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RECEIVED BY Tic  JUL 24 1972

1000 MW(e) MOLTEN SALT BREEDER REACTOR

CONCEPTUAL DESIGN STUDY
FINAL REPORT — TASK |

SUBCONTRACT NO. 3560
| WITH
~ UNION CARBIDE CORPORATION
NUCLEAR DIVISION
OAK RIDGE NATIONAL LABORATORY
'~ OAK RIDGE, TENNESSEE

UNDER

U.S. ATOMIC ENERGY COMMISSION
CONTRACT NO. W-7405-eng-26

 

.~

 

e = N O TIC E . -
1 “teport was prepared a3 sn account of wor
.Iptg:tsor:’d by - the ?Jtnitad, States Government, Neither
} the United States nor the United States Atomic Energy
Commission, nor sny of their ‘employees, nor any of
{ their contractors, subcontractors, or their employees,
-1 makes any warranty, express or implied, or assumes any
{ legal llabllity or responsibllity for the accuracy, com-
| pleteness or usefulness of ‘any information, spparatus,
| product or process disclosed, or represents that its use
would not infeinge privately owned rights.

 

 

 

MASTER

DISTRIBUTION OF THIS GOCUMENT IS URLIMITED

 

 

 
 

1000 MW(e) MOLTEN-SALT

BREEDER REACTOR

CONCEPTUAL DESIGN STUDY

Final Report - Task ~ I

Project Manager
T, A. Flynn, Jr. ¢

o £2.)

EBASCO SERVICES, INC,
NEW YORK, NEW YORK

February 1972

  

- Technica Direc;b;A
~ D. R, deBoisblanc

 

T T R b e -t

 
1.0

2.0

3.0

4.0

CONTENTS
Page
INTRODUCTION .. ........... e 11
1.1 Origin of Industrial Interestin MSBRs . . . . ... ........ 1-1
1.2 Project Organization . ......... e et e e e e 1-2
PURPOSE AND SCOPE OF TASKI . . . . . . 21
2.1 Purpose ... ........... e e e e e e e e e e e e e 2-1
2.2 Scope . ... e, e e e et e e e 2-1
REACTOR SYSTEM « . « « « v v v o v .. T 31
3.1  Graphite Moderator Element Design . . . .. .. e e e e 3-3
3.1.1 Mechanical Design . . ... ....... e e e 34
3.1.2 ArraysofCylinders . ........... e e 3-5
31.3 SolidBlocks ........ e e e e e e e e 3.6
3.1.4 Iris Array of Curved Slabs .......... P 3-6
315 SlabArrays . . . .. .o o s e 3.7
3.1.6 Graphite Utilization . ........ e e e e e e 3.9
3.1.7 Future Thermal-Hydraulic Conmdcratlons .......... 3.10
3.2 ReactorVessel . . . . . . . . i i i i it it e e e e 3-11
3.3 Reactor Drain System .. ... .... e e e e e e e e e s %
3.3.1 Drain Tank and Reactor Lining . . . . .. e e e e e 3-13
3.3.2 Piping and Component Arrangement . . . . . . . ... ... 3-15
SECONDARY SYSTEM . . . . it i v vt s v v v e on e ane s 41
4.1 General . . ... e E e e e e e e e e .. 4-1
4.2 Steam Generator - Reheater Concept_ Selection .......... . 4.2
4.2.1 Background e e Ce e 4-2
4.2.2 BascheslgnCntena S I,
4.2.3 SelectionCriteria ... . .. ... ..o . 44
4.2.4 Concept Evaluation . . .. et e e 45
4.2.5 Detailed Evaluatlon of Selected Concepts i e e e 4T
4.2.6 Reheater . . ... ... e e e e s e e e 410
4.2.7 Conclusions and Recommendatzons R R §

 

 

 
 

4.3 Primary Heat Exchanger Concept Selection . . ...........
431 Background ... ............... e
4.3.2 SelectionCriteria . . . . . . . v it
4.3.3 Concept Evaluation . .......... P
4.3.4 Conclusions and Recommendations . . . ... .......
4.4  Other Components . . . . . . v v v v v v v v v o v et e e e
441 CoolantSaltPump . ............... e e
442 RuptureDisk . . ................. e
44.3 CoolantStorage Tank . . . . . .. .o v v e ..
444 CoolantSaltFilter . . .. .. ... ... .. ... .....
445 DeMister . . . ...t v it e e e e e
4.4.6 CoolantSaltMeltTank .. ............ e e
. 5.0 STEAMPOWER SYSTEM . . . . . .« o i i it e e e e e e e e e e
o 5.1 General . . . . . . e e e e e e e e e e
s 5.2 Special Design Considerations . . . . . . v v v v v v v v v oo ...
5.3  Selection of Steam Cycle Conditions . . . . ... ..........
J 5.4  Design Criteria and Assumptions . . . . ... ... .. e
- 5.5  Description of MSBR Steam Power System . . . . . . e e e
- 5.5.1 Normal Operation . ............... e
™ 5.5.2 Standby Operation . . . .. ... ... ....... ..
5.5.3 Start-Up and Shutdown Operations . . ... .. e e
5.5.4 Cold or Initial Start-Up Operation . . . ... ... e .
H 5.5.5 HotRestartOperation . . . ... ... ... ........
5.5.6 Normal Shutdown Operation
5.5.7 Waste Heat Rejectlon and Recovery System Operatlon
5.6 Performance C e e e e e e e e e
bl 6.0  BUILDINGS AND STRUCTURES . . . ... ... R oL
: 6.1 Site e s s e --- s v . '-.,,', f-» -.'c.n * e v e e o . -- . e .--:- - a s e »
b 6.2 Buildings .............. . ... ... R
i-é 6.2.1 Classification of Structures, Systems o
and Equipment . . .. .. .. ...t ...
6.2.2 ReactorBuilding . . ... ......... e e e
ey | "
iv
7.0

8.0

9.0

10.0

105

CHEMICAL PROCESSING - .+ & v v vt e i i e e e e et e e e e e e e a 7-1
70 Gemeral . . ... e e e 7-1
7.2 DesignBases . . . .. it it e e e e e e e e e 7-1
7.21 ProcessFlowBasis . ............. e e e e e 7-1
7.2.2 SafetyBases . . . .. ... e e 7-2
7.2.3 Plant Maintenance Bases . . . . ... ... ... ... ... 74
7.3  Fuel Salt Chemical Processing . . . ... .............. 7-5
7.3.1 Purpose of Chemical Processing . . . ... ......... 7-5
7.3.2 ProcessDescription . . . .. ... ... .. 0., 7-5
74 IndexofDrawings . ... ..............c....... -~ 7-8
REACTOR OFF-GASSYSTEM . & + ¢ ¢ ¢ ¢ 4 e v o et o ot o e v v v o 8-1
8.1 System Requirements . . . . ... ... .. ... ... 8-1
8.2  BasicAssumptions . . ... ... e it e e e e 8-1
8.3  Reactor Interface System . . .. ... .. ..... e e . 8-2
8.4  PurgeGasCleanupSystem . ... .................. 8-2
8.4.1 Physical Holdup System . .. ................ 8-2
84.2 Charcoal BedSystem . .. ... ............... 8-3
GENERAL DESCRIPTION OF ELECTRICAL :
DISTRIBUTION SYSTEM . . . . . . . .. i i ittt it e e 9-1
FUEL-SALTDRAINSYSTEM . « ¢ ¢ v v v v v vt v v e v e e o e e e e 10-1
10.1 Purpose . e e e e e e e e e e e e e 10-1
10.2  Design Objectives of the Fuel-Salt Drain .
. Tank System .. ........... e e et Ve .. 101
 10.3  Design Objectives of Draln Tank Coolmg ' :
, System e e e 10-1
10.4 Design Criteria of Drain Tank ...... e 10-2
General Description of Drain Tank with L
~ NaK Cooled Heat Disposal System . . ... .... e . 102

 
 

| APPENDIXES

 

-

| A - Primary Piping Stress Analysis
fi B - Number of Steam Generators and Reheaters
C - Flow Stability of Supercritical Fluid in Steam Generators

D - Stress in Tubes Due to Differential Thermal Expansion

Vi
&

 

1.0 INTRODUCTION -

1.1 Origin of Industrial Interest in MSBRs

 

For a number of years Ebasco Services Incorporated (Ebasco) has had an interest in
advancing the technology of molten salt reactor systems as a natural consequence of its
continuing search for promising ways in which to serve its principal client, the Utility
Industry. This Industry, in its service to the public, constantly seeks to produce reliably, and

~ at lowest cost, the energy needed in the domestic and industrial sectors of the economy.

The desire to conserve and, if possible, extend the energy resources of the world to reduce
the environmental impact of energy generation, and to present the lowest hazard to the

- public are also strong motivations among the electric power companies.

In the summer of 1969 Ebasco announced the formation of a Molten Salt Group
bringing together all major industrial capability needed to conceive, design, manufacture,
and construct molten salt reactor systems for the utility industry and with utility companies
participating financially. The utility participants are:

Dallas Power & Light Company
Houston Lighting & Power Company
Kansas Gas and Electric Company
Middle South Services, Inc.
Minnesota Power & Light Company
Northeast Utilities Service Company
Texas Electric Service Company
Texas Power & Light Company
Union Electric Company

Industrial participants in that group are:

Ebasco Services Incorporated - Management, technical direction, nuclear

design, power plant technology.

" Babcock & Wilcox - Reactor vessel, primary heat exchangers, general reactor
technology, steam generators.

Continental Oil Cd., Inc. - Chemical processing, chemical engineering.

" Union Carbidé - Graphitg .techno_lo'gy. |
Cabot Corporation - Haspéllloy-N , special metal alloys, materials technology.
Byron-Jackson : Fus.eld salt l_mmps, general pump technoiogy.

The industrial corr.lpanies supplied senior technical personnel at their own expense to

work as a team under the management and technical direction of Ebasco to evaluate Molten
Salt Reactor Technology.

 
 

On September 30, 1970, Union Carbide Corporation, Nuclear Division, operators of
the Oak Ridge National Laboratory for the USAEC, issued a request for a proposal for an
independent Molten Salt Breeder Reactor Design Study. Ebasco and its group of industrial
companies responded with a proposal which was accepted by the Union Carbide
Corporation and the USAEC with Ebasco Services Incorporated as principal subcontractor
to Union Carbide Corporation and with the Molten Salt Group members as
sub-subcontractors to Ebasco. The official commencement date of this contract was March
8, 1971, and is expected to run 30 months.

1.2 Project Organization

 

The Molten Salt 1000 MWe Breeder Reactor Conceptual Design Study is under the
Corporate Cognizance, in Ebasco, of L. F. C. Reichle, Vice President - Nuclear. The Nuclear
Division (Figure 1.1) is responsible for the MSBR design study. This study is under the
technical direction of D. R. deBoisblanc, Ebasco’s Chief Nuclear Consulting Engineer and
Manager of Research and Development. Figure 1.2 shows the organization including the
sub-subcontractors.

The organization chart (Figure 1.2) shows five major divisions: Systems and
Components, Technical, Reactor Engineering, Plant Design, and Instrumentation and
Control. The Systems and Components Group is responsible for the conceptual design work
on the major functional components such as heat exchangers, reactor vessel, pumps, etc., in
the primary salt system and also for the development of the flow sheets and conceptual
design detail for the various subsystems.

The Technical Group provides the special consulting and conceptual input in the
areas of physics, chemistry, metallurgy, and graphite technology.

The Reactor Engineering Group is responsible for the overall nuclear engineering
design of the reactor including the reactor physics, thermal-hydraulics and the specifications
of the geometry of the graphite structures,

The Plant Design Group brings in the traditional power plant design disciplines of
mechanical engineering, electrical engineering, concrete-hydraulics engineering, architectural

and structural engineering and estimating.

The Instrumentation and Control r.Gl.'od'p will be responsible for the conceptual

“design of all instrumentation systems for the reactor and process systems in the plant. These
groups report to the Project Manager who in turn reports to the Technical Director.

Permission was obtained from the utility sponsors of the Molten Salt Group for a
continued activity of this Group until utility funds were expended. In addition, the
Technical Representatives of the industrial members of the group serve as an informal
advisory panel from time to time during the course of the design study. The availability of
this additional MSR effort within Ebasco and this advisory panel greatly enhances the
 

NUCLEAR

VICE PRESIDENT
L FCREICHLE

 

 

 

 

 

 

 

 

 

CHIEF . _
CONSULTING NUCLEAR MATERIALS
NUCLEAR LICENSING ENGINEERING
ENGINEER :
2 CHIEF ENGINEER MANAGER

MANAGER R & D

D R DeBOISBLANC

 

 

- 1

 

 

SPECIAL
PROJECTS

MOLTEN SALT
DESIGN STUDY

PROJECT
MANAGER

T A FLYNN

 

 

 

NUCLEAR
TECHNOLOGY

MANAGER

HCOTT

 

 

 

RADWASTE

&
WATER
TREATMENT

SUPERVISOR

S | GLASSMAN

 

 

 

" APPLIED
PHYSICS

 SUPERVISOR

' u
T A FLYNN

 

 

 

 

 

INFORMATION
SCIENCES
LABORATORY
DIRECTOR

W W BLACK

 

 

H OSLICK

 

 

 

I

NUCLEAR
LICENSING

 

SUPERVISOR

 

H OSLICK
(ACTING)

 

 

NUCLEAR
SAFETY
ENGINEERING

 

SUPERYVISOR

c oNDISHKO® *

 

 

NUCLEAR
ENVIRONMENTAL
ENGIN EERING

PRINCIPAL
NUCLEAR
| ENVIRONMENTAL
' ENGINEER

 

 

R M HARTMAN

 

 

| TeEcHNicAL
REPRESENTATIVE

 

 

"~ J D NORRIS

 

 

 

 *ACTING UNTIL COMMENCEMENT

OF PROPOSED DESIGN STUDY
**DECEASED

Figure 1.1

 

B E TENZER

 

-

 

MATERIALS
APPLICATIONS

CHIEF
MATERIALS
ENGINEER

W J FRETAGUE

 

1

 

CHIEF
WELDING
ENGINEER

LMPETRICK

 

 

QUALITY
COMPLIANCE
ENGINEERING

CHIEF
QUALITY
COMPLIANCE
ENGINEER

W C HERMAN

 

 

VENDOR
QUALITY
COMPLIANCE

 CHIEF-VENDOR

- QUALITY
COMPLIANCE

L E ELLISON

 

 

 

" EXPEDITING

CHIEF
EXPEDITOR

C R COUGHLIN

 

Ebasco Nuclear Organization

 

 

 

 

 

 

 
uoneziuediQ 1vfoig I(PmS HasS

z'1T 2y

I — g

 

 

 

TECHNICAL DIRECTION

D.R. DE BOISBLANC

 

 

 

 

 

 

 

PROJECT MANAGER

MOLTEN SALT STUDY GROUP

 

 

 

 

SYSTEMS &
COMPONENTS

 

T.A. FLYNN, JR.

 

 

| REACTOR VESSEL. -

B. MONG (1)

HEAT EXCHANGERS

STEAM GENERATOR
REHEATER

 

 

J. ACCIARRI (2)°

CHEMICAL
PROCESSING

 

 

| PRIMARY PUMPS

G. CLASBY (3)

SECONDARY PUMPS
OTHER PUMPS

1

TECHNICAL

 

_H.C., OTT

 

 

 

 

PHYSICS

 

 

 

 

W.W, BLACK
CHEMISTRY
] .
H.C. OTT
GRAPHITE
e

R. M, BUSHONG (a)

 

 

METALLURGY

 

 

 

 

R. MARTINEZ

RADWASTE

 

OFF=~GAS SYSTEMS '

 

T.K. ROCHE (5)

 

 

 

(1) BaBCOCK &
' WILCOX
(2) CONTINENTAL
ofL
(3) BYRON-
JAGCKSON

 

T.A. FLYNN, JrR,

 

 

 

1

 

REACTOR
ENGINEERING

. R, TRINKO

 

 

 

 

 

 

PLANT DESIGN

S. SPARACINO

 

 

INSTRUMENT A-

TION & CONTROL.  Lop

 

 

 

 

CORE DESIGN

R, TRINKO

 

MECHANICAL

R, L. FERRAR]

 

 

 

 

 

 

THERMAL~-
HYDRAUL.IC

R. MARTINEZ

 

ELECTRICAL

D, SANDIFORTH

 

 

K. DOWNES

REACTOR
I ———

K. DOWNES

PROCESS
PR

K. DOWNES

 

 

 

 

 

 

 

GRAPHITE
STRUCTURES

JOINT DESIGN
EFFORT

 

CONCRETE~HYD

B, C. GRANT

 

 

 

ARCHITECTURAL~
STRUCTURAL

E. BOVERA

 

 

(a) uNiON
CARBIDE
(5) cABOT CORF.

 

 

ESTIMATING

E.L. BETZ

 

 

 

 
 

overall management and technical direction of the project so as to maintain it within a
commercial framework which is one of the main goals of the design study.

The sub-subcontractors participating in the design study with Ebasco provide design
effort for a substantial portion of the overall MSBR Plant. Babcock & Wilcox provides the
design of heat exchangers, steam generators, and the reactor vessel. Babcock & Wilcox has
been active in the atomic energy business supplying these components to the Atomic Energy
Commission and to the Nuclear Navy since 1944 when they manufactured equipment for
the Manhattan Project. In addition, the company is currently designing and fabricating 13
nuclear steam supply systems for the nuclear industry.

Continental Oil Company, Incorporated is engaged in the design of the fuel salt
chemical processing system. CONOCO is among the world’s ten largest energy companies
selling over $2.5 billion worth of goods every year. CONOCO is involved in all forms of the
energy business; petroleum, coal, nuclear, chemical, and plant foods.

Stellite Division, Cabot Corporation, provides assistance in the design of advance
Hastelloy materials. Cabot Corporation is a diversified producer of performance chemicals,
energy, and engineered products. Its Stellite Division produces many types of high quality

“alloys developed to resist different conditions of wear, heat, and corrosion. They maintain a

staff of fully qualified chemical and metallurgical engineers.

Union Carbide Corporation, Carbon Products Division, provides assistance in the
design of graphite core structures and other graphite components. The Carbon Products
Division has a 75-year background in the manufacture of carbon, graphite, and related
ceramic, metal, and composite materials products. Carbon Products Division is the only

‘producer in the industry fabricating all varieties of carbon and graphite products. Its

production plants are supported by a completely integral technical center at Parma, Ohio.

Byron-Jackson, Division of Borg-Warner, provides assistance in the design of
salt-circulating pumps. Byron-Jackson is the leader in supplying pumps for the LMFBR
Sodium Reactor Program, and a leader in supplying pumps to the water-cooled nuclear
industry. Byron-Jackson conducts research on pump technology.
2.0 PURPOSE AND SCOPE OF TASK 1

2.1 PurEose

Task 1 had three main purposes:
a) to generate the bases for selection of an MSBR reference concept,

b) to develop an MSBR plant concept from these bases and the criteria
specified in the contract,

c) to identify other concepts meriting further study.

The reference concept developed in this study is one which the industrial team can
recommend to utilities for future construction. The plant should be licensable and operable
on a utility system with safe, economic, and reliable performance. This concept assumes
accepting advances in technology anticipated over the next fifteen years.

2.2 Scope
Emphasis of this study was limited to:

a) systems or components vital to MSBR success even though their cost
may not be significant, and

b) . systems ot components having high cost even though technical
feasibility may be reasonably well assured.

Based on these considerations, efforts \firere limited to the primary system (including
reactor and vessel), secondary system, steam system, off-gas system, fuelsalt drain tank
system, building and general arrangements, and design of a chemical process plant based on

‘the ORNL flow sheet.

‘Several graphlte element de51gns and methods of replacement were examined.
Several reactor vessel concepts were examined in some detail. Several primary system piping

-arrangements were laid out and stress analyzed. Several component arrangements were

examined. Design concepts for steam generators, reheaters, and primary heat exchangers

~were developed. A supercritical steam system with 700 F feedwater was analyzed and

designed. A layout of buildings and structures was developed and seismically evaluated. The
performance criteria of the off-gas system were re-examined. An alternate concept was

~ proposed, and an engineering design of the ORNL concept was developed. The electrical

distribution system was analyzed. Several fuel-salt drain tank systems were explored.
The initial effort in Task 1 was to establish what each system could and should do as
an integral part of the power plant. After cstablishing criteria (some of which were
borrowed from the Molten Salt Group Reports), we attempted to define a conceptual design
which would satisfy the criteria. In some cases, our design reached a high degree of
completion, e.g., the moderator element and the steam system. In other cases, this report
represents several alternates which must be evaluated in Task 2. In all cases, however, the
alternates are such that the overall concept resulting from the combination of each system is
a total power plant concept.,
 

¢

. r

3.0 REACTOR SYSTEM

The rcactor system is composed of the reactor vessel, the reactor control rods and
drives, the reactor internals, the circulating fuel salt, the fuel salt pump, the intermediate
heat exchanger, the fuel salt drain tank, and the piping system connecting these
components. Power is produced in the reactor when the fuel salt, 7LiF-BeF,-ThF4-UF4
(71.7-16-12-0.3 mole percent) is made critical by the moderating properties of graphite
structures through which it flows. The fuel salt is composed of low neutron absorption cross
section materials blended to produce a liquid having fluid properties suitable for developing
a breeder reactor, a high temperature reactor, and a reactor with inherent safety
characteristics. The fuel salt operates above its liquidus temperature of 930 F. The nominal
reactor inlet temperature of 1050 F provides sufficient margin to avoid concern over
freezing of the fuel salt in the system. The salt is quite viscous and so the highest AT is
established at 250 F, and the reactor outlet temperature is 1300 F, the upper limit being
based on the available mechanical properties of the piping systems as a function of
temperature and by a contractual requirement for Task I.

The pump cooling consideration results from the lined pipe concept in that the
bypass flow that cools the vessels experiences a rise of 50 F to 1100 F at the top of the
core. This stream flows in the outlet pipe liner at a rate of 30.5 gpm. The flow in the liner
rises 50 F in passing through the outlet pipe and up around the pump bowl. This allows us
to maintain the certerline injection temperature in the pump at 1150 F (as in the ORNL
4541 pump design).

The combination of the high operating temperatures and the extreme corrosiveness
of the molten fluoride fuel salt severely restrict the materials which may be used in the
system. The high moderating power of carbon (which is determined by the low neutron
absorption cross section of carbon coupled with its low atomic weight which renders
neutron scattering collisions efficient in slowing down neutrons) and the excellent high
temperature structural properties of graphite make this material uniquely suitable for the
reactor internal structures. The use of graphite in nuclear reactors has a history extending
from the same original point as that of uranium, the Stagg Field pile. The behavior of
graphite under radiation is complex in that normal graphite materials behave differently
along their crystallographic axes, i.c., they are anisotropic under radiation. This behavior is
even more complex from a mechanlcal design standpoint in that a cross section first exhibits

" a linear shrinkage after exposure to a fluence of 1.5 x 1022 nvt and then reverses the
~direction of change and swells. The studies at ORNL of graphite’s radiation induced
properties have led to a definition that the useful core lifetime of graphite is reached when

the regrowth of the ongmally contracting material equals the original dimension. For
graphites having suitable properties for MSBR applications this occurs at 3 x 1022 nvtata

“temperature of 700 C or about 4 years of operation at an 80 percent plant factor for the
-physics conditions leading to satlsfactory breeding. Our design study has initially accepted
* the ORNL physics design of the reactor (ORNL 4541) as a basic assumption for Task I. We

plan to run benchmark and survey calculations on physics uses in mid-FY 72.

3-1
 

N

A family of high nickel alloys, known as Hastelloys, has been developed which are
capable of containing the fluoride salts under pressure. Hastelloy was successfully employed
as the structural material for the MSRE. The high temperature radiation endurance of this
material has been established as a ground rule for the Task I design study. Curves giving
design stresses as a function of temperature have been supplied for Task I use. Modified
Hastelloy-N has been defined as having a value of Sm of 3500 psi at 1300 F. Since 3Sm is
defined by the nuclear power piping code as the maximum value of primary and secondary
stress acceptable for conservative design, the design study was originally restricted in the hot
leg portion of the plant to stress values of 10 500 psi for the combined sum of pressure,
weight, seismic and thermal forces. This restriction has had a very pronounced impact on
the mechanical and structural design of the reactor vessel, the fuel salt pump and the hot leg
piping system. It led to the Task I design effort to develop a lined and/or jacketed primary
system. The curve supplied by ORNL for Sm as a function of temperature shows a very
sharp rise in strength with decreasing temperature, e.g., Sm=13 000 psig at 1100 F.
Accordingly, the premium associated with reducing the temperature of the pressure
boundary is quite high. We found the thermal shock problem in the reactor outlet line so
severe that we concluded a bare system in this leg was only marginally feasible. Surprisingly,
the reactor inlet leg also required a protective measure when we considered the thermal
shock associated with loss of secondary cooling pump power. The temperature in the inlet
leg rose rapidly to 1300 F with a loss of strength and severe shock occurring. We decided to
avoid detailed design effort on this case by adopting a lined (or jacketed) inlet line.

A lined system is defined as one in which the pressure boundary of the fluid system
is insulated from rapid changes in temperature by either a stagnant or laminar region of the
contained fluid. A jacketed system is defined as one in which the pressure boundary of the
fluid system is separated from the contained fluid by a second fluid. The second fluid
(different from, but compatible with, the fuel salt) is maintained at a higher pressure than
the fuel salt so that leakage always occurs from the jacket fluid into the fuel salt across the
metallic “O” Ring seals.

The entire pnmary system whxch contacts the fuel salt is made of either Hastelloy or

graph1te The use of Hastelloy is dependent on its corrosion resistance and the ability of the

metal to maintain its ductility under neutron bombardment. All Hastelloys contain some

‘boron as an impurity. The botron tends to precipitate at the grain boundaries. Under neutron
irradiation, helium gas, via the 10B (n,a) 7Li reaction, is produced near and generates
~bubbles at the grain boundaries. The presence of bubbles in the grain boundaries is the
principle cause of the loss of ductlhty in standard Hastelloy-N. To overcome this problem
~ ORNL has developed modified alloys containing additives (both titanium and hafnium

appear promising) that have greatly improved rupture life and ductility under irradiation.

‘This material is assumed to have adequate ductility (defined as 4 percent minimum strain to |

f rupture) as a groundrule for Task I. Grain boundary attack was observed in the MSRE. Its
~ cause has not yet been firmly established. Oak Ridge is actively investigating this
 intergranular attack. As a Task I groundrule, we assume that both the cotrosion problem

and the embrittlement problem have been solved.

3-2
 

i

 

In the following scction we shall describe the reasoning which led to our Task 1
conclusions. The first consideration shall be the design of the graphite moderator, reflector
and internal structure of the reactor.

3.1 Graphite Moderator Element Design

 

The fabrication techniques and the properties of graphite for high temperature
radiation service are generally well understood. The excellent service of the graphite for the
lifc of the MSRE has led to confidence in the ability to design these elements. The condition
of breeding imposed on the MSBR design study, however, may make it necessary to have a
graphite that will not absorb xenon gas from the fuel salt stream as it flows over the graphite
surfaces. A contractual ground rule for the design study specified that the xenon poison
fraction be no greater than 0.005 neutrons lost to xenon per neutron absorbed in fissile
material (.5 percent poison fraction).This ground rule has been adopted as a criterion for the
Task I study effort. There are a number of ways of reducing the loss of neutrons to xenon.
We have studied the bubble generator system proposed by ORNL 4541 and, as an alternate,
have investigated a spray system. The results obtained in the salt simulation loop at ORNL
seem to indicate the pump’s action on the bubble rich fluid breaks the bubbles into a very
much smaller size range than anticipated. The effect on mass transfer of the smaller bubble
size has been calculated by several models and it appears to enhance xenon removal due
primarily to increased surface area. The spray chamber appears interesting because the ratio
of gas to fluid for a given surface is very much greater than in the bubble method. The high
gas to liquid ratio of the spray chamber appears to offer potential for removal of tritium as
well as of xenon.

The penetration of xenon into the graphite can be prevented by the application of a
pyrolytic carbon or graphite coating on the exposed surfaces of the graphite element. The
technology of applying exterior pyrolytic coatings to graphite is available today. The
technology of applying pyrolytic coatings to interior surfaces in long pieces of graphite is
less well developed. It is reasonable to expect that coating materials that will successfully
seal graphite against xenon intrusion will be developed within the 15 year projection of the
design study. Demonstration of the ability of exterior coatings to seal against gas
penetration for the 3 x 1022 nvt radiation lifetime of the graphite can be performed by test
irradiations of relatively small graph_:tc_ structures. Demonstrations of the seal for interior

coatings can only be performed by tests that utilize long pieces. No tests of this nature are

currently planned. The integrity of the coating under thermal cycling and mechanical
stresses in full sized pieces will not be demonstrated by extrapolating tests on small pieces.
The full verification of large scale pyrolytic coating under neutron irradiation will require a
considerable extension of the current testing programs which are not currently planned.
Thus, we estabhsh for Task I the followmg criteria for graphite element design:

1) Pyrolytic sealing materials_' are available.

2) Exterior coatings ‘that will resist handling fnctlon, stress, vibration and

temperature changes are available.

33
T

3)

Interior pyfolytic coatings should be avoided if possible.

3.1.1  Mechanical Design

The Molten Salt Groupr critique of the ORNL conceptual design led to several
serious questions with respect to the mechanical design presented in ORNL 4541. These
objections can be briefly summarized here as follows:

1)

4)

The practicality of lifting a large weight remotely under conditions of very
tight clearance on a nonroutine basis (i.e., every 4 years).

The desifability of using an unshielded machine with the risk of a cask
malfunction causing severe problems in recovery.

Graphite elements positioned outside of the central core zone achieve only a
small fraction of their allowable service life.

The feasibility of applying and inspecting a high integrity pyrolytic graphite
coating in small diameter holes running through long graphite elements may
be beyond a reasonable extension of technology.

Based on this critique, we established functional objectives for the design of a
moderator element:

1)

2)

4

Moderator elements shall have only external surfaces.
The maximum heat path from the interior of the log to the salt should not
exceed 0.7 inches. This results in a temperature-damage relationship at least

equal to the ORNL reference design element.

The saltlgraphife ratios for each core zone shall be rnaintairied at the ORNL

_conceptual design values to preserve the physics model to the greatest extent
- practical. (Task Iground rule established by Ebasco. )

The elements shall be '751zed through a trade-off between ease of handling and
the number of moves required to service a core. It was deceided not to

. reinsert underexposed elements because that would require, 1) reorificing

highly contaminated - assemblies, 2) possibility of seal damage and 3)

 variations in coefficient of friction during service life. While a hexagonal

geometry was selected (due to its self-standing characteristics), the results of

~ this trade-off apply approximately to other geometries.
 

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5) Raw graphite pieces shall not exceed the current maximum dimensions for
special graphite production. Manufacturing techniques are available that can
produce acceptable graphite uniformity and density if the distance from any
point of the section to the atmosphere of the graphitizing furnace does not
exceed 9 inches. For cylindrical shapes this implies a maximum diameter of
18 inches. The lengths available are unlimited for our purposes.

The geometries examined in the design study are presented in paragraphs 3.1.2 to
3.1.5. Concurrently with this geometric study, a unit element size selection study was

conducted. The results of the sizing study are presented in paragraphs 3.1.6 to 3.1.7.

3.1.2  Arrays of Cylinders

 

Cylinders were considered as the elements for the core pieces because of ease of
fabrication and the uniformity of heat removal across a section of the rods. The diameter of
the rods would have to remain small, 1-3/8 in. or less, in order to keep the graphite lifetime
in the range required by the contractual ground rules for the reference design. The large
number of these relatively long and slender rods does not appear attractive from the
standpoint of handling and fragility. It would be possible, of course, to use a central hole (1
in. LD, x 3 in. O.D.) and have both ease of handling and long lifetime, but the sealing
problem is not much different from the ORNL design.

Control of the salt fraction and improvement of the salt flow in the space between
stacked rods¥/could be achieved through use of raised helical rib, say 1/32 in. high machined
onto the surface of the graphite.

Unitizing of bundles of rods could be done by banding together hexagonal arrays of
rods at the tops and bottoms of their lengths with Hastelloy-N bands or with graphite. In
the case of the Hastelloy bands the coefficient of thermal expansion differential would
allow loosening of the bundle at operating temperature which would be undesirable. In
addition, there is some question whether the bundle could be held tightly enough to permit
handling as a unit without additional pinning or other means of fastening at the ends.

A possible unit considered consists of a hexagonal array of rods contained in a
hexagonal graphite box. Commerical production of graphite pipe in 10 in.-12 in. diameter is
state of the art, and box-ike extrusions in the same size range have also been produced. In
this configuration control of salt flow could be accomplished by cutaway portions of the
rods and box in the plenum regions.A cross section of a core unit of this nature is shown in
Figure 3.1. o - ' ‘ |

Y Experimental Investiga.t'ioh of Velocity Distribution and Flow Resistance in a

Triangular Array of Parallel Rods, W. Eifler and R. Wigsing (EURATOM) Nuclear

Engineering and Design 5, 1967) North Holland Pub. Co., Amsterdam.

 

35
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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Reactor Cross Section

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3.1.3  Solid Blocks

The use of machined solid graphite blocks was considered as an obvious method of
unitizing core sections into convenient size and weight units for replacement. For purposes
of this study it was assumed that molded blocks or hexagonals in sizes of about 18 in. across
flats by about 36 in. length could probably be produced in an MSBR grade, and that
extrusions 18 in. diameter by 15 ft long in MSBR grade were probably not feasible. As a
first approach to this design, a hexagonal block 9 in. across the faces with relatively large
holes for salt flow was considered (a sketch of this unit is shown in Figure 3.2). In this
sketch the 13 volume percent salt was obtained by using 19 holes approximately 1-5/16 in.
in diameter plus a spacing of 1/4 in. between adjacent hexagonal blocks. It is shown that
much smaller holes or slots would be necessary in order to avoid large distances for heat
transfer from the inner graphite to salt (and hence high graphite temperatures and reduced
life). It is thus shown that difficult and costly machining would be involved and that sealing
of the large number of small interior surfaces would be much more troublesome than in the
reference design. |

In addition, the necessity of several vertical segments to make up the core height
entails careful indexing at each joint and throughout the graphite machining operations to
insure continuity of the salt flow and poses a serious design problem to avoid stagnant salt
or ill defined salt flow regions at the joints.

The size effect on radiation damage is a further drawback. Although the incremental

dimensional changes, creep effects, etc., could possibly be kept in the same level, the
leverage of size is certain to cause breakage at some point in radiation dosage. This in itself
would not necessarily be catastrophic, but would complicate the replacement process which
we are trying to improve.

For these reasons the use of machined solid blocks in the core was considered
undesirable.

3.1.4  Iris Array of Curved Slabs

It would be desirable from the radiation damage (and power generation) standpoint
to have all graphite receive uniform dosage within any individual graphite unit. This is, in
practice, impossible, but an array of concentric cylindrical units built up of curved slabs in
an iris pattern was considered which appears to have some desirable features from the

symmetrical uniformity of radlatlon loading. A sketch is shown in Figure 3.3 which depicts

a section through two units of an iris array of curved slabs which would be unitized by

- a_t_tachment to graphite plates at the top and bottom. This configuration presents a gradually

increasing thickness of graphite (and salt channel) as the radiation intensity falls off and

thus a leveling of graphite temperature and lifetime. Hydraulically, this configuration
requires further de31gn enhancement before becoming acceptable.

3-6
 

 

 

    

 

 

 

  

 

  

 

 

 

 

 

 

 

 

  

 

 

 

 

  

 

 

 

 

 

 

 

   

 

 

 

 

   
  

  

 

 

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Graphite Elements

 

Iris Array
 

-

The curved slabs would be considerably more difficult to manufacture than straight
sided slabs but would be feasible to produce. Replacement of the central one or two
cylindrical units would appear to be feasible but the outer rows would have to be handled as
segmented rings and replacement of odd shaped and sized units is unattractive.

3.1.5 Slab Arrays

The important consideration of short heat flow paths from all areas of graphite in
the core is accommodated very well by use of relatively thin (up to perhaps 2 in. thick) slabs
or “boards” of graphite separated by salt flow channels. The manufacturing technology of
graphite plates is well established and needs only to be extended to MSBR grades. Sizing
considerations need only be consistent with reactor requirements and ease of handling (very
thin plates would be too fragile). A maximum size of plate for study purposes was taken as
2 in. x 12 in. in cross section by full reactor height (approximately ~ 15 ft).

a) Square

The simplest array considered was a square arrangement of plates
sized to the central control section (sketch shown in Figure 3.4).
This array would not be very symmetrical in a circular cross section
but might be entirely practical.The square cross section would require
some holding fixtures during replacement to maintain stability of the
remaining core elements after removal of the first few units.

b) Hexagonal (Task I Reference Selection)

 

The basic design selected for the Task I reference concept is shown in
Figures 3.5-3.8. This concept has all the desirable features mentioned
above, and in addition, is compatible with the overall dimension for
moderator element unit size which was being determined
independently. The element is composed of flat plates for which the
‘manufacturing and coating techniques are certain and with nubs or
ribs to provide the separation for salt channels. The Y-shaped yoke
forms the main structure both for normal operating conditions and
for removal. Flow control may be achieved by orificing in a bottom
end plate, a top end plate, or by a pronounced tapering of the flat
moderator- plates at the inlet and/or outlet. Assembly of the plates
into the hexagonal units is made by cementing and dowelmg All
cemented joints are made in low flux zones.

'The detalled de51gn developed for a typ:cal Zone 1 moderator |

element is shown in Figures 3.9 - 3.11. The basic slabs are extruded
and are machined to final shape. The Y Yoke slabs are then
cemented. The Y Yoke assembly is inserted in the bottom end box

3-7
                    

           

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which acts as a retaining jig during the element assembly. The ring is
free to slide on the milled shoulder in the yoke for 1 in. Dowels
inserted through the 5 in. support legs on the bottom of the yoke
picces hold the retaining ring on the element during lifting. While in
service the ring will float against the milled shoulders of the slabs.
The moderator slabs are then added starting with the closest to the
Y Yokeand working outward to the edge slabs. The top retaining and
support ring is then added. Both rings (top and bottom) are molded
graphite pieces formed out of graphite filament stock and machined
to fit up dimensions. The top ring is then doweled to the moderator

- plates and the edge plate is doweled to its inner neighbor. The
element is then placed in a furnace along with other completed
elements and the cemented joints are fired. After firing, the

. dimensions are checked and any final machining of outer surfaces
performed. The thickness of the moderator plates is varied for the

- central zone elements to accommodate the changing salt fraction
required to provide the axial blanket and plenum regions. The central

core region is composed of 13 percent salt and 87 percent graphite .

and has a height of 13 ft. The blanket region has a volume of 37
‘percent salt, 63 percent graphite and a height of 6 in. both top and
bottom. The plenum has a volume of 85 percent salt, 15 percent
graphite and a height, top and bottom, of 6 in. for an overall graphite
- element height of 14 ft - 11 in. There is one-in. clearance at the
bottom of the element. The distance from the top reflector bottom
- surface to the top of the bottom reflector is 15 ft. The blanket
- moderator elements which surround the core radially have a uniform
37 percent salt fraction for 14 ft, a top plenum of 6 in. and a bottom
plenum of 5 in. of 85 percent graphite and 1 in. bottom clearance of
100 percent salt.

This design produces 6 in. plena at both exit and entrance to the reactor core. These
plena are entered by four transition nozzles which convert the circular geometry in the exit
and entrance piping to a rectangular geometry compatible with the plenum’s thickness. In
addition, ‘these transition nozzles provide balancing points for the several sidestream flows
to originate and return to the main stream. The sidestream at the inlet transition nozzles
provides for:

1)  Bottom head cooling during drain tank dump flow
2)°  Reactor vessel coolmg flow.

At the outlet nozzle this sidestream divides to provide for:

1)  Flow into the nozzle liner-pipe annulus.
~2) An attemperatlon flow into the outlet pipe liner.

~ The entrance condmons for flow into and between the hexagonal elements are most

- constant over the element’s life at the bottom. The element never reaches the regrowth

portion of the graphite cycle due to low flux and low temperature. This makes the bottom

3-8
 

r

r

plenum especially attractive for flow control orifices. It may be necessary to interrupt the
nubs along the length of the element to prevent channeling and stagnatlon regions.

Handling of the moderator blocks has not been considered in detail other than to

consider the lifting requirements and their influence on the mechanical design. Normally the

moderator elements are lifted by rotating a three-pronged tool through the holes provided in
the. top plenum spacer lugs on the Y Yoke. These holes provide a total of almost 4 sq. in. of
lifting area or a load of 500 psi on the lug material. The graphite in the lug should easily
maintain sufficient strength to accommodate 2000 psi loading. Thus even with two defective

lugs the element w111 lift.

Conceptual sketches have been made of a backup tool consisting of three thin plates
which can slip down between the flow channels in the element, rotate and engage all plates
from below. This tool is considered a backup for mechanically damaged elements.

3.1.6  Graphite Utilization

The basic time interval selected for evaluating moderator handling requirements was
the 4-year interval between major turbine-generator overhauls. The basic trade-off
evaluation was between moderator element size and the number of element moves required
to service the core. The conclusion of the handling study, adjusted to the hexagonal
configuration developed in the geometry study, was that the “15 in.” element led to an
optimum total number, size, weight and number of moves for the element. Results of this
study were presented in the progress report dated June 1971.

The core configuration resulting from these studies is shown in Figure 3.1, which isa
cross section view through the midplane of the reactor. The central element is used for
reactor control and safety rods. There are 162 full hexagonal elements in six full hexagonal
rings and a partially filled seventh ring (24 full and 12 half hexagons out of a possible 42
elements). The reflector consists of 96 slightly wedge-shaped slabs 2 ft - 6 in. deep by 14 ft
long with an average width of 8 in. and a 2 in. taper over the width. The reflector is

~ connected to the array of full and 1/2 hexagons by 18 triangular and 24 flat wedge-shaped

spacers. The vessel is protected by a 2 in. thick neutron shield made either of borated

 graphite or a Hastelloy-N honeycomb structure filled with boron carbide. One purpose of

this shield is to attenuate the thermal neutron flux by a factor of 103 thus reducing the He

,generatlon rate in the reactor vessel. In addition, the thermal neutron shield forms a flow

‘baffle to divert a stream of inlet salt for the purpose of cooling the reactor vessel. The

- cooling salt annulus is approxlmately 1/2 in, thick. The reactor vessel is assumed to be 2 in.
* thick and is designed for a nominal pressure of 75 psi like that proposed by ORNL in ORNL

'4541. Salt enters the reactor vessel thzough four inlet nozzles, each of which is protected

“from rapid thermal transients by a liner similar to the top transition nozzle liner

- configuration. The cold (1050 F) salt from the heat exchanger forms a thermal sleeve to

protect the inlet pipeline from rapid changes due to transients in the secondary salt system.
This salt stream provides the normal and the emergency cooling flow for the bottom head.

39
 

et

L

3-10

'3.1.7  Future Thermal - Hydraulic Considerations

In Task 1 of the deSIgn study the reactor outlet temperature has been specified as a
maximum of 1300 F. The inlet temperature is chosen as 1050 F. These choices were based
o on reasoning similar to ORNL'’s, stemming from the high melting point of the primary salt,
. H | the high pumping power required by the viscous salt, the allowable stress values in
| Hastelloy-N material and our desire to preserve the ORNL reference design physics model
. for Task 1. The basic objective of the orificing provided is to maintain a 250 F temperature
__ u | | rise across the core. The hexagonal moderator concept developed in our design study utilizes
' a similar hydraulic control concept to the ORNL prismatic element, viz., unorificed flow in
- the flow channels between elements and orificed flow in internal channels. As we progress
H — further into the moderator element design, we must resolve several important questions on

hydraulics.

 

» H 1) The number of orifice zones required. In our configuration we would
anticipate that no more than seven zones (seven rings of hexagons)
would be required.

2) The need for interrupted spacer ribs to promote interchannel mixing.
u If studies in Task II show undesirable hydraulic effects, then the ribs
b can be interrupted. Also, holes can be drilled through the plates to
5 | - allow interchannel communication.

R 3) The actual method of providing the flow control. Control can be
- provided as part of the end boxes (either top or bottom) or as an.
o ~actual change in sections of the moderator plates at entry or exit.

. - Top onficmg would permit hydraulic compaction. Our design allows
: us maximum freedom of chmce based on the results of future
studies.

| 4) The flow perturbations resulting from dimensional change of the
| ~ graphite. These changes are functions of position, temperature, and
R A | | power history. Since the lowest temperatures and lower power
| | | | - history will be’ at the inlet, our current effort is directed at orificing
- there. :

5) - The hydraulic implications of the change in flow area required in the
- o | o ~ central region element to switch from 37 percent salt to 13 percent
e o salt and back to 37 percent salt again as the flow rises through the
: o C | -:core, thus formmg the axial blanket region. |

- An analys1s of the heat removal requirements for the radial reflector

blocks. Our present concept involves the use of large graphite blocks
7 - stacked to form a free-standing structure during handling operations
w ' | ~ (including initial loading). Design details must be completed to show
| | that adequate cooling flow between all blocks is provided. ‘

.
 

t-m—.m\

r

7) An analysis of the heat removal requirement of the top and bottom
~ reflectors. Our design provides a 6 in. plenum at the top and bottom
of the core to: a) improve the flow path from the inlet to the element
orifices and b) to reduce radiation damage to the reflector since we
desire both reflectors to last the plant life. The present top head
concept makes it possible to replace top reflector blocks. The
replacement of the bottom reflector would be a major operation. To
allow for this we are considering the use of a 2 in. thick inlay of
graphite over the central 11 ft radius of the bottom reflector. Since
the bottom reflector graphite runs cold (1050 F), it can take much
more fluence than the top reflector graphite. Thus it is unlikely that
the bottom reflector will require replacement. This inlay would be
replaceable - during the replacement of the four-year-life graphite
elements. -

° 8) The effect on the salt flow area of the 2 percent change in graphite
linear dimensions. At its maximum shrinkage, the interior portion of

the element experiences a change in graphite area to 83.5 percent

- from 87 percent. The corresponding change in salt area is from 13
percent to 16.5 percent. This effect is experienced in the ORNL

design also, but in that design the effect is absorbed by an overall
compaction in the outer flow channels. In our hexagonal design the

bridging effect of the radial rings strengthens the outer flow channel

structure and prevents the general compaction. The effect could
possible lead to a Task II physics opnmlzatlon that has a central
region less than 13 percent.

3.2 Reactor Vessel

The basic. concept of the reactor vessel for the MSBR design study is shown in
Figures 3.12 - 3.14. In each of these figures the same configuration of vessel is used with a

‘different head closure presented for each case. The vessel is a right circular cylinder 2 in.
(nominal) thick, having straight sides. The vesse! is closed by a bottom dished head of 3 in.-
~ (nominal) thickness forming a continuous membrane with the cylindrical vessel. The vessel
~ top head isa dlshed section of 2in. (nommal) thick Haste]loy-N

~ Three methods of closmg the top head onto the cylindrical vessel have been
~ developed on Task I and will be further evaluated in Task II. The first closure involves the
“use of a support skirt for the head extending from the top of the graphite reflector into the
" accessible areas of the reactor operating floor. Both vessel and head support skirt are bolted
together in a clean, cold area. This design requires the use of a large skirt for both the vessel -
‘and the head, which are fabricated of Hastelloy-N. This design may be unattractive for
. économic considerations. However, its design simplicity and the ability of effecting a closure

at low temperature in a nonradiation environment are tangible operating benefits. A cooling
flow would be introduced between the vessel and the head support skirts to give a

3-11
 

A
i

il

e

programmed temperature gradient in the distance from the top of the reactor to the

operating floor. The second head closure involves the use of asupport skirt for the head and
a vessel extension skirt again reaching from the top of the reactor to the operating floor, and
is shown in Figure 3.12. This more complicated geometry may be required to avoid

* streaming problems in the annular space formed by the 2 in. vessel, the 2 in. head support
skirt and the 1/4 in. void between them. The third head support scheme shown is Figure
~ 3.14, which involves the use of a flanged bolted head. A problem exists in this design in that

the space required between the outlet nozzle of the reactor and the flange on the vessel
produces an unnecessarily large top reflector region. To avoid the penalty of filling this
space with graphite, which would serve no useful neutronic purpose, we have provided a
graphite support structure which would hold the top reflector. We have several concepts for

this top structure. In one concept, it consists of an independent member which is removed

separately from the top head. In another it consists of a structure integral with the top head

- consisting of bars on which notched graphite members are -assembled and pushed into

position. The void above the structure supporting the top reflector would be filled with
inert gas, and possibly with injection tubes for salt, to cool the reflector.

There are four symmetrical inlet nozzles.The inlet nozzles terminate in a liner which
is located inside of the pressure bearing pipe coming to the reactor vessel to the head

exchanger outlet. The purpose of the liner is to protect the pressure bearing pipe from the

rapid changes in temperature associated with the loss of secondary coolant pumping power.
The loss of secondary coolant pumping power would produce a rapid rise in the temperature
of the salt exiting from the heat exchanger, and would constitute a serious shock to the inlet
pipeline and the reactor inlet nozzle. The liner protects the inlet pipe and the reactor inlet
nozzle from this shock by virtue of the fact that salt in laminar flow is an excellent insulator
and restricts temperature changes to very slow variations even when the fluid inside the inlet
pipe liner experiences a rapid temperature change. ’

The inler fluid enters the reactor at 1050 F. It flows through a 6 in. plenum
containing 85 percent salt and 15 percent graphite formed by the 5 in. support legs. A small
stream of salt is diverted down from the inlet to cool the bottom graphite reflector. This

“heated bypass remixes with the main stream in the bottom plenum. A very small salt stream

passes outside of the thermal neutron shield to cool the bottom head and is dumped to the

- drain tank by the lower head drain line. The reactor vessel cylinder wall is cooled by a
- bypass flow of 140 gpm which is diverted from the inlet nozzles and flows up in the gap
‘between the thermal shield and the reactor vessel wall. The vessel cooling flow rises 50 F to
‘1100 F as it is collected at the outlet nozzles. The 1100 F fluid is divided into four streams
of 35 gpm each, which is passed between the outlet nozzle and the transition nozzle liner
~ extension, This extension passes the cool stream of salt that has washed the vessel down to

the pump in an annular laminar flow to protect the outlet nozzle and pipe from the thermal

 shock of a reactor scram.Our calculations indicate that following a scram the fuel salt outlet
~ - tempetature falls from 1300 F to 1050 F in less than 10 seconds. For the thickness of pipe
" in the primary system, this is equivalent to a step change of 250 F. The top reflector
graphite is cooled by a stream of 1050 F salt injected through the control rod drive

penetr ation.

3-12
3-13

3.3  Rcactor Drain System

 

3.3.1  Drain Tank and Reactor Lining

 

Fuel salt is taken from the primary system both intermittently and continuously
(Figure 3.15). The intermittent drains from the primary system occur through the reactor
drain line, which is controlled by a freeze valve and located at the lower-most point of the
reactor vessel bottom head, and secondly in the event of a pipe break from the catch basin,
which is located beneath the primary system.

The catch basin is isolated from the drain tank by a rupture disk and a check valve.
The catch basm drain line is tied to the reactor drain line and penetrates the reactor cell
containment in a common penetration. The drain tank is situated in a separate containment
cell directly below the reactor containment.

e T f"'(

The continuous drains from the primary system are:
. 1) The reactor vessel bottom head drain line.
! 2) - The four pump overflow lines.
e 3) The four heat exchanger drain lines.
- 4;) The chemical plant return feed line.

;o The reactor dram line and the heat exchanger drain lines 1ntroduce fuel salt to the reactor
- cell batch tank #t a temperature of 1050 F. The chemical plant feed return line is also at
| relatively low temperature (probably 1100 F). The pump overflow lines are introducing fuel
D salt at a temperature of 1300 F. The volume of salt from the pump overflows will be held to
= a low value, just|sufficient to keep the line hot and in service. This line is available to allow
: overflows of larger quantities of salt'in the event surges occur in the pump tank. The major
flow of salt will l?e at the lower temperature, 1050 F,

The ratlc'b of hot to cold salt will be adjusted to produce a mixed-mean temperature
‘in the reactor cell batch tank of 1100 F. During surges’ thls temperature will be hlgher

o ~ The off-gas removed from the system by the gas stripping device w1ll provide an
: “ | intense source of decay heat. This gas will be introduced into the reactor cell batch tank
: where it will mix with the 1100 F fuel salt and be carried to the drain tank through the
-  reactor cell batch tank drain line. Since the fuel salt will be substantially below the 1300 F
| E | limit, which we are using in the design study for Hastelloy service, there will be sufficient
heat capacity in the nominal 600 gpm flow rate to contain the offgas decay ‘heat without

requmng external coohng of the line.

Fluid wi i be drawn from the drain tank by means of jet pumps and supplied to:
_. - 1) The chemical plant feed line.
‘\J 2) The primary system coolant pump suction line.
o 3) The chemical sampling system.
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A possible alternate on the return of the cooled fluid from the drain tank occurs in the fully
lined piping system. The concept described in the Task I report uses a stream of fluid leaving
the pump to serve as a thermal barrier between the turbulent inner stream and the pressure
containing pipe. Since this nominal layer of salt in drawn from the pump bowl,

“temperature. is 1300 F, and the allowab]e Sm is only 3500 psi. Relnjectlon of the return

flow by a booster pump into the Jacket between the hot salt and the pressure pipe on this
leg would allow a substantial upgrading in the stress capability of the Hastelloy. Only 2 small
portion of the return flow of approximately 150 gpm per loop would be required for this
jacketing function. A nominal value of 35 gpm has been used in Task I calculations. The
remainder would still return to the pump suction via the jet pump

In designing the reactor vessel bottom head for a dual drain system, i.e., one drain

~ for the bottom head space between the thermal neutron shield wall and the reactor vessel

wall, and a second drain for the primary salt to leave the reactor region and go directly to
the drain tank, we have provided flexibility that will enable a drain accident to be
accommodated without overheating of the reactor bottom head. In the drain, the fluid will
leave the reactor vessel passing through the bottom graphite space, enter the separate
concentric reactor drain tank line, and flow to the drain tank. The back flow from the heat
exchangers will pressurize the fluid in the plenum and also cause a flow to enter the space
between the reactor vessel and the thermal neutron shield, thus continuing to provide cool
fluid to the reactor vessel head in this region. The flexibility of this design allows us, if it
proves necessary, to incorporate into the system with minor modifications a third fluid
complete jacketing system except in regions where the Hastelloy can be kept in net
compression or easily replaced. In the present concept we have lined the system using fuel
salt at reactor inlet temperature and allowed the temperature of the liner fluid to rise 50 F
in passing up to the inlet nozzle liner and a further 50 F in passing from the nozzle linear to
the pump bowl. In the full jacket system a separate source of salt, possible L,B or a
uranium free salt mixture, would be injected into a sealed liner. Seals would be located at
the junction of the transition nozzles and the thermal neutron shield reactor vessel at the
outlet nozzle, thermal neutron shield junction, in the outlet nozzles and at the primary
pump piping bowl transition piece, and in the inlet line to the heat exchanger. These seals
would not be designied to be absolutely tight since the jacket salt would be selected on the

- basis of its compatibility with the fuel salt and would be maintained at a higher pressure
_than the fuel salt. Small inleakages could be tolerated. A total seal length for all seals
- required to fully jacket the system would be approximately 120 ft.Metallic or ring seals can

be expected to perform in this dimension to restrict inleakage to approximately 1 percent a

- day. In laying out the primary cooling system (we have assumed that we would use the

largest components feasible based on preliminary estimates by the Babcock & Wilcox

'Company), it appears that the entire intermediate heat exchange function can be
~conveniently provided in four units. Accordmgly, we have selected four loops to the reactor

vessel. The piping studies performed attempted to minimize the volume of primary salt

contained in the piping while at the same time accommodate the thermal expansion
requirements of the system. The first configuration studied is shown in Figure 3.16 (detail

A), and this is a minimum inventory system where the heat exchanger and the reactor vessel
and the pump are connected by the shortest possible pipe runs. The inlet and outlet nozzles

314
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Primary System Arrangements

 

 
 

 

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to the vessel do not fall in the same vertical plane but are d1splaced from each other so that
the piping system is categorized as the strut configuration. :

3.3.2

 

Piping and Component Arrangement

In developing our approach to the Task I piping for the reactor system, two design
considerations became paramount:

2

b)

The length of the intermediate heat exchangers far exceeded the
length of our reactor vessel.

The available high temperature strength of Hastelloy-N was urgently
needed for thermal transients and could not be spared for re31st1ng
large seismic forces.

The structural implications of the support requirements are more fully discussed in
Section 6.0. We will only summarize the implications of these studies as they affect the
reactor piping system in this section.

a)

d)

A decision was made to eliminate top support of the reactor and heat

'exchanger in part to reduce the seismic loads.

Even with the reduced magnification of the imposed seismic load due .

to lowering the building loading point by 50 ft, the piping could not
stand the differential response of the nonsynchronous reactor with
the four heat exchangers. -

A decision was made to provide a rigid horizontal coupling between
the reactor and the four heat exchangers and to anchor this intertie
into the building so that the building, the reactor, and the four heat
exchangers respond within the same forced horizontal motion, thus
eliminating any component loading on the pipe, i.e., the pipe must
withstand only its own and its contained fluids acceleratlons

The prov:sxon of a continuous metal support path from the bottom

of the containment through the top of the reactor and heat

exchangers greatly increased the thermal growth of these vessels with

.frespect to the operating deck and ‘its fixed support points. (The

pump mountlngs and the vessel flange)

~ The decay of reactor power followmg a reactor scram imposes the most severe

transient the plant can produce on the piping system at the point of highest temperature

and lowest allowable strength — the reactor outlet line. Scram will produce an outlet

_ temperature line change of 250 F in less than 10 seconds. We were not able to restrict the
~secondary stress produced by this thermal shock to allowable limits. To solve this problem
we adapted a pipe liner to provide a thermal barrier between the pressure containing pipe

3-15
A P

and the hot turbulent salt leaving the reactor. The liner has two basic functions: a) it
insulates the pressure pipe from the 1300 F outlet salt by trapping a stream of cool salt that
has bypassed the core to cool the reactor vessel; this salt is heated from 1050 F to 1100 F
at its flow rate of 30.5 gpm per loop; it constitutes a large thermal resistance to a radial
transfer of heat, and b) it restricts the outer pipe wall temperature thus making this wall
stronger. In Task I we have not utilized this second factor in the design since the actual
thermal and hydraulic performance of the liner requires further study.

The thermal transient produced in the system by a loss of secondary coolant
pumping power is nearly as severe for the reactor inlet line. If we assume no natural
circulation or bulk mixing of the secondary salt, this transient heats the salt leaving the heat
exchanger to 1300 F(from 1050 F) in under 30 seconds. To protect against this transient we
have provided a similar low velocity laminar flow through a pipe liner.

It should be noted that all the transients discussed here are assumed to proceed in
their most odious fashion. There has been no corrective action assumed. For example, scram
has been assumed to occur by firing all safety rods into the core at the highest velocity
possible. No coastdown was assumed for the secondary salt pumps. There are a great
number of corrective measures that will improve the plant’s ability to accept transients, e.g.,
programmed scram rates, pump flywheels, and pony motors. However, if the-plant can

accept the unmodified transient, it cannot be rendered unsafe by the failure of a corrective

measure, and so our design effort has been directed against deliberately chosen severe
conditions.

The inventory in the primary coolant piping system in the strut configuration was

lowest for the configurations studied. The pipe stress calculations performed on this

configuration resulted in extremely high values of both primary and secondary stresses and
bending moments, because the configuration proved to be extremely stiff. Accordingly, this
configuration was abandoned as being not practical. Calculations for the hot leg used a value
Sm of 3 500 Ibs/sq. inch. {This is the value at 1300 F.) |

~ The next configuration studied is shown in Figure 3.16 (detail B). In this
configuration (which we called the inine arrangement) the primary heat exchanger, the
pump and the reactor centerline all fall in a common vertical plane. The piping is necessarily
longer than in the strut configuration which results in a higher salt inventory penalty. Hot

" salt leaves the top reactor nozzle, flows in the bottom of the pump, is circulated through the

pump discharge line to the heat exchanger plenum, and returns to the reactor through the

bottom heat exchanger discharge line entering the reactor at the inlet nozzle. The inventory .
~ penalty in this configuration is equivalent to the distance from the centerline of the pump
~ to the reactor inlet nozzle in both horizontal and vertical dimension. This configuration also

proved to be extremely stiff and mast of the hot leg from the reactor through the pump to
the heat exchanger inlet was considerably overstressed. |

After studfing this :irr_ahgemérit we undertook a third configuration shown in Figure

© 3.16 (detail C). This is called the dog-leg configuration. The hot salt leaves the reactor outlet

nozzle, enters the pump casing which is located on a plane in common with the centerline of

3-16
 

the reactor, is rotated through the pump casing, exits from the pump in a direction at 90 F
from its entry, passes to the heat exchanger, flows into the heat exchanger top plenum
down through the heat exchanger tubes, is collected in the heat exchanger outlet plenum,
and is returned to the reactor through a pipeline that passes back under the pump at 90 F and
enters the reactor in the same plane as the reactor inlet nozzle and pump centerline. In other
words, the return line is everywhere below the exit line from the reactor, making a right
angle bend underneath the pump. This configuration has an even higher inventory of salt
than in the inline configuration. The total linear distance traversed in the hot line is
approximately the same. The total linear distance traversed in the return line is again
approximately the same, but there is a considerably greater quantity absorbed in the elbow
than in the straight reactor pipeline. This penalty is small in comparison to the increase of
the inline arrangement with respect to the strut arrangement. The stresses in this
configuration again proved to be over the allowables, that is over 3Sm in the hot line.
However, they were closer to allowable than they had been in any of the preceeding runs by
virtue of the increase in flexibility accounted for by the hot-leg configuration. It had been
noted that changes in elevation "of the pipeline caused by changes in the support
configuration below the reactor produce substantial variations, variations in fact that were
larger than those produced by the change in piping configuration in the pipe stresses.
Accordingly, it was decided to lower the reactor with respect to the heat exchanger. This
would provide more flexibility in the hot line which was the one that was most severely
overstressed and make the cold line or leg more rigid. A reduction of 12 ft was effected, and
a substantial improvement in pipe stresses resulted.The degree of overstressing present in the
system with this configuration was within the margin allowable in a Task I study, i.e., 3Sm
of 10 500 psi. The resulting stresses were less than 3 percent greater than the allowable. This
was considered satisfactory, and the configuration was selected for the Task I concept.
Additional details of stress analysis are presented in Appendix A.

317
 

4.0 SECONDARY SYSTEM

 

4.1 General

It is necessary to redundantly isolate the highly radioactive fission products
contained in the circulating fuel salt from the potentially dispersive forces that could be
caused by the high pressure contained in the steam system. This protection must be
afforded without interfering with the transport of heat from the fuel salt to the steam
system. To accomplish this decoupling an intermediate coolant salt system is provided.

The advantages of using a coolant salt loop are:

1)

3)

To provide a redundant barrier for

a) protecting the steam system from fission products and

b) protecting the primary system from steam system pressures.

To bridge the temperature gap between the fuel salt melting point
and the steam system feedwater temperature. |

To compress the tubes of the intermediate heat exchanger to
a) insure inleakage of secondary salt to the primary system and

b) to keep the Hastelloy-N tubes in compression to prevent
cracking associated with intergrannular corrosion.

A flow diagram of the coolant salt system is shown in Figure 4.1. The coolant salt
system is comprised of four independent loops each loop consisting of shell side of the
intermediate heat exchanger, the shell side of the steam generator and the reheater, a
coolant salt pump, a blowout device for overpressure protection, a coolant salt melt tank for
adding salt, a coolant salt storage tank, a coolant salt filter, a cover gas system, coolant salt
piping and valving. The selection of the coolant salt for the MSBR design study is confirmed
by a study by H. C. Ott for the Molten Salt Group which evaluated:

a) Cost of inventofy and makeup
b) Cost of heat transfer surface
c) Cost of pfimp and pumping power

d) Cost of fuel salt inventory in the Intermediate Heat
Exchanger (a function of heat transfer area)

41
 

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e) Chemical stability and vapor pressure (at temperature)
f) Corrosion characteristics
g) Compatability with the fuel salt in the event of leakage into

the fuel salt system
h) Effect on the migration of tritium to the steam system.

Based on this study, Ebasco has selected the eutectic of NaBF4-NaF as the Task I
reference coolant salt. A possible alternate selection for Task II study is He based on the use
of a fluidized bed heat exchanger. The use of He gives the potential of using a simple
chemical purification for tritium recovery. The increase in overall heat transfer coefficients
derived from the use of the fluidized bed decreases the size and design pressures required of
a helium coolant system to the point of practicality.

4.2  Steam Generator—Reheater Concept Selection

4.2.1 Backgzound

This section establishes concept recommendations to Ebasco for the steam
generators for use in the 1000 MW(e) MSBR reference concept. Based on experience, the
formulated basic design criteria establish the basic design features such as vertical
orientation, counter-current flow, and other preferred characteristics of the various designs
that have been studied. The section lists and discusses the selection criteria which were used
to evaluate the entire range of feasible designs.

After the designs are compared, four are selected for detailed evaluation. These four
concepts are arranged in order of preference from the component designer’s viewpoint, and
recommendations are made for further evaluation in conjunction with the system designer.

The design development, evaluatlon and selection for the steam generator are also
applicable to the reheater. Section 4.2.6 deals w1th a study of the reheater.

| 4.2.2 Basic Desigh ’Criteria )

In the context of this report, design criteria represents the procedures and methods

. used by the component vendor to meet the requirement of the MSBR system. When

developing the design criteria for future steam generators, we used experience gained
through design, fabrication, and operation of past units. Based on this experience with

high-temperature, high-pressure steam generators, several basic requirements have been

established. The following features provlde the simplest, most economical, and reliable
units: ,

1) The units will be basically shell-tube designs with the
high-pressure supercritical fluid inside the tubes and the
coolant salt on the shell side.
 

 

 

 

Most standard materials like Croloy and stainless steels are impractical for high

2)

4)

The heat transfer surface shall be arranged for
~ counter-current, once-through flow between coolant salt and

water; when using parallel (co-current) or mixed parallel and
counter-current  flows, the high-temperature difference
existing between the coolant salt inlet and supercritical fluid
inlet would imposc severe thermal stresses on the tubes due

to the large temperature gradients across the tube wall. The

thermal effectiveness of the heat exchanger is also increased
if counter-current flow is used.

The steam generator should be arranged so that the heated
fluid flows upward in the region of heat transfer, and the
cooled fluid flows downward. This arrangement minimizes
chances of unstable operation.

All bofindaries 'contéining' coolant salt are made of
Hastelloy-N or modified Hastelloy-N, but other boundaries
containing water may be made of Croloy-2-1/4 or stamless
steel.

cotrosion resistance on the coolant salt side. Hastelloy-N, which does have good cotrosion
resistance, is very costly compared to the standard materials; therefore, for the concept
selection the use of Hastelloy-N has been restricted to the coolant salt boundaries. The cost
of Hastelloy-N in the future may be comparable to standard material after sufficient
industrial use, and/or a less corrosive coolant may be developed.

After establishing the common design factors, it is necessary to formulate the
various possible designs for identifying the major parameters that can be varied. The design
parameters that were considered for this preliminary study are as follows:

1)

2) -

3

Tube Gebmetry

a)  Straight tube (mcludmg sine-wave tube, C tube, hockey-stlck tube) ,
b)  Utube -
- ¢)  Return-bend tube (plattens)
- d) Helical-coil tube

' Size of Units

a)  Large (4 units)

b) Module (8 or 16 units)

Unit Orientation

a)  Vertical

b) Horizontal

4-3
-Emmmn
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b

.

e

Using evaluations performed in conjunction with Ebasco and evaluations based on
the preliminary study of thermal- hydraulic performance and cost evaluation (discussed in

" Appendix B), we selected four large units for the reference 1000 MW(e) design. Selecting four
full-size units rather than eight units was for plant simplicity; otherwise, no significant

differences in performance, cost, safety, or reliability are evident.

Vertical units are preferred to horizontal units because the horizontal units are more
prone to stratification into temperature layers of coolant salt in the shell side. (One
exception is the U tube, Ushell with the hot leg vertically above the other leg.) This

- condition affects the performance of the supercritical fluid side and may cause reverse flow

and instability. Although it is possible to reduce this instability by orificing and baffling, it
is generally not practical; therefore, horizontal units are restricted to the U tube and U shell
(with hot leg above cold leg) types.

The combination of the design parameters and the limitations previously mentioned
provides several practical and feasible designs, from which four concepts that are based on

selection criteria are recommended.

4.2.3 Selection Criteria

The selection criteria that were applied to various feasible steam generator designs
are themal stresses; thermal-hydraulic performance; manufacturing; and inspection,
maintenance, and repair.

Thermal Stress

The thermal stress criteria for the steam generator are similar to those for the IHX.
The major problems in steam generators are thermal stresses (particularly during transient
operation) combined with the mechanically induced stresses caused by the high-pressure
supercritical fluid. The design of the tube bundle is most important; however, unlike the

 IHX, the steam generator tubes, tubesheets, and plenums are designed for very high
watersteam pressures. These sections are quite thick, are associated with slower transient

“response, and have nonlinear temperature gradients and higher thermal stresses. A

combination of tube geometry, tubesheet, shell, and plenum designs arranged for low

- combined stresses is desirable. -

o Thermal-Hydraulic Performance

e is 1mportant to study the thermal-hydrauhc performance of each of the steam
generator designs because the density dlfference between supercritical fluid inlet and outlet

. conditions ranges from 35 to 5 1Ib/ft3. A thermal system having such a large density
~ difference should have reasonably . well-balanced flow through the system. A large flow
~imbalance impairs the heat transfer performance, causes large temperature differentials in

different circuits, and enhances the possibility of flow and pressure oscillations. The
problem is more severe at low loads when the frictional pressure losses are considerably
reduced with the reduced mass flow rates, and when the pressure differences due to density

4-4
 

differential become predominant. The steam generator should have a stable operation at full
load and at expected minimum loads. The details of flow stability and reversibility are
explained in Appendix C.

Manufacturing

Each design was studied to determine the degree of difficulty in manufacturing the
components. For the IHX the assembly of the tube bundle and of the tube supports was a
problem because most of the preferred designs were variations of the straight- tube concept.
The manufacturing problems with straight-tube variation concepts and (to some extent)
with U tube concepts are discussed in section 4.3. The return-bend (Platten) concept and (to
some extent) the helical-coil concept are not as well established as the conventional heat
exchangers. The degree of difficulty when manufacturing these units can only be studied
qualitatively.

For the steam generator the tube configuration of several varieties should be studied
further. Each of the configurations requires an entirely different manufacturing approach;
expetience has proved that all these are feasible to manufacture but are manufactured with
some degree of difficulty, which is often directly related to in-service reliability.

It may be concluded that the manufacturing of the concept design should be simple,
and for reliability the final assembled unit should be accessible for inspection.

Inspection, Maintenance, and Repair

Selecting these criteria is probably the most difficult because the state of the art in
1980 is very difficult to assess due to rapid developments in the industry.

The inspection and maintenance of the steam generator are much simpler than those
of the THX and reactor internals because the supercritical fluid side is not radioactive. The
tubes are easily accessible from the manholes in supercritical fluid-side plenums with little
delay for the radioactive decay. If required, major repairs could also be accomplished with

the unit in place. Due to this relatively high degree of confidence the steam generators are
usually designed to have non-removable tube bundles and this condition enables the designer

to offer simpler designs.

It is expected that the various designs formulated earlier could use identical

inlet-outlet plenums. The major difference among the various designs is in the tube-bundle
region, which is fixed between the tubesheets and surrounded by shell. Since the routine
inspection and maintenance areas are outside this fixed region, the inspection and.

maintenance criteria are not significant when selecting a steam generator concept.
424 Concépt Evaluation -

Several steam generator concepts are evaluated. Some of the units that show obvious
drawbacks when subjected to these selection criteria are eliminated; for the remaining

4-5
 

 

r

rm:.—-r
b 1

.

i

concepts that may be used, all the merits and demerits in each category of selection criteria
are tabulated. The discussion that follows is in the sequence of tube geometry formulation.

Straight Tube and Variations

Figure 4.2 shows four different concepts based primarily on the straight-tube
principle. The straight-tube bundle (Figure 4.2a) is the simplest to manufacture and is
probably the most reliable from a manufacturing viewpoint, but from the thermal-stress
viewpoint, the straight-tube bundle is expected to experience severe transients and is

eliminated.

To withstand large thermal differential during transient operation, more flexible
tubes are needed. The sine-wave, hockey-stick, and C tubes, which are treated as variations
of the straight tube, are shown in Figure 4.2 (b, ¢, and d). Each of these variations can be
selected for the practical steam generator design; however, there are manufacturing
problems with the hockey-stick and the Ctube bundle — these are the shell of the
hockey-stick tube bundle and the support baffles of the C tube bundle. With the increasing
use of sine-wave tube bundles currently, major advances in the manufacturing of these units

are expected.

Figure 4.3 shows the bayonet tube arrangement, which is also a variation of
straight-tube bundles. Each tube consists of two concentric tubes fastened to separate
tubesheets at the top. The inner tube is open at the bottom, and the outer tube is sealed at
the bottom. Coolant flows down the inner tube and returns up the annulus. This
arrangement increases the size of the tube and the unit due to extra tube thickness required
in the larger-sized outside tubes. Compared to other straight-tube variation, the unit is
somewhat thermally inefficient; however, experience with this concept is quite limited and
it will not be pursued further.

Based on the previous discussion, we have selected the sine-wave tube bundle for

further evaluation.
U Tube Concept Variations

Figure 4.4 shows four different concepts using the Utube geometry. The
arrangement shown in Figure 4.4a has one serious disadvantage, which is the close proximity
of the tube-side inlet and outlet, which gives rise to very high thermal gradients. This is a

‘very poor arrangement when faced with high-temperature units and is eliminated.

The vertical U tube arrangémeht (Figuie 4.4b) is generally avoided because it is
potentially instable. The desirable feature is to have the heated fluid flow up and the cooled

' fluid flow down. With two equal legs this criterion is violated in one half of the unit. This

arrangement also causes maldistribution and is eliminated.

The horizontal, U tube, Ushell unit (Figure 4.4d) has both legs in the same
horizontal plane. This arrangement is prone to stratification due to temperature layers of

4-6
suopjenre) 1dasuo) aqny, ySreng
Z'y amdig

 

 

 

 

 

 

 

a. Straight.Tube

 

 

b. Sine-Wave Tube

 

 

 

 

C.

 

 

 

 

 

 

Hockey-Stick Tube

 

 

 

 

 

 

 

C~Tube
 

o

N

| S

N
-
-’
~

N
W

m‘fr i ‘

e

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Bayonet Tube

Figure 4.3
Bayonet Tube Arrangement
 

 

 

4

 

e

 

 

 

 

 

a. U Tube—Straight Shell b. U Tube -U Shell

 

 

 

 

 

Y

 

 

 

|fl

c. U Tube—U-Shell Hot Leg . d. U Tube—U-Shell
‘Above Cold Leg in Same Horizontal Plane
Vertical Plane

Figure 4.4
U Tube Variations
 

 

 

coolant salt in the shell side; however, this condition may be reduced considerably by
providing several baffles for mixing on the shell side. This unit is not pursued further.

The horizontal, U tube, Ushell (Figure 4.4c) has the hot leg above the cold leg in
the same vertical plant. This unit has potentially better flow stability characteristics and

many other advantages of the U tube; it is further evaluated.

Return-Bend Tube (Platten)

 

The return-bend tube (platten) concept meets all selection criteria requirements and
is evaluated further. (See Figure 4.7.)

Helical-Coil Tube

The helical-coil tube concept meets all selection criteria requirements and is
evaluated further. (See Figures 4.8 and 4.9.)

It may be concluded that all four steam generator concepts selected previously are
suitable for the 1000 MW(e) MSBR reference design. The discussion that follows indicates the
advantages, the disadvantages, and some special comments for each of these concepts.

4,2.5 Detailed Evaluation of Selected Concepts

Vertical, Once-Through, Straight Tube
With Sine-Wave Bend (Figure 4.5)

Advantages

1) The unit is the simplest to manufacture.

2) Much less flow modeling and experimentation are necessary. Experience in
the operation of the sine-wave tube bundle unit ensures stable thermal and

hydraulic performance (for supercritical operation).

3)  The degree of difficfilty of performing rhajor repairs may be considerably
reduced because the experience with straight-tube bundles may be utilized.

4) Thermal efficiency is increased because small tubes which have small wall
thicknesses can be used.

Disadvantages

1)  Designing sine-wave tube bundles for suitable envelopes is difficult;
preliminary calculations indicate that to achieve stable operation the length
of the steam generator should be about 65 to 75 ft. This length may not
match properly with the system arrangement. It is possible to provide several

4-7
 

|2
£
i

 

Supercritical Fluid Outlet

     

 

t_.....m
msnr e

—

 

LB

i - Coolant
= Salt Inlet

2755 Tubes (0.5"
OD, 0.281" ID)

Coolant
Salt QOutlet

 

Supercritical Fluid Inlet

Tube -Side AP, ~300 psi
Shell-Side AP, ~ 200 psi

Figure 4.5
Straight Tube With Sine-Wave Bend Concept

L
 

 

2)

1)

2)

1)

2)

3)

4)

3)

1)

orifices to increase stability without adding length, but this approach is not
efficient and would require further cost evaluation.

A large sine-wave bend is necessary to withstand large transients usually
associated with the steam generators. There are some problems in providing
tube-support baffles in this large, sine-wave bend region.

Comments

If the large envelope for the steam generator is acceptable to the system designer,
further evaluation of the sine-wave tube bundle concepts is recommended.

Horizontal, Once-Through
U Tube (Figure 4.6)

Advantages

The thermal efficiency is increased because small-size tubes can be utilized.

Stable units can be designed by providing large numbers of baffles for better
mixing of the coolant salt on the shell side. The propagation of coolant salt
flow disturbances to the water-steam side may be avoided.

Disadvan tages

Large U bends and large clearances are required to withstand steam generator
transients.

The average temperature difference between the hot and cold legs of the
U-shell is generally very high. A complex support arrangement is required to
avoid high thermal stresses.

The tube-support baffles in.the U bend region have proved to be a major

problem in recent designs.

The shipment of the large unit mziy be quite difficult.

The ménufacturing of the U bend region of the shell has to be in several
pieces and is complex.

Vettical Once-Through, Return Bend
(PlattenL(Flgure 4.7) '

Advantgges

The tube bundle is quite flexible; therefore, it can withstand large thermal
transients.

4-8
 

(Coolant Salt Inlet

 

 

 

r
)
=
)
|
AN

 

 

 

Supercritical
Fluid Outlet

2755 Tubes (0.5"
OD, 0.281" ID)

Supercritical
Fluid Inlet

 

 

 

 

|
AN\

 

 

 

 

Coolant Salt Outlet

 

 

 

H . - l< _ _ “45;

St ' ; Tube-Side AP, 300 psi-
- | |  Shell-Side AP, 200 psi-

 

;. H Figure 4.6
U Tube, U-Shell Concept With Hot Leg
| Avobe Cold Leg
 

it

3

oo

o

96!:

~ 125 Involute Plattens
(5' long. 30' high, 22
tubes each on 0.875"

‘ 40"
centers) 1no"

2750 Tubes (0. 5" OD,
~100' long)

 

 

Figure 4.7
Involute Return-Bend Tube Bundle
 

 

 

£
t

2) Small-size thermally efficient tubes can be used.

3) The steam generator can be designed with large varieties of envelopes.

Disadvantages

1) The manufacturing of the tube bundle is complex and more expensive.
2) The manufacturing and operating experience is limited.
3) The shell-side flow is virtually axial from top to bottom of the steam

generator; therefore, the circumferential mixing on the shell side is
negligible. Should there be any major shellside flow disturbance, the entire
length of the platten would be affected.

Comments

From experience it is possible to design plattens based on the results of the use of
the helical-coil or straight-tube bundle. The results also indicate that the obvious trend is to
use large numbers of thermally efficient small tubes; this use increases the manufacturing
cost.

Vertical, Once-Through, Helical-
Coil (Figures 4.8 and 4.9)

Advan tages

1) The tube bundle is quite flexible; therefore, it can withstand large thermal
transients.

2) With the additional parameter (helix angle), it is possible to design the steam
generator to fit any reasonable envelope.

3) The chances of experiencing coolant salt flow disturbances of large
magnitude in some tubes are quite limited. The tubes are coiled parallel to
the shell; therefore, the water, flowing inside the tube, experiences the
coolant-salt flow at various locations along the circumference. The
coolant-salt flow disturbance is shared by most of the tubes.

4) Manufacturing and operating experience is rapidly increasing.

Disadvantages

1) = Due to manufacturing difficulties, large numbers of thermally efficient
small-size tubes cannot be used.

2) The coiling of the tubes and the manufacturing of the tube bundle are
difficult.

[
 

iy

(

r. . . | | r P t

o

r .

 

Coolant Salt W

Inlet

 

Supercritical Fluid Qutlet

RN

 

 

815 Tubes (1" OD, 0.6"
' ID, min clearance betw
adjacent tubes = 0. 375")

 

 

 

 

 

Helix Angle, ¢ = 10°

 

40
Eff Tube-
Bundle
Height

   

Shell (10' OD)

 

Coolant Salt
QOutlet

Supercritical Fluid Inlet

Figure 4.8
Helical Tube-Bundle Alternate Concept
 

 

 

-

Y
-

ti A

T “

o
'

'

A

Supercritical Supercritical
Fluid Inlet Fluid Outlet

Tubesheet (~22" thick)

 

Coolant Salt
Inlet

815 Tubes (1" OD, 0.6"
ID, min clearance betw
adjacent tubes = 0. 375")

 

 

 

 

 

 

 

 

 

 

 

 

Helix Angle, ¢ =10°

 

40!
Eff Tube-
Bundle
Height

\
N
\
N
N
N\
N
N

N
N
\
N\
\,
N\
N\
\\\
N

 

 

 

 

———Shell {10' OD)

 

 

 

 

 

 

 

Figure 4.9 _
Helical Tube-Bundle Concept
 

 

42.6 Reheater

The recommended reheater arrangement is identical to the steam generator except
for size; it was selected for the following reasons:

1) Costs and time spans for delivery will be reduced because
similar design and fabrication processing can be used for the
steamn generator and the reheater.

2) Development of welding mockups, stress modeling, and flow
modeling will be minimized.

3) The helical-coil arrangement permits versatility in the final
selection of the number of reheater units.

4) The helical-coil unit appears to be the most attractive because
the selection criteria of the helical-coil unit are nearly the
same as those for the steam generator.

A conceptual sketch of the selected reheater arrangement is shown in Figure 4.10.
To determine the basic size of the reheater, a parametric study was performed.

A primary objective when selecting a reheater design is to provide an arrangement
with a low steam-side pressure drop at minimum cost. For this study Ebasco recommended
that the tubeside (steam) pressure drop be limited to 20 psi. The shell-side (salt) pressure
drop was set in the range from 10 to 20 psi.

Other design parameters which were evaluated by tradeoff studies were the
Hastelloy-N volume, the number of units, and the unit height and diameter. When identical
designs were compared, the Hastelloy-N volume was virtually independent of the number of
units. For a straight-tube unit with full-diameter tubesheets (chosen as a base for parametric

study), the amount of material was constant at about 210 tons for one unit or 25 tons for -

each of the eight units. The helical-coil unit utilizes significantly smaller tubesheets because
fewer tubes are required for a given heat transfer surface. For example, the recommended
concept uses four 10 in.thick tubesheets in lieu of a single 28in.thick tubesheet;
however, this weight saving is counterbalanced by additional weight caused by using a large
tube diameter (due to manufacturing considerations) and a larger shell diameter. The
estimated weight range of the helical-tube reheater is 220 to 240 tons.
Four units were selected for the design because of the following factors:
1)  Desirability for plant arrangement.
2) Insigriificant Hastelloy-N material volume differences.

3) Operational flexibility.

4) Performance requirements met with reasonable dimensions.

4-10
 

Stean’: Inlet Steam Qutlet

 

=i

Tubesheet (~10" thick)

 

 

t,@
NN
A\

Coolant Salt
Inlet

t |

 

 

 

 

S
o6

Bundle (50" ID)

 

 

 

 

 

   

1200 Tubes (1'' OD, 0.9168"
ID, min clearance betw
adjacent tubes =0, 375")

¢
)

VI Ll L L idl el L
77

1
. Helix Angle, ¢ = 20°

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

24!
Eff Tube-
Bundle
Height

A
¥

 

—Shell (9' OD)

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Coolant Salt Outlet

 

t e e
it

| - Figure 4.10
Helical Tube-Bundle Reheater Concept
 

4-11

4.2.7 Conclusions and Recommendations

From an overall view of the design development, the evaluation, and the selection of

L several feasible steam generator concepts for the 1000 MW(e) MSBR reference plant design,
[ four different concepts have been established. The first choice is the vertical, once-through,
e helical-coil unit, which should be developed in subsequent tasks of the design study. The
ke concepts in order of preference are as follows:
y 1) Vertical, once-through, helical-coil.
2) Vertical, once-through, straight tube with sine-wave bend.
e 3) Vertical, once-through, return bend (platten).
B 4) Horizontal, once-through, U tube unit with hot leg above
- cold leg but in the same vertical plane.

P It is recommended that these concépts be evaluated further by the designers of the
b system and the components to establish a mutual choice.

4.3  Primary Heat Exchanger Concept Selection

 

4.3.1 = Background

This section discusses the criteria by which a primary heat exchanger (IHX) concept
. is to be judged and presents a large number of standard and nonstandard concepts with
H | evaluations based on the criteria. The IHX concepts are evaluated according to the following

breakdowns:

 

1) Removable Tube Bundle } Primary fuel salt
2)  Nonremovable Bundle in the tubes

- 3) Primary Fuel Salt on Shell Side

'4)  Advanced Concepts

Concepts 1 and 2 are rélatively standard compared to those generally studied for
“LMFBR heat exchangers except that the hot primary fluid is inside rather than outside the
tubes. Another difference is'the smaller size of the primary plenums, which causes problems

of flow distribution. The extent of the problem of the even distribution of primary fuel salt
| , ~around a cylindrical tubesheet could not be determined. One (undeslrable) solution for
e |  better flow distribution is addmonal mlet and outlet nozzles in each IHX plenum.
B A much more detailed manufactunng evaluation is requued to properly rate the
a | ~ concepts with primary salt on the shell side. Only a brief evaluation is made of the advanced
w concepts since these have major effects that should be evaluated by a systems engineer.
The criteria by which the concepts are to be evaluated are discussed below. The
parametric charts used to select an optimized straight-tube and helical-tube bundle are
included. The selected tube-bundle sizes are indicated on the concept sketches in section
4.3.3.

4.3.2  Selection Criteria
The selection criteria that were applied to various IHX concepts are primary salt
inventory: thermal stresses; thermal-hydraulic performance; manufacturing; and inspection,

maintenancc, and repair.

Primary Fuel Salt Inventory

 

Designs with low primary salt volumes appear highly desirable due to the design
requirement to minimize fuel inventory. In most conventional designs of heat exchangers
the ratio of tube-side volume to shellside volume is low (Figure 4.11). When using these
conventional designs, the arrangement of primary fuel salt on the tube side is mandatory.
For example, the reference tube bundle design chosen has 3/8 in. diameter tubes and
5/8 in. pitch (or 1/4in. between tubes). Figure 4.11 shows that the ratio of shellside
volume to tube-side volume is about 4. The chart is based on a straight-tube bundle, but the
results can be applied to most tube bundle designs.

Figure 4.12 shows a typical arrangement of parameters for a straight-tube-bundle
IHX. This chart shows the tradeoff among length, bundle diameter, tube size, tube spacing,
and tube-side and shellside pressure drop. For easy comparison rough tube-side volume

numbers, corresponding to each tube size, are given. The advantage of the smaller sized

tubes can be easily noted when envelope and primary volume are considered.

For most tube designs, such as the sine-wave, the C tube, the hockey-stick, and the
J tube, the envelope of the heat transferring zone is about the same as that for a straight
bundle. There will be small differences, but most designs are assumed to have
3/8 m outside-diameter tubes on a 5/8 in. pitch, and the salt volume penalty will depend
on the bundle ends and not on the main heat transferring region.

The helical-tube bundle is different. It would be extremely difficult to use
3/8 in. diameter tubes, as the bundle height/diameter ratio would be quite small.
Accordingly, the helical-tube design is heavily penalized on a volume basis.

. Thermal Stress
The stress criterion, which is used as a measure of the adequacy of the various IHX

designs, is the magnitude of the tube stresses caused by differential thermal expansion. It is
not immediately obvious why this condition exists. When stress analysis is performed,

4-12
 

 

Volume Ratio of Shell Fluid to Tube Fluid

1.

 

 

  

Ligament = 0. 35"

/0.25"

 

 

[

 

.50 0.75 1. 00 125

Tube OD, in.

Figure 4.11
Straight-Tube Bundle, Ratio of Shell-Side
Volume to Tube-Side Volume
cm.m E3
oo s i,

ety
s

t_‘ r

- S
o —

R 0
—

£ r

Bundle Length, ft

100. O

90. O

 

80. o}

70.0I—

60. OF—

50, Of—

40. 0

- 30.0—

 

 

 

 

 

Ligament=0,25" .
— — — Ligament=0.375

sv=v=rv Shell-Side AP, psi

 

emusme—= Tube -Side AP, psi

 

  
      
    
  

Tube, 0.75 OD

Avg Tube-Side Vol=200 ft3
Downcomer Vol=150 ft3

— — — Avg Shell-Side Vol=525 ft3
Avg Shell-Side Vol= 325 ft3

 

Tube, 0.5 OD

 

Avg Tube-Side Vol=120 ft3
Downcomer Vol =80 ft3

« —~Avg Shell-Side Vol=525 ft3
Avg Shell-Side Vol =325 ft?

 

Tube, 0.375"0OD

Avg Tube-Side Vol=70 ft?
Downcomer Vol=50 ft3

— — — Avg Shell-Side Vol= 525 ft3
Avg Shell-Side Vol =325 ft3

F

g 1 | 25 | o '
60.0 70.0 80.0 90,0 100.0

 

 

Bundle OD, in.

Figure 4.12
Straight-Tube Bundle IHX,
Parametric Chart
tubesheets, shells, nozzles, and heads should also be considered. Preliminary analyses of
these components indicate that the thicknesses, diameters, and shapes of these are within
state-of-the-art technology. Flat, circular tubesheets will be relatively thick (9 to 13 in.)
to reducc stresses, but they are not a major factor in limiting the basic IHX design. Shells,
nozzles, and headers will not be overly thick (1 to 3 in.), despite the high design
temperature, because the design pressures are low. Thick sections should be avoided because
they are subject to high radial temperature gradients during transients, and the number of
cycles of these transients fatigue the metal.

Supports are not a problem. Support skirts for bottom or top support are easily
designed; hanger rods for top or middle support are also possible, and all these can be used
to isolate the structure around the IHX from the 1300 F temperature of the unit.

The stress problem, which has not been solved easily, is that of secondary stresses in
the tubes. These are caused by differential thermal expansion between the tubes and the
shell joining the tubesheets. If we assume that the primary fuel salt flows in the tubes to
minimize primary inventory, the average temperature of the tube wall will be between the
average temperature of the primary salt (1300 + 1050)/2 = 1175 F, and that of the
secondary salt (1150 + 850)/2 = 1000 F. If a linear axial gradient is assumed, the average
temperature of the tube wall will be (1175 +1000)/2 = 1087.5 F. In most designs the shell
that joins the tubesheets is insulated on one side; so its average temperature is the
temperature of the shell-side fluid, which is 1000 F. Calculations show stresses of
EaxAT2(26.3 x 106) (8.4 x 10'6) (87.5) = 19 250 psi in a straight-tube bundle when the
tubes are 87.5 F hotter or colder than the shell. The allowable stress intensity range for
Hastelloy is 10 500 psi (3 S, at the 1300 F, maximum tube metal temperature).
Consequently, the steady-state --thermal stress or stress intensity is almost twice the amount
which is allowed for all thermal conditions combined. Note that the ASME Code Section HI
definition of differential thermal expansion stresses as secondary stresses is used.
Mechanically induced stresses, such as pressure and vibration stresses, are not negligible, but
these are not as significant as the 'thermally. induced'stresses.

The straight-tube de31gn w1th opposed tubesheets, either fixed or floating, is not

~ feasible under the given steady-state temperature differences; however, there are a few basic
 solutions applicable to the IHX to reduce tube stresses. There are three methods commonly

used to allow for differential thermal expansion between the tubes and the shell joining the
tubesheets. These are discussed as follows.

‘The two methods that secm apphcable are flexible tubes and bypassing-flow along
the shell. Some flexible tube shapes which are considered are the U tube, the sine-wave, the

~ Ctube, the helical, the J tube, and the hockey-stick. The tube shapes which appear useful to

the 1000 MW(c) MSBR, their sizing, and their manufacturing problems have been discussed
previously. Using these tubes would produce smaller secondary stresses than those produced
in straight tubes.

4-13
 

o

The sccond method generally involves bypassing a small percentage of the tubeside

fluid along the surface of the shell that connects the upper and lower tubesheets. The rate
and amount of flow are controlled so that the axial temperature gradient in the shell nearly
matches that in a typical tube. The object is to minimize the difference between the thermal
growth of the shell and that of the tubes; therefore, the tubes need not be very flexible to
absorb the small differential growth that does occur. Straight tubes in a fixed bundle with
this bypass can be considered if the connnecting cylinder responds to thermal transients as
fast as the tubes respond. In comparison to bent tubes, straight tubes are desirable due to
the manufacturing simplicity and size economics.

The third method, one that has been used in the past but is not developed in any
detail for the MSBR component design study, is the use of the flexible shell. The most
obvious design for a flexible cylinder is to include bellows, such as those used in the shells of
the Hallam Nuclear Power Facility IHXs. Currently, the major objection to the use of
bellows is their reliability and maintainability relative to the alternative designs mentioned
previously. Experience has shown that the bellows are susceptible to loss of integrity if they
are cleaned improperly or if the surface is not entirely free of scratches and dents. For lack
of other options, bellows are sometimes used in nuclear plant applications, but even then
they are generally small-diameter bellows located in accessible areas where their failure
would not produce a major safety hazard.

We recommend a straight-tube IHX design to suggest what might be done if large,
reliable bellows could be obtained in the next several years, although it appears that the
advantage gained would not be a great one. There are other problems, such as remote
maintenance methods and high-temperature material behavior, that are more worthy of
immediate development.

Some tentative conclusions can be made about the thermal stresses expected in the

various tube shapes which were considered. The steady-state thermal stresses have been

calculated for several of these and are presented in Appendix D. An assumption made for
each tube shape is that the total desplacement of the tube is the same as the free differential

- thermal expansion of the shell.

~ The types of tubes consideréd are the straight, the sine-wave-bend, the hockey-stick,

~ the Ctube, and the hockey-stick/ sine (ORNL IHX concept).

The Straight tube viritl.lo:utfl flow _B_ypaiss is overstressed due to differential thermal

‘expansion, and this condition alone is sufficient to remove it from consideration. Although
‘the transients have not been specified for the MSBR, it appears that the straight-tube bundle

with flow bypass could not survive the transients because of the low allowable stress at
1300 F.

4-14
e
-

-

 

The stress levels in the other tubes are roughly the same, but the hockey-stick, the
C tube, and the hockey-stick/sine are disadvantageous because they need a vessel with a
significantly larger diameter to accommodate the tubes.

Stress analysis of the tube during transients can be done using standard techniques;
however, the stress analysis of tubes with a C, J, or hockey-stick shape will be difficult at

the bend region due to the uncertain effect of tube support systems.

Thermal-Hydraulic Performance

 

Experience with heat exchangers operating at high temperatures where flow
maldistribution could cause severe temperature maldistribution and high stresses has led to
the development of vertical shell-and-tube heat exchangers. Other advantages of the system
also contribute to the selection of the vertical arrangement and are mentioned elsewhere. As
with the steam generator, a basic criterion is the hot primary fuel salt flowing downward
over the heat transfer surface in countercurrent to the secondary coolant salt flowing
upward. This type of flow arrangement ensures that flow and temperature imbalances,
which may occur at part- flow conditions, will be minimized.

Other design criteria which are related to the thermal and hydraulic performance of
the heat exchanger are as follows:

1) Provision for tube support designs that permit adequate
support against flow- induced vibration and ensure good flow
distribution.

2) Provision for inlet plenums or special baffles to aid in initial

distribution of primary and secondary salt.

To evaluate general trends in unit performance, it is best to start with the
straight-tube bundle because it is geometrically simple. The straight-tube bundle, in a sense,
represents one end of a scale. At the other end we have the helical unit, which represents
almost pure crossflow as opposed to the straight-tube axial flow. The general trends can be
formulated for all other bundles by evaluating these two types of tube bundles.

A general parametric study for the stréight tube is shown in Figure 4.12. Each of the

three regions represents one tube size with ligaments of 0.25 in. and 0.375 in., respectively.

Across the regions are plotted the lines of constant pressure loss for the tube side and the
shell side. To facilitéte the establishment of general trends, some values of salt inventory
have been averaged over small ranges. Obviously, the tube-side salt inventory changes with
bundle diameter; but for the range of the plots they were averaged within 25 percent of

actual values.

4-156
The smaller tube sizes offer large reductions in bundle length while maintaining

~ similar AP values. This condition reduces both shell-side and tube-side volumes significantly. |

Another significant feature is that the minimum bundle diameter is fairly well fixed at about

'55 in. To provide for clearances and shell thickness, the minimum shell OD should be about

65 in. The maximum bundle OD is about 70 in.; this makes a shell OD of about 80 in.

Currently there does not appear to be any reason for exclusion of the small-diameter
tubes (a final economic tradeoff would determine exact values); therefore, bundle lengths
smaller than 30 ft are reasonable, and this would imply an overall unit length of about 40 ft.
The large number of smaller tubes may cause high manufacturing costs plus maintenance

problems; therefore, B&W considered the possibility of 0.5 inch diameter tubes, which
~ would add about 20 ft, thus offering overall lengths of about 60 ft. This unit was later

found to be too long and to add significantly to the primary salt inventory.

Some potential problems with small tubes are supports, buckling, and tube plugging.
First, the number of supports may be determined by the natural vibration frequency
required. Figure 4.13 shows that small-diameter tubes require smaller support lengths, which
increase costs and pressure losses. Second, the buckling characteristics of small tubes (say,
during a thermal transient) usually require very close support spacing. This requirement may
be a tighter restriction than that of the support spacing for natural frequency evaluation.
Third, remote tube-plugging methods have not been developed for the tube ODs less than
0.625 in.

A parametric study for the helical bundle (which is the only arrangement with
significantly different performance from the straight tube) is shown in Figure 4.14. It is
difficult to extract general conclusions due to the complex relationships within the bundle;
however, with experience some generalizations can be made.

The major problem is that the tube-side pressure loss is relatively high even at the

~ large diameters. Smaller tube sizes would decrease the length but not the diameter. The
graph is drawn with the constraints of the minimum distance between tubes in the bundle -

being 0.375 in. and the helix angle being 20 degrees. The diameter of the tube bundle could

be reduced by possibly packing the coils closer. The tube AP could be decreased by adding

more tubes, and the diameter could be decreased by increasing the helix angle Only a more

| complete tradeoff study could dehneate these effects

MethOds bf improving heat”tr'ansfer have not been actively pursued at this design
stage. The use of studs or small fins could improve the heat transfer, but the pressure losses
- could mfluence the results greatly. Since these features can be used for any tube de31gn, it
B _' would seem prudent to pursue the analyms at a later stage |

Manufacturing

Each design has been studied to determine the degree of difficulty in manufectufing
the components. In nearly all cases the problems are the assembly of the tube bundle and

4-16
 

90

  
      

l- 25.”
60

Tube, OD

40
Unsupported Tube Length, in.

 

 

 

60 prmmm

=y -ud;;g;qgj\ jo Aouenboi g [eanjeN A

L
10

10—

| Figure 4.13 , |
Span Length of Tubes Vs Natural Frequency - S
Effective Tube-Bundle Height, f{t

- 50. 00—

110. Op—

100. O~

90. Op—

80, O—

70. 01—

60. Of—

 

=== Shell-Side AP, psi

 

e

175

100

Tube -Side AP, psi

‘,300
g

200
150

 

50

0 gk L
70.0 ~  80.0

Tube, 1.5" OD

Tube -Side Vol= 925 ft?
Shell-Side Vol=1200 ft?

Tube, 1.25" OD

Tube -Side Vol= 750 ft3 _
Shell-Side Vol=1000 ft?

Tube, 1. 0" OD

Tube -Side Vol= 630 ft3

Shell-Side Vol= 870 ft3
100

90.0. 100.0 10,00

Tube Bundle OD, in.

 Figure 4.14
Helical Tube-Bundle IHX, Parametric Chart
o %,

the method of tube support.

One major question has been the smallest possible tube size which can be assembled

| and welded. Currently the smallest tube size used at B&W is 5/8 in.; however, it is felt that

with development this size could be reduced to 3/8 in. A further reduction may be possible
(say, to 1/4 in.) by 1980. Due to the large reduction of unit volume, the units have been
sized on the basis of 3/8 in. OD tubes.

The method of tube support preferred by B&W for the more conventional types of

~tubes is the broached plate or drilled plate shown in Figure 4.15. This arrangement provides

support to each tube and is easily made. The results of preliminary laboratory tests of the
vibration/wear characteristics are highly promising. This tube support method has also been
adopted in PWR steam generators.

In many instances the plate cannot be used; for example, in the assembly of a
sine-wave tube bundle the tubes cannot be supported with plates in the region of the
sine-wave, although they may be used elsewhere along the tube. The sine-wave section may
be supported by wrapping a metallic band around each row of tubes.

There are potential problems with the Ctubes, the hockey-stick tubes, the J-tube

“bends, and the hockey-stick/sine tubes. The C tubes can be pushed into the tubesheets, but

there is no obvious method of support. Currently it appears that the best technique is
placing the tubes in circular spacing and banding each ring of tubes as it is placed. The
J-tubes and the hockey- stick tubes have a bend support problem similar to the sine-wave
tube bundle. The hockey-stick/sine tubes can be assembled by stacking the support plates
on the bottom tubesheet and bending the tubes to fit into the top tubesheet (cantilevered
from the bottom tubesheet). After all tubes are in position, the support plates may be raised

- (possible assisted by vibration) until they are spaced propetly.

Inspection, Maintenance, and Repair

. The impossibility of direct contact maintenance at any time after the reactor has

been in operation, and related requirements of inspection and repair, are major
considerations in developing and selecting the recommended IHX designs. Maintenance can

be performed on the IHX in one of three following ways:

1) The unit or tube bundle can be removed to a hot cell facility
-which has been set up for maintenance on various
components of the reactor plant.

2)  The heat exchanger or tube bundle can be removed to a
decontamination facility; it can be decontaminated and
cleaned, and maintenance can be performed using standard
techniques. |

4-17
r r e r r r 1.t Y. . £ o { B SN

r r. (

    

Drilled Holes
for Flow

Broached

fi Area

0,008"

       

Broached Hole Pattern Drilled Hole Pattern

L

 

 

 

 

 

 

 

 

Tube Si%pport and Flow Baffle 3

 

¥ '

MT T

 

 

 

 

 

 

 

 

 

 

 

. W

. Note: Velocity through plate = ,
twice mean axial velocity,

' Figuredas
Broached or Drilled Plate Baffl_e
 

3) In-place remote maintenance can be performed if access is

provided to the areas which are likely to be inspected or
maintained at some time during th¢ plant design life.

Currently the trend in the light water reactor industries for heat exchangers and

steam generators is not to remove the tube bundle but to provide hand holes and access

space to ensurc repair, maintenance operations, or inspections with minimum exposure to
radiation. The trend for the liquid metal fast breeder reactor (LMFBR) plants (either
already constructed or in the design stage) is removable tube-bundle designs for heat
exchangers and steam generators. This method has been made possible by some inherent
advantages of the LMFBR design because the components are not as thick and massive as

- those used in water reactors. These components are not anywhere near the weight of those
used in the large PWR steam generators; therefore, provision for tube- bundle removal is not

difficult.

The reliability of the first generation plants with high design temperatures, potential
sodium-water reactions, and unproven heat exchanger designs is of concern; however, the
technology required for the MSBR heat exchanger design is an extrapolation which can be
easily made as more heat exchangers are designed and operated in high-temperature
environments. With this expected experience, providing for a removable tube-bundle design
may not be necessary because by 1980 a fixed tube bundle will be a conservative design
arrangement due to technological changes and advancements. The component manufacturer
must realize that prospective purchasers of this equipment may prefer a removable
tube-bundle design, for high reliability and availability can be achieved perhaps with the use
of a spare tube bundle. Current estimates of the cost of outage of a 1000 MW(e) reactor plant
range from $100 000 to $200 000 per day depending on the location of the utility in this

country. These figures are based on the 1971 dollar value; at these rates the cost of a spare

tube bundle may be insignificant to the plant operator.

B&W does not believe that the choice between a removable or nonremovable tube

bundle is clear-cut, and consequently the major goals for the removable and nonremovable -

4-18

bundle designs were considered. The considerations for maintenance were applied to each

design concept during the concept evaluation, and the criteria for judging each design are as

follows:

1) 'PrbfiSidns_'fo'r meeting the requirements of the ASME Code

‘Section XI — In-Service Inspection of Nuclear Reactor

Coolant Systems — although this code presently applies only |

to light water cooled and moderated reactor systems.

2)  Provision for access to flanges, seal welds, or other joints
which must be broken for maintenance on the heat
exchanger.
 

e

~3)  Provision for convenient access to the. tube ends and
tubesheets. | | |
4) Provision for removal of tube bundle with minimum amount. :.

of piping or shell cuts.

4-19

5)  Considerations for access in plenums, downcomers, or other

areas where remote tooling may be utlhzed for maintenance .

or ll’lSpC ction.

6) Consideration of remote tube-bundle mspectwn and pluggmg |
equlpment
7) Provision for arrangement of the tube bundle to ens'ure .

convenient draining of the radioactive primary fluid and also’
the secondary fluid, if necessary.

~ Maintaining a radioactive system has been demonstrated successfully for the Molten
Salt Reactor in the MSRE. During the operation of this reactor it was determined that

remote maintenance was feasible by using fairly simple tooling. The equipment
- manufacturer should work closely with the system designer to accommodate simple tooling
or any specialized fixtures that would be required to perform remote maintenance. In some

cases, however, the method of remote maintenance, utilizing the techniques that will be

- developed from now to 1980, may be limited by some of the design goals for the IHX. In
- each concept developed with primary fluid on the tube side, one major design goal has been

to minimize the tube diameter consistent with low primary salt inventory. The trend of
smaller and smaller diameter tubes necessitates a re-evaluation of the potential of

_remote-tube-plugging techniques which are currently being developed in the industry. A

specific technique is the explosive-tube-plugging method being developed by B&W and other

~ component manufacturers. This method entails the detonation of an explosive plug within - -
_ - the tube to seal off any leaks between the tube and the tubesheet or within the tube itself.
 The detonation causes the tube to weld to the tubesheet or-to the plug, depending on whtch L
. area is bemg sealed off; to accomplish proper. welding there must be a sufficient amount of . -
“explosive charge in the tube. Presently, it is questionable if this technique can be used with -

the small-d1ameter tubes wluch fulfill the des:gn objectlve of nnninuzmg the pnmary salt'

- "mventory

All deszgns bemg conmdered are emnsxoned to prmnde good access to the tubesheets, e

"'_““however, . some demgns are 1nherently ‘more flexible in accommodatmg dlfferent
. maintenance techmques For example the cylmdncal tubesheets which could be utilized to:
_ conserve the amount of - Hastelloy material, are very difficult to reach for pluggmg tubesor
- for performmg any inspections. In other cases it is necessary to require a long reach. of
- temote tooling, which then bends upward to gain access to a lower tubesheet. This approach |
" is more. difficult than an approach where the tubesheet can be reached darectly through a

- seal-welded hand hole.
* For further discussion on the IHX design concepfs; see section 4.3.3,

4.3.3 Concept Evaluation

_.1. " Removable Tube Bundle

The following date pertain to the optimized straight-tube bundle selected by B&W
for this study of the MSBR IHX, and the numbers apply to all concepts (except the
hehcal-tube bundle and the ORNL IHX concept) when the primary is in the tubes. |

Number of tubes 7000
Tube OD, in. - 0.375
- Tube ID, in, 0.259
Pitch (in heat transfer area), in. 0.625
Bundle OD, in.
With no downcomer pipe 55.5 .
With 18-in. downcomer pipe 59.0
Effective heat transfer length, ft 26
Primary AP, psi 125
Secondary AP, psi 75

Some units require 5 to 20 percent mote volume in the tubes because of tube shape,

and this requirement is discussed for each unit. The minimum amount of primary salt

inventory which is required is as follows:

Tubes, fe3 65
Two plenums, ft3 55 to 60
Downcomer (if used for primary

fluid), fe3/fe of ]ength 1.75

The ratings of the five main removable concepts are presented in two ways: first,

a full-51ze sketch follows a list of the advantages, dlsadvantages, and conclustons for each

° concept; second, the evaluations are reduced to letter ratings in a simple table (see Figure
" 4.22). The same process is used in the following section; however, the ratings of the two
| -':_sectlons are not intended to be compared The J-stick and the hockey-stlck shapes are

“included in that section because these are not rernovable types. A concept is selected for -

- each group.

4-20

~ The evaluations should mdlcate that a choxce is not sxmple for either of the twoiu S

| groups, however, after considering all cntena, we reeommend the ORNL IHX: de51gn from
- the removable tube-bundle design group. - _,
N

 

4-21

Removable Straight-Tube IHX With
Flow Bypass {Figures 4.16 and 4.17)

Advantz_lges

1) The design and the construction are simple.
2) The tube thermal stresses are zero at steady state.

3) Easy access to the bottom tubesheet is possible (through tubes or
downcomer) for repair, inspection, or plugging.

Disadvan tages

1) The response of the support cylinder to thermal transients may be slow, even
with the effect of the bypass flow; this condition causes high tube stresses.

2) The location of the primary salt outlet at the top adds SJ.§n1ficant inventory
in the center pipe and outside the return pipe (up to 90 ft extra).

Conclusion

If the extra primary salt inventory is acceptable to the designer of the system, this
concept should not be deleted; although realistically the extra inventory will be very costly
for four units. When information on thermal transients is available, the response of the
support cylinder should be investigated. If the design passes both conditions, it should be
considered further because of the advantages offered.

Removable Sine-Wave Tube IHX
(Figure 4.18) -

Advan tag es

1) Maximum tube stresses occur at the cooler end of the unit where the sine
wave is located and where the a]lowable stresses are higher.

2) Access to the bottom tubesheet is possible through the tubes or the
'downcomer - -

3) This tube shape has been used prevmusly on several 'LMFBR  heat
| ‘exchangers. |
 

 

Primary Salt—e .

Inlet 1

Bypass Flow (Betw
tubesheet support
cylinder and center
pipe; see Figure 4-7)

Secondary Salt® T
QOutlet r

 

26
Effective
Heat
Transfer
Length

 

 

 

 

 

 

/ Primary Salt
Outlet
™y Tubesheet {~ 11" thick)
I
\
—h 7000 Tubes (0.375" OD, 0.259"
: ID, 0.625" pitch)
N

 

Broached Tube Support and
Flow Baffle Plate

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

e———— Shell (65" 1D,
] 2" thick)
\ -
\ H
Er I
N 1 I
A | __——Disc and Doughnut Baffle Arrgt
'k (Unsupported tube span=25")
[ f Shroud (59" D)
Y :
[ Center Pipe (18" D)
'
fi : ,
_ - f 3 Shock Pads
77 | - / (Along circumf)

Secondary Salt

Inlet

- Figure 4.16
Removable Straight-Tube IHX With Flow Bypass
 

 

 

= Primary Bypass
- \

? ' Tubesheet/\

N

/

|_—~Bypass Flow
— Control

A

 

 

.. .t
. .
.+ .s
"y .
.

r
L
=~

_—~Double-Wall Pipe

 

\

U Tubes/ Y '

Z
N

 

 

 

 

 

 

 

 

 

 

 

 

’ LIl NY
A B ;
i = o~ ]
- 3 - | /
e ! § )

 

r
/4
—

=

 

 

 

 

 

 

 

 

| ; A
. . '
N\ ¥
z

 

 

 

 

 

 

 

 

 

"u Note: Bypass maintains average temperature of support cy-
| linder equal to average temperature of tubes. At

steady state bundle operates with zero thermal stress.
Figure 4.17
Bypass Flow — Design Details
 

   
  
  
  
  
   
  
  
  
  

Primary Salt

Primary Salt
Outlet

Inlet

Tubesheet (~ 11" thick)
Bundle (59" D}

7000 Tubes (0.375" OD, 0.259"

e | Secondary Salt
| ecorsary o2 ID, 0. 625" pitch)

Outlet
Broached or Drilled Tube
Support Plate

26'

b Effective
f Heat
Transfer

Length

Downcomer (18" ID)

b= T S o _ . Se_éonda_ry.Salt Inlet

: Figure 4.18
" Removable Sine-Wave Tube IHX
 

 

. .

r_

4.22

Disadvantagcs

1) The downcomer and the return pipe add significant inventory, as in the
removable straight-tube concept.

2) The assembly of tubes in the tube bundle and clamping tubes are
complicated compared to those of the straight-tube bundle.

3) The tube shapes require extra manufacturing time.

Conclusion

The large amount of extra primary salt inventory makes this concept costly, as is the

straight-tube concept.

Removable Helical-Tube
IHX (Figure 4.19)

Advantages

1) The tubes are flexible.

2) The 360 degree flow in the tubes eliminates thermal maldistribution of
shell-side fluid.
3) The large OD tubes make remote explosive tube plugging easier.

Disadvantages

1) The extra tube-bundle length, the center exit pipe, and the return pipe give
this unit the highest primary salt volume requirement of the concepts.

2)  Tube coiling and tube -sfip-portv are complicated inanufactufing processes. |

~ Conclusion

Excessive primary salt inventory makes this unit very undesirable.

ORNL IHX Concept (Figure 4.21)

o Advantgges '

1) The primary salt inventory is a minimum.
 

   
   
  
 
  
 

Primary Salt

Primary Salt
Inlet

QOutlet

Tubesheet {~15" thick)

925 Tubes (1. 0" OD, 0, 788"
ID, min clearance betw

Secondary Salt adjacent tubes=0,375 ")

Outlet

Helix Angle, ¢ =30°

Tube Support
Bar

62.5'
Effective
Bundle
Height

— Shell (74" ID,
2. 25" thick)

 

- Downcomer (18" ID)

 

 

 

 

 

 

 

 

L
g
X Lt Iy 1 1%
i

 

 

 

LAz
.-TSeébndary"Salt Inlet

Figure 4.19 |
" Removable Helical-Tube Bundle IH
 

 

. E::;:;;

 

 

 

Overall Length—38'

 

 

Effective Heat Transfer Length—26'

o

 

 

     
 
   
  
 
  
   
 
   

Primary Salt
Inlet

Inlet Plenum
Diffuser

Sliding Seals (Split ring
or diaphram type)

‘Bundle (59" OD)

 

Removable Shell
(94" OD)

Space (For gas or stag~
nant coolant salt)

Bayonet Downcomer Pipe
(Double -walled with =
insulation) ,

 

9.5' Max. OD

Primary Salt Outlet

Note: See Figures 4-19 and 4-20
for C-tube arrangements. )

4 Figure 4'.20 _
Removable C Tube THX -

 
 

‘?w-r‘m

  
     

- Secondary

Primary Salt Outlet

Salt Inlet

VSe;:'on-dary
Salt Inlet

: H - Priméry,
: Salt Outlet'i

Figure 4.21
ORNL IHX Concept
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Primary Salt Thermal . .
Concept 7 Volume Ratio* Stress Manufacturing | Maintenance
Straight
Tube -
bypass)
Sine-Wave + +
Bend Tube 2.1 B B A
Helical +
Tube 5.5 B B B
\
C-Tube 1.0 A B~ C
ORNL © Lo BY B c
Legend:

A = Very Good
B or B = Good
B~ = Fair
C = Poor :
*Primary salt volume ratio based on minimum volume of 120 ft3.

Figure 4.22
Removable Tube-Bundle IHX Criteria Rating

 
 

 

-

2) The tube thermal stresses are low.,
3) A cylindrical tubesheet at the hot end is thinner than a flat tubesheet would
be. |

Disadvan tages

1) The tube dimensions change with each circular row as in the sine-wave tube.
2) The tube layout is a complicated radial pattern as in the C tube.

3) Tube support and flow baffling in the 6 ft long sine-wave bend region will
require special designs.

4) The secondary salt inlet and outlet are at the same end.
5) The upper tubesheet is not readily accessible due to its cylindrical shape.
Conclusion

 

Even though the disadvantages cause some concern about the manufacturing and
design details of this concept, solutions to these problems appear to be available in the near
future. Because of its minimum primary salt inventory and low thermal tube stresses, this
unit is recommended as a first choice from the removable tube bundle designs. An estimate
of the additional maintenance costs related to the inaccessibility of the upper tubesheet is

 

required.

Removable C Tube IHX
(Figure 4.20)
Advantages
1) Cylindrical .tubesheets. are 't;mch thinnér than flat tubesheets.

"2)  The pnmary salt mvenfofy 1s very close to the minimum amount because

| annular plenums reduce the volume considerably.
3) The tube thermal stress;afiére low.
4y ;The tubes can be close-packed whxie maintaining large hgaments in the

' tubesheets
 

Disadvantages

1) The tube layout and the tube shapes are complicated.

2) Tube support and flow baffles throughout the bundle will be special items.
The use of broached tube- support plates or disc-and-donut baffles does not
seem possible because of the tube shape. Alternative supports, such as
banding circles of tubes, still leave the bend areas unsupported.

3) Without removing the bundle, cylindrical tubesheets and small plenums
make tube-end access difficult for inspection or repair.

Conclusion

The difficulties of remote maintenance, tube support, and tube-bundle assembly
could be overcome with careful design, more expensive manufacturing, and the use of a
shielded cask for tube repair on a removed bundle. If these additional costs are acceptable, a
unit using low primary inventory with flexible tubes is realized.

2. Nonremovable Tube Bundle

As mentioned in the previous section, the evaluation is presented in two ways: first,
a sketch follows a list of advantages, disadvantages and conclusions for each of the six
fixed-bundle concepts; second, the comments are related by letter ratings in a table for each
of the four major criteria (see Figure 4.32). From this evaluation the fixed sine-wave-bend
tube bundle is recommended as the best nonremovable concept. Again, the choice is not
clear-cut, and a change in the emphasis on one or more of the criteria could alter the
selection. For example, if remote maintenance becomes a very important item because of
AEC regulations, the selection could change. When considering tube thermal stress, the
reader should refer to Appendix D, which presents stress graphs of several basic bent-tube
shapes. Since all of the tube shapes could be made flexible enough to exhibit low thermal

stresses, a qualitative decision was made to determine the maximum size of the bundles.

This decision limited the size of the various tube bends and gave an approximate 1nd1catlon

of relative thermal stress levels.

Fixed Sine-Wave-Tube IHX
(Flgure 4.23)

'. Advantgg’ es
1) - The primary salt 'vo'lum'é is a minimum.

2) Tube thermal stresses are low and can be located at the cooler end of the
unit where the allowable stresses are higher.

3) This tube shape has been used on several LMFBR heat exchangers in recent
years.

4-24
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

|
[ SN
|
v
z
-
R;la:)‘:)tfo : Shielding Plug Stepped to Avoid
I Radiation Streaming
£ Removable Plug
3| o
Primary ¢ :
Salt Inlet ™ §
Insulation
Sl
Upper End Guides
S?eliogd:lr¥ %E— | \ Center Cylinder Gas or
a utle J Stagnant Coolant Salt
R 7000 Tubes (0.375" OD,
0.259" ID, 0.625" pitch)
o Bundle (59" OD)
26 '

 

  
  

Impact Flow Baffle_

Secondary
Salt Inlet

Note: This upper plug design can be
applied to the other fixed-bundle
IHX designs. o

 

I
Primary Salt
Qutlet

Figure 4.23 = B
Fixed Sine-Wave Tube IHX With Tubesheet Access
 

4) The tubesheets are relatively accessible for inspection and for remote tube
repair.

 

Disadvantz_xges

{ 1) The tube-bundle assembly is complicated as compared to that of a
. straight-tube bundle.
: [ - 2) Tube supports and flow baffles in the sine-wave bend region require special
. attention to avoid vibration problems. |

] 3) Tube shaping is expensive since a double curvature is used, and the
bl . circumferential bend varies with each tube circle.
4) Maintenance must be performed remotely, but it is no more difficult than
. b . that performed on other fixed-bundle designs.

Conclusion

For a cost of some solvable manufacturing and tubesupport problems, a
] nonremovable unit can be fabricated with minimum primary salt inventory and relatively
. low thermal-tube stresses during steady-state and transient operation. Of all nonremovable
: designs, this concept appears to be the most attractive.

bt e

 

-~ Fixed Helical-Tube-Bundle IHX
L (Figure 4.24)
- Ad g
| vantages

H 1) Thermal stresses are negligible due to small differential expansion between
the shell and the tubes.
H 2) The large tube size simp]ifies:_rem'ote explosive tube pl_uggihg. |
| | 3)  Theflow pattern e].iminates major flow maldistribution possibilities.

4) A very small number of tubes is requlred consequently, tubesheets can be
thinner.

Disadvantages-

1) The prlmary-salt mventory is very high, or about four times that of the |
 straight-tube or sine-wave tube bundles. ' |

2) The tube manufacturmg process is complicated, and the tube supports may
be costly.
 

o e .
BT i 3 i i
i

Emfl" .
et et

el

Primary Salt Inlet

 
 
  
  
  
 
  
  
 

Tubesheet (~15" thick)

925 Tubes (1. 0" OD, 0.788"

ID, min clearance betw
adjacent tubes =0.375")

Secondary Salt - _

Outlet
Helix Angle, ¢ =30°

62.5' Shell (74" ID,

2.25" thick)

Secondai'y Salt
Inlet

Primary Salt
Outlet

Figure 4.24 . |
Fixed Helical-Bundle IHX
 

 

3) The unit length means that more secondary salt volume is required; the
containment ccll must be deep, and transportation will be difficult.

4) Shortening the unit would require a significant increase in the primary salt
inventory, which is already high.

Conclusions

Even if the extra primary salt inventory is acceptable, the length of this unit makes
it highly undesirable. If the tube-bundle effective height is shortened from 62.5 to 35 ft
while maintaining the same primary side pressure drop, the helix angle decreases, and the
primary salt inventory increases from approximately 450 to 750 cu ft.

Hockey-Stick IHX (Figure 4.25)

Advantages

1) Compared to the other bent-tube shapes, the hockey-stick is easy to form.

2) Tube thermal stresses are generally lower than those of the sine-wave or C
shapes, and the maximum stresses are located at the cool end of the IHX.

3) The lower tubesheet is accessible through the tubes or through the primary
salt outlet plenum.

4) Simple broached or drilled tube-support plates may be used in the straight
portion.

Disadvantages

1) The insertion of tubes into the tubesheets presents problems snmlar to those
of the sine-wave.

2) The shell shape at the bottom bend requires many fit-ups and welds.

'3) In the bend region at low Joad operation, uncérta.inties about the flow-baffle
design, the tube-support design, and the means of effective heat transfer
exist. |

4)  Although the hockeyrstick. is similar in shape to the J-tube, it re:quires about
| 15 percent more pnmary salt inventory than the sine-wave or stralght-tube
- designs requlre a o
Conclusion

Due to the shell manufacturing and tube-support problems, this concept appears to
be less desirable than the fixed sine-wave-tube IHX concept.

426
Primary Salt Inlet

     
   

 

l Sfiv.w ] e i
- |
e

A

AN

 

i,

 

r—,—-:r-h
] et
"y

 

Secondary Salt
Qutlet

 

 

5

 

| Broached Tube Support
and Flow Baffle

 

 

 

 

—— 7000 Tubes (0.375" OD,
0.259'" ID, 0, 625" pitch)

 

= Overall
: Length
~ 35!

e, e Y e

. < T < 3 < = < S - s v

 

 

 

 

 

 

<U
\ )

Tube Bundle (55. 5" OD)

 

Shell (65' ID, 0.75'" thick)

 

 

< X e < . ¥ T -

¢———3,5'—»
| ‘_(zo" min)

 

 

 

— -

 

 

 

  
  

 

 

 

 

Tubesheet Access
Port

 

 

 

 

- Pi-imary Salt
Outlet

 

 

 

 

Figure 4.25
Hockey-Stick Tube IHX
 

tm —

|

-

4-27
J Tube IHX (Figures 4.26 and 4.27)

Advantages

1) Inserting tubes into tubesheets and into broached or drilled tube-support
plates is simple. ,

2) The lower tubesheet is accessible for direct but remote tube plugging due to
its inverted orientation.

3) Tube thermal stresses are low.
4) When the short return pipe to the reactor is considered (see Figure 4.27),

this unit offers a minimum primary salt inventory comparable to or less that
that of the sine-wave, depending on the location of the reactor inlet.

Disadvantages

1) Supporting tubes and controlling flow in the bend region are difficult.
2) The complicated shell shape requires many welds and fit-ups.
Conclusion

In most areas the features of this unit are comparable to the best features of several

other concepts; this concept would be comparable to the sine-wave concept except for the
complicated shell shape.

C Tube IHX (Figures 4.28, 4.29 and 4.30)
Advantages
1) The tubes are flexible.

2) Cylindrical tubesheets _c__an. be thinner than flat tubesheets, and the tube-hole T
spacing can be larger than the pitch in the bundle. .-

"3) - The annular plenum shapes help to minimize primary salt inventory.

_ Disadvantaggs

1) The tubes are difficult to support in the bend region.

2) Tube access for repair or inspection is very difficult because of the 360
degree tubesheets.
 

 

 

+ Primary Salt Inlet

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

!
h Secondary Salt
Outlet
3 (e N
. “-ef‘ A I W
H - NN ; - ; ST
- Broached Tube Support \\ /K - fi\_)
i : o o - N
;-i =) and Flow Baffle A :b: ":
b stett (657 ) s~
- ey 2
i View A-A
. el——7000 Tubes (0.375" OD, Eight locations in bend region. Tube
b k '0,259" ID, 0.625'" pitch) clamps attached to tubes only. Clamps
| raise natural frequency of tubes, act
i -3~ Bundle (55.5" OD) as flow baffles, and permit differential
‘ 1 thermal expansion.
Co ! | Primary Salt
H 26" | Cutlet
Effective || Rall Tubesheet Access
Heat i ; Port
P Transfer ||
H Length
Tubesheet (54"
V /' / / /I~ op. 9" thick)
b

Bypass Flow
Spoilers

!

, N ' Secondary Salt
. Overall A u Inlet

Length -

¥ R

 

 

 

- Figure 4.26
J Tube IHX
 

T

-

r
§a

 

Primary

 

 

 

ANANAN

 

Pump

-
3!
X
11!

 

J=Tube
IHX
26’
5" 15!

 

 

 

 

 

 

- Note:

 

 

 

(7~

 

 

 

 

 

 

 

 

 

This arrangement indicates that the J-tube con-
" cept may require less primary salt inventory
than the sine-wave unit of the same size.

N
2
N

<

 

 

Pump

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

React |
Outlet _ RSN
4 -
—~t __"l -
14! Sine-Wave
THX
1 .
React - 26"
Tnlet /\‘\ ‘
| 14!
)
e
5 AN NN
Figure 4.27
Primary Piping Arrangements for Two Fixed
Tube-Bundle IHX Concepts
e

Secondary Salt
Outlet

  

Guides

     

Cylindrical Tubesheet
Primary Salt (~ 25" high, 6" thick)

Inlet

Cylinder for Argon Gas or

Detail Views (See Stagnant Coolant Salt

Figures 4-19 and
4-20)

Top View (See
Figure 4-20)

7000 Tubes (0.375"
OD, 0.259" ID)

 

Primary Salt 1
Outlet

Bottom Supports

 

: Secondary Salt Inlet

Figure 4.28
Fixed C Tube IHX
bty

Circular pitches between tubes in bundle, P, and P;, can be smaller

than pitches between holes drilled in

cylindrical tubesheet, P; and P,.

Lines represent tube centerlines in the bend region.

    

Cylindrical
Tubesheet

 

- tubesheet row.

- Tube Support (May be achieved by

 

2z

Tubes Near Outer Radius (Should be .
spread out prior to placement in tube Py
holes due to large number of tubes) -

Tubes Near Inner Radius (Since inner .
circular tube rows contain fewer tubes,
several rows may be placed in smgle

banding each successive circle at
several levels in bundle..

 

  

  

Involute Bends
{May be used to
increase tube flex-

\% 0 QO ibility, particular-
0N P0 ly in outermost
\ tubes. )

 

 

 

 

 

 

 

 

 

 

..

t
v
: Inner
‘Cylinder
4
Figure 4.29

C-Tube Arrangement, Circular Pitches
map dog ‘Aeary sppung-oqn], aInjoau]
0¢" 231

 

Cylindrical
Tubesheet

Tube Holes (Can be
on triangular, cir-
cular, or square
pitch.) :

 

 

 

 

 

Pl |

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Constant Spacing Between Tubes (Constant spacing is main-
tained except in area’ close to tubesheets. If bundle is a
close-packed design (primary salt on shell side), the spread-
ing of tubes in this area permits reasonably large spaces be-
tween tube holes drilled in tubesheet. )

 

 
rv. -

Conclusion

Major problems are foreseen with manufacturing, tube layout, and maintenance; this

design is not recommended.

ORNL Concept, Fixed-Bundle
Version (Figure 4,31)

The comments are the same as those for the fixed C tube IHX; the conclusions are
not significantly different.

3.  Primary Fuel Salt on Shell Side

From the viewpoint of maintenance, inspection, and repair, it is advantageous to
locate the primary (fuel) salt on the shell side. The surface of the tubesheets will not be
highly radioactive as it is in the previously discussed concepts with primary salt in the tube;
this location of primary salt simplifies shielding during remote maintenance on the tubes
and tubesheets. This section discusses concepts with requirements and the methods used to
achieve this unit at a reasonable cost penalty.

To avoid a primary salt volume penalty (a primary-to-secondary ratio of 1.0
indicates no inventory penalty), the tubes should be closely spaced (see Figure 4.33). A
ligament between tubes of about 0.050 in. would make the shellside volume about equal to

_the tube-side volume. This close spacing immediately introduces two problems — support of

the tubes and thermal- hydraulic aspects.

Currently the only method that is envisioned for tube support is to wrap thin metal
bands (approximately 0.050 in. wide) around the tubes. This method would require that the
tubes be in a circular array as shown in Figure 4.34.

The effect of tube spacing on thermal-hydraulic performance is shown in Flgure
4.35. Placing the fnmary salt on the shell side (line A-B) requires a volume penalty of 20 fi3
(85 minus 65 ft”). The volume in the IHX plenums and pipes is approximately the same
regardless of the locatlon of the primary salt. To achieve the correct tube-side (secondary

salt) pressure drop, probably.?OOO to 8000 tubes are required.

For companson point C is shown this is a typical design with primary on the tube

-~ side and 1/4-1n spacing between tubes The primary salt volume in the bundle region is
- about 65 ft3 as opposed to 85 ft3 of the other concepts. Figure 4.35 is intended to show a
‘rough comparison of tube-side versus shellside concepts. For example, it is assumed that the

designs are straight tubes, that the bundles with primary salt on the tube side have
disc-and-donut baffles, and that the bundles with primary salt on the shell side have no

~ baffles but are banded as shown in Figure 4.34. Obviously, the actual method of support

should be studied before definite conclusxons are made on these trends; however, despite the
small primary salt volume penalty of 20 ft3, the use of close-packed bundles appears highly
promising.

4-28
 

E =+ Secondary Salt
{

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 
 
 
 
  

 

Outlet
0““"‘" Xy —_
Primary Salt '
Inlet ,” \1 m \
= A Y |
17 l
-
J / 4 |
QA ) /) J
O / /
[/
N / /)
1
L/
Gasstfé):::tor \ j A N Sine-Waye Bend
Coolant A " \ Region
Salt fl A ’//
N /I \
" |/
) [ — ~ Tube Support and
‘ Flow Baffle
7000 Tubes (0.375"
26! OD, 0,259" ID)
Bundle (59" OD)
j’_' Secondary Salt

“Inlet

Primary Salt Outlet

Figure 4.31
ORNL Design, Fixed Tube-Bundle Version
 

 

 

 

i

 

 

 

 

 

 
  
   
 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

: Primary Salt Thermal l . :
Concept Volume ratio* Stress Manufacturing § Maintenance
-
Sine-Wave + A
Bend Tube 1.0 A B
P
."
H.‘l’.f:t‘::l 4.5 Bt p* A
S
Hockey-
Stick 1.15 A B A
Tube
J-Tube 1.15 st B A
C-Tube 1,05 A B~ B
ORNL 1.05 st B B
Legend: A= Very Good ~ B~ = Fair
Btor B = Good ' - € = Poor

*Primary salt volume ratio is based on the minimum volume of 120 ft>.

Figure 4.32 |
Fixed (Nonremovable) Tube-Bundle
IHX Criteria Rating
 

.

Volume Ratio of Shell Fluid to Tube Fluid

5.

1.

.0

 

0

0.25"

0.15"

0.05"

 

 

 

 

 

 

-Tube OD, in.

Figure 4.33
Straight-Tube Bundle, Ratio of Shell-Side
~ Volume to Tube-Side Volume
 

Bands (Around tubes to
provide support for
buckling and vibration)

_ Figure 4.34
Close-Packed Tubes on Circular Pitch
 

| .

.

€

Number of Tubes

 

 

 

 

 

 

 

 

 

 

 

12, 000¢
A B C
A Bundle length, ft 25 27 28
Bundle D, in. 48 38 55
11, 000§~ Primary AP, psi 115 115 115
Primary salt, Shell Shell Tube
location side side side
Secondary AP, psi 37 95 50
10, 000 }— Primary bundle 85 85 65
vol, ft3
Baffles No No Disc &
: baffles baffles donut
No. of tubes 11, 400 6,300 7,000
9, 000
Line for Primary
8, 000}~ Salt on Shell Side
7, 000}— Co
Point for Optimized
Bundle With Primary
B | Salt on Tube Side
, 0
6 000 0. 05 , 0.10  0.15 0.20 0.25 0.30

- Tube VSpacing (or ligament), in.

Figure 4.35
Effect of Tube Spacing
 

 

 

. . r.

C

Several methods of close packing the tubes have been devised. Tube holes cannot be
drilled within 0.050 in. in the tubesheets. Figure 4.36 shows a removable C-tube IHX with
the primary salt on the shell side, and the following section lists the advantages,
disadvantages, and conclusions. Spreading the tubes out as they bend to enter the cylindrical
tubesheets achieves a reasonably large ligament between tube holes while keeping the tubes
closely packed in the vertical heat transfer region. The penalties seem to be manufacturing
(long cylinders with tube welds inside) and some excess primary salt volume.

Removable C-Tube With
Remote Repair Feature

 

Advantgges

1) With controlled bypass past the ring seals, the steady-state operating
temperatures of the tubes and tubesheet support cylinder are equal. Stresses
due to differential thermal expansion are nearly zero at steady-state

operation.
2) Manufacturing the IHX is relatively simple except for the tube bundle.
3) Easy and remote access for maintenance to both tubesheets is possible

through the center.

Disadvan tages

1) To achieve low primary salt inventory, the tubes will be placed about 0.050
to 0.080 in. apart. This placement gives support problems; however, one
possible method is to wrap bands around each row of tubes (circular pitch).

2) More tubes are required to meet the AP requirements, say about 8000 as
opposed to 6000 to 7000 for other units,

3) This design will have an increase in primary salt volume of about 20 fie3
relatlve to the demgns with pnmary salt on the tube side. '

4) This C tube arrangement requlres a complicated tube arrangement and has
several variable dimensions.

Conclusion

The major problems are the tube arrangement and supports. If these problems can

berovercome the deSIgn has good possfl:uhtles

A concept for reducmg shel]-stde volume with displacement cylinders or rods in a

straight-tube bundle is shown in Figure 4.37. This concept allows normal tube pitches for

the holes in the tubesheets and can be applied to a sine-wave, J tube, or hockey-stick tube
bundle with some chance that there will be problems when displacing volume in the bend

region.

4-29
 

i

 

 

N

 

 

 

 

 

 

 

 

 

< /
amiia
V] |
Primar}.r_'b rfl} \
Ilet hE
4

 

 

 

 

c

 

Pfimary

 

  
 

—e S A — 'y S 5 3

.._Secondary Salt
Inlet

 

~-———aum=Secondary Salt
Outlet

 

‘Close-Packed or
Swaged C-Tubes

Shell

 

 

 

 

 

 

 

 

 

— Slip Ring Sliding

 

&

Mechanical Seal

 

 

 

 

v L LA

 

 

Salt Outlet

Figure 4.36
Removable C Tube IHX With Remote
Repair Features
 

—

r

  
      
 
 

 

 

/ A\
'A"‘V‘V‘V.

VAYAV. ..V,

Tubesheet, Top View

 

Side Section

 

 

 

 

 

]
T
Ry, ]
5 S

 

 

~

 

   

(W'eld_ed caps and

hnders)

‘Volume Displacement Cylinders

(Not welded to either tubesheet)

. l represent any method of
Solid A -sealing and tying upper
ends of displacement cy-—

 

/] <

 

See Detail (Entrance
plenum is provided
above displacement
cylinders for shell-
side fluid)

 

 

 

 

 

 

 

 

 

 

 

  
 

end clips

 

Lower Tubesheet

(Thin pins, supporting
'displacement cylinders
and resting in holes in
lower tubesheet, create.
exit plenum)

Figure 4.37
Conventional Tube Bundle With Volume
- Displacement Cylinders
 

h

.

These two methods, spreading the tubes into tall, cylindrical tubesheets or using
volume displacers where there are flat tubesheets, locate the primary salt on the shell side in
all of the concepts of sections 1 and 2 with the possible exception of the helical-bundle
concepts. The helical-tube bundle concepts have been rejected because of excessive primary
salt inventory, and locating the primary salt outside the tubes will not alter this conclusion.

4,  Advanced Concepts

 

In some respects the concepts that locate primary salt on the shell side could be
called advanced concepts. There is essentially no experience to which to refer, whereas most
of the concepts in sections 1 and 2 represent types of units that have been bullt or studied
for the LMFBR program.

As mentioned in section 4.2, a method, allowing for differential expansion between

the tubes and the shell, includes a large-diameter, flexible bellows in the shell. Figure 4.38

shows this method used with a straight-tube bundle. The major disadvantage, as discussed, is
the unreliability and maintenance difficulty of large bellows. Since the LMFBR industry still
discourages the use of bellows after more than a decade of development, this concept may
not be attractive even with the additional 1980 technology.

Since minimizing the primary salt inventory is a major design goal, three advanced
designs were sketched; these eliminate part or all of the primary piping. The IHX and pump
are attached closely to each other and to the reactor vessel. In the first concept (Figure
4.39) the return pipe to the reactor is eliminated since the lower tubesheet is built into the
reactor. Also included is a concentric pipe-pump concept. Since the fluid in the pipe-pump
area is usually at 1300 F, and the inner pipe is not welded to the vessel wall, differential
thermal expansion is not expected to be a problem. The large outer pipe is more capable of
withstanding external forces and moments than a single, smaller pipe. Approximately 15
rows of 3/4 in. tubes run 360 degrees around the 20 ft diameter reactor vessel, and four
pumps feed these tubes. The advantages of the hockey-stick-shaped tubes are the low
thermal stresses during operation. The design concept also has the advantage of cooling the
vessel walls. As with other cylindrical tubesheet concepts, this annulus of tubes presents
inspection and repair problems compared to remote maintenance of a flat-plate tubesheet.

- Explosive tube plugging is a p0551b111ty for repalr but removal of the tube bundle is out of

the question.

In the second concépt of the coupled pump and IHX, the pipe ‘that leads from the
pump to the IHX is eliminated. Since some primary salt inventory is saved it would be

- possible to elongate the return and outlet pipes attached to the reactor to give them more
flexibility and lower stresses from external forces and moments. As in the previously
discussed concept, this unit has a cylindrical tubesheet, and this causes tube inspection and

~ tube repair problems (See F:gure 4.40, )

Figure 4.41 shows the third concept of a closecoupled pump with a standard
hockey-stick IHX. A J-tube bundle IHX would also be applicable. The problems foreseen
with this unit are not unique; with such short pipes the differential thermal expansion
during thermal transients is expected to cause high nozzle and pipe loads.

4-30
 

. .

Primary Salt Inlet

 

 

   
 

 

 

Double Concentric Bellows
(Leak detector between)

 

 

Double Split Ring
Mechanical Seals

 

- Secondary Salt

Outlet

 

 

 

 

 

 

 

 

— b . - x T = ~ = s::sv____Broached Tube
Support Baffles

 

 

 

 

 

 

Z
L [ S L L

 

 

 

 

 

Secondary Salt
Inlet

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

Primary Salt Outlet

Figure 4.38
Fixed Straight-Tube Bundle With Bellows
- ( B

| S

 

 

 

 

 

 

 

 

 

 

 

 

i Reactor Core Graphite
. Reflector
h Primary Salt
Outlet /3
¢/
L
: : ’
Lo . "
\ _ _—
, 7 A
- | Lower ’ ¢¢ '
H | Tubesheetp A "’ '
- ‘\\\\‘V// A\‘\\‘ : “ '
| : [
oo e
- 0t
; 7 l Secondary Salt
‘_ Inlet '
Reactor
Vessel
Wall

 

 

 

 

 

 

 

 

 

 

 

 

  
   

 

 

Concentric
o Pipe-Pump
g = J Concept
- e
‘* F 2, h \l 5 h- N .
X g Primary Salt
L Inlet
“

N .
N \ Upper Tubesheet
N
\ .
N
N 1 A Secondary Salt
[ Outlet
V(IR
N
\- o \
q )
VIR
N
N |[HH
\ N
q N

N
\ N
LN

N

N

| ' Top View Detail

 

   
  
 

~15 Circular Rows of Tubes
(Tube OD, 3/4") '

Figure 4.39

Close-Coupled Pump and IHX, Annulus of Tubes

Surrounding Reactor
 

 

 

 

 

 

Primary Salt
Inlet

 

Secondary Salt
Outlet

i

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

- o Secondary Salt4
: Inlet :

. .

 

 

 

 

 

NG Primary Salt Outlet

Figurc 4.40
_ Close-Coupled Pump and IHX,
| Pump Atop IHX
 

 

fi

e

4

Reactor

 

Pump (May be slanted to
allow room for reactor
vessel machinery)

 

 

Reactor

Inlet

o

g et

 

 

r—

Primary Salt

 

Inspection and Repair
Manway

 

- oD Upper Tubesheet
——-tq
|

' ~— — Secondary Salt
| Outlet

 

 

 

 

Lower Tubesheet

Figure 4.41
Close-Coupled Pump and IHX, IHX and Pump
Mounted on Reactor |
 

 

A
o

.o

| S

These concepts generally do offer the advantage of a smaller reactor-pump-IHX
containment area. This means less shielding and, perhaps, reduced plant capital cost. Further
study may show that the savings of primary salt inventory, of Hastelloy, and of plant cost
are not sufficiently great to warrant designing a close-coupled IHX and pump.

Figurc 4.42 depicts an advanced IHX concept with either primary salt on the tube
side, utilizing very small-diameter tubes, or primary salt on the shell side, using a closely
packed tube-bundle arrangement. Figure 4.43 portrays the design which utilizes primary salt
on the shell side. The unit consists of a number of fuel-element-type modular assemblies of
tubes which are attached to tubesheets. These tubes can be small in diameter yet be
manufactured easily in the boiler shops. The modules can be fabricated in a separate shop,
and the assembly operation can be performed by welding the modules in place; the concept
of curved tubesheets and the specially machined weld-access preparation would be used. The
small size of the tubesheets in each module permits small holes to be drilled without drill
drift which affects the design. The tube-to-tubesheets. The small size of the modules makes
it easy to perfrom the tubing operation by inserting the tubesheet on the tube rather than
sliding the tube through two tubesheets — an operation that would eliminate the possibility
of using swaged tubes. Using swaged tubes in this design permits very close packing of the
tubes; this close packing decreases the primary system volume. An exaggerated view of the
bottom of the unit, which shows sine-wave-type flexibility members, indicates the provision
for differential thermal expansion in the design. Another reason for using this design would
be its relatively easy manufacturing of small-diameter tubes to obtain a compact design with
primary salt on the tube side. In this design the primary salt would flow down inside the
tubes and out the bottom, and the secondary salt would enter in the sine-wave section and
flow on the outside of the modules and out the top of the unit.

As stated previously, the major advantage of this design over other units, which use
very small diameter tubes or use closely packed tube bundles, is that it will minimize the
amount of delicate shop labor which is normally performed in the manufacturer’s heavy
vessel shop. A design with small-diameter tubes can be achieved by coordinating the
technology used in fuel element manufacture and the technology used in the boiler shop.

Further evaluation, of course, would have to be conducted to determine whether this design

were a competitive design to the selected reference concepts.

"For an advanced concept that could be applied to any IHX, a short study was
conducted on the use of tubes with an elliptical cross section; circular and elhptlca] tubes
were compared on the basis of equal hydraulic diameter, which is equivalent to comparing
tube sections with equal pressure drop. The elliptical tube (for each calculation, the ellipse
was taken with a major/minor axis ratio of 2) gave 25 percent more surface per unit mass
flow in the tubes. In other words, the elliptical section- tube bundle could have about 25
percent less heat transfer area that the circular tube unit. One disadvantage with the
elliptical tube is that the wall thickness is about 20 percent greater than that of the circular
tube; however, most of the resistance to heat transfer is in the film coefficient, so that the
increased wall thickness has little effect. The elliptical tubes would reduce the volume of the
IHX but could introduce severe manufacturing problems.

4-31

 
 

 

 

a

o

o B

o |

Secondary Salt

Qutlet

Primary Salt
Outlet

  
  
  
  
      
  
    

Primary Salt Inlet
Detail A

Weld Metal

   

Internal

 

 

Fore Weld
LAA—1
(Area ma -
Tube chined f!at
Hole to permat
assembly
placement)
View AA
Exaggerated View
of Bottom
Detail B
0.875"

 

 

 

 

 

Swaged Tubes (To

achieve min liga -

ment of 0.25" or
1 1 less)

Straight Tubes
(With 0.10" liga-
ment and electron
beam weld)

Secondary Salt Inlet

Figure 4.42
Modular Tube Assembly IHX
 

4.3.4 Conclusions and Recommendations

 

The fixed sine-wave tube bundle and the ORNL designs are recommended as
first-choice nonremovable and removable primary heat exchangers, respectively. Further
study of the manufacturing problems will be required before a selection can be made among
the concepts that contain the primary salt in the shell side of the unit. These concepts do
have several attractive maintenance and inspection features.

o "

i‘n | It should be concluded also that four criteria were selected as the bases of evaluation
= for all units. The question of bundle removability was kept separate, and the four criteria
o considered, in order of their importance, were primary inventory, tube thermal stress,
& ~ manufacturing, and maintenance.

L Some conclusion should be made about the desirability of a removable unit versus a
E nonremovable unit. Current trends in the LMFBR program provide for tube-bundle removal
because of certain factors, which are as follows:

i 1) The technology in sodium component design is not advanced to the
stage where highly reliable units are possible.

A 2) It is undesirable to cut primary pipes because the reactor is not
* , drained for maintenance, and because remote maintenance

y techniques would be required when they are not currently
L - | contemplated.
1 3) The component designer has quite a large amount of leeway in design

 

arrangement because there are no requirements for minimizing
primary inventory.

™ There appear to be incentives to utilize a fixed (nonremovable) tube-bundle design
as a recommended reference concept. The incentives are related to system effects; therefore,
the component designer can only consider these in a limited sense. These con51derat10ns are

 

as follows:
i 1) It will probably be necessary to use remote maintenance and repair
- | equipment on other components besides the IHX.
“ | N S 2) Pump maintenance will require the cuttlng of primary pressure
;o boundaries. o
Li — 3)  The reactor can be drained down to the level of the bottom of the

IHX (as a mmlmum)

4)  The removal of a tube' bundle requires the sa:rnertyipe of handling as |
the removal of a complete unit because of fission product plate-out.

 
,“
il

4-33

5) Remote equlpmcnt for cutting, welding, and inspecting is being

developed by ORNL, and the use of this equipment has been
somewhat successful.

 

.- Even though fixed tube bundles are being considered, more advanced remote tools
| may not alter the trend toward tube-bundle removal. Often units have exhibited vibration or
maldistribution problems that could not be predicted analytically in advance. In these cases
removal of the bundle for addition of baffle plates or flow guides becomes a very desirable
feature. In concepts where removability can be achieved without any increase in the
minimum primary salt inventory, this feature becomes even more attractive.

|

At this stage of development of the MSBR, it seems prudent to aim development
toward the fixed-bundle design. By planning to conduct sufficient testing on flow, stress,
and vibration in prototypes and/or models to ensure a trouble-free design, and at the same
time by taking advantage of design development of heat-exchange equipment and analytical
techniques over the next 10 years, B&W recommends the fixed-tube-bundle, sine-wave-tube
[HX as the preferred approach.

-

r.

4.4  Other Components

r

44.1  Coolant Salt Pump

The MSBR design study coolant salt pump is located in the cold leg of the salt

L system for five reasons:
k 1) To minimize the thermal duty (maximum temperature and rate of change of
- temperature) experienced by the pump.
| u 2) To provide the maximum possible suction head for the pump.
, 3) To maximize the coolant salt pump pressurc delivered to the intermediate

heat exchanger. Since it has been established as a design condition that the
tubes of the intermcdiate heat exchanger shall be held in compression by the
coolant salt to control leakage and corrosion, location of the pump at this

5 “ | position reduces the cover gas pressure required during operation by about
- 200 ps:

H | 4) To provide head for the purification of the coolant salt by filtration.

L - ~5)  To allow storage and Iinjectiqn- of rrmake'up salt at the system cold-leg
i H . | temperature. .

,1 The coolant salt pumps are of a standard cenmfugal design utilizing salt-lubricated
b beanngs and either a salt-controlled leakage or gas injection seal. Both salt seals and bearings
v are believed attainable in the time period of the study. Bearings are actually believed
 

o
-

- rtC o e e

ey

attainable today. If problems occur in salt seals, the presence of small quantities of noble gas
(from a gas injection seal) in the salt should not deter the heat-exchanger performance. The
gas would come off in the surge space provided in the rupture disk housings.

A coolant salt cover gas system provides the following functions:
1) Pressurizes the rupture disk housing gas space and the drain tank gas space.

2) Provides the appropriate concentration of BF3 in argon to stabilize the
coolant salt concentration.

3) Transports tritium leaving the coolant salt to an off-gas system.

The coolant salt pumps are driven by fixed-speed pump motors located in the steam
generator cell. Since all leakage is out of the coolant salt system into the primary, the major
24
activity in the coolant salt will be 24Na formed in the intermediate heat exchangers. This
activity is reported by ORNL to be 65 cur1es When distributed through the volume of the
coolant salt, the resulting concentration of 2 #Na is about 8 millicuries/ft3,

The radiation level in the steam generator cell should thus be between one and 100
mR/hr and be relatively accessible. The coolant salt pump will deliver a flow of 15.50 x 106
Ib/hr at a pressure head of 350 psig.

4.4.2 Rupture Disk

A multipurpose rupture disk housing is located at the high point of the loop
adjacent on the steam generator inlet. This housing provides the following functions:

1) To hold a rupture disk that will blow out in the event a tube break in the
steam generator overpressures the system.

2) To provide a pressurlzatlon gas space for regulation of the coolant salt static
pressure head by the cover gas supply system.

'3) To accommodate thcrma]'expansion of the coolant salt by gravity overflow
through a line to coolant salt storage tank.

4) To provide a cover gas — coolant salt interface at the highest coolant salt
temperature in the system to promote diffusion of tritium into the cover gas
and absorption of BF 3 into the salt.

The detailed de51gn of thc disk its housmg and the steam generator cell w1ll be a

Task 11 effort. -

4-34
 

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44.3 Coolant Storage Tank

A coolant storage tank will be provided for each coolant salt loop. The tank will
provide:

1) Storage space for 2500 ft3 of coolant salt when drained from the loop.

2) A chemical treatment volume as the filtered salt is returned to the tank
through the cover gas (argon and BF 3) by means of a sparging ring.

3) A sump for makeup salt being returned to the loop.
4) A receiver for coolant salt additions to the system.

The tank will be built of Hastelloy-N material and be positioned in the steam
generator cell to allow gravity drainage from all coolant salt components.

444  Coolant Salt Filter

A coolant salt filter is provided on a small side stream from the pump discharge. This
filter will provide the following functions:

1) At plant start-up, remove particulate material left from construction.

2) After each maintenance, remove particulates resulting from maintenance
operations and the ingress of moisture and air reacting with the system
materials.

3) During normal operation, remove particulate materials resulting from wear

and leakage.

Flow to the filter is controlled by a freeze valve The flow rate is 1 perccnt of the

_system flow rate.

44,5 De-Mister

A de-mister consisting of loosely packed Hastelloy-N wool in a 4in, plpe section is

_provided on the cover gas vent line. The de-mlster prevents mist created by the sparging flow

from being carried into the off-gas system
4.4.6 Coolant Salt Melt '_I‘ank a

A 25cuft HaStelon-N tank is located on the roof of the steam generator cell.
Additions to the system are made by filling the tank with dry frozen salt and heating the

tank with flame heaters. The salt will flow by gravity to the drain tank after passing through

a plumber’s “U” seal which forms a freeze valve. The freeze valve is also flame-heated.

435
 

 

 

5.0 - STEAM POWER SYSTEM

51  General
The steam power system proposed for use in the MSBR p.lant concept is designed to:

1)  convert the thermal energy released in the multiple MSBR salt
coolant loops into electrical power at a high plant efficiency,

2) to operate as the reactor heat absorption sink under all modes of
reactor plant operation,

3) to permit the hot and cold start-up of the MSBR plant independent
of outSide steam supply, and -

4) to convert rcclalmed plant decay and chemical processmg heat
_energy to electncal power output.

 This system consists of a supcrcntlcal pressure steam cycle with once-through steam
generation, condensmg turbine generators, steam reheatmg and regenerative feedwater

heating. Reactor heat is converted into turbine-generator motive steam at the four steam
| generators and four rcheaters serving each of the four MSBR coolant loops. Plant decay and

chemical processing reaction heat is introduced into the steam power system in the form of

- low-pressure steam for feedwater heating utilization. When operating with 3515 psia and

1000 F main throttle steam, with 1000 F reheat steam, and exhausting to 1% in. Hg abs.,
the system delivers 1000 MW(e) nominal at an overall net plant efficiency of 43.0 percent.

| Except for the steam generator and reheater components, special feedwater and
reheat steam preheating components, and a supercritical auxiliary boiler, which are unique
to the system, the system utilizes designs and equipment normally associated with

conventional turbine-generator type power stations, Design of the proposed steam system
_concept should not limit the feasibility of the molten salt reactor plant concept since the
~_constituent components are either conventional in present-day ‘power station applications, -~ -
‘or are within the extension of present technology With the exception of the steam
generator and reheat components, which are located in high temperature steam cells, the

components of the steam power system are all located within the combmed Feedwatcr and

g 'Turbme-Generator Bulldmg or in the area adjacent to it.

"; - 5 Specml Desxgn Consxderatlons

 

Integratmg the proposed steam power system into the molten sa.lt reactor pla.nt"

T _._'.concept requires providing special means to avoid thermal shock and possible coolant salt
" freezing problems in the system steam generator and steam reheating components resulting -

from  the relatively high operating and liquidus temperatures of the sodium fluorborate
coolant salt. |
!
;

 

 

In order to safely accommodate in the system steam generator and steam reheat
components an average coolant salt operating temperature of about 1000 F, an estimated
maximum allowable tube wall thermal gradient limit of about 300 F, and a coolant salt
freezing temperature of 835 F, the feedwater and cold reheat steam flows to these
components have to be preheated to inlet temperatures much higher than normally used in
conventional steam power cycles. For the purpose of this study, the lower limit of the inlet
feedwater and cold reheat steam temperatures to these components have been selected to be
about 710 F and 650 F, respectively. These temperature limitations reflect the evaluation of
the effects of reduction in plant cycle efficiency with increase in feedwater and cold reheat
inlet temperatures, of turbine extraction limitations, of additional power and/or equipment
size penalties for increased pressure drop and flow rates, and plant cycle start-up preheating
requirements. |

Various methods for providing the necessary feedwater and reheat steam preheating
were reviewed, including- indirect heating with either supercritical-pressure steam or
coolant-salt utilizing high-pressure exchangers, and combination indirect heating and direct
mixing with subcritical-pressure steam utilizing vapor compression in a modified Loeffler
cycle. A rather unconventional approach, initially suggested in ORNL study report No.
4541, in which supercritical-pressure steam is utilized in series for the indirect preheating of
reheat steam and the final stage feedwater heating by direct mixing, was evaluated as the
more efficient and convenient method, and, as such, is selected for application in this study.

~This method dictates the use of supercritical steam in special reheat-preheater exchangers

and feedwater mixing chambers, and special feedwater supply pressure booster pumps, the
design of which, although not conventlonal are fea51ble and within forseeable technological

development.
5.3  Selection of Steam Cy'c-lé“Conditiorns'

In addition to providing a simple and effective means of achieving the necessary
feedwater and reheat steam preheating, the use of a supercritical-pressure steam cycle as

proposed, inherently offers a higher thermal efficiency than obtainable from a

subcritical-pressure cycle, and as such, offers considerably lower electric power production
cost, less fuel processing, less fission product accumulation, and less heat rejection to the
environment. Also, the use of a supercritical-pressure steam results in less mass flow and
provides superior steam-side heat transfer coefficients, such that, the cycle equipment
capital cost is estimated to be not more, and very likely less than the cost of a
subcnncal-pressure system for sumlar service. -

Reheat and nonreheat cycle steam systems were compared, and on the basis of a

thermodynamxc gain in heat rate and a gain in effective turbine efficiency a reheat steam

cycle is recommended for spec1fic application to the MSBR plant concept. Improvement in
turbine efficiency obtainable in the reheat cycle, ‘due to the reduction in moisture losses,
 

results in the substantial reduction in size of the turbine and steam generator units as well as
the salt, main steam, feedwater and condensate piping, cycle heaters and pumps, and cooling
water flow required for the rated duty. In addition, a 3600 rpm tandem-compound type
turbine could be effectively utilized in the cycle.

- These advantages are not offset in the nonreheat cycle by the elimination of the
need for reheater and reheat preheater units, reheater steam and salt piping, and their
respective controls. This is due primarily to the major disadvantage of the nonreheat cycle,
in that the turbine required would have to be of a more costly cross<compound type, with

an 1800 rpm low-pressure section, because of the relatively high moisture content of the

steam in the turbine low-pressure section (18 percent moisture leaving the last stage blades).
Also an expensive moisture separator is required between the turbine intermediate-pressure
and low-pressure sections to minimize moisture and to increase turbine efficiency. In
addition, the turbine building would have to be increased in size to accommodate the larger

nonreheat cycle layout.

" While neither a reheat or nonreheat 1000 MW(e) turbine-generator unit utilizing
superheated high-pressure steam has, to date, been placed into operation, units of this size

are presently under design andfor are within the capability of current technological

knowledge. Based on evaluation of data provided by a leading turbine manufacturer of units
of these types, the differential in estimated cost in favor of a reheat, tandem-compound,
single-shaft turbine generator over the nonreheat, cross-compound type should be in excess
of 4.5 million dollars. In addition to greater cost, the nonreheat cycle requires
approximately 273 additional Btu/kW-hr estimated on a net plant basis or 254 additional
Btu/kW-hr on a net turbine cycle basis. These values are differentially taken from the
reference base value of 7528 BtquW-hr (ASME-62-WA-209).

In order to effectively and e_¢onor_nically prqduce the required power to safely
maintain the essential support systems and equipment in operation upon loss of the main

turbine generator or a failure of the power grid, a dual admission, condensing type, standby

 turbine-generator unit is introduced as part of the concept cycle. This unit is capable of

operating on either high-pressure main steam or intermediate-pressure turbine crossover
steam, and is designed to operate in parallel with the main turbme to export power.

In the event of a main turbme trip and loss of plant load, an ample supply of
high-pressure steam is available, from either decay afterheat removal in the cycle steam

generator or from the auxiliary start-up boiler, to assure adequate standby power generation
‘to maintain reactor operation at a banked, minimum load condition for any desired

operating period. Sufficient afterheat disposal steam is generated to operate the standby

turbine for at least several hours. For longer periods of power loss, the auxiliary start-up
boiler can be fired up and placed on the lme to supply the necessary standby motive steam.
reny

The standby turbine-generator unit is required to always be available and kept at
temperature. However, for the majority of the time the steam power cycle functions
normally off the mains with no need for independent standby power. At such times, the

‘standby turbine generator is efficiently operated at full load on main turbine crossover

steam to contribute, together with the main turbine, to the total plant power output. When
needed for standby protection, the unit is automatically switched over to high-pressure
steam operation, and by means of load rejection devices, all nonessential electrical loads are
dumped. In this manner, the unit is always available for standby operation and at the same
time capable of economically generating a proportionate share of the total plant output,
making full and effective use of the unit.

54  Design Criteria and Assumptions

 

The proposed steam power cycle is designed to comply with the following criteria
basic to the MSBR plant concept:

a) The MSBR plant is intended for baseload operation.

b) With 3515 psia, 1000 F throttle steam and 1000 F reheat steam
. exhausting to 13 in. Hg abs., the rated net plant electrical output is
to be 1000 MW(e).

c) The power cycle is to be cepable' of providing means for reactor heat
| rejection under a]l modes of plant operation. -

d)  The plant is to be capable of hot start-up and indefinite banked load
'~ operation, as well as cold start-up operation independent of outside
- steam supply -

e) Due to use in the plant concept of the secondary coolant-salt as an
~ intermediary heat transport fluid, steam generated by the cycle is not
- radioactive and the steam power cycle portion of the plant may be

| des:gned to conventlonal codes |

- | f) Adequate river water is to be avallable for cyclc c1rcu1atmg coohng
’ ‘water appllcatlon |

_ The 'followin'g design assumptions, wluch are typxcal and reptesedtative of present
| 'practlces and procedures employed in the des.lgn of conventional steam power stations, are
utlhzed in the development of the proposed cycle -

- e) Turbme-generator units are of a slx-flow tandem-compound type
- rated on the basis of fully-loaded low—pressure steam end operation
with valves wide open.

5-4
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Eight stages of feedwater heating utilizing main turbine extraction
steam are used to provide feedwater at about 548 F. Seven stages are
of the closed type arranged in two independent, parallel,
half-capacity circuits which share a single full-capacity deaerating
heater stage in common. The heater units are of conventional

shell-and-U-tube type design selected on a basis of:

1) Five percent pressure drop allowance from turbine extraction
stage to heater inlet.

2) A 0 F feedwater terminal difference -in the three top
extraction heater stages (with the incorporation of
desuperheating sections) and a 5 F feedwater terminal
difference in the remaining closed type heaters.

3) All of the closed type heater drains cascade back to the next
lower pressure heater stage with a 15 F drain cooler
approach, except for the bottom low-pressure closed heater
stage which cascades back to the condenser with a 10 F drain
cooler approach.

Boiler feed and boiler feed booster pump units are to be tandem
driven from a dual admission, steam drive turbine utilizing
low-pressure, crossover extraction steam, or high-pressure main
steam, while exhausting to the main condenser. The deaerator feed
boiler feed booster pump is driven off the front end of the boiler
feed pump turbine driver through an appropriate speed reducer. The
boiler feed pump tiesin is located between the third and fourth

feedwater heater stages from the top in order to keep the

temperature of the feedwater at the suction to the pumps at a low
level of about 350 F. A combined feedwater pressure rise of 125
percent above main throttle steam pressure is to be produced by the
boiler feed and boiler feed booster pumps.

The main condenser is a three cell, single pressure, surface type
designed to limit cooling water temperature rise to 15 F and flow

velocity to about 7 ft per sec with less than a 20 ft head loss.

The maximum allowable total reheat steam pressure drop from

turbine extraction to intercept valve is limited to 10 percent of the
- turbine extraction pressure, i.e., about 60 psi.

55
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Course control of steam generator and reheater outlet steam

temperature is achieved by adjustment in heat input (i.e., coolant salt
pumping rate). Fine control is attained by desuperheating by means
of feedwater injection (attemperation).

Cycle feedwater makeup is to be provided by the demineralization of
raw water, and full flow capacity polishing demineralization is used
to maintain feedwater condensate purity in the cycle.

In order to further define performance, or to simplify the steam power system
operation and start-up analysis, the following developed design assumptions are included:

a)

b)

d)

g

Component steam generator and reheat-preheater tubeside
high-pressure steam pressure drops are limited to 300 psi and 100 psi,
respectively, in order to minimize pressure-booster pump power
consumption and its detrimental effect on plant cycle efficiency.

Minimum allowable flow rate through the cycle turbogenerator and
steam generator components are to be limited to 5 and 10 percent of
the rated flow rate, respectively. At lower flow rates there are
problems in insuring stable operation.

A separate variable 'speed motor-driven type pressure-booster pump
is to . provide final stage feedwater pressurization at each steam
generator unit.

A separate temperature - controlled, supercritical steam/feedwater
mixing chamber and a feedwater pressure boosting pump are to
provide the required feedwater flow at each of the four component
steam generator units. ‘

A separate. supetcriticel steam heated reheat-preheater unit is to serve
each of the four steam reheat units to prov1de the required cold
reheat steam preheatlng

Plant auxliary power consumption is evaluated at about 36.7 MW(e),

and total stray heat loss from the reactor system is estimated at 25

“MW(t).

Electrical power load required durmg standby or emergency

operatlon is evaluated at less than 18 MW(e).

5-6
s

h) A supercritical pressure auxiliary steam pressure system consisting of
a light oil-fired, 750 000 pound per hr, once-through type steam
generator, with motor-driven forced draft fan and auxiliary boiler
feed pump, start-up flash tank and fuel oil supply system, is selected
as the most effective means to provide the required start-up
feedwater and preheating design conditions as well as means for
indefinite standby plant operation.

5.5  Description of MSBR Steam Power System

 

5.5.1  Normal Operation

Basic data for full load conditions in the conceptual design steam system are
summarized in Table 5.1, and a simplified flowsheet is shown in Figure 5.1. Superheated
steam leaves the four once-through type steam generators at about 3600 psia and 1000 F at
a maximum total rate of about 10.5 x 10% Ib/hr. Coolant salt at 1150 F is supplied to the
steam generator units at a controlled rate to hold the steam outlet temperature to within a
few degrees of 1000 F. Individual outlet steam attemperators, or desuperheaters, supplied
with 700 F feedwater, assists in holding the outlet steam temperature to within tolerances.

About 3.28 x 10° Ib/hr of the steam leaving the steam generators is diverted for cold
reheat steam preheating and the last stage of feedwater heating; the remainder enters the
3600 rpm high-pressure turbine throttle valve at 3500 psia and 1000 F. Part of the
high-pressure turbine steam is extracted at a pressure suitable for the final stage of the
regenerative feedwater heating. The remainder of the steam in the high-pressure turbine

-expands to about 600 psia and 549 F before exhausting into cold reheat steam mains

leading to the reheat steam preheater. A portion of this exhaust steam is also used for
feedwater heating in the No 2 stage heaters.

The minimum temperature for the steam entering the reheaters is required to be 650
F. The 549 F high-pressure turbine exhaust steam is therefore preheated, or tempered, in
the shell side of surface type heat exchangers using prime steam at 3600 psia and 1000 F in
the tubes. The high-pressure steam leaves the tubes at about 3500 psia and 860 F and is used
for preheating the feedwater, as described below. The preheated “cold” reheat steam, now
at 650 F, then enters the four reheaters, whlch are supplied with coolant salt at 1150 Fat a
controlled rate to provide 1000 F steam at the exit. The reheated steam is supplied to the

 stop valves of the mtermedmte-pressure turbine section at about 540 psia and 1000 F.

" Part of the intermediate-pressure turbine steam is extracted to heat the No. 3 stage

feedwater heaters of the cycle. The remainder of the steam is exhausted as crossover steam

to the low-pressure turbine, the tandem driven boiler feed pump and boiler feed booster

. pump turbine driver, and to the standby turbine generator. Although these drive turbines

normally operate on crossover steam extracted from the intermediate-pressure turbine

sections, they can also accept 3500 psia main steam during start-up or other times when

extraction steam is not available from the intermediate-pressure turbine. Steam for the No.
4 stage feedwater heaters is also taken from the intermediate-pressure turbine exhaust.

-7
 

Table 5.1

DESIGN PERFORMANCE DATA

 

General Performance Reheat-Steam Preheaters

( o T XY oo . ¢ K. ¥ T . owo o oo ( K

i

 

geaclior h?aatl input to ;:$F)system, 2 MW(t) 3333.4 Number of units
ete CCU’IC‘ output, A € Total du , MW(t
gro§s electr}cfal generation, MW(e) ; 1036.7 Total he::ed ste(arin capacity, Ib/hr
tation auxiliary load, MW(e) 24.5 f heated F
Boiler-feedwater pressure-booster pump load, MW(e) 12.2 Temperature of heated steam,
Boiler-feedwater pump steam-turbine power output, Inlet
MW (mechanical) ~29.5 Outlet
Flow to turbine throttle, Ib/hr 7.18 x 106 Pressure of heated steam, psia
Flow from superheater lb/hr 10.5x 10 Inlet
Cycle efficiency, (1000/2303.4), % 43.4 Outlet
Cycle heat rate, Btu/kWhr 7677 Enthalpy of heated steam, Btu/lb
Net efficiency, station, 1000/2328, % 43.0 Inlet
Net heat rate, Btu/kWhr 7866 Outlet
Steam Generators Total heating ste:fm, 1b/hr
—————— Temperature of heating steam, F
Number of units 4 Inlet
Total duty, MW(t) - 1958 Outlet
Total steam capacity, Ib/hr 10.5 x 105 Pressure of heating steam, psia
Temperature of inlet feedwater, F 709.1 Inlet
Enthalpy of inlet feedwater, Btu/lb 785.3 Outlet
Pressure of inlet feedwater, psia 3950
Temperature of outlet steam, F 1006.3 Boiler-Feedwater Pumps
Pressure of outlet steam, psia 3650 -
Enthalpy of outlet steam, Bru/lb 1421.7 Number of units
Temperature of inlet coolant salt, F 1150 Centrifugal pump
Temperature of outlet coolant salt, F 850 Number of stages
Average specific heat of coolant salt, Btu/ib1 { F)'1 0.36 Feedwater flow rate, total, lb/hr
Required capacity, gpm (each)
Steam Reheaters Head, approximate, ft
Water inlet temperature, F
Number of units 4 Water inlet enthalpy, Btu/ib
Total duty, MW(t) 345.5
Total steam capacity, lb/hr 6.0 x 109 Steam-turbine drive
Temperature of inlet steam, F ~650 Power required at rated flow, MW (each)
Pressure of inlet steam, psia 580 Power, nominal, hp (each)
Enthalpy of inlet steam, Btu/lb 13226 Throttle steam conditions, psia/ F
Temperature of outlet steam, F 1001.0 Throttle flow, Ib/hr (each) extraction
Pressure of outlet steam, psia 560 Exhaust pressure, approximate, in. Hg abs
Enthalpy of outlet steam, Btu/lb 1519.1
Temperature of inlet coolant salt, F 1150 Boiler-Feedwater Pressure-Booster Pumps
Temperature of outlet coolant salt, F 850
Average specific heat of coolant salt, Bru/lb'l (Fy1 0.36 Number of units

Centrifugal pump
Feedwater flow rate, total, Ib/hr
Required capacity, gpm (each)
Head, approximate, ft :
Water inlet temperature, F
Water inlet pressure, psia
Water inlet specific volume, 3 f1b
Water outlet temperature, F
Electric-motor drive

Power required at rated flow, MW(e) (each)

Power, nominal, hp {each)

5-8

4
119.6
6.0 x 106

~547
650

~590
580

1254.6
1322.6
3.32x 108

1000
~860

3600

~3525

7.18 x 10°
8460
3360
347.8
319.5

14.7
20 000
3500/140
307 405
1%

4

10.5 x 106
8350
1250
699.9
3450
0.031
709.1

3864
4000

 

2 Does not include 25 MW(t) heat losses from reactor system.

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Simplified MSBR Steam System Flowsheet

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Each of the six ]ow-preSSure turbine cylinders has three extraction points for the
remaining stages of regenerative feedwater heating. About 4 x 106 Ib/hr is finally exhausted
from the low-pressure turbines into three surface condensers operating at about 14 in. Hg
abs. Hot well pumps circulate the 92 F condensate through full-flow demineralizers for the
condensate polishing necessary to obtain the high purity water requlred in a once-through
steam generation. The feedwater flow then splits into two parallel paths for successive stages
of regenerative feedwater heating and- deaeration. Booster pumps taking suction from the
cycle deaerator circulate the water through feedwater heater stage No. 4 and to the two
main boiler feed pumps. Each is driven in tandem with the boiler feed booster pump by a
steam turbine with a brake horsepower capacity of 19 400 hp.

The feedwater, now at a pressure in excess of 3600 psia, flows through the three top
regenerative heaters and leaves at 3500 psia and 548 F. Each of the 3.59 x 10° Ib/hr
parallel-flow streams then enters a mixing chamber where the steam, at approximately 3500
psia and 860 F, from the tube side of the reheat steam preheater is mixed directly with it.
The resulting mixture, actually compressed water at about 3475 psia and 700 F, then enters
the boiler feedwater pressure-booster pumps. They are also shown as motor-driven pumps,
but optimization studies would be likely to indicate an advantage for steam-turbine drives
for some of the units. The feedwater, now at about 3990 psia and 710 F, is returned to the
steam generator at a rate adjusted to the plant load by controlling the pumping rate.

5.5.2 Standbz Operation

Upon main turbine tripout, or during prolonged steam cycle operation in a banked
load condition, a standby cycle operation exists in which the main turbine is down, the
standby turbine-generator unit is operating to provide necessary power to maintain the vital
plant standby functions, and a minimum feedwater flow is maintained through the cycle
steam generator units thus permitting the reactor to be shut down (standby operation with
the reactor at low power is also possible without these additional steps).

In this mode, the steam system is initially operated on steam generated from reactor
decay heat until the warmed-up auxxhary steam generator unit can be brought on stream in

the following manner:

a) A cycle is established in which main steam is utilized to drive
the condensing type standby turbine-generator and boiler
feed pump drive turbine units to maintain feedwater flow
from the main- condenser, through the main steam heated
fecdwater mixing chambers, to the steam generator units.

b) Maln steam output not utilized for standby turblne-generator
and boiler feed pump operation, for feedwater heating in the -

mixing chambers, or for maintaining the condenser vacuum
seal, is desuperheated and rejected to the main condenser, or
if necessary may be rejected to the atmosphere.
o

c) Sufficient feedwater flow is then diverted through the
auxiliary steam generator to permit commencement of
burner firing with the consequent raising of boiler outlet
pressure and temperature.

d) When the auxiliary steam generator outlet conditions match
the main steam conditions the output flow replaces the main
steam outlet flow from the reheat-preheaters to the mixing
chambers serving cycle steam generator units. As the steam
output from the cycle steam generator units decays with the
decay in reactor afterheat generation, output steam from the
auxiliary steam generator is phased in to make up the desired
standby operation steam demand. This is achieved by
restricting feedwater flow to the cycle steam generators, by
means of the inline feedwater flow control valves serving
these generators, while proportionately increasing feedwater
flow through the auxiliary steam generator to maintain a
constant total standby steam output.

e) When the restricted feedwater flow can no longer be
effectively converted to main steam in the cycle steam
generator units, the full feedwater flow is diverted through
the auxiliary steam generator, the cycle pressure booster
pumps are bypassed, main steam flow bled through the
reheat-preheaters is diverted to the main condenser, and
sufficient output main steam flow is diverted, via the mixing
chambers, through the cycle steam generators and
reheat-preheater units to keep them at about 1000 F, in
order to minimized any thermal stress problems. Thus, the
auxiliary steam generator output is on stream and the cycle
steam generator capability is kept in a banked condition
ready for hot restart operation.

f) Adequate steam flow may be bypassed through the
coldreheat circuit and the extraction side of the regenerative
feedwater heaters and exhausted to the main condenser to
maintain them sufficiently warm to preclude thermal stress
problems on turbine restart.

5.5.3  Start-upand Shutdown Operations

The steam powef system start-up cycle_ depicted in simplified form on Figure 5.2 is
intended to supply the necessary system warm-up, deaeration and motive steam for MSBR

cold start-up, hot “black” restart and shutdown operations. This cycle is somewhat different

from cycles of similar function in conventional fossil-fired, supercritical-pressure steam
stations in that it utilizes a supercritical-pressure type auxiliary steam circuit to provide the

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required start-up steam. In order to avoid coolant-salt freezing and excessive thermal
gradient problems in the system steam generator and reheater units during start-up,
temperatures and pressures in these units should be maintained as close as possible to full
load operating conditions before any steam loading is attempted in these units. The
supercritical-pressure auxiliary steam circuit achieves this aim by providing a means to attain
at these units a 1000 F inlet fluid temperature, at pressure, during zero power operation, in
addition to providing a means to control the minimum inlet feedwater temperature to the
steam generator units to about 700 F, and the minimum inlet cold reheat steam temperature
to about 650 F, during steam loading in the power range.

This start-up cycle basically consists of an independently fired auxiliary
super-critical steam generator unit, with its associated start-up flash tank, pressure letdown
and regulation valves, auxiliary boiler feed pumps, FD fan and support systems, connected
to operate in parallel with the steam system and capable of supplying system start-up steam
to approximately 10 percent of rated plant load.

5.5.4  Cold or Initial Start-up Operation

The basic steps in the cold start-up operation of the MSBR plant are listed below in
which it is assumed all systems arc in a cold and empty condition and no outside source of
start-up steam is available.

Salt Systems Steam Power Systems

 

 

feedwater  circuit  from
condensate storage. Cycle feedwater
through demineralizers for initial cleanup.
Fire auxiliary steam generator to preheat
and deaerate feedwater.

Primary and secondary cells are Charge
electrically heated and helium circulated in
salt systems by primary and secondary
pumps. When at temperatures above
liquidus, fill salt loops and circulate
isothermally at flow rates required for zero
power mode operation.

~ Make reactor critical at zero power.

Power controlled by automatic neutron
flux level controller to maintain fuel salt .

temperature at 1050 F. Increase reactor

~power. Increase reactor AT. Increase
coolant salt flow rate to match feedwater

heat extraction, Increase reactor power.

Warm-up steam lines, turbines and
feedwater heaters while bypassing steam

- generator and reheater units. Roll and
- synchronize turbine-generator units and
‘roll ~ BFP  turbine - drive.  Load

turbine-generator units to minimum loads,

Fire auxiliary steam geneator to rated

output conditions. Preheat steam generator

and reheater units. Load BFP turbine and

‘phase out motor driven auxiliary BFP.

Reduce  feedwater  temperature by

tempering with auxiliary steam to about

700 F. Place pressure booster pumps in
operation and increase reactor output to
maintain output steam conditions and

5-11
 

-

At about 8 percent of full load, reactor
is placed in a temperature control mode in
which reactor outlet temperature set point

minimum load. Divert main steam flow
(from reheat preheater) to phase out
auxiliary steam generator output utilized
for feedwater heating. Take auxiliary steam
generator off the line.’

System is selfsupporting at about 8
percent load. At about 20 percent load
steam temperature control is activated in

is adjusted for subsequent load-following which reactor outlet temperature is

control. regulated as a function of load while the
steam temperature controller holds steam
temperature at 1000 F,

On a cold or initial start-up, makeup feedwater is introduced into the steam cycle at
the main condensor hot well from the site condensate storage facilities via condensate
transfer pumps. Hot well condensate is pumped in sequence through polishing
demineralizers, the lower 3 stages of regenerative feedwater heaters to the cycle deaerating
heater, where level is maintained and from which all feedwater flow originates.

A pre-firing feedwater flushing cycle is then established in which feedwater is
circulated in series, by means of the auxiliary boiler feed pumps, through the intermediate
and high-pressure regenerative feedwater heaters, the auxiliary steam generator, the start-up
flash tank, the main condenser hot well, the polishing demineralizers, the low-pressure
feedwater heater units and back to the deaerator supplying the suction to the pump units.
In this manner, a minimum feedwater cycle flow of about 10 to 30 percent of the rated
auxiliary steam generator full flow is continually circulated until the concentration of solids
has been reduced by demineralization to acceptable limits to provide initial feedwater cycle
cleanup.

At this point in the start-up cycle operation, the superheating section of the
auxiliary steam generator is bypassed so that the feedwater flows through the unfired
furnace section of the steam generator, through a pressure-regulating boiler extraction valve
station to a start-up flash tank. Pressure in the steam generator furnace section is controlled
by regulation of the set pressure of the boiler extraction valve station. Feedwater level is set
and maintained in the flash tank by means of a level control valve discharging to the main
condenser.

~ Upon completion of the preliminary cycle cleanup operation, an initial firing (black
start) cycle, which functions to bootstrap the feedwater temperature to suitable auxiliary
steam generator inlet and deaerator pegging values, is established as follows:

a) ~ While maintaining a minimum feedwater circulation rate, the
flushing cycle is modified so that the makeup flow to the
deaerator is blocked off and the flash tank drain flow. is
diverted from the condenser to the deaerator, thereby,
temporarily isolating the condenser from the cycle.

5-12
 

E
-

”~
-

-

c)
d)

The deaerator pressure is then set so that the corresponding
saturation temperature is commensurate with the minimum
allowable feedwater inlet temperature recommended for the
auxiliary steam generator (20 psia). At this point, cooling
water circulation through the condenser is established, and
the mechanical-atomizing start-up burner, or burners, is fired
at a low load operation sufficient to heat up the cycling
feedwater to the deaerator saturation temperature (~ 260 F),

When proper inlet feedwater temperature is achieved, the
condenser is reconnected to the cycle by directing the flash
tank drain flow to both the condenser and the deaerator and

‘readmitting condensate flow to the deaerator from the

condenser, thus completing the feedwater cycle. Control
valves in the drain lines to the condenser and deaerator
function to control flash tank level and maintain deaerator
set pressure respectively.

The firing rate is then increased, while simultaneously
increasing the boiler extraction valve station set pressure, to
raise the steam generator outlet heat content to permit the
controlled increase in pressure and temperature at the flash
tank. Any flash tank steam generated is admitted to the
deaerator.

With the auxiliary steam generator pressure set at about 3550 psi and with increasing
L flash tank pressure and temperature, deaerator pegging and steam cycle warm-up operations

can be initiated as follows:

U a)

Flash tank drain flow is diverted from the deaerator to the
extraction inlet side of the No. 2 stage high-pressure

feedwater heaters while the flash tank level control flow is

maintained to the condenser. The deaerator is then pegged at

“.about 30 psi with the drain flow from the No.2 stage heaters.

Steam -generated within the flash tank flows through the

~ auxiliary steam generator superheater section, cooling the
- superheater, and is utilized to warmup the steam lines,
~ turbine units and the tubeside of the cycle reheat-preheater .
~ units, and to seal the turbines to permit establishment of

condenser vacuum. Warm-up steam, however, is bypassed
about the cycle steam cells. '

As the flash tank pressure increases, steam is also admitted to

the extraction side of the No. 1 high-pressure feedwater
heater units completing feedwater heater unit warm-up.

5-13
e

ST T ( B S

5-14

d) Any excess flash tank steam is dumped to the condenser.

e) At this time, the coolant salt circulating system should be up
to operating temperature.

By increasing the feedwater flow rate to about 10 percent of turbine full load flow
and increasing the firing rate to raise the flash tank to its set point pressure of about 1000
psi, the warmed-up main and standby turbine-generatcr units may be rolled, synchronized
and initially loaded to at least 5 percent of their respective full load flow rating with flash

r

r-

i tank steam. Also, at this time, the boiler feed pump drive turbine may likewise be rolled and
- loaded sufficiently to permit establishment of a minimum boiler feed pump recirculation
-t flow about the deaerator.

-

Turbine extraction steam generated is used for extraction heater and deaerator
heating replacing flash tank steam as it becomes available. Reheat temperature to main
turbine is established and controlled by bleeding warm-up steam through the tubeside of the
reheat-preheater unit sufficient to heat the cold reheat extraction steam flow until the
reheater units can be brought on line.

Auxiliary steam generator unit outlet flow is brought to rated output conditions of

b pressure and temperature by directly pressurizing the superheat section above 1000 F from

the boiler section and increasing the firing rate to provide load at 1000 F rated output. The

flash tank is automatically isolated from the start-up cycle by the boiler extraction valve
H station upon increase in steam pressure above the 1000 psi set pressure.

L When the auxiliary steam system is at rated operating conditions, the standby steam
Y generator is brought up to full load operation and the boiler feed pump drive turbine units
| are brought up to matching load with the motor driven auxiliary boiler feed pump output
and feedwater circulation is transferred to the drive turbines while the motor driven units

U are taken off the line.

- Bypasses around the steam cells can now be closed and 1000 F preheating steam
: v - admitted to the system steam generator units. Inlet steam flow temperature to the reheater
| ‘units can be regulated to about 1000 F by indirect heating with auxiliary steam diverted
through the system reheat-preheater units servmg the reheaters.

~ With the steam cycle at pressute' and temperature and operating at low load, the
1 reactor is made critical. Some feedwater flow at the auxiliary boiler inlet is diverted to the
b - mixing chambers, by means of the inline control valve, to gradually reduce the steam
temperature to steam generator units to about 710 F. Simultaneously, the reactor load is
~ increased to maintain steam generator output temperature at 1000 F, and the boiler firing
| - ~ rate is reduced to match the resulting reduction in flow through the boiler. The reactor flux
| controller set point is regulated to keep salt temperatures at desired levels as the reactor
power is matched to the load.
 

 

 

" By diverting main steam flow from the tubeside outlet of the reheater-preheater
units to the system mixing chambers, while simultaneously phasing out auxiliary steam
generator output flow to the chambers, reactor generated main steam is substituted for
auxiliary steam generator output and the auxiliary boiler can be taken off the line. Auxiliary
steam generator loading and firing must be controlled during the phasing out of the steam
flow to the mixing chambers until switchover is completed and the auxiliary generator can
be shut down. Required steam cycle heat input is not completely supplied from reactor
output.

With pressure booster pumps in operation, auxiliary steam generator unit off the line
and cycle at pressure and temperature, reactor is matched up with subsequent increase in
steam loading to full load operation and the system is self-supporting. When the load reaches
about 8 percent of full load flow, the reactor can be switched from flux control mode to a
reactor outlet temperature control mode upon matching the set point temperature with the
existing outlet temperature. Steam temperature control is initiated at about 20 percent load
operation in which the reactor power control is regulated as a function of load while the
main steam temperature controller holds outlet steam temperature at 1000 F.

Extraction steam from the main turbine will maintain the cycle feedwater
temperatures and the flash tank is kept warm through interconnection with the deaerator.

5.5.5  Hot Restart Operation

 

Hot restart operation, as applied to the MSBR plant, consists of the ability to place
the steam power system on the line from a standby or banked operating condition without
the assistance of any outside power. Prerequisites for commencement of the hot restart
operation are, therefore, a steam system operating at standby power level with either
auxiliary steam at standby operating conditions, or the reactor steam at standby power
conditions and all components of the steam warmed-up ready to accept loading. All
necessary cycle feedwater, condensate and cooling water circulation is provided by power
from the cycle steam or standby turbine-generator unit.

If the plant is on standby turbine-generator power, then the main turbine should be
carefully preheated with system steam regulated to match the turbine bowl temperature
limitation set by the turbine manufacturer (about 950 F). Turbine stop valve and main
steam line drain valves should be opened for warm-up. Output from the auxiliary steam

- generator is sufficient to permit rolling, synchronization and loading of the main turbine for
“minimum load operation. At this time, steam flow from the auxiliary steam generator is

replaced by steam from the cycle steam generator in the same manner during a normal cold
start-up operation. Reactor power is matched to steam loading as the turbine is brought up
to full power. The auxlllary steam generator can be kept warm and placed on standby
service until stable operation is acl'ueved

5-15
 

€«

5.5.6  Normal Shutdown Operation

 

Essentially, the normal shutdown operation of the MSBR steam system consists of
reversing the start-up operation sequences until flow to the main turbine has been reduced
to zero, at which time the system is placed on standby operation and held until afterheat
rejection has decayed to a very low value, then the reverse start-up operations are continued
until steam generation is terminated and the system is allowed to cool.

As the main turbine flow is gradually reduced to zero, the reactor control is
transferred from output temperature and load control to flux control and the power output
reduced to zero. The standby turbine generator and boiler feed pump drive turbines are
switched to main steam operation on the reduction in extraction steam pressure and output
accompanying the reduction in main turbine loading. Generated steam not required for
reduced load, standby turbme and boiler feed pump drive turbine operation is rejected to
the condenser.

Operating the system in the standby mode, with the auxiliary steam generator unit,
standby turbine generator and boiler feed pump operating, assures maintenance of minimum
system power requirements and continued feedwater circulation for afterheat removal and
maintenance of desired salt temperature profiles. While operating on standby the reactor is
shut down and the main turbine is allowed to cool gradually by controlled admission of
steam through the turbine seals and warm-up system.

The auxiliary steam generator is taken off the line in a manner reverse to start-up. As
output steam pressure is reduced to 1000 psi steam is generated in the flash tank and steam
temperature and burner firing rate are controlled to maintain acceptable furnace gas
temperatures as load is reduced. The auxiliary steam generator may be allowed to cool or
intermittently fired and maintained warm ready for restart operation. Likewise, sufficient
enthalpy is maintained stored in the flash tank by bleed steam to keep it in a standby
condition to preclude the need for feedwater temperature bootstrapping operations of a
cold start-up.

 

5.5.7  Waste Heat Rej'ecti'on and Recovery System Operation

'In this plant concept, décdy heat rejected from the fuel salt drain tank cooling
system and the heat from the chemical processing cooling system are intended to be rejected

“to a low-pressure steam circuit which, in turn, may either reject this combined heat directly

to the atmosphere during emergency-dump operation, or into the plant steam power cycle,
as low-pressure steam, during normal plant operation.

Radiant boxler units convert : the heat from the individual NaK cooling circuits into
low-pressure steam. About 18 MW(t) of fuel salt decay heat and about 7 MW(t) of chemical
processing waste heat is converted continuously during normal plant operation. During
emergency-dump operation the combined waste heat converted will increase to as much.as
53 MW(t).

5-16
 

 

The schematic flow diagram, shown on Figure 5.3, illustrates this closed circuit,
low-pressure, waste heat rejection and recovery system which is connected in parallel to
both a water filled basin open to the atmosphere, and the steam plant feedwater
cycle.

During the normal plant operation, the low-pressure steam generated from the waste
heat is directly introduced into the plant steam system deaerating feedwater heater stage,
thereby, reducing the amount of turbine extraction steam needed for regenerative feedwater
heating. Level controllers serving the radiant boiler units regulate condensate return flow
from the plant condenser to maintain the required circuit flow. In this mode of operation,
the pressure of the waste heat steam generated is determined by the operating pressure of
the deaerating heater, changing as the extraction pressure changes with plant load. At this
time, a pressure activated control valve, set to open at pressures exceeding the deaerator
pressure at rated load, isolates the waste heat steam circuit from its eemergency-dump
circuit.

During emergency-dump operation, the steam cycle extraction steam demand
cannot absorb the increased waste heatsteam release, thus, causing the circuit pressure to
rapidly increase. When the control valve set pressure is exceeded, the valve automatically
opens to divert waste heat steam flow to the emergency-dump circuit where it is condensed,
stored and the condensate pumped back to maintain proper circuit flow. Thus, the waste
heat is rejected to the atmosphere by the evaporation of the basin water used to condense
the steam in the coils within the dump basin. Although cooling water regulated by level is

normally used to provide the dump basin makeup requirements, the 30 ft x 20 ft x 15 ft

dump basin proposed contains enough stored water to provide evaporative cooling for at
least three hours at rated emergency-dump heat input with no makeup. Any auxiliary
makeup source, such as a river feed fire hose, capable of at least 150 gpm flow is sufficient
to provide for indefinite operation of the dump basin, in the event of loss of normal
makeup, under these conditions.

5.6  Performance

Figure 5. 4 indicates the schematic heat balance, with major flow, turbine

“characteristics and cycle heat rate values, for the proposed reheat type steam power cycle at
~ rate load operation. A gross cycle generating capacity of about 1036.7 MW(e) (proportioned

on the basis of 1018.7 MW(e) from the main turbine-generator unit and about 18 MW(e)
from the standby turbine-generator unit) provides the desired 1000 MW(e) net plant power
output and about 36.7 MW(e) for estimated plant auxiliary electric loads. About 12.2
MW(e) of this auxiliary plant load is estimated to be required to drive the special pressure
booster pump units feeding the cycle steam generators. The equivalent thermal energy input

" needed from the plant reactor to deliver this output, with the cycle as shown, is about
- 2303.4 MW(t) which results in a net cycle thermal effic1ency of 43.4 percent and a net cycle
~ heat rate of 7677 BtulkW-hr. |

Although dependent on the nature of the final plant design, the estimated heat
losses from the reactor plant, exclusive of decay and chemical processing heat, should be in

5-17
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Waste Heat Rejection and Recovery System Flow Diagram
 

Steam Cycle - Heat Balance

 

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the order of 1 percent of the total reactor output. For the purpose of this study, this reactor
plant heat loss is estimated at about 25 MW(t), thereby, making the total required energy
output of the reactor about 2328 MW(t). This results in an overall net station thermal
efficiency of about 43 percent and a station net heat rate of 7866 Btu/kW-hr for- the

proposed cycle without the utilization of any plant decay or chemical processing heat which

has not been included as part of the thermal capacity of the reactor.

Among the factors influencing the stecam power cycle performance, the introduction
of reclaimed reactor decay heat and chemical processing heat into the cycle represents the

most attractive means for material improvement of cycle capability. The schematic heat

balance, shown in Figure 5.5, indicates the effects to the proposed cycle obtainable from

the method suggested in this concept for the introduction of the 18 MW(t) reactor decay

heat and the 7 MW(t) chemical processing heat into the cycle as an addmona] source of
low-pressure feedwater heating steam.

As shown, this method of cycle enhancement is capable of increasing the obtainable
gross generating capacity to 1042.4 Btu/kW-hr without any additional increase in reactor
power output. Assuming the allowances for auxiliary plant load and reactor plant heat load

remain unchanged, the net plant power output capability is increased to 1005.7 MW({e) by .

this enhancement. This results in a net cycle thermal efficiency of 43.7 percent and a net
cycle heat rate of 7635 Btu/kW-hr, or an overall station thermal efficiency of 43.2 percent
and a station heat rate of 7625 Btu/kW-hr.

Further improvement of the cycle performance can be expected from optimization
of the special pressure booster pump power requirements and the utilization of steam
turbine drivers in lieu of variable speed motor drivers used in this study for these units. Less
significant changes to cycle performance as indicated are possible from trade-off studies to
determine the most economical feedwater cycle type and number of feedwater stages,
arrangement of boiler feedwater pumping, and in cycle, extraction steam line pressure drop
allowance, etc. | -

5-18

 
— - .,

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6.0 BUILDINGS AND STRUCTURES

6.1 Site

The site defined by contract for the design study is the Atomic Enérgy
Commission’s standard hypothetical site as specified in NUS-531 - Appendix A, dated

‘January 1969. This site is a level grass covered riverbank 15 ft above mean river level, The
‘soil has a depth of 8 ft and is underlaid by a limestone strata 30 ft thick. The limestone has -

a bearing capability of 18 kp/ft2. The river has adequate cooling water to maintain 1% in.
Hg back pressure in the condenser. A plot plan of the site is presented in Figure 6.1. The site

“is served by a single source of transmission which is subject to occasional outages. This

requires the site to have redundant emergency power sources. Tornado frequency is -
specified so as to require Class I structural design. The seismic criteria for the Task I design
effort was specified as 0.07g horizontal ground acceleration. The site imposes no special
environmental criteria on the plant design other than the normal licensing requirements.

6.2 Build.inE

Each of the principal buildings is situated on a common concrete mat with sufficient
space between the walls of each building to allow for seismic displacement without
disruptive contact between the buildings. With this arrangement relative displacement
between the buildings would not threaten the integrity of the reactor building or of the
steam generator building.

Plan and elevation layout drawings of the reactor building are shown in Figures
6.2-6.7. Plan and elevation layout drawings of the steam generator building and turbine
building are shown in Figures 6.8-6.13.

6.2.1  Classification of Structures, Systems,
and Equipment

‘a)  Definitions of Seismic Classification

" Class I

 

Class I structures, systems, and equipment are those whose failure could cause
‘uncontrolled release of significant amounts of radioactivity, or those essential for
safe shutdown and immediate or long term operation following a design basis

" accident. When a system as a whole is referred to as Class I, portions not associated
with the vital function of the system are to be designated as Class II.

| Class IT

Class II structures, systems, and equipment are those whose failure would not result
in the release of significant amounts of radioactivity and would not prevent reactor
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General Arrangement - Steam Generator Building
b)

6-2

shutdown, but could interrupt power generation. These structures are to be designed
to conform to the requirements of the Uniform Building Code - Zone II. Class IL
structurcs, systems, and equipment shall not degrade the integrity of those
designated Class 1.

Class 111

 

Those structures and components which are not related to reactor operation or
containment. Earthquake is not considered in the design of these structures.

Seismic Classification of Structures, Systems, and Equipment

1) Class I Structures
a. Primary containment
1..  Reactor vessel support
2. Primary heat exchanger support
3. Primary pump support
4. Reactor cell liner support
5. Horizontal support structure

b. Reactor building

1. Reactor building mat

Reactor building enclosure

3. Reactor building crane, crane runway, and crane support
structure ’ '

&

Main control room and cable vault
Battery room -
Primary salt drain tank cell
Off-gascell
Freeze valve cell
Control rod storage cell
Emergency generator room
Off-gas heat-rejection ce
Airlock | |

- Offgas auxiliary equipment cell

At RS0 M oAn

2) _ Class II Structures
The following are considered to be Class II seismic structures:

a. Turbine building
b. Turbine generator pedestal
 

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Steam generator building

Off-gas ventilation stack
Circulating water intake structure
Coolant salt drain tank cell
Auxiliary equipment cell

Waste storage cell

Spent core cell

Spent heat exchanger cell

Hot cell and work areas

Chemical processing cell
Components storage and assembly areas
Radwaste building

Class III Structures

 

Buildings containing conventional facilities.

Class 1 Systems and Equipment

The following are considered to be Class { seismic systems and equipment:

1)

2)

3)

Reactor Fuel Salt System

W e an o

Reactor vessel and internals

Primary fuel salt pump

Primary fuel salt heat exchanger

Freeze valves

Control rod and drive system

Control rod drive housing

All piping connections from the reactor vessel up to and including
the first isolation valves

Fuel Salt Drain and Of&gas Holdup System

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Fuel salt drain tank and piping

Fuel salt storage tank and piping
Off-gas heat reject system equipment
Off-gas and jet pump piping

Fuel Salt Transfer Pum' '

6-3
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4) Standby Electrical Power Systems
1
Dy a. Station battery system
b. Standby motor generator system
i c. Heating, ventilating, air conditioning and lighting in reactor building
e control room
* 5) Instrumentation and Control Systems
a. In-core instrumentation
| b. Control rod drive system instrumentation and control
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B 7) Gas Treatment System
8) Chemical Processing System
"
Class II Seismic Equipment and Piping Systems
¢
b The following are considered to be Class II seismic equipment and piping systems:
s a) Turbine-generator system |
e b) Main condenser and circulating water systems
c) Turbine building cranes
- d) Secondary salt circulating system
b e) Condensate storage and transfer system
f) Station auxiliary power buses
g) Electrical controls and instrumentation (for above systems)
h) Radwaste system
i) Turbine system moisture spearators
}) Condensate demineralizer system
k) Station service water system
. 1) Compressed air system
H m)  Steam 'genferétc')rfs' a.nd reheaters |
n)  Coolant salt rupture discs
. o) All other piping and equipment
Class I1I Seismic Equipment and Piping Systems
- a) Conventional equipment, tanks and piping, other than I and II
classes. ' o
| bu d) All Class 1 structures are designed against the possibility of an onsite tornado
occurence.

 
6.2.2  Reactor Building

The rcactor building is the sccondary containment structure for the reactor primary
system containment cell, the fuel salt processing system cell, the hot storage cells, the
coolant salt drain and storage tank cell and the off-gas system cell during normal operation.
In addition, the reactor building forms the primary containment for all cells opened during
maintenance operations.

The reactor building is a single integrated reinforced concrete multistory structure
as shown in Figure 6.7. The reactor is housed in a cylindrical furnace cell within the reactor
building. Plan views at the five major levels arc shown in Figures 6.2-6.6. The building is
approximately 290 ft long by 160 ft wide and 200 ft high. The top of the foundation mat is
set at grade 100 ft in order to avoid negative numbers and is equivalent to elevation zero.
This is compatible with the site conditions having the top of the limestone formation about

8 ft below grade.

The first level (elevation 100 ft) of the reactor building contains cells and space for
chemical processing drain tanks, heat rejection equipment, off-gas handling, fuel salt drain
and storage tanks, emergency diesel equipment as shown in Figure 6.2. Each respective
elevation, as shown in Figures 6.3-6.6 provides the space and cells for the remaining
equipment necessary for nuclear plant operation. The structure also supports a traveling
bridge crane which would span the length of the building to service the reactor and other
equipment within the building.

During normal operation the building is maintained at slightly below atmospheric
pressure by a controlled ventilation system discharging through filters and up a stack. The
primary purpose of the reactor building normal ventilation system is to limit exposure of
personnel to airborne contaminants and to maintain appropriate temperature conditions for
operating personnel and equipment.

The reactor building normal ventilation system shall:

a) Migrate air from clean accessible areas to areas of
progressively © higher  contamination  or  potential
.contamination. |

b) Remove the normal heat losses from all equipment and

piping in the reactor building during station operation.

¢)  Filter outside air to limit the introduction of airborne
particulate matter to the station. '

d) Exhaust potentially contaminated leakage to the stack
through the ventilation air treatment system. '

6-5
1
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T

The reactor building normal ventilation system consists of a supply and exhaust side.
The supply side includes in the direction of air flow, outside louvers, dampers, filters,
}leating coils, and two supply fans cach sized for the full system capacity. The exhaust side
consists of two exhaust fans, each having full system capacity exiting through two treatment
systems. |

The main supply and exhaust ducts penetrate the reactor building, through two
butterfly isolating valves in series, which are automatically closed by a primary containment
isolation signal. The valves in the main supply duct are powered from different buses. This is
also true of the valves in the main exhaust duct. All isolating valves fail closed.

Supply air will be distributed by means of a duct system to provide equipment
cooling in various areas within the building as required. Air will be routed from clean areas
to areas with progressively greater contamination potential. Gravity dampers are provided at
key points in the duct distribution system to prevent backflow of air from contaminated to

- clean exhaust duct branches. All exhaust air will be routed through a return duct system

where the exhaust fans direct the exhaust air to the treatment system which monitors the
air and 1) discharges to the stack if of acceptable quality or 2) processes the air through
filters, charcoal adsorbers, and if necessary, through delay tanks if contaminated.

Operating personnel would have access to the major portion of the reactor building
at all times except during certain phases of maintenance operations. During these periods
remotely-controlled equipment can be viewed through shielded windows in the remote
maintenance control room wall at the crane bay level (elevation 200 ft).

The design criteria for the reactor building is as follows:

1) Withstand a 300 mph wind; a storm caused 3 psi negative
pressure differential, and a 2500 Ib missile, 15 in. in
diameter, traveling at a speed of 150 mph.

2)  Design basis earthquake equipment to 0.07g horizontal base
acceleration. |

Reactor Cell

The reactor cell isr_th.é containment for the primary system. It is about 72 ft ID by
50 ft deep. The design criteria for the reactor cell is as follows:

| 1) The atmosphere of the reactor cell is to be operated at about 13 psia and at
1000 F * 25 F and a He atmosphere.

2)  The cell is designed for 50 psig. This pressure would be supp]iéd from He
storage tanks to assist in fan coolers transferring the heat of deposited fission
products from the drained reactor and primary system.
 

t“

3) All wall construction must provide both thermal insulation and gamma
shielding to protect concrete structures form rising above 150 F.

4) Previous projects which had base accelerations of approximately 0.07g and
foundations in rock were therefore examined to obtain what could be
considered reasonable values. Building periods of around 0.4 sec and
maximum accelerations (at 80 ft from foundation) were about 0.10g.
Therefore, if the supported equipment within the cell is considered to have a
period of 0.05 seconds (rigid), the acceleration at 80 ft from the foundation
for the reactor piping system was taken as .25g and for the structure as .5g.

5) Prevent the escape of radioactivity both during normal operation and during
accidents.

The reactor heat exchangers and salt piping are all supported from the cell floor
whereas the fuel salt pump is mounted in the roof of the reactor cell and is supported on a
set of springs to allow for the relative thermal expansion of the other components. The
primary heat exchangers are mounted on roller bearings to allow for horizontal thermal
growth but are restrained by a three tier girder arrangement as shown in Figure 6.14 for
seismic protection of the piping, reactor vessel and heat exchangers. Calculations indicate
that the restraining members, if made of a metal closely matching the piping system thermal
expansion coefficient, will be of the wide flange type with maximum sizes as follows:

a. Horizontal tiers-flanges — 2 in. x 18 in. wide
web — 1 in. x 36 in. deep

b. Vertical support-flanges — 7/8 in. x 14-5/8 in. wide
web — 9/16 in. x 12-5/8 in. deep

Supporting the reactor vessel and heat exchangers from the reactor cell floor thus
facilitates construction as the primary containment may be erected first and post-weld heat
treated followed by a placement of concrete for reactor cell shield walls. Sizes of primary
containment vessel plates need not be restricted as in the top support method.

Of the materials investigated for use as structural members within the reactor cell,
two possibilities exist: type 304 stainless steel (ASTM 240) and Inconel 625 (ASTM B443).
Due to the environment within the rcactqr_cell, the following were considered in the

~ selection of the above materials:

1) Tensile strength, yield point, allowable stresses and their decrease in values at
elevated temperatures - |

2) Creep and creep rupture

3) Fatigue
    

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6-8

 

%} 4) NDT temperature, precipitation hardening and temper brittleness
5) Corrosion (carbide instability and oxidation of scaling)
6) Radiation effects on above items

Based on the more critical items above, it is anticipated the Inconel 625 will be
ultimately selected as it appears the more favorable of the two materials. Allowable stress or
™ stress intensity, with reference to the ASME Code Section VHI (Case 1409-1) is 26 000 psi
| (Grade 1 at 1100 F) and, with reference to ASME Code Section III (Case 1422) is 31 000
s psi (Grade 1 at 800 F) versus 8700 psi and 14 800 psi, respectively for type 304 stainless
- steel. With respect to creep rupture, Inconel 625 again offers the best possibility. The below
| information was obtained from manufacturer’s literature:

 

; o | Inconel 625 Type 304
Stress for Rupture (psi) (psi)
- ( 1 000 hrs) 94,000 (approx) 24 500
| (10 000 hr) 86 000 (approx) 18 300
- | Reactor Cell Future Study
o 1) Insulated pipe and equipment simplifies support, and wall construction
o complicates maintenance. ‘
 

§
t

- a1t

r

7.0 CHEMICAL PROCESSING

7.1 General

 

Task I objective for CONOCO was to develop a spatial layout for the Conceptual
Design Flow Sheet of the chemical processing system supplied by ORNL.

The information provided by Oak Ridge National Laboratory contained the basic
chemistry for their Conceptual Design Flow Sheet and the major values for processing
component sizes and stream flow rates.

An engineering analysis by CONOCO of the ORNL concept was performed to define
the pumps, valves, auxiliary vessels, and other processing components which would be
necessary to complete the chemical processing system. From these engineering flow sheets
the spatial layout was developed.

Section 7.2 gives the design basis from which the process flow and spatial layout
were developed. Section 7.3 gives a process description of the chemical processing plant.
Detailed process flow drawings described by the process description and layout drawings
which define the spatial requirements are given in Section 7.4,

7.2 Dcsi@ Basis

The basis for the process flow and spatial layout of the chemical processing plant
was developed both from meetings with the personnel of Ebasco Services Incorporated, and
Oak Ridge National Laboratory, and from reports issued by ORNL.

Radioactivity in the chemical processing cell, which is at a very high level, requires
all operations and maintenance be performed remotely; human access to the cell is not
possible after initial activation. Design is for a 30-year nominal life expectancy for the
reactor. The cell enclosure is heated to the average process temperature. All process
components are capable of being drained by freeze valves to tanks equipped with a fail safe

-cooling system. The cell area is preferably rectangular in shape with process components

located along the walls of the cell

More specific premises upon’ which the process flow diagram and spatial layout were
developed are contained in the following lists. The first three state prermses concerning (1)
process flow, (2) safety, and (3) plant maintenance. ,

7.21  Process Flow Bases
a)  The process is continuous except for the periodic operation

of the rare earth salt fluorinator and UF product removal
system.

71
7.2.2

k)

m)

The reactor salt to be processed is pumped continuously to
the chemical processing cell.

The control functions necessary to the steady state operation
of the process are indicated. Instrumentation is assumed to
be available and sufficient space is allowed for it.

All process lines in periodic use are capable of being blown
with argon which is at process temperature.

Overflow as a method of moving fluids is preferable to
employing pumps.

Positive displacement pumps are equipped with necessary
instrumentation to function as flow metering devices. Gas
pressurization as an alternate to pumping will be investigated.

It is possible to use control valves on all process streams.

Certain process components are at sufficiently low levels of
radioactivity that they may be located in a hot cell adjoining
the main processing cell.

A means of sensing a salt-bismuth interface is available.
The makeup of beryllium and thorium salts to reactor salt is
not done in line because of the possibility of solids plugging

lines.

The purified bismuth streams leaving the two
hydrofluorinators for recycle back to the extractors may be

mixed together in a common holding tank.

- B_lof:k:valves_which are used solely for isolating a component
- during maintenance will not require extensions through the
- cell wall, They will be opened and closed by the manipulators

used for malntenance

| Haste]loy-N nickel, graphite, and molybdenum are assumed

to be the only feasible materials of construction.

Safetx .Base's

a)

A drain tank system is provided to drain all process vessels
and lines except those containing aqueous solutions.
 

r—

b)

f)

g)

h)

k)

I

Aqueous solutions are disposed of by direct routing to the
Radwaste disposal system.

Direct mixing of H, and F, is not to be permitted. It is
assumed that H, and F, mixtures react violently.

The drain tank system provides a number of tanks to prevent
H, and F, from contacting and to maintain segregation of
the three bismuth phases which contain 0.2, 5.0, and 50.0
mole percent lithium concentrations.

A floor drain system drains spills to either a temporary drain
tank or a flush salt drain tank.

Freeze valves provide a fail-safe system for draining all vessels
and for bypassing pumps and control valves.

All process lines are sloped to be self draining during
shutdown.

The drain tanks are sized on the basis of double the volume
of the vessels which each must drain.

Each drain tank is vented to the vessels it drains through an
argon manifold. Argon is added and relieved as necessary to
compensate for changes in liquid holdup.

The fluids in the drain tanks can be pumped back to the
processing system.

All pumps, valves, and vessels are cooled with the NaK
operational coolant system.

_ The NaK operational coolant system operates with its own
pumps and surge tank.

The drain tanks are equipped with a failsafe NaK coolant

- -system for emergency use in addition to the operational
~ cooling system. -

A]]_off_—gasés' are dis;pdsed of through a central off-gas disposal
~ system. "

The process cell atmosphere is argon gas under slight negative
pressure. Argon seals are provided for controlled in-leakage.

7-3
 

7.2.3

Plant Maintenance Bases

a)

f)

g)

k)

A rectangular shape for the chemical processing cell is
compatible with the building concept which has 40 fc bay
widths.

Normal equipment entry into the process cell is through a

baffled airlock.

All electric motors for process pumps are located outside
processing cell. |

All maintenance and component replacement is done
remotely. Overhead cranes and wall manipulators are
provided for normal remote handling. Removable concrete
ceiling slabs allow the reactor crane access to the processing
cell interior if necessary.

Viewing of the process cell is by view windows, mirrors, and
television cameras.

All wall penetrations for pump and valve drive extensions are

below knee and above head heights.

Vessels are located in a plane parallel to the wall. A minimum
wall clearance of 1 ft and a minimum clearance between
vessels of 3 ft is required.

Pumps are located at a minimum of 3 ft for horizontal
spacing and 2 ft for vertical spacing.

All lines have 6 in. clearance.

All process 'compbnents are capable of being blocked off by
valves and removed. '

The location of each component in the cell, is indexed for
positioning of replacements. However, final alignment of

components and operation of tools is by some visual means.

The flanging of lines in a leakprdof manner is assumed.

74
 

7.3  Fuel Salt Chemical Processing
7.3.1  Purpose of Chemical Processing

The chemical processing system has the objectives of isolating protactinium-233
(233Pa) from regions of high neutron flux during its decay to uranium-233 (233y), and
removing fission and corrosion products from the reactor salt. The processing system
functions to achieve these objectives by first removing uranium from the reactor salt by
fluorination. The protactinium is then removed from the reactor salt to a bismuth stream by
reductive extraction. The bismuth stream in turn transfers the Pa into a decay salt phase
where the Pa is held for its decay to uranium. The major portion of the uranium removed
from the decay salt by fluorination is returned to the reactor salt while a small portion is
routed to a product uranium hexafluoride container.

After Pa removal, the rare earth fission products are reductively extracted from the
reactor salt into a bismuth phase. The rare earths are then transferred sequentially into a
LiCl phase, into another bismuth phase, and finally into a salt phase rich in rare earths. The
rare earth salt is batch fluorinated periodically, to remove any remaining uranium before
being transferred to storage tanks to await disposal.

The reactor salt begins reconstitution by purging a small amount of salt periodically
to be batch fluorinated with the rare earth salt, and adjusting the salt composition with
beryllium and thorium salts. Uranium hexafluoride from the fluorinators is returned to the
reactor salt by hydrogen reduction. The salt is lastly treated with a hydrogen and hydrogen
fluoride gas mixture and filtered to remove corrosion products.

7.3.2  Process Description

The chemical processing plant is divided functionally into seven sections (Sections
100 through 700). Section 100 is the reactor salt system which has the function of removing
the Pa and rare earths for further processing and preparing the reactor salt for its return to
the reactor drain tank. A small stream of 0.88 gpm reactor fuel salt is taken continuously
from the reactor drain tank to be contacted with fluorine in the reactor salt fluorinator
R-101 where about 95 percent of the uranium salt (UFy) is removed as gaseous UF¢. The
salt is then contacted with hydrogen in the UF: reduction reactor R-111 to reduce the

' remaining nonvolatile uranium fluoride salts back to the UF 4 salt.

The salt continues to the protactinium (Pa) extractor W-102 where the remaining
uranium and protactinium salts undergo reductive extraction into a liquid bismuth phase
containing 0.2 mole percent lithium. The salt and bismuth are contacted counter—currently
at a metal-to-salt volume ratio of about 0.125.The lithium and thorium dissolved in the
bismuth reduce the uranium and protactinium salts to the metals which are soluble in the
bismuth phase. The bismuth stream containing the Pa continues to Section 300 for further

chemical processing.

The fuel salt continues to the rare earth (RE) extractor W-103 for removal of the di-
and tri-valent rare earth salts from the salt phase by reductive extraction into a bismuth

 
 

T

phase containing 0.2 mole percent lithium. The bismuth and salt are counter-currently
contacted at a metal-to-salt volume ratio of 14. With the lithium acting as the reducing
agent, the rare earth salts are reduced to the rare earth metals which are soluble in the
bismuth. The major portion (12.39 gpm) of the bismuth stream leaving the RE extractor
goes to Section 200 to the rare earth transfer (RET) extractor W-201 while a small portion
(0.11 gpm) is the source of bismuth for Pa extractor W-102.

Reconstitution of the fuel salt for its return to the reactor system begins in a small
settler D-109. Periodically a small volume of salt is transferred to a mix tank where thorium
and beryllium salts are added to adjust the composition of the reactor salt. After mixing, the
salt is transferred back to the settler. A reactor salt purge of about 0.16 cu ft/day is taken
daily from the settler to the rare earth salt hold-up tank D-402 in Section 400.

From the settler, reactor salt flow to the UFg reduction reactor R-104. A gaseous
UF¢ and F, mixture from the fluorinators reacts with a recycle salt stream containing UF4
to form subfluoride salts such as nonvolatile UF 5. The salt is then contacted with hydrogen
to reduce the subfluoride salts to UF 4.

Lastly, the salt is counter-currently contacted with a gaseous mixture of hydrogen
and hydrogen fluoride in nickel gauze packed reactors R-106 and R-107 to achieve a desired
UF4/UF 3 ratio and to remove impurities such as bismuth, nickel, and iron fluorides. The
salt passes through porous metal filter F-108 on its return to the reactor drain tank.

Section 200 is the transfer salt system. Its function is first to transfer the rare earths
extracted into the bismuth stream in Section 100 into a lithium chloride salt phase and
second to transfer the di- and tri-valent rare earths into separate bismuth phases. The
bismuth stream from RE extractor W-103 carries rare earths to RET extractor W-201. The
bismuth is counter-currently contacted with a 33.4 gpm stream of lithium chloride (LiCl)
salt. The rare earth fission products are converted from dissolved metals in the bismuth
phase to their rare earth chlorides which are soluble in the LiCl stream. The bismuth is
recycled to Section 100 to RE extractor W-103 after a makeup mixture of 0.2 mole percent
lithium in bismuth is added to rep]acé the bismuth routed to W-102.

The rare earths are removed from the LiCl stream in two stages The tri-valent rare
earths are transferred from the LiCl phase to a bismuth stream (8.28 gpm) containing 5

~ mole percent lithium by counter-current contact in RE3 extractor W-203. The bismuth is
“circulated from the RE3 heat exchanger X-205 to provide sufficient heat removal. Daily a
~ portion of the bismuth is drawn off to Section 400 to the RE3 hydrofluorinator R-401 for

bismuth purification.” An equal volume of recycled bismuth at 5 mole percent lithium

_concentration is made up daily from a Bi-Li holding tank. The major portion of the LiCl is
" returned to the RET extractor W-201 while 0.69 gpm (about 2 percent of the total LiCl
" flow) is routed to RE2 extractor W-202. The di-valent rare earths are removed from the LiCl
~ phase by counter-current contact with a 0.23 gpm bismuth stream containing 50 mole
percent lithium. For heat removal purposes, the bismuth is circulated through RE2 heat

exchanger X-204. Daily a portion of the bismuth is transferred to the Pa-RE2

7-6
 

 

hydrofluorinator R-301 in Section 300 for bismuth purification, and an equal volume of
recycled bismuth at the 50 mole percent lithium concentration is replaced from a Bi-Li

holding tank.

Section 300 is the Pa salt system which accepts Pa and di-valent rare earths from
Sections 100 and 200, respectively. The bismuth streams from the Pa extractor in Section
100 and RE2 extractor in Section 200 flow to the Pa-RE2 hydrofluorinator R-301 where
the bismuth is contacted with the Pa salt and hydrogen fluoride gas. The protactinium,
uranium, and rare earths dissolved in the bismuth are oxidized by the hydrogen fluoride to
their salt forms which are soluble in the Pasalt. The purified bismuth is recycled to the RE
and RE2 extractors in Sections 100 and 200, respectively, after proper lithium addition in
holding tanks. The Pa salt flows to the Pa salt fluorinator R-302 at 0.68 gpm where it is
contacted with fluorine. Ninety mole percent of the uranium in the salt is removed by
fluorination to gaseous UF¢ which is normally returned to R-104 for reconstitution of the
reactor fuel salt. The Pa salt is then contacted with hydrogen in Pa salt reduction reactor
R-305 to reduce rémaining nonvolatile uranium fluoride salts to UF,4. The salt stream then
continuous to Pa decay tank D-303 where a salt volume of 130 cu ft is held to allow the
protactinium sufficient time to decay to the uranium salt UF4. The uranium resulting from
the decay of Pa is removed in the Pa salt fluorinator as the Pa salt is recycled back to R-302.
A daily purge of about 0.1 cu ft of Pa salt is sent to Section 400 to the rare earth salt
holdup tank D-402.

Section 400 is the rare earth salt system which serves to concentrate rare earths into
a salt phase for storage and to recover any remaining uranium before salt disposal. The
reactor salt purge from D-109 and the Pa salt purge from D-303 are held in the RE salt
hold-up tank D-402. This salt is continuously circulated to the RE3 hydrofluorinator R-401
to counter-currently contact the bismuth from the RE3 extractor of Section 200 and
hydrogen fluoride gas. The tri-valent rare earth metals dissolved in the bismuth are oxidized
by the hydrogen fluoride to their salt forms which are soluble in the RE salt. The purified
bismuth is recycled to RE3 extractor W-203 after lithium addition in a holding tank. Every
68 days 18.5 cu ft of RE salt is transferred to the RE salt batch fluorinator R-403. The salt
is contacted with fluorine for a day to remove remaining traces of UF4 as gaseous UFy,.
After fluonnatlon the salt is dlscarded to RE salt drain tanks of Section 500 for storage.

In Section 600 the'uramum productlon is 'achleved by routmg, for about an hour
daily, the gaseous UF¢-Fy stream from the Pa salt fluorinator R-302 to a NaF bed W-602.
~ 'This gas mixture normally returns the UFg to the fuel salt in reactor R-104. The NaF bed

absorbs about 133 grams/day of UF6 which is later desorbed and condensed in a UFg
product cylmder for shlpment

Section ‘500 contains a system of tanks equlpped w1th a fall-safe coohng system.
These drain tanks provide liquid storage and emergency cooling 8capab1ht1es for penods of
abnormal process operation. | |

Section 600 is basically the gas recovery system which provides the HF, F,, and‘le
gases for the process. The hydrogen and hydrogen fluoride gas mixtures from the process

7-7
 

'!m*}.flm
[

return to Section 600 where the HF is condensed and converted to H, and F, in an
electrolytic cell. The hydrogen gas is scrubbed with a KOH solution and dried for recycle to
the process.

Section 700 contains those process components, which because they are ata lower
level of radioactivity, can be located in a hot cell adjacent to the main processing cell.

tm semrn.,

 

7.4  Index of Drawings
- | Figure CONOCO 1.D.
| Title ' Number Number
W Block Flow Diagram | 74 SD1700-42-1-C
: H - Process Flow Diagrams:
'; Reactor Salt System (Sec. 100) 7.2 SD1700-42-2-D
o - Transfer Salt System (Sec. 200) 7.3 SD1700-42-3-D
v Pa Salt System (Sec. 300) |
RE Salt System (Sec. 400) 7.4 SD1700-42-4-D
o Drain Tank System (Sec. 500) 7.5 SD1700-42-5-D
e
Gas Recovery System (Sec. 600) 7.6 SD1700-42-6-D
Auxiliary Processing Area (Sec. 700) 7.7 SD1700-42-7-D
-
Plans and Sections 7.8 SD1700-54-1-D
] Process Equipment Arrangement 7.9 SD1700-54-2-D
™ | Sec. 100 Isometric and Miscellaneous
| Details 7.10 SD1700-54-3-D
|
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(Ref Dwg No. SD 1700-42-2.D)

 
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L Process Flow ‘Diagram

Pa Salt System (Sec 300), Re
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(Ref Dwg No. SD 1700-424.D)
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Gas Recovery System (Sec 600)
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‘Plans and Sections
(Ref Dwg No. SD 1700-54-1-D)

 
 

r 1 ('r Tro D ¢l r .. .- ( . " W

 

8.0 REACTOR OFF-GAS SYSTEM

 

8.1 System Requirements

The reactor off-gas system must perform the following functions:

e)

‘maintain sufficiently low 135xe concentration in the fuel salt,
femove tritium from the purge gas,

accommodate other forms of contamination such as noble metals,
particulate, or hydrocarbons which might be picked up by the purge
gas,

have 100 percent availability,

. provide positive confinement of all radiation under all normal modes

of operation.

8.2 Basic Assuthions

The following assumptions provide the bases for the off-gas system design:

a)
b)
c)

d)

f)

At least 90 percent of the xenon 135 produced enters the off-gas.
system. This amounts to 6.2 x 106 moles/second.

Virtually 100 percent of all longer-lived fission gases enter the off-gas
system.

Virtually 100 percent of the tritium (nominally 2400 curies/day)
produced enters the off-gas system.

The purge gés_'emerginrg’ from the primary systl:m will also contain
short-lived fission gases in the amount determined by ORNL
(ORNL-4541).

- The carrier gas may be either helium or argon. Helium is arbitrarily
~ selected as the tentative choice for the Ebasco reference concept.

The rate of decay heat production in the purge gas as it emerges from

“the fuel salt is about G.lMW/ft3(STP)_.
 

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8.3  Reactor Interface System

 

The off-gas system is interfaced to the primary system by some method or device
which injects clean purge gas and removes contaminated purge gas from the salt. It is
assumed that R&D efforts and ORNL will lead to such a satisfactory system or device.

8.4  Purge-Gas Cleanup System

 

Two methods of cleaning the purge gas have been proposed. One method involves
physical holdup of the purge gas for 135xe decay. The other involves actual cleanup of the
purge gas by passing it through charcoal beds. Each system has specific advantages and
disadvantages. The following sections present the merits of each system and provide the
bases for choosing a reference concept.

8.4.1  Physical Holdup System

 

The purge gas emerging from the fuel salt flows directly to a holdup tank. The
volume of the holdup tank, together with the volume flow rate of purge gas, determine the
holdup time. After the desired holdup interval, the gas in the tank is recycled directly to the
bubble generator. In this system, the concentration of fission gases will build up until the
decay rate of fission gases in this gas reservoir equals the rate at which the fission gases are
transferred to the gas reservoir. If the volume of this reservoir is small, the steady-state
fission-gas concentration will be high. In this case, the 135%e back pressure can impede
135Xe transfer from the fuel salt to the bubbles. If the gas volume is made sufficiently large,
however, the steady-state 139Xe concentration will be low enough that the back pressure
will not impede 139Xe transfer, Figure 8.1 shows the poison fraction and other parameters
as a function of the purge-gas volume. It can be seen that a low volume gives a high
fractional back pressure, and therefore, a poison fraction approaching 5 percent, i.e., the
case for which there is virtually no 139Xe removal. For a very high reservoir volume, the
back pressure approaches zero, and the 135Xe transfer is limited by the bubble-surface area.
The desired poison fraction can be attained by the proper choice of purge-gas volume,
together with the proper choice of surface area. If the core graphite is sealed, a purge-gas
volume in the range of 4000 cu ft, together with a bubble-surface area of about 15 000 ft
is adaquate. This surface area is slightly larger than assumed by ORNL, but is sufficient to

compensate for the much higher back pressure than assumed by ORNL. These values are

based on the assumptions that the graphite would be sealed and the bubbles fairly large (.02
in. diameter). It is now believed that the bubbles will be smaller than the surface area

~roughly ten times larger. The back pressure will therefore become more significant and

perhaps justify going to a somewhat larger purge-gas reservoir. If the gas reservoir were

 pressurized, then the same level of dilution could be achieved with a smaller volume. Two
~ atmospheres of pressure would permit the reservoir volume to be roughly the same as that

of the drain tank. If the drain tank were used, its cooling system would be available to

" remove fission-gas decay heat. Conversely, the fission-gas decay heat could be used to

continually drive the drain-tank cooling system and thereby demonstrate its operability and
availability. The decay heat of fission gases alone has been estimated to be about 10 MW. An
additional 10 MW might be generated from decay of noble metals which accumulate in the
drain tank.
 

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Y9, PURGE - GAS RESERVOIR (cm?)

Figure 8.1

 
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The holdup tank will contain about 600 000 000 curies of fission gases. This might
present licensing difficulties, especially if the tank is pressurized. Alternately, several smaller
tanks could be employed to reduce the likelihood of a large fission-gas release, to insure 100
percent availability, and to permit on-line maintenance and repair.

It is recognized that some purge-gas cleanup via charcoal filtration will probably be
required. This will certainly be true if there is a continuous flow of clean purge gas into the
contaminated purge gas, as, for example, the flow through the purge glands in the primary
pumps. In this case, the supply of clean purge gas must be replenished by cleaning and
recycling the contaminated gas. If, on the other hand, purge glands are not required, then it
seems reasonable to have two separate gaseous systems: One system containing only highly
contaminated gas and not diluted by clean gas. This would include purge gas and cover gas
only. The second system would include the reactor cell atmosphere and a clean circulation
system which purges cell penetrations and discharges into the cell. The two gas systems
would have to be kept separate by gas-tight seals. The primary pumps would therefore
require gas-tight oil seals in order to keep the two gas streams separate. Although the
reactor-cell atmosphere normally would not be contaminated, it is anticipated that it might
inadvertently become contaminated because of a seal failure. Thus, a charcoal filtration
system should be available in the event that it is needed.

8.4.2  Charcoal Bed System

 

- This section presents a design description of a concept proposed by ORNL. The
performance criteria assumed in this study are the same as those assumed by ORNL. It is

understood that these criteria are largely arbitrary and subject to change.

Assumptions

1) Fission gas %roduction rates are based on a reactor power of 2329 MW(t) and
a fuel of 233y,
2) The carrier gas is helium, with a total flow to the off-gas system of 11 scfm.

This total is the combination of flows from each of the four pump loops,

consisting of 2.25 scfm from each of the gas separators and 0.5 scfm of

purge gas for each of the pump shafts. Net flow of fission products and
. materials other than helium is at most 0.1 percent, or 0.01 scfm.

3) The atom flow rates of Kr and Xe into the off-gas system are based on
| calculated atom flow rates at the gas separator discharge, with appropriate
corrections for a 2-hr residence time in the fuel-salt drain tank. All solids

~ which are gas-borne at the outlet of the drain tank (including noble metals, -
~salt mist, and solid daughters of the noble gases) will be removed by a filter

before the gas stream enters the off-gas system. The total yield of tritium

(°H) from all mechanisms will be 2400 curies/day, and all tritium will

remain in the offgas stream, that is, for the purpose of studying the off-gas
system, the rate of diffusion of tritium through vessels and pipe walls is
assumed to be zero. '

8-3

 

 
 

rmm-u
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r—

4) The gas will enter the off-gas system at 10 to 15 psig. Nine scfm will be

returned to the bubble generators at 5 psig. Two scfrn will be returned to the
purge-gas header at 45 psig.

5) Shielding will be provided for attenuation of penetrating radiation to
permissible levcls. Instrumentation will warn of excessive leakage of gas or
penetrating radiation.

6) The target reliability of the system is 100 percent; that is, spare units will be
provided, and the maintainability of units will be such that predictable
failures in the off-gas system will not result in shutdown of the reactor or
loss of the contaminants to the environment.

The flow of gas in the off-gas system can be represented by two recycle loops, a
47-hr Xe holdup loop, and a long delay (~90-day) Xe holdup loop, as shown in Figure 8.2.
These holdup times do not include the 2-hr residence time of the off-gas stream in flowing
through the fuel-salt drain tank. The 47-hr loop circulates through the bubble generator and
gas separator to strip the 139Xe from the fuel salt. The long-delay loop carries the balance
of the gas flow in the fuel system. The two loops are joined together at the salt entrainment
separator and flow concurrently through the primary drain tank and the 47-hr holdup
system. :

The concurrent stream enters the primary off-gas system from the fuelsalt drain
tank and is cooled by means of a radiator or forced convention type air coolers, The
purpose of cooling the gas would be to increase the effectiveness of the 47-hr holdup
system. The drain tank will probably serve as an efficient collector of particulates in the gas,
but if it proves necessary, a particle trap, or filter, can be added, as shown in Figure 8.2, At
this point, the gas will have been stripped of non-gaseous components (noble metals, salt
mist, and non-gaseous daughters of the noble gases), so that the primary contaminants are
Kr, Xe, and 3H. About 2 hr will have elapsed since the gas first left the fuel salt system. The

84

gas first passes through the 47-hr Xe holdup system to provide a residence time for xenon

molecules sufficient to permit the 135Xe to decay to about 3 percent of the inlet amount.
The 47-hr holdup system will utilize charcoal for the dynamic adsorption and holdup of

krypton and xenon. The decay heat will be transferred to a forced c1rcu]at1ng water system

within the shells of heat exchanger units.

At the outlet of the 47-hr system the gas stream is dmded into the two recycle

loops In the 47-hr recycle loop, 9 scfm, or about 80 percent of the total flow, passes in -

succession through a chemical trap and alarm sytem, a surge tank, a compressor, and an

- accumulator. From the accumulator the gas is metered to the bubble generators at the four

circulating pumps. In the second recycle gas stream, 2 scfm, or 20 percent of the total flow,

passes first through the long-delay Xe holdup system where the residence time for krypton

and xenon are sufficiently long to allow all radioisotopes except the 10-yr 85Kr to decay to

insignificant levels. The gas then passes through a gas cleanup system which reduces the level
of any remaining contaminants 5Kr H, stable isotopes _of Kr and Xe, water,
e

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Off-Gas System - Flow Diagram
T

hydrocarbons, etc.) to an acceptable level, then through a surge tank, a compressor, and an
accumulator, and finally is returned to the primary system.

Design Criteria for the Gas Cleanup System:

 

e

e

2)
3)
4)
5)
6)
7)

8)

Carrier gas is helium at a flow of 2 scfm and an inlet pressure of 20 psia. The

design pressure drop is 4 psi.

The level of each contaminant in the effluent gas is not more than 1 percent
of the value at inlet. A

The gas contains some 13Imy.  which is negligible from a mass flow
standpoint, but which must be considered in the design of shielding and the
heat dissipation system.

The stable noble gases, as well as essentially 100 percent of the 85Kr and
3H, will be carried into the gas cleanup system at a rate equal to the rate of
production in the reactor (assuming that no tritium is lost to other parts of
the reactor system by diffusion through pipe and vessel walls).

The tritium oxidizer prehéater and aftercooler have heat loads of 3 kW and
designed for negligible pressure drop.

The tritium oxidizer is 2 in. ID by 3 ft long, is packed with 13 Ib of copper

oxide, and operates at 1500 F. The tritium flow is 0.036 cu ft/day with an

allowable A p of 2 psi. The CuO consumption at breakthrough if 60 percent
and the operating life of a unit is to be 1000 days. Development work will be
needed to confirm the efficiency and pressure drop estimates, however.

Each adsorber is made up of 16 pieces of charcoal-packed 8 in. pipe, with
1-1/2 in. interconnections. The total length of 8 in. pipe is 288 ft, arranged
in two branches to provide a Ap of 2 psi. The pipes are closely stacked inside
a 3 ft to 4 ft diameter pipe with a heated or cooled fluid circulated in the

 interstitial spaces, to provide an average on-stream operating temperature of
0 F and a temperature of 500 F when on the regeneration cycle. Using an

adsorption coefficient of 4.8 cu ft/lb the total charcoal requirement is about
3000 1b. The operatmg cycle is 8 days—4 days on stream and 4 days
regenerating. |

'Nn'rogen storage bott]es, similar to a 1.5 cu ft hlgh-pressure gas cylinder,

each container would be kept on line for 12 cycles, or 48 days, About 30 1b
of xenon, 6 lb of krypton, and 0.1 Ib of tritiated water would be
accumulated in each bottle. Each freshly filled bottle would contain about

240 curies of 89Kr, equivalent to a decay energy of about 0.4 watts per -

bottle. The bottle pressure after equilibrating to room temperature would be
1000 psi. About 230 bottles would be filled during the 30-year life of an

8-5
R

T

"

MSBR station. After each container was filled, it would be transferred to
long-term storage, where, after a period of about 100 years, the 3H 4nd
85Kr would decay sufficiently for the contents to be released or sold
without radiological protection.

9) The gas-cleanup compressor has a capacity of about 2 scfm helium with an
inlet pressure of 14.7 psia and an outlet pressure of 60 psia. A major
requirement for the compressor is to provide positive sealing for the pumped
fluid so that the highly purified gas is not recontaminated.

10)  The Off-Gas System pipe design will be such as to minimize the effects of
solids accumulations, such as at valve seats, pipe bends, etc., where fission
product decay heating would tend to cause hot spots.

11)  Wherever necessary, valves will be provided with welded bellows for positive
stemn sealing. Positive-sealed end connections, either buffered 0-rings or butt
welds, are used. Where necessary, provisions are made for remote
maintainance of valving, |

12)  Gas system piping and components are provided with a controlled-circula-
tion ambient air system, which assures prompt detection of gas leaks, and
the channeling of such leaks to an absolute filter system.

Description of Gas Cleanup System

Upon entering the gas cleanup system, as shown in Figure 8.2, the off-gas first passes
through a preheater, which raises the gas temperature to 1500 F. It then passes through an
oxidizer, which converts the tritium to 3H20, and then through a water cooled aftercooler
and a refrigerant cooled aftercooler reducing the gas temperature in two steps from 1500 F
to 100 F and then to 0 F. The function of the aftercoolers is to reduce the heat load on the
ensuing components. The off-gas then passes through a charcoal-packed adsorber which is
maintained at O F. The H20 and the kryptons and xenons are retained on the charcoal
while the carrier gas passes through the bed. After leaving the refrigerated adsorber, the
carrier gas is recompressed and recycled to the reactor purge system. In normal operation,
two adsorbers are alternated on a fixed cycle. A regeneration process is used to transfer the
adsorbed gases in the offstream unit to a receiver cylinder for permanent storage.

The helium gas used for regeneration is taken from the He purge header, and
preheated, if considered necessary. Duting regeneration, the gas flow is about 10 percent of
normal on-stream flow and moves through the adsorber unit in the opposu:e direction. After
leaving the heated adsorber bed, the regenerating gas, now laden with HZO krypton and
xenon, passes through a storage bottle maintained at a liquid nitrogen temperature of -325 F
b use of an external liquid nitrogen refrigeration system. The water, krypton and xenon are
trapped in the bottle and the purified effluent returned to the main carrier-gas stream: A
compressor is used to return the effluent of the gas cleanup system to the purge-gas cycle.

8-6
8-7

Design Criteria for the 47-Hour Xenon Holdup System

1) The residence time for xenon is 47 hours. This time is exclusive of the
volume holdup in the primary system drain tank and other vessels and ducts.
A 47-hour delay time permits 97 percent of the 9.14 hr 135%e 1o decay.

2) The estimated heat load is 2.14 MW, 42 percent of which is due to
daughter-product decay. The design capacity of the heat removal system is
125 percent of calculated, or 2.7 MW.

3) The 47-hr Xe recycle system is to be designed to operate on the available
pressure drop, so a compressor probably will not be required. However, if
one is needed, the flow will be 9 scfm and the compression ratio will be
fairly low, about 1.4 to 1. Positive sealing will be essential to prevent
outleakage of the highly radioactive gas. Other requirements will be radiation
resistance and remote maintainability.

4) A dynamic adsorption system is used for delay of the xenon. The adsorbent
is activiated charcoal with transfer of the decay heat to forced circulating
water. The design temperature of the charcoal pipe wall is dependent upon
the average circulating water temperature in the respective decay heat
generating regions, The average temperature of the charcoal in regions I and
IT is 285 F and in region Il is 254 F.

5) The assumed charcoal properties are:
Bulk density 30 Ib/cu ft
Thermal conductivity 0.03 Btu/hr sq ft. F/ft
Size range ~ 6 to 14 Tyler Sieve Series

(1/8 to 3/64 in.)

6) The average de_c_éy_.he'at distribution is obtained By selecting three regions at
holdup time intervals of 17, 14 and 16 hours, resulting in respective decay
heat rates of 1.1, 0.7 and 0.4 kW/min.

7)  The efficiency of the bed is assumed to decrease with time due to
accumulation of solid daughters. Thirty percent spare capacity is provided
and provision is made for replacement of modules by remote maintenance
techniques. | - |

8) Carrier gas flow is 11 scfm and the overall pressure drop is less than 5 psi. An
- estimate of the size of the charcoal bed is obtained by using the empirical
relationship developed by Browning and Boltal! | -

1/ W E Browning and C C Bolta, Measurement and Analysis of the Holdup of Gas
Mixtures by Charcoal adsorption Traps, ORNL-2116, July 1956.
 

m’m . iy
- A - s
3

€

th = km

where: t, = holdup time
m = mass of charcoal
f = volume flow rate of carrier gas at local conditions and

k is a proportionality factor which is known as the
adsorption coeffiaent and which varies with the carrier gas composition, the
absorbent, the adsorbate and the temperature. For typical commercial
charcoals, Ackley and Browning="have determined the following relationship
between k and temperature from xenon at temperatures between 32 F and
140 F.

 

k (Xe) = 3.2 x 107 ex cu ft/lb Eq. 2

Equation 1 indicates that the holdup time increases directly with k.
However, an increase in holdup time increases the heat generation which
results in an increase in charcoal temperature and decrease in k in accordance
with Equation 2. An increase in temperature causes an increase in f (local
flow rate), which results in a decrease in holdup time. For any given section
of the bed, k and t}, will seek equilibrium values which are a balance between

these opposing forces.

For the purpose of this conceptual design, the assumption was made that
Equation 2 is valid up to 300 F. Equation 2 indicates that this temperature
would be equivalent to an adsorption coefficient of 0.87 cu ft/lb for Regions
I and I and 1.2 cu ft/lb for Region III. For a holdup time of 48 hr and a
flow of 11 scfm, Equation 1 indicates that the required mass of charcoal
would be 12 900 Ib for Regmn 1,10 700 Ib for Region If and 8680 Ib for
Region III. ‘

Desanptlon of 47- Hr Xenon Charcoal Beds

Using the dynarmc adsorpt:on method of calculatmg charcoal reqmrements,

the average adsorption coefficient is affected by: 1) Inside wall temperature,

2) Pipe diameter, 3) Heat flux, 4) Gas flow. For a given heat flux, the inside

~ wall temperature can be changed by altering the coolant water temperature.
Because the heat flux is low, the fluid film temperature drop is negligible,

~ and since the metal thermal resistance is low, the inside wall temperature can
be assumed to be approxxmately equal to the coolant water temperature in

2/ R D Ackley and W E Browning ]r. Equilibirum Adsorption of Kr .and Xe on
Activated Carbon and Linde Molecular Sieves, ORNL internal correspondence
CF-61-2-32 (February 14, 1961).
 

that region. Straight tubes filled with charcoal inside a baffled-delay type
exchanger would thus transfer its decay heat to water in the shell side. This
water would then be cooled by a conventional cooling water system as
shown on Figure 8.3. The baffled-delay type exchanger would be similar to
standard shell and tube type heat exchangers except that the tubes would be
packed with charcoal. One 47-hr charcoal bed would consist of two
half-capacity vessels in series. The first vessel would be designated as Region
I, whereas Regions II and III would be common to the second vessel. These
regions were aribitrarily chosen in order to obtain better estimates of average
sectional heat decay rates.

Because of the higher heat load in Region I, the first vessel would consist of
1-1/2 in. tubes, whereas the second vessel (Regions II and III) would consist
of 2 in. tubes. It should be noted that, within limits, the average charcoal
temperature can be adjusted by the pipe diameter and heat removal
capability. Due to the complex interaction of variables, however, the
optimum system would not necessarily be the one with the smallest mass of
charcoal.

The size of the first vessel is estimated to be 22 ft long by 8 ft in diameter
with approximately 1000 tubes. The mass of charcoal is estimated to be
8400 1b and the length of tubing is 19 000 ft of 1-1/2 in. size. The second
vessel is estimated to be 21 ft long by 8 ft in diameter with approximately
1000 tubes. The mass of charcoal is estimated to be 12 700 1b and the length
of tubing is 17 000 ft of 2 in. size.

There would be six vessels, each pair capable of handling half capacity with
one pair of vessels acting as a standby in the event of operatlonal problems.
The vessels are oriented vertically.

A minimum of two containment barriers are provided to guard against
leakage of the radioactive fission gas into areas which would be hazardous to
personnel. The cooling water heat exchanger capacity is 2.7 kW which is 30
percent over the maximum estlmated heat load

i "The 47-hr charcoal holdup interval was based on a 97 percent attenuation of
' Xenon-135 concentration of the outlet purge gas relative to the inlet. This

criterion was arbitrarily chosen. It appears that this delay time could be

‘substantially reduced without affecting the poison fraction of the reactor.

Calculations have been performed that indicate a 16 hour ‘holdup would
provide adequate decay to allow satisfactory stripping of the xenon from the

~ fuel salt by bubble circulation. In Task II a trade-off study will be performed

to determine the optimum decay period for both the bubble reinjection
decay period and the clean reuse decay period.
 

 

DREMINERALIZED

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Aux Cooling Water System - Flow Diagram

 

 
 

Design Criteria for the 90-Day Long-Delay Charcoal Beds
1) Holdup time for xenon is 90 days.

2) The heat load is 0.25 MW. The average heat load is 2 x 10-3 kW/min holdup

time.

3) The physical properties of the charcoal are the same as those noted in the
description of the 47-hr xenon holdup system, Section 3-C.

4) The gas flow rate is 2 scfm at an inlet pressure of 5 psig. The design Apis 5
psi.

5) The gas composition is 99.9 percent helium, with trace quantities of
contaminants. Since noble gas daughters will be deposited on the charcoal
during operation, there will be gradual reduction on the effectiveness of the

charcoal. About 30 percent spare capacity is provided to offset thls loss in
effectiveness.

6) The heat will be transferred to cooling water. The average temperature of the
charcoal is 125 F.

Description of 90-Day Long-Delay Charcoal Beds

At the outlet of the 47-hr xenon holdup system, the off-gas flow is split into two

- streams, as shown in Figure 8.2. One stream of 9 cfm is returned to the primary system by

the way of the bubble generator and the other stream, of 2 cfm, is fed to the long-delay
charcoal beds. The assumed design residence time is somewhat arbitrary since whatever load
is not handled by the long-delay bed must be dissipated by the gas cleanup system. The
incentive, however, is to handle as much load as possible with the long-delay bed, since its
construction and operation is more simple than that of the gas cleanup system.

" The size of the long-delay bed was estimated using a method similar to that used for

the 47-hr xenon holdup charcoal bed. However, in this case, it is contemplated to use
 helically wound tubing within or external to the charcoal packed vessel. Each long-delay bed

would consist of two vessels in series and oriented vertically. There would be six vessels,
each pair capable of handling half capac1ty w1th oone pair of vessels actlng asa standby in the

“event of operatlonal problems.

The size of each vessel is estimated to be 22 ft long by 5 ft in diameter. The mass of

charcoal in’ each vessel is estimated to be 17 8000 Ib. The average charcoal temperature is
125 F with the decay heat transferred to the cooling water system. Thirty percent spare -

capacity is prowded and any unit may be isolated from the rest of the system..

8-10

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9.0 GENERAL DESCRIPTION OF ELECTRICAL DISTRIBUTION SYSTEM

The onssite electrical distribution system is designed to provide reliable sources of
electrical power during all modes of operation and shutdown conditions. Electrical
equipment rclated to normal opcration of the power station will be connected to their
respective normal electrical buses and the auxiliaries required for safety-features operation
will be supplied from redundant emergency buses. This engineered safety loading is small
and consists mainly of reactor instrumentation, emergency lighting and control-room air
conditioning.

" In the event that all offsite power is lost coincident with a nuclear accident, two
redundant diesel generator sets will provide on-site emergency power to the engineered
safety-feature auxiliaries. Reactor instrumentation will be powered from redundant 120V
a-c instrument buses which in turn will be supplied by inverters from the station batteries.

The electrical distribution system is shown on Figure 9.1 and consists mainly of two
half-capacity start-up transformers which provide start-up power and full-capacity standby
service in the event of auxiliary transformer nonavailability. An auxiliary turbine generator
provides power to hold the reactor in a standby condition. Under normal operation, the
electrical power it generates can be exported via one of the start-up transformers.
Switchgear interrupting capability precludes the simultaneous use of both transformers for
this purpose.

In the event of a loss of offsite power, it is planned that the auxiliary turbine
generator will be rapidly isolated from the electrical transmission system for any occurrance
which would result in a trip of the main station turbine generator. At the same time other
selected loads would be dropped, leaving only those which are necessary for shutdown and
which are within the capability of the auxiliary turbine generator.

9-1
 

 

   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

   

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

 

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.

10.0 FUEL-SALT DRAIN SYSTEM

 

10.1  Purpose

The principal purpose of the fuelsalt drain system is to provide a place where the
salt can be safely contained and cooled under any accidental or intentional situation.
Without impairing the above-mentioned principal function of the drain system, the drain
tank can be conveniently used for other purposes, such as a holdup volume for off-gases to
allow about a 2-hr decay time before the gasses are processed. This arrangement would
reduce the heat load on the off-gas system while at the same time providing assurance that
the cooling system is operable and could accomodate a major drain. The internal surfaces of
the drain tank, particularly the cooled ones, may also act as sites for deposition of noble
metals in the off-gas and will possibly eliminate the need for a particle trap in the
off-system. The drain tank also serves usefully as a surge volume to which salt can be
continuously overflowed from the primary pump bow! as well as transferring fuel salt to the
chemical processing facility for processing independent of reactor operation.

 

10.2  Design Objectives of Fuel-Salt Drain Tank System
The fuel-salt drain system is to be designed to handle:

a) A maximum heat load of about 40 MW(t), if about 7 minutes is
allowed for fuel drainage to take place.

b) A maximum fission gas inventory of about 2.5X108 curies

c) An overflow fuel salt rate to drain tank of 90 gpm at a temperature
of approximately 1200 F.

d) An off-gas flow rate of approximately 11 scfm

e) A maximum transient heat release of about 53 MW(t) which would
- . occur after a sudden salt drain.

f)  About 18 MW(t) during full-load operation.

 

'10.3  Design Objectives of Drain Tank Cooling System

a) It must be able to keep the maximum drain tank temperature well
within the safe operating range even under the worst condition of
‘transient heat loads. | :

b)  The system must be reliable, with a minimum of reliance on the
electric power supply or operator-initiated actions.

10-1
-

c) A double barrier be provided between the tank coolant and the fuel
salt so that leakage of the coolant into the salt would be highly
improbable.

d) The cooling system should impose a minimal risk for freezing of

either the fuel or the cooling system coolant.

10.4  Design Criteria of Drain Tank

 

a) Vessel to be designed for 40 psig, a wall temperature of 1300 F and
meet ASME Boiler and Pressure Vessel ‘Code, Section III, 1968
Edition, Class A requirements.

b) Fuel-salt storage volume of approximately 2500 cu. ft.

c) Internals and supports designed for a seismic disturbance equivalent
to 0.007g to 0.07g horizontal ground acceleration.

10.5  General Description of Drain Tank with NaK-Cooled Heat Disposal System

In addition to eliminating a relatively tall natural-draft stack, as probably would be
required for a salt-cooled primary drain tank system, it is believed a NaK-cooled system

‘would provide a more efficient, dependable and freeze-free emergency cooling system, as

well as offering the following improvements.

NaK can be heated to relatively high temperatures and can experience significant

‘radiation fluences without problems of dissociation or high vapor pressure. Its density and

viscosity variations with temperature are favorable for natural circulation in the system, thus
no auxiliary power or action by the plant operators is required to initiate and maintain
circulation. Use of NaK, and placing primary emphasis on radiant heat transfer (which varies
as the fourth power of the absolute temperature), accommodates the wide ranges of
temperature and heat loads which may be encountered between the normal off-gas heating
load and the maximum transient after a sudden salt drain. The NaK is compatible with
Croloy or stainless steel and does not require the more expensive Hastelloy-N used in the

‘salt systems. Since NaK can have a eutectic temperature such that it will be liquid at room

temperature, no preheating of the NaK circuits prior to filling is required. The only major

~ disadvantage of NaK is its poor compatibility if mixed with the fuel salt. However, the

probability of this happening can be greatly reduced by providing an isolation barrier of a

suitable salt between the fuel salt and the NaK. For this study sodium fluoraborate was
“selected as the isolation barrier salt, and 'a 47 percent NaK composition with a eutectic
temperature of about 60 F was selected as the coolant salt.

10-2
| .

E- o

A simplificd flowshcet for the NaK-cooled drain system is shown in Figure 10.1. The
drain tank is located below and to the side of the primary circuit such that the fuel salt will
drain by gravity whenever the freeze-plug type valve is opened. A small circulation of fuel

~salt is normally maintained in the drain line between the reactor and the freeze valve to

prevent overheating due to stagnant salt.

During normal operation of the reactor a quantity of fuel salt overflows from each
circulating pump bowl. The gases stripped from the fuel salt at the gas separator, laden with
highly radioactive fission-product gases and particulates, are combined with the overflow

salt from the pump bowls in the batch tank (see Section 3.3) before flowing to the drain

tank. The overflow gas-salt mixture, which reaches the drain tank at an estimated
temperature of about 1100 F, enters the top of the drain tank and is first directed beneath
the top head by means of a liner positioned under the head and then downward into the
body of the tank.

The drain tank design is for a vessel about 15 ft in diameter by 20 ft high with the
bottom head containing a plenum for salt-actuated jet pump operation. The jet pumps
transfer fuel salt back into the primary system and to the chemical processing facility. The
drain tank is constructed of Hastelloy-N with approximately 1000 Croloy or stainless steel
open-ended bayonet tubes and mating thimbles extending vertically downward into the
tank, as shown in Figures 10.1 and 10.2. Each thimble in turn is inserted in another thimble,
made of Hastelloy-N. The two thimbles are separated from each other by an annulas filled
with the barrier salt. The eutectic NaK circulates down the bayonet tube and up inside its
mating thimble by natural convection to radiant boilers. A reservoir about 60 ft above the
drain tank provides the thermal driving head for natural circulation.

The drain tank is surrounded by essentially two open-topped stainless steel vessels.
The inner vessel is filled with tightly packed copper rope, the purpose of which is threefold:
to minimize the salt volume which can occupy the annulus in the event a major fuel-salt leak
develops in the drain tank, to provide a good conductor of heat to the outer vessel and to
act as a gamma shield. The outer vessel provides the necessary cooling by acting as a channel
for circulation of the NaK.

It is assumed that about 60 percent of the heat generated within the drain tank is in
the form of gamma rays, much of which will be absorbed by the vessel walls, the copper

- rope and by the thimbles and thus be directly transferred into the NaK. Most of the

generated heat will then be removed by the cooling thimbles. Heat is transferred from the

“outer thimble wall to the NaK cooling system by conduction through the salt isolation

barrier. Since the outer thimbles and NaK cooling components are not in physical contact,
either in the drain tank or in the radlant boilers, a leak in any system is unlikely to
contaminate another. '

10-3
 

 

 

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| Figure 10.1
Drain Tank Cooling System Schematic
 

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Drain Tank Thimble Detail
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The NaK cooling system is arranged with several autonomous circulating loops so
that failure of onc circuit would not cause a scvere loss of cooling capacity and necessitate
an immediate shutdown of the plant. Twelve separate loops, 8 of them tied into the thimble
arrangement and 4 of them connected into the bottom plenum of the outer tank, were
assumed in this study. An clectromagnetic pump (acting as a brake) is installed in each of
the NaK circuits to retard or stop the NaK natural circulation as necessary to protect against
freczing of the fuel salt in the drain tank. This arrangement is particularly advantageous
during start-up or partial load operation.

The radiant boilers have NaK heated inner surfaces separated from the water-cooled
outer surfaces by an inert-gas- filled space. Heat exchange is entirely by radiation. Heat
rejected from the NaK coolant loops is converted to low-pressure steam for either
reclaimation by introduction into the steam power system or for rejection to the
environment by means of a water-cooled condensing basin as discussed in greater detail in
Scction 5.5.7.

104
 

 

 

 

 

 APPENDIXES

 

 

 

 

 

 

 

 

 

 

 

 
 

 

 

 

 

 
T

.. (" (-

APPENDIX A
STRESS ANALYSIS

 

A.1  Background Investigation

 

The investigations which were undertaken in determining the final layout of the
primary salt piping evolved as follows; basically, the system had to meet two criteria:

1) For functional and economic reasons, the overall length of primary piping
had to be maintained at a predetermined minimum.

2) Stresses had to be within limits prescribed by the applicable Codes and Code
cascs for all anticipated conditions.

Two layouts were considered initially; one with the reactor, pump and exchanger in
line and the other with the major equipment forming an approximate right triangle in plan
(the one finally selected).

Initial computer runs were made for the normal, shutdown, scram and
one-pump-out cases with the triangular layout using wall thicknesses of 1-1/2 in., 1-1/4 in.
and 3/8 in. for the outlet piping from the reactor, pump and exchanger, respectively. These
wall thicknesses were based on examination of previous studies.

At this time two horizontal trusses or lateral equipment supports against earthquake
were contemplated. The material for the trusses was to be such that they would exhibit the
same coefficient of thermal expansion as the piping. Despite this factor which tended to
minimize thermal constraints, the piping reactions on the truss were in excess of 165 000 Ib.
Furthermore, excessive stresses existed in the piping.

The wall thicknesses were reevaluated by calculating them for pressure and allowing
for the existence of thermal, seismic and other stresses. This resulted in 1/2 in. for the
reactor outlet, 1 in. for the pump outlet and 3/4 in. for the exchanger outlet. This latter
increase (3/8 in. to 3/4 in.) was an. attempt to ‘“balance” the system which appeared
overstiff in the upper piping section.

Using these new thicknesses the normal, shutdown, scram and one-pump-out cases
were run for both the in-line and triangular layout. Reactions on the truss were reduced to
65 000 Ib with a lesser degree of overstress in the piping. The in-line layout showed higher
(145 000 1b) reactions on the truss and further piping overstress.

Analysis of these calculations made at this time indicated a preference for the
triangular layout and an “overstiff”’ upper piping section for both layouts. Additionally, it
was becoming apparent that the system was sensitive to minor changes in temperature and
geometry.

A-1
 

ity

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I
o St el

r-

The elevation of the reactor relative to the pump and exchanger was varied, while
keeping the overall volume of piping constant, and the four cases plus seismic effects
reevaulated for both the in-line and triangular cases. A position for the reactor was found
which minimized reactions of piping on the trusses and generally gave the best stress results.

At this point, the triangular layout appeared to be superior from a reaction and
stress viewpoint as well as having a minimum overall layout configuration. Further studies
were therefore directed toward the triangular layout.

All of the above studies neglected the growth of the reactor and exchanger supports.
With the reactor now optimally positioned, studies were made to evaluate these factors with
the reactor cell at temperatures from 950 F to 1100 F since holding the cell temperature to
an exact temperature of 1000 F would be virtually impossible—and the system was so
sensitive to these minor differences.

This additional, seeming minor constraint, on the “tight” layout produced high
stresses for certain cases and lead to the following conclusions:

1) The one-pump-out case would be taken care of by the use of pump spares
and bypass lines within the secondary salt coolant loop.

2) The reactor cell temperature would be limited to 1000 F minimum.

With these restrictions, the piping system was within tolerable stress limits for all
combinations of loading conditions. However, the bottom of the reactor support was 1 ft 9
in. above the bottom of the exchanger support. This complicated the layout of the cooling
plenum beneath the reactor cell.

When the system was rerun for the applicable cases with the reactor support at the
same level as the exchanger support, resultant stresses were in excess of the allowable for

SOIMe Ccascs.

The pipe from the pump to the reactor was increased by appro:nmately 7 in., and
length of the pipe from the exchanger to the reactor was increased accordingly. At the same
time, it was thought desirable to use a side outlet from the exchanger to facilitate drainage

of the primary system rather than use the bottom elbow outlet as originally designed. The
- vertical distance between the inlet and outlet on the reactor was also revised to 14 ft 9in.

from the 13 ft 9 in. originally used in the calculations to account for a change in reactor

e head de51gn Appropnate vertlcal plpmg length changes were made to suit. -

All cases were rerun to ascertain the effect of these changes The results 1nd1cated

that the stresses were reduced to within allowable values of the code. The feasibility of the
system, from the viewpoint of the primary salt lqop, was thus amply demonstrated.

A-2
.

It should be noted, however, that the system remains quite sensitive to minor
changes and that more than ordinary care must be taken with fabrication and erection. In
addition, very minor modifications which on an ordinary plant might be undertaken
without extensive study must be investigated carefully in order not to induce an
unanticipated overstress.

A.2 Basic Assumgtions

a) Limit fucl salt volume to minimum permitted by rcasonable pressure drop

and required piping flexibility.
b) Reactor vessel wall temperature to be maintained at 1050 F.

c) Ambient temperature of reactor cell is 1000 F minimum, with possible
temperature variations of 1100 F.

d) The piping system design pressures are as follows (Figure A.1):

Points 4-7 75 psig
- Points  9-10 250 psig
Points 15-20 125 psig
e) Equipment restraint members have the same coefficient of expansion as that
of the piping.
f) Localized stress raisers such as welds and changes in wall thickness at vessel

connections have not been taken into account.

g) Design basis earthquake equivalent to accelerations of 0.25g horizontally,
and 0.16g vertically acting on the piping.

‘h) The piping contains thermal sleeves to minimize thermal transients.

 

- A3 General Description

The piping was analyzed for three o?er_ating conditions, namely: a) normal

~operation, b) shutdown, and c) scram. The temperature conditions associated with these
“modes of operation are all shown in the figure. The loss of a secondary coolant salt pump

was analyzed as a separate case. The results of this latter analysis indicated that piping

~ stresses considerably above allowable would be introduced into the primary piping system.
" Therefore, for this study it is assumed that provisions will be available whereby the loss of a

secondary coolant salt pump will not affect primary piping stresses. Such provisions could
consist of pump spares and by pass lines within the secondary coolant salt loop.
 

|

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Figure A.1
Primary Piping Isometric

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C

The piping was analyzed for conformance to equations 9 and 10 of Section I-705 of
the Code for Pressure Piping, ANSI B31.7, Nuclear Power Piping. The analysis was made on
an elastic basis using the basic stresses for Hastelloy-N per Code Case 1315-3 dated April 25,
1968 and the appropriate paragraphs of Code Case 1331-4 dated August 15, 1967,

Using the equipment and piping temperatures shown on Figure A.1, the analyses
were made for a reactor cell temperature of 1000 F with possible variations in cell
temperature of plus or minus 50 F.

All points of the piping system meet the requirements for primary membrane stress
intensity (equation 9). Under operating conditions the system meets the requirements for
primary plus secondary stress intensity range (equation 10) with the reactor cell at 1000 F
except for a slight overstress of 346 psi at point 10.

With the reactor cell at 950 F, an overstress of 1000-2000 psi exists at points 5, 9,
901 and 10 under the operating condition. Additional analyses under “Simplified
Elastic-Plastic Discontinuity Analysis,” may show these points to be within tolerable limits
without major piping changes. However, rather than subject the system to further analyses
at this time, the reactor cell environment will be limited in this report to a minimum
temperature of 1000 F. |

One fact revealed by the piping analyses performed to date is the piping system
sensitivity to small displacements. For example, small changes (one to two ft) in the relative
heights of supporting skirts under the reactor and the primary heat exchanger cause sizable
changes in the piping stresses. Similarly, fluctuations in reactor cell temperatures which
effect component skirt and horizontal restraint expansions act in a like manner. Satisfactory
conformance to stress intensity limits indicated herein are therefore predicated upon the
assumed shell temperatures of the major equipment. Final designs must be checked against
actual shell temperatures obtained by a more detailed study.

Further, the deflection sensitivity of the system dictates that care must be taken
with fabrication procedures to minimize “locked in” erection stresses.

A thermal stress analysis was conducted to investigate the maximum thermal
transients that might occur in the primary piping system. The transient thermal stress
conditions were, of necessity, established for the worst possible cases - e.g., (1) assuming
that the changing of conditions between operating (1300 F) and scram (1050 F) from

points 9 to 10 was almost instantaneous (approximately 10 sec); (2) assuming that upon the
loss of a secondary coolant pump, the reactor fuel salt inlet temperature would rise from

1050 F to 1300 F in approximately 30 sec.

The computer program eniplojr_ed, uses input information describing the geometry,
thermal properties, and boundary conditions for an orthotropic or axisymmetric body and
calculates the transient temperature field resulting from specified heat sources on a finite
element formulation basis. The temperature distributions so generated were used to

A4
T T 1 L ek et it ATt e
1

 

fm-m pretemeny
W et

 

calculate the absolute AT temperature stresses in accordance with equation 10.

For Case 1, an additional thermal stress of approximately 23 000 psi occurred after
14 sec between points 9 and 10. For Case 2, a maximum thermal stress of approximately
20 000 psi occurred after 31 sec between points 15 and 20. Stresses of these magnitudes,
superimposed on the other stresses to be considered for equation 10, would far exceed the
allowable values for primary plus secondary stress intensity range of 3 Sm. Based on these
conclusions, a thermal sleeve design was selected and analyzed using the same program as
previously described, but with a few modifications. The thermal stress analysis was

- performed assuming a bypass flow of 30 gpm contained within a 1/2 in. annulus, the

annulus being formed by the L.D. of the primary piping and the O.D. of a 3/8 in. thick
thermal sleeve. To keep the fuelsalt inventory constant, the thermal sleeve sizes were
limited from 16 to 21 in. in diameter. The 30.5 gpm fuel salt bypass flow was not
considered as adding to the fuel-salt inventory.

With the thermal sleeves incorporated in the piping system, the maximum
temperature difference across any pipe wall caused by any transient is less than 3 degrees
and results in negligible stresses.

A summary of the pressure containing pipe sizes, allowing for thermal sleeves, are as
follows: |

 

 

Wall
Points Sizes Thickness
47 23 in, O.D. 4 in.
9-10 19 in. O.D. 1 in.
15-20 18% in. O.D. 3/4 in.

The above pipe wall thicknesses have been shown to be adequate for the conditions

“and analyses cited above, but may be modified by future optimization studies.

¢

A5
 

 

 

 

 

 

 

- APPENDIX B

 

- Nunr_l_bei", ,67f Steam Geriera_tofs and Reheaters IR L

 
Py

 

A study was conducted to assist the system designer to determine the optimum
number of steam generators and reheaters in the system. For the steam generator, 4, 8, and
16 units were sized, the pressure-drop ranges were calculated, and the amount of material
was estimated (see Figure: B.1). A similar study was performed on the reheater but1,2,4
and 8 units were considered. -

The results of these studies can best be understood if a few basic points are
explained. Components are generally designed to utilize the maximum pressure drop
allowed by the system’s designer, and this design tends to ensure a most compact unit. On
this basis, whether the design is for 4 units or 16 units, the allowable pressure drop will be
the same. If the tube size and tube spacing are the same, the use of the same pressure drop
will ensure identical fluid velocities in all units. Accordingly, if 4 units are designed to the
same pressure drop, tube size, and tube spacing as 16 units, the lengths will be equal; the
difference will lie in the diameters. Another way of interpreting this is to imagine 4 units,
cach with 1000 tubes. These would obviously be equivalent to 16 units, each with 250 tubes

of identical size and arrangement. Although the optimum selection may be slightly different

for each number of units, the studies showed that this analysis gives an accurate picture of
the problem. The conclusions are as follows:

1) The steam generator can be built and shipped in 4, 8, or 16 units.

The amount of material required is not dependent on the number of

units and is approximately constant. The major disadvantage of 4

units is that the tubesheets tend to be very thick (~25 inches). The

major disadvantage of 16 units is that the total manufacturing costs

- will be slightly greater due to extra numbers of components. Also the
larger number of units (8 or 16) w111 cause extra plpmg problems,

2) ‘The reheater can be built and shlpped in 1, 2, 4, or 8 units; the same
conclusions that apply to the steam generator apply to the reheater.

3) - After dlscussmns with Ebasco on the results of the study, it was

“mutually agreed that it became basically a choice for the system
- designer, and consequently 4 units were chosen for the steam
_generator and the reheater, -
3AIND) DLIIFWEIE] ‘SIOIRIDUIL) WESIS JO JIqUINN

1°q 2mdig

Tube Bundle Length, ft

 

 

 

 

 

 

 

 

 

30

r . . . E E © r e r r
70
0.25" Ligament
) — = — — (0,375" Ligament
60 Tube-side AP, psi
| sremosmeer Shell-side AP, psi
Hastelloy-N inventory/unit, ft?
' 60
50
~N
~ .50
~
-~
- 20
40 s 8Q0
350 g s =10 1000
S ~ -
v S~
. ~ Z ~
~ 5 ™~ AT
30~ ~ --.
A~ =nX
os?foo ’
20 :
16 Units 8 Units 4 Units
10 : i ] L 1 1 i ] L I
40 50 60 70 80 90 100 110

Tube Bundle OD, in.

120

   

 

 
 

  

 

 

  

 
 

 

E | e e T

P EREOAT e TR meme T

ef- -

- -, |
- - APPENDIX c

H TR o Flow Stabxhcy of Supercntlcal Fllnd

 

 

 

 
 

-

T (

r

One of the major problems in steam generators is the flow instability of the heated
fluid. There are two major types - dynamic instability and flow reversal. Dynamic instability

s characterized by rapid pressure and velocity fluctuations and can take place over a wide

range of loads. Flow reversal may occur at very low loads and result in some tubes having
fluid flowing opposite to the desired direction.

The dynamic stability characteristics of a steam generator can only be evaluated by
using highly complex analysis; however, a preliminary analysis can be made by assuming

 quasi-steady conditions. The static stability curve is shown by the example in Figure C.1 and

is drawn for a given load condition (say, 20 percent load, for example). Upon receipt of
operating temperature conditions, the steam generator designer can then calculate the

required mass flow to satisfy load and temperature conditions; this is indicated by Mpin-

the figure. The next step is to find out how the P varies with small changes in mass flow. If
the gradient is positive (such as C to D) in the region of MA, the load, flow, and temperature
conditions will most probably be stable. If the gradient is negative (such as B to C), the

- dynamic instability will probably occur during operation.

Rough calculations of the MSBR units have been made on the static stability
aspects. It was assumed that the mass flow was proportional to load or, in other words, a
constant AT across the unit. The results indicated that the minimum flow would be in the
range of 10 to 15 percent to avoid dynamic instability.

Flow reversal generally occurs at very low loads and is caused by the density change
in the up-leg of a loop, as shown in Figure C.2. The cold leg and hot leg may be inside the
steam generator shell, or the cold leg may be part of the piping system outside the unit. The
heated up-leg creates a change in density such that the outlet pressure (P3) is higher than the
inlet pressure (Pq1). A situation is created where it is possible for some tubes to have flow in
one direction while adjacent tubes have flow in the opposite direction. The solution is to
increase the mass flow until the friction AP equals or exceeds the pressure gain caused by the

density change.

The initial calculations were based on the same assumptions as previously stated;

that is, the AT across the unit was con_staht at all _loads.' It was found that to avoid flow
‘reversal, the minimum mass flow would probably be in the range of 10 to 15 percent of full
“load. o L

| . To avoid these large mass flow requirements the graph in Figure C.3 was prepared.
This graph shows that flow reversal can be avoided with mass flows of 5 percent if the inlet

~ temperature is increased to 750 F. The reason is that the flow reversal problem is aggravated.
~ by high density changes. Since a large portion of the density change occurs between 700 and

750 F (34 "lbl_ft3 to 12 lb/ft3);,, the solution is made easier by increasing the inlet

temperature to 750 F.

As it has not been checked, it can be stated with a high degree of assurance that the

conclusion of Figure C.3 can be applied to the dynamic stability analysis; in other words, -

the unit will by dynamically stable with smaller mass flows if the temperature range is
decreased.
 

ey
g

o

e

- K

-

r e

Tube -Side AP

 

 

Constant Load

 

 

 

- Mass Flow

If the unit is bperéted at mase flow corresponding to point
A, instability occurs. Any perturbation has the potential
to cause oscillation from region B to region C.

- Note the rnulti«\-v_jaified mass flow of AP,. This is a char-
acteristic of flow which has a high density change.

Figure C.1
Characteristic Flow Instability
 

N

ittty

rr. r.

o

1
Cold Leg Hot Lleg
Downcomer Riser
L
static pressure,
" | density,

 

 

friction factor,

o M O W
"

®

dismeter,

velocity head.

]
H

L
PZ = Pl + PI-ZL - (f ‘l—)' VH)I—Z.

L
Py = P = pa-3L - (f D VH)Z—B.
Py = Py = L(p1-p — p2-3) - | (EZ ¥ +(ekv
'P3 - P1 = L(p12 — P2-3. 5 Vi) _, 5 V) |

At low flow rates [ ] > 0.

. P3 - P} = L(py1_2 - P2-3).

If pé;3 < p)-2, then P3 > Py~ (reverse flow).

Figure C.2
Typical Flow Circuit, Schematic Diagram
 

r

. .

Inlet Temperature, F

900

800

700

Outlet Temperature = 1000F

Full Load AP = 300 psi

   

Full Load AP = 200 psi

5 10 15 20

‘Mass .FIOW, % full load

: | Figure C.3
Minimum Mass Flow Required to Prevent Flow
Reversal at Low Load

 
 

e

u

 

The conclusions are as follows:

1)

2)

_3)

"With an inlet ‘te‘mpera"turre of 700 F.and an outlet temperature of

1000 F, the steam generator will need minimum tube-side mass flows
of 10 to 15 percent of full flow to ensure freedom from flow reversal
and dynamic instability problems.

If the inlet temperature is raised to 750 F, the steam generator may
be free from flow reversal and stability problems at mass flows as low

~ as 5 percent of full flow.

‘The relatively large variation in outlet temperature does not affect

the minimum mass flow requirement for a stable operation because
the density - variation near the outlet conditions is negligible
compared to the inlet conditions.

C-2
 

.

 

 

 

 

 

 

 

 

5 .;._;:APPENDIX D

Stress 1n Tubes Due to D:fferentlal
Thermal Expansmn )

 

 
 

_-1). General

This appendix gives a comparison of the thermal stresses in the types of tubes which

~are considered for the IHX. These types are straight, sine-wave-bend, hockey-stick, C, and

hockey-stlck/sme tubes.

All tubes are stressed due to the difference in the average temperatures of tubes and
shell. It is assumed that the axial temperature variations in tubes and shell are linear and that
the tubes are forced to expand the same amount as the shell would expand with no
constraints. In other words, the tubes’ stiffness is negligible compared to that of the shell.

- All types of tubes have the same cross-sectional dimensions.

2)  Straight Tube

 

 

L

The formula for the axial force acting in a straight tube is

F = AE a AT

o0 = F/A = E o AT
= (25 x 10%) (10 x 1076)(87)
= 21,750 psi

elastic modulus,
a = coefficient of thermal expansion,

AT = temperature difference between gshell and tubes,

where | E

o = gtress,
A= cross-sectional area of the tube.

-‘The value of 1 5Sm at 1300 F is 5250 p51 This value indicates that a strmght tube wfll be

thermally overstressed.

3) 'lsifie-Waiie-Behd' Tubes -

 

 

 

 

 
i
N

r

ot

o i

8

r "

|

D-2

 

 

e E a AT 5
L1+L2+R(1+2"+3R
. I

where _ E = elastic modulus,

o = coefficient of thermal expansion,

AT = temperature differential between
shell and tubes,

A = cross-sectional area of tube,
I = moment of inertia of tube,
v = Poisson's ratio,
Ly, Ly, R = defined in picture,
Ly + Ly + 4R = 300 inches.

The maximum bending stress in the tube occurs at the midpoint of the sine wave. For this
reason the wave will be located in as cool a region as possible. The maximum bending stress
and the P/A stress for different values of R are shown in Figure D.1, By comparing these
stresses to 1.5 Sm at 1100 F, it is shown that R should be greater than 4 inches.

4)  Hockey-Stick Tubes

My
V—-/

L
R

 

 

 

 

L

The assumptions are that the forced tube displacement is the same as the free

thermal expansion of the shell at point A, and that there is no rotation at A. A force V and
 moment M are developed at A to satisfy these assumptions. The maximum stress occurs at
_point A. Graphs of stress versus R, Ly, and L, are shown in Figures D.2, D.3, and D.4. The
_tube dimensions and T are the same as for the sine-wave tube. The stress for a tube with a

different OD is obtained by multlplymg the stress shown on the graphs by the ratio of the

new to the old OD. By
Mg

I

The graphs show that the maximum bending stress varies significantly with R and L,
and is practically independent of Ly and R should be greather than 25 in. to have
stresses below 1.5 Sm at steady state. This leaves an additional 1.5 Sm for stresses that occur
during transients since the total stress intensity range allowable for thermal secondary type
stresses is 3 Sm.

 

5)  C Tube
Ll | | El
R _ .
. Ly -t Lo ———-I

 

 

 

A C tube can be thought of as being made up of two hockey-stick tubes. Each end
will be forced to displace about one-half as much as the end of the hockey-stick tube. Figure
D.4 shows that the stress in the hockeystick tube is practically independent of Lo;
therefore, the maximum stress developed in the Ctube will be about one-half of that
developed in the hockey-stick tube. This means tléat_ the sum of L; and R should be greater

than 12 in. to have maximum stresses below 1.5 °m.

6) Hockey-Stick/Sine-Wave-Bend Tube

The model for the combination hockey-stick/sine-wave-bend tube is shown below;
the distance L, + L3 + 4R is 300 inches.

Point 1
Point 2 —r
- . Ly

"R R

R\, | |
L‘r Lzr — ‘J

 

 

 

 

 

'Lra"-

“The stress at points 1°and 2 is ShoWn '.i'n. Figures D.5 and D.6. It is evidefit that the stress at

 point 2 is greater for a given combination of R and L. Figure D.6 shows that the sum of L,

and R should be greater than 14 in. to have stresses less than 1.5
 

Stress, ksi

80

70

60

50

40

30

 

 

 

 

 

“— VINE

L; + L, + 4R = 300 in,
Tube OD = 0, 375 in,
t = 0,06 in.
A = 0.0594 in,2
1 =0.00076 in.*
Ea = 250 1b/in.2-°F
AT = 87°F

1.5S., at 1100°F

 

 

 

 

Stresses in Sine-Wave Tubes Vs Radius of Curvature
 

Stress, ksi

 

 

 

  
   

L1= 15, Lz= 50
~L, =15, Ly= 300 L1123, Lz=50
/"L1= 5, Lzz 300

  

L1=30,__Lz= 50
'/L1=30, Lz-"; 300

 

 

 

R, in.

- ~ Figure D.2
Stress in Hockey-Stick Tube Vs Radius of Curvature
 

 

’l. 585, at 1300°F

 

/

il
Stress, ksi
KN

  

/

      

 

    
       
 
        
 

R=15, L,= 50
& R=15, Ly=300

/-R=5, Lz=50
R=5, Lz=300

-

R =30, L,=50
R =30, L, =300

 

 

o 5 10

15 20 25 30

Length L, in.

Figure D.3 ‘
Stress in Hockey-Stick Tube Vs Length L,
 

 

/-Ll=15, R=5

 

/1. 5 Sm at 1300°F

 

Stress, ksi

L1=30, R=5

 

L,=15, R=30

 

 

. L,=30, R=30

 

 

 

05 100 T - 200 , - 300

\ v, Length L;, in.

Figure D.4
Stress in Hockey-Stick Tube Vs Length L,
Stress, ksi

 

Ll=20

 

 

/'1. 5Sm at 1300°F

 

 

Ll= 15

  

Ll =30

L, =10~

 

 

R, in.

Figure D.5 -
Stress in Hockey-Stick, Sine-Wave Tube Vs Radius
of Curvature, Point 1
 

Stress, ksi

 

L1=25

Ll - 20

1.5 Sy at 1300°F

 

 

L, = 30

 

L1=15

  
 

\

L1 =5
L, = 10~/

 

 

- R, in.

- Figure D.6
Stress in Hockey-Stick/Sine Tube Vs Radius of
: | Curvature Point 2