-, and 1%-in.-OD tubes
and the 37%-in.-OD headers. For field welding molyb-
denum we have modified a commercial orbiting-arc
welding head and developed procedures for making
tube-to-tube welds. We continued studies on three
iron-base brazing alloys for molybdenum (Fe—15%
Mo—5% Ge—4% C—1% B, Fe—15% Mo—4% C—1% B,
and Fe-25% Mo—4% C—1% B). We determined the
optimum clearance for back brazing welded and me-
chanical joints in the molybdenum test stand and also
found that the braze alloys could be used to repair
cracked welds. The shear strength and percent elonga-
tion were measured for the Fe—15% Mo—5% Ge—4%
C—1% B and Fe—15% Mo—4% C—1% B compositions at
both room temperature (30,000 psi and 10%, respec-
tively) and at 650°C (18,000—29,000 psi and 40%
respectively). Molybdenum metal seal couplings were
procured and evaluated by helium leak checking before
and after thermally cycling the joints from room

temperature to 650°C. Two techniques, magneforming

and roll bonding, have been investigated for mechani-
cally joining molybdenum. Several helium-leak-tight
joints (<5 X 107% std cc/sec) have been made by roll
bonding.

We continued to evaluate the compatibility of several
- container materials and brazing alloys in pure bismuth
and bismuth-lithium. solutions. In quartz .thermal-
convection loop tests (700°C maximum temperature,

Xviii

100°C temperature difference) three different grades of
graphite showed no significant attack except that the
more open grades contained bismuth in their pores.
Four iron-base filler metals (Fe-C-B) that contained O,
15, or 25% Mo, and in one case 5% Ge, were also tested
in pure bismuth under similar conditions. No increase in
iron concentration of the bismuth was found, but the
alloys picked up significant amounts of bismuth and
exhibited surface layers that were high in iron. These
alloys showed no attack when exposed to an LiF-BeF, -
ZrF,-UF,;-ThF, mixture for 1000 hr in a 304L
stainless steel thermal-convection loop operating at a
maximum temperature of 690 and a 100°C temperature
difference. A quartz thermal-convection loop contain-
ing molybdenum and TZM samples was operated for
3000 hr with a Bi—100 ppm Li solution (700°C
maximum temperature and 100°C temperature dif-
ference), and only very slight attack of the very fine
grains along the surface was found.

We have continued to coat stainless steels and
nickel-base alloys with tungsten and molybdenum by
hydrogen reduction of WF, and MoF. Both coatings
were found to be adherent to the base materials during
thermal cycling between 25 and 600°C, and tungsten
coatings exhibited tensile bond strengths of 20,000 to
35,000 psi. Objects of various size and shape were
coated with tungsten to demonstrate the applicability
of the process. We have found that we can coat
cylindrical shapes of up to 4 in. in diameter and 48 in.
long with little difficulty. The optimum conditions for
applying smooth, adherent molybdenum coatings were
found to be about 900°C and an H, /MoF ratio of 3 to
6. Bend tests have shown that the molybdenum
coatings are more ductile than the tungsten coatings.

Studies have also continued to develop a technique
for depositing molybdenum on iron-base substrates
from MoFg in a molten fluoride salt mixture. Corrosion
reactions resulting from too high a concentration of
MoF¢ in the salt have limited the success of experi-
ments conducted thus far.

PART 5. MOLTEN-SALT PROCESSING
AND PREPARATION

17. Flowsheet Analysis

An improved flowsheet was developed in which
protactinium is isolated from the fuel salt of an MSBR
and held for decay in a secondary salt stream that is
physically and chemically isolated from the reactor. A
processing plant based on this flowsheet should be
much easier to control than one based on the earlier.
flowsheet, and a considerable saving in capital equip-.

ment cost should result. A method for combining and
fluorinating the various waste streams from the metal
transfer process and the protactinium isolation system
was developed. This will eliminate several potential
routes for loss of fissile material from the system.

Oxide precipitation is being considered as an alterna-
tive to the fluorination—reductive-extraction method
for isolating protactinium and for subsequently re-
moving uranium from the fuel salt of an MSBR. Two
possible flowsheets -based on oxide precipitation were
developed, and the effects of several parameters on
operation of the processes were investigated. In the first
flowsheet, the isolated protactinium is held in a
secondary salt from which the uranium is removed by
fluorination. In the second flowsheet, the isolated
protactinium is dissolved in processed fuel carrier salt
from the metal transfer process, and the resulting
stream is recycled through a protactinium decay tank to
the protactinium oxide precipitator. The precipitator
efficiency required for these flowsheets is 60 to 80% for
the first flowsheet and about 96% for the second.

Although the MSBR processing flowsheets considered
thus far have uniformly resulted in very low uranium
inventories in the processing plant, several potential
processing systems may result in uranium inventories as
large as 5 to 10% of the reactor inventory. A processing
plant uranium inventory of 5% of the system fissile
inventory would increase the fuel cycle cost by about
0.015 mill/kWhr and would increase the system dou-
bling time from 22 to 23.1 years. While there is
incentive for maintaining a low uranium inventory in
the processing plant, it does not appear that a uranium

inventory as high as 5 to 10% of the system fissile
inventory would rule out an otherwise attractive proc-
essing system.

18. Processing Chemistry

Studies in support of the development of the metal
transfer process for removing rare-earth and other fission
products from MSBR fuels were continued. The distri-
bution of several actinide elements between molten LiCl
and liquid bismuth solution was determined at 640 and
700°C. Additional information was obtained at 640°C
on the effect of LiF concentration on the distribution of
several solutes between LiCl-LiF solutions and liquid
bismuth. This information confirmed earlier indications
that LiF in concentrations of less than about 4 mole %
has little effect on the behavior of di- and trivalent
solutes but that the thorium—rare-earth separation
factor decreases with increasing LiF concentration.

Xix

Measurements of the solubilities, both individual and
mutual, of thorium and rare earths in lithium-bismuth
solutions were also continued. At each temperature in
the range 400 to 700°C, the solubilities of both
thorium and neodymium increased as the lithium
concentration in the solution increased from 0 to 25 at.
%. The mutual solubilities of thorium and neodymium
were measured in lithijum-bismuth solutions at 640°C
and appeared to be much higher than those required in
the stripping of rare earths from the LiCl acceptor salt
into lithium-bismuth solutions.

Investigation of oxide precipitation as a means for
isolating protactinium from MSBR fuel salt was con-
tinued. Two methods were used to estimate the
solubilities, at various temperatures, of Pa, Qs in mol-
ten LiF-BeF,-ThF,-UF, (71.8-16-12-0.2 mole %) that
was saturated with UO,. These solubilities define the
lowest protactinium concentrations attainable without
attendant precipitation of uranium and thorium oxides.

19. Engineering Development of
Processing Operations

The second engineering experiment on the metal
transfer process for the removal of rare earths from
single-fluid MSBR fuel salt has been completed. The
experiment operated satisfactorily for about three
months. before it was shut down for inspection. During
that period, more than 85% of the lanthanum and more
than 50% of the neodymium originally in the fuel
carrier salt were removed and deposited in a lithium-
bismuth solution. There was no measurable accumula-
tion of thorium (<10 ppm) in the lithium-bismuth
solution, thus demonstrating that the rare earths can be

- removed without significant removal of thorium. The

thorium-lanthanum decontamination factor was about
10%. The distribution coefficients for lanthanum and
neodymium between the fluoride salt and the thorium-
saturated bismuth were relatively constant and in
agreement with expected values. The distribution coef-
ficients for lanthanum and neodymium between LiCl -
and thorium-saturated bismuth were somewhat higher
than expected. When the vessel was disassembled for in-
spection after the run, the condition of the interior of
the vessel was generally good; however, some corrosion
had occurred on the components made of thin carbon
steel (the lithium-bismuth container and the sparge
tubes).

The third engineering experiment for development of
the metal transfer process is presently being designed.
This experiment will use flow rates that are 1% of the
estimated flow rates for a 1000-MW(e) reactor. Mechan-
ical agitators will be used to promote efficient contact-
ing of the salt and metal phases. The three vessels
required for the experiment will be made of carbon
steel; one of the vessels has already been fabricated. A
test vessel has been fabricated that will be used with salt
and bismuth for testing the proposed agitator drive
unit, the shaft seal, the performance of a vapor-
deposited layer of tungsten on the vessel interior, and a
new technique for applying nickel aluminide coatings to
the exterior of carbon-steel vessels.

A program has been initiated for the development of
mechanically agitated salt-metal contactors as an alter-
native to packed columns presently considered for the
MSBR processing system. Studies to date have been
concerned primarily with selection of a contactor
design for the third metal transfer experiment. Tests
with mercury and water using a four-bladed flat-paddle
agitator located in the mercury-water interface fre-
quently resulted in a stable dispersion of very small
mercury droplets and entrainment of water in the
mercury. We have begun studies of a contactor which
has a paddle operating in each phase, well away from
the interface, at a speed that does not result in the
dispersion of either phase.

We have continued to study the extraction of
uranium from molten salt by countercurrent contact
with bismuth containing reductant in a packed column.
Two successful runs, which included six periods of
steady-state operation covering metal-to-salt flow ratios
ranging from 0.75 to 2.05, have been made. The data
could be correlated by an HTU model based on the
assumption that the uranium transfer rate was con-
trolled entirely by the diffusive resistance in the salt
phase. The observed HTU values ranged from 0.77 to
2.1 ft and were inversely proportional to the metal-to-

sait flow rate ratio. Flooding data obtained during

countercurrent flow of salt and bismuth in the packed
column have continued to show good agreement with
predictions based on studies with a mercury-water
system. .

Preparations were begun for mass transfer experi-
ments in which the rate of exchange of zirconium
isotopes will be measured between salt and bismuth
phases otherwise at equilibrium in a packed column.
The first experiment of. this type had to be terminated
because of a salt leak, which necessitated replacement
of the salt feed-and-catch tank.

Studies of hydrodynamics in packed columns during
the countercurrent flow of high-density liquids are

being made in order to evaluate and design contactors

for processing systems based on reductive extraction.

XX

Data obtained by an MIT Practice School group show
that the slip velocity with nonwetted packing is
dependent on the —0.167 power of the continuous-
phase viscosity rather than being independent as was
previously assumed. Knowledge of the dependence of
slip velocity on the continuous-phase viscosity allowed
calculation of the power dependence of slip velocity on
the difference in densities of the phases. A power depen-
dence of 0.5 was calculated rather than the previously
assumed dependence of 1.0. Studies with packing wetted
by the metal phase indicate a substantial reduction in
interfacial area between the phases and an increase in
the slip velocity.

Twentyseven runs were made with the simulated
continuous fluorinator to determine heat generation
rates in a column- of nitric acid (used as a stand-in for
molten salt), in the pipe surrounding the acid column
(representing the fluorinator wall), and in induction-
coils of three different designs. Air was bubbled
through the nitric acid at rates up to 2.16 scfh, resulting
in bubble volume fractions as high as 18%, to determine
the effect of bubbles on the heat generation rate in the
acid. Efficiencies predicted for heating molten salt in a
fluorinator using the best coil design tested to date
ranged from 37.0% with no bubbles to 32.9% with a
bubble fraction of 15%. These efficiencies are suffi-
ciently high to allow operation of a fluorinator having a
4.5-in.-diam molten zone and a 1.5-in.-thick frozen salt
layer by using an available rf generator.

Rate constants for the corrosion of Ni-200 and
Ni-201 in gaseous fluorine at l-atm pressure were
calculated from literature data, assuming that the
reaction follows a parabolic rate law; the calculated rate
constants were correlated with temperature, assuming
an Arrhenius temperature dependence. The resulting
data were then used to estimate average corrosion rates
when the NiF, protective film is periodically destroyed.
Average corrosion rates of 2.9 and 0.97 mils/year for
Ni-200 and Ni-201, respectively, were estimated at
450°C (a typical fluorinator wall temperature) when
the NiF, film was assumed to be destroyed 52 times
annually. These corrosion rates are acceptable for
long-term fluorinator operation.

Studies of the effect of column diameter on axial
dispersion in open bubble columns were extended to
include a 6-in.-diam column, and additional data were
obtained on the effect of the viscosity of the liquid.
The dispersion coefficient values obtained with a
6-in.-diam, 72-in.-long column were about three times
the values measured at the same superficial gas velocity
in a 3-in.-diam column; the data show little dependence
on superficial gas velocity. The dispersion coefficient in
a 6-dn.-diam column is not noticeably affected by
changing the viscosity of the liquid from 0.9 to 12 cP.
The same change in viscosity in a 2-in.-diam column
results in a 20% decrease in dispersion coefficient.
However, since the diameter of continuous fluorinators
will be 6 in. or larger, the viscosity of the liquid will not
affect the dispersion coefficient.

An engineering-scale experiment to study the precipi-
tation of uranium oxide from MSBR fuel salt, from
which the protactinium has been previously removed, is
being designed. In this experiment, uranium will be
precipitated from 2 liters of fuel salt in a single-stage
batch operation by contact with an argon-steam mix-
ture. The design of the precipitator vessel is described.

Very little additional design work on the processing
materials test stand and the molybdenum reductive-
extraction equipment has been done pending develop-
ment of molybdenum fabrication techniques. The

X X1

plastic model of the head pot was tested and modified,
an improved design is ready for final testing.

20. Continuous Salt Purification System

To date, 11 iron fluoride reduction runs have been
carried out in order to study the countercurrent contact
of molten salt with hydrogen in a packed column. In
the first three runs, iron fluoride mass transfer coeffi-
cients averaged 0.016 ft/hr. Coefficients for the suc-
ceeding runs are questionable because of inconsistent
iron analyses. Oxide was removed from the salt by
contacting the salt with H,-HF, first by countercurrent
contact in the column and second by bubbling the gas
through a static column of salt. Corrective measures
taken to alleviate the increased pressure drop in the
packed column are also described.
Part 1. Molten-Salt Reactor Experiment

P. N. Haubenreich

The postoperation examinations were completed
during this report period, and the plant was secured to
await ultimate disposal of the salt. This section outlines
the examination program, describes the on-site work,

and presents some of the results. Results of some of the
detailed analyses made on items removed from the
MSRE are described in Parts 3 and 4 of this report.

1. Postoperation Examinations

R. H. Guymon

The purpose of the postoperation examinations was
to round out the MSRE experimental program by
providing information not available during operation,
such as the condition of materials throughout the
system (for comparison with specimens from the core),
the condition of key components, the location and
nature of deposits in the pump bowl, and the exact
location and nature of a leak in the fuel salt drain line.

1.1 OUTLINE OF PROGRAM

Within a few weeks after the end of nuclear opera-
tions in December 1969, the salt was allowed to freeze
in the drain tanks, and other system conditions were
established so that routine surveillance could be main-
tained from the ORNL Central Waste Monitoring
Facility.! The operating crews were disbanded, and the
maintenance planning staff (one engineer and two
technicians) began development of the special tools and
detailed procedures that would be required for the
examination tasks that had been selected.? Specimens
of the coolant piping and radiator tubes were cut out in
July 1970.> but it was mid-October before craft
support was available for the final testing of the tools
for the fuel system. '

1. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, pp. 24 -25.

2. P. N. Haubenreich and M. Richardson, Plans for Post-
Operation Examination of the MSRE, ORNL-TM-2974 (April
1970).

3. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL+4622, pp. 119-33.

P. N. Haubenreich

The special tools included abrasive cutters with
vertical and horizontal wheels and a hydraulically
operated shear. These were intended for use on small
piping, and tests consisted simply of cutting pieces of
pipe in the maintenance practice cell, with the tools
operated from overhead through the portable mainte-
nance shield. The job of cutting into the pump bowl
was more complicated, and the preparations included
moving the MSRE prototype pump from the pump
development facility into the practice cell and going
through the operations of cutting away interfering
structure and attaching the bowl-cutting tool. This tool
was tested by cutting inclined ' -in. plates simulating
the top of the pump bowl. Preparations for the heat
exchanger work involved development by welding
specialists " of techniques for remotely operating a
plasma torch for cutting the shell and a Heliarc torch
for welding on a patch. The tools and fixtures for this
work were tested on a partial mockup of the heat
exchanger in the maintenance practice cell at the
MSRE.

Final preparations were completed by November 23,
and on that date the reactor cell was opened to begin
work on the radioactive systems. The main line of
activities, a sequence of operations involving use of the
portable maintenance shield, is outlined in Fig. 1.1.
Concurrently the coolant pump and the 5-in. coolant
piping outside the reactor cell were cut out and
removed for use in MSR technology development
facilities.

Descriptions of the major tasks are given in the
sections which follow.
" ORNL- DWG T1-6975

OPEN REACTOR CELL
REMOVE CONTROL ROD DRIVES, RODS

REMOVE STANDPIPE

REMOVE CORE ACCESS PLUG AND THIMBLES } CORE
VIEW INSIDE REACTOR VESSEL

REMOVE CORE GRAPHITE ELEMENT

CLOSE REACTOR VESSEL

TEST REMOTE MAINTENANCE PROVISIONS
CUT SAMPLER TUBE

DEC 1970

FUEL

\ A
CUT AROUND SAMPLER CAGE _ " PUMP

REMOVE SAMPLER CAGE
VIEW INSIDE PUMP BOWL

RETRIEVE CAPSULE
PATCH PUMP BOWL

5 3

P-d

g CUT HOLE IN HX SHELL | HEAT
REMOVE & TUBES EXCHANGER
WELD PATCH ON HX SHELL
REMOVE THERMAL SHIELD SLIDE } 5%’5‘&8'*
OBTAIN INSULATION SAMPLE INSULATION
SEAL HEAT EXCHANGER PATCH
FIND LEAKED SALT
CUT DRAIN LINES

- LEAK

DELIVER FREEZE VALVE FVv-105

FEB 1971

PLUG SALT TRANSFER LINES
SEAL CONTAINMENT CELLS

. Fig. 1.1. Outline of postoperation examinations.

1.2 REACTOR VESSEL AND CORE

Before the large access flange on the reactor vessel
could be opened, equipment above it had to be
removed. The first items, the three control-rod drives,
were removed quickly and uneventfully. The drives
were not inspected but were placed in the equipment
storage cell. Next the three control rods were removed
and viewed in the reactor cell. They appeared to be in
good condition, still flexible, with no perceptible
distortions, surfaces smoothly oxidized with only minor
marks from rubbing inside the thimbles. The poison
section of one rod was severed, put in a lead-shielded
pipe, and placed in the storage cell. (Gamma radiation
was 7 R/hr at the outside of 2 in. of lead.)

The 20-in.-diam standpipe attached to the core
specimen access flange came out only after some

flexible air lines were cut with a torch to clear the way.
The standpipe, reading 10 R/hr at 3 ft, was bagged and
buried. After electrical connections and the standpipe
ventilation line were detached, the lead which had
protected the rod drives was removed.

The bolts in the 10-in. access flange came out without
excessive torque. (The removal tool suffered a broken
universal joint pin but worked satisfactorily after
repair.) The access plug (which includes the control rod
thimbles) lifted out freely and was laid down on top of
the thermal shield, as shown in Fig. 1.2, for viewing.
The surfaces of the rod thimbles were uniform in
appearance, with a dull-gray surface. The lower 42 in.
of thimble 3 was severed with a Heliarc torch and
delivered for hot-cell examination (see Sect. 13.3). In
the strainer basket section that had extended into the
top head of the reactor vessel, the edges of the
perforations .were sharp, with no visible erosion. The
only salt evident on the whole assembly was small blobs
visible through a few of the perforations. There was no
heavy deposit in the region of the salt-gas interface, but
farther up there was a blackened area (visible near the
flange in Fig. 1.2). Inspection of the inside of the access
nozzle with a periscope showed a darkened area near
the top matching that on the plug. The area was
opposite the air inlet into the cooling jacket, suggesting
that the dark material was a surface deposit. Otherwise
the inside of the access nozzle and the upper head of
the reactor vessel appeared to be in excellent condition,
with only a few scattered droplets of frozen salt in
evidence.

The first view down through the access opening was

- as shown in Fig. 1.3. Several irregular chunks up to 2 in.

in size were seen lying on top of the core structure in
this area. Close examination with a right-angle scope
showed them to be broken pieces of graphite. Figure
1.4 is a typical view through this scope, looking
westward at a large chunk inside the metal centering
bridge and a small chunk resting on top of the bridge.
(These chunks are visible in Fig. 1.3 at the lower right.)
A wide look at the top of the core through a fisheye
lens disclosed six or seven more chunks of graphite
lying about. Although no gaps were noticed in the core
array, the characteristic pyramid shape on at least two
of the chunks showed that they were the upper ends of
vertical graphite bars. Figure 1.5 is a photograph
through the fisheye lens, blurred by vibration of the
long scope during the exposure. The direct view was
sharper, and gaps in the array would probably have
been noticed. The elements were not counted, however,
and the absence of some around the periphery could
easily have been overlooked.

PHOTO 101607

Y

1 CORE_ACCESS PLUG S8

N . T
“ ¥ CONTROL RODf
==

i 3

Fig. 1.2. Top of thermal shield during core inspection.
PHOTO 101622

Fig. 1.3. View of top of core through access opening.
PHOTO 101619

Fig. 1.4. Closeup view of broken pieces of graphite on top of core.

One piece of graphite visible in Fig. 1.3 was definitely
established as coming from the periphery of the core.
This was the one lying partly across the rod thimble
hole toward the bottom of the figure. (In this view the
lighting makes the flat surface which faces upward
appear bright.) The angles on the machined end of this
piece coincide with those on construction drawings for
one of the elements that was partly cut away for the
retaining ring that circled the core at the top. It is the
mirror image of the piece pointed out in Fig. 1.6.

The most likely explanation for breaking of the
graphite is that it occurred during cooldown. After salt

trapped between the Hastelloy N retaining ring and the
graphite froze, further thermal contraction of the ring
could produce bending loads on the upper ends of the
bars. (This is basically the same chain of events that
broke many of the graphite pieces in the first specimen
array exposed in the core.)® The chunks that got loose,
to float up to the outlet screen on a fill and subside
somewhere atop the core on a drain, were broken off
above the retaining wire. The larger pieces (see Fig. 1.5)

4. MSR Program Semiannu. Progr. Rep. Aug. 31, 1966,
ORNL4037, pp. 97-102.
PHOTO 101615

Fig. 1.5. View of top of core through a fisheye lens.

must have come from the rows inside the centering
bridge, where the graphite extended almost 3 in. above
the wire.

An attempt was made to retrieve some of the pieces
lying directly below the access opening, but all were
accidentally dislodged and fell down through the
thimble holes into the lower head where they would
have been very difficult to recover.

Except for the broken graphite (which could not have
caused any significant perturbations during operation),
the conditions in the upper head looked good, with
clean surfaces and sharp edges. An example is the weld
at the extreme right of Fig. 1.4.

The east removable graphite stringer was marked on
top to show its orientation and then was removed
without difficulty into a shielded carrier for delivery to
the HRLEL. (It was found to be in excellent condition,
as described in Sect. 13.1.) Another stringer was being
lifted out to provide a wider view of the lower head
when an improper movement of the crane broke the
graphite and the lower part fell back into the core.

There was sufficient clearance for a light and a
periscope to be let down into the lower head, but the
view was rather restricted because of the metal grid
structure. The bottom surface appeared to have a sandy
texture, but if there was a deposit, its depth was very
small compared with the dimensions of the drain line
entrance, which could be clearly seen. No foreign
objects were seen in the lower head.

After viewing inside the reactor vessel was finished,
one of the control rods was placed in a core channel
and the access was closed with a gasketed blank flange.

1.3 FUEL PUMP

The preliminary steps in cutting into the fuel pump
bowl involved testing some of the built-in remote
maintenance devices on the pump. The mechanism for
compressing the bellows in the sampler tube spool piece
was rusty, and a drive chain broke when its use was
attempted, so the spool piece had to be removed by an
improvised procedure. On the other hand, bolts in the

PHOTO 70660

Fig. 1.6. Core graphite during assembly, before installation of retaining rings and centering bridge. Arrow points to piece similar

to one found broken on top of core.

large flanges on the motor and the rotary element were
loosened without difficulty.

An oxyacetylene torch was used to cut away the
pump support plate, the cooling air shroud, and other
obstructions over the pump bowl. The sampler tube was
then cut off short, and the latch stop was tapped for
attachment of the bowl-cutting tool shown in Fig. 1
This special tool consisted of an air-motor-driven
abrasive cutter swivel-mounted on a bracket that
fastened between the sampler tube stub and a bolt hole
in the pump flange. Manual swiveling of the tool
produced a trepan cut like that shown in Fig. 1.7 in the
simulated pump bowl top. The actual cutting of the
pump bowl was delayed by cutting wheels breaking,
and when the fourth wheel broke after the cut was
nearly complete, a chisel was used to finish freeing the
sampler mist shield and cage. This assembly was then
removed for detailed examination in the HRLEL, as
described in Sects. 8.3 and 13.4.

Visible on the bottom of the pump tank in the area
that had been enclosed by the mist shield were the
copper bodies of the two 10-g sample capsules that had
been dropped in August 1967 and March 1968. The
capsules seemed to be coated with an irregular dark
deposit, neither had any sign of its steel cap, and one
had been mashed rather flat. Subsequently the flattened
sule was retrieved for inspection, but the other
capsule was accidentally knocked out of reach. Evi-
dently the recovered capsule was the older of the two,
whose corroded steel cap had been retrieved in May
1968 and whose body had been bent by the impact of
the heavy magnets used at that time.®

In the same area as the capsules there was a pile (S0
to 100 cm?®) of loose dark material, some as small
particles, some as lumps up to 1 cm, and some as thin

S. MSR Program Semiannu. Progr. Rep. Aug. 31, 1968,
ORNL-4344, pp. 26-29, 113-14.

proTo_tor336

Fig. 1.7. Tool for excising fuel pump sampler cage, mounted on mockup of top of pump tank.

sheets up to 2 cm across. On contact, the larger pieces
crumbled easily, and heat from the viewing light caused
the material to smoke. A sample was scooped up and
delivered with the capsule. Analysis of the loose
material and the deposits on the mist shield and cage
are described in Sect. 8.3. Results of the inspection of
the capsule are given in Sect. 13.5.

A periscope was used to view the interior of the pump
bowl, where throughout the years of operation oil had
decomposed, salt mist had floated about,® and noble
metals had concentrated.” Considering all this, surpris-
ingly little material was found deposited on the various
surfaces of the pump tank that were visible from the
opening (see Fig. 1.8).

The top head of the tank was covered with a dull,
dark, rather uniform film, seemingly like that on the
outside of the mist shield near the top. Here and there
thin sheets of this material had peeled and hung down.

6. 1. R. Engel, P. N. Haubenreich, and A. Houtzeel, Spray,
Mist, Bubbles and Foam in the MSRE, ORNL-TM-3027 (June
1970).

7. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL4622, pp. 2—4.

Some of the sheets could be seen swaying, presumably
in the thermal currents from the hot lamp.

Quite different were the deposits on the upper surface
of the sloping baffles in the upper part of the bowl.
These were nonuniform, rather rough or angular, and
reflected more light than did the deposit on the top
head. This type of deposit is illustrated in Fig. 1.9,
which is focused on the edge of the "s-in.-thick baffle
that slopes down from the volute support cylinder.
(There is some blurring in the image due to motion
during the exposure. The overflow pipe and its junction
with the lower head, seen in the background, are also
out of focus.) The same kind of deposit is seen in Fig.
1.10. This is a closeup of the end of the spray ring and
attached baffle. The baffle itself is "-in.-thick metal,
and the deposit seen here in edge view is less than half
as thick. The ripples in the weld on the end of the tube
are visible through the deposit. In the patch seen in Fig.
1.10 near the juncture of the baffle and the tube, where
the deposit had evidently flaked off, the exposed metal
was clean and smooth. Since it seemed that this kind of
deposit was also present on the excised mist shield, no
effort was made to obtain a sample from surfaces inside
the bowl.

SAMPLE
CAPSULE
CAGE
OVERFLOW
PIPE

ORNL-DWG 69~ 10172

OFFGAS
LINE
BUBBLER

T

Z
(“Amn\\\

40| J
20
ot . Z
LeveL = DISCHARGE
o
e JSALT

<= RADIOACTIVE GAS
<= CLEAN GAS

SUCTION

Fig. 1.8. MSRE fuel pump internals.

Fig. 1.9. View inside fuel pump bowl showing edge of spray baffle, with scattered deposits. The overflow pipe is visible in the

background.

10

puoro_iorees

101098

Fig. 1.11. View in lower part of pump bowl showing ports into volute suction.

The portion of the spray ring below the baffle was
clean, and the spray holes that were visible were sharp
and clear. One of the brackets supporting the spray ring
appeared white and seemed to be detached from the
wall.

Nowhere to be seen in the pump bowl were salt
droplets or the globular deposits often associated with
frozen mist, and nothing was visible in the salt-gas
interface zone to suggest that there had been any foam.
Except for the debris around the sampler region, the
bottom of the pump bowl was relatively clean, with
only a few areas showing a thin dark film or deposit.
The metal surfaces that had been submerged in the salt
were generally smooth dull gray. One exception was on
the lower extension of the volute where there was some
discoloration and minor roughness around the ports.
This is evident in Fig. 1.11. On the bubbler tube there
was a white area just above its penetration of the baffle
extending out from the volute, where gas that collected
beneath the baffle bubbled up through the clearance
gap around the tube.

After the viewing was concluded, the hole in the
pump bowl was sealed. A steel patch with silicone
rubber sealant around the periphery, held in place by a
jack bolt extending up to the rotary element flange, was
installed and proved leak-tight in a soap test at 5 psig.

11

1.4 HEAT EXCHANGER

Two heater units, HX-1 and -2, had to be removed to
permit cutting into the heat exchanger. Both had been
removed and reinstalled in 1968 without particular
difficulty,® but this time HX-1 hung on a freeze flange
clamp during removal and was pulled out of shape.
Both units were placed in the equipment storage cell.

A section of the '-in-thick shell of the heat
exchanger was cut out, using a plasma torch device that
had been developed and tested on a mockup of the heat
exchanger shell. Figure 1.12 is a view through the 10-
by 13-in. oval opening. The dull appearance of the
tubes is due to a thin coating which wiped off easily
during handling. The oily spot on the tubes at the
center of the opening is lubricant used in tapping a hole
drilled in the shell for attachment of the torch fixture.
The abrasive cutoff tool shown in Fig. 1.13 was used to
remove sections of six tubes for the detailed examina-
tions described in Sect. 13.6. Several other tubes were
cut and raked out into the reactor cell to permit
viewing inside the heat exchanger. The periscopic view
showed the inside of the shell, baffles, tubes, and

8. MSR Program Semiannu. Progr. Rep. Aug. 31, 1968,
ORNL4344, p. 31.

Fig. 1.12. Opening in primary heat exchanger, showing cut made with plasma torch and exposed tubes.
PHOTO 101333

Fig. 1.13. Abrasive cutoff tool for removing heat exchanger
tube section.

cross-lacing. There seemed to be a slight coating on
everything, but otherwise nothing unusual was ob-
served.

The opening in the heat exchanger shell was closed by
welding on a patch of %-in.-thick type 304L stainless
steel. A sound weld was obtained around most of the
periphery, but at some spots there was porosity. After
repeated attempts to obtain a good seal by adding filler
metal were unsuccessful, iron-filled epoxy putty was
applied. A soap test at 5 psig then disclosed no leaks.

1.5 LEAK AT FREEZE VALVE FV-105

When the leak appeared during the final shutdown on
December 12, 1969, evidence pointed to freeze valve
FV-105 as the probable location.” This was confirmed
during the postoperation examination as soon as the
heater-insulation units were removed from the salt lines.
None of the units was especially radioactive except for
the one covering FV-105, which was significantly
contaminated, and a visual scan of the lines showed a
blob of salt in the suspected area, immediately adjacent
to the cooling air shroud. Close inspection through a
periscope revealed the knobby lump shown in Fig. 1.14,
suggestive of a slow leak in a freezing environment. On
the other side of the pipe, there was, in addition to
more knobby material, a thin sheet of salt that had run
out on the flat base under the removable insulation
unit. The total amount of salt that had leaked was
estimated at between 2 and 3 in.% .

The section of 1%-in. pipe containing the freeze valve
was cut out with an abrasive cutoff tool and delivered
to the HRLEL. Examinations (described in detail in
Sect. 13.7) showed the leak to be a crack in the pipe
just outside the shroud, where large cyclic stresses were
produced by differential thermal expansion of the
shroud and pipe during operation. These stresses were
higher than in the development model that had been
tested because a late modification to increase cooling
air flow'® had the unintended effect of greatly
strengthening the shroud. This had been overlooked,
and the failure occurred after many fewer cycles than
had been considered allowable.

1.6 OTHER EXAMINATIONS

One of the uncertainties in the tritium balance in the
MSRE'' was the amount of tritium produced from
lithium in the thermal insulation around the reactor
vessel. In order to measure the lithium content, samples
of the insulation were obtained by taking out one of
the removable slides of the thermal shield. The cooling
water lines were cut with a torch, and the slide was
lifted out without difficulty. Working through the
maintenance shield, we chiseled the metal can open and
took three samples of the block insulation. Lithium
analyses ranged from 12 to 38 ppm, which meant that

9. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL4548, pp. 3, 14.

10. MSR Program Semiannu. Progr. Rep. Feb. 28, 1965,
ORNL-3812, p. 28.

11. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, pp. 9-10.

PHOTO 023071

Fig. 1.14. Closeup view of salt near freeze valve FV-105. The drain line and %-in.-diam thermocouple sheaths loom at upper left.

only a small fraction (less than 5%) of the 3 to 5 Ci/day
of tritium appearing in the cell atmosphere could have
been produced in the insulation; nearly all must have
diffused out through the walls of the fuel system.
Tritium in the samples indicated that the total retained
in the 500 kg of insulation was only 2 to 3 Ci.

The outside of the reactor vessel was examined, using
a periscope inserted through the thermal shield slide
opening. All visible surfaces of the vessel were smooth
and seemed to have remarkably little oxide, being
noticeably lighter in color than the 5-in. salt pipe into
the vessel, which had a uniform black surface. The drain
line, between the reactor vessel and its penetration of
the thermal shield, showed no visible effect of opera-
tion. The same was true of the heater sheaths and
thermal insulation cans.

Several sheathed thermocouples were removed from
various locations in the MSRE for testing to determine
effects of radiation and/or high temperature. Thermo-
couple TE-R52, which had been in a well in the core

access plug, was the most highly irradiated of those that
were recovered. Two thermocouples were chiseled off
the core inlet line and four off the heat exchanger, but
attempts to detach one of the thermocouples welded to
the outside of the reactor vessel were unsuccessful.

1.7 EVALUATION OF TOOLS AND PROCEDURES

It was not required that the postoperation examina-
tions leave the fuel system in operating condition or
able to be restored. The patches on the openings in the
fuel system were required to be reasonably leak-tight
but not necessarily resistant to very high temperature
and radiation. This simplified the tools and procedures
but limited the demonstration of radioactive mainte-
nance capability. Even so, the experience was valuable.

The technique used in the examinations was basically
the same as had been used for maintenance of the
radioactive systems: long-handled tools, operated

through a portable work shield set up over a hole in the
cell roof. Most of the tasks were new, however, and
required specially designed tools. In addition to the
usual socket wrenches, hooks, etc., there were air-
motor-driven abrasive wheels in several configurations, a
hydraulically operated shear, oxyacetylene cutting
torches, a ‘plasma torch, and Heliarc welding torches.
Most of the special tools are shown arrayed on top of
the drain tank cell in. Fig. 1.15. Not shown are the
periscopes and the plasma torch with its fixture for
mounting on the heat exchanger.

The removal of the control rods, rod thimbles, and
core graphite had been provided for in the reactor
design, and few difficulties were encountered. The only
one that would have been troublesome if the aim had
been to resume operation was the unexpected inter-
ference of some flexible air lines with the removal of
the containment standpipe.

The remote removal and replacement of the fuel
pump motor and rotary element had been tested in
1965, and the relative ease with which the flange bolts
were removed during the postoperation work suggested
that this task was still practical. Replacement of the
pump. bowl and piping was a job that had been
anticipated but not tried. The failure of the built-in
devices for removing the sampler tube spool piece
would have added to the already formidable propor-
tions of this task.

14

that the pipe ends were left clean and square, ready for
tapping or plugging. The oxyacetylene torch proved
quite satisfactory for rough cutting, including the 2-in.
steel support plate for the fuel pump, the cooling air
shroud, and the water lines to the thermal shield slide.

The plasma torch used on the heat exchanger shell
was a commercially available unit: only the device for
moving it over the surface was specially designed for
this job. Development tests on a mockup proved the
device capable of maintaining a constant separation
distance and established the proper control settings. In
the reactor cell, however, considerable difficulty was
met in starting the cut. On the first five attempts,
backspatter of molten metal ruined the torch nozzle,
which had to be replaced. This was complicated because
of radioactive contamination on the torch. Finally the

. gap between the torch and the shell, which had been set

Cutting Hastelloy N with long-handled tools was |

something new. This alloy is characteristically difficult
to cut, and under the conditions in the reactor cell, the
positive advance of the cutter necessary to avoid work
~ hardening could not be guaranteed. Abrasives work
satisfactorily on Hastelloy N, however, so the rotating
grinder (Fig. 1.7) was chosen over a hole saw for cutting
out the fuel pump sampler cage. When the ground-off
latch stop was tapped for attachment of the tool, even
though lubricant was applied, the torque required was
so high that elastic torsion of the pipe on which the tap
was mounted made the job difficult. During the
trepanning operation, grinding wheels broke more
frequently than in the mockup tests. It appeared that
bits of slag from the torch cuts on the obstructions
overhead were being dislodged by the vibration and
falling into the trepan groove. The cut took about 8 hr
grinding time, but replacement of three broken wheels
extended the working time to several days.

The grinding wheels worked well in cutting the
sampler tube, heat exchanger tubes, and the fuel lines
around the freeze valves. Two of the cuts in the last
area were made without difficulty through a pipe filled
with frozen salt. An advantage of cutting this way was

at "¢ in., was reduced to % in., and the cut was
made with no difficulty. The jet from the torch
extended far enough to sever a few tubes. Spatter was
confined to a zone near the cut, but there was a
widespread, nonuniform light coating inside the heat
exchanger that was probably condensed material vapor-
ized by the torch (see Sect. 13.6).

Patching the heat exchanger shell was a challenging
task. The original plan was to cut a ' -in.-thick piece of
curved plate fitting into the hole left by the plasma
torch. Then a Heliarc torch, guided around the path
followed by the cutting torch, would fuse the patch and
shell. After mockup tests showed that it was very
difficult to get an adequate fit, however, it was decided
to try an overlapping patch. The overlap unfortunately
ruled out use of the existing torch guide to maintain a
fixed gap, but successful welds were made by a welder,
watching through a periscope and controlling the gap
by a rope attached to the torch. Although this worked
well in the mockup, troubles were encountered in the
reactor. The shell around the hole was prepared with
brushes and sanding disks mounted in an electric drill,
and the patch was set in place without difficulty. When
welding was started, however, the edges of the patch
tended to melt and run off without adhering to the
shell. Changing the angle of the torch to put more heat
on the shell helped, and a good-looking fusion weld was
obtained around most of the periphery despite the
awkward working arrangement. In places, however, the
patch had melted back so far that gaps were left.
Attempts to close these up using welding rod were not
completely successful as shown by a pressure test. After
more fruitless efforts at weld repair, a good seal was
made with iron-filled epoxy, backed up by a “dam” cut
from a stainless steel plate that had been rolled to fit on
the heat exchanger shell around the patch. The radia-
tion level on the shell was 630 R/hr, which would not
cause deterioration of the epoxy for several years.

The severed fuel drain lines were plugged with devices
like the one shown in Fig. 1.16. A conical plug of soft
copper was forced into the end of the 1%-in. pipe by a
jack bolt through a bracket clamped to the pipe. A tube
connected to a hole through the plug extended up to
the top of the cell, where temporary connections could

be made for gas addition. Each of the three lines leading
to the salt tanks extended down into the frozen salt and
5o could be pressurized to test the plugs. There were
some problems in installation due to brackets breaking,
but satisfactory seals were obtained in every case.

The hydraulically operated shear worked well in tests
but was not used in the reactor because the anticipated
need did not materialize.

PHOTO  0227-71

B
WAL LABORATORY

Fig. 1.16. Device for plugging severed ends of salt lines.

2. Further Investigations

During operation of the MSRE there arose two
discrepancies which we hoped to resolve during the
postoperation examination. One was the discrepancy
between the reactor power indicated by the heat
balance and that indicated by the changes in ratios of
nuclides in the fuel. The other was a difference between
the amount of uranium that was stripped from the fuel
salt in 1968 and the amount that was later recovered
from the UF4 absorbers.

2.1 TEST OF COOLANT SALT FLOWMETER
AND CONCLUSIONS

C. H. Gabbard " P, N. Haubenreich

The power level of the MSRE was routinely obtained
from a heat balance on the reactor cell by the on-line
computer. Using the finally accepted value of the
specific heat of the coolant salt, this computation
indicated that the maximum power (heat removal) was
about - 8.0 MW. Serious doubt was cast on this value,
however, when analysis of the changes in uranium and
_ plutonium isotopic ratios over long periods of operation
indicated that the power had been 7 to 8% less than
that indicated by the computer heat balance. Heat
removal in the coolant salt was the preponderant term
in the heat balance (98% of the total at full power), and
suspicion was directed at the coolant salt flow measure-
ment. _ :
The volume flow rate of coolant salt was measured by
a venturi flowmeter in the 5-in. pipe near the radiator
inlet, There were two readout channels, each consisting
of two pressure taps in which molten salt transmitted
pressure through metal diaphragms to NaK-filled lines
leading out of the heated zone to a differential pressure
cell. Associated electronics produced a signal that
ranged from 2 to 10 V in proportion to the square root
of the pressure differential. In the computer this voltage
was converted to flow rate by a multiplicative factor
based on the manufacturer’s calibration of Ap vs flow
in the venturi and a voltage-Ap relation for the
instrument. that had been determined before operation.
Near the end of nuclear operation, a review of the
flow-Ap calibration disclosed a mistake that was in-
troducing an error of —2.9% in the indicated flow, and

17

K

'whose correction only increased the discrepancy be-

tween heat balance power and the nuclide indications.’
The Ap-voltage calibration of the flowmeter readout
instrumentation could not be checked until the salt
piping could be cut.

During the postoperation examinations, the 5-in.
coolant pipe was removed on either side of the venturi,
and a plug was installed in the converging section and
one upstream so that pressure could be applied to the
upstream taps. The pressure differential between the
upstream and throat taps was raised and lowered
through the normal range while Ap and voltages were
accurately measured. It was found that the voltage
which had been interpreted as 850 gpm was produced
by a pressure differential equivalent (according to the
corrected flow-Ap calibration) to a flow of only 793
gpm, or an error in the Ap-voltage calibration of +6.7%.

In addition to the errors of —2.9 and +6.7% in the
measured salt flow rate, an error of +0.4% in the salt
temperature rise had been indicated by an experiment
on the effects of isothermal system temperature
changes on thermocouple biases.? The net effect of
correcting - for these errors was to change the heat
balance power from 8.0 to 7.65 MW. .

Ragan had calculated nominal full power values of
7.30 + 0.10 MW from the 23°U depletion and 7.45 *
0.18 MW from the 2°® U buildup for a weighted average
of 7.34 + 0.09 MW.? Thoma and Prince inferred a value
of 7.41 MW from the changes in 2*°Pu/23°Pu ratios
and observed that changes in the 234 U/?33 U ratio were
in excellent agreement with this value.* The 7.65-MW
value from the finally corrected heat balance is higher
than the nuclide values by about 3%. Uncertainties in
cross sections, fission energies, salt density, salt specific
heat, venturi calibration, and d/p instrument calibration
are enough to account for the difference.

1. MSR Program Semiannu. Progr. Rep. Aug 31, 1969,
ORNL-4449, p. 12. _

2. C. H. Gabbard, Reactor Power Measurement and Heat
Transfer Performance in the MSRE, ORNL-TM-3002 (May
1970).

3. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, pp. 65-66.

4, Ibid., pp. 98—-102.
It is unfortunate that throughout the operation of the
MSRE, changes from time to time in the best estimate
of the power level required reworking of detailed
analyses of reactivity, heat transfer performance, and
fuel chemistry. There appears to be no prospect of
further reducing the uncertainty that still remains. The
nominal full power of the MSRE was probably between
7.4 and 7.6 MW. Analyses and reports that have been
made since the last few months of operation have
generally used a nominal full power of 7.34 or 7.4 MW,
and there is insufficient cause to redo this work.

2.2 INVENTORIES OF RESIDUAL URANIUM
AND PLUTONIUM

R. E. Thoma

As mentioned in the preceding section, we deduced
from comparisons of observed changes in the isotopic
composition of plutonium with nominal values that the

-maximum power output of the MSRE was 7.4 MW(t).

After obtaining this value, we computed a material
balance on the 228U + 23°U that was in the fuel salt
from 1965 to 1968. This balance indicated that the
amount that left the fuel salt during the chemical
processing operations in 1968 was some 4 kg more than
the amount that was recovered from the UFq ab-
sorbers.® As will be described in a forthcoming report
summarizing experience with the MSRE.® refinement
of that assessment resulted from isotopic-dilution anal-
yses, which showed that the amount of 23®U loaded
into the MSRE was actually about 2 kg less than was
credited in on-site records. The resultant material
balance is presented in Table 2.1.

Recovery of uranium was carried out by Goodyear
Atomic Corporation, Piketon, Ohio, by dissolution of
the NaF pellets in which the UF¢ was absorbed. Special
precautions were taken to minimize measurement un-
certainties,” and a review® by Goodyear of the oper-
ations did not reveal any reason for revising the
originally reported amount of uranium recovered. We
have reexamined the possibilities of such errors as might
be ascribed to misestimates of power output of the
reactor and the implications of reactivity anomalies and

5. R. E. Thoma, internal correspondence to M. W. Rosenthal,
Sept. 14, 1970.

6. R. E. Thoma, Chemical Aspects of MSRE Operations,
ORNL-4658 (in press).

7. Letter GAT-532-70-212 from W. B. Thompson, Goodyear
Atomic Corp., to R. B. Lindauer, ORNL, Sept. 2, 1970.

8. Letter GAT-510-70-153 from C. D. Tabor, General
Manager, Goodyear Atomic Corp., to R. V. Anderson, Manager,
Portsmouth Area, AEC, Nov. 11, 1970.

18

Table 2.1. Material balance for uranium in the fuel sait,
1965—-1968

Kilograms of

uranium

Charge at initiation of power operation 227.020
Additions as fuel replenishment +2:461
Transferred to flush salt -6.272
Removed in samples? ~0.256
Consumed in 9005 EFPH (equivalent full -3.594
power hours) of operation? I
Net at end of 235U operation 219.359
Retained and mixed with 233U charge€ —1.935
Removed during processing 217.424
Recovered from UF¢ absorbers (max) 214.776
Disparity 2.648

4]. R. Engel, MSRE Book Uranium Inventories at Recovery of
235U Fuel Charge, internal memorandum MSR-68-79 (May
1968).

bMSR Program Semiannu. Progr. Rep. Aug 31, 1969,
ORNL-4449, p. 25. ‘

“MSR Program Semiannu. Progr.
ORNL-4396, p. 131.

Rep. Feb. 28, 1969,

short-term trends in the results of chemical analyses;
their possible contribution to the disparity indicated in
Table 2.1 is negligible. Analysis of scrubber solutions
during the processing showed less than 1 g of uranium
discarded. The conclusion, therefore, is that the ma-
terial balance affords unequivocal evidence that some
2.65 kg of uranium (33.08 wt % 23°U) remains at the
MSRE, possibly in the chemical processing plant.
Although there was no direct evidence to support
uranium retention in the processing equipment, two
components have been identified as conceivable sites
for such retention. Engel noted that the particle filter (a
9-ft? filter designed to remove corrosion product solids
from the fluorinated salt before its reuse in the reactor)
in the line between the fuel storage tank and the
processing tanks could, after treatment of the flush salt
was completed, have contained an unknown amount of
zirconium metal, delivered to this location as the
processed flush salt was returned to the reactor system.
It is difficult to assign high probability to the events
which could have reduced the uranium from the fuel
charge as, subsequently, it passed through this filter in
such a.way that some 2.5 kg of uranium remained in
the filter; however, the possibility cannot be excluded.
Another possible site where uranium may have been
retained, as suggested by R. B. Lindauer, is the
high-temperature sodium fluoride absorber bed, which
is positioned between the fuel storage tank and the NaF
absorbers. The design temperature for operation of this
19

Table 2.2 Inventory of residual uranium and plutonium in the MSRE#

Uranium inventory

233U 234U 235U 236U 238U sU
Fuel circuit inveritory, run 20-1, kg 28.568 2.526 0.869 0.036 2.020 34.019
Total inventory, run 20-1, kg 31.052 2.746 0.945 0.039 2.196 36.978
Drain tank inventory, run 20-1, kg 2.484 0.220 0.076¢ - 0.003 0.176 2.959
Fuel circuit inventory, run 20-F, kg 28.406 2.533 0.866 0.036 2.011 33.852
Transfer to flush salt, run 20-F, kg 0.411 0.037 0.013 0.061 0.029 0.491
Charged into drain tank, run 20-F, kg 27.995 2.496 0.853 0.035 1.982 33.361
Drain tank residue, kg 2.484 0.220 0.076 0.003 0.176 2.959
Final drain tank inventory, kg 30.479 2.716 0.929 0.038 2.158 36.320

U/zU, wt % 83.918 7.478 2.558 0.105 5.941
Plutonium inventory

239Pu 240py 238,241,242Pu > Pu
Fuel circuit inventory, run 20-1, g 625.8 61.81 2.39 690.0
Total inventory, rtun 201, g 680.2 67.19 2.60 749.99
Drain tank inventory, run 20-1, g 54.4 .5.38 0.21 59.99
Fuel circuit inventory, run 20-F, g 615.6 65.43 2.37 683.4
Transfer to flush salt, run 20-F,P g 61.8 6.57 0.23 68.6
Charged into drain tank, run 20-F, g 553.8 58.86 2.14 614.8
Drain tank residue, g 54.4 5.38 0.21 59.99
Final drain tank inventory, g 608.2 64.24 2.35 674.79

Pu/ZPu, wt % 90.13 9.52 0.35

Weights are based on comparisons of analytical results and computed values. These comparisons indicate maximum power
output as 7.4 MW(t). Final estimates assume 4167 EFPH at 7.4 MW(t).
bThis item makes the simplifying assumption that the total amount of plutonium estimated to be transferred to the flush salt was

transferred during the final flush of the fuel circuit.

Table 2.3. Composition of fuel salt stored in the MSRE drain tanks

TLiF BeF, ZrF, 233.42yF, 239.11pyp,
Composition, mole %
Nominal 64.50 30.180 5.199 0.132 . 2.38 x 1073
Analytical® 64.53 30.43 4.90 0.137
Composition, wt %
Nominal 41.87 35.44 21.67 1.0195 0.0177
Analytical® 41.37 35.27 20.19 1.06

2Current calculations do not include corrections for transfer of carrier solvent residues to flush salt nor flush salt residues to fuel.
Disparity between nominal and analytical values for zirconium will be reduced by introduction of this correction factor.

bRun 17-20, average of 33 samples.

absorber is 750°F, based on previous laboratory
studies.” The laboratory studies indicate that this
absorber would not retain UF4 at the operating
temperature; only the possibility that temperature
gradients prevailed within the absorber at periods near
the end of fluorination operations and allowed the
retention of some uranium within the absorber gives

9. MSR Program Semiannu. Progr. Rep. Aug. 31, 1966,
ORNL4344, p. 321. -

any credence to the likelihood that uranium wou}d be
retained here. Attempts to locate uranium in the
processing plant are described in Sect. 2.3.

While confirmation of the amount of uranium in the
processing plant by direct experiment would be de-
sirable, the isotopic dilution analyses and the data
which were used to monitor the transfers of uranium
and plutonium within the reactor system appear to be
sufficiently reliable for estimating the amounts of 23°U
and 238U retained in the reprocessing system and for
computation of final inventory distribution. "

In preparation for phase Il of the program for
decommissioning of the MSRE, inventories of uranium
and plutonium in the stored salt charges were esti-
mated. Weights and isotopic composition of uranium
and plutonium in the drained fuel and flush salts are
listed in Table 2.2. Using these weights and a value of
4707.5 kg as the weight of the fuel carrier salt, the
composition of the fuel salt was computed. Nominal
values are compared with analytical data in Table 2.3.

2.3 SEARCH FOR UNRECOVERED %3°U
J. R. Engel

The data and analysis reported in the preceding
section clearly showed that the removable absorbers in
which the original charge of uranium was recovered as
UF¢ contained significantly less uranium than had been
in the fuel salt at the end of nuclear operation with
2357U. We sought to determine more accurately how
much uranium had been left behind and where it was
located. As a first step we hoped to examine the
particle filter and the hot NaF trap.

Since it was necessary to leave the processing plant in
an operable condition for ultimate disposal of the fuel
and flush salts, a nondestructive technique was required
to look for the uranium. Neutron interrogation has
been developed as a highly useful technique for assaying
the fissile-material content of closed containers. Nor-
mally the technique is applied to relatively portable

20

objects that can be examined under carefully controlled
conditions with high-sensitivity neutron detectors and
intense sources. For this particular application, the
interrogation would have to be performed in situ under
relatively crowded conditions in a concrete cell. Never-
theless, it appeared to offer the only hope for a
nondestructive search.

Preliminary investigations indicated that useful in-
formation might be obtainable with a relatively simple
experimental arrangement. Therefore it was decided to
proceed with an experiment using an isotopic neutron
source and a single neutron detector. The neutron
source to be used contained ~14 ug of 232Cf, which
produced about 3 X 107 neutrons/sec. The neutron
detector was a boron-lined chamber with an absolute
efficiency of 15 counts per nvt. Tools were developed
and built to permit positioning and manipulating the
source and detector in the processing cell near each of
the two components.

Because of the complexity of the geometries involved
and the expected importance of extraneous scattered
neutrons from the source on the detector response, a
simplified mockup of the salt filter was built to measure
the absolute sensitivity of the interrogration equipment.
A clean ORR fuel element was used as the uranium-
bearing target in the mockup. Measurements with this
system revealed that the sensitivity was too low to
provide reliable information about the components in
the processing cell, and the investigation was dis-
continued. Consideration will be given to other
methods of locating and recovering the unaccounted-for
uranium during the final phase of the MSRE de-
commissioning,
~Part 2. MSBR Design and Development

'R. B. Briggs

The design and development program has the purpose
of describing the characteristics and estimating the
performance of future molten-salt reactors, defining the
major problems that must be solved in order to build
them, and designing and developing solutions to prob-
lems of the reactor plant. To this end we have done a
conceptual design for a 1000-MW(e) plant, and the
report describing the plant is in press. A contract is now
being negotiated with an industrial group to do a
conceptual design of a 1000-MW(e) MSBR plant using
the ORNL design for background and incorporating the
experience and the viewpoint of industry. One could
not, however, propose to build a 1000-MW(e) plant in
. the near future, so we are doing studies of plants that
could be built as the next step in the development of
large MSBRs. One such plant is the Molten-Salt Breeder
Experiment (MSBE). The MSBE is intended to provide
a test of the major features, the most severe operating
conditions, and the fuel reprocessing of an advanced
MSBR in a small reactor with a power of about 150
MW(t). An alternative is the Molten-Salt Demonstration
Reactor, which would be a 150- to 300-MW(e) plant
based largely on the technology demonstrated in the
Molten-Salt Reactor Experiment, would incorporate a
minimum of fuel reprocessing, and would have the
purpose of demonstrating the practicality of a molten-
salt reactor for use by a utility to produce electricity. In
addition to these general studies of plant designs, the

design activity includes the assessment of the safety of
molten-salt reactor plants. Some studies related to
safety are in progress preliminary to a comprehensive
review of safety based on the design of the 1000-MW(e)
MSBR.

The design studies serve to define the needs for new
or improved equipment, systems, and data for use in
the design of future molten-salt reactors. The purpose
of the reactor development program is to satisfy some
of those needs. Presently the effort is concerned largely
with providing solutions to the major problems of the
secondary system and of removing xenon and handling
the radioactive off-gases from the primary system. Work
is progressing on the design of one loop facility for
testing the features and models of equipment for the
gaseous fission product removal and off-gas systems and
for making special studies of the chemistry of the fuel
salt. Design is nearing completion for a second loop
facility for studies of equipment and processes and of
the chemistry of sodium fluoroborate for the secondary
system of a molten-salt reactor. The steam generator is
a major item of equipment for which the basi¢ design
data are few and the potential problems are many. A
program involving industrial participation is being un-
dertaken to provide the technology for designing and
building reliable steam generators for molten-salt re-
actors.

3. Design

E. S. Bettis

3.1 SINGLE-FLUID 1000-MW(e) MSBR
DESIGN STUDY REPORT

Roy C. Robertson

The report' covering the design and evaluation
studies of a 1000-MW(e) molten salt thermal breeder
reactor power station in which the fissile and fertile

- materials are incorporated in a single fluoride salt has

21

received final editing and approval and is now in the
process of being published. Distribution is scheduled for
June 1971.

1. Molten-Salt Reactor Program Staff, Conceptual Design
Study of a Single-Fluid Molten-Salt Breeder Reactor,
ORNL-4541 (in press).
3.2 MOLTEN-SALT DEMONSTRATION
REACTOR DESIGN STUDY

E. S. Bettis

H. A. McLain
J. R. McWherter
H. L. Watts

C.E. Bettis
C.W. Collins
W. K. Furlong

3.2.1 Introduction

Design and evaluation studies of a 300-MW(e) molten-
salt demonstration reactor (MSDR) were continued.
This power station would be a first-of-a-kind prototype
to demonstrate the feasibility and delineate the prob-
lems of construction of a large-scale molten-salt reactor
power station. The prototype studies have concentrated
on concepts which would require a minimum of
development and would permit construction of the
demonstration plant in the near future. This aspect led
to our decision to design the MSDR as a low-power-
density converter rather than a breeder, since this could
be done without a significant penalty on the short-term
fuel cycle costs, yet would eliminate the need for core
graphite replacement during the 30-year life of the
" plant and for sealing of the graphite against gas
permeation to reduce the '*%Xe poisoning. The con-
verter could also substitute periodic salt replacement
for continuous fuel salt processing until a suitable

chemical plant was fully developed. The conversion

ratio without processing would be about 0.8.

The MSDR general flowsheet, plant layouts, reactor
vessel, primary heat exchangers, and drain tanks have
been described previously.? During the past report
period we decided to make several revisions to the
concept, however, the most notable of which was to
interpose an additional salt circulation loop between
the primary heat exchangers and the steam generators
in order to assure confinement of the tritium formed in
the fuel salt. The building structure was altered to
accommodate the new loop, the drain tank and
gas-handling systems were modified, and the systems
for heating the cells were revised. The revised flowsheet
is shown in Fig. 3.1.

No revisions have been made in the reactor concept
itself since last reported.? The reactor core is 21 ft in
diameter and 21 ft high and is surrounded by a
2Y, -ft-thick graphite reflector. The all-welded reactor
vessel is fabricated of Hastelloy N fortified with
additives to improve the resistance to radiation damage
(see Sect. 15).

2. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548.

22

3.2.2 Addition of Third Salt-Circulation Loop

Although studies of the amounts of tritium that could
diffuse from the fuel salt into the coolant salt and
steam system are not complete and the various meas-
ures that could be used to mitigate the migration have
not been fully evaluated, it was decided for the present
to incorporate into the MSDR concept an assured
method of tritium confinement. Chemical considera-
tions indicate that an efficient tritium trap can be
obtained by use of a nitrate-nitrite salt mixture in a
loop between the secondary system and the steam
system, as shown in the flowsheet, Fig. 3.1. The salt
would oxidize the tritium to water and prevent its
reaching the steam systemn. The nitrate-nitrite salt is
believed to be unsuitable for use in the secondary
system because it decomposes rapidly as the tempera-
ture is raised above 1100°F and it would react
vigorously with the graphite if it were to leak into the
primary system.
 Addition of another set of heat exchangers, circu-
lating pumps, and connecting piping has the obvious
disadvantages of added complexity and cost, but there
are desirable features of the new arrangement besides
trapping the tritium, and some of the aspects tend to
partially offset the capital cost of the equipment
needed for the third loop.

With the steam-raising equipment in a thlrd loop the
equipment in the secondary system can be closely
coupled, the inventory of secondary salt can be -
considerably reduced, and use of more expensive
coolants with higher melting temperatures can be
considered. The most obvious of these is the ’ LiF-BeF,
(66-34 mole % with 99.99% 7Li) salt mixture which
performed well as a coolant in the MSRE. Although it
has a high melting peint of about 850°F and a relatively
high cost of about $11.40 per pound (based on a "Li
cost of $120 per kilogram), it has the decided advantage
that a leak of the coolant into the fuel salt
(LiF-BeF,-ThF4-UF,) becomes of considerably less
significance. On this basis, " LiF-BeF, was selected as
the fluid for the MSDR secondary circulation system
shown in the revised flowsheet (Fig. 3.1).

The tertiary circulation loop, which transports heat
from the secondary heat exchangers to the steam
generators and reheaters, uses a commercially available
nitrate-nitrite heat transfer salt, KNO;-NaNO,-NaNO,,
with the eutectic composition 44.2-48.9-6.9 mole %. In
addition to being an effective chemical trap for tritium,
the salt has a desirably low liquidus temperature of
288°F, a relatively low cost of about 15¢/Ib, and good
heat transfer and flow properties (see Table 3.1). There
are three outstanding advantages to its use in transporting
ENTRAINMENT
SEPARATOR

TERTIARY
SALT PUMP

SECONDARY
SALT PUMP

1100°F
23,000 gpm

1000°F
27,180 gpm

HEAT REJECT TANK

ORNL—DWG 71—3633

GAS
SEPARATOR REHEATER
PRIMARY SECONDARY
HEAT HEAT
EXCHANGER EXCHANGER ~_700°F
1050° F 900°F
I__J'_"‘I
|
| | GAS CLEANUP -
| |~ SYSTEM
| .
) N
> - TO CHEMICAL
| ¥|| - «— PROCESSING SYSTEM
H
|
i1l PRIMARY
1 {| DRAIN TANK
ip3

[}

Fig. 3.1. Preliminary flowsheet for MSDR power plant.

900°/ 900°F .
2400 psi N
HIGH
PRESSURE
TURBINE

SUPERHEATER

MOISTURE
SEPARATOR

INT. AND LOW
PRESSURE
TURBINES

!-_'__'I

L

FEEDWATER
CHAIN

GENERATOR

]

i

|

CONDENSER

€T
heat to the steam generators and reheaters: (1) in the
event of small steam leaks, there is no chemical reaction
with water, (2) corrosion rates at high temperature are
very low with materials which are less expensive than
Hastelloy N, and (3) the low melting temperature
permits use of feedwater and cold reheat steam at
conventional temperatures. These three advantages are
discussed below. .

The steam pressure will exceed the pressure in the
tertiary heat transport loop; thus tube failures or
leakage would cause water to mix with the coolant salt.
At the operating temperatures and pressures in the
nitrate-nitrite loop, however, essentially all the contami-
nating water would exist as steam above the surface of
the salt, and the vapor could be purged, taking care, of
course, to prevent the escape of tritiated water. The salt
would not react with the water to produce highly
corrosive conditions.

Hastelloy N has been chosen as the material of"

construction in all previous designs of steam generators
and reheaters for molten-salt reactor power systems.
Use of the nitrate-nitrite heat transport salt would
make it possible to use less expensive materials. For
example, the data available to date indicate that the salt
would be compatible with Incoloy 800, a material that
has proven acceptable for service in high-temperature
steam systems. The substitution of Incoloy for Hastel-
loy N could make significant savings in the capital costs
of equipment.

The previous MSBR conceptual design employing
sodium fluoroborate coolant salt to transport heat to
the steam generators provided a minimum feedwater
temperature of 700°F. Although this minimum has not
been established experimentally, the 725°F liquidus
temperature of the salt is likely to dictate feedwater
temperatures well above those in conventional regenera-
tive feedwater heating systems. In the steam system
flowsheet for the reference design MSBR, the cold
reheat steam is preheated from 550 to 650°F by heat
exchange with 3500-psia 1000°F steam taken from the
throttle supply, and the feedwater is heated to 700°F
by direct mixing with the exit heating steam from the
reheat steam preheater. If the nitrate-nitrite salt, with
its liquidus temperature of 288°F, were used to
transport the heat, conventional feedwater and cold
reheat steam temperatures could be used. This would
effect important savings through elimination of the
reheat steam preheater, the pressure booster pumps,
and the mixing chambers employed in the special
MSBR feedwater circuit. Since the aforementioned
direct-mixing arrangement for feedwater heating fa-

24

vored use of supercritical pressure steam,’ elimination
of this aspect allows consideration of a 2400-psia cycle
on a more equal basis with the 3500-psia system,
although the latter is still probably preferred for the
MSDR. Although the startup of an MSDR has not been
studied to date, it seems probable that the less-
restrictive feedwater conditions of the nitrate-nitrite
coolant system would also make it possible to use a less
expensive type of startup boiler than was assumed in
the reference design MSBR.!

In summary, addition of a third salt-circulation loop
in an MSDR or an MSBR provides tritium control and
other important advantages. In particular, it could
simplify operation of the MSDR steam system. Were it
not for the tritium problem, however, the additional
loop probably would not be recommended because of
the added capital expense. A more definitive study is
now under way to evaluate the costs.

3.2.3 Salt Overflow and Gas Stripping Systems

A further major change in the flowsheet from that
previously reported? is elimination of the drain tank as
an overflow tank for the primary circulation system. All
salt volume changes in the circulation system are now
accommodated in the pump tanks. This revision was
made in the plant, because the jet pumps used to return
the overflow salt to the circulation systems would have
had to pump a large flow of salt against a head now
considered too large to be practicable in the MSDR.
Each primary pump tank has sufficient free volume to
accommodate a 10% total change in system salt volume.
In addition, each pump tank has an overflow line
connected to the reactor outlet plenum. During steady-
state operation, the cooling and fountain flows in the
pumps are returned to the circulation system via this
same route of low flow resistance.

3.2.4 Primary Drain Tank

With elimination of the continuous salt overflow from
the pump bowls, the drain tank could be modified to
provide only one entering salt line. This nozzle connects
with both the reactor drain line and with the reactor
cell catch basin drain line through valves that are
discussed below. A jet pump (activated by a salt flow
from a small salt-circulation pump located in the
chemical processing cell) is used for returning salt from
the drain tank to fill the primary system for startup.
The size of the drain tank, the drain tank crucible, and

- the arrangement of the cooling system are as previously

reported.?
3.2.5 Drain Valves for Salt Service

The drain valves in the main reactor drain line and in
the catch basin drain line are located in a small cell
between the reactor cell and the drain tank cell, as
shown in Fig. 3.2. The main drain valve for the reactor
is now visualized as a combination mechanical and
freeze valve. A bellows-sealed poppet, or plug, ap-
proaches a hard-faced surface but does not seat tightly
against it. A cooling liquid is then circulated through
the poppet to freeze an annular ring of stagnant salt
between the two faces to effect a tight shutoff. The
circulating liquid can be heated to thaw the salt to open
the valve. '

Two valves are used in parallel in the line from the
reactor cell catch basin to the drain tank. These valves
would not be opened except in the unlikely event of a
salt spill from the primary system and are sealed by a
thin membrane. If salt entered the pipe upstream of the
valves, a signal from an electrical conductivity probe
would cause a spring-loaded actuator to rupture the
membrane and permit the collected salt to enter the
drain tank. A normally open mechanical-type valve is
installed downstream from each of the rupture-type

25

valves to isolate them from the drain tank in the event
it becomes necessary to replace the rupture disks. The
bottom of the above-mentioned valve pit is also
provided with a catch pan, drain line, and rupture-disk-
type drain valve,

3.2.6 Heat Exchangers

Addition of the third salt-circulation loop, as dis-
cussed in Sect. 3.2.2, not only required a conceptual
design for the new secondary heat exchangers but,
through changes in the working fluids, made it neces-
sary to redesign the primary heat exchangers and the
steam generators. Also, in order to stay closer to MSRE
experience, we decided to reduce the maximum salt-
temperature in the MSDR from 1300 to 1250°F. A
further aspect was a change in the maintenance philoso-
phy to one in which we propose to plug failed tubes

. rather than to replace an entire tube bundle as was
previously planned.

The properties of the fuel salt and the two coolant
salts used in the heat exchanger calculations are given in
Table 3.1. The design data for the primary heat
exchanger are given in Table 3.2, and a sketch of the

Table 3.1. Physical properties of the fuel and coolant salts used in the MSDR

Fuel salt

Composition
Density, Ib/ft>

Viscosity, Ib hr 1 ft~!

Specific heat, Btu 1b~! (°F) -1
Thermal conductivity, Btu hr~! ft -1 (°F)~!

LiF-BeF,-ThF 4-UF 4 (71.5-16.0-12.0-0.5 mole %)
236.3 — 2.33 X 107 2T (°F)

7362

0.2637 exp ——————
637 eXP 9.7+ T CF)

0.324
0.75

LiF-BeF;, coolant salt

Compeosition

Density, 1b/ft>
Viscosity, Ib hr ™1 ft~!

Specific heat, Btu Ib~! (°F)"!
Thermal conductivity, Btu hr~! ft ! (OF) -1

TLiF-BeF; (66-34 mole %) (99.99+% "Li)
138.68 — 1.456 x 10727 (°F)
6759

0.2806 exp —— 2 ——
*P 459.7 + T °F)

0.57
0.578

Nitrate-nitrite coolant salt

Composition (eutectic)
Density, 1b/ft>

Viscosity, 1b hr ! ft!

Specific heat, Btu Ib~! (°F)-!
Thermal conductivity, Btu hr ! ft~1 (°F) 1

KNO3-NaNQ;-NaNQ 3 (44-49-7 mole %)
130.6 — 2.54 X 1027 (°F)
3821.6

0.1942 exp ————>—
P 4597+ TP

0.373
0.33

ORNL-DWG 71-5034

26

16-in. SLEEVE

PUMP COOLING, 2in.
SYSTEM DRAIN, 6in.

SYSTEM FILL, 2in.

OFF-6AS, 1Y in.

VALVE- CELL, 64-in. DIAM

{7
PaCiy!
%

0

EMERGENCY DRAIN VALVES

4...
a%eled

22

1
4
!
f
L
r
I
!

A
(Y g
KO

N

s

AN

-

———

MECHANICAL VALVE
{NORMALLY CPEN)

EMERGENCY DRAIN, 8in.

_-.h

THERMAL INSULATION

-—\J

SYSTEM DRAIN VALVE

|

CATCH PAN

<> XK A A
RIRIESK

TG

FROM FILL PUMP

\

TO DRAIN TANK

Fig. 3.2. Containment cell for valves in salt drain lines.
27

Table 3.2 MSDR primary heat exchanger design data

Type

Rate of heat transfer per unit
MW
Btu/hr

Tube-side conditions
Hot fluid
Entrance temperature, °F
Exit temperature, °F
Pressure drop across exchanger, psi
Mass flow rate, Ib/hr -

Shell-side conditions
Cold fluid .
Entrance temperature, F
Exit temperature, °F
Pressure drop across exchanger, psi
Mass flow rate, Ib/hr

Tube material

‘Tube OD, in.

Tube thickness, in.

Tube-sheet-to-tube-sheet distance, ft

Shell material

Shell thickness, in.

Shell ID, in.

Tube sheet material

Number of tubes

Pitch of tubes, in.

Total heat transfer area, ft2

Basis for area calculation

Type of baffle

Number of baffles

Baffle spacing, in.

Disk OD, in.

Doughnut ID, in.

Qverall heat transfer coefficient, U,
Btu hr™! ft72

Volume of fuel salt in tubes, ft3

U-tube, U-shell, countercurrent,
one-pass shell and tubes with
disk-and-doughnut baffles

125
4.2687 x 108

Fuel salt
1250

1050

127.4

6.588 x 106

Coolant salt (2LiF-BeF,)
900

1100

114.7

3.744 x 10°

Hastelloy N
0.375

0.035

29.96

Hastelloy N

0.5

26.32

Hastelloy N

1368

0.672 (triangular)
4023.6

Qutside of tubes
Disk and doughnut
47

7.66

19.0

18.6

700.7

20.78

exchanger is shown in Fig. 3.3. Use of the " LiF-BeF,
coolant salt in the shell side and the reduced tempera-
ture difference between the fuel salt and the coolant
salt altered the heat exchanger dimensions somewhat
from those of previous concepts. In order to keep the
shell diameter small, for ease of afterheat removal and
maintenance, we found it desirable to use two primary
heat exchangers in parallel in each of the three fuel-salt

circulation loops. The exchangers have a bolted, in- -

verted dished head over a seal-welded membrane. A
metal ring gasket in the head flange provides backup

containment in the event of a leak in the membrane.

This design makes maintenance easier than in the

previously reported concept that involved plasma arc
cutting of a 1-in.-thick heat exchanger head.

The secondary heat exchanger design data are given in
Table 3.3. The configuration is like that of the primary
heat exchanger. Since the LiF-BeF, heat transport
circuits will become much less radioactive than the
primary salt circuits, the remote maintenance require-
ments for this exchanger are not nearly as restrictive as
for the primary units.

The different parameters that can now be considered
in design of the steam generators were mentioned in
Sect. 3.2.2. We are investigating conceptual designs for
ORNL- DWG 71-5032 |

SECONDARY SALT

PRIMARY SALT INLET

1J\10-in. PIPE

}

—== | ft Bin, e

- 16 ft 73 in.

T~2ft 3%in.

13 ft ' -

HEAT TRANSFER RATE: 125 MW
TUBES: 1368 ’

TUBE SIZE: ¥g-0D x 0.035-in. WALL
TUBE PITCH: 0,67 4

Fig. 3.3. MSDR 300-MW(e) primary heat exchanger.

both 2400- and 3500-psi steam systems but as yet have
no results to report.

3.2.7 Building and Containment

The structures housing the MSDR are basically the
same as previously reported,? but the main contain-
ment building has been increased from 92 to 112 ft in
diameter and the heat exchanger cells from 23 to 33 ft
in diameter and to 38 ft in depth to accommodate the
third salt-circulation loop, as shown in Fig. 3.4. The
control room has been moved to outside the contain-
ment area, and some adjustments have been made in the
dimensions of other cells. The revised building plan and
elevation are shown in Figs. 3.5 and 3.6.

The reactor cell wall construction has been revised
and the cell heating system has been changed. The new
construction has % -in.-thick 304 stainless steel walls with
about 5 in. of blanket-type insulation between the wall
and the concrete biological shield, as shown in Fig. 3.4.
The flat tops of the cells are hung by Unistrut hangers
from the concrete. This design simplified the contain-
ment construction to provide a maximum of freedom
for thermal expansion and to present a smooth surface
on the inside of the cell. The integrity is more assured
both from the standpoints of inspection and access for
repair. As in the previous design, the cell will operate at
a slight negative pressure, but the structure would be
designed for a positive pressure of 50 psi.

Table 3.3. MSDR secondary heat exchanger design data

Type

Rate of heat transfer per unit

MW
Btu/hr

Tube-side conditions
Hot fluid

Basis for area calculation

Type of baffle

U-tube, U-shell, counter-
current, one-pass shell
and tubes with disk-and-
doughnut baffles

125
42687 X 108

. 2LiF-BeF, salt

Entrance temperature, °F 1100
Exit temperature, °F 900
Pressure drop across exchanger, psi 80
Mass flow rate, Ib/hr 3.744 x 10°
Shell-side conditions
Cold fluid Hitec
Entrance temperatgre, °F 700
Exit temperature, F 1000
Pressure drop across exchanger, psi 80
Mass flow rate, 1b/hr 3.815x 10 -
Tube material Hastelloy N
Tube OD, in. 0.375
Tube thickness, in. 0.035
Tube-sheet-to-tube-sheet
distance, ft 37.5
Shell material Hastelloy N
Shell thickness, in. 0.5
‘Shell ID, in. 30.5
Tube sheet material Hastelloy N
Number of tubes 1604
Pitch of tubes, in. 0.7188 (triangular)
Total heat transfer area, ft? 5904

Outside of tubes
Disk and doughnut

Number of baffles 52
Baffle spacing, in. 8.65
Disk OD, in. 22.0
Doughnut ID, in. 21.6
Overall heat transfer coefficient, '

U, Btu hr ! ft =2 - 501
Volume of 2LiF-BeF, salt in tubes,

ft3 30.5

Three external electrical furnaces and a circulating
hot gas system are used to heat the reactor, heat
exchanger, drain tank, and drain valve cells rather than
the electric resistance-heated thimbles formerly em-
ployed. The furnaces are located adjacent to the heat
exchanger cells, as shown in Fig. 3.5. The design data
for the heating system are shown in Table 3.4. The new
BUILDING

CONTAINMENT\

COOLING WATER

PIPES——__ |

L § [ ]
CONCRETE :
COOLING WATER OUTLET INSPECTION ACCESS PLUG PUMP SLEEVE
PUMP SLEEVE \ ( ,/ '

ORNL—-DWG 71—5036

[CONTNNMENT HEAD FLEXIBLE SUPPORT RODS

TO STEAM
GENERATING
SYSTEM

N

XA

&

3

P

% 4— CELL COOLING AND HEATING RETURN DUCT

THERMAL {NSULATION

CELL CONTAINMENT

[~

COOLING WATER QUTLET

HEAT EXCHANGER CELL

| /PIPE SLEEVE

REACTOR CELL

) CATATAN SN

HEAT EXCHANGER SUPPORT.

)

;/
.

N

COOLING WATER

PLENUM

COOLING WATER
INLET

COOLING WATER
INL

ET e £

COOLING WATER
CUTLET 4=

™ CELL CONTAINMENT

REACTOR DRAIN LINE

DA SO ST O TATAAA. ST T

Z

/

\\ DEFLECTOR

EMERGENCY
DRAIN PORTS

COOLING
e WATER
sl PIPES
¥
4 REACTOR SUPPORT THERMAL INSULATION -—*"
4 /
> il

N AVAVAV. O ATAVAVLN. 90 6 6 0 6 8 44

AX A X P AT AVAV.N

DX XK ~

XX

k(ZOOLING WATER PLENUM

Fig. 3.4. MSDR 300-MW(e) reactor and heat exchanger cells — élevafior_l.

| INSPECTION ACCESS PLUG (3)

4— COOLING WATER INLET {3)

CELL COOLING

—+—— AND HEATING -

_\SUPPLY DUCT (3)

= COOLING WATER
OUTLET (3)

d___TO VALVE ANC DRAIN

TANK CELLS

62
30

ORNL-DWG 71-3632

CELL HEATING SYSTEM

PRIMARY AND SECONDARY
HEAT EXCHANGER CELL

NoK-WATER
HEAT EXCHANGER

CONTROL ROOM OFFICES AND
AUXILIARY EQUIPMENT BELOW

STEAM GENERATING CELL

Fig. 3.5. MSDR 300-MW(e) reactor complex — plan section A-A.
31

ORNL-DWG 71-5035

[P 12 ft CONTAINMENT : |

16 ft -
176 ft

Fig. 3.6. MSDR 300-MW(e) reactor complex — vertical section X-X.
Table 3.4. Heater design data for MSDR containment cells

Estimated normal containment heat loss (total), kW

Design heat loss, kW

Number of heater cells

Capacity of each heater cell, kW
Circulating gas

Gas inlet-outlet temperatures at heaters, °F
Gas flow rate through each heater cell, cfm
Heater element

Heater length, ft
Number of heater elements per heater cell
Heater element arrangement per heater cell

Pressure drop in circulating gas, in. H,O
Heater cell width and depth, ft

600

1200

3

400 -

N2

1000-1100

32,500 (at 1050°F)

1-in.-OD cartridge with
Incoloy 800 cladding

3

150

12rowsof 12 or 13
elements on-3-in. A pitch

9.6

3.25%X 2.6

Thermal conductivities for cell wall (k), Btu hr™! ft™! (°F) !

: 1/2 -in. stainless steel cell liner
5-in.-thick fiber glass cell insulation
Prestressed concrete with 2-in. sched 40

carbon-steel water cooling pipes on 6 in. centers

located 4 in. from inside concrete face

12.4

0.034

1.12 (concrete)
25.9 (steel)

Assumed heat transfer coefficient in water pipes, Btu hr ™! ft =2 (°F)~! 7.9

Maximum concrete temperature, °F

150

arrangement reduces the number of cell wall penetra-
tions and gives better access for repair of the electric
heaters. The heater elements are inserted into double-
walled thimbles welded into the furnace duct and are
therefore easily removed for replacement. The same gas
circulation system can be provided with cooling coils to
remove excess heat from the cell.

3.3 INITIAL TEMPERATURE TRANSIENTS
IN EMPTY MSBR “REFERENCE DESIGN”
MSBR HEAT EXCHANGERS

J. R. Tallackson

Additional calculations® to estimate the initial after-
shutdown temperature transients in these heat ex-
changers were completed.® As in the earlier analyses,
40% of the noble metal fission products (Nb through
Te) were assumed to be deposited uniformly on the
heat exchanger tubes. Figure 3.7 shows the estimated

3. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL4622.

4. J. R. Tallackson, Thermal Radiation Transfer of Afterheat
in MSBR Heat Exchangers, ORNL-TM-3145 (March 1971).

after-shutdown temperature growth in the 563-MW(t)
“reference design” heat exchanger. If used as a design
guide, the estimated peak transient temperature,
slightly less than 2150°F, is on the high side and should
be regarded as conservative. This is because, during the
initial phase of the transient, the elements comprising
the exchanger, the tubes and inner shell, the inter-
mediate shell, etc., were considered to be isolated and
insulated from each other and from the outside. The
method used to calculate this transient is described in
detail in ref. 4. The curve in Fig. 3.8 applies to the
141-MW(t) unit. This curve was inferred from Fig. 3.7
and drawn by inspection. The peak temperature during
the transient is expected to be about 1800°F. These
peak temperatures, which take into account the effects
of heat distribution and heat capacity, supersede the
higher steady-state peak temperatures reported earlier.?

The Hastelloy N heat exchanger shells, if not highly
stressed, will safely withstand temperatures up to

2150°F for short periods. If the current estimates of

fission product deposition remain unchanged or de-
crease, it can be concluded that heat exchangers, not
dissimilar to the “reference design” and with ratings in
the 500-MW(t) region, can be designed to accommodate
safely this worst-case afterheat situation.
33

ORNL-DWG 71-576A

Curve A: Peak steady-state temperature computed for type 1 afterheat rates at the indicated
times and with the emissivity.of all internal surfaces = 0.2 and the emissivity of
the outer surface of the outer shelf = 0.8.

Curve B: Temperature growth in the inner shell and the tube annuius computed as if; (1) the
annulus and shell are perfectly insulated, (2) have a total heat capacity of 129 Btu/°F
per foot of height,and (3) generate 77% of the total afterheat.

Curve C: Temperature growth in the intermediate shell computed as if; (1) the shell is per-

fectly insulated, (2) has a heat capacity of 287 Btu/°F per foot of height, and (3)
generates 23% of the total afterheat.

TN | R AU R

CURVE B

CURVE C

2500
\\\CURVE A

2000 : /E;:;;ATED

o
3 TRANSIENT Y
L
1
o
l—
L=
o
wl
a.
s
i
1500 ve

1000 _TEMPERATURE OF INFINITE "BLACK" SURROUNDINGS, 100C°F
2.78 hr 27.8 hr 11.6 days
S T T A Y O AR R A
10° 10° 10* 10° 108

ELAPSED TIME AFTER SHUTDOWN (sec)

Fig. 3.7. Estimated initial temperature transient caused by noble metal afterheat in an empty 563-MW MSBR heat exchanger.
TEMPERATURE (°F)

34

ORNL-DWG 71-577A

Curve A: Peak steady-state temperature computed for type 2 afterheat rates at the indicated
times and with the emissivity of all internal surfaces = 0.2 and the emissivity of
the outer surface of the outer shell = 0.8.

Curve B: Temperature growth in the inner shell and the tube annulus computed as if: (1) the
annulus and shell are perfectly insulated, (2) have a total heat capacity of 32 Btu/°F
per foot of height,and (3) generate 7 0% of the total afterheat. _

Curve C: Temperature growth in the intermediate shell computed as if: (1) the shell is per-
fectly insulated, (2) has a heat capacity of 72 Btu/°F per foot of height, and (3)
generates 23% of the total afterheat.

3000 '

/ CLTTHHE T DTy TP rry 1T T TIT
CURVE B
CURVE C
2500
2000 \
ESTIMATED
TRANSIENT
1500 7 .
—
1000 TEMPERATURE OF INFINITE "BLACK" SURROUNDINGS, 1000°F
2.78 hr 27.8 hr 14.6 days
SO I 1 A A A 1 O R W B A 11
102 10° 10* 10° 10°

ELAPSED TIME AFTER SHUTDOWN (sec)

Fig. 3.8. Estimated initial temperature transient caused by noble metal afterheat in an empty 141-MW MSBR heat exchanger.

3.4 THE CONSEQUENCES OF TUBING FAILURE

"

An analysis has been started to determine the
consequences of tubing failure in one of the main
MSBR heat exchangers. The following four cases -are

being considered:

R. P. Wichner

=+ IN THE MSBR HEAT EXCHANGER

1. double-ended rupture near the fuel outlet,
2. double-ended rupture near the fuel inlet,
3. small leak into the primary system,

4. small leak into the secondary system.

For case 1, it is estimated that 3.5 lb/sec of NaBF,
will leak into the primary system until protective
action, such as pump cutoff, is taken. It is not known if
BF; gas will evolve in this case; though sufficient
solubility of BF; in fuel exists to maintain BF; in
solution, the kinetics of the mixing process are insuffi-
ciently known to preclude this possibility. Hence case 1
is subdivided into (@) assuming no BF;(g) evolution and
(b) assuming complete BF,(g) evolution. For case 15,
prompt pump cutoff is essential to prevent excessive
primary loop pressures and pump damage due to circu-
lating voids.

It is estimated that the poison effect of NaBF, is
—0.18% &k/k per pound of NaBF, smeared over zones I
and II of the MSBR core. This is overwhelmingly due to
boron, the effect due to sodium being approximately
0.06% as large. |

In addition, for case 1, an estimated 0.96 Ib,, [sec of
fuel will leak into the secondary system until protective
action is taken. For case 2, an estimated 7.4 Ib,,,/sec of
fuel will initially leak into the secondary system, and
there will be no coolant flow into the primary system.

3.5 TRITIUM DISTRIBUTION AND CONTROL
IN THE MSBR

R. B. Korsmeyer

The transport behavior of tritium in a 1000-MW(e)
MSBR was examined previously, and various hypothet-
ical methods for its control and containment were
considered.® In that study the tritium concentrations in
the fuel and coolant salts were assumed constant
around each circulation loop — a condition that is
closely approximated only when the fraction of tritium
that is removed per pass is small.

Since then this restriction was removed, and the
efficacy of preventing tritium from appearing in the
steam system by purging it from side streams of fuel
and coolant salts was examined, along with the effect of
a circulating bubble fraction in both fuel and coolant
loops.® It was found that a bubble fraction in the range
of 0.01 to 0.1 circulating in the salt streams had

35

negligible effect on the tritium transport. Increasing the -

purge gas flow rates increased the tritium removal from
the side streams and markedly reduced that reaching
the steam system for all side stream fractions. A purge
gas flow sufficient to remove essentially all the tritium
from a side stream of 10% of the salt flow through the
reactor reduced the amount of tritium reaching the
steam to 0.2% of the production.

5. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, p. 53.

6. R. B. Korsmeyer, The Effect of Purge Rate on Tritium
Distribution in the MSBR, ORNL-CF-70-11-5 (Nov. 16, 1970)
(for internal use only).

In all cases where the tritium appearing in the steam
was negligible, almost all of it was removed from the
fuel as TF, which is assumed not to diffuse through the
metal walls. Since these results depend strongly on a
high UF,/UF, ratio to convert T, to TF according to
the mass action relationship used and on the rates of
reaction, of TF with the walls being low, it will be
necessary to reexamine the supporting chemistry before
reliance is placed on these results.

In another study’ the effect of using wet helium
under pressure as the secondary coolant was examined
with reference to the distribution of the tritium flow in
the purge gas systems, the cell enclosures, and the steam
system by suitable modification of the transport equa-
tions used in the TRIPOR program.’> The results of
some 18 typical runs showed that:

1. With helium at high pressure and containing water
vapor ‘at a partial pressure of 1 atm, the tritium reaching
the steam is negligible provided the water vapor is
changed at a rate equal to or greater than 0.2 Ib/hr. If
radiolysis were to increase the hydrogen concentration
in the steam by a factor as great as 100, the water purge
rate would have to be increased to 20 Ib/hr to maintain
the low tritium loss. It would be desirable to keep the
purge rate below about 1 Ib/hr, but rates as high as 75
lb/hr might be acceptable. :

2. If the water vapor partial pressure is as low as 6 or
7 torrs, water purge rates as low as 0.02 lb/hr are
effective in preventing significant quantities of tritium
from reaching the steam. If radiolysis were to multiply
the hydrogen concentration by a factor of 100, the
purge rate would have to be increased to about 2 1b/hr.
A radiolysis factor as high as 100 is considered to be
very unlikely.

In a third study® the effect of adding HF to the
sodium fluoroborate coolant salt was examined with
reference to the distribution of tritium flow in the
coolant purge gas system and the steam system,
utilizing a simplified model of the secondary coolant
and steam systems instead of the more complete model
and transport equations of the TRIPOR program. Use
of the simplified model appears justified in that the
calculated fraction of tritium transported to the steam
in the absence of HF is comparable with that calculated
for the same conditions in the TRIPOR program. The
results of nine runs showed that:

7. R. B. Korsmeyer, Capture of Tritium in Helium—H,0
Coolant for MSBR's, ORNL-CF-70-12-9 (Dec. 11, 1970) (for
internal use only).

8. R. B. Korsmeyer, Suppression of Tritium Transport to
Steam in an MSBR by!HF Addition to the Coolant Salt, ORNL-
CF-71-1-19 (Jan. 15, 1971) (for internal use only).
1. A ratio of HF addition to HT diffusion into the
coolant salt of at least 50 is required to hold the

/ tritium reacliing the steam to 1% or less of that
entering the coolant, the balance being carried off
by the purge gas.

2. For large ratios of HF/HT addition the tritium
fraction reaching the steam is inversely proportional
to the ratio. :

3. Increasing the purge gas flow from 20 to 100 cfm, or
by a factor of 5, has a negligible effect on the
tritium transport, at least for large HF/HT addition
ratios.

The mean residence time for fluoride gases (HF +
HT) in the circulating coolant salt is an inverse
function of the purge gas flow and is independent of
the HF feed rate inasmuch as an increased feed rate
at constant purge rate merely increases the concen-
tration proportionally. For all cases studied, a purge
rate of 20 cfm established a residence time of 62.5
hr, and a rate of 100 cfm reduced the residence time
to 12.5 hr. For HF/HT = 50 the corresponding HF
concentrations in the salt were 0.60 and 0.12 ppm.

5. For an assumed corrosion rate equal to half of the
HF feed rate to the system and HF/HT = 5000, the
calculated uniform rate of metal removal is about
0.001 in./year.

3.6 INDUSTRIAL STUDY OF 1000-MW(e)
MOLTEN-SALT BREEDER REACTOR

M. 1. Lundin J. R. McWherter

Proposals were solicited from a number of industrial
firms to perform design studies of a 1000-MW(e)
molten-salt breeder reactor. An evaluation team visited
each group that submitted a proposal. The Ebasco
Services group, consisting of Ebasco, Conoco, Babcock
and Wilcox, Cabot, Union Carbide, and Byron-Jackson
companies, was selected as the one that could most
nearly meet our objectives. A subcontract is being
negotiated with them.

They will initially develop their concept of a
molten-salt breeder reactor plant. Using this concept as
a base, trade-off and parametric studies of the nuclear
steam supply system, the energy conversion system, and
the fuel processing system will be made. After incorpo-
ration of the results of these studies in the reference
concept, they will estimate the plant capital and
fuel-cycle costs. A review of the research and develop-
ment program will be made. An independent assessment
of chemical processing and a safety review of the
proposed plant will be conducted. Technical liaison is

36

being furnished by ORNL. All the work will be
reported.

3.7 MSBE DESIGN
M. 1. Lundin J. R. McWherter

3.7.1 General
J. R.McWherter W, Terry.

The design studies reported previously’ were con-
tinued. A report'® was issued outlining the objectives
of the MSBE and establishing the design bases of the
plant. The reference core configuration consists of a
cylindrical array of graphite bars. The core is 45 in. in
diameter and 57 in. high, with 15% of the horizontal
cross section open for salt flow. The core is in a 7.5-ft-
ID spherical reactor vessel. That part of the vessel
between the core and vessel wall, with the exception of
the graphite alignment structure, is filled with salt. This
configuration achieves a start of life breeding ratio near
1 (0.96) and the desired fast (£ > 50 keV) neutron flux
of 5 X 10'* neutrons cm > sec ™' at a thermal power
of 150 MW. Design data of the MSRE and MSBR and

‘that proposed for the MSBE are compared in Table 3.5.

We are conducting studies to determine the problem
areas .and to evaluate possible solutions. Layouts of the
reactor cell and the primary system are being made to
indicate how the operational and maintenance problems
might be handled in different configurations. The
current reactor cell concept is shown in Figs. 3.9 and
3.10.

Removal of the graphite array as a unit and removal
of individual graphite bars are being considered. The
reactor vessel closure problems associated with the
graphite maintenance are being evaluated. It is desired
that the graphite maintenance approach be similar to
that proposed for the MSBR. The thermal and hy-
draulic problems of the core are being examined for use
of cylindrical and prismatic graphite bars. The support
and alignment structure of the graphite bars is being
studied.

3.7.2 MSBE Core Design
W. K. Furlong  W. Terry

A brief investigation was made of alternatives to four
previously published”+'® MSBE design concepts, speci-

9. MSR Program Semiannu. Progr. Rep. Feb. 28, 1969,
ORNL-4396, pp. 71-75.

10. J. R. McWherter, Molten Salt Breeder ExperzmentDeszgn
Bases, ORNL-TM-3177 (November 1970).
fically: (1) the prismatic moderator elen..nts, (2) the
removal and replacement of all the elements as a unit,
(3) the unreflected core, and (4) the three-pass flow
pattern. The results of these investigations are sum-
marized in the following paragraphs.

A cyhndrlcal graphite moderator element design was
evolved. A 4-in.-diam cylinder with a 1-in. central hole
appears acceptable, based mainly on temperature con-

37

siderations. The cylinders are separated by ", ¢-in.-thick
collars at three axial locations. This eliminates the
low-flow cusp region at points of tangency (except at
the narrow collars) and permits interchannel mixing of
the salt. The center hole is required to have acceptable
temperatures with a 4-in. element. The salt fraction is
20.5% in an array of such elements with 4 in. triangular

pitch.

Table 3.5. Comparison of design data for the MSRE, MSBE, and MSBR -

MSRE MSBE MSBR
Reactor power, MW(t) 7.3 150 2250
Breeding ratio 0.96 1.06
Peak graphite damage flux (£, > 50 keV), neutrons cm ™2 sec ™! 3x 1013 5 x 1014 3x 1014
Peak power density, w/cc, core including graphite 6.6 114 70
Volume fraction of salt in core 0.225 0.15 0.13
Primary salt
Composition, mole %
LiF 65 71.5 71.7
BeF, 29.1 16 16
ThF,4 None 12 12
UF, 0.9 0.5 0.3
Z1F,y 5 None None
Liquidus, ° 813 932 932
Density, 1b /ft at 1100°F 141 211 210
Viscosity, lb ft™ hr! at 1100°F 19 29 29
Heat capacity, Btulb 1ept 047 0.32 0.32
Thermal conductivity, Btu hr ™! ft™1 ¢F)™! 0.83 0.71 0.71
Temperature, °F )
Inlet, reactor vessel 1170 1050 1050
Outlet, reactor vessel 1210 1300 1300
Circulating primary salt volume, ft3 70 266 1720
Inventory, fissile, kg 322 3962 1470
Power density, primary salt, circulating, average, W/cc 4 20 46
Number of primary loops 1 1 4
Primary pump capacity, gpm 1200 5400° 16,000
Secondary system salt LiF-BeF, " NaBF4-NaF NaBF;-NaF
Composition, mole % 66-34 92-8 92-8
Liquidus temperature, °F 850 725 725
Density, Ib,, /ft® at 1000°F 124:1 117 117
Viscosity, i, /ft ™1 hr ™! at 1000°F 28.7 3.4 3.4
Heat capacnty, Btulb,, ™' CF)™! 0.57 0.36 0.36
Thermal conductmty, Btuhr™! ft™! CF)! 0.58 0.23 0.23
Temperature, °F
Heat exchanger inlet 1015 850 850
Heat exchanger outlet 1075 1150 1150
Number of secondary pumps 1 1 -4
Secondary pump capacity, gpm 850 5300 - ‘ 20,000
Tertiary system Air _ Steam Steam
Inlet temperature, °F ~70 700 700
Outlet temperature, °F ~180 1000 1000
Outlet pressure, psia 14.7 3600 3600

9MSR Program Semiannu. Progr. Rep. Feb 28, 1 970 ORNL-4548 pp. 42-45.

2334 initial.
“For 200°F AT; 4300 gpm required at 250°F AT, -
38

ORNL~DWG 71-6T710

90-1t DIAM -
wa
v i L
AIR LOCK PRIMARY PUMP — SECONDARY PUMP } i
i A
HOT CELL STEAM
o BLOCK VALVES
SUSZS R .
[Ei I-lllr- 2
STEAM
GENERATOR
60 ft Oin

NN\

————

N ——

PRIMARY SALT DRAIN TANK

. k‘.".'.v'.‘v.'[ viwieratalsl
|

REACTOR VESSEL
HEAT EXCHANGER

SECONDARY SALT DRAIN TANK

Fig. 3.9. Molten-salt breeder experiment, reactor building section B-B, 150 MW(t).

A major feature of cylindrical elements is that they
do not interlock (as the prismatic elements do) and
hence can be moved sideways relative to one another by
a handling machine stationed in the center of the core.
This idea was pursued through the conceptual design of
a moderator element handling machine. The machine
has two telescoping arms operating on the jackscrew
principle. It can rotate azimuthally, move up and down,
and can operate such that elements are removed and
replaced either along a radius or offset 4 in. from a
radius. The latter position is required for a few elements
that will not move along a radius without forcing or
jamming them. Exact indexing in all degrees of freedom
is obtained by use of worm and sector gears. They are
driven either by cables or shafts from electric motors
located in a relatively cool, shielded area. The machine
is inserted into the core after first rcmoving the center
16 elements plus three control rod sleeves. These are
removed through a 20-in. flanged access port by a
straight lift using an expandable tool. The flange serves
to support and align the machine, which then removes
each element, including the noncylindrical “‘fence”

pieces, to a point where it can be grasped and lifted
clear of the vessel by an expandable tool. Further
refinements of the machine design may incorporate an
elevator mechanism in place of the “manual” expand-
able tool.

Several reactor and vessel layouts were studied using
the cylindrical elements both with and without the
need for the handling machine. In the latter case a large
(~48 in.) removable head is required. One such layout -
used a cylindrical vessel with a 1-ft graphite reflector. It
appeared feasible to build such a reactor, although,
once installed, the reflector would not be removable.
Because of this and several other considerations, the
decision was made to use an unreflected reactor in a
spherical vessel.! °

One-pass flow through the reactor vessel was used in
this most recent study. This has advantages over the
previously proposed three-pass (up through blanket,
down through outer part of core, up through center of
core) arrangement because (1) downflow in the three-
pass concept requires a velocity of 2.4 fps to overcome
the buoyancy effect for a salt AT of 125°F, (2) the
PRIMARY SALT
DRAIN TANK

90 ft Qin.

REACTOR VESSEL

39

——

ORNL-DWG 71-6711

GAS PROCESSING CELL

STEAM CELL

STEAM
BLOCK VALVES

|

- HEAT EXCHANGER

SECONDARY SALT
DRAIN TANK

Fig. 3.10. ' Molten-salt breeder éxperiment, reactor building plan A-A, 150 MW(t).

resulting mechanical layout is simpler, and (3) the
reactor pressure drop is lower. A drawback is the lower
velocity and correspondingly lower heat transfer coef-
ficients for cooling the vessel wall and the moderator
elements. '

3.7.3 MSBE Primary Heat Exchanger Design

H. A.McLain C.E.Bettis W. Terry

The heat transfer area required in the primary heat
exchanger was determined for core inlet and outlet
temperatures of 1050 and 1300°F respectively. It was
assumed that the primary salt was in the tubes and that
a secondary salt was on the shell side. A number of tube
and shell configurations are being considered in the
studies of the primary cell arrangement. In the heat
exchanger studies, the primary sait inventory is being
kept as small as practicable.

" It was reported previously that spirallyindented heat
exchanger tubes would reduce the inventory of primary
salt.' ' However, they are not being used in the MSBE
design for’ the following reasons: (1) there are no
experimental data for heat transfer to molten salts
flowing in spirally indented tubes at Reynolds numbers
of interest'and (2) initial fabrication work on enhanced
tubes of “standard” Hastelloy N was troubled with
cracking resulting from the carbide stringers present in
the metal.!? - B

The principal data for a 150-MW MSBE shell-and-tube
primary heat exchanger are given in Table 3.6.

11. MSR Program Semiannu. Progr. Rept. Feb. 28, 1969,
ORNL4396, p. 57.

12. MSR Program Semiannu. Progr. Rep. Aug. 31, 1968,
ORNL-4344, p. 289.
Table 3.6. Principal design data for 150-MW
MSBE primary heat exchanger

Thermal rating, MW

Tube-side conditions

Fluid
Tube OD, in.
Tube wall thickness, in.
Tube length, ft
Number of tubes
Temperatures, °F

Inlet

Qutlet
Mass flow rate, 1b/hr
Pressure drop due to flow, psi
Volume of fluid in tubes, ft3

Shell-side conditions
Fluid
Shell ID, in.
Baffle type
Baffle spacing, in.
Baffle cut, %
Tube pitch, in.
Temperature, °F
Inlet
Outlet
Mass flow rate, Ib/hr
Pressure drop due to flow, psi
Tube natural frequency, cps

Vortex shedding frequency, cps

Frequency ratio

Approximate overall heat transfer
coefficient, Btu hr™! ft 2 (°F) !

150

Primary salt
¥
8
0.035
. 28.0
1340

1300
1050

6.32 X 10°
111

19.0

Secondary salt
31.5

Disk and doughnut
7.3

40

1316 triangular

Considerable attention was given to reducing the
possibility of having a tube vibration problem. There-
fore, one design criterion was that the ratio of the
natural frequency of the tube to the shell-side vortex
shedding frequency must be 2 or greater. Because of
this, the tube pitch is greater and the baffle spacing is
smaller than are normally used in commercial heat
exchangers. The 28-ft tube length is about the same as
is proposed for the MSBR primary heat exchangers. Use
of the same tube length is desirable to permit investiga-
tion of such factors as corrosion, fission product
deposition on the tubes, bubble behavior, and transport
of tritium through the tube walls.

Layout studies of this unit are in progress, with some
consideration being given to the fabrication, reliability,
and maintenance. The maintenance approach is to be
able to locate a leaking tube and plug the tube with the
heat exchanger in place in the reactor cell. In addition,
it-is to be possible to remove the entire heat exchanger
for repair or replacement. These studies so far indicate
that the best arrangement is to have a U-shaped
horizontal primary heat exchanger located near the top
of the reactor cell.
4. Reactor Physics

A. M. Perry

4,1 PHYSICS ANALYSIS OF MSBR
4.1.1 Single-Fluid MSBR Reference Design

H.F.Bauman '

The nuclear performance of the reference-design
single-fluid MSBR has been calculated again to take into
account the effects of the recently adopted metal
transfer process on the concentrations of various
nuclides in the fuel salt. The last set of calculations for
the reference design (case CC-120), reported a year
ago,! was based on a reductive-extraction process which
removed protactinium on a 3-day cycle, europium on a
500-day cycle, and other rare earths on a 50-day cycle;
plutonium was assumed not to be removed, though in
fact it probably would be removed (in a hydrofluori-
nation step in the process). Thus the assumption is
equivalent to assumning recycle of plutonium as well as
uranium. The new set of calculations assumed that the
fuel salt circulates through the processing plant on a
10-day cycle, with removal efficiencies and effective
removal times that depend on the element. In addition,
a few minor data corrections and cross-section changes
were included in the later calculations.

In calculating the concentrations of noble gases and
daughters, we assumed, as in the earlier calculation, that
xenon and krypton are removed by gas stripping on a
50-sec cycle. Also, as before, a fixed poison fraction of
0.005 was assumed to allow for absorption of ! 3° Xe by
the core graphite. Except for the elements listed in
Table 4.1, there is complete removal of fission products
from the salt as it passes through the processing plant
on a 10-day cycle. Also removed completely are
plutonium and protactinium. Two important neutron
poisons, 237"Np and 23°U, were assumed not removed
in processing but were assigned 16-year removal times
to simulate their average concentrations over the
lifetime of the reactor. As in previous calculations the

1. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL4548, p. 58.

41

Table 4.1. Removal efficiencies and effective removal
times for rare earth and active metal elements
in the metal transfer process

Effective
Element Removal 16
efficiency removal time

(days)?
Rb 1.0 10.0
Sr 0.595 16.8
Y 0.333 30.0
Cs 1.0 10.0
Ba 0.595 16.8
La 0.452 22.1
Ce 0.602 16.6
Pr 0.333 30.0
Nd 0.333 30.0
Pm 0.340 294
Sm 0.372 26.9
Eu 0.195 51.2
Gd 0.333 30.0

9For a ten-day processing cycle.

effects of deposition of noble metals in the core and the
burnout of boron in the graphite were neglected, since
the latter slightly more than compensates for the
former. Carrier salt was discarded on a 4200-day cycle.
The description of the reactor given in Table 4.2 is
the same as in the last set of calculations (CC-120)
except for the absence of plutonium in the fuel.
There were several differences from the earlier cal-
culation besides the change in processing. A key-punch
error was corrected in the 233Pa decay constant (from
2.6279 X 1077 sec™! to 2.9279); this correction makes
a negligible change in the neutron balance. The same
spectrum-weighted cross-section set was used for the
annulus as for core zone 2; this set matches the neutron
spectrum in this narrow region better than the set
weighted for an infinite 100% salt region used pre-
viously. The same cross-section set was used for the
reflector as for core zone 1. This set matches the
spectrum in the reflector reasonably well and avoids an
error in the thorium resonance group cross section
which was uncovered in the set previously used for the
Table 4.2. Characteristics of the single-fluid MSBR reference
design with processing by the metal transfer process

on a ten-day cycle

A. Description

Identification
Power
MW(e)
MW(t)
Plant factor
Dimensions, ft
Core zone 1
Height
Diameter
Region thicknesses
Axial
Core zone 2
Plenum
Reflector
Radial
Core zone 2
Annulus
Reflector
Salt fraction
Core zone 1
Core zone 2
Plenums
Annulus
Reflector
Salt composition, mole %
UF,4 '
PUF3
ThF,
BCF2
LiF

B. Performance

Conservation coefficient, [MW(t)/kg] ?
Breeding ratio
Yield, % per annum (at 0.8 plant factor)
Fissile inventory, kg
Specific power, MW(t)/kg
System doubling time, years
Peak damage flux, £ > 50 keV, neutrons
cm™? sec”!
Core zone 1
Reflector
Vessel
Power density, W/cm?3
Average
Peak
Ratio

16.2
1.071
3.63
1487
1.51
19

3.5 x 1014
2.9 x 1013
3.3x 10!

22.2
70.5
3.18 °

Table 4.2 (continued)

Fission power fractions by zone

Core zone 1
Core zone 2
Annulus and plenums
Reflector
~ C. Neutron balance
Absorptions
232Th 0.9968
233p, 0.0045
233y 0.9242
234y 0.0819
235y 0.0758
236y 0.0074
237Np 0.0064
238py 0.0001
239py 0.0
240py 0.0
241py : 0.0
242py 0.0
Li : 0.0032
TLi 0.0161
Be 0.0070
19g 0.0203
Graphite 0.0510
Fission products 0.0150
Leakage® 0.0221
ne 2.2318
D. Fuel-cycle costs® (mills/kWhr)

Inventory
Fissile
Salt

Replacement salt

Processing

Fissile production credit

Total

0.791
0.150
0.050
0.009

Fissions
0.0030

0.8245
0.0004
0.0618

0.00444

0.382
0.060
0.034
0.293
-0.104

0.665

%(n,2n) reaction.
bInciuding delayed neutron losses.

€At 13.2%/year inventory charge on materials, 13.7%/year
fixed charge rate on processing plant, $13.8/g 233U, $11.9/g
235y, $12/kg ThO,, $120/kg Li, $26/kg carrier salt (including -

7Li). Excluding graphite replacement cost.
reflector. Both these cross-section changes would be
expected to enhance the absorptions in thorium relative
to uranium and hence to give a higher breeding ratio.
However, these changes were in minor regions of the
reactor, and the effect on the calculated reactor
performance is estimated to be small relative to the
effect of the change in processing.

The results indicate a higher breeding ratio, a lower
fissile inventory, and a higher conservation coefficient
for the reference MSBR with the metal transfer process,
compared to the same reactor with the reductive-
extraction process (1.071 vs 1.063, 1487 kg vs 1504 kg,
and 16.2 [MW(t)/kg]? vs 14.1 [MW(t)/kg]? respec-
tively). The most important differences in the neutron
balance are the absence of absorptions in plutonium,
reduced absorptions in fission products, and a higher
value of 7ne, because a higher proportion of the fissile
material is 233U,

The fuel-cycle cost in Table 4.2 is on a different basis
from the earlier calculation. The fissile material charge
is based on 235U valued at $11.9 per gram. The
inventory charge is taken as 13.2% rather than 10%, and
no allowance is made for graphite replacement [this
allowance has previously been estimated to be about
0.1 mill/kWhr(e)]. On the same basis, the fuel-cycle
cost for case CC-120 is 0.70 mill/kWhr.

4.1.2 Fixed-Moderator Molten-Salt Reactor
H. F. Bauman

The use of the ROD and the HISTRY codes for the
calculation of molten-salt reactors with batch proc-
essing was described in the last semiannual report.? An
iterative process is used, in which reaction rates per
atom for each type of nuclide are obtained from ROD
and supplied to HISTRY, where nuclide concentrations
as a function of time are calculated over one or several
batch processing cycles; averages of these concen-
trations over the reactor life are then computed and
returned to ROD for another calculation of reaction-
rate coefficients. The process is continued until ROD
and HISTRY converge on a common set of average
concentrations. This process has now been automated
by including the HISTRY code as a subprogram in
ROD. The average performance of a reactor with batch
processing may now be calculated in a single case,
rather than in four to six separate cases required by the
old method. The results obtained are the same, but the
chance of error in transcribing data in the iterative

2. MSR Program Semignnu. Progr. Rep. Aug. _31, 1970,
ORNL4622, p. 28. :

’

43

Table 4.3. Fixed-moderator 1000-MW(e) molten-salt
breeder reactor

Core graphite life, 30 years
Effective processing cycle time: protactinjum removal,
10 days; rare earth removal, 25 days

Plant size, MW(e) 1000
Thorium concentration in fuel salt, mole % 14
Case identification (SCC series) 201
Volume fraction salt in core
Zone 1 0.137
Zone 2 0.111
Zone 3 0.127
Thickness of core zones ? ft
Zone 1 8.46
Zone 2 3.64
Zone 3° 1.97
Annulus 0.14%
Core diameter, overall, ft 28.5
Fissile inventory, kg 2220
Breeding ratio 1.069
Conservation coefficient, [MW(t)/kg] 2 7.0
Fuel yield, %/year (at 0.8 plant factor) 2.34
Fuel-cycle cost,© mills/kWhr 0.85

9Three-zoned core with zone thicknesses and volume fractions
optimized for maximum conservation coefficient by ROD calcu-
lations in spherical geometry.

bNot optimized.

©Based on 235U at $11.9 per gram and an inventory charge
of 13.2%.

process has been eliminated. The cases in the following
section were calculated with the new ROD-HISTRY
program. |

A study of the fixed-moderator MSBR with con-
tinuous processing by the metal transfer process was
reported in the last semiannual report. The best
1000-MW(e) case from this study was SCC-201, a
three-zone-core reactor with 14 mole % thorium fuel
salt. The characteristics of this reactor are given in
Table 4.3. We have made a brief study of the
performance of this same reactor operated as a con-
verter with batch processing. We assumed plutonium
feed of a composition typical of first-cycle discharge
from a light-water reactor. We examined thorium
concentrations lower than used for the continuous
processing case, anticipating that this would give lower
fissile inventory and fuel-cycle cost. Since there were no
processing restrictions on the salt composition, we
elected to use the minimum-liquidus-temperature com-
positions given in Table 4.4. Batch processing times of
six and eight years were used; we know from earlier
cases that the optimum range of batch cycle time is
about four to eight years. The results of the four cases
in this study are shown in Table 4.5. The fuel-cycle
costs are based on a fuel and salt inventory charge of
13.2%, rather than the 10.0% used in previous cal-
culations.

Cases A-13 and -14 show the effect of the batch-
processing cycle time, with the salt composition held
constant. The reactor lifetime is assumed to be 30 years
at 0.8 plant factor, or 24 fuil-power years. It is
convenient to assume a refueling program of either
three 8-year cycles or four 6-year cycles. The 6-year
cycle appears at first glance better than the 8-year; it
has a higher conversion ratio, a lower fissile inventory,
and essentially the same fuel-cycle cost. However, it has

Table 4.4. Fuel carrier salt compositions for minimum
liquidus temperature

.Salt number 14/17 12/20 10/23 8/27
Composition, mole %
ThF4 14 12 10 8
BeF, 17 20 23 27
LiF ‘ 69 68 67 65
Liquidus temperature,? °C 495 485 475 465

9R. E. Thoma (ed.), Phase Diagrams of Nuclear Reactor
Materials, ORNL-2548, p. 80 (November 1959).

44

a higher peak damage flux and therefore a lower core
life (0.73 relative to case SCC-201), although this
apparent disadvantage could probably be overcome by a
slight change in salt volume fractions in the three core
zones.

Cases A-14 through A-17 have the same cycle time
but different thorium concentrations. The conversion
ratio, fissile inventory, fuel-cycle cost, and relative core
life all increase with increase in thorium concentration.
The lifetime fuel composition for the 12% thorium case
is given in Table 4.6.

The base reactor (SCC-201) was optimized with a salt
containing 14 mole % thorium so as to have a flat flux
distribution in core zone 1. When other salt compo-
sitions are substituted in this design, the flat flux
distribution is no longer obtained, and for this reason
the cases with lower thorium concentrations tend to
have higher peak fluxes and therefore shorter-lived
cores. We believe that this penalty on core life can be
largely avoided by reoptimizing the core design for the
desired salt composition; our purpose in this study,
however, was to calculate the effect of changing the
fuel composition in a reactor of fixed core design.

We conclude from this study that the fixed-moderator
MSBR design is attractive when operated as a converter

Table 4.5. Lifetime averaged performance of a 1000-MW(e) fixed-moderator MSBR operated as a converter reactor
with batch processing and plutonium feed

Base reactor design: SCC-2019

Identification
Thorium concentration in fuel salt, mole %
Carrier salt?
‘Processing cycle time, batch, years
Average conversion ratio®
Specific fissile inventory, kg/MW(e)
Peak damage flux (£ > 50 keV), 10!3 neutrons cm 2 sec™!
Relative core lifed :
Fuel-cycle cost, mills/kWhr
Inventory, at 13.2%
Fissile
Salt
Replacement®
Processin,
Fissile feed®
Total fuel-cycle cost

A-13 A-14 A-16 A-15 A-17
10 10 8 12 14
10/23 10/23 8/27 12/20 14/17
8 6 6 6 6
0.88 0.90 0.85 0.92 0.93
1.64 1.60 1.32 1.89 2.09
5.74 6.15 6.27 5.29 4.73
0.78 0.73 0.72 0.85 0.95
0.317 0.322 0.249 0.394 0.443
0.048 0.045 0.044 0.048 0.049
0.060 0.080 0.075 0.083 0.087
0.033 0.036 0.036 0.036 0.036
0.015 0.093 0.147 0.059 0.043
0.573 0.576 0.551 0.620 0.658

2Refer to Table 4.3.
bRefer to Table 4.4

€Assuming that the plutonium discarded with the fuel salt is not recovered.

4Based on the 30-year core life of the base reactor = 1.0.
€Replacement includes thorium and carrier salt.

fNormalized to 0.1 mill/kWhr for a 300-MW(e) plant on a 4-year cycle; unit cost assumed to vary inversely as the 0.3 power of

the average processing rate.

1

8First-cycle plutonium; composition 60% 239Pu; 24% 24%Pu; 12% 241 Pu; 4% 242 Pu. Valued at $9.88 per gram fissile.
45

‘Table 4.6. Lifetime fuel composition and conversion ratio for a fixed-moderator 1000-MW(e)
MSR with plutonium feed for four 6-year cycles

Case A-15, 12 mole % thorium

Time (full-power years)

Inventory (kg) Conversion
Percycle Cumulative 233pa 233y 234y 235y 236y 237Np 239py  240py 241p,  242py, ratio?
0 0 0.0 0.0 0.0 0.0 0.0 0.0 6835 2745 1378 . 46.1 0.768
1 1 58.7 4419 118 0.3 0.0 0.0 4752 4877 336.6  146.5 0.861
2 2 63.0 804.0 375 2.0 0.1 0.0 250.7 4049 3715  246.1 0.919
4 4 733 1291.1 1243 124 0.8 0.0 107.5 1946  229.0 390.4 0.909
6 6 75.0 1526.7 230.7 325 3.6  0.1b 89.8% 131.7% 141.5% 434.5% 0.880
0 6 0.0 16019 230.7 325 3.6 0.0 29.0 11.6 5.8 1.9 1.063
1 7 © 89.5 16062 2928 47.0 6.6 0.2 53.5 64.0 49.4 25.1 0.980
2 8 88.5 1659.4 3504 619 106 0.5 33.1 51.2 50.7 45.4 0.970
4 10 85.2 1699.2 4503 912  21.3 1.4 345 446 43.4 74.2 0.938
6 12 81.5 1698.1 527.6 1169 349  2.6b 41.7% 5220 4872 97.0% 0.913
0 12 0.0 17799 5276 1169 349 0.0 - 00 0.0 0.0 0.0 1.020
1 13 91.9 17135 563.5 1293 433 1.6 19.2 18.7 13.1 5.9 0.993
2 14 89.1 1726.7 5954 1403 521 3.0 18.8 22.8 19.6 ' 14.0 0.974
4 16 84.7 1725.1 6473 1588 703 5.9 27.2 32.3 28.9 33.4 0.941
6 18 80.8 1708.0 685.1 173.3 888  8.6° 37.4b  45.1b 407 5745 0.915
0 18 0.0 1789.1 6851 1733 888 0.0 0.0 0.0 0.0 0.0 1.009
1 19 90.7 - 1719.4 7023 1797 990 3.7 16.0 14.5 9.8 4.2 0.992
2 20 87.9 1730.3 718.3 1854 109.0 6.8 17.7 20.2 16.8 11.1 0.972
4 22 83.6 1726.7 7439 1948 . 1282 11.8 27.1 31.7 27.9 29.7 0.939
6 24 79.9 1709.0 761.6 202.0 146.6 1580  379Y% 453>  407P  s54.0% 0913

4Not adjusted for discard of plutonium,
bQuantity discarded at end of cycle.

with batch processing. The results suggest that if the
reactor is to be designed for dual-mode operation (e.g.,
initially as a converter with batch processing and
subsequently as a breeder with continuous processing),
the salt composition should be essentially the same for
either mode of operation. In particular, the thorium
concentration cannot be changed significantly without
altering the peak flux and the core graphite life.

4.2 MSR EXPERIMENTAL PHYSICS

4.2.1 HTLTR Lattice Experiments
G.L.Ragan- O.L. Smith

The High-Temperature Lattice Test Reactor (HTLTR)
is being used to study the reactor physics characteristics
of a lattice that simulates a typical molten-salt reactor.
Reactor measurements began in February and are to be
completed in May. The experiments are being spon-
sored by the Physics Branch of DRDT and performed
by Pacific Northwest Laboratory (PNL). They were
planned by ORNL, in consultation with PNL; data
reduction and analysis will be similarly shared.  Sub-
sequent use of the experimental results to test MSR
computational methods is the responsibility of ORNL.

Gulf General Atomic has cooperated with PNL on
similar lattice experiments related to high-temperature
gas-cooled reactor (HTGR) work. Although the HTLTR
lattices related to HTGR differ from those for the MSR
in several important details, they are closely related,
and we plan to study them carefully. To this end, there
are continuing consultations between Gulf General
Atomic and ORNL personnel.

The HTLTR consists of a 10-ft cube of graphite
surrounded by a gas-tight insulating oven. The central
5-ft-square by 10-ft-long section can be removed and
loaded with as much as necessary of the test lattice of
interest. The driver fuel and the control system are
loaded outside the test lattice. One measures the
reactivity changes aécom‘panying a ‘series of pertur-
bations made in a 7.5 X 7.5 X 24 in. test cell volume at
the center of the lattice. These and other data are used
to determine certain lattice characteristics, especially a
suitably defined lattice multiplication factor k. Meas-
urements can be made from room temperature to
1000°C. '

An extensive set of measurements is first made at
room temperature. The cadmium ratio of gold is
measured at several positions in and near the test cell
volume, as an indicator of spectrum matching. Loading
adjustments are made until the spectrum matching is
satisfactory. Foils of other materials are then activated
in various positions to obtain data needed in inter-
preting later measurements. One result of these acti-
vation measurements, coupled with a measurement of
k. for the main lattice, is an experimental determi-
nation of the effective value of 7(??3U). Because of its
importance in evaluating the breeding ratio of MSBR
design proposals, 7(33U) will be measured as precisely
as possible. However, the accuracy with which n can be
determined from these measurements is intrinsically not
very high, and it is not clear, at present, whether the
resulting value will be any more reliable than one
obtained from basic cross-section data and calculated
neutron spectra.

No attempt was made in designing the HTLTR
experiments to achieve an exact mockup of details of
current MSBR designs. Instead we use a simple lattice
having neutronic characteristics comparable with those
of current designs and make certain parametric vari-
ations on this basic lattice. Thus we obtain information
applicable to possible future changed designs and on the
sensitivity of the reactor- physics characteristics to such
parametric changes. For each configuration, we will test
our calculational techniques and cross-section data by
making critical comparisons between calculated and
measured quantities. These tests of calculational ability
constitute the most important objective of the ex-
periment.

Although the lattice finally adopted is neutronically
similar to that of a typical MSR, it is quite different
physically. Fuel channels are simple 0.786-in.-diam rods
spaced 1.875 in. apart in a square lattice array. Instead
of using molten salts as fuel, we use an appropriate
mixture of solid materials: pyrocarbon-coated particles
(containing ?*3UQ, and some ThO, ), powdered ThO,,
and powdered graphite. These are the same materials
that were used for the HTGR experiments but with
proportions chosen to give the 222 U-to-moderator ratio
and 233U-to-Th ratio typical of MSR designs. Com-
pared with the use of actual molten salts in the test
lattice, the advantages in availability and ease of
handling are evident. Equally evident are shortcomings
in simulating MSR fuels: the much lower thermal
expansion coefficients of the solid fuel and the
omission of significant absorbers such as Li, Be, and F.
To compensate for the omission of absorbers, the
effective cross section for each omitted nuclide will be
determined from central reactivity worth measurements
made on a sample containing that nuclide. We can thus
evaluate each absorber separately, instead of lumping it
with others. Fuel expansion differences are evaluated as

46

follows. By comparing the worth of a special test block,
containing fuel of the same composition but more
densely packed, with the worth of a test block
containing standard fuel, we will be able to determine
the fuel density coefficient of reactivity at each
temperature. This is useful information in itself and also
permits evaluation of the temperature coefficient of
reactivity that would have been obtained using a molten
salt fuel, with its known higher expansion coefficient.

Test temperatures selected are room temperature,
300°C, 627°C, and 1000°C. At each temperature the
reactivity worths outlined in Table 4.7 are to be
measured. Series A includes insertion of a standard
block into the central void (to get k), replacement by
a copper foil (for calibration), and replacement by
blocks having various perturbations in fuel density, fuel
channel geometry, and fuel volume fraction. Other
blocks simulate control rods of the “displacement”
type (graphite displaces fuel) and of the poison-
shutdown type (Hastelloy N plates). In series B,
material worth measurements are to be made on
samples containing Cu (for calibration); Be (as solid
rod); F (in Teflon rod, empirically CF,, at T < 300°C);
"Li (in solid " LiF salt, at T < 627°C); Hastelloy N (as a
foil); and ThO,, 233U, ?3°U, and 23°Pu (as dilute
dispersions in graphite powder).

The HTLTR experimental design is based upon a
series of XSDRN®>* neutronic calculations in which the
central portion of the HTLTR was mocked up in
one-dimensional cylindrical geometry. The results of
these calculations are given in Table 4.7 in terms of the
“expected” reactivity of the corresponding experi-
mental measurement. It must be emphasized that these
calculated reactivities are approximate and were in-
tended only as a guide in choosing experimental design
parameters. Numerous known approximations were
made. For example, fuel and ThO, particle self-
shielding effects were neglected. Though these effects
are small, they are not negligible and must be included
in the final analysis of the experiments. Special ana-
lytical tools are now being developed to treat the
double heterogeneity of particles in a fuel lump..

There are two series of experiments listed in Table
4.7. In both series the worth of the 7.5 X 7.5 X 24 in.
test cell (see Fig. 4.1a) is to be determined for a variety

3. XSDRN is used both to obtain group cross sections for
each region and also to perform multigroup one-dimensional
neutronics calculations with the S, transport method.

4. N. M. Greene and C. W. Craven, Jr., XSDRN: A Discrete-

Ordinates Spectral Averaging Code, ORNL-TM-2500 (July
1969).
of test cell compositions and temperatures. (The test
cell layout for item B1 is shown in Fig. 4.1 and will be
discussed in more detail later.) In series A, various
neutronic properties specifically applicable to the
MSBR are to be determined. In item Al, the test cell
will contain the same compesition and arrangement as
the main lattice fuel, which is designed to approximate
the principal properties of an MSBR fuel cell, that is,
the same spectrum, k,, C/U ratio, resonance absorp-
tion, etc. The overall cell temperature coefficient will
be determined from this item. The remaining items of

47

series A will simulate other important features of the
MSBR. For example, item A4 will be used to measure
the salt density coefficient. Items A5 and A7 will help
determine the fuel lumping effect on resonance absorp-
tion. Items A8, A9, Al10, and All will be used to
determine control-rod worth. Series B consists of
material worth measurements for the principal MSBR
materials. From both series A and B, we will obtain
information to test the basic nuclear data and the
reactor-physics computational techniques used for MSR
analysis.

Table 4.7. Summary of HTLTR reactivity measurements required at each temperature

Series A: Worths of blocks as described (7.5 X 7.5 X 24 in. volume) to be measured relative to same volume voided

[tem Descriptive L Precalcul?)ted
No. block title Block description? worth
(¢ at 900°K)
1 Standard Standard fuel in 16 holes 0.786 in. diam, 13.64 vol %
2 Calibrating copper A standard copper foil, lining cavity
3 V replacing Th Vary 1: V,03; replaces ThO, in fuel
4 Dense fuel Vary 1: fuel more densely packed
5 Coarse holes Vary 1: 4 holes, 1.572 in. diam, 13.64 vol %
6 37.1vol % fuel Vary 1: 4 holes, 2.578 in. diam, 37.1 vol %
7 Cruciform Retain 13.64 vol % fuel, geometrically rearranged
8 Solid graphite Contains no fuel +15
9 4-in. control rod Vary 8: one 4-in.-diam fuel rod, 22 vol % +3
10 6-in..control rod Vary 8: one 6-in.-diam fuel rod, 50 vol % -13
11 Shutdown rod Vary 8: insert four Hastel-loy N blades (2.5 X 0.25 X 23 in.) —88
Vary n: means “like item n, except as follows.”
bworths based on 0.018¢/g Cu per measurements at room temperature.
Series B: Worths of material samples to be measured relative to graphite rod (at center of Fig. 4.1)
ltem o Precalculated
No Description of sample?® worth?
’ (¢ at 900°K)
1 " Graphite rod, normal density (1.62) 0
2 Graphite powder, density ~1.0 -0.3
3 Void -0.8
4 Copper foil (10-mil, on 1.96-in.-diam graphite rod) -3.6
5 Beryllium rod . +1.7
6 Teflon rod (empirically CF;) at T < 300°C -0.7
7 7LiF (1.375-in.-diam rod; Li is 78 ppm ®Li) at 7 < 627°C -2.0
8 ThO, (250 g) dispersed in item 2 type powder —-4.4
9 233y + Th coated particles (7.5 g 233U) in same +1.9
10 235y coated particles (10 g 235U) in same +1.4
11 Pu coated particles (4.6 g 23?Pu) in same +2.0
12 Hastelloy N foil (11-mil) on item 4 type rod —-1.8
13 Main-lattice fuel mixture -6.1

9All samples 23 in. long and (unless otherwise noted) 2 in. in diameter.
bworths based on 0.018¢/g Cu per measurements at room temperature.
ORNL-DWG 71-6712

MAIN
L+ ATTICE CELL

| _——GRAPHITE

4 FUEL

TEST
SAMPLE ~_E
L:’).TS in, ——a—t—— 3 75 in, —=
(o}
|
|
% >
/////‘, ///’ Z 2,
, ‘ 72}% ,
' Z ;flf?’; WHITE BOUNDARY
7 g CONDITION
‘.43/!“;..{{,
D
R % ”
TEST
SAMPLE B

GRAPHITE

(&)

Fig. 4.1. HTLTR central core geometry and composition. (a)
Midplane, (b) as simulated in XSDRN calculations.

As an example of the precalculational methods used
to generate the approximate values given in Table 4.7,
we will describe item B1 in some detail. In Fig. 4.1a is
shown one quadrant of the central portion of the
HTLTR, including the test cell and one square ring of
adjacent main lattice cells. Zone 5 contains the material
whose reactivity worth is to be determined. In item B1
the material is graphite. The geometry of Fig. 4.1a is

48

represented in XSDRN as shown in Fig. 4.1b, where
zones 1 through 4 are homogenized circular annuli
representing the actual geometry. For example, zone 1
represents the main test lattice.

The white boundary condition at the outer edge of
zone 1 simulates the rest of the HTLTR. This is a good
approximation since the fluxes in the central portion of
the HTLTR are nearly flat. ‘

The direct result of the XSDRN calculation is the
multiplication factor k& for each of the assemblies
described by Fig. 4.1b. Since the reactivity of the
HTLTR with the various test samples listed in Table 4.7
(series B) will be measured relative to its reactivity with
item Bl in the configuration of Fig. 4.1z, we have
expressed our calculated multiplication factors as re-
activity changes relative to item B1. In item B2 of Table
4.7 we introduced a 5-mil-thick copper foil at the
interface between zones 4 and 5 and calculated the
difference in multiplication factor between items 1 and
2, as given by XSDRN for the geometry of Fig. 4.1b.
Using the experimentally determined worth of copper
at the center of the HTLTR (0.018¢/g), we normalized
our calculated reactivity worths for the remaining
samples, as listed in Table 4.7.

Status of the experimental and calculational program
at the end of February was as follows. Precalculations,
of approximate nature, had been made as a basis for
design of the experiment. PNL had prepared all lattice
blocks and all materials test samples, except for a few
special ones that were supplied by ORNL: "LiF, Teflon
(CF,), Hastelloy N foils and plates,-and some special
carbon powder for the densely packed fuel. PNL had
completed the room-temperature measurements and
had begun on those at 300°C. These measurements
indicated that good spectral matching was achieved and

- that the specified design gave k. and measured re-

activity worths in satisfactory agreement with the
approximate precalculated values. ORNL was preparing
to make more precise calculations on the as-built
lattice. Codes were being developed to permit inclusion
of grain self-shielding effects, neglected in precalcula-

- tions. Cross sections were being processed for the other

temperatures needed, the precalculations having been
done at 900°K (627°C).
5. Systems and Components Development

Dunlap Scott

5.1 GASEOUS FISSION PRODUCT REMOVAL
C. H. Gabbard

5.1.1 Gas Separator and Bubble Generator

Construction of the water test loop' for testing
MSBE-scale gas separators and bubble generators was
completed, and the loop was put into operation. Figure
5.1 shows the test loop with the venturi-type bubble
generator and the gas separator installed. The loop has
been operated primarily on demineralized water at
liquid flow rates of 200 to 550 gpm and at gas flow
rates of O to 2.2 scfm. It has also been operated with
water containing small amounts of n-butyl}alcohol and
sodium oleate and with a 41.5% glycerin-water mixture
which is hydraulically similar to fuel salt.

The performance of the 4-in.-ID gas separator was
greatly improved by installation of the annular takeoff
port in the recovery vane hub.? Straightening vanes
were added within the recovery hub to reduce the
pressure drop associated with the flow of gas and liquid
through the hub. Figure 5.2 shows the final design
operating on demineralized water containing about
0.3% voids at the entrance to the separator. The vortex
is firmly attached to the central cone within the
Plexiglas recovery hub, and only a trace quantity of
very small bubbles escapes the separator at any normal
operating condition within the gas flow range of O to
2.2 scfm. The improvement in performance can be seen
by comparing Fig. 5.2 with Figs. 5.1 and 5.2 in the
previous semiannual report.>

The .performance of the venturi-type bubble generator
has been satisfactory when operated with the axis
vertical. The bubbles were adequately dispersed and
were 0.020 in. in diameter or less at flow rates of 400
gpm or greater. There were no strong pulsations in the
bubble output, as had been observed in some of the

1. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,
ORNL-4622, p. 39.
2. Ibid., p. 38.

3. Ibid., pp. 36 and 37.

49

previous tests of smaller-scale bubble generators, but
hydraulic turbulence in the diffuser cone produced
small fluctuations in the bubble concentration. A
straight cylindrical mixing section about 2.5 throat
diameters in length was incorporated in the design to
permit the bubbles to be formed and partially dispersed
in a high-velocity region prior to reaching the diffuser.
However, in practice, the gas traveled the length of the
mixing section in relatively straight plumes of large
elongated bubbles. The small bubbles were actually
formed and dispersed by the turbulence in the diffuser
cone. Maintaining a balanced gas feed around the
periphery of the bubble generator has not presented a
problem with the bubble generator operating in the
vertical position and has not required a significant
pressure drop across the gas feed holes.

The loop has been operated for short periods of time
on water containing, separately, n-butyl alcohol, so-
dium oleate, and 41.5% glycerin. The n-butyl alcohol
was used in concentrations up to 160 ppm to study the
effect of a surfactant that would not otherwise alter the
properties of the liquid. The sodium oleate was used in
concentrations up to 40 ppm with 4 ppm of an
antifoaming agent to determine the effect of reducing
surface tension (from 72.5 dynes/cm for demineralized
water to about 45 dynes/cm). The 41.5% glycerin-water
mixture is hydraulically similar to the fuel salt. The
addition of n-butyl alcohol and sodium oleate may have
caused a slight decrease in the size of bubbles produced
by the bubble generator, but the most obvious effect
was to stabilize small bubbles, inhibit their coalescence,
and promote clouding of the loop fluid by bubbles
0.001 to 0.002 in. in diameter which were formed by
the pump. These small bubbles were not effectively
removed by the separator, and several minutes of loop
operation without gas input were required to clear the
fluid. Operation with the 41.5% glycerin-water mixture
gave similar results except that the small bubbles were
removed much more slowly. A small amount of
foaming occurred at the free liquid surfaces with water
containing the n-butyl alcohol; the sodium oleate
produced extensive foaming that was greatly reduced
but was not eliminated by addition of the antifoaming
50

PHOTO 78448

Fig. 5.1. Gas separator and bubble generator water test loop.

PHOTO 78450
GAS TAKEOFF LINE

RECOVERY VANES

FLOW i

Fig. 5.2. Four-inch-ID pipeline gas separator operating on demineralized water.

ORNL-—-DWG Ti— 6713

4 \
-3 \
£ \ \ 1.5% GLYCERIN -WATER
I
: N\
4 » ./ MSBR FUEL SALT
g WA
-
= :
2 \ X

\ Ty
o
0 0.004 0.008 0.012 0.016 0.020

RUBBLE DIAMETER (in.}

Fig. 5.3, Calculated separation length vs bubble diameter for
4-in.-ID gas separator and liquid flow of 400 gpm.

agent. No foaming occurred with the glycerin-water
mixture. Loop and pump modifications are in progress
to determine whether fewer small bubbles are produced
by operating the pump near its best efficiency point
and at a lower speed.

We infer that the circulating bubbles in the MSRE and
in the MSRE prototype pump test loop were not so
small because they separated out relatively quickly
under low or zero flow conditions. Attempts to
produce foams in molten salt* by bubbling gas through
a coarse glass frit produced bubbles 2 to 3 mm or larger
in diameter under normal salt conditions, and a foam
was generated only when moisture or hydrates were
added to the salt. Therefore we do not expect to
encounter significant quantities of the very small
bubbles in a salt system.

A computer program was written to calculate the
required length for separating bubbles in a pipeline
separator. The mixture of bubbles and liquid enters the
separator with specified axial and rotational velocities,
and the required length is the axial distance that a
bubble of a specified diameter will travel as it is caused
to migrate from the wall of the pipe to the central
vortex by the centrifugal forces. Calculations were
completed for a 4-in.-ID separator operating on water,
-salt, and the 41.5% glycerin-water mixture. The results
of these calculations are shown in Figs. 5.3 and 5.4.
Figure 5.3 shows the rapid increase in the required
separation length with decreasing bubble size and shows

4. MSR Program Semiannu. Progr. Rep. Feb. 28, 1969,
ORNL-4396, p. 137.

51

ORNL—DWG 71— 6714

6
z° N
£ a \\\BUBBLE DIAM =0.005 in.
pd
oE ~ N —
e ]
'."‘--..____ [ ——
g ~— T————_0.007
r 2 | B =—
: ———t— oo
—
[ ———
P =] 0.015
0.020
o} |
200 300 400 500 600

LIQUID FLOW RATE (gpm)

Fig. 5.4. Calculated separation length vs fuel salt flow rate for
4-in.-1D gas separator.

that the separation length is greater with salt than with

_ water. Figure 5.4 shows the effect of flow rate on the

separation length for several bubble sizes.

5.2 GAS SYSTEM TEST FACILITY
W. K. Furlong

The conceptual design of a molten-salt loop for
testing gas systems was completed, the conceptual
system design description was written, and work is now
beginning on the preliminary design. The facility will be
used for developing the technology of the fuel salt for
MSRs and in particular for tests such as the perfor-
mance of the bubble injection and separation compo-
nents and the off-gas handling system. In addition,
several other important tests will be run over the life of
the loop. These include measurement of surface tension
as a function of UF,/UF; ratio, the study of noble gas
and possibly tritium distribution, the study of bubble
dynamics, the measurement of heat transfer coeffi-
cients, and the measurement of the product of mass
transfer coefficient and bubble surface area.

The facility is scheduled for initial operation in the
first quarter of FY 1973.

5.3 MOLTEN-SALT STEAM GENERATOR
J. L. Crowley  R. E. Helms

Our program plan for obtaining reliable steam genera-
tors for the MSRP consists of the following three
phases.

Phase 1 — preliminary development of molten-salt
steam generator technology. Some conceptual designs
and proposals for development requiréments will be
prepared by industrial manufacturers of steam genera-
tors.> We will independently prepare a molten-salt
steam generator development basis report which will
also show which elements of the LMFBR program will
be of use. In addition, some preliminary information
will be obtained from some small-scale tests. Phase I is
expected to allow us to proceed with the engineering
development, procurement, and testing program for the
MSBE steam generator.

Phase II — completion by the industrial mamifac-
turers of a preliminary design of a steam generator
based on the conceptual design and completion of the
engineering tests defined in phase I These tests will
include the operation of some full-size steam tubes to
be supplied by the designer in a molten-salt steam
generator tube test stand (STTS). The completion of
phase II should place the manufacturers in position to
proceed with the detailed design of our MSBE steam
generator.

Phase [l — detailed design, fabrication, and operation
of a prototype steam generator in the MSBE or in a test
facility with a very large heat source. With the
experience gained from this phase, the manufacturer
should be capable of supplying reliable steam generators
for future needs of the MSBR program.

A discussion of the present activities of phase I
follows.

5.3.1 Steam Generator Industrial Program

Efforts continued toward obtaining conceptual de-
signs and development program proposals from a
qualified industrial manufacturer of steam generators. A
package consisting of a summary of the 1000-MW(e)
MSBR reference design, a proposed scope of work for
the industrial program, and a cover letter requesting an
expression of interest in receiving a request for proposal
was sent to nine industrial firms. Eight firms responded
affirmatively, and the request for proposal package was
sent to them. We hope to enter into a CPFF contract
with one firm during this fiscal year to begin work on
the four tasks of the industrial study.’

5.3.2 Steam Generator Tube Test Stand (STTS)

The conceptual system design description (CSDD)
was completed and issued as an internal ORNL memo.
This CSDD describes the 3-MW test facility to be built
as part of phase II of the molten-salt steam generator

5. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL4622, pp- 39-40.

52

program plan previously discussed. Further work in this
area has been suspended until late in FY 1972,

5.3.3 Molten-Salt Steam Generator
Technology Facility (SGTF)

The steam generator technology facility is being
considered for use in testing sections of steam generator
concepts or portions of single-tube heat exchangers
while generating steam at heat rates of 150 kW or less.
As an aid in preparing the requirements for such tests
we are examining several combinations of tube-in-tube
test sections to determine what would be most useful.
Calculations indicate that these small tests can be useful
for tubes up to % in. in diameter and 33 ft long
producing steam at 1000°F and up to 3600 psi
pressure. A conceptual design of the equipment to do
these tests is being prepared.

5.3.4 Development Bases for Steam Generators
Using Molten Salt as the Heat Source

Work has begun on preparation of a development
basis report for molten-salt steam generators. This
report will review the present technology of similar
systems using sodium as the heat source and will point
out elements of the LMFBR program which will be of
use in the molten-salt technology program. This report
will identify molten-salt steam generator problem areas
needing further fundamental study and development
and will outline a program for the development efforts
needed for designing and fabricating reliable molten-salt
steam generators.

5.4 SODIUM FLUQROBORATE TEST LOOP
| A. N. Smith

The test program was completed during the previous
report period.® Following the final shutdown of the
loop, the pump rotary element, the bubbler tube which
served as the BF; feed line and as an indication of the
salt level in the pump bowl, the spark plug salt level
probes, and the pump bowl pressure control valve were
removed from the system for inspection.

The general appearance of the pump impeller and
adjacent surfaces indicated that the 11,567 hr of
fluoroborate salt circulation did not cause excessive
corrosion of the Inconel system. Deposits of salt were
found in the upper part of the rotary element near the
thermal barrier, about 4 in. above the normal level of

6. Ibid., p. 41.
salt in the pump bowl. We concluded that the nature
and cause of these deposits were the same as for similar
deposits observed in May 1968.7 The BF; feed and salt
level bubbler tube (Inconel) was found to contain
deposits of several different corrosion products, includ-
ing one at the mouth of the tube which had severely
restricted gas flow. Metallographic and chemical analy-
ses were made to determine the nature of the deposits
and the extent of damage due to corrosion. Metallurgi-
cal analysis of the Hastelloy N spark plug salt level
probes is incomplete. Visual examination of the stem
and seat of the off-gas pressure control valve revealed
no evidence of corrosive attack. The stainless steel valve
was in service throughout the entire test program,
operating at room temperature with a C, 0of 0.01,a AP
of 24 psi, and a gas flow of 1.5 liters/min consisting of
helium with 3.5% BF;. The material below presents a
more detailed discussion of the observations and con-
clusions resulting from the examination of the pump
and BF; feed and salt level bubbler tube.

5.4.1 Pump Bowl and Rotary Element®

The pump rotary element was removed from the
pump bowl in September 1970. The appearance of the

PHOTO 75584

AFTER 187 hr SALT CIRCULATION
(APRIL 1968 )

53

pump bowl as viewed from the north side is compared
with the same view taken in April 1968 after only 187
hr of salt circulation (see Fig. 5.5). This time consider-
ably more debris, presumably corrosion products, was
present on surfaces visible through the fountain flow
windows. The material probably fell from the thermal
barrier region when the rotary element was removed.
The sharp edges on the metal surfaces at the entrance to
the volute, which is adjacent to the impeller tip (tip
speed of the 9-in. impeller was about 70 fps), is
interpreted as an indication that the combined corro-
sive-erosive attack was slight.

The as-removed appearance of the pump rotary
element is shown in Fig. 5.6. Here again, the sharp
edges at the impeller discharge openings are an indica-
tion of insignificant attack. The debris on the horizon-
tal surface just above the impeller probably fell to that
point from the vicinity of the thermal barrier. The
presence of the white cake of salt on the thermal barrier
surface (an annulus about Y, ¢ in. wide is formed by this

7. MSR Program Semiannu. Progr. Rep. Aug. 31, 1968,
ORNL-4344, p. 78.

8. Reactor Handbook, vol. 1V, Engineering, 2d ed., p. 824,
Wiley, 1964.

AFTER 11,567 hr SALT CIRCULATION
(SEPTEMBER 1970)

54

PHOTO 78058

THERMAL
BARRIER

GREEN
SALT
DEPOSITS

THERMAL
RADIATION
BAFFLES

GREY-BLACK MATERIAL
HAVING HIGH NICKEL
CONTENT

Fig. 5.6. Post-test appearance of salt pump rotary element, NaBF circulation test, PKP loop, 9201-3.
surface and the pump bowl neck) is an indication that
salt was forced into this region, probably during the
ingassing transients.” The dark material in the openings
immediately above the heat baffles is primarily green
salt similar to that observed in May 1968.7 The
chemical analyses of the two green salt deposits (Table
5.1) show the later deposit to have a higher iron and
chromium content, a lower uranium and thorium
content, and a significantly different U/Th ratio. The
region of the green salt deposits is about 4 in. above the
normal level of salt in the pump bowl. The appearance
and composition of the deposits seem to argue against
formation from vapor or mist. The inference is that the
green salt is a relatively insoluble phase formed by
combination of the flucroborate salt with residue of the
fuel (BULT-4) salt formerly used in the loop and that
the deposits were formed during periods of abnormal
salt level when the bulk salt came into contact with the
cool surfaces in the thermal barrier region. The only
known periods of abnormally high salt level were during
the several brief ingassing transients which occurred
during the operation with flush salt. We conclude that
the deposits found in September 1970 must have been
formed sometime between May 16, 1968, when opera-
tions were resumed after inspection of the rotary
element, and June 24, 1968, the date of the last
ingassing transient.

A gray-black piece of material found outside the
fountain flow window (Fig. 5.6) was magnetic and had
the following chemical analysis:

Element Weight %
Na 5.0
B 3.0
Fe 2.5
Cr 0.2
Ni 66.3
F Assumed remainder

We think that this deposit -precipitated from the
fountain flow (upper labyrinth seal leakage) stream at a
point near the impeller hub where the temperature was
below the bulk salt temperature. We can speculate that
the circulating salt stream was saturated with nickel at
normal circulation temperatures, since it has been
reported by others'® that the nickel ion has a low

solubility in the fluoroborate salt and tends to deposit

in low-temperature regions. The nickel-bearing corro-
sion products would thus tend to separate out at the

9. MSR Program Semiannu. Progr. Rep. Aug 31, 1968,
ORNL-4344, p. 76. '

10. MSR Program Semiannu. Progr. Rep. Féb. 28 1970,
ORNL-4548, p. 242, '

55

Table 5.1. Chemical analyses of green salt deposits

Analysis (wt %)

Element
September 1970 May 1968
Na 17.8 11.0
B 2.6 3.9
Li 0.17 0.21
Be 0.10 0.04
U 0.35 12.2
Th 13.9 26.5
Fe 9.45 1.5
Cr 8.9 0.27
Ni 0.08 0.03
F 45.3 43.5

first convenient cool surface external to the isothermal
circulating stream. None of the green salt was found in
this area, which implies that the bulk salt at 1025°F
was not saturated with respect to the green salt.

5.4.2 BF; Feed and Salt Level Bubbler Tube

The gas bubbler tube which was used for BF;
addition and salt level indication was removed from the
pump bowl in October 1970. The %-in.-OD by % -in.-
ID Inconel tube had been in service for the entire test
program from March 1968 through April 1970, during
which time salt circulated 11,567 hr. The tube was 13
in. long, and the bottom end contained one triangular
notch Y% in. wide by % in. deep. Although operating
conditions were varied briefly for special tests, during
most of the test program the total gas flow was 370
cm® /min, and the gas was helium containing 13.5% BF;
by voklu,me. During periods when the loop was shut
down ‘and drained, the BF; flow was stopped, argon
was used in place of the helium, and the flow rate was
somewhat reduced. The total volume of gas which
passed through the tube was about 0.26 X 10° liters
during salt circulation and about 0.4 X 10° liters during
the entire test program. Note that inert gas was
routinely added to the pump bowl through other lines,
so that the total pump bowl gas flow was normally
1500 ¢cm® /min. The loop was filled and drained a total
of 16 times during the test program, and the pump
bowl salt level varied somewhat from one operating
period to the next. Also, the level varied whenever the
salt temperature was changed. The average salt level was
2Y, in. above the bottom of the tube during the flush
salt operating period and 4% in. above the bottom
during the clean salt operation. In each case there were
variations of not more than *1 in. . '

During the final 20 days of loop operation, the salt
level instrument showed a steady increase of about 0.06
PHOTO 78258

INCHES

Fig. 5.7. Interior view of BF 3 feed and salt level bubbler tube, NaBF; circulation test, PKP loop, 9201-3.

in./day. This was interpreted as an apparent increase,
probably caused by an increase in flow resistance (plug)
in the bubbler line. During the final salt drain, when the
salt level was decreasing, the indicated salt level rose
and went off scale (100% = 20 in. H,0), an indication
of a plugged line.

The appearance of the exterior of the tube was as
follows: the opening at the bottom end was completely
filled with a silver-gray metallic-appearing magnetic
deposit; the lower 4 in. of the exterior surface had a
bright silvery-grainy appearance and was very rough to
the touch; the next 3 in. was dark, not as rough as the
lower region, somewhat grainy in the lower half, and
had a thin coating of bright green solid on one side; the
upper 4 in. was gray-black to black with a greenish cast.

The tube was cut into four pieces by cuts at 1, 4%,
and 7% in. from the lower end. The three lower pieces
were then slit axially. The inside of the tube appeared
as follows (Fig. 5.7). The plug at the bottom end was %,
to % in. thick; for about % in. above the plug, the wall
was covered with a thin black film; for the next % in.
there was a thin deposit which was mostly light gray
with a sprinkling of light green particles. From the
1%-in. level to the 4%e-in. level, bright green material
was deposited and from 4% to 7% in. there appeared
to be the same green material covered with a loose
black magnetic powder. The deposit was %, to % ¢ in.
thick for the most part, except that it was somewhat
thicker at about the 2-in. level and filled about 90% of
the tube cross section at the 5%-in. level. In cross

section, the plug at the bottom end of the tube had a
white, mottled appearance as though the metallic
deposit was interspersed with lumps of salt. The green
deposit had a yellowish cast.

Metallographic examination of the tube, under the
direction of Koger,'' produced the following conclu-
sions. About 5 mils of material was removed from the
Inconel tube. As expected, attack by salt and by vapor
was evident, but no void formation occurred; corrosion
product was left on the inside of the tube from vapor
attack and was deposited on the lower opening by mass
transfer from the bulk salt.

Identification of the various materials found in the
tube was made under the direction of Cantor.'?
Analyses by x-ray diffraction and electron microprobe
indicated that the green salt was Na; CrFg, the yellow
salt was predominantly NaNiF; , and the black, powdery
magnetic material was a mixture of nickel and iron in a
ratio stoichiometrically equivalent to the alloy Ni; Fe.
The plug at the bottom of the tube was a mixture of
metallic nickel and NaBF, in approximately equal
parts.

Cantor surmised that the NazCrFg and NaNiF; were
produced by reaction of the salt with moisture intro-
duced by the various gas streams. The reaction products

11. J. W. Koger, part 4, this report.
12. S. Cantor, sect. 9.3, this report.

probably accumulated as a scum on the surface of the
salt and were subsequently transferred to the inner
walls of the tube by some mechanism, most likely in
liquid droplets formed by the continuous breaking of
gas bubbles. Routine surveillance and control were not
maintained over the water content of the incoming
helium and BF;, so we do not have an accurate measure
of the quantity of water which was added to the system
in the gas feed. However, based on limited observa-
tion,'® the water content of the BF; is thought to have
been less than 50 ppm by volume. The helium was
supplied from a source which is known to be of high
purity (less than 1 ppm H,O). Brief tests with a
moisture monitor indicated that the helium stream at
the loop contained about 20 ppm, so we believe that it
certainly contained less than 50 ppm H,O. The
calculated rate of addition of water to the system
would be about 0.8 g of H, O per 10° liters of gas per
ppm H,0 by volume. Therefore, if we assume a
maximum concentration of H,O in the feed gas of 50
ppm by volume, the maximum amount of water added
during salt circulation is estimated to be 40 g total
injected into the pump, with 10 g entering through the
bubbler tube and the remainder through the shaft
purge. These numbers represent water input rates of
about 1 and 3.6 mg/hr, respectively, when averaged
over the 11,567 hr of operation but do not include the
10 g of water injected instantaneously in a.special
test.'* We do not know the weight of material which
was deposited in the tube, nor what fraction of the
total corrosion product this represented. However, by
making rough assumptions regarding volume and den-
sity, we estimate that the total amount of material in
the tube was probably less than 10 g, and on this basis
it appears possible that water in the gas feed could have
been an important contributing factor in the plugging
of the tube.

5.5 COOLANT SALT TECHNOLOGY FACILITY
A. 1. Krakoviak
A conceptual system 'design description (CSDD) for
the coolant salt technology facility (CSTF) has been

prepared (ORNL-CF-70-12-18), and preliminary design
of the facility and components has begun.

13. R. W. Apple to A. N. Smith, personal communication,
Mar. 5, 1971.

14. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,

ORNL-4622, p. 41.

57

We plan to use the piping (5 in. sched 40), the pump,
and other components from the MSRE coolant system
to upgrade the existing sodium fluoroborate test loop
(constructed of Inconel) to a facility constructed of
Hastelloy N. Where applicable, the existing drawings of
the PKP-1 pump test stand, service piping, containment
enclosure, cover-gas supply, and off-gas disposal facili-
ties are being reused. : o

The proposed salt loop will have a pump, a’salt-to-air
heat exchanger (to dissipate pump power), a drain tank,
and the necessary connecting pipe. An orifice will limit
the loop flow to ~600 gpm at a pump discharge
pressure of ~100 psig and a suction pressure of ~25
psig. The loop will be fabricated with matching sets of
nozzles on the discharge and suction lines to permit
later attachment of parallel salt circuits. These side
loops will be used to provide the additional information
needed to assure that the sodium fluoroborate—sodium
fluoride eutectic does indeed meet the requirements of
a secondary fluid for molten-salt reactors.

The new facility, with its side loop, will supply
additional information on .

1. boron trifluoride cover-gas addition to and removal
from the pump bowl,

chemical removal of water from the fluoroborate
salt at rates typical of small steam inleakage,

3. detection and removal of corrosion products that
would result from small steam leaks into the salt,

. control of the salt chemistry so that the coolant may
also serve as a sink for the tritium inherently
produced by lithium-based reactor fuels,

5. operation of a steam generator test heated by
molten salt,

. heat and mass transfer near the cold wall of a steam
generator,

7. operation of valves in molten salt.

In some cases, further fundamental chemical and/or
metallurgical studies are needed before an engineering-
scale test can be made. Although the main facility is
designed primarily for testing with sodium fluoroborate
salt, it would also be suitable for testing other fluoride
salts.

At the end of this report period, the coolant system
piping had been removed from the MSRE and the
mechanical design of the loop was ~10% completed.
The main loop is expected to be complete at the end of
December 1971. .
5.6 MSBR PUMPS

L. V. Wilson
A. G. Grindell

H. C. Savage
H. C. Young

5.6.1 MSRE Mark 2 Fuel Pump

As previously reported,' during the shutdown of the
MSRE Mark 2 pump, we found that the system salt
could not be drained into the storage tank because the
connecting line was plugged. Various attempts to
promote draining by increasing the temperature of the
salt piping system including the freeze valve and by
increasing the differential pressure between the pump
and storage tank were unsuccessful. We then replaced
the drain line by use of a procedure in which all the salt
was frozen except that in the pump tank and in three
pressure transmitters. After the new installation was
inspected, the salt was melted carefully by heating the
pump suction and then the pump discharge piping
progressively outward from the hot salt in the pump
tank. The salt drained satisfactorily into the storage
tank, where it was frozen, and the facility was secured.

The entire length of the plugged drain line was x
rayed, and suspect sections were removed, separated
into two symmetrical longitudinal pieces (shown in Fig.
5.8), and examined by the Reactor Chemistry Division.
The plug was located approximately 32 in. downsfream
of the junction of the drain line with the main loop
piping and 44 in. upstream of the freeze valve.

A considerable amount of relatively large (0.5 mm
long) zirconium oxide crystals,'® which may be seen in

15. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,
ORNL-4622, p. 6.

16. R. E. Thoma, Examination of Specimens from the MSRE
Fuel Salt Pump Test Stand, MSR-7061 (Nov. 17, 1960)
(internal correspondence).

SECTION I-J

Fig. 5.8, was found in section I-J of the drain line. The
zirconium oxide crystals, section I-J, are compared with
normal salt, section K-L, in the figure. We hypothesize
that the contaminants needed to produce the plugging
material were introduced into the salt system by the
argon purge gas and during the approximately 60 times
the system was opened to perform various maintenance
activities and that they precipitated in a cool section of
the drain line. Additional details of the draining
operation are presented in Incident Report ORNL-
70-36-Final.

5.6.2 MSRE Salt Pump Inspection

A visual inspection was made of the rotary element
and the conveniently accessible interior surfaces of the
pump tank in the MSRE coolant salt pump as the
rotary element was being removed from the coolant salt
system. The observations are described below. The salt
had drained well from all salt-exposed surfaces. The salt
flow passages in the impeller, volute, and suction line
were clean, and no evidence of wear of any kind was
noticeable. A black carbon-like residue was seen on the
exterior surfaces of the shield plug and its extension,
both of which are parts of the rotary element, down to
the salt-gas interface. There was a similar residue on the
pump tank surface directly opposite the shield plug and
on the cylinder which supports the pump volute down
to the salt-gas interface. Apparently oil leakage from
the lower shaft seal had drained from the catch basin,
run down the outside of the shield plug, and decom-
posed at the 1025°F temperature in the pump tank.
These observations confirm the suspected leakage of a
mechanical seal in the oil drainage system. This seal was
replaced with a seal weld in the spare rotary elements
for the fuel and coolant salt pumps. The rotating

PHOTO 78332

SECTION K-L

Fig. 5.8. Plugged drain line from Mark 2 pump test stand. Section I-J shows presence of zirconium oxide crystals in the salt;

section K-L shows frozen normal salt.

components, including the impeller, shaft, drive motor
rotor, and all bearings and shaft seals, turned easily and
smoothly to the hand, lending support to our belief
that the pump would have operated satisfactorily at
high temperature for many additional hours. Additional
inspection of the pump will be performed while it is
being prepared for use in one of the technology loops.

5.6.3 ALPHA Pump

Water testing of the ALPHA pump,'” a centrifugal
pump designed for capacities to 30 gpm, 250 ft head,
and temperatures to 1400°F, was completed. Data
indicate that the design hydraulic performance condi-
tions can be met at a shaft speed of approximately
6750 rpm.

The pump is to be used initially in molten-salt
forced-convection loop’® MSR-FCL-2. The pump satis-
fies the MSR-FCL-2 hydraulic requirements of 4 gpm at
approximately 110 ft head at a shaft speed of approxi-
mately 3900 rpm. Based on the water test results, the
final design of the pump tank and internal parts was
completed, and the parts were fabricated of Hastelloy N
for use in the pump unit for MSR-FCL-2.

Further water testing of the ALPHA pump is not
planned at this time because the only available rotary
element will soon be installed in MSR-FCL-2. However,
additional testing may be desirable in the future to
measure the fountain flow rate and to define in more
detail the shaft deflection characteristics at higher
pump capacities (speeds above 6000 rpm).

5.7 REMOTE WELDING

P. P. Holz W. A. Bird
C. M. Smith, Jr. W, R. Miller

Most of the work previously supported under the
MSRP remote welding program has been transferred
into an automated welding program sponsored by the
LMFBR program and will be reported under that
program in the future. The new LMFBR program is
directed toward developing suitable equipment, welding
procedures, and programmed controls to produce butt
welds with consistently high quality in pipe to meet
LMFBR needs for reactor construction and loop main-
tenance. It is planned to extend automated welding
technology to pipe sizes as large as 30 in. diameter and
to field test the equipment at construction sites. The

17. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,
ORNL-4622, p. 45.

18. Ibid., pp. 176—78.

59

LMFBR program makes use of the prototype machine-
ry and controls developed in the MSR program.19-21
We are continuing 'to sponsor the work necessary to
complete ‘and test the weld torch positioning mecha-
nisms and "control circuitry; to define and document
remote maintenance inspection, viewing, and alignment
criteria; and to investigate pipe cleanliness requirements
for maintenance welding in salt systems.

5.7.1 Automatic Controls

We are striving to develop a welding system that can,
without human intervention, handle all the perturba-
tions, misalignments, etc., that are normally encoun-
tered in welding. This effort should not be confused
with programming the controls for automated welding.
Programming simply provides preset values for specified
variables at a prescribed time. A self-adaptive system
must have the ability to correct for unexpected
conditions at any time during the execution of the
weld. Thus, the system mechanizes operator capabilities
and reduces the level of skill required of welding
personnel. Such a system is necessary for completely
remote welding in radioactive areas.

An automatic horizontal torch control and an auto-
matic torch oscillator control are two new devices that
are being added to our present equipment, which
already includes an operational automatic vertical torch
control. The automatic vertical torch control employs a
servo-drive motor mechanism to control the torch-to-
work spacing. The system steps the torch vertically
within a narrow range to keep the arc voltage nearly
constant and thus permits the wire feed rate control to
operate as a fine control vernier. This combination of
controls produces uniform weld metal deposits while
maintaining proper arc voltage for uniform weld pene-
tration. The automatic horizontal control will work in
conjunction with automatic vertical control to position
the torch at the optimum location between the beveled
edges of the joint prior to making the weld pass. A
second motor has to be added to drive the torch
assembly horizontally across the weld groove area. The
present equipment utilizes a dc motor with an adjust-
able eccentric mechanism to oscillate the torch in a
fixed pattern. Variations in the side wall of the groove

19. MSR Program Semiannu. Progr. Rep. Aug 31, 1969,
ORNL-4449, pp. 79-82.

20. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, pp. 74-178. A

21. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,
ORNL-4622, pp. 45-50.
sometimes produce insufficient tie-in or, at the opposite
extreme, contact between the electrode and the side
wall The improved system will continually sense the
arc voltage and control the oscillator drive motor to
keep the electrode within the proper distance from the
wall of the groove. |

5.7.2 Pipe Cleaning Tests

A -series of experiments to determine cleanliness

60

requirements for pipe joints to be welded and to check

into possible contamination effects from salt plugs in

the pipe near the weld zone produced promising results. -

We were able to make butt welds manually to mieet

nuclear code x-ray standards in ‘- and % -in. standard--

weight Inconel pipe that were plugged solid with
LiF-BeF,-ZrF,;-ThF,-UF,; (68.4-24.6-5.0-1.1-0.9 mole
%) by merely clearing a distance of 1 in. on both sides
of the weld joint. Special provisions were made for
purging the pipe interior. Pipe cleaning consisted in
drilling out the plug and then wiping interior and
exterior pipe surfaces with emery cloth to remove salt
films and scale and to produce a reasonable surface
finish. We were careful not to polish the surfaces
because this would be difficult to do under working
conditions in a reactor. Weld preparation and welding
were done in a glove box.??

22, P. P. Holz, Butr Welding Experiments with Salt Plugged
Pipes, ORNL-TM report (to be issued).
6. MSBR Instrumentation and Controls

6.1 DEVELOPMENT OF A HYBRID COMPUTER
SIMULATION MODEL OF THE MSBR SYSTEM

O. W. Burke

6.1.1 Introduction

As an extension of the work done by.Sides,' a hybrid
computer simulation model of the MSBR has been
under development for the past three months. The new
model of the reactor, primary heat exchanger, etc., up
to the steam generator, is essentially the same as that
used by Sides. The model of the steam generator is
similar to that reported in ref. 2. The bulk of the
discussion here will pertain to the _steam generator
model.

6.1.2 Steam Generator Model

The steam generator is'a countercurrent, single-pass
U-tube exchanger approximately 77 ft long and 18 in.
in diameter. At the steady-state design point condition,
water enters one end of the exchanger at a temperature
of 700°F and at a pressure of 3752 psia. Salt enters the
opposite end at a temperature of 1150°F. The salt
flows through the shell side of the exchanger ‘and the
water through the tube side.

The mathematical model of the system consists in the
conservation of mass, conservation of momentum, and
conservation of energy equations of the salt and water.
These equations are written in one space dimension, x
(the direction of water flow), and time, ¢. In this initial
model the variations in the density and velocity of the
salt are neglected, and hence only the conservation of
energy is considered for this part of the system. If it
appears that a more detailed model of the salt side of
the exchanger is warranted, it will be added later. The
following equations were used:

1. W. H. Sides, I1., Control Studies of a 1000-Mw(e} MSBR,
ORNL-TM-2927 (May 18, 1970).

2. C. K. Sanathanan and A. A. Sandberg, University of
[llinois, Chicago, and F. H. Clark, O. W. Burke, and R. S. Stone,
ORNL, “Dynamic Modeling of a Large Once-Through Steam
Generator,” to be published in Nuclear Engineering and Design.

Conservation of mass-(water),
ap , 0 A
3 e PV =03 | o | (1)

conservation of momentum (water),

a(pv)
or  ox ax

kop '
)= ey )
conservation of energ'y (wafer); '
) ;) _ ' |
B;(ph) +a(phv)—k1H(9 - T); : (3)

conservation of energy (sait),

(T—-06); 4)
the equations of state for water,
T="T(p, h); | )
p=pp, h). | (6)

"The definitions of the variables used in the above

equations are as follows:

T = water temperature,

p = water density, |

v = water velocity,

P = water pressure,

¢ = coefficient of friction, ~

k-= constant used to make units consistent,
h = specific enthalpy of water,
H = heat transfer coefficient, salt to water,

k, =ratio of the surface area of a tube to the water
volume in the tube, . :

k,=r1atio of the surface area of a tube to the salt
volume adjacent to the tube,
p, = salt density (assumed constant),

p

= gpecific heat of salt at constant pressure,
8 = salt temperature, '

w = salt velocity.

Boundary conditions on p and A are applied at the
water entrance to the exchanger, on # at the salt
entrance, and on v at the throttle. The critical flow at
the throttle is expressed by the following nonlinear
relationship among the system variables at a point just
before the throttle:

AT p

2
=M ,
o (Am) <1 + bT>

where A, is the instantaneous value of the throttle
opening, Ar o the initial steady state value, M the
critical flow constant, and 5 an empirical constant.
Ag o is taken as 1.0 and Ay is varied as a function of
time during transients.

It was determined in previous work? that a con-
tinuous-space, discrete-time model is most satisfactory
for this steam generator simulation. By simplification of
Egs. (1), (2), (3), and (4), and using the backwards
differencing scheme for the time derivative, the fol-
lowing ordinary differential equations are generated:

(7)

2 V—vV
dp__pvdv_ev POV (M)
dx k dx k k At
dh _ 1 h—hy
S ylkHE - D] -, (2M)
dv__vdp PPk (3M)
dx pdx pAt

Hk,(T-6) 6-16
d@z_ 2( )+ k (4M)

dx PsCp Vs v At
where v, is salt velocity.

In the above equations, the nonsubscripted variables
are the ones being iterated for the values at the end of
the (k + 1) time increment, while the variables with the
k subscripts represent their values at the end of the kth
time increment. The time increment is represented by
At.

By judicious choice of the direction of integration in
space, of the various dependent variables, an initial
value problem can be formed. Since the water enthalpy
h and the water pressure p are known at the water
entrance end of the exchanger (left end), these variables

62

will be integrated from left to right. For the same
reason the water velocity (it can be calculated at the
throttle) and the salt temperature will be integrated
from right to left. .

In the hybrid program the integrations are performed
by the analog computer. The digital computer calcu-
lates the terms of the derivatives of the differential
equations, provides control for the calculation, and
provides storage. The thermodynamic properties of
water are stored in the digital computer as two-
dimensional tables. An interpolation routine is used to
get values from these tables. In the present model the
derivative terms are updated at 1-ft intervals in the x
dimension.

The calculational procedure for a Ar step is as
follows: With left boundary values of water enthalpy
and water pressure as initial conditions, a left to right
integration of these two variables is started. As the
integration proceeds in x, the derivative terms of the
differential equations are updated at 1-ft intervals.
Values of p and & are stored at these 1-ft locations in x.
The integrations are stopped when the x location
corresponding to the water exit end of the steam
generator is reached. In a procedure identical to that
above, and with the current values of p and A, the salt
temperature and the water velocity are integrated from
right to left. The initial condition of the salt tempera-
ture is that at the point where it enters the exchanger.
The initial condition of the water velocity is the
calculated velocity at the throttle. When the right to
left integrations have proceeded to the left boundary,
they are halted. The left to right integration of p and h
is repeated, using the current values of p, 4, v, and 6.
The right to left integration is repeated, etc., until the
convergence is satisfactory. It appears that the A¢ can
be of the order of 1 sec.

The model of the reactor; primary heat exchanger,
piping, etc., is a continuous-time model similar to those
traditionally used on analog computers. It was felt that
the detail required in these pélrts of the system was not
as great as that required in the steam generator.

The steam generator model will be coupled with the
analog model of the remaining parts of the system in
the following manner. The analog model of the system,
exclusive of the steam generator, will be slowed down
in time by a factor of 10 or 100, as required. At
intervals of approximately 1 sec the water pressure and
temperature at the left end of the steam generator and
the salt temperature at the right end of the steam
generator are read and stored in the digital computer.
These values are used for a At step calculation on the
hybrid computer as described above. At the end of the

~
At calculation, output values of p, A, 8, T, and v are
stored and fed back to the analog simulation. This
procedure is repeated at time intervals of Ar. If the
analog simulation is slowed down by a factor of 100,
for instance, the 1-sec At of the hybrid calculation will
look like 1/100 sec to the analog model. By using a
short time constant first order lag on the discrete
outputs from the hybrid model, it is hoped that the
outputs from the hybrid model of the steam generator
to the analog model of the rest of the system will be
fairly smooth in time. {

63

The analog model of the system, exclusive of the
steam generator, has been developed and is ready to be
patched on the analog patch boards. The patching and
debugging of this model will require a few days.

The hybrid program for the steam generator has been
written, and practically all of the debugging has been
accomplished. It is expected to be running within a very
short time. The integration of the two models will
require some time. The model should be complete
before the target date of June 30, 1971.
7. Heat and Mass Transfer and Thermophysical Properties

H. W. Hoffman

7.1 HEAT TRANSFER
J. W.Cooke

Heat transfer experiments employing a proposed
MSBR fuel salt (LiF-BeF,-ThF,-UF,, 67.5-20-12-0.5
mole %) flowing in a horizontal tube have shown that
the local heat transfer coefficient varies along the entire
length of the tube in the Reynolds modulus range 2000
< Ng, < 4000 for heat fluxes from 0.7 X 10° to 3.0 X
10° Btu hr™' ft™2. It is hypothesized that a delay in
transition to turbulent flow could result in such a
variation in the transitional flow range.'™

To investigate the effect of flow development on heat
transfer, a new test section, shown in Fig. 7.1, was
installed in-the inert-gas-pressurized molten salt heat
transfer system. This test section consists of a 48-in.
length of 0.25-in.-OD by 0.035-in.-wall Hastelloy N
tubing, with three electrodes welded to the test section
so that the left half, the right half, or the entire length
can be resistance heated. At each end are “disk-
donut”-type mixing chambers.

Since the flow alternates in direction, six different
modes of operation of the heat transfer system are
possible. For the initial studies, the left half was heated
to provide a heat flux range from 0.5 X 10° to 4 X 10°
Btu hr™' ft™, and the right half was heated just
enough to make up for the radial heat loss to the
surroundings (~0.05 X 10° Btu hr™! ft™?). Thus with
flow to the left there is a 24-in. adiabatic-hydrodynamic
entrance length, and with flow to the right there is a
round-edge nozzle-type entrance.

Over 200 calibration and data runs have been con-
ducted with this new test section at two temperature
levels (~1070 and ~1250°F) and three Reynolds

1. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,

ORNL-4622, pp. 53-57.

2. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,

ORNL-4548, pp. 87-88.

3. MSR Program Semiannu. Progr. Rep. Aug. 31, 1969,

ORNL-4449, pp. 85-89.

64

J. J. Keyes, Jr.

Table 7.1. Thermophysical prdperty data for molten salt
mixture LiF-BeF,-ThF4-UF4 (67.5-20-12-0.5 mole %)
used in the present calculation

Uncertainty
(%)
u (b £t ™ hr1) = 0.187 exp {8000/T(°R)]? £25
k [Btuhr! ft7! CF)71] =0.69° 12
p (Ib/ft3) = 230.89 — 22.54 x 1073+ CF)? +3
Cp [Btulb™! CF)71) =0.324% +4

43. Cantor (ed.), Physical Properties of Molten-Salt Reactor
Fuel, Coolants, and Flush Salts, ORNL-TM-2316 (August
1968).

bMSR Program Semiannu. Progr. Rep. Aug. 31, 1969,
ORNL-4449, p. 92.

moduli (~3500, ~7000, and ~12,000) covering a heat
flux range from 0.2 X 10° to 3.4 X 10° Btu hr ™! ft72.
Data from 27 of these runs at the lower temperature
level have been thoroughly examined thus far.

The data have been analyzed as previously described®
with the exception of two fundamental differences.
First, the electrical heat generated in the tube wall
(corrected for heat losses), rather than the enthalpy rise
of the salt, was used to obtain the heat input to the salt.
Second, the local bulk salt temperatures were calculated
using the measured wall temperatures along the adia-
batic entrance length rather than the mixing chamber
temperatures. In addition, the more recent values of the
thermophysical salt properties given in Table 7.1 were
used,

Preliminary results for runs 1 through 27 are given in
Table 7.2. The odd-numbered runs were without (flow
from left to right) and the even-numbered runs were
with (flow from right to left) the 24-in. entrance length.
Examining runs 1 through 12 (5000 < Ny, < 8000),
one notes little variation in the heat transfer function,

4. MSR Program Semiannu. Progr. Rep. Aug. 31, 1967,
ORNL-4344, pp. 96—100.
65

ORNL-DWG 71-6715

2-ft HEA]}D SECTICN 2-ft ENTRANCE LENGTH
%

s
7 N
Lrél—ltingocoupLE //; . 3g-in. .
RADIUS (TYP)
] Z//;‘“n ‘_“.‘i‘L“‘,‘L‘“I‘:‘_- fh / \ EEE'Egé. ot
' II | } 7 11 \1 1
- || “; % WQ »\‘ rg\l iy
i A\*/,G-in. RADIUS (TYP) N £

0.50-in. OD x 42-mil WALL

0.25-in. OD x 35-mi
HASTELLOY N TUBING n X mil WALL

HASTELLOY N TUBING "DISK AND DONUT"
MIXERS

1-in OD x 60-mil WALL

HASTELLOY N TUBING

Fig. 7.1. Details of test section, entrance length, mixing chambers, and electrodes for pressurized flow system.

" Table 7.2. Experimental data for heat transfer studies using the salt mixture LiF-BeF,-ThF4-UF,4 (67.5-20-12-0.5 mole %)

Even-numbered runs were with and odd-numbered runs without an entrance length

Modulus® h

q/A ' Heat transfer
i‘:}n (Zi:n) ?olg)t Aon (Btuhr™! ft ™2 bliea:;e — p— = Btu hr! ft=3 CF)"! function,
' 5 x 1075) alance N, Np, Ny, [P0M ] Ng.r
1 1085.7 1089.2 10.7 0.24 0.24 5208 15.1 48.6 2234 19.72
2 1090.1 1093.7 10.2 0.24 1.26 5197 149 51.2 2356 20.89
3 1090.7 11022 44 4 0.80 0.78 5483 14.7 50.5 2322 20.50
4 1099.5 11113 34.1 0.80 1.15 5486 14.2 50.8 2338 20.83
5 1106.1 11299 74.2 1.69 1.00 6051 136 49.7 2285 20.32
6 1118.3 11424 74.6 1.69 1.19 6211 13.1 494 2273 20.49
7 11324 1156.1 70.1 1.69 1.006 6512 12.5 525 2416 22.15
8 11395 1186.6 1209 3.38 1.09 7056 11.8- 60.8 2798 25.70
9 11684 12154 150.8 3.38 1.01 7696 10.8°  66.3 3049 28.96
10 1181.3 12286 1112 3.37 1.08 7941 104 66.1 3039 29.24
11 12034 1226.7 59.9 1.69 1.03 8230 10.2 61.4 2823 27.89
12 1203.2 1227.6 62.6 1.69 1.16 7883 10.2 587 2698 26.64
13 10646 1077.6 67.3 0.59 1.00 3313 16.0 19.3 887 7.50
14 1068.8 10814 58.6 0.59 1.07 3152 15.8- 219 1006 8.57
15 1077.2 10999 105.8 0.99 1.01 3370 15.0 204 937 7.96
16 1082.7 1106.6 85.8 0.99 1.08 3264 14.7 25.1 1153 9.94
17 10934 1129.5 1468 1.5 0.98 3467 139 224 1031 8.86
18 11044 11422 131.0 1.5 1.09 3352 134 242 1113 9.75
19 11185 11557 136.1 1.5 1.02 3554 12.8 235 1082 9.60
20 1059.6 1093.0 1409 1.5 1.05 3251 156 226 1038 8.58
21 10763 1109.3 1474 14 1.00 3447 148 215 988 8.30
22 1089.3 11355 170.7 2.0 1.13 3591 13.8 2511 1157 9.87
23 1114.2 11636 1898 2.0 1.09 3749 127 234 1076 9.40
24 1125.8 12043 2544 29 1.11 3659 116 246 1134 10.01
. 25¢ -
26 1080.3 11509 253.5 2.9 1.02 3524 13.6 26.0 1197 10.02
27 1113.8 11874 2728 29 1.03 3787 122 235 1081 9.33

%These are the Reynolds, Prandtl, and Nusselt moduli, respectively, calculated using the average of the local heat transfer
coetg'lcient from the exi/t to within 5 in. of the test-section entrance.
— A7 a7 1/3 0.14
CNS-T=NNu/NPr (.U/.U-s) .
This run was deleted because flow was not constant.
66

ORNL-DWG 74-6716

16
|
] T
| —
||/
4N FROM W.M. McKAY, Npe =3260  Np =15.0 7
- CONVECTIVE HEAT _ 5 B
\\ AND MASS TRANSFER (REF.5) GA=1X10"  fgyiy =160
12 iy \ \sm_.__d
Y T e
\\//
AN
1.0 :_“:E-—
RUN 16
/
RUN 17—
0.6
1.4
==
G5 2 \ Npe=3525  Np, =13.6
Ol Q/A =3%105  hggiy=1300
[
Ak
8|3 A
2 40
oL -
& Lg \‘--.__ RUN 26 |
-l o8 N 1
Yi= ' RUN 27 ="
Q|5 06 :
O |
4l 14
4'::':
\
S Npe = 5500 Np, =14.4
G/A=08%10°  h,;4=2250
\ RUN 4
——TRUN 3
08 |
14
o
NRe=86500  Np,=12.8
"2 \ O/A=A7%40°  Fexiy=2250
/ P——
RUN 7
0.8 I
0 ' 20 40 60 80 100 120

DISTANCE FROM ENTRANCE
TUBE INSIDE DIAMETER

X/D,

Fig. 7.2. Variation in ratio of local to exit heat transfer coefficient with distance from the entrance for the salt mixture
LiF-BeF,-ThF4-UF, (67.5-20-12-0.5 mole %). Even-numbered runs with, odd-numbered runs without, an entrance length.
NS-T_[NS-T = (NNU/NEI’P)(“/MS)O'“" where NNu
and Np,  are the average Nusselt and Prandtl moduli
(based on all but the first 5 in. of the test section)
respectively; u/u; is the viscosity ratio, bulk fluid to
wall], beyond that which would be expected from the
variation in Reynolds modulus. However, for runs 13
through 27 (3150 < Ng, < 3787), the heat transfer
function averages 15% greater with a hydrodynamic
entrance length than without.

The influence of the entrance region on the heat
transfer coefficients can be more easily seen from plots
of the ratio of local to exit heat transfer coefficient as a
function of axial position, as shown in Fig. 7.2. (The
exit value obtained with the entrance length was used
for each pair of plots.) Also shown in Fig. 7.2 for
comparisons are the expected curves for similar en-
trance and flow conditions extrapolated from published
results for the simultaneous development of hydro-
dynamic and thermal boundary layers.® Examining the
lower-heat-flux, lower-Reynolds-modulus case first
(upper curves), one can see that, with an entrance
length (run 16) fully developed heat transfer coef-
ficients are obtained within an L/D of 30, which is in
agreement with expected behavior. However, with the
round-edged nozzle-type entrance (run 17) full develop-
ment was not obtained within the length of the tube.
With a higher heat flux, fully developed heat transfer
was not attained within the expected L/D of 40 even
with an entrance region (run 26). At a Reynolds
modulus of about 6000, there is less distinction
between heat transfer development with (runs 4 and 6)
or without (runs 3 and 7) the entrance region; and in
both cases the the length required for constant heat
transfer coefficients to be attained is in good agreement
with the predictions of ref. 5.

The undeveloped heat transfer behavior that has been
observed in this study at low Reynolds moduli is
difficult to explain in terms of commonly held theories
regarding the combined development of the hydro-
dynamic and thermal boundary layers in a circular tube
with a constant heat flux. In fact, the thermal boundary
layer should develop more rapidly for higher Prandtl
modulus fluids, and the local heat transfer coefficient
should be larger than the exit value until fully de-
veloped conditions are reached. However, a theoretical
study for incompressible flow has shown, in general,
that the flow of a fluid whose viscosity has a large
negative temperature dependence will be stabilized by

5. W. M. McKay, Convective Heat and Mass Transfer, pp.
186—-96, McGraw-Hill, New York, 1966.

67

neating and destabilized by cooling.® Specifically,
computer results show that the critical Reynolds
modulus (based on boundary layer thickness) for water
at 60°F flowing over a heated flat plate can be
increased by a factor of up to 22 above the expected
isothermal value, depending on the heat flux, Thus the -
distance required for the attainment of fully developed
turbulent flow in the molten-salt system (du/d¢ for the
salt at 1100°F is five times larger than for water at
60°F) could be greater than the length of the heated
test section. Unfortunately, except for studies of gas
flow over heated airfoils, no experimental results are
available to quantitatively verify those theoretical
studies. Our experimental system, however, with its
capability of reversible flow, matched temperature
measurements, uniform heat flux, and absence of
pump-induced turbulence, is uniquely suited to measure
precisely the effect of an adiabatic entrance region on
the heat transfer of a moderately viscous fluid near its
melting point. The local heat transfer coefficients and
temperature profiles obtained from these measurements
can then be used to check the stability theory.

7.2 THERMOPHYSICAL PROPERTIES
J. W. Cooke

7.2.1 Wetting Studies

The new technique' for determining the wetting
characteristics of liquids has been used to study the
wetting behavior of the molten salt LiF-BeF,-ZrF,-
ThF,4-UF,; (70-23-5-1-1 mole %) on a Hastelloy N
surface at 700°C (1292°F). The primary objective of
this study was to determine the extent to which the
wetting behavior of this salt could be controlled. The
wetting behavior of several molten fluoride salt mix-
tures has been found to be affected by addition of
various metals to the melt.” The controlling mechanism
is believed to be the oxidation state of the melt, which
affects the liquid-solid surface energy level.

In the present investigation, zirconium metal (in the
form of a-0.25-in.-OD Zircaloy rod) was added to the
melt to make it more reducing; nickel fluoride (in
powder form) was added to make it more oxidizing.

6. A. R. Wazzan, “The Stability of Incompressibie Flat Plate
Laminar Boundary Layer in Water with Temperature Depend-
ent Viscosity,” pp. 184-202 in Proceedings of the Sixth
Southeastern Seminar on Thermal Sciences, Raleigh, N. C.,
Apr. 13-14, 1970,

7. MSR Program Semiannu. Progr. Rep. Feb. 28, 1969,
ORNL-4396, p. 205.
Over a three-day period, 2000 observations of the
contact angle were made. Since the initial intention of
this study was to observe only the change from
nonwetting to wetting conditions, no attempt was made
to obtain precise values of the contact angle.

By comparing the actual pressure traces in Fig. 7.3
with the predicted traces,' we infer that the molten salt
mixture did not initially wet the Hastelloy N surface
(contact angle in the range of 140 to 160°). However, a
pressure trace, shown in Fig. 7.4, typical of wetting
(contact angle ~20°) was obtained within 20 min after
the introduction of the Zircaloy rod into the melt. Two
hours later, 1 wt % of nickel fluoride powder was added
~ to the melt. Five hours after adding the powder, the
melt was only partially wetting (contact angle ~90°), as
shown in Fig. 7.5. The melt became nonwetting again 7
hr after adding the nickel fluoride, with the pressure
traces again resuming the shape shown in Fig, 7.3.

Based on these results, a new apparatus was con-
structed to enable a more detailed study of the
electrochemical processes involved in the wetting be-
havior of metals in molten salts. To this end, a
micromanometer has been added to this system to
obtain more precise values of the contact angle. The
system is currently being used to investigate the wetting

behavior of the salt mixture LiF-BeF,-ThF,-UF,
ORNL-DWG 74-6717

[FY]

v

>

v

wn

L

v

a

TIME

Fig. 7.3. Pr&gsure trace with Hastelloy N unwetted by the salt
mixture at 700 C.

68

(67.5-20-12-0.5 mole %) when beryllium metal is added
to the mixture. -

ORNL-DWG 74-6718

L/

PRESSURE

TIME

Fig. 7.4. Pressure trace with Hastelloy N wetted by the salt
mixture at 700 C.

ORNL-DWG 71-6722

 —w

PRESSURE

TIME

Fig. 7.5. Pressure trace with Hastelioy N partially wetted by
the salt mixture at 700°C. '
7.2.2 Thermal Conductivity

Calibrations of the improved variable-gap thermal
conductivity apparatus were completed using water at
12°C. The plot of overall thermal resistance as a
function of gap thickness shown in Fig. 7.6 displays the
precision of this apparatus. Qur measured value of
0.00622 for water at 12°C agreed with published
values® to better than 5%, a factor of 3 less error than
was associated with water tests using the previous
apparatus. Since the apparatus is designed for operation
with corrosive fluids at temperatures up to 850°C, such
agreement with the specialized room temperature meas-
urements published in the literature is outstanding.

Soon after measurements were initiated with the
molten salt LiF-BeF, (66-34 mole %), a leak developed
in a weld in the concentrically guarded heat meter.
Attempts to repair this weld were unsuccessful, and a
new heat meter was fabricated. The eloxing process was
utilized in the fabrication of the heat meter so that the
welds could be eliminated. Conductivity measurements
for the salt mixture LiF-BeF; (66-34 mole %) have not
yet been resumed.

8. A. R. Challoner and R. W. Powell, “Thermal Conduc-
tivities of Liquids: New Determinations for Seven Liquids and
Appraisal of Existing Values,” Proc. Roy. Soc. A238, 90—106
(Dec. 4, 1956).

ORNL-DWG 7{-6719

[
P

. 30

T /
z

o

§ 25

o /

-

§ 20 e

a P

[ L

v

o /

w

T :

|

2 /

(14

w40 //

= L}

-

a

€ 5

wl

>

C
~

q% 0

0 002 004 006 008 010 02 014

Ax_ SPECIMEN THICKNESS (cm)

Fig. 7.6. Data obtained in measurement of the thermal
conductivity of water at 12°C using the improved variablegap
apparatus,

69

7.3 MASS TRANSFER TO CIRCULATING BUBBLES
T. S. Kress

7.3.1 Experiment

A transient technique is being employed in the
MSBR-related mass transfer studies to determine mass
transfer rates of a gas dissolved in a turbulently flowing
liquid to cocirculating helium bubbles.® Experiments
have been performed using helium bubbles having mean
diameters from 0.015 to 0.05 in. to extract oxygen
from five mixtures of glycerin and water (Schmidt
moduli of 419, 719, 1228, 2015, and 3446) over a
Reynolds modulus range from 8.1 X 103 to 1.6 X 10°.
Both horizontal and vertical flow in a 2-in.-diam
conduit have been studied.

It was previously reported'® that the bubble diameter
distribution is described adequately by a relation that,
when integrated, gives the interfacial area per unit
volume, a,, as a function of the volume fraction, ®, and
the number of bubbles per unit volume, V,:

ay=%Gn/2 B N3 @23 (1)

Note that it is necessary to evaluate a, in order to
extract mass transfer coefficients from the data relating
concentration to time. By making a small adjustment in
the coefficient to allow for the velocity due to
buoyancy, the volume fraction for flow in vertical
conduits is approximated by:

b= 0.73Qg/QL , (2)
where Qg/QL is the volumetric flow ratio of gas to
liquid. In horizontal flow, stratification of the bubbles
affects the volume fraction to the extent that the
vertical flow approximation applies only at high flows.
It has been found experimentally that the horizontal-
flow volume fraction can be correlated with the ratio of
the mean axial velocity of the fluid to the bubble buoy-
ant velocity in the radial direction, V,/V,. Assuming,
following Peebles and Garber,'' that the terminal
velocity, V,, is given by:

V,= (4g‘7/3c)1 2 >

9. MSR -Program Semiannu. Progr. Rep. Aug. 31, 1968,
ORNL-4344, pp. 74-75.

10. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL-4622, pp. 57-59.

11. F. N. Peebles and J. G. Garber, “Studies on the Motion of

Gas Bubbles in Liquids,” Chem. Eng. Progr. 49(9), 95 (Feb-
ruary 1953).
where
c=18.7/(V,dplw)°-**
the desired velocity ratio becomes:
‘Va/Vr = 0.0188QL(;J/,0)0'5 1 5/31 273

The variation of measured volume fraction with this
ratio is shown on Fig. 7.7 along with the corresponding
least-squares curve:

& =0.00179 + 0.021/(V,/V,) . (3)

Using Eqgs. (1), (2), and (3) to establish the interfacial
area, typical experimental mass transfer data for a 25%
mixture of glycerin and water (Ng. = 1228) can be
compared, as is seen in the upper portion of Fig. 7.8.
Those results include mass transfer which occurs down-
stream from the test section in the bubble separator.
Certain aspects of those data are particularly revealing.

ORNL-DWG 71-6720 .

T T T T
0.515, _1.272
BASED ON  I,/V,=0.0188 @ (u/p) J
0.008 F— u/ 3 H/p / ]
WHERE ©=LIQUID FLOW,gpm
p=LIQUID VISCOSITY, Ibm ft~' hr™!
p= LIQUID DENSITY, Iby/ >
0.007 & = BUBBLE MEAN DIAMETER, in. __|
0.006
=
o
G 0.005
<C
o
[
1]
= o
= 0.004 @ = 0.00179 +0.024/{V /W)
9] \
>
)
=]
0.003 <O
\—-u-o-—-——o———-o—-a-
o
0.002
% GLYCERINE IN WATER
0.004 4 0
o 375
o 50
)
0 10 20 30 40 50 60

l/u/Vr, RATIO OF AXIAL TO TERMINAL VELOCITY

Fig. 7.7. Horizontal flow volume fraction correlation.
Volume fraction occupied by bubbles with a gas-to-liquid
volumetric flow ratio of 0.003 in a 2-in, horizontal pipe.

For example, above a Reynolds modulus for which
gravitational forces on a bubble become negligible
compared with turbulent inertial forces, the mass
transfer coefficients for vertical and horizontal flow
become identical. The Reynolds modulus at which this
occurs should not be affected by the amount of mass
transfer occurring downstream of the test section.
These data also suggest that the vertical flow coef-
ficients may be approaching a constant value as the
Reynolds modulus is decreased. This is the value that
would be obtained for bubbles rising through a quies-
cent liquid and should also be independent of mass
transfer occurring in the separator.

The corrected mass transfer coefficients in horizontal
flow (after that portion of the total mass transfer which
occurs downstream from the test section is subtracted)
are shown on the lower portion of Fig. 7.8. For
horizontal flow the quantity to be subtracted was
directly measured by relocating the bubble generator
immediately upstream of the separator and repeating
the test under the original conditions with the bubbles
feeding directly into the separator region (i.e., without
passing through the test section). A correction for
vertical flow can only be estimated, however, because
the bubblie generator had to be reworked in the course
of the original vertical flow tests. The subsequent repair
altered its characteristics to the extent that reproduc-
tion of the vertical flow runs with the bubble generator
repositioned downstream from the test section was not
possible.

7.3.2 Theory

A theoretical description of the problem is possible if
each bubble is assumed to move at the local liquid
velocity (i.e., without relative flow). A mass balance
within an equivalent volume influenced by a bubble
having the mean diameter results in the nondimensional
equation given below:

00, _ 1+ u,/0 (aze*

2 00,
el + = = 4
aX* NSCNRe ) ( )

or,2  r, or,

with boundary conditions
1- 0*(0:r*)= 1 E]
2. 0,.(X«>0,d[2D)=0,

a0 d .
3. —a—"-‘-:Oatr*= 55 (* 1/®)'/3 .

.
71

ORNL-DWG T7i1-672{

% ! —t
4 | MASS TRANSFER IN TEST
ad SEGTION AND SEPARATOR
[ [a]ie;
Al A
> e
A P
&
7
T n
£ ._/’ N 7 MASS TRANSFER IN TEST SECTION 7
< - / /a ! (SEPARATOR MASS TRANSFER
E o p==—=1= 7] kuunten ol s ] . P ¢ - ,.,/ X SUBTRAGCTED)
S ESTIMATED VERTICAL® '@ fA f' /|
b FLOW BEHAVIOR; \ ?3; ;
§ ! \""'IZ? 5/3
o ye) | T I 1]
L A Qg/@_=0.003
) [ dc g L_ . —
z % GLYGERINE = 25 T
o {SGHMIDT MODULUS ~1228) ]
L, 0.5 CONDUIT DIAM =2.0in. -
(%]
g GRAVITATIONAL ORIENTATION MEAN BUBBLE DIAMETER
< VERTICAL HORIZONTAL (in.)
. o 0.045
n a] 0.02
A A 0.035
0.2 -_——
0.4 ] { ‘ ‘ |
103 2 5 104 2 5 10° 2 5 108

REYNOLDS MODULUS

Fig. 7.8. Measured mass transfer coefficient as a function of flow Reynolds modulus, bubble mean diameter, and gravitational
orientation. Comparison of data before and after adjustments for the amount of mass transfer occurring in the separator.

Using Eq. (4) as the basic analytical description, a
theory of mass transfer from a turbulent liquid to
circulating bubbles involves establishing a relationship
between the relative eddy transport coefficient (u,/10),
the fluid properties, and the turbulence characteristics
of the field; that is, a function is desired such that

U8 =frs, d, Ny, Nge) - ‘

Eddy diffusion coefficients for turbulent flow in
pipes have been measured. The correlation of Groen-
hof'? for the center-line region of the pipe is assumed

to apply:

u,=40Xx 10 J7, /oD,

12. H. C. Groenhof, “Eddy Diffusion in the Central Region
of Turbulent Flows in Pipes and between Parallel Plates,”
Chem. Eng. Sci. 25,1009 (1970).

which can be converted to

1,/ =0.04Ng Ng,'® (5)
While there have also been measurements of the
variation in the coefficients with distance from a
surface in shear fields with stationary no-slip interfaces,
these results would not be expected to apply to a
cocirculating bubble interface which may be either rigid
(no slip) or mobile (a moving interface with internal
circulation). Therefore an analytical expression had to
be developed relating the eddy diffusivity to the
concentration gradient and the individual velocity and
frequency components of a Fourier analysis of the
turbulence. Representing the turbulence field by super-
position of sinusoidal idealized viscous “eddy cells” (a
quasi turbulence), an integral expression has been
obtained for the variation of eddy diffusivity with
distance from an interface that may be either rigid or
mobile. This variation, along with Eqgs. (4) and (5), is
being programmed for solution on a digital computer.
Nomenclature

a4 = bubble interfacial area per unit volume
C = drag ¢oefficient
.d = bubble Sauter-mean diameter
D = conduit diameter
1) = molecular diffusion coefficient
| g = gravitational acceleration
N, = number of bubbles per unit volume
Qg = volumetric flow of gas bubbles
@, = volumetric liquid flow

r = spherical radial coordinate measured from center
of a moving bubble

V, = liquid mean axial velocity
V., = bubble terminal velocity

x = axial coordinate

72

Greek letters
@ = concentration of a dissolved constituent
6o = radial average concentration of a dissolved con-

stituent at x =0
M =liquid viscosity
M, = turbulent eddy diffusivity
p = liquid density
7,, = wall shear stress -

® = volume fraction occupied by the bubbles

Dimensionless quantities

0.=0/6,

re=rfD
xXe =x/D
Ny = V,Dolu

Ng. =ulpl)
Part 3. Chemistry

W. R. Grimes

The chemical research and development activities
described below are conducted to establish the basic
chemical information required for the development of
advanced molten-salt reactor systems.

A substantial fraction of these efforts continued to be
devoted to the transport, distribution, and chemistry of
fission products in the MSRE. Similar efforts seek to
establish the nature and control of the interactions of
tritium with molten salts, metal alloys, and graphite.
Investigations of fission product behavior have been
continued with specimens removed from the MSRE fuel
circuit and by investigation of the chemistry of molyb-
denum, niobium, and ruthenium in molten fluoride
mixtures.

Investigations into the chemistry of sodium fluoro-
borate have continued and have been extended in
efforts to identify the factors which will ultimately
determine the applicability of fluoroborates as coolants.

A broad program of fundamental investigations into
the physical chemistry of molten-salt systems was
maintained; from it are derived the basic data for
reactor and chemical reprocessing design. Within the
scope of these efforts are included research in solution
thermodynamics and phase equilibria, crystal chem-
. istry, electrochemistry, spectroscopy (both Raman and
electronic absorption), transport processes, and theoret-
ical aspects of molten-salt chemistry.

Studies of the chemical aspects of separations meth-
ods were continued. The results of these studies form
the basis for evolving modifications of methods for
reprocessing MSBR fuel salts. With adoption of a
reference design for MSBR fuel reprocessing which
effects transfer of rare earths from liquid bismuth to
lithium chioride as an acceptor solvent, emphasis has
been given to development of innovative means for
separation of rubidium, cesium, and the rare earths
from the fuel solvent and to the removal of solutes
from the liquid metal extractant.

The principal emphasis of analytical chemical devel-
opment programs has been placed on methods for use
in semiautomated operational control of molten-salt
breeder reactors, for example, the development of
in-line analytical methods for the analysis of MSR fuels,
for reprocessing streams, and for gas streams. These
methods include electrochemical and spectrophoto-
metric means for determination of the concentration of
U*" and other ionic species in fuels and coolants, and
adaptation of small on-line computers to electroanalyti-
cal methods. Parallel efforts have been devoted to the
development of analytical methods related to assay and
control of the concentration of water, oxides, and
tritium in fluoroborate coolants.

8. Fission Product Behavior

8.1 DETERMINATION OF TRITIUM
AND HYDROGEN CONCENTRATIONS IN MSRE
PUMP BOWL GAS

S.S. Kirslis  F. F. Blankenship

The dimensions of the problems caused by tritium
diffusion through metal walls of a molten-salt reactor
have been estimated using a particular mathematical
model.! The validity of the model could be confirmed
by measurements of the tritium concentration in .the

73

MSRE pump bowl gas. Therefore several different types
of sample were taken while the MSRE was operating,
and apparatus was constructed and tested for analyzing
those samples. Since hydrogen partial pressure strongly
influences the solubility and permeation rate of tritium
from a gas mixture into a metal, it was also necessary to

1. R. B. Briggs and R, B. Korsmeyer, Distribution of Tritium
in a 1000-MW(e)j MSBR, ORNL-CF-70-3-3 (Mar. 18, 1970)
(internal memorandum).
determine the hydrogen concentration in the pump -

bowl gas.

Several types of tritium sampling devices were used in
the MSRE pump bowl. Advantage was taken of the
unique properties of hydrogen and its isotopes in that
they are highly soluble in and permeate rapidly through
metals at red heat. The same properties would make it
difficult to obtain a representative gas sample using an
evacuated metal bulb with a resealable freeze valve.

The simplest sampling device was a solid nickel
cylinder, % in. in diameter and 6 in. long. The
solubilities of hydrogen and tritium are such that at the
expected partial pressures, the solid nickel bar at
solubility equilibrium at 650°C would contain about
ten times as much tritium per cubic centimeter as the
surrounding gas. In the several samplings, exposure
times in the pump bowl gas varied between 2 and 10 hr.
The bar was pulled up into a cool section of the
sampling line about 2 ft above the pump bowl and
allowed to cool in essentially the same atmosphere.
Each saturated nickel bar was taken to a hot cell,
decontaminated in dilute nitric acid and 1 M H,0,—1
M oxalic acid, and stored in a hood in a Dewar flask
under liquid nitrogen.

A variation of this sampling method was a sealed
thin-walled nickel capsule filled with nickel powder
which had been deoxidized with hydrogen and gvacu-
ated at red heat. The thought was that hydrogen and
tritium should saturate the capsule wall and powder
particles much more rapidly than in the case of the
solid bar.

A third sampling device was a thin-walled nickel
capsule containing degassed CuO “wire.” At 650°C,

ORNL—DWG 71-7244

TO GAS
MANIFOLD

BOURDON
GAGE
THERMOQ -
QUARTZ COUPLE
TUBE GAGE

.| U POWDER
UTs

TUBE /
FURNACE

L4
IONIZATION /
CHAMBER

Fig. 8.1. Calibration apparatus.

gaseous H, or HT permeating the capsule wall would
react with the CuO to produce H, O and copper metal.
The H,O or HTO could not diffuse back through the
capsule wall and would thus be trapped in the capsule.
Exposure times in the MSRE pump bowl gas phase
varied between 2 and 10 hr for this type of capsule.

A single capsule was prepared which was identical to
the one just described except that one hemispherical
end of the capsule was made of thin palladium sheet
welded to the cylindrical nickel capsule body. Since
palladium is much more permeable to hydrogen than
nickel, a faster collection of H, O and HTO should have
resulted.

All of the different capsules were decontaminated and
stored under liquid nitrogen in a Dewar flask, as
described for the nickel bar samples.

8.1.1 Calibration Apparatus

From the total amount of tritium and hydrogen
found in the sample capsules or bars, it is theoretically
possible to calculate how much tritium and hydrogen
were in the MSRE pump bowl atmosphere, using
capsule dimensions and literature values for the solu-
bility of hydrogen in nickel and for the permeability of
hydrogen through nickel at 650°C. However, in the case
of the nickel bars, there is doubt whether the interior of
each bar was fully saturated and doubt as to the
fraction of dissolved gas lost during cooling. The latter
doubt is of still greater concern for the nickel powder
capsule. A more basic concern is whether Sieverts’s law
(the solubility of hydrogen in a metal is proportional to
the square root of the hydrogen pressure) is valid at the
low pressures we are dealing with. It has been suggested

‘in the literature that the law fails at low pressure, since

for many gases the extrapolations of plots of hydrogen
solubility vs square root of pressure do not go through
the origin. It is desirable to check this point since it has
an important bearing on the effect of added hydrogen
pressure on tritium permeation through metals. Finally,
while no faults with the CuO method of sampling are
obvious a priori, it is desirable to check new analytical
tools and methods under conditions closely simulating
those of actual use.

For those reasons, a calibration apparatus (Fig. 8.1)
was built in which the sampling devices could be tested
in tritium and hydrogen concentrations similar to those
expected in the MSRE pump bowl gas.

The vertical quartz tube in which the sampling device
was suspended was part of a thermal-convection loop
constructed mainly of Pyrex tubing. Side loops through
which the convective flow could be directed contained
a tritium gas counter and a uranium powder—uranium

hydride—uranium tritide trap. The counter was an
ionization chamber operated at 500 V with the output
current read with a vibrating reed electrometer. The
uranium hydride—tritide trap could be heated to supply
a pure dry mixture of 98% H, and 2% HT to the
convecting gas. By cooling the trap the H, and HT
could be resorbed on the powdered uranium. Provision
was made for sampling the gas in the loop into a
removable glass bulb. A side arm from the loop
contained a Bourdon gage (vacuum to 15 psig) and a
thermocouple gage. All glass valves in the loop system
were of the greaseless O-ring variety in order to
minimize exchange of tritium with the hydrogen in
stopcock grease. Finally, the loop was connected to a
gas manifold by means of which the loop could be
evacuated or pressurized with dry helium, dry argon, or
dry hydrogen. The complete calibration apparatus was
installed in a hood in case of accidental tritium release.
A tritium “sniffer”” was installed very close to the loop
to warn of any tritium leakage. '

To date, the calibration procedure has been carried
out only a few times, using the solid nickel bar type of
sampling device. The procedure involved first evacuat-
ing the nickel bar suspended in the heated quartz tube
(650°C). Then dry argon was admitted to. nearly
atmospheric pressure and allowed to circulate through
the uranium tritide trap and the ionization chamber.
The temperature of the trap was adjusted to give the
desired H, and HT concentrations and the gas allowed
to circulate past the nickel bar for several hours. The
counter indicated the constancy of the HT pressure. A
sample of the circulating gas was taken in the glass bulb,
and the quartz tube and nickel bar were cooled rapidly
to room temperature. The uranium tritide trap was then
cooled to room temperature to resorb the circulating
tritium, and the loop was evacuated and flushed with
dry argon to remove the tritium. The ground joint at
the top of the quartz tube was then opened (with dry
argon flow blanketing the loop) and the nickel bar
removed.

The nickel bar was then placed in the quartz tube of
the extraction apparatus in which the tritium was
extracted from the nickel bar by heating it at red heat
for several hours in a 99.9% He—0.1% H, flow. The gas
then passed through a CuO trap at 500°C to convert HT
to HTO, and the HTO was finally caught in a series of
distilled-water bubblers. The water was analyzed for
tritium with great sensitivity by scintillation counting.

The glass bulb sample was analyzed for hydrogen by
mass spectrometry and for tritium using the same
extraction apparatus but omitting” the quartz heating
tube.

75

The original rather simple design of the calibration
apparatus had to be modified and refined repeatedly in
order to perform as desired. To achieve the required
leak rate and degassing rate (mainly H, O which would
exchange with and dilute HT), the apparatus had to be
leak-tight to the limit of a helium mass spectrometer
leak detector, and ordinary glass stopcocks had to be
replaced with greaseless glass valves. Before a gas was
admitted to the loop, it had to contain less than 10
ppm of water as measured by a Meeco electrolytic
water meter. It was necessary to pass the convective
flow through the uranium tritide trap and through the
ionization chamber to achieve controllable and steady
tritium concentrations.

The ionization chamber had an unfortunate memory
effect. After once introducing a high tritium pressure
(~100 p) for a special test, it was not possible to get rid
of a high background due to tritium dissolved in the
thick stainless steel wall of the ionization chamber.
Weeks of evacuation and flushing while warming the
chamber did not succeed in reducing the background to
a satisfactory extent. A Pyrex ionization chamber was
designed and built, with silvered interior surfaces for
electrodes. The silver film is so thin that degassing of
tritium should be rapid. The new ionization chamber
has not yet been checked with tritium gas. '

The most successful calibration tests with a nickel bar
sampling device achieved only a qualitative agreement
with the glass bulb samples. It is thought that the
failure to obtain exact agreement was due either to
changing hydrogen and tritium concentrations during
the exposures or to variations from Sieverts’s law for
hydrogen solubility at low pressures. This will be
checked in future tests by taking glass bulb samples at
the beginning and end of each test and by working at
different tritium and hydrogen partial pressures.

8.1.2 Analysis for Hydrogen

Analysis of ‘the sampling devices for tritium is much
simpler than analysis for hydrogen, since the nickel
contains only about 0.02 g of H, per gram and it is very
difficult to avoid picking up small amounts of water
from the surfaces of gas handling equipment. For this
reason an analytical method was devised which involves
a minimum of gas handling. The standard vacuum
fusion method is not sensitive enough. The nickel bar
and nickel powder samples will be heated in a quartz
tube containing hot CuQ to convert the extracted H, to
H,0. At one end of the quartz tube there will be a
simple dew-point measuring device by means of which
the H, O concentration in the tube can be measured.
The main difficulty with this method is expected to
arise in desorbing water sufficiently from the system
and sample surfaces in the early stages of heating under
evacuation, before the valve to the pump is closed and
the sample heated further to liberate its hydrogen.

In the case of the CuO sample capsules, the hydrogen
is already in the form of water, and it is only necessary
to puncture the capsule with a special puncturing valve
and let the inside gas into a dew-point meter.

The dew-point method does not destroy the sample,
so that the analyzed sample can be flushed through
water bubblers for tritium analysis. For the nickel bar
and nickel powder samples, a method for analyzing the
extracted gas for H, by mass spectrometer has been
devised in case the tramp water problem turns out to be
too difficult.

8.1.3 Tritium Diffusion Studies

An alternative reason for constructing the calibration
apparatus with its convective circulation, its tritium
counter, and its tritium-hydrogen supply is that it is an
ideal system with which to study the effect of hydrogen
pressure on tritium diffusion through Hastelloy N, using
realistic concentrations of tritium and hydrogen. The
system will also be useful in testing the effects of
various metal surface treatments on the diffusion
process.

8.2 EXAMINATION OF DEPOSITS FROM THE
MIST SHIELD IN THE MSRE FUEL PUMP BOWL

E.L.Compere E.G. Bohlmann

In January 1971 the sampler cage and mist shield
were excised from the MSRE fuel pump bowl by using
a rotated cutting wheel to trepan the pump bowl top
(see Sect. 1.3). The sample transfer tube was cut off
just above the latch stop plug penetrating the pump

bowl top; the adjacent ~3-ft segment of tube was

inadvertently dropped to-the bottom of the reactor ceil
and could not be recovered. The final ligament attaching
the mist shield spiral to the pump bowl top was severed
with a chisel. The assembly, of mist shield spiral sur-
rounding the sampler cage attached to the latch stop
plug, was transported to the HRLEL for cutup and
examination.

Removal of the assembly disclosed the copper bodies
of two sample capsules that had been dropped in 1967
and 1968 lying on the bottom of the pump bowl. Also
on the bottom of the bowl, in and around the sampler
area, was a considerable amount of fairly coarse
granular porous black particles (largely black flakes ~2

76

to S mm wide-and up to 1 mm thick). Contact of the
heated quartz light source in the pump bowl with this
material resulted in smoke evolution, and apparently
some softening and smoothing of the surface of the
accurmnulation. Periscopic examination of the interior of
the pump bowl is described in Sect. 1.3.

One of the sample capsule bodies was recovered and
examined as described in Sect. 1.3. A few grams of the
loose particles were also recovered and transferred in a
jar to the hot cells; a week later the jar was darkened
enough to prevent seeing the particles through the glass.
An additional quantity of this material was placed
loosely in the carrier can. Samples were submitted for
analysis for carbon and for spectrographic and radio-
chemical analyses. The results are discussed below.

The sampler assembly as removed from the carrier can
is shown in Fig. 8.2. All external surfaces were covered
with a dark gray to black film, apparently 0.1 mm or
more in thickness. Where the metal of the mist shield
spiral at the top had been distorted by the chisel action,
black eggshell-like film had scaled off, and the bright
metal below it appeared unattacked. Where the metal
had not been deformed, the film did not flake off.
Scraping indicated a dense, fairly hard adherent black-
ish deposit.

On the cage ring a soft deposit was noted, and some
was scraped off; the underlying metal appeared un-
attacked. The heat of sun lamps used for in-cell
photography caused a smoke to appear from deposits
on bottom surfaces of the ring and shield. This could
have been material, picked up during handling, similar
to that seen on the bottom of the pump bowl.

At this time samples were scraped from top, middle,
and bottom regions of the exterior of the mist shield,
from inside bottom, and from the ring. The mist shield
spiral was then cut loose from the pump bowl segment,
and cuts were made to lay it open using a cutoff wheel.
A view of the two parts is shown in Fig. 8.3.

In contrast to the outside, where the changes between
gas (upper half) and liquid (lower) regions, though
evident, were not pronounced, on the inside the lower
and upper regions differed markedly in the appearance
of the deposits.

In the upper region the deposits were rather similar to
those outside, though perhaps more irregular. The
region of overlap appeared to have the heaviest deposit
in the gas region, a dark film up to 1 mm thick, thickest
at the top. The tendency of aerosols to deposit on
cooler surfaces (thermophoresis) is called to mind. In
the liquid region the deposits were considerably thicker
and more irregular than elsewhere, as if formed from
larger agglomerates,
77

R 53633

Fig. 8.2. Mist shield containing sampler cage from MSRE pump bowl.

In the area of overlap in the liquid region this kind of
deposit was not observed, the deposit resembling that
on the outside. If we recall that flow into.the mist
shield was nominally upward and then outward through
the spiral, the surfaces within the mist shield are
evidently subject to smaller liquid shear forces than
those outside or in the overlap, and the liquid was
surely more quiescent there than elsewhere. The condi-
tions permit the accumulation and deposition of ag-
glomerates.

The sample cage deposits also were more even in the
upper part, becoming thickest at and on the latch stop.
The deposit on the latch stop was black and hard,
between 1 and 2 mm thick. Deposits on the cage rods
below the surface (see Figs. 8.4 and 8.5) were quite
irregular and lumpy and in general had a brown-tan
(copper or rust) color over darker material; some
whitish material was also seen. Four of the rods were
scraped to recover samples of the deposited material.

After a gamma radiation survey of the cage at this time,
the unscraped cage rod was cut out for metallographic
examination; another rod was also cut out for more
thorough scraping, segmenting, and possible leaching of
the surfaces.

The gamma radiation survey was conducted by
lowering the cage in ‘-in. or smaller steps past a
0.020- by 1.0-in. horizontal collimating slit in 4 in. of
lead. Both total radiation and gamma spectra were
obtained using an Nal scintillation crystal. The radia-
tion levels were greatest in the latch stop region at the
top of the cage and next in magnitude at the bottom
ring. Levels along the rods were irregular but were
higher in the liquid region than in the gas area even
though considerable material had been scraped from
four of the five rods in that region. In all regions the
spectrum was predominantly that of 367-day '°®Ru
and 2.7-year '258b, and no striking differences in the
spectral shapes were noted.
. R 54220

Fig. 8.3. Interior of mist shield. Right part of right segment overlapped left part of segment on left

79

R 54219

Fig. 8.4. Sample cage and mist shield.

R 54189

Fig. 8.5. Deposits on sampler cage. Ring already scraped.

80

Analyses of samples recovered from various regions
inside and outside the mist shield and sampler cage are
shown in Table 8.1. The samples generally weighed
between 0.1 and 0.4 g. The radiation level of the
samples was measured using an in-cell G-M probe at
about 1 in. distance, and at the same distance with the
sample surrounded by a %-in. copper shield (to absorb
the 3.5-MeV beta of the 30-sec '°®™Rh daughter of
106Ru). Activities measured in this way ranged from 4
R/hr (2 R/hr shielded) to 180 R/hr (80 R/hr shielded).
the latter being on a 0.4-g sample of the deposit on the
latch stop at the top of the sample cage.

Spectrographic and chemical analyses are available on
three samples, (1) the black lumpy material picked up
from the pump bowl bottom, (2) the deposit on the
latch stop at the top of the sample cage, and (3)
material scraped from the inside of the mist shield in
the liquid region. The material recovered from the
pump bowl bottom contained 7% carbon, 31% Hastel-
loy N metals, 3.4% Be (18% BeF,), and 6% Li (22%
LiF). Quite possibly this included some cutting debris.
The carbon doubtless was a tar or soot resulting from
thermal and radiolytic decomposition of lubricating oil
leaking into the pump bowl. It is believed that this
material was jarred loose from upper parts of the pump
bowl or the sample transfer tube during the chise! work
to detach the mist shield.

The hard deposit on the latch stop contained 28%
carbon, 2.0% Be (11% BeF,), 2.8% Li (10% LiF), and
12% metals in approximate Hastelloy N proportions,
again possibly to some extent cutting debris.

The sample taken from the inner liquid region of the
mist shield contained 2.5% Be (13% BeF,), 3.0% Li
(11% LiF), and 18% metals (with somewhat more Cr
and Fe than Hastelloy N); a carbon analysis was not
obtained.

In each case, about 0.5 to 1% Zr (~1 to 2% ZrF, ) was
found, a level lower than fuel salt in proportion to the
lithium and beryllium. Uranium analyses have not yet
been received, so we cannot clearly say whether the salt
is fuel salt or flush salt. Since fission product data
suggest the deposits built up over appreciable periods,
we presume fuel salt.

In all cases the dominant Hastelloy N constituent,
nickel, was the major metallic ingredient of the deposit.
Only in the deposit from the mist shield inside the
liquid region did the proportions of Ni, Mo, Cr, and Fe
depart appreciably from the metal proper. In this
deposit a relative excess of Cr and Fe was found, which
would not be attributable to incidental metal debris
from cutting operations. It is also possible that the
Table 8.1. Chemical and spectrographic énalysis of deposits from mist shield in the MSRE pump bowl

- Radiation level

Sample (R/hr @1 in.) Weight  Percent . T_l ;. Percent Li Percent Be Percent Zr* Percent Ni’ Percent Mo? Percent Cr? Percent Fe? Percent Mn?
Location , . (mg) C (dismin" g™ ")
(shlelded)
Pump bowl 10(5) 402 3.1 E10 6.00 3.42 0.5-1.0 20-30 2-4 1-2 0.5-1.0 0.5-1.0
bottom 25(12) 108 7.1
Latch stop 130(60) 291 1.85 El11 2.75 2.01 0.5-1.0 5-10 2-4 0.5-1.0 0.5-1.0 <0.5
180(80 365 28 _
Inside, 40(17) 179 4.7 E10 3.03 2,52 0.5-1.0 - 5-10 24 3-5 2—-4 <0.5
liquid
region
MSRE fuel - 11.1 6.7 11.1
(nominal)
Hastelloy N 69 16 7 5 ~1
(nominal)

aSemiquantitative spectrographic determination.

I8
various Hastelloy N elements were all subject to mass
transport by salt during operation, and little of that
found resulted from cut-up operation.

8.2.1 Tritium

Tritium determinations have been received on samples
of material from the pump bowl bottom (14 mCi/g),
latch seat (83 mCi/g), and inside liquid region of the
mist shield (21 mCi/g). Tritium may well have ex-
changed with the hydrogen of the oil entering the pump
bowl and thereby been retained with any tar deposits.
The loose material recovered from the pump bowl bot-
tom contained 8 X 107° atom of T per atom of C, and
the latch stop deposit contained 1.2 X 107 atom of T
per atom of C. The agreement of the two values is useful.

For comparison, assume oil (—CH,—) entered the
pump bowl at w g/day, equivalent to 4.3 X 10*%w
atoms of C per day and 8.6 X 10*?w atoms of H per
day. Also at full power, 40 Ci/day of tritium was
developed in the MSRE. If a fraction p does not pass
through walls but remains available for exchange, we
have 8.3 X 102%p atoms/day of T thus available. If a
fraction X in fact does ex'change,. then there are 0.96 X
1072 (p/w)X atoms of tritium which are exchanged per
atom of available hydrogen. If n hydrogen atoms are
attached to each atom of carbon, then

t T Xn
oM —096x 1072222
atoms C w

Because the observed value was ~107%, the ratio of
observed to calculated indicates that

pnX _

~

1072,
w

This does not seem unreasonable, since p is less than 1,
w somewhat greater than 1, and » is probably 1 or less.
We conclude thereby that X is of the order of at least a
few percent.

Probably more of this tar was carried out of the pump
bowl by off-gas than remained. In the two determina-
tions reported above we found 0.2 and 0.3 Ci of tritium
per gram of carbon. If a (low) few grams of lubricating
oil passed through the pump bowl and was associated
similarly with tritium by exchange reactions, then up to
a curie or so of tritium per day could enter the off-gas
in this way. Presumably much would deposit or
condense and not emerge from the off-gas system.
Tritium as HT or CH3T, etc., would of course pass
through. |

82

In an MSBR, utilization of exchange processes to
affect tritium behavior would depend on additional
factors which will not be considered here.

We now come to consideration of fission product
isotope data. These data are shown in Table 8.2 for
deposits scraped from a number of regions. The activity
per gram of sample is shown as a fraction of MSRE
inventory activity per gram of MSRE fuel salt, to
eliminate the effects of yield and power history;
materials concentrated in the same proportion should
have similar values.

We first note that the major part of these deposits
does not appear to be fuel salt, as evidenced by low
values of ®°Zr and '**Ce. The values 0.13 and 0.11 for
144 Ce average 12%, and this is to be compared with the
combined 24% for LiF + BeF, determined spectro-
graphically as noted above. These would agree well if
fuel salt had been occluded steadily as 24% of a growing
deposit throughout the operating history.

For '?7Cs we note that samples below liquid level
inside generally are below salt inventory and could be
occluded fuel salt as considered above. For samples
above the liquid level inside, or any external sample,
values are two to nine times inventory for fuel salt.
Enrichment from the gas phase is indicated. Houtzeel?
has noted that off-gas appears to be returned to the
main loop during draining, as gas from the drain tanks is
displaced into a downstream region of the off-gas
system. However, our c'ieposit must have originated
from something more than the gas residual in the pump
bowl or off-gas lines at shutdown. An estimate sub-
stantiating this is as follows.

With full stripping, 3.3 X 10'7 atoms '37chain per
minute enter the pump bowl gas. About half actually
goes to off-gas, and most of the rest is reabsorbed into
salt. If we, however, assume a fraction f is deposited
evenly on the boundaries (gas boundary area ~16,000
cm?), the deposition rate would be ~2 X 10'3f atoms
'37chain per square centimeter per minute. Now if in
our samples the activity is / relative to inventory salt
(1.4 X 10'7 atoms '37Cs per gram) and density is ~2,
then the time ¢ in minutes required to deposit a
thickness of X centimeters would be

.o L4X10V X IX2X X
2X 10'3f

A I
= 14X 1047X,min.

In obtaining our samples we generally scraped at least
0.1 g from perhaps 5 cm?, indicating a thickness of at

2. A. Houtzeel, private communication.
Table 8.2. Gamma spectrographic (Ge-diode) analysis of deposits from mist shield in the MSRE pump bowl

99TC 95Nb 103Ru 106Ru lZSSb 127mTe 137CS 9521.0 144Cea
Half-life 2.1 X 10° years 35 days 39.6 days 367 days 2.7 years 105 days 30 years 65 days 284 days
(after °5Zr)
Inventory, dis min ! g~! 24 uglg 8.3 E10 3.3 E10 3.3 E9 3.7E8 2.0E9 6.2 E9 9.9 E10 59E10
Sample activity per gram expressed
as fraction of MSRE inventory
activity/grams fuel salt?
Pump bowl bottom, 0.23 36 52 20 17 2.1 0.13 + 0.03 0.10
loose particles
~ Latch stop 265 364 1000 5.7 69 31 0 0
Top .
Qutside 122 167 328 104 98 9.5 0 0
Inside 42+ 14 237 £ 163 273 1000 69 4.3 0 0
Middle outside ’ 277 531 646 563 54 8.0 0
Below liquid surface -
Inside No. 1 354 164 224 271 143 0.6 0 0.13 +0.02
Inside No. 2 463 148 292 310 72 98 1.1 0 0 '
Cage rod 305 221 692 198 187 0.2 0 0
Bottom . '
Outside (0, <60) (0, <60) 1210 167 232 - 3.2 0 0
Inside 16 £ 9 159 189 51 27 0.32 0 0.11

?Background values (limit of detection) were as follows: °°Zr, 2-9 E10; '*%Ce, 1-2 E10; ' 3% Cs, 2-9 E8; ! '®Ag, 1-3 E9; ! 5*Eu, 1 E8—2 E9.

bUncertainty stated (as * value) only when anappreciable fraction (>10%) of observed.

€8
least ~0.01 c¢m, and [/ values were ~4, whence ¢ =
600/f.

Thus, even if all (f ~1) the !'*7chain entering the
pump bowl entered our deposits, 600 min flow would
be required to develop their ! *?Cs content — too much
for the 7-min holdup of the pump bowl, or even the
rest of the off-gas system excluding the charcoal beds.

It appears more likely that ! *7Cs atoms, from ' 37 Xe
atoms decaying in the pump bowl, were steadily
incorporated to a slight extent in a slowly growing
deposit.

The noble metal fission products, 35-day ?°Nb,
39.6-day '°3Ru, 367-day '°°Ru, 2.7-year '258Sb, and
105-day ‘2?7 Te, were strongly present in essentially
all samples. In all cases 35-day °3Nb was present in
quantities appreciably more than could have resulted
from decay of ®° Zr in the sample.

Antimony-125 appears to be strongly deposited in all
regions, possibly more strongly in the upper (gas) region
deposits. Clearly '*5Sb must be considered a noble
metal fission product. '?7™Te was also found, in
strong concentration, frequently in similar proportion
to the 23 Sb of the sample. The precursor of 27" Te
is 3.9-day '?7Sb. It may be that earlier observations
- about fission product tellurium are in fact observations
of precursor antimony isotope behavior, with tellurium
remaining relatively fixed.'

The ruthenium isotopes were present in quantities
comparable to those of *>Nb, '25Sb, and '?7" Te, If
the two ruthenium isotopes had been incorporated in
the deposit soon after formation in the salt, then they
should be found in the same proportion to inventory.
But if a delay or holdup occurred, then the shorter-lived
193 Ru would be relatively richer in the holdup phase®:
the activity ratio '°3Ru/!'°®Ru would exceed the
inventory ratio, and material deposited after an appreci-
able holdup would have an activity ratio ! °3 Ru/!°6Ru
which would be less than the inventory value. Examina-
tion of Table 8.2 shows that in all samples, relatively
less '°?Ru was present, which indicates the deposits
were accumulated after a holdup period. This appears
to be equally true for regions above and below the
liquid surface. Thus we conclude that the deposits do
not anywhere represent residues of the material held up
at the time of shutdown, but rather were deposited over
an extended period on the various surfaces from a
common holdup source. Specifically this appears true
for the lumpy deposits on mist shield interior and cage
rods below the liquid surface.

3. E. L. Compere and E. G. Bohlmann, MSR Program
Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, pp. 60—66.

84

Data for 2.1 X 10%-year ?°Tc are available for one
sample taken from the inner mist shield surface below
the liquid level. The value 1.11 X 10* ug/g, vs inventory
24 ug/g, shows enhanced concentration ratio similar to
our other noble metal isotopes and clearly substantiates
the view that this element is to be regarded as a noble
metal fission product. The consistency of the ratios to
inventory suggests that the noble metals represent ~5%
of the deposits.

The quantity of noble metal fission products held up
in this pump bowl film may not be negligible. If we
take a median value of ~300 times inventory per gram
for the deposited material, take pump bowl area in the
gas region as 10,000 cm?® (minimum), and assume
deposits 0.1 mm thick (~0.02 g/cm?; higher values
were noted), the deposit thus would have the equivalent
of the content of >60 kg of inventory salt. There was
~4300 kg of fuel salt, so on this basis deposits
containing about ~1.4% or more of the noble metals
were in the gas space. At least a similar amount is
estimated to be on walls, etc., below liquid level; and no
account was taken for internal structure surfaces (shed
roof, deflector plates, etc., or overflow pipe and tank).

Since pump bowl surfaces appear to have more (~10
times) noble metal fission products deposited on them
per unit area than the surfaces of the heat exchanger,
graphite, piping, surveillance specimens, etc., we believe
that some peculiarities of the pump bowl environment
must have led to the enhanced deposition there.

We first note that the pump bowl was the site of
leakage and cracking of a few grams of lubricating oil
each day. Purge gas flow also entered here, and hydro-
dynamic conditions were different from the main loop.

The pump.bowl had a relatively high gas-liquid
surface with higher agitation relative to such surface
than was the case for gas retained as bubbles in the
main Joop. The liquid shear against walls was rather less,
and deposition appeared thickest where the system was
quietest (cage rods). The same material appears to have
deposited in both gas and liquid regions, suggesting a
common source. Such a source would appear to be the
gas-liquid interfaces: bubbles in the liquid phase and
droplets in the gas phase. It is known that surface-
seeking species tend to be concentrated on droplet
surfaces.

The fact that gas and liquid samples obtained in
capsules during operation had '®3Ru/'°®Ru activity
ratios higher than inventory® and deposits discussed
here had '°?Ru/'°®Ru activity ratios below inventory
suggests that the activity in the capsule samples was
from a held-up phase that in time was deposited on the
surfaces which we examined here.
The tendency to agglomerate and deposit in the

85

less-agitated regions suggests that the overflow tank .

may have been a site of heavier deposition. The pump
bowl liquid which entered the overflow pipe doubtless
was associated with a high proportion of surface, due to
rising bubbles; this would likely serve to enhance
transport to the overflow tank.

The binder material for the deposits has not been
established. Possibilities include tar material and per-
haps structural or noble metal colloids. Unlikely,
though not entirely excludable, contributors are oxides
formed by moisture or oxygen introduced with purge
gases or in maintenance operations. The fact that the
mist shield and cage were wetted by salt nevertheless
suggests such a possibility.

8.3 SYNTHESIS OF NIOBIUM FLUORIDES
C.F. Weaver J.S.Gill

Various methods for synthesis of NbF; and NbF,
have been reported previously.*—8 In attempts to
synthesize the fluorides of lower oxidation numbers, we
have utilized the disproportionation of NbF,. Experi-
ments conducted by heating NbF, in closed evacuated
quartz tubes with one end at room temperature have
suggested that the NbF, disproportionates in the
temperature range 250 to 350°C under conditions
which maintain an NbF; pressure of a few hundred
microns. The disproportionation of NbF, at 350°C was
found to produce a compound identified by x-ray
diffraction as NbF3; (ASTM-9-168). This material is
extremely sensitive to air, reacting to form a vapor and
NbO, F, as identified by x-ray diffraction. The NbO, F
decomposed above 300°C in a vacuum to form NbOF,
vapor, identified with a mass spectrometer, and a
residue of niobium oxides. In this regard the dispropor-
tionation of the niobium compounds is similar to that
of the molybdenum fluorides.

4. L. M. Toth, H. A. Friedman, and C. F. Weaver, MSR
Program Semiannu. Progr. Rep. Feb. 29, 1968, ORNL-4254, p.
137.

5. L. M. Toth and G. P. Smith, Reactor Chem. Div. Annu.
Progr. Rep. Dec. 31, 1967, ORNL-4229, p. 64.

6. F. P. Gortsema and R. Didchenko, Inorg. Chem. 4,
182-86 (1965).

7. C. F. Weaver et al., MSR Program Semiannu. Progr. Rep.
Feb. 28, 1971, ORNL-4548, pp. 124-29.

8. C. F. Weaver et al., MSR Program Semiannu. Progr. Rep.
Aug 31, 1970, ORNL-4622, pp. 71-74.

8.4 REACTION KINETICS OF MOLYBDENUM
AND NIOBIUM FLUORIDE
IN MOLTEN Li, BeF, SOLUTIONS

C.F.Weaver J.S.Gill

The behavior of dilute solutions of trivalent molyb-
denum fluoride in molten Li,BeF, at 500°C was
summarized in ref. 8. The molybdenum left the system
by a half-order process which was unaffected by helium
flow rate, surface area of the copper container, presence
of UF,, presence of graphite, or quantity of molyb-
denum metal produced.

The behavior of similar solutions at higher tempera-
ture-9— 12 suggests that the mechanism of removal of
Mo>* from molten Li,BeF, is different in the range
600 to 700°C than near S500°C. A pair of experiments
conducted at 600°C emphasize this behavior. Both
solutions were prepared at 500°C. On increasing the
temperature to 600°C, molybdenum vanished from the
solution. These experiments were performed with iden-
tical procedures except that in one the solution was
held at 500°C for only one day (Fig. 8.6), while in the
second the solution was held for 1000 hr (Fig. 8.7) in a
copper container with ten times the surface area present
in the first case. Clearly, either the age of the solution
or the surface area of the container had a pronounced
effect on the rate of loss of molybdenum, although no
such effects were observed at 500°C.

Previous descriptions® of the behavior of niobium
fluoride solutions have indicated the existence of an
intermediate oxidation state, probably (IV), stable in
molten Li, BeF, at 500° for periods as long as a month;
this species was tound to ve reauciole by hydrogen only
with difficulty. A value of Py /Py, 1/2 %107 atm!/2
was observed for the reaction

NbF,(d) + 2H, = 4HF + Nb°

with a concentration of 1000 to 1600 ppm of niobium
in solution.

9. C. F. Weaver, H. A. Friedman, and D. N. Hess, Reactor
Chem. Div. Annu. Progr. Rep. Dec. 31, 1967, ORNL-4229, pp.
36—37. '

10. C. _F. Weaver, H. A. Friedman, and D. N. Hess, MSR
Program Semiannu. Progr. Rep. Feb. 29, 1968, ORNL-4254,
pp. 132-34,

11. C. F. Weaver, H. A. Friedman, and D. N. Hess, MSR
Program Semiannu. Progr. Rep. Aug. 31, 1968, ORNL-4344,
pp. 15455, :

12. C. F. Weaver et al.,, Reactor Chem. Div. Annu. Progr.
Rep. Dec. 31, 1968, ORNL-4400, pp. 34-39.
ORNL-DWG 71-7245

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TIME (hr)
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ORNL-~DWG 70-13508
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¢} 200 400 600  BOO 1000 1200 1400
TIME (hr)

Fig. 8.7. Removal of Mo from molten Li, BeF,.

We have now exposed these solutions at higher
temperatures and noted instability. The niobium con-
centration in molten Li, BeF, decreased at 700°C from
1200 to 950 ppm in 1190 hr (50 days) and at 900°C
from 950 to 55 ppm in 847 hr (35 days) (Fig. 8.8).
Since there was no evidence of corrosion of the copper
container, it is assumed that the loss was a result of
disproportionation of the niobium compound. The
expected increase in nobility of the niobium with
temperature was manifested by a value of =~1072
atm1/2 for Py F/PH2 1/2 3t 900°C and 3000 ppm of

ORNL-DWG 71-7246

2
4°<.\ 700°C ! 900°C
10} —ve— e
oy
i\
\. ]
\e

CONCENTRATION {ppm)

| \

102 \

0 400 800 1200 1600 2000 2400
TIME (hr)

Fig. 8.8. Removal of Nb* from molten Li;BeF,.

niobium in solution. The stability of these solutions
contrasts sharply with the ease of decomposition of
pure NbF,, as described in Sects. 8.3 and 8.5. The
observation that NbF,, when dissolved in molten
Li;BeF,;, seems to disproportionate in the operating
temperature range for molten-salt reactors indicates
that whenever oxidized niobium exists in the reactor
fuel some NbFs must exist in the gas phase even though
the species in solution has a lower valence.

8.5 MASS SPECTROSCOPY
OF NIOBIUM FLUORIDES

C.F.Weaver J.D. Redman

Earlier work® 12-14 on the mass spectroscopy of
niobium fluorides consisted in observing the pentafluo-
ride polymers, the associated oxyfluorides, the fluorina-
tion of niobium metal, and the disproportionation of
the lower fluorides. Both the disproportionation of
NbF, and the decomposition of NbO, F mentioned in
Sect. 8.3 were followed by mass spectroscopy. The

13. C. F. Weaver et al., MSR Program Semignnu. Progr. Rep.
Feb. 28, 1969, ORNL-4396, pp. 157—-62. ‘

14. C. F. Weaver et al., MSR Program Semiannu. Progr. Rep.
Aug. 31, 1969, ORNL-4449, pp. 113-21.
87

cracking pattern- for NbOF; was somewhat different The. tentative cracking patterns previously reported
from that previously observed at 800°C, with the for Nb, F,, and NbF; have been confirmed with minor
principal change being in the relative intensity of the refinement. The cracking pattern for NbF;, including

NbOF; ion. These small changes, shown in Table 8.3, the intensities for the doubly charged ions, is shown in
are probably temperature effects. Table 8.4.
Table 8.3. Cracking pattern for NbOF4 Table 8.4. Cracking Pattern for NbF; Monomer

Relative intensity Ion Relative intensity
Ion = = "
800°C - 450-650°C NbF, 100 .
NbF4 " 8
NbOF; ™" 50 68 NbF, " 13
NbOF, " 100 100 NbF* o
NbOF* 13 8 Nb* 5
NbO™ 8 NbF,2* 0.1
No* 14 1 NbF, 2" 10
NbFy 1 5 NbF, 2+ . 5
NbF, " 13 12 NbE2* .

9. Coolant Salt Chemistry and Tritium Control

The eutectic mixture formed from NaBF, and NaF
(92-8 mole %), melting point 383°C, is the proposed
MSBR coolant salt. In projected pump-loop experi-
ments (see this report, Sect. 5), engineering experience
with this salt will be gained over a period of several
years.

Although its cost and most of its chemical and
physical properties are favorable for its adoption as the
coolant, the fluoroborate salt suffers potential disad-
vantages, as compared with the ? LiF-BeF, coolant used
in the MSRE; the chief disadvantage is that the
fluoroborate coolant exhibits a significant vapor pres-
sure at operating temperatures and would exert an even
higher vapor pressure if there was accidental mixing of
the coolant and fuel salts.

The strong possibility that the coolant salt will have
to be the sink for the tritium produced in an MSBR
places an additional criterion on this or any other
coolant to be used in an MSBR; the coolant should be
able to contain a sufficient concentration of hydrog-
enous species to exchange isotopically with the tritium
passing into the coolant circuit. At the same time the
coolant with the hydrogenous additive should be
compatible with its alloy containment system.

Success in meeting the performance criteria requires
very carefully conducted research efforts. The investi-
gations described below are motivated by the desire to
meet these criteria. At the current stage of develop-
ment, none of these efforts has advanced to the point
where it provides a clear solution to the problems
associated with the application of fluoroborates as
coolants, although some results seem encouraging.
(R. E. Thoma)

9.1 STUDIES OF HYDROGEN EVOLUTION
AND TRITIUM EXCHANGE IN
FLUOROBORATE COOLANT

S. Cantor R. M. Waller

We seek to determine how much chemically bound
hydrogen can be retained in molten fluoroborate
coolant. To the extent that this hydrogen does not
corrode metals in contact with the coolant, it is
available for isotopic exchange with any tritium enter-
ing the coolant circuit of an MSBR,

88

In these experiments evacuated nickel capsules con-
taining NaBF, -NaF (92-8 mole %) and metal coupons
are heated within a silica vessel which is connected to-
gas-handling and pressure-measuring apparatus. Hydro-
gen (or tritium) gas diffusing through the nickel capsule
can be readily determined since vessels of fused silica
are virtually impermeable to hydrogen.

'In a previously reported’ experiment, hydrogen (in
concentrations equivalent to 60 ppm H,O) in the salt
reacted with chromium coupons at 500°C and was
completely converted to H,(g). The evolved hydrogen
was initially established by gas-chromatographic analy-
ses; the loss of hydrogen from the salt sample was
confirmed by a subsequent tritium tracer analysis.

Two experimental runs (see Table 9.1) performed
with nickel as the only metal in contact with the salt
indicated that low, but encouraging, levels of chemi-
cally bound hydrogen can be retained in the salt. In
both runs the salt contained hydrogenous impurities
incorporated during the recrystallization of NaBF,
from aqueous media. Although the postexperimental
examination of the salt after both runs confirmed
isotopic exchange, it should be noted that the salt
samples could not be stirred in these experiments and
perhaps had not achieved equilibrium. Further tests will
be necessary to verify isotopic exchange. Should iso-
topic exchange be confirmed, experiments in a circu-
lating salt system will be required to establish isotopic
exchange in the coolant as a practical method for
tritium control.

Five experiments were carried out with nickel cou-
pons in which NaOH and H;BO; were encapsulated
with the coolant salt. The amounts of hydrogen added
with the NaOH and H3;BO; were many times greater
than the hydrogen believed to be bound in the
fluoroborate itself. Two capsules containing 0.4 mole %
NaOH were maintained at 600°C, one for 192 hr and
the other for 387 hr. In both cases the evolved
hydrogen gas, determined chromatographically, was
equivalent to the chemically bound hydrogen loaded
into the capsules. :

1. S. Cantor and R. M. Waller, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL4622, pp. 79-80.
Table 9.1. Details of two experiments involving isotopic exchange in NaBF,-NaF

Run No.

Hydrogenous impurity

H,(g)
Conditions and procedure accounting
(ueq H/g salt)

Analytical examination of salt
sample after opening capsule

6, nickel capsule
and coupons

16, nickel capsule
and coupons

80 hr at 520°C; cool down over

102 hr at 600°C; during most

5.0 £ 0.5, all collected prior to
weekend; 10 hr at 600°C; the 22-hr evacuation.
during and between heating, gas

samples were collected and ‘

analyzed chromatographically.

Capsule and silica vessel then

evacuated for 22 hr at 600°C;

Ta(g), 0.1 Ci, introduced into

system, which was maintained at

600°C for 140 hr.

14 £ 1.4, collected prior to
of this time the silica vessel introduction of D,(g).
was evacuated by means of a

Toepler pump from which

exiting gas was analyzed for

H;(g); then D,(g) at about

22 torrs was introduced and

maintained at 600°C for 90 hr.

By counting tritium in the salt
and by prior mass-spectrographic
analysis of H,(g) and HT(g), it
was deduced that 0.5 + 0.1 ueq
of H per gram of salt had been
retained in the salt.

Infrared absorption spectrum of
15-mg pellet of salt showed an
—OD peak whose absorbance was
31/2 times greater than the —OH
peak; however, neither peak
provides a quantitative analysis
for hydrogen in the salt.

4Assuming a 10% relative standard error in the gas-chromatographic analysis.

68
Two capsules containing 0.5 mole % boric acid were
heated at 600°C, one for 435 hr and the other for 644
hr. In these time intervals the evolved hydrogen gas,
when analyzed, corresponded to about 60% of the
hydrogen introduced into the capsule. However, in-a
third capsule, with an initial concentration of 0.66 mole
% H3BO; and heated at 600°C for 300 hr, virtually all
the chemically bound hydrogen in the capsule escaped
as H,(g). The disparity between the first two H3BOj;
experiments and the third can probably be attributed to
inadvertent pumping away of H, in the first two
experiments.

Post experimental infrared spectral examination (by
John Bates) of the salts in these five latter experiments
revealed low OH concentrations, hardly different from
that found in fluoroborates, without added NaOH or
H;BO;, that had been subjected to similar treatment.
These infrared results, togethier with the quantities of
H,(g) collected during the experimental runs, suggest
that the NaOH and H;BO; did not augment the
capacity of the coolant salt to retain hydrogen. Further-
more, the higher level of protons increased the corro-
sion of nickel. After the experiments the salt and
capsule walls were examined and found to contain an
-easily observable yellowish skin, identified as NaNiF;.
With the fluoroborate salt alone, such corrosion prod-
ucts, though possibly formed, have not been detected.

9.2 REACTION OF SODIUM FLUORIDE—-SODIUM
TETRAFLUOROBORATE WITH WATER

H. W. Kohn

Attempts to refine our previous measurements of the
equilibrium constants for the reactions

NaBF, + H,0 = NaBF, OH + HF (1)

and -

NaBF;OH - NaBF,0 + HF (2)
by improving analytical techniques for the transpiration
method have so far met with failure. These have
included installation of thermal conductivity apparatus
to measure gas composition, sodium fluoride traps for
absorption of HF and BF; from the gas stream, and
azeotropic distillation from pyridine solutions of sam-
ples of the effluent gas collected from the transpiration
experiment. During the course of these expe}iments, we
have observed some HF present at all times (nickel
containers and copper lines and samplers are used).
Consequently, our approach has been to determine

90

more accurately the role of HF in stabilizing OH™ and
in producing H, O.

We have improved on some earlier experimental
results by using more refined analytical techniques to
measure the extent of reactions (1) and (2). By
collecting the off-gas containing BF; and HF in
aqueous sodium fluoride followed by CaCl, precipi-
tation:

BF;(g) + NaF(aq) » NaBF, , (3)

NaBF, + 2CaCl, + 3H,0 - H3;BO;

+2CaF,{ + NaCl + 3HCl, (4)

2HF + CaCl, — CaF,{ + 2HCI | (5)
followed by a double-end-point titration of the result-
ant HC1 and H; BO; -mannitol complex, we ascertained
that in fluoroborate to which H, O was added slowly as
vapor over a period of several hours, there was always a
small pressure of HF which would persist for two or
three days. The HF pressure measurements so obtained
are still erratic since, for example, any error in H3 BO,
determination is reflected threefold in the determina-
tion of HF, but the implication is that under our
experimental conditions (nickel containers, copper

lines, 400°C) the equilibrium content of HF in the

argon sweep gas is approximately 0.1 meq/liter. Part of
the difficulty stems from the apparent high solubility of
HF in fluoroborate at 400°C.

We believe that the HF is due to reaction (2) above.
Using the following numbers, oxide = 400 ppm ob-
tained by inventory of the H,0O, and HF = 0.1
meq/liter, we can calculate an equilibrium quotient =
[HF] [0*]/[OHT] = 10™* at 400°C (when HF is
expressed in atmospheres and the other quantities are
expressed in moles per liter). The OH is calculated from
the vapor pressure of water and the flow rate of argon
through a side stream. We assume all the water reacts,
since virtually none is found, under these conditions, by
Karl Fischer titration of the effluent gas. Using the
temperature variation of the equilibrium determined
previously,? this means that (at the same oxide and
hydroxide levels) the equilibrium pressure . of HF
required to stabilize hydroxide at 650°C would be
about 50 mm.

We have obtained preliminary values for the solubility
of HF in the NaF-NaBF, eutectic mixture with HF at

2. H. W. Kohn, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL4622, p. 81.
600 550

91

ORNL- DWG 71- 2732
TEMPERATURE (°C)

A 500 450 400
0.
| | | A
P B
/“
e
A
0.2 / SHAFFER - WATSON
E A ¢’
e / a}
_;:__-" /
Ny rY
?:‘ 0.1 2k
= A
C 0.08 e
- &
5 I
4 0.06
o
(/3]
" MEASURED BY ]
0.04 |—o O SATURATION AND STRIPPING |
® PRESSURE RISE AND TITRATION
A PRESSURE RISE CORRECTED
0.02 .
110 1.15 1.20 1,25 1.30 1.35 1.40 1.45 1.50
f000/7 (oK)

Fig. 9.1. Henry’s law solubility of HF in NaBF, —9% NaF eutectic.

0.29 atm in a nickel vessel. The HF was supplied in a
gas mixture composed of Ar, BF;, H,, and HF.
Solubility was determined by sparging the HF from the
saturated melt; the amounts of HF removed were
measured by alkalimetric titration. The removal of HF
was first order as expected, but the curve of concentra-
tion in the gas vs sparging time showed a long tail; that
is, the last traces of HF were difficult to remove. The
value of the Henry’s law constant at 400°C is 0.33
0.03 equiv liter™! atm™!. During the three measure-
ments the oxide concentration ranged from 500 to 300
ppm, determined by KBrF,, but the change in oxide
content did not seem to affect the solubility markedly
in this range.

Solubilities at other temperatures were measured in

three different ways: (1) by saturation and stripping,

(2) by saturation at 400°C followed by raising the
temperature and titrating the HF in the cover gas, and
(3) by saturation at 400°C, followed by raising the
temperature and measuring the corresponding pressure
rise and subsequently correcting it for expansion of the
cover gas and for the contribution from BF; overpres-
sure. In Fig. 9.1 the results of these three methods are
presented, along with a correlation developed by

Shaffer and Watson® which related HF solubility in
NaF-BeF, and in NaF-ZrF; with the free fluoride
content of the melt. The Shaffer-Watson formula is

given below.
K = ex A+BRC_D+ERC
P\ RT R
where

A, B, D, and E are empirical constants,

Cis thg fraction of free fluoride in the melt,

R is the gas constant,

T is the absolute temperature,

K is the Henry’s law constant.

One may extend this correlation to the NaF-BF;

system by using the vapor pressure of BF; over the
eutectic from Cantor’s data® to calculate the free

3. J, H, Shaffer and G. M. Watson, Reactor Chem, Div. Annu.
Progr. Rep. Jan, 31, 1960, ORNL-2931, p. 32.

4. S. Cantor (ed.), Physical Properties of Molten Salt Reactor
Fuel, Coolant, and Flush Salts, ORNL-TM-2316 (August 1968).
fluoride in the mixture as a function of temperature
and use these numbers in the Shaffer-Watson formula to
calculate HF solubility. The results of this calculation
are shown on the graph.

A more refined apparatus for measuring HF solubility
has been constructed and is presently being tested.

We have also used infrared spectroscopy to determine
the stability constant referred to before. This is
reported separately.

9.3 IDENTIFICATION OF CORROSION PRODUCTS
IN THE BUBBLER TUBE OF THE
FLUOROBORATE TEST LOOP

S. Cantor

For approximately 1.3 years, an Inconel pump loop
(operated by the Reactor Division under the supervision
of A. N. Smith) circulated molten NaBF,-NaF (92-8
mole %). Welded into the pump bowl was a % -in.-ID
Inconel tube through which helium and BF; entered
the circulating salt. After shutdown of the loop, the
lower part of this tube was cut off, sectioned, and split
axially. One half of the -axially split portion was
submitted to the Reactor Chemistry Division for
chemical examination of the corrosion and mass trans-
fer products deposited on the inner surfaces of the
tube.

A full description of the appearance of the bubbler
tube is given elsewhere in this report.’ Five corrosion
and mass transfer products were identified by x-ray
diffraction and electron microprobe analyses:

1. agreen salt, NajCrFg;
2. ayellow salt, NaNiF;;

3. a black powdery magnetic material with the stoichi-
ometry Ni; Fe;

. a porous, gray metallic mass, which plugged the
mouth of the tube, composed of nickel metal and
NaBF, in approximately equal parts;

5. a millimeter-thick layer, running through most of
the length of the tube, composed of nickel and
nickel oxide.

A plausible inference which may account in part for

92

the morphology of the deposit assumes that Na;CrFy, -

NaNiF;, and Ni; Fe were present in a surface scum on
the fluoroborate melt prior to migrating up the bubbler

5. A. N. Smith, “MSBR Design and Development,” this
report, Part 2.

tube; the migration possibly occurred when liquid
droplets were formed by bubbles breaking at the mouth
of the bubbler. A second transfer mechanism might
have been backflow up the tube; however, such events
were infrequent and of short duration.

The presence of the salts Na;CrFg and NaNiF; has
been noted previously as corrosion products from
molten NaBF,. Both are only slightly soluble in the
melt. They arise initially through oxidation of the
metal. In the case of Na;CrF,, the oxidation of
chromium in the loop probably arose from three
sources:

1. H,O entered as an impurity in the helium and
possibly in the fluoroborate salt charge (and also on one
occasion deliberately introduced®); the reactions may
be written:

H, O(g) + 2F “(in the salt) = 2HF(g) + O* (in the salt) ,
Cr + 3HF(g) + 3NaF(l) > Na;CrF + %, H,(g) .

2. Air trapped in lines when tanks of gas were
changed; besides the moisture thereby introduced,
0, (g) would also lead to oxidation:

Cr+ %,0,(g) + 3NaF(l)
+ BF3(g) - Na3 CI'F6 + 1/2 B2 03(1) .

3. Metallic ions already present in the melt oxidize
chromium; for instance, the fluoroborate charged into
the loop contained about 200 ppm iron; the reaction
with chromium may be written:

Cr+ % Fe™() > Cr¥*() + %, Fe .

The trivalent chromium subsequently precipitated as
Na; CrFg. '

In accounting for the formation of NaNiF;, reactions
similar to those given under 1 and 2 above seem likely:

Ni + 2HF(g) + NaF(l) > NaNiF, + H, ,
Ni+ %4, 0,(g) + NaF(1)
+ % BF;(g) > NaNiF; + 4B, 05(1) .

The reasons for the formation of Ni; Fe are not easily
ascertained. One can speculate that ionic iron and
nickel in the salt oxidize metallic chromium and

6. MSR Program Semiannu. Progr. Rep. Aug. 31, 1970,
ORNL-4622, pp. 41-44.
coprecipitate as an alloy:

%,Cr + 3Ni?*(1) + Fe?*(l) > % Cr**(1) +[Nis Fe .

Most of the Cr** would subsequently precipitate as
Na;CrF¢. The dissolved ionic nickel .and iron could
have originated by means of the oxidation by H, O(g)
and O,(g). .

The origin of the “extra” nickel in the tube (noted
above under products 4 and 5) is uncertain. The nickel
plug formed in a position where the salt was exposed to
lower temperatures; perhaps the reaction

NiF, (dissolved in melt) + H,(g) = 2HF(g) + Ni

permitted the deposition of nickel at the mouth of the
tube. ,

In summary, the corrosion products identified in the
bubbler tube originated with the introduction of
oxidants, the most important being water vapor. The
water vapor led to the formation of HF, which
corroded the Inconel. This corrosion, which implies the
dissolution of chromium, iron, and nickel ions, eventu-
ally led to the precipitation of Na;CrF¢ and NaNiF;,
some of which were carried on the surface of the
molten fluoroborate. Dissolved iron and nickel also
reacted further with metallic chromium and were
apparently precipitated out as the alloy Niz Fe. Some of
this alloy was also present on the surface of the molten
salt.

9.4 MASS SPECTROSCOPY OF FLUOROBORATE
MSR COOLANTS

C.F.Weaver J.D.Redman

With respect to tritium control in a molten-salt
reactor, F. F. Blankenship has recently suggested that
hydrogenous material might be added to and main-
tained in the proposed coolant salt, NaBF,-NaF (92-8
mole %), by incorporating 1 or 2% NaOH into the
liquid if the behavior of this system as such or its
corrosive attack on structural materials is acceptable.
Consequently, we have initiated studies of the vapors
over such materials using a time-of-flight mass spec-
trometer to detect the species effusing from a Knudsen
cell.

93

behaved in a more complex way. In addition to the
previously reported monomer and dimer, both the
trimer and tetramer of NaOH were observed. Water was

~ also evolved from material which had been dried and

The first experiments were made with the separate -

components BF;, NaBF,, and NaOH in nickel cells to
determine cracking patterns and to confirm that the
boron compounds behaved as commonly described.

The BF, yielded a cracking pattern essentially like
that in the literature, and no polymers of BFj3 were
observed. The NaBF, decomposed, yielding BF; mon-
omer as the only vapor species. However, the NaOH

outgassed at 200°C. This was attributed to the decom-

-position 2NaOH — Na,O + H,01. In addition a small

amount of H, was evolved, probably by Ni + 2NaOH —~
Na, O + NiO + H,, although the condensed products of
this reaction are not well established. By far the most
intense of the reduced species was sodium vapor,
attributed to the reaction Na, O + Ni - NiOJ + 2Na?t.
The Na vapor was detected by mass spectrometry and
the NiO precipitate by x-ray diffraction. Severe attack
of the nickel cell was also evident by direct observation.
The standard free energy change for the reaction as
written at 1000°K is +29 kcal, which implies that P, =
7 X 107? torr, well within the range of detection of the
mass spectrometer.

The possibility of adding hydrogenous material to the
fluoroborate coolant as water either intentionally or
accidentally as a steam leak suggested that the reaction
of moisture with NaBF, should also be investigated.

Two approaches were used to obtain reaction prod-
ucts of the NaBF,-H,O system. One was to add H,0
(0.1 g) to NaBF, (0.2 g) and to reflux at 300°C under
an atmosphere of helium for an hour. After refluxing,
the cell was cooled, evacuated, and studied in the
normal manner. Vapor over the temperature range of
25 to 750°C was quite similar to that over pure NaBF,,
even to the amount of water vapor evolved. No reaction
products had been formed, at least to the 300°
refluxing temperature limit; moreover, the NaBF,
displayed a very impressive, nondeliquescent character.

In the second approach to obtain reaction products
from the NaBF,-H, O system, steam was admitted over
the temperature range 25 to 750° at an approximate
leak rate of 1077 torr liter sec™!. The complexity of
the reaction was apparent. For convenience, an attempt
was made to assign fragments to possible molecular
precursors, almost arbitrarily (Table 9.2). The only

Table 9.2. Partial pressure of ‘assumed molecules
in vapor at 300° from the NaBF ,-H, O reaction

Molecules Pressure (torrs)
x 1072

H,0 1.0

BF4 2.5

HF 1.0

NaBOF, 0.008

BF3 'Hzo 0.2

NaBF, 0.003

? {mass 19) 0.01

criterion was to group fragments displaying the same
coefficient of temperature dependence with the same
precursor. Some fragments certainly had more than one
precursor. Nevertheless, at 300°, reaction products of
NaBF, and steam accounted for at least 30% of the
total vapor pressure, most of which was HF..

9.5 SPECTROSCOPIC INVESTIGATIONS OF
HYDROGEN- AND DEUTERIUM-CONTAINING
IMPURITIES IN NaBF, AND NaF-NaBF,
EUTECTICS

John B. Bates Harold W.Kohn
Jack P. Young Marvin M. Murray
George E. Boyd

Evidence for the existence of hydrogen-containing
impurities in NaBF,; was first obtained from near-
infrared spectra of the molten salt and in mid-infrared
spectra of pressed pellets of the “pure” crystalline
material. Bands observed in the region between 1.35
and 2.55 p (Fig. 9.2) are believed to correspond to
overtones and combinations of fundamental vibrations
of an OH species (either free OH™ or BF;OH ") in the
melt, Although the vibrational spectrum of crystalline
NaBF, has been thoroughly investigated,” ® no infrared
measurements have been reported in the region above
3000 ¢cm ™' with thick samples for transmission meas-
urements.

We have recently repeated infrared transmission meas-
urements on a single crystal of NaBF, supplied by L. O.
Gilpatrick. This crystal was polished to a thickness of
about Y% to Y% mm before examination with a
Perkin-Elmer model 621 spectrophotometer. A single
sharp band at 3641 cm ™! was observed which exhibited
about a 2:1 dichroic ratio as the infrared polarizer was
rotated through an angle of 90°. This band was assigned
tentatively to an OH-containing molecular species (most
likely BF3;OHT) which is effectively isolated in an
NaBF, matrix. Subsequent spectroscopic studies'® on
relatively pure samples of crystalline NaBF3;OH appear
to confirm this assignment.

Spectroscopic observation of isotopic exchange of
hydrogen with deuterium was investigated in solid-state
and in molten-salt exchange reactions. In the first
experiments, D, O vapor was bubbled slowly into a

7. H. A. Bonadeo and E. Silberman, Spectrochim, Acta 26A,
2337 (1970).

8. J. B. Bates, A. S. Quist, and G. E. Boyd, J. Chem. Phys.
54,128 (1971).

9. J. B. Bates, ibid., 54 (1971) (in press). .
10. J. B. Bates and A. S. Quist, work in progress (1971).

94

ORNL-DWG. 71-3460

!

&

5 W—’\:—

>

=

<I

" - |
T 3641 2688
y |
(0l

<

FREQUENCY (cm™)

Fig. 9.2. Transmittance and absorbance of a 0.15-mg pellet of
a D, -treated sample of NaBF,.

molten NaF-NaBF, eutectic at 400°C. The amount of
deuterium introduced into the melt correspondéd to
about 200 ppm D,0. Two sharp bands were observed
at 3641 and 2688 cm ™!, respectively, in the infrared
spectra of pellets pressed from a quenched sample of
this melt. The band at 2688 cm ™! corresponds to the
OD stretching frequency of an impurity assumed to be
NaBF;0D. A quantitative determination of the amount
of OH impurity in the starting NaF-NaBF, eutectic and
of the percent of isotopic exchange could not be made;
but, judging from relative peak intensities of the 3641-
and 2688-cm~' bands, the OD concentration was
estimated to be about 30% of the OH concentration.
Much larger amounts of the OD-containing impurity
were incorporated when D, 0O vapor was passed over
solid NaBF, at 377°C. The OD concentration in these
samples was estimated to be 3600 ppm by titrating DF
liberated in the reaction

NaBF, + D,0 - NaBF; 0D + DF .

The extent of reaction and exchange of DF with
NaBF, was studied by equilibrating molten NaBF, at
450°C with DF and argon. Subsequent infrared meas-
urements on pressed pellets of the solid material
obtained from quenching this melt also revealed two
sharp bands at 3641 and 2688 c¢cm™'. The infrared
spectrum of a ~0.25-mm-thick pellet of the DF-treated
NaBF, in the region above 2000 cm ™" is shown in Fig.
9.3. This spectrum is typical of others obtained with
ORNL-DWG. 71-3462

TRANSMISSION —

2688

3641

FREQUENCY (cm!)

Fig. 9.3, Infrared spectrum of a 100-mg pellet of DF-treated
NaBF4.

samples of NaBF,; which had been treated with D, 0.
The production of the OD-containing material in this
case could occur by reaction of DF with oxides (e.g.,
DF + BF,0™ - BF30D") and by isotopic exchange
with the OH-containing impurity.

Solid samples of KF, KBr, and NaBF, were treated
with DF at 350°C to check on the possibility that the
bands at 3641 and 2688 cm ™' may be caused by HF
and DF, respectively, dissolved in NaBF, . Alkalimetry
showed that the solids after treatment contained 5.46,
0.25, and 0.097 meq/g of acid respectively. No band at
2688 cm™! was detectable, although a weak band at
3641 cm™' appeared in the spectrum of the NaBF,
sample. The results from this experiment appeared to
confirm our earlier contention that the bands at 3641
and 2688 cm™' were not caused by dissolved HF and
DF respectively. : '

The results of the above experiments -demonstrated
that NaBF, contained an “OH impurity” (probably
NaBF3;OH) and that the deuterium-substituted form of
this material could be produced in molten NaBF, by
reaction with the BF,™ (D, 0) or with oxide impurities
(DF) and possibly by isotopic exchange. The extent of
isotopic exchange alone was investigated with a sample
of NaBF, prepared by S. Cantor which was sealed in a
nickel capsule and heated for about three days at
600°C in an atmosphere of D, gas (~25 mm). The
infrared spectrum of a pressed pellet of this sample is
shown in Fig. 9.4. The ratio of the absorbance of OD to
that of OH is about 3:1. The results of this experiment
appear to indicate an exchange of atomic deuterium for
the H in OH, although it is also possible to interpret
these results in terms of the reaction

D,(g) + Ni + 20%°(in the melt) =

20D “(in the melt) + Nj ,

95

ORNL -DWG, 71-346t

TRANSMISSION —>

|

2690

FREQUENCY (cm™h

Fig. 9.4. Band due to absorption by OD-containing species
in the infrared spectrum of molten NaBF,4 at 425°C.

The above experimental results demonstrated that
samples of presumed high-purity NaBF, contained OH
impurities and that exchange with deuterium atoms
occurs in the molten state. We thus desired to obtain a
direct measurement of the OH absorption in molten
NaBF, as a function of time. It was observed that the
bands due to overtones of OH disappear from the
near-infrared spectrum of the melt in a matter of hours.
However, pellets pressed from the frozen salt always
showed the sharp absorption at 3641 cm™". These two
observations can be reconciled if it is assumed that a
small, but unknown, concentration of OH or BF;OH™
is in equilibrium with NaBF,. Indeed, the rate of
decrease of the near-infrared OH peak does not appear
to follow simple first-order kinetics. Therefore, com-
peting, consecutive, or higher-order processes must be
involved. It is interesting also to note that the same
near-infrared spectrum is observed if either NaOH or
H, BO; is added to molten NaBF,, .

Near-infrared spectra of molten NaBF,; were meas-
ured with SiO, cells and a high-temperature furnace
described previously.!! For measurements in the mid-

11. J. P. Young, Inorg. Chem. 6, 1486 (1967),
infrared region, samples of NaBF, were spiked with
D;BO; because it was possible to study only the
absorption by OD rather than that by OH with the
available equipment. The infrared spectrum of a molten
NaBF,-D;BO; solution in the 2700-cm™' region at
about 425°C is shown in Fig. 9.4. The broad OD band
is centered at about 2690.cm !, in excellent agreement
with the frequency of the sharp band observed at 2688
cm ™! in the pellet spectra (Fig. 9.3).

A slight decrease in the absorption of the 2690-cm ™
band with time was observed. However, the band was
still observable after 48 hr in a 2-mm-path-length SiO,
cell. Under our experimental conditions, the OD ab-
sorption in the melt was stable in the presence of added
INOR-8 but unstable in the presence of pure chromium
metal. It is not known how much OH or OD was
actually observed in either the melt or pellet spectra.
Since it was demonstrated that OD species can be seen
spectrally in the melt, an apparatus which does not use
Si0, as a window material is being designed and
fabricated for use in further studies on molten NaBF,.

The results of these spectroscopic studies indicate
that an OH-containing impurity occurs in “pure”
samples of NaBF, , that exchange with deuterium atoms
takes place in the molten state, and that the OH species
persists at melt temperatures of 425°C over a period of
days. Additional experiments are under way to deter-
mine quantitatively the amount of OH-containing
impurity in NaBF, and the extent of isotopic exchange
of H for D or T in the molten state at temperatures of
about 600°C. '

1

9.6 RAMAN SPECTRA OF THE
HIGH-TEMPERATURE PHASE OF
POLYCRYSTALLINE NaBF,

Arvin S. Quist  John B. Bates
George E. Boyd

Raman spectra were measured with polycrystalline
NaBF, at temperatures just above and below the
dimorphic crystal transition (245°C). The results of this
study rationalize anomalies in x-ray diffraction data for
NaBF, and related structures. The spectrum of the
high-temperature solid phase indicates the presence of a
highly symmetric effective field about the BF,” ion
similar to that previously observed in melt spectra. A
complete disordering of the fluoride ion positions must
be assumed for the spectral results to be consistent with
recent x-ray data on crystalline NaBF,; powders at
265°C, which have been indexed in the hexagonal
system.

96

The room-temperature crystalline form of NaBF, is
orthorhombic, space group Cmcm, and is isostructural
only with the room-temperature forms of CaSO, and
NaClO,4. All three compounds undergo phase transitions
at elevated temperatures. Sodium perchlorate has a
cubic structure above 308°C which is said to be

" isostructural with the high-temperature phases of many

perchlorates, tetrafluoroborates, and sulfates. The high-
temperature phase of CaSO, (above 1210°C) was
reported to be hexagonal. At 245 + 1°C, NaBF,
undergoes a change in crystal structure to a form first
reported as monoclinic. Dworkin and Bredig,!? how-
ever, have expressed the view that a lowering of the
symmetry of the high-temperature form without a
lowering of the number of molecules per unit cell is
quite unlikely. They. have indexed the powder pattern
reported by Pistorius, Boeyens, and Clark'?® on the
basis that a mixture of the orthorhombic low-
temperature phase with a high-temperature hexagonal
phase was measured. Recently, Bredig'? obtained x-ray
powder patterns of NaBF, at 265°C with a Buerger
precession camera. The powder photographs were
indexed in the hexagonal system with parameters ¢, =
500 £ 002 and ¢ = 7.75 * 0.03 A. Acceptable
agreement between observed and calculated diffraction
intensities, however, could not be obtained for ordered
hexagonal space groups.

In view of the uncertainties in the structure of the
high-temperature phase of crystalline NaBF,, we have
measured the Raman spectrum of this material in an
effort to provide additional information regarding the
lattice symmetry. Static and dynamic field effects on
the vibrational modes of BF,” in the orthorhombic
(room temperature) NaBF, structure have been estab-
lished by previous studies.!5:'® The two BF, ions in
the primitive unit cell of this crystal occupy C,, sites,
so that all the components of the doubly degenerate
v,(e) mode and of the triply degenerate v;(f,) and
v4(f,) modes are observed in the Raman spectrum. An
increase in the site symmetry of the BF,” ion in the
high-temperature form may cause a collapse of the

12, A, S. Dworkin and M. A. Bredig, J. Chem. Eng. Data 15,
505 (1970).

13. C. W, F. T. Pistorius, J. C. A. Boeyens, and J, B. Clark,
High Temp.—High Pressures 1,41 (1969).

14, M. A. Bredig, ORNL MSRP Monthly Progr. Rep.
(December 1970—-January 1971), MSR-71-13, p. 21 (internal
memorandum).

15. J. B. Bates, A. S. Quist, and G. E. Boyd, J. Chem. Phys.
54,124 (1971).

16. J. B. Bates, ibid., 54 (1971) (in press).
splitting, and hence only one frequency for each of the
vibrational modes would be observed. The presence or
absence of such effects in the spectra of the high-
temperature phase would establish the symmetry of the
effective field about the BF,;” ion in the high-
temperature lattice and perhaps also resolve the existing
conflict in the x-ray data.

In the current Raman studies of the high-temperature
phase of NaBF, only the »,(a,), v;(e), and v,(f;)
modes were examined in detail;! 7 the intensities of the
v3(f,) modes are quite weak under these conditions. A
detailed comparison of the v, and v, bands at 240 and
254°C is given in Fig. 9.5, which clearly shows the
abrupt change in the Raman spectrum of crystalline
NaBF, as it undergoes the phase transition.

The experimental observations indicate that the effec-
tive environment about the BF,” ion in the high-
temperature crystalline phase of NaBF, has tetrahedral
(or higher) symmetry and that there is no measurable
(or allowed) dynamic coupling between neighboring

BF,” ions. This observation is consistent with the recent.

x-ray results,!* which led to a hexagonal structure only

if a complete disordering of the fluoride positions is
- assumed. With this complete disordering, the effective
field experienced by a BF,; ion would be highly
symmetric, and the lack of splitting of the », and v,
bands in the Raman spectra could be explained. This
structure also would be consistent with recent infrared
spectra obtained to 270°C in which the »; mode
disappeared in the high-temperature phase and the three
components of v, observed at room temperature were
reported to merge into a single component above the
transition point.!® Furthermore, the assumption of a
disordered lattice structure explains the similarity be-
tween the Raman spectrum of the high-temperature
crystalline phase and that of ihe melt.!® In the
high-temperature crystalline phase the BF,” ions are
undergoing large-amplitude random reorientations
(librations), which gives rise to the disorder, and a given
BF, ion “sees” an average symmetric field similar to
that experienced in the melt. Because the neighboring
BF, .ions are closer together in the high-temperature
solid than they are in the melt,?® the F—F nonbonded
repulsion is larger in the solid and a higher »; frequency
is observed.

17. A more extensive discussion of the results of the present
study is contained in a paper by the authors which is to be
published in J. Chem. Phys.

18. H. A. Bonadeo and E. Silberman, Spectrochim. Acta
26A, 2337 (1970).

19. A. S. Quist, J. B. Bates, and ‘G. E. Boyd, J. Chem. Phys.
54,4896 (1971). ‘

97

ORNL- DWG. 71-962

240°C
4.0 cm! slit
5x10%¢c/s

> 1 l ] | 1 | 1 l I | I I
5 550 530 510 370 350 330
=
L
2| me,
4.0 cm ' slit
5x102¢/s
1 ] ) | ) | ! | | I |
550 530 510 370 350 330

FREQUENCY (cm™!)

Fig. 9.5. Raman spectra of the v, and v, regions of
polycrystalline NaBF, at 240 and 254°C. A phase transition
occurs at 245°C.

A disordered hexagonal structure also can explain the
observation of optical isotropy of NaBF, in the
high-temperature phase.”" The fluorine atoms make the
largest contribution to the polarizability of the crystal,
and their orientation will determine the symmetry of
the crystal indicatrix. Complete disordering of the F
atoms would give a nearly spherical indicatrix, and the
crystal would appear to be optically isotropic.

20. The density of NaBF, is higher in the high-temperature
solid than in the melt, and the anion-anion separation is
generally known to increase in going from the solid to the
molten state (J. Braunstein, “Statistical Thermodynamics of
Molten Salts and Concentrated Aqueous Electrolytes,” in Jonic
Interactions. Dilute Solutions to Molten Salts, S. Petrucci, ed.,
Academic, in press).

21. George Brunton, ORNL, Reactor Chemistry Division,
private communication (1970).
9.7 ANEW METHOD FOR SYNTHESIS
OF NaBF, OH

L. O. Gilpatrick  C. J. Barton

Laboratory studies are continuing in attempts to
synthesize hydrogen-containing species for the reten-
tion of tritium in the MSR coolant salt. One of these
species is the first hydrolysis product of NaBF, in
aqueous solution, which has been reported to be
NaBF;O0H.?2>23 Ryss and Slutskaya® have reported a
successful synthesis of this compound in which boric
acid is reacted at 0°C with a saturated solution of
NaHF, as follows:

INaHF, + H, BO, % NaBF,OH + NaF .
2

This synthesis has the disadvantage that it produces a
mixture of NaBF;OH and NaF and in addition also
becomes contaminated with some NaBF,; produced by
a side reaction when attempts are made to recrystallize
it from water solutions.

A new synthesis has been devised for NaBF;OH
which employs the following reactions in a saturated
solution of NaHCO;:

BF, + H,0 "~ H(BF,OH) ,
H, O
H(BF, OH) + NaHCO, 3—2%» NaBF;OH + CO, + H,0 .

One mole of solid NaHCO; and BF; gas are added
progressively to 5 moles of water at 0°C, which
produces a clear solution. Four volumes of cold 95%
ethanol are then added. Storage at 0°C gives a good
yield of product which has a better purity than that
produced by the method of Ryss.

The identification of this material is relatively certain
and direct based on its properties. Sodium hydroxytri-
fluoroborate is very water soluble and shows rapid
hydrolysis to NaF, H;BO;, and NaBF, when stored as
an aqueous solution. It is birefringent to polarized light
and exhibits refractive indexes from a maximum of
1.350 to a minimum of about 1.343 at 25°C. The
crystals belong to the hexagonal system and are polar,

22. A. Travers and L. Malprade, Bull. Soc. Chim. 47, 788
(1930).

23. C. A. Wamser, J. Amer. Chem, Soc. 70,1209 (1948).

24. L. G. Ryss and M., M. Slutskaya, J. Gen, Chem, USSR 22,
45 (1952).

98

as has been shown by a recent x-ray structure study?®®
which makes positive identification convenient. ’

9.8 SOLUBILITY OF BF, GAS
IN FLUORIDE MELTS

S.Cantor W.T.Ward

The purposes of these measurements are to relate BF,
solubility to changes in the thermodynamic properties
of molten fluorides and to provide data relevant to the
use of BF; as a burnable neutron poison for purposes

~of reactor control. The single-vessel apparatus and the

procedures of measurement have been described previ-
ously.?®

Solubilities of BF; have been determined in five
molten-salt solvents composed of LiF and BeF,. The
chief results obtained thus far are:

1. In each solvent, the higher the temperature, the
lower the BF; solubility; that is, the enthalpy of
solution is exothermic.

. At constant temperature and pressure, the higher the
concentration of LiF, the higher the solubility of
BF;; indeed, the Henry’s law constant appears to be
linear with the thermodynamic activity of LiF.

3. Henry’s law is obeyed in almost all cases; however,
where -the concentration of BF; in the melt
approaches 1 mole % or greater, there are discernible
positive deviations from Henry’s law, '

The data and derived thermodynamic information are
summarized in Table 9.3. The magnitude and negative
sign of the enthalpy of solution suggest a strong
interaction of BF;(g) with the melt. There does not
appear to be a correlation of enthalpy of solution (of
BF3;) with melt composition. The approximate con-
stancy of AH in Table 9.3 signifies that plots of log K
(Henry’s law constant) vs reciprocal temperature (in °K)
are approXimately parallel.

As may be noted in column 3 in the Table, the higher
the mole fraction of LiF, the higher the Henry’s law
constant. In Fig. 9.6, the Henry’s law constant at 600°C
is plotted vs activity of LiF reported®”’ for these melts;
the excellent linearity of these two properties also holds
at other temperatures. An interpretation of this linear

25. M. T. R. Clark and H. Lyton, Can. J. Chem, 48, 405
(1970).

26. S. Cantor and W. T. Ward, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, pp. 78-179. '

27. B. F, Hitch and C. F. Baes, Jr., Inorg. Chem. 8, 201
(1969).
99

Table 9.3. Solubility? of BF3 in molten LiF-BeF, solvents; enthalpy and entropy of solution

Solvent Temperature Henry’s law b b
composition range measured K (mole fraction AH AS_, -1
(mole % LiF) ©o) BF 3/atm) at 600°C (keal/mole) (cal mole = deg ")

6364 470-653 0.00207 T —15.1 -29.6
66 520-725 0.00242 -15.6 -29.9
70 548-732 0.00354 -15.5 : -29.0
7SC 641857 0.00500 —15.9 -28.7
80 714866 0.00613 -15.6 -28.0

2Pressure range, 1.3—3.0 atm,

AS AH

bCalculated from the equation In K = — —

R

RT’

¢Measurements in this solvent composition not completed.
dIn the pressure range 1,3-3.0 atm, at temperatures below 520 °C. there are positive deviations from Henry’s law.

ORNL-DWG 71-7250

0.006

{mole % LiF)
80’

/

/

0.005

~J

[ ]

0.004

70,0

HENRY'S LAW CONSTANT {mole fraction BF3/aim)

0003

0.002 63 .V

04 0.2 03 04 05

ACTIVITY OF LiF

06

or

Fig. 9.6. Henry’s law solubility of BF3 vs activity of LiF in

molten LiF-BeF, at 600°C.

behavior is that the ionic reaction equilibrium

F~(1) + BF;(g) = BF, (1)
occurs, from which it follows that

aBF‘;_ - KfPBF3 ag -,

(1)

mole fraction BF3 dissolved /atm

ORNL-DWG T{-7254

00059 | ‘ ?
00057 |— ' /
* MEASURED {
o EXTRAPOLATED FROM »
HIGHER TEMPERATURE

0.0055 ,

S

00053
/

— / | *_IFHENRY'S LAW HELD __
00051 \ = |
EFFECT ON SOLUBILITY
’OF AN ERF\‘OR OF {°C
0.0049 f
0 1 2 3

PRESSURE OF BFz (atm}

Fig. 9.7. Solubility of BF; in 63-37 mole % LiF-BeF, at
517.6 + 0.5°C.

where ag . - and ag - are, respectively, the activities of
fluoroborate and fluoride jons; K; is the equilibrium
constant of the reaction (like any equilibrium constant,
K; varies with temperature); and Py is the equilib-
rium pressure of BF;. At constant temperature and
pressure, Eq. (1) explicitly states that app,- varies
linearly with ap-. In this interpretation of the phe-
nomena depicted in Fig. 9.6, the concentration of
dissolved BF; is identified with the activity of BF,4 in
the melt, while the activity of LiF is, of course,
identified with the activity of fluoride ion in the melt.

In the pressure range under study (1.3—3 atm) the
solubility of BF; is sufficiently low that the dissolved
molecules do not interact with one another. This -
noninteraction between solute molecules is the basis of
Henry’s law. In this investigation virtually all the
solubilities of BF; obey Henry’s law. However, Henry’s
law does not seem to hold at .relatively low tempera-
tures in the melt of composition LiF-BeF, (63-37 mole
100

%). At temperatures below 520°C the solubilities of
BF; become 1 mole % or greater. Figure 9.7 shows data
which do not obey Henry’s law; although the data
exhibit some scatter, the solubility of BF; per unit
pressure almost certainly increases with pressure; that
is, the deviations from Henry’s law are positive. Further
measurement will be made to confirm these deviations.

9.9 EQUILIBRIUM PHASE RELATIONSHIPS
IN THE SYSTEM RbF-RbBF,

L. O. Gilpatrick  C. J. Barton

An investigation of the equilibrium phase diagram for
the RbF-RbBF, system was completed. Final revisions
to the system diagram, as reported previously,*®
resulted from measurements with purer rubidium
fluoride than was previously employed.

New RbF has been obtained and purified by treat-
ment with dry H, and HF at temperatures up to and
above the melting point to remove water and hy-
droxide. Rubidium fluoride was produced which has a
melting point of 793°C, which is in good agreement
with the literature value of 795°C. Binary mixtures of
RbF and RbBF, were prepared at 5 mole % intervals
ranging up to 20 mole % RbBF,.

28. L. O. Gilpatrick and C. J. Barton, MSR Program
Semiannu. Progr, Rep, Feb, 28, 1970, ORNL-4548, p. 133.

Differential thermal analysis (DTA) was used to
examine these compositions, which gave the expected
higher temperatures in each case, lending support to the
belief that the irregularities in the original part of the
study were due to RbF of insufficient quality.

The phase diagram in its final form is shown in Fig.
9.8.

9.10 ACTIVITIES IN ALKALI FLUORIDE —
FLUOROBORATE MIXTURES

D. M. Moulton J. Braunstein

Thermodynamics of mixing in alkali fluoride — alkali
fluoroborate mixtures are of interest in connection with
proposed coolants for the MSBR. Deviations from
ideality, although small, are less well understood in such -
common cation binary mixtures (and have been less
thoroughly studied) than in common anion binary
mixtures. Endothermic mixing is frequently observed
rather than the exothermic effect expected from
electrostatic considerations, and may be associated with
changes in packing of large anions of differing size as
well as with the dispersion energy of the mixed
anions.?® Here we report some activity coefficient

29. J. L. Holm and O. J. Kleppa, J. Chem. Phys. 49, 2425
(1968); O. J. Kleppa, Annu. Rev. Phys. Chem. 16, 187 (1965);
J. Brdunstein in lonic Interactions, S. Petrucci, ed., p. 179,
Academic, New York, 1971. ‘

ORNL-DWG 70-30234A

800.
1 \l\ | l I
R 4 LIQUIDUS
\n\ o SOLIDUS
700 ~] e POLYMORPHIC TRANSITION -
\ IN RbBF,
600 \\ LIQUID
" b
LIQUID + RbF /n/.
T 500
Q
= LIQUID + & RbBF,
g ———o— d42°C ‘;\‘5/.* o]
& 400
14
a
= a RbBF, + RbF
E 300 |
- 247°C .
200
B RbBF, + RbF
100
0
RbF 10 20 30 40

50 60 70 80 90 RbBF,

RbBF,; (mole %}

Fig. 9.8. The system RbBF4-RbF.
calculations from available data in fluoride—fluoro-
borate mixtures.

We have used the phase diagrams of the alkali
fluoride—fluoroborate systems (Na, K, Rb) to calculate
the component activities along the liquidus lines. The
method was the usual one of integrating the Schréder—
van Laar equation,

to give the ideal liquidus and then comparing actual and
ideal mole fractions. A power series was used for the
heat capacity difference. Each activity refers to the
hypothetical pure liquid at the same temperature. The
sources of the data are Barton et al.,>° Kubaschewski
et al.,>! and Dworkin and Bredig.??

In all three cases the behavior of the activity
coefficients is qualitatively the same (see Table 9.4).
The fluoride shows positive deviation at first but then
becomes negative before it reaches the eutectic. The
fluoroborates always show negative deviation (with
sodium it is only very slight). In both cases the sodium
salt shows the smallest and rubidium the largest
tendency toward negative deviation.

A comparison with the activity coefficients of
Cantor,>® which were derived from decomposition
pressure measurements, shows that for NaF we agree
fairly well up to about 50 mole % but then deviate
increasingly. For KF the agreement is worse, and for

Table 9.4. Activity coefficients in alkali fluoride—
fluoroborate mixtures at 1000°K for Na
and 1100°K for K and Rb

Xpp, 02 04 06 07 08 09 095
YNg 110 120 123 119 112 0.89 J[ 1.00(-)
yna® 115 126 131 132 133 133 H 1.00(+)
Tk 1.05 1.08 1.02 0.97 H 0.98 099 1.00(-)
y® 114 124 132 13514101 1.00 1.00
Yrp 102 100 093W 094 097 099 099

dCantor,

[[ Eutectic: fluoride to left, fluoroborate to right.

101

the fluoroborates the direction of deviation is the
opposite. The activity coefficients are shown; they have
been put on the same standard temperature T, as
Cantor’s by saying that Yr, = 'yT(T/TO) (i.e., ideal
entropy of mixing), which does not change them much
and in particular does not affect the shape of the curve
or the direction of deviation.

The differences in the calculated activity coefficients
may indicate a significant temperature dependence of
the activity coefficients since the eutectic temperatures
lie some 500 to 700° below the temperatures of the
fluoroborate decomposition pressure measurements.
Deviations from ideality calculated from the NaF-
NaBF, and KF-KBF, phase diagrams®® are compared
with those calculated from the NaF-Nal and KF-KI
phase diagrams®** and are shown as the interaction
parameters for excess free energy at the eutectic in
Table 9.5.

In both the BF,-F~ and 1°-F~ systems, deviations
from ideality become less negative with decreasing size
of the common cation, as would be expected if changes
in anion packing contribute positive deviations.

9.11 HIGH-TEMPERATURE CRYSTAL
STRUCTURE AND VOLUME OF SODIUM
TETRAFLUOROBORATE AND
RELATED COMPOUNDS

M. A, Bredig -

High-temperature x-ray diffraction patterns of NaBF,
and CaSQ, (isostructural) reported by several authors

30. C. J. Barton et al.,, MSR Program Semiannu. Progr. Rep.
Dec, 31, 1967, ORNL-4191, p. 158.

31. O. Kubaschewski, E. L. Evans, and C. B. Alcock,
Metallurgical Thermochemistry, Pergamon, Oxford, 1967.

32. A. S. Dworkin and M. A, Bredig, J. Phys. Chem. 64, 269
(1960); A. S. Dworkin and M. A. Bredig, J. Chem. Eng. Data
15, 505 (1970).

33. 8. Cantor, MSR Program Semiannu. Progr. Rept. Aug. 31,
1968, ORNL-4254, p. 170; MSR Program Semiannu. Progr.
Rep. Feb. 28, 1869, ORNL-4344, p. 159.

34, E. M. Levin, C. R. Robbins, and H. F. McMurdie, Phase
Diagrams for Ceramists, 1969 Supplement, The American

~ Ceramic Society, Columbus, Ohjo, 1969.

Table 9.5. Interaction parameters for excess free energy at the eutectic

NaF-NaBF,
—0.53

Binary system
GE/x x5 (kcal/mole)

KF-KBF,
-0.73

NaF-Nal
+0.62

KF-KI
—-0.70

Table 9.6. Powder x-ray diffraction pattern of
high-temperature form of NaBF, at 265°C

hkl Intensity d, obs. (A) d, calc. (A)?
100 ww 434 433
002 3.90 to 3.88
101 v 3.77 3.78
102 m 2.89 2.89
110 w 2.50 2.50
200 2.165
112 m 2.097 2.100
202 w 1.884 1.890
104 w 1.770 1.770
203 1.659
210 w 1.645 1637
114 1.531
212 vs 1525 1.508
302 1.353

9Hexagonal: 2 = 5,00 + 0.02 &;¢c = 7.75 £ 0.03 A.

were critically reviewed and reinterpreted in earlier
~ reports.>®> A hexagonal crystal structure was found to
take the place of a monoclinic one suggested by
Pistorius et al.>® for NaBF, and of a cubic one given by
Floerke®” for CaSO,. The hexagonal structure seems to
be a novel one among compounds of the stoichiometry
M(ZX,). Somewhat like (NH,4)F, it appears to be
related to the hexagonal structure of ZnS, wurtzite, or,
much less likely, that of nickel arsenide (NiAs), in a
manner similar to the relationship between the cubic
high-temperature form of the other alkali fluoroborates
and the structure of rock salt.>®

After constructing a simple heating device, auxiliary
to the Buerger x-ray diffraction precession camera, the
hexagonal, bimolecular unit cell with 4 = 5.00 £0.02
and ¢ = 7.75 + 0.03 A has now been directly confirmed
on powder samples of NaBF, at 265°C, that is, about
20° above the transition point (Table 9.6). From this,
one calculates the molar volume as 50.6 * 0.6 cm?®/
mole and the specific gravity as 2.13 + 0.02 g/cm?.

When, on considering atomic positions, by analogy
with (NH4)F the space group Cg, (C6,mc) was used,

35. M. A. Bredig, Chem. Div. Annu. Progr. Rep. May 20,
1968, ORNL-4306, pp. 129—-30; Chem. Div. Annu. Progr. Rep.
May 20, 1970, ORNL-4581, pp. 116—18. '

36. C. W. F. T: Pistorius, J, C, A, Boeyens, and J. B. Clark,
High Temp.—High Pressures 1,41-52 (1969).

37. O. W. Floerke, Naturwissenschaften 39,478 (1952).

38. Ralph W. G. Wyckoff, Crystal Structures, 2d ed., vol, III,
Interscience, New York, 1963.

102

distances in the unit cell were too short to accommo-
date an F (or an O) as a bridge between B and Na (or S
and Ca), corresponding to the hydrogen bridging in
(NH,)F. For ordered orientation of the BF,  (or
S0,%") ions within the cation tetrahedra, no agreement
between diffraction intensities observed and calculated
(with the help of G. D. Brunton) was obtained thus far.
Attempts to reconcile the discrepancies by a highly
disordered model, or an even more modified structure,
are being continued. A high degree of (rotational)
disorder of the tetrafluoroborate ions would be in
agreement not only with the relatively large entropy of
transition,>® but might also reconcile the optical
isotropy, observed by G. D. Brunton, and apparent
degeneracies in the Raman spectrum, measured by
Quist, Bates, and Boyd (Sect. 9.6), with the hexagonal
symmetry based on the powder x-ray pattern.

The volume change, AV/V,, from room temperature
(Vo) to slightly above the transition point of NaBF, is
15.5%, which is slightly smaller than a previous, rough
estimate of 18%.*° If for comparison one takes the
more accurate high-temperature data for NaClO, of
Braekken and Harang®' instead of those given by
Wyckoff>® based on the results of Herrmann and
Iige,*? AV/V, for NaBF, is not similar to but is almost
twice as large as that for NaClO, .

Figure 9.9 shows the corresponding volume changes
for alkali tetrafluoroborates (partly based on Brunton’s
work®? and on my own recent high-temperature
measurements), for alkali perchlorates,®®>*! and the
sulfates of Ca, Sr, and Ba.’® The value of AV/V, for
the tetrafluoroborates is consistently larger (~30%)
than that for the corresponding perchlorates and
alkaline-earth sulfates. The earlier speculative estimate
of similarity in AV/V, (ref. 40) was based on a value
for the volume of the high-temperature form of KBF,
given by Finbak and Hassel,** which I found to be too
small by 6%. ,

Figure 9.9 demonstrates the large departure of the
two sodium salts from the trends found with the other

39. A. S. Dworkin and M. A. Bredig, J. Chem. Eng. Data 15,
505-7 (1970).

~40. Stanley Cantor, Dana P. McDermott, and L. O. Gil-
patrick, J. Chem. Phys. 52,4600—-4604 (1970).

41. H. Braekken and L. Harang, Z, Kristallogr. 75, 53849
(1930).

42, K. Herrmann and W, Ilge, ibid., 41-66 (1930).
43. George Brunton, Acta Crystallogr. B24, 1703-4 (1968);
ibid., B25, 2161-62 (1969).

44, Chr. Finbak and O. Hassel, Z, Phys. Chem. B32,433-38
(1936).
ORNL- DWG 71-7252

AV/Vyy, RELATIVE VOLUME INCREASE (%)

25 \
f N
’ KBF>\
20 /
/
[ srs0, 8™\ RbBF,
f —
AN
d ,/ KCIO:“.‘quo4
15 |NaBF, — .
/
RbCIO,
// \\CSBF4
10 A
%
J<I-Nacio, CsCl04
I/oc:aso4
/
5
09 10 141 12 13 14 15 16 17

CATION RADIUS (A)

Fig. 9.9. Relative volume increase (%) from room tempera-
ture through transition in alkali metal tetrafluoroborates,
perchlorates, and alkaline-earth sulfates as a function of
cation size.

six alkali and alkaline-earth metals. Recently, prelimi-
nary comparisons were made between the volume
changes shown in Fig. 9.9 and isothermal volume
changes AV, at the transition temperatures, calculated
from the measured transition entropies, AS,,, and
pressure dependencies of the transition temperatures,
dT,/dp,**>*® by means of the Clausius-Clapeyron
relationship. These calculations indicate that actually
the relative volume change on transition, AV, /V, for
NaBF, is quite similar to that for KBF, and RbBF,,
namely, 10%. Thus the apparent deviation for NaBF, in
Fig. 9.9 must be caused by the thermal expansion
coefficient in the low-temperature orthorhombic phase
being considerably smaller than in those of KBF, and
RbBF, of a different structure. This is reminiscent of
the difference in the thermal expansion of NaNO; and
KNO,. Additional x-ray diffraction measurements are
planned to confirm these observations.

45, C, W, F. T. Pistorius and J. B, Clark, High Temp.—High
Pressures 1,561-70 (1969).

46. Carl W. F. T. Pistorius, J. Phys. Chem, Solids 31, 38589
(1970).

103

9.12 HYDROGEN PERMEATION THROUGH
OXIDE-COATED METALS

R. A. Strehlow

The control of tritium flow in a molten-salt reactor
system involves both knowledge of permeation rates
through steam generator heat exchanger material and
the extent to which-permeation is decreased due to the
oxide coating which is expected to exist in contact with
the steam. It has been long recognized that oxide
coatings on metals impede hydrogen permeation.®”-*®
Occasionally large impedance to flow of hydrogen has
been observed and attributed or related to an oxide
film,*®+*? The impedance to hydrogen flow is often,
however, observed as an artifact of an experimental
determination of hydrogen permeability. It is reflected
as an aging or conditioning time often of subsidiary
interest. - A

An experimental program was begun to study hydro-
gen permeation under conditions closely analogous to
those which might be expected in a molten-salt reactor
steam generator heat exchanger. The apparatus is shown
schematically in Fig. 9.10. Since several modes of
operation are possible, note need be made only that the
diffusion membrane (a piece of tubing of normal wall
thickness) separates two regions, A and B. Region A is
analogous to the steam side of a heat exchange tubing
section in that the gas which can be admitted and
analyzed may be admitted wet or dry. Region B is
analogous to the inside of a heat exchange tubing
section only to the extent that a gas may be flowed
through the tube and analyzed. (A suitable mode of
operation would be with this side evacuated.) Salt has
not been used in permeation studies conducted so far.

A quadrupole mass analyzer was used variously to
assess leak-tightness of the apparatus, gas purity, and
approximate analysis, as well as to follow the changes in
gas composition in both regions A and B for different
experiments. The mass analyzer was equipped with a
source which permitted fairly high pressures to be used
(in the range of 0.1 to 0.3 torr). Under these conditions
an enhanced sensitivity for several species is observed
due to the occurrence of ion-molecule reactions.

Two stainless steel samples were selected for two
initial experiments in order to correlate this work with
earlier work on hydrogen permeation. Estimates of flow

47. C. J. Smithells and L. E. Ransley, Proc. Roy. Soc. A150,
172-97 (1935).
48. P.S. Flint, KAPL-659 (Dec. 14, 1951).

49, Tennyson Smith, J. Nucl. Marer. 18,323-35 (1966).
ORNL-DWG 71-7253

TO PUMP,

ARGON-HYDROGEN —=

/DIFFUSION SAMPLE

MASS-ANALYZER

M

GAS INLET FCR [

=== GAS INLET

(WET OR DRY)

—==TO VENT

I
TO PUMP AND
MASS-ANALY ZER

Fig. 9.10. Schematic diagram of permeation measurement apparatus.

were made using ordinary techniques of pumping speed
and volume measurement. Thermocouple gages were
used for measurement of pressures. The two experi-
ments studied (1) permeation of protium from a wet or
dry argon-hydrogen mixture in region A through a
stainless steel (304) tube into a deuterium atmosphere
in region B and (2) a similar measurement of the
permeation of protium through a stainless steel type
347 tube into evacuated region B, with monitoring of
hydrogen concentration in region A.

The first experiment was performed as an attempt to
follow the development of an oxide layer by watching
the fall of protium partial pressure in region B after
adding steam to the gas in region A. It was observed
that the protium partial pressure increased when water
vapor was added to the argon-hydrogen mixture. The
second experiment was conducted to clarify the ques-
tion of the hydrogen partial pressure in region A when
the gas was saturated with water vapor. Analysis of the
data indicated that the increase was less than 2 or 3%.

The first experiment was conducted at a temperature
of 579 +-2°C. The results are shown in Fig. 9.11.
Deuterium flowed through a Y, 4-in. tube 32 in. long
with a wall thickness of 0.010 in. Evacuation of region
A vyielded a log P vs time plot with a slope corre-
sponding to 0.018 cm?®/sec, a value of 1 atm for the
deuterium pressure, and 45 u for the steady-state
(corrected) pressure of deuterium in region A measured
by a thermocouple gage; a value of
1/2

3
0‘09cm mm atm

hr cm?

was obtained for the 304 stainless steel tube used here.
This compares quite well with the value obtained by
Webb,3® although it was not intended to be a precise

50. R, W. Webb, Permeation of Hydrogen through Metal,
NAA-SR-10462 (July 25, 1965).

ORNL-DWG 71- 7254
SCAN NUMBER

o 100 200 300 400 500 600
100

rp—\wmi? 3 (DHM)
O Y |
O S
o}

50 ’é;
S~

I MAaSS 2 (DY)

20 o R e :

MASS 5 (DpH™)

I* {arbitrary units)
)
T

5 A

Tl

]

I—- Ar, Hp, Hp0
2 — DRY Ar, H2 ON T DRY Al’ HZJ ]
EVACUATE REGION A J
D, OFF (REGION E!)

1

o 700 1400 2100 2800 3500 4200

TIME (sec)

Fig. 9.11. Protium permeation into deuterium from steam-
hydrogen-argon.

measurement of the permeability. Since mass analysis
was conducted at high pressure (150 u in the ion
source) numerous ion-molecule reactions occurred.
From deuterium, masses 2, 4, and 6 are obtained with
mass 4 (D,") dominant, mass 6 (D3") to the extent of
about 10% of mass 4, and mass 2 (D) to the extent of
about 1.5% of mass 4. The amount of protium
measured was about 0.6% of the deuterium peaks and
was therefore reflected almost entirely in changes of the
mass 3 and 5 peaks. These changes with alteration of
the experimental conditions are shown in Fig. 9.11. On
starting a flow of dry argon — 4% hydrogen, a rapid
parallel increase in the protium-containing peaks (3 and
5) occurred with no significant change in the mass 2

peak. After a steady state was reached (at about 700"

sec) the gas mixture was valved through a water
bubbler, which should have produced about a 2% water
content. However, instead of a decrease in protium in
region B, a slow increase was observed, which is
attributed to oxidation of the metal by the water. This
increase lasted about 12 min, after which a low decrease
was observed with a slope reflected both in the mass 3
and 5 curves of

dlog Py, ~ .
—r 03hr .

Since a slight decrease in mass 2 also occurred (with a
slope of 0.15 hr™"), the decrease of hydrogen permea-
tion into the steel tube (presumed due to the oxide
buildup) was probably about 0.2 for the value of the
slope. After 2000 sec the argon-hydrogen mixture was
valved so that the water bubbler was bypassed. A level
for protium was reached which was about the same as
the initial rate with dry gas. (The scatter of points for
mass 3 between scans 210 and 400 was due to a
recorder scale change.) The chamber was evacuated to
permit measurement of the permeability of the sample
as described above. :

The question of whether there was an alteration of
gas composition in region A during this study was
explored in a second experiment, A sample of stainless
steel (type 347, Y, in. in diameter, 11 in. long, 0.030 in.
wall thickness) was evacuated inside (region B) and
exposed to the same perturbations of the outside as
before (region A). The permeability measured (at
620°C) was

1/2

_, cm® mm atm

43 X 10

hr cm?

To determine whether there was an appreciable
change in the H,/Ar ratio when the gas was admitted
wet, use was made of the fact that although the
reactions

105

H, +H,0* >H,0*+H  —44 AH kcal/mole

and

H,*+H,0-H,;0"+H —202 AH kcal/mole
are exothermic and will occur, the ion-molecule reac-
tions involving the production of ArH* from Ar* do
not involve water or water ions, Hence the ratio of mass
41 (ArH*) to mass 40 (Ar*) should reflect the
hydrogen gas concentration.® !

The results summarized in Fig. 9.12 show an initial
condition with an ArH*/Ar* ratio of 2.29 * 0.05
initially and a precipitous drop on diverting the flow
through the bubbler due to the presence of air in the
bubbler volume. (N, has a proton affinity of 131
kcal/mole. Thus the ArH*/Ar” ratio dropped to 0.4
for 2 min.) The level of the ratio was reestablished
after .3 min and maintained at 225 * 0.02. Some
erratic behavior was observed during reestablishment
of dry conditions. The pressure level in region B
showed a 25% increase of pressure as compared with
45% observed in the first experiment.

Aside from the point of no appreciable hydrogen
being made by the admission of water is the sensitivity
for water shown in Fig. 9.12. Both by charge exchange
from the dominant gas, argon, and proton transfer
(principally from H,"), the water produces the domi-
nant ions in the spectra. While this is expected, it is
interesting to tabulate the spectra to show the enhance-
ment of sensitivity which is obtained by using ion-
molecule reactions. Table 9.7 shows the ionization
produced by these chemical ionization reactions and
compares them with the data from electron impact
ionization performed by increasing the mass spectrom-
etry repeller potential and decreasing the ion source
pressure to a few microns.

In summary, after establishment of steady permeation
of protium from an argon—4% hydrogen atmosphere
through types 304 and 347 stainless steel, switching to
wet conditions (about 2% H,O) yielded an increase in
protium permeation which is attributable to a corrosion
reaction. As the oxide coating increased in thickness,
the permeation rate slowly decreased. The protium

51. This approximation would not be valid at very high
water concentrations since the proton affinity of H,O is 169
kcal/mole and that of argon is but 54: fewer average coilisions
would be needed to be reflected in an ArH™ decrease. For the
ion source used here approximately 20 collisions are calculated
at 100 u.
106

Table 9.7. Comparison of chemical and electron impact
ionization for Ar—4% H, —2% H, O as percent of total

Argon H, ' H,O
(40 + 20 amu peaks) (2 amu) (17 + 18 amu)

Electron impact 92 5 3
ionization (40 amu) (41 amu) (18 + 19 amu)
Chemical (ion-molecule) 9 21 70
ionization '
ORNL-DWG 71-7255 increase is not due to an increase in hydrogen partial

pressure but to a direct charging of the hot metal with
hydrogen during the oxidation process. No diminution
of deuterium permeation was observed.

The questions on permeation processes still to be
l l resolved are:

acat, 19, H30* 1. the extent to which the slow oxide growth in the
K 3 presence of steam at any pressure up to normal
steam generator conditions will ultimately impede
18, "sz0+ ~ permeation,

o”

2. the extent to which independently applied coatings

A : -
sddage—p—a | Y can be used to decrease permeation rate under

+_ T Ad AedbAAA A 41
41, ArH 2 steady-state conditions,

3. the functional relationship between permeation and
partial pressure at very low partial pressure of
N ‘k hydrogen,

a trtrirenies 2 b : 4, whether the assumption of general equivalence of

\ protium permeation and deuterium permeation is
\ \ valid, that is, whether at any protium pressure
\ \ competition for desorption sites might impede out-
\ \ flow of deuterium under conditions like the first
experiment reported here,

ION CURRENT (arbitrary units}

19 | \ 5. the relations between alloy composition, structure
of the stable oxide coating, and effectiveness in
\ reducing hydrogen permeation,

6. the possible role of molten salts in the permeation

29(x ¥5) NoH*
process.

\ 28(x Y5} Nyt

w1
P

0 200 400 600 80C " {000
TIME (sec) .

Fig. 9.12. Hydrogen-argon ratios as monitored by masses 40,
41, 19 amu.
10. Physical Chemistry of Molten Salts

10.1 THERMODYNAMICS OF LiF-BeF, MIXTURES
FROM EMF MEASUREMENTS OF
CONCENTRATION CELLS

D.D.Sood!  K.A.Romberger? J. Braunstein

Electromotive force measurements of the molten-salt
concentration cell with transference

LiF ! LiF
Be I Be (A)
BeF, | BeF,
I II

have been reported previously along with their applica-
tion to refinement of the liquidus of the LiF-BeF,
phase diagram®»* and determination of the ionic
transference numbers,> when combined with pre-
viously reported emf measurements of concentration
cells without transference (The transference number of
an ionic constituent is the fraction of the electric
current carried by that constituent relative to a refer-
ence constituent.) Here we report some excess chemical
potentials and partial molar enthalpies of mixing
derived from such measurements as part of a study of
the thermodynamics of mixing in molten salts of
reactor interest. '

The electromotive force £ of the cell (A) may be
written®

1. Representative of the Indian Department of Atomic
Energy, on assignment at QORNL from Bhabha Atomic Re-
search Centre, Bombay, India.

2. Present address: Kawecki-Berylco Corporation, Hazelton,
Pa. .

3. K. A. Romberger and J. Braunstein, MSR Program
Semiannu. Progr. Rep. Feb. 28, 1970, ORNL-4548, p. 161; J.
Braunstein, K. A. Romberger, and R. Ezell, J. Phys. Chem. 74,
4383 (1970).

4. R.E. Thoma et al., J. Nucl. Mater. 27(2), 176 (1968).

5. K. A. Romberger and J. Braunstein, fnorg. Chem. 9, 1273
(1970).

6. B. F. Hitch and C. F. Baes, Jr.,

Inorg. Chem. 8, 201
(1969). :

107

(1

X duy;r »

where F is the faraday, ¢ ; is the transference number
of lithium ion relative to fluoride, x is the mole fraction
of BeF, in a BeF,-LiF mixture, and HpeF, and up;p
are the chemical potentials of BeF, and LiF. In order
to derive chemical potentials from isothermal measure-
ments of a cell with transference, it is necessary to
know the concentration dependence of the transference
numbers. In this system, it has been shown that t; =1
at temperatures between 500 and 700°C and composi-
tions between 0.3 and 0.7 mole fraction BeF,. (The
transference number probably is also equal to unity in
the concentration range between 0 and 0.3 mole
fraction BeF, since the transference number of a
constituent in a molten salt generally approaches unity
as its mole fraction approaches unity.)

The isothermal composition dependence of the excess
chemical potential may be obtained from emf measure-
ments of the cell (A) at closely spaced compositions in
the half-cell (II) (compartment I is a reference half-cell).
From Eq. (1) and the definition of the chemical
potential, the composition dependence of the excess

(a#fm> ) 3E -
x /o ox/) p
Measured values of AE/Ax at 500°C are shown in Fig.
10.1; the smooth curve is (0£/0x). Calculated values
of (aufiF/ 0x) are plotted against composition in Fig.
10.2.
The excess chemical potential of LiF along the

liquidus curve, calculated from the liquidus data,® can
be represented by the equation, fitted by least squares,

chemical potential is

X

1 +x

108

ORNL-DWG 74-3550

1600 I
jLIQUIDUS l .
! SYSTEM LiF-BeF,
H o
| et T =498 +0.5°C
1400 Z._’“ ..qm.; +—— SILICA CELL
hi H O —— METAL CELL
H
s LH H
S Tl H A DL
[Hg)
1200 : . S
H A,A‘H
2Y
iR
‘ ™,
T 1000 =
z \T
— ( ]
/—l- b_“‘
wlx
|0
= 800 “r\_q
e
\H
H HH
600
'_'t_u—&_k‘
W
~
400 -——\«\ =
— l\
200
0.30 0.34 0.38 0.42 0.46 0.50 0.54 0.58 0.62
X (mole fraction Bef,)

Fig. 10.1. Isothermal composition dependence of (AE/Ax)T for the concentration cell (A) with transference employing

beryllium electrodes.

ORNL-DWG 713551
- 1

10
I}
o —
£ O —==3 0 S
£ .
B \t E
= —
2 -0 \-\ pd -0 8
3 N s
=E Yo Th
mg: § -20 -20 w3y
S ——
-30 -30
0 0.1 02 03 04 05 08
X (mole fraction BeF,)
E kcal/mole
HLiF
S = —4.7141
X liguidus
—10.5916x — 61.282x* . (3)

The derivative [dufiF/dx] liquidus 2lso is shown in Fig.
10.2; it may be written

E E
[d'uLiF] _ (a‘uLiF>
dx 1 jiquidus ox /r
E
OML iF dT
+ _ . (4)
oT X dx liquidus

Fig. 10.2. Composition dependence of the excess chemical
potential of LiF in BeF,-LiF mixtures at constant temperature
and along the LiF liquidus.

Since (a,ufiF/aT) = _SEEF, the partial molal excess
entropy and enthalpy may be calculated as

E E
(af-"L iF/ax)T - [d”L iF/-dT] liquidus

[dT/dx] iquidus

(5)

3

(6)

~E _ E oE
Hi g =uLir t TSLiF -

These equations, based on isothermal emf measure-
ments and accurate liquidus data, lead to the estimates:

- x=03,T=530°C SE.p=+025eu
fifiF = —1013 cal/mole

E.r = —1215 cal/mole)

— — o cE _ l
x=033,T=450°C  §F..=086eu
IjE
HLiF
E -
LiF

= _-]1051 cal/mole
—1670 cal/mole)
Data at other temperatures and compositions are
being analyzed, and experimental work is in progress on
the development of an Li-Bi alloy electrode to obtain
additional thermodynamic data in this and other
systems.

10.2 EQUILIBRIUM PHASE RELATIONSHIPS -
IN THE SYSTEM LiF-BeF, -CeF;

L. O. Gilpatrick  C. J. Barton

A tentative phase diagram was presented earlier’ for
the system LiF-BeF,-CeF;. Additional data have been
obtained using the differential thermal analysis (DTA)
technique. No major changes in the proposed diagram
have resulted from this new information. Liquidus
temperatures were measured for six compositions rang-
ing from 5 to 20 mole % CeF;, which as a group
indicate that the liquidus contours within this composi-
tion region are some 50°C higher than those proposed
earlier on an a priori basis. Further revision of the
equilibrium phase diagram will be delayed until addi-
tional data become available.

10.3 ELECTRICAL CONDUCTIVITY OF MOLTEN
AND SUPERCOOLED NaF-BeF, (40-60 MOLE %)

G. D. Robbins J. Braunstein

Preliminary electrical conductance measurements of
supercooled molten NaF-BeF, (40-60 mole %) have
been obtained with an all-metal conductance cell,
described previously.® These extend from 418° to
310°C, 95° below the liquidus.®s'® The purpose of
these measurements is to relate the temperature de-
pendence of conductance in glass-forming beryllium
fluoride systems to their glass transition temperatures.

After loading the component starting materials (222.
g), always handled in a protective atmosphere, resist-
ance was monitored as a function of temperature and
measuring frequency for more than two weeks, during
which period the frequency dependence gradually
decreased until the change of resistance at constant
temperature and constant frequency remained ~0.01
Q/day for several days. Data obtained 17 days after

7. L. Q. Gilpatrick, H, Insley, and C. J. Barton, MSR Program
Semiannu, Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 89,

8. G. D. Robbins and J. Braunstein, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, p, 98,

9. R. E. Thoma, ed., Phase Diagrams of Nuclear Reactor
Materials, ORNL-2548, pp. 34135,

10. D. M. Roy, R. Roy, and E. F. Osborn, J. Amer. Ceram.

Soc. 36, 185 (1953).

109

ORNL— DWG 74-7519
12 I T

TEMPERATURE
(°C)

/O 3|O°M-‘

\
\

b 326°

\

| pzase

0 363°

|0 380°

Lo | — 393°
-_______—-O 404°
"] Lo—0—"] 0 HO°
2 W?—T" TP aig>_|

§g}@o——°w
—-—"'——-
1
0 -
0 0.004 0.008 0.012 0.016

Fig. 10.3. Measured resistance of NaF-BeF, (40-60 mole %)
as a function of the square root of the reciprocal frequency

IN/Y

initiation of the experiment are shown in Fig. 10.3,
where the measured resistance is plotted against the
reciprocal square root of frequency and extrapolated to
infinite frequency. These data were taken over a
two-day period in the order indicated on the Arrhenius-
type plot (Fig. 10.4) of the logarithm of the extrap-
olated resistance (R_ ) vs reciprocal temperature.
Calibrations of the all-metal conductance cell have
been carried out at 25°C in 0.01 and 0.001 N aqueous
KC1 solutions and 0.001 N dioxane (79 wt %)—H,O
solutions (each corrected for solvent conductance) to
determine the variations of cell constant with electrode
depth (the major variable), solution volume, dielectric
constant, and measured resistance. From the calibration
data, computer-fitted interpolation functions will be
used for final calculation of the specific conductance.
Based on a glass transition temperature of 117°C (see
following section), the data presented here encompass
the interval 7(°K)/T, = 1.77—1.49. From these prelimi-
nary results, apparent activation energies for specific
ORNL— DWG 71 - 7520

7(°C)
425 400 375 350 325 300
2.4 .
: | i | ) I
[ — 10
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2.2 : yi 9__ 9
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. 4+ — 4
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8 ! / SUPERCOOLED c
12 ' /£ uaub . —
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| / -~ 3
|
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1/
0.8 : ¥ 4 6
s 1,
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./
A
0.4 ;’fl
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0.2 :
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0 L {
.40 1.45 150 155 160 165 170 {75

1000/7 (oK)

Fig. 10.4. Logarithm of extrapolated resistance vs reciprocal
temperature CK).

conductance vary from 13 to 17 kcal/mole over this
108° interval. For the glass-forming pyridinium chlo-
ride—zinc chloride system'' at 58.4 mole % ZnCl, (T,
= 1°C), activation energies vary from 6.5 to 10.2
kcal/mole at T(°K)/Tg = 1.77-1.49, increasing to 20.9
kcal/mole at 1.26.

The molten fluoride data presented here contain a
large (relative) frequency extrapolation at the lower
resistances and should be viewed with appropriate
uncertainty (perhaps £20%). Studies are in progress at
higher BeF, contents, where considerably larger meas-
ured resistances are expected. However, in order to
obtain satisfactory data in the composition region of
the larger glass transition thermal effects, it will be
necessary to modify the cell design (electrode size) so
that a higher resistance is measured, as the low
measured resistance amplifies apparent frequency (elec-
trode surface) effects.!2+13

10.4 GLASS TRANSITION TEMPERATURES
IN THE NaF-BeF, SYSTEM

G. D. Robbins J. Braunstein

The relation of glass transition temperatures to
models of transport in molten salts has been discussed
in an earlier report.'® Experimental techniques and
results were reported for glass transition temperatures,
Tg, obtained by differential thermal analysis
(DTA),' 51 for NaF-BeF, mixtures of composition
50, 55, and 60 mole % BeF,. These studies have been
expanded to encompass what appear to be the limits of
the glass-forming region in this system. Evaluation of
these glass transition temperatures will permit compari-
son with “theoretical glass transition temperatures”
(believed to be near, but slightly below Tg”), as
indicated by the temperature dependence of activation
energies of electrical conductance, currently being
obtained.

Glass transition temperatures were measured in DTA
heating curve experiments with 3-g samples of NaF-
BeF, mixtures which had previously been heated for
~20 hr at 850°C in evacuated, sealed nickel tubes'®
and quenched in ice water. The samples were heated
slowly (2°/min); the transition temperatures observed
in the glasses are listed in Table 10.1, together with data
for the three previously reported mixtures. Column 2 of
Table 10.1 lists the temperatures at which a transition
occurs in the plots of temperature difference between
an Al,O; reference and a fluoride glass sample vs the
temperature of the fluoride sample. These thermal
effects are interpreted as resulting from glass transitions
— the relaxation on heating of the frozen glass into a
supercooled liquid. '

The new transition temperatures reported here were
measured with a calibrated Pt vs Pt—10% Rh thermo-
couple. Two thermal transitions which occur in this

11. A. J, Easteal and C. A. Angell, J. Phys. Chem. 74, 3987
(1970).

12. G. J. Hills and S. Djordjevié, Electrochim. Acta 13, 1721
(1968). :

13. G. Jones and S. M. Christian, J. Amer. Chem. Soc. 57,
272 (1935).

14. G. D. Robbins and J. Braunstein, MSR Program
Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 99.

15. Obtained with the apparatus of L. O. Gilpatrick, whose
technical advice is gratefully acknowledged.

16. L. O. Gilpatrick, S, Cantor, and C. J. Barton, “D.T.A.
Apparatus with Agitation and Sealed Specimens for Molten Salt
Phase Studies,” pp. 85--96 in Thermal Analysis, vol. 1, ed. by
R. F. Schwenker, Jr1., and P. D. Garn, Academic, New York,
1969.
Table 10.1. Glass transitions of NaF-BeF, mixtures

Mole %

BeF, Tg Q) AT (°C)
30
38 130, 127 .5 0.5,0.5
43 123.5, 124 1.1,1.2
46 120.5, 120.5, 120.5, 120.5, 0.7,1.1,1.1,1.1,0.5
118.5
50! 117.5,117 0.8,0.8
55! 116.5,118, 115,115 0.5,0.7,0.7,0.5
60 11’/'.5,1 118,' 117, 116.5,117 0.8,0.4,0.7,0.9,0.8
80 118,119.5 0.3,0.3
90 114.5,111, 113, 113 <0.1
1.00 No effect No effect

temperature range and which have been found useful
for DTA temperature calibration!”? (the crystalline
inversions of potassium nitrate at 128°C and ammo-
nium nitrate at 125°C) appeared 1.5° too low when
measured with this thermocouple. Hence a 1.5° correc-
tion was applied to all data obtained with the thermo-
couple. Redetermination of the previously measured
value of T, for Xp,p, = 0.60 (and adding 1.5°)
resulted in the last three values listed for that composi-
tion in column 2. These values agree well with the
previously reported values obtained with calibrated
Chromel-Alumel thermocouples. Although tempera-
tures in column 2 are reported to the nearest 0.5°C to
avoid round-off error in computation, resolution of the
transition temperatures more closely than 1°C is con-
sidered uncertain.

Column 3 contains an indication of the relative
magnitudes of the glass transitions, reported as the
maximum temperature shift at the transition from the
basedine DTA trace. On increasing the beryllium
fluoride content from the eutectic at 43 mole %
BeF,,'%:'® the thermal effect decreases until at 90
mole % BeF, the break is at the limit of detection. No
such thermal effect could be found on samples of pure
beryllium fluoride. However, since pure BeF, is known
to form a glass easily,?® it appears that the glass-
forming region extends to this end member. A
quenched sample containing 30 mole % BeF, did not

17. 1. Barshad, Amer. Mineral. 37, 667 (1952). .

18. R. E. Thoma, ed., Phase Diagrams of Nuclear Reactor
Materials, ORNL-2548, pp. 34—35 (November 1959).

19. D. M. Roy, R. Roy, and E. F. Osborn, J. Amer. Ceram.
Soc. 36, 185 (1953).

20. H. Rawson, Inorganic Glass-Forming Systems, pp.
235-48, Academic, New York, 1967.

111

ORNL- DWG 71- 7524

~ 122
2 ¥ .

-~ - : ~ \

- 1 - ¥ ~124 DR
?L { |‘ ~ . \\‘|
Ty i \ \ II
o \ 120.5 # —
- - N S J

120.5R:j7'

Y N Y (SO N E H N

80 90 100 MO 420 - 130 140 150 460
#(°C)

100 150 200 250 300 350 400
7 (°C)

Fig. 10.5. Differential thermal analysis traces of an NaF-BeF,
(5446 mole %) mixture.

exhibit a glass transition; instead only the expected
solid-state transitions’®:!® were observed. It thus ap-
pears that the region of glass formation, under the
conditions described above, extends from the mid-30%
beryllium fluoride region to pure BeF,. This may be
compared with the variously reported glass-forming
regions in this system?® of 60—100,2' 40—72,2%, and
45-70%3 mole % BeF,.

In one case we were able to observe the reverse glass
transition on cooling. Several of these DTA traces are
shown in the upper portion of Fig. 10.5 for XBBF2 =

046. Observation of T, while cooling is especially
difficult due to the slow rate of heat loss from the
furnace smearing out the transition. The data shown in
Fig. 10.5 were all obtained in experiments where the
heating rate was 2°/min, but with an average cooling
rate of 0.3°/min.

21. D. M. Roy, R. Roy, and E. F, Osborn, J. Amer. Ceram.
Soc. 33, 85 (1950). .
22. W. Vogel and K. Gerth, Glastech. Ber. 31, 15 (1958).

23. M. Imaoka and S. Mizusawa, J. Ceram. Ass. Jap. 61, 13
(1953), as reported in ref. 9.
ORNL-DWG 71-7522

NO GLASS
140 _f
' i GLASS
130 3 —
g ¥ B
i 120 | %fl $ §
4 & |
1 |
1o :
100
03 04 05 06 Q.7 08 09 1.0
XBer

Fig. 10.6. Glass transition temperatures of NaF-BeF, mix-
tures vs mole fraction BeF,.

The lower portion of Fig 10.5 demonstrates the effect
of heating a quenched sample beyond T, until the
supercooled liquid crystallizes, followed by the ex-
pected crystalline inversion in 2NaF-BeF, (320°), then
by the solidus (340°) and liquidus (~365°) transi-
tions.’®>!? Marked undercooling is observed on the
cooling cycle. Such extended heatings were used to
confirm the interpretation of the glass transitions.

Figure 10.6 shows a plot of the observed values of T,
as a function of mole fraction beryllium fluoride. The
open circles indicate experimental data, with duplicate
values represented as ears on the circles. The average
value of 7, at a given composition is shown as a solid
circle. The vertical bars represent +2° of the average T,
value, the estimated accuracy of these results being no
greater than this.

It appears from the data presented here and the
reported behavior of pure BeF,2° that glasses can be
formed from NaF-BeF, mixtures of composition 38 to
100 mole % beryllium fluoride. This region includes the
primary phase fields of BeF,, NaF-.BeF,, and
2NaF-BeF,.'%:'® For compositions less than 31 mole
% BeF, the precipitating phase is NaF,'®:'® and no
glass formation could be detected in the sample of 30
'mole % BeF,. The central portion of the glass-forming
region exhibits glass transitions at 117 + 2°C from 50
to 80 mole % BeF,, rising with higher NaF content and,
perhaps, decreasing near pure BeF,.

10.5 RAMAN SPECTRA OF BeF, 2 IN MOLTEN
LiF AND NaF TO 686°C

A.S.Quist J.B.Bates G.E. Boyd

Raman spectral studies are particularly well suited for
investigations on BeF,?”, because all of the four normal

112

modes of vibration of the tetrahedral anion are Raman
active, whereas only two modes are infrared active,
Molten Li,BeF, and Na, BeF, were studied to obtain
information on the effect of the nature of the cation on
the vibrational spectrum of BeF,?". In molten
Na, BeF; and Li, BeF; only the minimum fluoride ion
concentration necessary for complete 4-coordination of
beryllium is present; hence, it was of interest to
measure the spectrum of BeF,?  in an excess of F™.
Accordingly, a solution of 17 mole % BeF, in NaF-LiF
also was studied. Additional measurements were made

"on aqueous BeF,* solutions at 25°C to obtain spectra

under conditions of reduced interionic interactions.

Raman spectra of molten Li,BeF, were obtained
from 487 to 640°C.2*A typical spectrum taken at
533°C is shown in Fig. 10.7. Three of the four bands
are clearly visible at frequencies near 390, 550, and 800
cm!. The expected fourth band is located near 260
cm~! but is somewhat obscured in Fig. 10.7 by the
steeply rising background; this band is more apparent in
other spectra, such as in Fig. 10.8 for molten Na, BeF,
at 616°C. The vibrational frequencies observed with
molten Li; BeF, at several temperatures, with Na, BeF,
at 616°C, with the NaF-LiF-BeF, mixture at 645 and
686°C (no difference), with aqueous BeF,?™ containing
excess F~ ion, and with solid Li,BeF,; at 25°C are
listed in Table 10.2.

The frequencies of the BeF,? vibrations observed
from Raman spectral measurements on the molten salts
and in aqueous solution (Table 10.2) are relatively
insensitive (within the experimental uncertainty) to the
cation, the temperature, or the composition. The
polarization (Fig. 10.8) and intensity of the band at
550 + 3 cm™! in the Raman spectra make it logical to
assign it to the totally symmetric stretching mode,
v1(ay), of the tetrahedrally coordinated BeF4%  anion.
This frequency is essentially the same in aqueous
solution as in the melts; only in solid Li, BeF, does it
occur at a different (higher) frequency. The observed
frequency and assignment are consistent with previ-
ously reported values for the Raman spectra of solid
alkali metal tetrafluoroberyllates?> and for aqueous
solutions.?® Although the position of the », band in
the melt does not vary greatly with temperature or
cation, its half-bandwidth decreases considerably when

24. A more extensive discussion of the results of the present
study is contained in a paper by the authors which is to be
published in J. Chem. Phys. (1971).

25. A. . Grigorev et al., Dokl, Akad. Nauk USSR 152, 762
(1963) (English transl.).

26. R. E. Mesmer and C. F. Baes, Jr., Inorg. Chem. 8, 618
(1969). :
113

Table 10.2. Vibrational frequencies observed in the Raman spectra of BeF42 in melts in solid Li, BeF,,
and in aqueous solutions (frequencies in cm ™)

Mol NaF-LiF-BeF, 2.5M
Li, BeF 4 melts olten (53-30-17 (NH,),BeF, Li, BeF,(s), 25°C )
" - Na,BeF,4, . Assignment
487°C  533°C  640°C 616°C mole %), in 2 M NH4E Raman  Infrared
645, 686°C 25°C
255 240 260 265 255 257 va(e)
295
348 360
385 390 " 390 385 380 . 380 377 372 v4(f2)
. 402 405
440 435
475 463 a
. 500
547 550 545 550 552 548 563 vi(@y)
775 775
800 800 800 . 800 800 795 795 805 v3(f3)
' 850 860
2Vibrations assigned to the Li* sublattice moving against the F ~ sublattice.
ORNL-DWG. 70-5696R
! I ] [ | | [
LiyBeFy (y)
533°C :
90cm i SLIT ,
5 x 102 cps i :
I i i ,::"; -fé
T 3 +::_$- H
> : SR ;
E g : :
2 g, :
Ll HEC HIEHT H i
= +
= ‘
B
Y, B
v T JIHHJ ; H
. SR SEhE o At 1
i e e N M NS HERERE I B HRE R R S S e it e
1 I I | | I | | I
1000 900 800 700 600 500 400 300 200

FREQUENCY (cm™)

Fig, 10.7. Raman spectrum of molten LiyBeF, at 533°C.
114

ORNL-DWG. 71-2575R

Hay

NozBeF4
ei6°C B
10 e SLIT
1x10° cps
[=] P
T 4880 A jifitig
. Iy i o
E[F'. RhByak] -
> T
= iR 8 ;i
%) et Pl i
Al TR MNENEY
=z W AN ] L LV
—— ; H _:]f!i LY. A “ |‘ :;: | i [|;
' R e R N PR ER
e SR [ o R
SR AT YT |
TN AL S R i
g i P AR i 53 T H R
| 1 ]

500
FREQUENGY (cm™)

300 100

Fig. 10.8. Raman spectrum of molten Na,BeF,; at 616°C. Incident light polarized (a) perpendicular, (b) parallel to plane

containing the slit and laser beam.

lithium is replaced by sodium. Thus, in molten Li, BeF,
at 582°C the v, half-bandwidth is about 100 ¢cm™,
whereas in molten Na, BeF, at 616°C this value is near
50 cm™". The origin of this effect appears to be cation
dependent since any significant change in anion-anion
interaction would probably have been reflected in a
frequency shift of v, .

The triply degenerate va(f,) and v4(f;) vibrational
modes of tetrahedral ions are both Raman and infrared
active. Infrared studies of solid alkali metal fluoro-
beryllates have led to the assignment of bands near 800
em™! to the »32%:2772% and of the bands near 380
cm™! to v,25°2%2° vyibrations. The Raman spectrum
of solid Li, BeF, exhibits three bands in each of these
regions, consistent with a lowering of the T; symmetry
of the BeF;% ion when it is in a hexagonal unit cell
lattice.>® In molten and aqueous tetrafluoroberyllates,
only single peaks at 800 + 10 and 385 £ 5 cm ™' are
observed for the »; and v, vibrations, respectively,
indicating a symmetric environment for BeF,%  (no
complexing) in these solutions. Half-bandwidths for v,

27. R. D. Peacock and D. W. A. Sharp, J. Chem. Soc., 2762
(1959).

28. 1. LeComte, C. Duval, and C. Wadier, Compt. Rend. 149,
1991 (1959).

29. E. Funck, Ber. Bunsenges. Phys. Chem. 68, 617 (1964).

30. J. H. Burns and E. K., Gordon, Acta Cryst. 20, 135
(1966).

and v, in Na, BeF, melts are approximately 110 and 75
cm ™! respectively.

The double degenerate v,(e) vibration has not been
reported previously. In solid Li, BeF,, peaks at 257 and
295 cm™! are assigned to this vibrational mode. In
melts this band occurs at 255 * 10 cm™!; it was not
detected in aqueous BeF,?  because of the high
background in the spectra. The half-bandwidth was
approximately 90 cm ™' in molten Na, BeF, .

No other bands were observed in the Raman spectra
of the melts. However, the spectrum of solid Li, BeF,
contains additional bands near 440 and 475 cm ™! and
possibly other very weak bands near 125, 165, and 210
cm™'. The latter may be caused by external lattice
modes. The strong infrared bands at 500, 463, and 435
ecm™! correspond to the weak bands observed in the
Raman spectrum at 475 and 440 cm™. We have
assigned these bands to vibrations of the Li* sublattice
against the sublattice of F atoms. Similar bands
observed at 400 to 550 cm ™ in solid Li, CO53! were
found to shift to higher frequencies when °Li was
substituted for 7Li. These bands were assigned to
stretching vibrations of LiO, tetrahedra. The Li* ions
in Li; BeF, also are tetrahedrally coordinated;?° hence,
the assignment of the bands between 430 and 500

.em™ (Table 10.2) to motions of “LiF, tetrahedra”

seems reasonable.

31. P. Tarte, Spectrochim. Acta 20, 238 (1964).
The results of the study reported here indicate that
the BeF,%  anion retains tetrahedral symmetry in its
aqueous solution and in Li; BeF; and Na, BeF,; melts.
Both the frequencies and force constants of “free”
BeF,?” were well characterized from these studies. A
comparison of the frequencies and half-bandwidths for
v, in molten Li,BeF,; and Na,BeF, showed that
cation-anion interactions are important, while effects of
anion-anion interactions were not apparent.

10.6 BUBBLE FORMATION BY IMPINGEMENT
OF A JET STREAM ON A FLUID SURFACE

H. W. Kohn J. R. Tallackson

During the operation of the MSRE, the void fraction
was observed to be somewhat dependent on flow rate
of the primary salt and on the salt level in the pump
bowl. A qualitative investigation of this phenomenon
was made using two spray jets of water and flash
photography. Two mechanisms were postulated for
bubble formation, depending on whether the stream
was continuous or had broken into drops. An air sheath
which surrounds the liquid jet is apparently dragged
beneath the surface of the water, where it breaks up
into bubbles. When individual droplets hit the surface,
holes may be punched which are then closed over,
forming bubbles, a mechanism established by Worthing-
ton®? over 70 years ago. A proportionality of total void
fraction to stream velocity was noted.

A more quantitative study of the phenomenon was
made by an MIT Practice School team®® and will be
issued as a Practice School report. A preliminary
abstract follows.

Bubble formation by impingement of a liquid jet on a fluid
surface was studied. During the Molten-Salt Reactor Experi-
ment helium bubbles were formed in the pump bowl in this
fashion. However, variations in the void fraction were observed
with changes in salt flow rate and salt level in the pump bowl.
The purpose of this study was to determine the dependence of
bubble surface area, mean diameter, and average residence time
on surface tension, viscosity, angle of impingement, flow rate,
tube diameter, and stream length. ’

A two-level half-factorial experimental design was performed
using an air—aqueous-solution system to determine the effect of
each of the six independent variables on the three responses.
Additional data were obtained using an air-water system and
varying only flow rate, tube diameter, and angle of impinge-
ment.

The instantaneous bubble surface area was found to be
proportional to a Weber number based on stream velocity and
tube diameter to the 1.2 power. Bubble mean diameter was

32. A. M. Worthington and R. S. Cole, Phil. Trans. Roy. Soc.
London A189, 137 (1897); A194, 175 (1900).

33. J. T. Boepple and J. B. Cabellon, Bubble Formation by
Impingement of a Liquid Jet on a Fluid Surface, CEPS-X-122.

115

found to be proportional to a Reynolds number based on
stream velocity and tube diameter to the 0.6 power. The
average residence time was found to be proportional to the
product of the stream velocity, tube diameter squared, and the
square of the sine of the angle of impingement to the 1.3
power,

Bubbles were observed forming via the sheath mechanism
proposed by Tallackson. A sheath of air was dragged by the
liquid jet stream beneath the fluid surface and entrained as
bubbles when the flow patterns covered the air with fluid from
the tank. Examination of high-speed photographs taken at 1000
frames per second shows that entrainment of an extended air
sheath and its subsequent breakup into bubbles is a compara-
tively rare occurrence.

10.7 THE SOLUBILITY OF HYDROGEN
IN MOLTEN SALT

J. E. Savolainen  A.P. Malinauskas

In .order to determine, and then to control, the
transport of tritium which is formed in the operation of
a molten-salt reactor, it is desirable to ascertain the
extent to which this species interacts with the molten
salts in the various loops of the reactor. To achieve this
end, we have initiated a program of study of the
solubility of hydrogen and its isotopes in molten salt.

Because no adequate theoretical treatment of gas
solubility -in salt miXtures exists, our experimental
program was formulated on the assumption that the
hydrogen—molten-salt interaction was similar to that
between neon and salt, within a factor of 10 or so. This
assumption necessitated that we obviate or minimize
the effects of two properties of the system to be
investigated. The first .of these, as is apparent from a
consideration of Table 10.3, is the significantly higher
solubility of hydrogen in candidate apparatus materials
compared with the expected solubility in molten salt.

Table 10.3. Hydrogen solubility values at 700°C

and 1 atm saturation pressure

108 X solubility

Solvent (moles H, /cm3 solvent) Reference
Fe 651 1a

Ni 2579 14

Cu 196 _ le

Mo 98 1d

Pt 87 o le

Au 39 2
Li,BeF, (10) 3

1. S. Dushman, Scientific Foundations of Vacuum Tech-
nique, Wiley, New York, 1962, 2d ed. a, p. 522; b, p. 525;¢, p.
528;d,p. 533;e,p. 535.

2. W. Eichenauer and D, Liebscher, Z. Naturforsch, 17a, 355
(1962).

3. Estimate.
The second complicating feature, which is illustrated by
the data of Table 10.4, concerns the transparency of
most metals to hydrogen at elevated temperatures.
(Note in particular the rate of transport through gold as
opposed to the rate of transport through quartz. Until
recently, gold was thought to be impermeable to
hydrogen.) -

Because it appeared unlikely for us to employ a
technique and construction material which would com-
pletely eliminate solubility and transport effects, we
decided to consider what at first seemed to be a totally
illogical approach, viz., to employ Hastelloy N, which,
though compatible with the molten salts of interest in
this work, is both quite soluble and transparent to
hydrogen. This approach, however, led to the apparatus
design which is sketched in Fig. 10.9.

The apparatus itself is essentially that first employed
by Grimes, Smith, and Watson,>* but appropriately
modified for use with hydrogen. In brief, the technique
involves saturating a molten-salt supply in one chamber
of the apparatus and then transferring part of the
saturated liquid to another chamber where the gas is
stripped from the salt and collected for measurement.
Hydrogen losses from the salt in the saturator are
eliminated by doubly containing the salt in this region

34. W. R. Grimes, N. V. Smith, and G. M. Watson, J. Phys.
Chem. 62, 862 (1958).

116

and by filling the annulus between the containers with
hydrogen. Thus, in the saturation procedure, the
hydrogen is bubbled through the salt in the saturator
and then made to flow around the salt container before
being vented.

Once saturation has been achieved, a freeze valve,
located at F in Fig. 10.9,is opened, and part of the salt
is transferred into the stripper chamber. The stripper,
like the saturator section, is doubly contained also, but
in this case the annulus between the containers is

Table 10.4. Rates of transport of hydrogen at 700°C and 1 atm
pressure through a 1-mm-thick plate of 1 cm? cross section

]

Material (mi)(l)es :157;;“) Reference
Fe 83 la

Ni 112 la

Cu 3.16 la

Mo 1.98 la

Pt 7.85 la
Au. 0.71 2
Quartz?3 0.04 " 1b

1. S. Dushman, Scientific Foundations of Vactuum Tech-
nique, Wiley, New York, 1962, 2d ed. a, p. 573; b, p. 497.

2. W, Eichenauer and D. Liebscher, Z. Narurforsch. 17a, 355
(1962).

3. Probably similar to Li,BeF,4 relative to hydrogen trans-
port,

ORNL-DWG 71-1632

—><}— VACUUM
=P
<1 R GAS
GAS  —rd i
BURET
AND I
VACUUM™] ¥
vent |

STRIPPER

SATURATOR

\

F
|

(

]

2!

Fig. 10.9. Sketch of the hydrogen solubility apparatus.
continuously pumped and the gas collected for meas-
urement. The hydrogen which does not permeate the
Hastelloy N container during the transfer and subse-
quent strip operations is stripped directly from the salt
through the use of a xenon purge. Separation of the
hydrogen and xenon is readily accomplished in a trap
cooled with liquid nitrogen. :

Several experiments had been conducted to demon-
strate our ability to recover virtually all of the hydrogen
which is delivered to the stripper compartment. The
data from one of these experiments are presented in
Table 10.5. The sample of hydrogen which was col-
lected at the conclusion was found to contain 29.2%
impurity as determined by mass spectrometric analysis.
Moreover, xenon appeared to be virtually the only
significant impurity (27.4%). On the basis of this
analysis, the experiment yields an effective recovery of
about 95% after 98 min of operation. Since the
amounts recovered from the annulus are entirely free of
xenon, we estimate that 3.57 cm® (STP), or about 65%
of the original amount of hydrogen admitted, was
actually recovered through the Hastelloy N!

Several preliminary experiments have been performed
to determine the solubilities of both helium and
hydrogen in Li;BeF,. To date, the reproducibility of
the data is poor (within a factor of 2). Faulty
temperature control and measurement have been dis-
covered as a possible cause, and this condition is

Table 10.5. Hydrogen recovery experiment

Hydrogen admitted into strip section at 500°C
and at ¢ = 0 min: 5.42 cm?® (STP)

Time Re 'or’1 Amount recovered Rate

(min) & [cm3 (STP)] [cm3 (STP)/min]
28 Annulus 1.41 0.050
41 Stripper 3.27 .
78 Annulus 1.64 0.033
83 Stripper 0.41
98 Annulus 0.52 0.026
Total 7.25 % 0.708 = 5.13

Table 10.6. Solubility of helium and hydrogen in Li, BeF,

G Temperature? Solubility
e O (moles gas/cm? salt-atm)
X 1077
He 630 1.89
H, 620 248
H, 617 - 3.29

2These values may be too large by about 30°C.

117

currently being' corrected. Typical results of the pre-
liminary experimentation are listed in Table 10.6. It
must. be noted that these results may be in significant
error (also within the factor indicated above).

10.8 ENTHALPY OF UF, FROM 298 TO 1400°K
A. S. Dworkin

The unusually low literature values for the enthalpy
and entropy of fusion of UF,; (102 kcal and 7.7 eu
respectively) are suspect because the calorimetric meas-
urements®> from which they are derived were prema-
turely terminated when leaks developed in the con-
tainer. A recent paper>® reports a polymorphic solid
transition in UF,; involving a large amount of heat to
explain the low fusion value. However, neither the
calorimetric study®® nor the many investigations of
UF, at thislaboratory, including a recent DTA study by
Gilpatrick,®” have shown any indication of such a tran-
sition. We have, therefore, remeasured the enthalpy of
UF, from room temperature up into the liquid phase in
an attempt to clarify this matter.

Our results give no indication of a transition. What
was taken to be a transition was possibly caused by an
impurity present in the material used in ref. 36. The
following equations represent our measured enthalpy
data for UF, in cal/mole:

Hyp — Hyg =—9650 +29.53T+1.15X 107°T?
+2.21 X 10% 77! (298--1309°K),

AHg, o = 11,230 cal/mole
AStusion = 8-6 eu (1309°K) ,
Hp— Hygg =—9420 + 39.57T (1309—1400°K) .

Our entropy of fusion is somewhat but not signifi-
cantly larger than that in the previous measurement.?3
It is still quite low for a compound with five atoms.
Table 10.7 shows a comparison of the entropies for the
isostructural (monoclinic) compounds ZrF,, UF,, and
ThF,. The entropy of melting for ZrF, of 12.7 eu is

35. E. G. King and A. U. Christensen, U.S. Bur. Mines, Rep.
Invest. 5909 (1961).

36. L. A. Khripin, Y. V. Gogarinski, and L. A. Lukyanova,
Izv. Sib. Otd. Akad. Nauk SSSR, Ser. Khim. Nauk 1(3), 14
(1965).

37. L. O. Gilpatrick, MSR Program Semiannu. Progr. Rep.
Feb. 28, 1970, ORNL-4548, p. 148.
Table 10.7. Comparison of entropies of ZrF,;, UF,, and ThF,

Tyr  ASy 8% 98 S°1400—5°208 51400
(°K) (eu/mole) (eu/mole) (eu/mole) {eu/mole)
ZiF, 1205 12.72 25.0% 59.7¢ 84.7
UF, 1309 8.6 36.39 56.0 92.3
ThF, 1383 ? 34.0¢ ? ?
aGee ref. 37.
bSee ref. 38.
CSee ref, 39,
dgee ref. 40.
€See ref, 41,

considerably larger than the unusually low value of 8.6
eu for UF,. However, the absolute entropy of UF, at
1400°K is larger than that for ZrF, by about 7 eu. This
is approximately what would be expected considering
the differences in size, mass, and magnetic character-
istics of the cations.

An attempt was also made to measure the enthalpy of
fusion of ThF,. However, experimental difficulties
were encountered due to the high temperatures in-
volved, and the measurements were terminated to
prevent irreparable damage to the calorimetric equip-
ment. Enough qualitative data were collected to in-
dicate that, as expected, the fusion behavior of ThF,
approximated that of UF, rather than that of ZrF, .

10.9 ABSORPTION SPECTROSCOPY OF MOLTEN
FLUORIDES: THE DISPROPORTIONATION
EQUILIBRIUM OF UF; SOLUTIONS

L. M. Toth

An investigation of chemical factors which affect the
stability of uranium trifluoride as a dilute species in
molten fluoride mixtures was initiated sometime ago
utilizing absorption spectroscopic techniques. In this
research, molten fluoride solutions of dilute UFj3,
contained in graphite, are examined to determine the
effects caused by changes in the concentration of
fluoride ions in the solvent. That such changes affect
the coordination number of fluoride ions around U(IV)
was demonstrated previously®® by results which
showed that the fluorine coordination number de-
creases from 8 to 7 as the fluoride ion concentration in
the solvent decreases. Such results allow measurements
of the spectrum of U(IV) in a molten solution to be

used to estimate the F~ concentration of that solution

38. L. M. Toth, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL4622, pp. 103—5.

118

relative to others and to facilitate studies of equilibria
which are affected by these concentrations.

Although the stability of UF; in molten fluorides has
been the subject of previous investigations,>® none of
the experiments employed graphite as a container
material. From thermodynamic data it was inferred that
solutions in which up to 60% of an initial 1 mole % of
UF, is converted to UF; should be stable in the
presence of graphite. It was not the purpose of the
current work to reinvestigate the stability of UF;, but
rather to develop laboratory procedures for spectro-
photometric studies of UF; molten fluoride solutions
which employ graphite spectrophotometric cells.
Anomalous behavior of UF; was observed, however,
when it became apparent that the stability of UF; in

~the molten fluoride solvent mixtures was less than

predicted originally and, in addition, when results
showed that the stability of UF; was more sensitive to
the acid-base properties of the solvents than was
anticipated. These observations motivated the investiga-
tion described below, which is still in progress.

The equilibrium

2C(graph1te) + 4UF3 = 3UF4 + UC2 (1)

has been identified as the disproportionation reaction
of dilute UF; solutions in graphite by two approaches:
(1) the forward reaction in which UF; is allowed to -
disproportionate and reach equilibrium and (2) the
back reaction in which UF, and UC, are equilibrated,
forming the same approximate (see discussion follow-
ing) U(LII)/U(IV) ratio achieved by the forward reac-
tion. These measurements were performed in the
diamond-windowed spectrophotometric cell*® at tem-
peratures up to 750°C in a helium atmosphere with
samples of <1 cc volume. The LiF-BeF, solvent system
has been used to effect changes in F~ activity by
varying the concentrations of solvent components from
48-52 mole % LiF-BeF, to 66-34 mole %. The
temperature effect on the equilibrium has also been

studied. Uranium(lIT} and uranium(IV) concentrations

were measured spectrophotometrically, taking account
of absorption coefficient changes due to both solvent
composition and temperature. Uranium dicarbide, as a
product of reaction (1), was identified by Debye-

39. (@) G. Long and F. F. Blankenship, The Stability of UFj;.
Part 11, Stability in Molten Fluoride Solution, ORNL-TM-2065
(November 1969). (b) F. F. Blankenship et al., unpublished
data on “The Stability of UF; in Molten Fluoride.”

40. L. M. Toth, J. P. Young, and G. P. Smith, Anal. Chem.
41, 683 (1969),
Scherrer patterns of scrapings®! taken from the walls of
the graphite cells.

119

To study the back reaction, pyrolytic-graphite-coated .

UC, microspheres®? were crushed and used as the
reductant for UF,
analysis of the microspheres revealed the presence of
graphite, UC,, and approximately 10% UC. The effect
of this UC phase is estimated to produce UF; at a
concentration which represents an upper limit for the
disproportionation equilibrium.*® The UF; concentra-
tion produced by this mixed reductant was found to be
somewhat greater than that achieved by the forward
reaction.

Although reaction (1) is identified as a dispropor-
tionation reaction, the overall reaction rate which is
measured experimentally need not necessarily be sec-
ond order as is found for elementary homogeneous
disproportionation mechanisms. The rate for reaction

(1) is expected to be typically first order, being con-
trolled by diffusion to the surface of the graphite. This
reaction is expected also to be surface-to-volume depend-
ent so that caution should be exercised in making
comparison of rates between systems of different sizes.
The only meaningful data in support of the dispropor-
tionation mechanism are then theresults of the tests of
equilibrium performed in the two experimental ap-
proaches, together with the overall material balance,
which showed that in solution 4 moles of UF; produce
3 moles of UF,. It is impossible to account quantita-
tively for the insoluble components of the equilibrium
under the conditions of these experiments.

The immediate results of this work show that not
60% (approximately) of the total 1 mole % uranium is
stable in solution but instead approximately 6%. This
number is, however, very dependent upon temperature
and fluoride ion concentration, approximately doubling
for F~-deficient solutions at high temperatures (700 to
750°C) and decreasing to much less than 6% for F “-rich
solutions at lower temperatures.

Data which will more clearly describe this stability are
currently being evaluated and will be given later in
detail. It is interesting to note that the MSRE solution
composition, LiF-BeF, -ZrF, (64-29-5 mole %), behaves

41. Grateful acknowledgment is made to H. L. Yakel, Metals
and Ceramics Division, and H. W. Dunn, Analytical Chemlstry
Division, for these determinations,

42, Grateful acknowledgment is made to J. L. Scott and W,
M. Proaps for supplying this material.

43. W. R. Grimes, Chemical Research and Development for
Molten-Salt Breeder Reactors, ORNL-TM-1853, pp. 37-39
(June 1963). '

in solution. A Debye-Scherrer

as an F-deficient solution,**® whereas the proposed

MSBR composition, LiF-BeF,-ThF, (72-16-12 mole
%), appears to be quite rich in F~ as determined by
UF, coordination number determination — much more
than would be expected for a solution with 12 mole %
ThF,. '

10.10 THE OXIDE CHEMISTRY OF Pa**
IN MSBR FUEL SOLVENT SALT

C. E. Bamberger R.G.Ross C.F. Baes, JIr.

While we have demonstrated that the addition of 0%
to a fluoride melt containing Pa** at 563°C precipitates
Pa, 05 quantitatively,*s it was also of interest to study
the behavior of Pa** in the presence of variable amounts
of 0. _

The experimental procedure consisted in stepwise
additions of ThQ, microspheres to 200 g of LiF-BeF,-
ThF, (72-16-12 mole %) containing 1800 ppm protac-
tinium, maintained in the tetravalent state by a hydro-
gen atmosphere. The protactinium solution was
prepared by extensively hydrofluorinating, at 700°C, a
mixture of the salt with Pa,Os and ThOQ,. This
treatment was followed by a lowering of the tempera-
ture to 560°C, and since this produced no detectable
precipitation of Pa, Qs, it was concluded that the oxide
removal was essentially complete. The protactinium-
containing melt was then extensively reduced by
sparging with hydrogen at two temperatures; because of
the negative temperature coefficients for the reduction
by hydrogen of metallic oxidizing impurities*® (i.e.,
Ni**, Fe?*, Cr**), they were removed by the hydrogen
treatment at the higher temperature, 720°C. The
subsequent hydrogen reduction was carried out at a
lower temperature, 560°C, where the protactinium was
presumed to be more easily reduced to Pa** (see Fig.
10.12). -

The equilibration with omde was conducted by
stirring both phases under a flowing hydrogen atmo-
sphere, while monitoring the content of HF in the
effluent gas. This kept the system sufficiently reducing,
while the rate of HF evolved provided an indication of
whether the system was accidentally contaminated with
oxidizing impurities (i.e., air, moisture).

44. Private communication with D. L. Manning, Analytical
Chemistry, is gratefully acknowledged.

45. R. G. Ross, C. E. Bamberger, and C. F. Baes, Jr., MSR
Program Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p.
92.

46. C. F. Baes, Jr., Symposium on Reprocessing of Nuclear
Fuels, Nuclear Metallurgy, vol. 15, CONF-690801 (1969).
ORNL-DWG 71— 7523

663 °C
5 S
£ ~ -
—— ~ 'I-.-
o ~A
oy patt AT 569°C
™~ -
S \‘~‘§
5 T
3
<
& L Pa”* AT 668°C
\
S \
2 AN
g b,
£ o0 N
= \
g - \o \
PeS* AT 563°C N
\1
L Y
\\
\ .
0
0 10 20 30

millimoles of 0%~ added /kg of melt

Fig. 10.10. Precipitation of Pa** and Pa® oxides from
LiF-BeF;-ThF4 (72-16-12 mole %) by addition of ThO,. The
data for Pa”" are from ref. 45; the solid line corresponds to the
quantitative precipitation of PayOs. '

This experiment is still in progress. The data presently
availablé are shown in Fig. 10.10 as millimoles of Pa**
in solution per kilogram of melt vs millimoles of 0%
added per kilogram of melt. For comparison we have
also plotted the behavior of Pa®* reported earlier®® at
nearly the same temperature. It can be seen that the
behavior of the two oxidation states of protactinium is
quite different; the Pa*'-containing phase appears con-
siderably more soluble than Pa, Og. Although we have
not yet isolated and analyzed the solid Pa**-containing
phase, preliminary material balance calculations would
indicate that it is a (Pa-Th)O, solid solution. A less
likely possibility is that it is an oxyfluoride, PaOF,.
Work continues in order to characterize the solid phase
and to study the effect of temperature and of the
composition of the solid phase on the distribution of
Pa** between the fluoride melt and the oxide phase.

10.11 THE REDOX POTENTIAL OF
PROTACTINIUM IN MSBR FUEL SOLVENT SALT

R.G.Ross C.E.Bamberger C.F. Baes, Jr.

It has been previously reported*®” that the precipita-
tion of protactinium oxide from MSBR fuel salt holds

120

promise for a very efficient and simple way of removing
protactinium from the fuel stream of an MSBR. Our
previous work indicated that under oxidizing conditions
pure or nearly pure Pa,O;s precipitates by addition of
O*" to the molten fluoride mixture and that Pa,Os has
a considerably lower solubility than UQ,.

In order to select the best conditions for removing
protactinium from MSBR fuel salt by the oxidative
precipitation of protactinium as Pa,Os, values for the
redox potential of the couple Pa**/Pa®* are required.
We have previously reported limits estimated from
direct measurements of the equilibrium

PaFs(d) + 1/2H,(g) = PaF,(d) + HF(g) ,

(1)
Ql :Xpa4+.PHF/(XPas+.PH21/2) ,

where d, g, and X; denote, respectively, dissolved and
gaseous species and mole fractions. In these measure-
ments the amounts of Pa** and Pa*" present during the
hydrogen reduction were estimated by a material
balance based on the amount of HF evolved. The
reduction of impurities such as Ni?* and Fe®* (which
also produced HF) occurred simultaneously, however,
necessitating corrections which could be estimated only
roughly. Hence only limits for the values of Q, could
be obtained.

In order to overcome the above-mentioned difficulties
we chose to measure the equilibrium quotient of

1/2Pa, O5(c) + 5/4ThF,4(d) + 1/2H, (g)

= PaF4(d) + 5/4ThO,(¢) + HF(g) , (2)
Q2 =Xp,¥, 'PHF/XThF45/4 Py, v
and to combine Q. with @5, measured previously,
1/2Pa,; Os(c) + 5/4ThF,(d)
= PaF;5(d) + 5/4ThO,(c), (3)

Qs = XPaFS/XThF4S/4 -

Reaction (2) has the following advantages:

1. the concentration of PaFs is held constant by the
presence of the saturating phases Pa, O and ThO,,
or more likely, by Pa, O5 and a dilute solid solution

of Pa** in ThO, ;

47. R. G. Ross, C. E. Bamberger, and C. F. Baes, Jr., MSR
Program Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p.
92.
2. the concentration of PaF; is known from Q5

3. the concentration of PaF, can be easily estimated
by measuring the concentration of total protactin-
ium dissolved in the melt and subtracting the small
concentration of PaF calculated from Q5

4. as long as the partial pressure of HF is measured
under equilibrium conditions, it is irrelevant whether
the HF is generated by the reduction of PaFg, by
reducible metallic ions, or by both. '

The experimental procedure consisted in equilibrating
under a hydrogen atmosphere the solid phases, Pa; Os
and ThO,, with the melt, LiF-BeF,-ThF,; (72-16-12
mole %), in a stirred vessel assembled inside a glove box
suited for work with alpha-active elements. Since the
reaction studied (2) seemed to involve slow kinetics, the
hydrogen was not flowed through the system, as in a
transpiration experiment; rather its pressure was held
constant, and only the hydrogen lost by diffusion
through the walls of the container was replenished.

At various intervals, approximately one-fourth of the
gas was withdrawn into an evacuated copper vessel and
then flushed from it into a small titration vessel with
argon. Since the HF pressures measured were in the
order of 1072 to 10™* atm, the titrations were carried
out with care in order to avoid CO, contamination and
to obtain sharp end points. The melt was also sampled
at the time of the gas sampling. Since hydrogen was
present, we used a special shroud to surround the
Teflon seal of the salt sample tube with a flowing
atmosphere of argon. This prevented either air from
leaking into the equilibration container or pure H,
from leaking out into the glove box.

All the results obtained are presented in Fig. 10.11.
On the left side of this figure the PaF, concentration of
the melt is shown as a function of time in each of two
runs at each temperature. The amount of PaF, was
increased between runs by protonged hydrogen reduc-
tion. At the two higher temperatures the product
Xpar, Pup. plotted on the right, remained constant
within the scatter of the data in spite of the changes in
the PaF, concentration; that is, the partial pressure of
HF was found to be inversely proportional to the PaF,
concentration in accord with equilibrium (2).

At the lowest temperature (584°C), however, an
obvious change in behavior had occurred in that the

PaF, concentrations in the two runs tend to approach a.

single value while giving an Xp,p, Py product much
lower -than would be expected by extrapolation of the
otherresults in Fig. 10.11. It is presently believed that at
584°C a major fraction of the Pa** was in the form of a
(Pa-Th)O, solid solution, which should have been

121

3 ORNL-DWG 74-7524 -6
10° = = = —= 10
= 732°C =
O~ = = —]
l '-_._-‘-o — = —
Lo Lo
102 — — :d:' ——-&)—__.—E = 10—7
L PR NS = R _]
10’ 1078
10> = = = 107®
= 659°C =
- = — w
: H-.\‘A\_d_“— = = a ] o
2 10% = Bk o S ) NS b s e i N [T,
— — - o | o
E = = = f— o.
o s oy L = - —_ x
=Y — e —daed b — — —]
— — —
103' 1078
10° = o = — 107~
= 5_84E =
10% = s = L S T
— = Ea+% =
Ehfl_‘ p— E p—
l— "u--..___. - = -
10‘ - -8- 10— T [|]" 10—9
0 60 120 180 O - 60 120 180
TIME {hr} TIME {hr)
Fig. 10.11. Variation of Pa* concentration and of

Xpaa+* Py 2s afunction of equilibration time for two different
initial concentrations of Pa® at each temperature.

appreciably richer in Pa*" than the solid solutions
formed at the higher temperatures. With most of the
Pa** present in a (Pa-Th)O, solid solution, even a dilute
one, the composition of this phase would have re-
mained relatively fixed during the last two runs at the
lowest temperature. This, in turn, would tend to fix the
value of Xp,p, by the equilibrium

PaF, (d) + ThO,(ss) = PaO, (ss) + ThF,(d) ,
Xpao, XThF, )

XTh02 Q4

XPaF4 -

The relatively constant level of Xp g, thus produced
apparently was lower than that required by equilibrium
(2), and hence it appears that Pa, O5 was dissolving, but
at too slow a rate, owing to its low solubility, to permit
equilibrium with respect to reaction (2).

At 659 and 732°C, however, it appears that equilib-
rium with respect to reaction (2) was maintained,
presumably because the Pa*®* present in ThO, solid
solution was but a small fraction of the amount of PaF,
in the fluoride phase. The composition of the solid-
solution phase therefore could be controlled by the
Pa** in the melt and, hence, by reaction (2).
ORNL-DWG 71- 7525

0.1
/l'
0.05
L7
Q
<
L 0.02
X
e
(=]
[+ N
Ny /
LL¢ .
2 0.0f /
x 7 /
H]
o //
0.005 //
/
0.002 :
0.95 1.00 1.05 1.10 145
1000/ (o)

Fig. 10.12. Variation of 0 = (XpaF,/XpaFs)PHE/PH, /)
as a function of 1/7(C K).

Figure 10.12 shows a plot of log Q, as a function of
1/T(°K) estimated for the two higher temperatures
only. This plot indicates that molten fluoride solutions
of PakF, are oxidized more easily by HF at the higher
temperatures and hence that solutions of PaF are
reduced more easily by hydrogen at the lower tempera-
tures. Due to the importance that the redox potential
of Pa*/Pa®* has in an efficient protactinium-removal
process as well as in the control of MSBR fuel
chemistry, the magnitude of Q, and its temperature
dependence will be investigated by means of other
reactions.

122

10.12 THE CRYSTAL STRUCTURES |
OF COMPLEX FLUORIDES

G. D. Brunton

The development of automated techniques for rapid
collection of diffraction data from single crystals has
greatly improved the precision of crystallographic meas-
urements. As a result, studies of the crystal structures
of the complex compounds encountered in molten salt
reactor technology have advanced to the point that the
stoichiometry of such crystal phases can often be
determined independently and unequivocally by single
crystal diffraction methods. Application of this capa-
bility to MSR R&D programs provides a powerful
means to advance understanding of chemical phe-
nomena which involve liquid-solid interactions.

A brief review of the results obtained in recent studies
of the structures of complex fluorides is included here
to illustrate some details of the coordination chemistry
which typifies many of the complex compounds which
are formed by the heavy metal fluorides.

10.12.1 The Crystal Structure of CsUF, s*®

The complex fluoride CsUgF,s crystallizes with
space group P63/mmc; a, = 8.2424(4) A, c,
16.412(2) A, Z =2, and the calculated density = 7.0013 .
g/cc. The U* ion is coordinated by 9 F~ ions at
distances of 2.267(8) to 2.54(4) A and the Cs* ion by
12 F~ at distances of 3.12(2) and 3.45(4) A.

10.12.2 The Crystal Structure of a-KThF, 43

Crystals of a-KThgF,s are hexagonal-rhombohedral
R3m with hexagonal axes ao = 8.313(2) and co =
25.262 A, Z = 3, and the calculated density = 6.281
g/cc. The structural units of a-KTheF,s (Fig. 10.13)
are identical to those of CsUgF,s; two rings of Th-F
polyhedra surrounding a central K ion. The structural
units are stacked ABCABC along ¢, in @-KThgF, 5 and
ABAB in CsUgF, 5 (Figs. 10.14 and 10.15).

10.12.3 The Crystal Structure of Li, MoF*?

Crystals of Li,MoF, are tetragonal P4,2,2 with ay =
4.6863(7) and c, 9.191(2) A, Z = 2, and the
calculated density = 3.687 g/cc. The Li* and Mo** ions
are octahedrally coordinated. The Li-F distances range
from 2.017(2) to 2.102(7) A and the Mo-F distances
range from 1.927(2) to 1.945(2) A. The MoF4? ion is
coordinated by ten Li* (Fig. 10.16).

48. Acta Crystallographica (in press, 1971).
49. J. Crystal Growth (in press, 1971).
123

ORNL-DWG 71-1795

Fig. 10.13. KThgF, 5 structural unit.

ORNL-DWG 71-2330

Fig. 10.14. KTh¢F, s structural stacking.

ORNL-DWG 71-2331

Fig. 10.15. CsUgF, 5 structural stacking.
124

ORNL-DWG 71-1686

Fig. 10.16. Li,MoFs.

ORNL—DWG 70-13164

Fig. 10.17. RbTh;F ;.

10.12.4 The crystal Structure of RbTh, F, ;%8 of these ions has 9 nearest-neighbor fluorine jons at the

: : . corners of capped trigonal prisms. The Th-F distances
The complex fluoride RbThFy; crystallizes with 00 trom 2 32(3) to 2.48(2) A. The rubidium jon has

space group P2,ma; aq =8.6490(5), by = 8‘17§(2)’ and nearest-neighbor fluorine jons in the range 2.79(3)

Co = 74453(4) A. There are two formula weights per t0 3.37(2) A (Fig. 10.17). The final discrepancy index is

unit cell, and the calculated density is 6.488 g/cc. There 0.0710 for 2134 Ag Ka reflections >a.

are two thorium ions in the asymmetrical unit, and each
Fig. 10.18.

10.12.5 The Crystal Structure of RbsZr, F, *®

The. complex fluoride RbsZr,F,, crystallizes with
space group P2,; 2o = 11.520(5), by = 11.222(5), and
co = 7.868(2) A; cos § = —0.1445(3). The calculated
density = 3.930 gfcc, and Z = 2. The structure was
solved by a tangent formula procedure and refined by
Fourier and least-squares methods to a final discrepancy
index (R) of 0.0508 for 1376 observed Cu Ka;
reflections. The F~ coordination polyhedra are dif-
ferent for each of the four crystallographically inde-
pendent Zr ions. Seven F~ ions are nearest neighbors to
Zr(1) at the corners of a pentagonal bipyramid, and the
interatomic distances are 1.99(2) to 2.20(2) A. The
second Zr ion has eight nearest-neighbor F~ [1.98(2) to
2.24(2) A] at the corners of an irregular antiprism. The
third Zr ion_is octahedrally coordinated by six F~ at
distances of 1.90(2) to 2.10(2) A. The fourth Zr ion has
seven F~ nearest neighbors at the corners of an irregular
antiprism with one corner missing. The interatomic
distances Zr(4)-F are 1.95(2) to 2.19(2) A. The Rb-F
distances range from 2.68(2) to 3.21(3) A. The struc-
ture is composed of cross-linked chains of Zr-F poly-
hedra. The chains are connected by Zr(2)-Zr(4) edge-
sharing polyhedra, and the space between the polyhedra
is filled with Rb ions (Fig. 10.18).

10.12.6 Discussion

With the elucidation of these structures it becomes
increasingly evident that the complex fluorides have a
number of common features from which some generali-
zations can be made. A few examples are noted, as

125

ORNL-DWG 70-14373

s
| D

follows. The heavy-metal-to-fluorine bond distances are

generally found to range from 1.9 to 2.5 A, dependent
approximately on the charge/size values for the cation;
the fluorine coordination numbers for the heavy metals
are consistently large, one almost never found to be less
than 6, and may be as large as 9. Coordination numbers
are not infrequently variable within a single structure.
Although such a phenomenon does not necessarily
imply the existence of variable quasi-anionic character
among structurally distinguishable cation-anion clusters,
it does suggest ‘that success in the development of
theoretical models of liquid-state behavior may be
dependent upon cognizance of some of the charac-
teristic details of such structures as are cited here. '

10.13 NONCRYSTALLINE BeF, AT 25°C:
STRUCTURE AND VIBRATIONAL MOTION

A. H. Narten J. B. Bates

A program of investigation of molten salts by x-ray
and neutron diffraction was resumed in support of
spectrographic researches with these materials. Prelimi-
nary to the study of complex systems, initial studies .
were devoted to glasses of BeF, and SiO,. These
compounds form highly viscous melts near the fusion
point, and the structure of the melts, though of
considerable interest, is poorly understood. Upon
quenching, the melts solidify into glasses, and the
structure of these is thought to be similar to that of the
liquid phases. A study of noncrystalline BeF, and SiO,
at room temperature by x-ray diffraction has yielded
new insight into the average configuration and motion
of atoms in these phases.
The polymorphism of crystalline SiO, and, to a lesser
- extent, of crystalline BeF, is well known. All crystal-
line modifications of these compounds contain SiO,
and BeF, tetrahedra joined at the vertices so that an
oxygen or fluorine atom is common to two tetrahedra.
The crystalline modifications differ in the packing of
these tetrahedra.

The diffraction pattern of noncrystalline SiQ, and
BeF, is characterized by the absence of Bragg peaks,
and hence no long-range order is present in these
glasses. Atom pair correlation functions derived from
the diffraction data show very sharp peaks charac-
teristic of near-neighbor interactions, but all positional
correlation is lost about 8 A away from any starting
point. The average distances between near-neighbor
atoms in vitreous Si0, and BeF, are equal to those
found in the high-temperature crystalline modifications
of both compounds (8-quartz). As in the crystalline
forms, deviations from ideal tetrahedral geometry are
very small in the noncrystalline phases at room temper-
ature. In contrast, the average coordination numbers are
slightly lower in the glasses. Intensity and correlation
functions calculated for a model based on the B-quartz
structure are in quantitative agreement with those
derived from the diffraction data. The slightly lower
density and coordination number found for the non-
crystalline phases are described by the model in terms
of random vacancies in the tetrahedral networks (12%
for SiO, and 6% for BeF,), and these defects result in
the progressive loss of all positional correlation with
separation.

With a structural model thus available, it was possible
to calculate the vibrational modes to be expected.
Based on the f-quartz structure, the k = o (optical)
crystal vibrations of SiO, and BeF, were calculated
using an eight-parameter modified valence force field.
Both the optical selection rules and calculated frequen-
cies were in good agreement with the vibrational spectra
of Si0, and BeF, measured in this laboratory and with
spectra reported in the open literature. In particular,
the theory predicts that the only totally symmetric
vibration of the -quartz structure is a bending mode of
the 0-8i-O (in SiO, ) or F-Be-F (in BeF,) valence angles.
The strongest bands observed in the Raman spectra of
Si0, and BeF, were observed at 437 and about 280
cm ™!, respectively, and were therefore assigned as due
to the totally symmetric bending modes in each case.
The vibrational spectra measured on the glasses agreed
with the calculations both in frequency and in the
selection rules. This agreement strengthens the con-
fidence which can be placed in the structure deduced
from diffraction. Neither method can give an unambig-

126

uous structure, and it is therefore useful to employ
them together.

10.14 RELATIONSHIP BETWEEN ENTROPY AND
SONIC VELOCITY IN MOLTEN SALTS

'S. Cantor

A correlation of the entropies of crystalline and liquid
ionic compounds with molar weight and with molar
volume was reported previously.®® The correlating
equation is essentially a method for estimating a
quantity € which is related to the distribution of
acoustical vibrations in a crystal. For crystals, 8 is
usually determined from calorimetric data, but it can
also be obtained from elastic-constant data; @ is related
to the elastic constants through the mean velocity of
sound:

(1)

where C is the velocity of sound; V is the molar volume;
n is the number of atoms in the chemical formula of the
compound; and k, h, and N are, respectively, Boltz-
mann’s, Planck’s, and Avogadro’s constants.

For a molten salt, we may use entropy, either
predicted or measured, to provide values of § by means
of an equation derived from the Debye theory of lattice

vibrations:
7]
T >

where Si(liq) is the entropy of the molten salt at
absolute temperature 7 and R is the gas constant
[1.987 cal mole™ (°K™')]. The 6 calculated from Eq.
(2), when substituted into Eq. (1), permits the calcula-
tion of the sonic velocity of the molten salt. The sonic
velocity, thus calculated, is compared with experi-
mental values of several binary salts in Table 10.8.

The agreement between calculated and experimental
sonic velocity is quite good. Thus it would seem that a
“liquid” @ derived from entropy is an effective method
for predicting sonic wvelocities of binary molten salts.
The deeper significance of the good agreement in Table
10.7 may be that “solid-like” models of molten salts
lead to useful relationships between physical properties;
when these models succeed they ought to be refined
and applied to other physical properties:

Sr(liq) = 3nR (g — 1n (2)

50. Stanley Cantor, MSR Program Semiannu. Progr. Rep.
Aug. 31, 1970, ORNL-4622, pp. 95-98.
Table 10.8. Velocities of sound in molten salts (km/sec)

1000°K 1200°K

Calculated Experimental? Calculated Experimental?

LiF 2.57

(2.72) 2.48 2.45
LiCl 1.88 1.93 1.83 1.77
NaF _ 1.78% 2.05%
NaCl 1.45 (1.81) 1.38 1.63
NaBr 1.15 (1.34) 1.12 1.21
KF 1.59 (1.94) 1.53 1.76
KI 0.802 1.09 0.787 1.02
CsCl 0.908 1.10 0.844 0.970
CsBr 0.731 0.955 0.680 0.826
CaCl, 1.71 (2.10)

SrBr, 1.19 1.54

BaCl, 1.44¢ 1,72¢

2The experimental data were reported by J. O’M. Bockris et
al. and M. Blanc et al. and are summarized in G. J. Janz, Molten
Salts Handbook, p. 251, Academic, New York, 1967. Paren-
theses indicate experlmental data extrapolated into super-
cooled-liquid region,

bat 1300°K.

CAt 1235°K.

What value of @ predicts the sonic velocity of more
complex molten compounds such as nitrates? When
molar entropies are substituted into eq. (2), the
calculated sonic velocities are too low by about a factor
of 10. However, in these compounds the total molar
entropy includes contributions from internal degrees of
freedom — interatomic vibrations and internal rota-
tional modes of the complex ion. When we subtract
these entropy contributions, representing them as ‘“‘gas-
like” behavior (i.e., unhindered rotational and internal
vibrational entropies of NO;~ calculated from the
standard partition functions of an ideal gas), we obtain
a residual entropy associated only with “lattice” contri-
butions. Correspondingly, the complex ion is countéd
as one particle; for instance in the case of KNO;,n =2
in Egs. (1) and (2). This “lattice” entropy, when
substituted into Eq. (1), yields sonic velocities in
excellent agreement with experimental values:

Velocity of sound at 700°K (km/sec) .

Calculated Experimental
NaNOs(liq) 1.76 1.67
KNO;(lig) 1.62 1.76
AgNO;(liq) 1.37 1.42

These results do not prové that nitrate ions execute
unhindered rotation in the melt. What has been shown
is that the propagation velocity of sound waves in
molten salts containing nitrate ions appears to depend

127

only on “lattice” vibrations. The entropy associated
with the “lattice’ vibrations is approximated by sub-
tracting entropy contributions assumed to be caused by
free rotation and internal vibrations of the complex
ions.

10.15 A REFERENCE ELECTRODE SYSTEM
FOR USE IN FLUORIDE MELTS

H. R. Bronstein

Further study of the galvanic cell comprised of the
half-cell Be/LiF, BeF, (67-33 mole %) and the Ni-NiF,
reference electrode assembly, as illustrated in Fig.
10.19,51733 led to findings believed to be of consider-
able interest and significance.

Deterioration of the initial cell voltage with time was
found to be caused by the depletion of the small NiF,
content of the reference half-cell, as evidenced by its
conversion into an equivalent amount of finely dis-
persed nickel metal. Since a static atmosphere of highly
purified argon or helium gas was maintained in the
quite limited volume of the apparatus, an impurity in
this small amount of cover gas could not be held

_responsible for the reduction of the NiF, . This conclu-

sion was verified by subsequent experiments which
established the actual mechanism involved. )
This phenomenon continued to occur with all ex-
ternal measuring circuits disconnected, as demonstrated
by measuring the potential at periodic intervals. Also,
beryllium metal has no measurable vapor pressure at
600°C, the temperature of the experiments. The con-
clusion had to be drawn that an internal “short” in the
cell supplied the electrons for NiF, +2¢~ = Ni® + 2F .
Obviously, the beryllium metal, the only material
present capable of causing the reduction, must have
reached the NiF, in some manner. Therefore, one must
not only invoke the transport of -beryllium metal to the
reference electrode cell through the salt melt, either by
dissolution or by the.so-called “chunking effect,”4:%%
but also an electronic component in the conductance of
the single-crystal LaF;. The beryllium arriving at or in

51. H. R. Bronstein and D. L. Manning, MSR Program
Monthly Report for December 1969 and January 1970,
MSR-70-10, p. 28 (internal memorandum).

52. D, L. Manning and H. R. Bronstein, MSR Program
Semiannu. Progr. Rep. Feb. 28, 1970, ORNL-4548, p. 184.

53. H. R. Bronstein, Chem. Div. Annu. Progr. Rep. May 20,
1970, ORNL4581, p. 119.

54. The Encyclopedia of Electrochemistry, ed. by C. A.
Hampel, pp. 44—51, Reinhold, New York, 1964.

55. H. Aida, J. Epelboin, and M, Garreau, J, Electrochem,
Soc. 118,243 (1971).
ORNL-DWG. 71-2238

|<—3/4in.—)1 :

— %
BORON NEAN Ni
NITRIDE —= R || - .
\ __'.:.'-? |...|F
SINGLE 2\ i 13 in.
N \ .
CRYSTAL—Hh M\ ,'::.'Ez‘d’
\\\\\ ! Z(C)
Ni FRIT

Fig. 10.19. Reference electrode assembly.

the nickel frit would act as a beryllium atom, if the
“chunking effect” were valid, or perhaps as a subvalent
Be™ or Be,?* ion, if the solution assumption is valid.
Nevertheless, the mechanism would be as follows:

Be® - Be?* + 2¢”
if as a “chunk of beryllium,” and

Be* - Be** +e”,

Be,** = 2Be?" + 2e”

if as a subvalent species dissolved in the melt, with
transport of the electrons through the LaF; crystal and
reaction with the Ni** ions to form nickel metal. The
excess fluoride ions are also transported (in the op-
posite direction) through the LaF; crystal®*!»®? into
the main bulk of the melt to maintain electrical
neutrality. The overall reaction is then simply Be +
NiF, (from the electrodes) to give nickel metal in the
LaF; cup and beryllium fluoride in the main melt.

A series of experiments was conducted to verify this
mechanism. First, a beryllium electrode was mounted
through an inverted BeO crucible. Since solid beryllium
metal is less dense than the melt, any solid metal
detached from the electrode by the so-called ‘“‘chunking

128

effect” would rise to, and be trapped at, the melt
surface within the crucible. Dissolved metal, on the
other hand, would still reach the reference electrode.
The observation was that under these conditions the
NiF, in the LaF; cup was still transformed into Ni
metal. _

Second, with the same crucible-encapsulated beryl-
lium electrode, the reference electrode assembly was
modified by. cutting the LaF; crystal below the cup
portion and placing a thin disk of CaF, Y4 in. thick
between the two portions of the LaF;. It was hoped
that the CaF, crystal would perhaps have little or no
electronic conduction and so act as a filter of electrons.
This cell maintained the theoretical voltage for a period
five times longer than the single-crystal LaF; before
onset of deterioration. The CaF, apparently had some
small electronic conductance of its own.

A third, and definitive, experiment was made. The
reference electrode assembly was remade without the
CaF, disk. To the main melt, NiF, was added as a
“getter” for the dissolved Be. The beryllium electrode
was now encased in a somewhat porous BeO crucible of
Y% in. wall thickness by % in. ID containing the pure
melt LiF-BeF, (67-33 mole %). The crucible prevents
the mixing of the two solutions, its porosity allowing
ion migration. Any Ni** ions entering the crucible are
immediately reduced to Ni metal by the saturated
beryllium solution. The same result occurs if any
dissolved beryllium diffuses into the main melt sat-
urated with NiF,. By this means the theoretical cell
potential was maintained for a period of three days.
Termination of the experiment apparently was caused
by the loss of ion migration through the BeO crucible
walls plugged up by the deposition of nickel metal.

The voltage recorded in all these cells before onset of
deterioration was 2.045 + 0.003 V. According to the
compilation of Baes®® for the free energy of formation

~of NiF, and of BeF, in BeF,-LiF (33-67 mole %), this

cell should yield 2.041 + 0.025 V at 600°C.

These experiments have established that beryllium
does dissolve in the LiF-BeF, (67-33 mole %) melt to
some small extent, yet to be determined, and that,
somewhat surprisingly, the conductance in the LaF,
crystal has a small electronic component.

‘With this knowledge, the technique described can and
will be used to obtain reliable thermodynamic data on a
variety of systems of interest.

56. C. F, Baes, J1., “Nuclear Metallurgy,” vol. 15, AIME
Symposium on Reprocessing of Nuclear Fuels, pp. 615—-44,
1969,
11.

11.1 EXTRACTION OF RUBIDIUM AND CESIUM
FROM MSBR FUEL SOLVENT INTO BISMUTH
BY REDUCTION WITH LITHIUM AT 650°C

D. M. Richardson  J. H. Shaffer

The experimental determination of the distributions
of the alkali metals in the metal transfer process was
initiated over a year ago to assure that the behavior of
these metals as fission products, contaminants, or
intentional diluents could be predicted.. The distri-
butions of sodium and potassium were previously
reported for the complete metal transfer process.'
Distributions in the chloride/bismuth extraction were
reported for cesium,? and the distributions of rubidium
and cesium in the fluoride/bismuth extraction at 650°C
are reported here.

The reductive extractions were performed indi-
vidually in 4-in. IPS 304L stainless steel vessels with
mild steel liners. Two %, -in. steel pipes extended to
within % in. of the bottom of the vessels and provided
separately for adding lithium directly to the metal
phase and for withdrawing bismuth samples in graphite
ladles. Additional ports were provided for material
additions, inserting beryllium electrodes, salt sampling
with hydrogen-fired copper filter sticks, gas exhaust,
and thermocouples. The vessels were each charged with
approximately 3000 g of bismuth and were hydrogen
sparged at 600°C for 8 hr.

The fluoride salts consisted of LiF-BeF,-ThF,
(72-16-12 mole %) in batches of approximately 3000 g
that were individually treated with carrier salt and
radiotracer: 1.48 X 1072 mole % RbF (with *Rb)and
0.95 X 1072 mole % CsF (with '27CsF). These salts
were treated in nickel preparation vessels at 650°C
according to standard hydrofluorination procedures,
followed by hydrogen sparging. Approximately 3000 g
of each salt was transferred at 650°C under flowing
argon to the extraction vessels.

1. D. M. Richardson and J. H. Shaffer, MSR Program
Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 107.

2. D. M. Richardson and J. H. Shaffer, MSR Program
Semiannu. Progr. Rep. Feb. 28, 1970, ORNL-4548, p. 172.

129

Chemistry of Molten-Salt Reactor Fuel Technology

Reductions were performed by successive additions of
0.25 or 0.5 g of clean lithium metal, and samples of salt
and metal phases were taken after a minimum of 5 hr of
argon sparging at 1 liter/min.

Analyses of lithium, thorium, and the corrosion
metals were performed by the Spectrochemical Labo-
ratory and by the General Analysis Laboratory. Anal-
yses of rubidium and cesium were made by gamma-
spectrometric counting of the 1.077-MeV gamma of
86Rb and the 0.661-MeV gamma of '27Cs. Over a
week was allowed for decay of thorium daughter
activities in the bismuth samples. Thorium activities in
the salts were corrected by means of a salt blank
without radiotracer.

The equilibrium quotient obtained for rubidium was
K, = (Dry/Dpp) = 8.62 + 1.26, the mean of three
samples at the 0.95 confidence level. The corresponding
distribution equation for this composition of salt was

where X . is the mole fraction of lithium in bismuth.
The cesium salt preparation was used twice for a
reductive extraction into bismuth. In the first instance
there was large scatter of the sample data obtained, and
the spare samples were expended in the course of
measurements of phase segregation of solute in frozen
bismuth pellets.? Subsequently, this salt was transferred
to a second extraction vessel containing freshly pre-
pared bismuth, and the reductions were repeated. The
equilibrium quotient obtained in this experiment was

Ko =(DcgDLj)=15.05 £081,

the mean of 15 samples at the 0.95 confidence level.
The corresponding distribution equation for this com-
position of salt was log D¢ = log X ; +1.320.

The relative extractabilities of the alkali metals from
fluoride salt into bismuth containing lithium at 650°C
areCs :Na:Rb:K:1:0.80:0.57:0.53.

3. D. M. Richardson and J, H. Shaffer, sect. 11.4, this report.
11.2 DISTRIBUTION OF THORIUM BETWEEN
MSBR FUEL SOLVENT AND BISMUTH
SATURATED WITH NICKEL AND THORIUM
AT 650°C

D. M. Richardson  J. H. Shaffer

Occasionally, in reductive-extraction experiments
where various metals were reduced from LiF-
BeF,-ThF, (72-16-12 mole %) into bismuth at 650°C,
there were instances where the solubility product
constant of ThNiBi, in bismuth was reached during the
later stages of reduction. It was of interest to examine
the experimental data obtained under these conditions
to determine whether or not the normal distribution of
thorium prevailed.

The other metals whose distributions were indi-
vidually measured in each of these experiments were
cesium,® potassium,® rubidium,® and sodium.® The
thorium and nickel solubility data of concern here are
plotted in Fig. 11.1, together with the derived line
(with theoretical slope): Xt X Xy; = 2.36 X 1076,
where X is the mole fraction in bismuth. The experi-
mental data shown in Fig. 11.1 were obtained from
spectrographic analyses and provide confirmation of the
reported solubility product constant® at 650° of 1.5 X
107°.

The thorium and lithium data for the same samples
are plotted in Fig. 11.2, together with the derived line
{(with theoretical slope):

XTh/(XLi)4 =47 X 10—7 s
where X is the mole fraction in bismuth. The corre-

sponding log,, distribution equation for this compo-
sition of fluoride salt is

log Dt =4 logDyp; +8.022.
The distributions of thorium and lithium for this
fluoride salt at 600°C and 700°C have been reported’

and are described by the derived equation

log [Dyp/(D;)*] = 8987(1/T°K) — 1.4844 .

4. D. M. Richardson and J. H. Shaffer, sect. 11.1, this réport.

5. D. M. Richardson and J. H. Shaffer, MSR Program
Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 107.

6. F. J. Smith and L. M. Ferris, J. Inorg. Nucl. Chem. 32,
2863 (1970).

7. L. M. Ferris and J. }, Lawrance, MSR Prograrh Semiannu,
Progr. Rep. Feb. 28, 1969, ORNL-4396, p. 284.

130

_2 ORNL— DWG 7{— 7526
10
> N
8 Tl
¢ PN
PR N
2
[ ]
- 9
z \ b
-3 .
210 /A b
= _ — G
5 Xpp* Xy =2.36 X140
@ L 4
Z .
= 5 ™
]
z
O
I
-
2
104

10"4 2 5 1073 2 5 1072
NICKEL IN BISMUTH (mole fraction)

Fig. 11.1. LiF-BeF,-ThF, (72-16-12 mole %) and bismuth at
650°C. Five experiments at nickel and thorium saturation.

102 ORNL—DWG 74— 7527
S
U
- L4
LJ
/ .
2
/’
L
L 4
£~ .
1070 -fi'.l.
- L]
5 . Il

X
=47 x40
(XLI)
2 /

XLi

Fig. 11.2. LiF-BeF,-ThF4 (72-16-12 mole %) and bismuth at
650°C. Five experiments at nickel and thorium saturation.
The resulting estimated distribution equation for 650°C
is

log Dy, =4 log Dy ; +8.251 .

Considering the scatter of the present data and the
inaccuracy of simple interpolation of log K, there is
close agreement of these distributions at 650°C. Within
the limitations mentioned, the present data show that
the distribution of soluble thorium is unaffected by the
presence of ThNiBi, in the bismuth phase.

11.3 BISMUTH-MANGANESE ALLOYS AS
EXTRACTANTS FOR RARE EARTHS
FROM MSBR FUEL SOLVENT

D. M. Moulton  J. H. Shaffer

Bismuth-manganese alloys have been used to extract
cerium from LiF-BeF,-ThF,; (72-16-12 mole %). Man-
ganese forms reasonably low-melting eutectics and no
strong compounds with thorium and rare earths, and it
seemed that adding it to the bismuth might raise the

thorium solubility without any undesirable effects. An

extraction was carried out at 600° with 24 mole % Mn
in the Bi, using Th metal as the reductant. Then more
Mn was added — enough to increase the concentration
to 36.5 mole %, but the solubility is reported to be only
32% — and the extraction was continued. Finally the
temperature was raised to 700°, where all the Mn
should have dissolved, and more additions were made.
Analysis of the samples was by gamma counting and
spectrographic techniques.

Thorium solubility increased substantially over that in
bismuth alone, to 1.36 and 2.38 wt % at 600 and 700°,
and it is not certain that saturation was reached even
then. The maximum distributions (0.075 and 0.13)
were not correspondingly high because of the lower
average atomic weight of the metal. Although thorium
distribution continued to rise, that of cerium did not; it
leveled off as if saturated, but its concentration was so
low (0.1 wt % in the salt) that it does not seem likely
that it formed a primary solid phase.

The distributions are shown in Fig. 11.3, and the
equilibrium constants are given in Table 11.1; these
latter values are not very different from the figures in
pure bismuth. The two manganese concentrations gave
the same results and are not distinguished. If the tailing
off of the cerium, so far unexplained, turns out not to
be real, then it should be possible with this solvent to
reduce the amount of liquid metal circulating in the
reductive-extraction process without much penalty in
the separation. |

131

ORNL-DWG 71-7528

i R T T | N W T T T
[ o 700°
05 [-—® 600:] MAXIMUM POINTS IN PURE Bi
—— ® 700° ) (MSRP SEMIANNUAL Pz
| FEB 1969, p. 191) —
/ jel
0.2 5]
-g/
//
S 0.4 Y )
Q . S
oe”
v
0.05 1
& a
SGASLoPE = V3
0.02
0.01 '
0.001  0.002 0.005 001 002 005 04 0.2
Dy,

Fig. 11.3. D, vs Dy, into Bi-Mn solutions from LiF-
BeF;-ThF, (72-16-12 mole %).

Table 11.1. Logs of equilibrium constants in Bi-Mn solutions

0 Dy/Dc, Dy*/Dyy, D¢ /Dy
600 —7.77+0.18 ~10.17+0.19  +0.81 + 0.09
700 —6.74+0.15 —8.90 + 0.29 +0.26 + 0.39

11.4 REMOVAL OF SOLUTES FROM BISMUTH
BY FRACTIONAL CRYSTALLIZATION

D. M. Richardson J. H. Shaffer

The occurrence of sometimes large variations in solute
concentrations was noted in reductive-extraction exper-
iments where metal samples were taken in graphite
ladles directly from the bismuth phase. Efforts to
determine the causes for this have resulted in finding
evidence that solutes are strongly expelled to the
surface of small bismuth pellets during the process of
freezing. This effect could be utilized in a process for
concentrating solutes and, conversely, for the purifi-
cation of bismuth.

The graphite “ladle” is a 7 4-in.<diam, 2% -in.long
rod that is screwed to a Y, -in. metal rod at its top end.
Near the bottom end and near the middle there are
drilled holes at a 45° angle that are 74, in. in diameter
and ', in. deep. When pushed down into the molten
bismuth and then withdrawn, the ladle removes two
specimens of the bismuth phase that are usually
between 1 and 1'% g each. Before reduction by lithium
132

ORNL-DWG 71-7529

102
5 \CHROMIUM
\—{ MANGANESE \
\ NICKEL~ LN\ \ Y
wITHIUM \ \\ NJHORIUM AND CERIUM
a4 . ) \ \ \\
— N N NN
o CESIUM\\ \\ \\
Tl 10 A S A
1By N \\ ~
- N RN ~
S ~ SN NN
— \\ QQ\‘ \\\ \
5 NN
~ \\:k \t\
b
~ IR
\h\\\
2 \\\\\\
A
IS
10°
2 5 1072 2 5 10! - 5 e
W —#h
W

Fig. 11.4. Distribution of solutes in bismuth.

metal additions is started, the pellets are of bright
metallic appearance and are easily dropped from their
holes. As the reduction proceeds, however, the pellets
become black, and gentle tapping of the ladle may be
required to dislodge them.

Initially, the differences in specific radioactivities
within pairs of bismuth pellets were speculated to be
the result of small and variable contamination with the
salt phase. However, it was found that gentle rinsing
. with salt solvents (acetone, methanol, or verbocit
solution) was not effective in equalizing the specific
activities.

The possibility that the black surface was a reaction
product with graphite and was responsible for a variable
loss of solute from the pellet surface to the ladle was
investigated by using a 28-kc sonic scrubbing bath to
physically remove varying amounts of the material from
the surface. For this case uniform solute concentrations
should be found inside the pellets. The most vigorous
scrubbing was obtained in verbocit solution, and after 1
hr a black pellet would become brightly etched and
have the appearance of a single crystal.

The samples examined in this manner were from the
first fluoride/bismuth extraction involving ! *7Cs® and
the chloride/bismuth extraction that involved po-
tassium and '%%Ce.?>!'® For the radioisotope solutes
the specific activity of single bismuth pellets was
measured before and after each of several scrubbings to

remove material from the surface. The change in weight
of the pellet was then used to compute the average
specific activity of the material removed, which was
then normalized as a ratio by dividing by the average
specific activity of the original unscrubbed pellet. These
ratios are plotted as the ordinate in Fig. 11.4 against the
fraction of original weight removed by scrubbing as the
abscissa. In the case of thorium, chromium, nianganese,
nickel, and lithium, the ratios were deduced by com-
parison of the two pellets within a sample pair and
consequently are less precise.

For all of the metals studied there was a significant
segregation of solute in the outer layers of the pellets.
For cerium, chromium, and thorium the average con-
centration in the outer 1% by weight was 60 or 70
times greater than the average concentration of the
original pellet. Expressed differently, this means that
two-thirds of all of these solutes was located in this
outer shell. Although it has not been proved that
high-concentration pellets were not the result of in-
homogeneities in the liquid bismuth phase, it is ap-
parent that solutes were strongly expelled to the surface

8. D. M. Richardson and J. H, Shaffer, sect. 11.1, this report.

9. D. M. Richardson and J. H. Shaffer, MSR-70-62, p. 11
(Sept. 2, 1970) (internal memorandum).

10. D. M. Richardson and J. H. Shaffer, MSR Program
Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 107.
and, as a consequence, that low-concentration pellets
might be expected.

In the second fluoride/bismuth experiment involving
137Cs,® it was found that the top position pellets were
consistently lower in thorium and cesium concentration
than the corresponding twins from the bottom position.
The median difference was 20% for thorium and 5% for
cesium. This result can be rationalized by the argument
that the pellet in the bottom position is usually the first

133

to freeze since it is farther from the hot mass of the
metal sampling rod and since it is upstream in the
convected cover gas that sweeps up the hot sampler
when it is raised into the standpipe. In this case the
slower. cooling of the top position pellet would allow
more complete expulsion of solute and a greater
likelihood that solute would be lost to the graphite
ladle.
12. Development and Evaluation of Analytical Methods

for Molten-Salt Reactors

A.S. Meyer

12.1 ELECTROANALYTICAL STUDIES OF
TITANIUM(IV) IN MOLTEN LiF-NaF-KF
(46.5-11.5-42.0 MOLE %)

F.R.Clayton' D.L.Manning G.Mamantov?®

We are investigating the electroanalytical behavior of
titanium(IV) in molten LiF-NaF-KF, NaBF,, and LiF-
BeF,-ZrF, . Titanium is present in the structural metal
(Hastelloy N) of the molten-salt reactor and, from free
energy considerations, should be somewhat more re-
active than chromium,. '

Voltammetric studies of the reduction of Ti(IV) in
molten LiF-NaF-KF (46.5-11.5-42.0 mole %) at
~500°C were carried out using a graphite cell enclosed
in a Pyrex jacket under a helium atmosphere. Platinum
(A = 0.1 cm?) and pyrolytic graphite (PG) working
electrodes, a large Pt wire counterelectrode, and an Ni
wire quasi-reference electrode (QRE) were employed in
the experiments. Titanium(IV) was added as K, TiF,
(obtained from Harshaw Chemical Company or D. E.
LaValle, ORNL). The cyclic voltammograms (scan rate
= 0.1 V/sec) obtained prior to the addition of K, TiFq
showed two small reduction waves at —0.20 and —0.48
V (vs Ni QRE) at both the Pt and PG electrodes
{(probably due to iron and chromium impurities in the
melt). In addition, a larger impurity wave (possibly the
reduction. of OH™) was observed at the Pt electrode at
—0.95 V (PG electrode cannot be used at negative
potentials in LiF-NaF-KF due to predeposition of
potassium). Addition of K, TiF¢ [Ti(IV) concentrations
~0.1 M] resulted in a large well-defined reduction wave
(£t 2 =+0.05 V vs QRE) at both Pt and PG electrodes.
The peak current (ip) for this reduction wave reached a
maximum after 1 to 2 hr and then decreased to ~0.1 of
maximum i, after ~24 hr. Further Ti(IV) additions
restored the wave, although it decreased with time. At

1. Student participant, University of Tennessee, Knoxville.
2. Consultant, Department of Chemistry, University of Ten-
nessee, Knoxville.

the Pt electrode a second, larger and steeper, wave was

. observed at —1.45 V. This wave was less reproducible

134

than the wave at +0.05 V. A stripping-type reoxidation
peak clearly related to the wave at —1.45 V was
observed at —0.80 V. The reduction wave at —1.45 V
was reasonably stable over a four-day period. Estimates
of the number of electrons, »n, involved in the two
waves (from logarithmic plots as well as cathodic to
anodic peak potential separation) gave n = | for the
+0.05-V wave and n = 3 for the —1.45-V wave
respectively.

Parameter ip/v”2 values for the wave at +0.05 V
were independent of the scan rate (up to 50 V/sec) and
Ti(IV) concentrations <0.1 M, indicating no kinetic
complications.

The above experimental observations suggest that
Ti(IV) is reduced reversibly to Ti(III), which in turn is
reduced to the metal at a potential ~1.5 V more
cathodic than that of reduction of Ti(IV). The disap-
pearance of Ti(IV), but not Ti(III), with time suggests
that Ti(IV) is reduced by graphite and/or the nickel
QRE. The examination of the Ni QRE at the end of the
experiment showed noticeable attack. A second experi-
ment in a graphite container using a Pt QRE gave
similar results. Again, Ti(IV) could not be kept in
solution for more than one to two days. In this
experiment all potentials were more cathodic by ~350
mV due to the shift in the QRE. Half-wave potentials
for polarograms constructed from current-time curves
for the two Ti waves were —0.30 V [Ti(1V) — Ti(III)]
and —1.75 V [Ti(Ilf) — Ti(0)]. Additional evidence
for n = 1 for the first wave was obtained from the peak
current ratio for derivative vs normal voltammetry. The
experimentally observed ratio of 0.500 is in very good
agreement with the theoretically predicted value of
0.508 for n =1 and v= 0.1 V/sec at S00°C. '

In view of the difficulty encountered in maintaining a
stable solution of Ti(IV) in a graphite cell, we con-
ducted further experiments in a platinum cell using
only platinum electrodes. In this cell, Ti(IV) was stable
and could be kept in solution for several weeks. The
Ti(IV) to Ti(IIT) voltammogram was well defined and
reproducible. The peak current was found to be
proportional to concentration of Ti(IV) and to the
square root of the scan rate (up to ~10 V/sec). The
Ti(IIT) to Ti(0) wave was preceded by what appears to
be a predeposition wave, probably due to alloying of Ti
with Pt. Presence of strong alloy formation is supported
by the unusually large separation between anodic and
cathodic peak potentials for this wave. The determina-
tion of #, from the ratios of voltammetric peak currents
and limiting currents for polarograms constructed from
current-time curves, gave n,/n; of 3. The half-wave
potentials for the two.waves (measured at a potential
corresponding to 0.85ip) with respect to an
Ni/NiF (sat) electrode contained in boron nitride were
found to be +0.15 and —1.45 V respectively.

Chronopotentiograms for the reduction of Ti(IV) to
Ti(IIT) were reasonably well defined. The ir'/? product
(r = transition time) was essentially constant in the
current range 1 to 10 mA. The diffusion coefficient,
calculated from Sand’s. equation, was 5.3 X 107
cm?/sec (at 480°C). Excellent verification of n = 1 for
the first wave was achieved from the ratio of voltam-
metric ip/v” 2 to chronopotentiometric ir'/* (for this
determination of #, concentration, diffusion coef-
ficient, and electrode area do not need to be known).

Placement of a titanium electrode in the melt resulted
in the expected conversion of Ti(IV) to Ti(Ill), as
. evidenced by the disappearance of the first Ti reduction
wave and a gross shift of the Pt QRE with respect to the
Ni(IT)/Ni BN electrode (the QRE potential before the
immersion of the Ti electrode was +585 mV; after the
immersion it eventually became —515 mV). Voltammo-
grams run with the Ti electrode as the reference
[presumably poised to Ti(III)/Ti] showed that the
oxidation of Ti(IIT) was occurring at about +1.5 V and
the Ti(III) reduction at 0.0 V.

We are now studying the behavior of titanium in
Li, BeF, and NaBF, .

12.2 REFERENCE ELECTRODE STUDIES
IN MOLTEN NaBF,

D. L. Manning

A stable and rugged reference electrode is needed for
the in-line analysis of the MSR coolant salt. Providing a
fixed potential reference for voltammetric scans (re-
placing the quasi-reference electrode) will simplify the

135

measurement of corrosion products. Also for in-line
measurement on the coolant salt technology loop, it
can be used to obtain a “‘redox” potential of the
coolant which can perhaps be correlated with corrosion
rates or proton concentration in the coolant. Accord-
ingly, a model of the Ni/NiF, (LaF,) electrode,® which
has been found to serve as a practical reference for
electroanalytical measurements-in fuel melts, was tested
in molten NaBF, at ~400°C. _

Upon immersing the reference electrode in the melt
and monitoring the emf vs a platinum electrode
(assumed to be poised at the equilibrium potential of
the melt), the platinum electrode measured approxi-
mately —150 mV vs the Ni/NiF, (LaF,) reference. The

-drift and unsteadiness of the emf, however, was much

more pronounced in molten NaBF, than was experi-
enced in LiF-BeF,-ZrF, melts.* A nickel electrode
immersed in the melt measured ~ —175 mV vs the
reference.

Nickel fluoride (450 mg) was next added to the melt
(80 g) to establish the Ni/NiF, redox potential.

~ Assuming the Ni/NiF, couple is the poised system in

the melt inside the LaF; crystal, the magnitude of the
emf between the reference electrode and the Ni
electrode in the molten NaBF, should decrease as the
NiF, dissolves. This was not observed. Within the
scatter of the measurements, the emf of the nickel
electrode and the platinum electrode did not signifi-
cantly change. Voltammograms recorded at a pyrolytic
graphite electrode did not reveal a reduction wave that
could be attributed to the NiZ* + 2e — Ni® electrode
reaction, and only a very small stripping wave for Ni°® -
Ni** + 2e could be developed. This -suggests that the
solubility of NiF, in NaBF, is very limited (<25 ppm).
The apparent limited solubility of NiF, may in part
explain the erratic-behavior of the reference electrode. -
The Ni/NiF, couple was probably not the emf-
determining system; consequently, the electrode mostly
resembled a quasj-reference. At this time, an Ni/NiF,
(LaF;) reference- with a saturated solution of NiF,
inside the LaF; crystal for use in NaBF, melts does not
appear promising. Further tests are planned, however,
with an alternate model of the electrode® which makes
use of a solid Ni/NiF, pressed pellet as the potential-
determining couple. Other reference couples will also be
tested in the LaF, electrode. '

3. D. L. Manning and H. R. Bronstein, MSR Program
Semiannu, Progr, Rep, Feb, 28, 1970, ORNL-4548, p. 184.

4, D. L, Manning and F. R, Clayton, MSR Program Semi-
annu. Progr, Rep, Aug, 31, 1970, ORNL4622, p. 115.
12.3 ELECTROCHEMICAL-SPECTRAL STUDIES
OF MOLTEN FLUORIDE SALT SOLUTIONS

J.P.Young F.L.Whiting'  Gleb Mamantov?

The present phase of the electrochemical generation
and spectral characterization of unusual oxidation
states of solute species in molten fluoride salts has been
concluded. This work has been summarized in the Ph.D.
thesis of F. L. Whiting. The solvent system has been
mainly, but not exclusively, LiF-NaF-KF. The study
was carried out to survey the utility of this technique as
a preparative and investigational tool. Our original
assignment of the absorption spectrum of superoxide,
0,7, in molten LiF-NaF-KF has proved to be in error.’
Superoxide does exist in this melt; it exhibits an
ultraviolet absorption peak at 234 nm and a Raman
spectral peak at 1107 cm™'. The yellow solution
obtained by the addition of NaQO, to the melt, however,
is due to the oxidation of chromium impurities (<10
ppm) to CrO, %", It has been reported that O, dissolved
in liquid NH, also forms a yellow solution with an
absorption peak near 360 nm.® [t was not possible to
confirm this; rather, the addition of NaQO, to liquid
NH; yielded a colorless solution with an absorption
peak at 252 nm. Other investigations have also indi-
cated the error of the original liquid ammonia study.”
In other results of this work, we found it possible to
generate electrolytically Mn(III), Cu(Il), U(III), and
CrO4* and to obtain their spectra in the melt.
Although Co(II) could be oxidized electrolytically in
the melts studied, the product, probably Co(IIl), was
not stable under the experimental conditions of the
study, and no spectrum could be obtained. A method
was developed for removing dissolved chromium and
other reducible metal ions to less than ppm amounts by
treatment of molten LiF-NaF-KF with elemental sili-
con. Futher, the generation and spectral measurement
of CrO,?*", by the oxidation of Cr(Il) or (III) in the
presence of excess oxide ion, can be used for ‘the
spectral determination of chromium to less than 1 ppm
in LiF-NaF-KF. Likewise, it would seem that the
generation and spectral measurement of CrO,%” by the
oxidation of excess chromium in the presence of oxide
ion could be applied to a colorimetric oxide analysis in
amenable fluoride melts.

5. J. P. Young, F. L. Whiting, and Gleb Mamantov, MSR
Program Semiannu, Progr. Rep. Aug. 31, 1969, ORNL-4449, p.
159.

6. J. K. Thompson and J. Kleinberg, J. Amer. Chem, Soc. 73,
1243 (1951).

7. G. Czapski and B. Halperin, Israel J, Chem, 5, 185 (1967).

136

12.4- SPECTRAL STUDIES
OF MOLTEN FLUORIDE SALTS

J.P. Young

As a result of the graduate research of F. L. Whiting
(reported above) it was apparent that oxygenated solute
species can indeed exist in fluoride melts, and the kinds
of such species are perhaps much more numerous than
had previously been expected. During this period,
various spectral studies of these oxygenated species in
molten fluoride media were carried out, One of the end
results of the study of oxyanions culminated in a
cooperative spectral study of the reaction of OH™ and
OD™ with molten NaBF,. The results of this study are
reported in another section of this report.®

Because of the chromate (CrO,27) formation in
LiF-NaF-KF, the generation and spectral studies of
chromate in other melts were undertaken. It was found
that CrO,*” dissolved in LiF-BeF, melts. The spectrum
is altered from that seen in LiF-NaF-KF, and it is
believed that the dissolved species is dichromate
(Cry0,%). Again, reduced chromium can be oxidized,
with MnF;, in the presence of oxide ion to Cr,0,*".
Chromate dissolves in molten NaF-NaBF,, and the
spectrum indicates Cr,0,%” is the resultant species. It
has not been possible, however, to generate Cr,0,° in
NaBF, by the oxidation of chromium in the presence
of added oxides. The nonreactivity could result from
several causes, such as insolubility of chromium fluoride
or the absence of free oxide ion via the formation of
BF,0".

The ultimate sensitivity of any possible spectral
determination of O*” in molten salts is limited by the
molar absorptivity of the resultant oxidized oxyanion.
The molar absorptivity of the most intense peak of
CrO,; %", at 372 nm, in molten LiF-NaF-KF is approxi-
mately 4000 M~ at 500°C. The molar absorptivity of
the isoelectronic VO, 3™ at its peak maximum, 272 nm,
should be near 8000 M~ ¢cm™; VO, 3 is soluble in
molten NaF-NaBF, or LiF-NaF-KF and exhibits the
expected spectrum. The oxidation of lower-valent
vanadium to VO, has yet to be studied.

The spectrum of nitrite (NO,") has been observed in
molten NaBF,. Nitrite exhibits a peak at 350 nm; the
sensitivity of this absorption is not accurately known,
but the peak is sufficiently sensitive to detect NO,™ at
concentrations less than 1000 ppm. If further consider-
ation is given to the use of a nitrate-nitrite secondary

8. John B. Bates, J. P. Young, H. W, Kohn, M, M, Murray,
and G. E. Boyd, sect. 9.5.
coolant for the MSBR, the results of this spectral study
of NO, can be applied to a possible determination of
the secondary coolant in molten NaF-NaBF,. It was
also observed that NaNO, dissolves in LiF-NaF-KF; the
spectrum of NO,  in this melt is quite similar to that
seen in molten fluoroborate melts.

It was observed in earlier studies of NaBF, melts that
many -salts, such as some transition-metal fluorides,
rare-earth fluorides, and actinide fluorides, did not
possess a sufficiently high solubility to be detected
spectrally. For this reason it was thought that analyt-
ically, spectral measurements might not be of much
value for determinations of solute species in this
proposed coolant salt, Earlier spectral data, further-
more, on what turned out to be impure NaF-NaBF,
indicated that the melt absorbed strongly in the
ultraviolet below 300 nm. If NaBF, is purified,
depending on the method of purification, the salt is
quite transparent even at wavelengths of 200 nm or less.
Based on chemical analyses, this increase in trans-
parency is caused by the removal of Fe(II) and/or
Fe(II) from the salt. Obviously, then, Fe(II) or Fe(IIT)
in ppm amounts could be spectrally determined by the
ultraviolet absorbance of molten NaBF,.

Other solute species could interfere; in this region of
the spectrum one observes allowed transitions of many
ions. The extent of such interference is being investi-
gated. Presently, the extent of melt purification can be
followed by absorbance measurements in the ultraviolet
region of the spectrum below 300 nm,

12.5 ANALYTICAL STUDIES OF THE NaBF,
COOLANT SALT

J.M.Dale R.F.Apple A.S.Meyer

In the previous report® it was noted that S. Cantor
had raised questions concerning the meaning of the Karl
Fischer (KF) determinations of “water” in samples of
NaBF, coolant salt. We cited earlier analyses in which
comparable results were obtained by the direct KF
titration and by titration of the water separated by
azeotropic distillation from coolant samples as evidence
that these analyses represented some protonated species
in the salt, It was found, however, that azeotropic
distillations of current samples yielded titrations much
lower than those obtained by direct titrations, and
therefore a critical evaluation of the azeotropic method
was made.

9. R. F. Apple et al., MSR Program Semiannu. Progr. Rep.
Aug, 31, 1970, ORNL-4622, p. 116,

137

In the azeotropic procedure, about 450 ml of
pyridine (1000 ppm maximum H, O) is placed in a glass
still fitted with a Vigreux column. After a period of
total reflux, about 50 ml of the distillate is discarded,
and successive 3-ml test portions of the distillate are
titrated with coulometrically generated KF reagent
until a constant titration establishes a blank for the
apparatus. The sample of salt is then added to the stiil
pot through a standard taper joint, and titration of the
distillate is continued until the titration drops to the
original blank value, usually three to four test portions
of distillate. : |

It was noted that the blanks could contribute as much
as several hundred micrograms to a sample titration
(depending in part on the relative humidity in the
laboratory), a value sometimes in excess of the net
titration for .smaller samples. Also, the addition tech-
nique was subject to some uncertainties. Although no
apparent introduction of water was observed during
simulated sample additions, the pulverized samples were
briefly exposed to the atmosphere during an addition.

We have modified the technique by eliminating all
unnecessary glass joints, minimizing the length of
rubber connectors, substituting a molecular sieve
column for the Ascarite—magnesium perchlorate drier
on the still vent, using manual rather than automatic
termination of the titration current, and using pyridine
reagent only from selected manufacturers’ lots. With
these modifications the blank titrations were reduced to
about 25 to 50 ug, at least a fivefold improvement.

Also, a solid sample injector was developed to
eliminate contamination during sample additions. This
injector is basically a plunger with a recess that can be
charged with approximately 1 g of salt in a dry box and
sealed in a Teflon sleeve., The loaded assembly after
weighing is in turn sealed in a metal standard taper joint
which couples to the still pot. After the system is
blanked the plunger can be depressed to inject the
sample without exposure to the atmosphere. With this
injector, about 95% of the water from a sodium tartrate
dihydrate standard is recovered.

With the improved procedure, negligible titrations
have been obtained on current coolant samples.
Samples taken from the FCL-1 thermal-convection loop
are conservatively reported as <50 ppm vs 1000 ppm
by direct titration. It should be noted that the blank
titrations reported cannot explain the much higher
azeotropic values (up to 2500 ppm) that were obtained
on earlier samples, nor do the present low titrations
completely eliminate the possible existence of protons
in these samples. It will be necessary to perform
azeotropic distillation on salt of known proton concen-
tration (e.g., NaBF3;OH when prepared) to evaluate the
meaning of these negative results.

We are proposing to use the azeotropic technique for
the determination of “water’ that is present as hydroly-
sis products in NaBF, cover-gas streams, The sensitivity
of the direct KF titration of these gases is limited by
the high blank resulting from the reaction of BF; with
the constituents of the reagent. In the proposed method
the water and BF; would first be stripped from the
“cover gas by sparging it through cold dried pyridine;
subsequently, the water would be separated for titra-
tion by azeotropic distillation,

In a preliminary test of this approach, H. W. Kohn!®
found that - negligible blanks. were obtained for gas
streams that contained only BF,; but when HF
(required for his reprocessing studies) was absorbed in
the pyridine, an unacceptably high blank was titrated in
the distillate. This is apparently caused by the reaction
of pyridinium hydrofluoride with the glass apparatus.
In KF titrations at room temperature, excess pyridine
eliminates the interfering reactions of HF, but these
reactions are obviously accelerated at the temperature
of boiling pyridine. A metal azeotropic still has been
designed and will be tested for this application.

12.6 IN-LINE CHEMICAL ANALYSIS

J. M. Dale T. R. Mueller

The project for making in-line measurements of
U(IV)/U(III) ratios and corrosion products in MSR fuel
salt is in the stage of testing of the individual
components involved in the experiment. The thermal-
convection loop, designed by members of the Metals
and Ceramics Division, has been delivered to the X-10
site, and heaters are being installed prior to a salt
flushing operation to ensure that the internal surfaces
of the loop are clean. Five electrode ports have been
provided in a test chamber of the loop. Three of these
will be used for the electrodes for the voltammetric
measurement of the U(IV)/U(III) ratios and corrosion
product concentrations in the melt, and the other two

10. Reactor Chemistry Division, unpublished work.

138

will be used for a reference electrode system for
following the redox potential of the melt. These
measurements will be made during the studies to
correlate the effect of melt potential on the corrosion
of test specimens to be installed in the loop. '

A newly constructed cyclic voltammeter, which pro-
vides several new capabilities for electrochemical studies
on molten-salt systems, will be used in this work. A new
plug-in module was developed to differentiate the
current-potential voltammeter output used for measure-
ment of U(IV)/U(III) ratios. The differentiator has 11
ranges for input signals of different rise time which are
designed to achieve optimum signal-to-noise ratio with
minimum wave-shape distortion. The new instrument
also includes an integrating circuit, provisions for
compensating that portion of the circuit resistance
(mostly in the cell) not compensated by the potentio-
stat, a preselect cyclic counter, a circuit layout that
minimizes cross talk existing in the prototype instru-
ment, a chopper-stabilized sweep generator for “drift-
free”” operation in the “cyclic” or “hold” modes, and a
potentiostat capable of delivering 500 mA of cell
current (100 mA in prototype). The voltammeter,
which can be directly operated by the PDP-8 family of
computers, has a terminal that permits ground isolation
between the computer and the experiment if this is
needed.

A PDP-8/1 digital computer purchased for application
to in-line analytical methods will be used ‘in this study
to allow unattended measurement of the U(IV)/U(III)
ratio in the melt. A computer program’! developed by
M. T. Kelley and R. W, Stelzner will be used in a
modified form for this purpose. Previous work on this
measurement was made in an electrically ungrounded
cell. Before installation on the thermal loop, it is
planned to check out the compatibility of the com-
puter-voltammeter system on a melt in a grounded
system. The experience derived from this study will be
most valuable for application to computer-operated
in-line measurements involving other types of chemical
transducers.

11. M. T, Kelley, R. W. Stelzner, and D. L. Manning, MSR
Program Semiannu, Progr. Rep. Aug. 31, 1969, ORNL-4449, p.
157.
Part 4. Materials Development

13. Examination of MSRE Components

" H. E. McCoy

Operation of the MSRE was terminated after success-
fully completing its mission. Further information can
be gained by examining selected parts of the system.
The examination of the coolant circuit was reported
previously,! and work on the primary circuit has been
partially completed during this reporting period. The
components examined include a graphite moderator
element from the core, a Hastelloy N control rod
thimble, the sample assembly from the pump bowl, a
lost copper sample tube, the shell and several tubes
from the primary heat exchanger, and a freeze valve
that failed as operation was terminated. All of these
parts have been examined visually, and metallographic
samples of most parts have been examined. Work is still
in progress to define the concentrations of fission and
corrosion products on many of the graphite and metal
surfaces. This work is partially completed and will be
reported at a later date. Concentration gradients near
the metal surfaces that are indications of corrosion
behavior are being evaluated by the electron probe
microanalyzer and will be reported later.

13.1 EXAMINATION OF A GRAPHITE
MODERATOR ELEMENT

B.McNabb  H. E. McCoy

One of the five removable graphite core moderator
blocks was examined in HRLEL after completion of the
operation of the MSRE. Visual examination revealed
machining marks still plainly visible in the fuel channels
and on surfaces not exposed directly to flowing fuel
salt, indicating little if any change during operation.
Figure 13.1 is a photograph of the fuel channel facing
the center of the reactor showing the excellent condi-
tion of the core block. The small thumbnail flaw in the
fuel channel about 39 in. from the top is believed to be

1. H. E. McCoy, MSR Program Semiannu. Progr. Rep. Aug.

31, 1970, ORNL-4622, p. 119.

a small crack that was in the bar during fabrication,
since tool marks are on the top of the flaw and skip
over an area at the end of the flaw. The few scratches
on the surface are believed to be caused by handling
and are not very deep.

There was a crack in the graphite around the
Hastelloy N lifting stud at the top of the bar that
extended about halfway around the bar, as shown by
the short arrows in Fig. 13.2. The long arrow indicates
the same channel in the different views. The small
punch mark visible in the top view was made to indicate
the fuel channel toward the center of the reactor, with
a control rod on one side and the surveillance sample
basket on the other side.

The crack in the graphite would not affect the
integrity of the bar, since it was on the end and several
threads remained to secure the lifting stud. The crack
was probably caused by the approximate factor of 2
difference in thermal expansion of graphite and Hastel-
loy N. '

Measurements were made of length, width, and
bowing. There were no significant changes since all
measurements were within the original tolerance for
fabrication of the core block. The measured length of
the approximately rectangular portion of the bar was
64Y, + Y, in. Thus, there was no measurable change in
length. Width measurements across the flat parallel
surfaces were made using 1-to-2-in. micrometers and
read through the Kollmorgen. The original dimensional
tolerances were 1.998%0-99% i Measurements shown
in Table 13.1 were made at ~1-ft intervals from top to
bottom of the side facing the center of the reactor.
Measurements were made near the top, middle, and
bottom at 90° to this on the side that had 0.075 in.
relief for flow between the channels starting 2 in. from
either end, top = 1.9226 in., middle = 1.9231 in., and
bottom = 1.923 in. All measurements were within
tolerance. No actual measurements of the individual

bars were recorded before operation, but assuming the

139
PHOTO 0370-71

Fig. 13.1. Photograph of one of the removable graphite moderator elements after operation in the MSRE, showing the fuel
channel facing the center of the reactor. Machining marks are plainly visible in the fuel channel, showing the excellent condition of
the core block.

(U2}
141

PHOTO 0371~ 74

Fig. 13.2. Photograph of the top of the graphite moderator element, showing the crack originating at the INOR-8 lifting stud and
extending partway around the graphite block shown by the short arrows. The long arrow indicates the same channel in the different
views. The crack was probably caused by the approximate factor of 2 difference in thermal expansion between the INOR-8 and the
graphite during thermal cycling of the reactor.

maximum and minimum of the fabrication tolerances
there were no significant changes or observable trends
(Table 13.1).

Measurement of bowing was made by stretching a
string tightly between two C-clamp-type fixtures at
either end of the bar and measuring the distance
between the bar and the string with a cathetometer,
then rotating the bar 90° and sighting on another bar
surface and measuring again. The maximum observed
bowing was 0.012 in. with the string being closer to the
bar in the center, possibly indicating slight sagging of
the string. It was concluded that there was no appreci-
able bowing. No tolerances for bowing were in the
specifications, and the bar was not supported on a flat

Table 13.1. Dimensional measurements
of graphite moderator element?

Change from original

Measured width dimensions, in inches,

Locaiea (in.) assuming —
Maximum size  Minimum size
Top 1.9947 ~0.0033 0.0017
1t 1.9954 ~0.0025 0.0024
2t 1.9944 ~0.0036 0.0014
36t 1.9948 ~0.0032 0.0018
4ft 1.9942 ~0.0038 0.0012
Bottom 1.9971 ~0.0009 0.0041
+0.000

@Dimensional tolerances were 1,998 in.

'~0.005

PHOTO

Fig. 13.3. Comparison of the bottom positioning stud on the graphite moderator element when initially installed in the MSRE and after operation. The large
picture was made before the reactor was operational, and the small picture in the lower right corner was made in the hot cells after the moderator element had been
at temperature for over three years.

(44"
143

PHOTO 0372-71

ETCHED

5008

Fig. 13.4. Hastelloy N thread that was in contact with graphite for about three yearsat 650°C. The modified layer is likely due to

working effects from machining the thread. Reduced 70.5%.

surface; so absolute measurements were not made or
considered necessary since the bar removed easily.
Photographs were taken of the assembled graphite core
blocks before inserting in the reactor core. Figure 13.3
shows the bottom of core block 1184 C19 in the
assembled position, and the inset shows the bottom of
the same core block as it appeared in HRLEL after
operation for the full life of MSRE. There is no
apparent change, as the small cracks and the small chip
missing from the bottom are apparent in both photo-
graphs and the numbers are just as legible as before
operation.

Core-drilled samples were taken from the fuel flow
channel near the midplane and used as electrodes in the
mass spectrograph, but the results have not been
analyzed. Core-drilled samples were taken at the top,
middle, and bottom of the same fuel channel for Auger
analysis of the surface layers for fission product
deposition, but the results have not been received yet.

A cross section of the Hastelloy N lifting stud in
contact with the graphite is shown in Fig. 13.4. There
appears to be little if any reaction with the graphite.
There is a surface modified layer about 1 or 2 mils deep

similar to that observed on surveillance samples and
attributed to cold working of the surface and subse-
quent heat treatment during service. There appears to
be very little difference in the side of the thread in
contact with the graphite and the root of the thread,
which probably was not in intimate contact.

13.2 AUGER ANALYSIS OF THE SURFACE LAYER
ON GRAPHITE REMOVED FROM THE CORE

OF THE MSRE

R. E. Clausing
The detection and quantitative analysis of thin layers
of fission products deposited on the graphite and metal
surfaces of the MSRE is necessary to predict the
requirements for afterheat removal. Auger electron
spectroscopy is a new technique which offers unique
capabilities in this respect. It provides, under suitable
conditions, excellent sensitivity (down to 0.1% of a
monolayer) and resolution down to one to three atom
layers.”> The technique is applicable to all elements

2. L. A. Harris, /. Appl. Phys. 39, 1419 (1968).
3. L. A. Harris, J. Appl. Phys. 39, 1428 (1968).

except hydrogen and helium and, when combined with
sputter thinning, can provide profiles of concentration
as a function of depth below the original surface.

13.2.1 Auger Electron Spectroscopy

If a vacancy is created in an inner electron shell of an
atom by electron bombardment or any other method,
the excited atom will revert to the ground state by
emitting characteristic x radiation or alternatively by
emitting an electron as the result of a radiationless
transition called an Auger transition. This Auger elec-
tron has an energy which is determined by the energy
levels in the parent atom. Figure 13.5 shows schemat-
ically the Auger process for the creation of the KL; L4
(2144 eV) and L,VV (~149 eV) Auger electrons
characteristic of sulfur. The energy of the Auger
electron is obtained by subtracting the energy of the
two final vacancies from the energy of the initial
vacancy. The most useful Auger electrons have energies
between 10 and 3000 eV, and when they originate
more than a few atom layers below the surface they
cannot escape from electrically conducting solids with-
out energy losses. This effect produces the excellent
depth resolution.

ORNL-DWG 71-6723

h KVV
AUGER
KL3 ELECTRON
AUGER
ELECTRON —]

} VALENCE

ZZ@ZZ%ZZ ELECTRON
VIS

My ! ‘\ ; LEVELS
L3 > = 164

Lo e

oA~
//
K [

VACANCY CREATION
SULPHUR (Z=16)

2472

Fig. 13.5. Energy level diagram showing schematically the
emission of KL 3L 3 and L, V'V Auger electrons from sulfur.

144

Special techniques are required to detect Auger
electrons stimulated by electron bombardment of solids
because the Auger electrons represent only a small
signal superimposed on a large continuous back-
ground.?’? Figure 13.6 shows how this small signal can
be recovered from an energy distribution curve through
differentiation. The differentiated curve is the one
normally reported.

The differentiated curve is usually obtained using the
instrumentation shown in Fig. 13.7. Electrons from the
target are energy analyzed using a retarding potential
applied to the second and third grids of a set of three
grids spaced between the target and a hemispherical
electron collector. The retarding potential is modulated
by superimposing an ac¢ voltage (K sin wt) on the
retarding voltage E£,. The amplitude of the ac compo-
nent (K''sin cwr) in the collector current is proportional
to the number of electrons in the energy range between
Ey + %K and £, — LK. A plot of K’ as a function of
Ey then gives an energy distribution curve. The
amplitude of the ac component (K" sin 2wt) is

ORNL-DWG 7t-6724

I CURVE A
I
I
=
Lo REGION
251! REGION
G|
- = | | REGION
= f ‘ .3
| !
| |
| T 1 J1J
1 I |
i I i I 'V\’ ]
A !
| i P! |
3 s |
AUGER PEAKS APPEAR
N THESE REGIONS |
------------------------------------------------------ |
[ ] [
o I : | CURVE B I
p-d i
b O N }
q - |
-5 | 8
z @ ¥
A= !
w2 |
w -
c%
N o ! u
JY !
<= W
Q l [ }v\v | [ |

0 50 100 1500
V (eV)

ELECTRON ENERGY SPECTRA

2000 2500

Fig. 13.6. Method of enhancing the visibility of Auger
spectra.

145

? TARGET

ORNL- DWG 71 - 6725

RETARDING GRIDS

= 210V
¥ I(E) .
Kk sin w?
6 LOCK—IN-AMP ISOLATION

%’0 Q — TRANSFORMER
il = = (1:1 RATIO)
- Y INPUT

X-Y RECORDER 0-1500 V dec

VOLTAGE SUPPLY

|

X INPUT

(AUTOMATICALLY
SCANNED)

Fig. 13.7. Instrumentation for Auger electron energy analysis using the retarding-potential techniquel

proportional to the derivative of the energy distribution
curve.® The lock-in amplifier output, which is propor-
tional to these amplitudes, can thus be used together
with a signal proportional to £, to plot directly either
an energy distribution curve or its differential by tuning
it to the fundamental of the modulation frequency or
its second harmonic respectively.

13.2.2 Results and Discussion

An apparatus has been built specifically for the Auger
analysis of radioactive samples from the MSRE, and
preliminary results are reported below. Figure 13.8
shows the apparatus. In addition to the basic equipment
for Auger analysis, an ion gun is provided for sputter
thinning the sample, and provision is made to heat the
sample by electron bombardment. A movable shield is

4. P. W. Palmberg and T. N. Rhodin, J. Appl. Phys. 39,2425
(1968).

_provided to minimize the contamination of the appara-

tus. :

Figure 139 shows a curve obtained from MSRE
sample 6. This sample is from a core surveillance
specimen removed in April 1968, after run 14. A
number of Auger peaks are present. They have been
tentatively identified as indicating the presence of
carbon, oxygen, sulfur, molybdenum, niobium, tellur-
ium, technetium, uranium, lithium, and beryllium.
Contributions of these elements to the spectra are -
indicated in the figure caption. There is some overlap of
spectra which complicates the interpretation; however,
the use of better resolution for intense peaks and
increased sensitivity for less intense peaks will minimize
this problem.

Table 13.2 shows how several of the peaks changed
during sputter thinning. These peak heights are propor-
tional to the concentration of the parent atoms at the
surface of the sample. Since no suitable standards were
available, no absolute concentrations can be given yet.
Additional samples and standards are now available, and
quantitative information will be available soon.

f
146

ORNL~ DWG 70-6794

RETARDING
POTENTIAL GRIDS

S
COLLECTOR «& ‘_
§

ELECTROSTATIC
SHIELD
SAMPLE
MOVABLE
SPUTTER SHIELD
FILAMENT FOR
STATIONARY ELECTRON BEAM
SPUTTER SHIELD HEATING

{ON GUN FOR
SPUTTER CLEANING

Fig. 13.8. Apparatus for Auger analysis of low-level radioactive samples.

ORNL-DWG 70-6801R

] |
.y Vad \“*MET;M 510
300 : Wjim
I r/ [ 430 "‘-‘\_‘\
"‘-u-h_
QEGV\/ L37 240\
i
e
N | 222 E.B. CLEANED
3 / / /HOA/|<200 2 min, 7mA, 3.4 kY
N . 800°C
© / { /( I/\182 MODULATION 10 VPP
120 | ATTENUATION AS MARKED
. | 270
160
I /
fM4B
0 100 200 300 400 500 600 700

ELECTRON ENERGY (eV)

" Fig. 13.9. Auger electron spectra from MSRE graphite sample 6 after heating to 800°C for 5 min by electron bombardment. The
following elements have been tentatively identified as making contributions to the peaks: § 148; 0 510, 490; C 270, 253, 240; Mo
222, 185, 160, 149, 137, 118; Nb 200, 160, 92; Tc 250, 220, 180, 145; Te 490, 410, 278; U 92 (86), (72}, (60); Li (39); Be 92.
Values in parentheses are from a low-energy scan not shown. '

147

Table 13.2. Auger electron intensity as a function of depth
below the original surface of sample MSRE 6 (film side)

Normalized to carbon (270) equal 100

Accumulated Estimated

Technetium,

argon ion depth - sulfur, ;Zi;g;g:;?; Niobium
bombardment (atom molybdenum (200 eV)
(X 10'5/ecm?) layers) (148 eV) (182 ¢V)
0 29 9 .2
5 1 9 712 4
13 6 16 3
23 5 7 19 3
= 58 10 5 23 4
136 25 4 16 3
280 55 4 12 3
540 110 3 12 2
1240 250 2 10
1886 380 2 9
3464 700 nd? 9
3570 710 2 14
3700 740 2 9
6764 1350 8
11,907 2400 nd?

9hd = not detected.

13.3 EXAMINATION OF HASTELLOY N
CONTROL ROD THIMBLE

B.McNabb  H. E. McCoy

Control rod thimble 3 was cut off just above the
midplane of the reactor core. Figure 13.10 shows the
electric arc cut made at the left and enlarged views of
the spacer sleeves. Some areas show the possibility of
some wetting of the Hastelloy N by the fuel salt near
the center of the reactor, where the témperature and
flux were higher, as shown around the left and center
spacer sleeves, but not at the bottom, as shown on the
right sleeve. Some of the variations in light and dark
areas are due to not obtaining even lighting for the pan
photographs made in the hot cells. The control rod
thimble was fabricated from Hastelloy N heat Y8487
with chemical composition of Ni—70.80%, Mo—
16.78%, Cr—7.32%, Fe—4.1%, Mn—0.3%, C—-0.05%,
Si—0.17%, S—0.0075%, P—-0.004%, Cu—0.03%, Co—
0.1%, Al-0.16%, Ti—0.25%, B—0.007%; the melt was
made and fabricated in 2-in.-OD tubing by the Inter-
national Nickel Company.

Tensile tests were conducted on some Y-in.-wide
rings cut from the 2-in. control rod thimble, using the
fixture shown in Fig. 13.11. The circular ring is placed

in the fixture when it is closed, and the two halves are
pulled apart in the tensile test. Most of the strain occurs
in the two sides between the grips, but the gage length
changes during the test and makes it impossible to
define elongation as a percent. Therefore crosshead
travel and the reduction in area at the fracture are
reported. Table 13.3 compares the properties of the
unirradiated and irradiated samples tested in the same
manner. It shows that there is little change in the yield
and ultimate strengths. The room-temperature reduc-
tion in area is reduced from about 45% in the
as-received, unirradiated condition to 28% for the
irradiated rings. Some of the ductility at 25°C can be
recovered by annealing at 871°C for 8 hr but is not
improved further by annealing 141 hr at this tempera-

“ture. The elevated-temperature reduction in area was

reduced more severely but still was 9.7% at a fairly slow
deformation rate of 0.002 in./min. Similar changes were
noted in the properties of the surveillance samples
reported previously.

The control rod thimbles were exposed to fuel salt
and fission product deposition only on the exterior of
the tube, and the irradiated rings only exhibited
fissuring or cracking on the exterior of the ring, as
shown in Fig. 13.12. The ID of the tube was slightly
148

R-54116

Fig. 13.10. Portion of Hastelloy N control rod thimble removed from the MSRE. The cut end was near the axial center of the
reactor. The bottom end was near the bottom of the core. The sleeves are Hastelloy N spacers and were held in place by a small weld

bead.
Table 13.3. Tensile data on rings from control rod thimble 3 (heat Y-8487)
Specien : Bosticradiatios Test Crosshead  Yield  Ultimate  Crosshead  Reduction in

No. Condition ot temperature speed stress stress travel area

“©0) (in./min) (psi) (psi) (in.) (%)

5 Unirradiated 650 0.05 39,100 76,700 0.37 28.7

6 Unirradiated 25 0.05 52,000 114,400 1.20 445

T Unirradiated 25 0.05 56,900 117,500 L17 46.0

8 Unirradiated 25 0.05 58,000 124,300 118 43.0

9 Unirradiated 650 0.002 40,700 62,300 0.20 20.6

2 Irradiated None 25 0.05 54400 105,100 0.54 23.2

3 Irradiated None 25 0.05 53,300 102,500 0.51 29.7

4 Irradiated None 25 0.05 60,500 110,800 0.42 285

6 Irradiated None 650 0.05 38,300 51,200 0.099 121

7 Irradiated None 650 0.002 34200 38200 0.061 9.7

5 Irradiated 8 hr at 871°C 25 0.05 49,300 100,500 0.36 34.9

8 Irradiated 141 hr at 871°C 25 0.05 48,700 104,900 0.39 344

“Based on 0.002 in. offset of crosshead travel.

149

oxidized by the reactor cell atmosphere, and the oxide

cracked during testing, but the cracks did not propagate
- into the metal wall. Figure 13.13 shows a cross section
of the control rod thimble and spacer sleeve near the
midplane of the core. The cross section was made
through the button weld that kept the spacer sleeve
from sliding up or down the thimble. The OD of the
sleeve was exposed to flowing salt and the ID to salt in
a 0.005-in. clearance or crevice between the sleeve and
the thimble. The OD of the thimble at this point was
exposed to the same conditions as the ID of the sleeve,
and the ID of the thimble was exposed to the control
rod thimble cooling atmosphere. The spacer sleeve was
made of heat 5060, and the weld bead was made with
heat 5090 weld wire, both air-melted heats of standard
Hastelloy N. There appears to be very little attack on
any of the materials, with only the 2- to 4-mil surface
modification similar to that noted previously. Surface
layers of the thimble material are presently being
dissolved for fission product analysis at the midplane

and bottom of the core, but results have not been
received. Fig. 13.11. Fixture used for pulling Hastelloy N rings cut
from the control rod thimble.

R-54166

0.07 INCHES
T 50x

o

5

Fig. 13.12. Hastelloy N ring from control rod thimble stressed at 25°C. The cracked surface was exposed to the fuel salt, and
the oxidized surface was exposed to the cell environment of 2 to 5% O, and

AS POLISHED

AS POLISHED 100X

ETCHED

100X

150

FHOTO 0373-7

[FrT—

ETCHED
QD OF SLEEVE

500X

ETCHED
D OF SLEEVE

500X

E

ETCHED.
0D OF THIMBLE

500X

ETCHED
1D OF THIMBLE

500X

Fig. 13.13. Cross section of Hastelloy N control rod, spacing sleeve, and weld bead. Reduced 72%.

Figure 13.14 is a photomicrograph of the cross
section of the tube wall, near the midplane of the core
showing the OD in the as-polished condition, showing a
grain boundary network by reliéf polishing, 4 or 5 mils
deep. Figure 13.15 is the same area in the etched
condition showing the modified etching characteristics
of the surface layers.

13.4 EXAMINATION OF THE
SAMPLER ASSEMBLY

B.McNabb  H. E. McCoy

One of the Y-in.-diameter rods from the sampler
enricher cage reported elsewhere® was sectioned for
metallography and fission product deposition analysis.

5. E. L. Compere and S. S. Kirslis, MSR Program Monthly
Progress Report December 1970 and January 1971, MSR-71-13,
p. 16 (internal memorandum).

A piece was cut from the bottom of the rod in the
liquid or fuel salt region of the pump bowl and one at
the top in the vapor or gas region above the liquid level
Figure 13.16 shows a photomicrograph of the rod
above the liquid level or in the vapor region of the
pump bowl. There is very little if any attack in this
area, either as polished or etched. Figure 13.17 is a
photomicrograph of the bottom of the rod in the liquid
region. There was quite a bit of deposited material on
the lower end of the rod in the liquid region, but there
was little attack on the rod as shown in the micrograph,
unless it was even removal of all constituents of the
surface. Electron microprobe examination and electro-
Iytic dissolution of the surface layers are planned for
one of the rods, and the electrolytic dissolution of the
specimen from the mist shield in the liquid and vapor
regions is planned for determination of fission products
and the deposited material on the rods.

151

Shaas
o

0.007 INCHES
500X

o

™

Fig. 13.14. Photomicrograph of the cross section of the Hastelloy N control rod thimble showing the outside diameter in the
as-polished condition.

\

0.007 INCHES
T
500%

©
~ - a
- \
-~ ¥ ls
! A
i Kf e i)
S ,
Sl | fi
1 g4 ; °s - \\
{ {
F I xPe co@ D> O - \\ o il
y . e \.

Fig. 13.15. Photomicrograph of the cross section of the Hastelloy N control rod thimble showing the outside diameter in the
etched condition. Etchant: aqua regia.
0.035 INCHES
T> 100X

(a)

o010 ™.

0.035 INCHES
T 100X

foo30m.

Fig. 13.16. Photomicrographs of a Hastelloy N rod from the sampler assembly located above the normal salt level. (a) As
polished: (b) etched with aqua regia

153

(a)

too0 .

0.035 INCHES
100X

T

to030 .

Fig. 13.17. Photomicrographs of a Hastelloy N rod from the sampler assembly. Normally located below the salt level. (a) As

polished: (b) etched with aqua regia.
13.5 EXAMINATION OF A COPPER
SAMPLE CAPSULE

B. McNabb H. E. McCoy

A copper sample bucket that was probably in the
sampler cage of the pump bowl from August 1967 to
January 1971 was examined. It had been flattened,
probably during retrieval attempts in April 1968, when
heavy magnets were used with no cushioning effects of
salt and with the pump bowl at high temperature.
Figure 13.18 is a photomacrograph of a cracked area of
the sample capsule showing some deposited material on
the outside of the capsule. Figure 13.19 is a photo-
micrograph of the cracked area polished down below
the surface, showing the intergranular nature of the
cracks and the extent of them. The capsule was made of
OFHC copper and hydrogen fired 2 hr at 1200°F
before use. Figure 13.20 is a photomicrograph of a
cross section of the copper capsule, showing the thick
deposit of material on both the OD and ID of the
capsule.

An electron probe microanal
grain boundaries and x-ray scans for the distribution of
nickel, iron, chromium, and copper in the surface layer

sis of the material in the

154

Fig. 13.18. Photograph of the crack formed in the copper
sampler that was retrieved from the pump bowl. The outer
surface was coated with deposited material.

AW U

R-54647
< g N

)|

Fig. 13.19. Section of copper capsule showing extensive intergranular cracking.

155

PHOTO 1183-71 in the area around the large metallic particle on the OD

were made. Figure 13.21 shows the distribution of

these elements, and it appears that the surface layers are

made up mostly of copper and nickel with some areas

containing high concentrations of chromium and mo-

lybdenum. The ID appears similar to the OD of the

capsule. The grain boundary precipitates were only 28%

copper, and no other element could be detected. This

observation indicates that the remainder is a light

element that could not be detected by the instrument.

Beryllium and lithium are possibilities for the unidenti-
fied element.

The surface deposits were mechanically ground away

e 3 2 S from two small pieces about %4 in. square by % in.

' : / == long for electrodes for the mass spectrograph. We do

- ; -~ not know the exact depth of sampling, but suspect that

it was near the surface. The analysis of these is shown in

_ Table 13.4. Several important points can be made from

S5 this analysis. The spectrograph was not adjusted to

detect lithium, but the presence of uranium, zirconium,

e beryllium, and fluorine indicates that the fuel salt was

present. More specifically, the isotopic concentrations

of U correspond to those of the first fuel charge rather

than the later fuel charge, where the uranium was

predominantly 2**U. The weight percents of various

elements present in the initial fuel charge were approxi-

4 mately 10.9 Zr, 6.3 Be, 67 F, 5.1 U,and 10.8 Li. Using

: U as a base, Be should be higher by a factorof 2.1,and F

higher by a factor of 13. The analytical values in Table

13.4 show the factors based on U to be 0.7 for Be, 6.6

for Zr, and 0.7 for F. Thus, using U as a base, the

material in the Cu capsule is enriched in Zr and

Table 13.4. Composition of copper capsule

Bement  Coneentration || L Concentration

(ppm) (ppm)

As 1 Si 70

Be 1000 sn ~20

Ca ~5 5t -

Cr 30 Te <s

c P 233y 3

Fe 300 2353 500

% ~5 o 1000

Mg 10 v 0.3

Mn \ Zn 6

e 30 Z 10,000

L ! F ~1000
Fig. 13.20. Photomicrographs of copper MSRE fuel sam- N 4500
pler showing layers on OD (top) and ID (bottom). ;b ;

CuKa

FeKa X-RAY SCAN Crka

Fig. 13.21.
photomicrograph showing the deposi

X=RAY SCAN

X-RAY SCAN

156

Y-106552A

MoLa X=-RAY SCAN

iews of the surface of the copper sampler in the electron probe microanalyzer. The upper left picture is an optical
d metal crystals. The other pictures show the intensity of x:

fluorescence for various ele-

ments. These displays are a mirror image of the first optical picture. Note that certain areas are enriched in Ni, Fe, Cr, and Mo. Some

copper from the capsule is present in the deposited material.

depleted in F and Be. There are significant quantities of
Fe, Ni, Mo, and Cr, the main elements in Hastelloy N.
There are also detectable quantities of several other
elements, many of which may have been present in the
copper initially

The copper is currently very brittle, and the present
analytical results do not shed much light on what
element is responsible for the embrittlement. Further
samples will be analyzed deeper into the material where
the salt did not penetrate in an effort to answer this
question.

13.6. EXAMINATION OF THE PRIMARY
HEAT EXCHANGER

B.McNabb  H.

McCoy

An oval piece of the heat exchanger shell approxi-
mately 10 X 13 in. was cut out for examination. Figure
13.22 is a photograph of the inside of this piece
showing the thin bluish-gray coating probably caused

Fig. 13.22. Section 10 X 13 in. cut from the primary heat
exchanger shell. The piece was cut with a plasma torch, and
the center stud was used for guiding the torch.

by the cutting operation. The bottom of the stud used

for guiding the cutting 'operation shows in the center,

with an area around it where the cutting fluid used in
threading the hole has washed away the powdery
coating. Some of this coating was brushed off and
gamma scanned in HRLEL. It. contained '°®Ru and
125Gk, but a detailed analysis has not yet been made.

Two '%-in.-diam cylinders were cut from the shell,
one for metallographic examination and one for fission
product deposition examination, which is presently
being conducted. Figure 13.23 shows photomicrographs
of the inner surface that was exposed to fuel salt. There
appears to be little attack, but there is a surface
modified layer with different etching  characteristics
approximately 1 mil deep. The shell was fabricated
from heat 5068, a standard air-melted heat with
chemical composition of 16.5% Mo, 4% Fe, 6.45% Cr,
0.45% Mn, 0.05% C, 0.58% Si, 0.03% P, 0.008% S,
0.02% Cu, 0.1% Co, 0.27% V, 0.01% Al, 0.01% Ti, Ni.
Figure 13.24 is photomicrographs of the OD of the
shell showing the oxidation of the surface and the area
under the oxide that etched differently to a maximum
depth of approximately 10 mils, likely due to depletion
of chromium from the surface.

Figure 13.25 is a photograph of the six heat ex-
changer tubes as they appeared after removal. The
longest tube was cut by the plasma torch when the shell
was cut through, and the others were cut by an abrasive
cutoff machine. The tubes had a light powdery coating
on one side similar to the inside of the heat exchanger
shell. This would be the case if the deposit was caused
by the plasma torch cutting operation, since the sides of
the tube away from the torch would have been
shielded. :

One of the tubes was cut up for metallography and
fission product analysis. Figure 13.26 shows a longitudi-

157

nal section of the tube OD (exposed to fuel salt) in the
as-polished and etched conditions, and Fig. 13.27 shows
the transverse section of the tube. Note that the OD in
the as-polished condition appears to have a grain
boundary network near the surface to a depth of 3 or 4
mils, but this- was not present in the as-polished ID
exposed to coolant salt.

Tensile tests were conducted at 25°C on some of
these tubes, and Fig. 13.28 shows photomicrographs of
a longitudinal section of the tube wall near the fracture.
Note that fissuring or cracking occurred to a depth of 4
to 6 mils only on the OD which was exposed to fuel salt
and fission products. The ID of the tube etched
differently to a depth of ~5 mils, but this might be
attributed to fabrication history similar to that of the
radiator tubing;® we have suggested that lubricants
containing carbon left carbon residues that were dif-
fused into the tubing and caused carbide precipitation
during the operation of the heat exchanger. The
inhibited grain growth on the ID also supports this
premise.

The heat exchanger tubing was ', -in.-OD by 0.042-
in.-wall Hastelloy N, heat N2-5101, with chemical
composition of 16.4% Mo, 6.9% Cr, 3.9% Fe, 0.45%
Mn, 0.06% C,0.60% Si, 0.001% P, 0.009% S, 0.01% Cu,
0.10% Co, 0.33% V, 0.06% W, 0.01% Al, 001% Ti,
0.006% B. The results of tensile tests on tubes in the
as-received condition and those removed from the heat
exchanger are summarized in Table 13.5. The strength
parameters and the fracture strain were slightly lower
for the tubing removed from the heat exchanger. The
property changes are rather small and are likely due
primarily to the precipitation of carbides during service.

6. H. E. McCoyh and B. McNabb, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, pp. 120-26.

Table 13.5. Tensile data on heat exchanger tubes (heat N2-5101)

. Crosshead iel i
Specimen Condition ect - T ol pge  Frewe  Reduonin
’ (in./min) (psi) stress (psi) ¢
6 From heat exchanger 0.05 650 44,300 67,400 21:3 22.4
5 From heat exchanger 1.0 25 66,300 118,000 37.0 20.5
4 From heat exchanger 0.05 25 64,800 122,000 39.0 29.0
3 As received 0.05 650 53,900 74,600 40.0 14.5
1 As received 0.05 25 73,000 127,400 51.1 426
2 As received 0.05 25 70,900 126,200 50.1 40.0

158

=

0.007 v

=

(a) =

=

0.007 INCHES
500

Fig. 13.23. Photomicrographs of the inside surface of the primary heat exchanger shell. The surface was exposed to fuel salt
and the modified structure to a depth of about 1 mil is apparent. (a) As polished; (b) etched with aqua regia

159

0.007 INCHES

&

(a) Lz

0,007 INCHES

Fig. 13.24. Photomicrographs showing the outside edge of the primary heat exchanger shell. This surface was exposed to 2 to
5% 0, and Na. (a) As polished, showing the selective oxidation that occurred; (b) etched view showing that the metal in the
oxidized layer was completely removed. Etched with aqua regia

R-54185

Fig. 13.25. Photograph of the '-in-OD Hastelloy N tubes
from the primary heat exchanger. The different shades arise
from a dark film that is thought to have been deposited when
the shell was being cut. The film was deposited on the side of
the tubes facing the shell

13.7 EXAMINATION OF FREEZE VALVE 105
B.McNabb  H. E. McCoy

The MSRE utilizes several freeze valves where a
section of tubing is air cooled to freeze a salt plug that
prevents salt or gas flow through the pipe. A schematic
of such a valve is shown in Fig. 13.29. The tubing is
usually flattened, and a thin housing for the coolant
(air) is welded around the flattened section. The
portion of the housing that is parallel to the tubing is
kept thin so that it will accommodate the differential
expansion between the pipe and the housing when the
valve is being thawed or frozen.

One of these valves (FV 105) failed when the MSRE
was shut down in December 1969. This valve was in a
1%-in. pipe, and the failed portion is shown in Figs.
13.30 and 13.31.

The salt was partially removed to reveal the crack
shown in Fig. 13.32. The crack begins at the spot on
the weld and proceeds about 1 in. parallel to the weld.

A metallographic sample was taken of the cracked
area. An as-polished view is shown in Fig. 13.33 and an

160

etched view in Fig. 13.34. The weld was on the tubing
OD and supported the Y-in. end plate of the cooling
shroud. The crack began near the weld outside the
coaling shroud and penetrated the tube wall. A com-
parison of the etched and unetched views shows that
some attack occurred along the crack and the tubing
OD and that the metal was completely removed from
the attacked region by etching. The attack was likely
due to the simultaneous exposure of the Hastelloy N to
salt and moist air. The attack likely involved the
solution removal of Cr and Fe, leaving metal that was
heavily attacked by the etchant. No such attack
oceurred on the ID, where only salt was present.

The failure that released salt was due to thermal
fatigue. The cooling shroud was initially 0.020 in.
thick, but the added cooling tubes increased the
thickness to 0.083 in. on the bottom side. This made
the shroud relatively rigid on the bottom side. During
freezing and thawing. the outer part of the shroud
changed temperature more rapidly than the wall of the
salt-containing tube. Whereas the outer part of the
shroud was originally thin enough to deflect to accom-
modate the differences in length of the two members,
the repairs made the bottom portion quite rigid. The
result was that differences in temperature imposed a
stress on the cooling shroud that was transferred by the
rigid (%-in.) end plate to the tube wall. A crack was
nucleated at the surface and propagated through the
pipe wall during the numerous cycles.

Parts of three lubes are visible in Figs. 13.30 and
1331. The large tube that is in relatively good
condition is the type 304 stainless steel air inlet tube,
the center tube is the original Hastelloy N air outlet
tube with two thermocouples visible, and the hole is the
remains of the type 304 stainless steel outlet tube, The
original Hastelloy N air inlet line was capped off and is
hidden by the salt. The attack of the type 304 stainless
steel by salt when air was present is as expected. The
relative nobility of Hastelloy N in this environment is a
strong argument for the use of Hastelloy N where salt
may be present.

161

[—

=

T~

7 INCHES

500%

Fig. 13.26. Photomicrographs of a typical longitudinal section of the Hastelloy N heat exchanger tubes. Outside surface, exposed
to fuel salt. (a) As polished; (b) etched with aqua regia

162

=y

=

0.007 INCHES

o

&

:

=

0.007 INCHES
500X

™

Fig. 13.27. Photomicrographs of a typical cross section of the Hastelloy N heat exchanger tubes. Outside surface, exposed to
fuel salt. (@) As polished; (b) etched with aqua regia.
163

R-54783
/

tooi0m.

0.035 INCHES

To010 .

0.035 INCHES
100X

too30m

Fig. 13.28. Longitudinal views of a heat exchanger tube pulled in tension at 25°C. The cracking is on the OD where the tube
was exposed to fuel salt. (a) As polished; (b) etched with aqua regia
ORNL-DWG 71-2450R
1% in. SCHEDULE 40 PIPE

~00239-in. THICK
SHIM STOCK //
- 0005-in THICK /N i )
. el L Y
2
“INSULATION \0.125-in. SHEET

L7 in-WALL TUBING

THERMOCOUPLES ~__ — Y5-in, ODx 0042~

(a) ORIGINAL DESIGN

THERMOCOUPLE
__— PATCH PLATE

LOCATION
OF FAILURE

| 0800-in 0D x0020-

LTJ’ in-WALL TUBING

AR INLET OUTLETS

{6) MQDIFICATION

Fig. 13.29. Diagram showing possible cause of freeze-valve failure after modification.

R-54204

Fig. 13.30. Photograph of the bottom portion of freeze
valve 105.

Fig. 13.31. Closeup of the bottom of freeze valve 105 showing the two gas outlet lines and one inlet line of type 304 stainless
steel that was corroded completely. The crusty substance is the salt. Note that there is no salt in the air annulus.

Fig. 13.32. Photograph of the weld at one end of the cooling shroud. The crack begins at the round salt stain and proceeds for
. about 1 in. along the pipe. The arrow points to the salt stain with the crack at 90° to the arrow.

PHOTO 118471

Fig. 13.33. Photomicrographs of tube wall of freeze valve 105,
where the failure occurred. As polished, 50X . Reduced 18%

166

PHOTO 1485-71

Fig. 13.34. Photomicrographs of tube wall of freeze valve
105, where the failure occurred. The weld supports the end of
the cooling shroud, and the failure occurred outside the
cooling shroud. Etchant: aqua regia. 50X . Reduced 24%

14. Graphite Studies

W.P. Eatherly

The purpose of the graphite studies is to develop
improved graphite suitable for use in molten-salt reac-
tors. The graphite in these reactors will be exposed to
high neutron fluences and must maintain reasonable
dimensional stability and mechanical integrity. Further-
more, the graphite must have a fine pore texture that
will exclude not only the molten salt but also gaseous
fission products, notably ***Xe.

The general evaluation of commercial grades of
graphite has been completed. Vendor materials are still
being irradiated where the material appears immediately
promising for reactor application or is of technical
interest because of raw material or fabrication tech-
nique. '

The program on graphite fabrications within our own
laboratory is accelerating. Raw isotropic cokes and
pitch binders are currently the basis for the studies.
Slurry techniques of mixing are being utilized. The
resulting molded and graphitized bodies show excellent
densities, good to excellent microstructures, and appear
to approach the monolithic structure desired. From the
chemical point of view, the effect of foreign atoms with
covalent tendencies (S, O, N) on precursor materials
and their coked structure is being evaluated.

Black-based graphites appear to have reasonably good
stability under irradiation at elevated temperatures. A
heat treatment series of black-based carbons and graph-
ites is being irradiated. Data on electrical resistivity and
X-ray parameters are being used to characterize the
degree of crystallinity and its effect on damage.

Irradiation results on pyrolytically impregnated and
sealed samples continue to be mixed. It is becoming

increasingly clear that the substrate is affecting the-

structures, and proper conditions for the pyrolytic
decomposition have not yet been defined.

Lattice distortions around an interstitial cluster have
been calculated on the basis of continuum elastic
theory for future application to the single-crystal
damage interpretation. Also, the phonon dispersion
curves for graphite have been measured by neutron
scattering as a future basis for investigating the
phonon-phonon scattering in thermal transport.

167

14.1 GRAPHITE IRRADIATIONS IN HFIR
C. R. Kennedy

Irradiation of bulk graphite in HFIR is proceeding.
The recent irradiation experiments were primarily to
extend the fluence levels of graphite grades previously
irradiated. The major emphasis has been to determine
the lifetime expectancy of materials derived from
blacks and several new commercial grades which have
structural physical characteristics similar to the Poco
materials. Also included is a series of experimental
mesophase graphites.

The justification for the reexamination of lampblack
grades has been discussed previously.! The fluence
levels on several of these grades have now been
extended to values adequate to indicate life expectancy.
The grades irradiated are given in Table 14.1, and the
irradiation results are given in Fig. 14.1. Behavior of the
lampblack grades is similar to the more conventional
graphites, but with several significant differences. The
first is that the initial densification is considerably more
rapid. The second is that the later expansion rate is
equivalent to or less than that for the best previously
observed graphites. In all cases the materials are
extremely isotropic, and the life expectancy appears to
be greater than the more conventional grades. The rapid
densification rate was anticipated due to the increased
dimensional change rates observed for the less-graphitic
materials. The reduced expansion rates, however, were
surprising and encouraging. They corroborate the earlier
results on black additions to normal filler materials.

In conventional graphites the effect of the binder was
not apparent in the dimensional changes of the graph-
ite. However, in Fig. 14.1 we observe large differences
which appear to be attributable to the binder type. This
is very likely a result of the fact that almost twice the
amount of the binder is required in black-based graphite
compared with conventional coke graphites. This also
probably accounts for the large effects due to heat

1. C. R. Kennedy, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL-4622, pp. 145-48.
ORNL- DWG 71-6976

6
9/
/0
a P g
T-v a A/
£ 2 - Rl
- g -681 L~
[ y / o
z A7 a P
5° T I
pe SA-45 8| 2 -
= — - s / / A
& -2 81 —
! N o././ A-680
A
4 Iy
- L/:
2/1} a
-6 21
0 5 10 15 20 25 (x10%")

FLUENCE (neutrons/cm?)

Fig. 14.1. Results of irradiating black grades in HFIR at
ns°c.

Table 14.1. Lampblack grades irradiated in HFIR

Maximum

Grade Source Filler Binder temperature
O
T-V Y-12 Thermax Varcum 2300
TITX Y-12 Thermax Isotruxene 2300
A-681 Stackpole (Proprietary) Pitch 2800
A-680 Stackpole (Proprietary) Pitch 1000
SA-45 UCC Lampblack Pitch 3000

treatment temperatures, since the blacks are themselves
relatively unresponsive to thermal treatments. This
behavior is at present being studied by the irradiation of
grade A-680 heat treated to temperatures ranging from
1500 to 2800°C. We are also receiving samples of these
materials where the density has been raised to 1.96
g/cm® by impregnation. This will increase the percent-
age of residue carbon similar to the binder carbon and
further dilute the lampblack content.

A new type of graphite made by Great Lakes Carbon
Company has also been irradiated. These materials are
called mesophase graphites, in that their raw materials
are derived from the so-called liquid-crystal state in
pitches. These first results, shown in Fig. 14.2, indicate
a behavior very similar to hot-worked grades. There is
very little densification, with fairly significant aniso-
tropic linear growth of the samples. These materials
might be of considerable theoretical interest if they can
be fabricated with varying degrees of anisotropy.

Three commercial materials, grade HL-18 from Airco
Speer Carbon Company, P-03 from Pure Carbon Com-
pany, and grade 1076 from Great Lakes Carbon

CRNL-DWG 71-6977

i
®, 0 H413
A 5 H414 u

B, 0 H415 /
3 [«

PARALLEL DIRECTION

9
" /
(]
z
T
(] /
I
E .
o / o
Ll-' .
: ™
—_— -
. :———_____-\A
TRANSVERSE DIRECTION \ .
A
T =u
G 21
q 6 8 10 12 (x40°"}

FLUENCE (neutrons/ecm?)

Fig. 14.2. Results of irradiating mesophase graphites in HFIR
at 715°C.

CRNL-DWG T1-6978

3 - d/
HL ~18 /.
2 :”
S y
g = 7
2 HL-18 AND . s o
< {076 L
5 }A/fl_—--ifl—&‘/
T 8 o
= O "N
e I3 it
L ®0 HL-18
| - A
-1 P-03 As P-03
*9 4076 l
-2 21
0 5 10 15 20 25 (x10%")

FLUENCE (neutrons/cm?)

Fig. 14.3, Results of irradiating grades P-03, HL-18, and 1076
in HFIR at 715°C.

Company, have been irradiated because of their mono-
lithic structure, similar to that of the Poco grades. The
results are given in Fig. 14.3, indicating a dimensional
stability potentially similar to the Poco materials. None
of these grades have been irradiated to a fluence level
high enough to give a real indication of expected life,
although grade HL-18 appears to be expanding. All of
these grades are available in meaningful sizes for the
MSBR.
14.2 GRAPHITE FABRICATION
C.R. Kennedy  W. P. Eatherly

For the graphitic materials we have irradiated in
HFIR, two general criteria have been correlated with
good resistance to radiation-induced dimensional
changes. These are isotropy and a monophase structure.
The isotropy .is required to avoid accelerating the
apparent damage in the preferred c-axis direction. The
monolithic-type structure is required to avoid the
weakness of the binder phase and presumably provide
for plastic flow between domains in preference to crack
failure. Further, one would prefer the isotropy to be on
a very fine scale, so the radiation-induced stresses
cannot build up over extended volumes.

These structural requirements also appear to be linked
to high thermal expansion coefficients. That radiation
stability and expansion coefficient should be linked is
not surprising, since both are certainly related to the
angular differences in orientation between neighboring
crystallites. ' '

Ordinarily graphites are fabricated from calcined or
graphitized coke flour and an aromatic binder, com-
monly pitch. The flour is relatively stable in volume
during heat treatment, whereas the binder will tend to
shrink 60 to 70% in volume during graphitization. This
obviously leads to microcracking in the binder phase
with subsequent weakening of the entire structure. An
obvious partial solution is to employ uncalcined cokes,
which themselves shrink about 35% in volume. An

additional attraction is that such cokes are chemically-

169

active and react with the binder to produce a more
monolithic structure.

The binder-filler mismatch can be further alleviated
by the use of superfines (coke dust) or blacks to fill the
interstices of the coarser flour. These additions to the
filler have some effect on the binder demand but are
more important in decreasing the web thickness of
binder between the filler particles.

In view of these considerations, it was decided to
concentrate our fabrication program on raw air-blown
(Robinson) coke and Thermax as filler materials. The
coke is derived from oxidized asphaltines and is
isotropic over a submicron size domain. The Thermax
was chosen mostly for convenience, since it is relatively
inert on heat treatment.

For binders we have elected to use pitch, partially
because it is a well-characterized material and partially
because it is highly aromatic. Slurry blending is em-
ployed both for convenience and because it permits
highly uniform blending and working of the binder into
the pores of the filler particles. Benzene is being used as
the slurry agent and is evaporated before the blend is
molded. ,

Early studies have concentrated on determining the
proper ratios of black to coke. Typical results are
shown in Fig. 14.4, where the Thermax-to-coke ratio is
varied but the pitch content is held fixed at 30 parts per
hundred of dry blend. The maximum in density is an
indication of optimal packing between the black and
coke flour components and for this binder content is
1.85 g/fem?.

ORNL-DWG 7{-6979

2.2 ’ ‘
o 0.925 in. DIAM
2.0 a 1.60 in. DIAM
GRAPHITIZED 4 CALCULATED
e 8 1 6
Sis > A— —
3 n-.-.-\\
% 1.6 BAKED R —
ul "‘-,_,""'--.,____..
o B g S M.
f _--—‘___.—-‘*.\:_‘.\
c:)n 1.4 (/ % \.:'-:-:.:é’
/3/1 GREEN
1.2 //5
ol
0 0] 20 30 ‘ 40 50 60 70 80 20 100

THERMAX (wi%)

Fig. 14.4. Bulk densities of Robinson coke—Thermax graphites.
The microstructure of the graphite is quite good,
exhibiting no clearly identifiable binder phase although
some segregation of the Thermax is apparent. The pore
texture is small and uniform, although not as good as
‘we have seen in some commercial grades.

Samples will be characterized more fully when the
process is closer to optimum and the material more
uniform. Preliminary samples are currently being irradi-
ated in HFIR.

14.3 GRAPHITE DEVELOPMENT — CHEMISTRY
R. A. Strehlow

Application of high-temperature centrifugation to
pitches which was reported earlier? was found to
produce two distinct types of carbon structures. The
denser structure had a microscopic appearance of highly
isotropic fine-grained material, while the supernatant
liquid produced typical needle-like structures. This
method of characterizing the orientation of product
carbons is potentially of great utility in attempting to
maximize the isotropicity of the carbons. It has been
found by Labaton, Jenkins, and Kilner® that the
presence of hetero atoms in various carbon precursors is
useful in developing isotropic graphite structures. Het-
ero atoms are defined as atoms other than carbon and
hydrogen present in a pitch or coke. Hetero-atom
concentration (Ng, N, and Ny) is defined as the
number of hetero atoms (S, O, N typically) divided by
the total number of atoms in a sample of material.

Experiments have been begun in the attempt to
prepare suitable mixtures by hetero-atom additions for
characterization by the centrifugation technique. The
principal objective of this part of the chemical study is
to determine the extent to which chemical alteration of
the binder system might be used to produce the dense
highly isotrépic carbons needed for molten-salt reactor
applications. Sulfur, oxygen, and, to an extent, com-
pounds with five-member rings have been used by
Jenkins et al.> to increase both isotropicity and coking
yield as well as to minimize the volumetric shrinkage
during coking. The work reported here concerns the
reaction of sulfur with a soft coal tar pitch (type 15V,
Allied Chemical Company). Sulfur was chosen as a first
material of study since earlier work with decacyclene
showed that a sulfur content of even 0.8% was found to
decrease the acicularity (needle-like or platy structure)

2. R. A. Strehlow, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL-4622, p. 135.

3. V. V. Labaton, M. J. Jenkins, and T. Kilner, The Role of
Hetero-Atoms in Carbonization, TRG Report 1738(c) (Aug. 9,
1968).

170

of a derived carbon relative to that formed from a
sample with less sulfur impurity. A study of desulfuriza-
tion by heat treatment at the low temperatures of 150
to 250°C was undertaken.

The phenomenology of sulfur reaction with aromatic
high-molecular-weight compounds has not been exten-
sively studied, but the reaction is reported to be similar
to that with oxygen in that quinone-like materials are
among the products. Polymerization occurs, yielding
mono- or disulfides, along with hydrogen sulfide as the
primary gaseous product. At higher temperatures fur-
ther molecular condensation along with continued
evolutions of hydrogen sulfide occurs. At temperatures
in excess of 1000°C carbon disulfide is evolved.

The changes in fluidity which affect the usefulness of
a modified binder as well as the ease of ultimately
removing the sulfur (for nuclear reasons) are dominant
questions with which this study must deal.

For this work batches of pitch and sulfur of various
compositions are heat treated in an inert atmosphere
for periods up to 100 hr. An ordinary 1-liter resin kettle
is used with a mantle heater. Nitrogen is used as a cover
gas, and mechanical stirring is employed in order to
determine the best heat treatment and composition for
preparation of a sample for centrifugation. A series of
preparations was conducted, some of which are shown
in Table 14.2.

Samples were taken periodically during the prepara-
tion and submitted for carbon, hydrogen, and sulfur
analyses. The product was assessed for its benzene-
insoluble content according to ASTM procedure D23-
17. The results of these measurements are shown in
Table 14.3. (The usual reproducibility of this measure-
ment has been found in our laboratory to be better
than £0.2%.) This benzene-insoluble fraction frequently
parallels the coking yield of a carbon precursor.

Results of chemical analyses are shown in Fig. 14.5
for preparation of RKB-2 and RKB4. They show a
slow decrease of rates calculated as d log N /dt. The
RKB-3 preparation was terminated after 40 min be-
cause it produced very quickly a highly viscous material
which prevented the stirrer from operating and, conse-
quently, frothed extensively.

The values of the rates for the desulfurization of the
preparations analyzed so far are puzzling, since it
appears that the greater the concentration of sulfur the
slower it reacts. A possible key is related to the heating
rates up through the melting point of sulfur to the
softening point of the pitch. For these two batches, the
rate of heating was 1.8 C°/min for the RKB-2 (the more
readily desulfurized) and 2.9 C°/min for the RKB4
171

Table 14.2. Composition, heat treatment, and desulfurization rates
for three pitch-sulfur preparations

Batch No Percent S, Ng Temperature Time Rate,
2 : initial (%) ©C) (min)  d log Ng/dt (hr™")
RKB-2 17.58 5.00 175 + §° 330 8§ x 1073
RKB-3 17.58 5.00 225° 40
RKB-4 9.64 2.32 175 £ 10° 475 1.8 x 1072
ORNL-DWG T71-69B0
18 Table 14.3. Benzene-insoluble (BI) content measured
. for three sulfur compositions
11
\7 Batch No. Bl (%)
BATCH RKB-4
2.0 _ o 2 15V (original material) 14.68
= N | . o : RKB-2 36.05
g \\ ‘\ ® L RKB-3 67.48
” g P RKB-4 30.32
k- \ - ™ =
£  |— BATCH RKB-2 15 &
W \ o :
8 AN yield to reach a maximum at about that concentration.
\\ In summary, sulfur-pitch compositions are being
N prepared for centrifugation at coking temperatures. The
7 desulfurization is being studied as a function of time,
0 100 200 300 . 400 500 .. "
TIME (min] temperature, and composition. In addition to sulfur,

Fig. 14.5. Sulfur remaining in pitch-sulfur composition after
heat treatment at 175°C.

preparation. This possible added parameter will be
examined.

The difference between these rates was reflected in
the  hydrogen-to-carbon  ratios dropping from
CiooHg2.4 to C 4oHsg ¢ for batch RKB4 and to
C,00Hs¢ g for batch RKB-2.

Thin-layer chromatographic analyses of the sulfur
compositions have yielded no indication of additional
low-molecular-weight species. One may infer that poly-
merization reactions rather than substitution reactions
are indeed dominant in the preparation of these
materials. No model compounds have been subjected to
oxidation with sulfur, but the work in ref. 3 showed
that the largest effects are with the complex mixtures
found in pitch. _ '

Fabricability of sulfur compositions exclusive of the
desulfurization question will probably hinge on the
increase of viscosity that accompanies gas evolution
occurring during carbonization. Jenkins® has said that
the forming yield decreases markedly for hetero-atom
concentrations greater than 4%. He reported the coking

4. M, Jenkins, private cornmunication.

other reactants suitable for normal-pressure applications
such as air blowing (an O, treatment superficially
analogous to sulfurization) require study. Other react-
ant systems can, of course, be considered for applica-
tion at higher-pressure carbonizing conditions.
14.4 GRAPHITIZATION STUDY OF A
LAMPBLACK-PITCH CARBON

O.B.Cavin W.H.Cook I.L.Griffith

Until recently no polycrystalline lampblack materials
had been considered in the HFIR irradiation materials
program as potential MSBR graphites because of the
extremely large neutron-induced crystalline growth
rates reported by Bokros.® Then, it was demonstrated
that propylene-derived pyrocarbons are quite stable in
neutron irradiation environment even though the appar-
ent crystallite sizes are quite small.5 Therefore,
Kennedy began an irradiation study of lampblack-based
materials. He observed that a pitch-bonded lampblack
material, grade A681,” exhibited relatively good dimen-

5. 1. C. Bokros and R, J. Price, Carbon 5(3), 301 (1967).

6. D. M. Hewette [l and C. R. Kennedy, MSR Program
Semiannu. Progr. Rep. Feb. 28, 1970, ORNL-4548, pp.
215-18, _

7. Manufactured by the Stackpole Carbon Co., St. Marys, Pa.
sional stability when irradiated at 715°C to a modest
fluence of 1.5 X 1022 neutrons/cm?® (E > 50 keV).®
This encouraged us to examine these materials in more
detail in an effort to correlate starting properties with
the irradiation damage. As previously reported,’® we
have three commercial grades of these materials, SA-
45,'° A681,7 and A680.7 The vendors reported that
the first two had been fired to 2800°C and the last to
~1000°C. The low-fired grade, A680, offered an
opportunity to study the irradiation effects as functions
of its structure resulting from higher firing tempera-
tures. For irradiation specimen stock, Kennedy fired
pieces for 1 hr at 1500, 2000, 2200, 2400, 2600, and
2800°C in purified argon. These samples were then
characterized by bulk density, electrical resistivity, and
x-ray diffraction. '

The effects of the different firing temperatures on the
resultant internal structure are reflected to some degree
by electrical resistivities and bulk densities. These data
were obtained on two sets of 0.25-in.-diam rods
machined parallel with the 6'%- and the 1%-in. dimen-
sions, respectively, from a 1'%4 X 53, X 6% in. plate.
The rods were measured and then fired along with the
HFIR stock to the next higher temperature in sequence.
Therefore, each bulk density and resistivity point
plotted in Fig. 14.6 was obtained on the same rods.
There were slight differences in the values determined
for the mutually perpendicular rods. This suggests that
there is probably a slight preferential particle and
_porosity alignment which may mean the material is
slightly anisotropic.

The electrical resistivity values suggest that between
2000 and 2200°C the crystalline development begins to
accelerate appreciably, and even at 2800°C it has not
saturated.

It has been reported'' that the lampblack-pitch
material, grade A681, has an average bulk density and
an average electrical resistivity of 1.61 gfcm?® and 2900
u§2-cm, respectively, vs 1.63 g/fcm?® and 3050 p2-cm for
A680 fired to 2800°C. These data suggest that grade
A681 has more crystalline development than grade
A680 fired to 2800°C. This was verified by the x-ray
analyses discussed below. It would be of interest to
determine if A680 could be made to approach the

8. C. R. Kennedy, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL-4622, p. 146.

9. W.H. Cook, MSR Program Semiannu. Progr. Rep. Aug. 31,
1970, ORNL-4622, p. 135.

10. Manufactured by the Carbon Products Division of the
Union Carbide Corporation, 270 Park Avenue, New York.

11. W. H. Cook, MSR Program Semiannu. Progr. Rep. Aug.
31, 1970, ORNL-4622, p. 136.

172

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RESISTIVITY: PETROLEUM 5" i
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& a THREE ROD SPECIMENS WITH AXES Il STOCK
o 1¥a DIMENSION .
1000 1400 1800 2200 2600

TEMPERATURE {°C)

Fig. 14.6. Effects of firing temperature on the electrical
resistivity and bulk density of the lampblack body, S20.

relative crystallinity of A681 by firing it to 3000°C or
higher.

A plot of the electrical resistivity as a function of
firing temperatures for a pitch-bonded petroleum coke
material reported by Okada,.Sekiguchi, and Ishii'? is
shown for comparison. Both show the characteristic
stepwise decrease in the vicinity of 2000 to 2400°C,
and, of course, the petroleum-coke—pitch shows the
greater propensity to graphitize. This is particularly true
at temperatures above 2300°C.

The samples for x-ray diffraction, approximately
0.030 in. thick by %, in. square, were machined from
the HFIR specimen stock. Each was a separate sample
for the particular temperature considered.

The step-scan data were obtained by using a General
Electric XRD-5 spectrogoniometer and single-
wavelength copper irradiation from a doubly bent
graphite crystal monochromator. The data were cor-
rected for background scatter, sample transparency,

12. J. Okada, A. Sekiguchi, and T. Ishii, “Effect of Rapid
Heat Treatment on the Properties of Carbon,” pp. 497-502 in
Proceedings -of the Fifth Conference on Carbon, vol. 1,
Macmillan, New York, 1962.
Lorentz polarization, structure, and temperature fac-
tors. Low-fired graphites are poorly crystalline and
generally do not show more than the (002) diffraction
maximum, which is very broad. Hence we could not use
the more sophisticated Fourier analysis approach for
determining the crystallite sizes, particularly of the
low-fired material. To make all crystallite size determi-
nations comparable, we used the peak width of the
(002) and the Scherrer equation:

_ 0.9A
L= Bcos8’

where
A = x-ray wavelength, A,

B = corrected peak width, radians, at one-half max-
imum intensity,

6 = Bragg diffraction angle.

The c-axis determination was made primarily from
the (002) peaks, but the (004) was also used when
sufficiently intense. Calculations for the g axes were
made using the (110) peaks and were found to be
relatively constant at 2.45 A. Relationships between the
¢ spacings and crystallite sizes as a function of
temperature are shown in Fig. 14.7. On the same .plot
are shown data obtained by Okada et al. from a
petroleum coke.'? Petroleum coke graphitizes more
easily than lampblacks; therefore the ¢ parameter of the
A680 never reached the minimum value obtainable for
well-graphitized material. A close parallel between the
two materials is noted, however. Growth rates of the
crystallites (L,) are considerably retarded until a
temperature of 1500°C is reached, after which the rate
increases up to 2000°C but remains constant at higher
temperatures. The ultimate crystallite size obtained was
determined by the heat treating témperature and did
not approach a limiting value over the range tested.
There is no observable discoritinuity in the apparent
crystallite size to correspond with that in the lattice
parameters.

The values shown at 3000°C are from a different
lampblack, grade A681, but these fit the extrapolated
portion of both curves.

As soon as data are available, the irradiation behavior
of these materials will be compared with these data. A
check of the crystallographic anisctropy. will be made
to see if any change can be correlated with the
-maximum rates of change with temperature above
2300°C.

These studies will be enlarged if the results are useful
toward interpreting the irradiation damage effects and

173

ORNL- DWG 71— 6453R

7.02 . 320
6.98 280
. f 3
. r4
\ I,

C - /

6.94 | \ #—1 240
<>
© =\ Ne——A680 LAMPBLACK \/

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6.78 / / o~ 80
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=

g

-

<

L., APPARENT CRYSTALLITE SIZE ()

6.70 0
1000 1400 1800 2200 2600 3000

TEMPERATURE (°C)

Fig. 14.7. Lattice paramétets and crystallite sizes as a

“function of heat treating temperature of two low-fired carbons.

the higher-fired specimens' continue to show relatively
good dimensional stability through fluences of at least
3.5 X 1022 neutrons/cm? (E > 50 keV). Additional
incremental measurements will be included such as pore
entrance diameter spectra, gas permeability, strength
plus crystallographic anisotropy, and crystallite sizes as
determined with X rays. |

14.5 REDUCTION OF GRAPHITE PERMEABILITY
BY PYROLYTIC CARBON SEALING

C. B. Pollock

Graphite to be used in the core of an MSBR must be
able to exclude fluoride salts and pgaseous fission
products. We have been studying techniques to seal the
surface. of the graphite with pyrolytic carbon. Two
techniques have been used: a vacuum pulse impregna-
tion technique in which surface pores of the graphite-
are closed by plugging them with pyrolytic carbon and
a coating procedure in which a continuous surface
coating of impermeable pyrolytic carbon is deposited
on the graphite. Both techniques have been successful
in sealing commercially available graphite suitable for
use in the core of an MSBR.
174

Table 14.4. Impregnated samples

Sample No. Permeabi;ity before PermeabQility after Fluen;:e . AL/L
(cm*/sec) (cm*/sec) (neutronsfcm* X 10¢%) (%)
1231 2.3% 107° 5.1x107° 12.11
1208 1.3x107° 22x 1078 8.76
1236 22x 1077 1.0x 107 7.82
1211 1.5x 107° 6.2x 1077 12.45
HR 20 1x107° 23x 1078 5.59
HR12 1x 1078 6.5% 1078 11.35
1205 4.4x%x 10710 7.2% 107° 17.06 0.7
1216 59x10°1° 9.8 % 107° 21.19 2.2
HL32 1.6%x 1078 9.1x107° 20.60 0.8
1182 1.7x 1078 1.5% 1072 31.24 6.0
1181 5.1x 107 9.0%x 1072 36.97 11.0
1163 7.0x 107° 29x 107 25.67 2.0

In conjunction with the fabrication program we have
also been conducting an irradiation testing program in
which we subject graphite samples that have been sealed
with pyrolytic carbon to MSBR conditions of neutron
fluence and temperature. An HFIR irradiation experi-
ment containing a number of coated samples and
impregnated samples was recently completed. This
experiment contained 10 graphite samples coated with
pyrolytic carbon and 12 graphite samples that were
impregnated with pyrolytic carbon.

Turning first to the coated samples, four of these have
been cycled twice through HFIR and now have a total
neutron dose of 2.4 X 10%? neutrons/cm? (£ > 50
keV) at a temperature of 715°C. Three of the samples
appeared to be intact except for some small cracks on
their ends, but independent of this, the helium permea-
bilities went from less than 10~ cm?/sec to greater
than 107> cm?/sec. The remaining sample was exten-
sively cracked over its entire surface. The remaining
coated samples have been cycled through HFIR for the
first time and received total neutron doses of up to 1.3
X 10%2 neutrons/cm? (£ > 50 keV) at 715°C. Five of
these samples cracked badly and had helium permeabili-
ties of greater than 107 c¢m?/sec. One of the samples
appeared to be unaffected by the experiment.

The only significant difference between these two sets |

of coated samples is the coating thickness. The first
group had coating thicknesses ranging from 4 to 6 mils;
the second set had coating thicknesses of 2 to 3 mils.
The one sample of the second set that survived had a
coating that was 3 mils thick. One has to infer that

coating thickness is the significant variable, but even a

2-mil -coating should not have been affected by the
irradiation condition of this test.

The results on the 12 pyrolytically impregnated
samples are given in Table 14 4. Six of these have been
recycled in HFIR to fluences of up to 3.7 X 10%%;six
new samples have received only their first irradiation up
to 1.2 X 10%2. The new samples appear to be somewhat
better than previous samples irradiated to similar
fluences in that they are opening up to about half the
rate. The pore size distribution of these samples will be
examined by mercury porosimetry.

It has been observed that some of the better samples
contain relatively little pyrolytic material. The amount
of pyrolytic carbon required for sealing can be con-
trolled by varying process parameters other than time.
Helium permeabilities of less than 1078 cm?/sec have
been obtained in graphite samples whose weight in-
crease was less than 4%. Processing time at 750°C was
only 2 hr. Graphite samples sealed in this manner
should behave more like unimpregnated base-stock
graphite. A number of samples containing the minimum
amount of carbon needed for sealing have been pre-
pared for the next experiment in an attempt to control
the neutron-induced expansion of the graphite.

As noted in Table 14.4, one of the samples has
increased its length by 11%, or substantially more than

~would be expected from the unimpregnated base stock

graphite at similar fluences.

14.6 FUNDAMENTAL STUDIES OF RADIATION
DAMAGE MECHANISMS IN GRAPHITE

S.M.Ohr T.S.Noggle

The in situ studies of radiation damage in graphite
resulting from the displacement of carbon atoms by the
175

PHOTO

1231-71

20 min

30 min

Fig. 14.8. (1011) dark-field micrographs showing the development of damage clusters in graphite irradiated with 200-kV
electrons at 600°C. Examples of black and white spots are outlined by circles. Both black and white spots are found to be dislocation

loops of interstitial type.

electrons in the illuminating beam of a 200-kV micro-
scope had established that the damage clusters pro-
duced are in the form of dislocation loops of interstitial
type. Figure 14.8 shows the development of the damage
structure under irradiation with 200-kV electrons at
600°C. In dark-field images such as in Fig. 14.8 (images
formed from diffracted electrons), the damage clusters
frequently appear as roughly equal numbers of black
and white spots. Stereoscopic study of these spots has
shown that the black and white spots are present as
several alternating layers parallel to the surface of the
specimen. The behavior of this black-white structure

with variation of the diffraction condition has been
found to be consistent with and hence attributable to
the stacking fault present in the prismatic dislocation
loop formed by the clustering of interstitial atoms.
Other aspects of the contrast of these clusters, although
qualitatively consistent with the behavior expected of
interstitial-type dislocation loops, suggest that the strain
contrast is less than would be expected on the basis of
diffraction contrast theory employing the displacement
field derived from isotropic elasticity theory. The stress
and displacement fields for a prismatic dislocation loop
have been derived using anisotropic elasticity theory,

176

and contrast calculations will be made for comparison
with the experimental observations. Close correlation of
the experimental contrast with the theoretically pre-
dicted contrast is necessary for quantitative determina-
tion of the true loop size and for estimation of the
minimum loop size detectable.

The generation of experimental information on the
interstitial clusters for comparison with the theoretical
model which treats the kinetics of the nucleation and
growth of clusters requires accurate measurement of the
experimental parameters of temperature and electron
dose. Since the irradiations are typically carried out on
areas less than 1/1000 in. in diameter, a special
technique has been developed for measuring the current
density in the electron beam. This technique takes
advantage of the magnification of the electron micro-
scope to form a highly magnified image of the beam
incident on the specimen and to systematically translate
this image over a Faraday cup so that the distribution
of current in the beam can be measured.

Experimentally, the technique employs a Faraday cup
mounted in the place of a viewing window of the
microscope such that the cup may be inserted to
coincide with the optical axis of the microscope when
current measurements are made and retracted for
microscope operation. The cup has an acceptance
diameter of 7 mm which allows probing the electron
beam over regions less than 2 X 10™° cm? (electron
beam magnified ~15,000X). The beam is probed
systematically by translating the magnified image over
the Faraday cup using the magnetic deflection system
which is used for centering the illumination during
normal microscope use. The currents in the deflection
coils were measured with a potentiometer, and the
calibration of deflection vs current indicated reproduci-
bility of beam position to better than #1 X 10™° mm.
Measurement of the total beam current at the specimen
level using an insulated specimen holder as a Faraday
cup, and intercepting the full beam in the cup in the
viewing chamber, has given the same total current. This
indicates that in' the absence of scattering and/or
limiting apertures in the microscope, currents measured
at the viewing chamber are an accurate estimate of the
current at the specimen. A further check involves an
integration over the full beam profile which gives within
experimental error the total current measured at the
specimen level.

Figure 14.9 shows current density profiles of the
electron beam as a function of the condenser lens
setting. It may be, noted that the detailed profiles are
sensitive to the condenser lens setting. At the center of
the beam, changes in the current density by a factor as

ORNL-DWG T70-11742

o
o]
Pl

— CROSSOVER

CURRENT DENSITY {amp/cm?2)
O
[o2]

0.4 | C OVER-FOCUS
’D.JJ-'O-. '
- ©
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0 39 - Bl Y
0 5 10 {5 20 25 30

DISTANCE (10 *em)

Fig. 14.9. Electron beam profiles as a function of condenser
lens setting. The profile labeled “crossover” corresponds to a
condenser lens setting which gives a focused image of the
electron source in the plane of the specimen. The profiles
labeled *‘over-focus™ and ‘‘under-focus” correspond to con-
denser lens settings which place the focused image of the source
approximately 5 mm above and 5 mm below the plane of the
specimen, The total current in the electron beam, as well as the
current density profiles, is sensitive to the condenser lens setting
due to changes in angular aperture which accompany the
changes in focal length of the condenser lens.

much as 4 occur for barely noticeable changes in the
brightness and apparent shape of the beam observed on
the fluorescent screen. This experience emphasizes the
need for careful monitoring of the current density in
studies of in situ damage in high-voltage electron
microscopes.

14.7 LATTICE DYNAMICS OF GRAPHITE
R. M. Nicklow  N. Wakabayashi  H. G. Smith

An investigation of the lattice dynamics of graphite is
currently being carried out on very high quality (mosaic
spread about ¢ axis ~0.5°) pyrolytic graphite samples® 3
using coherent neutron inelastic scattering techniques.

13. The samples were obtained by C. J. Sparks of the Metals
and Ceramics Division from the Union Carbide Corporation
Laboratory at Parma, Ohio.
177

The experiments are being performed on the triple-axis
neutron spectrometer at the HFIR. Pyrolytic graphite
consists of thin sheets of graphite crystallites randomly
oriented about a common ¢ axis, and one may expect
that unambiguous measurements of normal mode fre-
quencies by neutron scattering techniques are possible
only for the longitudinally polarized normal modes
with wave vectors in the ¢ direction. Such measure-
ments have been previously reported by Dolling and
Brockhouse'* for a sample of rather poor quality
graphite (5° mosaic spread); however, they also found
evidence of the transverse acoustic mode in the ¢
direction, although the phonon peaks were very broad.
In addition to measurements of the longitudinal pho-
nons in the ¢ direction, we observe well-defined
transverse acoustic phonons in this direction when
appropriate incident neutron energies are chosen. Pre-
sumably, this improvement over the earlier measure-
ments is due to the better sample available for this
study. The initial slopes observed for both of these
c-axis branches are consistent with the elastic constants
Cy, and C, 4 measured by Seldin.!®

We have also observed well-defined peaks in the
neutron scattering arising from the transverse acoustic
and the lowest-frequency transverse optic modes having
wave vectors in, and polarizations perpendicular to, the
basal plane. These measurements have been carried out
for wave vectors

0.50 47

lql < ;
V3 a
and they are possible because the dispersion curves for
these modes are very isotropic. The frequencies of the
small |q| longitudinal acoustic and lowest frequency
longitudinal optic modes with wave vectors in the basal
plane were less well determined. However, the measure-
ments performed strongly indicate that frequencies for
both of these modes increase very rapidly, reaching
approximately 12 THz for igl ~ 0.08(47/an/3 ). These
results are also consistent with the elastic constant
measurements of Seldin.!® The striking feature of these
measured dispersion curves is a very rapid increase in
the frequency with increasing wave vector q.

The analysis of the data has been carried out in terms
of an axially symmetric Born—von Karman force-
constant model. In order to reproduce both the shapes

14. G. Dolling and B. N. Brockhouse, Phys. Rev. 128, 1120
(1962).

15. E. S. Seldin, Proceedings of Ninth Biennial Conference on
Carbon, Chestnut Hill, Mass., June 18-20, 1969, p. 59.

ORNL- DWG 70-2244

[1070]

FREQUENCY (102 cps)

70y
/[TA

TA

oL I N R
04 02 O 01 02 03 04 05

(UNIT : 2m/c,) (UNIT : 47/43aq,)
REDUCED WAVE VECTOR COORDINATE

Fig. 14.10. Calculated and observed dispersion curves for the
I-A and I-M planes of the reduced Brillouin zone.

of the observed dispersion curves and the (4, elastic
constant, it is necessary to include interactions which
extend to the third-nearest-neighbor atoms in the basal
plane. Such a model, while perhaps not as physically
appealing as the usual bond-bending model,!® does
provide a very satisfactory description of these neutron
scattering results, elastic constant measurements,'® and
Raman'? and infrared'® measurements. In its present
form the bond-bending model includes explicitly inter-
actions only to second-nearest-neighbor atoms in the
basal plane (in the terminology used for Born—von
Kdrman models) and gives C44 = 0.

The calculated dispersion curves are shown in Figs.
14.10 and 14.11 together with the corresponding
observed points. The calculated density of states as a
function of frequency is shown in Fig. 14.12.

16. A. Yoshimori and Y. Kitano, J. Phys. Soc. Japan 2, 352
(1956); James A. Young and Jaun Koppel, J. Chem. Phys. 42,
357 (1965).

17. F. Tuinstra and J. L. Loenig, Bull Amer. Phys. Soc.
15(3), 296 (1970).

18. E. Burstein, private communication.
FREQUENCY (10% cps)

ORNL-DWG 70-2245

M K r M
[ | ] [ |
[1120] [to10]
L |
|
|
|
| | —]
|
|
|
- | —
|
| -
I I ——
|
|
¢ |
. ! —
| ()
L e,
)I//// N
- %
7/ | \
Ve | X
7 .
L™ 7 |
7 [
e
b
|1 |
05 04 03 02 O0f o 0.2 0.4
(UNIT: 4m/a,) (UNIT : 47/¥3a,)

REDUCED WAVE VECTOR. COORDINATE

Fig. 14.11. Calculated and observed dispersion curves for the
I'-K-M planes of the reduced Brillouin zone.

178

ORNL-DWG 70-2246

G () (arbitrary units)

l | !

f ]

5

20 25 30
FREQUENCY (10 ¢ps)

Fig. 14.12. Graphite G(v) calculated from four-neighbor axi-

ally symmetric model fitted to neutron data.
15. Hastelloy N

H. E. McCoy

The search for a chemically modified composition of
Hastelloy N with improved resistance to irradiation
damage continues. The elements of primary concern are
Ti, Nb, Hf, Zr, Si, and C. We continue to approach the
problem by initially making small laboratory melts and
then procuring 50- to 100-lb commercial melts. This
gives us some feel for the problems that will be
encountered in scaleup of these alloys to production
size (10,000 Ib or greater). This work involves mechan-
ical property studies on unirradiated and irradiated
samples.

Our compatibility programs are involved with the
corrosion of Hastelloy N in several fluoride salts and in
steam. The salt of primary concern is the new proposed
coolant salt, sodium fluoroborate. Several thermal-
convection and two pump loops are committed to
studying corrosion in this salt. The possibility of
holding tritium up in this salt is also being studied.
Steam corrosion work continues at two facilities.
Hastelloy N is currently being exposed in the unstressed
condition, but one facility is being modified to allow
dynamic stressing of the samples.

15.1 STATUS OF LABORATORY HEAT
POSTIRRADIATION EVALUATION

C. E. Sessions

The influence of alloy composition on the postirradi-
ation mechanical properties continues to be our pri-
mary area of study in developing advanced nickel-based
alloys for MSBR applications. During 1970 our studies
were directed toward evaluation of various alloying
additions on small laboratory heats after irradiation at
760°C to 3 X 10?° neutrons/cm? (thermal} and testing
in creep at 650°C. Since that time we have attempted
to reproduce certain beneficial alloying effects using
commercially supplied 100-1b heats. These latter results
are discussed in a subsequent section of this report.

Table 15.1 gives the range of laboratory heat compo-
sitions that we studied during 1970. The postirradiation
creep ductility is assessed either “poor,” “good,” or
“excellent” depending on whether the ductility was
<5%, 5—10%, or >10% after irradiation and testing.

179

For most of these alloys the creep strength was at least
as good as that for standard vacuum-melted Hastelloy N
under the same set of test conditions. The creep
ductility has been the property of primary concern as
regards possible reactor application. The results indicate
that intermediate Ti levels and Hf concentration above
0.5% provided good ductilities. For multiple additions
of these strong carbide formers the ductilities were
classed as either good or excellent. For heat 287 (0.1%
Ti, 0.6% Nb, 0.14% Si), the ductility was ~1%, and thus
this composition was rated as poor. Based on the
laboratory heat results the most promising alloying
elements for producing excellent postirradiation duc-
tility would be additions of either Ti-Hf, Hf-Nb, or
Ti-Hf-Nb. '

Table 15.1. Results of postirradiation creep tests
of laboratory melts?

o Concentration (%) Creep
Addition  Alloy No. T F Nb S ductility?
Ti 107 1.0 Poor

291 2.0 Good
292 24 Poor
Hf 299 0.5 Good
302 1.5 Good
Nb 285 0.5 Poor
298 . 2.0 Poor
Ti-Hf 309 0.5 04 Excellent
184 1.2 1.2 Excellent
Ti-Nb 287 0.1 0.6 0.14 Poor
303 0.5 0.8 Good
181 0.5 2.0 Good
Hf-Nb 307 0.8 0.9 Good
308 0.6 1.2 Excellent
Ti-Hf-Nb 310 0.1 0.5 0.6 Excellent
311 04 03 0.7 Good
314 08 07 13 0.3 Good

%Base composition is Ni—-12% Mo—7% Cr—0.2% Mn-0.05%
C. All alloys were annealed 1 hr at 1177°C, irradiated at 760°C
to a thermal fluence of 2 to 3 X 1029 neutrons/cm?, and tested
at 650°C.

b«poor” is less than 5% strain, “good” is 5 to 10%, and
“excellent” is greater than 10%.
15.2 POSTIRRADIATION CREEP TESTING OF
HASTELLOY N

C.E. Sessions H. E. McCoy

Because we have developed stronger alloys by the
addition of Ti, Nb, and Hf to our nominal Hastelloy N
composition, we must load to higher stresses in order to
obtain rupture in a reasonable time period. If these
higher stresses are above the yield strength of the
particular alloy, a significant amount of strain can occur
instantaneously on loading. The postirradiation test
loading procedure that we have used in the past is to
apply the stress and then begin to record strain as a
function of time. At stress levels above the yield stress
this method is in error and gives strains that are too
low. The complicating factor that makes it impossible
to record the strain on loading is that the extensometer
in the hot cells actually records as strain the relative
movement between the two specimen grips of the creep
machine. This movement in fact includes some immeas-
urable movement due to the seating of the specimen in
the grips. Then ideally what we would like to do is to
(1) apply enough load to seat the specimen but not
enough to cause significant plastic strain, (2) begin
recording strain, and (3) apply the remainder of the
load. However, the machines are not constructed to
allow this to be done easily.

180

The following example involving alloy 69-641 (1.3%
Ti, 0.7% HIf) stressed at 47,000 psi at 650°C illustrates
the severity of the problem. The three samples shown in
Fig. 15.1 were irradiated in the same experiment at
760°C to a thermal fluence of 3 X 10%2% neutrons/cm?.
In the first test (curve 4, Fig. 15.1) we applied the
maximum load and then began to record strain: failure
occurred in 9.5 hr with 1.3% strain. In the second test
(curve B) we added the 18 weights (needed to obtain
47,000 psi) one at a time in 50-sec intervals, and the
strain was recorded throughout: failure occurred in 8.5
hr with 5.5% strain. In the first test the measured strain
was too low and in the second test it was too high
because it included the movement due to the specimen
seating in the grips. In a third test (curve C) weights to
give 10,000-psi stress were added before we began
recording strain. The additional load was added to give
47,000 psi with the strain being recorded. After one
intentional interruption to measure the strain which
had occurred in 2 hr, we reloaded the test to 47,000 psi
and it subsequently failed in 4 hr with 6.5% strain. This
third loading technique (curve C) has been adopted as
our standard test loading procedure for postirradiation
creep-rupture testing. It offers a reproducible technique
for establishing a “zero” strain reading under a stress of
10,000 psi, which is significantly below the yield stress
of the Hastelloy N compositions that we are investi-

ORNL-DWG 70-8151

.
C/*
6 /
/*
—————_-J
/—_—_ /
"7 s —1
_--/
]
S 4
2 A NORMAL PROCEDURE, STRAIN RECORDED AFTER APPLICATION
E 3 OF 47,000 psi STRESS.
B STEP LOADED TO 47,000 psi IN INCREMENTS OVER 15
MINUTES, STRAIN RECORDED FROM START OF LOADING.
C SAMPLE STRESSED AT 10,000 psi, STRAIN WAS MEASURED
2 AT THIS POINT AND RECORDED AS A FUNCTION OF TIME
AFTER THE STRESS WAS RAISED TO 47,000 psi,
’ *
; /,/
A/
O el
0 1 2 3 4 5 6 7 8 .9
TIME (hr)

Fig. 15.1. Effect of loading technique on the postirradiation strain-time curves for Hastelloy N heat 69-641 at 650°C. Samples
solution annealed 1 hr at 1177°C and irradiated at 760°C to a thermal fluence of 3 X 10%° neutrons/em?.
181

gating in our test program. It also provides us with a
measure of the strain that occurs on loading which can
be significant for certain alloy compositions if the stress
is above the yield stress. ‘

Unfortunately, some of our test results on the

laboratory compositions discussed in the preceding
section were obtained using technique A4 in Fig. 15.1, so
that we missed recording the loading strain. This might
possibly affect the conclusions and relative ductility
values given in Table 15.1. This same difficulty also
influences some of our preliminary results for commer-
cial heats (100-lb size) discussed in the following
sections; but, we have made an attempt to estimate for
the commercial heats the magnitude of the strain on
loading which we failed to record.

15.3 THE UNIRRADIATED MECHANICAL
PROPERTIES OF SEVERAL MODIFIED
COMMERCIAL ALLOYS

H. E. McCoy  B. McNabb

Based on our studies of small laboratory melts, several
small commercial melts of selected compositions have
been procured. These melts all contain various additions
of Nb, Ti, and Hf. The first 13 were double vacuum
melted and were 50 Ib in size. The last 4 were melted
by the electroslag remelt process and were 100 1b. All
were finished to Y-in.-thick plate; their chemical
compositions are given in Table 15.2. The unirradiated
mechanical properties of these alloys were discussed

previously,! but testing has continued, and more
information has become available.

The results of stress-rupture tests at 650°C for the
double-vacuum-melted alloys are shown in Fig. 15.2.
The rupture lives cover about two orders of magnitude
in time at a given stress level. The minimum creep rates
shown in Fig. 15.3 show a similar variation, with all
alloys being stronger than standard Hastelloy N. The
properties of the electroslag remelted (ESR) alloys are
shawn in Figs. 15.4 and 15.5. These alloys all contain
about 0.5% Si, and heats 69-688 (no addition) and
68-689 (0.36% Ti) have properties about equivalent to
those of standard Hastelloy N. Heats 69-344 (0.77% Ti,
1.7% Nb) and 69-345 (1.05% Ti, 0.88% Hf) have
properties superior to those of the standard alloy.

The creep strengths of the alloys at 650°C are
compared in Fig. 15.6, where the stress to produce a
creep rate of 0.01%/hr is shown for each alloy. This is a
relatively high creep rate, but the trends will likely hold
at lower creep rates. Alloy 69-648 (0.92% Ti, 1.95%
Nb, 0.043% C) is the strongest. Three additional melts
of this same nominal composition were made. Alloy
69-714 (0.8% Ti, 1.6% Nb, 0.013% C) is weaker and
clearly illustrates the beneficial effects of carbon. Alloy
70-835 (0.71% Ti, 2.6% Nb, 0.052% C) is similar in
composition to heat 69-648 and has almost equivalent
strength. Heat 69-344 (0.77% Ti, 0.54% Si, 1.7% Nb,
0.11% C) was made by the ESR process and has

1. H. E. McCoy and B. McNabb, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 161.

Table 15.2. Compositions (%) of experimental alloys

Alloy No. Mo Cr Fe Mn Si

Ti Zr Hf Nb C

70-785 12.3 7.0 0.16 0.30 0.09 1.1 0.012 <0.003 0.097 0.057
70-727 13.0 7.4 0.05 0.37 <0.05 2.1 0.011 <0.01 <0.01 0.044
70-796 12.5 7.5 0.054 0.64 0.02 0.04 ~  0.024 0.79 0.04 0.04
69-468 12.8 6.9 0.3 0.34 0.05 0.92 0.005 <0.05 1.95 0.043
69-714 13.0 8.5 0.10 '0.35 <0.05 0.80 0.028 <0.01 1.6 0.013
70-835 12.5 79 . 0.68 0.60 005  0.71 <0.005 0.031 2.60 0.052
70-786 12.2 7.6 0.41 0.43 . 0.08 0.82 0.024 <0.003 0.62 0.044
69-641 13.9 6.9 0.3 0.35 0.02 1.3 0.021 0.40 <0.05 0.05

70-787 12.3 7.0 0.18 0.43 0.09 0.90 0.038 0.77 0.12 0.041
70-795 13.7 8.3 0.035 0.63 0.03 1.5 0.018 0.42 0.005 0.05

70-788 12.1 7.3 0.43 041 0.1 1.4 0.020 0.30 0.67 0.027
70-797 12.7 7.0 0.29 0.38 0.02 0.59 0.040 0.78 0.98 0.049
70-798 13.5 7.9 0.26 0.53 0.02 0.71 0.012 0.28 0.94 0.036
68-688 14,3 7.1 4.6 0.46 0.38 0.01 <0.050 <0.05 <0.05 0.079
68-689 13.7 74 4.6 0.46 0.53 0.36 <0.05 <0.05 <0.05 0.081
69-344 13.0 7.4 4.0 0.56 0.54 0.77 0.019 <0.1 1.7 0.11

69-345 13.0 8.0 4.0 0.52 0.52 1.05 0.038 0.88 <0.01 0.078

182

ORNL-DWG 71-6981

80 Hr— —
70-795 - | \ ] | ‘
70-796
69-641 69-848
] b Q\ +
70 R Xy 70-727
' N "N 70-788
NS T 70-797
N N
. . »
60 AN
DN |
b
~ STANDARD of [ e ) \tx
3 HASTELLOY N N f \ ™\
S 50 N N N
N 70-787 — N
e LloRRey oAt [T
p TR
@ A N
i L, a5
40
u'_') \\ ‘T‘fih -2
N 69-714 "
Ny 70-786
30 AN
o 70-785 & 69-641 \
& 70-727 @ 70-787 \
a 70-796 ~ A 70-795 N\
20 o 69-648 m 70-788 ;
vV 69-714 & 70-797 N
o 70-835 ¥ 70-798
¢ 70-786 .
" RN
10° 10! 102 103 104

RUPTURE TIME (hr)

Fig. 15.2. Stress-rupture properties at 650°C of several double-vacuum-melted experimental alloys (see Table 15.2 for the
compositions). All alloys annealed 1 hr at 1177°C prior to testing.

ORNL—-DWG 71-6982

80 T T 7 T0 711
70-835 I
70- 788 69-714~,
70 5 o SN / B
A 1 A1 /
69-648 Pd A AL // //
Py A b
< > LA LT L
LA
" vl // | A / />~ 70-795
- o | L 4REYL & A | Y —ca
1| A T | A 69-641
S 50 1 // /// A // //
e o / 'z L/ d
= C N
- ) Paird F e 20 | LA sTanoaro
* A / e /] HASTELLOY N
w Py /
x ” / M A
= 40 SRSy T
w L 4 y ‘///
%0 70-798 M
T 70-797 A
70— 796 // (o] .70—785 * 69-641
A a 70-727 ® 70-787
o 70-796 A 70-795
20 ¢ 69-648 m 70-788
. v 69-714 e 70-797
< 70-835 v 70-798
' & 70-786
o NI
1073 1072 107" 109

MINIMUM CREEP RATE (%/hr)

Fig. 15.3. Creep rates of 650°C of several double-vacuum-melted experimental alloys (see Table 15.2 for the compositions). All
alloys annealed 1 hr at 1177°C prior to testing.
80

70

60

STRESS (1000 psi)
)
o

Y
(@]

30

20

183

ORNL-DWG 71-6983

1 N
N N |
N N\ 69-344, 69-345
N N
\ \\
N ™
<> ® b
\\ \\
68-688, 68-689 N
I~
N \\
4 P
STANDARD HASTELLOY N \ \\
N ™
N N
™
"N
N
\ >
* 68-688 \
+ 68-689 \
* 69-344 N
® €9-345 C
10° 102 103 10%

RUPTURE TIME (hr)

Fig. 15.4. Stress-rupture properties at 650°C of several electroslag remelted experimental alloys (see Table 15.2 for the
compositions). All alloys annealed 1 hr at 1177°C prior to testing.

80

70

60

50

STRESS (1000 psi )

40

30

20

ORNL-DWG 7i-6984

/ v/
. /
,'
AT
69- 344, 69-345
T e
s o
' _ -
3! /,/5 68-688, 68-689
/'
i 2841112
v T
ya
rd "4
%
* f 7 p1 ;/ \STANDARD HASTELLOY N
/fl‘/ A / .
ey e r “{ ‘/
L B "/ /rv
-~ ///
- A7
"
b 4
Ediy - 68-688
v /]
S L + 68-689
- /|
_ Y * 69-344
® 69-345
3 102 107" 10° 1of

MINIMUM CREEP RATE (%/hr)

Fig. 15.5. Creep rates at 650°C of several electroslag remelted experimental alloys (see Table 15.2 for the compositions). All
alloys annealed 1 hr at 1177°C prior to testing.

184

ORNL-DWG 71- 3737
80

@®
<
iy
3
70
o)
o
X . 3
MO | 2 ~ ~_®
60 |~ i oO— o
o ™~ o~ M~
F g < LS
TR -
o r~ R % ~ o~ <+ W
. P~ — ! ~ ) q-_w,.._
=~ 50—t 4 H H - o—~—NF—wd H H Mn—m
7] (@) 9 o O ! !
a = < — W K~ [ )]
o — -— ! i ! (s3] )3 (s} <«
o ~ ) o} ol lo @ @ — M
Q @ 2 " 0| | ¢ 9
Z 40 - HH A H HH HHeHYHS o—o—,HOH
” = sl e 6 & |9 o
0 gl o] |o sinlElE
E . o L ol |a| |2 o O_ o
L] el Hopldsd e g2 = IS gl NN -1 S 5, B
. o| |m| |v] |© A W S L o @
| (2118 o] [e] [e] |2 I~ |5 [Z ~| (B
cHiets i 18] el 9] |2 |2 @ Sl 1= IS
Ol (el 19 el 9] 174 |o] [of [« [ ol |H 1412
20 H HoHoHSHZHHeHZHPHEH-H - HEHT m_w_(m_-u—)a
NI 21 12| || [« | || |®]| |o Sl 12 o
A | ol |©o] || I~ M~ @l (9] &
ol |8 ¥ o | |w . . ~| (N2 N o !
el (2] [ 1741 =] [« |2 (=] |s] (=] (2] 18] |e ol 9] |s
| (o] 1] 1= |17 |45 |o i of |~ |.]] =1 =1 = -
10 HOl- ..__H__.._I_|—_i:fii:_.__.:__.__g.__._l—_,_ m__y—__l—_i:_q
Zl [F] F] o] |o — Jof =] 1L E] E] o] |- o lo| I~ |,
— - | ™~ ol | M~ N o | | S B et Ml ™ ™~ o
0| 1< (o] O] (O] |of o] |o] [« o] [ |<] o] o Ol o] O] |=
0]

Fig. 15.6. Stress to produce a creep rate of 0.01%/hr at 650°C for several experimental alloys (see Table 15.2 for compositions).

significantly lower strength. We attribute this lower Table 15.3. Properties of several modified alloys at 704°C
strength to the presence of the coarser MyC-type and 35,000 psi ‘ :
carbide caused by the presence of Si. Alloys containing See Table 15.2 for compositions
titanium were reasonably strong and increased progres-
sively as the concentration of Ti was increased (alloys Rupture  Minimum  Fracture  Reduction
70-785 and 70-727). Alloys 70-786, 69-641, 70-787,  Alley No. (111,13 T @ ey
and 70-795 contained additions of Ti and Nb or of Ti ’ ’ ' ’
and Hf and had relatively low strengths although they Standard 70.0 0.20 16.0
were stronger than the standard alloy. Alloy 70-796 70-785 141.4 .0.111 45.40 45.0
(0.79 Hf, 0.040% C) and alloys 70-788, 70-797, and ;3;3; g;;i g-(l)gs ;ig ;g(l)
70?798- with multiple additions of Ti, Hf, and Nb had 69-648 732:2 0:025 21:6 18:0
quite high strengths. | 69-714 236.6 0.098 44.8 80.0
As shown in Table 15.3 the same trends continued at 70-835 898.1 0.032 62.3 51.0
a test temperature of 704°C. The differences become 70-786 205.9 0.049 57.60 56.2
somewhat less with a variation of one order of 69-641 171.7 0.135 45.6 - 647
magnitude in the rupture life and the minimum creep 70-787 193.4 0.090 38.5 44.9
. 70-795 162.6 0.11 §3.2 70.6
rate at a common stress level of 35,000 psi. The 20-788 408.8 0.072 70.8 64.7
modified alloys are all stronger than the standard alloy. 70-797 550.9 0.043 60.7 57.0
- The difference in properties at 760°C is even less. ' 70-798 302.3 0.061 55.2 64.8
We attribute these property variations to the various 68-688 40.0 0.24 19.8 - 198
carbide distributions that are obtained in these alloys 68-689 65.2 0.267 434 36.7
. . o 69-344 181.6 0.048 41.5 39.0
by solution annealing at 1177°C and subsequently - 69-345 400.3 0.053 457 48.0

testing at 650 to 760°C. The carbide structures depend
upon the composition and coarsen as the test tempera-
ture is increased. This convergence of properties at
760°C probably marks the temperature at which the
carbide precipitate coarsens in all alloys.

15.4 THE WELDABILITY OF SEVERAL
MODIFIED COMMERCIAL ALLOYS

B.McNabb  H. E. McCoy

Some observations on the relative welding character-
istics of commercial heats of modified Hastelloy N were
reported previously.> Mechanical property specimens
were prepared from the welded plates, as described by
McCoy and Canonico.® Side-bend specimens were
prepared by sawing Ys-in. strips across the welds from
the '-in-thick plates, and these were bent around a
Yy-in. radius. Most of the commercial heats bent at
room temperature did not develop any cracks in the
weld metal, but some did develop cracks, as reported
previously. Figure 15.7 is typical of those heats that did
not crack on bending. Figure 15.8 is a photograph of
heat 70-796 and is typical of those heats that cracked
severely on bending. We attribute this cracking to the
combined high Hf and Zr concentrations in those
alloys. Figure 15.9 is a photograph of heat 69-345
(additions of 0.88% Hf and 0.038% Zr) plates that
cracked during welding. Welding was discontinued, and

2. B. McNabb and H. E. McCoy, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4662, pp. 160-61.

3. H. E. McCoy and D. A, Canonico, “Preirradiation and
Postirradiation Mechanical Properties of Hastelloy N Welds.”
Welding J. 48(5), 203-5-211-s (May 1969).

the plate was examined metallographically. The cracks
were only in the weld metal and did not penetrate the
base metal. Both heat 69-345 and 70-796 plates were
welded successfully with a filler metal of standard
air-melted heat of Hastelloy N, heat 5090, having the
chemical composition Ni-16.04% Mo—7% Cr—3.7%
Fe-9.5% Mn—0.56% Si—0.03% C. Figures 15.10 and
15.11 are photographs of these heats after being welded
with this dissimilar filler metal and being bent without
cracking. This demonstrates clearly that the welding
problem is associated with the weld metal and could be
circumvented by adjusting the composition of the filler
metal.

Numerous mechanical property tests have been run
on small transverse specimens from the weld. The tests
are not complete, and we shall only generalize the
trends observed to date.

1. The weld metal has a higher yield stress than the
base metal at 25, 650, and 760°C. The ultimate
tensile stress is not affected much.

2. The fractures usually occur in the weld metal, and
the fracture strain is much lower than for the base
metal.

3. In creep-rupture tests at 650 and 760°C, the weld
samples compared with the base metal have a lower
minimum creep rate, a shorter rupture life, and a
lower fracture strain. The magnitude of these effects
is less at 760 than at 650°C.

Y-98657

INCHES

Fig. 15.7. Side-bend specimens (% in. thick bent around a %-in. radius) of heat 70-727. The alloy was welded with filler metal
from the same heat. Dye penetrant has been applied, and no flaws are visible.

186

Y-102370

INCHES

Fig. 15.8. Side-bend specimens (Y in. thick bent around a Y-in. radius) of heat 70-796. The alloy was welded with filler metal
from the same heat. Dye penetrant has been applied, and several flaws are visible.

Y—98539

INCHES

Fig. 15.9. Plate of 69-345 that cracked during welding. The plate was % in. thick and was being welded with filler metal from
the same heat.

187

Y-105388

INCHES

Fig. 15.10. Side-bend specimens (% in. thick bent around a -in. radius) of heat 69-345. The alloy was welded with a standard
Hastelloy N filler metal (heat 5090). Dye penetrant has been applied, and no cracks are visible.

Y—105387

INCHES

Fig. 15.11. Side-bend specimens (' in. thick bent around a %-in. radius) of heat 70-796. The alloy was welded with a standard
Hastelloy N filler metal (heat 5090). Dye penetrant has been applied, and no cracks are visible.
188

Table 15.4. Effect of postweld annealing on the tensile properties at 25°C2
Al Yield Tensile Fracture
N0°§ Anneal stress stress strain

) (psi) (psi) (%)
5065 Base metal 64,000 125,000 55.5
As welded 81,400 112,000 28.7
8 hr at 870°C 57,700 108,000 31.9
69-641 Base metal 44 900 117,000 72.6
As welded 73,800 114,000 29.6
8 hr at 870°C 65,500 116,000 36.7
1 hr at 982°C 60,000 104,600 27.7
1 hr at 1093°C 47,800 109,700 442
1 hrat 1175°C 45,700 116,800 60.9
69-648 Base metal 44,700 116,900 74.2
As welded 75,300 118,600 333
8 hr at 870°C 65.400 111,800 31.0
1 hr at 982°C 60,100 110,000 33.3
1 hr at 1093°C 46,900 119.900 62.9
1 hrat 1175°C 46,000 112,900 62.4
2Strain rate of 0.05 min~!.
bThe chemical compositions of these alloys in weight % are:
Heat No. Mo Ctr Fe Mn Si Ti Nb Hf Zr C
5065 16.5 7.2 39 0.55 0.60 0.01 " 0.065
69-641 139 69 0.3 0.35 0.02 1.3 <0.05 040 0.021 0.050
69-648 128 69 0.3 034 0.05 0.92 195 <0.05 0.005 0.043

The ability to improve the properties of the weld by
annealing is also different for the modified alloys than
for standard Hastelloy N. A postweld anneal of 8 hr at
870°C improved the tensile and the creep-rupture
properties.> The tensile properties at 25°C of two
experimental heats are shown in Table 15.4. The anneal
of 8 hr at 870°C reduces the yield stress of the weld in
heat 5065 (standard alloy) to that of the base metal.
The same anneal had little effect on heat 69-641, and
an anneal of 1 hr at 1093°C was required for significant
_recovery. Similar behavior was noted for heat 69-648.

The property changes with welding that we have
observed to date vary considerably from heat to heat.
Our data are too fragmentary to date to attempt
conclusions about the effects of composition on these
changes, but this is the goal of our work,

15.5 POSTIRRADIATION PROPERTIES OF
SEVERAL COMMERCIAL ALLOYS

C.E. Sessions  H. E. McCoy

Based on our work with small 2-1b laboratory melts,
several 50- and 100-b commercial melts have been
procured for evaluation. The compositions of these
alloys are given in Table 15.2, and other properties have
been discussed in previous sections. One of the items of

most importance for nuclear applications is how the
mechanical properties change during irradiation. We
anticipate that future systems will operate at about
700°C, but have made our irradiation conditions a bit
more stringent. All of the specimens to be discussed in
this section were irradiated at 760°C and tested at
650°C, a condition that we had found to give low
fracture strains and that is not unrealistic for a
transient. The thermal fluence was 2 to 3 X 10%°
neutrons/cm?, and the irradiation time was about 1100
hr.

The postirradiation stress-rupture properties of thése
samples are given in Fig. 15.12. Alloys 70-785 (1.1% Ti)
and 70-786 (0.82% Ti, 0.62 Nb) had properties about
equivalent to those of standard vacuum-melted Hastel-
loy N. Alloys 69-714 (0.8% Ti, 1.6% Nb), 70-796
(0.79% Hf), 68-688 (no addition), 68-689 (0.36% Ti),
and 70-798 (0.71% Ti, 0.28% Hf, 0.94% Nb) have
properties intermediate between those of vacuum- and
air-melted Hastelloy N. All of the other alloys show
some improvement in rupture life, with alloys 69-345
(1.05% Ti, 0.88% Hf), 69-344 (0.77% Ti, 1.7% Nb), and
69-641 (1.3% Ti, 0.40% Hf) being superior.

The creep properties of these same alloys are shown
in Fig. 15.13. The trends that exist here are quite
similar to those shown in Fig. 15.12 for the rupture life.
STRESS (4000 psi}

Fig. 15.12. Stress-rupture properties at 650°C of several small commercial alloys (see Table 15.2 for composmons) Samples were

189

ORNL~-DWG 71-6985

70
4
60
..
[ R 4 L] LS
\"-
50 NJ| STANDARD, AIR MELTED
<9 [ ] [~ T O L ]
..."-u
40 ¢ R - \"h\-oeamc —c -
q L v <
[ ‘\..__
o} \ ’ L B K
\\ \
oy ey
30 o ?r
™ \"-.
o™ o< | 4 - ML
\ ™
o 470-785 ® 470-795 S h\
. A
20 | & 470-727 A 470-788 e =
o 470-796 » 470-797 . . N
o 69-648 ® 470-798 STANDARD, VACUUM MELTED \-.._\‘. ¥
v 69-714 v 468-688 ™
10 o 470-786 * 468-689 ]
& 69-641 + 469-344
tt 470-787 * 469-345
5 Lo L
ot 2 5 1o° 2 5 1 2 5 10 2 5 10° 2

RUPTURE TIME (hr)

irradiated at 760°C to a thermal fluence of 2 to 3 X 1029 neutrons/cm?.

70

STRESS (1000 psi)

Fig. 15.13. Creep properties at 650°C of several small commercial alloys (see Table 15.2 for compositions). Samples were
irradiated at 760°C to a thermal fluence of 2 to 3 X 1029 neutrons/cm?.

-ORNL-DWG 71-6986 -

*
60
‘ * < A L 2l
: /‘/‘
%0 By (Al Sl d L=
4
d L
o]
I | T 1L+
dol | voel a4 LN e
40 M S A el
o %] o4~ STANDARD AIR
L _ MELTED
30 oo el
b1
// o | v|¥c L7 ‘
Hl e /——‘/ ‘ o 470- 785 ® 470-79%5
: o = 70-T27 4 470-788
20 — A4
A1y /</ o 470- 796 . 470-797
L STANDARD VACUUM o 69-648 + 470798
L1 MELTED v 69-714 v 468- 688
10 © 470- 786 «468-689 |
+ 69— 641 + 469-344
# 470- 787 * 469- 345
0 _ 111 L1 )bl IR
1074 1073 1072 10! 10° 10'

MINIMUM CREEP RATE (%/hr)
190

ORNL-DWG 74- 6987

PO T T ]
0 470-785 @ 470-795
& 470-727 A 470-788 v
30 [T o a70-796 .= 470-797
o 69-648 & 470-798
v 69-714 v 468-688
o5 |1l © 470-786 & 468-689
¢ 69-641 + 469-344
# 470-787 % 469-345
I
20
z
<[
et Ly
w
a !
2 1
e a.
*
<>
10
- [ 'y
P .
v <
O
5 -
* 3 %
T ] A : L 4 o]
o ° o 4 o ¥
5 10° 2 5 102 2 5 ' 2 5 10° 2 5 o 2

MINIMUM CREEP RATE (%/hr)

Fig. 15.14. Fracture strains of several commercial alloys creep tested at 650°C (see Table 15.2 for compositions). Samples were
irradiated at 760°C to a thermal fluence of 2 to 3 X 1029 neutrons/cm?.

One notable exception is alloy 69-648 (0.92% Ti, 1.95%
Nb), which has a very high creep strength but did not
have an exceptionally long rupture life.

The fracture strains of these samples are shown in Fig.
15.14. Under these same conditions, standard vacuum-
melted Hastelloy N had fracture strains of 0.2 to 0.3%,
and air-melted Hastelloy N had fracture strains of 0.5 to
2.5%.% The scatter in these data is quite large and is
influenced at high stresses by our inability to get a true
measure of the strain on loading. The limited number of
points for each heat is also a factor that makes
conclusions quite questionable. However, there are
several heats with fracture strains below 2%: 70-786
(0.82% Ti, 0.62% Nb); 70-796 (0.79% Hf); 70-798
(0.71% Ti, 0.28% Hf, 0.94% Nb); 70-787 (0.90% Ti,
0.77% Hf, 0.12% Nb); and 70-788 (1.4% Ti, 0.30% Hf,
0.67 Nb). Few of the alloys seem to consistently
exhibit exceptional fracture strains, The trend of
decreasing fracture strain with decreasing creep rate is
also disturbing, since these lower rates would be used in
the design for steady-state operation.

4, HI. E. McCoy and R. E. Gehlbach, “Influence of Irradiation
Temperature on the Creep-Rupture Properties of Hastelloy N,”
J. Nucl. Appl. Technol,, accepted for publication.

The scaleup experience is best conveyed by consid-
ering some specific compositions for which we have
both laboratory and commercial melts. One laboratory
melt, 184 (1.2% Hf, 1.2% Ti, 0.14% Si), had very good
properties. Two double-vacuum-melted alloys (70-787,
69-641) and one ESR melt (69-345) of the same
nominal composition were procured from commercial

~ vendors. The comparative stress-rupture properties in

Fig. 15.15 show that the laboratory and commercial
melts had similar properties, all considerably above
those of standard Hastelloy N. The postirradiation
stress-rupture properties of these alloys are shown in
Fig. 15.16. Heat 184, the laboratory melt, had good
rupture lives and excellent fracture strains. Heat
69-345, the ESR melt, had good rupture lives, but the
fracture strain seemed to decrease with increasing
rupture life. Heat 69-641 had good rupture lives and
good fracture strains. Heat 70-787 had stress-rupture
properties equivalent to those of standard air-melted
Hastelloy N. However, the fracture strain decreased
drastically with increasing rupture life.

Heats 181, 69-648, and 69-344 had the nominal
composition of 0.5% Ti and 2% Nb and were made by
laboratory melting, double vacuum commercial melting,
and ESR respectively. The stress-rupture properties in
191

ORNL-DWG 71- 6988

70 & v
.\\\ \\\
\\ \\
\ N
60 N N
N N
Oy
AN N
I~

50 \\ ™

- STANDARD \\ ‘N\

a

8 40 oo

o N

0 © 184,1.2 Hf, 1.2 Ti "N

& 30 | & 70-787,077Hf, 0.90Ti ™,

& a 69-641, 0.4 Hf, 1.3 Ti \

” o 69-345,0.88 Hf, 1.1 Ti, 0.5 Si N
20 \\\
10
0

10° 10! 102 10° 107

RUPTURE TIME (hr}

'Fig. 15.15. StIess -rupture propertles at 650°C of several
modified alloys having the nominal composition of Ni—12%
Mo-7% Cr1—0.2% Mn-1% Hf-1% Ti—0.05% C. All material
annealed 1 hr-at 1177°C.

ORNL—DWG 70~ 12643R

70 ‘
| ™
\
15
o]
50 12
46 || 98 N
N
\\
h 50 a6 [~ 28 L STANDARD, UNIRRADIATED [ ]
B 1 \ [
a
STANDARD, L :
9 49 IRRADIATED ™[I ©lz N
8 . ) -.\\ B - ¢
= g ' B i \\\ 203
7 TS AIR MELTED - ™ I
w 30 ! n\.\ N
: IS
= [
w oy
o
o o184, 12 HF. 12 Ti \..,\\{_,VA(ISUUM MELTED A
' A& 70-787, 0.77 Hf, 090T| ‘
o ™~
69-641, 0.4 Hf, 1.3 T|' ‘ S At e504c
10 o 69-345, 0.88 Hf, {4 Ti, 0.55i
Y (] ’ ‘ i ‘ 2 ’ 3
10 10", 10 10

RUPTURE TIME {hr)

Fig. 15.16. Comparison of stress-rupture properties at 650°C
of modified Hastelloy N irradiated at 760°C to a thermal
fluence of 3 X 1022 neutrons/cm?2.

the unirradiated condition were widely variant, with the
rupture life at a given stress being highest for heat
69-648 and lowest for heat 181; however, all heats were
superior to standard Hastelloy N (Fig. 15.17). The
postirradiation- properties -are shown in Fig. 15.18. If
one ignores the two points for heat 69-648 at 40,000

ORNL-DWG 71-6989

70 O &
\\ \ '\\ \\
° ™~ \ \Q\
60 - ] 5
. . ™
\\ \)\ \\\\1\\
50 N \Q; \
\ ~ ~
= \r\ g \ d \\_
a \\‘ \\
@) N
8 40 TN
; STANDARD HASTELLOY N
0 ™
& ™
30 ™,
» : AN
o 184, 0.5T:, 2.0 Nb
. s 69-648,0.9Ti, 2.0Nb \
20 | o 69-344,0.8Ti, 1.7 Nb, 0.5 Si \\
10
0>
10° 10' 102 10° o o

RUPTURE TIME (hr)
Fig. 15.17. Stress-rupture properties at 650°C of several

modified alloys having the nominal composition of Ni—12%
Mo-7% Cr—0.2% Mn-0.5% Ti—2% Nb—0.05% C

ORNL-DWG 70-12612R

70
) 1\\
60. .
26 \\;TANDARD, UNIRRADIATED
|
50 : 4 | S N 2.
- STANDARD, 9 a4 Y
2 IRRADIATED-4| - vl
g 40 . | N 0.5 o
2. M~ ’ 01.4 \T\H‘L * N
é' | TNNaevACUUM MELTED [[™~ fi\‘ 0.2 | |
30
Ll ~ Tony
: ~ eSS
7 ~ AIR MELTED NN
20—+ . T =
" o181, 0.5Ti, 2.0 Nb! S
1o |- & 69648, 0.9Ti, 2.0 Nb
o 69-344, 0.8 Ti, .7 Nb, 0.5 Si
o L LU L L. 4
10° 102 10° 10

RUPTURE TIME (h_r)

Fig. 15.18. Comparison of stress-rupture propertles at 650°C
of modified Hastelloy N irradiated at 760°C to a thermal
fluence of 3 X 1029 neutrons/cm?.

psi, all three heats fall within a rather narrow band and
have fracture strains of 0.2 and 4%.

Thus our experience with laboratory meits is forming
a rather consistent picture, but the small commercial
melts have inconsistent behaviors.
15.6 STATUS OF DEVELOPMENT OF A
TITANIUM-MODIFIED HASTELLOY N

H.E.McCoy C. E. Sessions

Our early studies showed that alloys containing as

little as 0.5% Ti had excellent postirradiation properties
when irradiated at 650°C, but deteriorated rapidly as
the irradiation temperature was increased.® The good
properties were associated with the presence of a finely
divided MC-type carbide and the poor properties with
the formation of a coarse M, C-type carbide. Titanium
has several advantages as an alloying element compared
with other elements, such as Hf and Zr; titanium (1) is
relatively cheap, (2) is a common alloying addition to
nickel-base alloys, (3) seems to cause no welding or
fabrication problem in concentrations up to at least
2.1%, and (4) diffuses very slowly in Hastelloy N and
should not contribute significantly to the corrosion
rate. Its single known disadvantage is that it promotes
the formation of a hard, brittle intermetallic com-
pound, Ni3Ti. The exact level of Ti required to form
this compound is unknown, since most commercial
alloys contain both Ti and Al that will form this
compound. However, the numerous advantages of Ti
have led us to study alloys with Ti concentrations in the
range of 1 to 3% with a watchful eye toward the
formation of the brittle intermetallic.
. The chemical compositions of the various alloys that
we have studied with Ti concentrations in the range of
1 to 3% are shown in Table 15.5. The alloys with two-
and three-digit numbers are 2-lb laboratory melts, and
the last three melts are 50- to .100-lb heats from
commercial vendors. All of the alloys were annealed 1
hr at 1177°C after final fabrication and irradiated for
about 1100 hr at 760°C to a thermal fluence of 2 to 3
X 10?° neutrons/cm®. They were then creep tested at
650°C.

The results of the postirradiation tests are shown in
Fig. 15.19. Alloys 107 (1.04 Ti) and 75 (0.99 Ti) had
properties intermediate between those of air- "and
vacuum-melted standard Hastelloy N. Alloy 76 (1.04%
Ti) had a higher .carbon level (0.117%) and had
properties superior to those of standard Hastelloy N,
Although the data scatter somewhat, there is a definite
trend of increasing rupture life and fracture strain with
increasing Ti concentration. However, alloysi292 (2.4%
Ti) and 293 (2.9% Ti) have improved rupture lives, but
the fracture strains are low. We have likely exceeded the

5. H. E. McCoy, Influence of Titanium, Zirconium, and
Hafnium Additions on the Resistance of Modified Hastelloy N
to Irradiation Damage at High Temperature—Phase 1, ORNL-
TM-3064 (January 1971).

192

Table 15.5. Compositions (%) of Ti-modified alloys

Alloy Ct Fe Mn Si Ti C
designation

107 120 7.0 ~ 0.08 0,20 <0005 1.04 0.056
75 11.8 79 <0.05 0,20 <005 0.99 0.062
76 11.8 7.9 <0.05 0.20 <0.05 1.04 0.117
290 002 1.2

327 1.63 0.044
328 1.87 0.80
291 11.6 776 4,22 0.21 0.02 1.97 0.055
292 113 7.26 424 022 0.02 24 0.060
293 11.0 796 4.11 0.19 0.02 29 0.052
67-548 12,0 7.1 0.04 0.12 0.03 1.1 007
70-785 12,3 7.0 0.16 030 0.09 1.1 0.057
70-727 130 74 0.05 037 <0.05 2.1 0.044

Ti level required to form Ni; Ti, although we have not
specifically identified the compound. Commercial alloy
67-548 (1.1% Ti, 0.07% C) has two points at 40,000 psi
with low fracture strains; otherwise, the properties are
about equivalent to those of standard air-melted Hastel-
loy N. Alloy 70-785 (1.1% Ti, 0.05% C) has properties
about equivalent to those of standard vacuum-melted
Hastelloy N. The data.are inconclusive on heat 70-727.
The properties of this alloy should be equivalent to
those of heat 291. One of the data points on heat
70-727 at 40,000 psi indicates that this is the case;
however, the other point falls far short.

Thus the results from our laboratory melts indicate
that an alloy containing about 2% Ti would have
acceptable properties. The data currently available from
commercial alloys indicate that the laboratory and
commercial alloys have similar properties when they
contain about 1% Ti and that the properties may be
equivalent at the 2% level. We plan to pursue further
the development of an alloy containing 2% Ti.

15.7 CORROSION STUDIES
J. W. Koger

The success of a molten-salt reactor system is strongly
dependent on the compatibility of the materials of
construction with the various fluids in the reactor. The
experiments discussed in this section are being con-
ducted to determine the behavior of reactor materials in
a molten fluoride salt environment. Because heat is
transferred to and from the salt, the most prevalent
form of corrosion is temperature gradient mass transfer. )
This can effect the removal of selected alloy constit-
uents, which may compromise certain favorable prop-
erties of the alloy, and the deposition of dissolved
193

ORNL-DWG 71-3738

100 ;
RN,
Ti (%) C(%) Ti (%) C (%)
90 | o 107 1.04 0056 ___ = 291 2.0 0.055
S a 75 0.99 . 0.062 + 292 2.4 0.060
o 76 .04 0.117 . 293 2.9 0.052
80 — ¢ 290 1.2 0.050 | < 67-548 - 1.1 0.070
. 327 1.6 0.044 A 70-785 1.1 0.057
s 328 1.9 0.080 X 70-727 2.1 0.044
14 1.4
[ 1§ >
= 60
S WONIY 1.4 1.4
Fanai i ]
o} aly
i >0 1.28 O 6& 9.8 . 19156
# Il ™2 T I 68
T 0.5 M~ 44 |49|1|0.9
@ % ok hH\._ IR . T lezo
AT s | 33w | 5
1.6 \\ 1.7 | \ 5.3
30 < 0.9 o N
A | [1e |75 | 7.8
| 6.3 I~_ 5.0 | T2
20 e h ~0
\..___\\
10 ~~\‘\.
Y ‘
107" 10° 10 10 10° 10?

RUPTURE TIME (hr)

Fig. 15.19. Postirradiation stress-rupture properties at 650°C of several modified alloys containing titanium. Each specimen was
annealed 1 hr at 1177°C and irradiated for 1100 hr at 760°C to a thermal fluence of 2 to 3 X 1029 neutrons/cm?.

corrosion products can also restrict flow in the cold
part of the system. Important variables which may
affect the mass transfer process are (1) impurities in the
salts, (2) the nature and amounts of alloy constituents,
_ (3) salt velocity, and (4) temperature. These effects are
being studied in both isothermal and polythermal
systems with conditions based on molten-salt breeder
reactor design parameters. '

Past work® has shown that of the major constituents
of Hastelloy N, chromium is much more readily
oxidized by fluoride salts than Fe, Ni, or Mo. Thus
attack is normally manifested by the selective removal
of chromium. The rate-limiting step in chromium
removal from the Hastelloy N by fluoride salt corrosion
is the solid-state diffusion of chromium in the alloy.
Several oxidizing reactions may occur, depending on
the salt composition and impurity content, but among
the most important reactants are UF,, FeF,, and HF.

6. W. D. Manly et al., Progr. Nucl. Energy, Ser. IV 2, 164—79
(1960).

In fluoroborate salt systems, which in some cases have
been contaminated with H, O, we have found that
elements other than chromium may be oxidized by the
salt, and as a result the corrosion rate is higher and the
attack is more uniform.

Experiments have continued on alloy compatibility
with molten salts of interest, and extensive work is
under way to determine the kinetic and thermodynamic
properties of the corrosion systems. The alloys of
interest (both nickel and iron based) are exposed to
flowing molten salts in both thermal-convection loops
and in loops containing pumps for forced circulation of
the salts. The typical thermal convection loop (Fig.
15.20) is operated by heating, with clamshell heaters,
the bottom and an adjacent side of the loop after which
convection forces in the contained fluid establish flow
rates of up to 8 fpm, depending on the temperature
difference between the heated and unheated portion of
the loop. The salts of interest are LiF and BeF, based
with UF, (fuel), ThF,; (blanket), and UF, and ThF,
(fertile-fissile) and an NaBF,-NaF mixture (coolant
salt).
194

ORNL-OWG 68-3987 The status of the thermal-convection loops in opera-
tion is summarized in Table 15.6.

STANDPIPE

15.7.1 Fuel Salts

Loop 1255, containing an MSRE-type salt with 1
mole % ThF,; and constructed of Hastelloy N, has
operated for 8.8 years. The loop contains insert
specimens (only removable by dismantling the loop) of
standard Hastelloy N and a Hastelloy N + 2% Nb alloy.
The latter Nb alloy was developed for improved
weldability and mechanical properties, although similar
compositions are now under consideration because of
increased resistance to neutron damage.

Loop 1258, constructed of type 304L stainless steel,
has operated about 7.6 years with the same salt as loop
1255. The hot leg of the loop contains tab specimens
that can be removed without interrupting loop opera-
tion. This loop provides the capability for testing the
compatibility of iron-base alloys with this particular
| salt. We recently replaced the heaters and thermo- -
| couples of both loops. Operation of these loops had
| become increasingly difficult as heater and thermo-
I 0f ‘ couple failures became more frequent. When the heaters
I
|
!

BALL VALVES

, were removed, we found many green chromium oxide

= :> CLAMSHELL crystals on the loop exterior directly under the heaters.

HEATERS These crystals were the result of a fairly common

phenomenon in Nichrome-wound heaters where chro-

mium, because of its high vapor pressure, is selectively

vaporized from the heater wire over a long period of

time. This vaporization changes the heating character-

istics of the wire and certainly contributed to some of

our operating problems. Both loops are agam operating
at design temperatures.

Loop NCL-16, constructed of Hastelloy N and con-
SAMPLER taining removable specimens in each vertical leg, has
operated with MSBR fuel salt (Table 15.6) for over
three years. The largest weight loss of a Hastelloy N
specimen is —2.1 mg/cm?® after about 24,000 hr. The
largest weight gain over the same period is +1.6
mg/cm?. Assuming uniform loss, the maximum weight
loss is equivalent to 0.03 mil/year. The chromium
content of the salt has increased 335 ppm in 24,000 hr.
We continue to show in this system, and several others,
that titanium-modified Hastelloy N specimens (12%
Mo, 7% Cr, 0.5% Ti, bal Ni) show smaller weight losses
FLUSH DUMP than standard Hastelloy N specimens (16% Mo, 7% Cr,
' TANK 5% Fe, bal Ni) under equivalent conditions. We attrib-
ute this difference to the absence of iron in the
modified alloy.

o [ Illf

- :/INSULATION
J
|
|
|
[
|
\

CORROSION
SPECIMENS

Fig. 15.20, MSRP natural circulation loop and salt sampler.

-
Table 15.6. MSR Program natural circulation loop operation through February 28, 1971

Salt Maximum Operating
Loop . . ... AT .
No Loop material Specimens Salt type composition temp:rature ) : (o 0) time
. (mole %) O (hr)
1255 Hastelloy N Hastelloy N + 2% Nb® b Fuel LiF-BeF,-ZrF,-ThF,4 704 90 77,040
(70-23-5-1-1)
1258 Type 304L SS Type 304 L stainless Fuel LiF-BeF,-ZrF4-UF4-ThF4 688 100 66,170
steel?¢ (70-23-5-1-1) , :
NCL-13A Hastelloy N’ Hastelloy N; Ti-modified Coolant NaBF,4-NaF (92-8) plus 607 125 20,380
Hastelloy N controls©4 tritium additions
NCL-14 Hastelloy N Ti-modified Hastelloy N&d Coolant NaBF,4-NaF {92-8) 607 150 28,955
NCL-15A Hastelloy N Ti-modified Hastelloy N; Blanket LiF-BeF,-ThF, (73-2-25) 677 55 22,215
Hastelloy N controls®9
NCL-16 Hastelloy N Ti-modified Hastelloy N; Fuel LiF-BeF,-UF4(65.5-34.0-0.5) 704 170 26,575
Hastelloy N controls®:4
NCL-17 Hastelloy N Hastelloy N; Ti-modified Coolant NaBF,4-NaF (92-8) plus 607 100 . 14,615
Hastelloy N controlsc'd steam additions :
NCL-19A Hastelloy N Hastelloy N; Ti-modified Fertile-fissile LiF-BeF,-ThF4-UF,4 (68- 704 170 9.00s
: Hastelloy N controls©4 20-11.7-0.3) plus
bismuth in molybdenum
hot finger _
NCL-20 Hastelloy N Hastelloy N; Ti-modified Coolant NaBF,4-NaF (92-8) 687 250 10,515

Hastelloy N controlsc.d

4Permanent specimens,

bHot leg only.

“Removable specimens.
9Hot and cold legs.

S61
15.7.2 Fertile-Fissile Salt

A fertile-fissile MSBR salt has circulated for over
8800 hr in Hastelloy N loop NCL-19A, which has
removable specimens in each leg. The test has two
purposes: to confirm the compatibility of Hastelloy N
with the salt and to determine if bismuth will be picked
up by the salt and carried through the loop. The
bismuth is contained in a molybdenum vessel located in
an appendage beneath the hot leg of the loop. Hastelloy
N specimens in the loop have shown maximum weight
losses and gains of —0.50 and +0.66 mg/cm?, respec-
tively, in 8800 hr. Assuming uniform loss, the maxi-
mum weight loss is equivalent to 0.024 mil/year. Figure
15.21 shows the weight changes of specimens in the
loops as a function of time and position (temperature)
and illustrates the mass transfer profile discussed
previously. A modified Hastelloy N alloy (12% Mo, 7%
Cr, 0.1% Fe, 0.4% Ti, 2.0% Nb) has lost less weight than
a standard alloy at the same position. The chromium
content of the salt has increased 134 ppm in 7400 hr,
and there is no detectable bismuth in the salt.

15.7.3 Blanket Salt

Loop NCL-15A, constructed of standard Hastelloy N
and containing removable specimens in each leg, has
operated 2.5 years with a blanket salt (LiF-BeF, with
25 mole % ThF,;) proposed for a two-fluid MSBR
concept. Mass transfer, as measured by the change of
chromium concentration in the salt, has been very
small. Specimens exposed to this salt tend to be
“glazed” with a coating (probably a high-melting Th

ORNL-DWG 71-6930

o

e |
~ 0
- R NI S S
W ¢ _~MODIFIED HASTELLOY N
Z 0

L e T — — — a4 685°C
S .“"“"3-.'—.7___—-.—£ :—-_:7050(:‘
= - * T —*705°C
I

1]

W —1

z

-2

0 2000 4000 6000 8000 10,000 12,000

OPERATING TIME (hr)

Fig. 15.21. Weight change vs time of Hastelloy N specimens
in NCL-19A exposed to LiF-BeF,-ThF,;-UF, (68-20-11.7-0.3
mole %) at various times. Liquid bismuth is in contact with the
salt at the bottom of the hot leg.

196

compound) that is impossible to remove without
damaging the metal. However, metallographic studies
also show little, if any, mass transfer.

15.7.4 Coolant Salt

Loops NCL-13A and NCL-14, constructed of stand-
ard Hastelloy N and containing removable specimens in
each ‘leg, have operated for two and three years,
respectively, with the fluoroborate mixture NaBF,-NaF
(92-8 mole %). The maximum corrosion rate at the
highest temperature, 605°C, has been 0.7 mil/year for
both loops. Corrosion has generally been selective
toward chromium, but there have been short periods
when impurities have caused general attack of the
Hastelloy N. The latter periods were caused by leaks in
the seals of ball valves which are located above the surge
tanks. Without such leaks the mass transfer rate during
a given time period has been low. The areas in which
the leaks are found are exposed to a mixture of He and
BF3 gas and not to salt. We attribute the leaks both to
heavy use of the ball valves and corrosiveness of the gas
mixture after air inleakage. Although the overall corro-
sion rates in these loops were not excessive, the rates
observed in the absence of ball valve leaks have been an
order of magnitude lower than the average rate.

Loop NCL-17, constructed of standard Hastelloy N
and containing removable specimens in each leg, is
continuing to operate after an addition of steam to the
coolant salt.” This experiment was undertaken to
determine the effect of steam on the mass transfer
characteristics of the fluoroborate salt mixture. During
one part of this report period the specimens were found
to have undergone large weight changes. Careful exami-
nation disclosed a leak in a cover-gas line leading to the
loop expansion tank. Along with the line repairs, ball
valve seals were also replaced. In an attempt to remove
adsorbed impurities after the repair, we evacuated the
loop while heating the salt to just below the melting
point. At present, some heat is required on the cold leg
to maintain salt flow and the desired AT. Radiography
has disclosed precipitated material in the bottom of the
cold leg.

Immediately prior to the leak in the cover-gas system,
we noted that the specimen in the hottest position
(607°C) had corroded at an overall rate of 2.5 mils/year
since steam addition (9000 hr) and at a rate of <1.0
mil/year over the last 3000 hr. The overall mass transfer

~ rate had steadily decreased since the steam addition.

7. 1. W. Koger, MSR Program Semiannu. Progr. Rep. Aug. 31,
1970, ORNL-4622, p. 170.
ORNL-DWG 71-6991

10

&

E 5
H

w - e[ 460°C

- e

R B

g 0] .-.._3_::-—___.__-__---

T —— o

fi \.t‘. 660°C )
g ~¢685°C

o -5

=

-10 ~
o 2000 4000 6000 BOOC = 10,000 12,000

OPERATING TIME {hr)

Fig. 15.22, Weight change vs time of Hastelloy N specimens
in NCL-20 exposed to NaBF4-NaF (92-8 mole %) at various
temperatures.

Loop NCL-20, constnicted of standard Hastelloy N
and containing removable specimens in each leg, has
operated for over 10,000 hr with the fluoroborate
coolant salt at about the most extreme temperature
conditions considered (687°C max and 438°C min) for
the heat exchanger of the MSBR secondary circuit.
Forced air cooling of the cold leg is required to obtain
this AT. After 8500 hr the maximum weight loss in the
hot leg was measured to be —4.3 mg/cm? (0.2 mil/year
assuming uniform loss) and was accompanied by a
maximum gain in the cold leg of +2.1 mg/cm?®. Figure
15.22 shows the weight changes of specimens in the
loop as a function of time and temperature. The
chromium content of the salt has increased 77 ppm,
and the oxygen content of the salt has increased about
300 ppm. The weight changes of the specimens and the
chemistry changes of the salt indicate that the mass
transfer is the lowest yet attained with the fluoroborate
mixture in a temperature gradient system and also
confirm the importance of salt purity on compatibility
with structural material and on operation.

197

Figures 15.23 and 15.24 are optical and scanning

electron micrographs of specimens from the hot and
cold legs of NCL-13.® These specimens were exposed to
NaBF,-NaF (92-8 mole %) for 4180 hr at 580 and
476°C respectively. The hot-leg specimen lost 3.3
mg/cm?, and the cold leg specimen gained 3.2 mg/cm?.
The scanning electron micrograph of the surface of the
hot-leg specimen, as seen in Fig. 15.235, shows that the

8. J. W. Koger and A. P, Litman, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1968, ORNL-4344, pp. 264—65.

Hastelloy N has been attacked nonuniformly and is
heavily contoured. Figure 15.24 shows that the cold-leg
specimen is coated with a discontinuous deposit which
is essentially analyzed as Hastelloy N with a little more
nickel and molybdenum.

Figures 15.25. and 15.26 are optical and scanning
electron micrographs of specimens from the hot and
cold legs of NCL-17. The specimens, at temperatures of
593 and 500°C, respectively, were exposed to NaBF,-
NaF (92-8 mole %) for 1050 hr, at which time steam
was injected into the loop. The specimens then re-
mained in the loop for an additional 9000 hr. The total
weight loss of the hot-leg specimen was 40.9 mg/cm?,
and the weight gain of the cold-leg specimen was 16.0
mg/cm?. In Fig. 15.25a we see that the surface has
receded uniformly, although a Widmanstatten precipi-
tate was left in relief. Figure 15.26a shows a dark
deposit on the cold-leg specimen, and the deposit
exhibits a layered structure when viewed from above
(Fig. 15.26b).

15.7.5 Analysis of H, O Impurities in
Fluoroborate Salts

Over the last few years, we have been concerned over
the behavior of H,O in molten fluoride salts. The
disposition of H, O in these salts has implications both
for corrosion studies and for tritium removal schemes.
In corrosion work, we have long recognized the need
for analyses which would allow us to distinguish
between water and (1) compounds which contain H*
ions, (2) compounds which contain O* ions, and (3)
HF (highly corrosive). In the past, a complementary
indication of mass transfer (besides weight changes and
metal analysis of the salt) was the analysis for H, O and
O, in the salt. On the basis of our tritium injection
experiment (subsequent section on retention of tritium
by sodium fluoroborate) and work in Analytical Chem-
istry and Reactor Chemistry Divisions, we now feel that
the results of the H,O analyses included many other
constituents of the salts, including oxides. Thus it was
possible that a result indicating a large amount of H, O
really only showed that amount of oxide. The lack of
H,O in the salt is completely reasonable on the basis
that all the hydrogen in impurity compounds that enter
the salt would eventually cause oxidation of the metal
wall and then would itself be reduced to H, gas which
would diffuse out of the loop. However, increased mass

-transfer (weight losses and weight gains of specimens)

by virtue of H, O impurities is still accompanied by an
increase in the oxide content of the salt. Thus the oxide
content is still somewhat indicative of the corrosion of
the loop system.
198

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[ > 5 Sl : - It

4 -' T - . ? - . : i
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(a) 4 . . -

- . * ¥ L4 ' .

Fig. 15.23. Hastelloy N specimen from NCL-13 exposed to NaBF4-NaF (92-8 mole %) at 580°C for 4180 hr. Weight loss, 3.3
mg/cm?. (a) Optical micrograph, S00X. Etchant, glyceregia. (b) Scanning electron micrograph, 1000X.

199

Y-105980
- - £ €
. . - o
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s ' g
kg S
. .. . IS
: v 0 L
g ., P
v F 5
R Sl . g - 18
(a) . .

Fig. 15.24. Hastelloy N specimen from NCL-13 exposed to NaBF3-NaF (92-8 mole %) at 476°C for 4180 hr. Weight gain 3.2
mg/em?. (a) Optical micrograph, S00X. Etchant, glyceregia. () Scanning electron micrograph, 1000X.

200

Y-105985

5.

(a) = \‘Y: =

Fig. 15.25. Hastelloy N specimen from NCL-17 exposed to NaBF4-NaF (92-8 mole %) at 593°C for 10054 hr (1054 hr before
steam addition). Weight loss, 40.9 mg/cmZ. (a) Optical micrograph, S00X. (5) Scanning electron micrograph, 1000X.

201

Y-105982

toooT .

T

16003 i

0.007 INCHES.
500X

16005

looo7 ™

Y-103330

]

Fig. 15.26. Hastelloy N specimen from NCL-17 exposed to NaBF;-NaF (92-8 mole %) at 500°C for 10054 hr (1054 hr before
steam addition). Weight gain, +16.0 mg/cmZ. () Optical micrograph, S00X. (b) Scanning electron micrograph, 1000X.
15.8 FORCED-CONVECTION LOOP
CORROSION STUDIES

J.W.Koger H.C.Savage W.R. Huntley

The MSR-FCL-1 forced-circulation loop is being
operated to evaluate the compatibility of standard
Hastelloy N with NaBF,-NaF (92-8 mole %) coolant
salt at temperatures and flow rates similar to those
which existed in the MSRE coolant circuit.? Salt velocity
in the '%-in.-OD, 0.042-in.-wall Hastelloy N loop is
nominally 10 fps. The maximum and minimum salt
temperatures in the loop are 588 and 510°C respec-
tively. Hastelloy N corrosion test specimens are exposed
to the circulating salt at three temperatures: 510, 555,
and 588°C. MSR-FCL-1 has accumulated & total oper-
ating time of over 11,000 hr.

A second forced-circulation loop, of improved design,
MSR-FCL-2, is in the final stages of assembly. This test
will study the corrosion and mass transfer of standard
Hastelloy N in sodium fluoroborate coolant at typical
MSBR salt temperatures of 620°C maximum and 454°C
minimum and at velocities up to 20 fps.

15.8.1 Operation of Forced-Convection
Loop MSR-FCL-1

~ Operational problems and equipment failure limited
the loop operation-during the report period to 1342 hr,
with a total accumulated time at design conditions of
10,335 hr as of February 28, 1971. o

The sixth run was abruptly terminated on October
19, 1970, after 1241 hr duration, when oil from a
broken pump cooling oil line ignited when it came in
contact with the pump bowl at 510°C and adjacent
piping. The loop was automatically transferred to
standby conditions by the control system, and an alarm
from the loop was received by the plant shift superin-
tendent. In standby condition the salt circulation is
stopped, the main loop heaters are turned off, and
sufficient heat is supplied to the loop to maintain the
temperatures around the loop circuit above the salt
liquidus temperature (385°C). We extinguished the fire
and turned off all loop heaters, which allowed the salt
to freeze in the loop piping.

The pump cooling oil line failure occurred in a %; -in.
sched 40 brass pipe at a point near the rotary oil union
mounted on top of the salt pump (model LFB). This
pipe is believed to have been in service for the entire

9. W. R. Huntley et al., MSR Program Semiannu. Progr. Rep.
Aug. 31, 1970, ORNL-4622, p. 174.

202

operating life of the loop (11,371 hr at the time of
failure). During this time it had been subjected to a
low-amplitude high-frequency vibration from the rotary
oil union, and the failure appears to be a fatigue type.

Damage to the loop as a result of the fire was
confined to electrical wiring, thermocouples, service
piping, and gages in the vicinity of the pump. The
helium cover-gas overpressure was lost as a result of the
fire, and air inleakage to the loop and drain tank
probably occurred.

We replaced the electrical wiring, thermocouples, and
service piping necessary to melt the salt in the loop
piping and transfer it to the drain tank. This was the
first time that we had attempted to melt salt in the
entire loop circuit, although salt had been melted in the
cooling coil without difficulty on several occasions. In
attempting to melt the NaBF,-NaF (92-8 mole %) salt
in the loop, a rupture occurred in the main loop piping
(*5-in.-OD by 0.042-in.-wall Hastelloy N) at a point
near the U-bend in the loop and adjacent to one of the
metallurgical specimen holders, as shown in Fig. 15.27.
After the rupture occurred, the salt in the loop was
frozen as quickly as possible by turning off all loop
heaters. However, approximately 77 g of salt was lost
from the loop. Prior to the rupture, the salt in the bowl
and cooling section of the main loop had been melted
without difficulty, and the temperature of the main
loop piping had reached temperatures varying between
371 and 482°C around the loop circuit, and the drain
line temperatures were above 426°C.

In reviewing the loop melting procedure to determine
the cause of the rupture, we concluded that uneven
thicknesses of thermal insulation caused considerable
variation in the temperatures. In particular, the tem-
peratures under the clamshell emergency heaters (where
the thermal insulation was ~1 in. thick as compared
with ~2 in. thick on the remainder of the tubing) were
~80°C below temperatures in the more heavily in-
sulated sections of tubing. Thus the salt in the ruptured
area melted but was not able to expand properly
because of frozen salt plugs under the emergency .
heaters. Another problem was the difficulty in control-
ling the rate of heatup of the main loop by resistance
heating. Even though the lowest setting on the loop
heater was used, we still had to turn off the heater
supply power for short periods to maintain temperature
control.

We removed the metallurgical specimen assembly and
ruptured section of tube and replaced it with new
tubing. Figure 15.28 is a photograph of the ruptured
section of tubing. Measurements of the outside diam-
eter of the remaining loop tubing were made and
203

ORNL-LR-DWG 64740R2

POINT OF RUPTURE DURING
SALT MELTING AFTER RUN 6,

’ METALLURGY )
SAMPL
110 kVA POWER SUPPLY (OETAIL A |

> (MAIN POWER)

1600-amp BREAKER

I
K

POINT OF RUPTURE
WHEN LOOP WAS
OVERHEATED DURING
RUN 7~

,4/,) 7
,%swssncv HEATERS

METALLURGY
SAMPLE

HEATER
LuG™A"

\ .
DRIVE MOTOR DETAIL 8

coiL <
; (DETAIL B) J\\\ >
>

-”\_/T "2““‘“‘

N 10KkVA POWER SUPPLY

FREEZE VALVE —
EEZE VALVE (COOLER PREHEAT)

DUMP TANK

Fig. 15,27, Molten-salt corrosion testing loop and power supplies, loop MSR-FCL-1.

Y-103986

Fig. 15.28. Photograph of tube rupture during salt melting
after run 6, loop MSR-FCL-1.

o 0.5 1.0
¥ | 1 1 1 |

INCHES
compared with the original outside diameter to deter-
mine if excessive permanent strain had occurred at
other locations. Generally all measurements were in the
range of 0.500 to 0.505 in., as compared with the
original nominal diameter of 0.500 in.

The emergency heaters (which are no longer in use)
were removed, and the loop was insulated uniformly.
The resistance heater power supply was modified to
reduce the loop heat input, and the salt in the loop was
again melted without apparent incident and drained to
the sump tank. The remaining corrosion test specimen
assemblies were removed for weighing and metallurgical
examination.

Corrosion test specimen assemblies were reinstalled in
the loop, the pump was repaired by installing new
bearings and seals (these are routinely replaced every
2000 hr), and a catch basin was installed for the pump
lube oil to prevent oil contacting the high-temperature
loop in the event of a lube oil line rupture or leak. We
returned the loop to design conditions for the start of
run 7 on January 8, circulating the same NaBF,-NaF
(92-8 mole %) coolant salt used since the start of loop
operation.

"On January 12, 1971, while repairs were under way
on the loop high-temperature protective instrumenta-
tion, an operating error resulted in a portion of the loop
tubing being heated to temperatures above 1100°C.
This caused one failure of the loop tubing at a point
adjacent to one of the electrical power input lugs (see
Fig. 15.27) and another about 12 in. downstream. Part
of the salt in the loop (~2 liters) was discharged into
the loop secondary containment. Temperatures were
immediately reduced, and the facility was shut down.
The tubing obviously melted at the rupture. However,
at 1100°C the strength of Hastelloy N is very low, and
the salt vapor pressure would approach 1000 psi. Local
boiling of the salt would also occur, and overheating
could reduce the tube strength until rupture occurred.
At the time of failure the loop had accumulated 10,335
hr of operation at design conditions (101 hr during run
7). A photograph of the tube failure is shown in Fig.
15.29.

The overheating of the loop was caused by the

following sequence of events:

1. While attempting to repair a thermocouple con-
nected to the high-temperature cutoff instrument, the
thermocouple circuit was first opened and subsequently
shorted. A fail-safe feature in the high-temperature
cutoff responded to the open circuit and caused the
loop to be transferred from design conditions to
standby conditions. The subsequent shorting of the
thermocouple rendered the high-temperature cutoff

204

inoperative. In standby condition the salt pump is
stopped and loop heater power is reduced such that salt
in the loop is maintained in a liquid state (~400 to
540°C). . .

2. The loop temperature controller had been tempo-
rarily placed on ‘““manual” to prevent power input
variations in the event the control thermocouple
(located near the high-temperature cutoff thermo-
couple) was disturbed while repairs were attempted.

3. As quickly as possible after the loop was trans-
ferred to the standby operating condition, the loop
operator proceeded to activate circuits necessary to
return the loop to design operating conditions. He did
not reduce the main loop heater power control before
activating these circuits, with the result that about 70
kW of heating was applied to the loop piping (2 or 3 kW
is sufficient to maintain loop temperatures in the
standby condition without the air-cooled radiator in
service). With both the loop temperature controller and
the high-temperature cutoff instrument inoperative,
loop temperatures reached the failure level before the
main heater power had been manually reduced (in ~2',
min). .

We removed the corrosion test specimens for weighing
and metallurgical examination. All loop tubing external
to the air-cooled radiator (the radiator was not over-
heated or damaged) was also cut out for metallurgical
examination of selected sections.

Before loop operation can be continued, it will be
necessary to replace the main loop tubing (~25 ft of
Y,-in.-OD by 0.042-in.-wall Hastelloy N). In addition,
we plan to modify the instrumentation to prevent
accidental startup of the main loop heater at high
power levels. '

15.8.2 Metallurgical Analysis of MSR-FCL-1

After the oil fire which ended run 6 and the
subsequent rupture of the tubing while attempting to
melt the salt from the loop, the corrosion specimens
were removed, weighed, and examined. Surprisingly,
weight losses were found for all specimens. The weight
loss rates found for specimens exposed to the salt at
588 and 555°C were (assuming uniform losses) ~1.9
and 0.5 mil/year, respectively, for the 1240 hr of run 6.
The corrosion rate ratio for these specimens (several
times greater loss at 588°C than at 555°C)'® was fairly
consistent with past findings. However, in earlier runs,
specimens at 510°C had gained weight. During run 6,

10. W. R. Huntley, J. W. Koger, and H. C. Savage, MSR
Program Semiannu. Progr. Rep, Aug. 31, 1970, ORNL-4622, p.
175.
205

Fig. 15.29. Photograph of tube rupture caused by overheating during run 7, loop MSR-FCL-1.

specimens at this temperature level lost material at the
rate of 1.9 mils/year, a rate that was the same as the
rate for specimens at 588°C. This weight loss anomaly
and the overall higher weight losses are attributed to the
problems encountered during this run and probably
attest to a local high-temperature condition in the
cooler section along with oxidizing impurities in the
salt. A salt analysis showed small increases of chro-
mium, iron, and nickel, all of which reflect the
increased corrosion rates.

Following the rupture of the loop tubing which ended
run 7, the corrosion test specimens were again removed
for weighing and metallurgical examination. The speci-
mens adjacent to the ruptured area (hottest section
under normal conditions, 588°C) were badly warped
and partially fused to the tube wall. Thus only
metallographic analysis was possible. The specimens in
the intermediate zone, 555°C, appeared to be un-

harmed with regard to physical appearance. The average
weight loss for these specimens was —1.3 mg/cm? in the
101 hr of run 7. Expected loss for these specimens in
100 hr would be approximately —0.1 mg/em?. Thus, if
we attribute the balance of the corrosion only to the
period of the temperature excursion (~1000°C), the
loss rate would be —1.2 mg/ecm? in 2.5 min or 10,000
mils/year (10 in./year).

Weight losses were also found for the specimens
downstream of the cooler. Although the specimens
were adjacent, one specimen lost 0.3 mg/cm? and the
other 3.0 mg/cm?. Again the severity of the overheating
was evident. Certainly a large part of these losses may
be attributed to moisture contamination of the salt and
air oxidation after the loop rupture. Nonetheless, the
corrosion rate was quite high.

Figure 15.30 shows micrographs of specimens from
FCL-1 after run 4 (9600 total hr exposure to the

206

. Y-99704

TEMPERATURE, °C 588
WEIGHT CHANGE, mg/em? -21.3

555 510
-4.6 -04

Fig. 15.30. Micrographs of Hastelloy N specimens from FCL-1 exposed to NaBF;-NaF (92-8 mole %).

fluoroborate mixture). The left micrograph shows the
surface of the specimen at the hottest position (588°C).
Average weight losses of the specimens were —21.3
mg/cm?. Microprobe analysis disclosed a concentration
gradient of chromium and iron for a distance of 0.6 mil.
There was less than 1.0 wt % Cr and only 1.0 wt % Fe
0.4 mil from the surface. This depleted area is repre-
sented by the darkened area extending from the surface
for approximately 0.5 mil. The center micrograph
shows the surface of a specimen at 555°C. The average
weight loss of these specimens was —4.6 mg/cm?.
Microprobe analysis of these specimens disclosed only a
small concentration gradient for a distance of 0.1 mil.
The micrograph to the right shows the deposit on the
surface of the specimens at 510°C. The average weight
change of the specimens was —0.1 mg/cm?, but this
overall loss stems from a great deal of corrosion on the
leading edge and is not indicative of the whole
specimen. Microprobe analysis of this surface showed a
large amount of iron and a little more nickel than usual
This situation existed for a distance of 0.5 mil, after
which the concentrations of both elements approached
that of the matrix. The average deposit is only 0.2 mil
thick, and it appears that portions of the deposit have
diffused into the sample. Prior evidence in studies of
deposits resulting from temperature gradient mass
transfer have disclosed similar findings." '

The tubing which surrounded the specimens was
removed after run 6 (11,371 hr) and examined both by
optical metallography and scanning electron micros-
copy. Figure 15.31 shows the tubing exposed to salt at
588°C. In Fig. 15.31a (the optical micrograph) the
surface roughening (some salt is still in place) can be
seen, and the attack at the grain boundaries is evident.
In Fig. 15.31b (the scanning electron micrograph) the
delineation of the grains is visible. Figure 15.32 shows
the tubing exposed to the salt at 555°C. Much less
attack is noted. Figure 15.33 shows the tubing down-
stream of the cooler in the coldest position (510°C) of
the loop. The deposit is visible in the optical micro-
graph and in the scanning micrograph in the form of
nodules. The upper portion of Fig. 15.33¢ shows a side
view of the nodular deposits.

15.8.3 Forced-Convection Loop MSR-FCL-2

Assémbly of a second molten-salt forced-convection
loop, MSR-FCL-2,'? is nearing completion. This loop,
constructed of Hastelloy N, will be used to study the

11. J. W. Koger and A. P. Litman, MSR Program Semiannu.
“eb. 28, 1970, ORNL-4548, p. 244.

R. Huntley et al., MSR Program Semiannu. Progr.
31, 1970, ORNL-4622, pp. 176-78.

Fig. 15.31. Hastelloy N tubing from FCL-1 exposed to NaBFs-NaF (92-8 mole 7%) at 588°C for 11,371 hr. (a) Optical
micrograph, S00X. (b) Scanning electron micrograph, 1000X.

208

T

|
looo7

(a) =\ : Salia B aSil

Fig. 15.32. Hastelloy N tubing from FCL-1 exposed to NaBF;-NaF (92-8 mole %) at 555°C for 11,371 hr. (a) Optical
micrograph, 500X. (b) Scanning electron micrograph, 1000X.

209

Fig. 15.33. Hastelloy N tubing from FCL-1 exposed to NaBFs-NaF (92-8 mole %) at 510°C for 11,371 hr. (a) Optical
micrograph, S00X. (b) Scanning electron micrograph, 1000X. (c) Scanning electron micrograph, 3000X.
mass transfer properties of Hastelloy N in fluoroborate-
type coolant salt systems at conditions proposed for the
MSBR.

All primary loop piping and components have been
installed. The installation of auxiliary heaters, thermo-
couples, thermal insulation, and service lines has been
completed except for those required for the pump
bowl—auxiliary tank assembly. Checkout procedures
for acceptance of the installation of the loop facility are
in progress. :

Performance tests of the ALPHA pump,'? to be used
for circulating salt in loop MSR-FCL-2, have been
completed and indicate that the pump satisfies the
MSR-FCL-2 hydraulic requirements of 4 gpm at
approximately 110 ft head at a shaft speed of approx-
imately 3900 rpm. Based on the results in a water test
stand, the final design of the pump tank and internal
parts was completed, and these parts were fabricated of
Hastelloy N for use in the pump for MSR-FCL-2.

A draft of the loop operating manual has been
prepared. The manual includes design criteria, a descrip-
tion of the loop and its components, procedures to be
used for system checkout, and detailed operating
procedures.

The instrument and control panels have been installed
and have been functionally checked and accepted. We
have incorporated safety features in the control system
to prevent accidental overheating of the loop piping,
which caused the failure of loop MSR-FCL-1 described
above.

15.9 RETENTION OF TRITIUM BY
SODIUM FLUOROBORATE

J. W. Koger

Because of the interest in the disposition of tritium!*
in a molten-salt breeder reactor, we conducted an
experiment to determine the extent to which tritium
(injected as a gas) would be retained in a fluoroborate
salt mixture circulating in a standard thermal-convec-
tion loop. The loop used for this experiment was
NCL-13A, which is described in Sect. 15.7. Tritium was
added to the loop along with a mixture of He and BF,
gas. The BF; addition was required to prevent a
high-melting NaF plug at the end of the tube through
which the tritium gas was added. The tritium was sealed
in a glass ampul and was placed inside a l-in.-diam

13. A. G. Grindell et al.,, MSR Program Semiannu. Progr.
Rep. Aug. 31, 1969, ORNL-4449, p. 78.

14. J. W, Koger, MSR Program Semiannu. Progr. Rep. Feb.
28, 1970, ORNL-4548, pp. 53-57.

210

copper cylinder which had Y;-in. copper lines leading
from each end. One of the lines was connected to the
He-BF; supply, and the other was lowered into the
flowing salt stream through a standpipe. Prior to adding
the tritium, a flow of He + BF; gas was established into
the salt. Then the loop was depressurized, and the glass
ampul containing 2 Ci of tritium was broken, and the
tritium was carried into the loop. The copper tubing
through which the He, BF;, and T, was flowing
remained in the loop until it was felt that all the tritium
had passed into the salt and until the gas pressure over
the circulating salt increased to ~10 psig. The tritium
addition, if it had gone totally into the salt, would have
produced an average tritium concentration in the salt of
740 uCi per gram of salt.

During tritium injection the loop was completely
shrouded in plastic which was connected to a hood
system over the loop. A tritium detection meter was
placed in the stream of air being pulled through the
shroud, and no tritium was detected. No tritium was
found in samples of insulation of the loop. Tritium was
detected whenever we opened our standpipe connection
above the surge tank and ball valve.

Salt samples were taken periodically and analyzed to
determine the amount of tritium in the salt. The results -
are shown in Table 15.7. The tritium concentration
decreased sharply in the first 40 hr and then decreased
much more slowly.

We conclude from these results that little, if any,
hydrogenous material, with which the tritium could
exchange, existed in the salt. From various analyses it
appeared that the tritium initially equilibrated between
the salt and gas space above the salt in accordance with
Henry’s law and that some of the tritium then escaped
from the vapor phase as salt samples were taken. The
tritium continued to partition between the salt and
cover gas, so that the concentration in the salt would

Table 15.7. Tritium concentration in salt as a
function of time

Concentration of tritium " Time after
Microcuries per Pob tritium injection
_gram of salt P (hr)

47.7 4.8 1
20.2 2.0 15
10.2 1.0 40
6.4 0.64 162
4.7 0.47 210
4.1 041 349
2.8 0.28 566
0.2 0.02 1340

continually decrease. It is probable that some tritium
also diffused into and through the loop wall, although
at a rate below that which could be measured by the
monitoring system around the loop. After 40 hr the
rate of loss of tritium from the salt was small, and some
exchange with hydrogen in the salt could have been
taking place. However, at this time the amount of
tritium in the salt was in the lower parts per billion
range. .

We plan to conduct an additional experiment in
which a stable hydrogenous compound will be added to
the salt prior to the tritium addition.

15.10 SUPPORT FOR COMPONENTS
DEVELOPMENT PROGRAM

J. W, Koger

15.10.1 Metallurgical Examination of Inconel
Bubbler Tube from PKP-1 Pump Loop

Loop PKP-1 has been used to test a molten-salt pump
with the coolant mixture NaBF,;-NaF (92-8 mole %).
After loop shutdown an Inconel bubbler tube which
admitted BF; to the salt mixture was examined to
determine its behavior in the loop environment.'® Of
particular interest was the fact that before shutdown,
plugging of the tube was suspected (see Chap. 5).

The 13-in.-long tube was 0.625 in. in outside diameter
and 0.500 in. in inside diameter (62.5 mils wall
thickness), and the nominal composition of the Inconel
600 was 76% Ni, 16% Cr, and 8% Fe. The inside of the
tube was exposed to a mixture of He and BF; (13.5 vol
% BF;). This gas mixture was supplied to the salt
during the experiment to avoid formation of high-
melting NaF, which would result if BF; were lost from
the NaBF,-NaF mixture and not replaced. The outside
of the tube was exposed to the fluoroborate salt
mixture at ~550°C (1025°F). The time of gas and salt
exposure was ~11,500 hr (1.3 years). '

Fluoride salt attack generally manifests itself as
selective removal of the alloy constituent which forms
the most stable fluoride. Thus the order of removal in
Ni-Cr-Fe alloys is chromium, iron, and nickel. These
elements form corrosion products which are soluble to
some extent in the melt. In highly impure salt, all of the
elements can be oxidized simultaneously, while in a
purer salt, only chromium is oxidized. In the presence

15. R. B. Gallaher and A. N. Smith, MSR Program Semiannu.
Progr. Rep. Feb. 28, 1970, ORNL-4548, pp. 69-72.

211

of a temperature gradient, oxidation of the alloy
constituents occurs in the hot section with concomitant
reduction of the fluoride in the cold leg. This leads to
weight losses and gains in the resperctive loop sections.
If the solubility of any of the corrosion products is
exceeded, then the fluorides will deposit, usually in the
cooler portion. Nickel fluoride apparently has little
solubility or stability in NaBF,-NaF (92-8 mole %)
since conditions leading to the oxidation of nickel
consistently lead to deposition of nickel-rich fluoride
compounds. Under the same conditions large concen-
trations of chromium and iron fluorides are observed in
samples of the salt. |

Figures 15.34 and 15.35 show the appearance of the
outside and inside of the Inconel 600 bubbler tube after
removal from the PKP-1 loop. We observed that the
bottom 4 in. of the outside of the tube had a silver
appearance, while the upper 3.5 in. was much darker.
The interface between these areas is probably indicative
of the average liquid level. On the inside of the tube,
the lower 0.5 in. of the tube was almost completely
plugged with a black material, the next 1 in. appeared
to be relatively clean, and the rest of the tube contained
green material (probably Na;CrF¢) with smaller
amounts of black material.

From the above, we concluded that the material
found below the liquid level inside the tube was very
probably a nickel-rich corrosion product resulting from
salt attack (not necessarily at that point). On the other
hand, the material above the liquid level undoubtedly
formed in place as the result of attack by H,0
impurities in the BF; and was comprised of chromium,
iron, and nickel fluorides.

There is a zone approximately 4 mils deep in the
etched photomicrograph of Fig. 15.36 where the
metallographic etchant has heavily attacked the surface.
The behavior of this zone suggests that it is almost pure
nickel and that the iron and chromium have been
preferentially removed by the salt. These results were
confirmed by x-ray fluorescent analysis.

Figure 15.37 shows a cross section of the tubing wall
near the liquid-gas interface, This section had an average
measured thickness of 57.5 mils, compared with a
nominal thickness of 62.5 mils before test. A zone of
metallic crystals is visible on the ID to a thickness of
approx 20 mils, In view of the apparent loss of wall
thickness in this area, these crystals may represent
vestiges of the original metal surface. Figure 15.38
shows a higher magnification of the ID surface, both as
polished and etched, and tends to corroborate that the
crystals are a result of material leaving the surface
rather than material deposition.
In conclusion, our investigation disclosed that about 5 tube as a consequence both of BF, attack and of
mils of material was removed from the Inconel 600 deposition of nickel-rich deposits at the lower opening.
tube. Attack by both salt and BF; vapor was evident. The latter deposit caused a restriction in flow.
Corrosion products were produced on the inside of the

PHOTO 78235

INCHES

Fig. 15.34. Photograph of the outside of Inconel 600 bubbler tube from PKP-1 pump loop exposed to NaBF;-NaF (92-8 mole
%) for 11,500 hr at 550°C.

PHOTO 78258

INCHES

Fig. 15.35. Photograph of the inside of Inconel 600 bubbler tube from PKP-1 pump loop exposed to 13.5 vol % BF3—bal He for
11,500 hr.

213

Y-103138

500X

Iooor ™

(a)

Y-104199

foo0rm

0.007 INCHES

looor

Fig. 15.36. OD surface of Inconel 600 bubbler tube from PKP-1 pump loop exposed to NaBF,-NaF (92-8 mole %) for 11,500
hr at 550°C. (@) As polished, S00X. (5) Etched with aqua regia, S00X.
214

Y — 104140

tooom.

0.035 INCHES
100X

o030 m.

oD 1D

Fig. 15.37. As-polished cross section from near liquid-gas interface of Inconel 600 bubbler tube from PKP-1 pump loop, 100X.

215

Y—103437 =

16003 i

6605w T 500X

(o) &

Y—104200 __

0.007 INCHE,

looo7m T

Fig. 15.38. 1D surface of Inconel 600 bubbler tube from PKP-1 pump loop exposed to 13.5 vol % BF3—bal He for 11,500 hr. (2)
As polished, 500X. (b) Etched with aqua regia, S00X.

15.11 CORROSION OF HASTELLOY N IN STEAM
B.McNabb  H.E.McCoy

The sample chamber was removed from the Bull Run
facility for modification after 6000 hr of operation at
538°C in 3500-psi steam. Figure 15.39 is a schematic of

the specimen chamber modifications being made to

facilitate stressing of specimens in the steam environ-
ment., The facility as originally installed was described
previously.'® The specimens will be double-walled
tube-burst specimens with an annulus between the walls
into which the inner tube will burst during the test. The
annulus is- connected to the steam condenser by a
capillary tube. A thermocouple is connected to this
capillary to sense failure by an increase in temperature
when steam is introduced into the annulus by failure of
the specimen. Pressure measurements will be made in
the specimen chamber by a Leeds and Northrup

216

pressure transmitter with a stainless steel Bourdon tube

16. B. McNabb and H. E. McCoy, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1969, ORNL-4449, pp. 205-9.

AN N

TUBE BURST
SPECIMEN

CLUSTER OF FOUR
TUBE BURST SPECIMENS
CLAMPED TO THERMO WELL

= 11

sensing element, having a 0-to-5000-psi pressure span.
The specimen wall thickness will be machined to obtain
the desired stress. The outer thick wall of the double-
walled specimen is designed so that it will not collapse
into the annulus during operation in the 3500-psi
steam, By being able to continuously stress specimens
in this way, we can assess the compatibility of Hastelloy
N with steam and determine if a stress corrosion
problem exists,

There appears to be no problem with the corrosion of
unstressed Hastelloy N. Specimens exposed at 538 and
593°C are still following the trends noted previously
(Fig. 15.40). The rates of air-melted and vacuum-melted
Hastelloy N are converging at 13,200 hr at 593°C,’
whereas air-melted Hastelloy N previously had a slightly
higher rate. The present unstressed specimens will
continue to be used for weight gain measurements and
metallographic examination of the penetration of the
oxide on selected specimens, but the ability to stress
specimens and retain the stress will be a valuable
addition to the facility. The modification should be
complete in about two weeks and ready for reinstalla-
tion,

ORNL-DWG 683-3995A

THIN-WALL
TUBING

CAPILLARY
TUBE

CAPILLARY
TUBE

§ N
N0 ) \\\IW -
C =3 BN
EEESE | et A e
Fl N\ ) < 72 ////,V/ —L\\\\\\
LTER
waCoen SRR

SAMPLE VESSEL

Fig. 15.39. Schematic of steam corrosion facility at Bull Run Steam Plant showing modifications for dynamically stressed tubes,
IS

WEIGHT CHANGE (mg/cm?)

217

ORNL-DWG 70-49332RA

Corrosion of Hastelloy N in Steam.

Fig. 15.40. Corrosion of Hastelloy N in steam.

2Ya Cr STEEL-SSEy
l/
/. | mpy_~
) et
-~
b -l
/o //
1/ = N, AR -
i L MELTED 593°C
/ - 0.5 mpy L.
~ L =
1 _ — /n [
/ L~ L1 T N, VACUUM
> T — MELTED 593°C ]
AP == e
~ Lo L e Qo |
P — |_—538°C, N, AIR AND
z P a =" VACUUM MELTED
2000 4000 6000 8000 10,000 12,000
TIME (hr)

14,000
16. Support for Chemical Processing

J. R. DiStefano

This portion of our materials program is in support of
the program to develop techniques for processing the
molten-salt fuel to isolate protactinium and remove
fission products. We have continued to concentrate our

efforts on the materials which appear most promising.

for containment of liquid bismuth at 500 to 700°C,
since it is an essential element in the reductive-
extraction process now being developed. Our program is
divided into two parts. One is the fabrication of a
molybdenum loop for the Chemical Technology Divi-
sion to obtain hydrodynamic data relating to the
reductive-extraction process. During the reporting per-
iod considerable progress was made in fabricating and
joining molybdenum components.

The second part of our program is research to study
the -compatibility of molybdenum, TZM, graphite,
tantalum, T-111, and iron-base brazing alloys with
liquid bismuth and bismuth-lithium solutions and to
investigate possible techniques for coating iron- and
nickel-base conventional alloys with corrosion-resistant
metals such as tungsten or molybdenum.

16.1 CONSTRUCTION OF A MOLYBDENUM
REDUCTIVE-EXTRACTION TEST STAND

J. R. DiStefano

We are fabricating a molybdenum loop for the
Chemical Technology Division to obtain engineering
data on a reductive-extraction method of MSBR fuel
processing. Details of the design of this loop have been
reported previously.' .

The loop consists of a 5-ftlong, 1%-in. packed
column through which bismuth and salt will circulate
countercurrently. Two 37-in.-diam head pots will feed
bismuth and salt to the column, and two 37%-in.-diam
.containers will serve as the end sections of the column
where bismuth and salt will be separated. The loop will

H. E. McCoy

be interconnected by %-in.-OD by 0.020-in.-wall, %-
in.-OD by 0.025-in. wall, %-in.-OD by 0.030-in.-wall,
and %-in.-OD by 0.080-in.-wall tubing.

Welding and brazing techniques are being used to
fabricate and assemble the loop. However, mechanical
couplings are also desirable since they would allow us to
easily replace certain internal components in case of a
failure, and would also allow a convenient access port
for removable corrosion specimens. Experimental mo-
lybdenum metal seal couplings were obtained from
Stanley Corporation and Aeroquip Corporation. Both
of these couplings use molybdenum seal rings, but a
threaded nut applies the force that seals the Stanley
joint, while an external compressive force must be
applied to seal the Aeroquip joint and then a gate or pin
is used to maintain the compressive force on the metal
seal ring. This latter design was aimed at remote
application and is attractive because molybdenum
components easily gall, and the threaded nut design is
particularly susceptible to this problem. However,
experimental problems in sealing the Aeroquip joint
caused cracking of one of the molybdenum compo-
nents, and we have not yet obtained replacements.

Helium leak tests were performed on the Stanley
coupling as well as the coupling shown in Fig. 16.1.

ORNL-DWG 7{-3181

MOLYBDENUM

L L L L L

C.25in.

SO\

GRAFOIL GASKET

1. W. F. Schaffer, E. L. Nicholson, and J. Roth, “Design of a
Processing Materials Test Stand and the Molybdenum Reductive
Extractive Equipment,” MSR Program Semiannu. Progr. Rep.
Aug, 31, 1970, ORNL-4622, pp. 11213,

MOLYBDENUM

Fig. 16.1. Molybdenum mechanical coupling.

218

This latter coupling also has molybdenum components.

as indicated, but the seal gasket is made of Grafoil.?
Each joint was helium leak-tight (<5 X 107® std
ccfsec). They were then thermally cycled ten times by
heating to 650°C and cooling to room temperature. At
this point the helium leak rate of the Stanley joint was
5 X 107° std cc/sec at both room temperature and at
650°C. However, the joint shown in Fig. 16.1 remained
helium leak-tight (<5 X 1078 std ccfsec). We disas-
sembled the Stanley joint and returned it to the
company for further evaluation. The joint in Fig. 16.1
was disassembled, resealed, and then thermally cycled
several more times, It was helium leak-tight as before
and therefore was again disassembled. This coupling was
more easily disassembled than previous joints using
molybdenum nuts because of its modified, rounded
thread design.

In addition to joints which can be disassembled,
mechanical tube-to-header connections are being devel-
oped to afford a method of attaching certain of the
loop dip lines. We feel that welding is not satisfactory
because of the weld joint design, and brazing is
uncertain because of the requirement that the braze
alloy be resistant to HF at 650°C as well as to bismuth
and salt. Two techniques, magneforming and roll
bonding, have been investigated. In magneforming,
magnetic pressures of up to 50,000 psi resulting from a
current discharged through a coil from a capacitor bank
are applied in pulses of 10 to 20 X 107® sec duration.
No direct mechanical contact between the maching that
applies ‘the force and the work is involved. Four
tube-to-tube joints were fabricated, and the best joint

produced showed a helium leak rate of 2 X 107° std -

cc/sec at room temperature. In roll bonding, a tool
containing expandable tubes mechanically forces the
two surfaces together. The operation can be done at
temperatures up to about SO0°C. Three helium-leak-
tight joints (<5 X 107® std cc/sec) have been made. If a
joint of this type is further strengthened by back
brazing with one of the iron-base filler metals discussed
below, we feel that it will be satisfactory for our
application.

We will continue to evaluate mechanical joints that
can be disassembled as well as methods of mechanical
sealing as discussed above. We feel that development of
alternative techniques for joining molybdenum will be
essential to the use of this material for future chemical
processing applications.

More detailed information on this loop is presented in
Part 5 of this report. Progress on welding, brazing, and

2. Registered trademark of Union Carbide Corporation.

219

fabrication of molybdenum components is reported in
this section.

16.2 FABRICATION DEVELOPMENT
OF MOLYBDENUM COMPONENTS

R. E.McDonald  A.C. Schaffhauser

We are supplying molybdenum components for a
chemical processing test stand. This work involves
development of fabrication processes for some compo-
nents, procurement of materials that are commercially
available, and coordinating the machining of test
components. .

We have previously described our development of a
back-extrusion process® for fabrication of 37%-in.-diam
closed-end vessels for the bismuth and salt head pots
and the upper and lower disengaging sections of the
extraction column. For welding studies, we have back
extruded and machined eight vessels of both hemispher-
ical and flat-end geometries with bosses for supporting
inlet tubes. These vessels were back extruded at 1350 to
1450°C with a 750-ton force on the plunger. The
as-extruded parts, shown in Fig. 16.2, demonstrate the
excellent internal surface produced by back extruding
over a ZrO,-coated plunger, whereas the external
surface is very rough due to interaction with the
tool-steel container during extrusion. The cross section
shown in Fig, 16.2 also shows the deformation pattern
produced by back extrusion of a square grid network in
the original blank. A finished machined vessel is shown
in Fig. 16.3.

The cracks resulting from interaction- of the outside
of the back extrusion with the extrusion press container
have limited the length of vessel we can back extrude to
about 4 in. To overcome this problem we have designed
and procured a ZrQO,-coated liner for our 5.6-in.-ID
extrusion press container. This tooling should allow us -
to back extrude vessels & in. deep.

The extraction column of the test stand requires a
5Y, -ft-long Mo pipe 1.16 in. in outside diameter and 1
in. in inside diameter. We have fabricated three pieces
of this pipe by floating mandrel extrusion of two
4-in.-diam billets with 1-in. holes at 1600°C at a
reduction ratio of 29:1. The second extrusion produced
a pipe 11', ft long that was concentric within 0.007 in.
with excellent outside and inside diameter surfaces.

All of the tubing required for the test stand ranging in
size from 7 in. OD and 0.080 in. wall thickness to ¥,

3. R. E. McDonald and A. C. Schaffhauser, MSR Program
Semiannu, Progr. Rep. Feb. 28, 1970, ORNL-4548, pp.
253-54.
in. OD and 0.020 in. wall thickness has been procured
and inspected. A prototype of a tee, which is one of the
many fittings required, was designed and fabricated
from Y,-in.-thick Mo plate.

Y- 105581

INCHES

Fig. 16.2. Back-extruded molybdenum vessels. The cross
section shows the deformation pattern produced by back
extrusion of a square grid network in the original blank

¥-103375

INCHES

Fig. 16.3. View of back-extruded molybdenum vessel after
machining.

220

16.3 WELDING MOLYBDENUM
A.J.Moorhead  T.R.Housley

In order to develop welding procedures for fabrica-
tion of the molybdenum loop for chemical processing,
we have been working on three major types of joint:
tube to header, tube to tube, and header to header. A
discussion of our initial work on the first two of these
joints and our basic fabrication philosophy was de-
scribed in an earlier report.* Since that time, we have
greatly improved our capabilities for joining molyb-
denum, which is inherently very difficult to weld. Most
significantly we have found that stress relieving the
components at 925°C prior to welding and preheating
the larger weld joints minimizes weldment cracking.

Five different sizes of tubing must be welded to the
various pots in order to fabricate this system. These are
Y-, Y- Y-, Y-, and 1%-in.-OD tubes. All of these
welds are made by the electron-beam process using a
trepan on the inside of the header. Essentially, the weld
is made on an edge joint between two components of
equal thickness. Figure 16.4. shows this type of weld on
a mockup of the 1%-in.-diam packed column. This
weld had a helium leak rate of less than 1 X 107 atm
cm? sec™. We have also welded three tubes of this size
into back-extruded molybdenum headers. The first
weld" did not pass helium leak inspection, and dye
penetrant revealed that there was extensive fusion-zone
cracking. We were unable to eliminate these cracks by
rewelding. Part of this failure may be due to contamina-
tion from dye (left in the initial cracks), which is very

4. A. J. Moorhead and T. R. Housley, MSR Program
Semiannu, Progr. Rep. Aug. 31, 1970, ORNL4622, pp.
185-89.

ORNL-DWG 71-7064
TREPAN GROOVE

14 in

"NOMINAL

—-o0szin 'F°°:" VA
4
50 in.—

N

(@) (8

16.4. (a) Electron-beam-welded tube-to-header joint

i
Y liam, 0.062-in.-wall tube); (b) joint design.

¥-106004

INCHES.

Fig. 16.5. Tubes (0.375 in. diam, 0.025 in. wall; 0.500 in.
diam, 0.030 in. wall) electron-beam welded into a back-
extruded molybdenum header.

difficult to remove. A defocused electron beam was
used to preheat the part prior to making the next two
1Y-in.diam tube-to-header welds, and both of these
welds passed an inspection with fluorescent dye pene-
trant. Three other sizes of tubes (%-, %-, and %-in.-
diam) have also been successfully electron-beam welded
into back-extruded headers. Figure 16.5 is one of these
headers with a %- and a Y-in.-diam tube.

Procedures have been developed for making tube-to-
tube welds outside the customary glove box by using
the orbiting-arc weld head. This commercially available
device has been modified by adding small rubber boots
which surround the tube on either side of the head to
provide a better atmosphere for welding. Helium-leak-
tight welds have been made joining %-, %-, and %-in.
tubes. The modified welding head and some typical
tube-to-tube “field” welds are shown in Fig. 16.6. Note
that some of these welds were made without a weld
insert. However, in order to control the amount of weld
bead protruding inside the tubes, we plan to use these
inserts in all tube-to-tube joints in the loop itself.

Electron-beam girth welds were made between two
back-extruded headers and short rings which repre-
sented other headers. The ring joined in the first
assembly had been machined from the same extrusion
as the header. The assembly was preheated with the
defocused electron beam, but the preheat temperature
was not measured. This weld bead looked good visually,
but penetrant inspection revealed fusion-zone cracking.
The second assembly, which had a ring machined from
bar stock, was similarly welded except that more

221

preheating was done. This weld was defect free when
inspected by fluorescent penetrant. This part is shown
in Fig. 16.7, together with a sketch of the weld-joint
design.

16.4 DEVELOPMENT OF BISMUTH-RESISTANT
FILLER METALS FOR BRAZING MOLYBDENUM

N.C.Cole

Commercial brazing alloys for molybdenum are gen-
erally not compatible with bismuth, To develop filler
metals that have satisfactory resistance to both bismuth
and molten salts in the range 500 to 700°C, we have
formulated several Fe-Mo (Fe base) alloys containing
such elements as C, B, and Ge to depress the melting
temperature. These iron-base alloys have shown excel-
lent wettability and flowability on molybdenum below
1200°C.

It is also desirable to braze at or below the recrystalli-
zation temperature of molybdenum to retain as much
of its ductility as possible: therefore we have deter-
mined the ductile-to-brittle transition temperature of
arc-cast molybdenum sheet after relatively short times
at temperature. We heat treated sheet specimens at
time-and-temperature combinations similar to the braz-
ing cycles for the iron-base alloys. The heat-treated
sheets were then bend tested, and those that bent 90°
without cracking were called ductile. Specimens heat
treated at temperatures up to 1180°C and for times as
long as 40 min were ductile. Specimens heat treated 3
min at 1200°C bent 90°, but those heat treated at
1200°C for 10 min achieved only a 50° bend. Since the
iron-base alloys all braze below 1180°C and the Fe-Mo-
Ge-C-B alloy brazes below 1100°C, we should have
enough of a safety factor in case the working history of
other shapes or sizes, such as tubing, pots, etc., causes
the base-metal recrystallization temperature to be
slightly lower.

To evaluate the mechanical properties of the brazed
joints, we have shear tested two of the most promising
iron-base alloys. Shear-test specimens made of molyb-
denum were brazed with 42M [Fe—15 Mo—5 Ge-4
C—1 B (wt %)] and 35M [Fe—15 Mo—4 C—1 B (wt
%)] . They were pulled at a strain rate of 0,002 in./min
at both room temperature and 650°C. Average values of
the test results are shown in Table 16.1. The shear
strengths are excellent for both brazes at room tempera-
ture and very good for 42M at 650°C. The shear
strength for 35M at 650°C is acceptable. The elongation
for both brazes is acceptable at room temperature and
excellent at 650°C.

222

Y-104203

Y-104357

#16 With Insert
(welded twice)

#12 Without Insert
(welded once)

#17 With Insert
(welded once)

Components

Fig. 16.6. (a) Modified orbiting-arc head for tube-to-tube welds; (b) typical ‘/yin.diam welds made using this device.

223

Table 16.1. Mechanical properties of brazed molybdenum joints

Shear strength (psi)

Elongation (%)

Brazing alloy (wt %)

Room temperature 650°C Room temperature 650°C
42M (Fe—15 Mo—5 Ge—4 C—1B) 30,000 29,000 10 50
35M (Fe—15 Mo—4 C—1B) 31,000 18,000 i1 42

ORNL-DWG 71-7085

MACHINED BACK—EXTRUDED
RING HEADER

i

i
0188in. % K\\\} N\ § 0425 in.
I L\

0425 im——l

FUSION
ZONE

(&)

Molybdenum joints brazed with the experimental
iron-base filler metals have been tested numerous times
in static bismuth and have been reported previously.>*®
They now have also been tested in a thermal-convection
loop which operated 2100 hr with a maximum tempera-
ture of 700°C and a AT of 90°C. Visually, the filler
metals survived extremely well with little, if any,
removal of the braze. Details of this test are given in the
following section,

5. N. C. Cole, J. W. Koger, and R. W. Gunkel, MSR Program
Semiannu. Progr. Rep. Feb, 28, 1970, ORNL-4548, pp.
255-56.

6. N. C. Cole and J. W. Koger, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 189.

Fig. 16.7. (a) Electronbeam girth weld on back-extruded
molybdenum header: (5) joint design.

In addition, molybdenum joints brazed with three of
the experimental iron-base alloys, 35M, 42M, and 16M
[Fe—4 C—1 B (wt %)], were corrosion tested in a
molten-salt environment. They were exposed to LiF-
BeF,-ZrF4-UF,-ThF; (70-23-5-1-1, mole %) for
1032 hr at approximately 670°C in a type 304L
stainless steel thermal-convection loop which operated
at a maximum temperature of 690°C and a AT of
100°C, Each of the samples showed a slight weight gain
of approximately 1 mg/cm®.” Metallographically, we
did not see any evidence of attack. Figure 16.8 shows a
typical as-brazed sample compared with the same braze
after test,

In building a complex structure such as the chemical
processing testing loop, repairs of cracked components
may have to be made. To determine whether an
iron-base brazing alloy could be used for repairs, a tube
which was not helium leak-tight because of cracks in
and along the weld was selected for repair. The brazing
filler metal, 42M, was attached (by wire) near the
cracks and heated to its flow temperature. It flowed
over and along the surface of the tubing, completely
covering the weld and cracked area, and after brazing,
the tube assembly was helium leak-tight (<1 X 107
atm cm? sec™'). Figure 16.9 is a photomicrograph of a
section through the tube. The bottom portion is an
insert used to enhance welding. To the left and out of
the picture is the weld. The top portion is the tube wall.
Note how the braze alloy flowed along the surface of
the insert into the 0.060-in. gap between the insert and
the tube and into the small crack (0.002 in. wide) in the
insert.

7. 1. W. Koger, private communication.

Y—105018 __

°

2

IS
" - 8 4%
s e

She

2z

H

ol

3

g

g

IS

Y—103078

Fig. 16.8. Molybdenum lap joint brazed with 35M [Fe—15 Mo—4 C—1 B (Wt %)]. (a) As brazed. (b) After testing in fluoride
salts at 67°C for 1032 hr.

e s

Y—105621

fooiom.

00X

0.035 INCHES
T

[CLED

Fig. 16.9. Portion of a cracked molybdenum tube that was braze repaired.

16.5 COMPATIBILITY OF MATERIALS
WITH BISMUTH

0.B.Cavin  L.R. Trotter

We are evaluating the compatibility of potential
structural materials and braze alloys with bismuth and
bismuth containing up to 2 wt % lithium. The tests are
conducted in thermal-convection loops that operate at a
maximum temperature of 700°C and a AT of 95 * 5°C
and in static capsules at 700 + 5°C.

Static capsule tests of iron-base braze alloys (see Sect.
16.4) developed for joining molybdenum indicated that
some of these alloys were quite resistant to dissolution
by bismuth.® Hence we subsequently tested the same
alloy compositions in a quartz thermal-convection loop
(No. 10) for 2100 hr to determine their mass transfer
characteristics. The loop specimens consisted of molyb-

8. N. C. Cole and J. W. Koger, MSR Program Semianni.
Progr. Rep. Aug. 31, 1970, ORNL4622, p. 189.

denum tabs brazed with four different iron-base alloys.
In addition to iron, the braze alloys contained 4 wt %
C, 1 wt % B, varying amounts of molybdenum (0, 15
and 25 wt %), and in one alloy (42M) 5 wt % Ge. We
found little evidence of attack of the braze alloys from
either chemical analysis or visual examination. How-
ever, each of the braze alloys was covered by a surface
layer. Photomicrographs of alloy 35M (Fe—15% Mo—
4% C—1% B) in the before- and after-test conditions are
shown in Fig. 16.10. An electron probe analysis was
made to determine the distribution of iron, molyb-
denum, and bismuth in the braze. The areas examined
are outlined by the black rectangles in Fig. 16.10 and
are shown in Figs. 16.11 and 16.12. Certain areas of the
braze are seen to contain small quantities of bismuth,
and it is interesting to note that they are associated
primarily with regions of higher molybdenum concen-
tration. As seen in Fig. 16.12, bismuth has penetrated
completely through the braze fillet, but none was found
in the base metal. This association of bismuth with
molybdenum cannot be explained on the basis of
known solubility data but could be related to the

Y—-105023 -
o
g
g
ki
g
sz
Y—104431

035 INCHE!
100X

% B). (a) Before test; (b) after 2100 hr in flowing

Fig. 16.10. Photomicrographs of braze alloy 35M (Fe-15% Mo-4% C
bismuth at 670°C. Enclosed rectangles 1 and 2 are shown in Figs. 16.11 and 16.12 respectively

Y-105042

BACKSCAT

-

ERED ELECTRONS

MolLa FeKa

Fig. 16.11. Electron beam scanning images of an Fe-Mo braze (35M) of region I in Fig. 16.10b. Lighter regions indicate greater
concentrations of metals shown

LTt

Y— 406001

BISMUTH La X—RAYS

Fig. 16.12. Electron beam scanning images of an Fe-Mo braze
(35M). This is region 2 of Fig. 16.10h. Lighter regions indicate
greater concentration of Bi.

formation of a complex intermetallic compound that is
high in molybdenum and bismuth. The continuous
surface layer on the braze is high in iron, and no
significant differences between samples from the high-
and low-temperature regions of the loop were observed.

Figure 16.13 shows photomicrographs of an Fe—4%
C—1% B (16M) braze before and after test. Electron
beam microprobe results of the area enclosed by the
rectangle in Fig. 16.13b are shown in Fig. 16.14, and
we can see that the surface layer is again rich in iron.
After test this braze was found to contain a relatively
high concentration of bismuth. The area denoted as “I"
in Fig. 16.14 contained approximately 88% Bi, 5% Mo,

Table 16.2 Results of graphite test after 3000 hr
in flowing bismuth

Test Weight Spectrographic
Grade® temperature change analysis
o) @x 107%) (ppm Bi)
AXF-5QBG 615 0 5
620 -13 50
685 20
700 20
AXF-5Q 620 500
630 200
ATIS 650 <2
675 200

“Order of increasing pore volume.

and <% % Fe, with the balance unknown. The area
denoted as “II” in Fig. 16.14 contained approximately
41% Bi, 23% Mo, 4% Fe, and balance unknown.
Although the braze has been generally depleted of iron,
one small area, designated as “IlI” in Fig. 16.14, was
predominantly iron (approximately 70%).
Semiquantitative spectrographic analyses of samples
taken from the bismuth drained from the loop indi-
cated concentrations of iron and molybdenum to be
less than the detectable limit of 3 ppm. These levels are
considerably lower than those found in static capsule
tests of most of the alloys. We plan further tests to
determine why the apparent solubilities were lower in
the thermal convection loop test and also to determine
the source of the iron-rich layers found in the loop test.
Quartz thermal-convection loop 8, which contained
specimens of three grades of graphite in both the hot
and cold legs, completed a scheduled 3000-hr test, and
test data are given in Table 16.2. The samples from this
loop were relatively free of retained bismuth except for
a very few surface particles that were mechanically
dislodged. Poco grade AXF-5QBG has the smallest
amount of open porosity of the three grades tested, and
we did not metallographically observe a significant
amount of bismuth intrusion into this sample. The
small amount detected by semiquantitative spectro-
graphic analysis was probably due to some of the small
particles lodged on the surface that were not removed.
The more open grades, AXF-5Q and ATIJS, both
contained bismuth in the pores as evidenced by
metallographic examination and chemical analysis. A
part of the observed weight loss is probably due to the
loss of machining dust, which is difficult to remove
from sample surfaces. The maximum loss is equivalent
to about 2 mg/cm? per 3000 hr or, assuming a uniform

229

Y—126026

too0 .

0.035 INCHES

too%0m.

(6)

Fig. 16.13. Photomicrographs of 16M braze (Fe—4% C—1% B). (a) Before test; (b) after testing in flowing bismuth for 2100 hr at
685°C. Microprobe analysis of area enclosed in black rectangle is shown in Fig. 16.14
Y-106000

MoLa

Fig. 16.14. Electron beam scanning images of 16M braze (Fe—4% C—1% B). Compositions of regions I, II, and 111 are given in
text. Lighter regions indicate greater concentration of the metal indicated.

0€T

surface removal, 1.2 mils/year. Chemical analysis of
four different samples taken from the bismuth drained
from the loop did not indicate any increase in the
carbon concentration. These results indicate that graph-
ite to be used in chemical processing applications
should have small pore entrance diameters (<1 u) and
should not have interconnecting porosity.

Since removal of rare earths from MSBR fuels
involves bismuth that contains up to 50 mole % lithium,
we are testing the various potential containers in
bismuth-lithium solutions. Static capsules containing
two grades of graphite have completed a 500-hr test in
bismuth containing 0.5 and 3.5 wt % lithium but have
not been examined yet.

Two all-metal thermal-convection loops will soon be
placed in operation. The first of these was fabricated
from gun-drilled low-carbon, low-oxygen molybdenum
bar stock and contains molybdenum tensile samples in
both the high- and low-temperature regions. The second
- was fabricated from T-111 (Ta—8% W—2% Hf) alloy
tubing and contains T-111 tensile samples. The bismuth
in each of these loops will contain 2.5 wt % Li and will
operate at a maximum temperature of 700°C. A
graphite loop is currently being designed and will be
operated under conditions-similar to those discussed
above,

We are planning to continue testing these three
potential structural materials (Mo, T-111, graphite) in
various concentrations of lithium in bismuth to deter-
mine their physical and mechanical property changes.
One point of interest is the possible formation of
lithjum carbide or intercalation compounds with graph-
ite. Future tests will also include the compatibility of
braze alloys with bismuth containing lithium.

16.6 CHEMICALLY VAPOR DEPOSITED COATINGS
| J. 1. Federer

Although fabricable into equipment for reprocessing
MSBR fuel, most iron- and nickel-base alloys are
attacked by liquid bismuth. Tungsten and molybdenum
are resistant to liquid bismuth, but they are much more
difficult to fabricate. Therefore, we are investigating the
use of tungsten and molybdenum coatings on iron- and
nickel-base alloys as a possible solution to this difficult
containment problem. The coatings are being applied
by chemical vapor deposition techniques using hydro-
gen reduction of WFs and MoF4 at about 600 and
900°C respectively. We have continued to characterize
coatings on small sheet-type specimens, and we are
attempting to demonstrate the applicability of the
coating process by coating a variety of test vessels and
more complicated shapes.

231

Tungsten coatings do not adhere to stainless steels
unless the steels are first plated with a thin layer of
nickel. Nickel is readily applied to simple shapes by
electrodeposition, but complicated shapes are difficult
to plate by this method. As an alternate method, we are
investigating electroless nickel plating from an acidified
phosphate bath. In our first experiment, bend speci-
mens of types 304 and 430 stainless steels measuring 10
X % X Y, in. were electroless plated with nickel,
which was then bonded to the base materials by heating
to 800°C for 30 min. Duplicate specimens of each steel
were coated with about 0.005-in.-thick tungsten, and
another pair' was coated with about 0.005-in.-thick
molybdenum. The tungsten coated smoothly, but the
molybdenum coating had numerous blisters. Moreover,
when subjected to a bend test, all of the coatings were
less adherent than had been previously observed when
an electrodeposited nickel layer was first applied. The
electroless nickel plate contained about 8% P, which
may have affected coating adherence. We are continuing
this investigation by testing nickel plates of lower
phosphorus content.

The largest object that we have coated was a
4%, -in.-OD by 36-in.-long Monel vessel closed on both
ends except for ¥-in.-OD tube extensions. The entire
inner surface was coated with 0.006- to 0.010-in.-thick
tungsten except -for the ends, where the coating was
thinner. A 4-in.-long section of the coated vessel was
thermal cycled 25 times between 25 and 600°C.
Examination after the 25 cycles revealed the coating
was intact and there was no evidence of spalling. A dye
penetrant revealed only a few edge cracks, probably
caused by cutting the section from the original vessel.
The entire section was slightly out of round after
thermal cycling, possibly as a result of the stress
produced from the difference in thermal expansion
between tungsten and Monel. Similar results would be
expected for tungsten coatings on other materials such
as Inconel 600, Hastelloy N, nickel, and nickel-plated
steels. '

We have deposited molybdenum on the inside surface
of a nickel capsule and tested it in bismuth for 700 hr
at 600°C. No attack of the coating nor the nickel
substrate below was found.”

Figure 16.15 is a schematic representation of a
Hastelloy N corrosion loop that is to be coated with
tungsten so that it can be tested in flowing bismuth.
Since the nickel-base alloy substrate would be readily
attacked by bismuth, a severe test of the integrity of
the coating will be obtained. The loop was modified as
shown in Fig. 16.15 for coating. A cross member was
removed to form a U-shape, thus providing both a gas
inlet and outlet. The U-shape will be coated first and
232

"HZ-WFG EXHAUST *

ORNL—DWG 74— 7066

Hy—WFg EXHAUST*

33 in,

0.65-in, OD «x

HASTELLOY N

AS FABRICATED

3Y%—-in. OD x ‘ (A A
4—2in. 12 ’ i \J) \ @/
in. AC POWER AC POWER
HASTELLOY N SUPPLY - SUPPLY
[ FURNACE ]
N\
N [\ 24 in
\ 5in.
) N
5 n. N
~_ A
WELD

0.065-in. WALL

33 in.

N

“

AS MODIFIED FOR COATING

Fig. 16.15. Arrangement for coating bismuth corrosion loop with iungsten.

then the cross member will be welded in place and
coated.

_Prior to coating the actual loop, a U-shaped piece of
%-in.-OD by 0.072-in.-wall Hastelloy N tubing was
coated. The long sides of the U-shape were 46 in. long
and the short side was 18 in., giving a total length of
110 in., of which 90 in. was heated and coated. The
tubing was heated to 400°C by self-resistance. Hydro-
gen and WF4 flowed through the tubing for 6 hr in one
- direction, followed by 6 hr in the opposite direction.
By using a temperature of 400°C (rather than the usual
550 to 600°C) for a long time, depletion of WF, was
minimized, and most of the heated length of tubing was
uniformly coated. The coating thickness ranged from
about 0.005 to 0.006 in. except for about 2 in. in each
end, where the thickness was only 0.001 to 0.002 in.
We expect to use similar conditions to coat the actual
loop.

A metal transport demonstration vessel has also been
coated with tungsten for protection against bismuth
corrosion. The low-carbon steel vessel is about 3 in. in
inside diameter by 21 in, long and is closed on one end
except for a Y-in.ID by 6-in.-long pipe extension.

Approximately 10 in. of the vessel and the 6-in.-long
pipe were coated. First, the vessel was nickel plated by
electrodeposition, and the nickel was bonded by heat-
ing to 800°C in vacuum. Calrod heating elements were
then fitted to the vessel to provide heat for coating.
However, the elements were not properly spaced to
heat the vessel uniformly. A temperature gradient from
about 400 to 650°C occurred in the section to be
plated. Consequently, the coating thickness ranged
from 0.004 to 0.020 in., with the thickest coating being
in the vicinity of the Y;-in.-ID pipe opening. Other-
wise, the coating appeared to be satisfactory and should
provide the intended protection.

16.7 MOLYBDENUM DEPOSITION FROM MoF,
J. W. Koger

Studies have continued on developing a technique for
depositing molybdenum on iron-base substrates from
MoF, in a molten fluoride salt mixture, Experimental
problems have occurred as a result of too much MoF
in the salt prior to deposition. This has caused attack of
the container vessel and precipitation of metal fluoride
corrosion products along with molybdenum.® An analy-
sis of the coating formed in the previous experiment is
given in Table 16.3. Stoichiometric calculations show
that all the Li, Be, Cr, Fe, and most of the Ni could
exist as fluorides. This is in agreement with our
interpretation of the microprobe analysis, where we
found nickel, molybdenum, and various fluorides. We
have completed an additional experiment with results
similar to those reported above. In subsequent tests we
will limit the amount of MoF¢ to slow down the

9. J. W. Koger, MSR Program Semiannu. Progr. Rep. Aug. 31,
1970, ORNL-4622, p. 197.

233

reaction. At this time the experimental system is being
modified to accomplish these objectives, and tests will
begin-again shortly.

Table 16.3. Chemical analysis of deposit

Element Weight percent
Mo 71.3
F 16.6
Ni 6.4
Li A 2.7
Fe 2.0
Be 1.2
Cr 0.6

Part 5. Molten-Salt Processing and Preparation

L. E. McNeese

Part 5 deals with the development of processes for the
isolation of protactinium and the removal of fission
products from molten=salt reactors. During this period
we have continued to evaluate and develop a flowsheet
based on fluorination—reductive-extraction for protac-
tinium isolation and the metal transfer process for
rare-earth removal. The portion of the flowsheet dealing
with protactinium isolation was simplified considerably

by retaining the isolated protactinium in a secondary .

salt rather than in bismuth at a point intermediate in
the protactinium extraction column. This change results
in a process that will be much easier to operate and will
effect a significant reduction in plant cost. Calculations
on the flowsheet continue to be promising.

We have begun considering oxide precipitation as an
alternative to the fluorination—reductive-extraction
method for isolating protactinium and for subsequently
removing uranium from MSBR fuel salt. Two possible
flowsheets are presented. The utility of these flowsheets
is largely dependent on the development of methods for
separating oxide precipitate from molten-salt streams.

Additional data were obtained on the effect of the
LiF concentration on the distribution of solutes,
including La, Nd, and U, between LiCl-LiF solutions
and liquid bismuth. This information confirmed earlier
indications that LiF in concentrations below about 4
mole % has little effect on the behavior of di- and
trivalent solutes but that the thorium—rare-earth separa-
tion factor decreases with increasing LiF concentration.
The mutual solubilities of thorium and neodymium in
lithium-bismuth solutions at 640°C appear to be much
higher than those required in the extraction of rare
earths from LiCl by contact with lithium-bismuth
solutions.

Our second engineering experiment for the study of
the metal transfer process for removing rare earths from
MSBR fuel carrier salt was successfully completed
during this reporting period. The experiment, which
demonstrated all of the important aspects of the

234

process, was operated for about three months before it
was shut down for inspection. During that period more
than -85% of the lanthanum and 50% of the neodymium
originally in the fuel carrier salt were removed and
deposited in a lithium-bismuth solution. There was no
measurable accumulation of thorium (<10 ppm) in the
lithium-bismuth solution, thus demonstrating that the
rare earths can be removed without significant removal
of thorium. The thorium-lanthanum decontamination
factor was about 10°. The distribution coefficients
between the salt and metal phases in the system were in
reasonable agreement with expected values. A third
engineering experiment, which will use flow rates that
are about 1% of those required for a 1000-MW(e)
MSBR, is being designed and fabricated.

Our work on contactor development was continued
successfully during the reporting period. Flooding data
obtained during the countercurrent flow of bismuth
and molten salt in a 24-in.-long, 0.82-in.-ID column
packed with % -in. molybdenum Raschig rings continue
to be in good agreement with flooding rates predicted
from studies with mercury and water. Data on the
extraction of uranium from salt by countercurrent
contact with bismuth containing reductant were effec-
tively correlated by an HTU model that assumed the
major resistance to uranium transfer to be in the salt
phase. Preparations were begun for additional mass
transfer studies in packed columns; these studies will
involve measuring the rate of exchange of zirconium
isotopes between salt and metal streams otherwise at
equilibrium.

We have continued our efforts to develop a frozen-
wall continuous fluorinator. Studies of heat generation
and heat transfer in a simulated fluorinator system have
shown that radio-frequency heating can be used to
provide the internal heat source required for studies
with nonradioactive systems. A relatively large con-
tinuous fluorination experiment is planned.
17. Flowsheet Analysis

A flowsheet in which fluorination is used for re-
moving uranium and reductive extraction is used for
isolating protactinium from MSBR fuel salt has been
described.'’? However, it has been found that a
considerable simplification in the flowsheet and a
significant reduction in partial fuel cycle cost can be
achieved by holding the isolated protactinium in a
secondary salt phase rather than in bismuth as pre-
viously considered. A flowsheet for this improved mode
of operation has been developed, and the partial fuel
cycle costs corresponding to several sets of operating
conditions have been calculated. We have also observed
that the waste streams from the protactinium isolation
and the rare-earth removal portions of the new flow-
sheet can be conveniently combined for uranium
recovery prior to disposal, thus decreasing the proba-
bility of loss of fissile material and eliminating the need
for adding “Li to the protactinium decay tank in order
to obtain an acceptably low liquidus temperature. A
method for combining the waste streams is described.

Oxide precipitation is being considered as an alterna-
tive method for selectively removing protactinium from
MSBR fuel salt and for subsequently removing uranium
from the fuel salt prior to the removal of rare earths.
Two conceptual flowsheets based on oxide precipi-
tation are presented, and the importance of several
system parameters is shown.

17.1 PROTACTINIUM ISOLATION USING
FLUORINATION AND REDUCTIVE

EXTRACTION
M. J. Bell E. L. Nicholson
W. L. Carter L. E. McNeese

Analysis of the fluorination—reductive-extraction
flowsheet for isolating protactinium from MSBR fuel
salt has revealed several undesirable features and has
suggested an improved method for removing fission
product zirconium and for retaining *?°Pa during its
decay to *23U. In the flowsheet described previously,
zirconium was extracted into the bismuth stream
exiting from the lower column of the protactinium
isolation system and was removed from this stream by
hydrofluorinating a small fraction of the bismuth in the

1. L. E. McNeese, MSR Program Semiannu. Progr. Rep. Feb.
28, 1970, ORNL-4548, pp. 282—88.

2. M. 1. Bell and L. E. McNeese, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, pp. 199-202.

235

presence of salt that was withdrawn from the system.
Since the bismuth stream also contained protactinium
and uranium, the portion of the stream that was
hydrofluorinated represented a compromise between
(1) maintaining an acceptably low zirconium concen-
tration in the bismuth in the lower part of the column
and (2) transferring acceptably small amounts of
protactinium and uranium to the waste salt, from which
these materials must be recovered. The remaining
bismuth was hydrofluorinated in the presence of salt
that was recycled to a point ahead of the fluorinator in
order to remove the uranium as UF4. This operation
also resulted in the recycle of zirconium, which, under
operating conditions of interest, was oxidized and
reduced several times before its removal. Such recycling
resulted in a significant increase in the quantity of
reductant required for isolating the protactinium.

We have observed that the flowsheet can be simplified
by hydrofluorinating the entire bismuth stream in the
presence of a secondary salt stream, as shown in Fig.
17.1. Salt is withdrawn from the reactor on a ten-day
cycle and is fed to a fluorinator, where about 99% of
the uranium is recovered. The salt is then fed to an
extraction column where protactinium, zirconium, and
the remaining uranium are extracted into a bismuth
stream containing reductant. The bismuth stream is
then hydrofluorinated in the presence 6f a second salt
stream, which -results in transfer of the extracted
materials to the salt. Reductant is added to the
recovered bismuth, and the resulting metal stream is
recycled to the extraction column as in the previous
flowsheet. The secondary salt stream is circulated
through a hydrofluorinator, a fluorinator, and a protac-
tinium decay tank. The fluorinator is used to maintain
an acceptably low uranium concentration in the protac-
tinium decay tank. Salt is withdrawn from the decay
tank periodically to remove zirconium and other fission
products that accumulate in the tank. The salt is held
for a sufficient period before final discard to allow
233Pa to decay to 233U, which is recovered from the
salt by batch fluorination.

- These flowsheet modifications offer the following
advantages over the earlier flowsheet:

1. The bismuth inventory in the system is greatly
reduced, thus avoiding a significant inventory
charge.

2. The protactinium decay tank can be fabricated from
a nickel-base alloy rather than molybdenum, which
236

L,- ORNL-DWG T7O-2811B
SALT UFg~ PROCESSED SALT

r—
PURIFICATION REDUCTION )
~ f ‘/
Hz

SALT CONTAINING RARE EARTHS

]
|

=

—— et o o —— —— —

f ’ EXTRACTOR
\

R UFG N '
| REACTOR l COLLECTION

1
|
|

- . . _}
EXTRACTOR '

LiCl r—— Bi- '
(0 5 mole fraction L|)

EXTRACTOR
UFg

J'I

HYDRO- | |FLUORINATOR] | o DECAY ] L =+ DIVALENT RARE

FLUORINATOR

—— et e - - —— — —— —— i e i T e — —— ————— —

I . 1 B N EARTHS
' ! 1 ) , bam— i
I HF F ’ l (0.05 mole fraction Li)
! . Pa DECAY | _ . I
B'I| 2~ FLUORINATOR 6 | cxTRACTOR :
}
' |
| | Bi-
I SALT ot TRIVALENT RARE
: To L EARTHS
Lo REDUCTANT | _ WASTE = ™—— ™~ - 3
ADDITION

t

Li

Fig. 17.1. Improved flowsheet for processing a single-fluid MSBR by fluorination—reductive extraction and the metal transfer
process. The isolated protactinium is held for decay in a secondary salt stream,

will result in a considerable saving in the installed 7. Very efficient hydrofluorination of the bismuth
equipment cost. stream would permit the initial salt inventory in the
protactinium decay tank to contain natural lithium

3. Control of the protactinium isolation system is iy
rather than ‘Li.

greatly simplified.

4. Zirconium will not be recycled in the lower part of Representative operating conditions and partial fuel
the protactinium isolation system, thus reducing the cycle costs have been determined for the revised
consumption of reductant. protactinium isolation system. Both the quantity of

reductant required and the rate at which fuel carrier salt
must be removed from the reactor to compensate for
the LiF added by the protactinium isolation system
depend on the fraction of the uranium that is removed
_ from the fuel salt by the primary fluorinator. For a
6. The flowsheet is simplified; recycle of fuel salt uranium removal efficiency of 95%, the lithium re-
containing uranium, zirconium, and protactinium to ductant requirement is 370 moles/day and fuel carrier
the primary fluorinator is avoided. Operation of the salt must be withdrawn at the rate of 0.3 ft*/day. For a
secondary salt circuit is restricted only by heat uranium removal efficiency of 99%, the reductant
removal and inventory considerations. requirement is 200 moles/day and salt must be dis-

5. The isolated protactinium is retained in such a
manner that maloperation of the extraction column
cannot return large quantities of protactinium to the
reactor.
carded at the rate of 0.16 ft®/day. In the former
instance, the chemical and inventory charges for the
fluorination—reductive-extraction system amount to
0.047 mill/kWhr. These charges are reduced to 0.034
mill/kWhr for the higher fluorinator efficiency. In each
case, the salt inventory in the protactinium decay tank
is 150 ft>. The uranium inventory in the tank is about
0.1% of the reactor inventory.

17.2 COMBINATION OF DISCARD STREAMS
FROM THE PROTACTINIUM ISOLATION
SYSTEM AND THE METAL TRANSFER SYSTEM

M. J. Bell L. E. McNeese

The flowsheet shown in Fig. 17.1 requires that about
40 moles of LiF be added daily to the protactinium
decay tank in order to obtain a suitable liquidus
temperature. Lithium fluoride purchased for this addi-
tion would increase the power cost by 0.0014
mill/kWhr; however, we have observed that an accept-
able liquidus temperature can also be obtained by
hydrofluorinating the lithium-bismuth stream from the
divalent rare-earth stripper in the presence of the salt
from the decay tank. This operation adds 50 moles of
LiF and about 1.1 moles of rare-earth fluorides to the
decay tank per day. With this addition, the composition
of the salt in the decay tank is 72-24-3 mole %
LiF-ThF4-ZrF,, 1 mole % divalent rare-earth fluorides,
and 360 ppm trivalent rare-earth fluorides. The salt has
a liquidus temperature of 570°C; at this temperature,
the rare-earth fluoride concentration is well within the
rare-earth solubility.

We have also observed that it is possible to combine
all waste streams from the metal transfer system and
the protactinium isolation system for uranium recovery
prior to disposal as shown in Fig. 17.2. In this
operation, waste salt from the protactinium decay tank
would be combined with the fuel carrier salt discard
stream. The lithium-bismuth stream from the trivalent
rare-carth stripper would be hydrofluorinated in the
presence of the resulting salt, and the combined stream
would be held for protactinium decay. The protac-
tinium concentration in the combined streams would be
only 500 ppm initially, and the specific heat generation
rate would be acceptably low. The salt in the waste
holdup tank would be fluorinated before discard to
recover uranium. The composition of the discarded salt
would be 74.7-13.5-9.50.8 mole % LiF-ThF,-BeF,-
ZrF4, 1.2 mole % trivalent rare-earth fluorides, and 0.3
mole % divalent rare-earth fluorides. Although the
liquidus temperature of the salt is near 500°C, the salt
temperature would have to be maintained at about

237

600°C so that the trivalent rare-earth fluorides would
not precipitate. This processing scheme would require
that salt be discarded at the rate of 60 ft* every 220
days.

17.3 PROTACTINIUM ISOLATION USING
OXIDE PRECIPITATION

M. J. Bell L. E. McNeese

Ross, Bamberger, and Baes® have shown that protac-
tinium can be precipitated selectively as Pa, O5 from
MSBR fuel salt by the addition of oxide to salt
containing Pa®*. Bamberger and Baes* have also found
that uranium oxide can be precipitated from fuel salt
from which protactinium has been previously removed,
with the attendant precipitation of only a small fraction
of the thorium. Further, Bell and McNeese® have
calculated that greater than 99% of the uranium can be
recovered from fuel salt by countercurrent multistage
contact of fuel salt with UO,-ThQ, solid solutions,
using relatively few stages. The composition of the
oxide product was calculated to be greater than 90%
U0,.

Hence, oxide precipitation is being considered as an
alternative method to fluorination and reductive ex-
traction for the isolation of protactinium and removal
of uranium from the fuel salt of an MSBR. A possible
flowsheet and typical operating parameters are shown
in Fig. 17.3. Fuel salt is withdrawn from the reactor on
a three-day cycle, and protactinium is removed by
precipitation as Pa, Os. Part of the salt is processed on a
30-day cycle for rare-earth removal by the metal
transfer process. Most of the uranium must be removed
from this stream prior to removal of the rare earths.
The separated uranium is then recombined with the
processed salt leaving the metal transfer system and is
returned to the reactor. The Pa,Q; precipitate is
hydrofluorinated in the presence of a captive salt phase,
which circulates through the protactinium decay tank
and through a fluorinator in order to maintain an
acceptably low uranium inventory in the decay tank.

- Part of the salt in the decay tank must be returned to

the reactor to compensate for salt that is transferred to
the tank with the Pa, Q5 precipitate.

3. R. G. Ross, C. E. Bamberger, and C. F. Baes, Jr., MSR
Program Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622,
pp- 92-95.

4. C. E. Bamberger and C. F. Baes, Jr., J. Nucl Mater. 35,
177 (1970).

5. M. 1. Bell and L. E. McNeese, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, pp. 202-8.
238

ORNL-DWG 7i- 2858A
Li-Bi + DIVALENT

RARE EARTHS

50 moles Li/doy UFe
1 f 1
Bi CONTAINING HYDRO- Pa
Pa,U,Li,Th,Zr —m = FLUORINATOR » DECAY
.FLUORINATOR
FROM Pa ISOLATION . TANK Li—Bi+ TRIVALENT
l T 1 RARE EARTHS
410 50 moles Li/day
Bi TO HF-Hz Fa moles salt/day
RECYCLE 0.089 H’3/d0y
| |

FUEL CARRIER

SALT DISCARD WASTE SALT HYDRO-
240 moles /day HOLDUP TANK 1 FLUORINATOR
0.16 f13/day
Bi TO
RECYCLE
UFg
BATCH SALT
FLUORINATOR - TO WASTE
0.27 ft¥day
¥
Fa

Fig. 17.2. Method for combining waste streams from protactinium isolation and rare earth removal processes. Flow rates are
shown for an assumed uranium removal efficiency in the primary fluorinator of 99%.

ORNL-DWG TO-14086A

k RECOMBINER 2
[
u
8.3x109 2.75
moles/day URANIUM | moles/day U METAL RARE
REMOVAL ! TRANSFER —# EARTHS
275 99% SYSTEM Sr, Ba, Zr, U
moles /day U

ACTOl

RE R
2250 MW (th)
1680 13 UFg

= -6
Xpqr, = 8-4%10

8.3x10°
moles/day

Pa DECAY TANK
395 moles PaFs AND
10.0 moles UF,
0.19x10° moles
X e, =56%1078

4

PREClFP)IGTATOR 5.2 moles/day OXIDE

60%  [7200 moles/day SALT

Xqu4=20x10‘5 EFFICIENCY 0.66 moles/doy UF,

X, =0.0033
4

Y

HYDRO-
FLUORINATOR

FLUQRINATOR

Fa

200 moles/day
0.4 moles/day PaFy,

0.19x10% moles/doy §

Fig. 17.3. Flowsheet for protactinium isolation using fluorination and oxide precipitation. Fuel salt is withdrawn from the
reactor at a rate corresponding to a three-day cycle.
As shown in Fig. 17.4, the protactinium removal time
for this system depends on the precipitator efficiency
and the rate at which fuel salt is transferred to the
protactinium decay tank along with the Pa,05. A
removal time of about five days can be obtained if 60%
of the protactinium is removed from the salt in the
precipitator, provided the salt transfer rate to the
protactinium decay tank is as low as 10 to 20
moles/day (a salt-to-oxide flow rate ratio of 2 to 4). For
the same removal time, a precipitator efficiency of
about 80% would be required if the salt-to-oxide flow
rate ratio were as high as 600. The uranium inventory in
the decay tank depends on the efficiency of the

ORNL-DWG 70-14098A

15 I | I
RATIO OF SALT -
FLOW RATE TO
E OXIDE FLOW RATE
o {2
T \ \600
=
=L\
2 ° NN N
g \\ 400
5 \
Ll \ N
200
NN
Pa CYCLE TIME: 3 days \\\
FLUORINATOR EFFICIENCY: 95% \
20 30 40 50 60 70 80 90 100

Pa PRECIPITATCR EFFICIENCY (%)

Fig. 17.4. Effects of precipitator efficiency and fuel salt
transfer rate on protactiniom removal time for the fluorina-
tion—oxide precipitation flowsheet, ‘

239

fluorinator in the protactinium isolation loop and on
the amount of fuel salt transferred to the protactinium
system with the precipitate, as shown in Fig. 17.5. This
uranium inventory will be only a small fraction of the
uranium inventory in the reactor, and the associated
inventory charge will be less than 0.001 mill/kWhr for a
wide range of operating conditions.

Isolation of protactinium by oxide precipitation can
also be achieved without the use of a fluorinator, as
shown in Fig. 17.6. In this flowsheet, the Pa,0Os
precipitate is hydrofluorinated in the presence of
uranium-free salt leaving the metal transfer system. The
resulting salt stream then passes through a decay tank,

ORNL-DWG 70-14099A

[ i 0.0008
X .26 |-90% PRECIPITATOR EFFICIENCY /
Z 1 DAY SALT RESIDENCE TIME
fi% Xpag, = 0.0022 -
W6 0.24 ' REACTOR FISSILE INVENTORY = | 00007 £
o 8 1460 kg =
0= 10 % INVENTORY CHARGE =
Zw . £
£ 0.22 7 £
&e } w
= &
Z5 00 FLUORINATOR / 0.0008 4
wgo EFFICIENCY v 5
£ e .
56 90 % x
(@]
wa 0.18 ] o =
s / ~ / =
< 95 % w
2 L L [ — 00005 2
2 =z
016 '
o) 100 200 300 400 500 600

RATIO OF SALT TRANSFER RATE TO OXIDE FLOW RATE
{moles salt /male axide)

Fig. 17.5. Effects of secondary fluorinator efficiency and fuel
salt transfer rate on uranium inventory for the fluorination—
oxide precipitation flowsheet,

ORNL-DWG 70-1{40874A

[
56 f+/doy
RECONSTITUTION
273 moles/day U 3
5500 moles/day SALT (4 £1°)
2.75
REACTOR 3 URANIUM METAL
2250 MW (th) 80 1700y _ | cEMovaL moles/day U_| - ANSFER | R
1680 ft3 999, EFFY SYSTEM Sr,Ba, Zr,U

i Xpop, = 8.4X10°
Xpgr,=20%10

a 560 ft3 day
Xyp =0.0033 /
a4

6

—

Pay05
| PRECIFITATOR |20 moles/doy FL R oR - ?SE(O:??Y TANK
) . FUEL SALT UGRI U INVE Y = 114
96 BEFFY | , FU motesday ORY =11.4 kg
X 250X10°6
3 Ho—HF uFy
60 ftdoy XpqF, =0.0020

Fig. 17.6. Flowsheet for protactinium isolation using oxide precipitation.
where it is held up two to four days before being
recycled to a point ahead of the precipitator in order to
return 233U to the reactor. Since the salt flow rate
through the protactinium decay tank must be relatively
large in order to limit the uranium inventory in the
decay tank, high precipitator efficiencies are required; a
precipitator efficiency of about 96% would be required
to obtain a protactinium removal time of five days.
Uranium inventories in the protactinium isolation sys-
tem using this flowsheet will be 0.5 to 1.0% of the
reactor fissile inventory; corresponding inventory
charges will be 0.002 to 0.004 mill/kWhr.

17.4 STRIPPING OF RARE-EARTH FISSION
PRODUCTS FROM LiCl IN THE METAL
TRANSFER SYSTEM

M. J. Bell

The flowsheet described previously':* for removing
rare-earth fission products from MSBR fuel salt using
the metal transfer process employs contact of the LiCl
with lithium-bismuth solutions for removal of the rare
earths and other fission products. In the reference
flowsheet, the trivalent rare earths are removed by
contacting the LiCl, at the rate of 33.4 gpm, with an
8.1-gpm recirculating bismuth stream having a lithium
concentration of 5 at. % in a single equilibrium stage.
Bismuth containing extracted rare earths is withdrawn
at the rate of 5.7 gal/day, and an equal amount of
lithium-bismuth solution is added.

Early data indicated that mutual solubility problems
might be encountered between thorium and trivalent
rare earths in bismuth having a lithium concentration as
high as 5 at. %.° Although this has been found not to
be the case, we have obtained data on the effect that
varying the lithium concentration in the lithium-
bismuth alloy has on the thorium concentration in the
metal and on reactor performance. These results are
presented in Fig. 17.7, which shows the effect of
increasing the flow rate of the lithium-bismuth with-
drawal stream while -holding constant the amount of
reductant fed to the system. The thorium concentration
in the metal is reduced from 420 ppm at the reference
withdrawal rate of 1000 moles/day to 140 ppm at the
discard rate of 3000 moles/day. The effect on reactor
performance is slight; the breeding ratio decreases from
1.063 to 1.060. It would be possible to compensate, in
part, for this loss in breeding ratio by operating the
trivalent-rare-earth stripper as a once-through batch

6. F. J. Smith and C. T. Thompson, ibid., pp. 207-8.

240

ORNL-DWG 71-2859A
mole fraction Li IN METAL

0.05 0.033 0025 0.020 0.0167

& 400 0063
- -
22 300 S, zo

2 N 0062 Z2
= 7 -—— \—-—b I
= \ r
5 200 ] 0061 22
g e N o
< 3 100 —50 moles/day Li METAL < 0060
»5 IN Bi s
E7 olesc | | | 0059

0 1000 2000 3000 4000

Li-Bi FLOW RATE {moles/day)

Fig. 17.7. Effect of lithium-bismuth discard rate on thorium
concentration in discard stream and on MSBR performance.

contactor and increasing the number of stages. This
mode of operation has also been considered as a means
for accommodating the high -specific heat generation
rates of the trivalent rare-earth fission products.

In the reference flowsheet, the divalent rare-earth
fission products are removed from the LiCl by con-
tacting 2% of the LiCl leaving the trivalent stripper with
0.56 gal of 50 at. % Li-Bi per day. In order to deal with
the high heat generation rates expected in the resulting
lithium-bismuth stream, we have considered operating
this contactor as a continuous column with a high
recycle rate of the metal stream. A decay tank could
then be placed in the recycle stream to dissipate a large
fraction of the fission product decay heat.

17.5 IMPORTANCE OF URANIUM INVENTORY
IN AN MSBR PROCESSING PLANT

M. J. Bell L. E. McNeese

The MSBR processing flowsheets considered thus far
have uniformly resulted in very low uranium inventories
in the processing plant, that is, inventories that are
usually below 1% of the fissile inventory in the reactor.
Since several potential processing systems might result
in uranium inventories as large as 5 to 10% of the
reactor inventory, we have investigated the importance
of uranium inventory in an MSBR processing plant. The
major effects of an increased uranium inventory are: (1)
an increase in inventory charges on fissile material and
(2) an increase in the reactor doubling time. The system
fissile inventory (which includes the ?32Pa in the
processing plant) was assumed to be 1504 kg, the value
of 2?3 U was taken.to be $14 per gram, and the capital
charge rate was assumed to be 10%jfyear. The calculated
system doubling time for the limiting case of a zero
uranium inventory in the processing plant was 22 years.
A processing plant uranium inventory of 5% of the
system fissile inventory would increase the fuel cycle
cost by 0.015 mill/kWhr and would increase the system
doubling time from 22 to 23.1 years. A uranium
inventory of 10% of the system fissile inventory would
result in a fuel cycle cost increase of 0.03 mill/kWhr

241

and an increase in doubling time from 22 to 24.2 years.
Thus, while there is incentive for maintaining a low
uranium inventory in the processing plant, it does not
appear that a uranium inventory as high as 5 to 10% of
the system fissile inventory would rule out an otherwise

attractive processing system.
18. Processing Chemistry

L. M. Ferris

Studies of the chemistry relating to the metal transfer
process' > for the removal of rare earths from MSBR
fuel salt were continued during this reporting period.
The equilibrium distribution of several actinide ele-
ments between molten LiCl and liquid bismuth solu-
tions was determined, and additional data were ob-
tained on the effect of LiF concentration on the
distribution of several elements between LiCI-LiF solu-
tions and liquid bismuth. The mutual solubility of rare
earths and thorium in lithium-bismuth solutions was
also investigated further. In addition to the metal
transfer process studies, work was continued on a
method for selectively precipitating protactinium, as
Pa, O, from MSBR fuel salt.

18.1 MEASUREMENT OF DISTRIBUTION
COEFFICIENTS IN MOLTEN-SALT-METAL

SYSTEMS
F.J. Smith J. C. Mailen
C.T. Thompson J.F.Land

Distribution coefficients,

D.. = mole fraction of M in bismuth phase
M mole fraction of M in salt phase

(1)

were measured, using molten LiCl as the salt phase, for
several transuranium elements at 640 and 700°C and
for uranium at 675°C. In addition, data were obtained
for uranium, neodymium, and lanthanum at 640°C,
using LiCl-LiF salt solutions, to further elucidate the
effect of fluoride concentration on distribution be-
havior. These latter measurements should be useful in
predicting how the metal transfer process would oper-
ate if the LiCl acceptor salt were contaminated with

1. L. E. McNeese, MSR Program Semiannu. Progr. Rep. Feb.
28, 1970, ORNL-4548, p. 277.

2. D. E. Ferguson and Staff, Chem. Technol. Div. Annu.
Progr. Rep. May 31, 1970, ORNL4572,p. 1.

fluoride fuel salt. As we showed earlier,>™ the distri-
bution coefficients at a given temperature can be
expressed as

log Dy =nlog Ny ; +log Ky *, (2)
in which Ny, is the mole fraction of lithium in the
bismuth phase, 7 is the valence of M7 in the salt phase,
and log Ky, * is a constant.

Values of n and log K, * obtained for the respective
solutes using the various salt phases are given in Table
18.1. The values for uranium at 675°C and for the
transuranium elements at 640 and 700°C were obtained
from isotherms, represented by Eq. (2), using the
general technique described previously.®* Data for
uranium, lanthanum, and neodymium, using LiCI-LiF
solutions as the salt phase, were obtained in a separate
experiment. Initially, the chlorides of uranium and
lanthanum, along with '*7Nd tracer, were dissolved in
pure LiCl in the presence of a pure bismuth phase at
640°C. Lithium-bismuth alloy was then added to the
bismuth phase until the lithium concentration in this
phase was high enough to give readily measurable
distribution coefficients for each of the three solutes.
Several samples of each phase were taken, and values of
log K, * were calculated, using Eq. (2), from the
analyses of these samples. Weighed amounts of LiF
were then added incrementally to the system to
gradually increase the LiF concentration in the salt
phase. After each addition, at least 24 hr was allowed
for equilibration before several samples of each phase
were taken; analyses of these samples not only con-
firmed the LiF concentration but also were used in the
calculation of log Ky *. In these calculations, it was
assumed that » had a value of 3 for each of the solutes.

The effect of LiF concentration on log K * for
several elements is shown graphically in Fig, 18.1. The -

3. L. M. Ferris et al., J. Inorg. Nucl. Chem. 32,2019 (1970).

4. L. M, Ferris et al., MSR Program Semiannu. Progr. Rep.
Feb. 28, 1970, ORNL4548, p. 289.

5. L. M. Ferris et al., MSR Program Semiannu. Progr. Rep.
Aug. 31, 1970, ORNL4622, p. 204,

242
Table 18.1. Values of n and log K4 * obtained in measurements
of the equilibrium distribution of several solutes between
bismuth solutions and molten lithium halide salts

" Composition
of salt phase  Temperature
Element (mole %) ©0) log Kps*
LiCl LiF
La 100 0 640 3 7.796 £ 0.4
Nd 100 0 640 3 8.374 £ 0.4
U 100 0 640 3 11.197%04
Np 100 0 640 3 10.33 £0.1
Pu 100 0 640 3 10.126 0.1
Am 100 0 640 ~3 ~10.16 £0.2
Cm 100 0 640 3 9.406 £0.1
Cf 100 0 640 ~2 ~5.38%0.2
La 99.5 0.5 640 3 7.607 £0.4
Nd 995 0.5 640 3 8.568 0.4
U 99.5 0.5 640 3 10954 04
La 98.0 1.99 640 3 7.878 04
Nd 98.0 1.99 640 3 8.064 £0.4
U 98.0 1.99 640 3 11.329%04
La 974 2.65 640 3 7.789 £04
Nd 97.4 2.65 640 3 8.310X04
U 97.4 2.65 640 3 11624 t04
La 95.7 4.33 640 3 7372104
Nd 95.7 433 640 3 7.805 0.4
U 95.7 4.33 640 3 11480*04
La 91.2 8.82 640 3 7.182*04
Nd 91.2 8.82 640 3 7.521*04
U ' 91.2 8.82 640 3 10340%04
La 82.6 1741 640 3 6458 +0.4
Nd 826 1741 640 3 6.739 04
U 826 17.41 640 3 8.604 £0.4
La- 73.9 26.07 640 3 5494 £0.4
Nd 73.9 . 26.07 640 3 6.354 0.4
U 73.9 26.07 640 3 8.758%04
U 100 0 675 3 10.11 £0.3
Np 100 0 700 3 10369 0.1
Pu 100 0 700 3 10.223 £0.1
Cm 100 0 700 3 9.589 £0.1
Cf 100 0 700 2 543 10.2

values of log K, * for lanthanum, neodymium, and
uranium are those given in Table 18.1; the data for
thorium and europium were reported previously.*’ As
seen in Fig. 18.1, the values of log K, * for the di- and
trivalent species did not vary markedly when the LiF
concentration was in the range of 0 to about 4 mole %.
However, the value of Ky, * decreased by about two
orders of magnitude as the LiF concentration was
increased from 0 to 4 mole %. These data indicate that
contamination of the LiCl acceptor salt with fluoride
fuel salt in amounts that are equivalent to 4 mole % LiF

243

ORNL—DWG 71-2861A

16
A\
14 &
\\A\\\
12
at
0’0'9— \‘___ Th4+
[o] \ -‘——.—_.____A
10
‘\\\N\\s__ U3+
* e © S
o '.— =
6 \_ ® A
L03+ ®
4
Eu2+
2 P — —
0
0 0.04 0.08 0.42 0.6 0.20 0.24 028

LiF IN SALT PHASE {mole fraction)

Fig. 18.1. Variation of log K,,* with salt composmon Salt
phase was an LiCI-LiF solution; data obtained at 640 °C except
that for europium, which was obtained at 650°C.

or less will not affect the extent to which the rare
earths are removed in the metal transfer process. The
presence of fluoride in the acceptor salt would, how-
ever, cause a decrease in the thorium—rare-earth decon-
tamination factor.

- Some '°*Ru was present as an impurity in the
neodymium tracer in the experiment involving uranium,
lanthanum, and !'#7Nd. Distribution coefficients for
ruthenium could not be obtained, however, because the
ruthenium was completely extracted from the LiCl into
the bismuth phase before the lithium concentration in
the bismuth reached 5 wt ppm.- At this lithium
concentration, the ruthenium distribution coefficient
was greater than 1000,

In the last semiannual report,® we presented values of
log Kp,* that were obtained in an experiment in-
volving several LiCl-LiF solutions as the salt phase, and
also estimated the neodymium-promethium separation
factor (Dyg/Dpy) to be 3 + 2. Inspection of the
gamma spectra of samples taken in this experiment has
revealed that a significant amount of '*7Nd was
present along with the promethium. Analysis of these
spectra yielded the neodymium-promethium separation
factors at 640°C given in Table 18.2. Assuming that the
average of these values (8.3 + 1.8) is also valid when
LiCl is the salt phase, we estimate log Kp,* = 7.45 %
Table 18.2. Neodymium-promethium separation
factors obtained at 640°C with LiCI-LiF
solutions as the salt phase

LiF concentration Nd-Pm
Sample in salt separation
(mole %) * factor
0 8.32+1.82
2.45 6.27P

1 3.22 10.1
2 3.22 8.4
3 3.22 8.5
4 6.59 6.1
5 6.59 124
6 6.59 10.9
7 10.22 7.1
8 10.22 6.8
9 10.22 6.8
10 17.56 7.4
11 17.56 8.5
12 17.56 8.9

dEstimated; see text.

bCalculated from values of log Kng™ and log Kp * derived
from the respective isotherms.

0.4. This estimate indicates that promethium would
behave much like lanthanum in the metal transfer
process. ‘

18.2 SOLUBILITIES OF THORIUM AND
NEODYMIUM IN LITHIUM-BISMUTH SOLUTIONS

F.J.Smith  C.T. Thompson
J.F. Land

In the metal transfer process'*? being developed for
the removal of rare earths from MSBR fuel salt,
lithium-bismuth solutions having high lithium concen-
trations (5 to 50 at. %) would be used-to strip the rare
earths (along with any thorium present) from the LiCl
acceptor salt. Therefore, we are determining the solu-
bilities, both individual and mutual, of rare earths and
thorium in lithium-bismuth solutions.

Solubilities of thorium and neodymium in several
lithium-bismuth solutions were measured using the
procedure and apparatus described elsewhere.® The
results are summarized in Tables 18.3 and 18.4 and are
shown graphically in Figs. 18.2 and 18.3. At each
temperature below 750°C, the solubility of thorium in
lithium-bismuth solutions increased by about a factor
of 3 as the lithium concentration in the solution
increased from O to about 40 at. %. Similarly, the

6. C. E. Schilling and L. M. Ferris, J. Less-Common Metals
20,155 (1970).

244

Table 18.3. Solubility of thorium in liquid
lithium-bismuth solutions

Thorium
Li concentration in Temperature solubility, log
solution range Sp (Wt ppm) =

(at. %) ©C) A+ B/T (°K)

A B
0 350800 7.708 —3852
54 350-800 7.073 -3205
8.0 350-800 7.073 -3205
19.3 400-750 7.133 —-3164
40.4 450-750 7473 -3333

Reference 6.

Table 18.4. Solubility of neodymium in liquid
lithium-bismuth solutions

Neodymium
Li concentration in Temperature solubility, log
solution range Sng (Wt pp;n) =
(at. %) &) A +B/T (°K)
A B
04 500-700 6.68 -2200
8.32 400-700 7.19 —-2531
17.4 450-700 7.11 —2435
25.0 450-700 7.07 -2379
38° 475-700 6.26 -1809

aPreliminary data.
bReference S.

solubility of neodymium increased regularly as the
lithium concentration was increased from O to 25 at %.
However, when the lithium concentration was increased
from 25 to 38 at. % at 640°C, the solubility of
neodymium decreased to about the same value that was
obtained when pure bismuth was used as the solvent
(Fig. 18.3). We are redetermining the neodymium
solubilities in lithium-bismuth (40-60 at. %) to confirm
this behavior.

In a series of experiments, we measured the mutual
solubilities of thorium and neodymium at 640°C in

- lithium-bismuth solutions having lithium concentrations

ranging from O to 20 at. %. The mutual solubility limit
was approached by making incremental additions of
neodymium to the thorium-saturated liquid lithium-
bismuth solution. At the end of some experiments
when the system had been saturated with neodymium,
several small additions of thorium were made. The data
obtained in this series of experiments were somewhat
scattered, due primarily to the difficulty in analyzing
ORNL— DWG 71— 2883A
TEMPERATURE (°C)

o' 800 700 600 500 450 400
ST T T =
- LITHIUM CONCENTRATION _|
L IN SOLUTION —]
— (at. %) —
| -——- 0 ]
° 5.4
10° §\_ ta, o 8.0 |
= NGO A 19.3 =
< — \3\;\ . A 40.4 —]
- [ \ A : ]
5 [
: C \\\z\{ _
> \ h
[ | — —
5 N
S 0! \3\\ \\
J — ' p—
? — N ]
= — \s \ X —
o — p—
z — AN \ ]
Q AN
= u \ (»] —]
1072 AN AN
— N
[ N
— \ =
[ N
10 > L .
9 50 1" 12 13 14 15 16
10,000/7- (°K)

_Fig. 18.2. Thorium solubility in lithium-bismuth solutions.
Data for O at, % lithium taken from ref. 6.

for thorium at a relatively low concentration (less than
0.35 wt %) in the presence of neodymium at a high
concentration (up to about 2 wt %). However, the
following general conclusions can be drawn from the
data:

1. Neodymium, in concentrations below about 1 wt %,
appears to have no effect on the thorium solubility
(about 0.35 wt %) in lithium-bismuth solutions at
640°C; as the neodymium concéntration was in-
creased from 1 to about 2 wt % (saturation), the
thorium concentration decreased almost linearly
from about 0.35 to about 0.1 wt %. I

2. The effect appeared to be reversible. Additions of
thorium to lithium-bismuth solutions saturated with
neodymium caused a decrease in the neodymium
concentration to about 1 wt % as the thorium
concentration increased to about 0.35 wt %.

The above results indicate that the mutual solubility
of thorium and rare earths in the solution, lithium-
bismuth (5-95 at. %), proposed’*? for stripping of the
trivalent rare earths (La, Pm, Nd) from the LiCl
acceptor salt is much higher than required. Measure-

245

ORNL-DWG Ti-2872A
TEMPERATURE (°C)
o 700 650 BOO 550 500 450 400 350 300

I I I I I I I
T L) T T ¥ ] L

|
T

%
v

NEODYMIUM SOLUBILITY (wt %)
o
// /z/
/

1o~ LITHIUM N I\

. CONCENTRATION A
IN SOLUTION
(at, %) th,

[T TTIIT

10 M 12 13 13 15 16 17 18
10, 000/7 (o)

Fig. 18.3. Solubility of neodymium in lithium-bismuth solu-
tions. The solubility of neodymium in lithium-bismuth (38-62
at. %) was taken from ref. 5. The solubility of neodymium in
bismuth (lithium = 0 at. %) should be considered preliminary.

ments of the mutual solubility of thorium and a
divalent rare earth (europium) in liquid lithium-bismuth
(40-60 at. %) are in progress. '

18.3 OXIDE PRECIPITATION STUDIES
| " J.C. Mailen |

We are investigating oxide - precipitation methods
(instead of fluorination and reductive extraction) for
the isolation of both uranium and protactinium from
MSBR fuel salt. It has been shown” that Pa*" dissolved
in a molten LiF-BeF,-ThF, solution can be oxidized to
the 5+ state either by hydrofluorination or by reaction
with NiO. Furthermore, the addition of oxide to a salt
solution containing Pa®" results’ in the precipitation of
pure, or nearly pure, Pa, Os. The solubility product for
Pa, Qs appears to be much lower than that for UO,;
hence Pa,0s; can be precipitated in preference to
U0, .>*" Measurements of the solubility of Pa, 05 in
MSBR fuel salt that is saturated with UO, are being
made at various temperatures. Accurate measurements
are extremely important in that the Pa, O5 solubility

7. R. G. Ross, C. E. Bamberger, and C. F. Baes, Jr., MSR
Program Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p.
92.
246

Table 18.5. Precipitation of Pay Qg from LiF-BeF,-ThF4UF, (71.8-16-12-0.2 mole %)
Initial weight of salt: 100 g

Pa concentration Total UO,
Temperature s 6 3 added to log @,
Sample CC) in salt 10°NpyFs 10 NUF4 system (footnote a)
(wt ppm) (mg)
1 600 111 304 2.68 0
2 600 110 30.1 10.8
3 600 100 27.5 10.8
4 600 66.9 18.3 2.37 22.8
5 600 14.2 3.87 31.6
.6 600 2.5 0.69 2.39 41.6 -2.91
7 600 2.7 0.73 . 51.0
8 655 114 3.12 2.33 51.0 -2.24
9 600 118 32.3 2.49 0 ‘
10 551 4.8 1.32 2.21 113 -2.59
11 577 9.7 2.65 2.10 113 -2.26
12 605 3.1 0.86 2.08 113 -2.74
13 632 7.4 2.03 2.13 113 —-2.38
14 659 7.8 2.12 2.05 113 -2.34
- u5/4
a NPaF SNU02 .
='—““—s‘/‘4“_;NU02 was about 0.95 in each case.
N UF,
b System was extensively hydrofluorinated after sample 8 was withdrawn.
defines the lowest protactinium concentration attain- for which
able at a given temperature without changing the
uranium and thorium concentrations in the salt. Solu- Np,g < (d)
bilities of Pa, 05 at UQ, saturation of the salt can be 0, = ~
obtained in two ways. One method consists in the N, ThF4(d)
direct measurement of the equilibrium quotient (Q) for
the reaction and
% UF4(d) + %4 Pay05(c) =PaFs(d) + % UO0,(ss), (1) _Nrwr @) Nuo,ss)

for which

5/4
0 Npap ) NS, s0)
=

5/4
MR, @)

if Pa,O5 is present as a pure solid phase. In these
expressions, N, d, ¢, and ss denote mole fraction,
dissolved species, pure crystalline solid, and ThO,-UO,
solid solution respectively. The other method involves
combining the equilibrium quotients for the following
two equilibria:

%ThF,(d) + Y, Pa, Os(c) =PaFs(d) + % ThO,(c), (2)

UF,4(d) + ThO,(ss) = UO,(ss) + ThF,(d) , (3)

3 = .
NyF, ) VTho, (ss)

Previously,® we reported preliminary data from an
experiment in which Pa, 05 was precipitated from
LiF-BeF, -ThF,-***UF, (71.8-16-12-0.22 mole %) that
initially contained about 110 wt ppm of 23'Pa tracer.
Complete data from this experiment, along with the
derived values of log Q,, are given in Table 18.5. In the
first part of this experiment, small amounts of UO,
were added incrementally at 600°C to about 100 g of
salt that had been extensively hydrofluorinated to
ensure that the protactinium was in the 5+ oxidation
state. The concentration of protactinium in the salt
decreased regularly with each addition of UQ, until it
reached a steady value of about 2.5 wt ppm (Table
18.5, samples 1 —7). (The protactinium concentration in
the salt should become practically constant when the
Table 18.6. Values of log @, calculated at NUF4 = (.0022 from measured values of log 0, and log Q5

'Pa® concentration at ThO,

Temperature saturation of salt 6 log 03 log Q)
°0) 3 1070, log Q2 (footnote g) (footnote b)
wt ppm 10 NPan
654 1.6 04376 6.20 -5.207 - 2.812 © ~1.692
678 14 3.829 542 —4.265 . 2.730 - -0.852
738 62 16.96 240 -3.619 2.536 —0.449

404 evaluated f;lt NUF4 =0.0022.
Plog Q1 = % log Q3 +log 0.

oxide concentration in the salt becomes high enough to
cause precipitation of a UQ,-containing solid.) The
temperature of the system was then increased to 655°C,
and sample 8 was withdrawn. After this, the system was
again extensively hydrofluorinated, and about 113 mg
of UO, was added. The temperature of the system was
varied between 550 and 660°C, and filtered samples of
the salt were taken at each temperature (Table 18.5,
samples 10—14). A period of at least 24 hr was allowed
for the attainment of equilibrium at each temperature.

The 233U concentrations in the salt samples were
determined by the alpha-pulse-height method; the
“protactinium concentrations were determined by count-
ing of the gamma rays emitted by the 23?Pa and by the
alpha-pulse-height method for ?3!Pa. The latter two
analyses generally were in good agreement. In calcu-
lating values of log @, from these analyses, it was
assumed that: (1) the system was at equilibrium in each
case; (2) the protactinium in the salt was in the 5+
oxidation state; and (3) the solid phase was a mixture
of pure Pa,0Os and a UO,-ThO; solid solution whose
composition could be calculated from the uranium
concentration in the salt and the expression

Nyo,(ss) VThF 4 (d)
Nyr, @) VTho, ss)
2101 + 550 Nyo, (ss)
- T

log Q3 = log

(4)

reported by Bamberger and Baes.® This calculation
showed that NUOQ(ss) varied only from 0.93 to 0.97,
the average value was 0.95. The average UF, concentra-
tion in the salt was 0.22 mole %. We strongly suspect,
but cannot prove, that equilibrium was not achieved at

8. C. E. Bamberger and C. F. Baes, Jr., J. Nucl. Mater. 35,
177 (1970).

551 and 577°C and, therefore, that our values of log Q,
at these temperatures (Table 18.5) are much too high.

We also measured the solubility of Pa,0s in LiF-
BeF,-ThF, (72-16-12 mole %) in the presence of excess
ThO, and NiO using the general technique just de-
scribed. These solubilities and the values of log (),
derived from them are summarized in Table 18.6. Our
values for the solubility of Pa, Os are higher than those
of Ross, Bamberger, and Baes’ at temperatures above
about 675°C; however, since our values appear to be
more dependent on temperature, they indicate lower
solubilities at the lower temperatures.

Table 18.6 also gives values of log Q, that were
obtained from the measured values of log 2, and values
of log Q3 that were calculated from Eq. (4) using

ORNL-DWG 71-2862A
TEMPERATURE (°C)

750 700 650 600 550
0
T T T T 1
o NO URANIUM IN SALT
o ® Ny = 0.0022 ]
xS i
n = uw _1
u” I"’\"‘Z:)
£
=z
—
on
o
- O
N
¢ -2
g .
L ° ¢
L
L
®
-3
95 100 105 #O H5 120 125
10,000/ oy

Fig. 18.4. Values of the equilibrium quotient for the reaction
1iPa, 05(c) + Y, UF4(d) = PaF5(d) + %3 UO, (ss) when the UF,
concentration in LiF-BeF,-ThF,; (72-16-12 mole %) is 0.22
mole %.
NUF4(d) =0.0022. This value ofNUI:4 was the average
UF, concentration in the experiment discussed above
in which we obtained direct measurements of log Q, .

A plot of all our values of log @, vs 1/T (°K) is shown
in Fig. 18.4. A least-squares fit yielded log O, =8.849
— 9859/T (°K), with a standard deviation of +0.47.

This expression gives values of log Q, that are signifi-.

248

cantly different than those estimated by Ross et al.,”
particularly at temperatures below about 650°C. In
addition, the uncertainty in the values of this important
equilibrium quotient is much larger than that required
for a detailed analysis of the oxide precipitation process
for the isolation of protactinium from MSBR fuel salt.
Work on this problem is continuing.

19. Engineering De'velopment of Processing Operations

L. E. M¢cNeese

19.1 ENGINEERING STUDIES OF THE METAL
TRANSFER PROCESS FOR RARE-EARTH
REMOVAL

E. L. Youngblood L. E. McNeese

The second engineering experiment (MTE-2) for the
study of the metal transfer process for removing rare
earths from single-fluid MSBR fuel salt has been
completed. The main objectives of the experiment
were: (1) demonstration of the selective removal of rare
earths from fluoride salt containing thorium fluoride,
(2) collection of the rare earths in a lithium-bismuth
solution, and (3) verification of previous distribution
coefficient data. All of these objectives were accom-
plished. The experiment was performed in the 6-in.-
diam carbon-steel vessel shown schematically in Fig.
19.1. The vessel had two compartments that were
interconnected at the bottom by a pool of bismuth
saturated with thorium. One compartment contained
fluoride salt (72-16-12 mole % LiF-BeF,-ThF,) to
which 7 mCi of '*7Nd and sufficient LaF; to produce
 a concentration of 0.3 mole % had been added. The
other compartment contained LiCl, as well as a cup

ORNL-DWG 70-12503RA

~— LEVEL
ELECTRODES
ARGON INLET 5 |
AND VENT —=o{ [T

{ —]
| E— ——l

| o——— CARBON-STEEL
PUMP WITH MOLTEN
CARBON-STEEL Bi CHECK VALVES

PARTITION

Lw— 6-in. CARBON-
STEEL PIPE

Fl 24 in.
72-16-12 mole % [

FUEL CARRIER N |
salT - ———— N\ - - |§

Fig. 19.1. Carbon-steel vessel for use in the metal transfer
experiment.

249

Table 19,1, Material used in metal transfer
experiment MTE-2

Volume (cms) Gram-moles
Fluoride salt 789 40.7
(LiF-BeF,-ThF4-LaF3,
72-15.7-12-0.3 mole %,
7 mCi 1*7NdF;)
Bismuth saturated with thorium _ 799 36.9
LiCl 1042 36.6
Li-Bi (35 at. % lithium) A 164 9.5

containing a lithium-bismuth solution. During opera-
tion, LiCl was circulated through the lithium-bismuth
cup at the rate of about 25 cm?/min. The concen-
tration of reductant (35 at. % lithium) in the lithium-
bismuth was sufficiently high that at equilibrium
essentially all of the neodymium and lanthanum would
have been extracted from the LiCl that was circulated
through the lithium-bismuth cup.

The quantities of materials used in the experiment are
given in Table 19.1. The LiCl was purified prior to use
by contact with bismuth saturated with thorium at
650°C. Both the carbon-steel vessel and the bismuth
were treated with hydrogen at 650°C to remove oxides.
The argon used as cover gas for the experiment was
purified by passage through a bed of uranium turnings
at 600°C and a bed of molecular sieves. The equipment
and operating conditions used for run MTE-2 were
similar to those used in the first experiment (MTE-1),!
with the following exceptions: (1) a carbon-steel pump
with molten-bismuth check valves, instead of a quartz
pump used previously, was used for circulating the LiCl;
(2) gas-lift sparge tubes were used in the fluoride, LiCl,
and lithium-bismuth compartments to increase the
contact between the salt and bismuth phases; and (3)
about four times as much '*7Nd was used in the

1. E. L. Youngblood et al, MSR Program Semiannu. Progr.
Rep. Aug 31, 1970, ORNL4622, p. 217.
250

second experiment to facilitate determination of the
neodymium concentration in the various phases.

.The experiment was operated for about three months
before it was shut down for disassembly and inspection.
During this time, the pump was operated 441 hr, and
702 liters of LiCl was circulated through the lithium-
bismuth container. The sequence of operations carried
out consisted in pumping the LiCl through the lithium-
bismuth container for 3 hr at a flow rate of 25
cm?®/min, then stopping the pumping, and allowing a
period of 4 hr for the system to approach equilibrium
before filtered samples of the salt and metal phases
were taken. This sequence was repeated three shifts per
day during the first week and two shifts per day during
the second week of operation. For the remainder of the
three-month period, the pump was operated during the
entire day shift and the system was allowed to approach
equilibrium at night and on weekends. During the
pumping periods, the bismuth-thorium phase was
forced back and forth between the fluoride and
chloride compartments at the rate of 10% of the metal
volume every 7 min to promote mixing in the bismuth-
thorium phase. Eight days before the end of the
experiment, 1 vol % fluoride salt was added to the LiCl
compartment to simulate entrainment of fluoride salt in
the bismuth. The temperature of the system was
maintained at about 650°C throughout the entire run.
The pump and other components of the system
operated satisfactorily.

During the experiment, the lanthanum .and the
neodymium that had been originally added to the
fluoride salt transferred to the lithium-bismuth solution
as expected. The rate of accumulation of rare earths in
this solution is shown in Fig. 19.2. There was essentially
no accumulation of lanthanum or neodymium in the

ORNL-DWG 71- TCA

80 ]
- 70 e LANTHANUM .
@ o NEODYMIUM Yloele
3560 L Fr
> METAL TRANSFER | '
~, 50 | EXPERIMENT-MTE-2 . o
= o
O 40 * 3
- [
[ ]
S 30 -~ &
= » 6
Z 20
2 I
E:_' 10 ) o o
a 6;919’
3
0 lewtym
0 100 200 300 400 500 600

VOLUME OF LiCl PUMPED (liters)

Fig. 19.2. Rate of accumuiation of lanthanum and neodym-
ium in the lithium-bismuth solution.

solution until operation of a gas-lift sparge tube in the
cup was initiated (after about 50 liters of LiCl had been
circulated through the lithium-bismuth compartment).
After 400 liters of LiCl had been circulated (about
two-thirds through the run), about 50% of the lantha-
num and 30% of the neodymium originally in the
fluoride salt were found to be in the lithium-bismuth
solution. During this time, 70 to 100% of the lantha-
num and neodymium initially charged to the system
could be accounted for by filtered samples taken of
each phase, indicating that most of the rare earths-
remained in solution. During the last third of the run,
the rare earths continued to accumulate in the lithium-
bismuth solution, but the rate of accumulation could
not be determined accurately because a leak developed
in the lithium-bismuth cup, allowing about 30% of the
solution to flow into the area between the cup and
holder. The extent of removal of lanthanum and
neodymium from the fluoride salt is shown in Figs.
19.3 and 19.4, which show that more than 85% of the
lanthanum and more than 50% of the neodymium had
been removed at the end of the experiment. Approxi-
mately 10 to 20% of the rare earths present were
removed from the LiCl in its passage through the
lithium-bismuth compartment. There was no measura-
ble accumulation of thorium (<10 ppm) in the lithium-
bismuth solution during the experiment, thus demon-
strating that the rare earths can be deposited in this
solution without significant amounts of thorium also
being deposited. The thorium-lanthanum decontami-
nation factor was about 10°. The distribution coeffi-

ORNL—-DWG 7{—63A

(R
L

0.5 e

0.2

FRACTION OF LANTHANUM REMAINING

METAL TRANSFER EXPERIMENT -MTE -2

04

0 100 200 300 400 500 6C0
VOLUME Of LiCi PUMPED (liters)

Fig. 19.3. Rate of removal of lanthanum from the fluoride
salt,
ORNL-DWG 71-64A

.- 1.0 -
z % .
% . -o.“ -
e é s .o ;
= a P .- .!
S5 b ),
205 :
8a
9 o
Z 5
S ¥ ‘
g g METAL TRANSFER EXPERIMENT-MTE-2
£ o
b .
@
(T
.
0.2
0 100 200 300 400 500

VOLUME OF LiCl PUMPED ({iiters)

Fig. 19.4. Rate of removal of neodymium from the fluoride
salt.

cients for lanthanum and neodymium remained rela-
tively constant during the experiment. The average
values of the distribution coefficients for lanthanum
(0.05) and neodymium (0.06) between the fluoride salt
and the bismuth-thorium solutions are in good agree-
ment with expected values. The average value of the
distribution coefficients between the LiCl and the
bismuth-thorium solution was somewhat higher than
expected for lanthanum (3.1) but was near the ex-
pected value for neodymium (4.8). During the run, the
lithium concentration in the lithium-bismuth solution
decreased from 35 to 13 at. %. Only a small fraction of
this decrease was due to the reaction of lithium with
rare-earth chlorides; the principal reason for the de-
crease has not yet been determined. The bismuth-
thorium phase remained saturated with thorium
throughout the run, indicating that there was no
excessive loss of thorium from the bismuth phase
during the experiment.

The concentrations of beryllium, thorium, and fluo-
ride in the LiCl were determined periodically after the
addition of fuel carrier salt to the LiCl near the end of
the experiment, as shown in Table 19.2. During the
93-hr period in which the LiCl was not circulated
through the lithium-bismuth container, the beryllium
concentration remained constant at the initial value of
490 ppm, and the thorium concentration decreased, as
expected, from the initial value of 9480 ppm to a value
of 644 ppm because of transfer of thorium into the
thorium-bismuth solution. When circulation of the LiCl
was resumed, the beryllium concentration in the LiCl
began to decrease, probably because of reduction of the
Be?* by the lithium-bismuth solution. After 27 hr of
operation, a beryllium concentration of 135 ppm was

251

observed. The thorium concentration in the LiCl at this
time was 171 ppm.

After the experiment was completed, the carbon-steel
vessel was cut apart for inspection. The general con-
dition of the inside surfaces was good. The salt and
bismuth phases were sharply defined, and there was no
evidence of accumulated impurities at the salt-bismuth
interface as had been observed in the first experiment
(MTE-1).! The salt and bismuth phases from the
fluoride compartment after the experiment was com-
pleted are shown in Fig. 19.5; Fig. 19.6 shows the LiCl
compartment after most of the LiCl had been removed.
The upper surfaces in the fluoride compartment were

- covered with a black powder having the composition

shown in Table 19.3. The upper surfaces of the LiCl
compartment were covered with a white powder con-
sisting of LiCl containing 0.6 wt % Bi. These materials
are believed to have resulted from vaporization or

Table 19.2. Variation of beryllium, thorium, and fluoride
concentrations in the LiCl after addition of 1 vol %
fuel carrier salt

T;g];i :;f)t;[ Puglnl:;ng Fluoride Beryitium Thorium
(hr) (hr) (wt %) (ppm) (ppm)
0 0 0.98¢ 4904 94804
19.8 0 1.49 490 7200
45.4 0 1.08 500 1100
69.7 0 1.81 320
93 0 2.03 490 644
115.9 5.9 2.34 470
163.9 20.3 1.07 120 1400
188.8 27.4 0.91 350
188.8 27.4 1.7% 135% 1712

%Calculated values based on the amount of fluoride salt
added. ,

bSalt taken from vessel after experiment was concluded.

Table 19.3. Composition of black powder on upper
surfaces of fluoride salt compartment

Element Weight percent
Li 2.4
Be 2.2
Bi 70.2
Th 1.1
Fe 0.02
F. 19.9
Total 95.82

252

PHOTO 102034

Fig. 19.5. Salt and bismuth phases from fluoride salt compartment on completion of metal transfer experiment MTE-2.
253

PHOTO 102035

Fig. 19.6. View of LiCl compartment after removal of most of the LiCl following metal transfer experiment MTE-2.

entrainment of salt and bismuth in the gas streams 10% La, and 2% Th. The mechanism by which this
passing through the compartments. A Y-in-thick layer ~ material was deposited has not been determined;

g of gray material had deposited on the lip and overflow ~ however, it may have resulted from wetting of the
spout of the lithium-bismuth container. The material ~ container wall by the lithium-bismuth solution and
had the composition (in wt %) of 23% LiCl, 59% Bi,  subsequent flow of the solution up the wall.

19.2 DESIGN OF THE THIRD METAL TRANSFER
EXPERIMENT

E. L. Nicholson = W. F. Schaffer
L. E. McNeese E. L. Youngblood
H. O. Weeren

The third engineering experiment for development of
the metal transfer process for removing rare earths from
MSBR fuel carrier salt is being designed. The experi-
ment (MTE-3) will use flow rates that are 1% of the

254

estimated flow rates for a 1000-MW(e) reactor. In the -

two previous engineering experiments,”*? the salt and
bismuth phases were only slightly agitated, resulting in
a low rate of transfer of rare earths from the fuel carrier
salt to the lithium-bismuth solution.

The planned experiment, shown schematically in Fig.
19.7, will use mechanical agitators to promote efficient
contact of the salt and metal phases. Fuel carrier salt
containing rare-earth fluorides will be circulated be-
tween one side of the salt-metal contactor and a
fluoride salt reservoir. Lithium chloride containing
rare-earth chlorides will be circulated between the other
side of the salt-metal contactor and a rare-earth
stripper, where the rare earths will be extracted into a
lithium-bismuth solution.

2. E. L. Youngblood et al., MSR Program Semiannu. Progr.
Rep. Aug 31, 1970, ORNL-4622, p. 217.

'3. Seesect 19.1°of this report.

The experiment will use approximately 35 liters of
fluoride salt, 6 liters of thorium-bismuth solution, 6
liters of LiCl, and 5 liters of Li-Bi solution having a
lithium content of about 40 at. %. The system will
require three process vessels, each of which will be
made of carbon steel. The largest vessel will be the
fluoride salt reservoir, which will contain approximately
32 liters of salt. The remaining 3 liters of fluoride salt
will be contained in the salt-metal contactor. The
fluoride salt will be recirculated continuously from the.
reservoir to the contactor at the rate of about 33
cm’/min by a pump similar to the one used in the
second metal transfer experiment. The salt-metal con-
tactor will be a 10-in.-diam, two-compartment vessel
having a mechanical agitator in each compartment. The
agitator will consist of two flat-bladed paddles mounted
on a common shaft, with a paddle operating in each of
the salt and metal phases in a manner such that the salt
and metal phases are not dispersed. The thorium-
bismuth solution will be captive in the salt-metal
contactor and will form a seal to isolate the fluoride salt
from the LiCl. The thorium-bismuth solution will be
recirculated between the two compartments of the
salt-metal contactor by utilizing the pumping capability
of the agitators. The third vessel, which will be similar
in design to one compartment of the salt-metal con-
tactor, will contain the lithium-bismuth solution. The
LiCl will be circulated between the salt-metal contactor
and the rare-earth stripper at the rate of about 1.25

ORNL-DWG 71-147RA
AGITATORS

LEVEL
ELECTRODES

VENT

T]

HJ

ARGON

SUPPLY

LEVEL
VENT ELECTRODES
FLUORIDE
SALT PUMP i
ARGON - i
SUPPLY
33 cm3/min
N N
/ N

72-16-12 mole Y% -
Lif-BeF,-ThF,

FLUORIDE :
SALT SALT-METAL RARE EARTH
RESERVOIR CONTACTOR STRIPPER

_Fig. 19.7. Flow diagram for metal transfer experiment MTE-3.
liters/min by varying the gas pressure above the LiCl in
the rare-earth stripper. All of -the vessels will. be
operated at 650°C. The carbon steel will be protected
from external air oxidation by a nickel aluminide
coating.

Development work is under way on the salt-metal.

contactor, using a water-mercury mockup to determine
optimum geometry and agitator speeds. Dispersion of
the salt or bismuth will be avoided in order to prevent
entrainment of bismuth in either of the salt phases or
transfer of fluoride salt into the chloride salt' via the
metal phase. This work. is described in detail in Sect.
19.3. :

The fluoride salt reservoir has been fabricated. The
design and fabrication of an agitator drive unit and a
small test vessel have also been completed. The test
vessel duplicates the geometry of the agitator shaft
assembly and will permit a test of the agitator drive unit
and shaft seal at 650°C using bismuth and salt. The
shaft is sealed with an inert-gas-buffered, water-cooled
nonlubricated shaft seal using Teflon Bal-Seal rings. The
agitator paddle and shaft parts that are exposed to salt

and bismuth are made of molybdenum. The interior of -

the carbonsteel test vessel has been plated with a
vapor-deposited coating of tungsten over an electro-
plated nickel bonding layer in order to evaluate the
performance of this type of coating in an agitated
salt-bismuth -system. Provisions have been made for
sampling the salt and metal phases. One-half of the
vessel exterior was spray coated with stainless steel
prior to application of the nickel aluminide in order to
evaluate this alternate coating technique. We have also
started purifying LiCl and preparing rare-earth fluorides
to be used in the experiment.

19.3 DEVELOPMENT OF MECHANICALLY
AGITATED SALT-METAL CONTACTORS

H. O. Weeren ~ L. 'E. McNeese
- J S. Watson

A program- has been initiated for the development of
mechanically agitated salt-metal contactors as an alter-
native to packed columns presently considered for
MSBR processing systems. This type of contactor is of
particular interest for use in the metal transfer process
since designs .can be envisioned in which the bismuth
phase would- be a near-isothermal, internally recircu-
lated captive phase. It is believed that such designs will
require less molybdenum fabrication technology than
would a counterpart system based on packed columns.

Studies to date have been concerned primarily with
selection of a contactor design for the third metal

255

transfer experiment (MTE-3), discussed in Sect. 19.2,
which will have bismuth and salt flow rates that are
about 1% of the estimated rates for a 1000-MW(e)
MSBR. Several scouting tests were carried out with
water and mercury in vessels having diameters of 4 to 7 .
in., and a mockup of the contactor proposed for
experiment MTE-3 was built for additional study with
mercury and water. The mockup consists of an 8-in.-
diam vessel having a central partition that extends to
within % in. of the bottom of the vessel. The first tests
with the mockup were made with a flat four-bladed
paddle located in the mercury-water interface. The
paddle, 1.5 in. in diameter, was located inside a
3-in.-diam, 3-in.-high shroud contammg four 0.3-in.-
wide, 3-in.-long baffles. The design was chosen to
maximize the extent of dispersion of the mercury in the
water, thereby maximizing the interfacial area between
the two phases. The system was tested with agitator
shaft speeds up to 1600 rpm. It' was found that, under
these operating conditions, a stable dispersion of very
small mercury droplets was frequently formed. Also, at
all but the lowest mixer speeds, a dispersion of water
droplets in the mercury was formed, and these droplets
were pumped from one chamber of the mockup to the
other. Such pumping cannot be tolerated in the metal
transfer process since it would result in the mixing of
chloride and fluoride salts. The tendency of the
salt-bismuth system to form emulsions may be quite
different from that of the water-mercury system;
however, it was concluded that the contactor should
operate under conditions that minimize the hkehhood
of formation of emulsions.

The contactor design that has the greatest potential
for ac}uevmg good mass transfer with minimum dis-
pers1on appears to be the Lewis cell® — a contactor
W1th a paddle 1n each phase, located well away from the

interface as ‘shown in Fig. 19.8. With such a contactor
the phases would be agitated: as vigorously as’ possible
without actually causing dlSperSIOI'l of one phase in the
other. The contactor mockup has been modified in
order to study the hydrodynamics of this contactor.

The mass transfer performance of this type of
contactor is known for a number of ag1tated nondis-
persed two -phase systems, and the ‘results can be
correlated in terms of the Reynolds numbers and
viscosities of the two phases. Overall mass transfer
coefficients for the salt-bisinuth system predicted by
this correlation appear to be adequate for MSBR
processmg operations.”

4. 1. B. Lewis, Chem Eng Sci. 3, 248—59 (1954).
ORNL—DWG T1—2864A

/AGITATOR

LIGHT PHASE ~_ |

i
|
iiiii! :/ Eii.i!i;ii [ INTERFACE
T
1

HEAVY PHASE —__ |—= il | /_/

e

"~ BAFFLES

Fig. 19.8. Proposed salt-metat contactor.

19.4 REDUCTIVE EXTRACTION ENGINEERING
STUDIES

B. A. Hannaford C. W. Kee
L. E. McNeese

We have continued to study the extraction of
uranium from molten salt (72-16-12 mole % LiF-BeF, -
ThF,) by countercurrent contact with bismuth con-
taining reductant in a packed column. Two successful
runs (UTR-3 and 4), which included six periods of
steady-state operation covering metal-to-salt flow ratios
ranging from 0.75 to. 2.05, have been made.

Prior to run UTR-3, 122.5 g of thorium metal was
added to the treatment vessel, which contained about
10% of the salt (1.5 liters) and most of the bismuth
(~17 liters) present in the system. The thorium
dissolved in the bismuth at a very low rate (~0.3%jhr),
which was nearly equivalent to the rate observed
previously® in the bismuth feed tank prior to run
UTR-2. In order to improve the contact between
thorium and the bismuth, an additional 119 g of
thorium (contained in-a l-in.-diam perforated steel
basket) was suspended in the bismuth. The resulting
dissolution rate was about ten times the earlier rate;
about 85% of the thorium dissolved in a 22-hr period.
Samples of the bismuth phase showed a range of
thorium concentrations and suggested the existence of a

5. B. A. Hannaford, C. W. Kee, and L. E. McNeese, MSP
Program Semiannu. Progr. Rep. Aug. 31, 1970, ORNL-4622, p.
211 ’

256

concentration gradient in the bismuth pool. The final
thorium concentration in the bismuth is believed to
have been 1200 ppm. The thorium-saturated bismuth
and the salt were then transferred to the feed tanks; the
final uranium concentration in the salt feed tank was
2100 ppm.

During run UTR-3, bismuth was fed to an 0.82-in.-ID,
24-in.-high extraction column at the rate of about 200
cm? /min, and salt was fed to the column at the rates of
100, 160, and 220 cm?/min. Under the first combi-
nation of operating conditions, five sets of samples were
taken of the salt and bismuth streams leaving the
column; however, only one set of samples was taken
under each of the other combinations of conditions,
after a salt volume equivalent to three column volumes
had passed through the column. As shown in Table
19.4, the fraction of the uranium extracted from the
salt decreased from 0.92 to 0.73 as the salt flow rate
was increased. ' '

The thorium dissolution rate in the treatment vessel
prior to run UTR<4 was about the same as that observed
prior to run UTR-3. Some nonuniformity in thorium
concentration was observed in the bismuth. The final
thorium concentration in the bismuth feed tank was
about 1100 ppm, and the uranium concentration in the
salt feed tank was 1680 ppm. During run UTR4, the
system was operated at three sets of salt and bismuth
flow rates. Eleven pairs of samples were taken from the
salt and bismuth streams leaving the column. The
fraction of extracted uranium ranged from 63 to 74%,
as shown in Table 19.4.

Watson has noted® that the uranium extraction data
from runs UTR-3 and 4 can be correlated in terms of
the height of an overall transfer unit based on the salt
phase (HTU) if several assumptions are made. The chief
assumption is that the rate at which uranium transfers
to the bismuth phase will be controlled by the diffusive
resistance in the salt film when the extraction factor is
high and when the salt film is composed largely of
nontransferring ions. In runs UTR-3 and 4, a small
amount of uranium was added to the salt after it had
been equilibrated with the bismuth phase; no significant
transfer of lithium and thorium occurred in the column.
In this case, the overall transfer coefficient based on the .
salt phase is equal to the individual salt film transfer
coefficient. By definition, the HTU and the number of
overall' transfer units based on the salt phase (NTU)
developed in the column are related as

H=HTU-NTU,, (1)

_6. I. S. Watson, ORNL, personal communication.
257

Table 19.4. Data obtained from uranium mass transfer experiments in an 0.82-in.-I1D,
24.in.-high column at 600°C

Uranium concentration in salt (ppm) Salt

. Bismuth Metal-to-salt Fraction -Fraction of Fraction
. Maximum. flow flow flow of U
Salt Salt equilibrium rate rate rate Of. U extracted remaining
feed, X ; effluent, X, value,? (cm®/min)  (cm>/min) ratio flooding from salt in salt
. X*
UTR-3 2100 159 3.6 100 205 2.05 0.87 0.92 0.076
445 9.4 160 195 .22 1.04 0.79 0.212
561 10.3 220 200 0.91 1.23 0.73 0.267
UTR-4 1680 524 9.7 149 140 0.94 . 0.85 0.69 0.312
439 7.2 118 117 1.0 0.69 0.74 0.261
626 9.7 210 157 0.75 1.07 0.63 0.373

%Calculated as the concentration that would be in equilibrium with the observed concentrations of reductant (lithium) and

uranium in the bismuth efflqent.

where A = column length. If it is assumed that uranium
is the major component transferring from the salt and
that the controlling resistance to-transfer is in the salt
phase, HTU can be written as

V

HTU =—, (2)
ka

where ¥V is the superficial velocity of the salt in the
column and ka is the product of the overall mass
transfer coefficient based on the salt phase and the
interfacial area between the salt and bismuth phases per
unit column volume. It has been previously” observed
that the dispersed-phase holdup is approximately pro-
portional to the flow rate of the dispersed phase except
at conditions near flooding. We would, therefore,
expect the interfacial area between the salt and the
bismuth phases to be proportional to the bismuth flow
rate, so that

ka=k'Vg,, G

where k' is a constant and Vg; is the superficial velocity
of the bismuth in the column. The number of overall
transfer units, based on the salt phase, developed in the
column is defined as

X

X |

NTU=f e (4)
S X

7. 1. S. Watson and L. E. McNeese, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 214.

where

X = uranium concentration in the bulk salt,

" X* = uranium concentration in the salt in equilibrium
with the bulk bismuth phase,

X; = uranium concentration in the salt fed to the
column,

X, = uranium concentration in the salt leaving the
~olumn.

As shown in Table 19.4, the value of X* at the bottom
of the column is much smaller than the value of X ,;
thus one would expect that X* < X throughout the
column. In this case, Eq. (4) would become

X, - ,
NTU=-In—. :
n— (5)

I

Combining Egs. (1), (2), (3), and (5) and rearranging
yields the relation

X, Va;
In— = —k'H —,
n (6)

H

which states that a semilogarithmic plot of the fraction
of the uranium remaining in the salt vs the bismuth-to-
salt flow rate ratio should yield a straight line having a
slope of —k'H. This line would pass through an ordinal
value of 1.0 at a bismuth-to-salt flow rate ratio of zero.

As shown in Fig. 19.9, the data from runs UTR-3 and
4 are well represented by Eq. (6). For these data, the
constant X has the value of 0.0529 in.”!. The product
ORNL—DWG 74— 2865A

\ © UTR-3 DATA
\ ® UTR—4 DATA

0.5 \\

o N

N\
\

X

N\
\

o} 0.5 1.0 1.5 2.0 2.5 3.0
¥,/Vy, BISMUTH TO SALT FLOW RATE RATIO

FRACTION OF URANIUM REMAINING IN SALT
o
N

0.05

Fig. 19.9. Mass transfer of uranium from LiF-BeF,-ThF,
(72-16-12 mole %) to bismuth containing reductant at 600°C in
an 0.82-in.-diam, 24-in.-long packed column.

of the overall transfer coefficient, based on the salt
phase, and the interfacial area is given by the relation

ka=0.0529Vy, , | (7)

where ka is the overall rate constant based on the salt
phase, sec™', and Vp; is the superficial velocity of the
bismuth in the column, in./sec. Values for the overall
rate constant for the present data range from 0.012 to
0.021 sec™ and compare favorably with a preliminary
value of 0.0076 sec™ measured® for the transfer of
uranium from a 96.2-3.6-0.2 wt % Cd-Mg-U solution to
a molten salt (50-30-20 mole % MgCl,-NaCl-KCl) at
temperatures ranging from 560 to 610°C.

The HTU for the data from runs UTR-3 and 4 is
given by the expression

18.9

HTU= ——,
VBi/ VS

®)

where HTU is the height of the overall transfer unit
based on the salt phase, in., and Vg;/V; is the
bismuth-to-salt flow rate ratio. The HTU values range

8. T. Johnson et al., Chem Eng Div. Semiannu. Rep.,
January—-June 1965, ANL-7055, p. 44.

258

Table 19.5. Predicted performance and required
" height for the protactinium isolation column

. Required
NTU Fraction of Pa column height
not extracted
(ft)
1 0.368 | 10.5
2 0.135 21.0
3 0.0498 315
4 0.0183

42.0

from 0.77 ft at a flow rate ratio of 2.05 to 2.1 ft at a
flow rate ratio of 0.75. It is of interest to note that the
flow rate ratio presently considered for the protac-
tinium isolation column in the reference MSBR flow-
sheet® is 0.15; the predicted HTU at this flow ratio is
10.5 ft. The predicted column performance and the
required column height for NTU values ranging from 1
to 4 are shown in Table 19.5. It is apparent that column
heights considerably greater than the presently assumed
height of 10 ft would be required for removing 95% of
the protactinium. However, it should be kept in mind
that these predictions were obtained by extrapolating
preliminary data to a point well outside the range
covered by the data. If these predictions are valid, the
processing rate could be increased by about 10% in
order to compensate for the lower extraction efficiency
expected with a column height of about 20 ft. A second
possibility would be the use of an alternate type of
contactor that would produce high extraction effi-
ciencies at low bismuth-to-salt flow rate ratios.

In order to measure mass transfer rates in the column
under more closely controlled conditions and under
conditions where the controlling resistance is not
necessarily in the salt phase, preparations were begun
for experiments in which the rate of exchange of
zirconium isotopes will be measured between salt and
bismuth phases otherwise at equilibrium. Residual
reductant was removed from the treatment vessel by
hydrofluorinating the salt and bismuth for 20 hr with
70-30 mole % H,-HF. Following the usual H, sparge
for reduction of iron fluoride, the two phases were
transferred to the feed tanks. A hydrodynamic experi-
ment (UTR-5) was then carried out (1) to remove from
the system any bismuth having a high thorium concen-
tration and (2) to test the sampling and analysis
techniques under conditions where no mass transfer
should occur. Surprisingly, the reported uranium con-
centrations of salt samples removed from the column

9. Seesect 17.1, this report.
259

Table 19.6. Flooding data obtained during countercurrent flow of molten
salt and bismuth in a packed column during runs UTR-5 and -6

Bismuth  Salt
Run No. Flow rae | Superficial o Superficial
3, . velocity 3, . velocity
{cm” /min) (£t/hr) {cm” /min) (ft/hr)
UTR-5 ‘115 65.8 123 70.2 Nonflooded
219 125.2 50 28.6 Nonflooded
146 83.4 161 92.0 Flooded
276 157.7 90 51.4 Flooded
UTR-6 117 66.9 125 714 Nonflooded
209 119.5 150 85.7 Flooded
. ORNL— DWG 74 —-2840A
effluent varied by £35% from the average value, which 5OSAL4To;LOW2(()%|/rgi3)c> 400 500 700
was in excellent agreement with the indicated uranium L2 | l | ] T 700
concentrations of the salt feed and catch tanks. = PREDICTED FLOODING CURVE  _| g00
An additional run (UTR-6) was made in order to s V2 + V%2 =19.0 (ft/nn)¥2 500
further test the sampling and analysis techniques and to N 0 l
obtain hydrodynamic data. In this run, salt and bismuth 5 16 \o o FLoopen ] 49° -
were fed to the column at flow rates of 125 and 117 3 \ . ® NON FLOODED | 55 E
cm?/min, respectively, for a 38-min period. During this Z. e .\o.. . £
period, seven pairs of samples were taken of the salt and o O\ ] — 200 2
bismuth streams leaving the column. Analysis of these E 0 i PN . E
samples showed a scatter in uranium concentration > . ¢ Jioo E
similar to that observed in run UTR-5. The flows of salt T . %
and bismuth were stopped following the initial oper- % 5 . — 50 @
ating period in order to freeze the drain line leaving the o N
specific gravity pot. Operation was then resumed witha <o
bismuth flow rate of 209 ¢m?3/min and a salt flow rate
of 149 ¢m3/min. Salt samples were taken periodically o

for bismuth analysis. The apparent dispersed-phase
holdup, as indicated by the pressure drop across the
column, stabilized at 30%. A bismuth concentration in
the salt of about 5 ppm was observed during this
period. When the salt flow rate was increased to 151
cm?® /min, the apparent bismuth holdup increased to
about 60% before the specific-gravity pot began to fill
with bismuth, a clear indication that the column was
flooded. A bismuth concentration in the salt of about
10 ppm was observed during the latter period. The
flooding data obtained during runs UTR-5 and -6 are
shown in Table 19.6. Figure 19.10 shows these data,
along with flooding data from previous runs. The data
are well represented by the predicted flooding relation
(denoted by the line in Fig. 19.10) resulting from work
with mercury and aqueous solutions (see Sect. 19.5).
Both phases were returned to the treatment vessel
after run UTR-6, and sufficient thorium was added to
extract about 50% of the uranium and zirconium from
the salt phase. The thorium dissolution rate was about

1 10 15 20
V2, SALT SUPERFICIAL VELOCITY (ft/hn)"2

Fig. 19.10. Summary of flooding data with salt and bismuth
in an'0.82-in.-diam, 24-in.-long column packed with 1/.;;-in.
Raschig rings.

one-half that observed prior to runs UTR-3 and 4. A
draft tube was installed in. the treatment vessel in order
to improve mixing of the phases. The basket containing
the thorium was removed after about 156 g of thorium
had dissolved. The treated salt and bismuth were then
transferred through the system in order to bring the salt
and ‘bismuth in the entire system to equilibrium before
the start of the ®7Zr tracer experiments. After both
phases had been returned to their respective feed tanks,
we withdrew samples that showed an average zirconium
concentration of 165 ppm in the salt and a zirconium
distribution ratio of 1.0, which are satisfactory con-
ditions for the tracer experiments.
The initial attempt at producing 17-hr °7Zr by
irradiation of ?°Zr0, resulted in-an activity about four
times that expected for °7?Zr. The discrepancy was
traced to an outdated value for the thermal-neutron
cross section for °°Zr, that is, 0.05 barn as compared
with a more recent value of 0.2 barn. After a suitable
interval, the original 20-mg charge of %¢Zr0, was
reirradiated for 2 hr to produce a calculated ®7Zr-* "Nb
activity of 13.7 mCi. Counting rates for ®’Zr were
measured for samples taken from the salt feed tank at
intervals following the addition of tracer to the tanks.
Mixing of the tracer in the 15-iter salt volume was
about 75% complete within 2 hr. The irradiation yield
of *7Zr->’Nb was about 30% higher than expected.
‘The °7Zr counting rate, which was greater than
200,000 counts/min per sample, indicates that counting
precision in future experiments should be good.

At the time the first ®7Zr tracer experiment was
begun, a leak appeared in the salt feed line at a location

260

near the salt feed tank. An attempt to transfer the salt.

from the tank by an alternate line disclosed a second
leak. The resulting salt spills caused the Calrod heaters
on the tank to fail and extensively damaged the nickel
aluminide coating on the exterior of the tank.

The tank was removed from the system for examina-
tion and disposal since the damage from the salt made
salvage impractical. Examination, by the Metals and
Ceramics Division, of a metal specimen cut from the
outer shell of the tank revealed that, although two years
of operation at about 600°C had produced considerable
graphitization of the steel, ductility and, apparently,
tensile strength were not significantly impaired. A new
salt tank was fabricated and installed in the system. A
2-in.Jong, 0.5-in.-diam sleeve was attached to each of
‘the sections of the ¥-in.-diam transfer lines on which
the mechanical fitting ferrules seat in order to prevent
occurrence of the earlier type of failure which resulted
from thinning of the transfer lines by distortion under
the ferrules. Insulation was removed from all transfer
lines to allow inspection, and all lines that were more
than moderately oxidized were replaced.

19.5 CONTACTOR DEVELOPMENT: PRESSURE
DROP, HOLDUP, AND FLOODING
IN PACKED COLUMNS

J. S. Watson L. E. McNeese

Studies of the hydrodynamics in packed columns

during the countercurrent flow of high-density liquids

are being made in order to evaluate and design
contactors for processing systems based on reductive

extraction. We have previously'?:!'! reported studies in

which mercury and water were used to simulate
bismuth and molten salt in columns packed with % 4-,
Y-, Y-, and % -in. Raschig rings and with %- and Y%-in.
solid fight circular cylinders. These studies have shown
that the dispersed-phase holdup and the column
throughputs at flooding can be correlated'' on the
basis of a constant slip velocity in the following
manner:

Vc Vd
Tt x " Vs (1)
Vc,flfz + Vd,f”2 = Vsl/2 , (2)

where

V. = superficial velocity of the continuous phase,
V4 = superficial velocity of the dispersed phase,
V, = slip velocity,

X = dispersed-phase holdup,

f=subscript denoting superficial velocities at flood-
ing.

These relations were previously'! extended to cover
dispersed-phase holdup and throughput at flooding with
salt-bismuth systems by assuming that for a given
packing size the slip velocity was (1) independent of the
viscosity of the continuous phase, (2) proportional to
the difference in the densities of the phases, and (3)
proportional to the packing void fraction. Although the
resulting relations predicted flooding rates that were in
excellent agreement with flooding rates measured with
bismuth and molten salt, it was realized that the
agreement did not constitute verification of the as-
sumed effects of the continuous-phase viscosity and the
difference in the densities of the phases.

During this reporting period, data showing the de-
pendence of slip velocity on continuous-phase viscosity
were obtained by an MIT Practice School group. A
2-in.-diam, 24-in.-long column packed with %;-in.
Teflon Raschig rings, which were not wetted by either
phase, was used in the study. The experimental system
was modified so that water or water-glycerin solutions
could be recirculated through the column at constant
temperature by installation of a heat exchanger and an

10. J. S. Watson and L. E. McNeese, MSR Program Semiannu.
Progr. Rep. Feb. 28, 1970, ORNL-4548, p. 302,

11. ]. S. Watson and L. E. McNeese, MSR Program Semiannu.
Progr. Rep. Aug. 31, 1970, ORNL-4622, p. 213,
261

Table 19.7. Variation of slip velocity with continuous-phase
viscosity and wetting of packing

Stip velocity, Standard Percent
System g deviation, deviation,
(ft/hr) o o/V,
1 ¢P, nonwetted packing 1129.3 +105.2 9.3
7.5 ¢P, nonwetted packing 847.6 +88.2 +10.4
15 cP, nonwetted packing 673.5 +49.0 +7.3
1 ¢P, wetted packing 1582.7 +267.6 +16.9

aqueous-phase surge tank. Data were obtained with
glycerin solutions having viscosities of 7.5 and 15 cP,
and earlier experiments with water (which has a
viscosity of 1 cP) were repeated since repacking a
column can alter the slip velocity by as much as 10%.

The results obtained were similar to those observed
previously for mercury and water. The dispersed-phase
holdup could be correlated in terms of a constant slip
velocity, as shown in Table 19.7, for the three cases
involving nonwetted packing. The relative standard
deviations for the slip velocities were about *10%; no
dependence on the flow rate of either phase was noted.
The variation of slip velocity with continuous-phase
viscosity is shown in Fig. 19.11, which indicates that
the slip velocity is proportional to the —0.167 power of
the continuous-phase viscosity. As expected, the de:
pendence is not large. However, neglect of this effect
was significant in the earlier extrapolation from a
water-mercury system to a salt-bismuth system since
the continuous-phase viscosity changed by a factor of
12.

After the dependence of slip velocity on the con-
tinuous-phase viscosity was known, it was possible to
reevaluate the dependence of slip velocity on the

ORNL— DWG 71— 2867A

20

ol .--""---._I

f\ ‘---—-.'-.--.._l\

= {0 - i
o T -.;

! SLOPE = —(VISCOSITY EXPONENT -

2 = —b~ 0467 i
ko

- 5

=

Q

[0

-

ul 4

-

a 2

-

wy

1 - .
(oX| 0.2 0.5 { 2 5 10 20

CONTINUQUS PHASE VISCOSITY (cP)

Fig. 19.11. Variation of slip velocity with continuous phase
viscosity in a 2-in.-diam column packed with ¥%-in. Raschig
rings.

difference in the densities of the two phases by using
the two data points afforded by the mercury-water data
and the salt-bismuth data. If it is-assumed that the
dependence of slip velocity on the difference in
densities is a power-type dependence, a power of 0.5 is
calculated. This result is interesting in that it is the same
as the dependence of drop terminal velocity on the
difference in densities in the inertial region, where the
drag coefficient is essentially constant. The final rela-
tion for predicting the variation of slip velocity with
packing void fraction, the difference in the densities of
the phases, and the continuous-phase viscosity is, then:

N € u \ 0167

= Ys,Hg-H,0

* HRE €ref “Hzo

A 0.5
X 2 , )
ApHg,HgO-
where
V¢ =slip velocity,
Vs Heg-H,0 =slip velocity for mercury-water for the

packing size considered,
€ = packing void fraction,

€of = void fraction for packing for which
Vs,Hg-H»zO was determined,

M = viscosity of continuous phase,
Hy,0 = viscosity of water at 20°C,
Ap = difference in the densities of the phases,
ApHg_H20 = difference in tl;e densities of mercury
and water at 20 C.

Slip velocity values calculated from Eq. (3) can then be
used with Egs. (1) and (2) for determining the
throughputs at flooding and the dispersed-phase hold-
up.
The effect of wetting of the packing by the nominally

dispersed phase was also evaluated by packing the
column with %-in. copper Raschig rings that had been
etched with nitric acid to cause the packing to be
wetted by the mercury. Complete wetting of the
packing was obtained by the mercury, which was
saturated with copper. (The solubility of copper in
mercury is quite low, and no important changes in
other physical properties should occur.) After a few
hours of operation, solids could be seen at the
water-mercury interface below the column; these solids
were believed to be copper oxide formed as the result
of the reaction of dissolved copper with oxygen in the
system. Periodic additions of nitric acid to the system
quickly removed the solids.

The interfacial area was decreased substantially when
the packing was wetted by the mercury. No dispersion
of the mercury was observed; the interfacial area was
essentially the packing surface area. The-slip velocity
(and hence the flooding rates) was considerably greater
than with nonwetted packing, as shown in Table '19.7.
It was not clear whether the data on metal-phase
holdup with wetted packing could be correlated on the
basis of a constant superficial slip velocity. Only nine
metal holdup measurements were made, and a quantita-
tive analysis of the relation between metal holdup and
the water and mercury superficial velocities was not
possible.

19.6 DEVELOPMENT OF A FROZEN-WALL
FLUORINATOR

J. R. Hightower, Jr.  C.P. Tung

An experiment to demonstrate operation of a con-
tinuous fluorinator having a layer of frozen salt
deposited on its walls for protection against corrosion
will use high-frequency induction heating to provide an
internal heat source in the molten salt. Estimates of the
performance'? of a frozen-wall fluorinator having an
induction coil embedded in the frozen salt near the
fluorinator wall indicate that this may be an acceptable
heating method. Because of uncertainties in the effect
of bubbles in the molten salt and in the amount of heat
that will be generated in the metal walls of the
fluorinator, heat generation rates are being measured in
a simulated fluorinator. In this simulation a 31 wt %
HNOj solution, which has electrical properties similar
to molten salts, is used to simulate molten salt in the
fluorinator vessel. Three induction coil designs have

12. J. R. Hightower, Jr_.,:et al., MSR Program Semi&npu
Progr. Rep. Aug. 31, 1970, ORNL-4622, pp. 219-21.

been tested, and the effect of bubbles in the nitric acid
has been determined with the best coil design tested
thus far. 7

Description of equipment for induction heating
studies. The induction-heated simulated fluorinator
consists of a 5-in.-OD, 5-ft-long glass tube inserted in an
induction coil and placed inside a 5-ftdong section of
8-in.-diam 304 stainless steel pipe. The nitric acid inside
the glass tube represents the molten salt zone of the
fluorinator, the space between the glass tube and the
pipe wall (which contains the induction coil) represents
the frozen salt layer, and the pipe represents the
fluorinator vessel wall. As shown in Fig. 19.12, the acid
is circulated through the glass column, where it is
heated by the induction coil, and through a heat
exchanger, where the heat is removed. The heat
generation rate in the acid is determined from a heat
balance on the acid as it passes through the column.
The pipe representing the fluorinator wall is equipped
with a jacket through which cooling water passes. The
heat generated in the pipe is calculated from the change
in temperature of the water as it passes through the
jacket. - :

The three induction coils that were tested consisted
of a number of smaller coils, each of which was 5.6 in.
in inside diameter, 3 in. long, and connected in parallel
to headers carrying water and rf current to each small
section. The characteristics of these coils are given in
Table 19.8. Coils I and I1I had adjacent smaller sections
wound in the opposite directions; coil II had all smaller
sections wound in the same direction.

Experimental results. Twenty-seven runs have been
made with the simulated continuous fluorinator to
determine the heat generation rates in the column of
acid, in the pipe surrounding the acid, and in the
induction coils. The results were expressed in terms of
effective resistances of each component (acid, pipe, and
coil). The effective resistance is defined as the ratio of
the heat generation rate (in watts) to the square of the

Table 19.8. Characteristics of induction coils
that were tested

Number of Diameter Diameter
Number of
. . turns of of
Coil Material small .
coils in each conductor header
small coil (in.) (in.)
I  Monel 17 ~6% Y %
Il 304-LSS 18 6 % Y,
Il 304LSS 18 6 % A

OFF—GAS

ORNL-DWG 70—14713A

SIGHT GLASS
v

)

5 o ———— —

Fl-4 ' ===
QR
Qj,_ _____ ‘o]
V-4 6LASS COLUMN L, \Q 0
S NI &
INDUCTION COIL\T&,Q S

COOLING WATER
TO INDUCTION COIL Q R
Q >
1L e}
JACKETED PIPE—]
Q fe]
Q rel
FI-2 q >
. L fi Jb 1 r

3

HEAT EXCHANGER

COOLING WATER
TO HEAT EXCHANGER

COOLING WATER
TO JACKETED PIPE

DRAIN TANK

Fig. 19.12. Flow diagram for induction heating experiments.

coil current in amperes (rms). The efficiency of heating -

the acid is found by dividing the resistance of the acid
by the sum of the resistances of the acid, pipe, and coil.
The results obtained in the runs are summarized in
Table 19.9, which shows: (1) the average effective
resistances at 20°C of the acid, of the coil, and of the
pipe; (2) the efficiency of heating the nitric acid; and
(3) the predicted efficiency of heating molten salt in a
nickel fluorinator after correcting for differences in
conductivities between the simulated and the actual

systems. In some of the runs with coil III, air was .

bubbled through the nitric acid at rates up to 2.16
scfm, resulting in bubble volume fractions as high as
18%. Typical fluorinators will operate with a bubble
volume fraction of about 15%. Also shown in Table
19.9 is the effect of bubble volume fraction on the
efficiency of heating the nitric acid and on the
predicted efficiency of heating molten salt in a nickel
fluorinator. _ ‘

The predicted efficiencies for heating molten salt
shown in Table 19.9 are for a 5-ft-long fluorinator
having a 4.5-in.-diam molten salt zone, a 1.5-in.-thick
264

Table 19.9. Summary of results from induction heating tests in simulated fluorinator

Effective Effective Effective .. Predicted
Bubble . . . Efficiency of g
. resistance resistance resistance . efficiency of
Coil volume . . . heating .
fracti of acid of coil of pipe acid (%) heating
: Taction (2) () () o molten salt (%)
I 0 0.0143 0.0617 0.0121 16.2 . 28.8
II 0 0.00845 0.0514 0.0076 12.5 24.7
11 0 0.0168 0.0559 0.0106 20.2 ' 37.0
111 0.1 0.0150 0.0559 - 0.0106 18.4 . 34.3
111 0.15 0.0141 - 0.0559 0.0106 17.5 ’ 32.9

frozen salt layer, and a 50°C temperature difference
across the salt layer. In such a fluorinator, about 1265
W must be generated in the molten salt; this requires a
total coil current of 243 A and a generator capable of
generating at least 3900 W. Such requirements are quite
reasonable and could be fulfilled by the present
generator.

We are planning to test another coil design before
proceeding to the next phase of experimental work, in
which we will attempt to form and maintain frozen salt
films in equipment containing a static salt volume.

19.7 ESTIMATED CORROSION RATES IN
CONTINUOUS FLUORINATORS

J. R. Hightower, Jr.

Nickel is the preferred material of construction for
fluorinators in the MSBR fuel processing plant since it
exhibits greater resistance to attack from gaseous
fluorine than any other material considered. The
resistance to attack by fluorine results from a tightly
adhering film of the NiF, corrosion product through
which fluorine must diffuse to react with the metallic
substrate. The rate of reaction with mickel is greatly
reduced, although not to zero, once a protective NiF,
film is formed on a nickel surface exposed to gaseous
fluorine. The purpose of the frozen salt film in a
continuous fluorinator is to prevent the protective NiF,
film from being washed away by the molten salt. (No
benefit is assumed for any added resistance to fluorine
diffusion which may be offered by the frozen salt film.)
However, it is likely that the protective NiF, film will
be destroyed unintentionally several times during the
operating life of a fluorinator, resulting in relatively
high corrosion rates during the period required for the
formation of a new NiF, film. Therefore, we have
estimated the expected corrosion rates which result

from periodic destruction of the NiF, film during
operation of a continuous fluorinator.

Data for the corrosion of Ni-200 and Ni-201 in
gaseous fluorine at 1 atm were collected from the
literature! 3~ 18 and were used to calculate rate con-
stants for the reaction between nickel and fluorine at
l-atm pressure in the temperature range of 360 to
700°C; the reaction was assumed to follow a parabolic
rate law. It has been shown'? that the reaction of
high-purity nickel with fluorine initially follows a
parabolic rate law from 300 to 600°C but that after a
certain exposure time, the reaction rate decreases and
follows a third- or higher-order rate law. Thus the
assumption of a parabolic relation takes into account
the time dependence of the reaction rate but should
yield conservatively high estimates of the extent of
cOrrosion. _

The extent of corrosion of nickel in fluorine was
assumed to be described by the equation

d=kJ/t, ()

13. R. L. Jarry, W. H. Gunther, and J. Fischer, The
Mechanism and Kinetics of the Reaction between Nickel and
Fluorine, ANL-6684 (August 1963),

14, Chem. Eng Div. Summary Rep., July, August, Septem-
ber, 1958, ANL-5924 (1958).

15. Chem Eng Div. Semiannu. Rep., January—June 1964,
ANL-6_900 (August 1964). .

16. W. H. Gunther and M. J. Steindler, Laboratory Investi-
gations in Support of Fluid-Bed Fluoride Volatility Processes.
Part XIV. The Corrosion of Nickel and Nickel Alloys by
Fluorine, Uranium Hexafluoride, and Selected Volatile Fission
Product Fluorides at 500°C. ANL-7241 (December 1966).

17. F. T. Miles et al., Progress Report of the Reactor Science
and Engineering Department, BNL-176, pp. 17—18 (March
1952). : '

18. P. D. Miller and W. E. Berry, A Survey of Corrosion in
the Fluidized-Bed Volatility Process, BMI-X-362 (November
1965). '
where

d = depth of nickel attacked by F,, mils;

t=time of exposure of nickel metal to gaseous
fluorine, measured from the time when no NiF,
film exists; '
k = parabolic rate constant, mils hr~1/2,

Rate constants were calculated for 41 measurements
for which the exposure times ranged from 5 hr to 960
hr. Most of the exposure times were in the range of 30
to 150 hr. Fourteen measurements had been made to
determine the rate of corrosion of Ni-201 in the
temperature range 380 to 700°C; exposure times varied
from 5 to 132 hr. The calculated rate constants are
given elsewhere.'® The best least-squares representa-
tions of the data, which showed considerable scatter,
are given below for Ni-200 and Ni-201 respectively:

3691
Ink=03773 — 221 2)
T
7836
Ink =4.3083 - ——, | 3)

where & has units of mils hr~!/2 and T has units of °K.
The largest deviations of individual data points from
these two equations were about an order of magnitude
higher and an order of magnitude lower.

If n is the number of times per year that the NiF,
film is destroyed, the extent of corrosion experienced
each year (8760 hr) is given in mils by

o /8760
nk
n

D

(4)

and is the average corrosion rate per year. The average
corrosion rates for film lives of one month, one week,
and one day are shown in Table 19.10. If the NiF, film
were destroyed 52 times per year, the average corrosion
rates at 450°C (the approximate wall temperature that
will be used in a frozen-wall fluorinator) would be 2.9
mils/year for Ni-200 and 0.97 mils/year for Ni-201. If
the film were destroyed 12 times annually, the average

corrosion rates would be 1.4 mils/year and 0.47

mil /[year for Ni-200 and Ni-201 respectively. According
to these results, Ni-201 seems to be more resistant to
corrosion than Ni-200; however, either material shows

19. L. E. McNeese, Engineering Development Studies for
Molten Salt Breeder Reactor Processing No. 9., ORNL-TM-3259
(in preparation).

265

Table 19.10. Estimated average corrosion rates in a frozen-wall
fluorinator having a wall temperature of 450°C

Corrosion rate

, Average life (mils/year)
of film
Ni-200 Ni-201
12 1 month 1.4 0.47
52 1 week 2.9 0.97
365 1 day 7.66 2.58

satisfactory corrosion resistance if the NiF, film is kept
intact for periods having an average length as great as
one week. It appears that the anticipated corrosion rate
would be influenced much more strongly by the length
of time that a protective NiF, film is absent in the
presence. of fluorine and molten salt than by the
frequency of. destruction of the NiF, film in the
absence of fluorine. '

19.8° AXIAL DISPERSION IN SIMULATED
CONTINUOUS FLUORINATORS

J. S. Watson L. E. McNeese

Axial dispersion is important in the design of con-
tinuous fluorinators, which are envisioned as open
columns through which fluorine is bubbled counter-
current to a flow of molten salt. We have previously
reported data?®~22 showing the variation of dispersion
coefficient with changes in gas and liquid flow rates,
physical properties of the liquid, column diameter, and
gas inlet diameter. During this reporting period, studies
of the effect of column diameter were extended to a

. 6-in.-diam column, and additional data were obtained

on the effect of the viscosity of the liquid. Most of the
experimental data were obtained by students of the
MIT Practice School.

Dispersion coefficient data measured in a 6-in.-diam,
72-in.-long column are shown in Fig. 19.13. The
dispersion coefficient values were about three times the
values measured at the same superficial gas velocity in a
3-in.-diam column. The data show little dependence on
superficial gas velocity. Although superficial gas veloci-
ties up to 10 cm/sec were used, the 6-in.-diam column
was never operated in the “slugging’ region. There was,

20. MSR Program Semiannu. Progr. Rep. Aug. 31, 1969,
ORNL-4449, p. 240.

21. MSR Program Semiannu. Progr. Rep. Feb. 28, 1970,
ORNL-4548, p. 307.

22. MSR Program Semiannu. Progr. Rep. Aug 31, 1970,
ORNL-4622, p. 216.
ORNL~-DWG 71~2868A
1000 Y I D N

— T
I — o WATER (0.9¢P)

500 | —— & WATER-GLYCERIN {1.8 ¢cP)
& WATER-GLYCERIN{12.1cP)

[]
[
1

DISPERSION COEFFICIENT (cmZ/sec)

Fa
200 - . e e B
e a& A
- i
100
oX| 0.2 .05 { 2 5 10

SUPERFICIAL GAS VELOCITY (cm/sec)

Fig. 19.13. Variation of dispersion coefficient with super-
ficial gas velocity and liquid viscosity in a 6-in.-diam bubble
column.

however, considerable coalescence of gas bubbles in the
column. Data on the effect of liquid. viscosity were
obtained with water (0.9 cP) and water-glycerin mix-
tures having viscosities of 1.8 and 12.1 cP. No effect of
viscosity was observed.

Dispersion coefficient values measured in a 2-in.-diam
column for a range of liquid viscosity values are shown
in Fig. 19.14. The dispersion coefficient was decreased
by about 20% when the viscosity was increased from
0.9 to 12.1 cP. Dispersion coefficient data measured in
a 1.5-in.-diam column are shown in Fig. 19.15 for
viscosity values of 0.9 and 1.8 cP. Although this effect
is small, there is a definite decrease in the dispersion
coefficient values as the viscosity of the liquid is
increased.

A continuous fluorinator will not use a small-diameter
central gas inlet, as was used in the present ‘studies,
because of the need for protecting the gas inlet from
corrosion by maintaining a layer of frozen salt on the
inlet surface. The gas inlet envisioned for a fluorinator
consists -of one or more large-diameter tubes that are
inclined upward at a 45° angle to the vertical. The
diameter of the tube(s) will be of sufficient size that a
frozen layer of salt can be maintained on the tube
surface without blocking the tube, and a gas-liquid
interface will be present in the tube. This type of inlet
was shown to perform satisfactorily previously in a
system in which molten salt and an inert gas were in
countercurrent flow.2? Dispersion coefficient values
were recently measured with this type of gas inlet in a
3-in.-diam column using air and water. The coefficients
were lower by as much as 10% than values measured
with a single central gas inlet.

23. MSR Program Semiannu. Progr. Rep. Feb. 29, 1968,
ORNL-4254, p. 252,

266

ORNL-DWG 71-2B869A

T T

| o WATER (0.9 cP)
o0 WATER~GLYCERIN (1.8 cP)
4 WATER - GLYCERIN {12.1cP)
100 |— s WATER-GLYCERIN (15.0¢P) =
80 |— ' -

200

:\\
TSX

L~
/‘
T

DISPERSION COEFFICIENT (cm/sec)
fon]
lo]

02 05 i 2 5 10 20 50
SUPERFICIAL GAS VELOCITY (cm/sec)

Fig. 19.14. Variation of dispersion coefficient with super-

ficial gas velocity and liquid viscosity in a 2-in.-diam bubble
column.

ORNL— DWG 71—2870A

——u 1 ] i 17T L ] I T TTTT
@ ® WATER (0.9cP)
o 200 |—
= o WATER—GLYCERIN (1,8 ¢cP) y
L2 HEAVY LINE REPRESENTS BAUTISTA'S o7
= MEASUREMENTS WITH WATER o~ |°
Z 100 — L, §
w . Loy
9 4
& - 2
w50 A
3 A
g .‘/
7] ./)
i 1
W 20 -~
17, L
a e

10

0.2 0.5 { 2 5 10 20 50

SUPERFICIAL GAS VELOCITY (em/sec)

Fig. 19.15. Variation of dispersion coefficient with super-
ficial gas velocity and liquid viscosity in a 1.5-in,-diam column.

In summary, recent studies have demonstrated that
the dispersion coefficient is a strong function of column
diameter in the bubbly and the transition regions. The
dependence of dispersion coefficient on liquid viscosity
increases with decreasing column diameter. The dis-
persion coefficient in a 6-in.-diam column is not
noticeably affected by changing the viscosity from 0.9
to 12.1 cP. The same change in viscosity in a 2-in.-diam
column results in a 20% decrease in the dispersion
coefficient. Since the diameter of continuous fluori-
nators will be 6 in. or larger, the liquid viscosity will not
affect the dispersion coefficient. However, the effect of
viscosity on dispersion in small columns is of interest
because our only data on uranium removal efficiency in
a continuous fluorinator were obtained in a 1-in.-diam
column.

19.9 ENGINEERING STUDIES OF URANIUM
REMOVAL BY OXIDE PRECIPITATION

M. J. Bell L. E. McNeese

Oxide precipitation is being considered as an alterna-
tive to the fluorination—reductive-extraction method
for isolating protactinium and removing uranium from
MSBR fuel salt prior to rare earth removal.?* An
engineering-scale oxide precipitation experiment is
being designed to study uranium removal from fuel salt
from which the protactinium has been previously
removed. In this experiment, 0.3 mole of uranium will
be precipitated from 2 liters (100 moles) of salt in a
single-stage batch system. The source of oxide will be a
steam-argon mixture that will be introduced through a
draft tube to promote contact of the solid phase with
the liquid phase. We will obtain information on (1) the
rate of precipitation of UQ, from fuel salt, (2) the
chemical composition and hydrodynamic behavior of
the precipitate, (3) steam utilization, and (4) the
general characteristics of precipitator operation. We
have planned a subsequent experiment in which multi-
stage countercurrent operation of a UO,-ThO, precipi-
tator will be studied. _

Operation of a precipitator was simulated in a
4-in.-diam glass column (using air, a 50 wt % glycerol-
water solution, and iron powder as the gas, liquid, and
solid phases) to assist in the design of the draft tube for
the vessel. Experiments with this system indicated that
the sides of the precipitator vessel should be tapered
near the bottom to bring the solids under the draft tube
-and that the tube should extend to within a few
millimeters of the bottom of the vessel. An all-nickel
precipitator vessel incorporating these features (see Fig,
19.16) has been designed. Precipitation will take place
in the lower part of the vessel, which consists of a
section of 4-in. low-carbon sched 40 nickel pipe. The
upper part of the vessel is designed to permit deentrain-
ment of salt from the gas stream and is fabricated from
a 6-in.-long section of 6-in. sched 40 L nickel pipe. Gas
is introduced into the system through a section of
Y, -in.-diam nickel tubing that extends to within % in.
of the bottom of the vessel and then flows upward
through a draft tube made of 1-in. sched 40 pipe.

A baffle is mounted on the gas inlet tube to prevent
gross quantities of entrained salt from reaching the
upper part of the vessel. Heat shields are also mounted
on the gas inlet tube to prevent excessive heat loss to
the head of the vessel. The gas inlet tube is fitted with a
" -in. nickel-plated ball valve through which samples of

24. See sect. 17.3 of this report.

267

- ORNL-DWG 71—28T71A
SAMPLE LINE

Y2~in-NICKEL
PLATED BALL VALVE
Ho—HF '
ARGON—H20 >

OFF-GAS

><]

ARGON

DE—ENTRAINMENT
SECTION 6-in. sched
40 PIPE x 6in. LONG

HEAT SHIELD =i

SALT BAFFLE
TRANSFER LINE

CRAFT TUBE i-in.
sched 40 PIPE
x 1o in. LONG

CALROD HEATER\

PRECIPITATOR
SECTION

4-in. sched 40
PIPE x 12 in. LONG

INSULATICN

Fig. 19.16. Schematic diagram of precipitator vessel for the
first uranium oxide precipitation experiment (OP-1).

salt and oxide may be taken. Several methods for
obtaining samples of the solid phase were tested with
the glycerol-water-iron system, and results indicate that
samples of the precipitate can be obtained satis-
factorily. _

The remainder of the system will consist of a salt
feed-and-catch tank; the argon, hydrogen, and HF
supply systems; the argon saturation system; the off-gas
disposal system; and scrubbers for measuring the HF
concentration of various gas streams. The equipment,
now being fabricated, will be installed in Building 3541.

19.10 DESIGN OF A PROCESSING MATERIALS
TEST STAND AND THE MOLYBDENUM
REDUCTIVE EXTRACTION EQUIPMENT

E. L. Nicholson = W. F. Schaffer, Jr.
J. Roth

Very little additional design work has been done
during this report period pending development of
molybdenum fabrication techniques by the Metals and
Ceramics Division (see Sect. 16 of this report for
fabrication development details). No changes have been
made in the conceptual design described in the pre-
ceding progress reports.?®-2¢

Conceptual design sketches were prepared for the
containment vessel nozzles that require transition joints
between the molybdenum tubing inside the vessel and
the external carbon steel or stainless steel process and
instrument lines. Fabrication of these transition joints is
being investigated by the Metals and Ceramics Division.

The full-scale transparent plastic model of the bis-
muth head pot and the top portion of the extraction

25. M. W. Rosenthal et alL, MSR Program Semiannu. Progr.
Rep. Feb. 28, 1970, ORNL-4548, pp. 289-300.

26. M. W. Rosenthal et al., MSR Program Semiannu. Progr.
Rep. Aug 31, 1970, ORNL-4622, pp. 212-13.

268

column was tested with mercury and water to simulate
molten bismuth and salt. Gas-lift performance was
satisfactory over the desired mercury flow rate range of
about 0.05 to 1.1 liters/min. Performance of the head
pot was generally satisfactory in that liquid flow rate
fluctuations in the gas-ift discharge were smoothed
out and a quiet pool of liquid was maintained above the

'metering orifice. However, entrainment of liquid in the

exit gas occurred, the liquid flow capacities of the weirs
in the sieve trays were low, and gas venting between
trays was insufficient. The sieve trays and deentrain-
ment baffles have been modified. The revised design
uses one less sieve tray and will be simpler to fabricate
from molybdenum. The mockup of the head pot and
the column is ready for final testing.
20. Continuous Salt Purification System

R. B. Lindauer

Following the previously reported’ flooding tests
with molten LiF-BeF, (66-34 mole %) and hydrogen or
argon, sufficient iron fluoride was added to the salt to
increase the iron concentration to about 440 ppm. The
first two iron fluoride reduction runs (runs 1 and 2)
were then made with the packed column operating at
700°C. Reasonable values for the mass transfer coeffi-
cients were obtained, as shown in Table 20.1. The
values are about twice those obtained by MIT Practice
School - students in 1968 with a 3.38-in.-ID column.

1. M. W. Rosenthal et al., MSR Program Semiannu. Progr.
Rep. Aug 31, 1970, ORNL-4622, p. 224,

Column operation during these two runs was erratic,
and the pressure drop across the column doubled during
the runs. The increased restriction was believed to have
resulted from precipitation of BeO on the column
packing as the result of oxide accumulation in the
system. S

Sufficient LiF and ThF, were added to the system to
produce a salt composition of 72.0-14.4-13.6 mole %
LiF-BeF,-ThF,, which is approximately the MSBR fuel
carrier salt composition. The newly prepared LiF-BeF,-
ThF, salt was then countercurrently contacted with a
hydrogen—10% HF mixture in the column at 600°C.
The salt and gas flow rates were 103 cm®/min and 5.6
liters/min respectively. Analysis of the column off-gas

Table 20.1. Data from iron fluoride reduction runs

Column temperature, 700°C

Gas flow Salt Analysis of Percent Mass
Run rate flow filtered samples of transfer
No. Date (std. liters/min) rate? (ppm of iron) batch coefficient,
i, AT (cm® /min) Foed Product contacted (ft’/‘;u)
1 7/27 20 100 425 307 71.4 0.018
2 7/28 13.5 100 307 228 72.9 0.016
3 10/22 16.6 100 220 158 75.3 0.015
4 11/3 14.6 210 158 373 84.7
5 11/4 18.2 161 373 137 68.8
6 11/5 4.5 4.5 106 137 110 81.8 0.011
7 11/6 3.9 3.9 142 110 70 81.8 0.030
8 11/9 24.0 103 70 75 86.5
9A 11/13 4.5 4.5 93 75 55¢ 79.4¢ 0.012
9B 11/13 3.5 3.5 136
10 11/17 14.1 105 69°¢ 77¢ 81.8
11 11/19 3.1 4.5 117 v 207 339 81.2
104¢

%The first two runs used LiF-BeF,; the remaining runs used LiF-BeF,-ThF,.

bBased on outside surface of Raschig rings only, since salt is considered to be nonwetting; &, is 78% of these values if entire area

of the Raschig rings is used as the basis.

€Average of flowing-stream samples; no batch samples were taken after the run.

dA value of 100% is used in calculating %, since samples are of flowing-stream type.

€Calculated from flowing-stream samples from previous runs,

269
indicated that 100 ppm of oxide was removed from the
salt with an HF utilization of 15%. Salt samples taken
before and after the oxide removal operation contained
119 and 121 ppm of iron, respectively, indicating that
little’ (if any) of the reduced iron had been present in
the packed column. The hydrofluorination operation
reduced the pressure drop across the column, with a
5-liter/min argon flow, from 12.0 to 10.3 in. H, O.

The extent of iron reduction was reasonable in the
third run; however, column pressure drop became
excessive (37 in. H,0). To provide an opportunity for
oxide in the column to be removed as it was dissolved
and to allow dissolution of the BeF, thus produced, the
column was filled with molten salt and an 88-12 mole %
H,-HF mixture was bubbled through the salt for 18 hr
at the rate of 6 liters/min. Column temperatures of 600,
650, and 700°C were used, and oxide was removed
from the salt at a rate equivalent to about 15 ppm/hr.
The maximum HF utilization was about 5%. This
operation was effective in reducing the column pressure
drop to the initially observed value. The column

270

pressure drop has remained reasonably low during the
subsequent runs.

Our analyses of salt leaving the column in the
remaining reduction runs have been erratic, and the
mass transfer coefficient values are questionable. Iron
analyses of salt leaving the column in four of the
succeeding runs showed an increase in the iron content
of salt, and the iron concentration in two of the .
samples exceeded the total iron concentration believed
to exist in the system (218 ppm). The most reasonable
explanation for the erratic iron analyses is that the iron
particles are small enough to pass through both the 50-u
main-stream salt filter and the 25-u sample filter. Since
the first three runs showed the expected extent of iron
reduction, it is possible that the reduced iron from
these runs saturated the available nickel surface in the
system and that the iron from subsequent runs ap-
peared in varying amounts, along with iron fluoride, in
the salt samples. Studies are under way to resolve these
discrepancies, and further studies of iron fluoride
reduction will be carried out.
271

OAK RIDGE NATIONAL LABORATORY
MOLTEN-SALT REACTOR PROGRAM

FEBRUARY 2B, 1971

M.W. ROSENTHAL, OIRECTOR ]
A.B.BRIGGS, ASSOCIATE DIRECTOR D
P.N. HAUBENREICH, ASSOCIATE DIRECTOR R

H. R, BEATTY.," BUDGET B l
ATERIALS
DESIGN STUDIES COMPONENTS & SYSTEMS DEVELOPMENT INSTRUMENTATION & CONTROLS PHYSICS M MSBR PROCESSING DEVELOPMENT CHEMISTRY
. : . 1. R.WEIR, JR. M&C
E.S.BETTIS | R DUNLAP SCOTT** R S.J. DITTO" 13C A.M.PERRY ] L. E. McNEESE w. A GRIMES® RC
M. 1. LUNDIN R - : :
I | I MSRE PHYSICS HASTELLOY N STUDIES CHEMICAL DEVELOPMENT
J. R.ENGEL R H.E.McCOY** MEC L. M.FERRIS T
. J.C.MAILEN T
MSBR DESIGN STUDIES PUMP DEVELOPMENT PROCESS INSTRUMENTATION j"s ssg'é:?_c" ::g 1 SMITH s REACTOR CHEMISTRY
C.E.BETTIS” GE A.G.GRINDELL™ R A.L.MOORE"® 1&C MSBR ANALYSIS C.E. SESSIONS M&C J.F.LAND cr F. F.BLANKENSHIP RC
W.L.CARTER'" cT P.G. HERNDON"® 18C 0.L.SMITH A G. M, SLAUGHTER"® M&C €. T. THOMPSON cT E. G. BOHLMANN RC
z- 5 z‘ofi:‘:f““ " L. V. WILSON" " R B.McNABB MaC E. L. COMPERE fC
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163.
164,
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171,
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178.
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181.

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183.
184.

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187.

188.
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190.
191.
192.
193.
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UC-80 — Reactor Technology

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274

223. J.Roth 243.

C. E. Stevenson 264. J. R. Weir
224. J. P. Sanders 244. R. A. Strehlow 265. W. 1. Werner
225. H. C. Savage 245. R. D. Stulting 266. H. L. Whaley
226. W. F. Schaffer 246. D. A. Sundberg 267-271. M. E. Whatley
227. Dunlap Scott 247. J. R. Tallackson 272. J.C. White
228. J. L. Scott 248. 0. K. Tallent 273. R.P. Wichner
229. H.E. Seagren 249. E. H. Taylor 274. D. Wilson
230. C. E. Sessions 250. W. Terry 275. L. V. Wilson
231. J. H. Shaffer 251-252. R.E. Thoma 276. G.J. Young
232. E.D. Shipley 253. L. M. Toth 277. H.C. Young
233. W. H. Sides 254. D. B. Trauger 278. J.P. Young
234. M. J. Skinner 255. W.C. Ulrich 279. E. L. Youngblood
235. G. M. Slaughter 256. W.E. Unger 280. F.C. Zapp
236. A.N. Smith 257. D. C. Watkin 281. Biology Library
237. F.J. Smith 258. G. M. Watson 282. ORNL — Y-12 Technical Library
238. G. P. Smith 259. J. S. Watson Document Reference Section
239. O.L. Smith 260. H. L. Watts 283-285. Central Research Library
240. A.H. Snell 261. C.F. Weaver 286—335. Laboratory Records Department
241. Din Sood _ 262. H. O. Weeren 336. Laboratory Records, ORNL R.C.
242. 1. Spiewak _ 263. A.M. Weinberg

EXTERNAL DISTRIBUTION

337. J. A. Acciarri, Continental Qil Co., Ponca City, OK 74601

338. D.T. Ahner, General Electric, Bldg 53, 1 River Rd., Schenectady, NY 12029

339. P. R. Allison, Public Service Co. of Oklahoma, P. O. Box 1, Washita, OK 73094

340. J.S. V. Andrews, Atomic Energy Attache, UKAEA, British Embassy, Washington, D.C. 20008
341. R. C. Armstrong, Combustion Engineering, Inc., P.O. Box 500, Windsor, CT 06095

342. D. M. Axelrod, Public Serv. Elec. & Gas Co., 80 Park Place, Newark, NJ 07101

-343. B. L. Bailey, Great Lakes Carbon Corp., Pine Ave. & 58th, Niagara Falls, NY 14302

344. W.K. Barney, Argonne National Laboratory, 9700 S. Cass Ave., Argonne, IL 60439

345. N. W. Bass, Brush Beryllium Co., 17876 St. Clair Ave., Cleveland, OH 44110

346. R. G. Bernier, Gulf General Atomic, P. O. Box 608, San Diego, CA 92112

347. Gottfried Besenbruch, Gulf General Atomic, P. O. Box 608, San Diego, CA 92112

348. J.M. Black, Northeast Utilities Service Co., P. Q. Box 270, Hartford, Conn. 06101

349. B. E. Blackman, Huntington Alloys, Thelnternatlonal Nickel Co., Huntington, WV 25720
350. J. C. Bowman, Union Carbide Technical Center, P.O. Box 6116, Cleveland, OH 44101

351. R. M. Bushong, UCC, Carbon Products Div., 12900 Snow Rd., Parma, OH 44130
352. R. H. Chastain, Southern Services, Inc., Birmingham, AL 35202

353. C. G. Chezem, Gas & Electric, Wichita, Kansas 67200

354. Gary C. Clasby, Byron Jackson Pump, P.O. Box 2017, Los Angeles, CA 90054

355. Paul Cohen, Westinghouse Electric Corp., P. O. Box 158, Madison, PA 15663

356. D. F. Cope, Atomic Energy Commission, RDT Site Office, ORNL, Oak Ridge, TN 37830
357. L. G. Cook, Esso Research and Engineering Co., P. O. Box 45, Linden, NJ 07036
358. Raymond L. Copeland, Tennessee Valley, Authority, Chattanooga, TN 37401

359. J. D. Corbett, lowa State University, Ames, IA 50010

360. J. J. Costantino, Great Lakes Carbon Corp., 299 Park Ave., New York, NY 10017
361. J. F. Cox, Foster Wheeler Co., 110 S. Orange Ave., Livingston, NJ 07039

362. P. V. Crooks, Atomic Energy Attache, Embassy of Australia, Washington, D.C. 20036
275

363. C. B. Deering, Black & Veatch, P.O. Box 8405, Kansas City, MO 64114
364. D. R. deBoisblanc, Ebasco Services, Inc., 2 Rector St., New York, NY 10006
365. A. R. DeGrazia, USAEC, DRDT, Washington, D.C. 20545
366. Edward Dempsey, Mobil Research & Development Corp., Box 1025, Princeton, NJ 08540
' 367. D.E. Erb, Battelle Memorial Institute, 505 King Ave., Columbus, OH 43201
368. F. V. Fair, Airco Speer Research, 4861 Packard Rd., Niagara Falls, NY 14302 _
369. Martin Fate, Jr., Public Service Co. of Oklahoma, P.O. Box 201, Tulsa, OK 74102
) 370. J. 1. Ferritto, Poco Graphite, P.O. Box 2121, Decatur, TX 76234
371-375. T. A. Flynn, Jr., Ebasco Services, Inc., 2 Rector St., New York, NY 10006
376. J. E. Fox, USAEC, DRDT, Washington, D.C. 20545
377. W. A. Frarks, S. M. Stoller Corp., 1250 Broadway, New York, NY 10001
378. L. W. Fromm, Argonne National Lab., 9700 S. Cass Ave., Argonne, IL 60439
379. A. E. Goldman, UCC, 270 Park Ave., New York, NY 10017
380. B. J. Goulding, Babcock and Wilcox, P.O. Box 1260, Lynchburg, VA 24505
381. W.J. Gray, Battelle-Northwest, P.O. Box 999, Richland, WA 99352
382. Norton Habermann, RDT, USAEC, Washington, D.C. 20545
383. Hans-Jiirgen Hantke, Brown Boveri-Krupp Reaktorbau GmbH, Otto Beck Str. 27, Mannheim, West
Germany
384. J. E. Hard, ACRS, USAEC, Washington, D.C. 20545
385. R.J. Herbst, W. R. Grace & Co., Clarksville, MD 21029
386. Irving Hoffman, USAEC, DRDT, Washington, D.C. 20545
387. A.Houtzeel, TNO, 176 Second Ave., Waltham, MA 02154
388. J.S. Iyer, Pioneer Service & Engr. Co., 2 No. Riverside Plaza, Chicago, IL 60606
389. Ralph Jamieson, Brush Beryllium Co., Cleveland, OH 44110
< 390. S.J. Jaye, Gulf General Atomic, P.O. Box 608, San Diego, CA 92100
391. T. R. Johnson, Argonne National Lab., Argonne, IL 60439
392. W. R. Kannie, General Electric Co., P.O. Box 8, Schenectady, NY 12301
393, H. H. Kellogg, Henry Krumb School of Mines, Columbia Univ., New York, NY 10027
394. L. R. Kelman, Argonne National Laboratory, 9700 S. Cass Ave. Argonne IL 60439
395. E. E. Kintner, USAEC, Washington, D.C. 20545
396. B. W. Kinyon, Combustion Engineering, 911 W. Main St., Chattanooga, TN 37402
397. J. C. Kosco, Stackpole Carbon Co., St. Marys, PA' 15758
398. H. N. LaCroix, Foster Wheeler Corp., Livingston, NJ 07039
399. Kermit Laughon, AEC, RDT Site Office, ORNL, Oak Ridge, TN 37830
400. R.C. Lindberg, Continental Oil Co., 30 Rockefeller Plaza, New York, NY 10020
401. J. R. Lindgren, Gulf General Atomic, San Diego, CA 92112
402. J. M. Longo, Esso Research & Engineering Co., P.O. Box 45, Linden, NJ 07036
403. D. E. Lyons, Combustion Engineering, 14 Office Park Circle, Birmingham, AL 35223

404. W. B. McDonald, Battelle-Pacific Northwest Laboratory, Hanford, WA 99352

405. T. W. McIntosh, AEC, Washington, D.C. 20545 '

406. P. McMurray, Jersey Nuclear Co., 777 106th Ave., NE, Bellevue, WA 98004

407. M. S. Malkin, NUS Corp., 2351 Research Blvd., Rockville, MD 20850

408. A.J. Marie, Philadelphia Elec. Co., Pittsburgh, PA 19105

409. H. E. Marsh, Stellite Div.-Cabot Corp., P.O. Box 746, Kokomo, IN 46901

410. C. L. Matthews, AEC, RDT Site Office, ORNL, Oak Ridge, TN 37830

411. D. N. Merrill, Public Service Co. of New Hampshire, 1087 Elm St., Manchester, NH 03105
412. R. H. Meyer, Northeast Utilities, P.O. Box 270, Hartford, CT 06101

413. A.J.Morse, Continental Oil Co., 30 Rockefeller Plaza, New York NY 10020
414. G. A. Muccini, Ashland Oil Inc., R&D Bldg., Ashland, KY 41101 _
415. C.P. Murphree, Poco Graphite, Inc., P.O. Box 2121, Decatur, TX 76234
= 416. W. Newbury, Combustion Engineering, Chattanooga, TN 37401
417. W. A. Nystrom, Stackpole Carbon Co., St. Marys, PA 15857
418. E. H. Okrent, Jersey Nuclear Co., 777-106 Ave., N.E., Bellevue, WA 98004
|
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ford, CT 06101

419. J. F. Opeka, Northeaét Utilities Serv. Co., P.O. Box 270, Hart
420. R.S. Palmer, General Electric, 310 Deguigne Drive, Sunnyvale, CA 94086
421. F. J. Patti, Burns & Roe., Inc., 320 Fulton Ave., Hempstead, NY 11550

422. F. N. Peebles, Dean of Engineering, University of Tennessee, Knoxville, TN 37976

423. David R. Perkins, United Nuclear Corp., Grasslands Rd., Elmsford, NY 10523

424. A.J. Pressesky, USAEC, Washington, D.C. 20545

425. R. A. Proebstle, General Electric Co., Knolls Atomic Power Lab., P.O. Box 1072, Schenectady, NY 12301

426. Karl H. Puechl, NUMEC, Apollo, PA 15613

427. D. W. Rahoi, International Nickel Co., Inc., Guyan River Rd., Huntington, WV 25720

428. M. V. Ramaniah, Bhabha Atomic Research Centre, Radiological Laboratories, Trombay, Bombay-85
AS, India :

429. David Richman, Research Division, USAEC, Washington, D.C. 20545

430. R. K. Roche, Stellite Division, Cabot Corp., 1020 Park Ave., Kokomo, IN 46901

431. A.P. Roeh, Idaho Nuclear Corp., Box 1845, Idaho Falls, ID 83401

432. W. E. Rosengarten, Philadelphia Elec. Co., Pittsburgh, PA 19105

433. H. M. Roth, AEC-ORO, Oak Ridge, TN 37830

434. Allen S. Russell, Alcoa, Research Lab., P. O. Box 772, New Kensington, PA 15068

435. G. A. Rutledge, Combustion Engr. Inc., 911 W. Main St., Chattanooga, TN 37402

436. R. O. Sandbert, Bechtel, 220 Bush Street, San Francisco, CA 94119

437. J. C. Scarborough, NUS Corp., 4 Research Place, Rockville, MD 20850

438. W. Schrock-Vietor, Kernforschungsanlage Julich, 517 Julich, Germany

439. Fred Schuellerman, Brush Beryllium Co., 17876 St. Clair Ave., Cleveland, OH 44110

440. Hans-Georg Schwiers, Brown Boveri-Krupp Reaktorbau GmbH, Otto-Beck Str. 27, Mannheim, West
Germany

441. R.N. Scroggins, USAEC, Washington, D.C. 20545

442. M. Shaw, USAEC, Washington, D.C. 20545

443. W. M. Sides, Northeast Utilities Service Co., P.O. Box 270, Hartford, CT 06101

444. Sidney Siegel, Atomics International, P.O. Box 309, Canoga Park, CA 91304

445. R. H. Simon, Gulf General Atomic, P.O. Box 608, San Diego, CA 92112

446. E. E. Sinclair, USAEC, Washington, D.C. 20545

447. W. L. Smalley, AEC, ORO, Oak Ridge, TN 37830

448. A.D. Smart, Detroit Edison Co., Detroit, MI 48200

449. E. O. Smith, Black & Veatch, P.O. Box 8405, 1500 Meadowlake, Kansas City, MO 64114

450. T. M. Snyder, General Elec. Co., 175 Curtner Ave., San Jose, CA 95125

451. N. Srinivasan, Bhabha Atomic Research Centre, Trombay, Bombay 74, India

452. A. Stathoplos, Combustion Engineering, Inc., Prospect Hill Rd., Windsor, CT 06095

453. R.C. Steffy, Jr., Tennessee Valley Authority, 540 Market St., Chattanooga, TN 37401

454. C. L. Storrs, Combustion Engineering, Inc., Prospect Hill Rd., Windsor, CT 06095

455. J. J. Sullivan, Charles T. Main, 441 Stuart St., Boston, MA 02100

456. E. J. Sundstron, Dow Chemical Co., Freeport, TX 77541

457. A. E. Swanson, Black & Veatch, P.O. Box 8405, 1500 Meadowlake, Kansas City, MO 64114
458. J. A. Swartout, UCC, New York, NY 10000

459. M. J. Szulinski, Atlantic Richfield Hanford Co., P.O. Box 250, Richland, WA 99352

460. B. L. Tarmy, Esso Research & Engr. Co., P.O. Box 101, Florham Park, NJ 07932

461. R. W. Taylor, Dow Chemical Company, Freeport, TX 77541

462. D. R. Thomas, Commonwealth Associates, Inc., 209 E. Washington Ave., Jackson, MI 49201

463. R. A. Thomas, Southern Services Inc., Birmingham, AL 35202

464. W.N. Thomas, Virginia Electric and Power Co., Richmond, VA 23209

465. Ulrich Tillesson, NUKEM, 6451 Wolfgang, Postfach 860, West Germany

466. R. E. Tomlinson, Atlantic Richfield-Hanford, P.O. Box 769, Richland WA 99352
467. S. N. Tower, Westinghouse Elec. Corp., P.O. Box 158, Madison, PA 15663

468. J. R. Trinko, Ebasco Services, Inc., 2 Rector St., New York, NY 10006

469. V. A. Walker, Detroit Edison Co., Detroit, MI 48200
470.
471.
472.
! 473.
474,
475.
476.
4717.
478.
479.
480.
481.
482-705.

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Ward, Northern States Power Co., Minneapolis, MN 55400

E.C.

C. H. Waugaman, Tennessee Valley Authority, 303 Power Bldg., Chattanooga, TN 37401
R. F. Wehrmann, 10062 Betty Jane Lane, Dallas, TX 75229

M. J. Whitman, USAEC, Washington, D.C. 20545

M. P. Whittaker, Great Lakes Research Corp., P.O. Box 1031, Elizabethton, TN 37643

J. H. Williams, Tennessee Valley Authority, 503 Power Building, Chattanooga, TN 37401
L. A. Wilson, Middle South Services, Inc., P.O. Box 61000, New Orleans, LA 70160

C. E. Winters, UCC, 777 14th St., NW, Washington, D.C. 20005

R.J. Zoschak, Foster Wheeler Corp., 110 S. Orange Ave., Livingston, NJ 07039

L. R. Zumwalt, North Carolina State Univ., P.O. Box 5635, Raleigh, NC 27607

Laboratory and University Division, AEC, ORO

Patent Office, AEC, ORO

Given distribution as shown in TID-4500 under Reactor Technology category (25 copies — NTIS)