ND-74-66.txt

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FOSTER WHEELER CORPORATION

NUCLEAR DEPARTMENT

110 South Orange Avenue, Livingston, New Jersey

Best Copy Available

Page 6-92 Missing
FWC FORM 172 - L4

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NOTATIONS IN THIS COLUMN INDICATE WHEERE CHANGES HAVE BEEN 1

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT s LIVINGSTON, N. J.

CHARGE NO. DOCUMENT NO. ISSUE DATE

TASK I FINAL REPORT
DESIGN STUDIES OF STEAM GENERATORS

FOR MOLTEN SALT REACTORS
REPORT ND/7Li/66

P‘,J(‘ "'r.‘:‘:“‘i.".'."’v Bl [ |

UCC PURCHAGE ORukil SUBCOUNTRACT NO., 21X-68070C

Approved by:

J. F. Cox
Project Manager

Pt ) n
/’% Lo / (/

W. Wolowodiuk
Program Manager

- 7 / ,

fl /‘/i/QGd‘

D. H. Pai
Chief Engineer

FOSTER WHEELER ENERGY CORPORATION
EQUIPMENT DIVISION-NUCLEAR DEPARTMENT
110 SOUTH ORANGE AVENUE
LIVINGSTON, NEW JERSEY 07039

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¥WC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO. 8-25-2),31 | DOCUMENT NO. Np/7L/66 ISSUE 1 DATE 12/16/7L

DISTRIBUTION LIST

Union Carbide Corporation

J. L. Crowley (30)

Foster Wheeler Energy Corporation

R. 0. Barratt
to J. K. 0'Donoghue
to Nuclear Files {Retain)

W. Wolowodiuk
to J. F. Cox (Retain)

D. H. Pai
to C. Nash
to H. Levy (Retain)

S. M. Cho
to H. L. Chou (Retain)

J. Anelli
to C. Holderith (Retain)

W. R. Apblett (2)
to E. D. Montrone (Retain)
to G. V. Amoruso (Retain)

M. J. Kraje
to J. G. Whelley (Retain)

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND/7./66

ISSUE

1

DATE 12/16/7L

"LEGAL NOTICE"

"This Report was prepared as an account of Government sponsored work.

Neither the United States, nor the Commission,

on behalf of the Commission:

nor any person acting

() makes any warranty or representation, express or implied, with
respect to the accuracy, completeness or usefulness of the in-
formation contained in this report, or that the use of any in-
formation apparatus, method or process disclosed in this report
may not infringe privately owned rightsy or

(b) assumes any liabilities with respect to the use of, or for damages'

resulting from the use of any information a

process disclosed in this report.

pparatus method, or

"As used in the above, 'person acting on behalf of the Commission'
includes any employee or Contractor of the Commission or employee
or Subcontractor of such Contractor to the extent that such employee
or GContractor of the Commission or employee or Subcontractor of such
Contractor prepares, disseminates, or provides access to, any infor-
mation pursuant to his employment or contract with the Commission

or his employment or subcontract with such Contractor, "

BY

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO., 8-25-2L31 | DOCUMENT NO. ND/7L4/66 | ISSUE 1 DATE 12/16/7L

NOTPATTONS

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Section No. Title Page No.
I Title and Signature Page Cover
II Distribution List a
ITI Legal Notice b
IV Table of Contents c
1.0 Abstract 1-1
2.0 Introduction and Concept 2-1
Arrangement Study
3.0 Mechanical Design Report 3-1
4.0 Thermal /Hydraulic Design Li-a
5.0 Structural Feasibility Analysis G-a
6.0 Hastelloy N Steam Corrosion 6-a
7.0 Manufacturing Engineering T-a
Appendic¢ies
A Design Calculations
B Thermal /Hydraulic Calculations
C Structural Calculations

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N, J.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. Wp/7)/66 ISSUE 1 DATE 12/16/7)

1.0 ABSTRACT

This Task I Final Report of Design Studies of Steam Generators
for Molten Salt Reactors presents the work done by Foster Wheeler
in accordance with Task I of the Work Plan Detail by Foster
Wheeler for Design Studies of Steam Generators for Molten Salt
Reactors under Purchase Order Subcontract No. 91X-88070C during
the period October 1971 through December 197Lh. The effort was
conducted under two different Foster Wheeler contract numbers
due to a termination for the convenience of the government on
January 31, 1973. FWC Contract No. 2-25-1352 covered the work
performed from October 7, 1971 through January 31, 1973. FWC
Contract No. 8-25-2,31 covered the work performed from May 17,
1974 through December 31, 197L4.

The Foster Wheeler design concept presented in this report is
an "L" shaped tube and shell heat exchanger to be used in a
steam generator system consisting of four wnits for a 1000
MW(e) Molten Salt Breader Reactor power plant.

The design utilizes an all welded construction with 100% radio-
graphy of all pressure boundry welds. The pressure shell is
designed to minimize the annulus between itself and the tube
bundle. Where the shell must be of a larger diameter a flow
shroud is used to maintain the flow over the bundle. Suitable
water, steam and molten salt inlet and outlet nozzles are
provided. The mechanical design provisions for the steam
generator are covered in Section 3.0.

The design work is supported by thermal/hydraulic and structural
analysis as reported herein. The thermal/hydraulic analysis
described in Section L.0 demonstrates that the present unit size
is sufficient for the intended service with minimal possibility
of the molten salt solidifying on the cold end tubesheet or tubes.

Structural analysis of the proposed design has been conducted

and is reported in Section 5.0. An elastic analysis on major

components was conducted. Also, simplified inelastic analysis
was conducted on the salt inlet nozzle, the outlet tubesheet,

the shell and the tubes. Based upon the stress analysis con-

ducted, the design is satisfactory.

Outline versions of a possible manufacturing and inspection plan
and maintenance procedures were prepared and may be found in
Section 7.0. '

References are provided as applicable and supporting calculations
are provided in the Appendices.

BY

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J,.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND/7./66 ISSUE 1 DATE 12/16/7L

2.0 INTRODUCTION AND CONCEPT ARRANGEMENT STUDY

2.1 INTRODUCTION

This Task I Final Report presents the work done by Foster Wheeler
during the period from October 7, 1971 to December 31, 197L on
Design Studies of Steam Generators for Molten Salt Reactors.

This work was performed in two segments due to a termination for
the convenience of the govermment on January 31, 1973. The
first segment ran from October 7, 1971 to January 31, 1973.

The second segment was begun on May 17, 197L and ends with this
report. This second segment limited the work scope to the
completion of Task I.

This document is submitted in compliance with the requirements
of Article I - Statement of Work of Purchase Order Subcontract
No. 91X-88-70C dated October 7, 1971 as ammended by Supplemental
Agreement No. 3 dated May 17, 197L. This Statement of Work is
further defined in the Work Plan Detail by Foster Wheeler for
Design Studies of Steam Generators for Molten Salt Reactors

dated October 7, 1971 and attached to the Subcontract. All

codes and standards applicable to this effort have been established
by mutual agreement under Sub-Task I of Task I of the work plan
defined above.

Portions of the information contained herein have been submitted
to Union Carbide in progress reports 1 through 12 which were
submitted as required under the subcontract.

BY

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/7L

2.2 CONCEPT ARRANGEMENT STUDY

A concept arrangement study was performed to arrive at a design
concept that offered the most promise. The starting point for
this study was to establish the design requirements and criteria
through consulting with Union Carbide and then to evaluate posi-
ble surface arrangements that might meet these requirements and
criteria. For clarity and completeness,the design requirements
and criteria established for Task I are presented below.

DESTGN REQUIREMENTS AND CRITERIA

1. The inlet and outlet salt temperatures, the maximum pressure
drops, and the total heat transfer capacity of the steam genera-
tors must conform with the overall system operating conditions.

a) Full load operating conditions (one of four coolant salt
loops - thermal duty = L83 MW):

Coolant Salt Water/Steam
Inlet temperature, F 1150 700
Outlet temperature, F 850 1000
Flow rate, Lb/hr 15,280,000 2,517,000
Inlet pressure, psia 235 3800
Outlet pressure, psia 175 3600
Maximum pressure drop, psi 60 200

The design temperatures and pressures are left to the discretion
of the designer.

More than one module per coolant salt loop is permissable. If
modules are used in parallel, they will share the coolant salt
loop duty and flows equally. The maximum water/steam side

pressure drop of 200 psi may be relaxed if found to be unneceg-
sarily restrictive.

b) Part load operating range at constant steam generator outlet
pressure is defined as any condition between 20 and 100% of full
load thermal duty. In this load range the coolant salt flow will
be varied linearly from 30% flow at 20% load to 100% flow at 100%
load. The feedwater inlet temperature will be maintained cor-
stant at 700 F over this load range while the water/steam flow
will be varied in proportion to load. The turbine inlet steam
temperature must be maintained at 1000 F + 15 over this load range.
If an attemperator is required to maintain the turbine inlet
temperature, then the attemperator design becomes a part of the
steam generator design.

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-0);31 | DOCUMENT NO.ym/7),/66 | ISSUE 1 IDATE 12/16/7L

c) Startup operation of the MSER power system is defined as

zero to 20% of full load. However, the primary fuel salt
requires that initial zero power operation must begin with both
the fuel and coolant salt systems circulating isothermally at
1050 F.

d) The steam generator must be able to operate at all loads with
tolerable thermal stresses and without freezing the coolant salt.
Alternately, operation with a frozen salt film may be permissible
if desirable and if operation can be shown to be stable,

2. The type of steam generator, the general location of nozzles,
the height of the unit, and the minimum tube diameter must be
compatible with various design, layout, fabrication, maintenance
and inspection considerations.

a) The steam generator shall be a once-through, shell and tube
heat exchanger with the coolant salt on the shell side and the
water/steam in the tubes.

b) For purposes of the steam generator design, the coolant salt
system is assumed to have forced circulation during all expected
operation. Natural circulation on the shell side, although a
desirable feature, is not a design requirement. The tube side
may have forced or natural circulation under decay heat removal
conditions.

c) There is no height limit on the steam generator.

d) As a guideline, a minimum tube ID of not less than 0.375
inches shall be used.

e) There are no physical layout limitations.

f) The shell side of the steam generator shall be completely
drainable. ,

g) A tube plugging allowance of 5% but not exceeding 25 tubes
shall be used for preliminary sizing purposes.

h) Heat transfer surface may be arranged for vertically up
or down tube side flow.

3. The steam generator should be arranged for relatively easy
tube bundle replacement or modular replacement. Both seal-welded
flanges and cut and weld removal techniques shall be considered
for easy tube bundle replacement.

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FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7,

L. The space requirements for the installation of the steam
generator should be as small as Practical commensurate with
economic considerations., It will be necessary to have some 3
data for incremental building and excavation costs, | f

5. The volumes of both the salt and water/steam in the unit should
be kept as low as practical.

6. The use of cover gases to protect structural members should
be avoided.

7. The following items shall be considered and accomodated in
the design of the steam generator:

a. Baffle and tube supports, as necessary, to prevent
damaging vibration.

b. Relative expansion between the shell and tube bundle and
between individual tubes.

c. Tube side flow stability under all load conditions.

d. Protection of nozzles and tubesheets against excessive
thermal stress due to transients.

e. Fabricability of the design including the ability to
radiograph all containment welds.

f. Relief of the salt-steam mixture resulting from a double~
ended steam generator tube failure to limit damage to the
steam generator. The remainder of the coolant salt system
can withstand a contimuous pressure of 220 psi without
damage. '

8. The steam generator shall be designed as a Class 1 vessel in
accordance with the applicable portions of Section ITT of the
1971 ASME Boiler and Pressure Vessel Code with addenda and
the other design standards listed in the corrected Appendixes
A and D of the Company's RFP, Enclosure 2. The design shall
include the Code consideration for operation in excess of 800
F over a 30 year design life at 80% plant factor.

9. The design of all portions of the steam generator in contact with
the salt shall be based on Hastelloy N material. The physical {
Properties of this Hastelloy N are listed in the corrected Appendix
C of the Comapny's RFP, Enclosure 2,

10, The steam properties on which the steam generator design is
based shall be obtained from the 1967 ASME Steam Tables.

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FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,,31 | DOCUMENT NO. ND/7./66 ISSUE 1 DATE 12/16/7L

11. The steam generator design shall be based on sodium fluoroborate
as the coolant salt. The physical properties of sodium fluoro-
borate are listed in the corrected Appendix B of the Company's
RFP, Enclosure 2. The substitution of a different salt with
a higher temperature melting point may be the subject of a
later study, :

12. A corrosion allowance of % mill/year on the salt side and
7 mill /year on the steam side shall be used in gelecting
material thickness.

13. It will be permissible to divide the steam generator into two
subunits in series if the design study indicates an advantage.

1L. The design of the steam generator shall be based on a 30 year o
Plant life with the steam generator experiencing 10,000 cycles
of 20% or more in power over its life., Normally the load will
be changed at a maximum rate of l%/min. The details of number
or reactor scrams etc. and the resulting temperature Tramps
will be made available.

As an additional requirement not bresented sbove, it was felt that
the use of bellows in the shell of the steam generator to accomodate
differential thermal expansions was to be avoided.

Within the restrictions set forth above, a large number of surface
arrangements were possible. Therefore, a second set of seven basic
design areas were setup. These seven areas are, in order of importance:
l. Adequate thermal expansion provision - does thig design have |
sufficient tube flexibility to allow for large temperature
difference between tubes and shell or between adjacent tubes?

2. Stratification Problems - are there posgible areas within the
unit where predicting flow will be difficult?

3. Good Utilization of Volume - does the design allow for all the
tubing to be used as active heat transfer surface?

i. Better Mechanical Arrangement - do the mechanical details appear
to be difficult?

5. Equal Active Circuit Length - will the tubes have approximately
equal heated lengths?

6. Tube Side Inspection - can the tubes be easily nondistructively
examined after the unit has been in service?

1BY | APPROVED | PAGE~ 2-5

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT _ LIVINGSTON, N. J.

CHARGE NO. 8-25-2};31 | DOCUMENT NO. ND/7../66 ISSUE 1 DATE 12/16/7)

1. Relative Complexity - do the uncertainties of the design appear

to lend themselves to straight forward solutions?

Page 2-7 shows the various possible surface arrangement candidates

and how they evaluated against the seven basic design areas. On

the chart,no line indicates that the candidate did not affirmatively
bass that design area., A solid line indicates that the candidate does
bass that design area. A dash line indicates that the candidate may
be able to pass that design area with difficulty. When a candidate
did not pass a basic design area (no line), it was dropped from
further consideration.

The above describe evaluation reduced the possible number of candidates
to 7. These seven candidates and some of their more difficult detailed
problems are given in on Page 2-8. As Page 2-8 shows the principle
problems relate to the fabrication of the U-bend or Elbow shell.

Pages 2-9 through 2-12 further outline the fabrication problems of

the U~bend shell and the e¢lbow shell,

As can be seen from the foregoing discussion it became apparent that
an elbow unit showed the most promise in meeting the design require-
ments and criteria with the least amount of problems. Thus the
elbow unit was selected as the reference design.

BY

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MOLTEN SALT S7EAMGENERATOR SURFACE ARRENGEMEN T

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Page 2-7
FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

(CHARGE NO. 8-25-21,31 | DOCUMENT NO. ND/7h/66 ISSUE 1 DATE 12/16/7L

DETATLED PROBLEM AREAS

1) The sine wave expansion bend will

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be excessively large.
2) Complicated tube supports/vibration

problems or the introduction of a

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large ineffective volume.
1) The fabrication of the U-bend is
-
very difficult.
w 2) Complicated tube supports/vibration

problems or the introduction of a

large ineffective volume.

1) The fabrication of the U-bend is twice
-
_ as difficult as the J tube unit.

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'2) There is no reasonable way of avoi-

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ding the U-bend, vibration, built-up
—
support problems.

1) The fabrication is simplified compared
with the other candidates but is still

difficult.

BY

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NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND /7 /66 ISSUE 1 DATE 12/16/L7

U-BEND SHELL FABRICATION

- Long shells are longitudinally split members.

- Long seam welding performed in situ,

- Weld distortion will cause the shells to go out of round.

. Subsequent Teé-preparation of circle seam welds requires

Special development.

+ Clam shell members for the U-bend will be positioned and

the long seam welded in situ. Circle Seam welds will require

re-~-preparation using similar special equipment.

- Final closure at each tubesheet will be done by the addition

of transition pieces which are custom machined and include

hand holes for completion of NDT procedures and repair if

necessary.

- These problems are somewhat simplified when a small shell

diameter are considered.

BY

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FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2),39 DOCUMENT NO. ND/7L /66 ISSUE 1 DATE 12/16/7)

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

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FWC FORM 172 -

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FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2);31 | DOCUMENT NO.ND/7L/66 | ISSUE 1  |pate 12/16/7L

ELBOW SHELL FABRICATION

1. The fabrication process utilizes shell sections without
recourse to clam shells, thereby avoiding the problems

of making and inspecting long seam welds in situ.

2. The problems associated with suppofting the tubes in the
bend area is not considered any more difficult than in the

case of the U-bend shell.

3. The installation of the tube bundle will require threading
of the tubes through the shell after the supports have been

installed.

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| BY APPROVED | PAGE  2-11

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NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

CHARGE NO. 8-25-21,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7L

s

ELBOW SHELL ASSEMBLY

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CHARGE MO g_o5_p),31 | DOCUMENT M0y 7,66

ISSUE

DATE 15 /16/7),

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| SECTION 3.0
. MECHANICAL DESIGN REPORT
PRELIMINARY DESIGN
by

Charles H. Holddtrith

Ajpproved By: .,

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FOSTER WHEELFR CORPORATION

CHARGE NO o . 3, ., | DOCUMENT NOw o . |ISSUE . DATE, 5 /16 /71,
TABLE OF CONTENTS
{ 3.1 Design Description Summary 3-L
!3 3.2 Steam Generator Design Data 3-5
i 3.3 Steam Generator Arrangement 3-6
; 3.4 Shell Assembly 3-6
% 3.4.1 Upper Shell Sub-Assembly 3-6
3.4.2 Middle Shell Sub-Assembly 3-7
3.4.3 Lower Shell SubfAssembly 3-8
3.5 Bundle Assembly | 3-8
3.6 Channel Assembly 3-11
é 3.7 OSpecial Design Features R 3-11
I 3.8 Steam Generator Drawings (Figure No.'s)
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3.1

3.2

3.3
3.4
3.5
3.6

3.7
3.8

Head and Tubesheet....

LIST OF FIGURES

. General Arrangement.....

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Shell Transition.............l...l..l..l.

Irllet and Outlet NOZZle-uoooooooo-oo-oooo

Irlteml Shroud-l....ll........t.l........

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Tube Vibration SUpPpPresSOTreccccascocssacsss

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3.1
3.2
3.3
3.4
3.5
3.6
3.7
3.8

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GHARGE N08-2€—éh31, DOCUMENT NO.N_D_7h=66

ISSUE

DATE

12/16 /7L

o . S . S T J————

3.0 Molten Salt Steam Generator Preliminary Design

3.1 Desgign Description sSummary

The Foster Wheeler Steam Generator is a "L"

shaped fixed tubesheet heat exchanger as shown in Figure 3.1.

The steam generator design is based on a system of four

units per loop. The steam generator has 101l tubes which

are attached at each end to a fixed tubesheet.

The steam generator design is such that the flow of

molten salt and super critical fluid through the unit is

in the direction of natural circulation.

Molten salt enters

through an inlet nozzle (Figure 3.l4), which is perpendic-

ular to the tube bundle, chamges direction 900 and decends

approximately 100 feet before exiting through an outlet

nozzle (Figure 3.4;). Supercritical fluid enters through

an inlet nozzle located in the center of the lower head

(Figure 3.2) and rises through the heat transfer tubes and

exits through an outlet nozzle in the center of the upper

head. Vent nozzles are positioned at each tubesheet to

shell juncture to prevent the possibility of trapped gases

resulting in localized hot spots on the shell (Figures 3.1

and 3.2).

BY

C.H.H. APPROVED

PAGE 3.}, OF

FWC FORM 172 - 4

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CHARGE No 8-25-2431

DOCUMENT NO.ND-7L-66

ISSUE 1

DATE12/16 /7L

-

3.2 Molten Salt Steam Generator Design Conditions

A.SUMOE.

Section III Class I Pressure Vesselx

Primary Side (Shell)

Design Temperature
Design Pressure
i Joint Efficiency

; Allowable Stress

i Shell Material

Shell -Inside Diameter

Secondary Side (Tube)

E Design Temperature

Design Pressure
Joint Efficiency
Allowable Stress
Number of Tubes
Tube Pitch

Tub; Size
Effective Length
Tube Material

Number of Tie Rods

1150°F
300 P.S.I.G.

100%

9500 PSI (Primary Membrane)

Hastelloy-"N"

39%n

1120°F
3800 P.S.I.A.

1009%

11,600 PSI (Primary Membrane)

1014

1-1/8" Triangular

3/L4" 0.D. x .125 min. wall

140 ft.

Hastelloy-"N"

13

¥ Designed to A.S.M.E. Section III Class I with Addenda thru
Winter, 1973

BY C.H.H.

APPROVED

PAGE 3-5 OF

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CHARGE NOg_p5-5),31 | DOCUMENT NO.yp_7,-66 |ISSUE 1

DATE 12/16/7h

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3.3 Steam Generator Arrangement

The steam generator shown in Figure 3.1 consists of:
1. 'Upper Shell Assembly
2. Middle Shell Assembly
3. Lower Shell Assembly
L, Bundle Assembly
5. Channel Assembly
The steam generator is installed with the curved short
section in the horizontal plane, and the long section in
the vertical plane. Support lugs or saddles have not been
considerea at this time. A functional description of the

steam generator can be found in Section 4.0 of this report.

3.l Shell Assembly 1

3.4.1 Upper Shell Assembly

The upper shell sub-assembly consists of: expanded shell
section, inlet nozzle, intershroud, baffle plate and a shell
transition. The upper shell sub-assembly is welded to the
tubesheet and the middle shell sub-assembly see Figure 3.1.

The expanded shell section is made from rolled and butt
weld Hastelloy - "N" plate. A 20" molten salt inlet nozzle
is located midway in this expanded region. This nozzle is

a saddle type which was selected so that the nozzle to shell

BY

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OHARGE NO 8-25-2,31 | DOCUMENT KO.nNp-7),-66 ISSUE 1

PATE 12/16/74

weld is removed from critically stressed shell to nozzle
intersection. The use of a saddle type nozzle permits, with-
out the use of special fiechniques, the radiography of the
nozzle welds.

Within the expanded shell section is a second cylinder
which is the intershroud, its purpose is to chammel the flow
of molten salt to the heat transfer tubes. This intershroud
has an impingement baffle which protects the heat transfer
tubes from excess flow errosion. The shroud is perforated
with holes to allow the molten salt to enter from the expanded
shell secfion and decend through the unit as shown on Figure 3.1.

The lower section of the intershroud and the shell tran-
sition are of single piece construction as shown in Figure 3.3.

1
A single piece design is recommended to avoid the use of welds
in the high stress area of the junction discontinuity.

The upper baffle plate is positioned between the molten
salt inlet nozzle and the tubesheet to act as a thermal barrier

to prevent the tubesheet and the I. B. W. welds from being

thermally shocked.

3.4.2 Middle Shell Sub-Assembly

The middle shell sub-assembly consists of rolled and butt

welded sections of Hastelloy - "N" plate. The middle shell

BY

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~acts as a continuation of the intershroud which channels salt

flow around the heat transfer tubes.

3.4.3 Lower Shell Sub-Assembly

The lower shell sub-assembly consists of: expanded shell
section, outlet nozzle, intershroud, baffle plate and a shell
transition. The lower shell sub-assembly is welded to the
lower tubesheet and to middle shell sub-assembly.

The expanded shell section is identical with that of the
upper shell except the 20" nozzle is used as an outlet for

the molten salt.

3.5 Bundle Assembly

The bundle assembly consists of (2) tubesheets, heat
;]

transfer tubes, tie-rods, support plates and vibration sup-

. pressors. The complete bundle assembly is welded to the shell

~assembly, for manufacturing sequences see section of this

report.

The tubesheets are 15" thick Vacuum Arc or Electroslag
Remelt Forgings of Hastelloy ~ "N" material. The tubesheets
are designed to meet the code requirements for stress in per-

forated plates. The outer rim of the tubesheet is machined

- to a countour which minimizes the thermal stress by having

this rim respond more wniformly to thermal transients with

s, Wk P T

1216 T ...

BY

APPROVED PAGE 3.8 OF

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CHARGE NO8-26-21,31 | DOCUMENT NO.wp-7,-66  |ISSUE1 | DAYE 12/16/7),

L Al R W R . e 1L W A

~ the tube hole ligaments. The lower face of each tubesheet

P S I,

is machined to provide spigots for welding of the heat transfer
tubes as shown on Figure 3.Z2.

i The heat transfer tubes are full penetration welded to
each tubesheet by using an Internal Bore weld technique. The

!

.

! . heat transfer tubes are of Hastelloy - "N" material designed
:

{;, 2 in accordance with A.S.M.E. tubing specifications. Allowances
;}; % for manufacturing scratches, thinning and corrosion have been
KL
fj; ? included in the tube wall thichness.
Eé’ | The heat transfer tubes are supported in the vertical
o) .
%gt { position by perforated tube support plates commonly used in E
ff! i heat-exchanger equipment. The support plates are of the "Drilled ;
g% é Plate" type as shown on FiguEF 3.6. This arrangement was !
é} 2 chosen in order to provide circulation between the heat trans- |
;it i - fer tubes, themwly reducing possible stratification problems
;;» _‘ - and preventing solid collection.
g% % The tube support plates and the vibration suppressor
y |
% ? grids are supported by (13) 3/L" diameter tie-rods which are
;; ? located in the tube pattern as shown on Figure 3.6 and 3.7.
@ab These tie-rods are threaded into the tube sheets, located at
)
%i both ends of the unit.
Eg The tie-rods are threaded into the upper tubesheet. The
support plates are spaced using sleeves which are 7/8" inside
1
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~diameter. The sleeves are dimpled to assure the proper location

concentric to the tie-rods. This type of assembly is joined in
the curved region by a turnbuckle as shown on Figure 3.8. The
curved region uses a vibration suppressor grid to properly
position the tubes in this area. The function of the grid

is to allow expansion of the tubes while restraining them against
vibration as shown on Figure 3.7. The tie-rods are secured at

the last guppressor grid by means of a heavy hex nut which is

‘seal welded in place.

The vibration suppressor grid is comprised of an outer
ring that ié roiied thru the "Y" axis and is machined to accept’
a grid of flat bars as shown on Figure 3.7.

| The turnbuckle used to j?in these grids is shown on Figure

3.8. Each turn buckle is positioned within the grid and accepts

.at each end a curved tie-rod., This threading assembly is made

‘_possible by use of a right and left hand thread within the turn-

buckle,

The by-pass of liquid to the outside of the support plates
and the suppressor grids is minimized by the manufacturing toler-
ances that can be maintained. The inside shell diameter can be

manufactured to 394" +3/16" -0" and the supports can be held to

an outside diameter of 39 3/8" +1/16 -0.

BY

APPROVED PAGE 3 10 OF

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CHARGS NO8-25-2),31 | DOCUMENT NO. ND-7),-66 ISSUE 1 j DATE 12/16/7u

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CHARGE NO§-25-2),31

DOCUMENT NO.ND-7L-66 | ISSUE 1

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3.6 Channel Assembly

The channel assembly is comprised of the bundle assembly

(described in the preceding section) and the upper and lower

hemispherical heads.

means of a full penetration weld.

head is a 16" nozzle

allows for full radiography of the closure weld.

in the lower head serves as the inlet nozzle which carries super-

Each head is joined to the tubesheet by

of the saddle type. This type of nozzle

Tocated in the center of each

PATE42/16/74 |

A el LY T W R

The 16 nozzle |

critical fluid into the plenum chamber (lower hemi-head) and

then through the heat transfer tubes.
the high pressure steam is collected in the upper plenum (upper
hemi-head) and allowed to exit through the 16" outlet nozzle

as shown on Figure 3.

2.

Saddle Nozzle

) 3.7 Special Design Features

Upon exiting the tubes |

The saddle type nozzle offers two advantages not found in i

nozzles of a different design.

the radiography of the nozzle to shell weld without the use of

special techniques.

weld from the critically stressed nozzle to shell intersection.

Internal Bore Weld

The I.B.W. welding process allows design of a tubesheet

Secondly, the saddle nozzle removes the

First, the saddle nozzle permits

1By

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CHARGE NO 8-25-21,31 | DOCUMENT NO. ND-7L-66 |ISSUE 1

~ that eliminates the crevice created by the standard "through"

tubesheet design. The I.B.W. weld is a full penetration weld
that can be fully radiographed.

Minimal Flow By-Pass

The close manufacturing tolerances that can be maintained
during manufacture of the intershroud, shell cylinder and tube-
supports minimizes the by-pass of molten salt to 2 percent of

the total flow.

BY

R p L A . A Tyl T 0w . Wb Sy g

LIL’{ TP 12/16 /7L

WL W e R e, S L s

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|

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[

APPROVED PAGE 312 oF
NOTES

1. DO WOT SCALF TH!S DRAWING. USE FIGURE
DIMENSIONS ONLY.

2, ABBREVIATIONS USED ON THIS DRAWIKG ARE IN
ACCORDANCE, WITH AMERICAN STANDARD "AB-
BREYLATIONS FOR USE ON DRAWINGS.”

1671.0.x [ 3 THK,

STEAM INLET
NQZZLE

o DESIGN _DATA

PRIMARY SIDE (SHELL)
-+, DESIGN TEMPR.  LISO'F
DESIGN PRESS. 200P3lC

i ALLOW. STRESS 9500 FS!
WELD EFF, 100 %
@

SECONDARY SIDE (TUBE)
DESIGN TEMP  LI20°F
DESIGN PRESS. 3800PSIA
ALLOW.STRESS 11B00E3I
WELD EFF. 100 %

MATERIAL

VENT

T SHELL HASTELLOY N
TUBES HASTELLOY N

TUBE_DATA
NO.OF TUBES, _I0id
TUBE SIZE 94 OPx25WALL
TUBE PITCH 1 )" 60°
EFF.LENGTH  140FT,

MOLTEN SALT
QUTLET NOZZLE
20"0_9,,3,8THK.

STEAM QUTLET

NOZZLE
1610~ 2 THK.

i
r

l 1,
‘.
541.0.
37582 #/‘i‘jfi?\
r\"
.
!

SHROUD

VIBRATION )
SUPPRESSOR
REF. DWG,

F.W.C. ND-740-497

LE

2070.D.x¥g THK.

ADDED T SieN DATA
ADDED T/E RODS & SPACERS
AR, OF TUBES /414 wAs025
! e wes 2-5"
1570, wRS @8 LD
7= 3% " TUBE SHEET S i
ARV S 0T

4 32 58Z MAd, wihS 31850
B o™, | 5476 wss 25¥a £
GEYe P waS TT VYR,
702 HeE WAS TR
g TR, wRS T TRA.
£rE LaTH 140 WRS FI107

MOLTEN SALT

INLET NO

TUBE SUPPORTS—"

(101a) TUBES 34 00.

x 125 MIN. WALL
018 8 PITCH,
EFFECTIVE LGTH.
140FT.

~di
1.D. SPACERS
/,‘

0313 014, TIE RODS
WITH i3 OD. «78"

A lzis7ix]| 30V 1D was 3@}

LETTER DATE

135 1

[
s DESCRIFTION

APFROX
B B

REVISIONS

STEAM GENERATOR
2= GENERAL ARRANGEMENT

MOLTEN SALT

DRAWING NUMNER I SCALE: 3y =ieow

ND-720-153

MITIAL oATE CONT, NO.

REVISION

OWG, MADE &
APPROVED BY
DRAFTSMAN WA e-27-78 QRDER MUMBERS
- CHECKER Coit {70774
SQUAD LEADER | s
SEGTION WEAD

MECHAN{CAL ' | PROJECT CONT SUPY
DESIGN MANAGER This Drawing is the Proparty ot tha

- - : FOSTER WHEELER CORPORATION
hally cdt “Hrs 7 Pl alind @ 140 50. ORANGE A¥. LIVINGSTOH. N. ..

.

STRESS MANUFACTURING A+t 18 LENT WITNOUT CONSIDKRATION OTHER THAN THL
ANALYSIS ENGINEER

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THiS DRAWING SUPERSEDES
DWG‘ No' THIS DRAWING SUPERSEDED BY

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Fig. 3.1

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8. 791" R. MAX.

49" DIA

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APPROX., |

(4) UNITS REQLIRED

DESIGN PRESSURE
DESIGN TEMP. 2 120° F

IRCO Psi

MECHANICAL | PROJELT
DESGW MANAGER,

o 25 .
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REVISED LOCATION OF 15 YE DIM
ADDLD A.ps,z * DA, ,vam- 19

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1RO E WA 18O
74 | OELETED, MAT‘.." aTE

LETTER

DATE DESCRIPTION

REVISIONS

(?) FIELD CONSTRUCTION

(6) FABRICATION

(5) PREPARATION OF SHOP DETAILS

(4) CUSTOMER COMMENTS

@ PURCHASE OF ALL MATERIALS

(2) PURCHASE OF MAJOR MATERIALS

(1) PRELIMINARY ARRANGEMENT

D'WG. REY,

DATE ISSUED FOR

HEAD AND TUBESHEET

DETAIL

MOLTEN SALT

This Drawing is the Property of the

FOSTER WHEELER CORPORATION
110 SOUTH ORANGE AVE., LIVINGSTON. N. J.

AMND 18 LENT WITHOUT CQHIIDI.A'ION OTHER THAN THE
AT IT SHALL NOT BE AE-
PRODUCID COPIID LINY Ol DISFOSED OF DIRECTLY OR
INDIRECTLY MOR USED FOR ANY PURPFOSE OTHER THAN
THAT FOR WHICH IT IS SPECHICALLY FURNISHED. THE
APPARATUS FHOWN IN THE DRAWING I8 COVERED @Y
PATENTS.

VENT

DRAWN BY:

JAK [7-26-72] scaLE 3" = V-0

ORDER NO.

CHECKED AaY:

CVhH 19-12-7

APPROVED BY: |7

7(7/%‘ ND"'722"64

REV]

DWG. NO.

THIS DRAWING SUPERSEDES
DRAWING SUPERSEDED BY

THIs

FORM 138-47-C

Fig. 3.2

»

i

(16)-7,, DIA.HOLES

EQ.SPACED

N N N

TOP BOTTOM
DIM  lrrasiTion [TRANSITION
A 541, D. S61.D.
B 285" 29 ot
C 26" 27"
NO. REQD 4 4

DES/GN PRESSURE - 300 F7 S/
DESICHN 7EMP

= //50°% F

MECHANICAL | PROJECT
DESIGN MANGER,
7 7 ~ B
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10-12-72 39Y%2 1.D. WAS 40" 1,D.

- '5 a4
A | wn ! 1:):(. WAS '¥g THK.
LETTER DATE DESCRIPTION
REVISIONS

(7) FIELD CONSTRUCTION

(€) FABRICATION

(5) PREPARATION OF SHOP DETAILS

(4) CUSTOMER COMMENTS

(3 PURCHASE OF ALL MATERIALS

(2) PURCHASE OF MAJOR MATERIALS

(D) PRELIMINARY ARRANGEMENT

DW'G. REY. DATE

ISSUED FOR

SHELL TRANSITION

MOLTEN SALT

W

This Drawing is the Property of the

FOSTER WHEELER CORPORATION
110 SOUTH ORANGE AVE.. LIVINGSTOM, N, J.

AND I8 LENT WITHOUT CONSIDERATION OTHER THAM THME
BOAROWER'S AGREEMENT THAT T BHALL NOT BE RE-
PRODUGCRD, COMED, LENT, OR DISFOSED OF DIRECTLY OR
INDIRECTLY NOR UBED FOR ANY PURPOSE OTHER THANM
THAT FOR WHICH [T I8 SPECIFICALLY FURNISHED. THE
APPARATUS SHOWN IN THE DRAWING |9 COVERED BY
PATENTSO.

DRAWN BY:

PN scaLx  3I¥ = 10"

ORDER NO.

CHECKED BY:

APPROVED BY:

LD 702 /6 7

REV,

DWG.

No.

THIS DRAWING SUPERSEDES
THIS DRAWING SUPERSEOED BY

FORM 138-47-C

Fig. 3.3
. MAX,

INLET OUTLET
Dim NOZZLE | NOZZLE

A 27" 28"

B 4% av,"

NO. REQ'D a4 4

ol

DESIGN PRESSURE = 300 PSI
DESIGN TEMP, = WBO°F

MECHAR ICAL
DRS\GWN

PROJECT
MAMAGER

Choalac 737

!
BNALYS\S

ENGINEER

{ ""[t

9-19-74 | ADDED: DiM. TABLE € LETTERS
LRSS [ REMOVED: MAT '
A DESIGM TEMP, |{S0°F WAS1100

%9 2o 141500,
L4

A R,

tog !
Lhooud oo

LETTER DATE DESCRIPTION

REVISIONS

@ FIELD CONSTRUCTION

(6) FABRICATION

(5) PREPARATION OF SHOP DETAILS

() CUSTOMER COMMENTS

@ PURCHASE OF ALL MATERIALS

@ PURCHASE OF MAJOR MATERIALS

(1) PRELIMINARY ARRANGEMENT

ISSUED FOR

O'W'G. REY. DATE

SALT INLET / OUTLET
NOZZLE DETAIL

MOLTEN SALT

This Drawing is the Property of the

FOSTER WHEELER CORPORATION
110 SOUTH ORANGE AVE. LIVINGSTON, N. J.

AND IS LENT WITHOUT CONSIDERATION OTHER THAN THE
BORROWER'S AGREEMENT THAT IT BHALL NQT BI RE-
PRODUCED, COPIED. LENT, OR DISFOSEID OF DIRECTLY OR
INDIRECTLY NOR USED FOR AMY PURFPOSE OTHER THAN
THAT FOR WHICH IT |9 SPECIFICALLY FURNISHED. THE
:::::;"I'Ul SHOWN IN THE DRAWING i§ COVERED BY

DRAWN BY: JAK |8. T2 sCALE - = 10"

REV

ORDER NO.

CHECKED BY: |C. #4912 24 ND-722-165

APPROVED BY: F 42

DWG. No.

THIS
THIS

DRAWING SUPERSEDES
DRAWING SUPERSEDED BY

FORM 138-47-C

Fig. 3.4
r.So“

VIEW BRoR"

, 25" THK, PL.

TOP BOTTOM
DI M. SHROUD | SHROLD
A g2" 54
NO, REQ'D 4 4

A DA,

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0000000000 000 OO0 O P OO O OOO0O0 0000000
0000000000 000 COQO P OO O OGO GO0 T
0000000000 OO0 O OO0 G OO O OOOOOO 000 000
0000000000 OO0 OO0 P OOOOCOOO0O000000

0000000000 000 QOO0 O OO O OQOOOQO000000
24" 00000000 0000000 \

100000000 0000000

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[40.010,0.016 0 Bl +— ~ T O000000

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00000000 0000000

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24" Q. x ., 50" THK.

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10000000000 0000000 G- 0000000000000

40"0,0%,25"THK,

I

e (82)-.75"DIA. HOLES PER ROW

(2O)ROWS SPACED 2.25"' FROM ToP
APPROY. 1640 WHOLES

PROJECY
B Cork

MANULEACTUR WY
EMGINEER

R S R
Lo ]

v

verren | oate |

DESCRIPTION

REVISIONS

®

(7) FIELD CONSTRUCTION

FABRICATION

@ PREPARATION OF SHOP DETAILS

{4) CUSTOMER COMMENTS

@ PURCHASE OF ALL MATERIALS

@ PURCHASE OF MAJOR MATERIALS

@ PRELIMINARY ARRANGEMENT

D'W'G. REY. DATE

ISSUED FOR

INTERNAL SHROUD

MOLTEN SALT

This Drawing is the Property of the
FOSTER WHEELER CORPORATION
110 SOUTH ORANGE AVE.. LIVINGSTON, N, J.

AND IS LENT WITHOUT CONSIDERATION OTHER THAN THE
T T

THAT FOR
PATENTS.

HAT IT BNALL NOT BE RE&-

]
PRODUCED. COPIED, LENT, OR DISPOSED OF DIRECTLY OR
IMDIRECTLY NOR UBKED FOR ANY PURPOSE OTHER THAN

WHICH IT 18 SPECHICALLY FURNISHED. THE

APPARATUS SHOWN IN THE ORAWING I8 COVERED BY

DRAWN BY: LS.

o [somx T —re

ORDER NO.

cuxckeo wy: (O )4

APPFROVED BY:

Z#ND-742-17!

DWG. NoO.

THIS DRAWING SUPKRSEDES
THIS DRAWING SUPERSEDED BY

FORM 3847 - C

Fig. 3.5

NNNNN
WOT SCALE DRAWING.
DIMENSIONS
ANBREVIATIONS USED OM THIS DRA
ACCORDANCE RICAH
RAEYVIATIONS DRAWL

MOLTEN SALT

(o ) 2
(13) 2" DA, TIE RODS o : —
WITH | L 0D. SPACERS : : ‘j —
3%¢ g ; ND~720-125
‘.“:__ s uane s | DATE |CONT. NO. 2 -2 4 -1409
nnnnnnn

NOTES

o,

P

1. DO NOT SCALE THIS DRAWING. USE FIGURE
DIMENSIONS ONLY.

2. ABBREVIATIONS USED ON THIS DRAWING ARE 1N
ACCORDANCE WITH AMERICAN STAMDARD ™AB-
BREYIATIONS FOR USE ON DRAWINGS.”

y/d 2 /4

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NO. ND-740-197 S

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VIEW "TRaIR"

SECTION A=Al

Tl ROD
TURNBUCKLE

TrHiS DRAWING 1S THE PROPERTY OF THE

FOSTER WHEELER CORPORATION
110 SO. ORANGE AVE., LIVINGSTON, N.J.

AND IS LENT WITHOUT CONSIDERATION OTHER THAN THE
BORROWER'S AGREEMENT THAT IT SHALL NOT BE RE-
PRODUCED.COPIED,LENT. OR DISPOSED OF DIRECTLY OR
INDIRECTLY NOR USED FOR ANY PURPOSE CTHER THAN
THAT FOR WHICH IT IS SPECIFICALLY FURNISHED. THE
APPARATUS SHOWN IN THE DRAWING IS COVERED BY

1'=1"—0""

MECHAN\C AL PROJECT
DESIGN MAMAGER
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FORM 285-63.C

Fig. 3.8
FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2/;31 DOCUMENT NO., ND/7.L/66 ISSUE 1 DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FWC FORM 172 - |4

SECTION 4

THERMAL/HYDRAULIC DESIGN

BY

%*~.I; (lkuv¥
H. L. CHOU
Thermal /Hydraulic Task Leader

Approved by

/‘«ZL‘*}‘H. C;A,JN
~Dr. S. M. Cho

Manager
Thermal /Hydraulic Engineering

BY APPROVED PAGE l~a

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2},31 DocmWMMPNO.Nm/7u/66' ISSUE 1 DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE. CHANGES HAVE BEEN MADE

FWC FORM 172 - |

TABLE OF CONTENTS

PAGES
L.0 THERMAL/HYDRAULIC ANALYSTS L~1
4.1 THERMAL PERFORMANCE L=2
L.1.1 OVERALL HEAT TRANSFER COEFFICIENT L-2
Lh.1.2 HEAT TRANSFER SURFACE REQUIREMENTS L-6
L.1.2.1  UNCERTAINTY ANALYSIS OF THERMAIL PARAMETERS L=7
4.1.2.2 THERMAL DESIGN MARGIN 4=9
L.1.3 DETAILED PERFORMANCE CALCULATIONS L9
L.1.L PART LOAD PERFORMANCE L-18
4.1.4.1 METHOD 1 4=19
L.1.4.2 METHOD 2 L~20
L.2 PRESSURE DROP CALCULATTIONS | L4-29
4.2.1 STEAM/WATER SIDE PRESSURE DROP L~-29
Lh.2,2 SALT SIDE PRESSURE DROP L~33
L.3 STABILITY L=33
L4.3.1 STATIC STABILITY L-36
4.3.2 DYNAMIC STABILITY 4=39
Lh.h SYSTEMS RELATED TO STEAM GENERATOR L=L6
L.b.1 START-UP SYSTEM AND WATER CHEMISTRY L-L6
4.4.1.1  START-UP SYSTEM L-l6
4.4.1.2  WATER CHEMISTRY L=-L9
Lh.h,2 PRESSURE RELIEF SYSTEM L-50
4.5 REFERENCES L4~55

BY APPROVED | AGE b

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION

LIVINGSTON, N. J.

CHARGE NO. 8-25-2);31 DOCUMENT NO. wND/7)/66 ISSUE ¢ DATE 12/16/7L

L.0 THERMAL/HYDRAULIC ANALYSTS

It is the intent of this analysis to obtain a conceptual design of
a steam generator which will operate with a molten salt system and

a supercritical steam-power cycle. The steam generater design is

of a once-through, counterflow, shell-and-tube type with salt
flowing downward on the shell side and water/steam flowing upward

in the tubes. Only one steam generator unit is intended for each

of the four heat transport circuits which are connected, in parallel,
to the Molten Salt Breeder Reactor.

An axial (or long) flow approach was utilized after several unsuc-
cessful attempts of meeting the design criteria by a cross flow
scheme. The difficulty of solving the problems of tube vibration
and the excessive pressure drop on the shell side similtanteously,
forced the cross flow approach to be abandoned. However, it is
noted that the advantage of cross flow approach is not so signifi-
cant in a supercritical unit, due to the fact that the thick tube
wall, necessary for high pressure, becomes a dominant thermal re-
sistance (about 50%) of over-all thermal performance.

The analyses of the basic thermal/hydraulic performances, design
uncertainties, flow stabilities, and related systems have been
performed and the results are presented in detail in the following
sections.

BY

APPROVED PAGE  }1~1

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-24,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7L,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FWC FORM 172 - |

L.1 THERMAL PERFORMANCE

‘The thermodynamic ang transport properties and thermal conductances
of all heat transfer media were first determined in order to obtain
the steady-state thermal performance of the steam generator,

L.1.1 OVERALL HEAT TRANSFER COEFFICIENT

The heat is transfered from the molten salt on shell side to the
steam on tube side. The thermal conductances of molten salt, tube
wall, supercritical steam and steam-side fouling were considered.

The molten salt gide fouling effect was neglected at the direction
of Oak Ridge National Laboratory. The Molten-Salt Reactor Experi-~
ment (MSRE) in 1960's denoted no evidence of fouling in the MSRE
heat exchanger. However, the coolant mixture chosen for that
application was BF5 with 66 mole % of LiF, and the coolant chosen
for MSBR is NaF with 92 mole % of NaBF),.Evidence of the corrosion
product, Na CrF6’ has been found in loops circulating sodium
fluoroboratg, and this corrosion/product is expected to deposit

on the outside surface of the steam tubes if not removed by some
means (Ref.1 ). The effect of molten salt side fouling on the
thermal performance of the steam generator was reported in the
FWEC Monthly Progress Report #10 (Ref. 2).

0

The steam side fouling coefficient of 6667 Btu/hr—ftz- F was
used and is considered to be a reasonable value for the single-
phase flow and the five~year period of tube cleaning.

The design Properties of the tube wall material, Hastelloy N
(Nickel-Mblybdenum—Chromium—Iron Alloy), are tabulated in Re—
ference 3 :

A, Steam Side Coefficient

The correlation by H. S. Swenson (Ref.lL ) was recommended
for super-critical water/steam flowing inside circular
tubes exposed to heat flux at the wall. This correlation
is expressed by

0.92 0.61 0.231
L - oougy (20 )P mm Ay ) ()
M

T; -

BY APPROVED JPAGE -2

FWC FORM 172 - |}

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J,

CHARGE NO. 8-25-2),31

DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/7L

where

B. Salt Side

heat transfer coefficient inside tube, Btu/hréft2—oF
inside diameter of tube, ft

thermal conductivity of fluid inside tube, Bbu/hr—f+-CF
mags velocity of fluid, lb/hr—ft2

Viscosity of fluid at temperature of inside surface
of tube, 1b/hr-ft -

enthalpy at temperature of inside surface of tube,
Btu/1b .

enthalpy at temperature of bulk fluid, Btu/1b
temperature of fluid at inside surface of tube, T
temperature of bulk fluig, OF

specific volume of bulk fluid, ft3/1b, and

specific volime of3f1uid at temperature of inmide
surface of tube ft-/1b

Coefficient

The Dittus Boelter correlation was applied to determine the
molten salt heat transfer coefficient. The correlation is

expressed

Nu

where Nu

Re

Pr

by

0.8 .
= 0.023 (Re) "~ (pr)°"l
h, De

BY APPROVED AGE  L=3

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,,31 | DOCUMENT NO. WD/7L/66 |ISSUE 1 [DATE 12/16/7h

h = heat tranafgr coefficient of molten salt,
Btu/hr-ft°="F

De = equivalent diameter, ft

Kb = thermal conductivity ofomolten salt at bulk
temperature, Btu/hr-ft-"F

2
G = mass velocity of molten salt, 1b/hr-ft

f‘b = viscosity of molten salt at bulk temperature,
1b/hr-ft

Cp, = specific heat at constant pressure of molten
salt at bulk temperature, Btu/1b-°F

A1l the conductances of overall, steam side, molten salt
side, tube wall, and steam side fouling at full load are
plotted versus length of steam generator, as measured from
cold end, in Fig. 4.1. The overall conductance was calcu-
lated based on outside surface of tube and maximum tube
thickness (0.125" + 7%) since the steam is in the super-
critical thermodynamic state, special care must be taken
to insure that the evaluations of thermodynamic and trans-
port properties are accurate. For given load conditioms,
the tube length was divided into a sufficient number of
sections, so that the specific heat, Cp, of steam in each
gsection could be treated as a constant with negligible
error, and the concept of logarithm mean temperature 4if-
ference could be applied.

The length-averaged mean values were calculated for each
conductance to determine the overall performance of. the
unit as full load conditions and are tabulated below.
The method in Ref. 5 was utilized. The attributions to
overall resistance of each are also shown.

BY

APPROVED PAGE L=l

v | o
IR RRX TN SONSY SEe:

Eaws e

Ste

|
!

-

.

condu

e -

|
|

The

ins.gm

:
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ances of
it

i

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DAGE L5

———— e

Lo

e &
)

d -
o)

4

o
S
\O

31J-IYy/n3}g 90UBLONPUOD TBUWIDYG

4,000

2,000

1,000

oo | Y0 - [
Cold Eng, ft

» Toot
Fig.ohi1

60

Tube Len

20

FWC FORM 172 - I

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTQN, N. J.

CHARGE NO. 8~25-2),31 DOCUMENT NO. ND/7L/66 ISSUE 1 |DATE 12/16/74

Conductance % of total

Btu/hr-f+°-°F resistance
Steam side LWho1. 11 12.0
Molten salt side 1079.21 31.50
tube wall | 499.79 h8.58
steam side fouling 6667.0 | 7.92
overall U 340 100
overall log mean temp. difference '205.670F

L4L.1.2 HEAT TRANSFER SURFACE REQUIREMENTS

The design criteria at full load operating conditions are ag

follows:

Salt Water/Steam
Inlet temperature, °F 1150 700
Outlet temperature, °F 850 1000
Flow rate, 1b/hr 15,280,000 2,538,000
Inlet pressure, psia 235 —
Outlet pressure, psia — 3600
Maximum pressure drop, psi 60 , 200

Thermal duty = 483 MW(t) = 1.65 x 107 Btu/hr.

A basic thermal/hydraulic performance computer program was de-
veloped to calculate the sizing of steam generator using the
correlations mentioned in Section ;.1.1. This resulted in the
basic thermal/hydraulic design of the steam generator as 1000
tubes with length of 120 ft. The specifications of this basic
design are summarized below:

BY APPROVED AGE  L-6

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 __|DATE 12/16/7)

No. of tubes per unit 1000
Total length, ft. 120
Tube O. D. in. 0.75
Tube thickness, in. max. wall 0.13375
Tube I. D. in min, ID ~ 0.4825
Transverse tube pitch, in 1.125
Longitudinal pitch, in 0.974
Heat transfer rate, Btu/hr 1.6L46 x 107
Total effective surface area, ft2

(based on tube OD) - 23561.8
Material Hastelloy N

However, the uncertainties of heat transfer correlations, varia-
tion of tube thickness, flow by-pass, and even the flucturations
of flow rates will require surface margin to accomodate all the
possible deviations from design conditions. This was best ac-
complished by statistical approach.

L4.1.2.1 UNCERTAINTY ANALYSIS OF THERMAL PARAMETERS

In order to obtain an appropriate surface margin to maintain a
high confidence level of the design, a statistical computer pro-
gram, entitled SIMPAK, (Ref. 30), was utilized.

SIMPAK is a package of subroutines which read in the date, draw
random numbers, relate data to probability distributions, facili-
tate the Monte Carlo simulation, compute the mean and standard
deviation of resultant probability distributions, and print out
details of the probability distributions of a few variables of
interest.

Uncertainties were considered for heat transfer correlations
used in the thermal sizing of the steam generator. For the
steam conductance correlation, Swenson reports a standard devi~
ation of T 7.2% which leads to % 21.6% for 3¢'-deviation. This
deviation was used for this study. The steam side fouling con-
ductance was assumed to have a range of + 25% (37 —deviation).

BY

APPROVED ' PAGE L-7

LIVINGSTON, N. J.

FWC FORM 172 -

NOTATIONS IN THIS COLUMN INDICATE WHERFE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 DOCUMENT NO. ND/7)./66 ISSUE 1 DATE 12/416/7L

The heat transfer data of molten salt obtained with the forced-
convection loop FCL-2 are in good agreement with the empirical
correlation of Sieder and Tate (Ref. 1). In Rerf. 6, the Seider-
Tate type correlation was developed for fuel salt as follows:

Nu = 0.023}4 Re0.8 Pr1/3 (jfifi_)o-1h

S

with a standard deviation of 6.2% for Re > 12,000 (18.6% for

3¢ -deviation) and all the Properties read at bulk temperature,
except M s at wall temperature. Since no better information

is available, this Sieder-Tate correlation was compared with

the Dittus-Boelter correlation, which was used in sizing the
steam generator, to predict the uncertainty of coolant salt con-
ductance computed in the performance computer brogram, It was
found, within the temperature range of the shell side fluid, the
conductance calculated by Sieder-Tate correlation was about
90.06% of the one calculated by Dittus-Boelter correlation.
Therefore, for this uncertainty analysis, the molten salt con-
ductance obtained by Dittus Boelter correlation was first multi-
plied by 0.9006 and then assumed to have a standard deviation

of 8% for more conservation. For thermal conductivity of
Hastelloy N, the data obtained from Haynes Stellite were about
L.19% higher than the values used in the performance computer
program, (Ref. 7). Therefore, the thermal conductivity of tube
wall wag first multiplied by 1.022 and then assigned a Standard

deviation of 0.7167%. All the above parameters were agsumed to

have normal distributions o simplify the statistical analysis.
The variation of tube wall }hickness was limited to the manu-
facture range of 0.125" +??.

=/c

The modeling method for overall performance of multi-stage heat
exchangers used in Ref. 5 was adopted here to significantly
reduce the Monte Carlo computation time. All the heat transfer
correlations were included as part of the SIMPAK program so that
the inter-dependent variables such as inside diameter of tube,
flow rate, mass velocity, heat transfer coefficients, could be
related to one another and computed gimultaneously. The input
data were uncertainties of heat transfer correlations, variation
of tube wall thickness, and overall performance parameters of
basic design at full load conditions., The SIMPAK did 500

Monte Carlo trials and the cumulative probability of the total
surface area of steam generator, based on constant duty and

BY

APPROVED AGE L8

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,,31 | DOCUMENT NO. ND/74/66 | ISSUE 1 __|DATE 12/16/7L,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

temperature profile, was obtained as shown in Table }.1 and
Fig. 4.2. It is noted that the flow rates of salt and gteam
were held constant. When one considers the malfunctions of
pumps, flow rates of salt and steam may fluctuate. Agsuming
t 10% variation of flow rates, a second calculation was made
and the results are shown in Table 4.2 and Fig. };.3. These
Tesults are only approximate in the sense that the large fluc-
tuations of flow rates would eventually change the total duty
and temperature distribution along the unit.

4.1.2.2 THERMAL DESIGN MARGIN

The above analysis indicates that the confidence level of the
basic design of 1000 tubes with length of 120 ft (area =
23562 £t<) is about LO%. This is considered to be too low

a level of design confidence and therefore additional surface
must be provided. For the case of constant flow rates (Table
4.1, Figure ,.2), the req%ired surface to give the 99.9% con-
fidence level is 26002 £4° which is corresponding to 101L
tubes with length of 131 ft. For the case varying flow rates,2
the 99.9% confidence level requires a surface area of 26611 ft
which corresponds to 101l tubes with length of 134 ft. The
flow by-pass on shell side being 1.2% would require additional
1 ft long.

Therefore the reference design steam generator is designed to
contain 1014 tubes, 140 ft. long, plus 13 tie rods. This design
gives additional 18% of surface ares over the basic design,

ang has the highest confidence level (99.9%) to achieve the
specified thermal duty and design criteria,

4.1.3 DETAILED PERFORMANCE CALCULATIONS

The heat transfer surface requirements shown in Section L.1.2
were obtained by a performance computer code. This computer
code could be easily modified for all the situations and was
extensively used for the thermal hydraulic performances. Si-
mulating the unit by this performance computer code, the unit
was divided into 63 sections in length. For each section the
thermodynamic properties, transport properties, pressure drops,
and heat transferocoefficients were determined by iterations, ,
From 700°F 4o TS0 F of steam temperature in the inlet region,
the change in steam enthalpy per section wag limited to not
more than 10 Btu/lb. From 750 F to final steam outlet tempera~
ture, the change in steam enthalpy per section did not exceed
20 Btu/lb. The 1967 ASME steam tables were used for alil steam

BY APPROVED | AGE ;-9

FWC FORM 172 - |
NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

IR’
Table 4.1 Summary Statisties for Overall Surface Area of Steam Generator g
Flow Rates of Salt/Water are held constant :—é
2 o
Mean 23,756 ft 8
R
Standard deviation 736 jch
o
g Confidence §
level 0.0 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00 %
Surface ?‘;;
area, ft 21,958 122,830 | 23,134 | 23,374 23,573 123,719 | 23,889 |2L,140 | 24,370 24,729 | 26,002 .
‘ Ei
=
S~
oy
o\
H
:
)
o =
5 2
T B
© o
' I~
)
=

LNAXIEVAIA VI TONN

‘I N ‘NOISHNIAIT

NOILVEOd¥0D XO¥ANH YHTITHM WALSOL
Confidence Level

Overall Surface Area of Steam Generator vs.

2

F

Flow Rate of Salt and Steam/Water are held constant

|
|is: | {
S SN RN N | R
e L. .ylt\vlyvl -
|

T R

|

: |

_ fid]

| i {
e oy LAEE! Louet sonat|
m

E
¢

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=111 SR M B

S

’l.“‘ =

T

H

: L P T S B S
| | \ H
i . : | % W Y R M
.. s s et
A4 5 [ ST GV IR DR L 20 oL 1
\ | ] ' | i | |
\ , t IO {
LR

\
S S————

-: -'r-;'.. ;_.__. b

EEIILGHN ANTa Ll sl 3.3
. |
1B 2
. ' fosmt fima s
| ]
) ) ]
! 1 . 1 !MJ
h. :
ol .J.Zv» A4S ol < !

e e T ——

. _ o
Hobbq_monwWHmdno

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

xq

Table L.2 Summary Statisties for Overall Surface Area of Steam CGenerator

Bach Flow Rate of Salt and Steam/Water entering Unit is Changed

"ON EDYVHD

Lef2—a¢=-9

THAOYddY

Mean 23762 £t°
Standard deviation - T79
Confidence
Level 0.0 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00
ourface 5
Area, ft 21,779 {22,789 | 23,089 {23,341 23,568 | 23,760 |23,930 {2L,1L7 2L,432 | 24,779 [ 26,611

qOVd

ARl

99/ML/QN *ON INEWADOT

dNSST

A

ML79V72T Gava

LNIWINVdId VA TONN

‘L *N ‘NOISOHNIAIT

NOILVIOQY0D ZOWANI YATATHM YALSOL
Confidence Level

Overall Surface Area of Steam Generator vs.

Fi

Each Flow Rate of Salt and Steam/Water entering Unit is changed

I
V |
|

| T EESS 5 RSP SRS R S
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- — - ;Oau-.

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L

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{
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{

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e g
jes

|

PAGE L=13

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

CHARGE NO. 8-25-2);31 DOCUMENT NO. wND/7L/66 | ISSUE 1 DATE 12/16/7),

properties.

This detailed performance analysis established the heat transfer
surface requirement and pressure drops at full load operating
conditions, in accordance with the design criteria (section
L.1.2). -

The pressure drops at entrances, exits, tube support plates and
vibration suppressors have minor effects on the thermal perfor-
mance of the steam generator. Temperature and pressure profiles
as a function of position in the steam generator for steady state
operation at full load are shown in Figures L.l to L4.6. The
results of the thermal/hydraulic analysis of basic design at full
load conditions are also summarized below:

Salt Water/Steam
Inlet temperature, oF 1150 - 700
Outlet temperature, OF 850.6 -~ 1000
Inlet pressure, psia 235 ' 3770.5
Outlet pressure, psia 296 3600
Flow rate, 1b/hr 15.28 x 106 o 2.538 x 106
Mass velocity, lb/hr~ft2 | 3.363 x 10° 1.999 x 106
Static pressure difference, psi - 61 + 170.5
Total net pressure loss, péi Lo.?2 158.1

Thermal duty = 1.6L6 x 109 Btu/hr
The pressure drop calculations are presented in Section L4.2.

There are two vent nozzles positioned at each tubesheet-to-shell
Juncture to vent trapped gases. These vent nozzles will also
assist in preventing the salt from freezing at the juncture of

the cold leg. The low feedwater temperature of 700 F angd stag-
nation of salt at the cormer of cold leg are the possible causes
of salt-freezing. For the basic design, the salt temperature at
the outside tube surface in the activ heat transfer region

near the lower tubeshegt is about 800 F at full load conditions
and is higher than 800 F at part loads. In the stagnant (inactive)

BY APPROVED AGE L-1]

t 2 52
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S st - '} a8 Sl o)

| Ey iy ;

= - et ey {
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foil e ld
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—

=

PAGE L4-15

..Afil

Water/steam

e s et =L ML LS IEEUR SR CSRRS ---.PL

‘fil .P_fi
L sl
B3 Ry |

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FWC FORM 172 - J

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 | DOCUMENT o, ND/74/66 | ISSUE 1 |patE  12/16/71

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE _

L.1.4

region, the poor heat transfer coefficient on the salt side
might create a situation in which the salt-side tube wall
tempergture could be lowered below the salt freezing point

of 725°F, However, the above possibility is remote considering
the fact that the steanm generator cell of the generating station
building is maintained at temperature of 1000°F (Ref. 23 )

by external heat source. The salt at the tubesheet-to~ghell
Juncture would absorb additional heat through the heat conduction
from heated cell to the shell of steam generator. Therefore,
the salt temperature near the cold tubesheet would not be below
the salt freezing point of 725°F.  The vent nozzles will contin-
uously vent out a small amount of salt to keep the salt always
in motion (also for higher heat transfer coefficient) instead

of stagnant in this region. From the preceding discussions,
the unit will be free from a salt freezing problem. Detailed
analysis was not undertaken of this problem in the present study
but should be 2 part of future studies.

PART LOAD PERFORMANCE

Partial load operation is defined as any condition between 20
and 100% of full load thermal duty. The operation from zerp
to 20% load is designated as startup operation.

Two major limitations, of high priority, when the plant un-
dergoes changes in load are: (a) turbine throttle temperature
to be held at 1000°F due to turbine limitations, and (b§ the
coolant salt temperature at steam generator outlet and fuel
salt temperature at reactor in%et to be held above the salt
freezing points of 725 and 930 F, respectively, at all loads.
The primary fuel salt gystem will operate at constant flow rate
and constant reactor inlet temperature of 10500F, with the re-
actor outlet temperature controlled as a function of load., It

tions. If the coolant salt decreases linearly with load, the

reduced coolant salt flow rate would decrease the coolant salt
temperature at the steam generator outlet excessively with the
danger of being lower than freezing point (Ref. 10). However,

duction of load (percentagewise), there is a chance that the
steam outlet temperature might exceed the limitation of 1000°F.
This indicates the difficulties of control during the part
load operations.

BY

APPROVED PAGE  }~18

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2);31 | DOCUMENT NO. ND/7L/66 ISSUE 1 [DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |,

Two principal methods for control of part load operation are
(1) Varying the coolant salt flow with the flow reduction less

than the load reduction (percentagew1se) and thus_allowing the

L.1.4.1 METHOD 1

Coolant salt flow was varied linearly from 30% flow at 20%

thermal load to 100% flow at 100% load. The associated inlet

and outlet temperatures of coolant salt are tabulated in Table

L.3 (Ref. 3). The steam inlet temperature was held constant at
700°F, while the steam flow varied in proportion to load (slight
deviation from linearity exists due to water bypass for the attem—
peratgr). The steam outlet temperature was allowed to rise above
1,000 F at part loads and was subsequently ttemperated with

the bypassed fgedwater at conditions of T00°F and 3700 psia to
maintain 1,000°F turbine inlet temperature.

The results presented in Figures L.7 to .13 indicate that this
method is satisfactory. Figure 4.7 shows the uncontrolled ang
attemperated final steam temperatures versus Percent of full
load. Figure L,.8 shows the amount of attemperator flow versus
percent load. Figure ;.9 shows the amount of salt and water
flow rates entering the unit versus percent of full load. Pi-—
gures L.10 to .13 show the temperature and bressure profiles
of salt ard water/steam at part loads,

4.1.L.1.1 FEASIBILITY OF USING A SPRAY ATTEMPORATOR AT THE OUTLET
OF THE STEAM GENERATOR

It has been Foster Wheeler's experience that utiligiss do not
prefer to use spray attemporation between the final stage of
superheat and the turbine. However, such an arrangement is
accepted only when there is one stage of superheat,

It has been‘standard practice at Foster Wheeler that, for
large fossil-fired steam generators, two Spray attemporator
locations are Provided between stages of superheat. This

———

BY APPROVED PAGE )19

FOSTER WHEELER ENERCY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND/7L/66 ISSUE 1_IDATE 12/16/7),

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

arrangement allows the superheaters to be designed with lower
alloy steels and lower metal temperatures. Also, this arrange-
ment provides rapid steam temperature control over a wide load
range since the steam travel time between point of spraying

and point where the temperature is being controlled (the other
side of a stage of superheat) is smell.

With the present reference design, there are two possible lo-
cations for a spray attemporator; (1) at the inlet of the

steam generator and (2) at the outlet of the steam generator
Placing an attemporator at the inlet of the steam generator

is not recommended because it would increase the probability

of salt freezing at the cold end of the steam generator where
the higher pressure (above 3800 psi), colder (below 700°F) water
would result.

Locating a spray attemporator at the outlet of the steam genera-
tor is quite practical. The inlet feedwater can then be used

as a source of spray water with proper pressure head and tem-
perature for this application. The calculations using this
approach that are reported herein indicate that the spray flow
quantities that would be required over the load range are
reasonable and in line with values required on fossil fired
steam generator.

It is noted that the change of moisture carry-over to the
turbine that could damage the high pressure stages does not
exist because of high steam temperatures (1000°F and above)
and supercritical pressures (above 3600 psi). Another factor
alding this situation is the long length of piping that will
undoubtedly be required between the steam generators and the
turbine.

L.1.4.2 METHOD 2

Partial decoupling of the secondary coolant salt loop from the
primary fuel salt loop could be accomplished by short-circuiting
a fraction of the coolant salt around the primary heat exchanger.
This would require a throttling devicé which is not presently
developed. The present steam generator design is not based on
this scheme, however, this method would provide useful infor-
mation for control study.

In this study, the steam %nlet and outlet temperatures were held
constant at 700 and 1000 F, respectively, during part load opera~--
tions for which steam flow rate changes linearly with load. The
salt flow rate and salt inlet temperature which would maintain

the salt outlet temperature at 850°F * 15° were to be determined.

BY APPROVED AGE  Jy-20

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO,8-25-2L,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7L

FWC FORM 172 - )

Table .3 Coolant Salt Temperatures and Coolant Salt flow with

Varying load.
Salt inlet o | Salt outlet o

Load, % Flow, % Temperature, F Temperature, F
100 100.00 1150 850

90 91.25 1147 851

80 82.50 114, 853

70 73.75 1139 855

60 65.00 1134 857

50 56.25 1127 860

L0 47.50 1117 865

30 B 38.75 1106 | 87L

20 - 30.00 1091 891

- Base on ORNL Reference Design Heat Exchanger and operation of the fuel
salt system at constant reactor inlet temperature of 1050°F,

BY APPROVED PAGE )21

PAGE }-22

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FWC FORM 172 - |,

WHERE CHANGES HAVE BEEN MADE

NOTATIONS IN THIS COLUMN INDICATE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT

LIVINGSTON, N, J.

CHARGE NO. 8-25-2),31 DOCUMENT NO. ND/7)/6

6| ISSUE 1

DATE 12/16/7l,

Therefore,

the salt flow rate and galt inlet

for each salt outlet temperature of 835,
temperature were computed

850, 865°F,

throughout the range of 20 to 110% of rated load,

The results are Summarized in Table 4.l ang
constant salt inlet temperature lines pre-~
were generated from Table L.l by using a

and Fig. 4,15, The
sented in Fig. 4,15
third order interpolation routine to
information for control analysis.

part load range of about 6L tQ 97%,
temperature from 1040
more than 100%

h.2

provide valuable overall

Fig. 4.15, shows that for )

The performance computer program used in sizing the steam ge—

drops between inlet/outlet nozzles,
region were adde
nozzle overall pressure differences
fluids,

4.2.1  STEAM/WATER -SIDE PRESSURE DROP

The absolute roughness, 0.00006 ft,

of tube, 0.4825", were used for

The pressure
tube support plates ang the

d to estimate the nozzle-to-
for both shell angd tube side

and minimum ingide diameter

tube side pressure drop calcu=

lation. Sigce the flow is in turbulent region (Re = 1.9 x 10

to 1.0 x 10

), the friction factor was calculated by the

Colebrook-White semiemperical formula (Ref. 11) which is ex-

Pressed as:

1 ks
— + 2 log ):
J f D
friction factor

absolute roughness, ft

D: inside diameter of tube,
Reynold number

1.1 - 2 log (1 + 9.35 )

Re { £ K

ft

BY APPROVED

PAGE  Li=29

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2),31 | DOCUMENT NO. ND/7)./66 ISSUE 1 |DATE 12/16/7)

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

Table L.l Part Load Performance — Method II
Salt Flow Rate and Salt Inlet Temperature

Salt Outleto Salt Inlet Salt Flow

Temperature, F Temperature °F Rate, % : Load, %
835°F 1001.8 35.60 19.81
1005.9 52,18 29.7L

1013.8 66.58 39.68

1026.3 77.8L L9.6L

1045.2 85.08 59.63

1071.0 88.55 69.6L

110L. 1 88.83 79.67

1145.6 86.69 89.72

1193.2 83.55 99.78

850°F 1001. 3 39,28 19.81
_ 1004.2 57.86 29.7hL
1009.7 ~ 74.51 39.68

1019.1 88.00 L9.6L

1033.3 97.57 59.64

1053.3 ' 102.6) 69.64

1079.8 10L.02 79.68

1113.0 102.39 89.7L

1152.6 98.91 99.81

865°F 1000.9 L3.74 19,81
1002.9 64,.69 29.74L

1007.0 83.85 39.69

101).1 99.92 49.65

1025.0 111.87 59.65

1040.5 119.04 69.66

1061.6 121,62 79.70

1088.., 120.50 89.76

1121.5 116.77 99.85

1160.2 11.75 109.97

Steam inlet temp. = 700°F
O
Steam outlet temp. = 1000 F

Steam flow rate = changes linearly
with load

BY APPROVED AGE 4-30

LE-1 @Ova

120

Percent load vs. salt inlet temperature at constant salt outlet temperatrue

steam inlet temp. = 700°F

o
steam outlet temp. = 1000 F

constant salt outlet temperature

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Percent salt flow vs. percent load at constant salt inlet and outlet temperatures

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110

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANCES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2L,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7L

h.2.2

L.3

The pressure losses occurred between the inlet nozzle and the
lower tubesheet, and between the upper tubesheet and the exit
nozzle were calculated per FWEC fluid flow formulas (Ref. 12).
The results are summarized in Table 4.5, with hand calculations
attached in Appendix B. Pressure profiles of steam generator
at full and part loads are also presented in Section 4.1.l.

SALT SIDE PRESSURE DROP

There are 23 tube support plates in vertical and horizontal
portions, three vibration suppressors in bend region, one shroud
in inlet region and four shrouds in outlet region. The pressure
drops occurred at nozzles, shrouds, tube support plates and vi-
bration suppressors were hand-calculated per FWEC Standard Ma~
nual (Ref. 12). These values were added to the pressure drops
calculated by the active heat transfer region performance code.
The results are summarized in Table L.6. Hand-calculations are
attached in Appendix B.

The Colebrook-White formula was used for friction factor cal-
culation. Unit pressure profiles in shell side are shown in
Section 4.1.3 and L4.1.l.

STABILITY

Historically, the early studies of flow instability were de-
veloped from operational difficulties with fossil-fired boilers.
A number of boiler~tube failures and thermal performance de-
gradation were attributed to water/steam flow ingtability.
Units designed in supercritical pressure region also suffered
the same problem as well as in the subcritical region.

In boiling systems, fluctuations are always present because of
variations of the rate of bubble formation and population, of
flow regimes, of the heat transfer coefficient, etc. Conse-
quently these fluctuations may induce the flow instabilities.

In the supercritical region, rapid changes of thermophysical
properties are observed in the vicinity of the critical point.
The propagations of variations of properties, in particular

of the density and of the enthalpy, through the system introduce
time and space lags of transformation which under certain condi-
tions can cause unstable flow.

Two major classifications of unstable phenomena are defined as
static and dynamic instabilities. Both of them could be analyzed
by conservation equations of mass, momentum, energy and the

BY

APPROVED | PAGE 1~33

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE. CHANGES HAVE BEEN MADE

Table 4.5 Steam-side Pressure and pressure drops 0

: =

o

=P at nozzle, P. leaving [AP between P. entering [AP at nozzle,

Entrance tubesheet, tubesheet, tubesheets, tubesheet, tubesheet, EBxit N

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FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J. -

CHARGE NO. 8-25-21,31 | DOCUMENT NO. ND/7L/66 |1Isste 1 |paTE 12/16/7L

4.3.1
4.3.1.1

proper equation of state. Static instability lies in the steady-
state laws, while dynamic instability is time-variant phenomena.

Zuber (Ref. 13) described three mechanisms which could induce
thermohydraulic oscillations at supercritical pressure. One is
caused by the variation of the heat transfer coefficient at the
pseudo critical point, which is defined as the point where C
reaches its maximum value. The second is caused by the effegts
of large compressibility and the resultant low velocity of sound
in the critical region. The third mechanism is caused by the
large variation of flow brought about by density variations of
the fluid during the heating process.

Both static and dynamic stabilities for the present steam
generator are discussed in detail in the following sections.

STATIC STABILITY

INTRODUCTION

Static instability is an amplification of steady state distur-
bances which encompass tube circuit configuration, heating
imbalances, flow rate perturbations, etc. The static insta-
bility of primary design importance in steam generators is

the excursive instability, which at supercritical pressure,

is the equivalent of the "Ledinegg" excursive instability

in boiling steam at subcritical pressures. A flow is sub-
jected to a static instability when the flow conditions,
changed by a small perturbation, will not return to original
steady state conditions (Ref 17).

The significance of the static stability is best analyzed by
plotting the pressure drop-~flow characteristic as schematically
shown in Fig. L.16. A system of many parallel heated tubes

is considered with attention focused on only one tube where
various levels of heat input are allowed. The quantity of
heat input depends gualitatively on the situation of heating
medium distribution among the heated tubes. Demand curves

Q1, Q2, and Q3 denote increased levels of heating medium
quantity surrounding the concerned heated tube. A constant
inlet-to-outlet pressure difference is imposed as indicated

by the horizontal line H. Intersections with curve @1 showing
the possible operating points for a constant pressure drop
supply system (or any pump characteristics) are indicated by
C, D or E. Operation at point D or E will be stable whereas
that at point C will be unstable. For example, if at point

BY

APPROVED [PAGE  L-36

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,31 | DOCUMENT NO. ND/7L4/66 |Issue 1 |paTE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

either D or E the flow is perturbated to increase (decrease),
the pressure drop of the heated tube increases (decreases),
i.e., the demand of the system is larger (less) than the ex-
ternal supply, and consequently the flow will return to its
original value. However, if the flow is perturbated to
increase (decrease) at point C, the external system supplies
more (less) than that required to maintain the flow. Conse-~
quently the flow rate will increase (decrease) until the new
operating point E (D) is reached. Therefore, the shape of
curve Q1, especially at point C, as shown in Fig. ;.6 should
be avoided to insure static stability within the possible
range of load operations (Ref. 21). TFigure L.16 also explains
the sensitivity of flow maldistribution in the same system.
For the sake of argument, assume the operating point is at

E. With an increase in heating medium flow around the local
tube to Q2, the flow decreases monotonically to point A, Pur-
turbations in any of the system variables can cause a flow
excursion or rapid deceleration to a stable point B. Further
increase in heating medium surrounding the tube to Q3 results
in operation at point ¥. Therefore, the heating imbalance
among circuits will induce flow maldistribution in a system
of many parallel heated tubes (Ref. 18).

Another phenomenon which should be considered is the potential
of flow reversal (Ref. 18). In a long, vertically-oriented
unit, the large hydrostatic head of the steam columm may lead
to flow reversal. This can occur when the difference in
hydraulic heads between two parallel downflow tubes exceed
the friction pressure drop. It has been recognized that a
superheater with heated downcomers undergoes a potentially
critical period during start-up, because at initial low flow
the friction pressure drop may be less than the hydrostatic
head (Ref. 20). However, the static head can be an important
factor in stabilizing upward flow.

L.3.1.2  ANALYSIS METHOD

For a constant pressure drop supply system, the preceding
introduction leads to the statement that the operating point
is stable if the derivative of the pressure drop -~ flowrate
curve is positive. The mathematical form is (Ref. 22)

S0P
& W

flow rate 1b/hr.
= pressure drop psi

70

where W

d
I

BY APPROVED PAGE =37

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE. CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2/;31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE12/16/7L;

4.3.1.3

L.3.1.4

‘This criterion for the flow excursion stability is well

known and the prediction techniques have been developed
which are based on the solution of the steady-state con~
servation equations for mass, momentum, energy and the
equation of state.

The static stability aspect of the reference design steam
generator was analyzed for conditions at 100, 60 and 20%
load. At each load, the water/steam and molten salt inlet
conditions were kept constant and the flow rate of water/
steam was perturbed. The effect of flow maldistribution
of molten salt, resulting in the variations of the heat
input to the individual tube, was also considered. This
investigated the flow sensitivity of water/steam flow mal-
distribution in a system of parallel heated tubes. The
thermal hydraulic performance computer code and the "Steam
and Water Pressure Drop Computer Program" (Ref. 1) were
used to generate pressure drop-flow characteristics for
each load. The potential of flow reversal was also
examined. For the calculations of part-load conditions,
Method 1 of Section 4.1.4.1 was applied.

RESULTS AND DISCUSSION

The results of static stability analysis are presented in
Figures L.17 to L.20. TUsed in these figures is the relative
flow rate which is the ratio of flow rate under perturbation
to that at normal operating condition of a specified load.
Curve @ denotes the condition of molten salt surrounding a
local tube under the normal flow distribution condition,

and curves Q+ and Q- denote the 110% and 90% of the normal
molten salt distribution surrounding the tube under consi-
deration respectively. Pressure drop was defined by cal-
culating the pressure difference from normal water/steam
inlet (bottom) to outlet (top) plenums regardless of the
flow direction. Fig. 4.20 is a continuation of Fig. L.19,
(20% load) to show the peak &P of 0.58 psi for the curve
Q- of the reverse flow.

CONCLUSIONS

Based on the comparison of results with criteria of incep-
tion of static instability, the analysis leads to the con-
clusion that the reference design steam generator is
statically stable.

BY

APPROVED PAGE  [,-38

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WIERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8ot 0]:31 | DOCUMENT NO.p/7/66 ISSUE_ 1 [DATE 12/16/7L

The static stability is insured for the slopes of all the
curves are positive ({2P/§W >0). These curves also
indicate the unit is insensitive to the flow maldistribution
among the parallel circuits. The onset of flow reversal is
limited to very small pressure difference of 0.58 pei (Fig.
4.20). This possiblity may not exist due to the fact that
the external pressure supply system is expected to prac-
tically always be operating at a much higher range.

4.3.2 DYNAMIC STABILITY

h.3.2.1

The dynamic stability work was performed by Gulf General Atomic
in 1972 under a contract to FWC. The GGA report Ref (15) in
its original form is presented in the Appendix B for further
reference. An abstract of the GGA report is presented here.

INTRODUCTION

Dynamic instability encompasses the possibility of small
density perturbations in the steam producing sustained and
growing disturbances within the steam generator. In this
regard, density wave perturbations are analagous to the
velocity perturbations in incompressible flow which can
give rise to sustained flow disturbances and eventually
produce a transition from laminar to turbulent flow. Since
the compressible flow of the steam is already turbulent,
unstable density wave perturbations will not lead to a flow
transition but will lead to other undesirable effects such
as mechanical vibration or thermal cycling of the tubes.
Because the perturbations and possible instabilities are
time variant phenomena, the instability is classified as
dynamic as opposed to static instability.

L.3.2.2 ANALYSIS METHOD

The steam generator for the Molten-Salt Breeder Reactor has
been analyzed at approximately 100, 80, 60, LO and 20 per—
cent of the rated load.

For this investigation an existing code, DYNAM, was modified
to permit analysis of the dynamic stability characteristics
in the supercritical region. The DYNAM code is based on a
method in which the governing equations are derived from
conservation principles for mass, momentum, and energy.
These time-dependent equations, simplified by considering

a single spatial coordinate along the tube axis, are

BY

APPROVED PAGE -39

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FWC FORM 172 - |

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

CHARGE NO. 8-25-2J;31 DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/74 |

linearized, Laplace—transformed, and integrated over small
spatial increments. The resultant linear perturbation
equations are examined within the framework of feedback
control theory to determine if the design is stable or un-
stable. Specifically, +the Nyquist stability criterion
has been used to predict the stability characteristies of
the steam generator. The steps in this procedure were
discussed fully in GGA report (Ref. 15),

4.3.2.3 CONCLUSIONS AND RECOMMENDATIONS

The results of the CGA report indicate that the System is
highly stable and will not amplify naturally occurring small
scale perturbations, Analyses of the effects of inlet ori-
ficing, exit orificing, and pressure level indicate trends
opposite to those observed in suberitical, two-phase systems.
Stability is enhanced for this supercritical flow by increasing
exit orificing, reducing system pressures toward the critical
pressure, reducing flow rates and reducing heating rates. By
comparison the stability of suberitical two-phase flows is
enhances by increasing inlet orificing and increasing system
pressures.

The results of the analysis tend to agree with the ides that
stability can be qualitatively checked by considering the
density ratio between the inlet and the outlet conditions as
a function of the systenm bressure. 1In the supercritical re-—
gion, an increasing density ratio with Pressure tends toward
instability and a decreasing density ratio toward stability.
Below the critical point, the ratio of the dengity of saturated
water at the inlet to the density of the saturated or super-
heated steam at the outlet decreases with increasing pressure
toward the critical point. Above the critical point, the
ratio of the demsity of the supercritical steam at the inlet
to the density of supercritical steam at the outlet increases
with increasing pressure for a constant heat input. Hence
the opposite trend of the effect of pressure on stability
above and below the critical point is not surprising. Similar
arguments can be made for the effects of other parameters by
considering their effect upon the density ratio. Quantita-
tively, systems with density ratios less than 50, which cor-
respond to the pressures greater than 600 psia for water
(Ref. 16), are generally stable. In the present case, the
density ratio is much lower than 50 and the system is highly
stable.

BY

APPROVED | PAGE )-L,5

FWC FORM 172 - I

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-20,31 | DOCUMENT NO. wp/7L/66 | ISSUE 1 DATE  12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHEREF. CHANGES HAVE BEEN MADE

Although both the quantitative analysis and the qualitative
discussion indicate that the system is highly stable, it
must be pointed out that experimental tests are required to
confirm the system behavior, Test data are available on
the behavior of subcritical systems, however data on super-
critical systems are quite limited. It is recommended that
tests in the supercritical region be performed in order to
confirm the stability of the system as determined by the
analysis.

L.L SYSTEMS RELATED TO STEAM GENERATOR

L.L.1 START-UP SYSTEM AND WATER CHEMISTRY

L.L.1.1 START-UP SYSTEM

Theofreezing temperatures of the fuel angd coolant salts are
930°F and 725 F respectively (Ref. 23). The start-up system
must therefore provide for the initial coupling of the steam
generator to the steam power system without freezing of salt
and with a minimum imposition of thermal shock. The salt °
systems must be filled angd circulating isothermally at 1000 F
before heat withdrawal can be initiated by decreasing the
coolant salt temperature. The temperature of the feedwater
must reach 1000°F utilizing a startup boiling before engering
the steam generatgr, and then it will vary between 1000 F at
zero load and 700 F in the 5 to 100% load. The 5% initial
load operation requires a startup boiler of some 225,000 1b/hr
steam capacity and a further increase to a 10% initial load
as presently envisioned by ORNL would require a 450,000 1b/hr
capacity boiler (Ref. 2,).

Additional equipment is necessary to provide the feedwater
conditions for starting, hot standby, and shutdown. This
inclydes an auxiliary start-up boiler capable of producing
1000°F supercritical steam, an auxiliary boiler feedpump,

a desuperheater and a steam dryer. The overall MSBR steam
plant start-up and shutdown system is shown in Figure L.21.
The startup procedures for the salt and steam systems are
outlined in the following sections.

Lh.ly.1.1.1 SALT SYSTEMS

The primary and secondary cell electric heaters are turned
on, and the primary and secondary circulation pumps are
started to circulate helium in the salt systems, When the
temperature of the secondary system reaches 850 F, the loop

BY APPROVED PAGE  L-l6

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2)31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

is filled with coolant salt from the heated drain tank, and
saltocirculation is started. When the primary system reaches
1000°F, it is filled from the fuel salt drain tank, and salt
circulation is commenced. Both salt systems will continue to
be circulated isothermally at 1000 F until power generation
is started. The primary and secondary-salt flow rates are at
the levels required for the zero-power level. The reactor is
then brought critical at essentially zero power and salt cir—
culating ig both systems, including the steam generators at
about 1000 F. -

L.L.1.1.2 STEAM POWER SYSTEM

Concurrent with the salt systems being electrically heated,

the steam system is also being heated. TFeedwater is cir-
culated through the mixer, pressure booster pump, attemperator,
boiler extraction valve (BE), desuperheater, condenser, de-
minearalizer, low-pressure feedwater heater, and deaerator.

A fraction of the feedwater is circulated through the auxiliary
boiler, while the remainder is circulated through the high-
pressure feedwater heaters before returning to the mixer to
complete the cold clean up circuit. Circulation of the feed-
water continues in this manner until the chemical requirements
of the feedwater for cold cleanup have been met. Cold cleanup
of the steam system is accomplished with all four of the steam
generators by-passed. When cold clean up is completed, the
feedwater flow through the heater gtring is diverted from the
mixer and recirculated back to the hotwell or through the

shell side of one of the high pressure heaters before passing
to the condenser. Feedwater flow through the auxiliary boiler
to the mixer is adjusted to the startup value,

The auxiliary boiler is then started. As the auxiliary boiler
load is raised, the steam produced is used to supply the main
turbine seals and deaerator. The steam downstream of the
boiler extraction valve (BE) passes through the desuperheater.
This steam is used for heating the feedwater in the high-~
pressure feedwater heaters, for warming and rolling the

boiler feed pump drive turbine, for warming the steam piping,
and for rolling the main turbine.

The steam/feedwater temperature is held below SOOOF until
feedwater requirements for hot cleanmip have been met. When
the auxiliary boiler reaches full pressure and temperature,
the steam at about 3600 psia and 1000°F at the discharge of
the mixer can be admitted to the steam generator. The steam

BY APPROVED PAGE  L~L7

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. §_o5-0),31 | DOCUMENT NO. wND/7L/66 {1sstE 1 |patE 12/16 /7L

generator bypass flow is then decreased until the full
auxiliary boiler flow passes through the steam generator.

When the steam system is ready to take on load, the control
of reactor is adjusted as required to maintain the desire
salt temperatures as the feedwater flow is increased. While
the steam/water flow is being established in the steam ge-
nerators, the temperature of the feedwater recirculaing to
the condenser will be raised to 550°F at the discharge of
the last high-pressure heater. The thermal load on the
steam generator is then incgeased byolowering the feedwater
inlet temperature from 1000 F to 70Q F by mixing this feed-
water from last heater and the 1000 F steam from the auxiliary
boiler in the mixer. When the 700°F feedwater temperature
is reached, the boiler feed booster pumps are started and
the feedwater pressure to the steam generator is raised to
about 3800 psia, which permits the use of the exit steam
from the steam generator passing through reheat steam pre~
heaters to heat the feedwater in the mixer instead of the
auxiliary boiler. The auxiliary boiler system is then taken
off line making the system self supporting.

The load is gradually increased and the reactor power is
adjusted accordingly. At this point in the startup procedure,
part of the steam generator output is going to the mixer via
the reheat steam preheater, and the remaining steam is going
through the boiler extraction valve (BE) to drive the main
boiler feed pumps, etc. If the load is about 5%, the main
turbines which have previously been warmed, can now be gra-
dually brought up to speed and temperature, first using
steam from the hot standby equipment (steam dryer). This
steam will give a turbine valve opening equivalent to that
at about 20% load with 3600 psia throttling conditions, so
that the throttle pressure rise may occur without having to
move the turbine control valves.

As the steam load is slowly increased, the reactor power is
matched to the load, and salt temperatures are kept at the
desired level. The load is held essentially constant until
the system comes to equilibrium, at which point the reactor
outlet temperature set point is adjusted to meet the require-
ments for subsequent load-following control.

As the load increases, the main turbines will use steam
taken directly from the steam generator. The boiler-turbine
valve (BTV), or a control-type bypass is gradually opened

BY

APPROVED PAGE L-08

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2,31 | DOCUMENT NO. ND/7L/66 | Issug 1 |pate 12/16/7L

while the feedwater flow is increased until, with a wide open
boiler-turbine valve, the throttle pressure is 3600 psia and
the load is about 200%. At thig power level the normal control
gystem regulates the reactor outlet temperature as a function
of load, and the steam temperature controller holds the steam
temperature at 1000°F,

L.4.1.2  WATER CHEMISTRY
The steam power system of the MSBR plant will not require
special water treatment. Aside from the steam generator,
material of construction shall not differ from present-day
fossil fueled supercritial cycles (Ref. 20) of startup boiler.
The recommended limits for feedwater conditions at the econo-
mizer inlet are given in the following table for normal opera-
ting conditions and quring start-up (Ref. 25).
Normal Operation Start-up
Total dissolved solids - ppb 50 —
Total iron ~ ppb B 50
Total copper —~ ppb 5 20
Total silica -~ ppb 20 30
Dissolved oxygen -~ ppb 5 10
pH 9.3 - 9.7 9.3 ~ 9.7
Conductivity - mmhos 0.5 1.0

During the start-up period, there will be variations in the
concentrations of the various feedwater contaminants due to
changes in temperature and flow conditions and placing into
service of cycle components. The feedwater conductivity
must be below 1.0 mmho before lighting the burners of the
startup boiler. Also after firing has started, the fluid
temperature at the roof outlet of the startup boiler mugt
not be permitted to rise above 500 F until the iron content
of the feedwater entering the economizer is less than 50 ppb.
The limits given in the above table for the startup condi-
tions are for continuous operation. Transient values higher
than those given in the table may be tolerated for a short

BY

APPROVED PAGE L~L,9

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

i

CHARGE NO. 8-25-21,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7h

time only, and if they do not decrease, firing rate and
temperature may have to be reduced until satisfactory
values are obtained.

There are two basic approaches to keep the various con-
stituents to the levels below which they will not cause
problems.

a. Minimize the corrosion that takes place within the
system: This is accomplished by removal of oxygen
and maintaining the pH of the condensate at the spe-
cified level. Oxygen is removed by chemical scaven-
ging with hydrazine and mechanical deaeration in the
condenser. A nominal hydrazine residual of 0.020
ppm is maintained at the economizer inlet. Feedwater
pH is controlled by adding ammonia or a volatile amine
such as morpholine or cyclohexylamine.

b. Remove corrosion products and leakage salts from the
system: Corrosion products, silicon and salts from
condenger leakage are removed by the full flow con-
densate demineralizer. Both ionized and suspended
matter are removed as the unit acts ags a highly ef-
ficient filter as well as an ion exchanger.

L,.,.2 PRESSURE RELIEF SYSTEM

The steam generator was designed to the sPecifications of
the ASME Boiler and Pressure Vesgsel Code Section III, Nuclear
Vessels for class A vessels. The design conditions are:

(Ref. 26)
shell side tube side
Design temperature, OF 1150 : 1120
~ Design pressure, psia , 300 3800
Allowable stress, psi 9500 11600

There is no violent exothermic reaction between the coolant
salt and steam, however, the mixture of salt and water is
corrosive to the material of steam generator (Hastelloy N)
(Ref. 1).

The pressuré relief system was designed to protect the steam
generator and coolant salt system from the overpressure and

BY APPROVED AGE  ];-50

AU P

1000

3GOO P 1000 F ; - LT A .
s Y ‘—-—-[l.z MILLIAL 12U ]-—\c-p ;
COOL ANt SALY i
WLt = DOVLIER
— -— N Pulap
M | KEHEATER < |“" s ) ‘\ -’ .
COOLANT 4 St —,
—) . I
Wau b <70 P B l ConuLhn ~ l . .
' - Y 9 1y Ay
[T At g B,
wol ° ’ = LP
. , Wi ’.T; EXTHACTION STEAM e
1 AS ] 3600 P uvo ¥ i FRom el
b Ry N s (St [ L R "e l
: e = FumP &
KEOEAT STUAM clapk ATEIY 1050 P )
oo = "o E ——
Wt | 2 l g y "l‘ &= ui Pue
o o Nm'e S % - ' o ong Tuhihg
< ud] l Z ut ! puoncn ¢ |so HFb G
(2) 30 temer nnton : :"c"‘?.,"_.___ ¥ lle: ‘l’f‘:\CllON 4 Jhy DEAERATOR
\ 3 . ’ oft Y S1AK0D T QR
Wy, ke 3 ’\_—-——M— : PORER TUihE (S
Bow b ‘ 04 SiLnam - (VR TR RV ]
L N} - DESUNER- o —————) -
Vag vt fil nEATLK z
: > ;
COULANT SALT 180 F '
i p ! 110U P 556 F 2
l 4 N L PUXTHACTION STEAM DUKING MEAT L ECTION
(\ ! STANT =ulk AND STANDLY THTICE vaew
v . EXTHACTION S1EAM
1 [
5nmo- s 3490 P ( )-dp-—‘—-'— = = 4
' <2 ~ wansfF
(e 1 \’( S w4 ey R s X ™
- S 1R " 1 G PiE S5, uumnltn HEATENS fun it g = '7
COULANT SALT 150 | r e MIER - I retivntn
sG 0F LOOLTI I Pume A nEaliny 1 huosten
STEAA CUNERATON

—

IumMp
> .
AUA O ER A p
12000300 1L/ ) Aux ur g
MOTUN DIIVE
MSIR steam plant startup and shutdown system, .

Re:.(23)

Fig. L.21

PAGE L-51

FWC FORM 172 - |

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J..

CHARGE NO. 8-25-2),31 { DOCUMENT fio. ND/7L/66 ISSUE 1 DATE 12/16/7),

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

highly corrosive mixture, should steam tube suddenly rupture.
It was assumed that the coolant salt system could withstand
a continuous pressure of 220 psi without damage (Ref. 3).

The present ASME Code Section III, which governs becauge of
emergency cooling consideration, accepts only relief valves

ag primary relief devices. However, the corrosive property

of the mixuure of salt and steam, requirement of rapid pressure
relief, are against using relief valves as an appropriate pri-
mary relief device. A new subsection NH under Section IIT is
in preparation for inclusion in the new edition of the Code
which is expected to recognize rupture disks as primary relief
devices for certain types of overpressure conditions (Ref. 27).
The Molton-Salt steam generator would be listed as an over
pressure condition to which the rupture disk or equivalent
device is applicable.

There are three basic types of rupture disks; prebulged disk,
reverse buckling assembly and the snap-over assembly. In
addition, the rupture disk for the British PFR is a hinged
plate supporting a nickel membrane, which serves as the sodium
seal; the plate itself is held by a shear pin, designed to

fail approximately twice the normal operating pressure (Ref.
28). However, it was learned that regponse time observed in
testing the PFT design was about 10.25 milliseconds slower than

- with the prebulged disk. All of these four approaches are

shown in Fig. .22 (Ref. 17). The evaluation and comparigon

of these four types were discussed in Ref. 27 and Ref. 29.

The reverse buckling assembly was recommended and therefore
used for this design. The reverse buckling disk has several
advantages. It is under compression rather than tension, and
the controling factor, elastic modulus, is insensitive to en-
vironmental conditions. The reversing pressure can be pre-
dicted accurately, and collapse occurs within £ 24 of the normal
pressure rating. It has 10 - 15 times the life expectancy of,
and is thicker than, the prebulged type. It can be used at
system operating pressure up to 90% of the rated burst pressure.
It needs no vacuum support and withstands repeated pressure-
vacuum cycles. :

The reverse buckling digk, for a given geometry buckling, is
controlled by elastic_modulus of the material. Within the
operating range (1150 F - 850°F) the elastic modulus of
Hastelloy N varies only about 6%. Therefore the Hastelloy N,
which is also, chemically, very compatible with coolant salt,
could be suggested be the material for the reverse bulking disk.

BY

APPROVED PAGE )52

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25~-21,31 | DOCUMENT NO. ND/7L/66 ISSUE 1 |DATE 12/16/7hL

There will be no cover gases to protect structure members.
The steam generator cell of the generating station building
is maintained at about 1000 F to prevent salt from freezing
(Ref. 23).

Reversed buckling disks can be used at gystem operating pres-
sures up to 90% of the rated burst pressure. The rated burst
pressure was determined to be 350 psia. A 20 in. diameter
disk (thickness = 0,08 in) would provide adequate relief

area 1o dump the mixture of reaction products, thus preventing
system from continuous pressure rise. One rupture disk as-
sembly is recommended at each of the inlet and outlet salt

piping.

Following a tube rupture, the block valves of water/steam

side would be closed. The mixture would be dumped through

the rupture disk to a dump tank. The connecting pipes would
be of the same size of rupture disk, Each steag generator
would have one dump tank of capacity of 2500 ft” with ample
freeboard to hold the mixture and vent out the steam. The
tanks are essentially conventional tanks with internal spiral-
guide vanes. These guide vanes cause the mixture to exper-
ience centrifugal forces that separate most of the liquid,

or solid reaction products from the gaseous reaction products.
The gaseous productions are discharged to the separator. The
separator is usually installed to remove liquid/solid products
from the gaseous discharge, and ensure no reaction products
are discharged to the atmosphere.

BY

APPROVED PAGE  [;-53

embly

S

A
i 2
.o

Reverse Buckiing

&

1

LoV
Shear Pin Corncept

TSR (2
PN A S A0 VAN W S \\\\ f
IR . Y - NRA U X N b .\\\\\\\.
AL AR R G W N, NSNS NS TS
PN W Jo i QPN NS N NN 't \\\\\C\
~/., ..v“/t. ¢/, 2 & ...— .\”.\ 4
) ) e NN '+

Prebulged Disk Assembly

Snap Over Assembly

PAGE L~-5L

Various Rupture Disk Concepts

Fig. L.22

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2);31 | DOCUMENT NO. ND/7L/66 ISSUE 1 |DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

L.5 REFERENCES

1. The Development Status of Molten-Salt Breeder Reactors,
ORNL - L4812, August 1972, Oak Ridge National Laboratory.

2. Design Studies of Steam Generator for Molten Salt Reactors
Monthly Progress Report #10, August 1 -~ August 31, Foster
Wheeler Corporation.

3. Proposed Scope of Work and Requirements for Design Studies
of Steam Generators for Molten Salt Reactors, Enclosure 2,
Union Carbide Corporation, March 1971.

;. Heat Transfer to Supercritical Water in Smooth - Bore
Tubes, H. S. Swenson, J. R. Carver, C. R. Kakarala,
Journal of Heat Transfer, November 1965.

5. GGA - HTGR Steam Generator Evaluation of Surface Mismatch
Using Probability Methods, L. Rianhard, R. M. Costello,
P. J. Prabhu, W. T. Klamm, Foster Wheeler Corporation,
July 1970.

6. Molten~Salt Reactor Program, Semiannual Progress Report,
Period ending August 31, 1972, ORNL - 1,832, March 1973,
Oak Ridge National Laboratory.

. Private communication from W. Apblett to J. Polcer on
thermal conductivity data of Hastelloy N (6/20/7L).

8. Design study of steam generators for Molten Salt Reactors
Monthly progress Report #8, May 20 -~ June 30, 197L,
Foster Wheeler Corporation.

9. Design study of steam generators for Molten Salt Reactors
Monthly Progress Report #10, August 1 - August 31, .197L,
FToster Wheeler Corporation. '

10. A general summary of the ORNL 1000 MW(e) Molten-Salt
Breeder Reactor Reference Design, Enclosure 1, Union
Carbide Corporation, November, 1970.

11. Fluid Dynamics, J. W. Daily, D.R.F. Harleman, Addison-
Wesley Publishing Company, Inc., 1966.

12. Engineering Department Manual, Volume I, Basic Design,
Foster Wheeler Corporation.

13. An Analysis of Thermally Induced Flow Oscillations in the
Near-critical and Super-critical Thermodynamic Region,
Novak Zuber, General Electric Company, May 25, 1966.

BY - APPROVED | P PAGE L-55

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT | LIVINGSTON, N. J.

CHARGE NO. 8-25-21,31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/Th

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

1);. Engineering Department Manual, Volume I1T, Computer Programs,
Foster Wheeler Corporation.

15. Analysis of Dynamic Flow Stability in Steam Generators for
the Molten-Salt Breeder Reactor, B. E. Boyack, Gulf General
Atomic, Gulf-GA - A12416, November, 1972.

16. Lecture Series on Boiling and Two-phase Flow for heat transfer
engineers, University of California, Berkeley, California,

17. Review of Two-Phase Flow Instability, J. A. Boure,

18, Static and Dynamic Stability of Steam-Water Systems Part 1,
Critical Review of the Literature, L. E. Efferding, General
Dynamics, GA-5555, October 196l.

19, Supercritical Furnace Design, Foster Wheeler Corporation.

20. Recent Experiences with Radient Superheaters in Central
Steam Generators and their Effect on Design Criteria,
R. P. Welden, H. H. Pratt, FWC, American Power Conference,
Volume XXIV, March, 1962.

21. Test and Evaluation of Alco/BLH Prototype Sodium-Heated
Steam CGenerator, S. M. Cho, etc., LMEC~-Memo-70-20,
January 1971, Liquid Metal Engineering Center.

00, Performance Changes of a Sodium-Heated Steam Generator,
S. M. Cho, K. A. Gardner, etc., 71-HT-15, ASME.

23, Conceptual Design Study of a Single~Fluid Molten-3alt
Breeder Reactor ORNL-4541, ORNL.

o),. Startup Steam ORNL 1000 MW(e) MSBR, C. R. Clark, Service
Department, FWC, Inter Office Correspondence, September 29,

1972.

25, Operating Instructions, Section 3, Feedwater and Cycle
Cleanup, IWC.

26, Design studies of Steam Generators for Molten Salt Reactors
Monthly Progress Report #11, September 1 - September 30,
1974, FWC. |

BY APPROVED PAGE 56

FWC FORM 172 - 4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION

CHARGE NO. 8-25-2,31 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/7h

27.

28.

29.

30.

State-of-the-art of Rupture Disks for LMFER Application,
J. P. VerKamp, NEDM-13981, GE, July, 1973.

Status of ILMFBR Reheat in Wegtern Burope - 1972, WASH-1219,
AEC, March 1973.

Evaluation and Procurement Guide For Three Types of Metallic
Rupture Disk Assemblies, ORNL~-TM-4OL6, ORNL, January, 1973.

Computer Programs For The Analysis of Complex Decision
Problems, Stanley I. Buchin, 8-171-070, EA-C-905, Harvard
University, September, 1970.

BY

APPROVED PAGE  4~57

LIVINGSTON, N. J.

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

DATE 12/16/7,

CHARGE NO. 8-25-2431 | pocumENT NO. ND/7L/66 ISSUE 1

SECTION 5
STRUCTURAL FEASIBILITY ANALYSIS
BY

K Tonse
DR. ¥. J. LEVY

Approved by

4 31

L
s -7 7 - ’
v T Cpran F

C. F. Nash
Structural Analysis Section Manager

BY APPROVED

PAGE G-

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2L31 | DOCUMENT NO. WD/7L/66 ISSUE 1 DATE 12/16/7)

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

¥WC FORM 172 - |

TABLE OF CONTENTS

Pages
5.1 Stress Analysis of the Salt Inlet Nozzle 5-1
5.1.1 Introduction 51
5.1.2 Summary and Conclusion 5-1
5.1.3 Stresses Due to Pressure 5-6
5.1.L Stresses Due to Temperature Transients 5-8
5.1.5 Simplified Inelastic Analysis 5-10
5.1.6 Fatigue Analysis and Creep Fatigue Interaction 5-11
5.2 Stress Analysis of the MSBR Tubesheet-Header Assembly 5-~13
5.2.1 Introduction and Sumfilaxy 5-13
5.2.2 Loading Conditions - 5-16
5.2.3 Some Details of the Finite Element Model 5-19
5.2.4 Simplfied Inelastic Analysis 5=30
5.2.5 TFatigue Analysis and Creep Fatigue Interaction 5-33
5.3 Stress Analysis of the MSBR Shell 5-35
5.3.1 Introduction and Summary 5-35
5.3.2 Shell Stresses at Location A o-11
5.3.3 8hell Stresses at Location B 5-42
5.3.4 Shell-Shroud Juncture Stresses IS—QB,
5.3.5- Simplified Inelastic Analysis 5-49
5.3.6 TFatigue Analysis and Creep Fatigue Interaction 5-50

BY APPROVED PAGE 5-%

FWC FORM 172 -~ |

NOTATIONS IN TH1S COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE -~

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION

LIVINGSTON, N. J.

DATE 12/16/7L

CHARGE 0. 8-25-2131 | pocument wo, ND/7L/66 | Tssuz

5.4 Stress Analysis of the MSBR Tubes

5.4.1 Introduction and Summary

5.L.2 Tube Primary Stresses

5.4L.3 Thermal Stresses

5.4.4 Fiow Induced Vibration

5.4.5 Simplifieg Inelastic Analysis

5.4.6 Fatigue Analysis

5.5 Tube Rupture Analysis

Pages
5-52
5-52
5-5L
5~59
5-62
5-6L
5-6L
5-66

BY APPROVED

PAGE 5_o

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NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO 8-25.2),31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE 12/16/7)

5.1

5.1.1

5.1.2

STRESS ANALYSIS OF THE SALT INLET NOZZLE

INTRODUCTION

This report discusses the analysis of the molten salt inlet nozzle
of the steam generator, During its service life, it is expected
that the nozzle will be subjected to thermal transients which are

no worse than those described in ORNL-TM-3767, HYBRID computer
simulation of the MSBR. In the interest of conservatism, the tran-
sient shown in Figure 2, which is a combination of most severe up
and down transient, was assumed for this analysis. Dead weight

and thermal expansion loads due to the molten salt feed Pipe were
not included because they are not available at this time, Internal
bressure loadings, however, were considered. Temperature distribu-—
tions due to the assumed transient condition, and stresses due to
these temperature distributions and pressure loads, were obtained

by using Foster Wheeler's version of Wilson's Finite Flement Program.
The finite element model is shown in Figure 1. Inelastic strains
were calculated using Bree's method & FWC's simplified elastic-plastic
creep computer program.

SUMMARY AND CONCLUSION

The purpose of our analysis was to determine the feasibility of

the sodium inlet nozzle under the severe thermal transients and

other loads anticipated. Based upon analyses specified in Section IIT
of the ASME Boiler and Pressure Vessel Code, and in Code Case 1331-5,
it was concluded that both the nozzle and shell required a thermal
liner. With the liner, stresses were calculated to be within the
allowable limits. Table I summarizes the results. Based on the
table, it is concluded the design is feasible. Final conclusions

are dependent on the magnitude of the piping loads (which are
presently unavailable).

BY

APPROVED PAGE 5-1 OF

FWC FORM 172 - |4

et . T T . el e s 0 s i 13 S5 ol 4T i

e ey S

e

NOTATICNS IN THIS COLUMN INLICATE WHERE CHANGES HAVE BEEN M4DE —— -

N

FOSTER WHEELER CORPORATION

CHARGE NO 8_25_2)31

DOCUMENT NO. Np/7L/66

ISSUE

1 DATE 12/16/74

TABLE 1

STRESS SUMMARY, MSBR MOLTEN SALT INLET NOZZLE WITH LINER

Calcuiated Allowable
Section Stress Stress Range Stress
(Location) Category (psi) Limit (psi)
1-1 P (design) - Lok3 S, = 9500
; | & @ 1150 F
' (Shell) P (operating) 3170 Syt = 7500
' @1%50 F 30 yrs.
‘ (PL+Pp+Q)p 19019 (EL.19) | 35m = 67,500
; 2-2 P, (design) 5100 9500
E (Shell at nozzle) Pm (operating) 1,000 7500
; (PL+Pp+Q) o 26,480 (EL. 158) | 67,500
| 3-3 P (design) 5360 9500
| | (Intersect. P, (operating) 4200 7500
: Shell & Noz.) .
; - (PLPb+Q)R 33,396 (El. 187) 67,500
f L=l P (design) 5236 9500
; (Nozzle at Shell) P (operating) 1,100 7500
| (PLB,+Q)R 19,108 (E1. 215) | 67,500
; 5-5 P (design) 2199 9500
(Noz. at Start P (operating) 1960 7500
of Reduced Wall) | ™
(PP +Q)y 10,79 (E1. 258) 67,500
6-6 P (design) 6415 9500
(Nozzle at P (operating) 5025 7500
Feed Pipe) n
, - (PLPb+Q)R 9,278 (E1. 336) 67,500
BY APPROVED

PAGE ¢_, OF

FWC FORM 172 - )

—-—NOTATI

——

OF

A 8 e M . ke e vams an e e e,

"
.

3 HAVE BEEN ¥

~
AT

ATE WHERE CHAN

-~
s

&AL

T

NS IN THIS COLUMN INI

T T T e+ttt it ot 8 e

-
s

. T

FOSTER WHEELFR CORPORATION

-_

CHARGE NO 8_p5_2),31

DOCUMENT NO. Np/71/66

ISSUE 1

DATE

T R o i . Ao e e e £ s

ks L S e+ o g ot s+ e+ e

T e A g

T bt At - el o R vt b S %1 0 e+

TABLE 1 (CONT'D)

m
(PP, +Q)R

does not control

Section otress Calculated Allowable Stress
(Location) Category stress Range (psi) Limit (psi)
-7 P ({design) 7840 9500
m
(Feed Pipe) P, (operating) 61,0 7500
‘ (PLPb+Q)R does not control ——
8-8 P (design) pr = 2700 9500
m 2t
(Shell) P (operating) 2100 7500

Note: Allowables are at 1150 F

See Section 5.1.6 for creep fatigue

interaction and inelastic strain analysis

See Appendix C-2 for material properties and allowables.

BY

APPROVED

| Page . .

OF

12/16/74

MOUTEN SHLT RAET NOZ2LE KITH LIMER

USEA:  AOYT, SULLJven CEPT. 156

t i'w
net

cair AN

SHELL

rr—— ——

3\
\,
EL. 18T -
| L] \2
N
N ————— L
4 4 -
9
EL. 215 - 1 | MOLTEN SHLT INLET
0ERLE .
| NOZ2LE WITH LINEQR
L
ORI gl WAL GEOMETRY
EL.258 . [ T 1
5 |
Ly B

6 7i
i o
FL. 33 — | FIGURE |
- i FINITE ELEMENT
s e MODEL
7 I

ORICINS.  CETETAY

Paoca O},
o °
< SSUME TEMp. HELD KT €S5S¢
= / AsSsumM M P EL . €S0
F / URTIL T = 185 Sec. THE L
1200 ( QEfiDLIELL\{ RelSED To 11Sc°F
{
\.\
\
900 | \
—=3
TeEMP. 0F SECONDRRY SaLT AT
00 | STERM €CEN. INLET V.. TIME
300 |
L e 1 1 L L 1 T‘ME(SEC.)
20 40 60 80 100 izo
FIGURE 2
J . _RSSVUMED TeANSIENT LOAD CONDIT|ON
¥ ¥ (steapy c<TaTe Temep. ERILT M 2 s iy

e
FWC FORM 72 - |,

NOTATIONS

N b . e et e

THIS COLUMN INDICATE WHFRF CHANGES HAVE BEEN MADE - —e

R i 4 T b S i s = gy ¢

-

FOSTER WHEELER CORPORATION

CHARGE NO8-25_2),31

ISSUE

1 DATE 11/16/7h

DOCUMENT NO. ND /71, /66

5.1.3 STRESSES DUE TO PRESSURE
From the finite element computer program, stresses at various
locations of the nozzle due to a design pressure of 300 psi
were obtained. They are shown below. (Ref. 9)

; section 1-1

;

| Element Oy g O Cig

|

g 86 1692 701 1,078 1137

| 87 1584 737 4003 1186

E 88 1508 740 3928 1209

2 89 14,62 702 3850 1210

| 90 11,59 629 3772 119),

f 91 1509 lE}l 3693 116l
Avg, 1536 - 670 3887 1183
Average principal stresses are
Jq = 2362‘L

g 8, = -156 > S = LoL3

f C73 - 3887 j

t

! |

§ pection 2-2

! . - —

! Element Cp 0, U Ory,

? 126 3862 2146 61,21 24,30

5 127 2149 159), 5813 14h3

} 128 1285 1032 Sl56 8L3

i 129 559 L89 5143 396

! 130 -17), -2 1,810 19

| 156 -701 ~571 LSLT -88

157 ~1002 -773 4321 -583
158 -1683 -667 4198 ~657

Avg. 537 1,05 5092 L5
Gp=9% (,=-8 03 = 5092

5 S = 5100

BY APPROVED [ PAGE ¢_¢  OF

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

Section 3-3 |

Element G r G’ 2 6T Grz
141 1868 3841 6308 2112
142 1761 1875 5868 1123
143 1354 793 5592 545
1L, 900 71 5L11 132
145 465 -453 5302 -163
184 62 -769 5237 -223
185 236 -812 5322 -1),8
186 -175 -934 5303 -398
187 -375 -988 5358 ~352

Avg 677 292 5522 259

O, =808 =162 - 5520
S = 5360

Section )}

Element ‘j‘r Gz GT { )rz
210 222 2701 5102 566
211 299 1686 4989 986
212 Lol 1174 5021 1083
213 146 632 L875 898
21) 47 30 1,81.8 658
215 =253 ~722 L725 235

Avg, 128 917 4922 738

(51 =130 (5= -314 .<55 = 1,922
S = 5236

Section 5-5

Element 61‘ 62 : GT Grz :
254 . -60 ~-750 1698 -193
255 - =3 26 2008 0
256 ~-63 91,8 2361 115
257 -131 1949 2746 137
258 -210 3056 3180 L0

Avg. ~-100 1046 2399 20

S = 2499

BY APPROVED | PAGE c_7 o

FWC FORM 172 - i

CHANGES HAVE BEEN MADE ——— — —

IMN INDICATE WHIRE

-r
£

by v —

c0

NOTATIONS IN THIS

PR

e —.

FOSTER WHEELER CORPORATION

From the finite element computer program, stresses due to tempera-

ture transients at various locations of the nozzle were calculated

and are given below. Note that pressure stresses do not enter -
into the stress intensity range calculation. (Ref. 10)

CHARGE NO 8-25-2),31 | DOCUMENT NO. ND/7l/66 | ISSUE 1 DATE]2/16/7h

i Section 6-6

Element U r ‘5 2 6 T 61‘2
! 332 182 5872 6772 207
f 333 76 4890 6509 3
; 334 -43 L223 6336 -120
5 335 -168 3645 6189 -149
| 336 279 3051 6011 -8l
{
i
? Avg. -L6 4336 6369 -29
| .
é S = 6415
§5.1.h STRESSES DUE TO TEMPERATURE TRANSIENTS
i
g
|

WS A e e e e e e

Section 1-1

U T C- P (YT O-I‘Z
Down Trans 13570 ,9L9 204,90 828l 5153 Sec)

i Up Trans 2L 73 _1361 _3°51 33 Sec)
! Range 13323 L3876 19129 7933
| 0, =1800 (,=110 (=192
; S = 19019
; Section 2-2
i
| Down Trans 18390 3625 25200 8870 (153 sec)
! Up Trans - =3016 -829 1351 =1299 (33 Sec)
| Range 211,06 LSl 238L9 10,169
Q, = 26,170 6, = -310 63 = 23819

S = 26480

{
BY APFROVED '
PAGE _ . OF

FWC FORM 172 - L

e . bt 25 g e Sl 2 1 ACBPkraO s T

NOTATIONS IN THIS COLUMN INDTICAT®E WHERE CHANGES HAVE BEEN MADE —— ——

—X

FOSTER WHEELER CORPORATION

CHARGE NO g8-p5-2,31 | DOCUMENT NO. yp/7l/66 |ISSUE DATE 12/16/74
| SECTION 3-3
s U’I‘ Gi o'r GI‘Z
| DOWN TRANS 1859 9190 31300 L725 (153 SEC)
| UP TRANS -503 -1852 1787 -822 (33 SEC)
| RANGE 2362 10L2 29513
i U o=T7287 i =3883 (4=29513
| 0=33,3%
|
| SECTION L-L
| DOWN TRANS -30 17030 21410 -56l (153 SEC)
f  UP TRANS -191 ~2060 2206 267 (33 SEC)
' RANGE 161 19090 192 3], -831
|y =19126 (=126 C4=1923)
f S=19108
|
! SECTION 5-5
f DOWN TRANS -167 9300 12180 -27
. UP TRANS -158 1677 1395 35
¢  RANGE -9 7623 10785 -62
f
? a l=7623 Lp==9. \J3=10785
i $=10,79L
! SECTION 6-6
r 6{. GT Grz
DOWN TRANS ~73 10900 10460 -25
UP TRANS -218 1),78 3867 -52
RANGE 145 9L22 6593 27
G,=9k23"
Co=1L5 S=9278
BY APPROVED
| PacE ., oF

FWC FORM 172 - L

INDICATE WHERE CHANGES HA

NOTATIONS IN THIS COLUMN

S J—

A R o LTI M M e b T M L N ol -y bt el v ST sl AR Y. g . e ¢ e . Pt 4 s . e

FOSTER WHEELER CORPORATION

e .,

e . a—_ a8

5.1.5

CHARGE NO g_pg5_o),37 | DOCUMENT NO. ND/7./66 ISSUE 1 DATE 12/16/7.

SIMPLIFIED INELASTIC ANALYSIS

Inelastic strains were calculated by Bree's simplfied method
(see Chapter VI of the IMFBR Piping Design Guide). In the
Bree analysis, it is assumed that there are a total of 200
severe themal cycles (sum of load scrams and reactor scrams).
With a 30-year design life, the time per cycle is 262,800/200
= 131L hours.

The derivation of the creep law used in the Bree analysis is
given in the tubesheet report. The Bree analysis has been
computerized by Foster Wheeler, and the following results
were obtained:

Location Total Strain/Cycle Total Strain
1-1 Negligible Negligible
2-2 Negligible Negligible
3-3 1x107 0.02%
L=l | Negligible Negligible
5-5 | Negligible Negligible
6-6 Lx10™k 0.08%

Ref.: Computer Run NHJLOSF

BY

APPROVED [ PAGE ¢ .~ OF

FWC FORM 172 - 4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE REEN MADE —— —— e

.-

e e L A

R

e = e WL S St s i A LS, - PR 30 = ogrr Bt somo a3t ot

FOSTER WHEELER CORPORATION

g —

CHARGE NO

DOCUMENT NO.

8-2H-21131 ND/7L/66 | FSSUE 1 DATE 12/16/ 74

5.1.6 FATIGUE ANALYSIS AND CREEP FATIGUE INTERACTION

A preliminary fatigue analysis was made with the aid of Code Case
1331-L. With a maximum stress range of 33,0 Ksi, or an alternating
stress of 16.70 Ksi, the Code Case gives an infinite number of
allowable cycles. Thus, fatigue should be no problem.

CREEP FATIGUE INTERACTION

Foster Wheeler has developed a computer program which performs
elastic-plastic-creep analysis of a cylinder subjected to time
dependent pressure, temperature, and axial loads. To use the
program, one constructs a histogram of the loading (see below).
The user alsosupplies subroutines describing the creep, fatigue,
and stress to rupture properties of the material. Creep properties
were obtained from "Data for Nickel-Molybdenum-Chromium~Iron Alloy,
Iron 8", June 1, 1961, while fatigue and stress to rupture prop-
erties were obtained from "Bases for Design of MSER Systems for
Temperatures to 1300 F", ORNL Central Files Number 73-1-23. The
program automatically computes stresses and strains as a function
of time, along with creep and fatigue damage. Creep and fatigue
damage are computed in accordance with ASME Code Case 1331-5.

Section 3-3 proved to have the highest loading and only that
section was analyzed. The nozzle was idealized as a cylinder
with 1.625" wall thickness and approximately 9" inside radius. -
The loadings and histogram are shown below.

e LT
Rea 7o . AT=-12F
e v

i't
%“rfi £ ;
U SEE B oo
ks /r F Load ' U
[ Serave T .

L AT-217°F

« i Cydes 2o hes,
r . =l . R e e e L e e —c———
B2 = 5360 >I

solving for p
P = 950 psi

E-AT _ 31,300 (up transient)

2(1-1) -1852 (down transient)

AT = 217°F

~13°F

where E = 26.2 x 109, «= 7.72 x 10-6

BY

APPROVED PAGE 5_37 OF

FWC FORM 172 - )

.

e s e o s i e e s e o

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE ~ — —- e e

FOSTER WHEELER CORPORATION

|

L.

DATE 15 /16/71,

s s e

PR

A . e A o A A e, = o e

Tt ——

CHARGE NO g.p5.p),31 | DOCUMENT NO. wp/71,/66 ISSUE 1

PROGRAM RESULTS

There are a total of 50 cycles, each cycle consisting of 3 load
scrams and 1 reactor scram. The program was run for 5 cycles and
the strain for each of the remaining cycles was estimated by

taking the difference in strain between the fourth and fifth cycles.
Maximum total strain = 0.29 + 45 (0.01) = 0.7L4%

Creep damage = 0.00222 + L5 (0.0L8 x 10_2) = (0.0238

Maximum strain range = 0.2.8%

-l
Fatigue Damage = SO/lO5 = 5 x 10 (maximum temperature = 1150 F)

Ref: Computer run EBHJLO6C

BY

APPROVED ' PAGE 512 OF

FWC FORM 172 - 4

UMN INDICATE WHERE CHANGES HAVE BEEN MADE — - — —— oo

e v Cr e <4575 1t

L

!

e
LN,

NOTATIONS TN THIS

T e e et 0 A e+ ot e o e

A P T v 3t b5 - P e 2 oo

T ——— A A . + P £

FOSTER WHEELER CORPORATION

CHARGE NO 8.p5_p),31 | DOCUMENT NO. ND/7L/66 |ISSUE 4

DATE 15 /16/71,

e o s e o 1y ks hw

FER et o T Y. o b

5.2

5.2.1

STRESS ANALYSIS OF MSBR TUBESHEET-HEADER ASSEMBLY

INTRODUCTION AND SUMMARY

The major tool in analyzing the tubesheet header assembly in Foster
Wheeler Corporation's Finite Element Computer Program (Reference 1 ).
A finite element model of the structure was made and is shown in
Figure 1. This model was used to determine both pressure and thermal
stresses. These stresses were compared to the allowables, using
ASME Code Case 1331-5 as a guide.

As seen in Figure 1, the upper (header) side of the tubesheet is in
contact with steam, while the lower portion of the tubesheet is in
contact with molten salt. The severest thermal stresses occur during
transients specified in Reference 2. The severest up and down tran-
sients were found to occur during a Ramp Change in Load Demand from
100 to LO%Z in 3 seconds (Figure 2) and Insertion of Two Safety Rods
(Figure 3), respectively. Outlet steam transients were found to be
more severe than inlet. Therefore, because inlet and outlet tube-
sheet geometries are identical, only the outlet tubesheet was analyzed.

Eight locations of possible high stresses were analyzed and are shown
in Figure 1. Primary membrane, local membrane plus primary bending,

and local membrane plus bending plus secondary stress intensity

ranges were obtained at 11 locations. The stress intensities obtained,
together with allowables, are shown in Table 1. Except where indicated,
primary stresses are due to a design pressure of L0OOO psi. Allowables
are based on operating (30 year life) conditions. This combination

of loading and allowables is conservative.

Although, according to Code Case 1331-5 there is no limit on primary
plus secondary stress intensity, the 3Sm limit of Section IIT of the
ASME Code is given for reference purposes.

It i1s seen that all primary stress intensities are within the allow-
able limits, Inelastic strains, computed by Bree's simplified method,
are also seen to be within allowable limits.

Computer plots of the stresses due to temperature transients are
shown in Figures 5 to 13.

BY

APFROVED
—J PAGE .. OF

TABLE 1

STRESS SUMMARY - MSBR TUBESHEET - HEADER ASSEMBLY

Calculated

Allowable Maximu
Location Stress Stress Limit Stress Range | Temper
(Section) Category (PST) (PSI) ture,
1-1 P St = 13,000 8,570 1075
(Center of T.S.) -
(Operating Pressure) Pr+P KS, = 1L,100 13,040
(PL+Pb+Q)R 38, = 69,600 23,360
2-2 P 13,000 12,160 1075
m
(Edge of T.S.) P+ 13,200 8,080
- (P_+F +Q)p 69, 600 1,220
3-3 B, 8,000 6,150 1150
(T.S. - Shell (P +P_+Q)p 67,500 49,810
L b
Juncture ) |
i P 19,000 6,550 1000
(T.S. - Header (P_+P_+Q) 71,500 20,310
L b g
Juncture)
5-5 P 19,000 7,150 1000
(T.S. Header -m
Juncture ) (P_+P +Q) 71,500 .60, 8L0
L b 'R
6-6 P 19,000 8,470 1000
m
(Header Wall) (PL+Pb+Q)R 71,500 31,320
7-7 P 19,000 12,970 1000
m
(Pr+P +Q) 71,500 37,090
b R
8-8 P 19,000 12,530 1000
| (P_+P +Q) 71,500 does not control
f Ly ™R -
NOTE: 1. There are no primary bending stresses at Section.3_3through 8-8

2. See Section 5.2.5 for creep fatigue interaction.

3. See Appendix C-2 for material Properties and allowables,

*Primary plus secondary stress ranges conservatively include peak portions.

Dn Fa . Tal E’._-I l'l
MSBA HEADER-TUBESHEET ASSEMSLY

USEA:  SHEAMAN [P DEPT. 156
] 2.2
notsme H————
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& HEADER

Flcuke | - (o)— R —(®
FINITE ELEMENT MODEL OF iypyssany
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ASSEMBLY | ~

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PN TAV AN
175 158 163
181\ N 168 Y o
1 t e 167 161
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72 NJes 15 153
158 »
164 e D N
156 16 N
13 0
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2
2 T "
X

TUBESHEET ®| I !

l® | @)\ = g #l 2.,"—@

| L) LAl A8
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; = = il ux WCIN I 7 W7 5
- 3, (=1 )]
h » n = b 385 )
g B B ™ P2 Lk 37 — @
: - © n n * »2\ w3 -
= e - d=d I-¥ o
2z /@
w0 wl w2 ws ™ | ey -
ho har - - - T2
1 il_r"‘z ¥

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.

i~ /&8 _t ,__Jl
| IGINAL u:n-flrw@ Sf/E[L

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

GHARGE NO 8-25-21,31 | DOCUMENT NO. p/7l,/66 |ISSUE 1 DATE 12/16/7L

5.2.2 LOADING CONDITIONS

The msximum operating pressure on the steam side of the tubesheet

is 3600 psi, while that on the salt side is 175 psi. A design con-
dition of 4OOO psi on the steam side with zero pressure on the salt
side was conservatively chosen. The primary stresses shown in Table
1 are due to a L0000 psi steam side pressure.

In order to establish the maximum ranges of primary plus secondary
stress intensities at the various locations of possible high stress,
the thermal transients of Reference 2 were examined. The following

transients, shown on Figure 2 and 3, were determined to be the most
severe areas:

Severest Up Transient: Ramp Change in Load Damage from 100 to 1,0%
in 3 Seconds (Load Scram)

Severest Down Transient: Insertion of Two Safety Rods (Reactor Scram)
The following load combinations were computed:

1 Pressure + reactor scram + steady state

2) Pregsure + load scram + steady state

3) Pressure + reactor scram

l}) Pressure + load scram

With pressure always acting, the maximum primary plus secondary
stress intensity range is established by one of the following:

A) Condition (1) minus condition (2)
B) Condition (1) alone (or condition (2) alone)

' C) Condition (3) alone (or condition (4) alone)

Appendix C-1 includes the calculation for all 3 cases. However,

in Table 1, only the highest case is reported.

APPROVED PAGE g6 OF

TEMPERATURE , F

LT1-9 afeg

FrGUE:. ; QUTLET STEAM TRANSIEN] DUE 7O CHANGE OF LOAD FROM
(007 70 4C% IN 3 _SECONDS, ALL CONTROLS OPERATING
1200+
196.2
276 0.1/ ——
" STEAM TRANSIENT
/501
ooy 7€/0=70 26175 2@ 4.5-¢ 3€/2 &6 ' |
A -l S 2 o N P NN b YT
s b =3.5 =30 =S$o =S50 =/50.C
NCTE - HIGHEST /1TETAL TEMP [N LFIGAMENT ODOCRING TRANSIENT
ol 1S /047 F ceccurrED A7 2zo Secs. AFTER
STALT CF TRANSIEN T 1047,
MEAN TEMPERATURE OF T.s. UGAME—NTJ.H‘/
013.9 fi‘r‘_’ St -t
1000 - Suil + + - + —— I
o 2o 30 40 0 €o 20 220

TIME , SECONDS

TEMPERATURE , F

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PEACTOR SCRAM
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STEAM TRAMSIENT
¢ - J '
+ ~ i 4
30 50 100 /50 2oo 219

JIME, SECONDS

FWC FORM 172 - |,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATIO

N

CHARGE NO 8-25-2/,31 | DOCUMENT NO. ND/7L,/66

ISSUE 1 DATE 12/16/71,

5.2.3 SOME DETATLS OF THE FINITE ELEMENT MODEL

Figure 1 shows the finite element model of the tubesheet-header

-assembly. The model was used for both pressure and thermal analysis.

The average temperature Teésponse of the perforated region of the

tubesheet wasg determined from g separat

Figures 2 and 3, respectively. The elemen
of the complete finite element model (Fi

have the above temperature-time histori
constants were used (as determined frog

BY

ts in the perforated region
gure 1) were "forced" to

es. Equivalent elastic

Section III of the ASME Code)
in the perforated region of the tubesghee .

APPROVED

| Page . __

or

,/
;, /:‘2&.5 ) ,E
e SN ‘l
K , | r—-‘
: Cir |
/ (lv' |
o 2 on
> ST
- R sl
/ | /
bt A % | l
7 e /) - ¢”(‘c’(“ St . J
g Ch=87 wern /Axue-)
- N e R AL
0 8 INLET SALT TEMP, = 11507 r~
s \\ C CrERATING  Cenp :
A\ i

ny ?
STEAN TEMP = foca = @ CLER, cone

| FIGURE 4
FILM COEFFICIENTS

Page 5-20

MSBA MEROEA-TUBESHEET ASSEM3LY
USER:  SHERMAN [P DEPT. 156

Ao soRr
I

f HEADER

{ 09893

FIGURE 5 | st

DESIGN PRESSURE _STRESS INTENSITY

! 1% 2
! /7’400 PS,-'/I o 1 m =

28,060 PSI.

(MAX.) \

« A L S -

- 3 v L1 L

@.01¢ IR IwTEvElTY
T e,

Page 5-21
MSBA REACER-TUBESHEET RSSEMBLY

FIGURKE &

¢ HEADE R

!

DESIGN. PRESSURE STRESS |

INTENSITY CONTOUR

USER: ~ SHERMAN [F- OEFT. 156

noT seRc t—!‘ia

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28,060 PS/.

(MAX.) L et

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=

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T 0.0

ONTOUR LEVILS Re XX _BeS_L050 _000 AN, (18 _2il) =100, J 110
FIX AW AN w0 ANED, OO SNKE_S1N NIV AAZN0 15220, RN AIND, 250, K26l ~e71C. 1250

Page 5-22

FIGURE T

DEFORMED _SHAPE FOR.

DESIGN FPRESSURE

& HEADER

MPCA HERDER-TUSLSHEET ASSEVALY

—

USER:  SrErveN IP CEPT. 156

nECT XA r_—.ia

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Page 5-23

MSE= ZTEmP OUY BT »Se0ER TRANGIENT TEMFFPATURE DISTRIBLTION
USER:  SHERMRN 1P CEPT. 1S6

g .25
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HEADER E?fi;}[ i
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TEMPERATURE _ CONToL iR AT | <L
END OF 220 SEC CF REACTOR

SCRAM TRANSIENT I

: - |

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T W =~ B MSEA STEAM OUTLET r=rRi¥R - - -—

USER:  SHEFMRN IP DEPT. 156
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P 5641 10
STRESS INTENSITY FOR e
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OPERATING FRESSUKRE PLUS se—

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Al et os

|
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38,900 PS/ —,.|

27800 FPS/.

52800 PS).

T (MAx.)
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| £.01¢ I TN TY
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MZ233 STLAZ SUTLET mePLiR
USER: SHERMRN P DEFT. 156

AT AL g‘_"‘la

¥ HEADER FIGURE 1O
STRESS INTENS |TY CONTOURS FOR OPERATING
PRESSURE PLUS REACTOR SCRAM
TRANSIENT
38900 Ps!. -

Sngoc Pol.

K777 e

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£.31¢ I JODDITY
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ORTAA OV Bef). Bedd, Cofd, D33, Frlld. FRif) Celd M6, 1-1N
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T3BA STERM OUTLET HEROER TRANSIENT TEMPERATURE DISTRIBUTION
USER:  SHERWWN TP DEPT. 156

norsns F—°

|
4 HEADER
FIGURE Il |
TEMPERATURE CONTOUR AT
END OF 220 SEC. OF LOAD |
SCRAM TRANSIENT ‘
|
| 1 O4OF —
|
AT = 40°F
3

I QDENT TOesTUN
' The: 19,0

Tl ' OPFOA LEMLS: Reifl  betie, ColOD] D008, C«1DL], Pei0LE, Col02], Me1025, [~tO3)
PAEE, ReiOML, LeiONG, P1OS[. e lOBA

Page 5-27

FIGURE |2

e———
—

STAESS INTENSITY FOR
CLERATING PRESSURE PLIS
LOALD SCRAM TRANSIENT

I
|
|
|

! /4-/ (OO PS/. *"'mm. -

|

HEADER

1SBR STERM OUTLET HEHRR

156

SIEFFN 1P OEPT.

not SXArc r—__:ln

USER:

1

NN WS

oo 1 2262211 110

SER LTS T

J ISR

/8.88Q £S5/

(MAX.)
= - ”oow s = 5
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n 19 - -
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n bl R 2
= w » .x
w . » o
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D01 AT INTOBLTY
T, 150,00

PSAF STIepr TOTLET MEACEF
USER:  SrEFRm P CEPT. 156

noT XRe L—‘—'is

FIGUKRE |73 |
STRESS INTENSITY CONTOURS
OPERATING FRESSURE PLUS
_LOAD SCRAM TRANSIENT

/4,100 Ps|. —m

=
e
=

{/
e

;_J’——/

i /

‘/
S, N &0 i

&.817 Py JENTY
Tie, 15,

CMIBF LT S ailE 7R, Rt T T BT i T, e

e H - .;\‘-.“ -u’s ‘1 b b ¢ ¢ | o ¢ ’ ¢ o ¢ ¢ ’ ¢ '.-.:\.os?,-o.'u——!o-"l.nz .Mfizmws:g9 .,. .

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2/,31 | DOCUMENT NO. ND/7L4/66 |ISSUE ~ 1 DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |4

5.2.4 SIMPLIFIED INELASTIC ANALYSIS

Inelastic strains were calculated by Bree's simplified method (Re-
ference 3 ). In the Bree Analysis, it is assumed that there are a
total of 200 severe thermal cycles (sum of load scrams and reactor
scrams). With a 30-year design life, the time per cycle is 262,800/
200 = 1314 hours.

The derivation of the creep law used is given on the next page. The
Bree Analysis of Reference 3 has been computerized by Foster Wheeler.
The following results were obtained (Reference: Computer Output
NHJLOST) :

Primary Secondary ,
Location Stress,ksi Stress,ksi Total Strain/Cycle,% Total Strain

-1 g.g 23.36 10~ 0.02%
2-2 12.16 14,22 10~k 0.02%
3-3 6.15 149.81 5 x 10~k O 0.1%
L-L 6.55 20. 31 0 0
5-5 7.15 60.8L 8 x 107k 0.16%
6-6 8.47  31.32 2 x 10~k  0.0L%
7-7 12.97 37.09 8 x 107k 0.16%

BY APPROVED , PAGE 5_30 OF

S BTN Sl CREER

e e e —

CONSTANTS fFoR

RELERNTE Frg. 1o~ @F "
CHL M N - IKRON AL

SECR FELIR = MECTR &

P
i
N
X
N

LOHERE © D, =

creer vare (AhAeceonrs <
FOR SECONDARY CREES EFF

J = S7eary - SrA7& APPLIED

FPRIMAR Y LOAD |
A 7 = EREE,L CONSTANTS ©,
YATERIAL AT TEMPERATL

LATA FOR NICKEL - MOLY BLENL/ 1T —
y, INOA-8& *

A4 (7.9 x0°)" O

"

A Cadiutsro® )7 @)

FCX _TEMI = 120 “F

72(_ & B fer = A (“/f 'V)N

,‘ o Lo -5 7 - @ V4
‘//C:‘r,;, Xy = fi(pX/O)

M
Iy
>
Y
o>

i

22
I

. _‘? n
M8 G Wap " w. A (X267
. C“ —5 .’ y '7
o= T e w A L2 xdE7)

/0 ¥ s (:;_5")71

7,
(".((/ & Z.2
PP e 3
G ¥ /0 -2
A = =L o R2.98 X lg T2
2 x8¥) " =

Page 5-32

FWC FORM 172 - 4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2431

DOCUMENT NO. wD/7L,/66

ISSUE

1

DATE 12/16/7L

5.2.5 FATIGUE ANALYSIS AND CREEP FATIGUE INTERACTION

A preliminary fatigue analysis was made with the aid of Code Case

1331-L.

obtained at 1100°F.

Location

1-1
22
3-3
L-L
5-5
6-6
I

S Ksi
33.62
22.76
49.81
20. 31
60.8)
31.32
37.09

Ref. Page

C-1-5
Cwl~9
C-1-10
C-1-11
C-1-13
C-1-15

C~1-16

The following results were obtained.

Salt

16.81
11.38
2L.91
10.16
30.42
15.66
18.55

Allowable cycles were

N all
15000
.c"m
2,500
e
1,300
20,000

10,000

With only 200 severe thermal cycles, maximum fatigue damage is 200/

1300 = 0.15 which occurs at location 5-5.

the allowable value of 1.0.

This value is well below

BY

APPROVED

PAGE _; OF

FWC FORM 172 - L

S COLUMN TH¥DTICATE WHERE CHANGES HAVE BEEN MADE

NOTATIONS IN Tt

Tk e e o rem—_a

FOSTER WHEELER CORPORATION

T U

— -

——r

CREEP FATIGUE INTERACTION

Creep fatigue interaction was performed by the method described in
the nozzle pertion of this report. Location 7-7 proved to be the

most severely loaded section, and only that section was analyzed.

The section was idealized as a cylinder with 8" inside radius and

L" thickness. The following loading was applied:

PR/t = p x 8/L = 12970 x 3600/L000 (refer to Appendix C-1 for source
of 12,970)

solving for p, p= 5836 psi

2218l AT = {156
2 (1-v) -8362 } i-59

, o= 7.43 x 107

=
£
o
H
I

26.7 x lO6

2
=
®
|_$
®
t=
]

PROGRAM RESULTS

The same type of loading histogram was used for the tubesheet as
was used for the inlet nozzle. There are a total of 50 cycles,
each cycle consisting of 3 load scrams and 1 reactor scram. The
program was run for five cycles and the strain for each remaining
cycle was estimated by taking the difference in strains between
the fourth and fifth cycles.

Maximum total strength = 0.432 + 45 (0.01) = 0.88%

Creep damage = 0.0135 + L5 (.0026) = 0.13

Maximum strain range = 0.381% .
(Maximum temp = 1000 F)

Fatigue damage = 50/500 = 0.1

Total creep fatigue damage = 0.23<1.

Ref: Computer Run EBHJLO&C

BY

APPROVED PAGE o_.,  OF

FWC FORM 172 - I

o By 2 e T

COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —- —-—

~
-’

NOTATIONS IN THIS

FOSTER WHEELER CORPORATION

i

CHARGE NO 8-25-2),31 | DOCUMENT NO. Np/7), /66 ISSUE 1 DATE 12/16/7),

.3 STRESS ANALYSIS OF THE MSBR SHELL

5.3.1 INTRODUCTICN AND SUMMARY

The MSBR shell was analyzed for both pressure and thermal transient
stresses. Three locations of possible high stresses were selected,
and are shown in Figure 1. |

At locations A and B, pressure stresses were determined by hand
calculation. Temperature distributions and thermal stresses .
were determined using the finite element models shown in Figure
2. At location C, the finite element model shown in Figure 3 was
used to determine pressure stresses, temperature distributions,
and thermal stresses.

Table 1 gives a summary of the stresses obtained at the three
locations. Primary stresses are due to a design pressure of

300 psi or an operating pressure of 235 psi, while primary plus
secondary stress intensity ranges are due to an operating pressure
of 235 psi, the severest up thermal transient, and the severest
down thermal transient. Examination of Reference 1 indicates

that a "Ramp Change in Load from 100% to LO% in 3 Seconds"

(Load Scram) is the most severe up transient, while "Insertion

of Two Safety Rods™ (Reactors Scram) is the most severe down
transient (Figure L).

The stresses in Table 1 are categorized in accordance with the
rules of ASME Code Case 1331-5. Although the Code Case does
not place a stress limit on the primary plus secondary stress
intensity range, the 35, 1limit was calculated for reference
purposes.

The effect of the tube thermal expansion loads on the shell was
calculated by hand and found to be negligible (<200 psi).

BY

APPROVED PAGE ¢_35 OF

FWC FORM 172 -

NOTATIONS IN THIS COLUMN INDICATE WHERE C

{GES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO g_pe_) 3

DOCUMENT

NO ND/7L/66

DATE 12/16/7),

TABLE I ~ STRESS SUMMARY - MSBR SHELL

See Appendix C-

Stress Calculated
Location Category Stress Intensity Allowable

g | a P 8.2 55=9.5

E PI+Pg 8.2 1.55,=14.25

% B Py 0.3 S6=9.5 _

© PL+Pp 0.3 1.53‘7‘40: 25

&

b c  Pp 8.2 S0=9.5

A P[+P, 10.3 1.55,=1%4.25

%

Mo oA P, 6.4 Spt=7.5

=) PL+Pp, £.% KS+=7.8

£ P +P,+) 38.9 35,=67.5

()

- PL+Pp, .2 KS,=9.3

é‘ P +P,+Q i1 35=67.5

3| ¢ Py 6.4 Spp=7.5
P1FP, ] ;5.0 350=67.5

NOTE: All allowables are at 1150°F

2 for material Properties and allowables.

See Section 5.3.6 for creep fatigué interaction.

BY

APPROVED

PAGE 5-36 OF

Ave. ) = 307
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FIGURE 72 - flmyTe clecMENT MoDELL

Page 5-38

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.

B e e QATE S-caiine et R L L A O WGP, .. S SR SHEETWNOE ... OF:-avae
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///: "f

(f"/fQ' '~ 7 A 7 S Af=- 7 y Y ¥ N g o= o -
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& wnir

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STRESS WIENSITY = 38 990 —//0

NCTE ' MAX, KEACTCE ScRAM STRESS
AFTEN STANT cfF SCRAM |
MAX. LOAD - SCRAM STRESS
AFTER Srakr CF SCRAM .,

———— .

/7 -5-5'0-//? z.

DES/GN FLeSSURE STRESS _J
=3 300 A r7. 78 '
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-
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L SANSIENT TEMFERATIIKE STRESSES
J < Uz Q—T GE& STIRESS IN
Ne—
“SMKREAC 1tR: SCrRAM e (S 80C 3/, 690 -20 3/ &/¢
“LOAD  ScrA A = 20 —3 /&0 -/ 30 /o 7 0%
O +@ + OPER, PLESS (to /7,560 3& 770 B¢ 3§ £&o

- #~ —30)'_—_] %

38,880 ps/.

OCCLRS AT ENO OF £5 sec,

Ar.=/7/)Fx
CCCURS AT END OF /18 S&C,

AT =38 °F.
Page 5-41

< SHELL STRESSES A7 locAar/ionN "B

LPES/IGN  PRESSURE STRESS

Ng =-3co PS/. (AT rye SURFACES )

TKANS/IENT TEMPERA TURE STRESSES

I8y J& 07 Urz STRESS |
L AtACTOR SCRAM —+ /55 2282 > 2278
G LOAD SCRAM -3 -1267 ~ 1837 -/ /833
L@+ OF PRESSUKE ¥ 28 0.5 1 /_7 3 ¢// <

2 2 -+
S, 93 = (2508 .+ 7)1 [(2&62"7) *(3)]_&

\J'

/%ol £./)397

—
=

= 28085, 7

Je .G = Gl

INTENSITY = ¢//7 - 7 =4 /l2 PS/.

STKESS

NOTE : MAX. KELTI1EN SCRAM STRESS €Ceures AT ENDO O©OF @7 S&C,
AFTER STAKT C€F SCrApr) . A7 = Vi
LCAD SCRAM STIPESS ECCURS Ar END oOF 7.5 SEC

MAax.
AT = /12°F

AFTER STAKT CF ScrRAM,
Page 5-L,2
FOSTER WHEELER CORPORATION
o YT de

110 SOUTH ORANGE AVE., LIVINGSTON, N,

BY. P DATE. LML sUBJECT e W Syl 5 o~ . SHEET NO...... ____OF._._.
CHKD. BY _______ RATE.. 2o Ui e SN i s S S i X JOB RO i h o ol bl
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16390

Page 5_.)43 FORM (285).47
FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.

) AL S SR e S RDATE e . oL, SUBJECT. SHEET NO..--‘.O}.--..OF

Pt TRZ Kl <L SE8U8R ki @ \ISo'F
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Fo=98.97< kSg=|.26»8-0.25x¢.43 ='8.39

FORM (205).47

Page 5al

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FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.
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CHKD. BY .______ DATE. cduetonn. | el fo O AU | o e cnnstne g O o o S JOB NOG e satcn-dikae
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FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON, I

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FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON, N.

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Page 5—h8 FORM (28%).47

FWC FORM 172 - L

T
s

NOTATIONS IN THIS COLUMN IND

CATE WHFERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORFORATION

CHARGE NO 8-25-2431

ISSUE 1

DOCUMENT NO. ND/7L/66 DATE 12/16/7L

e

s

5.3.5  SIMPLIFIED INELASTIC ANALYSIS
Inelastic strains were calculated by Bree's simplified method
(Reference 3). In the Bree analysis, it is assumed that there
are a total of 200 severe thermal cycles (sum of load scrams
and reactor scrams). With a 30-year design life, the time per
cycle is 262,800/200 = 131l hours.
The derivation of the creeplaw used in the Bree analysis is
given in the tubesheet section of this report.
The Bree analysis of Reference 3 has been computerized. The
following results were obtained: '
Location  Primary Stress Secondary Stress Strain/Cyclef Total Strain %
A 8200 38880 5x107L 0.1<1%
B 300 Lo 0 0
C 8070 55000 8x101 0.16<1%
Ref.: Computer Run NHJLOSJ
BY APPROVED

| PaGE 519 OF

FWC TORM 172 - 4

NOTATTIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

-

s e ——

FOSTER WHEELFR CORPORATION

CHARGE NO 8-25-2),31 | DOCUMENT NO. ND/7../66 ISSUE 1 DATE 12/16/7L

5.3.6 FATIGUE ANALYSIS AND CREEP FATIGUE INTERACTION

A preliminary fatigue analysis was made using the fatigue curves
from Code Case 1331-4. The following results were obtained:

Location S, Ksi Salt Na11
A 38.9 19.5 6000
B Ll 21 e
C 54.3 27.2 1500

With only 200 cycles of significatn stress, maximum fatigue damage
is 200/1500 = 0.13.

CREEP FATIGUE INTERACTION

Creep fatigue interaction was performed by the method described in
the nozzle section of this report. The load histogram is also
shown in that section. The shell shroud juncture had the highest
loadings, which are given below.

pr = 8.07; p = 0.75 x 8.07 = 0.31 Ksi

t 20

EasT = 40,8 AT = F (down transient)

2 (1-v) =13.9 F (up transient)

where E = 27 hS X 12 = 7.175 x 10‘2 (down transient)
E = x 10 = 7.8l x 10 ~ ( up transient)

In addition, aal of 20°F was applied to simulate the pressure bending
PROGRAM RESULTS gtress.

There are a total of 50 cycles, each cycle consisting of 3 load
scrams and 1 reactor scram. The program was run for five cycles

and the strain for each of the remaining cycles was estimated by
taking the difference in strains between the fourth and fifty cycles.

Maximum total strain = 0.487 + 45 (0.041) = 2.33%x

Creep damage = 0.05L7 + L5 (0.0120) = 0.595

 Maximum strain range = 0.33L7

BY

APPROVED | | PAGE c_zy  OF

FWC FORM 172 - |

NOTATICNS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —— ———————

A e i 3 i ki 2 st

FOSTER WHEELER CORPORATION

ISSUE ;

DATE 12/16/7L

L i ey i et P

Fatigue damage = 50/1000 = 0.05

Total creep fatigue damage = 0.645<1

Ref.: Computer Run EBHJLO1C

at 1150 F

% Some slight thickening in the area of the shell shroud juncture
may be necessary to reduce the total strain below the allowable

value of 2%.

BY

APPROVED

PAGE

| g iy | OF

FWC FORM 172 - §

NOTATIONS IN THIS COLUMN

— e

NDICATE WHERE CHANGES HAVE BEEN MADE

I

FOSTER WHEELER GORFORATTON

!

CHARGE NO §_»c_5)3) | DOCUMENT NO: wn/7u/66 | TSSUE 1 DATE 12 116/7),

o s
———t s g

L p—
-

T . AT - S Al 5. 1=

o~

5.L STRESS ANALYSIS OF THE MSBR TURES

5.L.1 INTRODUCTION AND SUMMARY

The MSBR tubes, constructed of Hastelloy ® material, were analyzed
for pressure, radial temperature gradient, and restrained thermal
expansion stresses. The temperature difference between the salt
and steam is greatest during the steady state condition (rather
than the transient condition) and, therefore, the highest. radia]
gradient thermal stresses occur during steady state.

Restrained thermal eXpansion stresses occur in the tube because
of the fact that the average shell temperature is different than
the average tube temperature. These stresses were calculated
with the aid of the computer program STRUDL.

Table 1 gives a summary of the results, Stresses were catagorized
in accordance with the rules of Code Case 1331-5. Design stresses
are due to a 3600 psi design pressure and restrained thermal
expansion. Operating stresses are due to a 3,25 psSl pressure
differential, radial temperature gradients during steady state,
and restrained thermal expansion. Although there is no limit

on primary plus secondary stress intensity range in the Code Case,
the BSm limit was calculated for reference purposes.

Natural frequiencies in the bend area of the tubes, where flow
induced vibrations are possible, were calculated. Tt was
found that natural frequencies were at least 50% higher than
the vortex shedding frequencies, which is satisfactory.

BY

APPROVED | PAGE ¢ oo oF

FWC FORM 172 - |}

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-21,31 | DOCUMENT NO. yp/7}, /66 ISSUE 7 DATE 15 /16 /7),

TABLE 1 - SUMMARY OF RESULTS - MSBR TUBES

Stress Calculated
Condition Category Value (ksi) Allowable (ksi)
Design Pm 11.67 S =11.8 @1120°Fx
PL+P 11.67 ~ 1.58,=17.7
Operating Pm 1. Spt=11.8
PL+Pb 11.4 Kst=12.2
(P +Pp+Q)g | 29.3 38 =67.5

*Maximum average through the wall tube temperature.

See Section 5.4.6 for creep fatigue interaction.

See Appendix Cw2 for material properties and allowables.

BY

APPROVED | | PAGE ¢ cy oOF

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.!
SHEET NO..__._._.. O ovisrs

BY st e DATE oo o v i SUBJECT.
CHKD, BY v s PATE Sacasidl | Anses v s Wit i s m s 5 S w5 s o S ol JOB NOicassovssenenasesnns
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L & - a*-
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FORM (285).47

Page 5-54
FOSTER WHEELER CORPORATION

__________ L IR Y s A
_______ DATE. &2 s T S S e M S S B R SS e e o e NS M B b e i i
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” i ‘_, (.,‘1' | = x\
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‘ el 00(,3:‘) F
“')".{ . "\)4?“,,
- = 4.0 ‘SP T 05+ 38 = 16.6 N30
P MengnkAnge PRESSURE STRESSES
T~ =0.40N -
) T % 3
- .04 ¥

Page 5-55

110 SOUTH ORANGE AVE., LIVINGSTON, N.

SHEET NO

FORM (285).47

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO g_p5_p),31 | DOCUMENT NO. ypy /) /g  |ISSUE 4 DATE 15 /16/7)

POy PHERMAT, EXEANGTON SR

Because of the temperature difference between the tubes and shell,
thermal expansion stresses will exist in the tubes. Below are
tabulated tube and she¢ll temperatures.

Tube - Shell | AT

Inlet Temperature 1083 1150 67
Outlet Temperature 756 850 oL
Average AT = 80.5

A STRUDL model of a typical tube was prepared (Figure 1). Applying an
average temperature difference of 80.5 F between the tubes and shell,
a maximum stress (at joint 29)of 6L00 psi was obtained.

Reference: FWC Computer Run EBHLEVYg, July 8, 197h.

BY

APPROVED | PAGE _ ., OF

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.

) A RS DATE .. ocoee. TR | oy e A L U A SHEET NO..._).. L SO s
CHKD. BY o _____ 5 TR v, PR, S JRm S N S e s L I I JOB. MO oo onsesmbomncass
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F'(7U'\€‘ g—I}\\ D M)DEL FOR TUBE FORM (203).47

THERMAL & xPANSIoN SSTRESSES
Page 5-57

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON, N.|
LR SR DATE cciivasns BUBDEGY.S s e a n AR S b M Ao SHEET NO..ooonono. OF: 5.

-

K DESIGHI CONDTT Lo
0r==0.417 3,600 = - |5ksi
- Max <1 = |[.67]
= 252« : =38 A : o
B2 \-048+3 (0p +€-4 = |6.1T
DVESATING  LonNDITioN Pz (3.6- 0.\‘\5Y(.043+ 5417 J46.4 = ||«
k\ \-:.." *,v‘\ tJ" \ C“’\J‘—r\"'\’ .hwvl = 3/‘4‘ 2)0 7(0(( = ‘.q7¥ 105 Lrs.
J,._,\ Ban ) ™ vy )\di‘r'qjk ’Tq\Q \_,\“ -T&"‘P("‘:TV‘“Q - ‘UjgoFo
Bor dhd € Son ® Il 80k ot WASF any 2le hes
. I , 5
e F‘— \Jo—T ‘v‘s\ cordiTicn Jing = 4> Ho yeary = 0.4¢x 0 ‘\.—( 5
¥ A e el Tule wall FapereTore = WITF:
& 3 « _ <
b, = LV 2Bae™ JauKsi N WZo F asd 9.6€% |0 ket
~he S h'&"\' -S"o;“‘,,.\
- 'Y
,'f. X = JQ).,N‘ D.6C= 10 -
T : i I'= ot "2 o °.9%.< ‘

FORM (2085).47

Page 5-58

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.

B Y e n s mmb s DATE caceeee . BUBIECT. 2 o via ¢l it 5 is'e ey ol ien s s sy e ke SHEET NO 7 OF

D I

STEADY STATE TUBE TEMPERATURE DISTRIBUTIoN

v ]
/ 0) '®
hi)Ti —/ Tos
¢
% _I\/_J
Q-2 5.
0.315 .
7~
¥
.T“:,: 1710 (STEAM OUTLET) k= 1.0 @ (080°F
TFEE To (STSe™ WL T) ‘<" J).08 @ 775'0‘:—

& . ;
- = XS$a ( Sk:'b‘\_)r"‘\ o~\.TL(—-T>
5= WSe ( Sobirumy INLET )

he= 4941 ( STEAM wieT - Sebium ou TS
hiz 1439 ( sTEAM 0/TLET- SobIvM INLET)

hs= 4772 (STEPM InLeT = Son (™ ObTLET)
ho= 1134  ( STEAM ouTiET - So)\yM nm.e-r)

- % =
-T- (L. - ‘LWk/‘y(fr’
.- A
g B

FORM (2885).47
o/ r

/__S’ S{ | l'\ N i
l, - T = *} [mk;r; T AWK 3 xwk,r,fik

FOSTER WHEELER CORPORATION

110 SOUTH ORANGE AVE., LIVINGSTON, N.,
Y. A~ . BATE B BT s e i s T L Ay SHEET NO.. 5.\ . OF......
CHKD. BY —___ L AU RO S S A SN T LYV o TESEY oo N e NN, SO NS B oS as ol
g ~ e ‘V\ r‘,/,_‘ ' =)
;z-,-t(:’T B PRl e el R "L\oro’k
RN T R O e
PR Tl S e R et o B
— I 2 5 L\,, ™ _E'_ L\b o =)
[ 2 M= AT AT 1['+ k \hf;"'k-rgl

T.-T;‘[T’-TL%K[H',\‘( e ¥ - -\

< G [y, o o kofl lnr"/r,;

Foe STEr vy \mlG T = Sedyumy ouTLE T

-1
e 7 4§4|u0‘.5/2 g 454 1
‘ 4
Te =510k
172 %0.315 /12 3 972 -
T ® B epsella s et EgeE |
T2 114 F T s 756F '
J.o5=12 - 9.0§=I1Z -2
= 1§55 { - 5 4&',,‘_& P AR 972 %0375 |nl-§
T- o= 7CF
6% oo (T Te ) = FFREI—(90) = 10,¢00 s

Fon $Terr v leg- Sobiur) INGET

o a9 w0 25712 1439
/\C:|o\u‘-|-l?.-\_|'§' \n\-g -1

TS 3 34

T2ty B

— [ ~‘o.’§?§/i: % (134 .

|Q: “ ? ‘4)[‘+ T \hl.g"' "2’ ‘4.30) FORM (285).47
= = Page 5-60

.= o TP W [7, 4= 1

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,,

- ), (R, SRR s DATE e . SUBJECT

.o=12 \l.ox |2

3}
1439 0.25-In1.g 134%0375> |nlS

T:——T\“— \'-\a_, [ I +

€« 2 b8 2¢.3»7.6 ) :
et b D e (€2)= 7420 psi

LIVINGSTON, N.J

.......................................... SHEET No.-__é-l---OF.-----.

-\

FORM (285).47

Page 5-61
FWC FORM 172 - | -

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

e # A ey . S - 3

5.3.4 FLOW INDUCED VIBRATION

There is a possibility of flow induced vibration only in the bend
area of the tubes. The direction of possible vibration is shown

in Figure 2. Supports have been provided as shown, to increase

the natural frequency to 50% above the vortex shedding, and minimize
the possibility of vibration.

The natural freguency is calculated below.

BY APPROVED [ PAGE _ . oF

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE.,, LIVINGSTON, N..
SHEET NO.occoue-

S

BY e casasannes soa DATE - cccncass; SUBJECT
CHEKD: BY cauesxa DA TE: nencois, | ol et aOenC ansE s e das ot aem e e e O h o a E X JOB NOi s ncmesonpvamespavs
g - N’K‘st\\* ‘\ 1w
- o\ Dy g
2} 5
7 ‘\ W5 S8, \\)"})
t '|\‘ =
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(v Ty yioaesd
v Mot L or I3 i uu\-.\’ L
T: s u\) t) .\d)
X
x—1
i:_l\ :'_ T-/I\fa%‘ A 5 g\/r‘(. .T'(-) [ B(’.'é p.rcm
/ “
o T AP
| el 1 — /hgt =1 v 0.0 24¢C, 3FC. 4 - -~ D
\‘\- \J AN N :.)77? - L\,‘)‘J‘}"Q
| - ~ —i/ - ¢
L \ by BMGT =) & 7 2 Q0.4
\A + Siri Pl 4"‘ 2 T)Ef‘*f\
- -"‘ - -
L. - \‘<1 = s 7 C.” -
's &) - Flas 384 0y
( -
. e - > ( - (L_‘b Wi r‘
NI n X
= B {38+ 7B )22 C
e Vo T » 0 MSpp e Freaa€r o = 4y g
- 4 U ] - S F v Pk ' < M\.rE ‘.r-’/"lxj EJ ‘7\' h . : b ;_.:‘\ FORM (285).47
Toatabe Mo T SHRDOD e FrSQuEreY, apaen  Flevs
‘U}J .4* o ) Be M"\}|r"’lL 3 Page 5-6

FWC FORM 172 - |}

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARQGE N08-25-2h31 DOCUMENT NO°ND/7h/66 | ISSUE 1 DATE]E/lé/Th

5.3.6

5.3.5 SIMPLFIED INELASTIC ANALYSIS

Inelastic strains were calculated by Bree's simplified method
(see Chapter VI of the IMFBR Piping Design Guide). 1In the Bree
analysis, it is assumed that there are a total of 200 severe
thermal cycles (sum of load scrams and reactor scrams). With

a 30-year design life, the time per cycle is 262,800/200 = 131l
hours. :

The derivation of the creep law used in the Bree analysis is
given in the tubesheet report.

The Bree analysis has been computerized by Foster Wheeler, and
the following results were obtained.

P PLP.* Qg & /cycle % €7

11.1 ksi 11.4+10.6 = 22.0 3x10~L 0.06< 194

Ref: Computer Run NHJLOTP
In computer run, input data is conservative.

FATIGUE ANALYSTS

A preliminary fatigue analysis was made with the aid of Code Case
1331-4. With a stress range of 22,0Ksi, or an alternating stress
of 11.0 Ksi, the Code Case gives an infinite number of allowable
cycles. Thus, fatigue should be no problem.

BY

APPROVED [ PAGE 5 g OF

FWC FORM 172 - |,

NOTATIONS IN THI

N e o L TR St I B 17, s e ke e i S - B 3 AP B e bk

S COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —— ——

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2,31 DOCUMENT NO. ND/7),/66 ISSUE 1

ATE 10/16/7)

WL i Lo —— oo #

At B bt . A eyt

YR £ Senr A

CREEP FATIGUE INTERACTION

Creep fatigue interaction was performed by the pethod described in
the nozzle portion of the report. An internal pressure differential
of 3L25 psi was applied for the 30~year 1ife of the unit. The
steady state radial temperature gradient is 52°F, In addition, a

28 F radialAT was applied to simulate the tube thermal expansion
stresses. During the transient conditions, the radial temperature
gradients in the tubes are milder.

PROGRAM RESULTS

Maximum total strain - 0.65%

Creep damage = 0.095

i

0.258%

Maximum strain range
Fatigue damage = 0.05 (Maximum temperature = 1100 F)
Total creep fatigue damage = 0.15<1.0

Ref: Computer run EBHJLO1C

BY

| APPROVED PAGE

c_gz OF

FWC FORM 172 - 4

NGES HAVE BEEN MADE -

NOTATICNS IN THIS COLUMN INDICATE WHERE Ci

Samb L ————— . ia m— R =~ o,

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2),31 | DOCUMENT NO. yp/7L/66 |ISSUE 1

DATE 12/16/74

O AL i Bl et v b R SR 1 ¢ oo i

5.5 TUBE RUPTURE ANALYSES

Tube rupture effect analyses were performed by the Gulf General

Atomics Company. Detailed descriptions of the analyses are contained
in Reference(8)which is ineluded in its original form in Appendix
C-3 of this report.

The analyses covered two areas of concern: the effect of a tube

rupture on the shell cylinder and the adjacent tube. The effect

of a tube rupture on the shell was examined by modeling the shell
in a state of plane strain loaded by internal pressure of varying
profile. Numerical calculations were done by using the finite

- difference code system, PISCES, developed by Physics International

Company. The bilinear stress-strain relation of Hastelloy N at
850°F was used. The results indicate a maximum stress of 42,900
psi which is well above the yield strength (27,600 psi), but only
one-half the ultimate strength (85,500 psi) of the shell material.
Tt also predicted an increase of 3.6" (or 9%) in the diameter of
the shell.

The effect of a tube rupture on the adjacent tube was examined by
modeling the adjacent tube as a lumped-mass continuous beam with

an appropriate elastic-plastic moment-curvature diagram derived from
the stress-strain curve of the tube material (Hastelloy N at 850°).
The loads acting on the beam model consisted of a transverse inertial
force and the resistive force of the molten-salt. The results
indicate that the ultimate strength is not exceeded and, therefore,
the adjacent tube would not rupture. However, a deflection of 2L4.5"
was calculated which was about 4O percent of the length of the

'span between two consecutive supports.

As indicated in Reference (8), the conclusion reached concerning

the shell response due to a tube rupture was conservative in that

the applied pressure was chosen as an upper bound of the data
available on ruptures. The assumption of a plane strain state

was alse conservative in that the pressure load resulting from

the tube rupture was assumed to act along the entire axis of the
shell. However, for a more accurate prediction, much knowledge
concerning the pressure profile is needed experimentally and
analytically. On the other hand, the analysis concerning the

effect of a.tube rupture on the adjacent tube is incomplete.

Due to the extremely large deflection calculated from the beam
model,it is felt that the applied loads have to be reexamined and
large deformation theory should be considered. In the presence

of large deflections, the analysis should also include the interactim
of adjacent tubes, and perhaps the associated problems of instability.

BY

APPROVED [ PAGE ¢_gg OF

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE ——— —-——-——

FOSTER WHEELER CORPORATTON

CHARGE NO 8-25-2,31 | DOCUMENT NO. Np/7,/66 |ISSUE 1 DATE 12/16/7L

LV P

5.6 REFERENCES

1.

Foster Wheeler Computer Program R10l5 by R. E. Nickell, revised
by M. B. Hsu.

Hybrid Computer Simulation of the MSBR by O. W. Burke, ORNL-TM-3767.
LMFBR Piping Design Guide by C. F. Braun & Company.

Bases for Design of MSBR Systems for Temperatures to 1300 F,
ORNL-73-1-23.

Tubesheet computer runs EBMSBR2A (pressure), EBMSBREC (steady state
temperature), EBMSBRCF-1 (load reduction tran51ent), EBMSBRCT-2
(reactor scram transient)

Shell computer runs, EBMSBR2C-1 (Pressure), EBMSBR2C-2 (thermal stress)

Data for Nickel ~ Molybdenum Chromium - Iron Alloy, Iron-8, June 1,
1961.

J. J. Johnson and D. A. Wesley, Tube Rupture Analysis of a Counter
Flow Heat Exchanger, Gulf-Ga-& 1241k, Nov. 20, 1972, Prepared for
Foster Wheeler Energy Corporation.

Nozzle computer runs - EBRS156C (pressure), EBRS156C-1 (temperature
stresses)

BY

APPROVED | PAGE ¢ ¢; OF
FWC FORM 172 - L4

: FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 |ISSUE 1  |paTE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

SECTION 6

HASTELLOY N STEAM CORROSION

GEo&gf V. AMORUSO

Approved by

E. D. Montrone
Assistant Chief Metallurgist

\D.Q.(g@»
"W. R. let¥, Jr.

Chief Metallurgist

BY APPROVED AGE  6-a

- FWC FORM 172 -~ L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT | LIVINGSTON, N. J.
CHARGE NO. g-25-243] | DOCUMENT NO. Np/74/66 | ISSUE 1 DATE 12/16/74
TABLE OF CONTENTS

PAGES

6.0 REVIEW OF HASTELLOY N STEAM CORROSION 6-1
6.1 ABSTRACT 6-1
6.2 SUMMARY 6-2
6.3 INTRODUCTION 6-4
6.4 DISCUSSION 6=5
6.4.1 EARLY STUDIES 6-5
6.4.2 OAK RIDGE NATIONAL LABORATORY STUDIES 6-11
6.4.2.1 GENERAL CORROSION TESTS 6-11
6.4.2.2 TUBE BURST TESTS 6-15
6.4.2.3 DUPLEX TUBING TESTS 6-18
6.4.3 CRITICAL REVIEW OF STEAM GENERATOR TUBING 6-21

MATERIALS
6.4.3.1 DESIGN REQUIREMENTS AND CRITERIA 6-21
6.4.3.2 CORROSION RESISTANCE TO MOLTEN FLUORIDE 6-23

SALTS
6.4.3.3 CORROSION RESISTANCE TO SUPERCRITICAL STEAM 6-25
6.4.4 CRITICAL REVIEW OF HASTELLOY N AND DUPLEX 6-26

TUBING |
6.4.4.1 HASTELLOY N 6-26
6.4.4.2 DUPLEX TUBING 6-27
6.5 CONCLUSIONS/RECOMMENDATIONS 6-28
6.6 REFERENCES 6-31
6.7 TABLES 1-26 INCLUSIVE 6-43
6.8 FIGURES 1-45 INCLUSIVE 6-71
<y / L a

' (/nstuucel APPROVED PAGE 6-b

BY%%%’)
J

FWC FORM 172 - |,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, Nf J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

6.0

6.1

REVIEW OF HASTELLOY N STEAM CORROSION

ABSTRACT

The literature survey described herein included a review
of high temperature water/steam material investigations
publ%s?ed in Corrosion Abstracts, ASME Transactions,
GEAp (1 , ORNL(2), gyrarc(3) and miscellaneous technical
papers concerning the corrosion behavior of various low
alloy steels, ferritic and austenitic stainless steels
and nickel base alloy subjected to water/steam environ-
ments at high temperatures and pressures.,

Particular attention was focussed on the corrosion
behavior of Hastelloy N in high temperature steam ser-~
vice since the alloy is the preferred material of
construction for the reactor vessel, piping and primary

and secondary heat transfer equipment in the Molten Salt
Breeder Reactor.

The literature search included a limited evaluation of
alternate alloys or combinations of alloys that could
be used for the steam generator tubing in the Molten
Salt Breeder Reactor.

The survey indicates that there are no single commercial
alloys or combinations of tubing alloys other than those
Presently under investigation at the Oak Ridge National
Laboratory that can satisfy the tubing rYequirements for
long term elevated temperature strength, stability,
ductility, creep and corrosion resistance.

The data indicates there are several areas which require
further investigation particularly with respect to the
evaluation of Hastelloy N under constant stress and new
duplex tubing processes used to provide increased corro-
sion protection to tubing surfaces.

(l)General Electric Atomic Power Equipment Department
Vallecitos Atomic Laboratory, San Jose, California
(Studies prepared for U.S. Atomic Energy Commission under
Contract AT(04-3)-189 Project Agreement 13.

(Z)Oak Ridge National Laboratory, Oak Ridge, Tenn. Studies
prepared for U.S. Atomic Energy Commission under Contract
W 7405-eng 26.

(3)Joint Europe Atomic Energy/U.S. Atomic Energy Commis-
sion studies performed under Contract # 089-62~7 RDB.

BY

APPROVED PAGE -1

FWC FORM 172 - |}

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, Nf J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1  IDATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

SUMMARY

The literature survey shows there 1is insufficient data
Presently available to properly assess or predict the
pPerformance of Hastelloy N in high temperature/pressure
Supercritical steam environments under constant and

cyclic loading conditions anticipated in Molten Salt
Breeder Reactor service.

The survey shows that aside from the data accumulated
in ORNL material evaluation programs conducted at two
Steam corrosion test facilities located at TVA Bull
Run Steam Plant in Knoxville, Tenn. and the Bartow
Plant of the Florida Power Corp., little data exists
concerning the corrosion behavior of Hastelloy N in

high temperature water/steam environments.

The published data indicates that in pure(l) deoxygenated‘
(< 1 ppb 02) supercritical Steam environments at tempera-
tures of 1000°~1100°F and pressures of 3500 psi pressure,
unstressed standard and modified Hastelloy N alloys con-
taining titanium exhibit satisfactory general corrosion
resistance,

The average thickness of the oxide layer formed after
15,000 hours exposure varies from .01-.04 mils. The
extremely thin oxide formed is tenacious, and non
spalling containing complex oxides consisting primarily
of NiO2, Mo02, a spinel and small quantities of Cr303.

The alloy exhibits a proneness to the formation of
irregular localized oxidized nodular/blister type
penetrations varying from .4-.8 mil in depth which ORNL
investigators have attributed to possible iron particle
deposition resulting from mass transport from ferritic
tubing in the System.

However, in other short term Studies (2,400 hours) con-
ducted in oxygenated (3-4 PPm) supercritical steam
environments @ 1022°F and 3000 psi, the presence of pits
having depths of 1.5 mils were noted after oxide lavyers
were electrolytically removed from the surfaces of
Hastelloy N test Specimens indicating that larger sized
nodules may have been formed as a result of higher oxygen
contents in the supercritical steam. ‘

(DPure Steam as used in this text isg actually high purity
Steam normally used in commercial once-thru boilers.

BY

APPROVED PAGE 6-2

FWC FORM 172 - L

NOTATICONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

The study also showed that unstressed modified Hastelloy
N compositions containing low molybdenum (12%Z) and iron
(.05%) contents exhibit deep localized intergranular
oxide penetrations up to 10 mils after 10,000 hours
eéxposure in pure deoxygenated supercritical steam environ-
ments. Intergranular penetrations of this type have not
been observed in modified alloys of similar compositions
to which additions of .5%-2.1% Titanium have been added.

The published data'indicates that constantly stressed
Hastelloy N alloys may be Prone to stress .corrosion
cracking in pure and impure supercritical steam environ-
ments.

Some data shows that Hastelloy N in the ground and
annealed condition is relatively immune to stress corro-
sion cracking processes when subjected to impure steam
environments containing small quantities of oxXygen and’
sodium chloride while tube burst tests conducted by
ORNL at the Bull Run Steam Plant and General Electric
at the Vallecitos Atomic Laboratory indicate that
constantly stressed Hastelloy N material in the as
received, and machined condition is susceptible to
Stress corrosion cracking in pure as well as impure
Steam environments.

Evaluation of the published tube burst test data indi-
cates that the erratic behavior may be due in part to
the technique used to apply constant stresses to thinly
machined (.010"-.020") tube wall specimens and failure
to consider the effects of residual stresses induced
during tube cold drawing processes and/or the effects

of macro/micro regidual Stresses superimposed by machin-
ing operations.

Summarily, the data indicates that there is no single
tubing material that can meet all prerequisite require-
ments to guarantee satisfactory material performance
for long term service in molten salt/supercritical
Steam environments.

The use of duplex tubing to provide increased corrosion

Protection to the outside and inside surfaces of the
steam generator tubing appears to be a prudent compro-
mise requiring further consideration and test evalua~-
tion.

BY

APPROVED | PAGE 6-3

FWC FORM 172 - 44

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1 |paTe 12/16/74

6.3

INTRODUCTION

A large amount of the technology incorporated in the
present Molten Salt Breeder Reactor program was derived
in numerous material investigationgof developmental

- nature which were conducted during the Aircraft Nuclear

Propulsion (ANP) and Molten Salt Reactor Experiment
(MSRE) programs at the Oak Ridge National Laboratory 7
(ORNL) in Oak Ridge, Tenn. during the 1950's and 1960's.

The thrust of these programs led to the development of
INOR-8 (Hastelloy N) as the containment material for
molten fluorides at service temperatures up to 1500°F.
This alloy was also designed to meet other requirements
for good oxidation resistance, moderate strength and
ductility at high temperature; long term stability at
Service temperatures; ability to be melted and fabrica-
ted into complex shapes and ability to be joined by
welding and brazing.

The satisfactory performance of the Hastelloy N alloy
in the short term MSRF program led to its preferred
sStatus as the material of construction for the core
vessel, heat-exchangers and the associated piping in
the Molten Salt Breeder Reactor.

However it was recognized that a material with greater
resistance to irradiation effects was required for
those parts of the MSBR that would be subject to long
term neutron irradiation during the 30 year design 1life.

Considerable improvement had been achieved by adding

small amounts of titanium and/or hafnium or niobium to
a slightly altered base material to obtain modified
Hastelloy N of the range of compositions shown in
Table 1.

The stress rupture life and the ductility of the modi-
fied Hastelloy N varies considerably with variations

in treatment and in the amounts of some minor consti-
tuents. In general, studies had shown that irradiation
also decreases the rupture life and ductility of the
modified alloy, but, for irradiation fluences of about
1021 neutrons/cm2 (fast and thermal) at temperatures

to 1380°F, its properties were about equal to the
standard alloy when unirradiated,

BY

APPROVED PAGE 6-4

FWC FORM 172 - |}

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

On this basis, and on the assumption that the reactor
equipment would be made of modified Hastelloy N, the
extensive data on the Properties of the unirradiated
Hastelloy N were used in ORNL design studies for the
reactor equipment. These properties are shown 1in
Table 2.

The maximum allowables Stresses shown in Tables 3 and
4 were established and reviewed by the ASME Boiler and
Pressure Vessel Code Committee and Stress values
approved for use under Case 1315 for Unfired Pressure
Vessels and under Case 1345 for Nuclear Vessels,

Since this alloy was developed exclusively for high
temperature molten salt service it possessed relatively
little other commercial value and was cursorily dis-
missed for steam tubing applications due to its high
cost, relatively unattractive fabrication/drawing pro-
Perties and low chromium content.

Chromium has long been recognized as the principal
elemental constituent required to confer increased
oxidation and corrosion resistance to alloys subjected
to air and steam environments at low and particularly
at elevated temperatures. '

In the Molten Salt Breeder Reactor the Hastelloy N
tubing must provide Simultaneously maximum corrosion
resistance to high temperature coolant molten fluoride
salts (1150°F) on one side and supercritical steam at
temperatures of 1000°F and 3800 psi pressure on the
other.

of Hastelloy N and to a limited extent other possible
alternate steam geénerator tubing materials for use
in the Molten Salt Breeder Reactor.

6.4 DISCUSSION
64,1 EARLY STUDIES
The earliest data involving the corrosion behavior of
Hastelloy N in high temperature water/steam environ-
ments was found in a material investigationl conducted
at the Battelle Memorial Institute at the request of
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behavior served as a model for other investigators in

the International Nickel Co. 1in early 1959. This
investigation was undertaken Principally to evaluate
the cracking behavior of Inconel 600 reported earlier
by Coriou and his coworkers2 at the Center of Nuclear
Studies at Saclay, France in 1959,

In this srydy which involved the use of several other
nickel base alloys for comparative purposes, an
annealed INOR-8 (Hastelloy N) strip specimen was
stressed beyond its yield point by a three point jig
shown in Figure 1 and subjected in a five liter
capacity autoclave to high purity, neutral, degassed
water at a temperature of 662°F for 4000 hours.

Visual examination of the stressed Hastelloy N test
specimen after exposure showed the surfaces to be
frosty but free of any evidence of cracking.

Earlier in 1956, W. K. Boyd and H. A. Pray3'reported the
results of an investigation which was undertaken to
determine the corrosion resistance of twelve alloys (1)
subjected to degassed supercritical water at tempera-
tures ranging from 800°F-1350°F and Pressures to

5000 psi.

While this investigation was concerned primarily with
corrosion properties of the various alloys in the
absence of applied stress, a few tests were conducted
to determine the effects of constant stress{(2) op

two of the stainless steel alloys (316 and 347).

In this study, the constant stress samples consisted
of six inch lengths of 1/2" 0.D. tubing sealed in a
jig (Figure 2) so that the pressure differential at
temperature between the outside and inside of the tube
provided a constant load, By varying the wall thick-
nesses of the test specimens over a 2" gauge length
the desired stress levels were attained.

The concept and tube jig configuration used by Boyd and
Pray to apply constant stress to evaluate material

future materials investigations.

(1)AMS 5616, 410, 302, 309, 310, 347, Armco 17-4PH,
Armco 17-7PH, Allegheny A 286, Inconel X, Hastelloy F ang
Hastelloy X.

(2)Stress was 907 of stress necessary to cause rup-
ture in 1000 hours; 15,000 psi for 316 and 12,000 psi
for 347.

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In a literature survey published in Jan. 1962 by C. N.
Spalaris and coworkers® a review was made of avail-
able data concerning the corrosion behavior of ferritic,
austenitic steels and nickel base alloys in order to
select candidate alloys for use in high temperature
nuclear superheated steam environments.

In this review, a number of alloys were recommended for
further study for nuclear superheat applications. These
alloys included Incoloy, Inconel, modified 300 austeni-
tic series such as 310, 304 L types but with low
nitrogen and carbon, Hastelloy X, AISI 406, and RA 330.
Also recommended for preliminary screening tests because
of attractive high temperature mechanical properties
were Hastelloy N, Ni-o-nel (modified), IN 102, R-20,
Discaloy and 17-14 Cu-Mo.

In July 1963, F. A. Comprelli, and coworkersspublished
the results of the follow up materials investigation
conducted to evaluate the relative performance of
various alloys in actual and simulated superheated
steam environments.

The purpose of the investigation was to select sheath
or cladding materials for use in the design/fabrication
of fuel elements which would be subject in service to
nuclear superheat environments at temperatures of 550°F
to 1350°F; pressures up to 1500 psi; oxygen contents

in steam up to 20 ppm; stiochiometric hydrogen, high
velocity (200 ft/sec. max.) steam, containing moisture
up to 1% with small amounts of solid impurities; stress
levels up to the yield point (often applied cyclically)
and neutron fluxes up to 1022 nvt accumulated total.

In this study, unstressed and constantly stressed tubular
test specimens diagramically shown in Figure 3 were
placed in the CL-1 superheat test facility loop shown

in Figure 4 and exposed to 1050°F superheated steam
containing 20 ppm oxygen and flowing past the test
specimens at a velocity approximately 20 feet per second.

Additional stressed specimens coated with chloride salts
deposited on surfaces were inserted into the Auxiliary
Coupon Section (ACS), a low flow bleed off from the CL-1
steam -supply. The specimens in this loop were exposed
to 1160°F static superheated steam environments contain-

ing levels of oxygen and hydrogen similar to the CL-1
steam supply.

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The Hastelloy N. tubular test specimens containing
applied stresses of 19,300 and 21,850 psia were
exposed in both the CL-1 and ACS corrosion test loops
for up to 3000 hours. The examination after exposure
showed no indications that the applied stresses
exceeded the creep-rupture strength of the alloy.

However metallographic examination showed the constantly
stressed specimens subjected in the CL-1 test loop @ '
1050°F to flowing superheated steam exhibited localized
intergranular cracking extending to a-depth of 4 mls
after 1000 hours exposure and a uniform scale penetra-
tion depth of .16 mils after 3000 hours exposure.

The examination of the constantly stressed specimen
coated with chloride deposits prior to subjection in

the ACS loop @ 1160°F static superheated steam exhibi-
ted complete intergranular penetration after 266 hours
exposure and a uniform scale thickness of .4 mls after
the 266 hour exposure period. Another chloride deposit
coated unstressed specimen in the ACS test loop exhibited
intergranular penetrations to a maximum depth of 2.5

mls after 1119 hours exposure.

The localized intergranular Penetrations of the Hastelloy
N specimens were attributed primarily to a highly pre-
ferential oxidation mechanism associated with the low
chromium content of the alloy rather than the applied
stress. The photomicrographs illustrating the appear-
ance of the intergranular cracking which occurred in

the failed Hastelloy N specimens are shown in Figure 5.

As a result of this behavior, Hastelloy N was eliminated
from further consideration as a material for Superheat
Fuel Sheaths. .

The final selection of candidate alloys was made and
included the following alloys in order of preference:
Incoloy, Inconel (at low stresses) Hastelloy X, Ni-o-nel,
AIST 406 and 310 stainless steel. Other alloys elimin-
ated from further consideration in addition to Hastelloy
N were 2-1/4 Croloy and 5 Croloy Ti due to their
excessive scale buildup in 1050°F steam.

A summary of the various materials investigated and the
tests used to determine their relative performance 1is

shown in Table 5.

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This study established the basis for evaluatin compara-
tive material behavior in subsequent studies®~10 yhich
ultimately led to the selection of Incoloy 800 as the
reference fuel-cladding material for use in superheat
reactor (SHR) systems.

In a paperl! published in 1963, H. J. Pessl of the
General Electric Co. (HAPO) describes the results of
autoclave corrosion testing of Hastelloy N and various
iron and nickel base alloys after 100 days exposure to
oxygenated (3-4 ppm 0y) superheated steam @ 550°C
(1022°F) and 3000 psi pressure.

A graph excerpted from this paper and illustrating the
relative behavior of the various materials subjected

to the test environment is shown in Figure 6. The data
shows that Hastelloy N test specimens exhibited an
average oxide depth of penetration of approximately

+16 mils and were pitted to depths of 1.5 mils.

In Nov. 1965, T. T. Claudson and R. E. Westerman of

the Pacific Northwest Laboratory published12 the results
of another investigation to evaluate the corrosion
resistance of several high temperature alloys including
Hastelloy N for nuclear applications.

The investigation consisted essentially of subjecting:
ten candidate alloys to short term (100-300 hours)
screening tests involving exposure at temperatures of
815°C~-1150°C to low pressure water vapor (15 torrs in
helium) to simulate corrosive environments in high
temperature nuclear applications.

The results of the investigation which are summarized

in Table 6 and graphically illustrated in Figures 7

to 9 show that the Hastelloy N alloy exhibited excellent
short term high temperature corrosion behavior in an
oxidizing environment.

Another paper, published in Nov. 1965 by T. T. Claudson
and H. J. Pessll3 describes the results of an investi-
gation conducted to evaluate various iron and nickel
base alloys including Hastelloy N for medium and high
temperature reactor applications.

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The investigation included extensive autoclave test-
ing of selected alloys in oxygenated and deoxygenated
superheated steam environments at temperatures of
1022°F and pressures of 1000, 3000 and 5000 psi for
2400 hours exposure.

This study established the following tentative group-
ing of alloys with increasing steam oxidation
resistance:

1. Type 304, 316, 406 and 430 SS showed weight gains
from about 50-300 mg/dm? or over .2 mils of penetra-
tion.

2. TFe-25 Cr-3A1-0.6Y, Hastelloy N, PDRL-102 and
Inconel X750 showed weight gains from about 8-75
mg/dm2 or 0.1 to 0.2 mil penetration.

3, Incoloy 800, AISI 446 3S, Hastelloy X280, Hastelloy
R 235 and Fe-24 Cr-5A1 showed weight gains below 15
mg/dm2 or less than 0.1 mil penetration.

The oxygen content and the pressure of the steam
appeared to affect the various alloys in different
ways. While the ferritic stainless steels and the
Fe-Cr-Al base alloys seemed to offer better resis-
tance in oxygenated steam, the austenitiec and
martensitic stainless steelsand the nickel base
superalloys were more resistant to deoxygenated steam.

The weight gain data obtained for the iron and nickel
base alloys in superheated steam at 1022°F and 1000
and 3000 psi pressure are summarized in Table 7 and
graphically plotted in Figure 10 A. Barographs
illustrating the corrosion of various alloys after
100 days exposure in deoxygenated (< 50 ppb) and
oxygenated (3-4 ppm) steam @ 1022°F and 3000 psi are
shown in Figure 11. Additional barographs illustrating
the descaled corrosion penetrations of the various
alloys after 100 days exposure in deoxygenated steam
@ 1022°F and 3000 and 5000 psi pressures are shown in
Figure 12,

A portion of the Hastelloy N to steam corrosion
described above was reviewed and summarized by H. E.
McCoy and J. R. Weir, Jr. in a report14 published in
June 1967. The summary is shown in Table 8.

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6.4,2 0AK RIDGE NATIONAL LABORATORY STUDIES

6.4.2.1

GENERAL CORROSION TESTS

The need for advanced materials technology required in
4 proposed Molten Salt Breeder Reactorl5 led to a
series of material development programs one of which
involved tests to determine the compatibility of
Hastelloy N in supercritical steam.

In this service, the outside surfaces of the steam
generating tubing material contacts the molten coolant
salt while the inside surface contacts supercritical
Steam having a maximum temperature and pressure of
1000°F and 3800 psi respectively. Because of the
higher pressure of the Steam, tube failure results in
Steam being forced into the coolant salt circuit with
deleterious effects arising therefrom.

The ORNL material evaluation programs involving
Hastelloy N and other steam generators alloys are
currently being conducted at two steam corrosion test
facilities located at TVA's Bull Run Steam Plant in
Knoxville, Tenn. and at the Bartow Plant of the
Florida Power Corp. in collaboration with Southern
Nuclear Engineering (SNE) in Dunedin, Florida.

The TVA Bull Run Steam Plant Corrosion Test Loop
Facility was designed primarily to ewaluate the
behavior of. standard and modified Hastelloy N alloys
and a number of other alloys in supercritical steam
environments @ 1000°F and 3,500 psi pressure. The
corrosion facility at Bartow was designed to evaluate
(a) the corrosion behavior of weldments in conven-
tional superheater alloys for advanced steam generator
applications and (b) the general corrosion and stress
corrosion cracking propensities of various steam
generator alloys.

The TVA Bull Run Steam Plant'® is a coal fired prant
with a supercritical steam cycle and a power genera-
tion capability of 980 MW. The facility is located
approximately 18 feet upstream from the turbine. The

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steam 1s extremely clean at this location containing
< 1 ppb O09; < 3 ppb Na; < 5 ppb Cuj; < 15 ppb S105;
and < 6 ppb Fe,

Hydrogen is added to scavenge oXxygen and the Ph is con-
trolled at 9.40-9.45 with ammonia. The electrical
conductivity in the condensed steam is less than

3 x 10°7 ohm™! cem-1

In the test loop, which is schematically illustrated
in Figure 13 and shown photographically in Figure 14,
the steam enters the 4" diameter schedule 160 type

316 stainless steel test chamber at a rate of approxi-
mately 16-17 1lbs/min. flowing longitudinally past the
Sspecimens, through the filter and the flow restricter,
and into the condenser. The steam pPressure is

reduced to approximately 1 psig in the flow restricter.
A thermocouple well was installed in the specimen
holder to monitor temperature and this tube also
supports the sample holder. The Grayloc flange is
removable with the specimen holder. The specimens

are coupons made from sheet material of 0.010",
0.020", 0.035", and 0.060" in thickness.

The corrosion test coupons are 0.5" and 2.0" with
0.1875" holes at each end for mounting on the bolts
in the specimen holder. The working volume of the
facility is about 2" x 2" x 2" and approximately 100
specimens were included in the first loading. This
was subsequently increased to 140. The velocity of
the steam across the specimens is approximately 20
ft/sec and the mass flow rate is about 1000 1b/hr.

Specimens from some commercial heats of standard adir
melted and vacuum melted Hastelloy N and from several
two pound laboratory melts and larger commercial melts
of modified Hastelloy N were included in the first
loading of the facility which went on stream on 8/7/69.
The results of Hastelloy N behavior in the super-
critical steam corrosion tests were systematically
teported by B. McNabb and H. McCoy in ORNL MSR
Semiannual Progress reports,(17-26)

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The authors recently published2’ a summary of the over-
all general corrosion data obtained for various modified
and standard unstressed air melted and vacuum melted
Hastelloy N alloys as well as various low alloy

ferritic and maraging steels; stainless steels; and
nickel base alloys which were subjected to the super-
critical steam at 1000°F and 3500 psi for exposure times
up to 15,000 hours.

The general corrosion data indicates that unstressed air
method and vacuum melted Hastelloy N alloys are compati-
ble with steam and exhibit corrosion rates that vary from
approximately .01 to <.04 mil after 15,000 hours exposurel.

The low corrosion rates are probably due to the low
(< 1 ppb) oxygen content of the superheated steam at
the Bull Run Steam Plant Facility. It was noted that

several specimens of similar heats exposed at 1100°F

for 15,000 hrs. in the Bartow Plant Facility exhibited
slightly higher corrosion rates which were equivalent
to oxide penetration depths of .1 mil.

The previously reported work of Claudson and Pess1l3
conducted @ 1050°F and 3000 psi pressure for 2400
hours showed that the Hastelloy N exhibited rates of
08 mil in deoxygenated (< 50 ppb) steam and .16 mil
in oxygenated (3-4 ppm) steam and pits extending to
depths of 1.5 mils.

The chemical composition and pertinent fabrication
data for all alloys tested is shown in Table 9 while
the weight change data recorded for all alloys exposed
to supercritical steam at temperatures of 1000°F and
pressures of 3,500 psi are shown in Tables 10 to 21
inclusive.

Graphs illustrating (a) the weight changes of several
heats of Hastelloy N and several alloys of Hastelloy N
modified with titanium are shown in Figures 15 A, B;
(b) the corrosion of various modified compositions of
Hastelloy N exposed to supercritical steam @ 1000°F
and 3,500 psi and similar alloys exposed to air at
1000°F are shown in Figures 16 A, B; (c) the effect

of surface finish on the corrosion of Hastelloy N and
the effects of cold work on the corrosion of Hastelloy
N and type 201 stainless steel in steam at 1000 psi and
3,500 psi are shown in Figures 17 and 18 respectively.

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LOW ALLOY FERRITE STEELS

For comparative purposes, the evaluation included
exposure of five low-alloy ferritic steels containing
chromium contents varying from 1.1%Z to 8.7% to super-
critical steam @ 1000°F and 3,500 psi for 14,000
hours. The data which is graphically illustrated in
Figures 19A, B, indicate the maximum and minimum weight
changes for this group of alloys may vary at any given
time by only a factor of 2.

The oxidation behavior of similar alloys in air is
quite different and is graphically illustrated in
Figure 20. The data shows that in an air environment,
approximately 8% chromium is requ1red to prevent
spalling of the oxide.

STAINLESS STEELS/NICKEL BASE ALLOYS

A graph illustrating the general corrosion of the
Croloys (1-97%Cr) and several stainless steel and
nickel base alloys is shown in Figure 21.

METALLOGRAPHIC EXAMINATION

A photomacrograph illustrating the appearance of the
surface of a typical Hastelloy N specimen exposed to
supercritical steam for 4,000 hours is shown in
Figure 22. A photomicrograph illustrating the
chemical composition of surface and matrix consti-
tuents of a standard Hastelloy N alloy exposed for
10,000 hours to supercritical steam is shown in
Figure 23A. The high iron content detected in oxide
nodules was attributed to iron mass transport from
low-alloy ferritic steel piping in the steam unit.
Composite photomicrographs shown in Figure 23B
illustrate the cross sectional appearances of nodule
blisters and intergranular penetrations observed in
several Hastelloy N alloys. It will be noted that
deep intergranular penetration occurs in a modified
Hastelloy N alloy having low molybdenum (12%) and
iron (.057%) content while no penetrations exist in a
similar alloy having 2.17% Titanium added.

The investigators indicate that the addition of con-
centrations as low as 0.57 Titanium prevent the deep
intergranular penetrations observed in Hastelloy N
alloys containing low molybdenum (12%) and iron (.05%)
contents.

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6.4.2.2

TUBE BURST TESTS

The earlier work of Comprelli and coworkers5 indicated
that Hastelloy N exhibited a Propensity to intergranular
cracking when exposed to superheated steam environments
in the constantly stressed condition below the yield
point. In addition the results of preliminary studies
conducted by J. Hammond at the Bartow plant indicated
the alloy appeared to be susceptible to stress corrosion
cracking when subjected to impure oxygenated steam |
environments containing small quantities of sodium
chloride.

Since steam generator tubing is constantly stressed during

unit operation, knowledge of Hastelloy N tubing material
behavior in the stressed condition is essential.

In early 1971, after 6000 hours of operation, the

sample chamber of the Bull Run Steam Plant Corrosion
Test Loop was removed for modification to facilitate
tests to observe the behavior of constantly stressed
Hastelloy N specimens subjected to supercritical steam
environments at temperatures of 1000°F and pressures of
3,500 psi. A schematic of the modified specimen chamber
is shown in Figure 24.

The constant stress tubular burst specimens were double-
walled with an annulus between the walls into which the
inner tube would burst during the test. The annulus

was connected to the steam condenser by a capillary tube
and a thermocouple attached to the capillary to sense
failure by an increase in temperature when steam is
introduced into the annulus as a result of specimen
failure. '

The wall thicknesses of the tube specimen were machined
to obtain the desired stress levels. The outer thick
wall of the double walled specimen was designed so that
it would not collapse into the annulus during operation
in the 3,500 psi steam.

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Two of the group of four Hastelloy N tube burst test
Specimens failed prematurely. The highest stressed
specimens (77,000 psi; 010" wall thickness) failed

in one (1) hour and the next highest stressed specimen
(52,500 psi; 0.015" wall thickness) failed in 3.7 hours.

These specimens which were removed after 1000 hours
exposure were stressed considerably above the yield

stress (40,000 psi) at 1000°F. The subsequent metallurgi-
cal examination of the failed specimen surfaces dis-
counted causes of failure due to the presence of flaws,
poor machining of gage sections or inaccuracies in
measurement of wall thicknesses.

A photograph of the two tube burst specimens and a
photomicrograph illustrating the cracking initiating
from the inside surfaces (steam side) of one of the
failed specimens are shown in Figures 25A, B.

The two failed tubular test specimens were subsequently
replaced by two similar type tube specimens stressed
@ 56,000 psi (0.014" wall) and 50,000 psi (0.016" wall).

Subsequently, four additional tube burst specimens were
installed in the test facility in order to increase

the rate of data accumulation. The specimen design was
identical to the other double-wall tube burst specimens
except that the annulus between the tubes was not con-
nected to the outside of the chamber by capillary tubing
for indication of failure. Internal diametral strains
were measured periodically and rupture times estimated
by comparison with the instrumented tests.

One of the specimens with the highest stress (58,000
psi) failed sometime during the first 1000 hours of
exposure. A small crack developed in the inner tube
and pressurized the annulus between the tubes. When
the plant steam pressure was reduced, the pressure in
the annular region collapsed the tube, as shown in
Figure 26. The specimen had a hairline crack extending
almost the entire length of the reduced section of the
tube wall. Numerous cracks formed on the inside and
outside surfaces of the tube.

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Examination of the Plant records at the Bull Run Plant
indicated that failure could have occurred when a
Pressure excursion to 3750 psig momentarily occurred,
and as a result the pressure dropped to zero in about
thirty (30) minutes. This 3750 psig would have been
equivalent to a stress of 62,000 psi on the Specimen.
Assuming this was the case, the failure occurred after
792 hours exposure. The appearances of the numerous
intergranular cracks formed in the failed sample are
shown in Figure 27.

" In the semiannual progress report23 period ending

8/31/72, B. McNabb and H. McCoy indicated that ten
instrumented tube burst specimens were in test at
Stresses varying from 28,000 to 72,000 psi (specimen
wall thicknesses varying from 0.0108" to 0.0302"),

Three additional specimens stressed at 66,000, 56,000
and 55,300 psi had failed at 4.0, 27.4 and 99.7 hours,
respectively. The investigators reported that the two
more highly stressed specimens had failed with con-
siderable deformation while the latter specimen failed
with less deformation.

The examination of the fractures showed that the speci-
men having the highest stress level exhibited the most
ductile fracture, being almost shear while the speci-
Men at the intermediate stress level had a mixed shear
and intergranular fracture characteristic of high
temperature failure.

The scatter in stress rupture data for tubes tested in
argon/steam environments led to the conduct of addi-
tional tests to determine if flawsg could be the cause
of the scatter. Control Specimens were ultrasonically
inspected and compared with an electrical-discharge
machined flaw 1 mil deep and 62 mils long.

O0f forty-two specimens inspected eleven gave indica-
tions equal to or greater than the standard. Specimen
wall machined thicknesses ranged from 0.011" to .020".
With an applied pressure of 3500 psi, a decrease of

1 mil in the wall thickness would increase the stress

in a 0.020" wall machined specimen by approximately

2000 psi and 6000 psi on a .011" wall machined specimen.

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NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1 DATE 12/16/74

A control tube burst test specimen was also tested at
1000°F at a constant stress of 40,300 psig in argon

with an argon internal pressure of 3500 psig. This
specimen failed after 565.2 hours which was considerably
less than a similar specimen exposed to steam still in
test after over 7000 hours.

The erratic behavior observed in test specimens was
believed to be due in part to the presence of undetec-
ted flaws in the thinly machined tube walls and, as a
consequence, no definite conclusions were drawn regard-
ing the effects of pure steam on the stress rupture
properties of Hastelloy N tubing material.

A photograph of the three Hastelloy N tube burst speci-
mens which failed after short time exXposure in the
supercritical steam environment is shown in Figure 28
while photomicrographs taken of the failure areas of
each of the specimens are shown in Figure 29, Photo-
micrographs illustrating the failure of the Hastelloy

N specimen tested in argon @ 1000°F are shown in Figure
30. The investigators indicated the specimen was not
ultrasonically inspected prior to testing and that the
numerous cracks observed on specimen inside surfaces
may have been present before testing.

A summary of the compiled tube burst test data is shown

in Table 21.

. 3

DUPLEX TUBING TESTS

Tests were also conducted at the Oak Ridge National
Laboratory to evaluate the properties of duplex tubing
for steam generator service. The duplex tubing which
was designed to provide each side of the tubing with
increased corrosion resistance for the molten fluoride
supercritical steam service was made by coextruding a
Nickel 280 sleeve over a wrought Incoloy 800 tube.
(The Nickel 280 alloy is pure nickel with .05% Al203
added for grain size control required for long term
elevated temperature service).

BY

APPROVED PAGE 6-18

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHEANGES HAVE BEEN MADE

The study also included an evaluation of the mechanical
and tube burst properties of the duplex tubing and
individual components as well as auxiliary tests to
evaluate the corrosion behavior of the Nickel 280 alloy
alone in supercritical steam environments @ 1000°F and
3500 psi pressure.

A twelve inch piece of 0.75 inch diameter tubing was
tensile tested @ 25°C at a displacement rate of 0.05
in/min. The test results indicated that the 0.2% off-
set yield stress was 37,600 psi; the ultimate tensile
Stress was 70,780 psi and the elongation, 50.5% in 2".
The room temperature results indicated that in the ten-
sile type test Nickel 280 and Incoloy 800 components
contribute equally to material strength.

A tube burst specimen was also tested @ 1000°F in an
argon atmosphere with an argon internal pressure and
ruptured in 3263 hours.

The results indicated that the Nickel 280 clad con-
tributed little to the strength of the tubing in this
type test. (The rupture life of 3263 hours for a
stress of 46,000 psi @ 1000°F is in agreement with data
published in Inco Technical Bulletin (T-40) for the
Incoloy 800 alloy).

A photograph illustrating the dye checked appearance
of the failed duplex Nickel 280/Incoloy 800 tube burst
specimen is shown in Figure 31A. A photomicrograph

of the Nickel 280 cladding illustrating the extent

and depth of cracking in the outer portions of the
nickel clad is shown 1in Figure 31B.

A photomicrograph illustrating the cracking occurring

on the inner Incoloy 800 tube and extending on through
the Nickel 280 clad is shown in Figure 32A. (The inter-
face between the two materials was metallurgically

sound and the Incoloy alloy was not cracked around the
circumference as was the Nickel 280 clad).

The nature of the profuse and preferential longitudinal
cracking in the Nickel 280 clad material led to the
conduct of additional mechanical tests to determine if
the prior working direction influenced the cracking
behavior.,

BY

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- FOSTER WHEELER ENERGY CORPORATION

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CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1 |DATE 12/16/74

Specimens from a Nickel 280 alloy sheet were cut
longitudinally and transversely to the rolling direc-
tion and creep tested at temperatures of 1000°F and
pressures of 2000 psi in argon. A plot of the percent
elongation v.s. time for specimens tested in the as
received condition is shown in Figure 32B. The data
shows that the longitudinal specimen is slightly
stronger with a longer rupture life and lower minimum
creep rate. However, fracture strains were high for
both specimens.

To observe the compatibility of Nickel 280 with steam
in the event of a leak, sheet specimens were exposed
to the supercritical steam @ 1000°F and 3500 psi for
2000 hours.

The specimens gained approximately 75 mg/cm2 and
appeared to be about completely oxidized indicating
that Nickel 280 is not compatible with steam under the
supercritical steam environments. A photograph of the
oxidized specimen is shown in Figure 33.

Additional tube burst tests were conducted on the duplex
tubing at 1000°F in an argon atmosphere with argon
internal pressure, One specimen with a pressure
corresponding to a stress of 46,000 psi on the Incoloy
800 (or 28,720 psi for the entire wall) ruptured at
3,263 hours with a diametral strain of 3.04%. The
Nickel 280 clad exterior was cracked profusely.

Another tube burst specimen was stressed at 40,000 psi
on the Incoloy 800 and was discontinued at 7075 hours
with a diametral strain of 1.147%. The dye checked
appearance of this clad specimen 1is shown in Figure 34.
A photomicrograph illustrating the appearance of the
cracking which occurs at relatively low diametral strain
is shown in Figure 35.

A graph illustrating several creep curves for Nickel
280 sheet at 1000°F in argon is shown in Figure 36,
The data indicates there is no clear cut effect of
rolling direction or creep properties, as the longi-
tudinal specimen had longer rupture times at the two
higher stresses and the transverse specimen had a
longer rupture life at 15,000 psi.

BY

APPROVED | PAGE 6-20

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

All the sheet specimens had rupture strains greater than
20% indicating much greater ductility than that shown
by the Nickel 280 on the duplex tubing.

Generally, the data indicates that the duplex tubing
can be produced with a nickel layer that possesses good
ductility.,

CRITICAL REVIEW OF STEAM GENERATOR TUBING MATERIALS

The literature search included a critical review of
alternate alloys for possible use as steam generator
tubing in Molten Salt Breeder Reactor service. This
involved a study of fundamental selection criteria and
review of all commercial and experimental alloys used
in the previously cited material investigations as
well as others described in the referenced literature.
1-108.

This also included technical discussions with tubing
manufacturers regarding the selection and usage of
commercial steam generator tubing materials; the
behavior of various iron and nickel base alloys subject
to long term exposure to high temperature supercritical
and molten salt environments; and the new duplex tubing
fabrication processes developed for special applica-
tions requiring increased corrosion protection for both
the inside and outside surfaces of tubing.

.1

DESIGN REQUIREMENTS AND CRITERTA

The steam generator is based on a 30 year plant life
and the use of tubing having a minimum I.D. of 0.375"
and corrosion allowances of 0.5 mil/yr. on the salt
side and 0.25 mil/yr. on the steam side.

The design operational conditions to which the tubing
will be subjected in service are summarized below:

BY

APPROVED B PAGE 6-21

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT

. LIVINGSTON, N. J.

CHARGE NO. DOCUMENT NO. ND/74/66

8§-25-24131

1

ISSUE

Coolant Salt

Inlet temperature, °F 1150
" Qutlet temperature, °F 850
Flow rate, 1b/hr. 15,280,000
Inlet pressure, psia 235
Outlet pressure, psia 175
Maximum pressure drop,psi 60

DATE 12/16/74

Water/Steam

700
1000
2,517,000
3800
3600
200

The coolant salt consists of sodium fluoroborate and
sodium fluoride salts having a eutectic composition

of 92-8 mole % (NaBF4—NaF).
below illustrates a liquidus

The phase diagram shown
temperature of 384°C

(723°F) indicating the coolant system must operate

above approximately 750°F.

ONNL-DWG 67-94231

y‘ & LIQUIDUS
i + SOLIDUS

1000 4 . : !

A~

1995°C e,

00 \ in CRYSTAL INVERSION |
. . ! !
. I D' H | | '
: : O~ | ! ;
00 : ‘ . , '\ 1 LIQUID | :
! ! ! 1
.~ ' ' l ' | .
s ' NoF + LIQUID e | | |
100 ) : } : : N ‘ f {
i ‘ ., | ‘
:l' . .
2
A ely! : ‘ '
a0 MNaiif, P |
o (HIGH- TEMPERATURE FORM) *\ IC\ !
VI g | ot L'OU’Di N N
— I, i
400 13847°C ——t . N
" NoF + NaDF, (HIGH - TE MPERATURE FORM)
300 ) ° ° 1 ’ r 1 r
243°C - _ ) .
F Fu (LOW-TEMPERATURE FORM)
200 oo NoF + NaBFg {(LOW-TE URE FORM} _ |
NoF 20 a0 60 80
NoBF, (mole %)
Two-Fluid MSBR Coolant Salt — The System
NaF-NaBF4,

PAGE 6-22

BY APPROVED

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FOSTER WHEELER ENERGY CORPORATION

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CHARGE NO. 8-25-2431 | DOCUMENT NO. np/74/66 ISSUE DATE 12/16/74

6.4.3.2

CORROSION RESISTANCE TO MOLTEN FLUORIDE SALTS

The best known alloy to contain molten fluoride salts
at temperatures up to 1500°F is Hastelloy N. This
alloy which was specifically developed for high
temperature molten fluoride service by ORNL as part of
the ANP program conducted during the mid 1950's posses=—
ses long term high temperature strength, thermal
stability, ductility, creep resistance and satisfactory
tube fabrication/joining properties.

In the early screening stages of alloy development
nickel base alloys were found to be the most promis-
ing and the commercial nickel-chromium group received
the most attention. However the alloys were subject

to selective chromium depletion processes which resulted
in the formation of subsurface voids illustrated in
Figure 37.

The corrosion mechanism consists of oxidation of alloy
constituents to their fluoride salts which are soluble
in the melt and therefore do not form passivating
protective films. This electrochemical type attack is
limited only by the thermodynamic potential for the
oxidation reaction and is more selective for chromium
since it is the least noble element in the alloy.

The possible corrosion reactions are expressed in the
equations below:

= N
2HF + Cr 2 H2 + CrF2

NiF, + Cr - Ni + CrF,

+ Cr <2 Ni + CrF,

The relative thermodynamic stabilities of the fluoride
composites are shown in Table 22 and illustrate that
molybdenum and nickel constituents are more stable and
more resistant to molten salt attack at high tempera-
tures.

BY

APPROVED | PAGE ¢_54

FOSTER WHEELER ENERGY CORPORATION
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CHARGE NO. 8-25-2431 | pOCUMENT No. ND/74/66 |11ssyE 1 |paTE 12/16/74

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FWC FORM 172 - L

In related developmental studies, Manley and coworker528

reported that while chromium in the newly developed
alloy was undesirable because of fused fluoride corro-
sion, the addition of at least 6% chromium was required
for oxidation resistance for long time service at
temperatures above 650°C (1200°F). The increased oxida-
tion resistance conferred by 6-8% chromium to an 80-20
Ni-Mo alloy is shown in Figure 38A. Also shown in
Figure 38B is a list of a number of alloys melted in
the course of developing the Hastelloy N (INOR-8). A
comprehensive review of the work conducted by ORNL and
other participating laboratories which led to the 29
development of the Hastelloy N alloy was published in
1969 by H. E. McCoy.

The results of corrosion tests obtained by ORNL in
continuing studies involving the behavior of Hastelloy
N and other materials subjected to various types of
fuel, blanket, fertile fissile and coolant molten salts
for various exposure times, temperatures, velocities,
At's in natural thermal convection loops (NCL) and
forced circulating loops (FCL) are described in
referenced ORNL MSR semiannual progress reports 30-38
and further discussion is beyond the scope of present
study.

Table 23 shows the status of the MSR program thermal
convection loops (NCL) as of 8/31/72.

Summarily the results of the continuing investigation

indicate that the Hastelloy N alloy is the most suit-

able material for high temperature molten salt service
having average corrosion rates of several tenths of a

mil/yr. '

The corrosion rate of the Hastelloy N material is
increased when the coolant salt (NaBF4—NaF, 92-8

mole %) 1is contaminated by the presence of impurities
particularly oxygen and moisture which results in the
formation of HF. However, analytical techniques are
being used to determine and monitor the oxygen and
water impurity levels while other techniques are being
developed to remove these impurities from the coolant
salt.

BY | APPROVED PAGE 6-24

FWC FORM 172 - L4

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

CHARGE NO. 8-25-243] | DOCUMENT NO. ND/74/66 ISSUE DATE 12/16/74

NOTATTIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

The corrosion reactions in the impure sodium fluoro-
borate salt mixture have not been well established
but they are reasonably expressed by the following
equations: | '

-
H,0 + NaBF, @ NaBF,OH + HF,

NaBF ;OH <  NaBOF, + HF,

6HF + 2Cr + 6 NaF = 2Na CrF6 + 3H

-
reactions with nickel ana iron),

2

9 (and similar
N

30, + 4Cr 2 2Cr203,

N ,
+ BNQBF4 + 6NaF 2Na3CrF6

FeFg + Cr & CrFy + Fe (and similar reactions with
nickel and molybdenum fluorides).

C 0 .+ 3NaBOF2,

ToV3

Graphs 1illustrating the relative corrosion resistance
of various alloys in a LIF—Ber—THF4-UF4 molten salt
are shown in Figure 39.

L] 3

CORROSION RESISTANCE TO SUPERCRITICAL STEAM

The best commercial alloy found for the supercritical
steam service is Incoloy 800. This alloy when exposed
to superheated steam at temperatures of 1050°~1100°F
forms a tenacious uniform and predictable oxide scale
and possesses high temperature strength, ductility,
creep resistance and thermal stability required for
long term high temperature/pressure steam service.

The alloy has been widely used for commercial steam
generating applications particularly where increased
strength and resistance to stress corrosion cracking
processes are required. The maximum allowable
stresses shown for the Incoloy 800 alloy in Table 25
were established and reviewed by the ASME Boiler and
Pressure Vessel Code Committee and the stress values
approved for use on 4/29/74 under Code Case 1592 for
Nuclear Vessels.

BY

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FOSTER WHEELER ENERGY CORPORATION

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CHARGE NO. 8-25-2431 | DOCUMENT NO.ND/74/66 ISSUE 1 DATE 12/16/74

Also shown in Table 24 for comparison purposes, are
the maximum allowable design stresses for 304 SS,

316 SS and 2 1/4Cr-1 Mo alloys. A graph illustrating
manufacturers data on the relative strengths of
Inconel 600, Incoloy 800 and Inconel 625 alloys are
shown for comparison purposes in Figure 40.

CRITICAL REVIEW OF HASTELLOY N AND DUPLEX TUBING

HASTELLOY N

The dilemma concerning the use of Hastelloy N in the
MSBR application is largely based on the fact that it
has no prior commercial use in steam generating appli-
cations and that the present nuclear application
requires the use of tubing that will be subjected to
steam side service conditions substantially more severe
than those previously encountered in conventional power
plant operations.

, 39
The recent work published by J. P. Hammond et al
indicates that the surface condition and prior heat
treatment has a pronounced effect upon the stress
corrosion cracking proclivity of welded and non
welded U~bend specimens subjected to impure, oxygenated
Steam environments containing small quantities of NaCl
introduced during cyclic operation from saturated
temperatures @ 544°F to superheat temperatures of
800-900°F.

A photograph of the U-bend specimen holder with
mounted specimens and two schematics illustrating
specimens/environmental conditions during cyclic
testing and the high pressure chloride SCC facility
at the Bartow Florida Plant are shown in Figures

41 A, B, C.

In these tests Hammond observed that the ground speci-
men surface condition promoted the highest suscepti-
bility to cracking in a number of alloys including
Hastelloy N while none of the ground and properly
annealed Hastelloy N specimens failed in the severe
test environment. ‘

BY

APPROVED PAGE 6-26

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. Np/74/66 | ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

These results which are summarized in the Table 25
indicate that the stress corrosion cracking sus-
ceptibility of the Hastelloy N material may be
reduced or eliminated by proper annealing treatment
prior to use. (The actual data as excerpted from
references 26 and 39 are shown in Tables 26 and 27
respectively).

The relieval of high levels of residual stress exist-
ing in as received material may provide additional
insight as to the behavior of the Hastelloy N tubing
material used by Comprelli5 at GEAP and H. McCoy at
ORNL in their constantly stressed tube burst tests.

In addition, it appears that the concept of machining
Hastelloy N specimen tube walls to thicknesses as low
as .010" to obtain required stress values should be
reevaluated to assess the effects of the magnitude
and distribution of macro/micro residual stresses
superimposed by the machining operations. The
magnitude of these stresses can be determined by
X-ray diffraction techniques.

16.4.4.2 DUPLEX TUBING

The use of duplex tubing to provide increased corrosion
protection to the outside and inside surfaces of the
steam generator tubing appears to be a prudent com-
promize requiring further consideration and test
evaluation.

In the duplex tube fabrication process developed by
Huntington Alloys in Huntington, West Virginia, a
Nickel 270 clad (pure nickel containing <.01%C)

sleeve made by powdered metallurgy techniques 1is

placed over a section of wrought Incoloy 800 tubing.
The composite is then sintered, coextruded and subjec-
ted to conventional cold drawing tube processing opera-
tions to obtain required cladding thicknesse$s ranging
from .010-.050" over final size Incoloy 800 tubing.

The duplex tubing section lengths vary from 85-120

feet in length depending upon the draw bench limitation
of the mill.

BY APPROVED PAGE g7

FWC FORM 172 -

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE  ; |DATE 12/16/74

The duplex tubing design strength is based on the ASME
Code approved strength levels for Incoloy 800 material
since the Nickel 270 clad does not contribute signifi-
cantly to the strength of the duplex tube,

The vendor recommends that the duplex tubing can be
welded by conventional standard welding practices using
Inconel 82 filler metal for the Incoloy 800 tube weld-
ments and Inconel 61 for the Nickel 270 clad weldments.

However, on the basis of the previously cited studies
conducted by J. Hammond and coworkers39 at the Bartow
Corrosion Test Facility it may be prudent to consider

for test purposes the use of duplex tube test specimens
containing Incoloy 800 tube weldments made from Inco A
welding rod and/or combinations of Inco A welding rod

and Inconel 82 filler metal since this combination appears
to be more resistant to stress corrosion cracking pro-
cesses when exposed to impure, oxygenated, supersaturated/

superheated steam environments containing sodium chloride
salts.

A vendor photograph illustrating a similar commercial
duplex tube composite consisting of a sintered nickel
shell on a Monel alloy 400 extrusion billet is shown
in Figure 42. Also shown are cross sections of the
clad tube sections at various processing stages.

Photomacrographs illustrating the transverse appearances
of experimental Nickel 270 and Nickel 280 claddings over
wrought Incoloy 800 tube sections are shown in Figure 43
in the unetched and macroetched conditions.

Photomicrographs taken of transverse sections of
coextruded duplex tube sections illustrating the micro-
Sstructural appearance in the interfacial areas of the
sound metallurgical bond between the Nickel 270 and 280
claddings on Incoloy 800 tube sections are shown in
Figures 44 and 45.

CONCLUSIONS/RECOMMENDATIONS

On the basis of the literature survey described herein
the following conclusions and recommendations are made:

BY

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FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO. 8-25-243]1 |} DOCUMENT NO. ND/74/66 ISSUE 1 DATE  12/16/74

1. There is insufficient published data to properly
assess or predict the performance of Hastelloy N in
high temperature/pressure supercritical steam environ-
ments under constant and cyclic loading conditions
encountered in Molten Salt Breeder Reactors.

2. There is no known single commercial tubing alloy
that can simultaneously provide the required long term
corrosion resistance to molten fluoride salts on one
side and supercritical steam on the other.

3. Standard and modified compositions of Hastelloy N
provide the best corrosion resistance to molten fluoride
salts than all other alloys. The corrosion resistance
depends essentially on the following:

a) alloy composition.

b) coolant salt temperature and velocity.

c) coolant salt impurity levels particularly oxygen and
water contents.

4. Published data concerning the performance of Hastelloy
N tubing in commercial steam generating power plant
applications is non existent.

5. Unstressed standard and modified Hastelloy N alloy
compositions exhibit satisfactory general corrosion
behavior in pure deoxygenated (<1 ppb) supercritical'
Steam environments @ 1000°F and 3500 psi.

a) The oxide formed is a thin, tenacious non spalling
type containing complex oxides consisting primarily of
Ni02, MoO9, a spinel and small quantities of Cr203.

b) The average depth of oxide penetration varies from
-01-.04 mil after 15,000 hrs. exposure. In oxygenated
(3-4 ppm) environments @ 1022°F and 3000 psi, oxide
depths of penetration of .16 mils were reported after
2400 hrs. exposure. -

c) Hastelloy N alloys exhibit higher corrosion rates in
Steam with increasing surface roughness. -

d) The corrosion resistance of standard Hastelloy N
alloys is improved by the presence of .5-.6% S1i.

BY

APPROVED PAGE 6-29

FWC FORM 172 - |}

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

e) The surfaces of standard Hastelloy N alloys are prone
to the formation of localized nodular type intergranular
penetrations possibly initiated and/or accelerated by
iron particle deposition.

f) Nodule depths of .4-.8 mils reported after 10,000

hours in deoxygenated (<1 ppb) supercritical steam
environments.

8) Pit depths of 1.5 mils were reported after 2,400
hours in oxygenated (3-~4 PPm) supercritical steam
environments.

h)  Modified Hastelloy N compositions containing low Mo
(12%) and Fe (.05%) contents exhibit deep localized
intergranular oxide penetrations up to 10 mils after
10,000 hours exposure in pure deoxygenated (<1 ppb)
supercritical steam environments. The intergranular
Penetrations are absent in similarly exposed modified
Hastelloy N compositions containing .5%7-2.1% Ti.

6. Material investigation indicates that the corrosion
behavior of stressed standard and modified compositions

of Hastelloy N under supercritical steam generating
conditions is uncertain.

a) The alloy in the as received and stressed condition
is prone to stress corrosion cracking processes when
subjected to pure and impure supercritical steam
environments containing small quantities of oxygen and
NaCl.

b) The Hastelloy N alloy in the properly annealed
condition exhibits immunity to stress corrosion cracking
in impure supersaturated/superheated steam environments
in the welded and non welded conditions.

7+ The use of duplex tubing to provide increased corro-
sion protection to the outside and inside surfaces of
the steam generator tubing appears promising and
requires additional testing/evaluation.

8. Due to uncertainties of material performance under
cyclic loading conditions,it is recommended that a
cooperative test program be considered to evaluate the
performance of Hastelloy N tubing material under commer-
cial high temperature/pressure steam service conditions.
The program should include the incorporation of selected

sections of Hastelloy N tubing in the superheater sections
of several commercial bollers.,

1BY

APPROVED PAGE 6-30

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

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CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

REFERENCES

1. STRESS CORROSION OF NICKEL-CHROMIUM ALLOYS IN 350°C HIGH
PURITY WATER, Warren E. Berry, W. Stiegelmeyer and F.
Fink. Special report prepared by Battelle Memorial
Institute for the International Nickel Co., 9/22/61.

2. STRESS CORROSION CRACKING OF INCONEL 1IN HIGH-TEMPERATURE
WATER, H. Coriou, R. Grall, M. LeGall and S. Vettier, 3e
Colloque De Metallurgie sur la Corrosion (1959), North
Holland Publishing Co. (1960), pp 161-169.

3. CORROSION OF STAINLESS STEELS IN SUPERCRITICAL WATER, W. K.
Boyd and H. A. Pray. Corrosion 1956, 375t-384¢t.

4, MATERIALS FOR NUCLEAR SUPERHEATED APPLICATIONS, C. N.
Spalaris, F. A. Comprelli, D. L. Douglass and M. B.
Reynolds. AEC Project Agreement #13 Contract AT (04-3)~189
GEAP 3875, Jan. 1962.

5. MATERIALS FOR SUPERHEATED FUEL SHEATHS-RELATIVE PERFOR-
MANCE OF ALLOYS IN SUPERHEATED STEAM ENVIRONMENTS, F. A.
Comprelli, D. F. MacMillian and C. N. Spalaris, AEC Project
Agreement #13 Contract AT (04-3)-189, GEAP 4351, July 1963.

6. GENERAL CORROSION OF INCOLOY 800 IN SIMULATED SUPERHEATED
REACTOR ENVIRONMENT, W. L. Pearl, E. G. Brush, G. G. Gaul
and G. P. Wozadlo, GEAP 4495, March 1964.

7. INCOLOY 800 FOR NUCLEAR FUEL SHEATHS (A MONOGRAPH), C. N.
SPALARIS, GEAP 4633, July 1964

8. STABILITY OF HIGH-NICKEL ALLOYS IN SUPERHEATED STEAM, F, A.
Comprelli, U. E. Wolff, GEAP 4745, Nov. 1964,

9. GENERAL CORROSION OF MATERIALS FOR NUCLEAR SUPERHEATED
APPLICATIONS, W. L. Pearl, E. G. Brush, G. G. Gaul, G. P.
Wozadlo, GEAP 4760, March 1965,

10. GENERAL CORROSION OF INCONEL ALLOY 625 IN SIMULATED SUPER-
HEATED REACTOR ENVIRONMENT. W. L. Pearl, E. G. Brush,
G. G. Gaul, and S. Leistikow, Nuclear Applications, Vol. 3,
No. 7., p 418-432; July 1967, BTR.

11. EFFECTS OF EXPOSURE TO SUPERHEATER - OXYGENATED STEAM OF
550°C (1022°F) AND 1000 - 3000 PSI PRESSURE UPON IRON AND
NICKEL BASE ALLOYS, H. J. Pessl - TID, 7674 Proceedings
of the Nuclear Superheat Meeting No. 8 3/20, 21, 22, 1963
Idaho Falls, Idaho.

BY APPROVED PAGE 6-31

FWC FORM 172 - L

IAVE BEEN MADE

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT No. ND/74/66 ISSUE 1 |pATE 12/16/74

12.

13.

14,
15.

16.

17.

18.

19,

20.

21.

22.

AN EVALUATION OF THE CORROSION RESISTANCE OF SEVERAL HIGH
TEMPERATURE ALLOYS FOR NUCLEAR APPLICATIONS, T. T. Claudson,
R. E. Westerman, P. 26(1965) Nov. Contract AT (45-1)-18390.
BNWL 155,

EVALUATION OF IRON*AND~NICKEL*BASE ALLOYS FOR MEDIUM AND
HIGH TEMPERATURE REACTOR APPLICATIONS, T. T. Claudson,
H. J. Pessl,; p, 60, (1965) November. Contract AT(45-1)~

183¢G. BNWL 154,

MATERIALS DEVELOPMENT FOR MOLTEN SALT BREEDER REACTORS,
H. E. McCoy, J. R. Weir, Contract W-7405-eng 26, 0ak Ridge
National Laboratory Report ORNL-TM-1854, June 1967,

DESIGN STUDIES OF 1000 Mu(e) MOLTEN-SALT BREEDER REACTOR;
ORNL 3996 (8/66), P. R. Kasten, E, §. Betlis, and R. (.
Robertson.

Tennessee Valley Authority, The Bull Run Plant, TVA
Technical Report 28, Knoxville, Tenn., 1967.

MOLTEN SALT REACTOR PROGRAM, PERIOD ENDING 8/31/69, Con-
tract W-74-5-eng 26 0Oak Ridge National Laboratory Report,
ORNL 4449, 8/31/69.

MOLTEN SALT REACTOR PROGRAM SEMIANNUAL PROGRESS REPORT,
PERIOD ENDING 2/28/70, Contract W-7405-eng 26, 0ak Ridge
National Laboratory Report, ORNL 4548, 8/70.

FUELS AND MATERTIALS DEVELOPMENT PROGRAM QUARTERLY PROGRESS
REPORT PERIOD ENDINGC 3/31/70 (Advanced Materials In Steam
Generators Pp 275-280), o0ak Ridge National Laboratory
Report, ORNL QEEQ, 3/31/70,

METALS AND CERAMICS ANNUAL PROGRESS REPORT PERIOD ENDING
6/30/70, Contract W-7405-eng 26, 0ak Ridge National Labora-
tory Report ORNL 4570, 10/70.

MOLTEN SALT REACTOR PROGRAM SEMIANNUAL PROGRESS REPORT
PERIOD ENDING 2/28/71, Contract W-7405-eng 26, 0ak Ridge
National Laboratory Report, ORNL 4676, 2/28/71.

MOLTEN SALT REACTOR PROGRAM SEMIANNUAL PROGRESS REPORT
PERIOD ENDING 8/31/71, Contract W~7405-eng 26, O0ak Ridge
National Laboratory Report ORNL 4728, 8/31/71.

BY

APPROVED | AGE 6-32

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO., 8-25-2431 DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

24. CORROSION AND MASS TRANSFER CHARACTERISTICS OF Na BF4-NaF
(92-8 mole %) IN HASTELLOY N, J. W. Koger, Contract W-7405-
eng 26, Oak Ridge National Laboratory Report ORNL-TM-3866,
10/72.

25. MOLTEN SALT REACTOR PROGRAM SEMIANNUAL PROGRESS REPORT,
PERIOD ENDING 2/29/72, Contract W-7405~-eng 26, H, E. McCoy,
J. R. Weir, Oak Ridge National Laboratory Report ORNL 4782,
2/29/72.

26, FUELS AND MATERIALS DEVELOPMENT PROGRAM QUARTERLY PROGRESS
REPORT PERIOD ENDING 12/31/72, Contract W-7405-eng 26 0ak
Ridge National Laboratory ORNL-TM-4105, 12/31/72.

27. CORROSION OF SEVERAL IRON AND NICKEL BASE ALLOYS IN SUPER-
CRITICAL STEAM AT 1000°F, H. E. McCoy and B. McNabb, 8/74,
ORNL-TM 4552,

'28. METALLURGICAL PROBLEMS IN MOLTEN FLORIDE SYSTEMS, W. D.
Manly, J. H. Cools, J. H. DeVan, D. A. Douglas, H. Inouye,
P. Patriarca, T. K. Roche, and J. L. Scott, Geneva Confer-
ence Paper (1958) p/1990, Reprinted from "Progress in
Nuclear Energy, Series IV, Vol 2 - Technology, Engineering
and Safety.

29. THE INOR-8 STORY, H. E. McCoy, Review, Fall 1969, ORNL
pp 35 - 49,

30. METALS AND CERAMICS DIVISION ANNUAL PROGRESS REPORT FOR
PERIOD ENDING 6/30/67, Contract W-7405~-eng 26 Oak Ridge
National Laboratory Report ORNL 4170, June 30, 1967.

31. MATERIALS DEVELOPMENT FOR MOLTEN SALT BREEDER REACTORS,
H. E. McCoy, J. R. Weir, Contract W-7405-eng 26, Oak
Ridge National Laboratory Report ORNL-TM~-1854, June 1967.

32. MOLTEN SALT REACTOR PROGRAM, SEMIANNUAL PROGRESS REPORT
FOR PERIOD ENDING 8/31/67, H. E. McCoy, Contract W-7405-
eng 26, Oak Ridge National Laboratory Report ORNL 4191,
Aug. 31, 1967.

33. AN EVALUATION OF THE MOLTEN SALT REACTOR EXPERIMENT
HASTELLOY IN SURVEILLANCE SPECIMENS-FIRST GROUP, H. E.
McCoy, Contract W-7405-eng 26 Oak Ridge National Laboratory
Report ORNL-TM-1997, Nov. 1967.

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

34.

35.

36.

37.

38.

39.

40.

41.

42,

43.

4254, Feb. 2», 1968,

TWO FLUID MOLTEN SALT BREEDER REACTOR DESIGN STUDY STATUS
AS OF 1/1/68, Contract W-7405-eng 26, Oak Ridge National
Laboratory Report ORNL 4528, 1/1/68,

MOLTEN SALT REACTOR PROGRAM, PERIOD ENDING 2/29/68, Contract
W-7405-eng 26, 0ak Ridge Natlonal Laboratory Report ORNL

MOLTEN SALT REACTOR PROGRAM, PERIOD ENDING 8§/31/68, Con-
tract W-7405-eng 26, Oak Ridge National Laboratory Report
ORNL 4344, 8/31/68.

MOLTEN SALT REACTOR PROGRAM, PERIOD ENDING 2/28/69, Con-~
tract W-7405-eng 26, Oak Ridge National Laboratory Report,
ORNL 4396, 2/28/69.

MOLTEN SALT REACTOR PROGRAM, PERIOD ENDING 8/31/69, Con-
tract W-74-5-eng 26, Oak Ridge National Laboratory Report,
ORNL 4449, 8/31/69.

CHLORIDE STRESS-CORROSION OF STEAM GENERATOR MATERIALS AND
POST-TEST ANALYSIS OF AN INCOLOY 800 LOOP by J. P. Hammond,
P. Patriarca and G. M. Slaughter of ORNL and W. A. Maxwell
of SNE. Paper 51 presented at NACE Conference 3/74,
Chicage, I1l1.

CORROSION STUDIES IN WATER AND VAPOR AT HIGH TEMPERATURES,
Contract 089-62-7 RDB Quarterly Report 1, 6/16- 9/30/62,
10/62, EURAEC 474.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM. Quarterly Report No. 2 Oct. 1 to Dec. 31 1962,
Societe d'Etudes, de Recherches et d' Appllcatlons pour
1'Industrie, Jan. 30, 1963. Work performed under U.S.
Euratom J01nt Res. and Dev. Program. A Translation.
1/30/62, EURAEC 552.

DEVELOPMENT OF NIOBIUM-BASE ALLOYS FOR USE 1IN A NUCLEAR
SUPERHEATER. Quarterly Progress Report. January-March
1963. John A. DeMastry, Arthur A. Bauer and Frank A,
Rough. Battelle Memorial Inst. April 1, 1963. Work
performed under U.S.-Euratom Joint Res. & Dev. Program,
Contract W-7405-eng 92, 12 pp. 4/1/63 EURAEC-626.

CORROSION TESTS IN STEAM. Quarterly Report No. 8,
January-March 1964. Work performed under US-Euratom Joint
Res. and Dev. Program, 27 pp (1964), Contract 067-62-4-RDF,
5/64, EURAEC-1062.

BY

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FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE ——

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

44,

45,

46 .

47.

48.

49,

50.

51.

52.

STUDIES ON CORROSION OF STEELS 1IN HIGH-TEMPERATURE WATER
AND STEAM. Quarterly Report No. 9, July-Sept. 1964,
Euratom Joint Res. and Dev. Program, 48 pp (1964),
Contract 089-62-7 RDB, 10/30/64, EURAEC—lZgl.

DEVELOPMENT OF NIOBIUM ALLOYS RESISTANT TO SUPERHEATED
STEAM. Robert D, Koester, John A. DeMastry, Warren E.
Berry, Arthur A. Bauer, Frank A, Rough, Battelle Memorial
Inst., Columbus, Ohio. Work performed under United States-
Euratom Joint Res. and Dev. Program, 44 pp (1964), Contract
W-7405~eng 92, 11/11/64 EURAEC-1285,

CORROSION STUDIES ON STEELS IN WATER AND STEAM AT HIGH
TEMPERATURE. Quarterly Progress Report No. 10, Oct.-Dec.
1964. Euratom Joint Res. and Dev. Program, 82 p.,
Contract 089-62-7 RDB, Jar. 1965, EURAEC-1308.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM, Contract 089-62-7 RDB, Quarterly Report 13 7/1-
10/31/65; 11/26/65, EURAEC~1500.

DYNAMIC CORROSION TESTS ON CARBON AND STAINLESS STEELS 1IN
PRESSURIZED WATER. M. Warzee, P. de Dorlodot, J. Waty,
CFSTI, Contract 089-62~7-RDB (EUR-2688) 12/6/65),
EURAEC-1546.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM, Societe d'Etudes, de Recherches et d'Applications
pour 1'Industrie, Contract 089-62-7 RDB; 31 p. Dec. 1965,
EURAEC-1581.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM, Societe d'Etudes, Contract 087-66-ITEEB (RD).
(EUR-2838) Dep. mn CFSTI (In Belgium), Apr. 1966,
EURAEC-1625.

CORROSION OF STAINLESS STEELS IN HIGH TEMPERATURE WATER

AND STEAM. M. Warzee, W. R. Ruston, P. deDorlodot,
J. Hennaut, J. - PH Berge; (Societe d'Etudes, de Recherches

et d'Applications pour l1'Industrie, Brussels (Belgium),
Contract 089-62-7-RDB, 14 p. (EUR-2857. e; CONF~660514~3) ;
Feb. 1966, LURAEC-1665,

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM, Quarterly Report No. 15, 4/16/30/66; 7/25/66,
EURAEC-1689.

BY

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FWC FORM 172 -

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

FOSTER WHEELER ENERGY CORPORATION

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

53.

54.

35.

56.

57.

59.

60.

61.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM, Societe d'Etudes, de Recherches et d'Applications
pour 1l'Industrie, Brussels (Belgium), (EUR-3304) 10/30/66,
Contract 087-66-1 TEEB(RD), 89 p EURAEC-1744,

BEHAVIOR OF STAINLESS STEELS IN SUPERHEATED STEAM. G.
Chaudron; Contract 059-65-7-~-TEEB(RB); 18 p (EUR- 3306) ;
Quarterly Report #5, 7/1-9/30/66; EURAEC-1749.

INFLUENCE OF SURFACE TREATMENT ON CORROSION OR CARBON AND
STAINLESS STEELS IN HIGH TEMPERATURE WATER AND STEAM,

PART II. CORROSION EVALUATION THROUGH DETERMINATION OF
THE HYDRCGEN FORMED IN THE REACTION. C. Sonnen, M. Warzee,
(Societe d'Etudes, de Recherches et d°' Appllcatlons pour
1'Industrie, Brussels (Belgium); Contract 089-62-7 RD- ~B;
35 p; (EUR-3319) Dep. mn CFSTI Dec. 1966, EURAEC 1764.

BEHAVIOR OF STAINLESS STEELS IN SUPERHEATED STEAM.
G. Chaudron; Contract 059-65-1 TEEB(RD): 18 p (EUR-3345);
Quarterly Report #16 10/1-12/31/66, EURAEC-1804.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM. Societe d'Etudes, Contract 087-1- TEEB(RD), 102 P
(EUR-3361); May 1967, EURAEC-1850.

CORROSION OF STAINLESS STEEL AND HICGH NICKEL CONTENT
ALLOYS IN HIGH~TEMPERATURE SUPERHEATED STEAM. M. Warzee;
C. Sonnen; J. Cremer; Ph. Berge; (Societe d'Etudes,
(Belgium); Contract 087-1-TEEB(RD), 74 p. (EUR-3387);
July 1967, EURAEC-1895.

OXIDATION OF SUPERHEATED MATERIALS IN STEAM: AUSTENITIC
STAINLESS STEEL, J. Board, G.Holyfield; J. Dalley (Atomic
Power Constructions Ltd. Hounslow, Eng ) pp 163-173 of
Journees Internatlonales d'Etude sur 1' Oxydatlon des Metaux
Brussels, Societe d'Etudes de Recherches et d' Applications
pour 1' Industrle (1965).

NEW STUDIES OF THE ISOTHERMAL CORROSION OF INCOLOY 800 BY
SUPER-HEATED - STEAM. S. Leistikow, Kernforschungszentrum,
Karlsruhe (West Germany) 5 p (l97l) March.

INVESTIGATIONS OF THE CORROSION BEHAVIOR OF AUSTENITIC
CrNi STEELS AND NICKEL ALLOYS IN SUPERHEATED STEAM.
S.Leistikow, 32 p (1972) (KFK~-1463) (N72-25534) EUR
FNR 954,

BY

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FWC FORM 172 - L4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN‘MADE

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 IDATE 12/16/74

62.

63.

64 .

65,

66.

67.

68.

69.

INFLUENCE OF COLD WORK ON THE CORROSION RESISTANCE OF
AUSTENITIC CrNi STEELS IN SUPERHEATED STEAM. 2. STATIC
AUTOCLAVE STUDIES ON INCOLOY 800 IN DIFFERENT GRAIN SIZES.
S. LEISTIKOW, E. Pott, (Kernforschungszentrum, Karlsruhe
(West Germany). Institut fuer Material und
Festkocrperforschung). 28 pp Dec. 1972 (KFK-1681).

UERFNR 1061.

INFLUENCE OF COLD WORK ON THE CORROSION RESISTANCE OF
AUSTENITIC CrNi STEELS IN SUPERHEATED STEAM. Part 3.
STATIC AUTOCLAVE INVESTIGATIONS ON MATERIAL 1.4301

(x 5 CrNi 18.9). S. Leistikow, E. Pott and W. Volz.
(Kernforschungszentrum, Karlsruhe (West Germany). Institut
fuer Material and Festkoerperforschung). 34 pp March 1973,
EURFNR 1087.

WATERSIDE CORROSION IN ADVANCED DESIGNS OF NUCLEAR STEAM
GENERATORS, R. Garnsey, B. Hearn, G. M. W. Mann. Journal
of the British Nuclear Energy Society, Vol II (1972)

p. 65-70.

METALLURGICAL PROBLEMS IN MOLTEN FLUORIDE SYSTEMS, W. D.
Manly, J. H. Cools, J. H. DeVan, D. A. Doublas, H. Inouye,
P. Patriarca, T. K. Roche and J. L. Scott. Pergamon

Press reprinted from "Progress in Nuclear Energy'".

Series 1V, Vol 2-Technology, Engineering and Safety,
Geneva Conference Paper (1958) P/1990 pp 164-179.

PROBLEMS DE CORROSION POSES PAR L'EMPLOI DES ACIERS DANS
LA CONSTRUCTION DES GENERATEURS DE VAPEUR POUR REACTEURS
NUCLEAIRES, H. Coriou, R. Darras, L. Grall, O. Konovalt-
schikoff, M. Pelras, J. Sannier, R. Teraube. Fourth
International Conference on the Peaceful Uses of Atomic
Energy, Geneva, Switzerland, Sept. 6-16/71.

MATERIALS FOR LIQUID METAL FAST BREEDER REACTORS VESSELS,
PIPING, INTERVALS AND CLASSING, S. H. Bush, 5/24/72. Conf.
7211104 Pacific Northwest Lab. Div. of Battelle Memorial

Institute.

RELATION OF TUBE MANUFACTURING VARIABLES TO CORROSION AND
CORROSION RELEASE RATES OF A NI-CR-FE ALLOY IN SIMULATED
P.W.R. PRIMARY WATER, G. C. Bodine Jr., J. W. Carter,
Paper #53 NACE 1974.

PWR WATER: VARYING CHEMISTRY AND "HOT" CONTAMINANTS.
W. D. Fletcher, Industrial Water Engineering. (1971)
March.

e S

BY

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FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHAWGES HAVE BEEN MADE —

FWC FORM 172 - |4

70,

71.

72.

73.

74.

75.

76.

77.

CORROSION IN WATER-COOLED NUCLEAR POWER REACTORS. P. J.
Shirvington, Australasian Corrosion Engineering (1971)
Jan-Feb.

STAINLESS STEEL REACTOR PRESSURE VESSELS, 27786, P. J.
Karnoski, Jr., (Brown and Root, Inc., Houston, Tex.),
W. J. Fretague, U. Potapovs, and L. E. Steele, Nucl.
Eng. Design, II, 347-67 (1970) April.

OXIDATION RESISTANCE OF SOME STAINLESS STEELS AND NICKEL~-
BASED ALLOYS IN HIGH-TEMPERATURE WATER AND STEAM, S.
Jansson, W. Huebner, G. Oestberg, M. Pourbaix, Brit.
Corros. J., Vol. 4, 21-31 (1969) Jan.

NEW DEVELOPMENTS IN MATERIALS FOR MOLTEN-SALT REACTORS,

H. E. McCoy, R. L. Beatty, W. H. Cook, R. E. Gehlbach,

C. R. Kennedy, J. W. Koger, A. P. Litman, C. E. Sessions
and J. R. Weir, Nucl. Appl. Technol., Vol. 8, 156-69 (1970)
Feb.

CORROSTION BEHAVIOR OF STAINLESS STEEL IN HIGH-TEMPERATURE
WATER AND STEAM. T. Maekawa, and M. Kagawa, Nippon
Kinzoku Gakkaishi, Vol. 31, 1213-19 (1967) Oct. (In
Japanese).

STRESS CORROSION CRACKING OF STAINLESS STEELS AND NICKEL
BASE ALLOYS IN CHLORIDE CONTAINING WATER AND STEAM AT
200-300°C; W.lubner, M. DePourbaix, and G. Ostberg,
Paper presented at the 4th International Congress on
Metallic Corrosion, Amsterdam, The Netherlands (1969)
Sept.

STUDY OF SEVERAL HIGH TEMPERATURE CORROSION PHENOMENA OF
AUSTENITIC IRON-CHROME-~NICKEL ALLOYS IN WATER AND SUPER-
HEATED STEAM. H. Coriou, L. Grail, and M. Pelras, Pre-
sented at the 4th International Congress on Metallic
Corrosion, Amsterdam, Netherlands, Sept. 7-14 (1969)

Order from NACE, 2400 West Loop South, Houston, Tex. 77027.

CORROSION AND FAILURE CHARACTERISTICS OF ZIRCONIUM ALLOYS

IN HIGI-PRESSURE STEAM IN THE TEMPERATURE RANGE 400 T0O 500°¢
A. B. Johnson, Jr., 25 p (1969) Sept. Contract AT(45-1)-
1830 CONF-681105-~2 CFSTI From Application-Related Pheno-
menon of Zirconium and Hafnium Symposium, Philadelphia, Pa.

BY

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FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO.ND/74/66 ISSUE 1 DATE 12/16/74

/8.

79,

80.

81.

82.

83.

84.

85.

86 .

87.

EFFECT OF SOMLE ENVIRONMENTAL CONDITIONS ON STRESS CORROSION
BEHAVIOR OF NI-CR-FE ALLOYS IN PRESSURIZED WATER. H. R.
Copson and G. Economy; Corrosion 24(3), (1968).

MECHANISMS OF SENSITIZATION AND STABILIZATION OF INCOLOY
NICKEL IRON-CHROMIUM ALLOY 825, E. L. Raymond; Corrosion,
24(6), (1968).

THE RESISTANCE OF VARIOUS TRON-CHROMIUM-NICKEL AUSTENITIC
ALLOYS AND AN IRON-40 PERCENT ALUMINUM ALLOY TO SUPERHEATED
STEAM. H. Coriou, L. Grall, C. Maniev and M. Pelras (A.T.B
Metallurgie, 6(2), 79-85, 1966.

CORROSION OF INCOLOY 800 AND NICKEL BASE ALLOY WELDMENTS IN
STEAM. J. P. Hammond (Oak Ridge National Lab., TN);

0. Patriarca, G. M. Slaughter, W. A. Maxwell and J. Weld
(N.Y.); Vol. 52, No. 6, pp 268s~-280s (1973) June.

FLUORIDE SALT CORROSION AND MASS TRANSFER IN HIGH TEMPERA-
TURE DYNAMIC SYSTEMS, J. W. Koger, Corrosion, Vol. 29,
do. 3 (1974). ‘

STRESS CORROSION CRACKING OF SENSITIZED STAINLESS STEEL IN
OXYGENATED HIGH TEMPERATURE WATER, Warren E. Berry, Earl
L. White, and Walter K. Boyd, Corrosion, Vol. 29, No. 12,
(1973). ' :

THE INTERGRANULAR STRESS CORROSION CRACKING OF NICKEL
ALLOYS IN PRESSURIZED WATER. J. Weber and P. Sury (Sulzer
Brothers, Ltd., Winterthur, Switzerland). Presented at
Corrosion/74.

STEAM GENERATOR AND HEAT EXCHANGER MATERIALS. (ORNL-TM-
4105), pp. 2.1-2.52 (1973) May.

INFLUENCE OF COLD WORK ON THE CORROSION RESISTANCE OF
AUSTENITIC Cr-Ni STEELS IN SUPERHEATED STEAM. Part 3.
STATIC AUTOCLAVE INVESTIGATIONS ON MATERIAL 1.4301

(x 5 CrNi 18.9). S. Leistikow, E. Pott, and W. Volz.
(Kernforschungszentrum, Karlsruhe (West Germany). Institut
fuer Material and Festkoerperforschung). 34 pp. (1973)
March. Dep.

INVESTIGATION INTO THE HOT STEAM CORROSION OF WELDED
INCOLOY 800 PLATE. S. Leistikow and W. Scheibe.
werkstoffe u. Korrosion Vol. 24, p. 269-273 (1973).

BY

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NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT _ LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

88.

89.

90.

91.

92.

93.

94 .

95.

96 .

97.

RELATION OF TUBE MANUFACTURING VARIABLES TO CORROSION AND
CORROSION PRODUCT RELEASE RATES OF A NiCrFe ALLOY 1IN
SIMILATED PWR PRIMARY WATER, G. C. BODINE, JR., J. W,
Carter, Paper # 53 NACE, 1974.

CORROSION OF ADVANCED STEAM GENERATOR ALLOY WELDMENTS 1IN
1100 AND 1200°F (595 AND 650 C) STEAM. J. P. Hammond,

P. Patriarca, G. M. Slaughter, and W. A. Maxwell, Proceed-
ings of the 26th Conference of the National Association of
Corrosion Engineers, Philadelphia, Pa., March 2-6, 1970,
pp 277-91, Houston, Texas, National Association of Corro-
sion Engineers (1970) (CONF-700324-1) Paper No. 46.

GENERAL CORROSION OF STAINLESS STEELS AND NICKEL BASE
ALLOYS EXPOSED ISOTHERMALLY IN SUPERHEATED STEAM. G. P.
Wozadlo and W. I. Pearl, Corrosion 21, No. 11 (1965) Nov.

WUCLEAR SUPERHEATED PROJECT TWENTY-SECOND QUARTERLY REPORT,
NOV. 1964 - JAN. 1965. W. L. Fiock, General Electric Co.,
San Jose, Calif., Atomic Power Equipment Dept., 163 p.
(1965) Feb. Contract AT(04-3)-189,

CORROSION-RATE-LAW CONSIDERATIONS 1IN SUPERHEATED STEAM.
E. G. Brush, Nucl. Appl., 1, 246-51 (1965) June.

CORROSION OF CARBON STEEL IN SIMULATED BOILING WATER AND
SUPERHEATED REACTOR ENVIRONMENTS. W. L. Pearl and G. P.
Wozadlo, Corrosion, 21 No. 8(1965) August.

CORROSION OF STEELS AND NICKEL ALLOYS IN SUPERHEATED STEAM,
W. E. Ruther and S. Greenberg, Journal of the Electro-
chemical Society, 3, No. 10, 1116-1121 (1965) Oct.

MATERTALS FOR BOILING WATER AND SUPERHEATED STEAM NUCLEAR
POWER REACTORS. W. R. Smith, Sr., Materials Research and
Standards, 4 No. 4, 170-174 (1964) April.

ZIRCONIUM ALLOYS FOR USE 1IN SUPERHEATED STEAM, S.Greenberg,
J. Nuclear Materials, 4, No. 3, 334-335 (1961) August/

September.

TESTING STEAM GENERATOR MATERIALS FOR PRESSURIZED WATER
REACTOR PLANTS, John W. McGrew, Corrosion, 18, No. 1,
27t-32t (1962) Jan.

BY

APPROVED PAGE 6-40

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO., 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

98.

99.

100.

101.

102.

103.

104,

105.

106.

107.

STUDIES OF STEEL CORROSION IN HIGH TEMPERATURE WATER AND
STEAM. Quarterly Report No. 6, Oct,-Dec. 1963. Euratom
Joint Research and Development Program, 70 pp (1964) Jan.
Contract 089 62-7 RDB, Order from OTS.

STRESS CORROSION CRACKING OF AUSTENITIC STAINLESS STEEL IN
SUPERCRITICAL STEAM. P, P. Snowden, Monthly Technical
Bulletin of International Combustion Ltd., 1V 17, Jan.
1963 pp 39-45.

CORROSION BEHAVIOR OF STEELS AND NICKEL ALLOYS IN SUPER-
HEATED STEAM. W. E. Ruther, R. R. Schlueter, R. H. Lee,

R. K. Hart, Presented at NACE 2l1st Conference in St. Louis,
Missouri 3/14~19/65.

REVIEW ARTICLE, THE AQUEOUS CORROSION OF REACTOR METALS,
J. N. Wanklyn and P. J. Jones. UKAEA Metallurgy Division,
Atomic Energy Research Establishment, Harwell, Didcot,
Berks, U. K. Received 12 March 1962.

METALLURGICAL EVALUATION OF SUPERHEATER TUBE ALLOYS AFTER
SLX-MONTHS EXPOSURE AT TEMPERATURES OF 1100° to 1500°F.,
C. L. Clark, J. J. B. Rutherford, A. B. Wilder, M. A.
Cordovi, Journal of Engineering for Power, Jan. 1960,

pp 35-67.

METALLURGICAL EVALUATION OF SUPERHEATER TUBE ALLOYS AFTER
12 AND 18 MONTHS EXPOSURE TO STEAM AT 1200, 1350, AND
1500°F. C. L. Clark, J. J. B. Rutherford, A. B. Wilder,
M. A. Cordovi, Transactions of ASME, July 1962 pp 258-288.

EFFECT OF ALLOY COMPOSITION ON STRESS CORROSION CRACKING
OF FE-CR-NI BASE ALLOYS, R. W. Staehle, J. J. Royuela,
T. L. Raredon, E. Serrate, R. V. Farrae, Corrosion,

Nov. 1970 451-486.

SCALING BEHAVIOR OF SUPERHEATER TUBE ALLOYS IN ASME HIGH
TEMPERATURE STEAM RESEARCH TESTS AT 1100-1500°F; F. Eberle,
C. H. Anderson; Journal of Engineering for Power, July

1962 pp 223-257.

CORROSION PROBLEMS IN NUCLEAR REACTOR POWER STATIONS, W. Z.
Friend, Reprint, Vol. XVIII Proceedings of the American
Power Conference, 1956.

STUDY OF SEVERAL HIGH TEMPERATURE CORROSION PHENOMENA OF
AUSTENITIC IRON-GHROM~NICKEL ALLOYS IN WATER AND SUPER-
HEATED STEAM. H. Coriou, L. Grall and M. Pebras, Pre-
sented at 4th International Congress on Metallic Corrosion,
Amsterdam, Netherlands, Sept. 7-14 (1964) .

BY

APPROVED PAGE 6-41

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 -

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

108.

PROBLEMES DE CORROSION POSES PAR L'EMPLOI DES ACIERS DANS
LA CONSTRUCTION DES GENERATEURS DE VAPEUR POUR REACTEURS
NUCLEAIRES. H. Coriou, R. Darras, L. Grall, O.
Konovaltschikoff, M. Pelras, J. Sannier and R. Teraube,
Presented at 4th U.N. International Conference on the
Peaceful Uses of Atomiec Energy, Geneva, Switzerland,
Sept. 6-16, 1971.

BY

APPROVED PAGE 6-42

IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

-~

s

ATTOR

FW. RN 172 - |

v

N

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 DOCUMENT NO. ND/74/66 ISSUZ 1 DATE 12/16/74

Fable 1 Nominal Chemical Composition
of Hastelloy N

Standard Alluyl 4 Modified /\Iloy9

Ulmint (Much as Used Recommended for

in MSRIE) MSBR'S
(wt ()9 (wi %)

Nickel Balance Balance

Molybdenum 15-18 12

Chromium 6-8 7

Iron 5 0-4

Muanganese | 0.2-0.5

Silicon | 0.1 max

Roron 0.01 0,001 max

Lituniom 0.5-1.0

HLinium or Niobium -2

Copper .35 3

Cobalt 0,2

Phosphorus 0015

Sullur .02 > 0.35

Curbon 0.04-0.08

lungsten 0.5

Aluminum + Titanium 0.5

“Single values are maximum percentages unless otherwise

specified,
Table 2 Physical Properties of Hastelloy N
R0°F 500°1 1000°F 1300°F - 1500°F
Density, Ib/m,” 0.3209
Density, Ib/ft* 553.0
I hermal conductivity, But he * 7! %! 6.0 7.8 10.4 12.6 14.1
Spectfic heat, B ||\_I T . (0.098 0,104 ().1154 ().136 (.153
Cocttivient of therntal exyansion per 18 WAL 2% {1 (o 207 007" RO T 92597107 Dy
Mg iae ot dlasniorgy iy, T AP T 27 710" 25 7 10" 440"
Prsa il resistanee, suvie i qm | ) 8¢ 123.7 1§ X 26 ()¢ 14 |4
; 125. 126. 124.1
\ \ 9 B altey e tS 1w i ur sy R
¢ g ; ‘ v " | iy ';HI)
' ! i, (Lt ip '
° Ny

2470 2555

Faken direcdy from rer, |0, All other values found from inte

cise information. rpolation of plots of ref. 10 data, See this reference for more pre

Average coefficient of expansion over 212 ¢ 3 iIs 8 X g
‘ an; <1210 1832°F runge is 8.6 X 107 per °FF b
“Ref 11 T T

BY APPROVED AGE 6-43

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO, 8-25-2431 | DOCUMENT No. ND/74/66 | 1getm 1 DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |}

ASME BOILER AND VESSEL CODE
CODE CASE: 1315-2

SECTION VIII

Table 3—Maximum Allowable Stress Values

Metal Temperatures Maximum Allowable Stress Values, psi
Not Exceeding

Deg. F. All Material Other Than Bolting Bolting

100 25,000 10,000

200 24,000 9,300

300 . 23,000 8,600

400 21,000 8,000

500 20,000 7,700

600 ' 20,000 - 7,500

700 19,000 7,200

800 18,000 7,000

900 18,000 6,800

1000 17,000 6,600
1100 13,000 6,000
1200 6,000 , 6,000
1300 ' 3,500 3,500

Tobie 4
Design Stress Intensity Volues, S,
in ksi

Temperature

100 26.7
200 25.3
300 24.2
400 23.3
500 22.6
600 21.9
650 _ 21.6
700 21.3
750 21.1
800 20.7

NOTE: Design stress intensity values are
basedon lesser of:

) % Specified Minimum Tensile Strength

) Y, Tensile Strength @ Temperature

) 2/3 Specified Minimum Yield Strength

) % Yield Strength @ Temperature

ASME BOILER AND VESSEL CODE
CODE CASE: 1345-2
SECTION III
APPROVED 3/9/72

BY APPROVED PAGE 6-44

FWC FORM 172 - L4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

ig

MATERIALS INVESTIGATED A

TABLE §

ND TESTS USED TO DETERMINE RELATIVE PERFORMANCE_

Tensile Properties

Structure

TIAOH4LY

Sv-9 HOV

Materlal Irradiations Uniform Corrosion at 1050 F Corrosion With Salts
6 % 1019 nvt Constantly Stressed 1075'“0(_) P 2
700-800 F Unstressed 0.1% Creep’in 1000 Hrs. 70 wgm. Cl1™ /in.
304 No deleterious Normal scale Some intergranular Intergranular failures,
effects growth stress ruptures stress-accelerated
Incoloy No deleterious Normal scale No failures. normal scale Normal scale formation
effects growth
Inconel Nol tested Normal scale Intergranular penetrations, Intergranular failures,
growth stress ruptures stress-accelerated
310 Not tested Normal scale Normal scale formation Ewbrittlement because of
growth sigma phase, Normal
scale.
316 No deleterious Not tested Not tested Not tested
effects
330 A No deleterious Not tested Intergranular attack Intergranular failure
effects
347 No deleterious Not tested Normal scale formution Not tested
effects
AISI 406 Loss in uniform Large oxide Large oxide layer

Hastelloy-X

Hastelloy-N
Ni-O-Nel

2} Croloy
5 Croloy Ti

elongation from
11. 4% to 4. 4%

Not tested

Not tested
Not tested

Not tested

Not tested

layer formation

Normal oxide
layer formation
Not tested
Normal scale
growth

Not Llested

Not tested

formation

Normal scale formution

Intergranular failure

Normal scale formation

Excessive scale formution

Excessive scale formation

Large oxide formation.
stress-rupture failure

Normal scale formation

Intergranular failure

Normal scale formation

Excessive scale formation

Excessive scale formation

Effect of Exposure

Not tested

Reduction in ductility
to 207. Recovery
after 3000 hours,

Slight changes in
strength and ductility

Not tested

Not tested
Not tested
Not tested

No effect up to
2000 hours

Reduction in ductility
from 427 to 107 in
5200 hours

Not tested

Slowly decreasing
ductility after 2000
hours

Not tested

Not tested

Microstructural
Changes

Not tested

Sensitization and
microstructural
instability

Coniplete
sensitization

Not tested

Not tested

Not tested

Not tested

No visible
changes

Continuous inter-
granular precipitate
network

Not tested -

Intermranular
precipitate net-
work

Not tested

Not tested

|

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"L N ‘NOISONIAIT

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FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

ig

THAO¥dAY

Zovd|

9%-9

' - TABLE6
AVERAGE WEIGHT GAISS FOR ALLOYS EXPCSED TO 15 TORR WATER VAPOR
a1 348, 1050, HwD 1138 °C T - Yo_. CLl ANY SO0 s
Tost Averape Weight Chans2 Tezt Avcrage Weight Crange
Temperature, Time, C k 428 Tempurature, Time, & &' hange,
Alloy C hr mn/cm - Allov C ar mglem
Inconel GO0 815 160 + 0.120% Hastellcy C 815 1G9 +0,1029
210 + 0.1617 209 +0.1194
302 + 0.1742 300 +0.1362
930 100 + 0.3:04 039 1C0 + 0.2730
200 + 0,4832 200 +0.3752
500 + C.%9%5 300 + 0. 4512
1033 150 + 0.9133 1033 1970 +8.5733
2C0 + 11,2831 230 + 0, 6029
300 + 1.%600 : 30 + 0.7:10
Inconel 625 815 100 + 0.1630 Hastzlloy N B15 1C0 + 0. 93531
200 + 0.2244 nc + 0.1023
270 + 0.2620 w50 + 0. 1037
930 16 * 0. ausd 230 160 +0.2370
202 + 0.8730 20 + 0. 2252
350 + 1.1004 300 +0.3133
1038 100 + 2,4505 1038 1€2 +0.5276
209 + 3.3270 200 40,6295
300 + 4.186 200 +0,7650
Inconel 702 815 100 + 0.2556 Hastelloy R-235 315 1G5 +0.7130
200 + 0.3238 260 + 0.£013
360 + 0.48935 300 +0.8343
930 1C0 + 01,1435 930 102 + 2,173
2C0 + 1.5732 222 + 3,0337
36 + 1.87232 300 + 3.5573
1038 169 + 1.7439 KY
20 + 1,521 MRk e Y
300 + 1.7300 300 s 63030
Hogrel 1 L 180 T .2.869 Hastelloy X-28 815 13 + 0. 0840
200 + 0.3832 - 200 by r
300 + 0.4013 gk e
g 560 + 0.0505
930 160 + C.9286 930 Y00 el g
200 + 1.3872 s P
200 + 1.6162 ‘ g; o0t g ks
1998 109 ~ 131208 1038 10) + 06701
200 - 5.0674 59 190, gats
300 - 13.8400 u >
s + 3
L el s i 0.5479 Haynes 25 81t 160 +0.1793
200 + 0.4827
200 +0.2147
300 + 0.5327 38 4 0.2281
o 109 L PIDR 930 100 +0.5122
290 + 1.2563 550 MR-
390 + 1.4634 Eon 0. 7506
1€38 100 + 1.5593 A - s £ 54
. 1038 1C0 4+ 0.8255
200 + 1.8669 :
300 + 2.3700 200 + 0.9939
. 300 + 1.0470

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FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

L

ig

THAOHALY

Lv-9 EOV

AT v ?

WEICHT GAaIN OF IRON- AND NICKE L- maASE Ar v 0
D e L L A S R e 2 SN

T B0 O 130199 Deor 100
AT 550 0C (1922 2 AXND _'-..1()";‘41(‘?‘.' s PRIESSURE

|

Aty
EXTOCURE rvnping 72 e ) iud 4 252 236 340
PiISSURE, psi 1009 3000 3c00 1000 2( 0 30060 Z:JOO 3;;;

ASTM A 212-RB 260.9 305.0 730.0

"L e .

n' ' 3 "1 SS 62.0 55.90 7e.0 4.0 117.0 121.0
\Y-' ~"t\ (141.) 123.0 (160.) i1 6.0 (2¢2.)) 177.0
ATSD 340 €8 3.2 5.3 4.1 7.6 6.3 2.3
!‘ SHRAL 122 0 16.0 ' "5; 0
VNI 2 Ss (AL (155.) 121.0 (126.) i 8.0 (252)) '—f":‘"
ISt 3 NS UCerp.) 5.0 2.2 " Fe

. { ) Wa e2d.u

Fe-2:C1-3.65 A) 2.2 "j’: - 3

Foe-25CK-2 41-0. 6Y 22.0 4(; ‘a. 5

Haste!lv X-220 4.5 6.0 =

Hoeswelloy N 12.0 .

Inconei X L. g

Inncolon A 6.4 l a3 =

PR <oz e.s 1L5 ) 12.4

YS IN SUI'ERHEATED STE

S.0

720 egy
1002 3900
2140
(225.)

5.7
1259.)

1052
3200
220.0**
230.0
11.0
5230
212.0
1900
310.0
{36.)
73.0
57.0
26.0
42.0

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FWC FORM 172 - L,

NOTATTONS IN THIS COLUMN iNDICATE WHERE CHANGES HAVE BEEN MADE

X4

Tzhle B8,

Summary of Deta on the Corrosion of Hastelloy

=
t
3
w
o
(D
W)
=3

THAOHddY

8r-9 ®=OVq

Weiznt Gain {(mz/cm?)

Brvirenment ard Test Conditions
cellev N Type 304L Has-allov N Type 304L
Hesteiles Steinless 3te=l R A “-zinless Steel
Surerhiested sTesn,
1022°F, 2000 psi, 1092 hr .57 2.20
Deciyzeneved stean, o
932°F, 3200 nsi, 2400 wr 0. 08 3.3
m \
022°F, 50500 zsi, 2400 nr 0.C4 0.14
i A . f“r-\ Ay
(< 50 ppb), 1022°F, 3000 psi, 0z C.33
2400 hr :
Oxycenated steam,
/3 tc 4 ppr), 1C22°F, 3000 psi,
2400 hr 0.15 .37
Helium plus 15 torrs water vapor,
0.109

1500°F, 300 nr, ~ 1 atm pressure

*ON HDYVED

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Table 9,

Chemical Analyses of Test Materials

Concentration, wt I

Alloy
Ni Mo Cr Fe Mn [+ 51 1 4 S Cu Co v W Al T B Nb :14 2r Other
Arwco Irom® Bal. 0.017 0,012 0.005 0.025
Low—alloy Perrvitic
1.1 Cr 0.23 0.49 1.1 Bal. 0.42 0.64 <0.05 <0.02 <0.05 <0,05 <0.02
1.9 Cr 0.20 0.54 1.9 Bal. 0.46 0.17
2.0 Cr 0.32 0.88 2.0 Bal. 0.40 0.25
4.2 Cr 0.36 0.47 4.2 Bal. 0.40 0.35
8.7 Cr 0.35 0.97 8.7 Bal. 0.44 G.50
12-5-3 Maraging 12.7 2.80 5.1 Bal. 0.05 0.10 ~0.3
Stalnleas Steels
Type 502° 0.5 5.0 Bal. 0.1
17-7 PH 7.10 17.0 Bal. 0.07 1.15
Type 201 5.23 16.55 Bal. 7.28 0.076 0.5 0.34 0,006 0.059 N
Type 3045 8.0 18.0 Bal. 0.03
Type 309 13.5 23.0 Bal. 0.2
Type 310. 20.5 25.0 Bal. 0.25 1.5
Type 316 13.01 2.8 17.0 Bal. 1.74 0.027 0.65 0.016 0.017 0.10 0.15
Type 321° 10.5 18.0 Bal. 0.08 0.4
Type 247" 11,0 18.0 Bal. 0.08 0.4 0.4 TA
Type 406 13.0 Bal. 0.15 4.0
Type 410° 12.5 Bal. 0.15
Type 646" 25.0 Bal. 0.20 | 0.25 N
N1-280 Bal 0.0002 0.002 ,0.003 <(Q.0001 0.34 . 0.005 <0.001 0.002 <0.0001 <0.0001 0.03 <0,0001 <0.0001 <0.0001
Monel 60 3.5 3.5 0.5 23 0.5
Copper® 0.02 9949
Inconel 600 18.0 14.5 7.0 0.05
Inconel 601 €0.5 23.0 14.1 0.5 0.05 0.25 0.007 0.25 1.35
Inconel 718 53.0 3.0 18.0 0.05 0.50 1.0 5.0
Incoloy 800 jl.3 20.1 46.2 0.84 Q.04 0.38 0,008 0.50 0.24 Q.36
Hastelloy B Bal 27.0 <0.2 5.2 0.96 0.3 0.01 0.48 0.2 <0.05 <0.01 <0.05
Hastelloy C Bal. 16.0 16.0 5.8 0.75 0.48 0.01 1.2 " 0.1 5.0 0.2 <0.01 <0.05
Haastelloy S Bal 14.7 14.5 0.90 0.04 0.007 <0,01 0.22 ' 0.2 0.01 B
Hastelloy W 6§0.0 25.0 5.0 5.5 0.08 1.0 0.30
Hastelloy X 8.6 22.0 19.0 0.64 0,60 0.02 2.0 0.05 0.5 0.2 <0.01 <0.05
Haynes Alloy 25 10.0 0.5 20.0 1.4 1.0 0.1 0.7 0.015 0.01 0.02 Bal <0.02 15.2 0.1 0.02 <0.05
Haynes Alloy 188 2.0 22.0 3.0 1.25 0.15 Bal. 15.0
Rene 62 Bal 9.0 15.0 22.0 0.25 0.05 0.25 1.25 2.5 2.25 0.01
Hastelloy N Modifications
185 Bal 11.0 5.9 3.8 0,46 0.05 0.10 <0.03 <0.1 <0.05 0.91 <0.1 0.98
186 Bal 10.0 5.4 3.5 0.45 0.05 0.09 <0.03 <0.1_ Q.84 0.88 <0.1 <0.05
188 Ral, 13.0 7.3 4.5 0.49 ¢.05 0.15 <0.03 <0.1 0.95 <0.02 1.1 <0.05
231 3al. 12,0 7.0 4.2 0.03 0.05 0.12 <0.03 <0.05 <0.02 <0.05 1.3 <0.05 1.2 Y
232 Bal 13.0 8.0 4.5 <0.02 0.05 .12 <0.03 <0.05 <0,02 <0.05 1.2 <0.05
234 dal. 16.0 7.2 4.0 <0.02 0.05 0.1} <0.03 <0.05 <(,02 <0.05 <0.1 <0.05
216 Bal 11.0 1.0 4,0 0.5 0.05 0.13 <(0.03 1.0 <0.02 <0.05 <0.05 0.5
237 Bal. 12.0 6.7 4.3 0.49 0.05 0.13 <0.03 <0.05 0.04 1.03 <0,1 <0.05
2477 Bal 16.2 7.0 4,2 0.055 0.057 0.047 0.008 0.004 0.01 0.05% <0.01 0.03 0.02 0.03 0.0002 <0.0005 <0.001 <0.001
5065 Bal. 16.5 7.1 4.0 0.55 0.07 0.58 0.005 0.004 0.007 0.05 0.20 0.1 €0.03 <041 0.001 <0.05 <0,1 <0.1
5067 3al. 17.2 7.4 4.0 Q.48 0.06 0,43 0.005 0.007 0.01 0.09 0.30 Q.6 a.01 0.01 0.004
5085 Bal. 17.0 7.0 3.6 0.64 0.06 0.65 0.004 0.003 0.01 0.15 0.20 0.07 0.05 <0.01 0.0G4 <0.05 <0.002
21541 3al. 11.6 7.3 0.04 0.16 0.05 0.017 0.001 0.002 0.01 <0.10 <0.10 1.98 0.03 0.005 0.0007 <0.005
21542 Bal 12.1 7.21 Q.041 0.16 Q.06 0.014 0.001 Q.004 0.01 <Q,10 <0.10 2.05 <0.10 <0.10 0.0005 0.96 <0.005
21543 Bal. 12.4 7.31 0.038 0.08 0.05 0.019 <0.001 0.004 0.01 <0,10 <0.10 <0.10 <0.02 <0.003 0.0002 0.70 <0.005
21544 Bal 12.6 7.3 <0.10 0.13 0.06 <0.03 <0.01 0.003 0.01 <0,10 <0.10 <0.10 <0.10 <0.10 0.0005 0.44
21545 Rlal 12.0 7.18 0.034 0.29 0.05 0.015 0.001 <0.002 0.01 <0.10 <0.10 <0.10 0.02 0.49 7.00007 0.01
21546 Bal 12.3 7.29 0.046 0.16 0.05 0.009 0.001 <0,002 0.01 <Q.10 <0.10 <0.10 0.02 Q.10 0.0002 0.005
21554 Bal 12.4 7.39 0.097 .16 0.065 0.01 0.004 <0.002 0.03 0.002 0.0002 0.35
21555 bal 12.4 7.18 0.065 0.16 0.052 0.008 0.003 <0.002 0.02 0,003 0.0007 0.05
M1566 Bal 16.0 7.5 5.0 0.5 0.06 a.5
68688 Bal 13.8 7.91 4.98 0.52 Q.079 0.38 0.042 <0.002 0.023 0.08 0.013 0.0002 <0,05 7 <0.05 <0,05
68689 Bal 13.7 7.6 4.8 0.47 0.081 .53 0.01 <0.002 0.02 0.075 0.36 0.0002 <0.05 <0.05 <9.05
69344 Bal. 13.0 7.4 4.0 0.56 0.109 0.5 0.001 0.004 0.013 0.06 <0.01 <0.01 0.24 Q0.77 0.00001 1.7 <0.01 0.001
69345 3al. 13.0 7.5 4.0 0.52 0.078 0.5 0.001 0.01 0.02 0.07 <0.01 0.01 0.27 1.05 0.00006 <0.01 0.92 0.3
69641 3al 13.9 6.9 0.30 0.35 0.06 0.02 0.001 0,003 0.01 <0.03 Q.02 <0.03 1.30 0.0001 <0.05 0.70 0.01
69648 lal 12.8 6.9 0.30 0.24 0.04 0.05 0.001 0.003 .02 <0.03 0.10 <0.05 0.92 0.00008 1.9% Q.08 0.02
69714 Bal. 12.4 8.0 0.10 0.35 0.012 <Q.05 0.001 0.004 0.05 0.05 <0.01 0.01 0.17 0.80 0.00001 1.6 <0.01 <0.01
10727 3al. 11.7 1.5 0.05 0.37 0.04 <(Q.05 0.004 0.001 <0.01 <0.01 <0.01 0.01 <0.03 2.1 0.00006 <0.1 <0.01 <Q.01
70785 lal 12.2 1.0 0.16 0.27 0.057 0.09 0.002 0.004 0.02 0.03 0.003 0.003 0.14 1.1 0.002 0.097 <0.003 0.01
70786 Bal. 12.2 7.2 0.41 48 0.044 0.08 0.002 0.01 0.02 0.05 Q,008 0.003 0.13 0.82 0.000% 0.62 0.003 0.06
10787 Bal. 12.5 7.0 0.18 43 0.041 0.09 0.002 0.004 0.02 0.05 0.003 0.003 0.17 0.90 0.0005 0.12 0.77 0.07
70788 Bal, 12.5 7.2 Q.43 0.43 0.027 0.10 0.002 0.004 0.02 <0.02 0.008 0.003 0.18 1.36 0.0005 0.67 0.30 0.01
70795 Bal. 13.0 7.8 0.04 0.63 0.054 0.03 <0.005 0.005 0.003 0.005 0.001 <0.005 0.06 1.49 0.002 0.005 0.42 0.017
10796 Bal. 12.4 7.3 0.05 0.64 0.043 0.02 <0.005 0.005 0.003 0.005 0.002 <0.005 0.10 0.04 0.0005 0.04 0.75 0.024
10797 Bal. 12.5 7.0 0.29 0.38 0.055 0.02 <0.005 0.003 0.003 0.005 0.002 <0.005 0.07 0.59 0.0002 0.98 0.75 0.035
70798 Bal. 12.8 7.5 0.26 0.53 0.038 0.02 0.002 0.005 0.003 0.005 0.003 <0.005 0.07 0.71 0.0001 0.94 0.32 0.011
70815 Bal 12.1 7.8 0.68 0.58 0.053 0.05 0.001 0.004 0.005 0.10 <0.01 <0.01 0,10 0.71 0.002 2.6 <0.01 <0.005
71114 Bal. 12.5 7.14 0.062 0.02 0.05 0.026 0.002 0.005 0.007 0.07 0.07 1.75 0.002 <0.03
71583 Bal. 12.4 7.35 0.13 0.03 0.05 0.055 0.002 0.004 0.015 0.10 0.2 1.44 <0.03
72115 lal 11.9 7.03 0.07 0.07 0.091 0.09 0.003 0.002 <0.001 0.72 0.2
72503 3al. 12.9 6.79 0.32 0.01 0.066 0.01 0.0008 0.002 0.1 1.94 <Q.001
712604 Bal. 11.5 6.69 0.07 0.07 0.09% 0.09 0.003 0.003 <0.001 0.30 0.02
Momiral coanngit!-- 311 p*hers a=e actual analyses
AT1 the alloys up to "Hastell i fi ‘ & i
oy N modifications’ i
y were commercial production heats. Heats 185 through 237 were

2-1b Taboratory melts.

melted and fabricated by commercial vendors.
commercial heats of standard Hastelloy N.

Hastelloy N heats 21547

ATl alloys were rolled to 0.035-in.-thick sheet.

relief; the finish of the exposed sam
long x 0.035 1in.
ferritic steels),

X 2

in.

through 72604 were small

50-

to 100-1b melts that were vacuum
Hastelloy N heats 2477, 5065, 5067, 5085 and M/566 were large

The rolling was done cold with intermediate anneals for stress
of cold-rolled sheet.

. ple surface was generally typical
thick were sheared,

cleaned, and annealed in argon for 1 hr at 1700°F
1 hr at 1900°F (the stainless steels), or 1 hr at 2150°F (all other alloys)

Samples

(the low-alloy

1/2

in.

wide

Page 6-49
FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

ig

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Table 10, Weight Change Data for Specimens of Low-Alloy Ferritic and Maraging Steels
Weight Gain, mg/cm2 at Various Times in hr
Area
Alloy Specimen (cm?) :
670 1000 2000 2482 4000 4482 6000 8000 10,000 13,000 14,000
Cr 1.128 172 13.6924 3.14 3.78 5.12 6,00 6.47 6.80 9.12
173 13.6444 3.21
205 13.6221 2.31 4,22 4.96 5.37 5.60 6.22 7.74
cr 1.92 202 13.5654 3.97 6.30 7.98
208 13.6392 4.28 7.95 8.53 8.89 8.89 g.78 11.71
163 13.6274 4.94
164 13.6355 4,93 5.98 7.57 8,90 9.31 9.81 12.03
Cr 2.02 166 13.6369 3.93
167 13.6287 3.40 4,32 5.80 6.93 7.37 7.63 9.60
203 13.563%8 2,90 4,63 5.91
209 13.5759 2.73 4,51 5,77 6.08 6.53 7.31 8.96
Cr 4.28 204 13.5999 3.76 6.14 7.74 8.07 8.64 9.97 11.88
169 13.6329 3.98 5.13 6.83 7.97 8.48 8.96 10.41
170 13.5935 4.49 6.89
cr 8.7° 175 13.6411 3.77
176 13.6668 3.73 4,56 5,98 7.03 7,43 7.68
b 206 13.6558 2.74 5.81 6,27 6,27 6.57 7.29 8.79
12-5-3 207 13.5928 3.97 6.61 8.91 9,73 10.22 11,33 13.99
178 13.5173 6.33 )
179 13.4833 6.16 7.37 9.48 11.04 11.77 12.13 15.48

99/%L /AN *ON LNEWADOQ

aLow~alloy ferritic steel annealed 1 hr at 1700°F in argon.

b12—5—3 maraging steel annealed 1 hr at 1500°F in argon.

Y

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NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

& Q
Table ]} . Weight Change Data for Specimens of Stainless Stecls E
ca)
- 2
Stainless Weight Gain, mg/cm? at Various Times in hr .
. , & Area

Steel Specimen Condition (em?) %
Tvpe - 1000 2000 3000 4000 5000 N
i
% 502 372 Annealed 13.7092 3.71 4,77 5.43 6.05 E
HJ 502 373 Annealed 13.0605 3.29 4,23 4,95 5.49 —

3 17-7PH 374 Annealed 13.9202 0.50 0.66 0.83 0.98
B 17-7PH 375 Anuealed 13.8598 0.33 0.51 0.67 0.84 =
201 352 Annealed 13.5782 0.71 1.16 1.81 2.20 Q2

201 353 Annealed 13.5727 0.81 1.35 2.08 2.49

201 354 Cold worked 50% 13.6137 0.04 0.04- 0.07 0.04
201 355 Cold worked 50% 13.7611 0.03 0.03 0.05 0.03 -
304 349 As received 13.6554 0.79 1.11 1.25 1.47 1.64 o
309 359 Annealed 13.7911 1.60 2.05 2.53 2.72 Z
309 360 Annealed 13.4095 1.69 2.06 2.48 2.67 N
310 361 Annealed 13.5377 0.57 0.91 1.10 1.20 a
310 362 Annezaled 13.5377 0.83 1.10 1.29 1.40 x

316 363 Annealed 13.5893 1.34 1.71 2.07 2.39

316 364 Annealed 13.1245 1.44 1.87 2.27 2.61
316 365 Cold worked 50% 13.5512 0.52 0.66 0.85 1.00 —
321 366 Annealed 13.4699 0.75 1.05 1.39 1.69 2
321 367 Annealed 13.4320 0.67 0.93 1.22 1.47 g

347 334 Annealed 13.2688 0.67 1.05 1.17 1.33 1.42

347 335 Annealed 13.3337 -0.55 0.78 0.92 1.09 1.19 —

406 368 Annealed 13.7580 1.56 1.92 2.28 2.51

S 406 369 Annealed 13.6301 1.25 1.54 1.86 2.09
= 410 336 Annealed 1700°F 13.7768 2.39 2.80 3.03 3.48 3.80 |B
= 410 337 Annealed 1700°F 13.5450 2.53 3.03 3.22 3.68 4,02 |8
o L6 370 - Annealed 12.8269 1.03 1.30 1.58 1.77 -
(:n 446 371 Annealed 14.0554 1,52 1.82 2.03 2.16 g
~ e
aAnnealed 1 hr in argon at 1900°F unless otherwise specified. I~

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FWC FORM 172 - )4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J,

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

Table |2. Weight Change Data for Specimens of Nickel 280
Annealed 1 hr in Argon at 1470°F

Weight Gain, mg/cm? at

, ‘ Area
Specimen (cm?)
¢ 400 hr 1518 hr
266 13.8262 4.44 54.43 -
267 13.8262 3.33 77.11
268 13.9680 4.30 72.95
269 13.7880 ' 5.25 75.29

BY APPROVED PAGE 6-52

WC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

ig

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:
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=
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Table 3. Weight Change Data for Specimens of Various Metals and Allovs 7
| IN
. e (%)
: Weight Gain, mg/cm? at Various Times in hr
: , ... 8 Area
Material Specimen Conditjion (em?) : S
o 1000 2000 3000 4000 5000 6000 9000 |Q
Armco Iron 356 Annealed 1700°F 13,6683 4.44  5.52  7.06 8.06 %
357 Annealed 1700°F 13.6241 4.33 5.37 6.88 7.87 -
358 Cold worked 50% 13.7063 4.66 5.68 7.22 8.24 o
Monel 332 Annealed 1470°F 13.6301 0.62 2,81 7.89 12,96 17.28

333 Annealed 1470°F  13.6301 1.19  4.59 9,50 14.42  18. 54 S
Copper 466 Annealed 1470°F 13.5783 0.34 —0.13 ~
467 Annealed 1470°F  13.5554  0.22 —0.18 =
Inconel 600 388 Annealed 2150°F 13.5130 0.26 0.30 0.41 0.46 o

389 Annealed 2150Q°F 13.5265 0.24 0.27- 0.37 0.43
Inconel 601 316 Annealed 2150°F 13.7337 0.06 0.11 0.17 0.17 0.19 0.25 0.36 —
317 Annealed 2150°F 13.7611 0.07 0.12 0.20 0.22 0.24 0.31 0.40 |
468 Cold worked 50%  13.7546  0.21 g

Inconel 718 390 Annealed 2150°F 13.5644 n.05 0.01 0.01
391 Annealed 2150°F 13.7195 0.05 0.01 —0.01 -

412 Cold worked 50% 13.3954 0.12 0.13 0.15 0.18
a =
Annealed 1 hr in argon at the indicated temverature, =
ho
T
o
S
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o
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.
E% Table ]14. Weight Change Data for Specimens of Incoloy 800 and Hastelloys B and C 'k
Do
E%; Weight Gain, mg/cm? at Various Times in hr
. .. . _a Area -
Alloy Specimen Condition ( cm? S
1000 2000 4482 6000 8000 10,000 13,000 %
Incoloy 800 198 Annealed 1900°F 13.6339 0.50 0.80 0.95
199 Annealed 1900°F 13.6200 0.41 0.71 0.81 0.81 0.84 0.90 0.92 i
200 Annealed 1900°F  13.6350 0.42 0.68 0.81 0.71 0.75 0.83 0.80 e
201 Annealed 1900°F 13.5310 0.51 0.81 =
469 Cold worked 50% 13.8543  0.17 3
Hastelloy B 376 Annealed 2150°F 12.6206 0.10 0.11 0.16 0.19 g
377 Annealed 2150°F 13.6480 0,12 0.16 0.21 0.25 A
407 Cold worked 507% 13.3631 0.16 0.29 0.49 0.62
Hastelloy C 378 Annealed 2150°F 13.52Z41 0.06 0.08 0.10 0.13
379 Annealed 2150°F 13.5621 0.04 0.06 0.07 0.08 —
408 Cold worked 50% 13.6220 0.15 0.17 0.21 0.21 g
®Annealed 1 hr in argon at the indicated temperature. -
HJj
’ -
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& ~
. S~
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lus] Q
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b
3
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Table 5. Weight Change Data for Specirens of Various Superallovs é
o
e ““ &
i - . . Area Veight Gain, mg/cm? at Various Tires in hr —
= Alloy Specimen Condition (em?)™
5 i 1000 2000 3000 40600 o4
9
Hastelloy S 417 Annealed 13.7229 0.09 0,14 .15 %
418 Annealed 13.7229 Nn.09 0.12 C.15 _
Hastellov W 380 Annealed 12,8533 0,04 0.05 0.09 0.10 -
381 Annealed 13.2794 0.04 0.06 0.09 0.07 o
409 Cold worked 507% 13.3134 0.11 0.16 0.20 0.23 %
Hastelloy X 382 Annealed 13.8476 0.06 0.06 0.09 0.11 A
383 Annealed 13,7953 0.04 0.04 0.05 0.05 Ny
410 Cold worked 50% 13.7063 0.12 0.14 0.17 0.21 &
Havnes 25 384 Annealed 13.4209 0,03 0.01 0.10 0.16
385 Annealed - 13,5889 0.07 .08 0.18 0.23
- Havnes 188 386 Annealed 13.2926 0.03 0.01 0.05 0.08 H
387 Annealed 14.0295 0.04 0.04 0.07 0.11 7
413 Cold worked 50% 13.76R84 0.23 0.27 0.30 0.33 é
Rere 62 392 Annealed 13.6678 0,18 0.16 0.19 0.22
393 Annealed 13.8791 0.15 0.14 0.17 0.18 —
411 Cold worked 13.7580 0.15 0.16 0.20 0.24
e nneals are for 1 hr in argon at 2150°F, ‘{3
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Table 6. Weight Change Data for Specimens of Modified Hastelloy N Laboratoryv Heats? .
[0 ]
A : o
Weight Gain, mg/cm? at Various Times in hr i
, Area . o
Heat Specimen (em?) : ~
E cm 1000 2000 4482 6000 8000 10,000 15,000 |
e _
é’ 185 87 13.6149 0.14 =
185 88 13.6431 0.21 0.29 0.49 0.54 0.59 0.65 0.74 |9
185 89 13.6603 0.20 0.32 é
185 90 13.6048 0.12 0.24 0.41 0.55 0.53 0.58
186 91 13.5903 0.21 0.24 o
186 92 13.6285 0.12 0.21 0.36 0.49 0.48 0.59 o
188 37 13.6213 0.29 0.40 0.62 0.57 0.68 0.71 0.78
188 38 13.5819 0.09 0.19 0.32 0.32 0.40 0.46 =
231 35 13.6267 | 0.12 0.21 0.28 0.29 A
231 36 13.6386 0.26 0.43 Yy
231 105 12.5198 0.27 0.38 0.58 0.62 0.65 0.69 0.72 |&
232 103 13.5416 0.18 0.22 0.39 0.43 0.44 0.47 0.55
232 104 13,5230 0.21 0.24 0.39 0.46 0.47 0.50 '
236 101 13.642¢ 0.11 0.17 0.35 0.41 0.45 0.52 0.60 |=
236 102 13,6070 0.18 0.26 0.43 0.46 0.49 0.54 é
237 99 13.6422 0.10 0.15 0.24 0.28 0.29 0.33 0.38
237 100 13.6359 0.10 0.14 0.23 0.32 0.26 0.32 —
;g ®Annealed 1 hr in argon at 2150°F. =
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Table {7 . VWeight Change Data for Specimens of Standard Hastelloy N Large Commercial Heats
. Weight Gain, mg/cm? at Varicus Tises in hr

. N Thicrness Area b [,

Heat Specimen Condition . 2 -
{in.) (c7) 1000 2000 4482 6000  BOOO 0,000 15,090

2477 23 0.033 13,6662 0.15 0.21 0.34 0.3 0.38 0.41 0.48
2477 24 0.03% 13.5224 0.13 0.25
2477 55 0.C10 12,6177 0.07 0,12 0.22 G.26 0.23 0.27 C.24
2477 56 0.010 12,7622 0,09 0.14% 0.4 0.29
2477 57 0.010 12,6151 0.10 0.17
2477 58 0.010 12,4519 0.67 0.13 0.18 0.26 0.22 0.26
2477 59 0.010 12.7940 0.47 0.09 0.25 0.26
2477 60 0.01% 12.6220 0.11 0.13
2477 61 Cold worked 507 0.010 12,6409 0.13 0.15 0.27 0.31 0.35 0.42 0.50
2477 62 Cold worked 507 0.010 12.9948 0.15 0.13
2477 63 Cold worked 50% 0.01C 12.9303 0.13 0.20 0.33 0.39
2477 64 Cold worked 50% 0.010 12.5853 0.10 0.17 0.28 0,36 0.36 0.45 0.52
5065 1 0.010 12,5543 0,14 0.20 0.29 0,32 0.29 0.2
5065 2 0.010 12.5932 0.10 0,14 0.33 0,28 1 0.30 0.23 0.35
5065 3 0.020 12.9448 0.15 0.20 0.30 0.32 0.35 0.37 0.41
5065 4 0.020 12.9358 0.14 0.22 0.33 0.35 0.33 0.36 0.41
5065 5 0.035 13.4914 0.16. 0,22 0.33 0.33 0.33 0.34 0.33
5065 6 0.035 13.5692 0.15 0.22 0.34 0.35 0.37 0.41 0.45
5065 7 0.060 4.5614 0.14 0.24 0.36 0.40 0.41 0.45 0.55
5065 8 0.060 14,5693 0.14 0.21 0.33 0.36 0.38 0.42 0.47
5065 9 0.035 13.5302 0.1 0.22 0.33 0.33 0.33 0.34 0.38
5065 10 0.035 13.6351 0.15 0.22 0.31 0.32 0.37 0.40 0.43
5065 11 Abraded 0.035 13.5960 0.29 0.50 0.56 0.60 0.63 0.70 0.81
5065 12 Abraded 0.035 13.5079 0.25 0.39 0.55 0.60 0.64 0.69 0.78
5065 13 Electropolished 0.035 13,3979 0.04 0.08 0.13 0.16 0.14 0.18 0.22
5065 14 Electropolished 0.035 13.2262 0.03 0.09 0.15 0.15 0.18 0.22 0.29
5065 15 0.010 12.5496 0.08 0.37 0.29 0.29
5065 16 0.010 12,5359 0.07 0.18
5065 17 0.010 12.5742 0.14 0,22 0.34 0.32
5067 18 0.035 13,3301 0.14 0.22 0.30 0.31
5065 19 0.010 12,4907 0.10 0.22 0.26 0.34
5067 20 0.035 13.3419 0.12 0.21 0.30 0.30 0.31 0.35 0.40
5067 67 0.035 13.2601 0.12 0.20 '
5067 68 0.035 13,1974 0.06 0.11 0.23 0.28 0.26 0.32
S08S 21 0.033 13.6078 0.12 0.23 0.32 0.32 0.32 0.37 0.44
5085 22 0.035 13.5278 0.13 0.15 0.34 0.34
5085 65 0.035 13.7014 0.11 0.17 0.34 0.36 0.37 0.43
5085 66 0.035 13.5821 0.13 0.18
M1566 394 0.035 13.6297 0.09 0.09 0.15P 0.18¢
1566 395 0.035 13.6678 0.11 0.12 0.17b 0.22¢

aUnless otherwise specified, annealed 1 hr at
b3000 hr.

c
400C hr.

2150°F in argon

s tested with the surface in the as-rolled

condition.

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o v
2
g Table [8. Weight Change Data for Specimens of Modified Hastelloy N Comnercial Fertg@ E
| 2
T T . 2 T =
Area elght Gain, mg/em® at Various Times in hr o
Heat Specimen (cm?) ) ‘ *
1000 2000 4482 6000 8000 10,000 14,000 o
——— 1
21545 27 13.4661 N.19 o
21545 28 13.4492 0.25 0.34 0.42 0.43 0.43 0.48 s
21545 97 13.3940 0.13 0.22 5
B 21545 98 13.4485 n.15 0.19 *0.34 0.42 -
i 21546 25 13.5637 0.23 0.31 0.40 0.37 0.41 0.44
3 21546 26 13.5364 0.18 , =
g 21546 95 13.6380 0.23 0.29 0.40 0.45 &
21546 96 13.5981 0.25 0.31
21554 83 13.5252 0.21 0.31 0.49 0.59 %
21554 84 13.5208 0.22 0.33
21554 85 13.5629 0.19 0.28 0.43 0.56 N.52 0.60 -
21554 86 13.5674 0.27 o
21555 79 13.4514 0.25 -
21555 80 13.5829 0.13 0.22 0.35 0.46 2
21555 81 13,5219 0.11 0.18 0.29 0.35 0.36 0.41 ~
21555 82 13.4692 0.29 n.39 z
68688 75 13.6022 0.12 0.18 R
68688 76 13.5285 0.10 0.14 0.24 0.33
68688 77 13.5245 0.04 0.13 0.24 0.30 0.27 0.33
68688 78 13.5056 0.10
68688 194 13.1127 -0.01 0.13 0.25 0.19 0.19 0.24 0.26P e
68689 71 13.2735 n.12 0.18 0.31 0.41 g
68689 72 13.5373 0.10 0.18 0.30 0.41 0.40 0.44 =
68689 73 13.4412 0.13 0.19 —
68689 74 13.4475 0.14
68689 195 12.9800 0.03 0.13 0.25 0.17 0.18 0.26 0.29°
69641 160 13.4860 0.10
= 69641 161 13.5752 0.09 0.12 0.24 0.22 0.25 0.39 0.40 2
b 69641 196 13.5937 0.04 0.17 0.18 0.21 0.27 0.32 0.38 =
& 69648 157 13,5881 0.12 0.14 0.26 0.34 0.32 0.36 0.52
o 69648 158 13.5307 0.10 , o
(fn 69648 197 13.6013 0.03 0.18 0.26 0.24 0.29 0.36 0.45 =
oo —_—— s (o)
a —~
Annealed 1 hr in argon at 2150°F =

T a7 et 2

b13,000 hr.

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Table 19 . Weight Change NData for Specimens of Modified Hastelloy N Commercial Heats®
Veight Gain, mg/cm’ at Various Times in hr
Heat Specimen Area
: (cm?) .

1000 2000 3000 4000 5000
21341 314 13.6926 0.12 0.14 0.19 0.22 0.18
21541 315 13.7063 0.11 0.18 0.20 0.23 0.20
21542 312 13.6926 0.02 0.06 -~ 0.09 0.13 0.12
21542 313 13,8022 0.10 0.12 0.16 n.17 N.16
21543 310 13.7611 0.09 0.12 0.15 0.15 0.14
21543 311 13.7748 0.12 0.14 n.17 0.19 N.14
21544 308 13.7200 0.24 N.35 0.36 0.31 0.19
21544 309 13.7959 0.17 0.26 0.31 0.36 "0.28
21545 343 13.4363 n.19 N.25 0.27 0.31 n.31
21546 344 13.6956 0.27 0.34 0.34 0.39 0.41
21554 345 13.6301 0.13 n.17 n,20 0.27 0.31
21555 346 13.3827 0.07 - 0,12 0.13 0.19 0.25
70727 342 14.1173 —0.05 —0.04 —0.05 -—0.03 —0.03
70785 306 13.7337 .05 0.11 0.17 n.23 0.18
70785 307 13,7063 0.12 0.12 0.18 0.20 0.21
70786 304 13.7063 0.07 0.11 0.16 0.20 0.20
70786 305 13,7474 0.09 0.N9 0.14 0.15 0.17
/0787 302 13.7337 n.10 N.15 0.24 N.25 0.27
70787 303 13.6926 0.07 0.15 0.23 0.26 0.23
70788 300 13.7337 n.N09 0.10 0.15 0.15 0.17
70788 301 13.7885 0.01 0.01 0.06 0.05 0.07
70795 298 13.6515 Nn.03 N.08 N.14 0.20 0.18
70795 299 13.6789 0.01 0.06 0.08 0.16 0.16
70796 296 13.8323 0.06 0.10 0.17 0.20 0.15
70796 297 13.8296 0.09 0,12 0.17 0.23 0.13
70797 294 13.7666 0.03 0.04 0.09 0.14 0.12
70797 295 13.7611 0.04 0.07 n.10 0.13 0.13
70798 - 292 13.7081 0.03 0.03 0.09 0 0.09

e £ R e LI i £ e B i s R+ 1eh

6000 9000
0.20 0.26
0.20 0.30
0.12 0.18
0.1¢9 0.21
n.15 0.22
n.17 0.24
N.23 0.36
0.27 0.41
0.19 0.30
0.23 0.26
0.20 0.33
0.21 0.33
0.28 0.44
0.25 0.38
0.18 0.30
0.08 0.18
0.18 0.27
0.16 0.28
0.13 0.25
0.12 0.28
0.12 0.19
0.12 0.20
0.09 0.16

*ON HDYVHD

1ev2-5¢-8

"ON LNHEWNO0C

99/%L /AN

HASST

1

HLV{T

wL/91/21

"L "N ‘NOISOHNIAIT

LNAWIIVdEId ¥VA'TIONN

R e N T

[ PN e L A N N T AT FTATIT AT e ey
FWC FORM 172 - )4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE — —

g

THACHAAY

Takle {9. (Continued)

——— T T e T T e ————

Area “eight Gain, mg/cm? at Various Times in hr
Heat Specimen (em?) — T
1600 2009 3000 4000 5000 6000 9000
70798 293 13.6877 0.04 0.08 0.12 0.16 0.15 0.15 0.25
70835 290 13.9187 0.05 0.02 . 0.06 0.08 0.07 0.07 0.09
70835 291 13,7819 0.04 0.04 0.09 0.09 0.09 0.08 0.14
71114 338 13.7502 0.10 0.17 0.17 0.23 0.28
71114 339 - 13.4699 0.10 0.15 0.15 0.20 0.23
71583 340 13.6410 0.08 0.11 0.13 0.20 0.25
71583 341 13.5512 0.08 0.13 0.15 0.21 "0.26
72115 396 13.7987 0.18 0.23 0.34 0.44
72115 397 13.7606 0.17 0.20 0.31 0.41
72115 414 13.7200 0.18 0.29 0.42 0.51
72503 398 13.8469 0.12 0.14 0.22 0.27
72503 399 13.9060 0.11 0.12 0,20 0.27
72503 415b 13,7870 0.04 0.02 0.02 0.04
72604 400 14.0059 .39 0.40 0.47 0.54
72604 401 13.9300 0.34 0.34 0.39 0.46
72604 416 12.4655 0.15 0.23 0.36 0.44

aUnless otherwise specified, annealed 1 hr at 2150°F in argon.

bCold worked 50%, /

[€%2-57-8 *ON "DavHD

ININIMVIEd Jvd100N

99/%L/AN *ON INIWADOT

NOILVEOJdd0D AD¥UANA YTATAHHM ¥HISOI

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YI79172T Gova
"N ‘NOILSONIAIT

*r
L

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE 3
P
2 5
|_4
3
=
S
Qo
i
N
7
- IN
% feble 20. Weight Change Data for Scecimens of “odified Hastellov N Commercial Heats® W
4"
2 T o
@ N eight Gain, mg/cm? at Various Times in hr pot
Heat Specimen (reg) - —_—
' cm 1812 3330 4330 5330 6330 8330 12,330
69344 258 13.6713 0.21 0.17 0.19 0.21 0.23 0.22 0.26 |3
69344 259 13.67N8 N.12 0.20 0.24 0.21 0.25 0.25 n.28 |°
68345 260 13.6722 0.18 0.13 0.16 0.20 0.26 0.21 0.26 =
659345 261 13.6106 0.15 .20 0.19 0.19 0.23 0.21 0.26 I3
69714 262 13.6283 0.16 .22 0.20 0.21 0.23 0.24 0.31 =
69714 263 13.6248 0.16 n.20 0.18 0.21 .30 0.28 0.34
70727 264 13.6558 .19 0.24 0.27 0.29 0.39 0.37
70727 265 13.6675 0.19 0.20 N.18 0.20 0.27 0.26 0.33
| o
a o .
Annealed 1 hr at 2150°F in argon. é
—t
: ;
3
Oh S
I s
() i
. [o)]
[~~~
Y
N

INIWILIVdEQ dvdETONN

"I N ‘NOISONIAIT

NOILVIO4Y0D AO¥HENA YHTIHHM YHLSOL
FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT | _ LIVINGSTON, W. J.

CHARGE NO. 8-25-24131 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

TABLE 21

SUMMARY OF TUBE BURST TEST DATA

{Machined

Date of 5 Heat Stress DiametralgExposure {FailureiTube Walil
Placement (1)! No. | (PSI) é Strain | Time (Hrs) |Thickness
§ § ‘ % (Hrs) (Inches)
2/28/71 5N25095*fi7,ooo - 1,000 1 0.010 |
2/28/71 N25095%52, 500 - 1,000 3.7 0.015
2/28/71 N15095 140,250 0.71 5,000 None -
2/28/71 N15095 28,000 ; 0.19 | 5,000 None | 0.0302%
8/31/71 N25101 56,000 0.57 | 4,000 None | 0.014
8/31/71 N25101 ;50,000 I 0.33 4,000 None | 0.016
2/29/72 N15095;58,000 f - 1,000 792% | -
2/29/72  IN15095153,000 | 0.25 - 1,000 None | -
2/29/72 ‘N15095I42,502 } 0.24 | 1,000 None | -
2/29/72 N15095 (42,400 |  0.14 | 1,000 None | -
(2) - 72,0004 - - - | 0.0108%
(2) - 566,000 - 1,000 4.0 -
| ;
(2) - 56,000 - 1,000 27.4 | -
(2) - 155,300 - 1,000 99.7 1 -

* Date Inferred

(1) ORNL 4676 Molten Salt Reactor Program, Semiannual
Progress Report Period Ending, 2/28/71.
ORNL 4728 Molten Salt Reactor Program, Semiannual
Progress Report Period Ending, 8/31/71.
ORNL 4782 Molten Salt Reactor Program, Semiannual
Progress Report Period Ending, 2/29/72,
ORNL 4832 Molten Salt Reactor Program, Semiannual
Progress Report Period Ending, 8/31/72.

(2) Between 2/29/72 and 8/31/72.

BY APPROVED | PAGE 6-62

FWC FORM 172 - L

NOTATTONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

*

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUR

1 DATE 12/16/74

Most Stable
Fluoride
Element Compound Stability*
chromium CrF, 72
Structural Metals iron FeF. 66
nickel NiF. 59
molybdenum MoF, 57
lithium LiF ' 120
sodium NaF 110
Carrier Salts potassium KF 108
beryllium BeF, 103
zirconium ZrF, 92
boron BF, 86
uranium UF, 92
Active Salts UF, 93
thorium ThF, 99

"Negative standard free energy of formation (@ 800°C,

Table 22 Relative thermodynamic stabilities of fluoride compounds.

BY

APPROVED

PAGE 6-63

FWC FORM 172 - |

NOTATTONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO.

8-25-2431

DOCUMENT NO. ND/74/66

I85UE

1

Laop No,

1258

NCL-TIA

NCL-14

NCL-15A

NCE-THA

NCL-TYA

NCL-20A

NCT-21

NOL-22

Table 2 &. Status of MSR program thermal-convection loops through

[ oop naterial

Type 304

stainless steel

Hastelioy N

Hastelloy N

Hastelloy N

Hastelloy N

Hastetioy N

Hastelloy N

Hastelloy N

[lasiclloy N

Hasteltoy N

Hastelloy N

ilastelioy N

Type 3G

statndess steel

Specihmens

Type 41 stanless

filL‘(‘l”'h

Hasetloy N3 T
modified Hasteloy
N wnluflsh‘(’

Timaodified Hastelloy

) .
N
Fi-maodaned Hastelloy
N Hastelloy N
Fl,(‘
controls
Tr-modilicd Hastelloy
N; Hastelloy N
controls e

ffastelioy N, one Te-

Hastelloy NoTi
modificd Hastelloy
N cnnlml.xh'(

I’ .
[lasteloy N He

Hastelloy NoTi-
modilicd Hastelloy
)
N controls™
[Lastelloy NoT-
modhiticd Hastelloy
he
N confroly”
Jastelioy Ng Tt
modified Hastelloy
)L
N cantrols

Hastelioy N™¢

Type 36 stndess
b
steel

ho
coated Hastelloy N

Salt type

Coolunt

Cooland

Rlankct

1ruet

l‘uel

Coolant

tertite-fissle

Fertde-fissile

Cuooliant

Coolant

MSRI fucl

Fertile-fissil

Lit-Beb 92l 4-Ul 4 ThY'y

Salt
compositionp
{mole %)

(70-23-5-1-D)
NaBil 4 -Nat (92-8) plus
tritium additions

NaBl 4-Nal- (92-5)

sugust 31, 197 2

Lil*-Bel 3 Thi g7 3-2-25)

Lit-lel - Ulig
(65.5-34.0-0.9)

Lil'-Bel-3-Ul
(65.5-34.0-0.5)

Nalily-Nal* (92-%)
plus steam additions

l,ii:-l)-t‘I“j“l 111‘.4 ! H‘-‘;
(6R-20-11.7-0.%1)
Lil-Bel - Thl 3 Uy

(68-20-1 1 7-0.1 plus
hisminth momelybdenum

hot Hneer

Naltl a-Nab (92 8)
NaBl4-Nal" (9.-%)

Al Bel  Zrvy Uty
(65 4-29.1-5.0.0.5)

LilBel - Uit Ut
(68-20 11.7-0.3)

-~

DATE 12/16/74
T YA i
coy (O (hn)
688 100 79.367
607 125  33.579
607 150 42,154
677 55 15416
204 170 37,9424
704 170 1,729
607 o 27,817
704 170 672
704 170 22.203
687 250 19.928°
687 250 LoH2
(56 116 G757
650 110 342

BY

APPROVED

PAGE 6-64

FPCPLP—.

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

WHERE CHANGES HAVE BEEN MADE

1200
l_) )“

I.,tln
1350

1 1)

IXATRIR
{350

1 2011

- I ) ‘l ,_,_,,,§ - .

1300
15540
]].Hll

LIVINGSTON, N. J.
CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 |Issum 1 |DATE 12/16/74
Table 24 A.
S, — Maximum Allowablo Stress
fntensity (ksi)
(For Design Condition Calculations)
Ni-Fe-Cr
Alloy 800H
{Solution
Temp.  F 304 SS 31655 Anncaled) 2% Cr-1Mo
R T N - R
750 15.0
SO0 i5.1 15.4 5.3 15.0
50 110 15.7 5.1 Jobd
900 16 5.5 1.8 13.1
951 11,3 15.1 11.6 1.0
1000 13.7 15.3 14.4 7.8
FO51 12.1 115 P 5.8
[ 100 9.7 o 1.5 .2
L1540 .7 04 - 1.2 3.0
[ 200 0.0 7.1 8.1 1.6
1250 1T 5.1 0.9
1300 3.7 £ 5.4
A5 24 3.0 1.5
1100 23 2.2 3.0
1150 1.8 1.7
15360 i |2
Table 24 B.
Mi-Fe-Cr {Alloy 800H), \mr Allowable Stress Intensity Values, ksi
F - CTTTTTTETIT T o T ;* - "'T"”f - B T A
‘_1 hr  10hr  30hr {100 hr FDD hr L 1000 hr } 3000 hr | 10,000 hr | 30,000 hr |100,000 hr | 300,000 hr
I 15.1 il 15 EER 3.1 | 151 T 15.1 15.1 15.1 15.1 15.1
148 18 158 0 LR { IR A B § 11.8 118 14.8 14.8 14.9
B R N A 1.6 116 0 1Lo |1 S TR R 146 14.6 11.6 14.0
} 111 L b1t ! j [ S O 1 I S O | 14.1 144 14.4
x 11,3 I Pes 143 5 1RD L4 14.3 11.3 14.3 14.3 13.4
L S R o I S R NN A I Y 1L L4 1 14.1 13.6 11.7 10.3
; 1349 2.0 15.9) 13.9 13,9 129 139 120 10.5 9.1 8.0
‘ 13,65 3.0 tsn 136 ° 1.8 125 1o 9.1 8.2 7.2 6.4
e B i_;.."»_‘ LR L3S ‘l_ !,._1__ 0.8 8.0 ) 0.0 5.8 5.1
13.2 152 121 .5 g T 6.9 6.1} 5.3 1.6 4.1
b2 R H.9 8.1 ! 6.1 5.0 1.9 4.3 3.7 3.3
: 1.0 9.2 ) 0 6O 3.2 L6 1 4D 3.5 3.0 2.6
~ Table 24C
Ni-Fe-Cr (A on SOOH) I — Allowable Stress lnten5|ty Values ksi
1'--__ — «;7— — :7 - T - T T W‘ - '
' thr { 10hr | 30hr | 100 hr? 300 hr T 1000 hrg 3,000 hr | 10,000 hr | 30,000 hr | 100,000 hr | 300,000 hr
L2000 2o 2000 200 } 20002000 20.0 20.0 200 | 20,0
108 1 1 o 10.8 s L1900 198 19.8 19.8 9.7 19.6
[ KETRS T RS RTINS TN VO B O I 190 19.6 1 196 19.5 19.3 19.2 y 19.1
U R T A TS S TS U 193 191 18.9 18.6 185 170
S L T S A T 1y 0 193 19.3 150 +  18.7 8.1 17.1 15.3 R EX
N B PO 180 8. 130 i 178 15.7 13.6 L7 10.3
B T E e T 163 b 10 12.0 103 9.1 8.0
IO B 173 160 117 125 1 109 0.1 8.2 7.2 6.4
A I L7 138 0 133 115 og o6 73 4 66 58 51
6 1 o115 12,1 L3 9.1 T 6.0 . 6.0 5.3 1.6 1
S 1t 0.1 0.1 T £ 5.0 5 1.9 1.3 . 3.7 3.3
126 2 5.0 0.3 fi.0) S0 1.6 1.0 35 3.) 2.6

%PAC—E 6-65

FWC FORM 172 - L4

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74
TABLE 25
PERFORMANCE OF HASTELLOY N SPECIMENS IN
HIGH PRESSURE CHLORIDE SCC FACILITY*
Specimen Type -Specimen No. of No. of
Alloy Weld |Non Weld Condition Specimens|{ Failures
Hastelloy N v - AR(l) 3 3
Y _ Ground(2)
Hastelloy N Annealed 3 0
Hastelloy N - v Ground 3 3
Hastelloy N! - v ar (D) 3 3
Hastelloy N - v Anneal 1(3) 3 0
Hastelloy N - v Anneal 2(3) 3 3
Hastelloy N - v Anneal 3(3) 3 0
Hastelloy N - v Anneal 4(3) 3 0
*Compiled from data reported in Ref. 39
(l)AR - As Received, Solution Annealed 2150°F.
(Z)Annealed in accordance with supplier of alloy.
(3)Anneal l - Ground and heated 1 hour at 2150°F; air
cooled.
Anneal 2 ~ Ground and heated 10 min. at 1600°F; air
cooled.
Anneal 3 - Ground and heated 1 hour at 1600°F; air
cooled.
Anneal 4 - Ground and heated 6 hours at 1600°F; air
cooled.
BY APPROVED PAGE 6-66

TABLE 26

Preliminary Results for Run 9 in the Chloride Injection Loop

Exposure consisted .of 10 weeks of thermal cycling between the super-
heated (700°F) and satureted (540°C) states, holding 24 hr at the
latter 3 times per week; 7 ppm O was injected continuously and

7 ppm NaCl during the 540°C sojourns.

Group Materials® Material Conditionsb Results®

1 Inconel 625 angd As-Turnished No cracking either at top
Inconel .5 fs-furnished and welded. or bottom location.
welded with Ground and welded.
itself. Ground, welded, and

annealed. As-furnished,
welded, and pickled.

2 Hastelloy X and As-furnished. No cracking either at top
Hastelloy X As-furnished and welded. or bottom location.
welded with Ground and welded.
itself. Ground, welded, and

annealed. As-furnished,
welded, and pickled.

3 Hastelloy N and As=-Turnished. All cracked to varying
Hastelloy N As-furnished and welded. degrees except for those
welded with Ground and welded, of ground, welded, and
itself. Ground, welded, and annealed condition.

annealed. As-furnished,
welded, ahd pickled,

4 Inconel #O0 and As-furnished. No cracking except for one
Inconel ¢0it As-furnished and welded. suspect specimen of the
welded with Ground and welded. welded and pickled group
Inconel .. Ground, welded, and located at the upper

annealed. As-furnished, position.
welded, and pickled.

5 Incoley 807N and As-furnished. No cracking either at top
Incoloy £ As-furnished and welded. or bottom location.
welded with Ground and welded.

Inconel B2, Ground, welded, and
annealed. As-furnished,
welded, and pickled.

£ Type 316 stainless As-furnished. No cracking except for one
steel and type 316 As-furnished and welded. suspect specimen of the
stainless steel Ground and welded. welded and pickled group
welded with itself. C(round, welded, and located at the bottom

annealed. As-furnished, location.
welded, and pickled,

7 Type 304 stainless As-furnished. No cracking except for one
steel and type 304  As-furnished and welded. suspect specimen of the
stainless steel Cround and welded. welded and pickled group
welded with Ground, welded, and at the bottom position.
type 308, annealed. As-furnished,

welded, and pickled.

g8 26 Cr—1 Mo-Ti Ground and welded. All cracked severely within
{electron-beam Ground, welded, and 4 weeks.
melted) welded pickled.
with itself.

9 1€ Cr— Mo-Ti As-furnished. No cracking either at top
steel and As-Turnished and welded. or bottom location.

18 Cr— Mo Ground and welded.

steel welded
vith itself.

Ground, welded, and
annealied optimumly.

TABLE 26
(Cont 'qd)

(iroup Materiagls? Material Conditiongb Results®

10 Super 12 Cr steel As~-furnished. All cracked severely
(HT-%} and super Ground, welded, and within 4 weeks of those
L' Cr steel annealed optimumly. annealed in the temper
welded with Ground, welded, and brittleness range.
itselrf. annealed in temper

brittleness range,

11 9 Cr—1 Mo steel and As-furnished. All cracked severely
9 Cr—1 Mo steel As-furnished and welded, within 4 weeks of those
welded with itself. tiround and welded. as-furnished and welded

(iround, welded, and and those ground and
annealed optimumly. welded.

iround, welded, and

annealed in temper

brittleness range,

1 7 Cr-1 Mo-Ti steel Ground, welded, and No cracking either at top
and 9 Cr—1 Mo -T4 annealed in temper or bottom location.
steel welded with brittleness range,
itself.

13 5 Cr-1/2 Mo steel As-furnished. . No cracking either at top
and 5 Cr—1/2 Mo As-furnished and welded. or bottom location.
steel welded with Ground and welded.
itselr, Ground, welded, and

annealed optimumly,

14 5Cr=1/2 Mo T4 steel As-furnished. No cracking except one in
and 5 Cr—1/2 Mo—T1 As-furnished and welded. the ground and welded
welded with itselrf, Ground and welded, conditions each at top

Ground, welded, and and bottom location;
annealed optimumly, one in ground and welded
Ground, welded, and conditions at bottom
annealed in temper location.

brittleness range.

15 " 1/4 Cr-1 Mo steel As-furnished. No cracking either at top
and 2 1/4 Cr—1 Mo As-furnished and welded. or bottom location.
steel welded with Ground and welded.
itself, Ground, welded, and

arinealed optimumly.
16 1/4 Cr=1 Mo—Nb As~-furnished. No cracking either at top

As-furnished and welded.
Ground and welded,
Ground, welded, and
annealed optimumly.
Ground, welded, and
annealed in temper
brittleness range.

steel (HTRX6) and
2 1/4 Cr—1 Mo—Nb
welded with itself.

or bottom location except
for ground and welded
condition for which two
cracked severely in

2 weeks at top position
and one at bottam
position.

aSpecimens were of t
wall tubing. They were s

he C-configuration t
trained 0.88% on the bore side

re annealed op
5 Cr-1/2 Mo-Ti, 2 1/4 Cr—1 Mo,
heating 45 min at 1350°F (
9 Cr-1 Mo were exposed to

1OO0°F and 5 ¢r—1/2

brittleness range by
Hastelloys,

c .
Specimens were test
in the autoclaves.

ed in triplicate for each surfa

Ype prepared from 3/4-1in.

; €6 Cr—1 Mo, 18 (r—2 Mo~

~0D X 1/16-in. -

incident to mounting.

tainless steels were
Ti,

n at 1400°F and 5 cpr1 Mo,

H

Cce condition and iocation
Cracklng Results on Stress Cors

TABLE 27

wlon Cracking Specimens Teated in Runa 6 Through 8

(rack

Specime ‘ . > ‘ in,
pecimen ‘ b ‘4\””_.“‘(‘ Numther Init it fon Crack Size, in Crack
Oroup Base Metal Filler Metaul Comdition' of Timed Lacat lon
No, 8 andicion Fatlures m Inittal® Final ’
{(weeks)
1} Type 304 §45R Type 10 S5 Ground 3 of 3 1-2 1/2 1/2 HAZ into WD
Cround ‘ ] of 3 [ 1/2 1/2 HAZ into WD
Ground and annealed 6 of 6 11 1/2 1/2 HAZ 1inte WD
Ground, annealed, and plckled 3 of 3] 10 1/2 1/2 M, HAZ inte WD
t
2 Type 410 48" Type 410 55 Ground 0 of 6
firound and annealed 0 of 3
Oround, anncaled, and pickled 0 of 6
1 Incoloy BOOF Inconel W) Ground 3 of 2 1-7 1/16-3/16  7/16  HAZ fnto FL
Ground 3 of 3 6 1/16 1/8-1/4 HAZ, HAZ (nto FL
Cround and annealed 1 of 3 7 1/16 1/4 HAZ into FL
Ground and annealed ot ) 612 1/16--1/8 1/4 HAZ
Gronnd, annecaled, and nickled 1 of 13 & 1/16 1/8 HAZ
4 inconel su0% Inconel K2 Graund 2 of 1 1-¢ 1/8-3/16 3/8-7/16 HAZ
Craound 2 of 3] 512 1/16-1/8 1/8-5/16 HAZ
Ground and annealed 0 of 6
Ground, anncaled, and plckled 0 of 3
5 Inconel 5258 Inconel 629 Ground 0 of 6
Ground and annealed 0 of 6
_ Ground, annealed, and pickled 0 of 3
6 iN-1028 IN-102 Cround 3 of 3 25 1/2 1/2 HAZ into WD
Ground J of 1] B-16 1/16-1/8 1/16—-1/8 HAZ into WD
Ground and annealed 3 of & 12 1/8-1/4 1/4 HAZ and Wn
7 Hastelloy x% Hastelloy X Ground 3 of 3 1 1/4-5/16 1/2 HAZ into WD, WD
Cround 3 of 3 16 1/8-1/4 1/4—-1/2 HAZ (nto WD, WD
Ground and annealed 0 of 6
Cround, annealed, and ptckled 0 of 3
8 Hastelloy N8 Hastelloy N Ground 3 of 3 4—11 1/2 1/2 HAZ into WD, WD
Ground and annealed 0 of 3
1
9 Type 104 5SS No weld Cround 3 of 6 9 1/32-1/8 1/16-1/4 E, (many)
Ground and solution annealed 6 of 6 >-18 1/64-1/8 1/64-1/8 E, 1 (many,
p auperiicial)
i .
I 10 Type JO4N &% No weld Cround of 6 9 1/664=3/16 1/32-1/4 E,1 (many,
Ground and solution annenled Jof b 14 1/16~3/16 1/16-3/16 Jsuperficial)
(Cont loued)
Crack
Spoectmen h Surlace Number lult;;tlon Crack Size, fn. Crack
Group Base Metal Filler Metal ’ ¢ of o L {on
N, 8 Condtt fon Fa{lures Time! Inttfat® Final ocatio
{weeka)
11 1B-18-2 SSL Ny weld Growund Jof 2 8 1/4-1/2 1/2
Ground and solutlon annealed 7 of 3 68 1/16-1/2 T/16-172 E
12 1823t Mo weld Cround 3 of 3 16 1/8-3/8  3/8-1/2 E
As received 3 of 3 518 1/16-5/16 1/16—1/2 E, !
11 x20-4 No weld Cround 1 of 3 18 1/64-1/16 1/64—5/16 E (many)
As recelved 3 of 2 18 1/64-1/16 1/64-1/16 E (many)
14 26 Cr=1 Mol No weld Ground 3 of ) 1--2 1/2 1/2
(B melted) As recelved 3 of 3 2—4 i/2 1/2
15 Type 410 55 No weld Ground 0 of
Ground and annealed 0 of 3
Criund, anncaled, and pickled O of 3
15 Super 12 crd No weld Ground 0 of 3
(HT-9) tGround and annealed 0 of 3
17 9 Cr-1 Mok No weld Ground 0 of 3
(SA-21179) Ground and anncaled 0ot 3
18 5 Cr-1/2 Mok No weld Ground 0 of 3 .
(5A-211T%) tround and annealed 0 of 3
19 JOl/4 (-] Mok No weld Cround 0 ot 1}
(SA=214T22) Ground and annealed 0 of 3
20 Incoloy 8OOt No weld Ground 1 of 3 16—18 1/16-5/32 5/32-1/8 F, I
Ground 1 of 3 16 1/8-3/16 1/8-3/16 [, E
Cround and annealed 1 of 6 16 1/16 1/16 E
21 20 Cr—45 NI-5 Mal  No weld Ground 3 of 3 1-15 Vie-1/4  1/6-1/2  E
Ground and annealed 3 of 3 4-16 1/32-1/8  1/32-7/32 E, 1
22 Inconel 600 No weld Ground 1 of 3 16 3/32 /32 E
Ground and annealed 0 of 3
As received 1 of 3 5 1/8 5/16 E
23 Inconel 601! No weld Ground 0 of 6
Ground and annealed J ool 6
24 Inconel 625° No weld GCround 0 of &
0 6

GGround and annealed

Page

6-69
TABLE 27
(Cont 'd)

G Crack
Specioen . Number Crack Size, in.
Lrunp Rane Metal Filler Hutnlb Cbusf:iv c of In;;lntion : L Crnik
Ng. Y on on Fallures me Inftial® Final ccation
(weeks)
25 Inconel 690} No weld Cround 0 of 3
(10 Fe—60 NI-3D Cr) . As received 0 of 3
26 16 Fe-32 Ni-32 crl No weld "Microduplexed and ground 0 of 3
{109 menh)
Microduplexed 0 of 3}
Hacroduplexed and ground 0 of 3
(1O mewsh)
Macroduplexed 0 of 3
27 Incoloy B09E No weld Hicroduplexed and ground 0 of 3
(19 Fe—44 Ni{-17 Cr) (100 menh)
Hicroduplexed 0 of 3
Macroduplexed and ground 0 of 3
(100 meah)
Macroduplexed 0 of 3
28 Hastelloy xi No weld Ground 0 of 2
As received 2 of 13 12 1/16~3/32 1/16-3/32 E (several)
29 Hastelloy c—2761 No weld  Ground 0 of 3
As received 0 of 3
30 Hastelloy cil No weld Cround P of 3
As recelved O of 3
a1 Hastelloy NI No weld Ground J of 3 1 1/2 1/2 {(Many smaller ones)
As received 3 of 3 1-2 1/2 1/2
Anneal 1 0 of 3
Anneal 2 3 of 3 7~14 . 1/16~-1/8 1/41/2 E
Anneal 3 0 of 3
Anneal 4 0 of 3

a

Specimens measured 3 1/4 » 1/2 % 1/16 {n. and were bent to a 1/2-1n. radtus {6.2% maximum strain) fncident to mounting.
b

Welda:nte were prepared by an automated gns8 tungdten-arc process.

“Cround = surfaces ground on a 100-mesh-grit belr. :

Cround and annealed = ground on 100-mesh belt and heated 10 min at 1800°F and covled 100°F/min except for ferricic steels, which
were ground and heated Just below thelr lower critical temperature (1400°F for type 410 S5, super 12 Cr, and 9 Cr~1 Mo steel, and
1350'F\for 5 Cr~1/2 Ho and 2 1/4 Cr-1 Mo) for 1 hr and gas cooled.

{Continued)

Cround, annealed, and pickled = glven the preceeding treatment and then pickled by the procedure recommended by supplier of alloy.

Ground and solution annealed = ground und heated 1/2 hr at 1950°F and cooled rapidly.
As recelved = solution annealed (Z150°F for Mastelloy S, G, and N; 2050°F for Hastelloy C-276; 1950°F for 183 Mn, 1900°F for X20-6;
1850°F for Inconel 600, and 1725"F for Iuconel h90) or normalized and temrered for fercritic alloys (except annealed at 1450°F and water

quenched for 26 Cr-1 Mo).
Anneal 1 = ground and heated 1 hr at 2150°F, alr cooled; and Anneal 2 = ground and heated 10 min at 1600°F, alr cooled; Anneal 1} =

ground und heated 1 hr at 1600°F, nair cooled; Anneal 4 = ground and heated 6 hr at 1600°F, air cooled.
Microduplexed = annealed at 1750°F and air cooled for 327 Cr—32 N{~36 Fe allay and at 1800°F and air cooled for 37 Cr—44 Ni-19 Fe alloy

(5-uz and J-im graln sizes, respectively),

Macroduplexed = heated at 2300°F and water quenched plus anncaled at 1750°F and alr cooled for 32 Cr—32 Ni—35 Fe alloy and heated at
2300°F and water quenched plus anncaled at 1800°F and air cooled for 37 Cr—44 Ni—19 Fe alloy (300-jm and 100-um grain eizes, respectively).

dCrack distance in a lateral direction at the time firat observed (specimens were inspected at one-week intervals).

ECrack distance in a lateral direction at the termination of test (after 16 or 18 weeka) .,

fHAZ = heat affected zone, WD = weld deposit, BM = basc metal, FL = fuston line, E = edge of specimen, and I = {nside.

Blot-rolled and descaled plate stock of 1/2-in. thickness.

Mn preparing weldment, 400°F preheating (continued through multipass welding) and a 1250°F postweld treatment was employed,

iAnnealed fheet or strip stock of 1/16-1in, thickness.

JPrepared as 1/16-t{n. strip by hot flattening a tube, austenttizing at 1925°F (atr cooled), tempering at 1435%°F (alr cooled), and
machining.

k

Furnlahed by supplier of tubes as speclal strip sample (anncaled).

1l‘rupnred as 1/16-in. strip by hot flattening a tube, and solution annealing at 1750°F (water quench),

Paae G_7n
FWC FORM 172 - |4

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66

155U

1

DATE 12/16/74

NOTATTONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

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BY 3-POINT LOADING (BATTELLE JIGS) AND
EXPOSED 4000 HOURS TO 350°C WATER

BY APPROVED

PAGE 6-71

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. g8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1  |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

Figure? ~Jig for stress-corrosion test,

BY APPROVED PAGE 6-72

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |,

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO., 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE  12/16/74

WELDS

\
\

e e e

INNER CHAMBER FILLED
WITH HELIUM AT ROOM
TEMPERATURE AND PRESSURE

————— QUTER SHROUD -
58" 0.0.x 0.032"" WALL

TEST SPECIMEN -
DIAMETER AND WALL
THICKNESS DETERMINED BY
TEST CONDITIONS

e ™™ e e ™,

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Figure 3. Constantly Stressed Coupon

BY APPROVED PAGE 6-73

FWC FORM 172 - |

NOTATTIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431

DOCUMENT NO. ND/74/66 ISSUE

DATE 12/16/74

c—

SUPERHEATED STEAM 1050°F
i

HEAT TRANSFER SEC TIONS

A AR

ey

e =

P

SUPERHEATED STEAM
COUPON SECTIONS

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)

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SATURATED STEAM
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PRIMARY STEAM
DRUM

HEAT TRANSFER ) * T
SECTION ——rt ;_
llj — ! RECIRCULATION TO BOILERS
1 !
SUPERHEATED STEAM 1150°9F
FROM BOILERS m:z:
1700-1
10 CONDENSER
Figure 4 Superheat Corrosion Facllity
BY APPROVED on 674

FWC FORM 172 -~ |}

FOSTER WHEELER ENERGY CORPORATION

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

NUCLEAR DEPARTMENT LIVINGSTON, N. J.
CHARGE r g o
| v
§ o
:
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Hastelloy-N Exposed For 1000 Hours In Steam At 1050 F

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Figure 5  Hastelloy-N Exposed For 266 Hours In ACS At 1160 F

BY

APPROVED ' PAGE 6-75

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STEAM (3 to 4 ppm O2) of 550 C (1022 F) AND 3000
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CORROSION OF IRON-AND NICKEL-BASE ALLOYS EXPOSED 100 DA
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FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

8-25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1

CHARGE NO. DATE 12/16/74
Hastelloy R-235 + Incornicl 625 & Hostelloy X-280
Inconel 702 X Haynes 25 O Hastelloy N
1.0 Inconc] 718 * Inconel €00
Incoloy 800 Vv Hast:liloy C

mg/emé

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FIGURY 7/

Oxidation of High Temnerature Alloys Exposad to 15 Torr Water Vapor
in Helium Flowing «t 1 {t3/vain at 815 °C

BY

APPROVED

PAGE 6-77

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

¥

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

¢ Iasteiloy R-235 + ia1conel (25 A Hastelloy X-280
& Inconel 702 X Lkaynes 25 O Hastelloy N
4.0 F— @ Inconel 713 * Inconel 60O
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FIGURE R

Oxidation of High Temperature Allovs Expcaad to 15 Torr Water Vapor
in Helium Flowing at 1 7i3/min at 930 °C '

BY APPROVED | PAGE 6-78

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

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Oxidation of High Temperature All>ys Exroszed to 15 Torr Water Vapor

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FIGURE |OA

Weight Gain of Iron- and Nickel-Base Alloys

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Weapht Gain Mg/t

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FIGURE O

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FWC FPORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431

DOCUMENT NO. ND/74/66

DATE 12/16/74

AISI 3041 S8

CERPIEEN

r—

T L T "
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N |

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0.1

Cerrosion of
FFollowing 100 Days in

T

BN

305/ AIST 4 0658

G.2
Penetration, mils

FIGURE | |

Various Alloys
-~ Q L, : .
000 " C - 3000 psi Steam

ponoi Deoxygenated (<50 ppb)

o) Oxygenated (3-4 ppm)

BY

APPROVED

PAGE 6-8]

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431

DOCUMENT NO. ND/74/66

ISSUE_ 1

DATE 12/16/74

ATST 450 55

E\\ NN

AJS) 206 ss (Vendor i3

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N Hastelloy N

Ffi_uu;oluy
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Peneiration, Mils

FIGURE | 2

(NSNS 3000 psi
77773 5000 psi

Descaled Corrosion IPenetrations of Various Alloys
Followin: 100 Days in 3000 psi and 3000 psi 500 °C Deoxygenated Steam

BY APPROVED

PAGE 6-82

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT ~ LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE ¢ DATE 12/16/74

ORNL-DWG 68-3995R

T —
STEAM SUPPLY V4
3500 psig, 1000 °F, {

16-17 1bs/min
' SAMPLE HOLDER
SAMPLE VESSEL\ SUFPORT AND
A THERMOCOUPLE
FILTER—"

'SAMPLE HOLDER

ReLier |.] - -FLOW RESTRICTER
VALVE el

WATER QUT

CONDENSER fi\;: WATER IN

et S /

RETURN TO CONDENSATE STORAGE J
s

P

Fig. 13 Steam Corrosion Facility at Bull Rup Steam Plant,

These cannot be viewed as control tests,

: but comparisons of the
alr and steam are useful in some instanc

es,

BY ' APPROVED IPAGE 6-83

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO., 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUR 1 DATE 12/16/74

X "“";',"J'r'wv""!"v-"

PHOT Q
powr Y OTQ 96290
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Fig. | 4 Photograph of the Stcam Corrosion Facility Before Instailation. The valve ties into the steam supply. The f'langc has
been unbolted to allow removal of the sample holder and the filter.

BY APPROVED | PAGE 6-84

FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT LIVINGSTON, N.

T -

R

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

with

ORNL~DWG 73-4140

T 0O 50685
& A 5067
E o 5085
0.5 1 u 2477 .
g " e , 8
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T
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550.2
J
=
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10° 2 5 10 v 2

TIME (hr)
Fig. /5A Weight Changes of Several Heats of

Hastelloy N Exposed to Supercritical Steam at 1000°F
and 3500 psi.

ORNL-DWG 73-4139

i
E:::o 21546 | 'l"j%“*i:j
S a 21545(0.50'N)V:lfPiA '
—© 69548(0.927i) [ | ]
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TIME (hr)

Fygfl@fl9 Weight Changes of Several Alloys of Hastelloy N Modified
Titanium and Exposed to Supercritical Steam at 1000°F and 3500 ps+

BY

APPROVED PAGE 6.g5

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74
{ ORNL-DWG 73-44
¢ ORNL-DWG 73 - 4436 | g )
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o 018si0s8tT, 09820 ¥ 0 185(0.91 Ti, 0.98 zr) ——
z . b186(0.88Ti, 0-84f“" * Q A 486 (0.88 Ti, 0.84 Al)
| 0 4BB(0.95 Al, {4 Hf) = o 188(0.95 Al 14 Hi) i
005 e ® 23143 Hf, 12Y) E—
T A 232 (1 2 Hf) 0.05 ® 23{({3 Hf, 1.2 Y) —
“‘ ® 234(0.75 Ce) [ A 232 (4,2 Hf)
W 236{(4.0 A1, 047 2Zr) ] B 234 (0.75 Ce)
v 237(4.03Ch) v 236 (1.0 AL, 0.472r)
0.02 : N v 237 (.03 Cb)
) 3 4
(0 2 5 10 2 0.02 L
TIME (hr) 103 2 5 104
TIME (hr)

Fig. /6A Corrosion of Various Modified
Compositions of Hastelloy N in Steam at
1000°F and 3500 psi.

Fig./68 Corrosion of Various Modified Cor
positions of Hastelloy N in Air at 1000°F.

BY APPROVED

PAGE 6-86

FWC FORM 172 - |

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT

LIVINGSTON, N. J,

1

DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE

CRNL-DWG 73-4435

1 N
-
— ] [
0.5 —— ] A O
F Y
fl-: a
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[ U
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T L 4 S
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© ELECTROPOLISHED
A AS ROLLED —
© ABRADED - 400 GRIT
Q.02
0.0t

F?g. /7 Effect of Surface Finish on the
Corrosion of Hastelloy N (Heat 5065) in Steam

at 1000°F and 3500 psi.

BY APPROVED

AGE 6-87

FOSTER WHEELER ENERGY CORPORATION

LIVINGSTON, N. J,

NUCLEAR DEPARTMENT

1 - 2LL WH0d OMd

~3 I
™ 4y
~ Q
: £ g
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Y N S o ™ =
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HCVI NHId JAVH SHONVED HYIHM HLVOIANT NWOTOD STHL NI SNOILVION

FC FORM 172 -~ 4

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74
. ] Sl Wi SO0 IN
A R T
A ;
-8 L | ‘ | \ /4‘ iR 5.0 i ‘
3 R g B R
S | | ! f 0 berrgo '
\J { : b, e lr 420 ]
I ! . ; e _)—’/"’ ,rl'cr 2.0 J
S N N R .2 S I
! | | | j;/y; L7 OCr 110
o ! |~ T /* - !
52 ! 1 - /‘ S T ! ‘
W | | -, ’ :7 ‘ ‘ :
Ny~ |
VA N
’JT - | ( l l ! L
! | : P i
[ I b A ,,,,_._[__ .
0 Ao 800 1200 1600
TiME (AF) ‘
Fig. /@A Oxidation of Chrominm Steels in a Steam

Fnviconment at 838 C and 1500 psi, The chromium concentra-

tion is shown by caclv symbol

ORNL-DWG 73-4433

WEIGHT CHANGE (mg/cm?)

o]

no

q & n
® &
~N O

1 .
103 2

TIME (hr)

Fig. /98 Weight Changes of Low-Alloy Ferritic Steels in Supercritical
Steam at 1000°F and 3500 psi.

APPROVED

PAGE 6-89

FWC FORM 172 - L4

Frg]
i

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MAD

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66

ISSUE 1 [DATE 19/16/74

WEIGHT CHANGE (mg/cm?)

|
P I

0 400

800

ORNL-DWG 70-6788R

1200

TIME {hr)

1600 2000

Fig.Z20 Oxidation of chromium steels in air at 1000°F., The chromium

concentration is shown for each curve.

BY

APPROVED

PAGE 6-9g

FWC ¥FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J

CHARGE NO.

8-25-2431

DOCUMENT NO. ND/74/66 ISSUD

1

DATE 12/16/74

WEIGHT CHANGE {mg/cm?}

ORNI-DWG 73-4138

0.5
0.2
0.4
Q.05 ¢ l o .
.1 & HASTELLOY X
14 © H-188 .
©® INCONEL 800
0.02 . L. a 347 STAINLESS STEEL +
‘| e INCONEL 718 .
Ly : S e
| . 1 R
0 0 P SN S S s | I
103 2 5 104 2 5 10°
TIME (hl’)

E

I
i
|
|
|
+
i

Fig. 2/ Corrosion of Several Alloys in Steam at 1000°F and 3500

BY

APPROVED

PAGE 6-91

§

Y-114392

A. 17 Fe [l (Standard,
]] Cr \
2 Mo B + .
50 Ni |- 1.8 mil
/ ‘ r
Bo 2 Fe -1t A
19 Mo 16.0% Mo ‘3o §
65 Ni 6.9% Cr o
4.1% Fe 20220 2
C. J3Fe 0.02 Ti
6 Cr L065Mn  22e2lg
12 Mo £.05 §1
86 Ni Sghg| o
d(Vacuum 73 °
D. 4Fe & melted) oo
o
7 Cr SZoal s
16 Mo -
72 Ni cooo
L2a&| °
Light Optics =T
. ' . A
Fig. 23A Photomicrograph of Hastelloy N (leat 2477) Oxidized in Steam ool v
at 1000°F and 3500 psi for 10,000 hr. The results of electron microprobe ngs A
analysis are shown at the left. 500%. g
Socol| O E
oo oo —
v-11e975  (Standard) =~—-—- g
' 16.5 Mo A
7.1 Cr 8500l o -
4.0 Fe BRI C
01 Ti aasol ol 7
(Air melted) S=ois 5|8
Heat 5065, 10,000 hr e 51 e
. o A A = &
(Modified) Z2g2| =35 2
12.4 Mo%  ~°F al 2
3_8 7.39 CY‘% gooé} té'
! of8 .046 Fe% s5g35| = | =
! B 10 Tis £
T . A =
L W O OO - £
I zx2Depth of  -532| 7| | F
o 982 penetration i
i 3 10 mils cooe 2z
i (:,OOL.) o -
} i 8889 Z
A AN
s o9
e 32| ¢
o
Heat 21546, 10, r . »
¢o! 0,000 h (T1tat1 um AAR
Modified) S o= 5
11.7 Mo% .
| - 7.5 Crb A A A
e e R RS gy .05 Fe? °eg22|
T AR e P 2.1 Ti% 2a =
Heat 70-727, 7,330 hr A
. . o Ol o
Fig. 233 Photomicrograph of Three Hastelloy N Specimens Following SOl I
Exposure to Steam at L1000°F and 3500 psi. oY
L’1 BY APPROVED PAGE 6-93 ;

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74
ORNL-OWG 6B~ 3995A
[HIN- WALL .
TUNING
CAPILLARY
- o
e e 1 iy iy
v
TG HI)H‘ST. ;
SPECIMEN STLAM
l—' INLE T ' _
;2;LJ’J;N$:7‘ay1thH NG, i ] 6\(i:bJ L‘uuyéd CAPIL LARY
CLAMPED 10 THERMO WELL J; h R o ( THBE
\ f J/ 8 Y,
PR P
winl R
(S
\ ‘\ {11 v ’uh ””j\ \L\L-...,‘. .. I j"‘[‘

SAMBPLF

HOLDER {/‘(‘ "\ 3 : ) SWAGE LOCK
A\

CONNECTOR

SAMPLE VIESSEL

Iig. 24  Schematic of steam corrosion facility at Boll Run Steam Plant showing modilications for dynamically stressed tubes.

BY APPROVED

PAGE 6-94

FWC FORM 172 - |,

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO,_gfgfifiQQQL,ArDOCUMENT NO. ND/74/66 _IISSUE
O ( l

S

INCHES

L IpaTE 12/16/74

Y-408427

HASTELLOY N TUBE BURST SPECIMEN

HASTELLOY N TUBE BURST SPECIMEN

FAILED IN 1.0 hr IN STEAM AT 77,000 psi FAILED IN 3.7 hr iN STEAM AT 52,500 p:

Fig. Z5A Hastelloy N tube burst specimens removed from the Buil Run Steam Plant after failure. The small tube in the te

section was exposed 1o steam at $38°C and 3500 psi.

0.035 INCHES
In 100X

T

Fig. 258, Photomicrograph of Hastelloy N tube burst speeimen. Failure occurred alter exposure to stcam for 3.7 hr at 52,500

psiat 538°C. Steam was on the inside of the tube.

BY APPROVED

PAGE 6-95

FWC FORM 172 - L4
--— NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERG: CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

Y- 1141910

INCHES

as stressed at 58,000 psi in steam at 538°C and failed in <1000 hr. The

Fig. Zfi Photograph of tube-burst sample that w
annular region, and the smaller tube was collapsed when the plant

smaller tube was inttially prcssurizcd, failed, and pressurized the

steutn pressure decreased rapidly.

i s gv——

BY APPROVED pacE 6-96

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

T

Y-112270

Fig.2 7 . Photomicrogiaphs of the inner tube shown in Fig. 14.30. Intergranular fractures occurred on both sides of the tube
because of the unusual loading sequence. (@) Cracks from the OD, (b)) cracks from the 1D, (¢) cracks from the 1D. Etched with

riveeria regld. Redueed 337070

BY APPROVED PAGE  6-97

s <t i vnc

FWC FORM 172 - L

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J,

CHARGE NO.

8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE DATE 12/16/74

Y-114994

HASTELLOY N TUBE BURST SPECIMEN
FAILED 1M 4.0 by IN STEAM AT 66,000 psi

Q10 NADIY
[ T T N I R O

PASTOLLOY MNOTUBL BURST SPECIMETS
FAG RD LN 2704 by R STEARY AT 56,000 psi

Fig.

assembly s exposed to the steam.

HASPELLCY P TUinl BURST SPECIMAET
FATLED B0 96,7 by 1ed STEARY AT 55 300 e

2,8 Photograph of Hastelloy N tube-burst specimens stressed in steam at 538°C. The inner annulus of the concentric tube

BY

APPROVED

PAGE 6-98

e st

FWC FORM 172 - |4
~—— NOTATTONS TN THIS COLUMN TNDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

Y-115741

Y-115742

-

0.035 INCHES
I 100X

Te-

Y-115739

(c)

I'ig.29 . Photomicrographs of the fractures of Hastelloy N tube-burst samples exposed to steam at 538°C. (a) Stressed at
66,000 paiund tailed i 4.0 hr, (h) stressed at 86,000 psiand failed in 27.4 hr, and (¢) stressed at 55,300 psi and failed in 99.7 hr. As

polished. .

BY APPROVED PAGE 6~99
NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES BAVE BEFN MATE

FWC FORM 172 - L

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J
3 L] +*

CHARGE NO, 8-25-2431 ]| DOCUMENT NO. ND/74/66 ISSUE 1 DATE 12/16/74

Y-1450“62 -

-

0035 INCHES
1N 100X

Ton

»
0.035 INCHES
I 100X

aphs of a Hastelloy N sumplé tested in argon at 538°C and 40,300 psi and failed in 565 hr. () Fracture

lig. 20 Photomicrogr

(1) Repion away hom Tacture showing Trequent eracks on the inside dinmeter ol the tube. As polished,

BY APPROVED PAGE 6-100

FWC FORM 172 - L4

Fig. 3/A. Specimen of duplex Incoloy 800~Ni 280 tubing tested at 538°C and 46,000 psi hoop stress in the Incoloy 800. The
specimen failed in 3263 hr with 2.3% diametral strain. The longitudinal markings arc due to dye penetrant that was absorbed by the

cracks.

NOTATIONS IN THIS COLUMN INDICATE WHERE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N.

Jl

CHARGE NO. 8-25-2431 DOCUMENT NO. ND/74/66

ISSUE 1

DATE 12/16/74

™

Y- 106620

0 0.5 1.0
I I N
INCHES

Y-106954

- a (
‘\ . ' L] [r*
. \ ‘
! " ‘
1 ¢ o
.- /o ’ . o
/ ‘._ v ¢ :’ . \'
: . . . . .
..’ "./ -[. » ! e ¢ ;' l' ) L
. ! - . - ‘ .‘
\ - ’ ‘ ‘
; . . P
i Y / ’ "‘ ) l‘
" '\ '\ . . L o
. . ¢ t o :
_ . e ! e
. 14 r nl@
b L]
t' hd N . 7 i
--I " {
- I ' :
. ; . \ o
. ' ! v ‘
' ! .
/
: P I |
¢ ) - :
d * f ’ ‘ -
. ’ - .-
L ‘s !
- ’ “‘
Fig. 3/8, Photomicrograph of the Ni 280 from the specimen in the previous figure. As polished. -
BY APPROVED PAGE 6-101

FWC FORM 172 - 4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO._

8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1

DATE 12/16/74

Iig. 32/-\ Photomicrograph of the primary failure region in the t

on the outside. As polished. Magnification 35X%.

ORNL-DWG Tt-13683

o | | L o h*
e e
| AT

: ‘g.&@#w e A

o o .
0] 50 100 %50 200 250 300 350 400 450 500
TIME (hr)

Fig. 32 8 Comparative crecp curves of Ni 280 shcet at
538°C and 20,000 psi in an argon atmosphcre. Samples taken
from longitudinal and transverse orientalion,

Y-106952

ube in Fig. 13.32. Incoloy 800 is on the inside and the Ni 280

BY

APPROVED

AGE 6-102

FOSTER WHEELER ENERGY CORPORATION |
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 DOCUMENT NO.ND/74/66 ISSUE 1 DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L

Y—106551

INCHES

Fig. 33 . Samples of Ni 280 exposed to steam for 2000 hr at 538°C and 3500 psi.

BY APPROVED PAGE 6-103

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - L4

FOSTER WHEELER ENERGY CORPORATION -
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE DATE 12/16/74

Y-111626

0 0.5 1.0
I | J

INCHES

Fig. 34— Photograph of duplex tube-burst specimen discontinued after being stressed for 7075 hr at 40,000 psi at 538°C. The
tube had a diametral strain of 1.14%. and the dye penetrant clearly reveals cracks in the outer nickel 280 layer.

BY APPROVED PAGE 6-104

FWC FORM 172 -

FOSTER WHEELER ENERGY CORPORATION

' NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO. §8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE DATE 12/16/74
M Y-112660
' “ i
: L
» ':,. '
(a) , s
' Y-112659

(&)

Fig. 35 Photomicrograph of the cross section of the specimen shown in Fig. 24
Humerous cracks are present. he Incoloy 800 is on the inside and is not cracked. As polished

0.0351MNLHES
iN"IC0X

. The nickel 280 is on the outside, and

BY

APPROVED

PAGE 6-105

FWC FORM 172 - 4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO.

8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE

DATE  12/16/74

ORNL-0WG 72~7¢ it
25000 pasi . ; |

—— LONGITUDINAL - ’ S
~—— TRANSVERSE

25000

i
20,000 . /

b /
N / I
: A ;
rd g i
o 715,000 :
|
i i‘
| |
.
, | ‘ = i | l
ol .. . o . . J R |

0 200 400 8OO BOO 1000 1200 1400 1600 1800 2000

TIME (hr)

Fig. 36 . Creep curves of nickel 280 sheet at $38°C in
argon.

BY

APPROVED

bAGE 6-106

FWC FORM 172 -

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER FNERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N. J.

CHARGE NO.

8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE

DATE 12/16/74

iple of INCONEL 600

Figure 37 Photomicrograph of a san
C0G Ni=13G Cr=6%Fer after exposird
in a pumped toop for 15000 howrs at 1300°F.
Voids near the surfuce are formed

as chromium is removed

selectively by the salt.

st fluoride salt

BY

APPROVED

PAGE 6-107

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT _ LIVINGSTON, N.

J.

30 ]
_ \ l 1,
25 2000°F '
— &
— ry
E ¥
o 20 5
0
o 1800°F o
\‘f s m———
0
E 15 // o \ * 14
St /— "
=
b /
()
= — 3
T
<0
L ]
z ___’_G_OOF |
P Ly — 2
- 1500°F \
5 A ™\
D '
1400°F
1200°F .
0] A i\ 0
0 1 2 3 4 5 6 7 8 9 10

CHROMIUM (wt %)
Figure 38AN ariution of the rate of axtdation of an 80% Ni=200 Mo alloy

wrth chromium content. Note the very sharp decrease tnorate at chromium

levels of o 7%.

Composition, % by weight (Base: Ni)

Alloy Mo Cr Fe Ti Al Nb W
“INOR-1 | 20

INOR-2 | 16 5

INOR-3 | 16 | 1.5 1

INOR-4 | 16 1.5 2

INOR-5 | 15 2 2

INOR-6 | 16 5 1.5 1

INOR-7 |16 6 ' 1 ]

INOR-8 | 16 6 5

INOR-9 | 17 5 3

F/GvR £ 388Several promising nickel base alloys
melted in the course of developing INOR-8,

CHARGE NO. 8.25-2431 | DocoMENT TMO. ND/74/66 l1ssoe 1 Ioame 12/16/74

"OXIDATION RATE {mils)

BY APPROVED | | ~ |PAGE 6-108

FWC FORM 172 - L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N, J.

CHARGE NO. 8-25-2431 | DOCUMENT NO. ND/74/66 ISSUE 1 |DATE 12/16/74

AY

20 . . ......-..’.. ‘

—|ncone| 606
0 b {Hastelloy N _.

-'_'_‘—-—Eh
Erfi‘ ______:-‘"
-20 | Sinconel 600~
. \nconel 601 -

-40 \ Py
| \ In conel%

-60

| Incoloy BIA
-80 \
-100 - \\

0_ 1,000 2,000 3,000
Cperation Time, Hi

Weight Change, Mg per Sq Cm

~120

FIGURE 39

. Graph illustrates the relative corrosion resistance of
various alloys in LIF-BeF,~ThFg4-UF, molten salt studies conducted
at ORNL @ 690°C(1,275°F).

BY APPROVED  [PAGE 6-109
FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N, J.

CHARGE NO. 8~25-2431 | DOCUMENT NO. ND/74/66 | ISSUE 1 DATE 12/16/74

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - |

For Metal Temperatures Not Exceeding, ° F

8

0501
0S1Lt.
ogzi
[£3 8

AR EEEEEERE
20,000 T T 1

|
19.000 ANNEALED MATERIAL -

=1 osvL

17000 \
\

=
g
et

A 2 \%

11.000 00 %
TN

\ o

°A 5
® Z

% 5
-]
AN

AN
Y \
oo \ \ '

%
Q
4
r
£
<
P
%

Maximum Allowable Stress Vaiues In Tension, psi
o
g

HIGH-TEMPERATURE
APPLICATIONS
900 to 1500 F

FIGURE 4O

By APPROVED | AGE 6-170

FWC FORM 172 - |4

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT

LIVINGSTON, N, J.

CHARGE NO. 8-25-2431

DOCUMENT NO. ND/74/66

ISSUE 1

DATE 12/16/74

Fi6- &L U-Bend SCC Specimen Holder with Mounted Specimens.

SUPERHEAT PHASE
OF TEST CYCLE
(AT 427 °C, 800 *F)

UPRIGHT
U-BEND
SPECIME NS

WELD

INVERTED
U-HEND
SPECIMENS

Flg.
During Cyciic Test,
NaCl) injected only during saturatlon phase (24 hr at S40°F three times
prr week),

DRY STEAM
CONTAINS 20pom O,
AND NEGLIGIBLE ClI-

_~SAME AS AROVE

¥=119400
QANL-0WG 73 3441

(— =~ $TEAM FROM FPC
SUPCRHEATER

A e e e e e t—— ——

BARRIER

A

|
|
|
|
|
|
|
Sy ! T ST SECTION

| rjmp
|
' THE RMOCOUPLE
| i (typ)
| 4
: {57 SECTION :
| |
[ |
! I
{ ' CONDENSER :

. b e e e — e e e — J

CHLORIDE :

TiON TANK CHLORIDE
sq o ANALYSIS
OXYGEN
é MUNITOR

oamNY

Schonioti;: of High Pressure Chioride SCC Facility

Flg. 4” Schematic of High-Pressure Chlortde SCC Facllity.

ORNL-DWG T2~ 722 1Ry
‘SATURATION PHASE
OF TEST CYCLE
(AT 282 *C, 540 °F)

WET STEAM
CONTAINS CI- AT
CONCENTRATION
TO /20 OF THAT
INJECTED AND Op
1%7 THAT INJECTED

i /m

WATER
CONTAINS Ci- AT
CONCENTRATION OF
442 TO 5 TIMES
THAT INJECTED AND
Q2 W20 THAT INJECTEQ

WELD

SAME AS ABOVE

}c Schematic Showing Specimens and Environmental Conditions

Cantinueus injection of 20 ppm 0,5 10 ppm €17 (as

BY APPROVED

[PAGE

6-111

FWC FORM 172 - L

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Photomacrographs at top (left and right) illustrate the appear-
ances of clad Ni 270/Incoloy 800 tubing at different magnifications in
the as received and macroetched conditions.

Photomacrographs at bottom (left and right) illustrate the appear-
ances of clad Ni 280/Incoloy 800 tubing at different magnifications in
the as received and macroetched conditions. page 67113

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPCRATION

LIVINGSTON, N. J.

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Photomicrographs taken of a transverse section illustrate

the microstructural appearances of the nickel
materials in the metallurgically bonded areas.

270 clad and Incoloy 800

L.‘ BY APPROVED

PAGE 6-114

FOSTER WHEELER ENERGY CORPORATION

NUCLEAR DEPARTMENT  LIVINGSTON, N. J.
" '
| | CBARGE NO. #-25-7%2/ | DOCUMENT NO. ~/5/7w/z: |TISSUE /  |DATE jz//c/v¢
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Photomicrographs taken of a transverse section illustrate the
crostructural appearances of the nickel 280 clad and Incoloy 800 materials
in the metallurgically bonded areas.

L’l BY APPROVED PAGE g_7115

FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE 0. 8-25-2431 | pocuMenT 3o, ND/7L/66 ISSUE 1 |[paTe 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE —

FWC FORM 172 -~ |

SECTION 7

MANUFACTURING ENGINEERING

Approved by

N &z G e
M. J. Kraje
Manager, Manufacturing Engineering

BY APPROVED PAGE 7-a |

FWC FORM 172 - ],

T UPEINTRE Lt e et . -7 B vl 2 03 0« AR el e - 1 < = a0

T o s o X 7. Ay W Rt B a5 i A AT

NOTATIONS IN THTS COLUMN INDICATE WHERE CHANGES HAVE BEAN MADE ——-—~— ———

e e b e 2o 3 o et |

N

FOSTER WHEELER CORPORATION

DATE 12/16/7),

CHARGE NO 8-p5-2),31 | DOCUMENT NO. ND/7L/66 ISSUE

TABLE OF CONTENTS

Pages
7.0 Manufacturing Ingineering 7-1
7.1 Maintenance Procedure T-2

% 7.2 Sequences of Operations For Manufacture Of e
;_ Molten Salt Steam Generator

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FWC FORM 172 - |

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2L31 | DOCUMENT NO.ND/7L/66 ISSUE 1 DATE

-12/16/74L

A,

7.0 MANUFACTURING ENGINEERING

The manufacturing engineering responsibility is to review the
drawings of the project and determine the manufacturing feasibility
of the design as it is presented; and make suggestions that will
enable the product to be manufactured and assembled to g1l control-

ling codes and specifications within the ability of the shops to
produce, '

The following assumptions have been made in order to make a continuous
review of the proposed steam generator.

1. The size of this unit precludes the shipment by rail, so it must

be shipped by barge. This can be accomplished at our Panama City,
Florida plant.

2. Our experience in internal bore welding and excellent intrepre-
tation of the gamma ray film should allow us to mass spectrometer
test the completed unit and void the requirement for helium leak
testing each individual tube.

3. The ability of Foster Wheeler to develop the internal bore weld-
ing in the "5G" position is only a matter of time as demonstrated by
the test welds made in our Mountaintop lab.

L. That the hastelloy-N tubes will be able to be produced by weld-

ing two short straight sections together and to bend the tubes so they
are defect free.

5. There will be no cladding of the tube sheet.

6. That the radius shell section can be purchased completely fabri-
cated except for girth seam welding by Foster Wheeler.

7. The parts of the shell will be manufactured of plates and forgings,

8. Supports at the curved portion of the shell will be made up in
pieces,

9. Stress relieving will not be required,

The summary of the results are detailed in section 7.2 where the
assembly procedure is presented.

BY

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FWC FORM 172 - }

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FOSTER WHEELER CORPORATION

CHARGE NO8-25-2,31 { DOCUMENT NO. ND/7),/66 ISSUE 1 DATE 12/16/7),

7.1

MAINTENANCE PROCEDURE

The following is an outline for detecting the leaking tube and the pro-
cedure for plugging it.

To find the tube that is leaking, it will be necessary to remove the unit
from the system. ' »

The unit may be removed by cutting the inlet and outlet steam pipes and
the salt inlet and outlet piping. The piping should be removed by cut-
ting through the field welds that were made to seal the piping to the unit.
By pressurizing the salt side of the unit with helium and "sniffing" the
tubes at the face of the tube sheet (steam/water side) the leaking tube
can be found. ‘

When the leaking tube has been identified and marked the following pro-
cedure will be used to plug the tube. This procedure starts when access

to both tube sheets has been obtained. The procedure is based on the
following:

1. The tube sheets are accessable to allow the insertion and weld-
ing of the plugs.

2. The unit has been decontaminated and cleaned making hands on main-
tenance possible.

7.1.1 Operational Procedure for Plugging the Tubesheet Tube Holes after
the Leaky Tubes Have Been Identified

1. Insert expandable plugs into both ends of the failed tube to pre-
vent foreign matter from entering the unit. These plugs must be in-
serted far enough to allow the various preparatory tube plugging oper-

ations to proceed. These temporary plugs are to be designed for con-
venient removal,

2. Plug both ends of the tubes of two adjacent rows surrounding the
failed tube with expandable plugs to avoid contamination. Inventory
plugs and record data. See sketch on page 7-17. for typical pattern
of tubes to be plugged with expandable plugs.

3. Cover the balance of the tubesheet holes on both tubesheets with
clean polyethylene and seal with tape.

i, Clean both ends of the failed tube hole, inside and adjacent sur-
faces with a clean stainless brush and acetone, then vacuum all sur-
faces clean.

5. P.T. area to be welded

6. Clean

7. Remove expandable plug prior to inserting permanent plug.

BY

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FWC FORM 172 - L

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DATE 12/16/7L

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13.

15.

Insert permanent plug as showh_on sketch
Weld plug complete using filler rings.
Inspect between each pass.

Clean éfter welding.

Cleanliness inspect.

Repeat operations 7.1.1-5 to 12 for other end of failed tube.

Remove plugs installed in Para. Tele.l=2 and polyethylene seal
installed in para Telel=3,

When the tfibe sineets have been exposed to plug a leaking tube,
the rest of the tubes can be checked for errosion by the use.
of U.T.

Each tube is U.T. examined al the time the unit is being
assembled and the wall thickness pattern is recorded. " The-
tubes that have been in service are reexamined by U.T. and
the wall -thickness pattern is recorded. The records of the
in service tubes are compared to the records of the tubes at

the time of unit assembly and the errosion determined by the
difference in these records.

BY

APPROVED PAGE 7-3

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FWC FORM 172 - 4

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NOTATIONS IN THIS COL

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FOSTER WHEELER CORPORATION

CHARGE NOB-25-2L31 | DOCUMENT NO. ND/7L/66 ISSUE 1 DATE _12)16/‘7h

7.2 SEQUENCES OF OPERATIONS FOR MANUFACTURE OF MOLTEN SALT STEAM GENERATOR

7.2.1 Purpose

To establish a preliminary procedure for fabricating the
Molten Salt Steam Generator.

7.2.2 General

Certain Assumptions were made in order to fabricate these
Hastalloy "N" Steam Generators and these are:

2.1 Tube Sheets will not be clad.

2.2 Tube to Tube sheet weld will be performed in the
2g ‘and 5g position respectively.

'3 2.3 We can purchasé the radius shell section completely fabricated

except for girth seam welding by Foster Wheeler.
2.k The parts of the shell will be made from plates and forgings.

2.5 Supports at the curved portion of shell will be made up in
pieces. '

2.6 Stress relieving will not be required.

7.2.3 Plan

et s e

Listed below is an index of procedures for component sub- |
assemblies and final assembly identifying the overall plan to
‘-fabricate the Molten Salt Steam Generator. Refer to sketches

shown on Pages 7=l and7~1l5 to assist in understanding these
procedures,

3.1 Straight Shell Course Subassembly "A"

3.2 Shell Section "B

3.3 Inner Shroud Section "C"

3.4 Thermal Sleeve Forging "D"

3.5 Straight Shell Section "A" through "D Subassembly .
3.6 Strdgight Shroud Section "EM

3.7 Radius Shroud Section "E"

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CHARGE NO8-25-2),31 DOCUMENT NO.ND/7L/66 ISSUE

DATE 12/16/7L

7.2.3 Plan (continued)
3.8 Radius Shell Section "Gv
3.9 Shell Section "HM
3.10 Tube Sheets "K"

3.11 Tubes "L

3.12 Adapter Ring "M"

3.13 Steam Nozzle "N"

—— o — s e

3.1k Bundle Assembly

T R o bl ™ A e s a1 kst

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CHARGE NO 8-25-2,31 | DOCUMENT NO.ND/7./66 ISSUE 1 DATE '12/16/7h

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7.2.3.1 Straight Shell Course Subassembly MmA"

The shell sections will be made in a number of courses, each with
one longitudinal seam and Joirled by girth seams.

1. TILayout and machine the weld preparations on each of four edges
in the flat,

2. Penetrant InSpec£ Weld Preparations,
3 Rdll to required inside diameter.
L. Set up and welq longitudinal sean,
5. Dress longitudinal weld inside and outside.
6. Check circularity and reround.:
7,. R%diograph longitudinal seam.
Re;eat sequences #1 through #7 as required for each course.
9. Set up for circle seam welding,
10. Weld circle seam;
11. Dress inside and outside girth sean.
12. Radiograph girth seam.

13. Repeat sequences #9 through #12 as required,

7.2.3.2 Shell Section "g" | | f

The shell section will be made in one course with one longitudinal
seam, :

1. Layout and machine weld preparations on each of three edges in the

flat. (Finish machine three edges.) Don't machine edge that butts to
adapter ring. ’ -

2. Penetrant inspect weld preparations.
3. Roll to required inside diameter.
Set up and weld longitudinal seam,

Dress longitudinal seanm inside and outside.

LAY & U

+ Check circularity and reround.,

| ]PAGE 7-6  OF

FWC FOZL 172 - L

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7. Install roundness retaining rings.
8. Machine cutout and weld-groove for inlet nozzle.
9. Dress balance of weld preparation.
10. Penetrant inspect weld preparations.
11. Weld nozzle to shell.
12, Dress inside and outside weld seams.
13. Radiograph longitudihal shell seam and nozzle to shell seams welds.
1h. Set up V.B.M. and machine girth seam weld preparation for "EB"
insert weld @ adapter ring., Tt may be necessary to weld build
up this section before machining.
15. Penetrant inspect girth edge preparation.

7.2.3.3 Inner Shroud Section "C"

The shroud section will be made in one course with one longitudinal
seam from perforated plate. :

1. Layout and machine weld preparations on each of three edges in the
flat, .

2. Penetrant inspect weld preparations.
3. Roll to required inside diameter.
L. Sét up and weld longitudinal seam.
5. Grind longitudinal seam inside and outside.
6. Check circularity and reround.
7. Penetrant inspect }ongitudinal seam,
7.2.3.4L Thermal Sleeve Forging "D
The thermal sleeve will be puréhased as & rough forging.
1. Layout and identify centerlines.
2. Set up on VBM and machine to configuration.

, -

3. Penetrant inspect weld preparatiafis.

BY

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CHARGE NO 8-25-2)31 | DOCMENT NO.ND/7L4/66  |1Ssop 1 DATE 12/16/7L

7.2.3.5 Straight Shell Section "A" through "D Subassembly

1. Set up for welding of shell course subassembly to thermal
sleeve forging "D,

2. Weld girth seams.

3. Radiograph inspect shell to forging girth seams.

L. Set up for welding Shell "B" to subassembly from sequence #3.
5. Weld girth seam and grind inside end outside.

6. Radiograph inspect girth séam.

7. Set up for welding inner Shroud "C" to subassembly from
sequence #6.

8. We%g girth seam and grind inside.
'9. Penetrant inspect shroud to forging girth seam.
7.2.3.6 Straight Shroud Section "EM
The shroud secfion will be fabricafied as two hélf shells from plate.

1. TIayout and machine weld preparation on each of four edges
« in the flat. '

2. Penetrant inspect weld préparations.
3. Roll half shells to required inside diameter.
L. Trial fit the two half shells.

7.2.3.7 Radius Shroud Section "F"

-

The radius shroud section will be fabricated from purchased formed
half elbows with all machined weld preparations.,

l. Penetrant inspect’weld preparations.

2. Set up for welding smaller radius half elbow sections together,
3. Weld girth seam.

L. Grind insigde and outside girth seam.

-«

5. Penetrant inspect girth seam.

BY | APPROVED | PAGE 7-8  OF

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6. Repeat sequences #1 through #5 as required to make complete
small radius section.

7. Set up for welding subassembly from sequence #6 to half shroud
"E" from #3.6. : :

8. Weld girth seam. -
9. Grind inside and outside girth seam.
10. Penetrant inspect girth seam. .

11. Repeat sequence through #10 to make Up complete separate larger
radius shroud section. ‘

7.2.3.8 Radius Shell Sections "G

The r%dius shell sections will be fabricated from two purchased
elbows*with the girth edges machined for "EB" insert welding.

1.  Penetrant inspect weld preparations.
7.2.3.9 Shell Section "g"
The shell section will be made in one course with one longitudinal seam.

1. TLayout and machine weld preparations on each of two edges in the
flat. (Finish machine two edges - Don't machine girth seam edges. )

2. Penetrant inspect weld preparations,
Roll to required inside diameter,
Set up and weld longitudinal seam.

3

L

5.. Grind longitudinal seam inside and outside,
] |

Check circularity and reround.

—q

Install roundness retaining rings.
8. Machine cutout and weld groove for outlet nozzle.
9. Grind balance of weld preparations.

10. Penetrant inSpect weld preparations.

w—

11. Weld nozzle to shell seamn.

BY

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CIARGE NO 8-25-2),31 | DOCUMENT NO.ND/7L/66 ISSUE 1 DATE 12/16/7),

- ——

12. Grind inside and outside weld seams.

13, Set up oh VBM and machine both girth seam weld preparations far
"EB" insert welds. Tt may be necessary to weld build up these
sections before machining.

1l Radiograph longitudinal shell seam and nozzle to shell seam
welds and penetrant inspect girth edges.

7.2.3.10 Tube Sheets "K"

The tube sheets will be purchased as forgings.

1. Set up on VBM and machine Both flat side suffaces and weld grooves,
2. Ultraéonic and penetrant inspect machined surfaces.

Layout and drill tube holes.

Machine spigots and counterbores,

Set up and machine for island removal between spigots.

o U W

Clean and deburr tube holes.
7. Penetrant inspect spigdts and island removal areas.
8. Radiograph spigots.

9. C(Clean.

7.2.3.11 Tubes "M

The tfibes will be purchased in various short lengths and welded
together,

1. Set up and machine ends.of tubes.

2. Weld tube to tube. |

3. Grind tube welds on outside diameter,
L. Penetrant inspect tube welds.

5. Radiograph tube welds.
6

. Helium test straight tubes.

.~

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| CHARGE NO 8-25-231 | DOCUMENT NO.ND/7L/66  |1SSUE 1 DATE 12/16/7%

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7. Set up bending machine for bending fubes.
8. Bend smallest radius row of tubes.
9. Trim tubes.to length and machine weld prep.
10. P.T. ends and inspect.
7.2.3.12 Adapter Rings "M".“

The adepter rings will be purchased as forgings and machined at
final assembly after welding Handhole nozzles.

7.2.3.13 Steam Nozzles "N"
The nozzles will be purchased as forgings.
1. Set up on VBM and machine complete.

2. Pefiétrant inspect weld preparations.

i Ay s e - B v =g, sl et -

7.2.3.14 Bundle Final Assembly

1. Set up for welding straight shell subassembly "D" to small radius
shroud half "Fn,

2. Weld girth seamn.
3. Grind girth seam.
i, Penetrant inspect girth seam.

5. Sgt up bundle assembly fixture and attach inlet and outlet
tube sheets "K™", :

6. Insert guide rods through steam outlet tube sheet into shell
and install full supportplates using the guide rods for alignment
purposes.

L - T A B P ™ A S A P T e et o o e . e bl . el s

7. Set up and install segmented support parts in the small radius
shroud for the smallest radius row of tubes. |

8. The guide rods will be attached to the longer tube leg and with-
drawn as the tube is inserted into the bundle. This will insure that the
tube passes through the proper hole in the support plate.

9. Thread the~tube using the rod, if required, and spring the tube
against tubg sheets.

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CHARGE NO8-25-2,,31 | DOCUMENT NO.ND/7L/66  |ISSUE 1 DATE 12/16,/7

. B DY i o s g . e B

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10. IBW weld tube to tube sheets for the smallest radius row of
tubes. '

11. Clean tube to tube sheet welds.
12. Radiograph inspect.

13. Install segmented support parts for assembly of the next
larger radius row of tubes. ' '

1L, Repeat sequences #8 through #1l, until all tubes have been
installed. | '

15. 1Inspect and check bundle for cleanliness.

16. Set up and weld larger radius shroud assembly to bundle,

17. Grind welds outside.

18. Penetrant inspect longitudinal and girth welds.
19. Slide shell sections "G" and "H" over shroud,

20. Weld girth weld of the shell to thermal sleeve and shell
course to shell course.

2l. Grind outside girth welds.
22. Radiograph girth welds.

23. Determine finished measurement of distance between tube sheet
and shell for adapter ring. f

2. Set up adapter ring mM" on-VBM‘and machine to suit,
25. Penetrant inspect weld preparations on adapter ring,

26. Slide adapter rings over tube sheet and wzld adapter
td tube sheets and shells.

27. Grind girth seams.
28. Radiograph girth welds.
29. Install test caps and test shell side.

30. Set up and- weld steam nozzles to “ube sheets.

31. Grind girth welds.

BY

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CHARGE NO 8-25-24,31 | DOCUMENT NO.ND/7./66 ISSUE 1

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32,
33.
k.
35.

DATE 12/16/7)

Penetrant inspect inside girth welds.
Radiograph girth welds.

Install test caps and test tube side.

- Pressurize unit and prepare for shipment.

BY

APPROVED PAGE

FWC FORM 172 - |4

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2/31

DOCUMENT NO.ND/71,/66

DATE 12/16/7L

a\

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NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

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BY

APPROVED

| PAGE 7-1, oF

NOTATIONS IN THIS COLUMN INDICATE WHFRE CHANGES HAVE BEEN MADE

FWC FORM 172 - L4

FOSTER WHEELER CORPORATION

CHARGE NO 8-25-2,,31 | DOCUMENT NO. wnp/7),/66

DATE 15/16/71,

U'r R
\LARGE RADIUS
SHROUD
SEGQGUENCE-23
|
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BY N APFROVED [ pace  7-15 oF

DATE 12/16/7L

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DOCUMENT NO. ND/7./66

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FWC FORM 172 - |

T MADR

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T3 COLUIY INDICATE WHTRE CaMnsg

HOTATZOMS T TH

FOSTER WHEELER CORPORATION

CHARGE N08-25-2L31 | DOGUMENT NO.ND/7L/65 ISSUE 1

DATE 15/16/7),

TYPICAL PATTERN OF TUBES TO BE PLUGGED WITH EXPANDABLE PLUGS

SK—4

FWC FORM 172 - L

S HAVE BXEN MADE

—
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HANG

4

NOTATTIONS IN TEIS COLUMN INDICATE WHERE C

NUCLEAR DEPARTMENT

FOSTER WHEELER ENERGY CORPORATION
LTVINGSTON, N. J,

CHARGE NO.

8-25-2431 | DOCUMENT NO. ND/7L/66 | ISSUE 1 DATE 12/16/7)

Contents:

(4-1)
(4-2)
(4-3)
(a-L)

APPENDIX A

Mechanical Design Calculations

ohell and Head
Approximate Weight
Tube Expansion

Tube Vibration

BY

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adies
FOSTER WHEELER ENERGY CORPORATION
NUCLEAR DEPARTMENT LIVINGSTON, N. J.

CHARGE NO. 8-25-2431 DOCUMENT No. ND/7L/66 ISSUE 1 DATE 12/16/7L

NOTATIONS IN THIS COLUMN INDICATE WHERE CHANGES HAVE BEEN MADE

FWC FORM 172 - )

APPENDIX B

Thermal /Hydraulic Calculations

Contents:
(B-1) Steam side pressure drops at inlet nozzle and tubesheet
(B-2) Steam side pressure drops at exit nozzle and tubesheet

(B-3) Molten salt pressure drops at tube support plates and
vibration suppressors

(B~4) Molten salt pressure drops at inlet nozzle and shrouds
(B—S) Molten salt pressure drops at outlet nozzle and shrouds

(B-6) Analysis of dynamic flow stability in steam generators
for the molten-salt breeder reactor
B. E. Boyack, Gulf-GA-A12),16
Gulf General Atomic
November 20, 1972

BY APPROVED AGE

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Calewhsds A olklel cham =:da Prassays &\vbY ‘\*hvo% %)

) 1’\_,'\‘;_\ \‘( .—\j oo ‘5';', be . 4o c‘.h;\ )DJ/. y«)w} c.m(kt'rws
_{;)':Kr*\)'np{ (e {-‘/.-uj.n'lkY{m) j:"(\;’O\YTn'\J.:‘ “'l;\nw\l y '\/u‘\M\M l P\'JC‘

el Themad ‘f\\!o\o\»\b«, Y««(\xmmu cwxx-»fir otk ovfsy\j.'

s "/tkm (’.Lriujw (Tlhl'(&i /&rvz\r\z‘gw«j

D(th' ‘\'} A

'y Vb7

\ 1) lev /, L S\

[AS ME

{A) = '{\ ‘,-_I E)f 1*’ I~ ‘b ,(4 o PA‘R

(! ~
Va= 40 = 1333}
-'5;\~ 2. C0pts §°) §

ND=- 726~ 1%

Fon Takleg

B Fwe

Fage B-2-/

LUV A LA

110 SUUTH URANGE AVE., LIVINGSTON

wWHooLLuLuon WCWUNTDUNRALLIULY
won DATE o a e e . BT 5 e e s s i S s i e o R e

Rt it g -

SHEET NO..._ 2. __oF

o v \ b 1
i At T - 1Y igLho 4l 4‘&
1y 2
= Y04 4oy 3 4‘»«
. by : )
TF - LY 2 235 10N 4y
N x WA -
—{- ; 0\.\; ‘
\ L ?
allys =5 ok, T
1
. o.cera | x —— x0139L x Sot-Qosy x TER
(+3113 b
Pz 4 4 \evnd z_? - : = 0,034 9
WA ARG N VAN = &
o [ PR
ol = ¢80 booied = o0b3l

.Z‘t\.)g C - ."\‘-‘\‘ ‘\ "’h);
e = /3.. —.lq
| el | gy LEREsR S
‘-"—{l‘ 5 l£\~ = } A ')

FORM (285]

Page B-8-2

rvoion WHobLEK CUKPURKATIUN 110 SOUTH ORANGE AVE,, uvngcs"ron

e S P 7 o 0L D2y PN LAt A\ e R S N SHEET NO... = ____ OF
CHKD. BY _____ e SISt SR TR, £ W W § e N JOB NOwooooooo .
I\ = ’ N \
.‘ — | N\ .\ ? \

o B
\ ¢ , _l._
aee = 130Ae Y Tan b X Sed Jesy x T4g = 71341k
e = 3beuelsb 4 703410 = 3het gosa
> }
= o liy)
\ !
SUIRES 20 .01 L 0k 6 .
J TN A P
C/“' = — - = N5, k4 4\1_ Se
Qeppatt x|
f .—\‘l.‘ t
— ¢y I '\.'x;
Rr 7
= e \
. 2y )
33 ) 9.b2y)
C‘z:l-l'\-‘____.Jm-l;\x;)
! A - i \. - |
Ao L | 269 ¥
Z — 0,144 ) } 2 \w{‘\l e ~|sVhy
. \ l — wd 1] 2 X b vy ~ \ > A ’('q_ !
; s = aovl - Lgiha = 3beC ¥3¥a
-
] \ 1,,. ¥
¢ \ )31 4 \ FORM (20885)..

Page B-2-3

............... DATE-----_--_. SUBJECT.--_-----------_-----------_---------_----
CHKD. BY._______ (oo it a0 2 T O e o AN
Vo
12713791 ((C,
H),
4
SSL ey f- tee
= r‘ (‘ \l A &4 LT '%i
b,
(W L b ¢ LRE ;‘\'.'.{
& TEI390 %, 245 Yot
¢ 5k ———— " = Y998, 4
' 4 .o Nloy
' ‘A
- - ot 5y bt REERL
‘ O =4y 1
' x) v ie) ® 2
et b piE Ve o= TR L i wigs
0 51 | i s 7 . \1
T ———— XSSy % ( USRI
-4 8, L4y \
, 12911
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4 Y
: - e A '\lA b AR} (
1
‘e \ \\_?j -\.’K\ 3
CnAXYy A5 )X by g a
T = bl
) ol
Bl ;- ' b 2179 = &.Y9%3
€= SWVaday +29958 < 3oy, 244) (S
! =
5
11 ::'
\ .\ R \

M AL UNMALLIUN

“v\‘a-%l,'/ = Shov

ok ?\I-\,k’)) J.,,>}-

v

SHEET NO..-.':t’.-- -OF

£ (O-l‘i“" 0+ 01192

0.14¢ |

< TH o= A9133

FORM (283) A7

Fage B-2-4
110 SOUTH ORANGE AVE., LIVINGSTON,

rudS1IER WHoBELER COKPORATION
Y S el R DATE e, L T . WY G S VA S s SHEET NO----5 ----- OF.
CHKD. BY . ____ PATE . sadsemsiy ¥ ndus drinansestbistbn DALEIn st MO s O St JOB. N i i 446 crstomarie

= Lot et Ny _mf‘ffis—e
"’{\ 2=t :\ 'E 1 '\“\
TK — I‘\' \. ,—_
& "R 't“ '
\ [}* - \ \1“\4‘ lb
1
Lip = UL SNEN -4
. '\}'"
) W\ - WP P /‘}V 4)
Q0. \b
NS ,)l;\ o
r, .= = 13se . ny A fy
A | 5\! )
| 154
= 3776 14T -
st nhx 13443 g
[\( = - = ——— = s ) ‘1';7!4&;
\ \‘J4'$I‘\
——o & 1. ST
-+ 2 ee )
(SRR ¥ "\‘ \l ! ) A %
;_.V = . X 0. ) b'|‘ x;‘“); Y%+ X {“' = o0.0\%%
-4 Fee i : . (e
\ 1.\')') 5N L"’-“
) -
i = i b s a s oa i oy fond o'c‘>'3c
JPhY ¥ N
t = 0
Ve 0¥+ Lie = 0,830
\
P S FORM (285)
) . "\;—. G ’\"
‘(;— \!._\\..:; 0>0{S ix)&\ (\A VR: D.l‘l)) \/L

Fage B-2-&

rvoion woLALER CUKPFUKATIUN 110 SOUTH ORANGE AVE., LIVINGSTON,

BY e DATE .cnauas - BUBIBCT. - indocsstioahata i Sl e s SHEET NO.._.©_____ OF.
CHKD. BY e DT adintoan | feamsnamanis e yonans Moaovmsior addosd oot st e, SO N st n i iii s adaians
_\_f_’_ ;y‘t..ft'aom ‘1!; c
(N = L3940
t s \
+) - ~ o A ¥ ¢ il by X < 2 . 0
lge. = V333 % g, A0V} X 390 g R (44 %ot 2
Lo= b Y 44042 = Yhoy wxso
Ve = Bt
PR NR Y (b
= g o ¢4 .
“Te Ly 2k C4 +vy 3 4@0,
|- v 5 DU
! \ ) .«_‘)
. L.-.‘i._ 21 U it
I\"}- £ !$|{\>i
! 2 {
) 3 N o 1 n g / =Y 2 5%
2\ Sy & = V3 \\;‘7; ” h v 1\\:'« ¥ | L‘-\—.q—q,q.l) XK o o;g\o"'
1\30-: e axy0 - C\Q)\U’) - ?\)U)\ iq’%’j
\D — W "'" ,"
‘:\’ =3 + R ‘ig.ll-—\' L(L

o
0 ) -~ 2
b, = e, S “"*'X;O\\‘J'bs] 4/\_,-\\

ih
- 4256 A

FORM (288).

Page B-2-
FUSIERK WHrELER COKPORATION 110 SOUTH ORANGE AVE., LIVINGSTON,

- STV e DATE ¢ oo v T e N e P P NPT T SHEET NO.--.? ..... OF..
CHKD. BY - ___ L e UL O R I S i AN e AP S JOB NO.cicasccasasani
< N (’fjl\\;‘k.';)‘\{\)\v‘*:):&
‘ = = ’)o&,ogq_. G2
u 0 RGNy
Lb = O ty\léy
pay i c.C\Y 2,;
/<.t'>(-.\.\h—\ ‘L = oSS ey )
: - .l \&‘.
e
\\-: 3‘_( Ve ,*-\jo E = 0.2\%b
’ 5 /badgs | s A "
i = Golyd x o o<zibh X 2 Sl x T = L83k
! ' C C\& 2 b .+
‘ SN .
aAY, = ez Q0 &
‘ . 300h v
a41 . sivy
— >
) . "HV ’ ! ’\S *
- \‘ = / WS > 0¢cl|lV
\ q’_g‘_

o< A6y & ougid 4 oollo =(?\"78b

“\'E = Abe3 SR 4 | FL = 3beX wrag
h‘h; = tubb F
- AL
= = v 2\hx l/k‘
T\C *.‘\_.\. ' ‘(“~"s\"-g.( ‘\ ‘);». ,t‘ s 2 %b C‘S.q}-l% == .;\) C\«'
: \ tofai

C. “,'\._l’\ ()g,; {

FORM (2089)

Fage B-2-7

110 SOUTH ORANGE AVE., LIVINGSTON,

rudioni WHoELERK COKPORATION
Y e e e PDATE oo cacasa: U B R Y i e I R e e i A dete s SHEET NO.._.&_____ OF.
CHKD. BY e DRI tissiias. | Psileanabbad s ins T et e G es s uah et ot s L JOBND . cococsnssensne
\ L) )/ -uu\
\ v/ 4 A ¢ oy % R 4
A 12 R CAL - Y
i'"/% O % F \('g,'-..&
.’T.}\ = 1oy (—
‘»h:— 0.3+ %\
L
WA = v.casly) ".“v"\k
V3L 229 \b
(:—r!., i, e PSS g3 /é"\y
3 15303
N \b"
= b Ase
g o TG I e e
. . = o™ 'J)\),’\’
_fr— = O, dWond 5 FK
\ = Qv J
5 Uds x | s i I
St 8 m—mmmmmae ¥ ror o) OHLNLY X e = oa(ed
’ \'g\ y d '
' )
> e = - == o U, 080
P, o 1) Aty gl 3 ‘
‘ i-.‘A\ =\
5P o anif ronded = wpdi]
- ’ - Vi AL ’ } - “ LI
~ \)‘;.) = oo, ew\) 'VK': o~k |
FORM (20t

Page B-2-§

FUSIERK WHRELER CORPORATION

110 SOUTH ORANGE AVE., LIVINGSTO)
) TSRS I DATE. ..oee. i L T e e SHEET No....1.____oF
CHKD. BY_______ OATE oot = Sraanstdinnnm e o S S A JOB NO

~ 4 . )
- C,_} ',‘NT'D‘\ sl ! \u._g'

P R

| -~ |

) T I X 7 B S 5 xX 0138 ) x 1Nl Kz Ty, = 3

(e = 133 Vin s 2182 % Ty >4y
:‘; 5 ‘,i)\v\aq’; >/ + -)‘%"\) > = sbvl-sg*z'

v ShlLee \y
' ) | : /s
= . : = Y].47¢v 4—5«-
').(1\ 2V v
- 7
g e v \(’ 5]

cy = =§13.15819 g

- 4
o
o ST A 0,1_-)“%:-‘ . ( $‘i.u-’)‘-§\l ) x = -~ 0O, S\‘S

to = 3bvz Si4l-osny = s 012

(}
|
|
)
<
}
>
o
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<
-5
"s
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-f

FORM (2085) .

Fage B-2 =0 &
rvwaion wnoL.huonmn CORPUKATIUN

- —— s e e et o o

L ———

110 SOUTH ORANGE AVE., LIVINGSTON, N.
SHEET NO

SUBJECT

T T T Rt e e e e - ke e

L N - -
- T T e T e e T N e e e e L e e e~ e i -—-- e R . kI S e W S -

) IR BRI L+ oo wr e » |
he = = Yosee’f >R
o ofs )|
"“':‘;“: o C.u - \ .\"'Q’Lr
e 1 :)!X

boLru g

\':T:_ =
om “;" f\ NG ’L}fl )‘::- C_\,"VL‘\"(’
RSN y _ ' . - ! A
= - o L SPRY I L{4. & TS = vl
s B GG e 5038 K g = e
iy |
- e 2 03
LMy VN
‘ \..‘1"1.03\-"
B f\‘--‘,l“'fi.‘l’wf 1 3_;..,7( \, x évl{-‘\;w ~

A

2 = ey vty A nvesy = e
‘:.: = -."-Ai?‘-’- Ll A UL‘??; = 3‘303“\?‘7
i
/ t ES 4 -’+ .';L' ‘\{
i ‘ & hall VI ‘;.!:“ ’{
-_,.\}‘_) - Vo 5, A l\)
{/ oy L A Thod 1 3
le o= thezenrt A L ogd b = 1bed jh g
S :
- e FORM (283;.47
o= TR

Pooe B-2-1p
rvoLlon wnoubbuom UWWURPFUKALLIUN

Y e e e DATE oot s nad: SUBJECT cous ianscn
CHEKD, BY uuuucan D A B e s mmiant st s e e o e e
E ‘l"-\ PAR QAL S {"‘V“" ¥ = )\ | \l \
\
,}: e o S -_i ¥ ‘.\
(h = 3be® i
—_— H (-
| iy = 1\ 0 Lf"’" =
_ A
Vo= 0 L2 h Y ) .
i/
W= V.es)) {v-4%
4 ‘Ve\)')bv )
th ! 34b3
= Vav, \x¥)
y Lo Pxdiud A4-3553
R, = = )oos¥4 Xys]
G © 33“»\

-~ 3/ ‘
; P‘ ALY N |

> --". — , r—

. 1538 b

!
. ‘\) ‘\1 = - 0 ogo L
y Oa:bd ¥ '

B
<5 L = 1
./’ ‘)I' - | r 's\ * L’,\‘_) >\,‘\j : D.c >5)

.\\& = \\t)“vx S] \\\_‘.‘a\_
AN
= 024 bt

110 SOUTH ORANGE AVE., LIVINGSTON,
SHEET NO.--).\.----OF.

\b
”4’ %,

‘ h/x‘-yzu

.aS

3 )

xy. 23 b% X \§o \sv’] X ol 0.00% )

FORM (202

Page B-2-

110 SOUTH ORANGE AVE,, LIVINGSTON, N,

rugion WHobLENK COKPORATION
BY. oo DATE oo oo . SUBJECT oot e SHEET No.--!_;L____oF-_l_:L:
CHKD. BY _______ AT E o e e JOB NOw_ o oo
|
Ly Lo v - {'T&L
v = ey
‘ ‘ LR
1 , VL i —— v E < | " P , - I
“ho = bEae A T b gy X R T W )
R e R L I LA TR
L -_— -
. 4
5 s Al
C =~ 0\ Lir\) ) ,‘b
_ Yo sthy i\
(—~ = ‘ - = = b e | ‘Jflf
< Fohu % Yoy e
Lot - a‘%
n (U S AL Ty
e

‘. ; f > \
S o\\f'l—\a] v ik}*L’J*S\) X m = —0.119 g

oo, Ti9b

'__.l
—
ft

‘lJ'[,. z  <bo Vo A
Voo ot b
- 1 1 *
4 Lt\T [ —Ab\bi (‘\\\L'\,f 41\&9(
" \ ¢ \-.—(‘:-gu \ bfjj ‘x -«
e IR I TR e
e Vb1 k
1L
= 1520 A s
m S EaLY)
AIVSL WY o @4en _
- . o o hEA £
r‘ ) \‘,u(ng' 13(0 _ FORM (285]-47

Psame R=2=}s
TN CURFUKATIUN 110 SOUTH ORANGE AVE.,, GSTO!
S N DATE....clc. e 0 A 6 e A i SHEET No.._'=2____ OF
CHKD. BY_______ i e B TS, LTS S o b

¢

:4"\,“.\ f\gt 3hox
t = v g,tle
v —\‘:( z p): o b\

\ 4 e 1S X (35,097 T X e = ey F3
— ‘_\ - e 01X ) X q\‘- b‘t > o 13 & X \% o O] ) lt‘* - -\+ q
) SRR

4 s e

=

fenby_carbh) x
o

= poel3
Yoy

ST ST LA S WY I 0.x24

f \ \ .
[0 vz ;)\ = 3Shea 1198 Yoy

= ] w3

'\,;{.-x 0-\-"L’)

—
r
I

o
11
S'o
i3
v
P
‘\J

FORM (208).47

PQgé' 8-¢~/3

FUSTER WHrRELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON,
DY it BATE et i U RS A LR L TR R sueer no.. T _or..

L) Ao £ Load

VA= YIbss buia
Th= 1o3s 3 R
e © R

Viy =

= ocishl

Lo Zis ‘ifln‘» e e e A)QLL(_;

')
Lo o

,.
A
)

(&
= 33003b. &b -t/‘\,l-,‘fq

: \v
< ‘Q‘ ‘. b‘\ol) \ \'}(L
Aol th x 13333 : .
= & A4S XY
N 0.0 %)) SO )
-.::— = 0. 050w 5 )
H

d “j"\ ol 13337 by 4
)
O e el B TR P
.\\. Yo b X Ty
-\i{'\\ : 0
it a Owng + V7317 = 0. 032 4

FORM (205)

P‘?ge‘ B':“u

FUOSTER WHERELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON,
A =k DATE . .t i L = P SR OO S ¥ SHEET No...' 2_.__oF.)
CHKD. BY_______ DATE s arassins.  otsevtsanbesmatinia o i o s i e Rie b des ke OB MO oovyamyosnsends

| .. %
ey Y Sy X L0 x Ty = 0117

Ve = b sih 4 a1 = Yo 3ins

S NS I
. L.'-'(.-u("]{"' \'e
(re 5 = — 5 (3 370 L5

1 b2l < b

2 )
W = = 1353 Taw MOV X (330 x T = ~0.055
(b= 39, L35~ vossy = Yhoo, 354k
\ = .1ty
.\_"fi) - J‘\. 5 —:.yj_‘» \.,. *‘\\‘i { “‘____t__._j-\:g
S \ b
."k -— :._....._...._._-.___. = ()bjTr)“.gg /‘\'k"r\f
D I Y LT
%
= 0%y e
"= V.o &Xs7)
61570 Lw x ool
L - = 17034).9
( d a ; ‘\ FORM (285).

Page B-2-§

FOSTER WHEELER CORPORATION

110 SOUTH ORANGE AVE., LIVINGSTON,
_______________ BRTR et antan « BUBIREE L 1 2 ot s bt ra i s dadadie. . SMEET O LR L
CHKD. BY ceee e o SRS SRS N S 5 o e SNPGRS JOB NOisasoserismnasna
— & C.re\ L4y
O
-4 = 00"
! .
:\l(\) Ve = dbosgxbh
v o= Ua Ny
-
il < nLLL
%
241 ) ; wb % :
*o = gl A q P em——— )(07.'13"\.' X\°\7.2\'&_§1 + lq,\* = 0 \332'
i &3 | 044
N 6.0 ¢
) \“1"} '7
Yo = —— — 2, 0.0 4]
-’ o.M X Tuy
' = U3+ Cous’y = ahsy

TJE = Ybww vebb + oy 3y = '}\wo.w",i"\

-&Ro-*‘hi. = o.q_'lo\

¥ Lo

FORM (283).

Page B-2-]

FOSTER WHEELER CORPORATION o 110 SOUTH ORANGE AVE,, LIVIN‘GSTON, N.]
BY-______E"_\_:';_C:_DATEf;‘i_f’_ff}? suBJECT..} __0.\_.4_"3_-:’.’:'(.J_:_-:%_E-_‘.{“.\‘.‘:‘l ..... ‘:‘.‘E‘&- SHEET NO..._. L ___OF..1.2.

e ‘ e .
CHKD. BY e = DATE ... -&_“i‘.{‘.i\\.dfi.\; - ;\1 \_‘(Z“Ré - .‘.\2‘_’\- _\"_‘}»1‘!:1 WA 3."-’:‘?1}1&.&5}.‘;\_\{) .- JOB NO. e eeeeeeae

Lt }n«‘f"f‘-fl E (;b)'\c,r.viu»t( ke t[)‘u«-‘s&w-u c;'\'w*\w UJL ho\’\'m _S.AJL*" PM&\‘n}

i \'\ hTRN f}‘fi‘\ .v-J;,\ Y . ‘ .‘" o fi" \ :d’} % (. r{/\ \/. i}Y[\j’l‘m ) %L‘\ é’(\“ A :\,,\.b‘(s

“\.\i‘ Jr(/ Gl Lh ;—:»}\ah(p,(_f\'f{hg:l a'._:{,[-?ml'\fm-‘;* t\flfi.r\M » UD\V\M I , -”;W C

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e ) i \fut T (,J, f~ s}i L\ \'o-l.-\."('-'i{., \u »(/{w L NS g _Cm?vxt-v L€A£_ 0"‘1—(?’:(_
1\""“ ‘—I.L \L{’« SL’\.{'T‘)‘S‘T {\“"‘\_’L . Qf\r‘flw ‘flg ~E N D - 1o R%S

Yy TL,\\;,& ‘v'(;&mvfi”\m (,_,u{r]hf{%c\/ ,'DYAM"‘} A ND - G0~ ]qrj

NS Vo L"] P\ G t (')’t-u\ ™ T‘5~L \4:";

. [ - .
I S A (5o
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. Y, 1 ‘ - I |
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(‘/o(,«i oo -‘tnf\%-p ..'i!vj, Voo ‘*L\‘H‘#, lve A% fl@‘a L3 JCFL ‘(D(LS

Feat 7l Loaw veee ol )fl.i’l‘& '\'-"-*i‘\“*"‘\’ Iz\'i = 0.6y X Lswe *A‘-

191y AT

|

\ ( 28 ~ ! A 3 Lo B \L
- A — l'n‘\ Dy 'S % A —_&“ u - b} T
A ST TR T AT ) T Y gl TE
_ ‘ « x‘\;(J\} - AN A & bv - UA‘J"’)’QQ\,,L*, FORM (285) -4
L =
= iy \3U1"fi \)\D -4 -\\J

Paae Be%-}
110 SOUTH ORANGE AVE, LIVINGSTO)
BY e miaaaa e DATE . ___. i) (T PO PRREL) SR COPE. ) SR Sy SHEET NO.---':‘T----OI
CHKD. BY o ____ e TR R JOB NO. .. oo .
3133w ¥ oy \*
“‘%? T L T i (T)
: R | A
\ ) .
T L v Ny
(>

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)

\b 1
T = e | 1,51\,“\"’ 7E7—- {-Tt

b
= 194, «x ] \41 W
:‘q IS~ A ks\\})m -4‘_'\.\ of | ~+ "‘h-:r P }\L ‘L\m\’lk .-r\jwr'x J\*\ J'Lt\b‘\ 4%—& ’\v\‘f?sr((’ ‘7‘&4‘1

'S Tabw A 4 3 oW

Lo Den v fibe ¢ L"\-’f”‘f fl'\+& 'S enlatured /6(\'77\ (
}\;

( v -'?\lj,r-.n\ ‘-.v-.-.) ‘)\.‘ U9 > v awn \uol-&(."\ )
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FORM (285

Fage 8-2-4

rveron wOowbhLoih UCWURFUKATIUN

NN
s :. H
A - 'fl
L
Lo N
A
£, Q
1.0 h
f“!' ‘
oo
Al
LY ™
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DR
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LN D

NS
L

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by ), -,.“
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s 73

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——— - ———

SUBJECT

T e A o E e e, — e o E - mw e mem .-

e e e - om o m

.\;rfi:_ut'w% “ UJ\N\'\

ia ‘_/
—-‘"'/lb

e - an e ee == e o A

Rl . X e

v &‘fl\j\w C_,p\{\r JALLBY §

FORM (2085).47

Pase B-3~3
FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON,

B Y i s e s DATE cacaauaiy B D T v s v o o5 om0 e i 0 0 SHEET NO... .{t-.---OF--

\ 2. a
\s SR A - DR g8 X '-)QQ.E.S"') X T‘l:L\( = 0’\85$°

‘ 2 3 = -q-\—-' —- 4 Q
) GV X T X owegd X8 857 x [ S 0. 135 )
ey > u
———- - + ,_.‘ Nl <
Yy > Undys % ¥ Gedy ¥ % S5 X @ S °o\¥ 3|
b e N

) ALY
LB = wgrs X 0— % oagh ok RSN A B = &2

"‘,L._ = L4 N
! A '/ e | < - ‘—t‘ —— 0\3'»;
EXe 5 sy g ST X NI T T P

. e o ’ A — \
.. i.)‘: 2. D.8Vy 2 pay ™ Qouszy XN )Q“ SS”) X VeV a 0, Q‘-k’)

1-'4..0}'&(.(. Sl S v ed v\j 4‘»1&( &a-~fie(" ¥ ‘vjls
o \ o J_;-,y"* \ .«:\1’1 + < .*\)5 =+ \ora, + &4 O(g"" 54{5

—~ 'f" l[) )( \’S\

™
- ¥

'\L’ =L \'-t"\“v‘~rr. e pp U 0
T

C Mdmen = A

/)

10 e o) et vt copprsen = 355 21X ) < 35300

| = \ )

( ‘.(-_, visn il ChERO /3‘ = _‘ifi \ ( M 1\‘1— ( 39-1v$ 52 }—_— (;_o.g‘pbl\j, “X’\,L

’%' Y ~ PR b _ /" LN ’{f A A B a\_,* -t'\ k‘( Lt "I"M

: - ! T 0.9\ *
) § - e ! ’ 4 'V‘\ ¥ ( )

s ) '\ \ i 3 H / - -~ — \\)0\) x %3 ———
| '

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‘ \ -

- ~ ‘ & / -
=~ ¥ ¢:qin) - : 1 Sl 3"\-')“& - ‘---ok 'q("“ FORM (28t

Fige B-2 ~Y

110 SOUTH ORANGE AVE,, LIVINGSTON,

FUSTER WHEELER CORPORATION
) STy g, 2 ¥ T Y ey S (o ey o P P e SHEET NO... . ____ OF..
CHKD: BY iaauoca PAT Rt aaniio, bl iwnS Slfie ca et AR RS TEoR (s s S otz s JOB NO.w e e

P | Jrv\kl\ Lo»o\u A"—? SJ"’(\%K\) '\'t»

A . 5Y 5589 ~ :
j - ; = © 53‘§/71\0 S o3

e
— N
)
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| | -—;— ( .
30 o < SR WA A |
‘b,' -

\ E ‘. ! (
L 2033 AN ) X GF =

z \)s 2 X 0Ax8r = o U(")Q—l)

7 ™ Y lu'L - Lo ~\\ Slones
3 ‘,")4,\) }.\ elwiie Ayt ‘ X .l.\ Ni\&\ ‘}L h{ ¢ u”xfij\ T lé.—‘.ts fi'w( V\.LVT\\N\ 40??’“5&0"1
: 43
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| 1 /? i "/. Loo- Cl\
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AR v L‘uv‘T " e e oy \ A “L\u "T/«i k/ l\’k)

FORM (288)

que g-2-4

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.]

v
BY e DATE _________ SUBJECT . . e e SHEET NO._.__ Y ____ OF._..%2.

T ; 5 vikarTom Sopprastirt

‘ '\ ;}bt) :
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FORM (288).47

Poae Pon~r,
FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON,
- ARSI AR DT e B B T o s s AR SN Rt S s o SHEET NO..._[ ____ OF.

L S e \
,S_‘ = >R - ‘ : et = QKS.S‘° \ 4 \“

2 )
) = OVE A T, X000¥s X hsSistlel x g = Olex

L
-2 \
Ny = BRI X W L L 2 bsC 1) X T 0, 12 b
) 1 __\' - e 1% R0
e ‘5 R TT 0 WP b4\ X U”\"‘I;.-\ o {15; S\“ K ey T < S
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‘;u = 0-3Vy X L"~—‘& « eoogl X L«&;.\}\o\ Ve o 18X

I - & S . 1 " —\—, % 3
i.-\_ = UsQ Yy X To e P G-oc [ X \.}S,S\s‘ X Yt o\ $‘
| b’ > \ & \" A
. 3 Sy ) : , 3 —_— A
.--:,‘1:’6 L T YR R VLY & PSS T Y L ) 3

U x._nj SO '._'\\')\"\\U‘l)

x g g g
Al 7 wovd > o X e d g bSS.sle) X RY = 0.1279

:- ._‘\’ ~ L\\O‘T'Ilf ; — \3\:)")0
>'a

\'_ ) %,:\'\ =1 h 'o‘\.\.'-\{.\ ’}v L: RS .\'"m"» \)lm AL '~~\ e L\'\o‘j’“"‘ 3"‘“!"155 ONS

\ \

FORM (203).

Page B~3~9

FOSTER WHEELER CORPORATION

BY e~ DATE ... ___. SUBJECT . . e
CHKD. BY .. DA TE. o e e
S co ;
..-‘— 5\"' ./"“ L ='L~-C‘\

SHEET NO

E )
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f fé.-f '*‘“4.!'{ ! iLmp. 'F "Pu k\“' V_\ e
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;
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p < ' . e ; = b f
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110 SOUTH ORANGE AVE,, LIVINGSTON, N.,
vy i

FORM (285).47

Pooo B=F~§
FOSTER WHEELER CORPORATION

110 SOUTH ORANGE AVE,, STON, N.
BYomo DATE oo, SUBJECT. - - - o e oo e SHEET NO..___t____ oF._ .2
CHKD. BY __.___ DATE - oo e JOB NO.o o oo
- \ 3 ER S b RERS : \B
(T = : = slb.apux t,l' Lo
{,14,\;}'\\3 x Ahto

A ) R l_ L
I M A Nox 6‘”'*.‘— X 0.0egY X i dbve X oy =0 "
4
. \ y . b - \ -
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' o L 3 R
P IV L R sih abvs x0Ty s 00 Ry
\, . kbl e )
. {tq - J. \.\y“-.) - ;a—-L Fe -deg P g\b. (-‘ byg K lq“f 2.4 h\_‘_"
) i ! ' 1N
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\:S = O 0wy X, $€{-‘U- L LU 4 '] * Clh.aglys  x “P{’ o ¥\
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K ' o s : —— « O.o%.
Vg = Ly X L4\ -% oeed R S \\31 q. b\«’S. A 1W ¢ X }

[
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e oV X vt x 9 L xS b v QW'
L gkt Stbbys x T o= 0bh b
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-;,\‘,- . —3\)_‘:“’_2 = L ‘,«0»%
i
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=0
v
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FORM (283).47

[0e B-3-9
SvVeaman woonLSf CURPURKATION 110 SOUTH ORANGE AVE,, LIVINGSTON
7 TS DATE. .o o 3 AT P S SR S N U sHEET No.. \?____of.

| b
SRR O B TN Spefec Vel
4,
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1 lb X 1/ i 4
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FORM (205) .4:

Page B~3- 10

FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, N.)

_______________ DATE ... SUBJECT. i mcccciciciiccccamem———————————— e SHEET N'O.--_{}‘_--_-OF \3

. e e Em e e - e = oo o o P e - ———— -

!}\"‘J‘; CZ. ERANEN

[ \b
T < = IV &N ‘/‘, L
T voav sk Kb > | {4 L

3 e ™

/ - - —_— 2y . . 4 = o<\
,_;;,ftz CAivy o TG ¥ au.;_.v? D SN 14§ 4—\\.\'\ 3 f‘f"\f v.oY4 LP

, | o R A ' ‘
= s A Y 0SS XN K S ek
| A~ i

i
<
p

£

»
Y

‘ ———— Y Ay : crappm——ry,

L )
} . - e O

Sl oo X g ¥ 0] X AN g s o4y

R : . e :

} : - | e
N SR i a.00¥y X TNy 1 TS 0.+ 44

= oofie
L i - . .
—= 1 e ‘\j-‘ L rt'-..T; Ve LY \\f‘fl Lafd
4 - ‘ o
ik o | ' < 0N mx - o.osb
PR EE LNA @»; Noveeg4 X 7. ‘-H*Q . Hc.il, -0

FORM (288} .47

Pgae 3-3-4/
FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, |

BY e DATE e BURIEC T oo PR, W SHEET NG, 185 o)
CHKD. BY ... -7y | WO O e .o . 2 R Sl o A
— . [
k - ) <4< »/3 1~'at-‘\

f r o i ¢
| Leaeliov Fa f lerp 'F SpReifee vdvmf
] ‘ -5‘(}/;,
Lb7 ' R a.00R4
4. \) , Yo '
751 : QLL
Tk 133 Vi L\,J\'m sv-“au sLavd
(x-) Yak 5 oa: R
(133 Y49 o
2133 4o o ve g
L )33 | pot!
12 3} | L=
3733 ,' Viay 0§41
- ’ ; ie¢hy |
<y ' 3b3 !
55 ion’l
£33 ~ Vo)
b h) T
Y735 X L
‘“3 AL
RRS &3
LM =877 !
3 B X1 {
12 43 VXS |
Ve \7d 4
' 3 3.3
v e
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(R \3 ¥

FORM (285)..

Page B-3-/:

ruUdDiIbl WHLALEKK CUKFURATIUON

110 SOUTH ORANGE AVE., LIVD!GSTON
BY e DATE ..o, SUBJECT __________________________________________ SHEET NO.--;.:?.----OF.
CHKD. BY ...____ i SRR s B T O AR N I FOB NOureiin e S
Q\ '.)‘ 4 Y 9 ‘
RS, = 12€.3%4% 2{‘,*‘_

y 1 .
O LY :L‘,_; A deids A .Zfi.;(/); e ey 0-(‘\‘0%

=0y = %oy X T «weegh A 2383 A JT o= ospn)

= g i
ot ORI 5 S SEPPRVINEICRTES B B ¥ & FEY 3 4P v = 0173

e i (R (R L AT DY %

o < \ ."\’
\) VDT \':{Y\ A\ ‘\’\\‘4‘ AWTS O

\__'- ) c\l ‘ -
_._,-r\ - 0.}-)_2)( ban ¥ Vo iYL x “;;x.>‘)'IS X VY = 0,\)“&\}

| \

| —
—

3 3 :
,_'j.l)L = (',-V& ) r;fi._; ¥ Q t.‘-‘lslk X 1,;k\,l. [k X 'w L O.O\LL &

)
£
T

J Tl = S 3184 2egig = eds) Y

FORM (283).

Page -3~

ruo.tnn wobisLLR bunruxuuu%“ llo SOUTH ORANGE AVE,, LIVINGSTON,
-l G WATA ___.\-_-_-__f&:& ..................... € sHEET NO..__ \____ OF.

- -.---------- - -

R- 4

'\‘L'\#'\X\‘Jl . &\:\M»j} *’\& T“O\tbh &e&* .@M.Livu O\“’F -\'\\n% iw\l/'i'

’\‘925\1 m\(“ ‘ .\\ges--c\_‘/

\

L’ fdvzwex 0 \Y G\fiu‘—uh'é CL MT\N.I \1\4\\\%& ] \)F‘\\AM X. PVJC/
\l‘. ’n\;‘“\,\}\ ’e\y(hl‘-h‘-‘- (Jlf'tgu»« neg L?fl\’\':V\-TA) C‘b(id Ot;*qwd—
RN G g T o L e A N L L
W Todieead Shed L AP-2%1-17)

sy WHY  AsME Siw Toblss

¥ ‘.5\,'\-( T = YLvL = LLhodaa “\“'.

2 _ » 5
Ax (" L ogth L 1% -4V = 308 = (.7211 ¢
K/
i > - 2 S\u“
Ao = 4T ( '-b!‘b’z) = & 6L\2 44
h U

Tt Te haeF f
2l P e | s a Shvomd.

A= g\ € o = g5 44 i

FORM (283

Pdge B-Y4~
FUOSTER WHEELER CORPORATION

110 SOUTH ORANGE AVE.,, LIVINGSTON,
) (R —— DATE ssssnns TRy SR o Ny NN R S PPINLL S SHEET NO.---.:.'...--OF Y
CHKD. BY e DATE teciias £ amuicnies b svewibl bk owides  ete o o Rl JOB NO:isocccancnncaus
C ’\ " .}) %
(55 = — LSS RS SN
% 3. k2
PA
= Jefg.4b2t, - At e
| Gipy 58 VML) X Lbey
i\ 5L e Y . = 4 bbyy X\~
/\A .-L.\s\.
sg 6 N0 S T\ )
2R e = O.0VuU 5
O -6 04
f - .o \ 3
R ¢ g ¥
e ) ‘Z i \ )
s (j_k = C/-L-~'.’Q. i\ )(! . Y — X ‘—:-' e lo"‘q.“,\{) e ”’* = 0-0")8
‘ [-bed L b4 ¥ 50
fe2 Max_ougg = 234 %F22
RER G . X aedionn ariia
(,.\‘\v LY i /S\\J (\ 4\/\4 ,! a O

: i L
e L ¢ b eva\ L= 48 = 4 A%

. ] / Y
e bdoan el chveed = 3 (5% - 40" ) = 9= 06833 4t

-

Rt «f h e Ag = 0578 % 4 = 31313 4%

~ , 1
N\g = 3 (1% g ato ): by xe ,%:r
yhda ts y 2
Gx = — SRS & R L s
' d-\?)'\fi\ |
b
= %0‘\.&1\§\' “\"\"“rfb

FORM (283)

Fage B-4~2
FOSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON,

B o o o DATE ..o .. BUBIECT. o a S R c e e dmmm s s R M o=
CHKD. BY . ___ OATE - cootrisre . orsadiccecisnnesatind Shuts S 3 Bt oaniamenn s e gins s e L~
c_ -
A TR — ) B s )
e b, Frann (P S .
&y, '\~ et
\-»n‘fi(?fi,)"vb(“:) :

3 ( ..-%).‘:”5_'\' - ( §o%. 543

— | =2 v B0k Yy AN
3997. G4 b >¢33.76>b ) 6

—" \7\bb’1'
) | ) £ Z X —‘_..
v o P . ” iy + o e Sy e o = ey ]‘
ol = v by g7 # § "X g e é"I ¥

v ¢ i | i )’. b zlx :
: - < 1a?Q Sl — K== x| S -~ .
= o8 LR X e T Wl 2o8q. b % ey ¥ V’* \8 5 L‘Pg‘u;? 2t A

P .
¥

o= R34 391 - o bSsY = 34168

< WA 4-x 13335 |
‘./\c; P g & ST e > = |l.c\§r Nt
P o 24 (a0 S-?)B) -
| 1013 L
=L
2L
_f\) — 3 ( \ v\« \+ _‘T 9\
= o 150D Nt
R > Ak - v
an et ——— - 'l |
De o X &
\(x = L‘-\W—b
\ & \
< rh: 2 49 X oy \—": X QO\&I—'LX X !'q_.:' O\,Q,)\
F\" = _EL\.)A.. 23 %.’,""Aé&_"):—* l‘.\*b- = lls ")bx‘ob FORM (283

W U 9
Pdge B-4-3

rusien WHeotLER CORPORATION 110 SOUTH ORANGE AVE.,

L e — DATE ..o, )0 ol ATISPRNE SN N A R L S SHEET NO..__ ' ____
CHKD. BY.ee___ DATE . «lukaice  onaonisaninsioss ey ot sk oot M S R o e 3
. ) - nonk
= - = O.oec2,45))
O {-21]L
k\ - G |
{ / L ovtm L8 ¥
Lowth ~1ave s 29 K7 2w xx115F = 1-10385 44

Ne B8 4 f

12 ' \‘V(?' ‘ .' X o S‘\ St
Sl (b Lo e iy i 6119 X 0.0\4
\’\." BE Y A ol X ek ~ V2 99 3 [y < b

- ;’:'Ll\)"\‘ A &""fiv‘(l V \n.“’i Y "(_) &&“ U\AL‘ = 124‘ ;)SK = 0. '3 27 ')

(.

= 133.369| pvya.

\ \)\’ ‘\M\« (\

P bGs  Rela o\ 0L D

. . l -
' ’\.’)'h,\.‘ AV i 4 ’((-"15 = \g(.to X :‘f\' (, A ) = C'.cgls Js‘tl:' Ao

| ‘:"0 ~ - Vo ~ '
T?*b\_\ Gox AN q‘k. \\k\\(T\A(,\ — M X ( "‘\'T)' ) )(* - 4‘88 ’,q %,L — At
A S B3IS veo =
5 bt ” = © \lob = (-’[; \)
A{_ 4\ FEN +
";": Osdb b
e
\(.7 PAR"
\s > ,“\&k %
Ll = = ————— \o].;;X/) AL e
{ € ’ X 3y &\T
- o). XY 1Y
5 " ‘ i 2l
= Lt TFe x G ¥ leb g £ = 3
' 1Y - R Q e FORM (285) ..
41 ate 4‘\‘*."‘-“‘ Taba ht"h(-”-(" — 3-.\1..”\4-‘1‘-0.cfl>- 132.\3’)§ T)H\, ¥

Page B-4-4

FOSTER WHELLER CORPORATION 110 SOUTH ORANGE AVE., LIVINGSTON, lN.J.
PR

BY. e DATE ___.____. SUBJECT. .

e - -

. . 7 i o, .
- -“\_“,\'(\ P"“ AN EA Vi L'{\; £ \ % 't~ t - .2.?_)\ - ngl *i g1y = V. vhl™ &)B\ -

th
i - i “ ':’ ;\ ¥
¥ t A 1
\}. v
WA + lO'l‘ "f\‘( lft
( { ik‘s/’
Lo e o | oy b
A - o= bl FEINNY g ay
‘( S - o+ b Lt
'J/’,
3
= ! Lo Ky
R IS R R T
.“o"-“"?' A S L , . -
- ' _ = 2N 42 LS
.= U = Jnt4r b s
i - (O
C h .
— T AN gy b
{0

RN { \ 2 B _ N
\‘\ > Y XK e A § X . v \f}gl+q X ‘\—\;‘; = 0.0‘ r)
3 S CEE L o 13y :
T SRR T T O Bk

N L S N CET T e
b
\"'3\.’»‘ \ \ ‘n
i = L SRR IS N VNN ¢ é Ty
- Lon oy

| W&y
= N0, 3b ) 4 Ke

FORM (28%).47

P&gfi’ B-u-&
FUSTER WHEELER CORPORATION 110 SOUTH ORANGE AVE,, LIVINGSTON, N.]
BY . ___ DATE . ____. SUBJECT. SHEET No.--__Gi--_-OF--.\-}f

R ST SR TS SN 4 L ox L srqso 1h7 - msl.‘i—’fj}
*/—”‘k“‘\' R R AT X 173249 Ay byr T ARAT 7T ,
= L Ny )=2 sz
S U4 o g
R S N R T R
e el
{ ) 2 } 5 -
o N AR “4 — = U b
";‘\fib SR bau BTy NSV \())) ol 4 g
e Geg L SR L o L
N U e S SN
' + b s
c
—— = v b g ._,\c’ ‘i‘
n
Jf\ . WL
T Vg y i | ' > 3 )
A X oo K e X o~ X L2360 A Tyl = O,C\DS
T4 il N G b N3 7 20N s
e *P = » -‘S'-»"}i
. ) : _‘_.'\“_,;“ .'..‘_-‘__ - f L%d:
T M&‘tm.‘r«z\ Chrewd = 23 s - 0 enb = il L8y
{;
SERRY
L b
:i:—‘, - - - ~ ey -’?)‘ l&.\\" S—tt‘y‘fb
[ é’»i_} ,’7"(\ p-’(! A ::,\3 Y]
\ 3 Wl Y T T bLev)
- L = VI x (‘tf—"f N iax x §i s L~ 1
:»_ | | — LR P BN y o N — o - 1— .\)"7 i
:)‘ Y é,h‘lg,‘n_\ b L‘\—ci o= it e — o = ok, b
: A
; : . . ! 4D ~ o7
—{; :’"") '\\ .‘\/( \."V\:{.‘( ; ;..’i“\»tr i\_‘ Lf})I » PR .\.r“l'\_ ‘:)S‘: D
o ‘ - Hlk, x.rat“k N o ? 74 FORM (263).47

Fage B-4-¢
FUDIER WHRELLER CORPORATION
BY

i

110 SOUTH ORANGE AVE,, LIVI}\TGSTON, N.J
I SUBJECT. - o oo e e SHEET NO..__! ____ oF.. L.
CHKD. BY e ___ DATE o e JOB NOue oo
- e ! !\4 /, e 0 \

; LLE Ui
IR
_— " I >/i
VoS &
'y
/’\..'\ - o ‘\:\ 3y '%;\,*’ 1\ ¥
/ — A 1Y R 12 .83
~Te ™ AR

- _."\Kl‘)il'yj:a{""q‘
y by
~
_—r:‘ o TR
" T
W "-Q’:“' ‘ “ ‘! rs
e L ) = et
- AT o 64 o SERS v
i - i 1 ) B
Ve = N - !

oy | i
<y bt ; T
1. = = 21y Hutbiiy A4
+ 05 N
= Sl 19y 41"5#&«
ir\, = J,"\ ) \‘g L

. « " )
L] - e f \ .\- l !
I T Gl e L I X TR -——(sk‘{:fi’:l =
A tefe s 7 VS e b X vy T bHes TS O 1 ¥ XRYS
LS R SRR LY, e

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FORM (285).47

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APPENDIX B-6
GULF GENERAL ATOMIC

REPORT GULP-GA-A12416
N\

GULF GENERAL ATOMIC

- Gulf-GA-A12416

FINAL REPORT

ANALYSIS OF DYNAMIC FLOW STABILITY IN STEAM
GENERATORS FOR THE MOLTEN-SALT BREEDER REACTOR

by

B. E. BoYack

Prepared under
P.0. N24013
Project No. 0540.0000
for
Foster Wheeler Corporation

under
Union Carbide Corporation, Nuclear Division
Subcontract No. 91X-88070C

under
Prime Contract No. W-7405-eng-26
with the
U.S. Atomic Energy Commission

Gulf General Atomic Project 0540 November 20, 1972

GULF GENERAL ATOMIC COMPANY
P.0. BOX 81608, SAN DIEGO, CALIFORNIA 92138
LIST OF SYMBOLS

Cross sectional flow area of a steam generater tube, ft2
Tube inside diameter, ft

Darcy friction factor

Local gravitational acceleration, ft/hr2

Newton constant relating force and mass, lbm-ftllbf-hr2
Steam enthalphy, Btu/lbm

Change in steam enthalphy, Btu/lbm

Number of tube segments

Static pressure, lbf/ft2

Normalized heat flow distribution

Complex variable replacing time

Time, sec

Mass flow rate per tube, lbm/hr

Spatial coordinate along axis of tube, ft

Greek Symbols

0 Tube inclination from vertical, degree
Steam density, 1bm/ft3
Time-dependent heat input, Btu/ft-hr

W Circular frequency, rad/sec

Subscripts

exit exit from steam generator tube

jnlet inlet to steam generator tube

j jth tube segment

iv
3-4,

3-5.

3-7.

3-8.

TABLES

Steady-state operating conditions for the MSBR
steam generator reference design . . . . . . . . .+ . o ..

Design properties for nickel-molybdenum-chromium-iron
alloy (Hastelloy N) - L ] L] L L] » » L] . L] L] L L] [ ] L] . L] L] L] [}

Open loop frequency response of reference design steam
generators for an MSBR operating at 99.68% of rated loads.

Open loop frequency response of reference design steam
generators for an MSBR operating at 79.95% of rated loads.

Open loop frequency response of reference design steam
generators for an MSBR operating at 59.97% of rated loads.

Open loop frequency response of reference design steam
generators for an MSBR operating at 39.89Z of rated loads.

Open loop frequency response of reference design steam
generators for an MSBR operating at 19.947% of rated loads.

Open loop frequency response of reference design steam
generators for an MSBR operating at 99.68% of rated
load. Tube length divided into 23 segments

Open loop frequency response of reference design at full
load, inlet orifice K=120. . . . . . . .+ .

Open loop frequency response of reference design at full
load’ exit Orifice K=20 . L] . . L] . L] LJ L] .. *

vi

20

21

22

23

24

27

29

30
1. INTRODUCTION

Design procedures for steam generators must include the analysis
of instability phenomena in heated tubes. The physical damage and per-
formance degradation associated with flow instabilities such as systen confrol
problems, mechanical vibration or thermal cycling of steam generator tubes
can be serious and must be avoided or reduced to minimize their effects.
Experimental prototype testing of each steam generator design 1s prohibitive,
and the complexity of the physical phenomena precludes simulation through
model testing. Consequently, analytical techniques are useful in evaluating
or predicting the onset of instability phenomena for conceptual and design
studies. A liquid flowing through a heated channel is susceptible to a
variety of destabilizing phenomena. It is useful to categorize these
phenomena at the outset in order to define the scope of this study.
Following Bouré, et al, (Ref. 1), two major classifications may be defined:
static instabilities and dynamic instabilities. Static instabilities
include the flow excursion or Ledinegg instability, boiling crisis due to
ineffective removal of heat from the heated surface, flow pattern
transition instability and the compound relaxation instabilities described
as bumping, geysering and chugging. The fundamental dynamic instabilities

are acoustic oscillations and density wave oscillations.

The static instabillity of primary design impoftance in steam generators
is the excursive ingtability. The criterion for onset of the flow excur-
sion instability is well known (Ref. 1), and prediction techniques have been
de#eloped which are based on the solution of the steady-state conservation

equations for mass, momentum and energy.
Dynamic instabilities are associated with the phvsics of wave phenomena,
density waves for density wave oscillations and pressure (acoustic)
waves for acoustic phenomena. In any real system, both kinds of waves are
present and interact, but their velocities differ in general by one or
two orders of magnitudé, thus allowing one to distinguish between the two
types of instabilities. The acoustic or pressure wave oscillations are
characterized by a high frequency, the period being of the same order of
magnitude as the time required for a pressure wave to travel through the
system. Acoustic waves are not the subjecfi of the pfesent investigation

and will not be discussed further.

Density wave oscillations are common in a variety of equipment
and have been extensively studied during the past fifteen years. These
are low frequency oscillations in which the period is approximately the
order of magnitude of the time required for adensity wave to travel
through the tube. The instability can be initiated by a temporary
reduction or perturbation of inlet fiow to the heated channel, producing
thereby an increase in the rate of enthalpy rise and a reduction in the
average density. The disturbance affects the pressure drop as well as the
heat transfer behavior. For certain geometrical arrangements, operating
conditions, and boundary conditions, the perturbations can require

appropriate phases and become self-sustaining.

Several computer codes exist which can be used to predict the onset
of density wave oscillations in heated channels. Codes developed prior
to 1965 were reviewed in a comprehensive testing program by Neal and Zivi
(Ref. 2). The STABLE-3 code of Jones (Ref. 3) was found to be the
most reliable, predicting the threshold of instability for loop experiments
within 20 percent for about 70 percent of the tests. It should be noted
that the study of Neal and Zivi was restricted to low quality steam
systems, the highest exit quality being approximately 0.2. The STABLE-4
program, modified to permit analysis of boiling tubes with superheat

and renamed DYNAM (Ref. 4), is in use at Gulf General Atomic Company
(GGA). To the author's knowledge, no computer codes have previouslv been
developed to permit analysis of forced circulation loops containing super-

critical steam,.

A conceptual design has been developed for a single-fluid 1000~-MW(e)
Molten-Salt Breeder Reactor (MSBR) power station by Oak Ridge National Lab-
oratory (ORNL). In a single-fluid MSBR the nuclear fuel is carried in a fuel
salt, a molten-salt mixture at temperatures above ~ 930°F. Four primary
shell-and-tube heat exchangers transfer heat from the fuel salt to a primary
coolant salt. Also contained in the primary coolant loop are steam genefators,
receiving controlled flow rates of the primary coolant salt to provide
1000°F outlet steam temperatures. The steam generators operate on a super-
critical pressure steam cycle which was selected because it affords a
high thermal efficiency and also permits steam to be directly mixed with
high pressure feedwater to raise its temperature to v700°F and thereby guard
against freezing of the primary coolant salt in the steam generators.

Foster Wheeler Corporation has been awarded a contract for the conceptual

design of the steam generators for the MSBR.

For this study, the reference design proposed by the Foster Wheeler
Corporation for the steam generators of a MSBR was analyzed to determine
stability with respect to density wave oscillations. This report describes
the method of analysis and summarizes the predicted dynamic stability char-
acteristics of the MSBR steam generator at 100, 80, 60, 40 and 20 percent
of full-rated load. Several parameters including inlet orificing, exit
orificing, and pressure level have been varied to determine their overall

relationship to the steam generator stability.
2. ANALYSIS

2.1 PROBLEM DESCRIPTION

The conceptual design of a steam generator for the ORNL 1000-MW(e)
reference steam cycle has been developed by the Foster Wheeler Corporation.
The unit, roughly the shape of an "L," is vertically oriented with molten
salt flowing on the tube side and steam in the supercritical state on the
tube side. Steady-state operating conditions for a single representative
tube have been prepared by Foster Wheeler (Ref. 5). The steam inlet and
exit conditions at the full and partial loads for which stability
calculations have been made are presented in Table 2-1. The tube
material is Hastelloy N having the properties outlined in Ref. 6 and
reproduced here in Table 2-2. Neither the properties of the molten
salt nor detalls of the tube layout are presented herein. The analytical
technique used to predict the onset of density wave oscillations considers
‘a representative tube only and does not explicitly require data describing

overall steam generator geometry or shell-side interactions.

2.2 LITERATURE SURVEY

The prediction of the onset of density wave oscillations in heated
channels or tubes containing two-phase flow has been reviewed in
Ref. 1. Previous studies have included both analytical and experimental
investigations to enable the prediction of the onset of the instability.
As part of this iInvestigation a literature survey was conducted
for the purpose of determining if previous'investigations have been
conducted to investigate the onset of density wave oscillations in systems
containing a fluid in the supercritical thermodynamic state. It is
emphasized that the literature survey conducted was limited and should

not be considered to be complete,
TABLE 2-1
STEADY-STATE OPERATING CONDITIONS FOR THE

MSBR STEAM GENERATOR REFERENCE DESIGN

Percent of Rated Load
99.68 79.95 59.97 39.89 19.94
Flow Rate/Tube (l1bm/hr) 2538. 1900.4 | 1378.1 910.6 463.1
Inlet Pressure (psi) 3730. 3690. 3658. 3632. 3613.
Exit Pressure  (psi) 3599.1 3599.9 3600. 3599.7 3599.8
Inlet Enthalphy (Btu/lbm) 770.5 771.8 772.9 773.9 774.6
1418.5 1466. 1490.9 1496.7 1485.

Exit Enthalphy (Btu/lbm)

TABLE 2-2
DESIGN PROPERTIES FOR NICKEL-MOLYBDENUM-CHROMIUM-IRON ALLOY

(HASTELLOY N) *
Allowable Modulus Mean Coefficient :
Temperature Stress of of Expansion Thermal
(°F) (psi) Elasticity (in./in.-°F) Conductivity
(psi) (Btu/ft-hr-°F)
X 1076 X 108 |
100 25,000  31.3 6.6
200 24,000 30.6
300 23,000 30.0
400 21,000 29.5 6.45 7.4
500 20,000 29.0
600 20,000 28.5 6.76 8.3
700 19,000 28.1 | |
800 18,000 27.7 7.09 9.2
900 18,000 27.2 | 9.3
1000 17,000 26.7 7.43 10.4
1050 26. 4 | |
1100 13,000 26.3 11.1
1150 26.1 12.1
7

1200 6,000 25, 7.81 11.7

Density, 0.317 1b/in.3 at room temperature

Specific heat, 0.095 Btu/lb °F at room temperature;
0.139 Btu/1b °F at 1200°F
A considerable body of literature has been developed which describes
the extreme property variations which occur near the critical point. A
recent review has been presented by Hall (Ref. 7). As the critical point
is approached large changes in density, specific heat, viscosity and
thermal conductivity are observed. It 1s precisely these variations
which have been postulated as forming the initiating agency for thermo-

hydraulic flow oscillations observed during several experimental investigations.

Mechanized (computer) literature searches were obtained from two
sources. The searches were conducted by the Nuclear Safety Information
Center (NSIC) located at Oak Ridge National Laboratory and the Heat
Transfer and Fluid Flow Service (HTFS) of the British Atomic Enérgy
Research Establishment, Harwell. Citations of publications dealing with
density wave oscillations in channels or tubes contailning steam in the
supercritical state were requested. Both mechanized surveys indicated

that the available data are very limited.

A series of studies have been conducted at Oklahoma State
University to investigate instabilities encountered during heat transfer
to a supercritical fluid. Although experimental programs at Oklahoma
State University have concentrated on natural circulation loops, associated
literature surveys have been more broadly directed to include both forced
and natural circulation test programs. Many fluids, including steam,
helium, hydrogen, etc., have been considered. Cornelius (Ref. 8) has
conducted an extensive review of the literature available prior to 1965
and reports that the literature contalns reference to two modes of
oscillations. The first 1s an acoustic oscillation while the second, having
a frequency several orders of magnitude less than the acoustic oscillations,
is attributed to a "boiling like" phenomenon. The nature of this second
oscillation is clarified by Walker and Hardon (Ref. 9), who focused on the
prediction of the threshold of these oscillations assuming that the '"density

effect" is the sole driving mechanism for the oscillations. They assumed the
density effect is fihe consequence of the non-linear physical relationship
between the enthalvhy and density of the fluid. The density effect model
formulated by Bouré (Ref. 10) was used to predict the flow instability
threshold, and excellent agreement with experiment was obtained on a natural

circulation loop with Freon-114 as the working fluid.

A similar investigation was reported by Zuber (Ref. 11) in 1966.
Zuber's report contains an extensive literature survev and describes an
analysis to predict the onset of oscillations in flow systems containing
fluid in the supercritical state. The method also follows Bouré (Ref. 10)
in that similar assumptions and formulations are used. The problem is
analyzed by perturbing the inlet.flow, linearizing the set of governing
equétions (conservation of mass, momentum and energy plus an equation
of state) and integrating them along the channel to obtain the characteristic
equation. Zuber describes three mechanisms which can induce thermo-
hydraulic oscillations at supercritical pressures. One is caused by the
variation of the heat transfer coefficient at the pseudo-critical point.

The second 1is caused b& the effects of large compressibility and the result-
ant low velocity of sound in the critical region. The third mechanism is
caused by the large variation of flow characteristics brought about by

density variations of the fluid during the heating process.

2.3 METHOD
2.3.1 General

The analysis of density wave oscillations in heated tubes can proceed
alofig one of several paths, depending upon the tvpe and detail of information
required. A requirement for extensive information concerning the time-
dependent physical processes associated with the instability would require
the simultaneous solution of a coupled set of non-linear, time-dependent
partial differential equations. Generally, such detail is not necessary and
only an answer as to whether an instability can occur under specified over-
ating conditions is required. If an instability is predicted, the design is

adjusted to eliminate such an occurrence.
For this analysis only a yes or no answer to the question of the possible
occurrence of density wave oscillations in the steam generators of the MSBR
is sought. The classical method of stability analysis is employed whereby the
governing equations are linearized and the behavior of the resulting equations
to small perturbations is determined (Ref. 12), The mathematical model
used to define steam generator instabilities is based on a single steam tube
located in an array of tubes connected in parallel between headers. It is
assumed that a flow perturbation might occur in this single tube while the
remalning tubes operate normally. Therefore, the pressure drop across the
tube under investigation will be constant since it 1is established by the
steady flow through the remaining tubes and all are connected to common
headers. In an actual steam generator it may be expected that once a reason-
able magnitude flow oscillation starts in one tube, others will be affected
and the mode of oscillation may be very complex owing to the many degrees of

freedom in the complete system.

The question of accuracy naturally arises when linearization is performed
on a set of equations to obtain a solution. In this case the linearization
leads to a conservative answer from the standpoint of safe performance. The
linearized analysis will predict the threshold of an instability due to a
small system perturbation. In the physical case, small perturbations tend to
be damped or lead to small limit cycle oscillations which are often difficult

to detect without precision instrumentation.

There are several methods for determining the stability of a system
described by a set of linear differential equations. A particularly useful
method 1s based on feedback control theory, and only a small part of this highly
developed fleld need be used for solving the problem of density wave oscilla-
tions in steam generators. The primarv reasons for choosing feedback control
theory are the completely systematized nature of the procedure and because the
method reduces the set of partial differential equations to ordinary differential

equations, a significant reduction in complexity for numerical analysis.
The steps in the solution of the density wave oscillation problem are:

The equations expressing the conservation of mass, momentum, and
energy and the equation of state are written in terms of suitable

variables.

The governing equations are linearized by assuming each variable

to be composed of a part which is at most a function of position

‘plus a perturbation term which is a function of vosition and time.

The Laplace transférm of the equations is taken,which has the effect
of transforming the original partial differential equations into
ordinary differential equations by replacing the time derivative

with a complex frequency variable.

The resulting ordinary differentiai equations are integrated over
small increments of length by assuming the system parameters are
constant over these small increments of length. It {is presumed that
the steady state solution for the steam generator is known since
parameters from the state of equilibrium are required to determine
the transient behavior. -Accuracy increases as the length of the

segment decreases.

The feedback control system representation of the steam

generator is set up. For the density wave oscillation problem,

the selection of input and output variables is arbitrary but
certain choices of these quantities turn out to be more convenient
than others. Specifically, pressure perturbations are used in this
study. The feedforward and feedback transfer functions are defined,

as is the dpen loop transfer function.

10
6. The magnitude and phase angle of the open loop transfer
function is calculated as a function of the input wvariable
which is assigned a magnitude of unity and frequencies
ranging from zero to a value considerably above the

expected frequency of density wave oscillations.
7. The values for the open loop transfer function are plotting
in polar form (Nyquist diagram) and the Nyquist criterion

1s used as the basis for assessing stability.

2.3.2 Method Development for Molten-Salt Steam Generators

Gulf General Atomic has developed a computer code which is used for
prediction of the onset of density wave oscillations in once~through steam
generators containing liquid, two-phase and superheated steam. The GGA
code is a modified and extended version of the'séries of STABLE codes
developed at ¥nolls Atomic Power Laboratory (Ref. 3). The STABLE codes
are applicable only to the analysis of boiling channels cbntaining single~-
phase liquid and two-phase steam. The STABLE~IV code was extended at GGA
to permit analysis of sfeam generators with superheated steam and the modified
code designated as DYNAM (Ref. 4).

In order to use DYNAM to predict the dymamic stability characteriétics
of the reference design steam generators of the MSBR, a modified version of
the code, hereafter referred to as the DYMSBR code, was developed. The
significant elements in the modification program are described in the fol-

lowing sections.

2.3.3 DYMSBR Code Structure

For the purpose of this analysis, the significant features of the tube-
side supercritical steam are that the fluid is compressible and can be
described as consisting of a single phase. The analytical formulation of

the DYNAM code permits analysis of steam generators having a single-phase

11
liquid at the tube inlet, two-phase flow through an intermediate region of
the heated tube and superheated steam at the exit. Thus, of the three flow

regimes only the superheat regime is a single-phase compressible fluid.

In order to develop an orderly code structure, that portion of the
DYNAM code associated with the superheat regime was extracted from DYNAM and
modified to include new input and output routines. As previously indicated,
the revised code was named DYMSBR and can be used to predict the dynamic
stability characteristics of a steam generator using supercritical steam on

the tube side.

2.3.4 Governing Equations

The equations governing the thermodynamic and hydrodynamic processes
occurring in a steam generator can be derived from the principles of con-
servation of mass, momentum and energy. An additional equation, the equation
of state,1s required if a compressible fluid is belng analyzed. If the
time-varying flow processes are considered to be one-dimensional in space,
considerable simplification of the governing equation is possible. The

resultant conservation equations are presented below:

3p . 1 oW
= tirap =0 (2-1)
2 £ 2
Jdp 1 w1 3 (W wl,_ s _
52 " g A ot ' g AZ 3z | b T 2g D A7 |5 T, P ee® (2-2)
oH oH |
Apat+W-§;—¢ (2-3)

12
Equations 2-1, 2-2, and 2-3 are derived from thé principles of conservation
of mass, momentum and energy, respectively. The variables W, p, p, and H
are mass flow rate per tube, density, pressure and enthalpy of the steam
flow. The variables z and t are the distance along the heated tube and time,
A 1s the tube cross-sectional area and D the tube inside diameter. The angle
6 is the tube inclination from the verfical, f 1s the Darcy friction factor
and ¢ is the heat input per unit length of tube. For this analysis the
equation of state considers the density to be a function of enthalpy only,

thus eliminating consideration of acoustic effects.

The procedure for deriving the final equation forms has been outlined in
Section 2.2.1. The equations are first linearized, the Laplace transform 1is
taken, and the resultant equations are integrated over small spatial increments.
The final equations obtained are identical to those reported in Ref. 4. The
final equation forms are algebraically complex and thus are not repeated here.
The final dependent variables are the perturbation quantities for pressure, mass
flow rate and density. The coefficients in the perturbation equation consist
of combinations of variables which include geometry factors, tube material
properties, and steady-state flow distributions. To this point analyses of
channels containing superheated steam and supercritical steam are identical.
However, the evaluation of the steady-state flow distributions and coefficients

in the perturbation equations marks the separation of the two analyses,

2.3.5 Steady-State Flow Distributions

A steady-state solution is required for evaluation of coefficients
appearing in linear perturbation forms of the governing equations. Several
features of the steady-state solution require comment. Since the tube-side
fluid in the steam generator 1s steam in the supercritical thermodynamic
state, special care must be taken to insure that the evaluations of thermo-
dynamic and transport properties are accurate. For this investigation a

set of computer routines published by the ASME (Ref. 13) were used. These

13
subroutines are based on the 1967 IFC Formulations which are described in

the 1967 ASME Steam Tables (Ref. 14). The effective film coefficient for
supercritical steam has been evaluated following the formulation of Ref. 15.
The friction factor for supercritical steam has been evaluéted using Deissler's
formulation (Ref. 16). |

For a given steady-state calculation the tube is divided into a number

of segments. For each segment the following parameters are determined:

1. Thermodynamic properties: enthalpy, temperature, pressure, and

specific volume.

2. Transport properties: viscosity, thermal conductivity, and specific

heat at constant pressure.

3. Pressure drop components: elevation, friction, momentum and orifice

losses.
4. Convective film coefficient, wall temperature.

The procedure for obtaining the steady-state solution is direct when
mass flow rate, inlet pressure, heat flow distribution, and enthalpy at
inlet and exit are specified. The enthalpy rise across the jth tube

segment can be calculated from

(2-4)

14
where F& is the normalized heat flow distribution and N is the number of
tube segments. To eliminate the need for an extensive iterative solution,
thermodynamic and transport properties are calculated using the enthalpy at
the midpoint of the j th segment and the pressure at the exit of the j-1 segment.
The wall temperature at the segment midpoint is calculated by an iterative
procedure which begins with the evaluation of the convective film coefficient
using values of wall temperature from the previous iteration. The heat

- transfer is calculated and compared to the known heat transfer, and, if
necessary, the wall temperature adjusted to start the next iteration. Once
calculated, the wall temperature is used to calculate the friction factor,
permitting evaluation of the friction pressure drop across the jth segment.
The procedure is repeated for each segment, and the required steady-state

distributions are stored for later use in calculating coefficients for the

perturbation equations.

Specific parameters from the steady-state solution required for
coefficient evaluation are the convective film coefficient, the derivative
of the specific volume with respect to enthalpy, specific volume, pressure
drop due to friction, specific heat, viscosity, thermal conductivity and

enthalpy.

15
2.3.6 Frequency Response Analysis

As previously indicated, once the governing linear perturbation
equations have been derived, a feedback control representation of the steam
generator 1s set up. An open loop transfer function is determined which
depends only on the parameters of the system (e.g., steady-state distributions,
tube geometry and physical properties, etc.) and is not related to either the
initial conditions or the forcing function. Since the transfer function
is.a complex variable quantity, the output generally differs from the input

in both magnitude and phase.

A steam generator is stable 1f the system will return to equilibrium
conditions after being perturbed by some external excitation. The stability
of a system can be completely determined from the transfer function by the
following rule: for a system to be stable, the transfer function cannot
have infinities (poles) in the right half of the complex S-plane. It 1is
noted that when Laplace transforms are taken of the original time-dependent
partial differential equations governing the flow of éupercritical steam
through the steam generator, the independent varlable time is replaced by
the complex variable S. The utility of the transfer function methodology
is based on the above rule since it is only necessary to determine the poles
of the transfer function to decide if a system 1s stable, rather than com-

pletely finding the solution of the governing equations.

Several methods are avallable for determining system stabllity using
the rule described above. The method of Nyquist is the simplest and most
direct method for determining system stabllity. However, a description of
the criterion and its relation to the transfer function representation of the
system of governing perturbation equation for a steam generator is beyond
the scope of this report. A simplified description of the method and its
application to steam generators is given in Ref. 12. The specific application
of the method to the system of equations solved in this analysis 1s presented
in Ref. 4.

16
In brief, the procedure for determining system stability is as follows:

l. Let S = iw with w the circular frequency and make a polar plot

of the open loop response as w takes on values from -= to +x.

2. Determine if a vector from the -1 point on the real axis to the
trace of the open loop transfer function makes one or more complete
revolutions as the trace of the open loop transfer function is

developed.

3. The system is unstable if the vector makes one or more clockwise

revolutions as w proceeds from -= to +=,

4. In many cases system stability can be evaluated simply by determin-
ing where the trace of the open loop transfer function crosses the
real axis. If the trace crosses the real axis to the left of the
-1 point the'éystem is unstable,and if to the right of -1 the system
is stable.

Example Nyquist diagrams are presented in Fig. 2-1 for the trace of the open

loop transfer function and 0 < w < + =,

It is emphasized that 1if there is any question about stability when
determined by the above rules, the complete Nyquist criterion must be used.
For each of the cases examined in this investigation, it was possible to

determine system stability using the simplified procedures listed above.

17
S-PLANE

_'1P

TRACE OF OPEN LOOP
TRANSFER FUNCTION

< W < oo

STABLE UNSTABLE

Figure 2-1. Example Nyquist diagrams for polar plots of the open loop
transfer function

18
3. RESULTS

The stability of the reference design steam generators for the Molten
Salt Breeder Reactor has been analyzed at 99.68, 79.95, 59.97, 39.89 and
19.94% of full rated load. Operating conditions and design parameters have
been provided by Foster Wheeler Corporation in Refs. 5 and 6. Individual
steam generator tubes are made of Hastelloy N having design properties
listed in Table 2-2, Tube length is 114 ft, tube inner diameter is 0.5 in.,
and tube wall thickness is 0.125 in, Inlet and exit orifice pressure
losses are taken equal to one velocity head or 0.85 psia at the inlet and
5.8 psia at the exit. Steady-state operating conditions for the reference
design at each load rating are listed in Table 2-1, The heating distribu-

tion along the tube for each load rating is given in Ref. 5.
3.1 TUBE-SIDE DYNAMIC STABILITY AT FULL AND PART LOAD

The open loop frequency response of the reference design is presented
in Tables 3-1 through 3-5 for 99.68, 79.95, 59.97, 39.89 and 19.94% of
rated load, respectively. In each case the reference design is stable
with respect to density wave oscillations as measured by the Nyquist
stability criterion. Nyquist plots are presented in Fig. 3-1 for the
reference design at 99.68 and 19.947% of rated load, It is clear that the
system is highly stable as the trace of the open loop transfer function does
not approach the region of the imaginary axis. Further, at reduced flow and
heating rates the trace of the open loop transfer function is shifted along
the real axis away from the -1 point, indicating an increase in stability
at the lower flow and heat rates. However, it is emphasized that the system

is very stable and that the above effects are not significant for the reference

design,

19
TABLE 3-1

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR OPERATING AT 99.68%

OF RATED LOAD

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .4105+04 .0000
. 2500400 .2650+04 .1680+00
. 5000+00 . 2650+04 .3360+00
.1000+01 .2650+04 .6719+00
.5000+01 .2650+04 .3360+01
.1000+02 . 2650404 .6718+01
. 2500402 . 2650+04 .1678+02
.5000+02 .2652+04 .3345+02
. 7500+02 .2656+04 .4989+02
.1000+03 .2660+04 .6600+02
.1500+03 «2674+04 .9681+02
. 2000403 .26914+04 .1251+03
.5000+03 .2860+04 .2143+03
.1000+04 . 3109+04 . 7495+02
(a)

Tube length divided into 60 segments

20

TABLE 3-2

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR OPERATING AT 79.95%

OF RATED LOAD

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .4876+04 .0000
.2500+00 .3119+04 .4114+00
.5000+00 .3119+04 .8228+00
. 7500+00 .3119+04 .1234+01
.1000+01 .3119+04 .1646+01
.1500+01 .3119+04 .2468+01
.2000+01 .3119+04 .32914+01
.5000+01 .3119+04 .82274+01
. 7500401 .3119+04 .1234402
.1000+02 .3119+04 .1645+02
.1500+02 .3120+04 . 2466402
.2000+02 .3120+04 .32874+02
.5000+02 .3128+04 .3156+02
.1000+03 . 3155+04 .1589+03
(a)Tube-length divided into 66 segments

21

TABLE 3-3

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR OPERATING AT 59.97%

OF RATED LOAD

Open Loop ReSponse(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .5698+04 .0000
.2500+00 .3533+04 .1030+01
. 5000+00 .35334+04 2059401
. 7500+00 .35334+04 .3089+01
.1000+01 .3533+04 .41194+01
.1500+01 .35334+04 .6178+01
.2000+01 .35334+04 .8237401
.5000401 .3533+04 .2059+02
. 7500401 . 3534404 .3088+02
.1000+02 . 3534404 .4116+02
.1500+02 .3536+04 .6169+02
. 2000402 .3538+04 .8216+02
.5000+02 . 3563+04 .2026+03
.1000+03 .3647+04 . 3860403
(a)

Tube length divided into 69 segments

22

TABLE 3-4

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR QPERATING AT 39,89%

OF RATED LOAD

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .6653+04 .0000
.2500+00 .4438+04 .2058+01
.5000+00 .4438+04 4116401
. 7500400 .4438+04 .6174+01
.1000+01 .4438+04 .82324+01
1500401 .4438+04 .1235+02
.2000+01 4438404 .1646+02
.5000+01 .4439+04 .4115+02
. 7500+01 . 4440+04 .6169+02
.1000+02 .44414+04 .8219+02
.15004+02 4446404 .1230+03
.2000+02 - 4453404 .1636+03
.5000+02 . 4530404 .3951+03
.1000+03 4777404 .69974+03
(a)

Tube length divided into 74 segments

23

TABLE 3-5

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR OPERATING AT 19.94%

OF RATED LOAD

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .8092+04 .0000
.2500-01 .6610+04 .4106+00
.5000+00 .6610+04 .8212+01
.1000+01 .6610+04 .1642402
.5000+01 .6614+04 .81944+02
.1000+02 .6627+04 .1628+03
+2500+02 .6716+04 .3887+03
.5000+02 .6990+04 .6602+03
. 7500+02 « 71331404 .7561+03
.1000+03 . 76374+04 .6931+03
.1500+03 . 7964404 .3466+03
.2000+03 .7992+04 .34254+02
.5000+03 .7626+04 -.5160+03
.1000+04 . 7156+04 ~.7167403
(a)

Tube length divided into 86 segments

24

S~PLANE

200 g~
99.68% OF
RATED LOAD
100 =
w = 1000 RAD/SEC
0 4 ' i I
0 2400 2800 3200
800 r
40O P 19.94% OF
RATED LOAD
0 A—
6000
-400 |
-800 L w = 1000 RAD/SEC

Figure 3-1. Nyquist diagrams of the open loop transfer function for MSBR
steam generators, reference design

25
In density wave oscillations, disturbances travel with the same
velocity as the fluid in contrast to acoustic type oscillations where
disturbances travel with the local speed of sound in the medium. The
period of a density wave oscillation in a steam tube can be roughly
estimated by dividing the tube length by the mean velocity of the steam
in the tube. For the full and partial loads listed above, density wave
oscillations, if they exist, are estimated to have a frequency varying
from approximately 0.5 radians/sec at full load to 2.0 radians/sec at
207% of full load. Thus, the frequency ranges as presented in Tables 3-1
through 3-5 clearly encompass the possible frequencies of density wave

oscillations.
3.2 SUPPORTING CALCULATIONS

The accuracy of the analysis is dependent upon the number of segments
into which the total tube length is divided. As the number of segments
becomes large, or conversely as the length of each segment becomes very
small, the accuracy of the method is expected to improve. The results
reported in Section 3.1 were obtained by specifying non-uniform segment
lengths identical to those reported by Foster Wheeler in Ref. 5. The total
number of segments used for the 99.68, 79.95, 59.97, 39.89 and 19.94% of
rated load cases were 60, 66, 69, 74, and 86, respectively. These tube
divisions were determined by Foster Wheeler Corporation to yield highly
accurate steady-state solutions. The effect of segment length on accuracy
was investigated by examining the 99.68% of rated load case with tube
length divided into 23 segments. The resultant frequency response is
presented in Table 3~6. It can be seen that the effect of the larger
segment length was to shift the trace of the frequency response curve along
the real axis away from the -1 point. Clearly accuracy is a second-order
effeét when compared to the increased stability predicted as flow and

heating rates decrease, Further investigation of the effect of segment

26
TABLE 3-6

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN
STEAM GENERATORS FOR AN MSBR OPERATING AT 99.68%
OF RATED LOAD. TUBE LENGTH DIVIDED INTO 23 SEGMENTS

Open Loop Response

Frequency
(radians/sec) Real Part Imaginary Part

.0000 .4129+04 0000

. 2500400 .2789+4+04 .1003+00
. 5000400 .2789+04 .2006+00
. 7500400 . 2789404 . 3010400
.1000+01 . 2789404 .4013+00
.15004+01 . 2789404 .6019+4+00
.2000+01 .2789+04 . 8025400
. 5000+01 .27894+04 .2006401
. 7500401 .2789+04 .3009+01
.1000+02 . 2789404 4012401
.15004+02 .27894+04 .6016+01
. 2000402 . 2789404 .80174+01
.5000+02 2791404 1994402
.1000403 .2797404 .39134+02

27

length on accuracy was not deemed necessary since the effect of increased
segment length was to indicate a more stable system. ‘Also, the segment
lengths reported by Foster Wheeler (Ref. 5), and used in this study, have

been found by Foster Wheeler to be adequate.

Although tube-side dynamic instabilities are not predicted for the
reference design, several brief parameter studies were run to determine
the effect of inlet and exit orificing and the effect of system pressure
level on system stability. The results for an inlet orifice resistance
coefficient K = 120 are presented in Table 3-7 and has an exit orifice
resistance coefficient of K = 20 are presented in Table 3-8. The corre-
sponding pressure decreases across the inlet and exit orifices are 102 psia
and 116 psia, respectively. It can be seen that the effect of inlet
orificing i1s to make the system less stable while the effect of the exit
orificing is to make the system more stable, This trend is just opposite
the observed stability trends for two-phase flows where inlet orificing
tends to stabilize and exit orificing makes the system less stable. An
additional.calculation was made with an inlet orifice value K = 180 (a
pressure decrease of approximately 153 psia), and the minimum real value
of the open loop transfer function was 15,64, Although the margin of
stability has been decreased by the inlet orificing, the reference design

remains very stable with respect to density wave oscillations,

Zuber (Ref. 11) in his analytical investigation of thermally induced
flow oscillations in the supercritical thermodynamic region predicts similar
trends for twd—phase and supercritical regimes. Thus, inlet orificing is
reported as stabilizing the flow and frictional pressure losses and exit
orificing leads to a less stable flow. Zuber further notes that his conclu-

sions and results are new and have not yet been verified against experimental

data.

28
TABLE 3-7

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN

AT FULL LOAD, INLET ORIFICE K=120

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 | .3480+02 -.0000
+2500+00 .23394+02 .8505-03
.5000+00 .2339+02 .1701-02
. 7500+00 «2339+02 .2552-02
.1000+01 .2339402 .3402-02
.1500401 .2339+4+02 .5103-02
.2000+01 . 2339402 .6804-02
.5000+01 .2339+02 .1701-01
. 7500401 .2339+02 .2551~-01
.10004+02 .2339402 .3401-01
.1500+02 .2339+02 +.5100-01
.2000+02 . 2340402 .6797-01
.5000+02 .23414+02 .1690-00
.1000+03 . 2347402 .3317-00
(a)

Tube length divided into 23 segments

29

TABLE 3-8

OPEN LOOP FREQUENCY RESPONSE OF REFERENCE DESIGN

AT FULL LOAD, EXIT ORIFICE K=20

Open Loop Response(a)

Frequency
(radians/sec) Real Part Imaginary Part
.0000 .8294+04 .0000
.25004+00 . 5303404 .3678+00
.5000+00 .5303+04 .7356+00
. 7500400 .5303+04 .1103+01
.1000+01 .3303+04 1471401
.1500+01 .5303+04 .22074+01
.2000+01 .23034+04 .29424+01
.5000+01 .5303+04 .7356+01
.7500+01 .3303+04 .1103+02
.1000+02 .3303+04 1471402
.15004+02 .53034+04 .2206+02
.2000+02 «2304+04 .2939+4+02
. 5000+02 .33104H04 . 7309402
.1000+03 .5333404 .1434+03
(a)

Tube length divided into 23 segments

30

Finally the effect of system pressure level on the stability of the
reference design was examined. It was found that the reference design was
less stable at lower pressures (3530 psia) than at higher pressures
(3930 psia), but that the effect was of second-order importance when
compared to the effect of either orificing or reducing flow and heating

rates. -

31
4. SUMMARY AND CONCLUSION

The DYNAM computer code has been modified to permit the analysis of
dynamic instabilities, specifically density wave oscillations, in the
reference design steam generator for the Molten-Salt Breeder Reactor.
Since the analysis is based on the solution of linear perturbation forms
of the conservation equations for mass, momentum, and energy, the primary
result of the analysis is to indicéte by a 'yes' or 'no' result whether
the reference design is stable or unstable with respect to density wave

oscillations.

It is emphasized that the available experimental data and analytical
studies are very limited. Thus, it has not been possible to verify the

solution against either test data or other analyses,

The reference design appears to be highly stable. Analyses of the
effect of inlet orificing, exit orificing, and pressure level indicate
trends opposite to those observed in two-phase systems. For the reference
design, increased inlet orificing and increased firessure level have been
found to decrease the stability margin., Exit orificing, reduced pressure

level or reduced flow and heating rates appear to increase system stability,.

32
lOl

11.

REFERENCES

Bourgé, J. A., Bergles, A. E., and Tong, L. S., "A Review of Two-
Phase Flow Instability," ASME paper 71-HT-42, ASME-AICHE Heat
Transfer Conference, Tulsa, Oklahoma, Aug. 1971.

Neal, L. G., and Zivi, S. M., "The Stability of Boiling-Water
Reactors,” Nuc. Sci. Eng., 30, 25 (1967).

Jones, A. B., "Hydrodynamic Stability of a Boiling Channel ,"
Knolls Atomic Power Iaboratory reports Part I, KAPL-2170,
Oct. 2, 1961; Part II, KAPL-2280, Apr. 20, 1962; Part III,
KAPL-2290, June 28, 1963; Part IV, KAPL-3070, Aug. 18, 196k.

Efferding, L. E., "DYNAM - A Digital Computer Program for Study
of Dynamic Stability of Once-Through Boiling Flow and Steam
Superheat,"” USAEC Report GAMD-8656, Gulf General Atomic, 1968.

Cox, J. F., Foster Wheeler Corporation, "Full and Part-Load

Operating Conditions for the Reference Design Molten-Salt

Steam Generator," unpublished data.

Cox, J. F., Foster Wheeler Corporation, "Properties of Materials
Used in Reference Design Molten-Salt Steam Generator,"
unpublished data.

Hall, W. B., "Heat Transfer Near the Critical Point," in
Advances in Heat Transfer, Irvine, T. F., Jr. and Hartnett,

J. P. (ed.), Academic Press, New York, 1971, p. l.

Cornelius, A. J., "An Investigation of Instabilities Encountered
During Heat Transfer to a Supercritical Fluid," Argonne National
Iaboratory Report ANL-7032, April 1965.

Walker, B. J., and Harden, D. G., "The Density Effect Model:
Prediction and Verification of the Flow Oscillation Threshold
in a Natural-Circulation Loop Operating Near the Critical
Point," ASME paper 67-WA/HT-23, ASME Winter Annual Meeting,
Pittsburgh, Pennsylvania, November 1967.

Boure, J. A., "The Oscillatory Behavior of Heated Channels,"
Part I and II, French Report CEA-R 3049, Grenoble, France, 1966.

Zuber, N., "An Analysis of Thermally Induced Flow Oscillations
in the Near-Critical and Supercritical Thermodynamic Region,"
National Aeronautics and Space Administration, Marshall Space
Flight Center Report No. NAS 8-11k22, May 25, 1966.

Katz, R., "The Analysis of Density Wave Oscillations in Steam

Generators by Means of Feedback Control Theory," Gulf General
Atomic Report Gulf-GA-A12228, Aug. 1, 1972.

33
13.

1k,

15.

16.

McClintock, R. B., and Silvestri, G. J., "Formulations and
Iterative Procedures for the Calculation of Properties of
Steam," ASME Publication H-17, American Society of Mechanical
Engineers, New York, New York, 1968.

Mayer, C. A., et al., "1967 ASME Steam Tables - Thermodynamic
and Transport Properties of Steam Comprising Tables and Charts
for Steam and Water," American Society of Mechanical Engineers,
New York, New York, 1967.

Swenson, H. S., Carver, J. S., and Kakarala, C. R., "Heat
Transfer to Supercritical Water in Smooth-Core Tubes," ASME
paper 64-WA/HT-25, Winter Annual Meeting, New York, Sept. 196L.

Deissler, R. G., "Heat Transfer and Fluld Friction for Fully
Developed Turbulent Flow of Air and Supercritical Water with
Variable Fluid Properties,” Transactions of the ASME, Jan. 195k,

pp. 73-85.

34
APPENDIX C-1
TUBESHEET HEADER ASSEMBLY

STRESSES AT SECTIONS 1-1 THROUGH 7-7
TUBESHEET STRESS AT TUBESHEET CENTER (SECTiQN_l—l)

(FOR DESIGN PRESSURE, p=4000 PSI)

Element 5. o - 5700 PS/.

391 4780 5111

72 ~3822 ~3826

68 ~2432 ~2440 I
440 - 418 - 439 ‘

64 435 424 \Y\
60 1410 1397

56 2408 2396

52 3490 3480

48 | 4731 4733  epeo pss.
344 5616 5946

-5900Ps/.

ME

| 6_7&0 PJ/.I

SS THROUGH TUBESHEET

PLOT OF STRE

6r(ToP) =-5700 psi
O . (BOTTOM) = 6700 psi
G'r(Avc.) = 1/2(6700-5700) = 500 psi

O, = * (5700 + 500) = * 6200 psi

GT(TOP) =-5900 psi
GT(BOTTOM) = 6900 psi
O (AVG.) = 1/2(6900-5900) = 500 psi

Gy = +(5900 + 500) = * 6400 psi

(Pm+3b)

Dooe (_1_1
(A) PRIMARY GENERAL MEMBRANE STRESS INTENSITY (SECTION 1-1)
(AT T.S. CENTER FOR DESIGN p=4000 PSI)

s; = P/h @b x /02 + (5

S5 P/Zh[vQAp X R/t)2 + (Er)i + 5} + 2p;h/p

Use the larger of 57 or Sj.

where: p/h = reciprocal of ligament eff. = 1.0/0.476 = 2.10
Lp = differential press. across plate = 4000 psi
pi = pressure inside tube hole = 4000 psi

R = radial distance from centerline plate to section of interest = 2.063"

t = plate thickness = 14"

Qv
I

radial stress averaged through the depth of the equivalent solid plat:

6,.(AVG.) + h/p ( (p-h) / h) p;

= 500 + 2096
= 2596 psi

S, = 2.1,/34.74 x 10% + 673.92 x 10% = 5590 psi

S, = 1.05 [ 2662 + 2596 + 3808] = 9520 psi

O
Maximum Average Plate Temperature < 1075 F

Page C-1-2
(B) PRIMARY MEMBRANE PLUS BENDING STRESS INTENSITY (SECTION 1-1)
(AT T.S. CENTER FOR DESIGN p=4000 PSI)

S = k p/h Gi

where: K = stress intensity factor
for top of t.s., § =0 /0y = 5700/5900 = .97, K = 1.0
for bottom of t.s., E5= 6700/6900 = .97, K = 1.0

dl = 5; orCTT, whichever is hrger.

W
I

1.0 x 2.1 x 5900

i

12,390 psi (TOP)

1.0 x 2.1 x 6900 = 14,490 psi ( BOTTOM)

W
I

Page C-1-3
(C) PRIMARY PLUS SECONDARY STRESS INTENSITY (SECTION 1-1)
(AT CENTER OF T.S.)

For Reactor Scram Transient Plus Opérating Pressure:

where K = stress intensity factor

for top of t.s., F>= -520/-780 = 0.67, K = 1,015
(E1. #390) |

for bottom of t.s., p,= 4220/4370 = 0.97, K = 1.0
(E1. #343)

61 = 5; 0rj5T, whichever is larger.

5]
il

1.015 x 2.1 x 780 = 1660 psi (TOP)

wn
il

1.0 x 2.1 x 4370 = 9180 psi (BOTTOM)

For Load Reduction Transient plus operating pressure:

Top of T.S. (El. #390), F’=( -6380 VA -6360

)

-2 1.0, K =

Bottom of T.S., (EL. #343), ¢ =( 6710 )/ 6747 )=1.0, K = 1.0

S 1.0 x 2.1 x 6420

13,480 psi (TOP)

1.0 x 2.1 x 6750

S

i

14,180 psi (BOTTOM)

Syange = 9-18 + 14.18 = 23.36 ksi

Paca (Ol

1

.0
(D) PEAK STRESS INTENSITY
(AT CENTER OF T.S.)

FOR TRANSIENTS DUE TO REACTOR SCRAM AND LOAD CHANGE PLUS OPERATING

PRESSURE .
S =Ypax B 04
WHERE: Yp.y =
Py =
0, -

FOR TOP OF T.S.,
(EL. #390)

FOR BOTTOM OF T.S.,
~(EL. #343)

S

S

1.43 x 2.1 x 11

NOTE:
FOLLOWING RESULT:
TOP OF T.S., N = =520 =
T 780

BOTTOM OF T.S., 3= L220

| L370
S =1.68 x 2.1 x 780 + 3600
S =1.43 x 2.1 x 4370 + 230

STRESS MULTIPLIER

PRESSURE ON THE SURFACE WHERE THE
STRESS IS BEING COMPUTED

PRINCIPAL STRESS HAVING THE LARGER

ABSOLUTE VALUE IN THE PLANE OF THE
EQUIVALENT SOLID PLATE

o =

= [j220 + 6710 = 10,930 =0.98, Y .
L370 + 6747 11,120

-520 - 6380 = 6900 = 0.97, Y
2780 - 6360 TALO max

1.43 x 2.1 x 7140 + 3600 = 25,040 PSI (TOP)

,120 + 230 = 33,620 PSI (BOTTOM)

FPOR REACTOR SCRAM TRANSIENT ALONE PLUS‘OPERATING PRESSURE WE HAVE

0.67, Ypax. = 1.68

= 0-979 Yma,x, = 1-’4—3

6350 PSI (TOP)

13,350 PSI (BOTTOM)

Page C-1-5

|

1.43

1.43
TUBESHEET STRESS AT OUTER RADIUS OF PERFORATED ZONE (SECTION 2-2)
(FOR DESIGN PRESSURE, p=4000 PSI ONLY)

e &
o - !
Element Se O~ - (¢oo PSt. ~3700 PS/ :
400 -858 -2661 r“—* | ]
75 - 91 ~1119 .
N, \,
71 145 - 217 \ -
\ | \\\. -
‘ ™~
443 1211 1476 N, | ~,
/
67 991 1368 £

63 1349 1799

59 1295 2128 R
55 1091 2319 !
51 946 2559

l
\
\\
\
\
| - A
,1.1 72 pP5). | 2750 PS/.'
353 861 2823 ' . F

PLOT QF STRESS THROUGH TUBESHEET
(Pp + PB)

L.

g_(TOP) = ~1400 psi
0, (BOTTOM) = 720 psi
0, (AVG.) = 694 psi

Py = %(1400 + 720) = * 1060 psi

O, (TOP) = -3700 psi

¢ (BOTTOM) = 2950 psi

Or(AVG.) = 1048 psi

P, = %(3700 + 2950) = +3325 psi

Maximum Average Plate Temperature < 10750F

Page C-1-6
(A) PRIMARY GENERAL MEMBRANE STRESS INTENSITY (EDGE OF T.S.)

Q
]

694 + 2096 = 2790 psi

2.1/1.7017 x 107 + 778.41 x 10"

= 10,460 psi

= 1.05 [ 4980 + 2790 + 3808 | = 12,160 psi

14p]
Mo
}

(B) PRIMARY MEMBRANE PLUS BENDING STRESS INTENSITY (EDGE OF T.S.)

~Top of T.S., (>= 6./67 = -1400/-3700 = 0.378, K = 1.04

Bottom of T.S., {3= 720/2950 = 0.244, K = 1.06

I
]

S =1.04 x 2.1 x 3700 = 8080 psi

S

1.06 x 2.1 x 2950

6570 psi

(C) PRIMARY PLUS SECONDARY STRESS INTENSITY

Top(E1l. 401) Bott(El. 354) Top(El. 401)7

Bott(El. 354)

Op.P. + R. Scram 6105 =581 . 3984
Op.P. + Load Change - 275 268 -2398
Steady-State 506 -2272 | . 710

1524

Page C-1-7
LOADING CONDITION (a) (P + R.S.)-(P+L.C.)

TOP BOTT TOP
6380 -849 6382

TOP OF T.S.:
(= 6380/6382 7 1.0, K=1.0

S =1.0 x 2.1 x 6382 = 13,400 psi (TOP)

BOTTOM OF T.S.:

£ = -8L49/-1068 = 0.79 K = 1.01
S=1.01 x 2.1 x 1068 = 2270 psi (BOTT.)

LOADING CONDITION (b) P + R.S. + 55

T0P BOTT 0P
6611 -2853 Leol
TOP OF T.S.: |

O = L69L /6611 = 0,704, K = 1.02
§ =1.02 x 2.1 x 6671 = 14,220 psi (TOP)

BOTTOM OF T.S.:

(3 = -1368/-2853 = 0.40 K = 1.03

s =1.03 x 2.1 x 2853 = 6170 psi (BOTT)

LOADING CONDITION (c) P + R.S.

TOP OF T.S.:
= 3984/6105

0.65, K = 1.015
S = 1.015 x 2.1 x 6105 = 13,010 psi (TOP)
BOTTOM OF T.S.:

g

S = 1.45 x 2.1 x 1524 = 4640 psi (BOTT)"

~-581/1524 -0.38, K = 1.45

BOTT

-1068

BOTT

-1366

Page C-1-8
PEAK STRESS INTENSITY
(at edge of tube sheet)

For transients due to reactor scram and load change plus
operating pressure.

For top of T.S., § = 1.0: ¥Ymax = 1.43

For bottom of T.S., @ = 0.79: ¥max = 1.50

P
Yma.x(h}fi'. + Ps

g =
S=1.43 x 2.1 x 6380 + 3600_= 22760 (Top)
S =1,50 x 2.1 x 6382 + 230 = .20333 (.Bottom)

Page C-1-9
SECTION 3-3  PRIMARY MEMBRANE
Element 9, O
36L 71 -1578
365 -8L46 -3897
366 ~1269 -~3658
367 1626 5858
368 ~2661 =7190
Ave. _1267 ~L1i36
111 = -28
O =-5676 S = 6150 = Py

PRIMARY PLUS SECONDARY STRESSES

ELEMENT 36l 3 -

1) OP. P. + R.S. 86L5

2) OP. P. + L.S. -1546

3) s.s. - 1710

comvrrioy (o) (0 - (&

10191

J, = 16052 0, = 116

S = 26360

compITIoN () (1) + (3)

20355
J, = 31506 0’2 = -286
S = 49,810

o)
Maximum Temperature < 1150 F

(Assume Also Pp)

10865
0’3 = ~18300

-6760
3543
-115L0

-10303

-18300

~15171

Page C-1-10
SECTION L-L PRIMARY STRESSES

Element S P Ty Crz
38) -L,06 2027 ~L0L 5013
385 -1098 L791 198 LL6T7
386 -1847 L}50 -205 2831
387 -1185 3698 8l 2929
388 -1107 . . 1144 -795 792
389 +106 -508 -68L 534
Avg. -922 2600 -301 2761
-
G, o= Ly Oé = =21436 3= -301
S = 6550
SECTION L-L  PRIMARY PLUS SECONDARY STRESSES
| g e - Oes
1 OP. P + R.S. -678 125], 11140 526l
2 O0P. P + L.S. ~609 L4377 1948 3974
3 s.s. 1363 1302 -1071 1111
Condition :f:'é,;} 1\ - 21
-69 ~3123 ~13088 1290
T, =-3595 (05 = L03 G4 = 13088
s = 13490
Condition (b) (1. + .3
, | 685 2556 -12211 61,05
0, = 809l 0, = L1852 05 = 12211
S = 20,310

Maximum Temperature < 1000°F

Page C-1-11
PRIMARY LOCAL MEMBRANE STRESS SECTION 5-5

El. 6 R 62 (jT A O’Rz.
420 | 5492 24660 7066 10250
421 3636 14480 3449 5126
422 2891 9002 | 1611 3000
423 | 1821 | 6679 | 689 1186
424 2126 4590 102 2404
425 1397 3811 - 240 | 460
426 1116 1778 - 931 850
427 379 - 684 | -1821 253
428 81 - 2693 ~2446 - 264
AVG. +2104 46847 + 831 + 2585
V1. T2 = %(2104 + 6847) # [((6847-2104)2)% + 2585%]%
= 4476 * 3508
= 7984, 968
T3="Tp =831

Stress intensity, S = 7984 - 831 = 7150 psi

o}
Maximum Temperature < 1000_F

Page C-1-12
SECTION 5-5  PRIMARY PLUS SECONDARY STRESSES
<% O +
1 OP. P. + R.S. 41110 21020
2 OP. P. + L.S. 9040 L6LO
3 §.S. 5603 3397
Condition aj, 1,2
32070 16380
14 = 11,054 ‘52 - 6396
S = 34,658
Condition (b 1 + '3
46713 2WaT
T, = 65985 (o = 5115
S = 60,8L0

i

28301
21691,

Page C-1-13
PRIMARY GENERAL MEMBRANE STRESS SECTION 6-6

El. & 8. | s}

YR z Jr
108 ~2550 4667 3805
109 -2236 5248 3990
110 -1773 5621 4166
111 ~1271 5840 4320
112 - 748 5974 4469
113 - 214 6070 4623
114 317 6165 4788
115 833 5360 4971
AVG. - 955 5736 4392
fl’fiz = 15(5736 - 955) * [ (5736 + 955/2)% + (-2599)2]%
= 2391 * 4236
= 6627, -1845
Gq =0p = 4392

Stress Intensity, S = 6627 + 1845 = 8470 psi (Section A-A)
Maximum Temperature < 1000°F

Stresses due to R.S. and L.S. alone are given below:

ir Oz Or
R.S. 3670 11130 23330
L.S. -1370 - 4150 - 8800

Page

CTRz

3214
-2946
~2696
-2485
~2344
-2291
-2336
~2477

-2599

C-1-1L
SECTION 6-6  PRIMARY PLUS SECONDARY STRESSES
(Tr ,:q“-
1 OP. P. + R.S ~1311 21,810
2 OP. P. + L.S ~3568 1139
3 S.s. 98 1968
Condition (a} (1, - 2)
2257 23671
0, = 25293 6, = 635
SI = 24658
Condition ié; 1+ kij
| -1213 26778
//
Ty = 28Ll G, = -2878
ST = 31322

18858

I

-6112

21,362

-7027
18858

Page C-~1-15
SECTION 7-7

PRTMARY STRESSES

Element Jr J di Qr?
180 -314) ~326), 13090 ~333
181 -2059 ~2809 12320 =475
182 -1129 -2265 11710 -687
183 ~369 -1656 11200 -917
181, 221 ~991 10760 ~1125
185 639 -281 10360 ~1279
186 887 168 9962 -1353
187 961, 1267 9571 -1328
Avg. -L499 -1191 11122 ~937
@} = 154
&, = 18l S = 12970
<33 = 11,122 | o
Maximum Operating Temperature < 1000 F

SECTION 7-7  PRIMARY PLUS SECONDARY STRESSES |

Or 5 o  Ora
1 0.P. + R.S 8L3 8181 35140 -5325
2 0.P. + L.S -L,228 -7165 2680 1576
3 S.S. -19 -50 -189 2l
Condition {a} 1 - 2

5071 153L6 32460 ~6901

G, = 18,812 0, = 1606 Cfé = 32160
S = 30,850

Condition (b)

does not control

Condition 9) S = 37,090 (max.)

Page C-1-16
SECTION 8-8

Primary stresses (due to design pressure of 1000 psi)

Element

332
333
33k
335
336
337
338

Average

5

==

-1908
LLo1

10623

Or Oz
-3629 L515
-29L1 L501
-2331 Lh9s
-1788 LL90
-1305 LL85
-87L LL8-
-Lg2 LL76
-1908 LL%1

S = 12531 psi

Negligible

Page C-1-~17
HASTELLOY

APPENDIX C-2

N MATERTAL PROPERTIES

A) ALLOWABLE STRESSES (REFERENCE ) )
S¢ Spt s
(Temp.  Ohr 10 hr 10° hr 10° hr 10* hr 10f hr 10 hr_ 10° hr 10® hr 10% hr 105 he O z
70 26,600 I ’ , , 29,000 26,600
Lo | 26,600 21,000 | 26,600
600 26,600 , 20,000 26,600
800 | 25,200 | ! | ] t 18,000 | 25,200
1000 | 23,850 ; : ' : ! ! 17,000 23,850
1100 | 22,950 | 18,000 34,000 2k,000 17,160 12,000] 22,950 - - 17,160 12,000 13,000 | 22,950
1200 22,050 | 34,670 24,000 16,000 12,000 7,260 | 22,050 - 16,000 12,000 7,260 6,000 22,050
1300 19,300 | 24,000 16,000 10,000 6,250 3,500 19,300 16,000 10,000 6,250 3,500} 3,500 19,300
1400 [ 16,600 | 12,000 7,670 5,070 3,330 2,000 12,000 7,670 5,070 3,330 2,000 16,600
B) OTHER PROPERTIES (REFERENCE 7 )
Modulus Mean Coefficient
- . . Thermal
Temperatiure of of Expansion e
° T - A . o Conductivity
(°F) Klasticity in./in.-°F (Btu/ft-nhr-°F)
(psi) (70° - T) |
X 1076 X 108
100 31.3 6.0
200 30.6
3_00 30.0 -
LOo £9.5 6.45 7.4
500 29.0
600 28.5 6.76 8.3
T00 28.1
800 2‘7'7 T|09 9.,?_
Q00 27.2 9.3
1000 26.7 7.43 10.4
1050 26. 4
1100 26.3 11.1
1150 26.1 1201
1200 25.7 7.81 1.7
Deusity, 0.3L17 1b/in.® at room temperature.
Spcecific heat, 0.095 Btu/lb °F at room temperature; PAGE C_2-1

O 130 Riba1/1W 9T b 1 A0 TS
APPENDIX C-3
GULF GENERAL ATOMIC

REPORT GULF-GA-A 1241,
LN

- GULF GENERAL ATOMIC

Gul f-GA-Al2414

FINAL REPORT

TUBE RUPTURE ANALYSIS OF A
COUNTERFLOW HEAT EXCHANGER

by

J. J. Johnson and D. A. Wesley

Prepared under
P.O0. N24013
Project No. 0540.0000
for
Foster Wheeler Corporation

under .
Union Carbide Corporation, Nuclear Division
Subcontract No. 91X-88070C

under
Prime Contract No. W-7405-eng-26
with the
U.S. Atomic Energy Commission

Gulf General Atomic Project 0540 November 20, 1972

GULF GENERAL ATOMIC COMPANY
P.O. BOX 81608, SAN DIEGO, CALIFORNIA 92138
ABSTRACT

The structural integrity of a counterflow heat exchahger subjected to
a steam tube failure was examined. The scope of the investigation was
limited to two areas of concern. The effect of the burst on the contain-
ment vessel was treatec as a plane strain problem, and the results indicated
that the shell could withstand the accident without failure. Stresses well:
above the yield strength of the shell result, however, and an increase in the
diameter of the shell of the order of 4,0 inches may be anticipated for the

worst case of failure of a tube immediately adjacent to the shell,

The integrity of a tube adjacent to the ruptured tube was considered
using a discrete method of analysis modeling the tube as a continuous beam.
The results obtained for the response of the adjacent tube indicate that
although very large beam deformations are predicted, rupture will not occur,
even for the very conservative assumption of no support from surrounding
tubes. While this is physically an unrealistic case, it may be Considered

an upper bound on maximum tube response.
ABSTRACT . ,

CONTENTS

LIST OF FIGURES . . . . . . . . . .. .« o e

INTRODUCTION .

ANALYSIS

2.1 Effect
2,1.1
2.1.2

2,2 Effect

RESULTS AND
3.1 Effect
3.2 Effect

CONCLUS | ONS

REFERENCES .

of a Tube Rupture on the Shell .
General Remarks . . . . . . . .
Models of the Problem . ., . . .
2.1.2,1 Model I , . . .., ...
2.1.2.2 Model II . . .....

of a Tube Rupture on an Adjécent

DISCUSSION L] [ ] - - L] L ] a L ] L ] L L ]
of a Tube Rupture on the Shell .

of a Tube Rupture on an Adjacent

. L] L] * L ® - Kl L] - L] - L ] L J L ] -
LIST OF FIGURES

|dealized stress-strain curve for Hastelloy N at 850°F
Grid plot for Model I of the shell analysis
Enlarged grid plot of tube rupture area of Model I

Pressure-time history of the steam after tube rupture

Grid plot for Model Il of the shell analysis
Pressure profiles applied to the shell

Points of application of the pressure profile-Model II
of the shell analysis

‘ldealized moment-curvature relationship of a tube

Effective stress~time history in Zone |
Effective stress~time history in Zone 2
Effective stress-time history in Zone 3

Effective stress-time history in Zone 4

Grid plot of the shell, time = 0 sec
Grid plot of the shell, time = 50 psec
Grid plot of the shell, time = 254 psec
Grid plot of the shell, time = 551 usec
Grid plot of the shell, time = 754 usec
Grid plot of the shell, time = 1 millisec

1.25 millisec

Grid plot of the shell, time

1.5 millisec

Grid plot of ‘the shell, time
Circunferential stress-time history in Zone |

Model and loading of a tube adjacent to the rupture

Maximum moment-time history in the tube

iv
1. INTRODUCTION

A history and general description of the molten-salt breeder reactor are
contained in Refs. | and 2 along with numerous references to more detailed
information on the individual components and the development of new materials
for their construction., The purpose of this study was to evaluate the struc-
tural integrity of one such component under a specified accident condition.
This involved investigation of the effects of a tube rupture on a counter-
flow heat exchanger. The vessel is a cylindrical shell 114,83 feet in
height, 41 inches in outside diameter, and .75 inch wall thickness. It
contains 1032 tubes of .75 inch outside diameter and .125 inch wall thick-
ness which are supported by tube sheets at 5 foot intervals (Ref, 6). The
shell-side fluid is molten-salt and the tube-side fluid is supercritical
steam. The inlet and outlet conditions of the steam and-molten-salt are
specified in Ref. 3. Both the shell and tubes are constructed of a nickel-
base alloy, Hastelloy N, with temperature dependent material properties

specified in Refs. 3-6.

The rigorous treatment of a steam tube burst on the sfructural integrity
of its containment vessel and the adjacent tubes is a complex hydrodynamic
problem, A mechanized literature survey obtained from the Nuclear Safety
Information Center at Oak Ridge, Tennessee, and one performed by the authors
verified the lack of completely general solution techniques available for
the indicated problem. Since inadequate time was available to develop
methods of analysis, it was concluded that the problem must be simplified
considerably and its solution sought by an appropriate available technique.
The basic philosophy of all such idealizations is to model the phenomena as

accurately as possible, using a conservative representation of the problem
where necessary. This allows one to qualitatively discuss the results with

a reasonable degree of confidence.
The scope of this investigation was limited to two areas of concern.
These were the effects of a tube rupture on the structural integrity of the
shell and the integrity of an adjacent tube. The shell was treated as a
plane strain problem using the PISCES 2DL computer éode. This is a finite-
difference program which is capable of treating hydrodynamic/structural
shock problems. The response of a tube adjacent to the rupture was
obtained from a lumped-mass model of a continuous beam. Due to reduction
of effort allowed, the treatment of effects of adjoining tubes was not
included, either in the overall beam response of the tube or the localized
response due to impact, nor was any consideration of scattering of the wave
nossible., The analysis and models are discussed in detail in subsequent

sections,
2, ANALYSIS

2.1 EFFECT OF A TUBE RUPTURE ON THE SHELL

2.1.1 General Remarks

The rupture of a tube adjacent to the wall of the shell was studied by
considering a typical transverse section through the shell and analyzing it
as a plane strain problem. Average values of temperature, 850°F, pressure
in the steam, 3700 psi, and pressure in the molten-salt, 206 psi, were
assumed, This assumption removes the longitudinal flow of the molten-salt
from consideration, which is reasonable in the time span being considered.
The problem was subsequently analyzed using a hydrodynamic computer code
named PISCES 2DL. | |

A series of computer programs identified by the name PISCES have been
develdped by the Physics International Company (Ref. 7) and are marketed
through the Control Data Corporation. The program PISCES 2DL is a two-
dimensional hydrodynamic code which utilizes the Lagrangian formulation
(Refs, 8 and 9) with a finite-difference approach. The programs contain
specific provisions for treating shocks and they allow fluid-solid inter-
action, a wide variety of equations of state, and non-linear stress-strain
models. The standard elastic-plastic model in PISCES is bi-linear and uses
the von Mises yield criterion with a Prandtl-Reuss flow rule. This model was
used in the present study for the idealized stress-strain curve of Hastelloy

N at 850°F shown in Fig. 1.

The actual computer runs were made at Lawrence Berkeley Laboratory at a
considerable reduction in cost over the Control Data Corporation's version,
The computer models were established after consultation with the staff of
Physics International Company since they were most familiar with the capa-

bilities and restrictions of the code.

3
90,000 , , | .
//////,
60,000 |- _
o
&
o = 27,625 PS!
: = 85,500 PSI
(V)
30,000 |- | | .
6 .
E = 27.45 X 10° PS|
;
0 ! | L J
0 0.1 0.2 0.3 0.4

STRAIN

Fig. 1. Idealized stress-strain curve for Hastelloy N at 850°F
2.1.2 Models of the Problem

2.1.2.1 Model I. The first'attempt at modeling the shell included the |
molten-salt, four steam zones (modeling the steam as an ideal gas), and four
artificially dense zones (representing the tube) to direct the flow of steam,
An overall view of the grid is shown in Fig. 2 and an enlargement of the
‘area containing the steam and artifically dense zones is shown in Fig. 3.
Several runs were made with this model for varying boundary conditions in
the steam. An accurate representation of the problem was not possible

using this configuration and an increase in the number of zones near the
rupture was indicated. Although this refinement would not significantly
increase the total number of zones, the corresponding reduction in the time.
step necessary to assure convergence made it economically infeasible. Also,
an initial increase in resolution of the mesh at the rupture did not guar-

antee the solution and a further reduction may have been necessary.

2.1.2.2 Model II. The alternative to this situation was to model the shell
excluding the molten-salt and the directed flow of steam from the tube.
The disturbance then had to be applied to the shell as a pressure distribu-

tion with a specified time and spatial variation,

Quanti tative results on tube ruptures are infrequent in the literature,
A comprehensive study was recently completed by Eiber EE_El: of Battelle |
Columbus Laboratories, Columbus, Ohio‘(Ref. 10) in which a series of exper-
iments were performed on cylindrical presSure vessels with through-wall
and surface flaws to investigate the many facets of a rupture. The failure
reported by Eiber propagated along the length of the vessel in contrast to
the mode of failure specified by Foster Wheeler Corporation (Ref. 12), i.e.,
a '"fish mouth! type opening in the tube. This difference may be attributed
Vto several factors such as the initiation of failure in the tube and the
fact that the tube is a thick-walled cylinder compared to the thin-walled

cylinders tested by Eiber,
Fig. 2.

Grid plot of Model I for the shell énalysis
|a»\l.n. \— r - . . u. i‘t\ Illrrc\-l(\r.fi \‘l\*“lfiulk\“flule
» ) on\'.l ‘. hodld -
Wt v J RSy o
e . M\»\\ll& fl. —~ wolPX
T, . - Pt IJIL.\ W..l\l\ ..»'x G
Y ) ' -~ X o J " VA\.-.. o,
.w.\- \).\I\-\J". fd\ll - ’A- -\I\\\ / \\\ * )

>

S - MOLTEN-SALT ZONE
H - HASTELLOY N ZONE

X - ARTIFICIALLY DENSE ZONE

e - STEAM ZONE

Enlarged grid plot of tube rupture area of Model I

Fig. 3.
The pressuré-time history shown in Fig. 4 was selected for the present
study. The initial tube pressure was allowed to act for 1.5 milliseconds
to reflect the aforementioned discrepancy in the mode of failure. The final
pressure was estimated from the results of a shock tube analysis assuming
interaction of two fluids with the pressure and other properties of the
molten-salt and steam, respectively (Ref. 16). The basic pulse shape and
duration are consistent with the measurements made by Eiber as well as
those utilized by Gulf General Atomic (Ref. 11) for the analysis of a
double-ended steam tube failure. |In both cases, the pressure profile of

Fig. 4 is an upper bound.

The source of the disturbance is assumed to act at a point 19.175
inches from the center of the shell, Fig. 5, which corresponds to the
inside edge of the tube nearest the wall (Ref. 6). The disturbance is
naturally modified as it travels through the molten-salt and interacts with
the shell. Two phenomena define the character of the pulse experienced by
the shell: divergence of the wave as it travels through the salt, and
reflection of the disturbance at the interface of the shell and molten-salt,
In close proximity to the source, the impedence difference of the molten-
salt and shell governs the definition of the disturbance. As the distance
from the source increases, spherical divergence plays an increasingly

important role.

It is not difficult to verify that the shell is a relatively fixed
surface compared to the molten-salt; i.e., a reflected wave of amplitude
approaching the amplitude of the incident wave would be generated upoh
normal incidence (Ref. 13). The maximum amplitude of the reflected dis-
turbance is thus equal to the amplitude of the incident disturbance and
occurs at a fixed boundary. This value was conservatively assumed in the
present case. This magnification of the pulse due to reflection decreases

rapidly as one proceeds along the shell due to the rapidly decreasing angle

of incidence.
PRESSURE (PS1)

3700

4oo

Fig. 4.

| 1 | 1 | i
] 2 3 4 5
TIME (mSEC)

Pressure-time history of the steam after tube rupture

SN

% PLANE OF SYMMETRY

Fig. 5.

Grid plot for Model II of the shell analysis
The effect of spherical divergence is two-fold: the disturbance is
Yspread out' in time, i.e., an increasing rise time with increasing distance
from the source, and a decrease in the magnitude of the disturbance occurs
(Ref. 13).

Several one-dimensional runs were made with PISCES 1DL (Refs. 14 and
15) in both spherical and plane symmetry to verify and quantify the preced-
ing observations. The results obtained led to the pressure profiles shown
in Fig. 6 with the corresponding zones of application depicted in Fig. 7.
The duration of the maximum pressure corresponds to the travel time of a
signal through the thickness of the shell. At points more remote from the
source, the rise time of the disturbance was assumed to increase as the
Square of its distance from the source and subsequently follow the pressure
profile of Fig. 5. ' The loadings were time lagged by the travel time through

the molten-salt.
The results of this analysis are discussed in Section 3.1,
2.2 EFFECT OF A TUBE RUPTURE ON AN ADJACENT TUBE

In addition to the treatment of the shell, the effect of a tube rupture
on an adjacent tube was considered. Those properties deemed important to
the phenomena were studied and modeled appropriately. In the present case,
the elastic-plastic properties of the tube together with the pfesence of
the molten-salt are necessary tQ_accurately represent the physical problem,

Thus, as in the previous case, a discrete method of analysis was required,

The tube was modeled as a beam, continuous over four supports, and
divided into a discrete number of elements with concentrated mass and stiff-
ness properties. The solution technique used is described by Biggs (Ref. 17).
In general, at each time step, an element is acted on by the applied load,
shearing forces (derived from the moments at the ends of the segment), an
inertial force, and a drag force which approximately represents the resistive
force of the molten-salt. The acceleration of the mass is then determined
from the equation of motion and integrated to obtain the velocity and dis-

placement of the mass. The moments are determined from a moment-curvature

11
7400 |

6600 |
= -
o 4
w3700 w 3700 |
= o
Y D
N w
Ll )
o il
a- a
ZONES 1 AND 2 ZONE 3
56 56
TIME (uSEC) TIME (uSEC)
9920 ¢ —_
z g 5180 {
w 3700 | < 3700 |
= 7S
w 7]
w Lt
";:-' o
a o
ZONE 4 ZONE 5
56 - S 56
TIME (uSEC) TIME (uSEC)
£ . e
N Lot w 3700
= 37001 s [
Iy w
» 0
ye o
o oo
ZONE 6 ZONE 7
56 56
TIME (uSEC) TIME (uSEC)

Fig. 6. Pressure profiles applied to the shell. Zone numbers refer to
Fig. 7. The pressure profiles follow Fig. 4 for times exceeding
56 sec. |

12
£T

7.

—t

. .%;_. -

Zones of application of the pressure profiles - Model II of the shell analysis
relation where the curvature is defined as the second central difference of
the displacements. An elastic-plastic moment-curvature diagram was derived
for the tube from the stress-strain diagram of Fig. 1 assuming a linear
distribution of strain through the cross section of the tube. The result

is shown in Fig. 8. The drag force was defined as in Ref. 18 with a minimum

value of the drag coefficient assumed (CD = 1.0). Several loading cases

were considered,

14
MOMENT (IN./LB)

6000

T 1 T I
(.933,5170)
5000 |- —
4000 -
3000 -
2000 |— -
(.136,1900)
1000 | ' -
(.00264,910) ]
|
0 | | l 1 ! L ] |
0 0.5 1.
CURVATURE (IN.”))
Fig. 8. Idealized moment-curvature rélationship of a tube

15
3. RESULTS AND DISCUSSION

3.1 EFFECT OF A TUBE RUPTURE ON THE SHELL

The pressure profile discussed in the previohs section was applied to
the shell and the solution obtained for 1.5 milliseconds. Effective stress-
time histories are plotted in Figs. 9-12 for the first four zones along theA
circumference of the sheli. The maximum stress (42,900 psi) occurs at 1.5
milliseconds and is one-half the ultimate strength of Hastelloy N at 850°F,
The computer solution was terminated at this point for several reasons.

The initial disturbance has made over two complete circumferential transits
of the shell and the remaining pressurization of the shell is slow in compar-
ison. At 1,5 milliseconds, the pressure profile 6f Fig. 5 begins to decrease,
approaching a quasi-static pressurization of the shell at a considerably

smaller pressure,

Grid plots of the shell at various times in the solution are depicted
in Figs. 13-20. Figure 21 is a computer plotted time history of the circum-
ferential stress in the first zone of the shell. An increase in the diameter

of the shell of 3.6 inches has occurred at 1.5 milliseconds.

3.2 EFFECT OF A TUBE RUPTURE ON AN ADJACENT TUBE

Several loading cases were considered for both a single span, simply
supported beam and a three span continuous beam, The model and loading
case reported here are shown in Fig. 22, A plot f the maximum mdment~time
history for the beam is shown in Fig. 23, The simple tube model considered
would not rupture. However, the results of the analysis exceed the appli-
cability of beam theory because very large deformations occur; e.g., a max-
imum mid-span deflection of 24,5 inches was calculated. Furthermore, the

geometry of the heat exchanger also limits the applicability of the model,

16
EFFECTIVE STRESS (PS1)

50,000 T I T T T T

40,000 |-
30,000 -
20,000 - : , —
10,000 —
0 1 . | 1 } ] ] 1
0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.
TIME (mSEC)

Fig. 9. Effective stress—time history in Zone 1

17
EFFECTIVE STRESS (PSI)

50,000

40,000

30,000

20,000

10,000

18

e
-
- n
1 i e I\ 1 i
0 0.2 0.4 0.6 0.8 1.0 1.2 1. i
TIME (mSEC)
Fig. 10. Effective stress-time history in Zone 2

.6
EFFECTIVE STRESS (PS1)

50,000

40,000

30,000

20,0001

10, 000

Fig. 11.

Effectlve stress—time history in Zone 3

19

—
p—
| ] | N I\ | 1
0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.
TIME (mSEC)

EFFECTIVE STRESS (PSI1)

50,000

40,000

30,000

20,000

10,000

—
-
b
\ L L 1 \ A
0 0.2 0.4 0.6 0.8 1.0 1.2
TIME {(mSEC)
Fig. 12. Effective stress-time history in Zone 4

20

T¢

0y
Ty -

H:

Fig. 13.

Grid plot of the shell, time

0 seconds
[

~ Fi
g
* 14
. Gr
id
pld
t
of
| the
sh
el
1
b ]
ti
me
= 50
pse
c

>
N
‘
N
W
W
Fig. 15. Grid plot of the shell, time = 254 usec
N
£~

Fig. 16. Grid plot of the shell, time = 551 usec
Fig. 17. Grid plot of the shell, time = 754 usec
. Fig. 18. Grid plot of the shell, time = 1 millisec-
Fig. 19. Grid plot of the shell, time = 1.25 millisec
Fig. 20. Grid plot of the shell, time = 1.5 millisec

o
STRESS (PSI)

55,000

t gt p oty r 2t
\
T T

k1

o
o

I Y TS T U0 D Y I AN N WO U N 2 P S Y BN N IS Y B

Fig. 21.

el
-
-
-
r
.
-
-
.
e

| S S N N N |
3

-
i
—t

I
I
T
fi—rfi—kr‘f—r_'_T_r‘T_*‘-'r—r_lerllrrrJl"T‘T"l'
200 600 . 1.0 E+3 1.4 E+3
TIME (uSEC)

Circumferential stress-time history in Zone 1

29

) —
1 TT

i e #

11,900 |

FORCE (LB)

TIME (mSEC)

(a) Force applied to Node 22

11,900}

FORCE (LB)

4
|

0 e+ 5
TIME {(mSEC)

(b) Force applied to Nodes 21 and 23.
Time lagged with a rise time of
0.2 milliseconds.

21 B
—(22
| T
AN 2 Z :
Lw 42 ELEMENTS AT 4.286 IN. = 180 IN. _}J

(c}) Model of tube

Fig. 22. Model and loading of a tube adjacent to the rupture

30
T¢

MOMENT (IN./LB)

6000

[ 1 I r o
Looo | —
_.—-——-""—'-—__”_N“”\i

e T ‘\\

2000 — —
0 | 1 _— SN I —t
0 1 2 3 4

TIME (mSEC)
Fig. 23. Maximum moment-time history in the tube

since after a very small amount of deformation, the tube would hit another
tube. This would increase its resistance to deformation (at least untjl
the combination of the pressure disturbance and impact loading of the
adjacent tubes results in essentially equal velocities). Also, localized
stress concentration at the points of impact would occur. Inadequate time

was available to consider these facets of the problem.

32
L., CONCLUS!IONS

The analysis of a counterflow heat exchanger under a specified accident
condition, i.e., a steam tube rupture, was performed in this investigation.
The results of the analysis indicate that the containment vessel will not
fail due to a tube failure. The analysis was a conservative one in several
regards. The applied load was chosen as an upper bound of the data available
on ruptures and, due to the idealization of the shell in two dimensions, was
assumed to act along the entire axis of the shell. This latter assumption
neglects the increase in strength due to axial stresses in the region of

the rupture.

The results of the analysis of a tube adjacent to the rupture are less
conclusive. A more sophisticated analysis is required including the inter~
action of several tubes and perhaps large deformation considerations. How-
ever, several observations can be made from the results of the present
analysis. The most significant one concerns the amount of deformation
sustained by the tube without failure. This is due to the stress-strain
relation of Hastelloy N and is an essential factor in this and ahy subseguent
analysis. Similarly, the effect of the molten-salt must be included in the

analysis,

Scattering of the disturbance, while not investigated, is believed to
be a beneficial effect reducing the intensity of the disturbance. Inadequate

time did not permit a detailed consideration of the phenomenon,

This investigation identified the need for greater research and under-
standing of the mechanisms of failure and related phenomena, e.g., a measure
of rupture area and pressure as a function of time. Only after such research
is completed will more sophisticated and accurate methods of analysis be

available,

33
REFERENCES

Nuclear Applications & Technology 8, 2, 105-219 (February 1970).

Robertson, R. C., ed., Conceptual Design Study of a Single-Fluid
Molten-Salt Breeder Reactor, ORNL-4541 (June 1971).

Cox, J. F., Foster Wheeler Corporation, "Basic Information - Reference

Design Molten-Salt Steam Generator," unpublished data.

Cox, J. F., Foster Wheeler Corporation, '"Properties of Materials Used
in the Reference Design Molten-Salt Steam Generator," unpublished
data.

Cox, J. F., Foster Wheeler Corporation, '"Bulk Modulus of Molten-Salt

and High Temperature Properties of Hastelloy N,' unpublished data.

Cox, J. F., Foster Wheeler Corporation, 'Final Drawings Steam Generator,'’

unpublished data.

Physics International Company, An Introduction to the PISCES System of

Continuum Mechanics Codes (October 1971).

Physics International Company, PISCES 2DL-General Description and

Finite-Difference Equations (1972).

Physics International Company, PISCES 2DL - Input Manual (1972),

Eiber, R. J., et al., "Investigation of the Initiation and Extent of
Ductile Pipe Rupture,' Battelle Columbus Laboratories Report BMI-1908
(June 1971).

Jones, D. J., (Gulf General Atomic), private communication,

Cox, J. F., Foster Wheeler Corporation, "Specification of Rupture Area

in a Tube," unpublished data.

34
REFERENCES {Continued)

Ewing, W. M., W. S, Jardetzky and F. Press, Elastic Waves in Layered

Media, McGraw-Hill Book Company, New York, 1957.

Physics International Company, PISCES 1DL - General Description and

Finite-Difference Equations (1971),

Physics International Company, PISCES 1DL - Input Manual (1971).

Liepmann, H. W. and A. Roshko, Elements of Gasdynamics, John Wiley and

Sons, New York, 1962.

Biggs, J. M., Introduction to Structural Dynamics, McGraw-Hill Book Co.,
New York, 1964, pp. 192-195,

Daily, J. W., and D. R, F. Harleman, Fluid Dynamics, Addison-Wes ley
Publishing Co., Reading, Mass., 1966, pp. 376-385.

35