ORNL-CF-56-8-208.txt

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! 56-8-208

OAK RIDGE SCHOOL OF REACTOR TECHNOLOGY

REACTOR DESIGN AND FEASIBILITY STUDY “> R

This document has been reviewed and is determined to be ,0 MW FUSE D SA LT H OMOGE NE OUS

APPROVED FOR PUBLIC RELEASE.

Name/Title: Leesa Laymance/ORNL TIO REACTOR POWER PLANT

Date: _10/04/2018

g‘;‘a‘fi;“* gfiiig‘“ fi By magctoriyg 1. ;

-'*’i 5 ;g‘#-;? /' -~ o~ 5/ -2 S /3

R G e S8k inld L[ RILUTZE
T e

Fu: B ¥ Vo, Ve
Lederzimy Motords Haskaw
GREL

NOTICE

This document contains information of a preliminary
nature and was prepared primarily for internal use
at the Oak Ridge National Laboratory. |t is subject
to revision or correction and therefore does not
represent a final report.

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DISCLAIMER

This report was prepared as an account of work sponsored by an
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makes any warranty, express or implied, or assumes any legal
liability or responsibility for the accuracy, completeness, or
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has decument consizts

Mo S 3 o __ G 3 T ———pages,

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2¢ )

L

OAK RIDGE SCHOOL OF REACTOR TECHNOLOGY

‘Reactor Design and Feasibility Study

&

"600 MW FUSED SALT HOMOGENEOUS REACTOR POWER.PLANT"

Prepared by:

R. W. Davies, Group Chairman
" D. H. Feener '

W. A. Frederick

K. Re Goller 4

I. Granet

G. R. Schneider

F. W. Shutko

August 1956

"This document contains rectricted cata as
oo A 4 - A A £ [a

L /s _transoun b e
; 9y manner to an unauinorze
person is prohibited.””

e

B

DISCLAIMER

Portions of this document may be illegible In
electronic image products. Images are produced
from the best available original document.
Date Declassified: March 4, 1957.

LEGAL NOTICE

This report was prepared as an account of Government sponsored work. Neither the
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As used in the above, “person acting on behalf of the Commission” includes any em-
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PREFACE

In September, 1955, a group of men experienced in various scientific
and engineering fields embarked on the twelve months of study which culminated
in this report. For nine of those months, formal classroom and student
laboratory work occupied their time. At the end of that period, these seven
students were presented with a problem in reactor design. They studied it for
ten weeks, the final perlod of the school term.

- This is a summary report of their effort. It must be reelized that, in
so short a time, a study of this scope can not be guaranteed complete or free
of error. This '"thesis" is not offered as a polished engineering report, but
rather as a record of the work done by the group under the leadership of the
group leader. It is issued for use by those persons competent to assess the
uncertainties inherent in the results obtained in terms of the preciseness of
the technical data and analytical methods employed in the study. In the
opinion of the students and faculty of ORSORT, the problem has served the
pedagogical purpose for which it was intended.

The faculty joins the authors in an expression of appreciation for the
generous assistance which various members of the Oak Ridge National Laboratory
gave. In particular, the guidance of the group consultants, E. S. Bettis and
D. A. Carrison, 1is gratefully acknowledged.

Lewis Nelson

for

The Faculty of ORSORT

ACKNOWLEDGMENT

The group wishes to express its appreciation to the many peéple of the
Oak Ridge National Labora£ory wvho gave us the benefit of their experience and
knowledge by supplying advice and information toward the completion of this
desigp study.

Our special thanks go to E. S. Bettis; our group advisor, for his guid-
ance and continuing interest in tfie project.

Finally, we want to thank the ORSORT faculty and staff for providing us
with the education and knowledge which made it possible for us to undertake

this design study.

ABSTRACT

2 The reactor is a fused salt, homogeneous, intermediate energy reactog

- that operates at a power density of 55 watts/cc with an enriched U-235 in;V
vehtory of 240 kilograms. The fuel bearing fused fluoride salt funpfiions
as neutron moderator and heat transfer medium. It 1s cfreulated through the
multi-pass reactor veésel and transfers the fission heat to a sodium loop
within the reactof vessel by an annular, U-tube heat exchanger which surrounds
the central core. This heat is then transferred from this radiocactive sodium
loop to an intermediate, non-radioactive sodium loop, which is used to generate
1000°F superheated steam at 1800 psig to drive a steam turbine-generatpr unit.

The net thermal plant efficiency is 38.2 percent at a maximum net electrical

output of 229 MW.

TABLE OF CONTENTS

Page
Number
CHAPTER 1. SUMMARY DESCRIPTION AND CONCLUSIONS ek 15
1.0.0 Introduction ' 15
1.1.0 Fuel 15
1.2.0 Materials 16
1.3.0 Reactor 17
1.4.0 ‘Sodium Coolant Systems 18 h.5
1. 50 Steam Pofier System l 20 h.é
1.6.0 Conclusions n .1 st il s 20 \ 4.7
CHAPTER 2. FUEL : 23 4.8
2.0;0 Ifitroduction ¥
2.1.0 Composition
2.2.0 Physical and Thermal Properties
2.3.0 Nuclear Properties

2.4.0 Availebility and Cost

2.5.0 Additiz;Hgf Uranium Fuel

2.6.0. ‘Fuel Reprocessing

CHAPTER 3.  STRUCTURAL MATERIALS

3.0.0 Introduction ’

L0 Reaétor Structursal Materialé
3:1.I Inconel . .
3.1.2 Nickel-Molybdenum Alloys
3.1.3 Nickel Clad Stainless Steel

3.2.0 Sodium System Materials

3.3.0 Steam System Materials

'-hp REACTOR ANALYSIS
Introduction
Critical Size and Fuel Concentration

sutron Flux

BAiy

of Reactivity
[ L
ID"'

AT EXCHANGER DESIGN

ary Heat Exchanger
"y
.1.1 Internal vs. External Arrangement
.1l.2 Internal Arrangement
~.l 3 Basic Design Criteria
.1.4 Parametric Study
5.1;5 Desigh Considerations

Reactor Vessel

>

1 Shell Design

2 Internal Arrangement

3 Structural Arrangement

.4t Fuel Circulating Pumps :

5 Pressurizer, Gas Removal and Expansion System
6 Effect of Volume Heat Sources

SODIUM AND STEAM POWER SYSTEMS

-

Introduction
Primary Sodium Coolant Loop

6.1.1 Schemes for Removing Heat from Reactor
6.1.2 Choice of Coolant
6.1 3 Sodium Coolant Activation

*

6.4.0

6.5.0

6.6.0
6.7.0

6.8.0

| .'
.

(cgn;tnued)

"Intermediate Sodium Coolant Loop

l'lfJuatification of Intermediate Loop

2 Choice of Coolant Fluid

3 Sodium-Water Isolation Problem -

4 Proposed Design for Sodium Water Isolation

Intermediate Heat Exchanger

6.3.1 Isolation of Radioactive from Non-
Radioactive Sodium

6.3.2 Calculations for Intermediate Heat Exchanger

Steam Gengratqrs

6.4.1 Once-Through Steam Generator

6.4.2 Natural Circulation Boiler-Separate
Superheater Steam Generator

Steam Generator Calculation Procedures

6.5.1 Heat Transfer
6.5.2 Pressure Drop

Summary of Heat Exchanger Data

Sodium Piping Systems

W2 g,

6.T.1 Pipe Length and Size

6.7.2 Pipe Slope

6.7.3 Expansion Tanks

6.7.% Cold Traps

6.7.5 Valves

6.7.6 Pressure and Instrument Taps
6.7.7 Drain and Charge Tanks

6.7.8 Cover Gas

6.7.9 Sodium Pumps

6.7.10 Heating the Salt and Sodium Loops

for Startup
Steam Power System

6.8.1 Steam Turbine and Steam Cycle Heat Balance

2 Calculation Procedure

.8.3 Auxiliary Power Requirements, Net Plant
Qutput and Efficiency

125

130

T-5.0
9.

Page

Number
PTER 7. POWER PLANT LAYOUT, OPERATION AND MAINTENANCE 133
7.0.0 Introduction , 133
.0 General Piping Arrangement 133
T-1.1 Primary Loop Piping 133
7.1.2 Intermediate Loop Piping 134
7.1.3 Steam Generator Connections 134
T-1.4 Steam and Water Piping 134
2,0 General Arrangement of Plant Components 13k
3.0 Shielding % i g ik 135
4.0 Maintenance b 139
T.4.1 Fuel Circulating Pumps 140
T.4.2 Primary Heat Exehanger Tube Bundles 140
T-4.3 Intermediate Heat Exchangers and Primary 141

Sodium Pumps

.0 Partial Load Operation 141
7.5.1 Mathematical Approach 142
7-5.2 Reactor Simulator Analysis 143
F0 figgnqmi; g 145
T7.6.1 Factors Reguiring Additional Investigation 151
T.6.2 Approximate Economic Analysis 152
hysical and Thermal Properties 15k
esign Curves 171
ieactor Analysis Calculations 176
rimary Heat Exchanger Calculations 189
[ntermediate Heat Exchanger Calculations 201
Steam Generator Calculations 207
falculations of Steam Plant Heat Balance 227

odium Piping Pressure Drop Calculations 234

'flfiyésflip (continued)
S
I. Partial Load Operation (Mathematical Approach)

)

 Reactor Simulator Analysis (Supplied by Dr. E. R. Mann)
.l

e, *Test of Double Tube Sheet Design for Water-Sodium Isolation
.< bols Used in Engineering Calculations
'Tejéia Used in Nuclear Physies Calculations

o'y I

Bi b 11 Qgraphy

¥ e

10.

Page
Number

237

o bm'f 3 2
" 2lgl

251
255
257
258

LIST OF ILLUSTRATIONS

Heat Tranéfer Diagramé - Fuséd Salt_Powér fieactor S&stems
Phase Diagram of the'Three;CofiponentfiNaF-ZrF4¥UFh'System
Phase Diagram of the Two-Component:NgFfszh-System. ’
Par£1a1 Pressureé of ZrF), . | |

Changes in Attack with Increasing Operating Time in
Forced-Circulation Inconel Loops

Corrosion of 85-15 Nickel-Molybdenum Alloy by Fused Fluoride
Salt No. 30 in a Thermal Convection Loop

Corrosion of Nickel.in Fused Fluoride Salt No. 30 in a Thermal
Convection Loop ‘

Variation of Mui%ifilication Constant K with Concentration
of U-235 and Central Core Radius

Total U-235 Invéntory Variation with Central Core Radius
Neutron Flux Distribution

Reactor Arrangement with Straight Heat Exchanger Tubes
Isometric View of Beactor

Prihgry Heat Exchanger Paramétric Study, 1/2-inch Tubes
. Primary Heat Exchanger Parametric Study, 5/8-inch Tubes
Primery Heat ExchangerlParametric Study, 3/4-inch Tubes

. Primary Heat Exchanger Parametric Study Summary, Variation
of Number of Tubes with Tube Ligament Ratio and Size of Tubes

Primary Heat Exchanger Parsmetric Study Summary, Variation

of Pump Power, Fuel Holdup Volume, and Heat Exchanger Surface
with Tube Ligament Ratio and Size of Tubes

Primary Heat Exchanger Tube Layout

Section Through Reactor Vessel

Plan View of Reactor Vessel

11.

Page

Number

19
2>
26
29
39

Lo

b1

50
52
62
63
69
70
T1
T2

T3

15
78

9

¢
*

Figure

Number

5.11
6.1

6.2
7.1
2
T-3
T.l

7.5

7.6

Pressurizer, Fuel Expansion and Gas Removal System

Suggested Test of Double Tube Sheet Design for Isolation
of Water and Sodium

Heat Balance Diagram for 600 MW Reactor-Steam Power Plant
Sodium and Steam Power System

Elevation of Power Plant

Plan View of Power Plant

Time Constants and Heat Capacity of Components in Once-
Through Steam Generator System

Partial Load System Temperatures without Auxiliary Heat
Exchanger Temperature Control

7 +

Partial Load System Temperatures with Auxiliary Heat
Exchanger Temperature Control

T 0 R Qemperature and Reactor Power Transients Due to Rapid

7.8

A.l
A.2
A.3
Ak
A.5

A.6

A.T
A.8

A.9

Increase in Load Demand without Auxiliary Heat Exchanger

" Temperature Control

Temperature and Reactor Power Transients Due to Rapid
Increase in Load Demand with Auxiliary Heat Exchanger
Temperature Control

Al 3

Thermal Conductivity of Selected Steels

Specific Heat of Water

Viscosity and Thermal Conductivity of Saturated Water
Density of Sodium, Potassium and Sodium Potassium Alloys
Viscosity of Sodium, Potassium and Sodium Potassium Alloys

Thermal Conductivity of Sodium, Potassium and Sodium
Potassium Alloys

opecific Heat of Steam at Constant Pressure

Thermal Conductivity of Steam

Viscoselty of Steam

150

155
156

157

Figure

Number
A.10 The
B.l Hea
Tur
B.2 Heal
B.3 Heat
B.h  Tube
C.1l Neut
G.1l Pres
J.l Simu
J.2  Simu]
J.3 Simul
Heat
K-1  Under
K-2 Defec
| Thermal Conductivity of Nickel

- Heat Transfer of Steam or Subsaturated Water in

Turbulent Flow

Heat Transfer from Flat Plates

Heat Transfer from Bare or Insulated Steel Pipe
Tube Count for Tubes Spaced on Equilateral Centers
Neutron Flux Distribution

Pressure-Enthalpy Relation of Expansion Line

Simulator Circuit fgf Fuel and Primary Sodium Loops

Simulator Circuit for Intermediate Sodium Loop

- Simulator Circuit for Driving Function in Intermediate
Heat Exchanger and Regulator for Steam Temperature Control

1
- Underside of Brazed Tube Sheet

i
Defective Brazing of Tube

1k,

LIST OF TABLES

Table Page
Number Number
h.1 Summary of Reactor Analysis Data 57
5.1 Summary of Primary Heat Exchanger Parametric Study s
5.2 Specifications for Fuel Circulating Pumps 88
6.1 Pripisry #eat Exchanger Specifications 11k
combines th
6.2 Intermediate Heat Exchanger Specifications 115 |
reactor des
6.3 Once-Through Steam Generator Specifications 116
reactor and
6.4 Convection Boiler Specifications 117
‘nology wher
6.5 Single Wall Superheater Specifications 118 | :
' S could opera
6.6 Double Wall Superheater Specifications 119
minimum of
6.7 Summary of Full Load Operating Data 132
coolants, h
T-1 Estimated Power Generation Costs 153
been made wi
A.l Thermodynamic Properties of Liquid Sodium _ 165 |
to fulfill 1
A.2 Selected Properties of Stainless Steels 169
ment work wi
A.3 Selected Physical Properties of "L" Nickel 170
into operati
C.1l Basic Microscqpic Cross Section Data Used to Obtain _ 178

3-Group, 34Region Code Constants intended to

c.2 Case Data and Computed Multiplication Constants

cC.3 Macroscopic Group Constants for 3-Group, 3-Region Code

H.l Sodium Piping Data

TABLE OF CONTENTS

SUMMARY DESCRIPTION AND CONCLUSIONS
Introduction

Fuel

Materials

Reactor

Sodium Coolant Systems
Steam Power System
Conclusions

FUEL

Introduction

Composition

Physical and Thermal Properties
Nuélear Properties
Aiailability and Cost
Addition of Uranium Fuel
Fuel Reprocessing

STRUCTURAL MATERTALS

"Introduction

Reactor Structural Materials

Sodium System Materials

Steam System Materials

I}

5.2.0

REACTOR ANALYSIS

Introduction

Critical Size and Fuel Concentration
Neutron Flux

Fuel Inventory

Fuel Burn-up

Uranium-233 Fuel

Fission Product Poisoning

Temperature Coefficient of Reactivifif
Decay Heating |

REACTOR AND PRIMARY HFAT EXCHANGER DESIGN
Introduction |

Primary Heat Exchanger

1l Internal vs. External Arrangement
2 Internal Arrangement

.3 Basic Design Criteria

4 Parametric Study

5 Design Considerations

Reactor Vessglv

Shell Design

Internal Arrangement

Structural Arrangement
Fuel Circulating Pumps

-

N AV AN O
PPPPOND
v FWw o

Effect of Volume Heat Sources
SODIUM AND STEAM POWER SYSTEMS
Introduction

Priméfy Sodium Coolant Loop

6.1.1
6.1.2 Choice of Coolant
6.1.3 Sodium Coolant Activatior

Pressurizer, Gas Removal and Expansion System

Schemes for Removing Heat from Reactor

Page
Number
CHAPTER 6.

6.2.0

6.3.0

6.4.0

6.5.0

6.8.0

O\ O\ O\ O\ O\ O\ O\ O\ O\ O©

(continued)

Intermediate Sodium Coolant Loop

1 Justification of Intermediate Loop

2 Choice of Coolant Fluid

3 Sodium-Water Isolation Problem -

4 Proposed Design for Sodium Water Isolation

Intermediate Heat Exchanger

6.3.1 Isolation of Radioactive from Non-
Radioactive Sodium

6.3.2 Calculations for Intermediate Heat Exchanger

Steam Generators

6.4.1 Once-Through Steam Generator
6.4.2 Natural Circulation Boiler-Separate
Superheater Steam Generator

Steam Generator Calculation Procedures

6.5.1 Heat Transfer
6.5.2 Pressure Drop

Summary of Heat Exchanger Data
Sodium Piping Systems |

Pipe Length and Size
Pipe Slope
Expansion Tanks
Cold Traps
" Valves
Pressure and Instrument Taps
Drain and Charge Tanks
Cover Gas '
Sodium Pumps
Heating the Salt and Sodium Loops
for Startup

L]
o

I, O, RN, IR QN T, T, UK AR,

(-]

=0 - OV W N

O

(-3
o

-] © [
L] -]

-]

Steam Power System

6.8.1 Steam Turbine and Steam Cycle Heat Balance
6.8.2 Calculation Procedure -

6.8.3 Auxiliary Pover Requirements, Net Plant
Output and’ Efficiency

99

100
100
102
102

102
10k

107

107
110

113
113

113
120
120
121
121
121
121
122
123
123

125
125

127
130

CHAPTER 7. POWER PLANT LAYOUT, OPERATION AND MAINTENANCE
7.0.0 Introduction
T.1.0 General Piping Arrangement
T-1.1 Primary Loop Piping
T.1.2 Intermediate Loop Piping
T.-1.3 Steam Generator Connections
T.1.4 Steam and Water Piping
T.2.0 General Arrangement of Plant Components
T.3.0 Shieldirg
T-4.0 Maintenance -
7.4.1 Fuel Circulating Pumps
T-4.2 Primary Heat Exchanger Tube Bundles
T.4.3 Intermediate Heat Exchangers and Primary
Sodium Pumps
7.5.0 Partial Load Operation
T.5.1 Mathematical Approach
T.5.2 Reactor Simulator Analysis
7.6.0 Economics
7.6.1 Factors Requiring Additional Investigation
7.6.2 Approximate Economic Analysis
APPENDIX
A Physical and Thermal Properties
B. Design Curves
C. Reactor Analysis Calculations

—

D. Primary Heat Exchanger Calculations

E. Intermediate Heat Exchanger Calculations
F. Steam Generator Calculations
G. Calculations of Steam Plant Heat Balance

H. Sodium Piping Pressure Drop Calculations

13.

Page
Number

133
1133
133
13
134
134
134
134
135
139
140
140
141
1

142
143

145

151
152

15k
171
176
189
201
207
227

234

APPENDIX (continued)

I. Partial Load Operation (Mathematical Approach)

J. Reactor Simulator Analysis (Supplied by Dr. E. R. Mann)

K. Test of Double Tube Sheet Design for Water-Sodium Isolation
Symbols Used in Engineering Calculations

Symbols Used in Nuclear Physics Calculations

Bibliography

1k,

Page
Number

237
2l
251
255
257
258

k4

15,

CHAPTER 1.

SUMMARY DESCRIPTION AND CONCLUSIONS

1.0.0 INTRODUCTION

This report is a study of the feasibility of using a fused salt fuel
reactor in a central station electric generating plant. The propoBed reactor
.combines the desirable features of various homogeneous and high temperature
reactor designs. The basic philosophy of the study has been to design a
reactor and pbwer plant system which could be built using'present day tech-
nology wherever possible; whicfi would be reliablé, safe, and efficient; which
could operate for long periods of timé; and which could be maintained with a
minimum of difficulty. The selection of the fuel, reactor materials, reactor
coolants, heat transfer system, steam turbine and associated equipment has
been made with the above objectives in mind, and the resulting design appears
to fulfili these conditions. This gtudy indicates, however, that some develop-
ment work will be necessary beforefifi plant of this type can be built and put

into operation. The basic plant arrangement presented in this report is not

| intended to represent a finished or optimum design, but it does appear to be:

reasonable and possible and warrants further study and consideration.

1.1.0 FUEL

The plant has been designed to utilize current technology in the deyelop-
ment of fused salt fuels. A reactor experiment has been conducted successfully
using a solution of sOdifim, zirconium and uranium fluorides for the fuel system.
This experiment indicated that a fused salt fuel of this type has the char-
acteristics required for the homogeneous reactor under consideration. The

moderating material in the reactor is the fluorine in the fused salt. With the

16.
composition contemplated in this design, the number of fluorine atoms is in-
sufficient to thermalize the majority of neutrons before fission takes place.
As a result, approximately 34 percent of the fissions occur in the intermediate
and fast energy ranges. |

A fused fluoride composition of 43 mol percent Zth and 57 mol percent
NaF was selected as the fuel base to obtain a compromise between a reasonable
fielting point (9329F), corrosion, and the ZrF), "snow problem". The concen-
tration of the uranium is low and has little effect on the over-all characteristics
of the salt.

Uranium is added to the salt in the form of (NaF),UF). Pellets, powder
or & dissolved solution of the compound can be added during operation to
replace uranium burn-up and to overcome the build-up of fission product and
corrosion poisoné° Economically, it probably would not be desirable to repro-
cess fuel to remove the solid fission products in order to minimize the amount
of uraniwm which must be added to override the effect of these poisons.

1.2.0 MATERIALS

Fused salts are extremely corrosive to conventional engineering struc-
tural materials. It has been shown that alloys containing large percentages
of nickel are the most corrosion-resistant metals to fused fluoride salts.
Inconel, nickel-molybdenum alloys and pure nickel were considered as possible
structural materials for the reactof.

It has been shown that the chromium in Inconel diffuses from the metal!
at a fairly substantial rate and, therefore, Inconel does not appear $0 be a
satisfactory structural material for a long life reactor.

Nickel-molybdenum alloys have been shown to have excellent corrosion

‘resistance to fused fluorides. A series of these alloys is being developed

e

17.
in an attempt to obtaifi a strong, fabricable material for use in high tem-
perature fused salt reactors. It appears that this development program will
be successful, but at the time this report was prepared, the exact properties
and fabricability of these new alloys were not sufficiently established to
presuppose a satisfactory reactor design.

The amount of available data on the corrosion of pure nickecl in fused
fluorides is limited; however, these data indicate little or no corrosion or.
mass transfer. On this basis, it appears that nickel has the corrosion resis-
tant properties for long reactor life. Unfortunately, at the operating
temperatures contemplated, the strength characteristics of nickel are inadequate,
and the nickel must be strengthened by suitable.structural materiglsa There-
fore, to present a design using present day technology, nickel clad stainless
steel was selected as the most suitable structural material for the reactor
in this study.

Because of the limited experimental data available, there is still some
question as to the suitability of nickel to resist corrosion to the degree
desired. The use of nickel cald stainless steel also poses fabricating and
inspection problems.

1.3.0 REACTOR

The reactor is a homogeneous circulating fuel reactor with the primary
heat exchangers an integral part of the reactor vessel. The heat exchangers:
are located in a baffled annulus surrounding the seven-foot diameter cylin-
drical core. Fuel flows upward through the core, through the U-tube hgat
exchafigers, énd through an annular downcomer at the periphery of the vessel.
Eight vertical axial flow pumps located symmetrically around the top of the

vessel maintain the fuel circulation in the reactor.

18. .

The reactor is designed to produce 600 MW of heat with fuel inlet and
outlet temperatures of 1050°F and 12009F, respectively. The pumps and the
entire reactor vessel, with the exception of the top, are subjected to the '
minimum fuel temperature.

The reactor vessel is a sixteen-foot diameter vertical cylinder with a ¢
dished bottom and protrusions near the top where the fuel pumps are located.
The only penetrations in the vessel are a series of drain connections in the
bottom; all other fuel and coolant 1inés enter the top of the vessel. The
heat exchanger baffles are supported internally, and the heat exchanger bundles
are supported from a circular cantilever arrangement extending radially inward
at the top.

The heat exchangers and pumps are arranged to permit removal.vertically
from the reactor. The layout of the equipment was arranged to provide for
replacement of components with a minimum of difficulty, since outages for
maintenunce should be nf short duration.

Provisions are made for the continuous removal of xenon, and space is
provided, external to the reactor, for the expansion of fuel. The reactor
will have a negative temperature coefficient of reactivity which will furnish
the only means of reactor load control. The design average temperature will
be controlled by the proper addition of uranium to the fuel solution.

"1.4.0 SODIUM COOLANT SYSTEMS

The sodium system is made up of two independent coolant loops. A heat
transfer diagram which illustrates the temperature drops through these loops
is shown in Figure 1.1. The sodium in the primary loop flows from twenty;
four heat exchanger bundles located in the reactor vessel to a common ring .

header above the reactor. Six independent circuits of the primary loop are

ORNL=LK=Dwg,=18169

19.
Figure % |
; HEAT TRANSFER DIAGRAMS
: FUSED SALT POWER REACTOR SYSTEM
1200 Fuel
Sodium o = S
3 Primary
, SRS 7 {Heat Exchanger
1100 Sodlum ~ N~ . o e
N SRR £ 48.0 x 106 1b/hr.
- i h-NEENEE NN
Water- . 45.7 x 100 1b/hr
1000 Steam A\ N 14
‘ \. - o e = e ¢
\ ‘\\( Intermedliate
g "Heat Exchanger
N \\
900 N\
/ N
HF s /
- Fx
:: i ; / N
ol ] A
1.2 T \
+ 5 T 800
T \ ! _ .
W [T Steam A
— \ Generator ]
SR L\ 15.9 x 10 1b/ar ]
- \\ o AT =
700 \‘
\
6003 \
\
\| 1.96 x 10 1b/nr
500
400
* Percent of Heat Transfer

20,
connected to the ring header. This division is used so that piping and pump
sizes in the primary loop are reasonable and practical. The six intermediate
heat exchangers are located in shielded compartments surrounding the reactor
shield.

There are two intermediate loop circuits, each connected to three inter-
mediate heat exchangers and one steam generator. Superheated steam from the
two steam generators is used to drive a steam turbine.

1.5.0 STEAM POWER SYSTEM

Fused salt reactors are especially attractive for central station appli-
cation because of their high temperature characteristics. In this study, the
steam turbine and associated equipment have been selected for operation with
1800 psig and 1000°F steam. A 250 MW tandem compound triple flow turbine will
be connected to a 3600 rpm conductor cooled generator. Five closed feedwater
heaters and a deaerating heater are included in the feed water heating cycle.

At raled load, the gross generator output is 245,400 kw, the plant
auxiliary power requirement is 16,400 kw, and the net thermal efficiency is
38.2 percent.

1.6.0 CONCLUSIONS

This design study of a fused fluoride salt reactor central station power
plant has shown that the proposed system is technically feasible. It is con-
sidered that such a plant can be built with a minimum of development work, and
that the developments would primarily be associated with the reactor.

The concept of using a fused fluoride salt as a homogeneous self-moderating
fuel solution has been established as practical. It has been shown that a large
fraction of fissions are caused by epithermal neutrons, which classifieg this

design as an intermediate reactor.

21.

No completely satisfactory materials are currently‘available for the con-
struction of the reactor vessel. Nickel clad stainless steel has many of the
necessary requirements; but fabrication and inspection techniques must be
improved. A series of nickelamolybdenufi alloys are in the process of develop-
ment which appear to have excellent corrosion and high temperature strength
properties. The success of the fused fluoride salt reactor may depend upon
satisfactory completion of this metallurgical development program.

The high temperature characteristics of the fused fluoride salt reactor-
make it possible to use a high efficiency steam turbine cycle for the con-
version of fission heat intoc electrical energy. The net thermal efficiency
of the proposed design is 38.2 percent, and it is possible to design for even
a higher efficiency if the over-all plant economics show that this is desirable.

The operation of the reactor is relatively simple. The reactor power
iS'controlled by its negative temperature coefficient of reactivity. Uranium
will be added to the fused fluoride salt to replace fuel burn-up and tc over-
ride the build-up of fission product poisons. Continuous fuel reprocessing
should not be necessary for extended periods of operation. |

The reactor has been deésigned to burn uranium-235. However, sincg thig
is an intermediate reactor, a number of advantages would result if uranium-233
would be bred during operation. The production of new fissionable material
vould reduce the fuel addition rate. The fuel bred ip the reactor could be
utilized without the neceséity of employing chemical separation techniques. ..

The mechanical design features of the reactor are novel. The reactor
vessel is & multi-pass cylindrical container. The primary heat exchanger
tube bundles, which are located inside the reactor vessel, and the fuel cir-

culating pumps can be removed for replacement by lifting them vertically

22,

from their sufiportan The critical components of the reactor and primary
sodium coolant circuits can be maintained with a minimum of difficulty. All
components in the system, with the exception of those inside the reactor -
primary shield, can be maintained directly.

It is the considered opinion of the authors that the fused salt reactor -
not only shows great promise as the heat source for a large central station
‘power plant, but that it is one of the most desirable reactor types for

this application.

23.

CHAPTER 2.

FUEL

2,0.0 INTRODUCTION

The wmost.outstanding advantage of fising a fused salt as the fuel solvent
and heatv exchanging medium is that it enables the attainment of high tempera-
tures without high pressures. 1In addition; the advantaggs inherent in homo-
geneous systems are possible; i.e., no neutron absorbing structural material
in the core proper; possibility of fission product removal, with the important
gaseous fission products {xenon) being especially easy to remove; nc fuel
element fabrication; extended periods of operation possible by the periodic
addition of a fuel conéentrate; and high permissible fuel burn-up.

The particular fused salt used should have a reasonable melting point

and must be stable at high temperatures and under high radiation. The

elemental constituents of the salt should have small neutron capture cross
sections. As in the cagse of the reactor in this study, it may be required
that the fused salt solvent also serve as the neutron moderator. In addition
to these specialized requirements, the salt must meet the more general con-
siderations of availability, reasonable cost, permissible toxicity, and
tolerable corrosiveness.

A good deal of thepretical and experimental work has been, and is.still
being done, by other groups in this field to select the constituents and the
composition of a fused salt that best meets the above or very qimilar require~
ments. The result of this work indicates that a fused salt composed of NaF,
ZrF), and UF) is satisfactory for use as a liquid fuel for reactors. A reactor -

has been successfully operated using this fuel system. After considerable

2L,
literature‘research and discussion with Oak Ridge National fiafioratory per-
sonnel; it was considered that this particular salt system did fulfill the
design requirements for the proposed reactor and also that this was currently
8till the most desirable system. By utilizing this salt system, there was
immediately available a considerable amount of experimental information,
physical properties data, and practical experience necessary for the intelli-
gent design of a plant using such a salt as the reactor fuel solvent.

2.1.0 COMPOSITION

Within the NaF -ZrF) -UF) system; there remained to select the optimum
composition of these constituents for this design study. Figure 2.1 shows
an isometric view of the temperature-composition-phase plot for the three
component NaF -ZrF) -UF), system. Nuclear calculaticns indicated that the frac-
tion of UF) required is small (<0.2 mol %), so that for many purpéses the
small salt system may be considered as a two-component system of NaF -ZrF) .
Figure 2.2 shows a plot of the temperature—composifiion-phase diagram for the
two-component NaF-ZrF) system. From both the two-component and three=component
plots; it can be observed tha£ there exist two relatively low melting eutectics
in the low UF), content region of interest. These occur at about 42 and 59
mol % Zr'Fh° In order to obtain the minimum melting point, it would be desirable
to specify the salt to be one of these eutectic compositions; however, there are
other factors that need to be considered in the composition selection.
Operational experience with NaF-ZrF) fused salt syétema revealed the
existence of a "snow problem" resulting from the low, but finite, vapor pres-
sure of ZrF) at operating temperatures. A simplified explanation of this
pPhenomenon is that ZrFu-vapor will escape from the fused salt to any existing

gas volume within the fused salt system. The subsequent sublimation of ZrF)

-25-

ORNL-LR-DWG 3398A

NaF

COMP. NO, 34
43 mol Y% ZrF4

Phase Diagram of the Three-Component
NaF -ZrF,-UF, System.

Figure 2.1

TEMPERATURE (°C)

_gz_

1000
ORNL-LR -Dwg. -12}J1A
N\ .‘ 7
o \V//’ //////
s /
| &% /
=
- O
\ %E | /
600 \\IJIQ //
Y N
S N
No3ZrF7—// i NaZrFg (METASTABLE)
400 - | il
Na,ZrFs. +— | I
O {O 20 30 40 50 60 70 80 90 {00

ZrF,4 (mole %)

rigure 2.2 Phase Diagram of the Two-Component NaF —ZrF, System.

I
I

27.
crystals on lower temperature surfaces results in the formation of snow-like
Zth crystals which can build up and accumulate to such an extent that plugging
of tubing or even filling the entire gas volume may result. This most undesir-
able phenomenon can be significantly imfroved by reducing the fraction of ZrFu
in the fused salt composition.

Other factors are in opposition to shifting the fused salt composition
in this direnfion" For nucleur considerations, it is desirable to have the
highest fraction of fluorine atoms possible in the fused salt solvent. .The
Zth_salt constituent has four times as many fluorine atoms for each metal
atom as the NaF constituent. Further, the corrosiveness of the fused salt is

decreased by increasing the fraction of ZrF) and decreasing the NaF. These

- factors both suggest having as high a mole percentage of Zch a8 possible.

For this design study, the relatively slight effect on corrosion and

nuclear properties by the small changes in composition that are possible were

‘considered of lesser consequence than the more significunt effect these small

changes in composition have on the "snow problem". For these reasons, then,
it is desirable to utilize a fused salt of composition close to the lower ZrF)
content eutectic (41 mole percent).

An examination of the temperature-composition-phase diagrams shows that
Just below the 41 mole % Zth eutectic, a large increase in melting point -
temperature results with only a slight_decrease in ZrF), composition. Such a
decrease would quite probably occur during operation, since the uranium fuel
concentrate is added as the complex salt (NaF),UF,. To safeguard against the
poesibility of the fused salt composition falling into this region after a

period of reactor operation, it was considered advisable to accept a very

slightly higher fraction of ZrFu than the exact eutectic. For these considerations,

28.
an already standardized composition, Composition No. 34 (Ref. 31) was there-

fore selected.

Comp. number Component Mole % Weight Liquidius Temp.
34 NaF 57.0 2l ,98 500°C, 932°F
ZrF), 43.0 75.02 -

2.2.0 PHYSICAL AND THERMAL PROPERTIES

The physical and thermal properties of this salt composition are as
follows (Ref. 17): |
fiensity:
3.86 gm/cc at room-temperature
€ (gm/cc) = 3.65 - 0.00088 x (°c) * 5%
€ (1v/£t3) = 228.8 - 0.0305 x (°F) * 5%
Heat Capacity:
0.27 cal/gm - °C at 700°C
Thermal Conductivity:

1.3 BTU/hr-£t-°F

Viscosity:
600°C =-mcnemcan 7.5 cp
(Ol — 4.6 cp
8009C --c-cemmeaa 3.2 cp

A good indication of the pfobable degree of the "snow problem" discussed
briefly above 1s given:by the vapor pressure of ZrF) that exists at the tem-
perature of operation. Figure 2.3 shows the variation of the partial pressure -
- of ZrF) with temperature for various compositions of the NaP -2xF), éystem (Ref. 26).
An interpolation of the plotted data to the 43% ZrF), compésition and extrapola-

tion down to the design operating temperature (1050°F, 565°C) of the fused salt -

1000 T - S
l h
10C N
748 N
CNIN
o O ‘*
100 891 T
13 — £Y7 W .
419 oI\ T SOO\\\ ‘:\‘
'&,\ < \\ N *\1‘ N \
N\ T \:\ "~ .
10— ~ ~C ~ T~ ~
s N I BN q
N R CNERN \\i\\.\ N
- N NN
E i \\ M ’\\ ~ DN
£ S ~ - - X
‘w - N S AN N
a A i e N j‘\\
o 283 \ ~ C% Pa
' <3 N O'S'-Z‘é .
N [ 207 Joy
X EONSENS
¢ ‘\|95 4 2 ™~
0.01}—x : : % -
-
\:\ ~
°-tfius
00— =8 80 82 ¢ B¢ 88 9C 97 97 9% 38 00 102 104 106 108 110 112 1€ 16 N8
’ : Reciprocal Tempaerature, -T'-?K—]
| : 1 e | b
1050 1000 900 800 ) TO00 600
" Temperature, C 565 C
1050 F

- FIGURE 2.3 PARTIAL PRESSURES OF Zrf, BASED ON ASSUMPTION THAT
ONLY Nof AND Zrf, EXIST IN VAPOR PHASE

Figures on plot denote mole per cent ZrF,.

Source: BMI -~ 1064; "Vapor Pressures of the Sodium Fluoride -
Zirconium Fluoride System ard Derived Information", by Karl A,
Sense, et, al. '

0LTQT= *3MI=HT~-TNIO

‘62

30.
—
at the location of the gas surface shows the partial pressure of Zth to be
less than 0.01 mm mercury. It was the considered opinion of personnel experienced
in this field that a "snow problem” would be virtually non-existent for Z1F),
vapor pressures less than 0.0l mm mercury, even for extended periods of operation
of the order of years (Ref. 18). In addition, mechanical appliasnces called
"snow traps" have been developed which prevent plugging of the critical areas,
and these, or similar innovations; could be adapted to this design if required.

2.3.0 NUCLEAR PROPERTIES

The nuclear characteristics of the NaF-ZrF) salt are not outstanding, how-
ever, they are tolerable. The absorption croes sections, particularly of the
sodium atom for thermal neutrons; are quite large. Neutron moderation is
primarily performed by the fluorine atom which, with an atomic weight of 19,
is rather poor in this respect. Alternate salts could be suggested (i.e., BeFp)
with superior nuclear properties to those of the NaF -ZrF) salt, but other con-
siderations such as corrosion, viscosity, stability; melting point, etc.; make
the selected Na¥F-ZiF) sult the most demirable at this time from an over-all
consideration.

2.4.0 . AVAILABILITY AND COST

The availability and cost of this salt were investigated (Ref. 19). The
NaF constituent is commercially available at $0.12 per pound. Hafnium separated
- ZrF), has been produced by Oak Ridge National Laboratory from Hf separated ZrCl)
at a total cost of $3.15 per pound. The NaF and ZrF) salt constituents are
blended as powders and then fused in a batch process at 1500°F to form the com-
position desired. This fused salt must then be processed to remove trace quantitiés
of sulfur, iron, nickel,‘EQO, oxides, chlor;des, and other impurities for corrosion
pufposes° This treatment is performed at 1500°F with hydrogen and hydrogen fluo-
ride'gaa° At the present time, ORNL can produce a hafnium-free, processed

NaF-ZfiFu fused salt ready for use as a reactor fuel solvent for approximately

31
$7-.50 per pound. It is the considered opinion of the responsible Oak Ridge
National Laboratory personnel that increased capacity and technical develop-
ments would very probably reduce this cost to $5.00 per pound, or less. -

The nuclear calculations for this project were based on the assumption
that the elemental zirconium constituent of the salt is pure, the hafnium
impurity having been removed. This has the effect of minimizing the uranium
inventory by decreasing parasitic absorption at the expense of an increase in
the cost of the salt.

2.5.0 ADDITION OF URANIUM FUEL -

-The uranium fuel burn-up is compensated by the periodic addition of a
uranium concentrate—(NaF)EUth From the three constituent phase diagram
(Figure 2.1), it is observed that as the uranium concentrate compound is dis-
gsolved in the composition number 3l salt solvent, only compositions of con-
sistently lower melting points are formed in the dissolution process. The
dissolution should, therefure, proceed without difficulty.

The physical addition of the uranium concentrate can be accomplishéd in
several ways. The concentrated uranium salt could be added directly to the
fused salt in the form of pellets. It could be added more or less continuously
as a powder, or it could be dissolved in a small, isolated quantif& of the
fused salt solvent in-a continuous or batch process and then injected into‘the g
reactor. The final selection oflthe preferred methodlwould need to be deter-
mined by a consideration of the hazardé involved and the ease of operation.

The quantity or rate of fuel concentrate addition is controlled by the
average temperature of the fused salt in the reactor core (1125°F design).

As fuel is burned up, this average temperature would be gradually reduced by
the reactor's negative temperature coefficient, which keeps the reactor "just

critical”. When this temperature falls below 1125°F by some predetermined

) 32.
temperature increment, a carefully controlled quantity of fuel concentrate
would be added or, if this is a continuous process, the rate of addition
would be increased. The average temperature would thereby be raised a regu-
lated increment above the average design temperature and the cycle would
begin again.

2.6.0 FUEL REPROCESSING

Fission products and corrosion products will accumulate in the fuel solu-
tion during operation. Additional uranium will have to be added to override
the poisoning effect of these nuclear poisons if they are allowed to remain in
the reactor. For this reason, it may be necessary to provide a means of fuel
reprocessing to remove some of these poisons.

8

Xenon has a very low solubility (n{lO moles/cc @ 1 atmos. of xenon) in
thg fused salt fuel solution and will therefore be removed by the off-gas system.
The gaseous iodine fission product which is the precursor of xenon will also be
removed by the off-gas system. If this is accomplished before the iodine;, which
has a 6.7-hour half-life, decays to xenon, the reuctor core will operate with
little or no xenon present.

Some of the metallic fission products (i.e., Ru, Mo) do not form fluoride
compounds and will not dissolve in the fused salt solvent. They will probably
plate out on the surfaces within the reactor. This has the advantage of placing
these nuclear poisons in regions of low flux. The effect of these plated metals
on corrosion and heat transfer will require further investigation.

The leasgt. soluble of the fluoride forming fission products is believed to
be that of the rare earth, CeF3, vhich is soluble up to 3 wt. %.At the tempera-

ture of this design study.(Ref. 32). Approximately one-third of the fission

products resulting frqm,fission are rare earths.. Even by making the pessimistic

33.
assumption that all the rare earths must be less than 3 wt. % in order tpgt
they will not precipitate, the solubiiity limit of the rare earth fluori@es
would not be reached for many years of operation (>20 years).

- The conclusion that may be derived from this is that the fused salt
would need to be reprocégsed for other reasons long before there would be
'any possibility of precipitating these fission products.

The nuclear poisoning of the dissolved fission products would eventually
cause the uranium fuel inventory required to attain criticality to become
excessive. A discussion of this effect is given in the chapter on Reactor
Analysis, Section 4.5, and Appendix C.6 and C.7.

Although it is expected that improved reprocessing techniques will be
developed in the near future, present technology would restrict fuel repro—.
cessing to the following method. Contaminated fused salt would be drawn from
the reactor system and the uranium recovered by the UF6 volatility process.
The uranium-depleted, fission product-contaminated fused salt solvent would |
then be discarded to a hot-waste storage system. The salvaged uranium wquld‘

- be dissolved in fresh, pure salt and returned to the reactor system. This
same salt might also provide a convenient solvent for the uranium which must
be added because of burn-up.

A possible alternative which might be developed would be to treat the
uranium-depleted, contaminated salt to remove some of the fission prohucts,,
This would permit re-using the fused salt, thereby eliminating the consider- -

- able expense of replaging the fused salt, and the problem and expense of
storing the radioactive, contaminated salt.

This treatment process might possibly be a solvent extraction type pro-

cess, using fused salts, or it might be accomplished by cooling the salt to

Just above its freezing temperature and precipitating some of the fiesion products.

3k,

The economically optimum extent of fuel processing is determined by the
relative costs of increased uranium inventory and the costs of processing.
Based on the currently available processihg method outlined above; an economic
analysis to determinelthe optimum length of processing cycle time was made
and is outlined in Appendix C.6 and C.7. Based on the assumptions made in “
this analysis, the reactor should be operated at a uranium inventory-fission
product equilibrium that would replace all the salt in the core system by a
continuous process over a 19-year period. This means that unless a superior

method of fuel reprocessing can be developed, it would be economically desir-

able to perform no fuel reprocessing with the lifetime of the plant (20 years). g

CHAPTER 3.

STRUCTURAL MATERIALS

3.0.0 INTRODUCTION

The major metallurgical consideration of a fused fluoride reactor is thg
choice of a structural material for the reactor vessel and primary heat exchanger
vhich will resist corrosion, be fabricable, and have satisfactory strength pro-
perties at elevated temperatures. The nuclear characteristics of the structural
materials in this reactor are not critical. Materials problems associated -with
the sodium cooclant system and the steam power plant are important, but previous
experience is available; therefore, the design_of these systems introduces no
serious problems. |

3.1.0 REACTOR STRUCTURAL MATERITAILS

The power plant design presented in this study is proposed for central
station electric generation, and it would be desirable to have a reactor which
would have an operating life eqfialAto the 20 or 30-year life expected from the
steam plant equipment. Present technology is not sufficiently advanced to in-
dicate that such an extended operating time can be obtained from fused.fluoride
reactors, but it is mandatory that a material be obtained which will resist fluo-
ride salt corrosion at elevated temperatures to the extent that these reactors
can be considered for this type of operation.

A homogeneous reactor of the contemplated design does not have as many
materials problems as other types of reactors. The fuel is in solution and

is self-moderating, and there are no fuel fabricating, cladding or moderating

material problems. The reactor vessel is a low pressure container and does not

have to be designed for high pressure operation. Since there are no structural

materials in the high neutron flux region of the reactor, the effect of the
nuclear characteristics of these materials on neutron economy is of seconda
interest.

Fused salts, however, are considerably more corrosive on materials that
many other reactor fluids. Experiments have shown that nickel base alloys |
the best corrosion resistance to fused fluorides. Inconel, nickel-molybdes
alloys and pure nickel have been considered as possible structural material
this reactor design.

3 1.1 Inconel

There: afe considerable data on the corrosion of Inconel in fused fluorl
at temperatures in excess of those contemplated in this design. An exmqflr
Inconel corrosion is shown in the photomicrographs in Figure 3.1 which ilh
trates the time dependence of the attack° Inconel corrodes in fluoride sal
by a leaching of chromium frofi the surface of the metal. New chromium the|
diffuses from the interior of the metal from which it in turn is leached.
voids remain and concentrate at the grain boundaries; these voids are not |
nected to each other-or to the surface.

Corroqiofi data indicate that after the first 50 hours, the(depth of
appearo to be linear, with a rate of between 3 and 4 mils per 1000 hours
operation. The photomicrographs and corrosion data were obtained from fo
circulation loops with 1500°F maximum fluid bulk fluoride temperature, a
1600-1620°F maximum wall temperature, 200°F temperature gradient, and 10,
Reynolds number. The fluid circulated was Fuel No. 30 (Ref. 48, p. 97).

There is also an initial impurity attack of approximately 3 to 4 mil
which continues until the UF) -Cr reaction reaches eqpilibriuo and the chr

content in the fluoride mixture remains constant (Ref. 48, p. 96, 97). T

37

tion will take place with each additional new charge of fluoride
fore, the long.range corrosion rate will depend upon the number

it the fluoride salt is changed.

psion resistance of Inconel to fused fluoride salts is not satis-
} reactor of this design.

kel -Molybdenum Alloys

nown Mickel-molybdenum alloy, Hastelloy B, has shown excellent
fstance to fused fluoride salts (Ref. 50). This alloy, however,
actory for reactor use because of its low ductility at elevated

and because of its undesirable fabricating characteristicso

el -molybdenum alloys are currently being developed which are in-
tain the negligible corrosion rates, but which will have satisfactory
ure and fabricating properties.

ogion resistance of an 85-15 nickel-molybdenum alloy in a fluoride
jermal convection loop is illustrated in Figure 3.2. These photo-
jhow that there is essentially no corrosion or mass transfer in

L a temperature of 15009E. Certain of these new alloys appear to
.*ary'physicair;:zperties for reactor use; however, it was felt
opment program is not sufficiently advanced to completely sub-
expected excellent properties. These new alloys were not

d this reactor study because of the philosophy originally estab-

bit a reactor design using present day technology wherever possible.

fckel Clad Stainless Steel
opears to have excellent corrosion resistance to fused fluoride
iic tests of nickel exposed i@ fltuoride:fuel.No. 30 Toryl00.Hours at

own no measurable attack (Ref. 49). A small amount of mass transfer

was * ."Ied-:.m the f;l.'rs-t 'r..lickv.al -thefmal"convection loop in which Fuel No. 30
u“flabjcirculatéd (Ref. 46, p. 63, 64). To confirm this single test, another
:niékéi'loop was cleaned with Nafosfiénd ogerated with high pfi}ity fuel. No
evidence of subsurfaceavoids,‘infergranular attack, dr mass fransfer’wés
observed (Ref. 47, p. 73;;7h)., Thé“phétomicrOQraphs'in figure 3.3 show the
absence of these effects in a nickel thermal convection loop operated for
500 hours at 1500°F with Fuel No. 30.

No convection or forced circulation loops have beern constructed which .
i1lustrate serious mass transfer properties of nickel, although the opinion

has been expressed that nickel theoretically should mass transfer.

Nickel does not have adequate strength to be used as a structural
material; howéver, nickelqca# be clad satisfactorily on stainless steel. This
duplex material will have the required cor?osion resistance and physical pro-
perties for reactor cofistruction. Nickel clad stainless steel can be obtained
at the present time as plate for the reactor vessel and as extruded tubing for
the primary heat exchangers (Ref. 51). The inspection of éhe cladding is, how-
ever, a problem.

A metallurgicaliibfifi is obtained between the nickél cladding and the stain
less steel backing material. Due to the intimate contact between tfie materials
and the adherence of the bond, there 1s essentially no thermal resistagce° The
efficiency in heat transfer propefties bf_this clad material is almost equal t&
that of solid stainless steel (Ref. 52).

The welding of this clad material will also present difficulties, but with

the proper weld area preparation and use of special welding rods, this problem

can be solved (Ref. 53).

T-8227
1000 HRS.

Figure 3.1 Changes in Attack with Increasing Operatlng Time in Forced-
Circulation Inconel LOODPS. Fluoride mlxture circulated, fuel

No. 30; maximun fluoride-mixture temperature, 1500 F; Temp-
erature differential, 200 F:; fluoride-mixture Reynolds

number, 5000.

6¢

- ————

7-8798

HOT LEG

250 X T-8799
FIGURE 3.2

corrosion of 85-15 Nickel-Molybdenum Alloy by
Fused Fluoride Salt No. 30 in & Thermal Convection
Loop. Meximm fluoride mixture temperature, 1500 F

Duration of test, 1000 hours.

COLD LEG

O

T-4233

COTTOBLUM Va Wy ==y ===

Tused Fluoride aalt No. 30 in a Thermal Convectlon

Loop. Meximm fluoride mixture temperature, 1500 F

Duration of test, 1000 hours.

250 X p-L23l
BOT LEG FIGURE 3.3

jckel in Fused Fluoride Salt No. 30

Corrosion of N

in a Theruml.ConNection 1.00D . Maximum.fluoriae

mixture temperature, 1500 F; puration of test,

500 houx's.

COLD LEG

“Th

L2

The choice of nickel clad stainless steel as the structural and primar
heat exchanger material for this reactor has made it possible to prepare a
* preliminary design which could be constructed with a minimum of development
work. -;t 1; recognized, however, that this duplex material has shortcoming
and that a sipgle material with all the desired propertieewwould be superio
Before a reactor of this basic deeign is constructed, it would be advisable
to examine the nickel-molybdenum system Very carefully to determine if those
alloys have the required properties to warrant substitution of the selectio
of nickel clad stainless steel with a nickel-molybdenum alloy.

3.2.0 SODIUM SYSTEM MATERIALS

Selection for the materials for.the'sodium system was based on the foll

ing considerationsi

1. Ability to resist corrosion and mass transfer in high temperature
sodium. _

2. Adequate mechanical properties at elevated temperatures for
extended period.

3. Stability under high radiation fields at elevated temperatures for
extended period. :
P
L. Fabricability into the required shapes.

5. Weldability

6. Reasonable neutron'ebsorptionneross section.

T. Availability and cost

At the present time, the austenitic stainless steel alloys more nearly
satisfy these requirements than other materials, except these alloys have a
relatively high neutron absorption cross section. However, since in this
design the heat exehangers carrying the sodium are outside the central core

and in a low flux region, this is of little consequence.

43,

There are three types of stainless steels, ndhely, Types 30k, 316 and 347,
vhich have been previously used in sodium systems with satisfactory results. -
There remains the problem of selecting the type that not only has the best
mechanical properties at the elevatéd temperature under consideration, but also
the type most readily weldable with no cracking of the weld and no embrittle-
ment after long-life high temperature service.

The material chosen fo contain the sodium in this reactor system was
Type 304-L stainless steel (0.03% maximum carbon content). Although the three
types of austenitic stainless steels show acceptable resistance to corrosion in
sodium up to 1000°F, columbium stablized Type 3h7 has been generally favored.
However, investigation at the Knolls Atomic Power Laboratory showed tha£ Type
304 stainlesfi steel is equally resistant to corrosion from sodium as Type 347
and that it has good strength properties and contains less strategic materials.
No mass transport difficulties have been encduntgred up to 1200°F, which is
higher than the sodium operating temperatures contemplated in this désignoi

Since the austenitic stainless éteels‘appear reasonably corrosion resis-
tant, namely, Types 304, 316 and 347, the strength versus ductility at elevated
tempefatures should be given considerable weight in selecting the material.
Type 304 stainless steel causes only slight loss in room temperature ductility
due to carbide precipitation (Ref° 56) , whereas Type 316 stainless steel,
under the same conditions and treatment, indicates rather severe decrease in
room temperature ductility. ‘ N

Table A.2 lists some of the pertinent properties of Types 304, 316 and

\

347 stainless steels at elevated temperatures. Therefore, from mechanical

properties consideration, Type 30h'appears to be the superior.
”~ ..

et
.

Lk,

Since there will be many welds in the sodium piping and components, the
corrosion effect of sodium on welded joints must be considered. The corrosion
resistance of welded 18-8 stainless steels is affected because of chromium
carbide precipitation at the grain boundaries in the heat affected zone. The
formation of chromium depletes the surrounding.area of chromium and thus the
area loses its corrosion resistance. The precipitation'of carbides can be pre-
vented by: (1) reducing carbon content below 0.03 percent; (2) using columbium
or other elements which form stable carbides with the carbon present; (3) cool-
ing the heat affected zone quickly through the critical range (1 to 1-1/2
seconds from 1500°%F to 900°%F); or (4) heating to above 1750°F following welding
80 as to redissolve chromium carbide, followed by rapid cooling to retain them;

Until recently the use of (2) above to prevent the precipitation of carbides
had been considered the superior method; however, in recent years, the use of
Type 347 weldments in tfiick sections for high temperature-long-time service has
resulted in cracking during welding and embrittlement during service. It has
been found that columbium free alloys do not become embritfled) even when Lhey
contain normal amounts of ferrite (Ref. 55). Also, the tendency for cracking
of the weld metal during welding is reduced with the columbium free alloys.

To prevent carbide precipitation; the use of a low carbon content austenitic
stainless steel appears now to be the better method. Rapid coqling and re-
heating to above 1750°F are not practical methods for projects of this scope.
The material which best meets the above requirements is Type 304-L. Therefore,
Type 304-L stainless steel was selected as the structural material for the
sodium systems.

3.3.0 STEAM SYSTEM MATERIALS

Since the temperatures and pressures of the steam system for this design

study are consistent with those of large conventional steam power plants today,

45,
the selection of materials %or the steam and water sysfems presents no parti-
cular problems. The main steam line piping would be fabricated from standard
- Cr-Mo alloy steel pipe, ASTM Spec. A-335, Grade P-1]1 or P-22. The higher
pressure boiler feed water, condensate, and extraction lines would be fabri-
cated from seamless carbon steel pipe, ASTM Spec. A-53, Grade B. Other
miscellaneous lines would be fabricated from standard seamless carhon stecl

pipe. All valves would be fabricated from the same material as that of the

piping in which they are located.

h6.

CHAPTER k.

REACTOR ANALYSIS

4,0.0 INTRODUCTION

The nuclear characteristics of this fused salt "burner" reactor are
relatively flexible and are not restricted by the stringent requirements
of breeding. In addition, fhe high solubility of UF) in the NaF-Zth fused
s8alt system allows considerable variation in uranium concentration which
provides a wide latitude for compromise between nuclear and engineering
considerations.

The basic reactor design consists of é vertical cylindrical core sur-
rounded by primary heat exchangér tube bundles. The nuclear requirements
of the reactor determine an optimum core diameter for minimum fuel inventory.
The core dimensions selected represent a compromise between these considerations.

In order to proceed with thé nuclear calculations, a preliminary arrange-
ment of the reactor vessel was selected. This section of the repurt describes
the reéégor analysis calculations that were performed to determine the optimum
core size and fuel concentration and other nuclear rgquirements, A summary of
the results is included in Table k4.1.

4.1.0 CRITICAL SIZE AND FUEL CONCENTRATION

A preliminary criticality calculatioh was made by W. K. Ergen, OaknRidge :
National Laboratory (Ref. 29), for a reactor similar to that of this project
using a bare sphere, Fermi Age approximation. The optimum size and fuel con-
centration for minimum fuel investment calculated by this method was 136 cm
(4-1/2 feet) in radius with a fuei concentration of 0.0208 gm/cc of uranium-235.
The fused salt used in these calcfilations was a 50-50 fiole percent composition

of NaF and ZfiFh with the uranium dissolved as UFh“

b7,

The reactor contains no moderating material other than the fused salt
itself. The atom having the smallest mass and existing in greatest abund-
ancé in the fused salt to be used in this reactor is fluérine, Fluorine;
then, is largely responsible for the moderating characteristics of thé fused
salt. With an atomic weight of 19, fluorine is a relatively poor moderator
compared with the lighter elements usuaslly used as neutron moderators. It
was suspected, and later calculations have shown, that this reactof was not
" a thermal reactor, but that a high percentage of fissions would occur from
epithermal neutrons.

.Hand calculations for a bare sphere with th?ee energy groups indicated
that approximately 50 percent of the fissions would involve epithermal nefitrons}

~With these preliminary hand calculations as background, a series of three
energy group, three geometry region (3G3R5 calculations were performed on the.
Oak Ridge National Laboratory digital computer (ORACLE) ..

The energy group limits were_éelected for reasons given in Appendix C.l.
The intermediate group extends from 0.5 ev to 100'ev,_.The fast group extends
from 100 ev to 2 mev. The central core of the reactor is region 1; the lrinch.‘
" thick baffle surrounding the core is region 2; and the heat exchanger and
downcqmgr volumes gre homogenizedlfor region 3.

The phygical properties and nuclear parameters were based on a salt of
composition 43 mole percent ZrF) and 57 mole percent NaF (Section 2.1.0).
‘The fuel is highly enriched uranium-235 (over 90% U-235). The fast and in-
termediate energy group nuclear cross sections were obtained from the "eyewash
code” (Ref. 27). Thermal cross sections were obtained from BNL-325 (Ref. 30).
The detailed,détg and calculations for the code input are given in Appendix C.

In order to minimize the number of computer calculations, 1t_was_de§ided

to vary only the radius of the central core and the fuel concentration. Other

48.

parameters such as the height of the reactor and the thickness of the heat
exchanger were held constant. Since previous estimates indicated a h-1/2 ft
radius core to be optimum, a 1C-ft core height was chosen to provide com-
patible arrangement with this core radius and the heat exchangers. The heat
exchangers contain a considerable amount of fuel, acting somewhat as a
reflector to the central core, which makes 1t desirable to have a central
core height larger than the central core diameter. Calculations were per-
formed with central core radii of 3, 3-1/2, 4, k<1/2, and 5 feet. For each
radius, calculations were made for fuel concentrations of 0.005, 0.010, 0.020
and 0,030 gm/cc of enriched uranium. |

The above information and the data in Table C.2 were coded (Ref. 28) for

the electronic computer (ORACLE). The electronic computer calculated the

multiplication constant (k) for each radius and fuel concentration (Figure 4.1).

*From these data, the optimum size and fuel concentration were determined (dis-
cussed in Section 4.3.0) as a 3-1/2 foot central core radius and 0.0088 gm/cc
of enriched uranium-235. A re-ecvaluation of the height of the central core
is probably desirable in light of the smaller central core radius. In the
final design, the height is shown to be 8 feet, which is probably a more com-
- patible figure. A concave head and bottom are used to obtain this reduction
in height and to direct the fused salt flow (See Figure 5.9). However, the
nuclear, heat transfer, and other cfilculations were not revised to reflect
this change. |

In order to check the inaccuracy introduced by using the l-inch thick
stainless stee; and nickel baffle plate as a diffusing medium (region 2),
two check calculations were made. One calculation considered the Baffle as
an absorbing "shell".between the central core and the heat exchanger regionms,

and the other eliminated the baffle plate e?tirely. The multiplication

49,
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51.
constants (k) and flux plots obtained by the three methods of treating the
baffle plate were almost identical (Table C.3). It was therefore inferred
that using the baffle plate as. a region did not in itself introduce any
appreciable error.

4,2.0 NEUTRON FLUX

The electronic computer (ORACLE) calculated the neutron flux distribution
for each encrgy group as a function of radius (Figure 4.3).

At 600 MW total power, the fast neutron flux is approximately 1.2 x lO15
neutrons/cm2 sec.; the intermediate neutron flux is approximately 6 x lOl,+
neutrons/cm2 sec., and the thermal neutron flux is approximately 1.6 x lOlu
neutrons/cm2 sec. (Appendix C.4). These values, coupled with the fission
cross section for each group, show the fissions to be distributed among the
three energy groups as follows: 9 percent in the fast group; 27 percent in
the intermediate group; and 64 percent in the thermal group (Appendix C.5).
This means that this reactor is what is often termed an "intermediate" reactor.

A nuclear calculation was prepared to utilize the 30 energy group "Eye-
wash" code on the UNIVAC digital computer. Unfortunately, the results of
this calculation were not available to be included in this report; but should
be shortly thereafter (Ref. 40). The results from this more detailed calcu-
lation will serve as a check on the 3G3R code calculation results. Also, and
the original reason this alternate calculation was prepared, the 30 energy
groups will provide a much better indication of the neutron flux energy spec-
trum in this reactor. This is of particular interest for the high neutron
energies, to better establish the possibility of utilizing this reactor as a
large flux source of high energy neutrons for purposes of radiation damage

research.

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4.3.0 FUEL INVENTORY

The critical uranium concentr;tion for each central core radius was
determined by the results of the three gfoup, three region calculations.
‘The critical uranium concentration decreases withran increase in the central
core radius. The volume obviously ;ncreases with an increase in the central
core radius. The combined effect fésmlts in s minimum uranium inventory
required for criticality at some value of central core radius. This vériation
‘of critical mass (uranium inventory) with central core radius is shown ifi
Figure 4.2, |

For 600 MW rating, the heat exchanger volume, and ihus the fuel volume
in the heat exchanger and downcomer; will remain nearly constant. The fused
salt solvent inventory is decreased by a decrease in central core radius.
Therefore, the central core'radius was chosen slightly smaller (3-1/2 feefi)
than that for the minimum fuel inventory indicated (Figure 4.2), in order to
allow some effect of the decreased salt inventory. The total uranium-235
fuel inventory in the entire fuel system for the hot, clean critical condition
18 approximately 240 kg.

4L.4,0 FUEL BURN-UP

The fuel burn-up at 600 MW will be about 780 grams of uranium-235 a day.
The fuel would be added on & temperature regulation basis to maintain an average
operating temperature of 1125°F.

4L.5.0 URANIUM-233 FUEL

Highly enriched uranium-235 is the fissionable fuel utilized in this
-design study. The fact that a large fraction of fissions is caused by epi-
thermal neutrons suggests that it may be desirable to use uranium-233 as the

fissionable fuel instead of uranium-235s Uranium-233 has a uniformly low

L
value for the ratio of capture-to-fission cross section (oC ) of 0.1 over
all neutron energles. This,méans that a considerably smaller portion of
non-fissioning neutron absorptions occurs in uranium-233 than in uranium-235,
particularly for neutrons in the intermediate energy range. The advantages
that would be. gained by using uranium-233 the fuel, if_availableg are that -
the uranium inventory required would be le:!!\hid the energy release for the
uranium burned-up would be greater.

4.6.0 FISSION PRODUCT POISONING

Fission products and corrosion prodficts will accumulate in the fuel solu-
tion and will add considerable nuclear poison to the core. This will require
an increase in the uranium fuel inventory and could become a significant
economic consideration.

An evaluation of the probable magnitude of this effect was made, and is
outlined in Appendix c;.6°

Based on the reference source utilized in this analysis (Ref. 33 and 34),
and the assumptions and approximations made to adapt the data of this refer-
ence to this design study, the following effects of fission product build-up
are predicted. The nuclear poisoning of the non-gaseous fission pfoducts
only (no corrosion products) will asymptotically approach a negative reactivity
of about 0.1 during the lifetime of the reactor. If corrosion should be appre-
ciable; the relatively high absorption cross section of the nickel atom could
appreciably change_the rate and magnitude of this poison build-up.

From Figuré L,1, it can be seen that a negative reactivity of 0.1l would
require an increase in uranium concentration from 0.0088 gm/cc to about
0.0115 gm/cc to override this poison and maintain criticality. This is a{

31 percent increase in uranium inventory from 240 kg to 314 kg.

03
Some form of fuel reprocessing could reduce the uranium required to over-
come the poisons. This possibility was investigated (Section 2.6.0., Appendix
€ C.6 and C.7), and it was determined that it would not be economical to provide
- for fuel reprocessing as limited by current technology.
o The increase in uranium inventoys. to override fission product-poisoning
would then progress gradually ov# life of the reactor, being added as .

required along with the uranium added for burn-out.

L,7.0 TEMPERATURE COEFFICIENT OF REACTIVITY

The temperature coefficient of reactivity of the reactor was not calcu-
lated, but is believed to besufficieqtly negative for the»reactor to be
stable. This is based on the fact that a reactor using a similar fused salt -
fuel has been operated, and this reactor had a negative temperature coEfficiefit
of reactivity of approximately 5 x 1072 Ak/k per F (Ref. 39), p. 1009). The
fact that this fuel has a desirable temperature coefficient of reactivity is
also indicated by the slope of the curves in Figure L.1.

4L.8.0 DECAY HEATING

The decay heating of a reactor during shutdown is dependent upon the
amount of negative reactivity injected into the reactor_and the delayed neu-
trons and gamma decaf rate.. In a reactor with control rods, a considerable
amount of poison can be inserted very quickly. In the fused salt reactor,
however,‘no control rods are present, and the ‘amount of negative reactivity
achieved depends upon the magnitude of the temperature rise above the average

L operating temperature.
In the event of complete loss of pumping power, the reactor must be shut
down and the decay heat removed. 1f it is assumed that no heat is removed and

the temperature is allowed to rise, the power generation would decrease rapidly,

1

56.

and the reactor period would level off to an 80-second period in a fraction
of a minute. At this time, the power generation of the reactor would be re-
duced to a few percent of operating power. The temperature rise under these i
conditions may be several hundred degrees within the first minute. This
condition of rapid temperature rise is undesirable, if not intolerablg_° Imme- ~
diately following the loss of pumping power, some heat would be radiated from
the reactor vessel; some heat would be removed by the reactor vessel surface
cooling; and some would be removed by the momentum of the sodium in the primary
loop.

A detailed calculation of the flow characteristics of the primary and
intermediate sodium loops after l&ss of pumping power was not made; however,
a simplified calculation showed that the inertia of the fused salt in the
reactor, and the sodium coolant in the primary and intermediate loops is
sufficient to maintain the flows in the turbulent range for approximately
one minute. This would be adequate to carry away a considerable amount of -

decay heat. and limit the fused salt temperature to a safe value.

o7
TABLE 4.1

SUMMARY OF REACTOR ANALYSIS DATA

1. Dimensions of reactor central core:

Radius - 3 1/2 feet

Height - calculations based on 10 feet
2. Critical fuel concentration:

0.0088 grams/cc (for hol clean critical)
3. Power density:

55.2 watts/cc at 600 MW

4. Flux and fissions for each energy group:

_Flux Fissibns:l
Fast {100 ev to 2 mev) 1.2k x 1017 neut/cc-sec 9%
Intermediate (0.5 ev to 100 ev)  5.96 x 10M* neut/cc-sec 27%
Thermal (up to 0.5) 1.62 x 101% neut/cc-séc 64%

5. Fuel Inventory:

240 kg (for hot, clean critical condition)

314 kg (for hot, fission product contaminated condition)
6. Fuel Burn-up:

780 grams/day at 600 MW

T. Temperature coefficient of reactivity:

Not calculated; but reasonable assurance that it is satisfactory.

58.

- CHAPTER 5.

REACTOR AND PRIMARY HEAT EXCHANGER DESIGN

5.0.0  INTRODUCTION -

The basic design of the reactor consists of a vertical cylindrical
pressure tight vessel to contain the homogeneous circulating fuel. The
size of the reactor depends on the nuclear characteristics of the fuel and
the method of removing the heat from the circulating fuel system. Critical
size and fuel concentration calculations have been discussed in Chapter &,
based on the design concept of using primary heat exchangers internal-tp
the reactor vessel. Several arrangements were considered prior to the |
selection of the internal arrangement and reasons for this decision are dis-
cusesed 1n the subsequent sections.

5.1.0 PRIMARY HEAT EXCHANGER

The function of the primary heat exchanger is to transfer the heat from
the circulating fuel to the primary sodium cooclant loop. The design of the
exchanger is based on a heat removal rate of 2°Oh8 X lO9 BTU/br for both &ides
of the exchanger with mixed mean fuel temperatures of 1200°F entering and
IOSOQF leaving, and mixed mean sodium coolant temperatures of 1000°F entering
and 1150°F leaving the exchanger. The minimum fuel temperature has been arbi-
trarily set at lOSOQF, approximately 100°F above the fuel melting point to
allow a temperature drop of 50°F‘for fuel processing and an additional 50°F
for cold spots in the system before reaching the solidification temperature. ~ -
The upper fuel temperature was set at 1200°F, which was considered the maximum
allowable for tolerable corrosion and mass transfer rates for a long life power

reactor. The sodium coolant temperatures were chosen to obtain a practical

09
e,
log mean temperature difference across the heat exchanger which assures reason-
able flows and heat exchanger surfaces. These temperatures provide sufficient
margin above the fuel melting temperature to prevent operational difficulties °
from possible cold spots caused by poor.distribution in the flow pattern

through the heat exchanger and reactor.

5.1.1 Internal vs. External Arrangement

In determining a practical arrangement for the reactor and primary heat
exchanger, the following two basic design concepts were considered: (1) Use
of internal primary heat exchargers with associated piping and pumping equip-
ment integral with the reactor. In this design, the fuel remains in the
reactor vessel and the heat is transferréd from the fuel circulating inter-
nally to the sodium coolant piped inside the reactor vessel. (2) Use of
external primary heat exchangers through which the fuel is continuously cir-
culated outside the reactor with piping and pumping equipment adjacent to,
but external to, the core vessel. Each arrangement offers certain advantages
and disadvantages, and the final selection depends on an extensive evaluation
of the economic operation and design factors involved, and their effect on the
design chgracteristics of the reactor and system.

- The important advantages of the interfiél primary heat exchanger arrange-
ment over the external arrangemeflt are:;

1. Lower fuel holdup because external fuel piping eliminated.

2. More compact arrangement résulting-in reduction of requirements for

shielding, piping, containment, building and services (i.e., crane
handling equipment,; etc.)

3. Less pumping power because of reduced piping friction losses.

4., Provisions for electrically heating the fuel containment vessel and
fuel piping simplified.

These 5&vantages must, in a complete economic and feasibility study,
overbalance the following apparent disedvantages:

1. Difficulties in maintaining and replacing heat exchanger tube bundles.

2. Structural problems in supporting integral tube bundles. |

3. Higher radioactive sodium loop caused by the higher flux level inside
the reactor increasing the desirability of an extra sodium lcop.

k. Problems in fabrication of the larger reactor vessel and complexity
of heat exchanger tube layout.

5. Difficulties in locating leaks and isolating the defective tube bundle.

Time did not permit a complete anslysis of this problem. However, based
on preliminary studies; it was concluded that a design employing internal heat
 exchangers and pumps could be developed wvhich would result in an economical and
practical arrangement for a high power reactor. Therefore, the design of an
internal arrangement was pursued from the outset of this study.

5.1.2 Internal Arrangement

Several schemes for the arrangement of the internal primary heat exchangers
were‘considered and are schematically shown on Figures 5.1 and 5.2. In all the
'schemes proposed, the requirement of counter-flow heat exchamge was followed in
order to minimize tube surface and to obtain the best heat exchanger character-
istics possible.

A. Separate Upper and Lower Header and Single Pass Arrangement (Filg° 5,.1)

The tube bundles are located in an annular space around the core section
with baffles to direct the fuel flow down through the exchanger on thg outside
of the tubes. The sodium cooclant is through the tubes. The tube bfindles ter-
minate 1n concentric toroidal headers; with the coolant outlet header located
at the top of the reactor vessel and the coolant inlet header locgted at the

bottom. This arrangement imposes several apparent disadvantages.

61.

1. To obtain adequate heat transfer surface using practical tube sizes
and a reasonable number of straight tubes would require tube lengths
considerably longer than the optimum length dictated by the nuclear
requirements for the reactor core size.

2. Use of a lower header would necessitate many large penetrations in
the bottom section of the reactor vessel, which is undesirable and
should be avoided if possible.

3. This arrangement requires pumping the high temperature fuel, which
tends to decrease pump life because of increased corrosion rate at
the higher temperatures.

Other design difficulties inherent in this arrangement are thermal stress
and expansion problems between the two headers of the exchangers, and the large
number of tube-to-header welds required to permit a practical length.

To overcome these disadvantages, Scheme II was devised and is the basis

for the final design.

B. Combined Header and Two-Pass Arrangement (Fig. 5.2)

The tube bundles consist of U-tubes with the coolant inlet and outlet
headers located annularly and concentrically at the top of the reactor vessel.
The tube sheets have the shape of truncated circular sectors. The fuel makes
two passes on the outside of the tubes, with the flow directed by concerntric
baffles into the suction of the fuel circulating pumps which are located on
the top outer periphery of the reactor vessel.

5.1.3 Basic Design Criteria

In establishing the design, certain criteria were followed to take advan-
tage of the present state of development of pumps, materials and heat exchangers.

1. The pumps would be positioned vertically and located symmetrically at
the top of the reactor and would pump cold fuel. This would minimize
seal problems and provide convenience in arrangement. Pumps would
be sized practically.

2. Annular vertical U-tube heat exchanger bundles would be considered
with duplex tubing, with 304 stainless steel as the structural
material and clad with type "L" nickel on the outside to resist
the corrosive fuel. Type 304 SS provides compatibility with the
rest of the Na coolant circuit.

-62- ORNL-LR-DWG-15339

ENRICHER LINE-—
SODIUM AXIAL FLOW PUMPS

T

—~ SODIUM HEADER

SODIUM SODIUM
TO HEAT EXCHANGER TO HEAT EXCHANGER

FIG 5.1- REACTOR ARRANGEMENT WITH STRAIGHT- TUBE HEAT EXCHANGER

PRIMARY MEAT

EXCHANGER
TUBE BUNDLE

BAFFL

FIG. 5.2- REACTOR

63.

ORNL-LR-Dwg. -18172

GRAPHITE REFLECTOR

ARRANGEMENT WITH U-TUBE HEAT EXCHANGER

PRIMARY SODIUM
COOLANT OUTLET

PRIMARY SCDIUM
COOLANT INLET

FUEL
CIRCULATING
PUMP

6.

3. The flow pattern through the reactor and heat exchanger would tend
to maintain the circulation of the fuel in the event of complete
loss of pumping power.

k. The heat exchanger tube bundles would be sized to minimize fuel
holdup and the number of tubes while maintaining practical fuel
and sodium coolant velocities for reasonable friction pressure
drops and heat transfer coefficients. This requirement necessi-
tated a parametric study for the primary heat exchanger.

5.1.4 Parametric Study

A parametric study was performed for the primary heat exchanger to deter-
mine the optimum tube size and number of tubes for the required duty. The
study was based on varying tube outside diameters with different ligaments
(minimnm distance between outside diameter of adjacent tubes) for a number

of cases, utilizing the following design conditions:

Heat exchanged 2.048 x 107 Btu/br
Fuel inlet temperature 1200°F
Fuel outlet temperature 1050°F
Sodium coolant inlet temperature 1000°F

Sodium coolant outlet temperature 11509F

Tube material: duplex tube, Type ALSL, 304 stainless steel (.06 € max)
clad with Type "L" nickel

Tube thickness: total 0.065, SS 0.042 in., Ni 0.023 in.

Fuel properties based on: 50 mol % NaF - 50 mol % ZrF)

Justification for the selection of these design conditions has been dis-
cussed previously. Tube sizes of 1/2 in., 5/8 in., 3/4 in. OD each with
ligaments of 1/4 in., 3/16 in., and 1/8 in. were studied, using a staggered
arrangement with triangular pitches of tubes in bundles. For the purpose of
this investigation, average values were used for the physical and thermal pro-
perties of the fuel, sodium and structural materials. This simplification was

considered valid over the temperature ranges involved. The complete derivation

65.
iof the formulae for the study is outlined in Appendix D; however, in the
development of the expressions, the foilowing basic_eqpétions-were used for
establishing the applicable heat transfer coefficients.

Tp simplify the study, the internal heat generation in the eirculating -
fuel and sodium caused by delayed neutrons and gamma heating was neglected.
Therefore, if a uniform wall heat flux is assumed, the Martinelli and Lyon
equation (Ret. 15, p. T3) is applicable for determination of the sodium

coolant tube side film coefficient,
hdaT7%0.025 (Pe)°°8 (5.1) ,
k- . _

To establish the shell side film coefficient for the fuel, the Dittus-Boelter

relationship (Ref. 6, p. 219) was used,

4, = (0.023) (Ry)0-8 (p.)0-% | (5.2)
. |

Evaluafion of fieynolds numbers for the various parameters studied indicated
values‘in the fange of 1000 to 4200 on the shell side. For a-fube side calcu-
lation, Reyndldé numbers of thip order 6f magnitude.would indicate that the
flow was in the transition range between laminar and turbulén£ flow. Keén
(Ref. 20, p. 137) indicates that for baffled heat exchangers, there is no dis-
continuity in flow at Reynolds numbers near 2100, such as occurs with:tu§e
side flow, and that the character of the flow on the shell side cannot be
evaluated on the basis of the Reynolds number only. Further, this,referenée
indicates fhat at Reynolds numbers‘as low as 100, fully turbulent flow can be
developed on the shell side of a baffled heat exchanger. The equation given ,
by Kern (Ref. 20, p. 137) for shell side film coefficients gives values cfin-

siderably larger than obtained from the Dittus-Boelter relation. Therefore,

66.

use of the Dittus-Boelter equation results in conservative values for design,
even though the bvaffling geometry is unknown.
Calculation of the tube wall resistance was based on averaging thermal
conductivity of the tube at the hot and cold mixed mean sodium coolant temperatures.
The over-all heat transfér-caefficient was obtained by combining the indivi-
dual resistances for the duplex tubes and film coefficients based on external

tube surface,

1 =1 +Rys +Rsg+ 1 {5.3)
Uo Bbs bya

The required heat. transfer surface was determined by using the equation for
~constant over-all heat transfer coefficient (Uy) and counter flow adiabatic

heat exchange between fuel and sodium (Ref. 6, p. 190),

q = Uy S (At,;) {(5.4)
For this study, the fuel and sodium flowse, the logarithmic mean temperature
difference, and the yuantity of heat exchanged are fixed by the design comditions.
It is possible by proper arrangement of these equations in conjunction'with the

one dimensional steady state continuity equation;

ws: AV (5.5)
to develop simultanecus equations for the dependent variable, the over-all heat
transfer coefficient (U,) with appropriate constants (C) as fumctions of the
number of tubes (N) and length of tubes (L).

The two equations developed in Appendix D are then,

1 .SR+ 1 4__1 (5.6)
U, - - c/N'® . c¥cenB |

67,
and
U NL=zC (5.7)
where the C's are independent constants

number of tubes

N =
L = length of tubes
R = sum of the individual duplex tube resistances

For an assumed length (L), these equations are then two simultaneous
equations in two unknowns which cen be solved graphically by plotting the
over-all heat transfer coefficient (Uo) against assumed values of N for each

case of tube diameter and ligament under consideration. For the equations to

B
L

be satisfied; there exists only one value of N which is determined by the in-
tersection of the curves of the two equations. The required number of tubes.
(N) to give the necessary heat transfer for each tube size and various ligamnets
is shown on Figures 5.3, 5.4 and 5.5.

The pressure drops for both shell and tube side of the exéhanger can be

. h
BT ot R L E LN N
ARy IR

determined from the Fanning friction pressure drop formula, once the requiréd
number of tubes for a particular tube size and ligament 1a-established°

ap (pst) = (2.0+f L)V P (5.8)

de 2g 1M
The friction factor f was determined from Moody's chart (Ref. 16,

Fig. 15) for smooth Pipe. The constant 2.0 in the above expression allows
for inlet, outlet, and U-bend losses in the respective circuits (Ref. 16, p. 21
et sequi). The additional fluid pressure loss across the tube bundle supports
on the shell side was not included since its method of determination is highly
empirical and depends on particular tube layout and design, and, in additionm,
it was considered that its contribution would not effect the selection of tube

size and spacing.

68.

A summary of the results of the parametric study is shown in Figures 5.6
and 5.7 and Table 5.1. These graphs ého?;cross plots for the number of tubes,
heat transfer surface} fuel holdup'volfimé and total fluid phmp powef a8 a
function of tube ligament ratio (ligament to tube OD) for the values of N, as
determined from figures 5.3, 5.4 and 5.5.

Examination of the results of this st;dy indicates that for minimum fuel
holdup, reasonable pumping requirements and a practical tube arrangement, the
selection of approximately 14,600 5/8-inch OD U-tubes with 1/8-inch ligament
(0.20 tube ligament ratio), and an effective length of 20 feet results in the
most serviceable and practical primary heat exchanger design. In determining
the optimmm tube size; it is desirable to minimize the number of tubes consis-
tent with reasonable pressure drops and fuel holdup volume. The fewer the

number of tubes; the less the probability of tube or tube-to-tube sheet failures,

201.5 Design Considerations

To determine the practicability of such a U-tube design in annular con-
centric sections with a 3-foot, T-inch inner core radius; a tube bundle layout
study was made using approzimatély 15,000 5/8-inch 0D tubes arranged on a
3/4finch staggered equiléfiergl triangular pitch. The result of the study
indicated that such an arrangement was feasible and is shown in Figure 5.8.
It was finally concluded that to provide additional wargin in the design for
contingencies, 24 tube bundles, each with 650 tubes to carry, or a total of
15,600 tubes with an effective tube lengfih_of 2i‘feet would be specified for
the primary hegt exchanger., |

Due to the difference in areas caused by the concentric sections,; some
crossing over of the tubes is necessary to retain the same tube density on

both sides of the tube bundle. Although this increases the fabrication and

69.
\
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Figure 5,7 Primary Heat Exchanger Parametric Study
Summary - Variation of Pump Power, Fuel
Holdup Volume and Heat Exchanger Surface

EgfifiEIZOOHHWWrTIMIuITHTH with Tube Ligament Ratio and Size of Tubes
s Total Fluid Pump Power HiEIN{HHIIHsE T I
ssccl URNL g =181 . 13-
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Tube Ligament Ratio (Ligament/Tube OD)

TABLE 5.1 :7‘5%{

SUMMARY OF PRIMARY HZAT EXCHANGER PARAMETRIC STUDY

Physical properties used for study:

FUEL
Absolute viscosity «¢ 190 lb/hr-ft
Density ¢ 209 1b/cu ft
Conductivity k 1.5 BTU/hr-ft-OF
Specific Heat Cp 0.285 BTU/1b

The study covered six arrangements using 1/2

type 304 stainless steel with ligaments of l/fi"

SODIUM Type ALSL - 304 Stainless Steel Type "L"
Nickel (.06 C Max)
A 0.53 1b/hr-ft
¢ 51 1b/cu ft
k  36.2 BTU/hr-ft-°F kK 12.9 BTU/hr-£t-°6° k  33.5 BTU/hr-ft-°F
p 0.3C BTU/1b

5/8", and 3/4" duplex tubes with 0.023" clad Ni on outside of 0.042" wall
s 3/16", ard 1/8".

C

6

Reactor heat load - 600 Mw Temperature - Fuel 1200°F inlet - 1050°F outlet Flow rate - fuel 48.0 x 10> 1b/hr
Sodium 1000F inlet - 1150°F outlet sodium 45.7 x 10 1b/hr
Tube OD (in.) 1/2" 5/8" 3/4"
Ligament (in.) 1/4" 3/16" 1/8" 1/4" 3/16" 1/8" 1/4" 3/16"  1/8"
Tube Ligament Ratio 0.5 0.375 0.25 0.40 0.30 0.20 0.33 0.25 .17
No. of tubes required with 20 ft 60,000 32,500 18,000 50,000 26,000 14,600 46,200 22,200 13,200
effective length
Velocity - Fuel outside tubes, ft/sec 0.50 1.3 3.6 0.50 1.3 3.4 0.46 X3 3.2
Sodium in tubes, ft/sec 5.6 10.3 18.6 27 7.2 12.5 2.6 5.4 9.0
Reynolds No. - Fuel 1310 2310 4150 1200 L 2350 4000 1070 L 2250 4050
Sodium coolant 5.95x10% 1.10x105 1.98x10° 5.34x10* 1.025x10° 1.78x10° 4.6x10 9.6x10" 1.61x107
Over-all heat transfer coefficient- 260 485 885 252 492 865 228 470 820
BTU/hr-sq ft
Heat transfer surface 157,000 84,700 46,500 164,000 84,000 47,500 181,000 87,200 51,000
Pressure drop - Fuel, psi X1 .89 8.0 .13 .89 7.8 .13 .84 Gl
Sodium, psi 2.5 1.2 22.0 .91 3.0 8.5 S 1.4 3.6
Pumping horsepower - Fuel, hp 1.8 14.9 134 2.2 14.9 130 2.2 14.0 102
Sodium, hp 163 468 1430 59.0 195 552 2k .0 91.0 234
Total, hp 165 483 1564 61.2 210 682 26.2 105 336
Fuel holdup in tube bundle, cu ft 2420 967 355 2470 950 365 2730 980 405

.flL

] /DIVISION LINE BETWEEN BUNDLES
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T
/ 9L R
/ e NOTES
- | TUBES IN AREAS A, AND A, ARE
A CONVENTIONAL U-TUBE BUNDLE.
FIG. 5.8 PRIMARY HEAT EXCHANGER 2 TURES INHRDA 8. SE OE 125
‘ OF A U-TUBE BUNDLE WHICH DIVIDES
TUBE LAYOUT AND CROSSES OVER BELOW THE
TUBE BUNDLE CONTAINS 650 TUBES. BUNDLE A,-A, INTO AREAS B,.

24 BUNDLES REQUIRED

T6.
assembly problems;, the heat transfer characteristics are improved. Since the
physical crossing of the tubes occurs in the lower portion of the bundle; the
fuel flow pattern through this section will be irregular; improving the mixing 1
of the fuel as it passes through the exchanger.

In order to obtain adequate cross sectional area for welding the tubes
to the tube sheets, it was initially planned to increase the pitch from 3/4"
to 7/8" and extend the tube sheet radially outward over the tube bundle to
obtain the additional area; however, an actual layout of this arrangement in-
dicated that the outer radius of the tube sheet became excessive and therefore
this plan was abandoned. To solve this difficulty, the following various alter-
nates were proposed: (1) place the tube sheets at angles from 45° to 90° to
obtain the required additional area. This method of arrangement offers the
disadvantage of imposing severe stresses on the tubes at the bend where the
tube enters normal to the tube sheet. (2) Allow the fuel to flow inside the
exchanger tubes and the sodium coolant ocutside with the tube sheet positioned
in the vertical. This arrangement offers some advantages such a&s reductlon in
the required heat transfer surface because of the larger over-all heat transfer
coefficient brought about by the fuel film coefficient; the main resistance to
heat flow in the system, being increased with the fuel inside the tubes. The
difficulties involved in draining the tube bundle completely and problems in
successfully cladding and inspecting the inside of the tubes ruled this arrange-
ment out. (3) Finally, it was concluded that the most expedient solution to
the problem would be to swage the tubes to 1/2" OD size Jjust before the tubes - .
enter the sheet; maintaining the 3/4" pitch the full length of the tube. This
allows a spacing of 1/4" between the tubes on the tube sheet with 5/8" OD tubes,
which is considered enough for a successful welding procedure. No difficulty

is anticipated in the swaging of the duplex tubes.

e

The final exchanger design comsists of 24 U-tube bundles with 650 tubes
positioned vertically in the reactor vessel; each with an effective length
of 21 ft. The size of the duplex tubes suggested is 5/8" OD with .O42" type
304 stainless steel wall clad with .023 type "L" nickel. Approximately a
total of 20 tube spacers and support plates will be required in order to in-
sure proper tube spacing and support. An arrangement with the tube spacers
staggered on 12" centers will result in a practical configuration. The tubes
will be swaged to 1/2" OD as they enter the tube sheet.

The method for supporting the individual heat exchanger assemblies is
discussed in Section 5.2.3.

5.2.0 REACTOR VESSEL

The reactor; shown schematically on Figures 5.9 and 5.10, is basically
a vertical cylindrical vessel with an over-all height of approximately 13 ft
and a minimum diameter of 16 ft for most of its height. The diameter in-
creases to 20 ft at the top of the vessel to allow for the installation of
eight equally spaced fuel circulating pumps. In order to reduce fuel holdup,
the minimum diameter is maintained for the height of the vessel; except at
the location where the fuel circulating pumps are positioned. At these loca-
tions, the vessel flares into protuberances to provide the necessary flow
area for the pump discharge transitional sections. An annular plenum chamber
to serve as a suction header is located at the top of the vessel directly
below the pump base plate. This distributes and permits flow to all pump
inlets in the event of possible failure of any single pump.

Both the bottom and top of the reactor vessel are concave hemispheroidial
sections to minimize fuel holdup, to allow for more effective length of the

heat exchanger, and to improve the flow characteristics of the fuel into and

2%

t FUEL CIRCULATING PUMP

8 REQUIRED
| PRIMARY
{ * ’ L CCIOLA'NT TO PRESSURIZER, GAS REMOVAL,
| PRIMARY * Y ? AND EXPANSION TANK

//—SF)DIUM COOLANT
. a:

Q| V7777 b o B Z

Bf 1B LR T I Sy H |
'1 Q.. 1. X L\'\OPENING FOR
T AN A O fq GAS ESCAPE
RETURN FROM C c H o e :
PRESSURIZER AND \
EXPANSION TANK [ \
by
N
N
\ REACTOR CORE
SWING CHECK VALVE
N
Bl

24 U-TUBE BUNDLES o

650- 5/8"0.D. TUBES EACH [ | { ‘
4" beer—
—— | 8- 27" - 7

2222722222 22272277 ELIELLLLL

: -~ ‘O - GRAPHITE REFLECTORd' EN
: £ o TN e : Pt e
OPENINGS FOR .
FUEL DRAINAGE FUEL DRAIN AND FILL LINE|

F1G-5.9-SECTION (A-A) THROUGH REACTOR VESSEL

—BL-

+ QMI-YI~INY0

V79241

ORNL-LR-DWG: 15228 A

PUMP DISCHARGE COLUMN

. QZ% SECTION C-C

/'/.

24 U-TUBE BUNDLES
650-5/8"0.D. TUBES EACH

RETURN FROM PRESSURIZER
AND EXPANSION TANK

TO PRESSURIZER,GAS REMOVAL
AND EXPANSION TANK

FUEL CIRCULATING PUMP

A

SECTION B-B

STRUCTURAL MEMBER

PUMP SUCTION
PLENUM CHAMBER

FIG. 5.10-PLAN VIEW OF REACTOR VESSEL

—GL_

PRIMARY SODIUM
COOLANT INLET

PRIMARY SODIUM
COOLANT OUTLET

-80-
i =S

ORNL-IR-DUG: 15348

PRESSURIZER, GAS REMOVAL,
AND EXPANSION TANK /

\_
TO GAS OFF-<
TAKE SYSTEM ~__
TO HELIUM
== S
24

FUEL CIRCULATING
PLIMP MOTOR

PRIMARY SODIUM
COOLANT LINES

LIQUID LEVEL
@ 1300F
@ 1125F

FIG: 5.11- PRESSURIZER, GAS REMOVAL , AND EXPANSION SYSTEM

81.
out of the centrai core section. Use of the concave sections reduces the need
for internal baffling and perforated plates to distribute the fluid flow,“

. To minimize neutron leakage at the top and bottom of the reactor, graphite
blocks are packed into the concaved sections to serve as a neutron reflector.
A side stream of the primary sodium coolant is used to remove the hesat caused
by the gamma and neutron interactions in the graphite. Similar provisions
could be uwmde fér cooling the feactor shell if further investigation indicates
that the gamma heating in the reactor shell is excessive.

)

5.2,1 Shell Design

-Since the fused salt in the reactor has arvéryAlow vapor pressuré at the
operating température (,Ol.mm Hg), the system is pressurized with helium at a,
pressure of only 10 psi (See Section 5.2.5). Because of this low pressure, -
the shell thickness reqnired for the Qessel-is relatively thin as compared to
most other types of high temperature reactors. By using a design pressure of

50 psi in the vessel, a maximum temperature of 1250°F of the fuel, Type 304

stainless steel as the structural material, and a joint efficiency of 80%,-a - .

shell thickness of 1.8 inch is required for the reactor vessel. This was
obtained byvapplying the ASME formula for cylindrical shells listed in Unfired:
Pressure Vessel Code (Ref. 35, p. 113). To allow an additional margin of
safety, the shell thickness 6f the vessel 18 designed for 2 inches and is
fabricated from Type 304 L stainless steel with 0;03“¢ maximum.

- The choice of Type 304 L stainless rather than Type 347 was made since
it has been demonstrated that the possibility of incipient cracking:at the
weld Jjoint is less brobable with Type 304 L. The weldability.of Type 304 L
compares favorably with Type 347 as long as the carbon content is limited to

0.03 maximum.

82,
The vessel is clad with approximately 0.20 inch of type "L" nickel at all
points which are in contact with the fused fluoride salt to prevent corro-
sion. It should be noted that in establishing the thickness of the vessel,
no increase in strength was allowed for the nickel cladding. This is in
accordance with the ASME boiler code. The sides and bottom cf the vessel
will be fabricated from 2-inch plate into sections that will be welded to-
gether to make the complete veséelo The assembly of the top of the vessel
is more complex since the heat exchangers are seal welded onto structural
members and the top ecover section.

5.2.2 Internal Arrangement

The internal arrangement of the reactor vessel consists of three prin-
cipal components: the central core section where mdst fissioning takes place;
the'primary heat exchangers in which the heat of fission is removed; and the
fuel- circulating pumps which provi@e the necessary head to maimtfiin continuous
circulation of the fuel. By referring to Figure 5.9, the fuel circulation path
can. be followed.

Fuel flows upward through the T-foot diameter cylimdrical core;, outward
radially at the top, and down through the first pass of the U-tube heat
exchangers which are located in fhe baffled apmulus surrounding the core. The
fuel reverses its direction at the bottom of the heat exchangers and flows up-
ward thfough the second pass of the heat exchangers. At the top of the second
pass, the fuel flows radially outward again to eight axial flow pumps located
symmetrically around the upper periphery of the reactor vessel. These pumps
force the fuel downward through an annular downcomer which surrounds the entire
vessel. At the bottom of the downcomer, the fuel flows radially inward and

‘back into the central reactor core to repéat the eircuit.

4)

<)

83.

The internal baffling is 7/8-inch Type 304 stainless steel base plate
clad on both sides with 1/16-inch minimum thickness Type "L" nickel. The
inner and outer baffles of the heat exchanger are connected to the reactor
vessel through internal structural members; not shown on the diagrams. The
center baffle .is welded to each primary heat exchanger tube sheet channel and
is supported at the bottom by the radial-support plafies of the heat exchanger
bundleu- Leakuge and streaming of the fuel between the hot and ¢old legs of
the heat exchanger bundles is minimized by using a tongue and groove arrange--
ment at the points where the center baffling ends between adjacent tube
tundlies. In addition; special internal baffling is required to prevent fuel
streaming between bundles in the voids caused by the radial structural members
at the topr of the vessel which support the heat exchangers. Perforatiofis are
provided at the bottom of the U-bend of the baffling to permit complete drain-
age of the fuel, and at the void spaces around the fuel pump discharge transi- .
tional sections to avoid overheating in this stagnunt volume.

Eight fuel drain and fill lines are provided at the bottom of the vessel
and are connected to. the fuelldump system through let-down valves and safetly
devices.

5.2.3 Structural Arrangement .

The reactor vessel is supported by eight vertical members embedded in

- the concrete floor of the containment vessel. These members are welded to

the outer structural ring beam which acts as the main support member for

the reactor vessel. Twenty-four equaliy_spaced beams extend radiaily inward
from the outer ring beam and are in turn welded to an inner concentric.ring

beam. The resultifig design is a ¢ircular cantilever functional arrangement

similar in appearance to a large wheel with 24 spokes between the hub and rim.

8l.

The primary heat exchangers are located between these radial members. The heat
exchangers are seal welded individually by radial and concentric seal strips.
The radial sea; strips are located on each side of the heat exchanger channels
and are welded to the radial beams running between the heat exchangers. The
concentric seal strips aré located at the iwmer and outer periphery of the heat
exchanger channels and are welded.to the concentric ring beams to complete the
closure. These seal strips are also designed to restrain the heat exchangers
from the upward force caused by the buoyant effect of the fused salt on the
tube bundles.

This design is unique in that it utilizes the inlet and outlet ch;nnels
of the primary heat exchanger as an integral part of the reactor veaéel°
- Ffirthermore, the primary heat exchangers can be removed by remotely cutting
out the seal rings of any particular bundle and lifting the entire bundle
out vertically after the primary coolant piping has been cut away.

5.2.4 Fuel Circulating Pumps

The fuel circulating pumps have several special reguiremcnte that must
be satisfied for a practical design. These requirements are as follows:
(1) absolute leak tightness, since any leakage of the highly radioactive fuel
would conteminate the area and would eventually cause a shutdown; (2) opera-
tional reliability with minimum maintenance for uninterrupted; continuous
service, since all maintenance must be accomplished remotely once the reactor
has been operated; (3) be easily replaceable as a unit since plant shutdown
is required for access; (4) vertical installation fof complete removal of-

pump internals without necessitating complete draifiage of the fuel; and

| {5) reasonable cost. | ‘

In selecting a design for the fuel circulating pump that would reason-

ably fit the above requirements, two types were considered sufficiently

85,
developed to warrant investigation for thié application: .(l) the totally
enclosed or camnned rotor pump; and (2) conventional motor drive coupled to .
the pump by shafting that has special provision for frozen seals to pre-
vent leakage.

The canned rotor pump met many of these requirements; however? such a
pump has not been fully developed to operate with fused salt at temperatures
in the required range of 1000°F to 1300°F, Thg bearings of a canned rotor
pump are generally lubricated by the fluid being pumped and cooled by a
secondary coolant. Application of a canned rotor pump to a fused salt system
would necessitate that the rotor; sxnd consequently the electrical igsulation,
operate at temperatures considerably ifi excess of those now considered prac-
tical. Coolant passages would require judicious design to avoid freeze-ups
in the coolant circuit. Further, the cost of development and fabricatiom would

be considerable for materials suitable to resist the corrosiveness of the fluo-=

ride salt. For these reasons, this type of pump was rejected. A veriical pump

cbupled to a conventional motor drive with special adaptations of freeze seals
and hydraulic, piston-type, self-lubricating bearings was selected as more
suitable for this application.

Calculations indicated that the total fuel flow is 32,000 gpm. This
requires a total dynamic.pump head of 11 ftor 15 psi. This load is distri-
buted among eight pumps; on the basis that this number resulted in a
éymmetrical arrangement in conjunction with the 24 primary heat exchaqgers;
and also that failure of one of eight pumps would not seriously reduce the
capacity of the plant. Further, the impeller‘aize for a LOOO gpm pump is
reasonable and fits into the chosen pump location. The required'drive‘motor.
rating for each pump unit of 60 hp is practical. A value of 60 percent was
used for the phmplefficiency in the design vhigh-is considered consgrvative

for axial flow pumps.

86.

To determine the pump classification; the specific speed was obtained

from the following expression (Ref. 21, p. 296),

Ns = I::) GPM - | | (5.9)

where Ns 1s the specific speed; N is the revolutions per minute of the
pump shaft, GPM 1is the capacity in gallons per minute, and H is the total
dynamic head of the pfimp in feet. |

Application of the above expression when considering a pump speed of
1750 rpm results in a specific speed of approximately 18,000, which is in
the axial flow pump range. Use of an axial £low pump simplifies the arrange-
ment of the pump discharge nozzle and transitional pieces since volute
diffuser sections are not required and,; therefore, both suction and discharge
can be axial. |

In order to reduce the possibility of cavitation for the top suction
pump design, it was decided to provide a pressurizer forrthe fuel in the
reactor vessel. The'pressurizer will maintain the fluid level several feet
above the pump suction.

The location of the pump in- the reactor vessel is critical since the
maximum shaft overhang tolerable is of the order of 18 inches. For this
reason,; the pumping element is located as near to the top of the vessel as
flafi permits. A new development in pump bearings, called the hydraulic,
piston-type bearing, permits the bearing to operate in the fluid and is lub- -
ricated by the fluid being pumped. This design would have application in
the pump installation proposed, as it is specially designed for high tempera-
ture corrosive fluids. To prevent leakage, a freeze seal is proposed.

Essentially, this consists of a-section of the pump shaft where cooling coils

87.
freeze the fuel around the shaft, thereby_preventing fluid leakage. These
-seals have been developed and are considered practical.

In order to prevent backflow through a pump which has been shut down; a
check valve is required at each pump discharge. The check valve is located
in the pump discharge downcomer and arranged so that it can be removed when
the pump internals are withdrawn. The check value must offer a minimum resis-
tance to flow sou tfiat natural circulation will not be restricted on complete
eleectrical failure of the pump motor drives.

The choice of suitable materials for the fiump internals is 1imited at
the present time to Inconel. Experiments (Ref. S57) have been performed with
an Inconel pump operating in a forced cireulation loop with fluoride fuel
Noo 30 at a cold cperating temperature of lQBOOF and with a maximum loop tem-~.
perature of lSSOOF° The pump was successfully operated for over 8000 hours, -
when an electrical power failure forced a shutdown and restart was prevented
by a pipe break. It was considered that the pump was cepable of operation o
for many more hours without mechaniéal failure. The pump was equipped with
a conventional seal of §ilver impfegnated graphitar against tool steel and
with commercial ball bearings. This seal and bearing arrangement is different
from that proposed above. On inspection of the pump impeller, it was con-
cluded that the effect of corrosion and mass transfer on the pump materials
vas negligible. On the basis of this experience, the use of Inconel for pump
internals is considered satisfactory for the design and operating conditions
planned for the fuel circulating pumps for this design study.

A summary of the specifications for the fuel circulating pumps is given

in Table 5.2.

TABLE 5.2

SPECIFICATIONS FOR FUEL CIRCULATING PUMPS, (8)

Type-Vertical-Axial flow B
Capacity 4000 gpm .
Temperature of salt entering pump 1050°F ‘
Total dynamic pressure head 15 psi

Speed - Constant | 1800 rpm

Pump horsepower | 56 hp

Motor horsepower 60 hp

Impeller outside diameter 10 in,

Hub diameter - 5 in.

Materials - Trim ' ‘ . Inconel

Seal Freeze Type

Bearing Hydraulic Piston Type

éfigo

5.2.5 Pressurizer (Gas Removal and Expgnsian System

The reactor vessel is pressurizéd to a maximum pfessure of 10 psi. The
pressure is limited to this value to keep the design pressure below 50 psi in -
the reactor vessel. Helium is used to pressurize the system because of its-
negligible solubility and inertnesé in the fuel sclutiom. Pure helium does
not become radiocactive and therefore can be vented di%ectly to the atmosphere.
(Ref. 15, p. 115).

The pressurizef syéfiem is required for several important reasons:

1) To maintain a positive flooded suction head on the fuel-circulating
pumps to prevent pump cavitation and vapor binding.

2) To provide gas pressure for the fission product gas removal systemok

3) To minimize the possibility of the "snow problem" described in
Chapter 2.

4) To reduce the tendency for formation of fission product gas voids
in the circulating fuel.

5) To provide a free ‘surface for release of the fission product geses
which will accumulate in the helium volume at the top of the pres-
surizer vessel.

6). To serve as a surge tank for changes in the fuel volume.

The pressurizers shown on Figure 5.11 consist of eight connected chambers,
adequately sized to accomodate the expamsion and contraction of the fuel fluid |
caused by the variation in dengity between the solidification temperature snd |
13OO°F° .Above the free surface in the pressurizer, édditional volume is avail-
able for collection of the fission product gases which are scavenged to the |
off-gas system. A pipe connection is provided between each fuel pump at the
hiéhest elevation of the reactor vessel for connecting to the top of the
pressurizer vessel for gas removal. To provide continuous circulation of the

low temperature fuel through the pressurizer, the return line is connected

from the bottom of the pressurizer to a cconnection at each fuel pumfi suction.

90,

5.2,6 Effect of Volume Heat Source

Baffles - and flow dividers are located within the reactor vessel to
di;ect the flow of the fused salt fuel through the heat excha.nger*s,;n.mzps_9
ddwncomer, and back to the central core. These baffles are positioned sc
that they are in direct contact with a volume heat source on’both sides.
Initially, this created some concern because of the possibility of devel§p=
ing excessive temperatures and thermal stresses in metal that is in contact
with a volume heat source. References 42, 43 and 44 analyze this phenomenon
f;n congiderable detail, and indicate that the magnitude of the effect is a
‘ffinqtion of the relative values of neutron flux and coolanfi velocity and
their distribution over fhe flow area. An authoritative source was consulted
on this problem, as it appiies to this design study (Ref. hS)o- It was con-
clfided that the neutron flux and velocity profiles in this reactor are such
that the temperature rise in the baffles due to the volume heat source effect

would be negligible.

9l.

CHAPTER 6.

SODIUM AND STEAM POWER SYSTEMS

*

) 6.0.0 INTRODUCTION

The ultimate use of the reactor hest is the generation of useable power
for distribution to the consumer. In order to accomplish this safely, economi-.
cally, and vith & maximum flexibility of control, it was necessary to investigate
this portion of the plant in detail. This chapter includes the primary loop;
intermediate loop, the steam generators,; the sodium piping, the steam cycle
and auxiliary power requirements. Other items of interest included are the
intermediate heat exchangers, the steam cycle, a plant heat balance and.tfie
plant efficiency. |

6.1.0 ~PRIMARY SODIUM COOLANT LOOP

The heat of fission in the circulating fused fluoride salt fuel can be
transferred to the steam system for the turbine unit by seversl alternate
schemes. The acceptance of a scheme depends on the limitation that under all
conditions of operation; the heat exchange media minimum temperature will
always maintain the minimum fuel temperature in eficess of.the melting point
to prevent local solidification. The minimum temperature qf the reactor has
been set at 1050°F, for reasons discussed in Chapter 5, whereas the boliler
feed water temperature is set at 450°F by the steam turbine cycle design;
resulting in a temperature difference of 600°F. This large temperaiure dif-
%: - ference must be held nearly constant over the entire load range in order to

maintain the system design temperatures which, in 50 doing, creates difficult

problems in controlling the heat transfer rates.

92.

6.1.1 Schemes for Removing Heat from Reactor

The possible schemes for transferring the heat from the fuel are dis-
cussed below.

A. Direct Heat Transfer by Boiling Water -

It is possible to circulate the fuel external to the reactor directly
through a once-through superheater and steam generator without interposing
additional loops in the system. The fuel would undoubtedly flow on the out-
side of the tubes and the steam inside in the heat exchanges. The following
are several apparent disadvantages which make direct heat transfer unacceptable.

l. Large Fuel Holdup - The over=-all heat transfer coefficient‘would be

lower than if transferred to another fluid. This results because the tubes

in the heat exchanger have to be thick, with probably double walls for added
safety, to withstand the high pressure steam and also because the film heat
transfer coefficient for the superheated steam is low. A low over-all heat
transfer requires more heat tranafer area and therefore a bigger heat exchanger,
which results in larger fuel holdup requirementis.

2. Reduced Safety - The larger the heat exchanger and the higher the

préssure, the greater probability of a leak. Any leak would be from the‘high
pressure wvater system into the low pressure, but high temperature fused salt.
- Such a leak would cause a violent reaction, increasing the hazards of plant
operation and containment.

3. Radiocactive Steam - The fuel will activate the steam, which increases

the problems in maintenance, shielding, and designing the turbine and feed -
systems completely leaktight.

4. Decomposition of the Steam - Due to the high radiation field, de-

composition of the steam results, which requires an extensive O,-Hp

recombination system.

93.

5. High Thermal Stresses - Because of the large over-all temperature

drops, excessive tube wall stresses will result.

6. Heat Exchanger Design Problems - Because of the large temperature

difference between the cold legs of the heat exchanger, the effect of film
boiling must be considered in the design with the resulting problems of
instability.

B. Heat Transfer to Liquid Metal

In order to avoid the disadvantages of direct heat transfer, the use of
a liquid metal for a primary coolant was considered. The liquid metal coolant
will flow inside the tubes and the fuel on the outside of the tubes, and the
heat exchanger will be internal to the reactor vessel. Because of the high
thermal conductivity of the liquid metal and low vapor pressure, the heat
exchangers will be much smaller, therefore decreasing holdup. Safety features
will be improved; decomposition and thermal stress problems will be reduced.
Furthermore, existing deeign information can be used with confidence in deter- -
mining the required heat transfer surface and the possibility of operational
instability no longer exists.

6.1.2 Choice of Coolant

In making the selection for the primary coolant, the properties of several
metals such as mercury, sodium, bismuth and NaK were considered. The general
requirements for a coolant, exclusive of nuclear requirements, are that the
density, thermal conductivity and heat capacity should be high, whereas the
viscosity, melting point and vapor pressure shofild be as low as possible. 1In
addition, the coolant should have a high degree of wetting ability, negligible

toxicity and, most important, should not expand upon freezing.

9l

Combining these properties for an over-all comparison, sodium proved to
be the most desirable coolant for the primary loop for the following, several
reasons: . : , . , o
1) Lowest pumping power required for a given rate of heat removal.

2) Superior heat transfer properties, resulting in the smallest
equipment sizes.

3) Contraction upon solidification.

L) Superior neutron economy. Since the fuel primary héat exchanger is
internal to the reactor; it is ilmportant to keep neutron absorption
losses to a minimum. The coolant also acts as a reflector to reduce
neutron leakage from the reactor core.

5) Complete stability under high irradiation. field.

A discussion. of these requirements is extended in Section 6.2.2 of this

report, in the presentation of the arguments for and against an immediate

heat exchanger.

6.1.3 Sodium Coolant Activation

As a result of the pfimary sodium coolant flowing through the primary
heat exchangefs in the reactor vessel; the sodium will become very radio-
active. When the sodium coolant captures a neutron, radioactive sodium-2k
is formed and will build up in the coolant system to an equilibrium value
which depends upon the particular design and flux level of the reactor. The
radioisotope sodium-24 has a half-life of approximately 15 hours, and emits
beta particles and two hard gamma ray photons of an energy of 1.38 and
2.75 mev (Ref. 8, p. 525). -The sodium activation imposes many limitations
on the freedom of the design of the plant since it affects shielding, plant ‘ -
layout, maintenanée, servicing, and personnel agcessibility to the radio-

active areas.

95.

Impurities in the sodium cause long-lived radioactive contaminants and
their effect must be considered in the over-all design, in particular when
draining the system in preparation for access to the shielded compartments.
Care must be taken to assure all such containmepts are completely flushed.
Some of the plant layout and design problems considered becausé of this acki-
vation are discussed in Sections 6.7.7 and T.3.0.

An approximate value for the level ot activity in £he primary coolant

sodium can be determined by applying the following expression when the exposure

time is much longer than the mean sodium-24 half-life (Ref. 15, p,_398).

Naah activity in primary coolant = ¢ SlaV curies

qb = Neutron thermal fiux per cm® * sec

S a - Sodium thermal macroscopic cross section cm™!
V - Total volume of sodium in primary exchanger
Considering only the activities from the thermal flux and uslng an average

flux in the sodium in the primary heat exchanger of 0.8 x’lO12 per cm2~sec,

obtained from the ORACLE calculations, gives an activation of 1.46 x 106
curies of the radioactive sodium. When this activation is diluted by the
approximately 2000 £t3 of sodium in the primary system, the resuitant activity
is 0.0258 curies per cc. |

Since this is considerable activity, it is necessary that all heat
exchangers, pumps and piping containing the primary sodiufi-coolant be ade~

quately shielded to protecthoperating personnel.

6.2.0 INTERMEDIATE SODIUM COOLANT LOOP

It is possible to transfer the heat directly from the primary loop to’

generate steam or to interpose an intermediafe loop between the primary and

(——

96.
steam portion of the system. Several excellent discussions of the advantages
and disadvantages of using an intermediate circuit are given in the literature
(Ref. 7, 15, 25). The primary basis for any such selection must be in terms
of the economics, safety, reliability and operation of the systen.

6.2.1 Justification of Intermediate Loop

In order to mske the decision to include such a loop in the system; the
following advantages and disadvantages of using an intermediate loop were
considered:

Advantages

1. Decreased thermal stress problems in the steam generators.
2. Increased flexibility of the system for partial load operation.

3. Sepafation of the primary heat transfer loop from the steam loop
in case of a leak.

L, Simpler maintenance of the steam loop outside of the shielding.

5. Avoldance of the possibility of film boiling by decreased At in
the boiler. .

6. Reduced shielding problems.

T. Reduced possibility of steam decomposition as a result of gamma
radiation from the primary sodium loop. As is noted in Section
6.4.1, the presence of oxygen and hydrogen in the steam loop would
be very detrimental from a corrosion standpoint.

8. Lighter reactor construction, since the reactor will not be called
upon to withstand the full steam pressure in case of a leak in the
steam generator portion of the system.

Disadvantages

1. Increased piping, pumps and valves for the intermediate loop.
2. More hesat egchangers, |

3. Increased system éize°

4, Larger holdup of heat transfer fluid.

5. More equipment which may be potential socurces of leaks.

97,

From the previousiy mentioned discussions (Ref. 7, 15, 25) and the list-
ing of the advantages and disadvantages, it is apparent that the main advantages
lie in the areas of safety, reliability and ofierafiility, while the principal
disadvantages lie in the increased cost of having such a loop. Safety and
reliability must of necessity govern the design of any reactor system or its
components;, and cost, vhile important, must be sécondary in the design philo-
sophy. On this basis, the choice was made to includé an intermediate loop in
the system.

6.2.2 Choice of Coolant Fluid

The choice of heat transfer fluid to use in the intermediate loop was
readily narrowed down to one of the liquid metals or a fused salt. Organics
or aqueous media were eliminated becauée of the gamma radiation and high tem-
peratures in this loop. Gamma radiation and high temperatures cause these
compounds to deccmpose or dissociate. Some of the variables that were con-
sidered in choosing the fluid were:

1. High heat transfer properties at reasonable velocities

2. Cost

3. Toxicity

k., Relative pumping power. For example, mercury requires from 5 to
10 times the pumping power of the alksli metals.

5. Corrosiveness of the fluid in contact with common materials at the
temperatures of interest.:

6. Tendency to dissociate, combine with oxygen or, in genéral, be
chemically degraded at operating temperatures and in the gamma
flux from the primary loop.

T- In case of a leak, the fluid must be compatible with the sodium
in the primary circuit and also with the fused salt fuel.

8. The materials of construction necessary to contain the fluid must
be compatible with those of the primary loop to prevent the possi-
bility of corrosion or mass transfer. .

98,

9, The melting point, heat capacity and thermal conductivity must all
exhibit reasonable values for a practical system.

10. Handling characteristics must be practical. For example, the
expansion of bismuth on solidification creates many extremely
difficult problems.

11. The boiling point and vapor pressure must be reasonable for this
high temperature application. Mercury, which has good heat transfer
characteristics, has a boiling point of 675°F at atmospheric pres-
sure. In order to take advantage of the high heat transfer
characteristics of this material, the system would have to be pres-
surized to maintain it in the liquid state.

Of the wany materials possible, the alkali metals appeared to be best
suited for this system. As is noted in the Reactor Handbook (Ref. 7, p. TT3-
777), the heat transfer, where the liquid metal is the sole fluid, does not
vary appreciably regardless of the fluid. Sodium was selected as the liquid
metal which fulfills more of the requirements listed above than any other
fluid. However, this material presents one major problem, namely, the

exothermic sodium - water reac;tion°

6.2.3 Sodium-Water Isolation Problem

In,order to preclude the possibility of sodium and water contact,
systems have been designed (Ref. 15) using such design details as double
wall tubes or a third fluid to separate the sodium and water. At present,
" one system (Ref. 5) contemplates the use of single wall tubes and éingle
tube sheets to separate these fluids. A relatively small mock-up of this
system will be tested in the near future and this test shduld provide valuable
information regarding the ability of such a unit to run for a period of time,
comparable to central station steam pracfice, without leaks.

As is noted in Section 6.4.2, it is possible to adequately treat the
feed water and toAproperly design a heap exchanger to preclude almost entirely
the péssibilitj of stress corrosion failure of the tubes or tube sheets. On

the basis that the previous statement represents the present state of boiler

_22,
and heat exchangér practice, then the tube-to-tube sheet Joint must be care-
fully considered. By exercising extreme care and quality control, units.havé
been built for nuclefir applications which, prior to service,'have been made- -
leakfree. fioweverj in service, the tube-to-tube sheet jJoint is subjected =~ -
to a complex stress patterh'and thermal cycling. Under these conditions,
there is always the distinct possibility of having some tube-to-tube sheet
welds fail. The other type of posoible fallure lies in the realm of faulty
tube material being inadvertently used. To prevent this, it 15 necessary to.

make some, or all, tests such as radiography, dye penetrant methods, physical

tests, ultrasonic inspection, helium leak tests and mass spectrometer, both '

prior to and after the tubing is installed in the unit.

6.2.4 ° Proposed Design for Sodium-Water Isolation

of the potent1a1 types of failure mentioned above, the tube-to-tube
sheet weld failure presents an area with the greatest possibilit&_of.occur-
rence in service. As has been mentioned previously in this section, safety
must be a prime consideration in this system. For this reason, the design
detail shown in Figure 6.1 is proposed to minimize the effect of a leak in
a tube-to-tube sheet joint. In essence, it consists of using single wallj

tubes with double tube sheets. A thin walled chamber separates the tube -

sheets and helium would be used to purge this chamber to preclude the possi-

bility of'é sodium-to-air reaction, and also to act ae a monitor to detect

a leak. Should a leak occur at either tube sheet, the heat transfer fluidg
would not mix and the magnitude of fihe leak would be monitored fiy the helium.
If leaks occurred simultaneously at both tube sheets, only small volumes of
materials would be involved and corrective action, such as isolation of the

~

component, could be initiated.

Lt el e TS
e Tt ST
Preliminary work has been initiated* to determine the feasibility of
this type of construction. This work consists of a test section of approx-
imately 400 1/2-inch OD tubes on 3/4-inch pitch braze welded to a 1/2-inch
thick tube sheet on both sides of the tube sheet. Metallurgical and physical
tests will be utilized to check the integrity of the braze welds.

A summary of this test work up to the date of publication of this report
is given in Appendix J.

6.3.0 INTERMEDIATE HEAT EXCHANGER

The reasons for placing the intermediate sodium loop and heat exchanger
in the cycle are given in Section 6.2.1.

In order to maintain reasonable equipment size, six intermediate heat
exchangers were used and therefore each heater transfers 100 Mw of heat from
the primary to the secondary sodium loop. The heaters are of thé U-tube type
to minimize thermal stresses.

6.3.1 Isolation of Radiocactive from Non-Radioactive Sodium

Cince the intermediate sodium loop extends beyond the secondary radiation
shield, the leakage of rédioactive sodium from the primary loop intc the non-~
radiocactive sodium of the intermediate loop through the heat exchanger is
intolerable. It was decided that the intermediate heat exchanger would be
designed as a two-fluid type rather than the more costly and difficult to
manufacture three-fluld type. The non-radicactive intermediate sodium will
be pressurized above the radioactive sodium in the primary loop. Any leakage
would then be into the radioactive sodium rather than out of it. The isolation
design, as discussed and proposed in Sections 6.2.3 and 6.2.4, would be appli-

cable and desirable for the intermediate heat exchanger.

* Work Order iésued by W. D. Manly to P. Patriarca, dated July 16, 1956

ORNL=LR=Dwg,=18178
101.

STEAM @ 1000°F

?QPV |
7z
7\l o8

IMPOSSIBLE TO BACK
BRAZE. PLATE TOO HEAVY.

z -

HELIUM [ ; :
: HOLES TO BE 0.006" LARGER
THAN TUBES iFO. LEAVE
0.003" CLEARANCE.
BRAZE r'_"—' '—"'-——-’— i Rl et __._.‘|
" '
7777 |
|
SODIUM | J|

— i = j__ i S b
/M
SUGGESTEDTEST ;

FABRICATE TEST SPECIMEN FROM 172" PLATE AND 172" 0.D. TUBE
SECTIONS MAKING LARGEST ASSEMBLY WHICH WILL FIT IN FURNACE
AVAILABLE IN BRAZING LABORATORY. ( APPROXIMATELY 400 TUBES.)

FIG.6.l -SUGGESTED TEST OF DOUBLE TUBE SHEET DESIGN
FOR ISOLATION OF WATER AND SODIUM

102.

6.3.2 Calculations for Intermediate Heat Exchanger

The calculations for the intermediate heat exchanger design are given
in Appendix £ and the specifications are listed in Table 6.2, . »
The equations and methods used in determining film coefficients, over-all
heat transfer coefficient, heat exchanger area and pressure drop are given in
Sections 5.1.4, 6.5.1 and 6.5.2.

6.4.0 STEAM GENERATORS

The design of a steam generator for a nuclear system presents several
unusual aspects when compared to conventional central stationm practice. The
choice between a once-through unit or natural circulation boiler with sepa-
rate superheater is not clearly defined for this application. Therefore; in
order to determine the proper selection for this system; both types of steam
generators were investigated.

6.4.1 Once-Through Steam Generator

In principle, the once-through steam generator represents the simplest
and most direct method of transferring the rcactor heat to the steam loop.
It consists essentially of a number of tubes in parallel with feed water
entering the tubes at one end and superheated steam leaving at the other.
The heat transfer fluid, which in this case is sodium from the intermediate
loop, is on the shell side of the unit flowing parallel to the tube axis.
The feed water flow to the unit may be adjusted so that either wet, dry or
superheated steam is obtained at the exit, dependent on the design condi-
tions. In the event that wet steam is produced, it can be dried by use of - »
a relatively small in-line steam separator, eliminating the need for a
steam drum. The principal advantages and disadvantages for this type of
unit as compared to the natural circulation boiler with a separate super-

heater are as follows:

103.

Advantages
1. qggactness - For a given set of design parameters, the unit is

20

more compact..

Cdst.- Since the requirement for a separate steam drum is eliminated
and the amount of interconnecting piping, tube sheets and shells is
reduced, the unit is inherently less costly. '

Control - By the relatively simple method of controlling feed water
flow to the tubes, wide ranges of load can be achieved.

Arrangement - Since-this type of steam generator does pot require a

steam drum and interconnecting piping to a separate superheater, it

is adaptable and ideal for a nuclear plant. layout. In particular, the
problems of shielding or isolation for reasons of safety are minimized,
resulting in simplified arrangements and substantial savings.

Internal Arrangement - The use of low pressure sodium on the shell
side and high pressure water cn the tube side results in the best
over-all heat transfer and a low pressure shell design which is
ideally suited to this reactor system.

Disadvantages

1.

Lack of Water Storage - In the event of complete loss of feed water.. .

flow caused by electrical or mechenical failure of the feed pump,

the once-through unit has essentislly no water storage and provision
must be made to insure the immediate aveilability of an independent
water supply to the unit to provide mesns of decay heat removal from
the reactor.

Water Treatment - Experience with high pressure once-through boilers
in Eurcpe has demonstrated that the major difficulty encounteres.in .
these units has been the inability o maintain extremely high purity -
wvater required for successful operation. According to Shannon
(Ref. 1), investigations have indicated that the oncemthrough unit

can be operated successfully only if the following conditicns are mets::

a. Solids content of feed less than 0.5 ppm -
b. Oxygen content of feed not to exceed approximately O 006 ppm
c. Removal of oxygen formed by the dissociation of steam

d. The absolute exclusion of chlorides from the system

e. pH control by use of aqueous ammonia

It is believed that these conditions can be met when operating the

plant (Ref. 1).

3.

Thermal Stress - The outlet of the evaporating section is particularly
subjected to conditions that lead to stress reversals and the possi-
bility of fatigue or stress corrosion failure resulting from the
intermittent heating and cooling of the tube walls by the finely
dispersed particles of water in the steam at this location.

OGS

104,

L. Flow Instability - In forced circulation boiling, the problem of flow
instability (Ref. 2, 3, 4) in parallel circuits can endanger the per-
formance of the unit. If an obstruction or scale builds up in a tube,
the flow rate in the tube may actually lead to an increasing flow
resistance due to the increase in the average specific volume of the +
tube contents. Ultimately, the tube may become steam bound and become
completely ineffective from the standpoint of constituting active heat
transfer surface. A further effect under these conditions is to de-
crease the total feed water flow to the unit, which also decreases .
its ability to perform properly.

5. Design - The calculation of the heat transfer surface required to
perform a specified duty with a complete change 6f phase is especially
difficult in those sections where there are high percentages of steam
by volume. At present, such calculations represent "educated guesses”,
and the reliability of such calculation procedures must be validated
against future tests such as are presently being contemplated by
APDA (Ref. 5).

6.4.2 Natural Circulation Boiler-Separate Superheater Steam Generator

The natural circulation boiler with a separate superheater requires more
basic components than the oncé-through unit and; therefore; represents a more
complex arrangement for the removal of the reactor heat. Inthe natural circu-
lation boiler, steam is generated on the shell side and low pressure sodium
is on the tube side. The steam-water flow in the boiler is directed by baffles
which also serve as tube supports. The steam-water mixture flows through the
risers to an external steam drum by natural circulation. As in conventional
practice, feed water is introduced into the system at the steam drum to avoid
‘the possibility of temperature shock to the steam generator. The feed water
is heated to saturation by condefising steam in the steam drufi and saturated
water is recirculated to the boiler by the downcomer system. Saturated steam,
mechanically separated in the steam drum, is Piped to the separate superheater.

The principal advantages and disadvantages 6f the natural circulation

boiler with a separate superheater, as compared to the once-through boiler, are:

105.

Advantages

lD

Reliability - This type of steam generation system has been used in
principle in several actual and proposed reactor installations.
Among these are the Mark I and Mark II STR units, the PWR, SIR,; SAR,
HRT, etc. The reliability has been established under actual and
test conditions of reactor operation.

Design - Design methods are available which enable the designer to
predict the performance of these unite to the required degree of
accuracy for engineering of the system.

Water Storage - This system has an inherently large water storage
built intu 1t in the steam drums. Thus, in the event of a forced
outage or the failure of a feed pump, the system is still capable
of removing heat from the reactor for several hours. In addition,
the effect of a perturbation of the reactor part of the system is
minimized on the steam portion of the system because of the damping
effect of the large amount of water storage.

Water Treatment - In both this unit and the once-through boiler,
careful control must be exercised in treating the feed water for
proper operation of the units; however, for this unit, the water
treatment problem is less stringent than for the once-through -
boiler since scaling and tube plugging are not as critical. The
problem of stress corrosion prevails in both types of unit and
needs to be carefully considered.

Inherent Stability - This type of design has established its in-
herent stability under various operational conditions. The type
of instability that occurs in forced circulation boiling does not
exist in the natural circulation system.

Disadvantages

ld

Cost - Because of the necesslty for a steam drum, piping from the
steam drum to the superheater, interconnecting piping in the steam
generator, multiple tube sheets, shells and heads, this type of unit
is more costly than the once-through unit.

Arrangement - The engineering design requirements of the natural .
circulation equipment make it less adaptable to either shielding
or compartment arrangement in a nuclear system.

Size - For a given set of design parameters, the natural circulation
system will occupy more volume than the once-through system.

The natural circulation design is predicated on the basis that the
high pressure water is on the shell side, which subjects the tubes
to collapsing pressure. This results in thicker tubing with an
attendant increase in weight and decrease in heat transfer capa-
bility per unit volume.

106,
5., Control - As described in Section 7.4.0, the control of this unit
presents problems in operating the system over wide ranges of load.
In order to accomplish the desired regulation; it is necessary to
provide elther external regulation of -the reactor or special opera-
tional control of the heat transfer circuits.

- In a previous section of this report (6.2), the problem of a possible
godium-water reaction has been discussed and a tentative scheme has been pro-
posed to minimi?e the pogsibility of this reaction by the use of double tube
sheets. In all of the systems built to date (Ref. 15), sodium and water have
been separated by at leagt two metal walls, with the one exception of the APDA
tests (Ref. 5). The boilers and superheateré for this present sfiudy'have been
designed on the basis of fising a single metal wall to separate the sodium from
the wa.t'.er.T In order to indicate the magnitude of the size change when two
metal walls are used, the superheater of the natural circulation boiler with
a separate superheater was calculated on two bases. The comparison design
was for two metal walls having perfect thermal contact (no interface resistance),
but not mechanically joined. Each wall thickness was taken as being separately
capable of withstanding the full design préssure°

After careful consideration of all the factors involved, it was not possible
| to select either the once-through or natural.circulation boiler with separate
superheater as having an over-all advantage over the other, for the proposed
application. Thefefore, for the purpose of this study, it was considered that
prelimingry designs for both units should be developed. The calculation proce-
dures used in making fhese designs are given below. These designs were made
on the basis that the units would conform to present technology wherever
possible, and no attempt was made to ofitimize the performance of this equip-

ment, as this optimization can be realistically done only after the system

has been finalized.

107.

6.5.0 STEAM GENERATOR CALCULATION PROCEDURES

6.5.1 Heat Transfer

A. Boiling

In making such design studies, it was apparent that the greatest'diffi-
culty was the inadequacy of the information available in the field of boiling
heat transfer. Since these units must operate with large temperature differ-
ences between the fluids, there is the additional problem of the possibility
of unstable film boiling occurring. The literature (Ref. 6, T7) indicates that
a temperature difference of 45CF between the surface causing boiling and the
saturation temperature of the fluid being boiled is the maximum value pogsible
to have nucleate boiling. To a first approximation,'this temperature differ-
ence is independent of pressure. When the above-mentioned temperature differ-
ence exceeds 459F, unstable film boiling results. At the present state of
development, it is Qirtually impossible to design engineering eqfiipment with
any assurance of predicting performance when in the unstable film boiling
region. In Appendix F, é check calculation was made to assure that these
units were operating in the nucleate boiling range.

A solution to the problem of designing & heat exchanger with variable
fiucleate boiling coefficients 1s available (Ref. 8, p. 699). It is exceed-
ingly difficult to utilize this for design, s}nce it 1s necessary to use
iterative techniques to obtain a single soiution° Therefore, an anaiysis
was made in Appendix F to determine a reasonable value of an average boiling
plus scale coefficient to use for fhe design of these units. It is to be
noted that the operating temperature différences in the units analyzed were
less than the temperaturé differences in this study. The numerical value of

the boiling plus scale coefficient was 2000 BTU /hr/sq ft/F,Aand since the

108,
temperature differences in the equipment will be greater than for the units
analyzed, this value is conservative for design. This value of 2000 was used
in all cases where there was nucleate boilifig,

B. Tube Resistance

The resistance of a tube wall to heat transfer based on the outside
tube surface is,
ro In ro : | (6°l)
k Ty '
where the thermal conductivity of the tube material is taken to be the value

at the average tube wall temperature, as given in Figure A.l.

C. Superheated Steam Resistance

Colburn (Ref. 9) has correlated the heat tranéfer data on many fluids
and his work (or winor modifications) has been the basis for most of the corre-
lations for fluids in turbulent flow inside tubes. The greatest difficulty |
in fising any heat transfer equation lies in the corfect values of the physical
properties of the.fluid to use. In order to correctly apply these équations,
it is necessary to use the same properties as the correlator used. At times,
this is difficult to ascertain, and in these instances,; the designer must
exercise Jjudgment in the choice of physical properties. lThere has been, and
still is (Ref. 10), a great deal of controversy regarding the properties of
superheated steam. It is the general opinion (Ref. 11, 12) that the properties
of steam given in Appendix A represent the "best" values available at this time.
.These values were used in a recent investigation (Ref. 13) of the flow of super-
heated steam in annuli. The recommended equation for the heat transfer

coefficient when superheated steam is flowing-turbuI§ntly in pipes is based

109.

on the properties in Appendix A, and ia,'

hd - .023 (gfi)lk} aveP-8 | (6.2)
k k A

where all properties are evaluated at the avefage bulk temperature. Since
this is‘based on the inside tube surface, it is necessary to multiply the |
h from the above equation by the ratio of the inside tube radius to the
outside tube radius to base h on the outside tube surface. The resistance
is the inverse of h . Equation 6.2 has been plotted in convenient form for
both subsaturated water and superheated steam in Appendix B (FigurelB.l)°
This form was suggested by Dr. G. E. Tate (Ref. 14).

D. Sodium Resistance

a. Sodium flowing inside tubes - In the natural circulation boiler,_.
sodium is on the tube side. In this case, the most accepted value of the
sodium film coefficient is given by the Lyon-Martinelll equation, which is
listed as Equation 5.1 in Section 5.1.4. It is noted that the sodlum coeffi-
cients are generally so high as to constitute only a smail part of the overQall
fesistance of the equipment.

b. Sodium flowing on the shell side - Several correlations have been
proposed for calculating the film coefficient when godium is floying turbu-
lently on the shell side of baffled and unbaffled heat exchangers (Ref.‘ 15, 25).
Kern (Ref. 20, p..136) notes the difficulty of calculéting the heat transfer
coefficient on the éhéll gide of a baffled heat exchanger, even when common
industrial heat exchange fluids are used. Kern gives a éemi-empir}cql equation
of the Colburn type (Ref. 9) and states that "....excellent agreement is
obtained if the hydraulic radius is calculated along instead of across the long

axis of the tube....". For the range of Reynolds' numbers involved, Dwyer

0.

(Ref. 25, p. T78) recommends the use of the Lyon-Martinelli equa.tion° This
was used and all physical properties were evaluated at thé average fluid
£ulk temperature.

The reciprocal of the sum of all the above applicable resistances is
the over-all coefficient of heat transfer for the equipment. In general; the
over-all coefficient of heat transfer will vary throughout the length of the
unit since it is temperature dependent. In such cases, Equations 8-17 of
Reference (6) should be used to determine the surface required for a given
set of design paraméters. For this study, a constant average value of
over-all coefficient of heat transfer was used, since the variation of
fluid‘properties is not great and the error so incurred is well within the
accuracy of the heat transfer correlations used. The surface was then calcu-

lated from Equation 6.k4.
q =Uy x8 x Aty (6.4)

6.5.2 Pressure Drop

According to Reference (15), "The use of liquid sodium and NaK as heat
transfer media involves no new problems in the application of fluid mechanics”,
Therefore, the calculations of the various pressure drops in the equipment
resolved themselves into identifying the various resistances involved and
applying known. -solutions. Pressure losses due to nozzles and interconnecting
piping are not included as part of the equipment pressure drop. These losses
are included in the piping pressure drops; however, the calculation procedures
of this section were used to evaluate the piping pressure losses.

A. Entrance (Contraction) Losses (Ref. 16)

The entrance loss was taken to be independent of Reynolds number and

also independent of the ratio of the'tube flow area to the header flow area.

111.°

.

This loss is given below (Eq. 6.5), where the:velocity is the average‘tube
velocity as determined from the one-dimensional, steady state continuity
- relation (Eq. 6.6),
AP - 0.5 !3 (psi) | (6.5)
2g 1 '
VoW | | | (6.6)

B. Exit Loss (Ref. 16)

It was assumed that there would be no regain in pressure at a tube
exit and the decrease in static pressure at the exit would be one veloéity

head,

AP - 1.0V2 @ (psi) (6.7)
2g 1LL ' _ _

C. Friction (Single Phase Incompressible) (Ref. 16)

The friction factor uscd in the Fauning equation is a function of the
relative roughness of the tubing and the Reynolds number. The relative
roughness of the tubing was taken to correspond to fhe smooth pipe curve of
the Moody diagram. The friction factor was taken directly from the Moody

diagram at the proper Reynolds number,

oL v @ (pst) | . (6.8)
AP = f 3 5z I pe : ,

D. Friction and Acceleration (Two Phase) (Ref. 7, p. 69)

The method of Martinelli and Nelson was used to calculate the friction 7
and acceleration pressure losses in the boiling section of the once-through

boiler.

112,

E. Friction and Acceleration (Single Phase Compressible) (Ref. 7, p. 69)

The acceleration and pressure drop for single phase compressible flow
such as occurs in the superheater can be calculated from Equation 23 of Refer-
ence (7), page (69). The preliminary calculations of these pressure losses
showed that they were of the order of 3 percent or less of the system pressure.
Therefore, the arithmetical average density was used and the flow was treated
as before (Section 6.5.2, Subsection C), neglecting the acceleration pressure
drop (Ref. 21, p. 234).

F. Bend Losses (Ref. 16)

The bend loss pressure drop is known to be & function of at least three
variables. These are relative roughness, Reynolds number, and the ratio of
the radius of bend to tube diameter. The value of this loss for 180° vend

(Ref. 16) was taken to be 1/2 velocity head,

p.o.5s¥ £ (pst) | (6.9)

Shell side preséure drops were caléulated ou the same baoio as the tuhe
side pressure drops above. The equivalent diameter concept (Ref. 6) was used
based on typical unit cells of the tube layout pattern. No attempt was made
to estimate the effect- of tube baffles on the shell side pressure drops.

The drum size, its location above the stesm generator, and the unit's
abllity to maintain an‘adequate‘circulation rétio vere investigated ofily from
-8 feasibility standpoint‘(Ref° 22). The results only of this pfeliminary
calculgtipn are tabulafied on ihe equipment data ;heet for thexnatural circu-
1ati§n uni#. | | |

In ;ll instances, a minimum distance of 1/4 inch was allowed between

adjacent tubes at the tube sheets, to permit the welding of the tubes into

113.
the tube sheets. Calculafiions of the péssible relative expansion of the tubes
and shells of these units (Ref. 24) indicated that a straight tube exchanger
with fixed tube sheets would not be suited for this application. Therefore,

the "horse shoe" shaped unit was used for all of these exchangers.

6.6.0 SUMMARY OF HEAT EXCHANGER DATA

The specifications for the primary heat exchanger, intermediate heat
exchanger, once-through steam generatér, convection boiler, single tube wall
superheater, and double wall superheater are included in Tables 6.1 through 6.6.

6.7.0 SODIUM PIPING SYSTEMS

The technology of sodium piping is well established (Ref. 15), and it is
considered that the piping systems proposed in this study can be designed satis-
factorily. Rough piping designs and layouts have been prepared and the follofiing
sections briefly describe the considerations which were used to Justify these
preliminary designs. Qertain design criteria, which are not shown on the lay-
out drawings or in the tabulations, are also discussed.

6.7.1 Pipe Length and Size

AY

In estimating pipe length, an attempt was méde to provide'a genefous
~ allowance for expansion loops. Figures 7.2 and 7.3 were used'to make,reasonmr
able estimates of over-all pipe lengths.

Pipe wall thickness was chosen on the basis of'commercially available
pipe. Sin;e the sodium system is a low pressure system, Scheduié 40 pipe was
generally selected. Pipe diameters were influenced by the-chqicé of sodium
velocity. -Since 30 fps to 40 fps sodium velocity is being cénsidered for desigfi
purposes (Ref. 15), all pifie slzes were selected so that the velocity_did not
exceed 40 fps and, in most casés, the velocity was near the 30 fps value.
Wherever excessive pressure drops were encountered in the initial estimate;

the pipe size was increased.

11"' L]
TABLE 6.1

EQUIPMENT FOR SPECIFICATION SHEET

Name of equipment: primary heat exchanger
Description: U-tube bundles

Service of Unit: fused salt to sodium

No. of units required: 2k

Connected in: parallel

Surface per unit: 2230 £42

PERFORMANCE OF ONE UNIT

Shell Side Tube Side
Fluid circulated fused salt sodium
Total fluid entering, 1b/hr 48 x 10° 5.7 x 106
Temperature in 1200°F 1000°F
Temperature out | 1050°F 1150°F
Operating pressutre 10 1b/sq in. 125 1b/sq in.
Velocity 3/7 ft/sec 12.8 ft/sec
Pressure drop (not including nozzles) 8.5 psi 9.1 psi
Heat exchanged, BTU/hr 9,09 x 107 =
Log mean temperature difference SOOF
Transfer rate design - - 815 BTU/hr-£t2-°F

¥ Based on a tube side surface of 2230 ft2 per bundle
CONSTRUCTION

Tubes: .O42 inch wall type ALSL 304 clad with .023 inch type "L" nickel;
No. 650; 0.D.,5/8 inch; I.D.,4#95; Effective Length 21 ft; Pitch 3/4 inch

Tube sheets: stationary

Baffles and tube supports: as required
Code requirements: ASME, TEMA and others
Remarks: shaped for expansion |

MATERTALS
See Chapter 3

115.
TABLE 6.2

EQUIPMEN? SéECIFICATION SHEET
N Name of equipment: intermediate heat exéhanger
Description: horseshoe shépe, P1xed tfibe shéét exchanger
Service of unit: sodium-to-sodifim |
No. of units required: 6
Connected in: parallel
- Surface per unit: 1910 f£t°

Shells per unit: 1

PERFORMANCE OF ONE UNIT :
Shell Side Tube Side

Fluid circulated. sodium sodium
Total fluid entering, 1b/hr 7/62 x 106 2.67 x 105
Temperature in 1150°F 675°F
Temperature out | 1000CF ~ 1100°F
- Operating préssure | 50 1b/sq in. 150 l‘b/sqvin°
Velocity - | . 12.29 ft/sec. 13.73 ft/sec
_ Pressure drop (not including nozzles) - 1.61 psi 9.02 psi
Heat exchanged, BTU/hr - | 3.41 x 10°
B Log mean temperature difference ' 1&7QF
) Transfer rate design o 1216 BTU/hr-ftz-QF
. o CONSTRUCTION
Tubes: No.:1100 0.D.:1/2 in. I.D.:O.h}6 in. Length: 13.3 £t Pitch: 3/4 in.
.~  Shell (approx. dia. & length): horseshoe shape, 30 in. dia.,, 16 ff developed length

Tube sheets: stationary
Baffles and tube supports: as required
Code fequirements: ASME , TEMA and others

Remarks: shaped for expansion

MATERIALS -

See Chapter 3.

116,
TABLE 6.3

EQUIPMENT SPECIFICATION SHEET
Name of equipment: once-through steam generator
Description: horseshoe shaped, fixed tube sheet exchanger o
Service of unit: sodium-to-water No. of units required: 2
Surface per unit: 10,600 £t° Shells per unit: 1 Connected in: parallel

PERFORMANCE OF ONE UNIT :
Shell Side Tube Side

Fluid circulated ' sodium water )
Total fluid entering, 1b/hr 7.93 x 10° 9.78 x 107 -
Temperature in 1100°F 450°F
Temperature out 675°F 1000°F
Operating pressure | 100 1b/sq in. 1950 1b/sq in.
Velocity 13.25 ft/sec  5.14-67.3 ft/sec
Pressure drop.(not including nozzles) | 48 psi 50 psi
Heat exchanged, BTU/hr | | 1,02 x 107 -
(preheat 178.5°F
Log mean temperature difference (evaporating 221 .0°F -
' (superheating 192, 5°F
- (preheat 690 BTU/hr-ft2 -OF
Transfer rate design (evaporating 700 BTU/hr-ft=-OF

(superheating 300 BTU/hr-ft€-CF

CONSTRUCTION
Tubes: No. 1,620 0.D. 1/2 in. I.D. 0.37 in. Length 50.ft Pitch 3/4 in.

Shall (Approx. dia. & length): 34 in. I.D. 26 £t - 6 in.
Tube sheete: stationary

Baffles and tube supports: as required

Code requirements: ASME, TEMA and others

Remarks: shaped for expansion

MATERIALS

See Chapter 3.

117.
TABLE 6.4
EQUIPMENT SPECIFICATION SHEET
Name of equipment: convection boiler
Déscription: horseshoe shaped, fixed tube sheet exchanger
Service of unit: sodium-to-water No. of units required: 2

2

Surface per unit: 10,600 ft Shells per unit: 2 Connection in: parallel

PERFORMANCE OF ONE UNIT

_ Shell Side Tube Side

Fluid circulated N water sodium
Total fluld entering, 1b/hr 9.75 x 10° - 8 x 10°
Fluid vaporized _ water — '
Temperature in L450°F 960°F
'Temperature out 635.8°F 675°F
Operating pressure 2000 1b/sq in. 100 lb/sé in.
Velocity ‘ | None | ' 21,9 ft/séé
Pressure drop (not including nozzles) cmmea 38.6 psi
Heat exchanged, BTU/hr 6.86 x 108 |

Log mean temperature difference 1359F

Transfer rate design ' 480 BTU/hr-fta-QF

- CONSTRUCTION
Tubes: No. 2000 O.D. .75 in. I.D. .42 in; Length 28.5 ft Pitch 1 in.
Shell (approx. dia. & length): L47 in. I.D. | 15.8 £t |
Tube sheets: sfiatiopary | |
Baffles and tube supports: as required
Code requirements: ASME, TEMA and others
Remarksf shaped for expansion |

MATERIALS

See Chapter 3.

TABLE 6.5

EQUIPMENT SPECIFICATION SHEET
Name of equipment: superheater, single wall

Description: horseshoe shaped, fixed tube sheet exchanger | -

" Service of unit: sodium-to-steam No. of units required: 2 X
Surface per unit: 3520 ft2 Shells per unit: 1 Connection in: parallel
- PERFORMANCE OF ONE UNIT
Shell Side Tube Side
Fluld circulated | sodium steam )
Total fluid entering, lo/hr - ' 8 x 106 9.75 x 107 -
Temperature in : . 1100°F 635089F
Temperature out 960°F 1000°F
Operating pressure 100 1b/sq in. 2000 1b/sq in.
Velocity 17 ft/sec T4.8 ft/sec
Preesure drop (not including nozzles) 1 psi Lk fisi
Heat exchanged, BTU/hr | 3.34 x 108 o ST
Log mean temperature difference | : 191°F )
Transfer rate design 500 BTU/hr-£t2-OF

CONSTRUCTION
Tubes: No. 1275 0.D. .5 I.D. .37 Length 22.25 Pitech .75
Shell (a.pproxo'dia,o & length?s 32 im. IOD; 2.5 £t | ;
Tube sheets: staticnary | _ -
Baffles and tube supporté: as fequired |
Code requirements: ASME, TEMA and others
Remarks: shaped for expansion |

MATERIALS ) B z‘_ -

See Chapter 3.

i
TABLE 6.6
EQUIPMENT SPECIFICATION éHEET
) Name of equipment: superheater, double wall
- Descriptions horseshoe shaped, fixed tube sheet exchanger
Service of unit: sodium-to-steam . No. of units required: 2 .

Surface per unit: 6400 fta, Shells per unit: 1 Connection in: parallel

PERFORMANCE OF ONE UNIT .
Shell Side Tube Side

— Fluid circulated ' . sodium . steam
Total fluid entering, ib/hr | 8 x 10° 9.75 x 105
Temperéture in 1100°%F 635.8°F -
Temperature out 960°F 1000°F
Operating pressure ' | 100 1b/sq in. 1950 1b/sq in.
Velocity | 10.2 ft/sec 63.5 ft/sec
Pressure drop (not including nozzles) b psi 4O psi
- Heat exchanged BTU/hr o 3.34 x 108
. Log mean temperature difference l9qu
| Transfer rate design | 275 BTU/hr-fta-oF‘

CONSTRUCTION

Tubes: No. 1500° 0.D. .666 I.D. .500 Length 25.75 £t Pitch .916
,500 .370

Shell (approx. dia. & length): 42 in. I.D. x 14.3 ft
Tube sheets: stationary
- Baffles and tube supports: as required

Code rgquirements: ASME; TEMA and others

Remarks: .shaped for expapsion

MATERIALS

See Chapter 3.

120.

ST

The important piping data, length, size, velocity and pressure drop are
given in Table H.1

6.7.2 Pipe Slope

In laying out the sodium piping system, a continuous slope of 5/8 inches
per foot will be used throughout. The purpose of this slope is toc facilitate
charging, venting, draining and washout. Where traps or pockets are required,
individual draifi lines will be provided at these locations. These drain lines
are not shown in the piping layout. All of the sodium piping, with the excep-
tion of the primary heat exchanger tube bundles and a portion of the feed lines,
can be drained by gravity.

6.7.3 Expansion Tanks

An expansion tank will be provided in each circuit of the primary and
intermediate sodium loops, since their design is based on & closed loop system.
The location of the expansion tanks will be at the highest point in the cir-
cuits and in the vicinity of the pump suctions to take advantége of pressure
drop in the system, to-minimize the cover gas pressure.

Computations were made to determine the size of tanks necessary to
accomodate the expansion of a full hystem of sodium from a temperature of
208% to 1200°F. This consideration indicated that tanks approximately 4 ft
in diameter and 5 1/2 ft long are required for each circuit. To reduce the
size of these tanks, provisions are included for draining sodium back to the
drain and charge tank as the system is brought up to operating teuperature by
providing an ovérflow line near the top of a reasonably sized expansion tank.
An electromagnetic pump will be used in this overflofi Bystém to add sodium
to the expansion ténk during operation to maintain a minimum safe fluid level

in:the system.

4

121.

6.7.4 Cold Traps

Cold traps are located in the cold leg of each closed sodium loop and
on the suction side of the pumps when pcssible. Either static or circulating
ccld traps could be used.

6.7.5 Valves

A minimum number of valves are shown on the General Piping Layout,

Figure 7.1, and the design presented is considered practical and operative;
however, the final philosophy of the design dictates the exact number and
location of valves in the Bodium piping. . A careful evaluation of the operating
and maintenance procedures to be followed would determine the optimum number of
valves and their location.

In general, low pressure drop valves.will be used; and freeze stgm valves
will be used in preference to bellows sealed valves because of the large pibé
gizes required. Bellows sealed valves may be desirable in some locations for
small pipe sizes. Check valves will require no stem seal.

6.7.6 Pressure and Instrument Taps

No consideration was given to pressure and instrument taps since this
would require going into more detail than time would permit.

6.7.7 Drain and Charge Tanks

Drain and charge tanks will be provided for eech circuit of the primary
and intermediate sodium loops. These tanks will be located below the lowest.
point in the circuit so that draining can be accomplished by using gravity
assisted by gas pressure.

All components of the sodium system are drainable, except the primary

heat exchangers and portions of the feed lines connecting the heat exchanger

bundles and the ring headers. It is considered that the sodium level remaining

1122,
in the feed lines after the system has been drained can be lowered sufficiently,
for maintenance purposes, by applying gas pressure to the upper ring header to
force the sodium into the ldfier ring header, from which it can be drained to
the tank.

Six tanks, approximately 6 feet in diameter and 15 feet long, will be
installed in the priméry loop;‘ These tanks will be located in shielded com-
partments directly below the intermediate heat exchangers. Valving and other
provisions will be included so that all radioéctive sodium can be drained and
fluéhed with non-radioactive sodium from the piping, intermediate heat exchangers,b
and primary loop pumps in the intermediate compartments during outages for main-
tenance work on the above-mentioned compcnents.

Three tanks, approximately 8 feet in diameter and 15 feet long, will be
installed in the intermediate ioop; These tanks will contain non-r;dioactive
sodium and will be located below the steam gefiérators outside the reactor con-
tainment vessel.

It has been estimated that the total sodium volume in the primery and
intermediate loops 1is approximately 2000 ft3 and 1500 ft3, fespéctively° At
208°F, the:melting point of éodium, this volume represents a total sodium
mass of 202,650 pounds which will be used in these two coolant loopé° Adai-
tional sodium will be kept in the drain and charge tanks to make up fortlosaes
during opefation and  to reduce the'thermal éhock involved in d:ainihg a hot
system by allowing the cold sodium in the tank tO‘fiix with the hot_sodium.
(Note: The above estimate of sodium volume in the intermediate loop was
made considering the straight-through steam generator only.)

L

6.7.8 Cover Gas

Cover gas, probably purified heiium,'is to be provided_over the sodium‘":

surfaces in the expansion tanks and drain and charge tanks. A gas equalizing

123
Iine will connect these tanks to facilitate gravity drainage. Valving will
also be provided in this connecting line to allow gas pressure to be directed
to the expansion or drain tanks for forced drainage or charging operatiqns°

6.7.9 Sodium Pumps

Centrifugal pumps will be used for pumping the large quantities of sodium
required in the coolant systems. Six pumps operating with a head of 200 feét
at design flow of 18,550 gpm will be used in the primary loop system. Two pumps
operating with a head of 140 feet at a design flow of 18,350 gpm are required
in the intermediate loop. As shown in Figures 7.l and 7.3, a spare pump is
to be installed in the intermediate loop. These pumps will have rated load
motor horsepower requirements df 1335 and 950, respectively.

It 1s possible that design changes can be made in the two coolant ld&ps
8o that identical pumps can be used in the primary and intermediate coolanfi
lecops. Such interchangeability would fie desirable from maihtenanee considera-~

tionc and stocking of spare parts.

6.7.10 _Heating the Salt and Sodium Loops for Startup

Because of the large quantities of fused salt and sodium required for
the various loops, there are many profilems involved in designing methods for
preheating the systems, charging the systems and maintaining approximately
design operating temperatures with thé reactor subcritical during startup
procedures. The design of the heatiné accessories ié bfised on the concepts
that adequate heéting capacity will be inclfided to: (1) melt the sait.mixe.
ture and sodium in the primary loop; (2) heat and maintain them at a minimum
temperature of lOBOqF while continuofisly‘circulated in tfieir réspeétive
systems;.and (3) melt tfie sodiu& fof the intermediate loop -and heat it to

approximately 250°.F°

124,

At the present time, induction heating is the only possible method of
preheating the primary sodium and salt mixture up to 1050°F, since this tem-
perature is above the temperature range of resistance heating. The use of
induction heating with an austenitic stainless steel piping system and reactor
vessel requires high frequencies and is quite inefficient (Ref. 15, p. 255).
Utilizing this method of preheating would require development work before
application.

Both systems will be charged by using pressurized helium. Extreme care
must be exercised to assure that all the oxygen has been purged from the system
prior to charging.

The following procedures were considered as possible methods for heating
the systems prior to initial operation:

1) Initially, the reactor vessel is internally heated to at least

250°F to prevent the sodium from solidifying in the primary
exchanger during the charging of the sodium.

2) The sodium for the primary loop is melted, charged, and then slowly
brought up to 1050°F by its heating system and friction heat of the
pumps in the systen.

3) The sodium is continuocusly recirculated for a long period until it
is determined that the reactor vessel internals have reached a tem-
perature of at least 1000°F by the sodium circulating through the
primary heat exchangers.

4) During this period, the salt mixture in the fuel dump tanks is
melted and the salt piping system and reactor vessel are heated

to 10509F, in preparation for charging.

5) When the complete salt system has reached a temperature of 1050°F,
the salt is charged into the reactor vessel.

6) The fuel circulating pumps are started and the fuel is circulated
in preparation for uranium fuel addition to bring the reactor to
the critical condition. During this latter operation, the salt
and primary sodium systems are maintained at 1050°F by their indivi-
dual heating systems.

125.

7) The intermediate sodium system is heated to 250°F and charged when
criticality is reached, and means of heat removal from the reactor
system is necessary. Charging the sodium in the intermediate loop
last, keeps the high temperature systems isclated from the steam

. system heat sink for the startup preheating. This reduces the pre-

heating requirement for this intermediate loop, since it will then

require only enough heating capacity to maintain the metal tempera-
tures above the melting point of sodium to prevent solidification

. _ during charging.

6.8.0 STEAM POWER SYSTEM

A conventional regenerative steam cycle was chosen for this fused salt
reactor power plant. In this type of cycle, steam is extracted from the tur-
bine in various stages of expansion and is used to heat feed water (Ref. 23).
This reduces the amount of heat converted to work in the turbine, but there
is an even greater reduction in the heat rejected to the condenser cooling
water. This results in an increase in efficiency over that of a cycle with
no steam extraction for feed water heating.

The over-all heat balance diagram and the conditions of the working
fluids are shown in Figure 6.2. The results of the heat balance are given
in Table 6.8.

6.8.1 Steam Turbine and Steam Cycle Heat Balance

The stefim'conditions entering the turbine were set at 1800 psig and
- 1000°F. These are standard and conservative. |
The turbine exhaust pressure was set at 1.5 in. Hé, which is easily
attainable with a reasonable condensef cooling water temperaturé° it is pro-
béble that in most locations the exhaust pressure.could‘be lowered to 1 in.
. ‘ Hg or-léss, with a resulting increase in the cycle efficiency. The temperature
and quantify of cooling water available are a major factor in determining the
- location of a steam plant. If a fiatural supply of cooling Qater is not a#ailm :
able, it is necessary to resort to a cooling tower, with a resulting increase

in capital cost and operating expense.

126,

The turbine-generator losses were found by the method deééribed in Refer-
ence 41. In order to apply this method, the type and size of the turbine must
be known. It was known that the size of the turbine would be between 200 and
250 Mw, and from this it wae decided that a 3600 rpm single shaft turbine of
the tandem compound triple flow type with 26-inch last stage blades would be
used for this study. This type of turbine has a high pressure section and
three low pressure sections on one shaft. All the steam passes through the
single high pressure secfiion and then is divided among the three low pressure
sections. This allows smaller wheel diameters than those of a single low
pressure section. A detailed econcwic study would have to be made to ascertain
whether the higher efficiency of a cross compounded double flow turbine would
Justify the higher capital cost. This unit would have a high pressure 3600 rpm
section on ore shaft and two low pressure sections at 1800 rpm on another shaft.

For a steam cycle in the EOOQESO Mw range, i%t is the practice to have six
or seven feed-water heaters; with the water leaving the last heater and enter-
ing the boiler at 450°F (Ref. 41). 8Six feed-water heaters and a final feed
water temperature of 450°F were assumed for this study. It is quite possible
that a final feed water temperature in excess of'h50°F would be desirable. A
detailed economic study would show if the increase in c¢ycle efficiency due to
a seventh feed-water heater justifies the additional expenditure for the heater,
controls, piping and maintenance. The temperature rise of the feed water was
taken’to be apprdximately the same in each heater.

The fioo 3 heater was made the open heater in this cycle. In an open heater,
the feed watef, extracted steam, and drains from higher pressure heaters are
mixed and any dissolved gases are driven out of the mixture of steam and satura-

ted water. These gases are remcved from the cycle by venting of -the atmosphere.

127.

All other feed-water heaters were taken to be of the closed type with no
contact between the feed water and extracted steam.

Two pumps were placed in the cycle. The hot well or condensate pump
takes the condensed steam from the condenser hot well and discharges it into
the No. 3 opefl.heater, after passing through Nos. 4, 5 and 6 heaters. The
hot well pump works against the difference in pressure;, the pressuré drop due
to friction, and the difference in elevatiom‘befiween the=No; 3 heater and con-
denser. The boiler feed-water pump takes the feed water from the No. 3 heater
and pumps it through Nos. 1 and 2 heaters into the boiler. This pump works
against the préssmre difference, friction drop, and elevation difference be-
tween the boiier and No. 3 heater.

6.8.2 Calculation Procedures

The method used to determine the cutput of the 600 Mw reactor power plant
is described below. Detailed caldulationé are given in Appendix Go_

The heat energy converted to work at the turbine wheel was found by cglcuo
lating the steam flow and enthalpy change between the throttle, extraction'
points and condenser (Ref. 23). The stea; flow is constant between the extrac-
tion points and the wérk done between extraction points is'given by,

work = steam flow ;g,x enthalpy change BTU o . (6015)
hr 1b '
The steam flow at the turbine throttle is found by a‘heat.baiance betweén.
the 600 Mw heat input from the reactor and the enthalpy differenée between the

steam and feed wafier at the boiler outlet and inlet.

throttle steam flow 1b s 600 Mw x 3413 x 103 BTU (6.11)

bhr AH BTU MWH

128.

In order to calculate the various extracted steam and water flows in
the feed water heating portion of the cycle, it was necessary to make the

following assumptions: ‘ .

1. The terminal temperature difference in the feed-water heaters was
taken as shown in Figure 6.2 (Ref. 23). This is the temperature
difference between the feed water leaving the heater and the con- =~ "
densed extraction steam leaving the heater and drained into the
next lower pressure heater. The open heater (No. 3) has no terminal
temperature difference since the feed water, extracted steam, and
drains are mixed.

2. The steam pressure in a heater was taken to be the saturation
pressure corresponding to the heater drain temperature.

3. No. 3 heater has no drain and the feed water leaving this heater
was assumed to be saturated.

4, A pressure drop of 8 percent was assumed between the turbine extrac-
tion points and the feed-water heaters.

5. No enthalpy correction was made for sub cooling.
6. It was assumed that there was no heat loss between the reactor and

boiler. This is closely approached in practice by proper insulation
of the heat exchangers and connecting piping.

T. The efficiency of pumps and motors was taken to be TO percent and
95 percent, respectively (Ref. 41).

The extracted steam flows required by all heaters, except heater No. 3,
were calculated from a heat balance between the heat gained by the feed water
and the heat given up by the extracted steam and drains from higher presauréi o
heaters. For No. 3 heater, both the extracted eteém flow and the entering
feed water flow were unknown. These were found from a heat balance and a
mass balance for the heater.
In order to find the enthalpy corresponding to the pressure at the tur- »
bine extraction points, a turbine expansion line with its end point was needed,
and this was found by the method described in Reference 41. Figure G.1l shows .

this pressure-enthalpy relation.

129,
) [ ]
The amount of work at the turbine wheel, as found above, had to be
corrected for heat put into the feed water by the condensate pump and voiler
. feed pump.. This heat caused a reduction in the required extraction steam .
- flow and & corresponding increase in the turbine steam flow and wbrk output, -
The heat gained from the pump per pound of feed water was found by,
Q BTU = work into pump BTU - work into feed water BTU - (6.12)

1b ' 1b 1b

The work put into the feed water is found from,
vork = (PpV, - P1Vy) BTU | (6.13)
1b
where PoV, and P;V, are the respective products of the outlet and inlet

pressure and specific volume. Since the change in specific volume is very

slight, Equation (4) may be rewritten as,

vork = (pp - p1) V; BTU ~ (6.14)
1b .

The work put into the pump is found from,

Ba

_ work in = work out BTU (6.15)
efficiency 1b :

The reduction in extracted steam is found from a heat balance between heat

from the pump and heat from the extracted steam,

feed water flow 1lb x heat from pump BTYU

W hr 1b (6.16)

H of extracted steam BTU/1b

130,

The turbine work gained due to the reduction in extracted steam is found
from Equation (6.10).

The corrected work done by the steam at the turbine wheel is; therefore,
the work as calculated initially plius the work due to the correction for heat
gained by the feed watef from the hot well and boiler feed pumps.

The work at the turbine Qheel, as found abcve, must be corrected for
exhaust losses in the steam leaving the last stage of the turbine and for
mechanical losses in the turbine and generator. These losses are found by
the method in Reference L,

The gross electgic output of the generator is equal to the work at the .
turbine wheel minus the turbine and generator losses, as described above. The
method used to determine thé net plant output is described in the next section.

6.8.3 Auxiliary Power Requirements, Net Plant Output and Efficiency

The method used to determine the groés electric power output of the
reactor power plaht was described in the previous section.

The net electric power output of the'plant is found by deducting from
the gross power all auxiliary power required by the plant.

Power required by the plant would 1nclude all pumps used in the cycle;
i.e., fuel circulating pumps, sodium pumps, feed water pumps, and the con-
denser ccoling water pump, and this power can be calculated. Additional
power is required for other pumps, air compressors, ventilating fans, light-
ing, etec. This power'is indeterminate for this study and an ‘arbitrary figure

of 2.45 Mw (1 percent of full load) was assumed.

-1€1-

TURBINE GENERATOR: }
1000F . 2I5F \ 250,000 KW TCTF,3600 RPM
Isalofi;?F :Is,ooopF ‘,%%%: jt8oopsig 15.6p 1800 psig, 1000F, 1.5 IN ¥g Abs~
oor 2 5/700,000 { | 5.850,000 ‘ 055000 1b/H- 1480.3H 1095 H / . \ M , GONDUGTOR COOLED, 26" BLADE
——fl N Ib /e Ib/mr AAA \ , ‘
3 f ]
FUEL| 2 3 3 245,420 KXW
SAUL : '
1050F|, * \
48,000,000 |b/Hr J \\ | 94ca
REACTOR fi00oF (O 675F N . ) 1024 H
Na NG b 3.2°P
2 \T 9£,800
INTERMEDIATE HEAT zlag -~ 15/ ne
EXCHANGER ol 8 [ 1.5 Hg N _mM
Nl‘"’a;
o padt I
~
. } { o Py A.1,600,000
REACTOR POWER-600,000 KW (HEAT) > DEAERATOR S Hwp ] Ib/Hr
GROSS ELECTRIC POWER - 245,420 KW 1_- I :- HEATER T ;‘r. z| 2
PLANT AUXILIARY POWER-16,400KW tags ol 59 # 3 MY aloo
NET ELECTRIC POWER- 229,020 KW a2l :gg- z|+ 2 olw
Tl o e 17 -
oFTo}" T~ OFTOY§ ~ SFTD  SFTD § & 5FTD
450F AV 90F ANAAR BFP AAMA 2I0F| 150F| ... [SL7F
W : 8Fe AT B T -
1950 p Ny 364.2h &2 fl[ 2000p ®a 8.1h a5 H7.9 h &6 §9.7h
430.1h 330F
l / 380F  300.Th 270F
450F 364.2h 238.8h 275F 2IEF IS5F
430.1h 243.9h 182.1h 122.9h
600,000 X 3413
GR HEAT RATE = - = 4 TU /KWH
0SS HE E 545,920 8,3¢58TU/
600,000 X 3413
: 2 : 8,935 BTU/XWH
NET HEAT RATE ® ;45,420 - 16,400
229,020
S22 - 38.29
NET PLANT THERMAL EFFICIENCY 500,000 Yo NOTE :
{1) BASED ON ZERO MEAT LOSS TO AMBIENT
. {2) TWO HIGH PRESSURE MEATERS AND DEAERATOR HEATERS
ZERO T.T.D. OTHER HEATERS 5F T.D,
{3} ENTHALPY NOT CORRECTED FOR SUBCOOLING,
, (4) STEAM EXTRACTION LINE PRESSURE DROP OF 8%
) ASSUMED FOR ALL HEATERS.
FIG.6.2-HEAT BALANCE DIAGRAM FOR 600 MW REACTOR-STEAM POWER PLANT.

V42251 : ONA-Y1~INHO
132.
TABLE 6.7

SUMMARY OF FULL LOAD OPERATING DATA

Turbine Throttle Pressure 1800 psig -
Turbine Throttle Temperature 1000°F )
Turbine Exhaust Pressure 1.5 in. Hg

Steam Flow 1.955 x 106 15/br

'Gross Generator Output 245 .42 Mw - -

Auxiliary Power Requirements:

Condensate Pumps : 0.33 Mw
Boiler Feed Pumps S.34 Mw
Condenser Cooling Water Pumps | 0.52 Mvw
Reactor Fuel Pumps - 0.36 Mw
Primary Sodium Pumps 5.98 Mw ’
Intermediate Sodium Pumps 1.42 Mw’ !
Miscellaneous 2.45 Mw
Total Auxiliary Power 16.40 Mw )
Auxiliary Power as Percent of Gross Output 69'{%
Net Station Output 229.02 Mw -
Net Station Thermal Efficiency '38,21:_

Net Station Heat Rate 8,935 BTU/KWH

133.

CHAPTER 7.

POWER PLANT LAYOUT, OPERATION AND MAINTENANCE

T.0,0 INTRODUCTION

In order to determine the suitability of using a fused fluoride salt
reactor in a central station power plant, it has been necessary to consider
all components of the reactor and power plant system. A pfeliminary piping

layout and drawings of a proposed arrangement of plant components have been

prepared to illustrate the details of this design.. The resulting power plant

‘design is feaaible in size and arrangement.

It wasralao necessary #o consider the operating characteristics; main-
tenance and'econfifiics-to.determine whether such.a plant would be attractive
to the power‘industry,; |

7.1.0  GENERAL PIPING ARRANGEMENT

The-génefél'piping arrangement shown in Figure 7.1 has been designed to
be safe, reliable, and to pro&ide flexibility in operation and maintenance.

T-1.1 Primary Loop Piping

The twenty-four primary heat exchanger bundles are connected to ring. \

héaders located aboie and outside the reactor vessel. Six sets of primary
circuit lines connect the ring headers to the intermediate heat exchangers.
The main purpose of this arrangement is to provide uniform cooling‘of the
reactor in case one or.more.intermediate heat exchangers are out of service.
The'pump for each primary circuit is located in the intermediate heat
exchanger compartment. The piping passes through the primary shield at an
éngle to minimize streaming of neutrons and gamma rays, and to retain the

integrity of the shield.

13k,

T-1.2 Intermediate Loop Piping

The six intermediate heat exchangers are distributed uniformly around
the priméry reactor shield. This symetrical arrangement minimizes shielding
requirements and the cost of the containment vessel. |

The six intermediate heat exchengers are divided into two groups, with
each group cennected to 8 steam Qeneratoro' Thé heat exchangérs-in each group
are connected in parallel to eaehvother inside the contalnment veésel, and a
common line is passed tarough the containmémt vessel to the steam generatoro
The connecting lines are designed to pr@#i&g equal firessure drops in order
to bélance flow and heat load among the intermediaté heat exchangers.

T.1.3 Steam Generator Connections

The steam generators are located as near the intermediate heat gxchangera‘
88 possible, outside the secondsry shield and containment vessel. The sodium
pumps for the intermediate circuits are located between the two steam generators

and are arranged so that one spare pump can serve either steam generator.

7-1.4 Steam and Water Piping-

In general; the piping in the steam turbine cycle is‘represefitative of
standard power plant practice. The differences are in the steam dump system,
vhereby steam way bypass the turbine and be condemsed in the condenser, in the
emergency cooling of the steam generators from head condensate tanks in case
of a complete power loss, and in the method of attemperating or controlling
the final steam temperature.

7.2.0 GENERAL ARRANCEMENT. OF PLANT COMPONENTS

The general plant layout is shown on Figures 7.2 and 7.3. The major pieces
of equipment are drawn to scale and gemerous allowamces have been made for pipe

expansion loops and clearances. The mean diemeters of the primary and secondary

. ) 135
circular shieldé are approximately 35 and 84 feetv, respectively. TheVSpace
between the primary and secondary_shields is divided into six compartments
by radial shields to contain radiosctive compbnefits of the primgfy and inter-
mediste circuits. The containment vessel, as shown,-is TO feet in diameter
and 110 feet in over-all height. The dimensions of the containment-vessel are
approximate only, since fio calculations have been made to estimate the volume
required to contain the gaseous fission‘products which would be reléaged in a
serious nuclear incident.

The reactor; primary heat exchangers and fuel dump tanks are located in-
side the primary shield. The:primary loop sodium circulating pumps, expansion .
tanks, drain and charge tafiks,'énd the intermediate heat exchangers are placed.
in the six shielded compartments. All radioactive components and materials are .
inside the sealed compartment vessel.

The steam generators and intermediate sodium pumps are outside tfie con-
tainment vessel and shielding, since they contain only non-radioasctive sodium
from the intermediate loop and water from the steam cycle. Isclation compart-
ments are provided around the steam generators and sodium pumps to prévent the ;:f*
spread of fires which could result from serious failures of the steam generators.

'The turbine-genérator, condenser, feed-water heaters and boiler feed pumps

are adjacent to the steam generators, as in a conventidnal.steam power plant.

T:3.0 SHIELDING

The dfiplex-compartmented shielding design for a reactor plant of this type
depends greatly upon the method of maintenance which is to be followed. The
basic premise’fér maintensnce in this design is that personnel will be permitted
access to the containment vessel only during periods when the reactor:is shut

down. . Based on this premise, the primary shield surrounding the reactor is

/36

[] TaNK

EXPANSION

INTERMEDIATE
AT

HE
EXCHANGER

——FLOWMETER
CHECK VALVE
SHUT-OFF VALVE

4 SAFETY VALVE

& POSITIVE ACTING CHECK VALVE

MOTOR OPERATED CONTROL VALVE

-Of Y o2X Z

COLD TRAP

GLOBE VALVE
DIAPHRAGM VALVE
NORMALLY CLOSED VALVE
EXPANSION' TANK

ORNL-LR-Dwg. -15263
CONDENSATE
STORAGE
TANKS
EMERGENCY
COOLING CONTROL
CONDENSATE
EMERGENCY COOLING
" STEAM RELIEF VALVE TURBINE o TUR BINE-
[ .. p STEAM S OM LVE GENERATOR
e . e
, ) / N —DESUPERHEATER UNIT
b ~ [l N “_o }—de
NCE —
THROUGH ]
STEAM GENERATOR
—)‘—N—N-d)gbo-i = A
I
STEAM - I
MP
l g‘:sm PRESSURE-TEMPERATURE l
» O ! REDUCING STATION
4 J1 1 LOW LEVEL
CONDENSER
CONTROL:
DESUPEREATER- JD Dé:::i-fl
T surenreaten PIEDWATER
i €7 . - HEATE 6] CONDENSATE
L) $ PUMPS
HIGH LEVEL
CONDENSER
X CONTROL:
L FEEOWATER I e |
REGULATING L A N ek ae
VALVES o Ak
Y 3
ATURAL 50 AJ_.'.LH F'... ' 1® I_i ® ©®
CIRCULATION K s > RO W N o O oy DL
BOILER PO o
+—
NOTE: FEEOWATER HEATERS OEMINER
FOR PURPOSES OF ILLUSTRATION -
A ONCE -TMROUGH 'STEAM GENERATOR FEEOWATER HEATERS ALIZER I SToRAGE
AND A NATURAL <IRCULATION BOILER TANKS
WITH SUPERHMEATIR ARE SHOWN.
CONDENSATE
MAKE-UP SUPPLY
FIG.T.1 - SODIUM AND STEAM POWER SYSTEM

"9ET

+

CONTAINMENT VESSEL

CRANE

3
1

TURBINE GENERATOR

INTERMEDIATE

HEAT EXCHANGER—]

Ne PUMB__|

\—FU!L DRAIN YANKS/

ZGADlOACYIVE SOCIUM DRAIN

AND

STORAGE TANK

FIG 7.2- ELEVATION OF POWER PLANT

° oo
STEAM GENERATOR ¥ . el
INTERMEDIATE No . O
Ne EXPANSION TANK COOLANT LOOP No EXPANSION TANK J. ¢ L
ING HEADERS [-—-]/- : | Sy
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FIG. 73— PLAN VIEW OF POWER PLANT

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7
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c9zGI- ‘Sma-y¥T1-TINHO

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139,
designed to protect personnel against decay gammas while perforuing direct-
maintenance within an intermediate compartment on the intermediate heat .
exchanger or primary loop pumps without'draining the fuel from the reactor
vessel. In addition, the primary shield serves as a neutron shield so that
the equipment in the compartments does not become radidactive°

Provisions are made for draining and flushing the sodium from individual,
intermediate, heat exchauge? ¢ircuits so thet no radiocactive sodium is within
the compartment where maintenance is to be performed. The radial‘compartment

partitions serve as shields and provide protection from decay gammas from the

.radiosctive sodium in adjacent compartments. This eliminates the necessity of

draining all the primary circuitsfwhen-oniy one cifcfiit requires maintenance.
The secondary shield,.located on the outside of the containment vessel,
protects plant personnel from excess radiation exposure while the reactor is
in cperation. The shield over the reactor will be made as a rotatiflg plug
with an offset removable section for access to the primary heat exchanger and
fuel circulating pump for maintenance. With this firrangement, the huge mass
of the top shield will not have to be handlgd before performing wmaintenance.
An aiternétive method would be to fabricate the top bhield;ifl sections and

.

remove only the sections necessary to provide access to the primary heat

exchanger or fuel circulating pumps.

The shielding shown in Figures 7.2 and 703.18 intended to be schematic
only.

T.4.0 MAINTENANCE

Maintenance of equipment outside the containment vessel will be conducted
in a conventional manner'since none of the equipment or materials to be handled
will be radioactive. Special provisions, however, are necessary to allow

-

. ‘-].-lfl'-’
maintenance of equipment inside the containment vessel and shielded areas.
The major items of equipment which will require special techniques are the
- fuel circulating pumps, primary heat exchanger tube bundles, intermediate
heat exchanger tube bundles, and primary circuit sodium circulating pumps.

7.4.1 Fuel Circulating Pumps

Access to the fuél circulating pumps and primary heat exchanger tube
bundles is through an opening in the top shield. This shield, which covers
the entire reactor vessel compartment, can be rotated until an offset removable
plug is centered over the desired location. The plug can be lifted with.the
Erane, and the pumps can be removed by remote handling eq_uipment° Spare pumps
will be kept on hand to replace faulty pumps. .

7-4.2  Primary Heat Exchanger Tube Bundles

In the event that a leak should develop in a primary heat exchanger tube
bundle, the sodium in the circuit will leak into the fused fluoride salt and
dilute the fuel concentration. The rate at which uraniufi is added to the
fused salt could serve as an indicator of tube leakage. 1t is felt that a
means of locating the leaking tube bundle can be developed. Since extra heat
transfer surface has been designed into the primary heat exchanger, ofie or
more leaking tube bundles can probably be blanked oéf and left in the reactor,
with little loss of power, until time is available to remove and replace the
faulty bundles. The bundles will be taken out of service by cutting the feed '
lines connecting the tube bundle to the ring headers and blanking them‘off. The
cutting and welding are to be done remotely, since the radioactivity will be too
intense for direct maintenance. The bundles will be removed through the top

gshield, in the same manner as described for removing the fuel circulating pumps.

lhl o

7.4.3 Intermediate Heat Exchangers and Primary Sodium Pumps

The intermediate heat exchangers and primary sodium circulating pumps
are located in shielded intermediate compartments. Priér to méintenance,
the radioactive sodium in the piping, pump and heat exchanger, in one compart-
ment, will be drained and flushed out with non-radioactive sodium by remote
operation. The equipmént can then be maintained directly. With the shielding
provided, the sodium in adjacent compartments wlll not have to be drained.
Sections of the top shield will be removed so that the crane can be used during
maintenance of, these componen‘fis°

T.5.0 PARTIAL LOAD OPERATION

The ultimate removal of the heat energy from the reactor system requires
the coupling of the fuel circulationvto the primary, intermediate and steam
loops. Operation of the system at some steady state partial load can be
achieved by variation of the full load design flow conditions in one or more
of the loops, use of control rods in the reactor, change of fuel concentration
with load, bypass control of the loop flows, fise of a supplementary heat
exchanger, or any combination of these procedures. It was apparent that this
multitude of fiossible combinations precluded a complete mathematical solution
of the problem in the time available for this study. Therefore, the decision

vas made to approach the problem of partial load operation in such a manner

as to give results in the time available. This was accomplished mathematically

by only studying the effect of varying the flow rate in the intermediate loop.
Another approach was used which involved the reactor simulator facilities of
the Oak Ridge National Laboratory. A description of the use of the reactor
simulator, the circults and the results are given in Section 7.5.2 and

Appendix J.

142,

T.5.1 Mathematical Approach

In a8 reactor system such as .the HRT or PWR, partial lcad operation is
achieved by aliofiing the steam pressure to increase as load decreases. This
method of control could only be used in a very limited manner for those reactor
systems in which high steam pressures are achieved at full load. Of the many
possible methods of control available, it was decided to'attempt a mathematical
solution of the system with a variable flow rate ih_the intermediate loop. The
reasons for this choice were the following:

1. Time limitations prevented anything but a single, variable study.

2. With the exeeption of the intermediate loop pumps and the.interu

mediate sodium coolent flow, all of the pumps and loop flows would
operate at constant values.

3. Temperatures would be held to reasonable values.

L. The problem of safely varying fuel concentration with load
is eliminated.

5. The solution thus obtained is also'useful if-bypass control of the
intermediate loop flow is utilized.

6. Control rods are not necessary for the achievement of steady state .
partial load operation.

7. The mathematical equations obtained would serve &as & basis for
further parameter studies of this system.

In order to obtain the mathematical solution to this problem, several
simplifications and assumptions were made. It is believed that individually,
or in the aggregate, these simplifioatioos.do.not invaiidate the analysis
over wide load~range5uf The mathematical analysis i1s given in Appendix I .
and is based upon the following considerations:

1. The flow rate of the fuel in the resctor and the sodium coolant in -
the primary loop remain constant w1th load

-2, Physical properties of the heat transfer fluids and the materials
of construction remain constant for the temperatures of operation..:

143,

3. Sodium heat transfer resistances are not the governing items in the
over-all heat transfer coefficients and do not vary with load. This
assumption is the most limiting restriction of the analysis given in
Appendix I and affects the load range, for which the derived equations
can be applied with reasonable accuracy.

4. Based on the‘foregoing,.the_0veffall héat_transfer coefficient in the
primary heat exchanger, the intermediate heat exchanger and the boiler
remains constant. The constant over-all coefficient in the boiler
further presumes constant average boiling and scale coefficients.

50- The final steam temperature and pressure arg constaht with load.

6. The average fuel temperature is constant with load}

T, The feed water temperature is h50°F at all loads.

8. At very low loads, some of the steém generated is bypassed directly
to the condenser. This does not enter the mathematical analysis
directly, but is an important element of control. ' '

Equations describing the behavior of the system at steady state, partial

load operation have been obtained for both the naturallcirculation boiler with
separate‘superheatér and the'once-through boiler. These equations are given
in Appendix I. It is noted that for the natural circulation boiler witfi sepa-
rate superheater,‘consideratibn (5) above, involving the constancy of the
final steam temperature with lcad, was abondoned at the end of the analysis
for this unit. It was found that a &ingle variable control cannot séfiisfy
all of the restrictions placed on the systefi in this section.

7.5.2 Reactor Simulator Analysis

In order to determine the partial load and'fransient.operétifié conditions
of this péfier plafit design; the circulatihg fuel find:sodium loofié were simfilafed
on the Analog Reactor Simulator at thé Cék Ridge National Labdratory.' The
analysis was limited to the design utilizihgfihe'qnéefthrofigh steafi generator.

The time constants and heat capacity of the reactor and sodium coolant

loops are shown on Figure 7.4. The simulator constants were computed by

1hk,

-Dr. E. R. Mann. These constants and the simulator diagram are included in
Appendix J.

‘The purpose of this simulator analysis was to determine whether the .
reactor would be stable at partial load and during operating transients, and
to determine the minimum load at which the final steam temperature could be -
maintained at design conditions. The analysis indicates that the reactor is
stable down to at least 17 percent of full load, the lowest partial load con-
dition investigated.

A final steam temperature control device was simulated by providing an
afixiliary heat exchanger in the intermediate sodium circuit, where heat is
transferred from the hot leg of this circuit to the cold leg. By varying
':the flow through this heat exchanger, the heat input to the steam generator
can be controlled. The results of the simulator analysis indicated that the
design temperature of 1000°F could be maintained down to 50 percent of rated
load. Below this load, the temperature increased gradually, but this excess
temperature can be controlled by the use of an attemperator, which is provided
in the steam cycle.

Figure 7.5 shows the steady state, partial load temperatures without the
use of the auxiliary heat exchanger mentioned above. The steam temperature -
increased from 1000°F at rated load to 1110°F at 17 percent of rated load.

The feed water temperature entering the steam generator is MSOOF at rated

load and decreased as the load is decreased. The resulting temperature differ-
ential between the sodium and water at the cold end of the steam generator is
considered to be too great from the standpoint of thermal stresses in the tubes.
For these reasons, it appears desirable to include the auxiliary heat exchanger _

in the circuit.

© 1k,
Figure 7.6 shows the steady state, partial load temperatfires with the
auxiliary heat exchanger in the circuit. The control signal used to adjust
the bypéss flow through the auxiliary heat exchanger circuit was set for plus
or minus 2°F of design steam temperature. The exact setting of these limits
would have to be studied carefully to eliminate hunting in the system. The
close tolerances used on the simulator are not required for satisfactory
steam turblne ovperatlon. The curves on Figure 7.6 indicate that thc steam
temperature remained constant to below half load and increased to only 10909F
at 17 percent of rated load. 'These temperature conditions are considerably
better than those obtained without the use of the auxiliary heat exchanger.
.The reactor power and tempe;ature transients, with the auxiliary heat
excfianger 1n_the infermediate loop, are shown ifi Figure 7.7. During this S
test, the load demand of the steam turbine . was reduc;d frém rated load to
half load in approximately 15 seconds. The initial effect of this load
demand 1is éeen in the rise of the sodium temperature entering the steam
generator. It is interesting to note how the reactor power and other tem- s
peratures varied and leveled out in a relatively short period. A series of
these tests vas:‘performed and all showed similar results, which indicate
that the reactor and system are stable during load changes. It should be
noted that the rate of these load changes is conaiderablylin excess of that
normally occurring in conventional steam power plant practice. These same
system transients withbut the auxiliary heat exchanger in the intermediate
loop are shown in Figure 7.8.

7.6.0 ECONOMICS

A complete economic study of the proposed power plant has not been attempted.

It must be realized that the large capital expenditure for a nuclear power plant'

74

1150F 1.3 Sec. 1150F 1100F L Sec. 1100F 1000F
. - ‘ ’ A
-[1200F VWA
| —— 1.6 Sec 1.1 Sec. 1l sec. 3 sec
1125F , -
: X Vo
1050F S
— v,
1000F 2.3 Ssec. 1000F  675F 5 Sec. 675F L5OF

* REACTOR INTERMEDIATE HEAT EXCHANGER BOILER-SUPERHEATER

_ ShelX Tube- Shell Tube
AT =150 F AT =150 F = hos F AT = 425 F

' = = - | = 887.5 F
T, avg - 1125 F Tavg = 1075 F Tavg 887.5 F Tavg 7.5
trgel (in core) 9.3 sec M © 21,990 BIU/ F tNa (exchanger) 1 sec
t sec
®ruel (in exchanger) °°° 5¢ Hopel1 = 170 BTU/F - Na (boiler) ~ S | .
' t _
% el (in transport ) 2.8 sec Hpipe = 9,190 BTU/F Na (transport)= 9 sec
Hovel (core) = 397290 BTU/F “xa (tuve) = L6 =€ Hy, : 1320 BIU/F
6,020 BTU/F ’ -
Ha (exchanger) / ty (transport) 3.6 sec Hy o = 190 BTU/F
" = 1,800 BTU/F 2 SN
S.S. (exchanger) - t e (she11) = - sec = 20000 BIU/F
= 20,800 BTU/F pipe

Hevel (exchanger) .

FIG. 7.4 TIME CONSTANTS AND HEAT CAPACIT® OF COMPONENTS

JN ONCE THROUGH STEAM-GENERATOR SYSTEM

= 550 F

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i51.
is one of the major factors which influences the cost of generating electricity.
These capital costs would have to be evaluated for various arrangements of the
basic design parameters in connection with the resulting reliability, ease of
maintenance, safety, efficiency, etc., and a careful analysis made of the re-
sults before such a study would have any real meaning.

7.6.1 Factors Requiring Additional Investigation

These studies would have to include the following considerations:

l. The selection of internal or external primary heat exchangers would
have to be evaluated carefully. It is considered that fuel inventory and
space requirements have been minimized through the use of the internal arrange-
ment, but these desirable features would have to be balanced against the over=-all
comparative costs of the two possible arrangements, their reliability, ease of
maintenance, etc.

2. Two sodium loops have been included in this study and are considered
a necessity, resulting from the internal arrangement of the primary heat ex-
changers. It is possible that the intermediate sodium loop could be eliminated
if external primary heat exchangers were used. However, again, such an analysis
would require considerable investigation to determine the lowest cost and most
reliable system.

3. If the reactor type using internal heat exchangers, as proposed in
this study, is considered to be desirable, further studies would have to be
made to determine the optimum arrangement of intermediate heat exchangers and
steam generators. It is quite possible that a unitized arrangement using an
equal number of intermediate heat exchangers and steam generators would result
in many desirable operational and economic features. This question could not

be resolved without a complete analysis of the entire system.

152.
4. As mentioned previously, the steam cycle arrangement also would
require much investigation. The question of steam conditions, reheat or
non-reheat, tandem compound of cross compound turbine generators, feed water )
heating cycle, turbine exhaust pressure, etc., would have to be settled from
a careful economic analysis. »

7.6.2 Approximate Economic Analysis

In order to examine the approximate economics of this plant, Reference 37
was used to estimate the capital costs. The capital cost for various types of
nuclear power plants listed in this reference varied from $183 per kw to $450
per kw. It appears that the fused salt reactor and steam power plant discussed
might cost in the range of $250 to $300 per kw. These two values will be used
in arriving at a powér cost.

For 240,000 kw generator output, 80 percent load factor, the total kw-hr
per year is 1.7 x 109.

There will be no operating cost for fuel processing. It is shown in
Section C.7 that the original fused salt charge can be used for at least 19
years before it is economically desirable to add a new charge of fused salt.
During this period, there is no fuel processing and uranium is added to over-
ride the effects of fission product poisons.

Using Reference 37, a figure of 1.2 mils per kw-hr for operation and main-
tenance is cited for 10 mil per kw-hr total power cost. Since this is the highest
value used in the reference, it will be used in this report as a conservative
figure.

The inventory charge is based on 150 percent of the hot, clean critical
mass of U-235. An additional 50 percent is allowed for poison override and .

fuel on hand. (See Section C.6.) Using 4 percent interest charge and an

153.
assumed cost of $20 per gram as the cost of U-235, the inventory charge is
0.2 mils per kw-hr.

Since some atoms are destroyed and no heat is derived by resonance absorp-
tions in the uranium, the fuel burn-up will be greater than 1 gram per megawatt
day. This figure is estimated to be 1.3 grams per megawatt day for this reactor.
Based on this value, the fuel burn-up at full load is 780 grams per day. For
80 percent load factor, the annual burn-up is 227,760 grams. At an assumed
cost of $20 per gram, the fuel cost per year is $4,555,200. This amounts to
2.6 mils per kw-hr.

These data are summarized in the following table:

TABLE T.1

ESTIMATED POWER GENERATION COSTS

$250/kw $300/kw

mils/kw-hr mils/kw-hr
Capital cost 5.4 6.4
Fuel processing 0.0 0.0
Operating and maintenance 1.2 1.2
Fuel burn-up 'l 25T
Fuel inventory 0.2 0.2

15k,
APPENDIX A

PHYSICAL AND THERMAL PROPERTIES

Figures A.1l through A.10

Tables A.1l through A.3 o

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ORNL~LR=Dug,=18185

156.

::::?I
: = / x
# i A
St R
f.
Sahasa
F :
H = ;’
1 l ;_
i ; i |
H
o
e .
asi
£ : .
Figure A,2:
Source: Thermodynamic Properties of Steam by 3
Keenan and Keyes; John Wiley and Sons
0 i o

157.

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600

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uctivity

-

Jele

300

rnal- Co:

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200

a

3rd. Edition

L]

VISCOSITY & THERMAL CONDUCTIVITY OF SATURATED WATER

Source: Heat Transmission,

100

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M
T

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1
T

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063 HF

d p *QT-£3TSODSTA -

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d, *3d tIg/mag
£yTATIONpUO) TRUILY]

Tenperature-oF

—_— :

ORNL=LR=Dwg,.~18187
158.
DENSITY OF SODIUM, POTASSIUM A SODIUM POTASSIUM ALLOYS
Source: Mine Safety Appliances Co. report dated Dec, 15,1953
curve @-202
58
56
($)
o,
udl Q’Q’%
N
N .{‘%Q\%&Q I
Yo Q"
K75 O0r'9,
PAes
AR
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< N %0
M . 48 G, 04
& ): 0,28 ¢§b
= | - ; .b »
E 50 F?JI/‘ HES oo "“‘%0 Qo‘? [
: N R =
1 1 % a‘ i ‘;E::
B éQ (] oD N i
K> S Ky (?Q’ 5‘5
g 2 O R, 25 K
[} % &> X, R4
g \e N 00 Q’é :
¢ 1,8 2 RS0 Y
(=] Q Q\? T 0 o3
< 4 %
(°4 Jo2 <,
W, 6 log
f = Q\,IA QJ &
Q‘[ O)
Q, § QJ E
16 o
do :
4
08, .
¥ Qf
Iy
! Figure A.l
I-I-Z i

0 1 2 3 L4 5 6 7 8 9o 10 11 12 13 1
Temperature-100°F

ORNL=LR=Dwg,.=18188

l .
VISCOSITY OF SODIUM, POTASSIUM AND SODIUM POTASSIUM ALLOYS
Source: Mine Safety Appliances Co, report dated Dec, 15, 1953
« L6 curve G-203
- 2nd Ed, Liquid Metals Handbook
) Journal of Metallurgy and Ceramics, Issue #3 May 19L9
')-Il-l— E L NRL report # C-28I
L2 =
. Lo !
- 38 ? ) : :
- 36 | : i
i | e e
.3 , N
. 32 521 128
}*\'O' [:;:% i
« 30 o e I
: i
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22 F O b
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H 5 &
1 B ] NG :
(B , s
e FH -
¢12 B set . odium
* iss AT 2 : ;
- HHH g i 156 NakK
of Figurs 4.5 =
578 NakK
FEEdE Potassium
o HHHH

IR ER R S s S T e e T L o]
Temperature-100°F

Thermal Conductivity-Btu/hr, sq. ft. °F/ft.

Temperature-100°F

ORNL=-LR=Dwg.=18189
160.
THERMAL CONDUCTIVITY OF SODIUM, POTASSIUM H
AND SODIUM POTASSIUM ALLOYS =
50‘ Source: Mine Safety Appliances Co. report EE;
dated Dec, 15, 1953; curve G-200 -
L8 '
1.6
Iy
L2
110
38 Sodium- 2nd Ed, Liquid Metals Handbook
HHH KAPL lemo LFE-6 ’ 19 51
MSA Tech, Report i
36 SEnaaE: H H
34F !
32 H
30 :
28 1 £
" +Potassium~ 2nd Ed, Liquid Metals Handbook
2 ] HE "
2
22
i -
20} s
Figure A,6
18 b o
16[0T 2 56 NaK
i | - SR 8 NaK
1y MSA Tech, Report # 2l
l2o0 2 2 3 L 858 6 7 8 9 10 1 12 13 1

SPECIFIC HEAT AT CONSTANT PRESSURE- BTU/[°F LB

'/Z/'

%

' SPECIFIC HEAT OF STEAM

AT CONSTANT i

IR ERENE NN

NN EEENE N

SN NN

: ' PRESSURE FROM STEAM TABLES.

|

L

ot

7
i

= T -

NN N
SN SN SNEEE N

i 1]

- Figure A,7

........................

Sources Thé;nddinanic Propefties of Stean by

0

100

200 300 400 500 600 700 800 900 (000 1100

TEMPERATURE- F

19T

0618T="* Anq=gT~TNHO

THERMAL CONDUCTIVITY - BTUju <er* e

/6 >—
. THERMAL CONDUCTIVITY OF STEAM BASED i
ON CALCULATIONS OF GRANET AND GOULD..
004
0.
0025 :
a0 | s
:
0.02C
00§
0010
0005 FlgureAB f 2 T
Source: Trans., ASME, Aug. 1950, Vol., 72
pp. 767-778

0

100

200 300 400 500 600 700 800 900 K000 |00

| )

TEMPERATURE-F

1200

*291

T6TQT=*3Ma=YT~TINHO

ol
olo -
Q0
o =7
m‘ g7,
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-
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) :
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O oo :
2 t
> |
003 8
uJ 1
- : H
30.02 :
Q
80.0’ HHH
< Figure A,Q [ VISCOSITY OF STEAM BASED UPON DATA
OF TIMROTH.
o FHHHFRHH R P FPPFFFPRFEEEEERET
- Source: Petroleum Refiner, May 1953, Vol. 32, HHi
No. 5, pp. 179-181 g
O 100 200 300 400 500 600 700 600 900 000 10O 1200

TEMPERATURE-°F

€91

26TQT=" SMq=NT~"TNHO

164,

ORNL~LR=Dwg.,=18193

i

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= m o+ e
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3
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n ~ ~ HH ~ ~t o on o on o0
saaun o UOUT/d;=»%d *IH/M3d ~£3TAT00pUO) TemISYY
1 ui

Source:

TABLE A.1 .

Specific Heat
Bx/1b - OF

COOOO0 C.O0 00 oRoNoNoN o] OCOO0OOCOC OO COCOC ooNeoNoNe

OCO0OOOO0

- 3117
.3115
.3113
-3110
3108

.3106
310k
3102
3100
3098

.3096
.3094
.3092
3090
.3088

. 3086
. 3084
.3082
.3080
.3078

- 3077
- 3075
.3073
.3071
. 3069

. 3068
. 3066
. 3064
.3063
.3061

. 3060
.3058
.3056
3055
3053

THERMODYNAMIC PROPERTIES OF LIQUID SODIUM

Lee, John F., "Thermodynamic Properties of Liquid Sodium",
NUCLEONICS, Vol. 12, No. 4, page T4, April 1945

Enthalpy
Btu/lb

1k1.
143,
1l ,
146,
148,

149,
151.
152,
15k,
155.

WMo\ OWV\VO F®

}—l
=
N
ONOWVOF WVEFODWE® WU —JHOGHW

165.
Table A.1 (continued)

Temp .

Specific Heat
Bx/1b - OF

0.3052
0.3050
0.3049
0.3048
0.3046

oNoNoNeoNe OO OCOO COOO0O0 COO0OO0O0O
W (7Y (WY) (oY
O o o o
| = o n
o — n \O

oNoRoNoNe
()
O
O
3

166,

Enthalpy

Btu[lb

195.7
197.3
198.8
200.3
201.8

167.
Table A.l (continued)

Teup . Specific Heat Density Enthalpy

Op Bx/1b - OF 1b/ft3 " Btu/1b
875 0.3004 52.21 256.2
880 0.3003 52.18 257.8
885 0.3002 52,13 259.2
890 0.3002 52.09 260.8
895 0.3001 52 .04 262.3
900 0.3000 52,00 263.8
905 0.2999 51 .96 265.3
910 0.2998 51.92 266.8
915 0.2998 51.87 268.3
920 0.2997 51.83 269.8
925 0.2996 51.78 271.3
930 0.2995 51.7T4 272.8
935 0.2995 51.70 27h.3
9ko 0.2994 51.66 275.7
9hs 0.2993 51.61 277.2
950 0.2993 51.57 278.8
955 0.2992 51.54 280.2
960 0.2992 51.48 281.7
965 0.2991 51 .44 283.2
970 0.2991 51.40 o84 .7
975 0.2990 51.36 286.2
980 0.2990 51.32 287.7
985 0.2989 51.27 289.2
990 0.2987 51.23 290.7
995 0.2988 51.19 292,2

1000 0.2988 51.15 293.7
1005 0.2987 51.10 295,2
1010 0.2987 51.06 296.7
1015 0.2986 51.02 298.2
1020 0.2986 50.98 299,7
1025 0.2986 50.93 301.2
1030 0.2985 50.89 302.7
1035. 0.2985 50.85 304.1
1040 0.2985 50.81 305.6
1045 0.2985 50.77 307.1
1050 0.298L 50.72 308.6
1055 0.2984 50.68 310.1
1060 0.2983 50 .64 311.6
1065 0.2983 50.60 313.1

-
3
O
o
g
w
\n
o
\n
\Nn
(VN
=
=
O

Table A.1l (continued)

Temp » Specific Heat Density Enthalpy
Op Bx/1b - F 1b/ft3 - Btu/ib
1075 0.2983 50,51 316.1
1080 0.2983 50.47 317.6
1085 0.2983 50.43 319.1
1090 0.2983 50.38 320.6
1095 0.2983 50,34 322,0
1100 0.2983 50.30 323.6
1105 0.2983 ' 50,26 325.0
1110 0.2982 50,22 ' 326.5
1115 0.2982 ) 50,17 328.0
1120 0.2982 50.12 329.5
1125 0.2982 50,08 . 331.0
1130 0.2982 50 .04 _ 332.5
1135 0.2982 49.99 334.0
1140 0.2982 | 49.95 335.5
1145 0.2982 49.91 337.0
1150 0.2983 49,87 338.5
1155 0.2983 49,83 340.0
1160 0.2983 49.78 341.4
1165 0.2983 : 49, Th 3L2,9
1170 0.2983 49,70 34k, 4
1175 0.2983 49,66 : 345.9
1180 0.2983 49,62 474
1185 0.298L ' 49,58 348.9
1190 0.298kL 49,54 350, 4
1195 0.2y84 49.50 351.9
1200 0.2984 | Lg.45 o 353. L
1205 0.2985 , 49.41 354.9
1210 0.2985 49.37 356, U4
1215 0.2985 L9.32 357.9
1220 0.2985 . 49,28 359.L4
1225 0.2986 : L9 .24 360.8
1230 0.2986 49.20 362.3
1235 0.2986 49,16 363.8
i240 : 0.2987 49.11 - : 365.3
1245 0.2987 49,07 366.8
1250 0.2987 49.03 368.3
1255 ‘ 0.2988 - 48.98 369.8
1260 0.2988 48,94 371.3
1265 - 0.2988 48.90 ~ 372.8
1270 0.2990 48.86 37h.3
1275 0.2990 48,81 375.8
1280 0.2991 W8.77 377.3
1285 0.2991 48.73 378.8
1290 0.2991 48.68 380.3
1295 0.2992 L8 .64 381.8

Type

Analysis

Carbon
Manganese
Phosphorus
Sulphur

- : Silicon
Chromium
Molybdenum

- Other Elements

Minimum Physical Properties
Tensile Strength, psi
Yield Point, psi

Elongation, % in 2 in.

Max Brinell Hardness

Creep Strength

pounds per
sq in.

-~  Rupture Strength

169,

load in pounds
- psi which lead

to rupture in
- 100,000 hrs

c——
TABLE A.2 '
SELECTED PROPERTIES OF STAINLESS STEELS
Source: Pressure Tubes and Piping - Timken Roller Bearing Company
Type 304 Type 347 Type 316
0.08 max 0.08 max 0.08 max
. 2.00 max 2.00 max 2.00 max
0.030 max 0.030 max 0.030 max.
0.030 max 0.030 max 0.030 max
0.75 max 0.75 max 0.75 max
18.0-20.0 17.0-20.0 16.0-18.0
- Cb lo X C min, l:o max MO 290“’390
Ni 8.0-11.0 Ni 9.0-13.0 Ni 11.0-14.0
75,000 75,000 75,000
30,000 30,000 30,000
35 35 35
200 200 200
Rate of 1%, hrs Rate of 1%, hrs Rate of 1%, hrs
100,000 10,000 100,000 10,000 100,000 10,000
800°F a-- - - -—- ——— -—-
QO0CF --- --- - -—- --- ——
1000°F 10,700 18,000 --- - 1k ,750 2k ;500
1100°F 7,900 13,000 14,250 21,000 12,000 21,000
1200°F 4,300 8,000 1,750 5,300 6,800 14,500
1300°F --- --- 660 2,000 4,300 9,200
1500°F 1,450 2,850 -—- -—- 1,800 4,200
" Load Load Load
900°F -- -- -
1000°F -- -- -
1100°F -- 16,000 --.
1200°F 7,200 6,200 12,500
© 1300°F 3,700 2,800 8,900
1500°F 1,700 -- 1,600

Source:

170.
TABLE A.3

SELECTED PHYSICAL PROPERTIES OF "L" NICKEL

Private Communication, Jr. J. J. Moran, Jr., The International
Nickel Couwpany :

Creep Properties

Condition: Cold Drawn-Anncaled

Stress, 1000 psi, to Produce a Secondary Creep Rate of:

0.00001%/hr 0.0001%/hr
1%.0 25.0
7.5 13.0
L.3 7.3
3.7 6.3
2.5 h.5
2.1 3.7
1.2 2.1
0.5 0.9

Ruptiure Properties

Condition: Cold-Drawn Annealed

Stress, 1000 psi, to Produce Rupture in:

1,0C0 hrs 10,000 hrs 100,000 hrs
27. - - -
20.5 -- -
17.5 14.0 11.5 -
1.5 11.3 8.7
11.3 8.7 6.7
7.8 5.4 3.7 -

b7 g

APPENDIX B

DESIGN CURVES

Figures B.l through B.k

17 32—

HEAT TRANSFER TO STEAM OR SUBSATURATED WATER IN TURBULENT FLOW
(INSIDE A TUBE) :
600
. T B e
Gk et i | F H
x®" 2
e 4 H
b { .2 - o E
& na®° = ¥ ( w000 a )° $ :
> £
100 o®” . W - fluid flow rate - 1bs,/hr,t
el e d - diameter - inches i
b‘:;]j a - flow area - square inches |
o H Derived from Equation 6,2
300 I]ll[I.IIIIIlllll(llll"ll|lil!llLlIllllll!‘lllllll!ll[llIlllllllllll
:'GOO Physical Properties from Appendix A
S iyin e J. .
%*’ [ o Go %bo t 1
% i e Do My Yo
9,(’ 1 & 'ZO u‘i: Gj TN 4
200 K)oc T %2 i "{I Ll‘oo p Q‘-{
- }%9'4 T »“’l.’ 7_iymi. % .?ie “ --1 'I,. St
LT 1500 pgi1g T Bepassaansett : 58
- aasail 1_!' j:"Zo H_H_IL JE EEE ] ii”i H““.“ }
100 EEEEEEE WS Average Bulk Fluld Temperature Fl‘igureB.'T!‘.
200 300 }:00 500 600 700 800 900 1000 1100 1200

H6TQT="3na-UT=TNHO

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17k.

ORNL'LRPM ° .18196

*
T TIT T wt._jl:, ASE] XA TE RS IREST TURR PR RN (R 4R : 23
SR ERSEL RRRRERaEas AEARE ARSH HaCSi T ] ft e544 L1141 301 1501 DSEEE FIST] ISOR) LIRS RSP EE: {8 M@ e b
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T H T 8 254 E2i0d Ehawd LONRE ENRRY KRR ERRLY ERERS RETH ERE N ! e . ks
T a3s s o jseed Said Shacd HERESRRRT MES H i L_m g = K 1
1 1] 1L ] * 1 ! 5 perrig ot d HOR B ) R i brd 24 o .
¥ s : R e Hit ] e e H v -
: SISO FAAFRANSEe Ses i = . e S
ans .t 4 L H o . ! = b 'Mf .m. Jf Pty .w i ik ! 0 @ ! . 1.l
= : . 0o BREd o “ 28 : FREL Gl Tt
U T kv i 45015451 isTHRERER) tskcARinE B e N0 Nt _ £ : K
sspujgpany oe - abaa s sl E=Rot 18! 14 r _ . o
— 0 + u SulgE *.. w\w 1 A* wA m + = ; 11 Ht ; A o
sasalS3and I2s %u e SO | i H 8 eRESEA LY Pl ¢ 1 Yo =
ssags RdEd bul ssmss o e w) - suses gauns g 111t 8 S AORA (i 1 1 ! = m W )
P D TR o ke cea: HHHEH x. L = " & ‘
”&H;H .MML ..x,rufiru.v leHJH L w ...r 1 AYHL J J 4 : b 1 :AA nm U w “rm m e "flw m : u
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Bow CJ _ +HH - TR AR . Lwfl f fi: 1 I By b o B |t Aol .va
SSSENERIST ERSRE RRR - all g8 se s HH BTG P - AEHEEN e kBT L . ._-
et st M g & [t L nTesetrr is o P “ PR BRI A (o
o ,. H Ll lo; ™ 1t -+ lfim ' m w fi hvwv. { “ t | il 0 O TS . ) . ....l.w.L.Lm
" L & gagse H flw. HAHHH 1 3RRR25800 i b , | 13 3 ERed LEns 4 L ..A
2 THHEITH L Had] jpdae 1 ! -t P AR o6 S dip R
.m @ [if B TH LM H1] seadsasd i ; = ” Seadrioiseesy pery ST R
G I | i o Bl RSt LR Ei Sl e e O m
1 = 4 - -
n - - 1T ITHE H T =8 EEREYFERS! 12
, H T H . un ngssdud SREAR FER (I8
o .Mu HH 1 m T HHHTTH . E_ 158881 ISEL1 HERS) FRNEIR-
. | e lJI 3! R e eE BEE: T vl ] “Hnfl T .Ww w* .”M ; : m
sastlass i TR e T 23005500 MRARROREE CEzed dnuzt punni Atl IR
- T RS- WSS SRS S o -t .
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sge & o + - Il T LY pEs 25 + H A TS0 AN UE RO > aal pid 1 3
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seeassieatis 1 8 o H &8s gigdpas: R PR T LU - Sy
TR ans 25| : 24 B r C Hit 38 b sSSasns ssbau iss 1 5 a8 8 \r ] 111
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= H4 J.1. » B @ L4 /B
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wags * e s ,:.Axv 1f1.fi. vI”A Av e 1 A I + 11! $ “ + 1
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THH [T - 3 1 [ o8 CRRSE R238 - $223
SR RS LR 828! P H 33 °ay/nyg=-uai LVIH ;
Wl* LT | | * i 4+ + N

/75—

TUBE COUNT FOR TUBES SPACED ON EQUILATERAL CENTERS
! (%ube on center)
Source: Unpublished derivation by I. Granet, 8/55

L619T="3rq=uT~TNHO

*GLT

g ,
H L
1
8
2
5
T |
:
=
2
"R-Radius to center of outermost tube
P-Tube pitch, center %o center of tubes
1 0 O O =%
Figure B.l
e H-
3 P R | -

APPENDIX C

REACTOR ANALYSIS CALCULATIONS

C.1l INTRODUCTION

The solution of the three group, three region (3G3R) problem has been
coded for the Oak Ridge National Laboratory's electronic digital computer,
ORACLE (Ref. 28).

The fast and intermediate energy group microscopic cross section data
used to calculate the input data for the code were obtained from the "Eyewash"
code (Ref. 27) and from BNL-325 (Ref. 30) by assuming a 1/E flux distribution
in these slowing-down groups. The thermal cross section data were obtained
from BNL-325 by correcting the tabulated values to the energy corresponding to
the average design temperature of the reactor, 607°C. Table C.l is a summary
tabulation of the microscopic cross section data used for each energy group.

The thermal energy cross sections from BNL-325 were corrected to the
average neutron temperature and for a Maxwell-Boltzmann flux distribution

as follows:

o = 0.885 5 (To)% = 0.48 o,
T

where, o and T, are the BNL-325 values for microscopic cross section
and the standard temperature of 293 A (20°C); and T is the average design
temperature, 880°A (607°C).

The intermediate energy group was selected to range from .Ql15 ev to 100 ev.
This upper limit of the thermal group and lower limit of the intermediate group
wvere selected to approximate the energy at which the discontinuity in the neutron

flux distribution occurs (Figure C.l). At this energy, the neutron flux changes

177.
from a near Maxwell-Boltzmann distfibution to an approximately 1/E distribu-
tion. This discontinuity occurs at an energy which is approximately six times

- the average thermal neutron energy. ©Small changes in this group enérgy limit'

(+ 1 ev) have only a negligible effect on criticality calculation results.

Maxwell -Boltzmann
e Distribution

Neutron
Flux

1 Distribution
E

E

|
|
[
l
|
I
|
T Neutron Energy —>

Figure C.1  NEUTRON FLUX DISTRIBUTION

The upper energy limit of 100 ev for the intermediate energy group waé
selected because the uranium-235 resonances cease at about this energy.

The upper energy limit of the fast energy group was selected to be 2 mev.
The 3G3R code calculation assumes a monoenergetic fission neutron source; and
2 mev is approximately the average energy of the urafiium-235 fission neutron
energy spectrum. |

Following the notation convehtion of Reference 28, the designation of
the fast energy group becomes Group 1, the intermediate energy group becomes
Group 2, and the thermal group becomes Group 3. The geometric region designa-
tions are: the central core is Region 1; the l-inch thick.baffle between the
central core and the heat exchanger is Region 2; and the heat exchanger, other

baffles and the downcomer homogenized to be Region 3. The numerical subscripts

|

[

'

3

I

i
|5
174

TABLE C.1l

BASIC MICRCSCOPIC CROSS SECTION DATA
USED TO CALCULATE 3-GROUP, 3-REGION CODE INPUT

Element — _Fast Group _ — o Intermediate Group _ ——  Thermal Group
(bighly enric§8d) g6_'01 . O—trl 0?1 ral f Tty 30trp 5—;2 6;2 1 3 J—1:1:‘3 0—5‘3 fi3
Fluorine 0.36k4 11.30 .- 0028 0.352 9.85 o 00T 11.7 --  .0051
Sodium 1.03 36.3 -- 0 0.286  9.70 -- .0102 12.0 - 0.258
Zirconium 0.163 22.5 -- 0005 0.135 18.6 o= ,0113 18.6 - 0.092
Uranivm 0.175 36.4 5.31 7.12 0.5 35 34,3 50 30 301 356
Nickel 0.459 36.3 -- 050k 0.592 49.6 - 275 L9 -- 2.35

§ Etl 321-.1-1 Zfl & ai § 21:2 3 gtra z 2 b3 a2 3 2“‘3 s 3 233

*Stainless 0.024 1.67T -- 0.005 0.036 2.937 - 0.017 2.89 - 0.1k0
Steel (304)

* Note, macroscopic, not microscopic, values for S.5.

QLT

-
-
-

iEE&
on the nuclear property notation refer to the energy group and geometric
region, respectively. For -example, 25312 "is the macroscopic absorption
cross section for eénergy group 1 apd geometric region 2.

C.2 SAMPLE CALCULATIONS

Calculation details of macroscopic group constant® for group 1, region 1,
for a 0.01 gm/cc uranium concentration are given as sample caleulations below.
All the cases for which calculations were made, and the geometric and uranium
concentration parameters of each, are given in Table C.2. A summary of all the

macroscopic code input data calculated for each case is given in Table C.3.

The fused salt fuel composition:
43 mol % ZrF),
57 mol % NaF

~0.1 mol % U-235

The variation of the density of this fused salt composition with tempera-

ture is given by the equation,
€ - 3.65 - 0.00088 x (°c)
Then at the average design temperature of 607°C (1125°F),
¢ = 3.12 gg/cc

The weight percent of each elemental constituent of the fused salt was
calculated to be:
Fluorine 45,49

Sodium 13.T%

Zirconium 40.9%

180.
TABLE C.2
CASE DATA AND COMPUTED MULTIPLICATION CONSTANTS (k)
Uranium-235
Core Radius R, Concentration
Case No. ft-in. cm - | gm/cm3 Remarks Computed, k
"RO CO by - 6" 137.16 0.020 Basic Case 1.34
1 " " 0.010 (Baffle Plate 1.13
: as Region 2)
2 " " 0.005 0.86
3 " " 0.030 1.42
R1 CO 4t - O" 121.92 0.020 1.31
1 " " 0.010 1.09
2 " " 0.005 0.83
3 " " 0.030 , 1.39
R2 CO 3 - 6" 106.68 0.020 1.26
1 " " _ 0.010 1.05
2 " " 0.005 0.79
3 " " 0.030 1.3%
R3 CO 5t - o" 152.40 0.020 ' 1.36
1 " "  0.010 . 1.16
2 o " 0.005 ’ 0.88
3 " " 0.030 B 1.44
RL CO 3 -0"  9L.uk 0.020 | 1.21
1 " " 0.010 0.99
2 " " 0.005 0.7T4
ROSCO Lt - 6" 137.16 0.020 Baffle Plate 1.35
as "shell"
ROOCO br - 6"  137.16 - 0.020 No Baffle Plate 1.3k

5,68x102cm”

181.
TABLE C.3
MACROSCOPIC GROUP CONSTANTS FOR 3-GROUP, 3-REGION.CODE
Conc. of U Region 1 Region 2 Region 3
- in gm/cc (Core) (1" SS Baffle) (Heat Exgha.ngerL
Group 1 Dy 0.906 cm Dyp 0.50 cm Dy3 0.783
Fast £ixy] 3-04x10-3ecm-1 £y, 2.79x10-3cm1 ax]_g 2°72"1°'3°“_’;-/l
S are 2.21x10"%em-l  Sai, 4.9x10-3cm™1 S a1 1.29x10 3cm-1
=005 { - gffi 1.68x10-%cm 1 1 IS fig 5.86x1072cm =L
a1 3.12x10 " eml  Sags 4.9x10-3cm1 b2 1.32x20"3em ™t
-010 Z VEfyy 3.35m10 ket 2Tl VE £13 L11x0  en"
T ~3em-1 8x10-3cm~1
s }.95x10" *cm k.9x107Fcm = 1.38x10 Jem_
-020 v E ?i’i 6.,"{0x10'1’cm'1 ‘ alz ——- S ?ig 2..35x10"‘*cm
s 6.77x10 %em"L £ 4.9x10"3cm™1 < 1.45x10 3cm*1
+030 { "’Zgii 1.00x10"*em1 12 --- -yi?ig 3~52x10"hcm”l
Group 2 Dsy 1.42 cm Dnn 0.308 cm . Dpz 0.818 cm
Intermediate fixgi 3.80x103en™t £ xgg 7.48x10-3cm ™1 é‘_.xag 3083x10_’3cm’l
-3 -1 -3..-1 -3 -1
Sayy 1.06x10 3cm Eapn, 1.86x1073cm Zapz 5-5x1073cm
-005 { 1)2;21 1.08x10"3cn ™+ a22 -—- VE 23 3.78x10-%em"1
1.70x1073cm™l  £aop 1.86x1073eml  £gnn 5.74x10"3cmL
-010 .ué ?Si 2.16x10-3cm™ -—— vE, f:3 7o55x10"l‘cm“‘1
2.98x10 3em~1 < 1.86x10-3cm~1 £,apz 6.18x10"3cm"1
.020 {‘U g::i 49322{10‘3(:"['1 322 - _VZ f23 1. 5]-.:(10‘“3cm'°:!
- 4.26x10"3en"! S, 1.86x103em™r < .. 6.63x10"3em™*
-030 é -.fg‘ ror 6.48x1073em=l T BR T 2 Z«;gg 2.27x10 3cn"
Group 3 1331 1.22 cm D3p 0.312 cm D33 0.788 cm
Thermal 3 2‘ 1 = 5 1
< 8.41x10"Jem-1 0.155 cm” 4. 76x10"cm”
o { V< :gi 9.48x10"3cm~1 "32 --- ’Uzggg 3032fl0-3cm-1
. = 1,30xio--2cm'l Zna, 0.155 cm L < ans 4.92x102cm-1
-010 {—ys :gi 1.90x10~2cm -t 832 - Vs ;gg 6.65x10 3em "1
< 2.21x102em~t &£ ___ 0.155 em~1 £ aas 5.24x10-2cm~1
a020 {'}’E;gi 3079x10-2cm-l ,. 8.32 - Wé-fgg 1033Xl0-2(3m-1
1030 3.13x20"cm~1 2832 0.155 cm~t Zoa33 5,56x10-2cm ™1

.= fgS 1.99x10-2cm 1

182.
The negligible quantity of sodium and fluorine in the (NaF)QUFh fuel con-
centrate was not included in the calculations.

The atomig concentrations in Region 1 are:

Npq =z £ (W% of F in salt) Ng = hb9 x 1022 atoms/cc
atomic wt of F

1.12 x 10°¢ atoms/cc

]

NNa 1

8.42 x 10°! atoms/cc

NZr 1

Ny 1 = (uranium conc.) Ny = 2.56 x 1019 stoms/cc @ 0.01 gm/cc
235

The total macroscopic absorption cross section,

' F Na zr U
Zau =NF10;1 + Nya 1 Zal ‘”'NZrla;:L +N 1%

3.12 x lO"'h cm 1

]

The total macroscopic fission cross section, multiplied by -/ , the

average number of neutrons released per fission (2.46 for U-235),

=k

fission neutrons/cc
@ 0.01 gm/cc U conc.

vEg, =¥y (VI7 ") = 3.35 x 10
The diffusion coefficient for Group 1, Region 1j

l = l ‘ o
Na "y U
3Zir 11 Np 33%aF + Mya 3% - + Ngp 13 % 2Ty 130

0.906 cm

The lethargy change in the fast group,

ev) = 9.9 units

Aul;ln 2x:l.06
100 ev

183.

The lethargy change in the intermediate group,

Auy = 1n (100 ev) - 5.3 units,
0.5 ev ' |

The effective removal cross section (Efx) was approximated as follows:

lethargy interval of group

AL

'E = average change in lethargy per collision

f_-TZs ¢ average no. of scattering collisions/cc-sec
Au/f avg no. of scattering collisions for neutrons to pass thru u

Approximating 238 by E;t’ the effective removal cross section; or effective
cross section, for a neutron to pass entirely through a lethargy interval, fi,

becomes,

The effective removal cross section for Group 1, Region 1,

= Ng 1(§ o—t) F 1t NNa 1(§ o_t)Na 1+ Nzr 3 fd’t) Zr 1"'_1_“‘1_;- '1(3 U'QU 1
Aul

| ixl

1

- 3.04 x 1073 emn~?

C.3 POWER DENSITY

For a central core radius of 3 1/2 feet and a core height of 10 feet,
Total volume of the central core - 1.09 x 107 cc

Assuming all 600 Mw of heat are generated in the central core,

Average power density in the core - 55.2 w/cc

18k,

C.4  NEUTRON FLUX

The average neutron fluxes in the central core for the hot, clean

+ eritical condition were computed as follows,

)

average rate of fission = 55.2 w/cc x 3.1 x 1010 f’issions/wasec

1.71 x 1022 fissions/cc-sec

average rate of fission also equalszzf 11¢11 +2f 21 ¢21 + 2-} 31 ¢31'

From Figure 4.3,

B, =088,

Py =038,

The values of 2 ¢ for the hot, clean critical uranium concentration

for a 3 1/2-foot radius central core ;, 0.0088 gm/cc, were computed to be,
] b oL
g gy = 120 x 10 cm

Sy =763 x 107 em”l

Seqs 6.80 x 1073 cm~1

Making the appropriate substitutions and solving these equations; one -
obtains for the hot, clean critical conditicn of a 3 1/2-foct radius central
core,

average fast flux in the central core, 100 ev to 2 mev:

¢ll = 1.2k x 1015 neutroms/cc-sec

'avera.’gel intermediate flux in the central core, 0.5 ev to 100 ev:

¢ 2] = 5.96 x 1014 neutrons/cc-sec

185,
average thermal flux in the central core, up to 0.5 ev:

§j31 = 1.62 x 103* neutrons/cc-sec

the percent of the total fissions occurring in each energy group:

=+ A x 100 . Se Bx 100

total fissions/cc-sec 1.7l x 101

Fast group -—- %
Intermediate group -~=-- 27%
Thermal group --- 64%

C.5 FUEL BURN-UP

Approximately 1.3 grams of-uranium-235lare consumed to produce 1 Mw day
of energy. At 600 Mvw power, the uranium burn-up will be appfoximately 780 g/day.
This is equiva}ent to burning up a éuantity of urapium approximately equal to
the total uranium inventory in the fuel system every year.

C.6 EVALUATION OF FISSION PRODUCT POISONING

An extended analysis of ihe fission prpduct nuclear poison buildwpp for
conditions similar to thosé\ef\phis design estudy has fieen perfofmea (Ref. 33
and 34). This analysis is for a reactor using ufanium-233 as fuel. The fission
product yield for U-235 is so nearly equal to that of U-233 that this factor
was neglected. Thi§'analysis assumes couplete removal of all.ggeeous fission
products, which should very nearly be the condition for the reactor of this
‘design study. The results of this reference are presented as the "poison
fraction" build-up with time of operator for the reactorvconsidered°

Poison fraction, P = ééz = £(t) .
EF
The fission product production, and therefore EE:P, is directly proportional

to the power density of the reactor fuel system, which in turn is proportional to

186.

the product of neutron flux and fission cross section, fo}?: If it is fiasumed
that the primary means of fission product removal is by neutron abscrption and
only a small portion is removed by radioactive decay, the quantity of fission
product poison, Zp, present is approximately inversely proportional to neutron
flux,Sd° | |

The poison fractiom build-up with time is then seen to be very nearly
independent of operating conditions of different néutron flux, fuel concen-

tration and power density,

P:2p =f(t) (£f¢)(—}3) = £(t)
>t E

The plot of poisoh fraction increase with time given in Reference 34,
page 14, should, therefore, be applicable to this design study's reactor as
é first approximation. |

The data presented in Reference 34 extends only to 150 days of operation.
A visual extrapolation of the curve for total fission product poisoning would
indicate that the curve asymtotically approaches a poison fraction of about
15 percent for very long periods of operation (>5 years). This is probably
a conservative estimate. Any plating of the metallic fission products in the
low flux, heat exchanger region (see Section 2.6.0) would help to reduce this
value.

Assuming little or no corrosion, and therefore corrosion product nuclear
poisoning to be small, the value of 15 percent poison fraction may be used to
approximate the negative reactivity introduced in this design study'’'s reactor

after very long periods of operation (:>5 years) .

187.

Converting poison fraction to units of reactivity (Ref. 8, p. 262),

z = Z_-u = 2 (for this design study)
Zn

reactivity, €= - P z ’-E' - 091'5 x 2 = -0.10 |
1+ z l1+2
From Figure 4.1l of this report, it may be seen that it will require a
uranium concentration of approximately 0.0115 g/cc to override'this negative
reactivity and maintain criticality after a very long period of operation.
This is equivalent to about 131 percent of the hot, clean critical
inventory.

C.7 REPROCESSING CYCLE TIME

The excess uranium fuel inventory that is added to override fission
product and corrosion product nuclear poisons should be the economic bptimum
that considers the costs of fuel reprocessing. Thé following econofiic analysis
is based on the fuel reprocessing method available today (Section 2.,6,0) which

requires contaminated salt solvent to be discarded and replaced.

Cost of replacement salt = $5 per 1b

. Cost of recovering uranium =  $1 per 1b
from contaminated salt by
UF), volatility process (Ref. 38)

Total = $6 per 1b

Wt of salt in fuel system = $1.85 x 10° 1b

- Cost of replacing all of the & e

solvent salt in the fuel system = $1.1 x 10 —

.

188,
. Assuming an interest charge of 4 percent per year on the cost of uranium

fuel inventory,

Optimum cost of uranium inventory in = $l.1 x 106 = $27.5 x 100

excess of that for the hot, clean 0.04 x Y Y
critical condition, so that the accumu-~

lated interest charge on this inventory

just equals the cost of fuel reprocess-

ing (replacing salt)

where, Y = cycle time, in years, to replace all the fused salt solvent in

the fuel system.

Asguming $20 per gram for highly emriched uranium-235, tkis is,

Optimum uranium inventory in excess of = 1.4 x 10° g
that for the hot, clean condition . Y

Volume of the fuel system = 2.7 x 10( cc

Optimum uranium concentration in excess == 0,051 g/cc

of hot, clean critical condition after Y ' !
reproces8sing equilibrium is attained
Uranium concentration for hot; clean = 0.0088 g/cc

ceritical condition

Equating the sum of these uranium concentrations to that required to
override the fission product poison (Section C.6),
The optimum cycle time to replace all = 19 years
the solvent salt in the fuel system '
It would initially require a period of time equal to the time of one

reprocessing cycle, during which no fuel reprocessing is performed, to attain

-this balance of uranium inventory and fission product poisons.

APPENDIX D

PRIMARY HEAT EXCHANGER DESIGN

D,1.0 DERIVATION OF FORMULAE FOR PARAMETRIC STUDY

The following basic formulae were used considering sodium coolant on
the tube side and fused salt fuel on the shell side.

D.1.1 Continuity Equation

w - Cav | - (D.1),

D.1.2 Flow Area on Shell Side for Triangular Pitch

»>
88

1/2 PT x 0.867 Pp - 1/2 rrao2 per 1/2 tube (D.2)
R

p—

. 5 o =
B867TP°2 - 7 a per tube
/ T 3 ° '

'D.1,3 Equivalent Diameter for Triangular Pitch for Shell and Tube Sides

A

Shell Side:
de = 4 ° flow area = 4 x l/2[(!86722r2 - QEL-doiZ | (D.3)
' wetted parameter 1/2 d, :
- vhere, Tqr =d, +s
- de = b 6867PT2 - I d02>
o Tdg
K Tube Side: 2 -

. de = 44 | ‘ | (D.&) .

190.

D.l.4 . Over-all Heat Transfer Coefficient for
Duplex Tube Based on Outside Surface

1 =1 41, /1lnrg + [ry 1n 1y To Ni + l, (D.5)
U bg k rijng |k ri/ss To 55 bya [Ti ‘

To
l =1 +ro Inrog+ 10 1N T+ T'o

Up by kyy ry kgg 11  hyg Ty

D.1.5 Film Coefficient between Fluid and Surface

Shell side - Dittus-Boelter Equation

Nu = h de = 0.023 [deV )0-8[ cp/0.h (D.6)
B3 P k
- Tube side - Martinelli and Lyon Equation

Nu = h dg = 7 40.025 (eig_‘_’_)(LB (31)4?0«:8 | (D.7)
& Iz k |

D.1.6 Heat Transfer Equation

q = Uy S (Atgp) - | . (D.8)

D.l1.7 Surface Area Required for Heat Transfer

So =Trd, LN ~ (Dp.9)

D.1.8 Combining the Above Basic Equations into Working Formulae

Then from (D.l) and (D.2) with constants for the conditions of this
design, the equations become:

D.1.9 Salt Flow

. wWg = 65[867 PT? ‘JL;L doajvs x N (D.10)

191,

D.1.10 Sodium Flow

, -
"Na = eNa _71?[_ d3” VNa N (D.11)

D.1.11 Salt Heat Transfer Coefficient

From (D.6),

| hg = 0.023 k (eac Vs)o -8 (Ef)o i , (D.12)
0.4 0.8 .
hs-0023k C&)OB( ) v . oy vg® S
| then,

bg = Cq Vg0 | ¢ 5

C, = 0.023 k_[Pde 0.8 4 A0k ' (D.1k)
| de A K | -

D.1.12 Sodium Heat Transfer Coefficient from (D.T)

hyag = 7 }_c_ 4+ 0.025 L edev 9"8 S.Ef‘o.a | (D.15)
de ‘ de /( k .
- hNa = C2 + C3 VQ-8
then,
=7 k_ and C3 = 0.025 .lf_ Cdg 0.8 32_@0"8 ' : | (D.16)
. de de M k , S

D.1.13 Over-all Heaf Transfer Coefficient in Terms of
: the Number of Tubes

From (D.8) and (D.9),

q =UoTTdg LN (Atgyg) (D.17)

192,

For a éiven Pp and d,, and for the flows required for heat transfer, from

(D.10),
VS = Ch.
N
and
YNa = 05
N
Cy = __¥s .
]
where,
CS = wNa
oo
from (D.17),
Ut.J N L =z Cg
where,
C6 = a
T, (BEeg)

from (D.5), (D.14) and (D.15),

= 1 + Ryy + Rgs +

1
Uo cl VSOoa

RNyi = ro 1n 1o

RSS = _1;9-_ln'rm .

C7 +Cg Vyal*O

(D.18)

(D.19)

(D.20)

(p.21)

(D.22)

(D.23)

To
CB :C3 I"i
To

Rewriting (D.24), inserting (D.18) and (D.19),

1
Yo

Cl c_hOos' c7+08 Eé 0.8
N : N

Therefore, U, is a function of N since,
1 = f (N), then also U, = f (N)
Uo

But from (D.22) and (D.30), then,

£f (N) NxL zCg

193.

-~ (p.27)

(D.28)

~(D.29)

(2.30

Now, as L is evaluated at some arbitrary value to agree with the requirements

of the problem,

f (N) N - Cg
L

r

- (D.31)

which is known, then there is only one value of N to satisfy the conditions

of the two simultaneous equations (D.22) and (D.29) with L established. These

two equations can be solved either by iteration or by graphical means.

latter procedure was selected for this study. -

The

19k,

D.l.14 Calculation of Pressure Drop and Horsepower Due to Friction

Sodium pressure drop,

Ap = (2.0+ 2L ) VW2 € . (D.32)
de 2g 14k

combining with (D.19),;

ADp = (200 +Clo) 09 (D°33)
N2
where,
¢ | | (D.34)
9 > ‘1L % 2g
Cig = fL x12 =202 (D.35)
de de
vhere L - 20 feet, do in inches.
Horespower,
Pnp =Ap x 144 x VNa | , (D.36)
CRa 250 :
where,
Cyq = b4 x wyg (1b/sec) (D.37)
Cyva 5% |

Salt pressure drop, combining with (D.20) and (D.32)

Ap = [&.o + 013] C1ow -~ (p.39)
where,
Cip = C° @ - | | | (D.40)

195.

Ci3 =fL_x 12'= 2ho t (D.k41)

where L = 20 feet.

Horsepower, '
where;
Cqp = 14h x fis (1v/sec) ' | | (D.43
} g 550

D.2.0 PRIMARY HEAT EXCHANGER SAMPLE CALCULATIONS

Solution of the two simultaneous equations discussed in Section D.1.13
results in a heat exchanger with 14,600 5/8-inch 0.D. duplex tubes with an
1/8-inch ligament and .065-inch wall as the optimum and most applicable to
this design. However, actual layout of the heat exchanger tube bundle indi-
cated that 2L éxchangers with 650 tubes each was a practical éf?aqééfienton
Nevertheless, the following sample calculation is based on é heat Efghéfiger
of 14,600 tubes with 0.065-inch duplex wall with 0.Ok2-inch Type 30hn??gin=
less steel and 0.023-inch Type "L" nickel clad.

D.2.1 Properties

Fuel Mixture Side
Salt No. 3%, 47 mol % NaF and 53 mol % ZrF),

Average Properties

C

' N

195 1b/ft3, /“= 19 1b/hr-ft, C, = 0.285 BTU/1b-°F,

1.30 BTU/hr-ft2-OF /£t

Inlet Temperature = 1200°F Outlet Temperature = 1050°F
- Sodium Coolant Side -
Average properties
© = 51.0 1/£t3, 4= 0.53 1v/hr-ft, C, = 0.30 BTU/1b-°F,
K

36,2:BTU/hr-ft2-9F/ft

196.
Inlet Temperature - 1000°F Outlet Temperature = 1150°F
Inlet Enthalpy = 338.5 BTU/1b Outlet Enthalpy = 293.7 BTU/1b

D.2.2 .Flows

Sodium

6

Wya = 600,000 x 3413 = 45.7 x 10° 1b/hr = 12.7 x 103 1b/sec

338.5 - 293.7

Salt
WS a
C, Ebfid

wvg = _ 600,000 x 3413 = 48.0 x 10® 1b/br = 13.3 x 103 1b/sec
0.285 (1200-1050)

D.2.3. Velocities

Sodium
‘ 2
v = VNa 4 =TT d N - No, of tubes
Na ETKI?E Ay e i
VNa = 45,7 X 106 x b x 14k - 12.8 ft/sec
3600 x 51.0 x (.625 - .13)2 x 14,600
‘ Salt
. VS = Vg Ap = .867 P T2 - Trd°2
C AoN
V. - 8.0 x 106 x 14k = 3.7 ft/sec

s = -
3600 x 195 x (.867 x .752 - _%;_x .6252) 14,600

197.

D.2.4 Equivalent Diameter

Sodium
N de = 44 = .625 - .13 = 0.495 in.
Salt
de = 4 (Ay) =4 (867TP2 - /4 a ®)
perimeter ‘ T dg
de = 4(.867 x .75° - 7T /4 x .6252) - 0.368 1n.

- T x .625

- D.2.5 Reynolds Number

Sodium i
Re = €de VNa
/-(
Re - 51.0 x 0.495 x 12.8 x 3600 = 183,000
12 x .53
Salt
Re - ede VS
Re - 195 x .368 x 3.7 x 3600 = 4200
12 x 19

D.2.6 Heat Transfer Coefficients and Wall Resistance

Sodium

hya de = 7 + 0.025 (Re)-8 (pr).8
’ X

Re x Pr - Pe

hya = 36.2 x 12 [7 4 0,025 (183,000) -8 (?30 X ,5€)°8;]; 10,750

36.2

198.
Salt

hg dg = 0.023 (Re)-8 (pr)-*
X

h =1.3x12 [0.023 (1:,2oo)°8 285 x 19 T = 1360
- 368 13 '

Wall resistance based on external tube surface

*

Kgg = 12.9 BTU/hr-rt°-F/rt
kg = 335 BIU/hr-ft°F/ft
Ryy =T Inrg =z .310 1o .312 z 0.595 x 107
kyy Tp 33-5 x 12 .289
Rys = 1
N 18,700
RSS - I'o ln I'm - 0310 ln 0289 - 3011" X lo-ll.
kg ri 12.0 x 12 257
Rgs = _L

D.2.7 Over-all Heat Transfer Coefficient Based.on External Tube Surface

- -l - ,
1 =1 +BRy +Reg+ 1 = 1 4.595x10"% 43,1k x10"+_ 1 (.2h7
U, bs LTS hyg Ty 1360 - 10,750 (.312
ro ——

1 =7.35x 10™* + .595 x 107 + 3.14 x 10°% + 1.18 x 10°%

Uo
2
U, = 815 BTU/hr-ft°-F

D.2.8 - Effective Length of Tube Required‘for Heat Transfer

q = Uy Sg (Btyp) Sg = Mo N L

L = q = 600,000 x 3413 x 12.
U, Td, N (atgg) 815 x T x .625 x 14,600 x 50

e
i

= 21 £t

P.2.9 Pressure Drop Through Tube Bundle

- Sodium
- (200 + fL )VNaa . e .
T 2 1

1

5.'
&
o
o
'

= 183,000, £ = 0.016 (Ref. 16, p. 15).

D>
d

= (2.0 4 £L) vg?
de 2SI&

_(20+.0h—0x2lx12)372xl95-85psi
368 2g 14

D
g

D.2.10 _ Pumping Power Required for Heat Exchanger Only

Php = aop x 14k = 9.1 x 144 x 12.7 x 103 = 594 hp
AP Na
%550 51 x 550

o vPhp =aAp x 14l wg = 8.5 x 14l x 13.3 x 103 = 152 hp
: . € 550 195 x 550

D.2.11 _Fuel Holdup in Tube Bundle
Salt area flow Ayg = [867]? - Tr a 2]n
’

 Salt volume Vol = Agg x L, assuming L = 21.0 £t

Vol -_-' .867 X .52 - 1T .625] 1k ,600 x 21
‘ B 1Lk

Vol = 385 £t

D.2.12 Sizing the Fuel Circulating Pumps

Pressure drop in tube bundle 8.5 psi
Outer flow periphery 1.0
Check valve 2.8
Contingency 2.7
Total 15.0 psi

Total pump power

With 8 pumps

Assume pump efficiency at 60%

Pnp = AD X 1k wg wg = 13.3 x 106 1b/sec

550 %

Php = 15 x 14k x 13.3 x 103 = 446 np
195 x 550 x .60

Horsepower each pump - 56 hp

Specify 60 hp motor driven pumps

200,

hg]
e

201.

APPENDIX E

CALCULATIONS FOR PRELIMINARY DESIGN OF INTERMEDIATE HEAT EXCHANGER

E.1 DESIGN PARAMETERS

The intermediate heat exchangers are designed to transfer 600 Mw of
heat from radioactive sodium in the primary loop to non-radioactive sodium
in the intermediate loop. The design 1f based upon the following, previously

- selected sodium temperatures and heat exchange rate:

) ~ Primary Na entering “tpy .11509F
Primary Na leaving thé 1000°F
Intermediate Na enterifig tél' 6?59F
Intermediate Na leaving tc2 1100°F
Total heat to be transferred --- 2.048 x 109 BTU/hr

Heat losses from the reactor vessel, primary ldop piping and immediate
heat exchanger have heen ignored.
The basic assumption was made to use six intermediateflheat exchangers.

E.2 PROPERTIES OF SODIUM AND 304 SS

Sodium
} Temperature, OF 1150 1100 1000 675
Density, 1b/ft3 49.87 50.30 41.15 53.92
) Enthalpy, BTU/1b 338.5 323.6 293.7 195.7
Specific Heat, BIU/Ib - 0.2983 0.2983 0.2988 0.3052
Thermal Conductivity, 35.7 36.4 37.7 42,1
. BTU/hr-ft2-F/ft
Viscosity, 1b/ft-sec x 103 0.136 0.141 0.152 0.201
304k Stainless Steel
Thermal Conductivity 13.2 13.0 12.6 11.3

BTU /hr-ft°-F /£t

E.3 SODIUM FLOWS

a. Primary loop flow - q = 2.048 x 107 BTU/hr - = 45.7 x lO6 1v/hr
OH  (336.5 - 293.7) BTU/1b

b. Intermediate loop flow =z g = _ 2.048 x 10° BIU/hr _ = 16.0 x 10° 1v/hr
AH  (323.6 - 195.7) BTU/br

E.4 SODIUM VELOCITIES IN HEAT EXCHANGER

Preliminary calculations indicated that 1100 1/2-inch 0.D. tubes with
0.042-inch wall appear to be satisfactory.
a. Primary sodium velocity (shell side)

Assume that the tubes are in a triangular pattern with a
3/4-inch pitch

Using Figure B.4 for 1100 tubes,

Radius of tube bundle = R = 17.h4
Tube Pitch P

R = (17.4)(P) = (17.4)(3/4) = 13.05 in.

Shell I.D. = tube bundle dia + tube dia + 2 in. free space

f

(2)(13.05) + 0.50 + 2.0

28.6 in. (Use 30 in.)

Shell cross section area - (O°785)(§(_)_)2 - 4,91 £t°
12

Total tube cross section area = (0°785)£005)2 (1100) = 1.50 £t2
144

Shell side flow area = 4,91 - 1,50 = 3.k1 £t2

6 7.62 x 100 1b/hr

Na flow per exchanger = 45.7 x 10
. 6

Average Na density = 49.87 + 51.15 = 50.51 1b/ft3
| 2

(7.62 x 106'lb/hr)

Shell side velocity

(50.51 1b/£t3)(3.41 £t2)(3600 sec/hr)
12.29 ft/sec '

Y

203 o
b. Intermediate sodium velocity (tube side)
Flow area per tube = 0.1358 1n®

Tube side flow area - (0.1358)(1100) = 1.037 £t°
=T

Na flow per exchanger = 16.0 x 106 1b/hr = 2.67 x 106 1b/hr
. =

Average Na density - 50.30 + 53.92 - 52.11 1b/ft3
)

(2.67 x lO6 lb/hg)
(52.1 1b/£t3)(1.037 £t2)(3600 sec/hr)

Tube side velocity

= 13.73 ft/sec

E.5 EQUIVALENT DIAMETER FOR SHELL SIDE

de = 4 x shell side flow area = (4)(3.41 £t2)(1kk in.®/ft) - 1.137 in.
- wetted perimeter (1100)(3.14)(0.5 in.)

E.6 REYNOLDS: NUMBER

Primary Sodium (shell side)

Average Na viscosity =(b,136 +0.152) x 1073 - 0.14k x 1073 1b/ft-sec

2 ./
Re = X#aL = (12.29 ft/sec)(1.137 in.)(50.51 1b/ft3) = 3.75 x 10
Z< (12 in/ft)(0.0001kk 1b/ft-sec)

Intermediate Sodium (tube side)

Average Na viscosity = £0.141 + 0.201) x 1073 : 0.171 x 10~3 1b/ft-sec

N 2 J
Re = #d€ =(13.73 ft/sec)(0.416 in.)(52.11 1b/ft3) = 1.45 x 107
)3 (12 in/£t)(0.000171 1b/ft-sec) _

E.7 HEAT TRANSFER COEFFICIENTS

Film Coefficient Primary Sodium (shell side)

Average specific heat 146.2983 4+ 0.2988\ = 0,2985 BTU/1b-F
| \ 2

Average thermal conductivity =(35.7 + 37.7\= 36.7 BTU/hr-ft2-F/ft
—

20k,

Pr ='&';F a (0.2985 BTU/1b-F)(0.00014k 1b/ft-sec)(3600 sec) = 0.00425

" (36.7 BTU/hr-£t2-F/ft)(hr)
Using an empirical correlation for shell side of unbaffled liquid metal .
heat exchangers (Ref. 15, p. 285),
Nu - 0.212 [(de)(Re)(Pr)]O°6 = 0,212 [(1.,13,7)(30*5 x 105)(o¢,00‘t+25fl°“6 .
Nu - Eg_: 19.08
X
hy =

19.08 (E_)= (19.08)(36.7 BTU/hr-£t-F)(12 in/ft) - 8402 BTU/hr-ft°-F

de (1.137 in.)

Film Coefficient Intermediate Sodium (tube side)

Average specific heat = 0.2983 + 0.3052 =z 0.3018 BTU/1b-F
2

Average thermal conductivity = 36.4 + 42.1 = 39.3 BTU/hr-ft-F
2

Using Lyon-Martinelli relation for liquid metals and a uniform wall

heat flux,
Pe - AXCC, = (0.416 1n.)(13.73 ft/sec)(52.11 1b/t3)(0.3018 BTU/1b-F)x3600 soc
X (12 in/ft)(39.3 BTU/hr-£t-F)
Pe - 685.7

Nu = 7 + 0.025 peQ-8

Nu = by = 7 + (0.025)(685.7)0-8 _ 11.62 _
K
h; = 11.62 (k) = (11.62)(39.3 BIU/hr-ft-F)(12 in/ft) -

(€)) (0.416 in.)

by, = 13,173 BTU/br-ft2-F

Tube Wall Thermal Conductivity

Average thermal conductivity = 13.2 + 13.2 + 12.6 + 11.3
]+ .

12.52 BTU/hr-ft-F

205.

E.8 OVER-ALL COEFFICIENT OF HEAT TRANSFER

do
1 dy +dp.1n3 +1
' Uo Mdy Kewe  Bo
| 0.50
- (0.50) + (0.50} 1n EII?TG'F 1

(13,173)(0.Lk16) (12.52)(12) 8402

it

0.0000912 + 0.000612 + 0.000119

0.0008222

U, = 1216 BTU/br-£t°F

- E.9 . HEAT TRANSFER SURFACE AREA

Logarithmic Mean Temperature Difference
At = (th2 - tep) - (tny - tep)
In thy - te;
(1000 - 675) - (1150 - 1100)

1n /1000 - 675
(1150 - 1109

= 147°F

Total Heat Transfer Surface

q '.'."Uo So At

. Sy =_ g =__(2.048 % 107 BTIU/hr) = 11,457 £t?
U, Ot (1216 BTU/hr-£t2-F)(147°F)

- Heat Transfer Surface per Heat Exchanger

S = 11,b57
_%_._

E.10 EFFECTIVE TUBE LENGTH

- 1910 ft°

L S - (1910 £t2)(12 in/ft) = 13.27 £t

- (N)(Tube perimeter) (1100 tubes)(0.50 in.)(3.1l4)

206.
E-11 SODIUM PRESSURE DROP THROUGH HEAT EXCHANGER

Pressure Drop for Primary Sodium (shell eide)

From Section D.6, the Reynolds number for the shell side is 3.75 x 107
and the corresponding friction factor is f = .01k,
From Reference 39, Chapter 1.5, the pressure drop is given by,

AP = £ L x Cy2

de 2g
AP .(.,011+ x 13,27\::(50,51 x 12.299)
1.137 /] | ek x 12 J
NP = 1.61 p{ai

Pressure Drop for Intermediate Sodifim (tube side)

From Section D.6, the Reynolds number for the tube side is 1.45 x 107
and the corresponding friction factor is £ - .0l7. In this case, the pres-

sure drop is given by Equation (5.8),

AP = £ £f1L\ x £V?
de 2g

where 2 1is the number of velocity leads allowed for the tube sheet entrance

and exit losses and the 180° bend.

AP =/2 4 .017 x 13.27 x 12\ x{52.11 x 130739\
316 oh.k x 164 J

OP - 9.02 psi

207,

APPENDIX F

STEAM GENERATOR CALCULATIONS

. F.l NATURAL CIRCULATION BOILER WITH SEPARATE SUPERHEATER

From superheater heat balance,
8boslers = 16,000,000 (127.9 - 42) = 1,372,000,000 BTU/hr
Assume two bollers genergting saturated steam-at 2,000 psié, steam
on the shell side and sodium on the tube side. Say désign is for 1000°F
- and 2000 psi external pressure, the fhickness of a tube for external pres-
sure can be calculated (Ref. 35, paragraph UG-31).
For 30k sS,
"S" allowable = 8800 psi (Ref. 35, Table UHA-23)
. X ¥ .19 (min)
".For 3/k in. 0.D. tubes, X min = .143
' Use .165 average wall (8 BWG)

For a Single Wall Separating the Fluids Wall Resistance,

ro Inrg = 0.375  1n .750 = .00Lk6
"k ry 12.hx 12 20

.Water Side

For present purposes, the assumption will be made that thé unit opefates
in the nucleate boiling region and latér on this assumption will bé verified.
In order to detérmine a reasonable value of the average boiling and scale
coefficients for design, resort will be made to both the-PWR and HRT designs.

mRY -

From ORNL-LR-Dwg-2795A, there are 251 3/8-1nch 0.D. x .065-in. wall type

347 SS tubes of 20 ft effective length.

e

208,

Surface (actual outside) = 77 x .375 x 251 x 20 - 494 sq ft
15

On Page (8) of Spec. HRT-1004, the fluid properties, flow rates and

temperatures are given,

572 -
T 4ok .5
471 )
—
Atol = 7705 = 5302
In 101
35
U(Design) = 1.71 x 107 = 652
kol x 53.2
Tube Wall
ro Inry = .1875 1n .375 = .00625
k rf 10.6 x 12 245
Primary Fluid
hd = .023 ( ) (dV€
Flow area = ( .245)2 x 251 - .0822 sq ft
Ny
Velocity = 1.79 x 102 = 11.k fps )

3.6 x 103 x .0822 x 53.2

.339 x .023 x 12 (1 23 x 266\“ ..%2*_5 x 53.4 x 11.4 x 3600\° - 5800

245 .339 / 266 /
h (based on outside surface) = 5800 x .245 = 3790 |
.375 ' e
1 = .000264

3790

209.

The over-all U for the HRT was based on using 3/4 x U clean.

R | =.00115 = 1
652 x 4/3 U clean

"

(wall and primary resistance) = .00890

R | = 000890 = .00026
. U clean
h (Boiling) - 1 = 5100; say, boiling coefficient
-00026 x 3/ of 5,000

A similar calculation on the PWR design (Ref. 36) indicates that these

units héd a degign boiling plus scale coefficient of the order of magnitude

- of 2000, This indicates a scale'coeffiéiént of the order of magnitude of
3300 average was used on the steam side of these units. For the present
design, an average value of scale plus nucleate boiling coefficient of 2000
is fised. .This should be conservative for design purposes, since the bigher .
average temperature differentials in this unit should yield higher bo;ling
coefficients than either the PWR or HRT designs.

For a natural circulation boiler, the steam water mixture will be
external to the tubes and the sodium will be internal to fhe tu.beg° On the
basis of preliminary calculations, it appeared that a reasonable size and
pressure drop would be realized in two units with approximately 2000 3/k-inch

0.D. tubes to carry in each.

i Sodium Side
hd = 74 .025 (d_vf cp>'8
*® ~
d = 0.42 ft
x 12
K = 40 BTU/hr-£t2-OF /ft

Cp = -302 (avg) BTU/1b-°F

V - 21.9 fps

/<= .178 x 3600 x 1073 - .64 1bs/ft-hr

52,7 lbe/ft3

o
<
m

11

h =12 x 40| 7 + .025 h2 x 21.9 x 52.7 x 3600 x 30%)
A2 Lo
h = 15,700 |
h based on outside tube surface = 15,700 x .42 = 8800
.15
1 - .000113
00
Flow area = 7T x (.42)° x 2000 = 1.93 sq ft
N 1hh
V = 8,000,000 = 21.9 fps
3600 x 1.93 x 52.7
1 - (Boiling 4 scale) = .0005
h
1 = 0.002073
U0
Uy = 483
Say, U, for design = 48O
960
675
635.8-
Atog = 285 = 135
in 324.2
39.2

' Surface Area (Boiler) = 686,000,000 = 10,600 sq ft
j4»8021{135

211.

377 x 2000 x L effective = 10,600 -
L8

L effective = 26.9 ft.
Say, total length (including tube sheets) = 28.5 ft

Reynolds number = L%g x 3600 x 21.9 x 52.7 :‘229,000

- : qan
£ - .015 |
| Ap - [2 4 .015 x 28.5 x 1%] x (21.9)% x 52.7 psi
_ _ [j 42 4 - 2g 14
AP = 38.6 psi

Check of assumption of nucleate boiling,
At inlet At - 324.2°F
(q/s)inlet - 324.2 x 480 - 156,000
As indicated (Ref. 4 and 6), this heat flux is still in the nucleate boiling

region. As a further check, assume the tubes to be. clean.

1 - .00208
L8o

1 - .000303
3300

(1/Uy) clean = .00LTT7
v Uy clean = 564 - |
q/S inlet clean - 324.2 x 56u = 182,500

At £ilm = 182,500 a 36.6°F
5000

This is less than the critical At of 45 tc®50 F (Ref. 6, p. 386). There-
fore, the assumption of nucleate boiling is Justified.
Size of steam generator: R/P : zu.9 (Fig. B.h)

Diameter of channel: 20.9 x 2 x 1 + 2.75 = 44.6, say, 47 in. at

ID of channel - -

212,

A preliminary check of the circulation characteristics of the steam
generator indicated that a reasonable number of risers and downcomers can
be provided to give adequate circulation for these units. The steam drums A
would each be approximately 66 inches ID x 45 feet over-all length to give
commercial steam purities with commercial type feed water treatment. The
center line of drums would be approximately 30 feet above the center line
of the exchangers. Drum supporting steel Qill be provided and the support
"~ saddles on the heat exchangers will be designed so that one saddle is fixed -
to the foundation and one is let free to move to accomocdate the expansion
of the unit and/or the piping.

Over-all Heat Balance: Boiler and Superheater

Assume boiler at 2000 psia and superheater outlet to be 1000°F and
1800 psia.
Total heat load: 600 Mw
Q BTU/hr - 600 x 3413 x 1000
H superheated oteam @ 10N0°F 4 1800 psia - 1480.8 BTU/1b
H feed @ 450°F + 1800 psia = 430.9 BTU/1b
AH - 1049.9 BTU/1b
Total 1bs of steam produced = 600 x 3413 x 1060 = 1,950,000 lbs/hr -

1049.9
Temperatures of intermediate loop are as indicated on heat balance

diagram Figure 6.2.
H sodium @ 1100°F - 323.6
H sodium @ 675°F = 195.7 <
Ah - 127.9 BTU/1b

600 x 3413 x 1000 = 16,000,000 1lbs/hr sodium (total)
127.9

AH Sodium/1b = 1,950,000 x 345.7 - 42 BTU/lb
16,000,000

AH superheater 1480.8

1135.1
AH = 3B5.7
A Eya entering superheater - 323.6
- 42.0
H = 281.6
t = 960°F

At 1100°F, "s" is 7500 psi Table UHA-23 ASME Unified Pressure Vessel
Code - Section 8. The minimum tube wall thickness required is,

‘%t = PRg (PUA-1) = 2000 x .25 = .0602(nomenclature according to ASME)

SE + &P 7500 -+ 800

Use 0.065 inch minimum wall tube.

If two walls are used to separate the water and sodium, and each wall
is required to withstand the full pressure, then the o;xter tube required will
be .6660D x .5001D with a wall thickness of 0.083 minimum. |

A. Single Wall Separating Both Fluids

Agsumed sodium flowing parallel to the tubes and that the unit is of

the counter flow type,

1100 o :
\ 960 E S
t [NOOO |

635.8

At g = 324.2 - 100 = 191°F
. . 32%.2
100 ' .
Tube resistance = rg ln ry (based on outside tube surface)
R 7
K - 12.4 BTU/hr-£t2-°F/ft

' Tube resistance - .25 1n .500 = .00050k4

12.4k x 12 « 370

21k,
Preliminary calculations indicated that a reasonable size of super-
heater could be achieved by using 1275 tubes to carry per unit. Two such
units will be required. : o #

Flow area/tube = %E_(°37)2 - .1075 'sq in.

Water (steam) side coefficient, . .

N o8
a2 zy (. ¥W__
1000a

. Where,
W = 1,000,000 1bs/hr per unit
b= _ 175 1,000,000 B 175 ,
(.370)-= 1000 x .1075 x 1275 .82 - o

(7.3)°8 - 1046 BTU/br-£12-OF
Based on cutside surface; this is,

h = 1046 x .370 = TT75
. 500

 Resistance = 1 = ,00129
175

Sodium Side

The free area/tube = .866 x °752 - T (1/2)2 = .290
E ’
The perimeter/tube = 77 = 1.57
: 2

=4 x Area = 4 x .290 = O.Th .in.
perimeter  1.57 -

The equivalent dia

For the sodium side heat transfer, the Liquid Metals Handbook indicates
that the Lyon-Martinelli or some modification is satiéfactdry9i Sincé'the
sodium is not the controlling resistance, and for the sake of consistency,

the Lyon%fiartinelli correlation will be used.

215,

hde = 7 + .025 /dgVP Cp -8
R R

. De - 007,4‘ ft
- 12

Pavg =20- 3'+ 51.67 = 50.9 lbs/cu ft

RE36.U4 'f;_ 38,2 - 37.3 BTU/hr-ft2-OF/ft
Cp = +299 BTU/1b-°F
Based on main bank geometry,

Free area (total) - 1275 x 0.29 = 370 sq in.

) W = f’x AxV
Vv = 8,000,000 x 14k - 17 ft/sec
50.9 x 3600 x 370 |
h = 37.3 x 12 /; + .025 7h x 17 x 3600 x 50.9 x °2950
i 373
- 8200 BTU/hr-£t2-CF
Resistance = 1 = .000122
8200
1 = .00050% 4+ .00129 + .000122
UO
Uo = 522

For design purposes, assume U, = 500

Surface required =z 1,950,000 x 345, 7 = 3520 5q ft

unit ; 500 x 191 x 2
Active tube length = = 3520 = 21.1 ft
) : 1275 x
e

Say, approximate tube length (including tube sheets) - 22 1/L ft
From tube count layout sheet,

- +  "R/P = 18.82

Dia to outer tube - 2 x 18.82 x 3/4 = 28.2

216,
Approximate dia of channel = 28.2 + .5+ .5+ 1 1/2 = 30.7
For clearance; etc.; say, ID of channel 32 inches.

Pressure Drop

The following calculations are for the pressure drop in the exchanger.
The pressure drop in the inlet and outlet nozzles is not included as part
of the pressure drop of the components. This 1s estimated elsewhere.

Sodium Side

Reynolds No. = DVE = ‘% x 1T x 3600 x 50.9 = 356,000
< St
f = .01k for smooth tubes

AP

(105+,01hx21,1x12\ v2 @ -63v2¢0
.T% ] 2g 14k 2g 1LL

603x%x%gg510psi

For bend use 1/2 velocity head,

OHP Bed - 0.79 psi

say, ll psl for exchanger only

Steam Side

Assume flow and friction pressure drop are essentially that for am
incompressible fluid. If the calculations on this assumption show a high
pressure drop, then the variation of specific volume with tube length may
have to be considered. The acceleration loss due to the change in specific
_ volume between inlet and outlet will be considered.

In terms of our nomenclature and the assumed flow conditions, the general

pressure drop equaticn becomes,

- 2 2
Py -Pp = f L Vayg Cavg + vi) 1 (Vo - V1) + say, 1.5 vel heads,
d 2g 1hk4 A) (3600)2g 144

entrance and exit 4+ .5 vel heads bend

217.

Cavg = 5»3';—203' - 3.8

v, =.188
v2 g 0M2 -

1,950,000 lbs/hr total

=3
e

- 010§5 x 1275 = .954% sq ft/unit
1

1,950,000 = 4.8 fps
3600 x .95% x 3.8

Vavg =

Reynolds No. = DVC - Qigl x T.8 x 3600 x 3.8 = 48,000
u .068

f = 021 for smooth tubes

021 x 22,3 x 12 x (7 u 822 x 3.8 + 2(74.8)° x 3.8

AP
.37 1L% 2g 1LL
/1,950,000 e
X 2 hh2 - .88
\3600 x 954 32 2
:172(7h8)2x38+(28h) x (k2 - .188) x 1
2g L 1R 32.2

39.4 + 4.4 = 43,8, say, 4k psi

Since this represents but a small part of the pressure of the system,
it will bé satisfactory to take the pressure drop as being this value. A
more accurate calculation, taking into account the variation of specific
volume with length, will give a higher pressure drop; but not significantly
so. For system design purposes, asaumé a nominal 50 psi pressufe drop on
the steam side, not including nozzles.

B. Two Walls Separating Both Fluids

" The assumption made in this calculation will be the same as for the

single wall separating the two fluids. Because of the thicker tube assembly,

the 3/4-inch pitch cannot be used. It will be assumed that a 1/k-inch clear

218.

ligament will be required for welding of these tubes into tube sheets. It

will further be assumed that a thermal bond can be realized between the inner

and outer tubes,; but that this bond is not a mechanical bond. Thus, the

failure of one fiall will not give rise to & complete failure of the unit. It

is further realized that radial temperature gradients; in Some instances; may

tend to cause separation of the two tube walls. However, in the superheater,

it may be possible to use two different types of stainless steels with

\

S

e

N

/

05007

coefficients of expansion in such a ratio as to

negate the temperature differentials.

Pitch = .666 +-°250 = .916 in.

For tube -sheet layout assume ;g in. pitch (.9375)
' 1

Log Mean Temperature Difference = 191°F

For Preliminary Purposes, neglect sodium coefficient

Assume 1500 tubes to carry

.370 {Based on OD of outer tube)

-]

510 BTU/hr-£t2-°F .(outer tube OD)

< 666 >
Steam Side
h - 1046 x'(iél;)°8 x
1500
h =
Resistance : 0.00196
Wall

Inner tube

Outer tube = .333

1
Uo

.00050k x .666 = .00067

. 500
1n .666 = .0006k

124 x 12

500

1 - .00067 + .00064 + .000196 - .000327.

219,

Uy = 306, say, 300. It will be necessary to compute the sodium side

afl this point.

[866 x_(.9375)° '%;L ( 666)"17& = 0.79 in.

77 X
W = Cpx A x V
A - 1500 x 413 = 4.3 sq £t
14k
V = 8,000,000 - =10.2 fps

50.9 x 3600 x 4.3
hd, = 7+ .025 /a (eVP CID"S

h -
0.79 37-3

h = 566 [7 + .025 (1780)] - 6500

1 = .000154

500

1 = .003k2

UO

U, = 282

Say for design purposes,

U, =

Sreqy = 1,950,000 x 345.7 = 6400 sq ft/unit

275 x 191 x 2

~.37°3X1?f/+( .025) (79x102x3600x509x 299)7

275 BTU/hr-fte-QF (outer tube outside surface)

Surface per unit = 1500 x 7T (.666) x L effective = 263 L effective

12

s
Leff s 6&00 s 24 ft, 5 in.
260 .

Say total tube length = 25 ft, 9 in..

From tube count layout sheet,

R/P = 20.3

Dia to L of outer tube - 2 x 20.3 x .9375 =

38;1 in.

Approximate channel dia = 38.1 +-2 5 = 40.6 in.
For clearance, etc., say ID of channe1==h2 in.

Pressure Drop

Sodium Side

.79 x 10.2 x 356,000 = 235,000
i 17

Reynolds No. =

f - 0015

AP = (? $+ .015 x 2k b2 x 12) x (?0,2 2 x 0.9 = 3.8, say, b psi

-T9 2g 1

Steam Side

Reynolds No. = 63.5 x 48,000 = 40,800
4.8

f - 0021

37 x 2g / | 1k

Ap friction + Ap acceleration = 021 (25 .75 x ) ( 63.5)°

+2635> x38+
g

x (M2 - .188) x 1

1% 32.2

- 32.4 + 3.2 - 35.6 psi

F.2 STRAIGHT-THROUGH BOILER

1. Steam FPlow

Steam out at 1900 psia, 1000°F, H = 1W77.7
Feed water in at 1950 psia, 450°F, H = 430.1
Change in enthalpy - 1047.6 BTU/1b

Steam flow = 600 Mw % 3%13 x 103 BTU/Mw-hr
1047.6 BTU/1b

Steam flow = 1.955 x 100 1v/hr

1,950,000 2
.2 x 1275

1k 3600 x 1500

say, 4O psi
for desigg

221. .

2. Sodium Flow

Na Temp. in = 1100°F
Y Na Temp. out = 675°F
At average Na temp.  (887.5°F), c, = 304 BTU/1b-°F

600 Mv x 3413 x 10° BTU/Mv-hr

o Sodium flow =
' -304 BTU x L250F
1b-OF
Sodium flow = 15.85 x lO6 1b
hr

3. Change in Sodium Temperature through Preheat, Evaporator

- and Superheater Zones

The change in sodium temperature in a zone is found from a heat
talance between heat gained by the water and heat given up by the sodium.
(Na flow) cp'At = (steam flow) AH

Dtyg = Woag

WN& C’P

The sodium temperature change calculation is tabulated below:

zone = e s
preheat 2%6 . « 3090 95
evaporator h73 « 3050 ' :190
superheater | 539 - - 3015 ” lég

| 425°F

4, Calculation of Over-all Coefficient of Heat Transfer (U) in Pre-

heat, Evaporator and Superheat Zones (Refer to Section 6.5.0)

A. Preheat Section

a. Since the water entering the boiler is subcooled, fiucleate
boiling will occur. A combined coefficient for boiling film .
" : and scale of 2000 B’I’U/hr £t2-OF was used (Section 6. 5 1).

222,
b. Tube wall coefficient
k - 11.2 BTU/hr-£t2-°F/ft for 304 S5 at T23°F

‘h =k = 11.2 - 1790 BTU/hr-£t2-OF

ft .25 In .2
12 185

c. Sodium film coefficient

k = 41.b BTU/hr-ftach/ft |

h zk [7 + (0.025 DV f’cp)°°8] (Ref. T)
d X

d = 0.0618 ft

V = 13.25 ft/sec

- 53.6 1b/ft3

0
'

0.3085 BTU/1b-CF

Ca
g
"

h = 41.4 {7+ 0.025 (.0618 x 13.25 x 53.6 x 3085)0"8

5400 BTU/hr-£t2 -OF

1= 1 + 1 4+ _1 = .001453
U 2000 x .185 5K00 1790 -
.25

U = 690 BTU/hr-£t°-°F

B. Bo@;ing Section

a. Use a boiling film and scale coefficient of 2000 BTU/hr-fta-QF !
b. Tube wall coefficient

k = 11.8 BTU/hr-£t2-OF /£t for 304 SS at 865°F

h=k _ 11,8 = 1890 BTU/hr-rt2-°F
£t .25 1n L2 ‘

12 .185

c. Sodium film coefficient

39.3 BTU/hr-£t2-OF /£t

w‘
1)

<
"

13.25 f£t/sec

oLy

o7

.0618 ft

AN
)

52.35 1b/ft3

p = 0.3047 BTU /1b -OF

h =k |T+0.025 (va’cp)°°é]
a k . |
- 39.3 |7+ 0,025 {.0618 x 13.25 x 52.35 x. .30k7)""
,0618 39.3
- 4460 BTU/hr-ft2-COF |
= 1 + 1 4+ 1 =-0.001k29
2000 x .185 - 1890 4REO-
- 25
U - 700 BTU/hr-ft2-OF

C. Superheater Section

a.

Steam film coefficient

hde = 0:023 [dev \O-® cp/4l/3
Kk ' = k|

Steam temp. out = 1000°F

Steam temp. in - 620°F

Average temp. sz 810°F
na®-2 . Y(w/1000a)°-8

W =2 x 10° = 625 1b/nr per tube

3200
Y - 165
d - dia, in. = 0.5 - 2 x 0.065 = 0.37 in.
A = flow area 1in.2 = 0.1076 in.° -

h x (,'37)0°2 ;'165 (625/,0716)0*’8

h = 1100 BTU/hr-£t€-°F

-Tube wall coefficient

h =k = __ 11.8  =.1890 BTU/br-ft>-°F
ft .25 In .25
12 165

)

223,

c. Sodium film coefficient

h -k [Tr + 0.025 (dvfcp)°°8]
d X
d - 0.0618 £t
V - 13.56 ft/sec
€ - 51.0 1w/rt3
k - 37.3 BTU/hr-£t2-OF /£t
Cp = 0.3015 BTU/1b/°F

h = 37.3 [? +0.025 (,0618 x 13.56 x 51.0 x 3015 x 3600)°

.0618

8450 BTU/hr-£t2-OF

37-3

.2 .
+ .25 1n jigg-f 1 - .003358

37.3

1= 1
U 1100 x :185
025
U - 300 BTU/hr-£t2-OF

5. Calculation of Zone Surface Required

8450

y

In a zone, the heat gained by water equals the heat lost by sodium and

this heat is trensferred by the equation, q = UAQAt, where &t is the log

mean temperature difference.

re U
4608 x 10° 690
9252 x 107 700
6624 x 10° 300

6. Number of Tubes

A

3,741 £t°

5,981 £t°

11,470 £t°

21,192 f£t°

‘hAs a result of checking pressure drop and steam velocities; a tube

size of 1/2-inch OD and nominal 50?f00t length was decided upon.

| 225,
Total tube area required from (5) above = 21,192 £42
Area per tube,TTDL = 6.54 £t°

No. of tubes - 21192 - 3240 tubes
6.5k

7. Pressure Drop in Tubes (Ref. 7, p. 36, l(5, 55, 69)(Section 6.5.2)

The fraction of the 50-foot tube length required by the preheat,
evaporator and superheat zones is proportionai to the percent of ares of
(5) above. | | |

8., Preheat zone

26,800 = 26,800 = 112,000

Re =
/A~ 239
£ = .019
Op = FV2 L e
2g de |
Ap = .019 x 5.14% x 8.85 x L47.1
6.4 x .0308 |
- Ap = 105 1b/£t° - .73 psi

b. Evaporstor zone {two phase flow)

Re = 26,800 = 26,800 = 132,500
M .202
f - .018
) AP = -018 x 6.13° x 14,1 x 39.5
bh.bk x .0308
fl APo = 190.4 lb/ft2 (single phase)
Appp = APy (APgps) + r (G)° (Ref. 7, p. T2)
AP, g
Appp = 19C.4 (6) + .2 x (2u2)°
32.2
APpp = 1506 1b/f1:2 z 10.5 psi (two phase)

c. Superheater zone (Ref. 7, p. 69)

Re - 26,800 = 26,800 = 413,000
/"{ 0065

d.

226,

£ - .015
AP = G2 flv +§2 (V2 - Vl)
2g de g _
Ap = (242)2  x .015 x 27.05 x .305 ¢ (242)2 (.4165 - .1950)
Bk x L0308 3.0

AP = 4059 1b/ft° - 28.2 psi
Total tube pressure drop

The total drop is the sum of a, b, and ¢ above.

(a) 0.73
(v) 10.50
() 28.20

Total Dfop = 39.43 psi

M 227 °

APPENDIX G

CALCULATIONS OF STEAM PLANT HEAT BALANCE .

The calculations below determine the‘heat balance and gross electric
& povwer output of the steam cycle described in Section 6,8°O‘of this report.
Thé over-all cycle and the conditions of the working fluids are shown in
Figure 6.2,

- G.l THROTTLE STEAM FLOW

The steam flow at the turbine throttle is calculated from Equation 6.11.

Steam flow = 600 Mw x 34:3 x 103 BTU/Mw-hr
(1577.7 - 430.1) BTU/1b

Steam flow

1.955 x 10° 1
hr

. G.2- EXTRACTION STEAM FLOWS

No. 1 Heater heat balance

1.955 x 10° 1b (430.1 - 364.2) BTU - Wy (1344 - 430.1) BTU
hr 1b 1b
Wy = .14 x 10° 1b
hr

No. 2 Heater heat balance

- 1.955 x 106 (364.2 - 300.7) = .14l x 106 (430.1 - 364.2)
+ W, (1286 - 36h4.2)
Wo = 1246 x 108 1b/hr

No. 3 Heater mass balance and heat balance

6

s I. 1.955 x 106 = (.11 + .1246) x 10 +w3'+Fow,

6

II. 1.955 x 106 x 300.7 = (.141 4 .1246) x 10~ x 36L.2 + 1222 w3

+238.8 F.W.

228,
The solution of Equations I and II above gives,

Wy z .089 x 10° 1
hr

6

F.W. = 1.6 x 10” 1b (feed water flow leaving No. 3 Heater)

No. 4 Heater heat balance

1.6 x 106 (238.8 - 178.0) = (1164 - 243.9) Wy

Wy = .1057 x 10° 1
- hr

No. 5 Heater heat balance

1.6 x 100 (178.0 - 117.9) = .1057 x 106 (243.9 - 183.1) + (1095 » 183.1) Ws

W5 = .0984 x 100 1b
hr

No. 6 Heater heat balance

1.6 x 106 (117.9 - 59.7) = (.1057 + .0984) x 106 (183.1 - 122.9)
+ (1024 - 122.9) Wg

WE = 0877 1b
hr

G.3 HEAT CONVERTED TO WORK IN TURBINE (Equatiop 6.10)

Position Steam Flow (1b/hr) OH (BTU/1b) Work (BTU/hr)
Throttle 1.9550 x 10° 136.3 266.5 x 106
No. 1 Heater 1.8140 x 10° 58.0 - 105.2 x 108
No. 2 Heater  1.689% x 10° 64 .0 108.1 x 106
No. 3 Heater 1.600% x 106 58.0 92.8 x 108
No. L4 Heater 1.4947 x 10° 69.0 103.1 x 106
No. 5 Heater  1.3963 x 106 71.0 99.1 x 108
No. 6 Hea£er 1;3666 x 108 82.6 107.1 x 106

Total of heat converted to work = 881.9 x 10° BTU/hr

¢

G.4t CORRECTION TO WORK FOR FEED WATER HEATING BY PUMPS

a. Correction for Hot Well Pump

» The work by the pump done on the feed water was fofind from Equation

6.14. Py is the pump inlet pressure of 1.5 inches Hg or .736 psi. P, is

the pump discharge pressure and is the sum of the three items listed below:-
No. 3 heater pressure 103.1 psia

Elevation difference (cond. to’ 17.3
 No. 3 Heater)

Friction drop 40,0

- : P, = 160.4 psia

The specific volume, Vi, is .0161 ft3.
1v
Work = (160.4 - .736) x 1hk x .0161
/ 778

Work - 476 BTU/1b

The work put into the pump was found from Equation 6.15.

Work = 476 - .676 BTU/1b
o7

The heat gain of the feed water from the pump was found to be .2 BTU/1b
by Equation 6.12.
The reduction in steam extracted for No. 6 Heater was found from
- -Equation 6.16.

_ W = 1.6 x 10 x .2 |
- (102h - 122.9) | | ’

W = .000355 x 10° 1b

hr

The increase in work at the turbine vheel was found from Equation 6.10.

Work = .000355 x 106 (1024 - 9hioh)

Work = .0293 x 106 BTU
hr

b. Correction for Boiler Feed Pump

This correction was made in the same manner as for the hot well

pump above.

Work into feed water = (2000 - 103.1) x 14k x .01766
178

Work into feed water = 6.20 BTU/1b
Work into pump = 8.86 BTU/1b
Heat gain of feed water - 2.66 BTU/1b

The reduction in extracted steam is,

W = 1.955 x 106.x 2.66
(1286 - 36L.2)
W = .00564 x 106 1b/hr

The increase in turbine work is,

Work = .0056L4 x 106 x {1286 - 9ki.h)

- 1.9 x 10° BTU
hr

Work

The turbine work calculated in Section G.3 was corrected as shown

below.

Uncorrected turbine work (Section G.3) 881.Y x 100 BTU/hr .

Hot well pump correction + .03
Boiler feed pump correction + 1.9

Corrected turbine work = 883.83 x 106 BTU/hr
This is equivalent to an ocutput of 258.96 Mw. |

G.5 CORRECTION FOR TURBINE -GENERATOR LOSSES

The losses found by the method of Reference 4l were applied to the

corrected turbine wheel output of Section G.4 as shown below.

P

Corrected output at turbine wheel . 258.96 Mw
Turbine exhaust loss , - 8.35 Mw :
Mechanical and generator losses Del »

: : Total losses 13.5%

Gross Electric Power Output 2h§°h2 Mw

231,

G.6 CALCULATION OF COOLING WATER PUMP POWER .

The cooling watér pump power was found from Equations 6.14 and 6.15.
A pressure drop of 10 feet was assumed in the condenser (Ref°’23.)° The
cooling water flow was calculated from a heat balénce between the heat
rejected to the condenser by the turbine exhaust and No. 6 Heater drain
and the heat gained by the cooling water. In calculating the required
cooling water.flow, a8 15°F rise in temperature was assumed. This could
be greater, or less, depending upon the tefiperature and quantity of cool-
ing water available and the type of condenser selected, i.e., single or
double pass. |

a. Heat rejected to condenser by iufbine éxhauat
q = WAHR

g = 1.3126 x 106 1b/hr (941.% - 59.7) BTU/1b

un

q = 1157 x 108 BTU/nr

131

b. Heat rejected to condenser by No. 6 Heater drain

.2914 x 106 1b (122.9 - 59.7) BTU/1b
hr

q

q = 18.4 x 106 BTU /hr

Co Totalzheat rejected to condenser. The total réjected heat is the

6

sum of Items (a) and (b) above, or 1175.4 x 10° BTU/hr.

d. Cooling water flow to condenser

W o= 1175.4 x 10® = 78.% x 10°
lx15 hr

e. Work done by cooling water pump

Work

(P -P,) VW

Work = 10 £t x 62.4 1b x .0161 £t3 x _ BTU _ x 78.4 x 108 1b
£t3 1 778 £t/1b hr

232.

Work = 1.002 x.10° BTU - .294 Mw
hr

f. Work put into cooling water pump. From Equation 6.15,

Work in = o29h~ Mw - 011-90 Mw

Q

G.7 POWER REQUIRED BY PUMP MOTORS (steam cycle)

The motor power was found by dividing the required pump power by the
motor efficiency.
a. Cooling water pump motor

pover = 490 Mw = .515 Mw
95
b. Hot well (condensate) pump motor

BTU 6 1b = Mw-hr 6
pover = 676 T 1.6 x 10 or X 3013 X 10° BTU

=95

334 Mw

pover

¢. Boiler feed water pump motor

power = 8.86 %%H x 1.955 x 106 %% X %!fi%g x 106 BTU
.95

power = 5.34 Mw
The total motor power required by the above pumps in the steam cycle

is 6.19 Mw,

I

ORNL=LE=Dwg,=18198

1000

IRV ENE N

| 1 10

Absolute Pressure- Psis

T e T T L e g

I O 5 0P 1 O 0 A 0 O O )19 O (O U 0 W Y 0 L

of

| 4 10 0
T

Pressure - Enthalpy Relation

Turbine Expansion Line

1 O O O O S U Y O O O O O

INEGFEN IEESNREEES EEE AR W NN AN NN

. Figure G,1!

1
T
]

*qT/n3g -Ldreyjuyg

23k,

APPENDIX H

SODIUM PIPING PRESSURE DROP CALCULATIONS

Sample calculations: (Ref. 7T)

H.1 PIPE LINE PRESSURE LOSS

Primary loop: 8 in Schedule 40, 304 SS, 15 ft long
T = 1150°F

€ - 49.9 1b/ft3

/A= .00013 1b/ft-sec

E/D = .00006

D = 0.665 ft

V = 30.5 ft/sec

Re

pv€ - .665 x 30.5 x 49.9 = 7.78 x 106
A -00013

L = sOkk

AP = £ L V2 € psi = .011 x 15 x (30.5)2 x 49.9 = 1.25 psi
D 2g 1Lk .665 ‘ 154

H.2 EXPANSION LOSS

Hot leg to heat exchanger. Assume one velocity heat loss for
expansion.

AP = K vl2 x ¥ psi =1.0 x (31+,h)2 x 49.9 = 6.4 psi
2g 1Lk oL. L 1LL

H.3 CONTRACTION LOSS

Heat exchanger to cold leg. Assume 0.5 velocity head loss for
contraction. 16 in Schedule 40, 304 SS pipe.

T = 1000°F

€

51.2 1b/ft3

1.227 £t

235,

V = 145,700,000 1b x 1 br x 1 £t3 x 1 = 33.5 ft/sec
- hr 3600 sec 51.216 1.227 x 6 ft°© '
OP = K V12 @ = 0.5 x (33.5)2 x 51.2 = 3.1 psi

. 2g 1Lh oh. I 1L

H.4 90° BEND LOSS

@ Assume 0.1 velocity head loss per 90° bend (Ref. 16).

AP =K V2 € psi = 0.1 x (33.5)2 x 51.2 = .6 psi
2g 1LL ' 6L, L N

100°F

51.2 1b/ft3

33.5 ft/sec

T
e
V'
K

- Ool

H=5  VALVE LOSS
o Assume 0.6 velocity loss per swing check valvé (Ref. 16).
T = 675°F
53.96 1b/£t3

= 26.2 ft/sec

€

v

AP =KV2 € psi £ 0.6 x (26.2)2 x 53.96 = 2.4 psi
2g 1L AT T1hL

236

TABLE H.1l
Sodium piping pata ) o
VALVES 90~ bends
Description No.of lines Length Pipe Schedule Design Number Type Pressure Number Pressure Line Pressure Total Pressure
ft. Size Number Velocity Drop Drop Drop Drop
ft/sec pai psi psi psi
A, PRIMARY LOOCP
1. Primary heat exchanger, 20 | 5/8 in| .065 1n] 12.8 15,7
tube. side wall
2, Feed line (hot)* 24 15 8 ir | 40 30,5 1.5 1,3 2.8
3. Ringheader (hot) 1 75 18 ir 30 26.4 - 8.3
ave . P
swing N
4e leader (hot) * 6 30 116 im | 40 344 check 3.8 3,3 1.5 8.6
5. Intermediate heat 14.8 12,2 12,2
exchanger, shell side
‘ swing _
6, leader (cold) * 6 50 16 ia | 40 35.5 check 3.7 4e3 2.5 10.5
7, Ringheader (cold) 1 75 18 123 30 25,7 8,7 8.7
8, Feed line (cold) 24 9 8 in _ | 40 29,6 1,4 0,7 2.1
7.5 10,5 26.9 68,9
‘ 9, Pump Pescription: 18,55D gpm., 200 [ft.head,| 176C rpg,, 1000 F, 1335 HP
B,  INTERMEDIATE LOOP
l, Intermediate heat
exchanger, tube side 1.8 1,2 i 065 in 15.0
wall
2, Hot leg 6 50 10 in, | 40 27,1 1,6 2.6 4,2
] 3, Secondary leader (hot) 2 65 18 in | 30 27.7 gate 0.4 1,2 2.8 A
J 4, Boiler, shell side 50 . 9.6 11.6
5, Secondary leader (cold) 2 80 18 in | 30 26,2 gate 1,2 2.8 2.9 9.3
swing '
check 2.4
6, Cold leg 6 40 1C in 1 40 25,5 1.5 2,0 3.5
4.0 7.1 10.3 48.0
7. Pump description: 18,350 gpm,, 140 Ft.head,|17€0 rpn., 675 F| 950 HP

-,

]

APPENDIX I

| ' PARTTAL LOAD OPERATION (MATHEMATICAL APPROACH)

The assumptions used in deriving the équations of this appendix are
explicitly given in Section T.4.1 of this report. In addition to the nomen-
clature given at the beginning of these appendices, other symbols are used

for the derivations of this section. These arc as follows:

- Subscripts - Genera}
1 Full load value of a variable
X Partial load value of a variable
p - Value of a variable
b Value of & variable in the boiler
5 Value of a variable in the superheater
f Value of a variable in the fuel loop

Subscripts - Temperatures

0 Cold mixed mean temperature of fuel leaving the primary
heat exchanger

2 Hot mixed mean temperature of fuel éntering the primary
heat exchanger

- 3. Inlet sodium temperature to primary heat exchanger
L Outlet sodium temperature from primary heat exchanger
) 5 Inlet sodium temperature to intermediate heat exchanger
6 Outlet sodium temperature from intermediate heat exchanger
s T Outlet sodium temperature from superheater
L Load fraction ratio actual load

full load

W Pounde of steam generated per hour at full load

M Weight flow of sodium in the intermediate loop, 1v/hr

zZ Logarithmic mean temperature difference

I.1l NATURAL CIRCULATION BOILER WITH SEPARATE SUPERHEATER

Since the average fuel temperature and flow rate are comstant,
Aty = 150 L | - (I.1)
and
tp average - 1125°F {1.2)
Solving Equations (I.1l) and (I.2) yields the following relations,

1125 - 75 L | . (1.3)

Q
"
"

to, = 1125+ 75 L (I,h)
As a consequence of the assumption of constant over-all heal transfer.
coefficient and constant flow rate of sodium in the primary loop, Equations

(1.5), (1.6), (I.7) and (I,B)'result,

Zpx = 50 L - (1.5)
tyyx - t3x .= 150 L | ‘ (1.6)
tox = t3x = 50 L (I.7)
toy -~ tyx = 50 L (r.8)

The simultaneous solution of Equations (I.6), (I.7) and (I.8) gives the
equations for tj, and t3x’

t3x = 1125 - 125 L (1.9)

1125 + 25 L (1.10)

Thx

The sodium in the intermediate loop transfers its heat to the supexr-
heater and the boiller, respectiveiy° At partial load, the over-all transfer
rate in the‘superheater will be,

1 = .000710 + 1 . | | (T.11)
Uox 775 (L)"8

where the constants in Equation (I.11l) are based on a design value of U,

1

of 500.

20

RC |

1000 SN

§t7

. 63508
Therefore, (1.12)
st = L x 500 X Zsl .000710 + . 1 . , a(téx - 1000) - (b'-(x - 635»8)
775 (L)-° 1n [ t6x - 1000
trx - 635.8

From the design conditions at full load,

The assumptions made for partial load operation include the constancy
of the over-all coefficient of heat transfer in the boliler. Based on this

and the fact that the surface is constant, Equation (I.13) follows directly.

a
A 635.8
135 L = by - toy - (1.13)
th - 635.8

The last equation needed to specify the conditions in the loops at -

partial load is obtained by considering the intermediate heat exchanger.

147 L =(t1+x - t6x) - @x - tey) | (14&)
in [tux - b6x
G3x - tSX)

Substituting the values for t3; and tj,, from Equations (I.9) and (I.10)

into (I.14) and simplifying, ylelds Equation (I.15),

WTL =150 L - (g - toy) (1.15)
In (125 ¥ 258 - 6z
1125 - 125L - te

The weight flow of the sodium in the intermediate loop is determined

directly by a heat balance,

L = M, 6‘6)( - toy) - - (I.16)
M) 425

Equations (I.12), (I.13), (I.15) and (I.16) completely determine the
steady state fiartial loadfioperation of the system based upon the simplifica-
tion and assumptions given in Section 7.4.l. However, a superheater heat
balance has nof been utilized in the foregoing. The superheater heat Salance
gives an,additionél:independent equation (I.,22)° The ability to write more
independent equations than there are unknowns indicates that too many con-
straints have been placed on the system by the assumptions made in Section
7.4,1. Therefore, the final steam temperature will not be taken to be con-
stant with load and the control of the final steam temperature will be
accomplished by using an attemperator. Equation (I.12) has to be adjusted
to reflect the change in this assumption, and instead of 1000°F, the'final
steam temperature should be indicated to be variable. In this case, Equation
(I.22) has to be written in terms §f the enthalpy of the steam side corre-
sponding to the unknown final steam temperature at the superheater ofitlet.

It is further noted thaf for a variable final steam temperature leaving

the superheater, Equations (I.11) and (I.12) are not correct as written. A

v

N

241,

'pumerical solution is necessary where the final steam temperature is assumed
for a given load. From a heat balance of the steam loop, thé weight of steam
is obtained and on is calculated. The procedure is then to solve the appli-
cable equations simultanecusly, and finally to check the assumed final steam
temperature for the load under consideration. A procedure such as outlined
above is tedious and time consuming.

1.2 ONCE-THROUGH STEAM GENERATOR

The basic difference in the mathematical approach to the problem of
partial load operation of the once-through boiler as compared to the natural
circulation boiler with separate superheater is that the sum of the surfaces

of the boiler and superheater sections is a constant, but each can vary indivi-

.Gually with load. Equations (I.3), (I.4)}, (I.15) and (I.16) are applicable

to this case, However, certain of the other relation§ were modified, and,
where necessary, new equations were developedol In general, the loéd fraction
(L) can be written for a given component as,
L = Uy Z, Sy | | (1.17)
Uy 21 5,

For the boiler and preheater section, Equation (I.18) is obtained by combining

Equations (I.17) and (I.13),

Spx = L x 135 x SBpy _ln*(t"{x - 635°9 ' (I,iS)
\th - 63508 '

(‘b"{x - th) ’

The same procedure follows for the superheat section, where Equation (I.17)

is combined with Equation (I.12). However, the design values for the once-

through boiler are different than for the separate superheater and the con-

stants in Equation (I.12) were adjusted accordingly. The design value for

242,

the steam side coefficient was 1100 and the over-all coefficient was

300 BTU/hr-£t°-OF. - |  (1.19)

1100 (L)

LxU; x191 [.002k2 + 1 %fisl = (t6x - 2000) - (b7x - 635.8) Sex.

1n tgx - 1000
‘t7x - 635.8

Solving for S, (1.20)

s L x 300 3 242 1 t6x - 1000
gx = x 3 x 191 [00 _ + 1100 (L).B (Ssl) in t7x _ 635.8
(t6x - 1000) - (7x - 635.8)

The sum of the surfaces at any load must be a constant. This sum at design

load is 21,192 square feet vhere,

Seqy = 11,470 sq £t
and
sbl' = 9722 8q ft
Therefore,
. 1 tgyx - 1000
L x 300 x 191 (.oozhz + = 8) 11,470 1n )
1100 (L)* t7y - 635.8
(t6x - 1000 ) - 6;7,( - 635.8) |
4+ L x 135 x 9722 ln(th - 635-8) = 21,192 (1.21)
. 'bsx - 635.8 - .

A heat balance of the superheater ylelds Equation (I.22),

~

My = 1ko L (I.22)
fiI t6x - t?;)'

Equations (I.22) and (I.16) are then combined to yleld a single equation

with the flow rates eliminated,

66;( - tox)e %2% (tsx - tr(; | | (1.23)

.1“":‘

B

2k3.
The solution of Equations (I.15), (I.21) and (I.23) simultaneously
gives the desired steady state partiaL load operation of the system for the
once-through boiler. . It is noted that the final steam temperaffire can be
set at 1000°F and the system is cbmpletelyldete;minate'fiith only inter-

mediate loop sodium flow rate control; this is not the case for the natural

circulation boiler with separéte superheater.

ok,

APPENDIX J

SIMULATION FOR 600 MW FUSED SALT REACTOR AND STATIONARY POWER PLANT

by
E. R. Mann

Instrumentation and Controls Division
Oak Ridge National Laboratory

Figures (7.%), (J.1), (J.2) and (J.3) are the elementary flow sheet
for the power plant and the road maps required for simulation, with the
excéption of that for the reactor kinetics which is described in Ref. 5k.

On Figure (J.1l) is the road map for the fused salt loop and the first
or primary sodium loop. Amplifiers 1, 2, 3 and 4 provide a unity gain cir-
cuit for the fused salt loop. Amplifiers 7, 8, 9 and 10 provide a similar
clrcuit for the primary sodium coolant. Amplifiers 5 and 6 generate the
coupling between these two flulds as the difference between the mid tem-
peratures of the two fluids in the heat exchanger.

The time constant of the metal in the heat exchanger was not included
in these simulation studies. Accordingly, about amplifier 5 are shown a
potentiometer setting "o£" and series resistor "R" to indicate that such
a time constant should be included. |

. This time constant can be calculated rather simply as follows. The
heat capacity of the fused salt in the heat exchangér is given by the
expression,

Ce = P

ST/St
where P 1s the power given up by the fused salt in passing through the
heat exchanger at design point steady state, and ST is the temperature

drop of the salt during the time St 1s in the heat. exchanger. P 1s in

5
2
’\

245,
BTU/sec, OT in degrees F, and St in seconds, and Cs is in BTU/F.

Now the mass of metal in the heat exchanger times its specific heat
gives the heat capacity Cyg g, of the heat exchanger. If then, [g¢ 1is the
time constant of the fluid in the heat exchanger, which here has beeéen
approximated by two first 6rder lags whose sum is the fluid tfansit time,

then the time constant of the heat exchanger is,

In transient, this circuit with the time constant 7:H,E.’ provided by
amplifier 5, gives the heat transfer delay between fluids due to heating |
or cooling the metal of the heat exchanger. |

The road map on Figure (J.2) shows, through amplifiers 13 and 14, the
intermediate sodium in the sodium-to-sodium heat exchanger. Amplifiers 15,
16, 19 and 20 show the sodium "going and returning” in a heat exchanger
between the hot and cold legs of the sodium piping. The flow sheet, Fig-
ure (7.4),shows how valves V1 and Vg will provide means tor returning part
of the cold sodium through this auxiliary heat exchanger which drops the
temperature of the sodium entering the boiler.

This auxiliary heat exchanger is one fieans available to liwmit thg
steam temperature on partial'loads to a maxifium temperature of 1000°F. It
was selected for simulation here primarily because the analog facility re-
quired less equipment io do thisrthan it would have required to vafy the
sec&ndary sodium flow rate. |

The road map, Figure (J.2), shows amplifiers 17 and 18 as the sodium
in the sodium-to-wa£ép andyvapor heat exchanger. It was assumed that the

load "L" was proportional to water flow, i.e,, the water entered the boiler

at constant temperature for all loads.

246,

Road wmaps, Figure (J.3), show by amplifiers‘ll and 12 how the driving
function "4 " between the two sodium systems in the sodium-fo-sodium heat
exchanger was generated, and by amplifiers 21 through 26, inclusive, how
the regulator action tb keep the steam fiefiperature constant with varying
loads "i" was obtained.

The quentities ".&" and "R", with amplifier 11, are determined by the
time constant of the metal in the heat eichanger for the sodium-to-water
heat transfer. The heat capacity of this heat exchanger was not used ifi
these simulations. -This time constant could have beén determined by the
same procedure proposed for that of the metal in the sodium-to-sodium
heat exchanger.

The output of amplifier 27, Figure (J.3), at design point should be
+.2,5 V. This output was recorded on a Brown recorder with a full scale
reading of 5 V. The output gafie a midscale reading. Two limit switches
about the midscale reading actuated a mofior‘to-turn.a ten-turn potentiometer
"E", Figure (J.3). The motor was reversible and: its-direction was determined
by the limit switch contacted b&‘deviation of the output of amplif;er‘ET;,.If
the deviation was positive and caused the recorder pen to move upscale into
the upfier limit switch, which would be true for decreasing load, then the
motor turned the potentiometer setting up and thereby "valved" more sodium
through the‘auxiliary heat éxcha.nger° Except for the relatively slow motor
speed, or ?ate of motion of fhé glide wire on the potentiometer due to the
géar reduction avallable, this device could hold the steam temperatfife con=-
stant for large variations in power. The curves show deviations in steam
temperatures durifig transients; but these‘comé from tfie relatively slow

response times of the actuator mecha.nism°

247,
Scale factors in the simulation studies were as follows:
1. Computer time was real time .
-2, One volt represented 10 megawatts of power
3. One volt represented 20°F
The circuits were not set up from a set of differential equations
representing the system. The set of equations can be derived from the
circuits by setting the currents at the amplifier inputs equal to zero.
Then by converting electrical units to power and temperature units using

the scale factors given above, one gets the equations of the system. Some

of these equations will be dependent.

+A,+2.5V

+T,456.25V

+A,+25Vv

+2.5V

66181~ "8m-dT1-"INHO

“gh2

¢

+ T,
+44.37V

+8,+9.375V

M .
M IM
I G-
+Tg & b
' I'M
-C R
‘FIG.—J.Z- SIMULATOR CIRCUIT FOR INTERMEDIATE SODIUM LODP

00281- "Emg-¥T1-"INYO

* 642

+B
+9.375V ‘ ‘ ‘ -9.375V

FIG. J.3-SIMULATOR CIRCUIT FOR DRIVING FUNCTION IN INT.ERMEDIATE HEAT

EXCHANGER AND REGULATOR FOR STEAM TEMPERATURE CONTROL

10281~ " SmI-d1-TINYHO

*062

J,\

251.

APPENDIX K

TEST OF DOUBLE TUBE SHEET DESIGN FOR WATER SODIUM ISOLATION

An experiment has been conducted to determine the feasibility of
using brazed joints in the fabrication of a double header for the heat
exchanger(s).

Procedure

Approximately 400 holes were drilled in a 1/2-inch thick type 347 SS
plate to receive short lengths of type 304 stainless steel tubes of nominal
1/2-inch diameter, .035-inch wall. The holes were reamed to allow .005-inch .
clearance on the diameter to permit easy assembly and satisfactory design
for brazing. Although none of the currently available elevated temperature
alloys have been found to be entirely satisfactory in elevated tempergture
sodium corrosion tests}/, the alloy Cast Metals No. 52 (84 Ni-4 Si-2B) was
selected as representative of the alloy system most iikely to yleld a éorro-
sion resistant alloy for extended service. This alloy system, however,
possesses a fundamental.disadv&ntage, in that Boron diffusion into the stain-
less steel grain boundaries results in a serious loss of dfictility. The
experiment was conducted; therefore; to merely verify the feasibility follow-
ing development of a nickel-base corrosion resistant alloy. The alléy V&8
applied as & precast brazing ring on one face of the tfibe shee? afound eachn
tube and fixed in place with a CM 52 pounder slurry. The tuhe;éheet was
heated in -lOOoF'dew point hydrogen to 1050°C at a rate of lOSOOC/hour, held
1/2 hour at 1050°C and cooled at 100°C/hour. This cycle was selected as

representatifie of commercial practice.

1/ CF-56-4-130, "Sodium Corrosion and Oxidation Resistance of High
Temperature Brazing Alloys", G. M. Slaughter et al.

252.
Results
The results of thie experiment are illustrated in Figures (K.l) and (K.2),

photographs of the underside of the tube sheet after brazing. Four tubes were

found to be only partially brazed, the most serious‘flaw being shown enlarged
in Figure (K.2).

In view of these results and in view of the research still required to
develop a brazing alloy suitable for extended service in sodium, the method of

fabrication suggested by this study is at best marginal and not recommended.

——
\

253

Underside of Brazed Tube Sheet
(Arrow points to brazing flow)

-1

Figure K

254

ONE INC

Figure K-2 Defective Brazing of Tube
(Close-up view of tube indicated
by arrow in K-1)

A

f(N)

255.

SYMBOLS USED IN THE ENGINEERING CALCULATIONS

Cross sectional area for flow of fluid ft2

Ai inside or tube side

A0 outside or shell side
Constant
Specific heat at constant pressure, BTU/1b-F

Tube diameter, ft

di inslide diameter

d outside diameter ‘
4~ diemeter of interface of cladding and base metal
d

equivalent diameter; 4 x flow area/wetted perimeter
Function of N |
Friction factor, dimensionless
Acceleratién due to gravity, 32.2 ft/sec?l

Heat transfer coefficient, BTU/hr-ft°-OF
hNa for sodium

hS for salt

Enthalpy, BTU/1b

Thermal conductivity, BTU/hr-ft>-OF
Geometrical length, ft

Number of tubes, dimensionless

Nuééeit number, dimensionless

Pressure, psi

Power delivered to the fluid, horsepower
Peclet number, Pe - Re Pr, dimensionless

Tube pitch, ft

Prandtl number, dimensionless

Rate of heat transfer, BTU/hr

256.
2 ,
R Resistance to heat flow, hr-ft~-°F/BTU
Re Reynolds number, dimensionless
r Tube radius, ft .
S Area of heat transfer surface, ft2
S{ referred to inside surface of tubes
So referred to ocutside surface of tubes K
s Ligament between adjacent tubes; ft
t Temperature, °F
T Absolute temperature, %k o
Uo Over-all heat transfer coefficient based on outside
surface of tubes, BTU/br-ft2-OF -
v Velocity of fluid, ft/sec
v Specific volume, ft5/1b |
Vol Volume of holdup, ft3
\Y Mass flow rate, 1b/hr -
x Thickness, ft
Greek
AP Pressure difference, psi
At Temperature différence, %
Aty Temperature difference between bulk mean
temperatures of two fluids ,
Atyy Logarithmic mean temperature difference -
72 Pump efficiency
/7 Avsolute viscosity of ‘fluld, 1b/hr-ft "
¢ Density of fluid, 1b/ft>
i; Notation signifying a summation, dimensionless ny, |
77 3.1416
Subscripts - . S ¢
Na Refers to sodium
S Refers to the salt

257.

SYMBOLS USED IN THE NUCLEAR PHYSICS CALCULATIONS

Atomic weight of an element .
Geometric buckling of the reactor, cm~2
Constant

Diffusion coefficient, cm

Multiplication constant of the reactor

" Atomic concentration, atoms/cc

No Avogadros number, 6.02 x 1023 atoms/gm atomic wt

Poison fraction,CZ{p 2{;, dimensionless

Radius of reactor's central core, cm or ft
Time, seconds

Temperature; oC‘or éF

Lethargy, logarithmic energy decrement

Time, years

Greek

§
7
¢
o
z

Average change in lethargy of a neutron per collision
Average humber of neutrons released per.figs;on
Nefitron flux, neutrons/cm®-gec¢

Density, gm/cc

Microscopic cross section, ém? or barhs

Macroscopic cross section, em~t

258.
BIBLIOGRAPHY
Ref. No.
1. Shannon, R. H., APDA, Private Communication, June 20, 1956
2. Ledinegg, M., "Flow Instability in Natural and Forced
Circulation"; "Die Warme", 1938
3. Leib, E. F., Discussion to paper by Van Brunt, Trans. ASME,
May 1941
b, Jens, W. H., "What is Known about Boiling Heat Transfer",
Metropolitan Section of ASME, April 7T, 1954
5. Morabito, J. J. and Shannon, R. H., "Test of a Once-Through
Steam Generator with a Liquid Metal as a Heat Source", ASME
Preprint Paper No. 55-A-189, 1956
6. McAdams, W. H., "Heat Transmission", 3rd Edition,
McGraw-Hill Book Co., 1954
T. U. S. Atomic Energy Commission,; "Reactor Handbook",
Engineering, Vol. 2
8. Glasstone, S., "Principles of Nuclear Reactor Engineering",
1st Edition, D. Van Nostrand Co., 1955
9. Colburn, A. P., "A Method of Correlating Forced Convection
Heat Transfer Data and a Comparison with Fluid Friction",
Trans. A.I.Ch.E., Vol. 29, 1933
10. Fong, J.; Research Dept.; Foster Wheeler Corpo, Private
Communication, June 1956
11. Granet, I. and Kass, P., "The Viscosity, Thermal Conductivity
and Specific Heat of Steam at Elevated Pressures and Temperatures"”,
Petroleum Refiner, May 1953 -
12, "Fourth International Conference on the Properties of Steam", 1954
13, McAdams, N. H., Kennel, W. E. and Addoms, J. N., "Heat Transfer
to Superheated Steam at High Pressures", Trans. ASME, Vol. 72, 1950
1k, Tate, G. E., "Notes - Advanced Heat Transfer", Polytechnic
Institute of Brooklyn, 1948
15. Jackson, C. B., "Liquid-Metals Handbook Sodium-Nak Supplement",
3rd Edition, A.E.C. and Dept. of the Navy, July 1, 1955
16. "Pipe Friction", Standard of Hydraulics_Inspitute, 1949

L

e

o

S

¢

17.
18.
19.
20.
21.

22,

230

2k,

25,

26,

27,
28.

29.

30.

31,

32,

33.

3k,

¢£22;
Cohen, Stanley, REED, ORNL, Private Communication, June 1956
Barton, C. J., ORNL, Private Communication, June 1956
Nessle, George J., Jr., ORNL, Private Communication, Jupe 1956
Kern, D. Q., “Progeés Heat Transfer”, McGraw-Hill Book Co., 1950
Binder, R. C., ”Flfiid.Meéhanies", Prentice-Hall, Inc., 1949

Granet, I., “The Computation of Nétural Circulation in Large
Boilers", M.S. Thesis, Polytechnic Institute of Brooklyn, ‘1948

' gaffert, Gustaf A., “Steam Power Stations", McGraw-Hill Book

Co., 1952 ~

Granet, I., "The Calculation of the Relative Expansion in
Fixed Tube Sheet Heat Exchangers", Petroleum Refiner, August 1955

Dwyef, 0. E., "Heat Exchange in a Liquid-Metal-Fuel Reactor
for Power", Nuclear Engineering - Part I, AIChE, 1954

. Sense, Karl A., et al, "Vapor Pressures of the Sodium Fluoride-

Zirconium Fluoride System and Derived Information", Baftelle
Memorial Institute, BMI 106k, January 9, 1956

Kinney, W., ORNL, Private Communication, June 1956

Bate, R. R., Einstein, L. T. and Kinney, W. E., "Description
and Operation Manual for the Three Group, Three Region Reactor
Code for ORACLE", ORNL, ORNL-CF-55-1-76, January 13, 1955

Ergen, W. K., "Uranium Investment in a Circulating Fluoride
Power-Producing Reactor", ORNL, ORNL -CF -56-4 <29, Classified

April %, 1956

Hughes, D, J. and Harvey, J. A., "Neutron Cross Sections",
Brookhaven National Laboratory, BNL-325, July 1, 1955

Barton, C. J., "Fused Salt Compositions", ORNL, CRNL-CF-55-9-78,

Classified September 16, 1955

Grimes, W. R., ORNL, Private Communication, July 1956

Gresky, A. T. and Arnold, E. D,, "Poisoning of the Core of the
Two Region Homogeneous Thermal Breeder: Study No. 1", ORNL,
ORNL-CF-53-12-165, Classified December 11, 1953

Gresky, A. T. and Arnold E, D., "Poisoning of the Core of the
Two Regior Homogeneous Thermal Breeder: Study No. 2", ORNL,
ORNL-CF-54~2-208, Classified February 22, 1954

260,

35. "Unfired Pressure Vessels" Section VIII of the ASME Boiler

and Pressure Vessel Code; 1952
" 36. "PWR Preliminary Design Report", Atomic’ Power Division,

Westinghouse Electric Corp., WAPD-112, Classified,
June 1954

37. Lane, James A., "Economics of Nuclear Power", Geneva
Conference Paper No. 590A (UN-476), 1955

38, Cathers, G. I., ORNLp Private Communication, August 1, 1956

39. U. S. Atomic Energy Commission, "Reactor Handbook", Engineering,
Vol. 2, Classified

LO. Charpie; R, A. and Silver, Ernest, ORNL, Private Communication,
August 1956

b1, Hegetschweiler, H. and Bartlett, R. L., "Predicting Performance
of Large Steam Turbine-Generator Units for Central Statioms",
ASME, Paper No. 56-SA-52, 1956

L2, Poppendiek, H. F. and Palmer; L. D.; "Forced Convection Heat
Transfer in Pipes with Volume Heat Sources within the Fluids",
ORNL, ORNL-1395, November 11, 1953 .

L3, Poppendiek, H. F. and Palmer, L. D.; "Forced Convection Heat
Transfer Between Parallel Plates and in Annuli with Volume
Heat Sources within the Fluids", ORNL, ORNL-1701, May 21, 195k

Lk, Poppendiek, H., F. and Palmer, L. D., "Application of Temperature
Solutions for Forced Convection Systems with Volume Heat Sources
to General Convection Problems", ORNL, ORNL-1933, December 23, 1955

45, Poppendiek, H. F., ORNL, Private Communication, July 1956

5k Mann, E. R.; "Electronic Analog Devices for Design of Reactor
Controls", IRE Transactions of the Professional Group on
Nuclear Science, Vol. NS-3, No. 2, March 1956

A 55 . Thomas, R. Davis, Jr., "Selection of Stainless Steel
Electrodes for Trouble-Free Welds", Metal _Progress,
Vol. 70, No. 1, July 1956 -

56. Woolsey, C. C., Jr., "Selection of Materials for Sodium
Graphite Reactor System", ASME Paper 55-S-16

57. Dreisbach; R. A., and Smith; A. G., "Gas-Fired Life Test",
Experiment No. 4935-6, ORNL, July 25, 1956

e V]

Energy Act of 1954, al or the dis

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