NUREG/CR-5314

EGG-2562

Vol. 3

LieAssmntPoeue

Life Assessment Procedures,:

for Major LWR Components

Cast Stainless Steel Components

Prepared by C. E. Jaske, V. N. Shah

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Prepared for

U.S. Nuclear Regulatory Commission
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NUREG/CR-5314

EGG-2562

Vol. 3

RM, R9

Life Assessment Procedures

for Major LWR Components

Cast Stainless Steel Components

Manuscript Completed: September 1990

Date Published: October 1990

Prepared by

C. E. Jaske, V. N. Shah

G. H. Weidenhamer, NRC Program Manager

Idaho National Engineering Laboratory

Managed by the U.S. Department of E1nergy

EG&G Idaho, Inc.

Idaho Fas D) 83415

Prepared for

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555.

NRC FIN A6389

Under DOE Contract No. DE-AC07-761D01570
:V i

ABSTRACT

7bis report presents a procedure for estimating the current condition and residual life of

-safety-related cast stainless steel components in light wa racos (LWRs). The procedure accounts for loss of fractur toughness caused by hrmal embrittlement and includes

the folloing: a review ofdesign and fabrication records, nserice inspection recods, and

operating history a fracture mechanics evaluation to determine the required tughness at

end-of-life using worstMoads and woast-law indications; current and future toughness estimates: and criteria regarding continued service, repair, or replacement of the component

being evaluated. The report discusses the available Charpy V-notch impact energy. fcture

toughness, tensile strength, fatigue sistance. and ftge-crac k growth data, and presents

two methods for assessing the degree of thermal embrittlement: metallurgical evaluation

and analytical modeling of inservice degradation.

FIN No. A6389--Components and systems Iv

iii
INEL FOREWORD

The United States was one of the first nations to use

nuclear power to commercially generate electricity

and, therefore, has some of the olde operating commrcial reactor As U.S. light water reactors (WRs)

have matured, problems associated with timo- or usedependent degradation (aging) mechanisms such as

stress corrosion, radiation embrittlement, fatigue, and

other edfcts have occurred and have raised questions

about the continued safety and viability of older nuclear plants Some of the recent aging-related problems include primary water sue corrosion cracking

of pressurized water reactor (PWR) pressurizer heater

sleeves and instrument nozzles and PWR steam genrator tube plugs, steam generator tube ruptures caused

by high-cycle fatigue and by a faied tube plug, significant wall-thinning of light water reactor metal containments caused by corrosion, fatigue failure of

boiling water reactor (BWR) recirculation pump internals (resulting tn potential damage to reactor pressure

vessel core and internals), catastrophic failure of a

"nonnuclear" portion of a PWR feedwater line caused

by erosion-corrosion, and through-wall thermalfatigue cracks in high-pressure safety injection lines

and a residual heat removal lie.

At the same time, with a continually increasing demand for electricity and limited new generating capacity under construction, the U.S. electric utilities are

motivated to keep their existing plants operating beyond the original design liMf at as high a capacity as

possible. The economics of plant life extension are

clearly favorable. Studies cosponsored by the U.S.

Department of Energy (DOE) and the Electric Power

Research Institute (EPRI) show that replacing any

single nuclear plant component can easily be justified,

if the life of the plant can be extended for a number of

years. Extending the life of a 1000-MW plant by

20 years is expected to realize a net present worth of

perhaps $1 billion.

Therdore, the potential problems of managing aging in older plants and the resolution of technical

safety issues in consideration of the development of

appropriate license renewal criteria have become a

major focus for the research sponsored by the U.S.

Nuclear Regulatory Commission (USNRC). An important part of the USNRC research effort is the

Nuclear Plant Aging R ch (NPAR) Pram that is

being conducted at several national laboratories, including the Idaho National Engineering Laboratory (INEL). One of the NPAR program tasks at the INEL

is to develop the appropriate technical criteria for the

USNRC to assess the residual life of the major PWR

and BWR components and structures. These

assessments will help the USNRC identify and resolve

safety issues associated with LWR aging degradation

and develop policies and guidelines for making operating plant license renewal decisions

Most of the effort for this life assessment task is

focused on integrating, evaluating, and updating the

technical information relevant to aging and license renewal from current or completed NRC and industry

research programs. A five-step approach is being pursuod to accomplish the life assessment task: (1) identify and prioritize major components, (2) identify degradation sites, mechanisms and stressors, and potential failure modes for each component, and then

evaluate the current inservice inspection (1SI) mthods, (3) assess advanced inspection, surveillance, and

monitoring methods, (4) develop life assessment models and procedures, and () support the development of

technical criteria for liese renewal.

A brief discussion of each of these steps follows:

1. Identification and prioritlzatlon of major

componuetr Virtually all major equipment

contained within a nuclear plant complex is

subject to some aging degradation and must

be evaluated in an aging and licensee renewal

program. From the USNRC's perspective of

ensuring the health and safety of the public,

the first step in this assessment task: was to

Identify those major components critical to

nuclear power plant safety. Components that

help contain the release of fission products

during normaL off-normal, or accident condhions were selected. The PWR components

(in rough order of importance) include the

reactor pressure vessel (RPV); the containment and basemat reactor primary coolant

piping, safe ends, and nozzles; steam generators; reactor coolant pump bodies; pressurizer

and associated surge and spray lines; control

rod drive mechanisms; cables and connectors; emergency diesel generators; RPV internals; RPV supports; and feedwater lines and

nozzles. A similar list was developed for

BWRs except that the containment was

ranked most important. Although some PWR

steam generator and BWR recirculation piping have been replaced, their replacement is a

major, time-consuming operation. In addition, steam generator tubes and recirculation

piping constitute a part of the primary pressure boundary, and therefore, aging

iV

I
evaluation of these components is included in

this ask. The lifethme of many of the smaller,

less expensive components such as the

pumps, valves, sensors, batteries, controls,

etc. is often less then the initial license period

of 40 years, and these components are usually

repaired, refurbished, or replaced relatively

frequently. Therefore, these components are

being studied in other NPAR tasks that are

focusing more on reliability, availability, and

maintainabiity than on life assessments.

2. Identation of the degradation sites, mechanism and stressors. and the potential failmm modes and evaluation of the current 1SI

methodr fmTe-oruse pendentdamageor

aging can be caused by one or several different mechanisms active within a component,

structure, or material and, if not recognized

and properly managed, may result in some

type of failure or impairment of function. The

degradation is often the result of interations

between design, materials, operational

stressors, and environments. Poor design,

improper material selection, severe environments, or inadequate maintenance practices

can accelerate the degradation. Therefore,

identification and understanding of the design, materials, stressors, environments, and

aging mechanisms is essendaL This step consists of identifying and qualitatively evaluating the stressors, potential degradation sites

and mechanisms, probable failure modes, and

current ISI methods for each of the selected

major coponients. This qualitative analysis

of the degradation sites and processes and

currentlSI methods Isessential to developing

a proper understanding of tie impact of aging

on safe operation of nuclear power plants and

identifying and prioritizing the unresolved

technical issues relevant to nuclear power

plant aging and license renewaL

3. Assessment of advanced Inpection, swoveUlance, and monitoring methodr Knowledge

of the current damage state of the material is

essential for a proper assessment of he residnal life of a component or structure. Inservice

Inspectns are performed to measure the current state of damage. However, many of the

standard nondestructive examination (NDE)

methods employed to satisfy current ISI requirnents were developed for the detection

and qualitative assessment of fabricationrelated flaws. These methods are not entirely

adequate for residual lifeassessment. Inspections for life assessment generally require

greater detection reliability and a more quanitative determination of defects and accumulated damage than tradital ISL. Therefore,

it is essential that emerging methods for insection be developed and evaluated and that

they accurately determine, for example, the

size, shape, location, orientation, and type of

both surface and internal flaws.

4. Develop (or evaluate) Ife assessment models

andprocedures: An important feature of this

NPAR task Is the development of models or

procedures to estimate fte remaining useful

life of the major, risk-sIgnificant LWR components and structures. The residual life

assessment models and procedures appropriate for various structures and components

may differ considerably. However, a general approach that may work well for many of the

reactor primary system pressure boundary

components, the containment, and possibly

other major components is as follows: (a)

evaluate the present state of the component,

(b) estimate the change expected during the

planned operating period, (c) identify appro- priate design requirements, and (d) compare

the estimated condition of the component at

the end of the planned operating period with

the design requirements.

The component state or condition of interest

might be one or more of its material properties. such as fracture toughness, fatigue

usage, strength, elasticity, etc., and/or one or

more geometry characteristics, such as flaw

size, wall thinning from wear, corrosion, and

erosion, tolerances, etc. These characteristics

can be determined by inspection, destructive

examination, analysis, or other methods.

However, note that the use of analysis to estimatekey matial properties (such as fracture

toughness or elasticity) and component geometry changes from wear, corrosion, erosion, flaw growth, etc., generally requires a

good knowledge of the initial material conditon and tolerances and a very good understanding of the relationships between

stressors and states. This understanding must

be based on appropriate data and physical

models.

The changes in the component state or condition during some futuae planned qperating petod (such as a license renewal period) can be

v
estimated based on conservative extrapoladon of previous ISI (for example, wall thickness data) and destructive examination

measurements or analysis. Again, the analysis tools should reflect a good undetanding

of the physical relationships between the operating stressors and the changes in the component state or condition.

The component or structural design requirements should be based on traditional engineering analyses (thrmal, hydraulic, stress,

neutronic, etc., as appropriate) or testing and

reflect appropriate margins between the expected loads and the calculated failure loads.

The original design analysis may be sufficient for some components. A revised design

analysis may be necessary for other components subject to additional cycles, previously

unknown stressors, etc. New information

may, in some cases, allow a reduction in or revised estimate of the margi

The final step, comparing the design requirements with the expected condition of the

component, would be followed, of course, by

a decision to replace or repair the component

or leave the component as is. The comparison

of design requirements with a component's

expected condition might also influence the

planned future operating period (license renewal period) and the frequency of future ISI

or destructive examination activities.

A somewhat different approach is associated

with components subject to environmental

qualification (EQ) requirements. These are

components that must function properly during a normal operating period of some given

duration and also during selected design basis

accidents (DBA). Here three general approaches might be considered: (a) compare

actual operating environments with the

preagingEQ environments, (b) compare estimated mechanical properties at the end of the

planned operating period with the preaged

EQ sample mechanical properties. and (c) remove samples from the plant, add additional

aging (to reflect the expected aging during the

planned future operaing period), and subject

the samples to DBA testing. The first two ap

proachesrequire modeling of the complex relationships between environmental stressors,

aging degradation (which are often material

property changes), and time. ITe third approach provides results with more certainty

but possibly at the highest cost.

The residual life assessment models or procedures will have a strong impact on the development of technical criteria for license

renewal. It is important that the limitations

and uncertainties associated with whichever

approach is chosen be assessed so that confidence intervals can be calculated and safety

margins properly assessed,

5. Development of technical criteriafor license

renewah The results and outputs obtained

from the activities discussed above will be

used to help the USNRC ident and resolve

technical safety issues associated with LWR

aging degradation and develop policies and

guidelines for making license renewal decisins The qualitative analysis of the degradation sites and processes and current IS1

methods (Step 2) will facilitate the identification and prioritization of the umesolved technical issues relevant to nuclear power plant

aging and license renewal. The qualitative

analysis will also help provide the USNRC

and industry a basis for lifa assessment model

(or procedure) development. The assessments of the advanced inspection, surveillance and monitoring methods could provide

a technical basis for acceptance or rejection

of any ISI associated with a life assessment

Most important, the use of acceptable life

assessment models or procedures could form

the technical basis for a revised license renewal safety analysis report (SAR) as well as

a revised NRC standard review plan and reg-

- guides.

Overall project results to date are as follows. The

first two steps have been completed, and results are

presented in a two-volume report: Residual Life

AsessmentofMajorLight WaterReactor Components

- Overview, NUREG/CR-4731. Some progress has

been made on Step 3. the assessment of emerging inspection, surveillance, and monitoring methods The

emerging inspectio monitoring and material evaluation methods are briefly discussed in Volume 1 of

NUREO/CR-4731. An indepth assessment of both

advanced fatigue monitoring and material evaluation

methods will be published in the near future. Indepth

assessments of improved acoustic monitoring, advanced ultrasonic testing, and eddy current techniques

are in progress.

Life assessment procedures will be published soon

for five major components as part of Step 4: PWR

Vi

I
reactor pressure vessels, reinforced concrete containments, cast stainless steel components, PWR stam

generator tubes, and metal containments. We plan to

publish these procedures In this multi-volume

NUREG report, as indicated below.

PWR Stea Generator Tubes - Volume 4

Metal Containments - Volume S

Life assessment niodels or procedures for the other

major components and structures will be presented in

future volumes of this report

PWR Reactor Pressure Vessels - Volume 1

Reinforced Concrete Containments - Volume 2

Cast Stainless Steel Components - Volume 3

P. E. MacDonald

V. N. Shah

September 1990

vii
EXECUTIVE SUMMARY

Many critical pressure boundary components in

commercial light waterreactors (LWRs) are composed of cast stainless steels Life assessment procedures are

needed for these components because cast stainless

steels are subject to thermal embrittlement during

long-term service exposure at LWR operating temperatures. The components of concern include pump

bodies, reactor coolant piping and fittings, surge lines

(in a few plants). pressurizer spray heads, check

valves, control rod drive mechanism housings, and

control rod assembly housings. These are made of

grade CF-8, CF-A, or CF-M stainless steel in U.S.

LWRs; grade CF-3 stainless steel also is used in some

foreign LWRs. The purpose of this project was to

review dte available data on thermal embrittlement of

cast stainless steels and to develop updated procedures

for life assessment of key LWR cast stainless steel

components.

Cast stainless steels have a two-phase microstructure consisting of ferrite islands in an austenite marx

With long-term thermal exposure at LWR operating

tempe e, other phases form in the ferrite phase

that cause it to become hard and brittle, while the austenite remains ductl If the amount of ferrite small

and if it is distributed evenly and finely throughout the

austenite, then the properties of the casting are not significantly affected by the thmal embrittlement of the

ferrite. However, as the amount of ferrite and the

coarseness of its distribution increase, the thermal embrittlement of the ferrite increasingly affects the properte of the casting.

The properties most affected by thtermal embrittlement are Charpy V-notch (CVN) impact energy and

fracture toughness (JIJ. Both of these properties decrease as the degree of thermal embrittlement increases. If these values become too low, the structural

integrity of a cast stainless steel component could be

seriously impaired The available data indicate that

thermal embrittlement does not have a large effect on

tensile, fatigue-crack initiation or low-cycle fatiguecrack growth behavior However, more fatigue-crackgrowth data are needed for CF-8 and for all cast

stainless steels in the high-cycle (low growth rate, near

threshold) regime. Thus, for life assessment of cast

stainless steel components, the main concern is loss of

fracture toughness and impact energy. Data and engi- neering models have been developed to help predict

the degree of embrittlement as a function of thermal

exposure history. For reactor internals, there also is a

concern that irradiation may embrittle the austenite

and add to the overall emzbriulementof the component.

No information on this potential problem was found

during this project, but it is an area that merits future

study.

The minimum CVN impact energy aft long-term

aging has been found to be proportional to the square

of the fraction of ferrite, the mean fernte spacing, and

a chemical-composition parameter. This model should

be developed further for application to the assessment

of components. A metparameter can be

used to define lower-bound trends to the available impact energy values for cast stainless steels as a function

of chemical composition and thermal exposure time,

In this project we popoase a model that uses that parameter to predict the impact energy decrease for any

particular lot of cast stainless steel. This predicted impact energy value or the predicted minimum impact

energy value is then used to estimate fracture toughness from correlations between impact energy and

fracture toughness at both room temperature and

2900C This proposed approach should provide a conservative estimate of frture toughness for use in the

structural integrity assessment of cast stainless steel

components.

Insnvice inspection (0SI) is needed to define type,

size, and location of defects in cast stainless steel components so that the structural integrity of these components can be evaluated. Use of radiography during ISI

Is less practical than during fabrication and is inefient. In addition, conventional ultrasonic testing (UI)

methods are not reliable to detect flaws in cast stainless

steel components because coarse grains and different

grain structures in cast stainless steel result in a low

signal--to-nise ratio. Advanced UT methods with enhanced signal-to-noise ratios are being developed. These methods have an improved capability to detect

flaws in cast stainless steel components and have been

used to inspect statically cast components in several

PWR plat The effect of grain structure on the propagation of ultrasonic waves in cast stainless steel material has been studied to increase the accuracy of

locating defects. Because of the difficultes with radiography and UT methods to detect and size flaws,

application of the acoustic emission technique to detect crack growth in cast stainless steel needs to be

evaluated.

Weld repair procedures are employed to remove

fabrication defects during the original production of

cast stainless steel components. Itshouldbepossibleto

vii
adapt, modify, and qualify those procedures for

performing weld repairs to components that are in

service in order to extend their useful life.

We developed a procedure for estimating the current

condition and residual life of key LWR cast stainless

steel components. The procedure is implemented in

-nine major steps. The first three steps involve the collection, examination, and storage of records for fabrication and construction, inservice inspection, and

operating history. Ike fourth step involves a conservative fatigue and fracture mechanics evaluation to determine the worst-case flaw size and the minimum

reqWred fracture toughness at the end of the next operatng period I the fifth step, the current condition of

the mateial is assessed using either a proposed analytical model, microstuctural daa or measured properties (or using some combination of those three

methods). In the sixth step, the results of the fourth and

fiffh steps are combined to evaluate the structural integrity of the component. The seventh step establishes

what actions (none, repair, replace, or shut down) are

to be taken, and the eighth step establishes the plan for

the next ISL From this point, the steps are repeated (the

component is reevaluated) as needed.

Future work is recommended to provide information for use in implementing the life assessment procedure. Guidelines of worst-case thermal emnrittlement

data should be developed. Thermal eonbrittlement surveillance programs should be initiated. More impact

and fracture toughness data should be developed for

material aged and tested at maximum LWR operating

temperature. Ihe aging parameter model should be

improved to better account for the observed mechanisms of thermal embrittlement. The work on evaluation of the relation of farite amount and spacing to the

degree of embrittlement should be continued. Highcycle faigue-ack-growth tests should be performed

on aged and irradiated materials, especially CF-8. The

use of field replication and magnetic measurements for

charactfizin cast stainless steels should be demonstrated and validated. Techniques involving the use of

substandard sized specimens and microhardness to

quantify the degree of thermal embrittlement should

be developed. Repair welding procedures should be

qualified and validated. Fracture-nechanics analyses

should be performed for representative combinations

of components and operating histories to estimate the

minimum fracture toughness required for safe operadon. ISI methods and programs should be modified to

inco te procedures and evaluations appropriate to

cast stainless steel components.

Lx
ACKNOWLEDGMENTS

Ihe authors of this report acknowledge the significt role that G. IL Weidenhamer of

the US. Nuclear RegulatoryCommission has played in providing programmatic guidance

and review for this report

The authors sincerely appreciate the critical review of this report performed by B. L

Landerman. The authors thank 0. K. Chopra and H. M. Chung of Argonne National

Laboratory and M. B. Lapides of Electnc Power Research Institute for several telephone

conversations and meetings related to aging of cast stainless steel components. The authors

also thank 0. K. Chopra for providing the data from the Charpy impact tests of aged cast

stainless steel specimens, and thank K E. Lapides for providing a summary of a draft EPRI

report on aging of cast stainless steel components. The authors thank C. W. Rainger of Sandusky Foundary & Machine Company for reviewing the report.

Finally, the authors thank D. R. Pack and D. L Bramwell of the Idaho National Engineering Laboratory for their technical editing assistance and the significant effort they put forth

in coordinating the production of this report, and to K. 0. Roberts and B. L. Thompson for

their special effot in composing the document.

I

I
CONTENTS

ABSTRACT ................ . . .. . '.'. . . iii

RqE3L F1REWVORD ............................ I.................. 11iv

EXECUTIVE SUMMARY viii

1. . . i ACNKN W E ON~il]EM ................................................................. .. x

1. DlR M OD UCnION ..... ............................ t-. ''---'*--.-..-....................-

1.1 LWVR CastStadnlessStee Conponents ............................, I

12. Physical Metallurgy ofCast Stainless Steels ......... ...................... 6

13 Benefits of Ferrite in Cat Stainless Steel ................................. ............ , 9

2. REVIEW OF INFOROATION O THTEREMAMLAELMBR]LEM NN ....... ................. 11

2.1 Basic Mechanisns . ...... . .*. ... . . . 11

22 Charpy Inpact Energy Data ........... * ................... 12

23 FracturtebughnesslDta ...................... 21

2.4 lensile nd FatigueProperties .....................................-. ;. 24

3. iNSERvicEINSPECTiONOF C ONPENTS . . ................................. 27

4. ASSESSMENT OF DEGREE OF MATERIAL EMBRITTLEMENT ......................... 29

4.1 Analytical ModelingofInservice Degradation ... ........ . 29

4.2 'metAurgicalEvaluationiMethods . .. 30

4.2.1 icostuctureC terization. 30

4.22 TIsting of Miniature Samples ................ 34

4.3 Useof Ultrasonic lsting toCharacterize Matrials . ................................. 35

S. FUELDREPAIR PROCEDRES . . ...................................... 36

6. IJFE ASSESSMIENrTPROCEDtURE . . .................... ............. 3737.... 

7. Sl.j0:UARY ........................ 42

8. RECOMMNDED FUEW FL.R.E ORK 1. .. . . .44

9. FER....ES.. . .. 45

xi
FIGURES

1. PW!R TypePcoolantpumnp ................................................. 3

2. PW TRlWpeEcoolantpump . ......................................................... 3

3. BWDR Type C coolantpump . ......................................................... 4

4. Typical microstructures of centrifugally cast austenitic-ferritic stainless steels, with islands of

ferriteinanaustenitematrix . ........................................................ 8

5. Effect of thermal aging on the room-tempeaue Charpy V-notch impact energy of CF-3

stainless steel .................................................................. 13

6. Effect of thermal aging on the room-temperature Charpy V-notch impact energy of CF-S

stainless steel ...................... 13

7. Effect of thermal aging on the room-temperature Charpy V-notch impact energy of CF-SM

staiesasteel ...................... 14

8. Effect of test temperatmr and thermal aging on the Charpy V-notch impact energy of CF-3

stainless steel, data from Chopra (1990)

a. Cast76-imm slab. Heat69 ...................... 16

b. Castpump impellervanes, Heat ...................... 17

c. Centrifugally castpipe, HeatP2 ...................... 17

9. Effect of test temperature and thermal aging on the Charpy V-notch impact energy of CF-8

stainless steel cast 76-mm slab, Heat 68, data from Chopra (1990) .......................... 1B

10. Effect of test tempeature and thermal aging on the Charpy V-notch impact energy of CF-SM

stainless stee, data from Chopra (1990)

a. Cast 76-4nm slab, Heat 70 . ....................................................... 18

b. Cast76-mm slab, Heat74 . ....................................................... 19

c. Cast76-mm sLb, Heat 75 . ........................................................ 19

11. Correlation between minimum oomn-teperature Charpy V-notch impact energy and material

parameter ' for aged cast stainless steels, adapted from Chopra and Chung (1989a), data supplied

by Chopra (1990c) ................................................................. 21

12. Correlation between room-temperature fracture toughness (Ja and impact energy for cast stainless

steels . .................................................................. 22

13. Correlation between rm pae tearing modulus (l) and fracture toughness (JIc) for cast

stainlesssteels .................................................................. 22

14. Correlation between the 2900C fractue toughness (Jic) and onom-tempera impact energy for

cast stainless steels ........................ 23

15. Correlation between the 290°C tearing modulus (1) and fracture toughness (Jic) for cast stainless

steels .......... 23

xiD

I
16. Fatigue crack growth rae of CF-SM stainless steel, in air and in a PWR environment, at a cyclic

frequency of 0.01711z . ............................................................. 25-

17. Predicted lowerbomnd Charpy-mpact values for cast stainless steels as a function of aging time

at 300 0C (5000F) . ................................................................. 30

18. Ferrite content of CF-8M castings as a function of secdon thickness .......................... 32

19. Comparison of ferrite content predicted from chemical composition with measured fete

contentusing lodel 1 of Aubrey etaL(1982) ........................................... 33

20. Comparison of ferrite content measured using a ferritescope with that determined metallographically

bya point-countproceduie .......................................................... 34

21. Generic p edure for he evaluation of LWR cast stainless steel components ....... ........... 38

TABLES

1. LW1Rcam stainlesssteelcomponents ................................................ 2

2. Castprimary coolantpiping materials .................................................. 5

3. Required chemicalcompositions of casttainless steels .................................... 7

4. Required room-temperature tensile properties of cast stainless steels ......................... 7

5. Results of electron microprobe analysis of two centrifugally cas stainless steel samples .......... 9

xiii
LIFE ASSESSMENT PROCEDURES FOR

MAJOR LWR COMPONENTS

VOLUME 3: CAST STAINLESS STEEL

COMPONENTS

1. INTRODUCTION

All opeatig commercial nuclear power plants in

the United States are light water reactor (LWR) systems; they employ a variety of cast stainless steel components in the primary pressure boundary. Cast

stainless steel is used because of its high resistance to

hotcracking, corrosion, and strss-corrosion cracking,

and because of its improved mechanical properties

(Solomon and Devine, Jr., 1983). In addition, use of

cast stainless steel provides benefits of the casting

process in manufacturing. Complex shapes, such as

pump housings, valve bodies, and fittings are statically

cast; usually the molten metal is poured into a fixed

sand mold. Cylindrical shapes, such as pipes are centrifugally cast; te molten metal is poured into a rotating sand-lined metal casing.

Maintaining the structural integrity of these

components is essential to the safe operation of the

power plants. Thus, any mechanism that may

degrade the mechanical properties of the cast

stainless steels used in these components

should be considered in efforts to extend the operating

life of an LWR. Thermal embrittlement at LWR operating temperatures is the major degradation mechanism of concern for cast stainless steel components.

Typical operating temperatures in pressurzed water -

reactor (PWR) primary coolant piping varies from 288

to 327°C (550 to 621F) (S hah and MacDonald, 1987),

and the PWR pressurizers operate at 3430C (650-F).

The typical operating temperature in boiling water

reactors is about 282?C (S40F) (Shah and

MacDonald, 1989). For service exposure in this relatively low-temperature range, the embrittlemrent occurs at a very slow rate over a period of many years.

As aging progresses, and the degree of embrittlement increases, the critical flaw size decreases; thus,

even initial defects that do not grow, as well as those

that may grow subeitically, may in fime become defects of critical size. During strucural integrity assessments of cast stainless steel components, one must

determine if the critical defect size will become too

small to be reliably detected during inservice inspection (ISI). Finally, leak-before-break &LB) analysis

should be perfcrmed to deternine if a detectable leak

will develop before a growing crack reaches a critical

size for unstable ppagation.

The concern about possible thermal embrittlement

becomes progressively more important as cast stainless steel components enter the second half of the

40-yr period of a nuclear power plant's original operating license, and as extending the operating license

for an additional period of 20 years or more is considerel The potential problem of thermal embriulement

is aggravated by the fact that the current nondestructive examination (NDE) techniques used for wrought

stainless steels usually are not as effective when

applied to cast stainless steels.

The objectives of this repo't are to review the available information on thermal embrittlement of cast

stainless steels, discuss methods of assessing the

degree of embrittlement for inservive cast stainless

steel components, discuss the possible effects of inservicerepair welding on thermal embrittlement, and suggest possible approaches for evahmuing the remaining

safe life of cast stainless steel components, approaches

that can be used in efforts to extend the operating license of a light water reactor. To provide background

for these discussions, the following sections of this intoduction briefly describe (a) the LWR cast stainless

steel components of interest and (b) the physical metalluy of cast stainless steels.

1.1 LWR Cast Stainless Steel

Components

Table I lists the eight LWR cast stainless steel components considered in this project The foLlowing paragraphs give brief descriptions of these components.

Reactor Cool=nPw*pBody. Three types of reactor

coolant pump bodies are used in LWR plants. Ues F

andE pumps are used in PWRs, nd Type C pumps are

used in BWRs. Pump bodies we manufactured using

CF-8, CF-SA, or CF-8M statically cast stainless steel

sections that are subsequently welded. The casting is

typically done in more dan one piece because of the

I
Table 1. LWR cast stainless steel components

Component

Reactor Coolant

Pump Body

CoolantPiping

Coolant Piping

Fittings

SurgeLines

Pressurizer

Spray Head

Check Valves

cfpeci

Casting

Static

Centzifagal

Static

Centrifugal

Grade

CF-8

CF-8M

CF-8A

CF-SA

CF-8M

CF-8A

Number

of Planti Comments

Type F in 61 plants

Type } in 15 plants

Type C in all BWRs

20

7

27

44% of Westinghouse

plants

All Westinghouse

plants

CF-8M SomeCombrustion

Engieing Plants

Static Westinghouseplants

Static CF-8M Some Westinghouse plants

Recirculation

Piping Fittings

Control Rod Drive

Mechanism Housing

Static

Static

CF-8

CF-8M

BWRplants

CF-8 Some Westinghouse plants

Care Inenals Static CF-8, CF-3 PWR and BWR plants

size of the pump body, and the pieces are welded

together to produce the fmal assembly (Shah and

MacDonald, 1989). Recent designs of Type F pump

casings decreased or eliminated the number of weldments as it became practical to manufacture larger

casting sections. The pump impellers are made of

CFM- cast stainless steeL Figures 1 through 3 show the

pump configurations (ASME Code Section III).

Figure 1 shows a Type F pump. Most Type F pumps

are manufactured by Westinghouse. Their wall thckness ranges from 100 to 200 mm (4 to 8 in.). These

pumps are used in Westinghouse and some Babcock &

Wilcox (B&W) plants. Type F pumps are used in a

total of 61 nuclear plants. The Bingham-Willamette

Company also makes Type P pumps that are used in

some B&W plants. Tlpe F pumps are manufactured

using an electroslag welding process with no postweld

heat reatment the lack of such treatment can leave

some high residual stresses at levels near yield

strength. (Section m of the ASME Boiler and Pressure

Vessel Code for Class I components neither requires

nor prohibits postweld heat treatments for austenitic

stainless steel welds.)

Figure 2 shows a Type E pump. 7ype E pumps are

manufactured by Byron Jackson Pump Division,

Borg-Warner Corporation. These pumps have wall

thicknesses in the range of 60 to 80 mm (2 to 3 in.).

They are used in 15 B&W and Combustion Engine"-

ing (CB) plants. These pumps are manufactured with

manual welds followed by a full solution heat treatment, which eliminates nearly all the residual stresses

if the assembly is slowly cooled from the solution

treating temperature. However, the weldedcastings are

quenched from the solution treatment temperatures

[-1038 to 11490C (1900 to 2100F)l to avoid sensitization, which can leave high residual stresses (Shah

and MacDonald, 1989). Any improvement provided

by solution heat treatment is limited because the sections of th pump body are thick. The optimum improvement will be at and near the surface, because the

2

I
rate of cooling is important. Tb minimize sensitization

and residual streasses In Type E and C pumps (described in the next paragraph), aI1 welding performed Discharge

after the solution heat treatment of the material is limited to a low-heat input no greater than 1900 J/cm nozzle

(5000 ) Crot) Horizontal f -4* ross

Figure 3 shows a Type C pump. Type C pumps are Section AA

manufactured by Byron Jackson Pump Division.

Borg-Warner Corporation, for use in all GE BWR ump shaft

plants. These are smaller than the pmnps used in PWR

plants. The smaller size accommodates the lower

system pressures and space limitations of BWR Type

C pumps also are manufactured with manual welds

followed by a full solution heat treatment. In most Closure

cases, these welded castings also are quenched from studs

the solution treatment temperature, which can leave

high residual streCSe.Upe

flangeo a e

Pump casing wall

Discharge casing

nozzle

Vertical cross section 9-3044

Weld

joint Figure 2. PWR Type E coolant pump.

Reactor Coolant Piping. he newer Westinghouse

plants use centrifugally cast CF-8A or CF-8M stainless steel for the main coolant piping (Shah and

MacDonald, 1987). CE, B&W, and older

Closure Westinghouse plants did not use cast stainless steel for

studs _main coolant piping. Table 2 gives the names, sizes,

operating age in years, and piping materials for the 27 Pump _Westinghouse plants with cast stainless steel primary

shaft Ipiping (gan et al., 1987; USNRC, 1988a). The plants

are listed in chronological order of their operating liDscha lcense dates. The typical diameter of te coolant piping is about 0.8 m (32 in.). Almost no problems have been

detected in centrifugally cast piping to date (Egan

Weld Het al., 1987). joint 4 @ ai

Pump Reactor Coolant Piping Fittings. The main coolant

casin \ 7 piping fittings, such as elbows, are statically cast

CF-A stainless steel. Al Westinghouseplants use this Impelle 9-046 material in the primary coolant pipe fittings. Generally,

a larger population and deeper defects are expected

Figure 1. PWRTypeFcoolantpump. in statically cast components than in welds or

3
I

centrifugally cast components This should be taken

into account when selecting sites for inservice

inspections.

steam space of the pressurizer. The cod-leg coolant

sprays through the spray head into the pressurizer to

reduce the pressure in the primary coolant system In

Westinghouse plants, the spray head is made of cast

stainless steel and is subjected to an opezating tmperature of approximately 3430C (6500

F), which is the

sanution temperature at the PWR operating pressure.

Check Valves. Check valves with cast stainless steel

bodies are installed in LWR safety systems having

stainless steel piping. For example, in Westinghouse

emergency core cooling systems (ECCS), there are

cast stainless steel check valves to prevent flow from

the primary loop hot leg or cold leg back into the

ECCS piping (USNRC, 1988b; Greenstreet et al,

1985).

Recirculation Piping Fiaings and Valves. The BWR

recirculation piping fittings, such as elbows and tees,

and pump suction and discharge valves, are statically

cast CF-8 and CF-8M stainless steel.

Convrol Rod Drive Mechanism Housing. The control rod drive mechanism (CRDM) housing constitutes

a primary pressure boundary. In certain Westinghouse

plants, some of the CRDM housings are made of CF-S

cast stainless steel (Shah and MacDonald, 1989).

CocImnternals. Core internals made of cast stainless

steel may be subjected to both thermal and irradiation

embrittlement during long-tem service. In CE plants,

the control rod assembly shrouds are made of centrifugally cast CF- stainless steel. In Westinghouse-type

plants, the lower support structures, cruciform instruments guides, and flow mixer plates are made of stainless steel castings. In GE plants, the orificed fuel

supports are made of CF-8, and parts of the jet-pump

assembly are made of CF-8 and CF-3 stainless steel

castings.

An evaluation of the effects of irdation on stainless steel castings should be made. This report reviews

the thernal embrittlement of such castings. The thermal embrittlement is caused by embrittlement of the

ferrite phase in the ferrite/hustenite microstructure that

is inherent in these alloys. The austenite phase is ductile and remains so after thermal exposure; thus, when

the fraction of ferrite is small, the degree of overall

embrittlement of the casting is low. However, if the

components used in internals are subjected to high

levels of irradion, the austet phase could become

embrittled as well. In such a case, where both the austenite and ferrite phases are embrittled, the overall

structural integrity of the casting could be in question.

4

FIgure 3. BWRIypeCcoolantpump.

SurgeLine. The surge line connects the reactorcoolant piping hot leg to the pressurizer. As the temperature of the reactor coolant fluctuates, the volume

increases or decreases, causing water to flow into or

out of the pressurizer through the surge line. Typically,

the surge line is made of austenitic stainless steel.

However, in some Combustion-Engineering plants the

surge line is made of CF-8M stainless steeL This line

is typically 300-mm (1-in.) diameter Schedule 160

pipe (Shah and MacDonald, 1989). The pressurizer

end of the surge line has operating temperatures as

high as -3430C, which is the highest operating temperature of the major PWR primary coolant pressure

boundary components.

Pressurizer Spray Head. T pressurizer spray head

is the termination of the spray line and is located in the

I
Table 2. Castprimary coolantpipingmaterials

(Ean ct al., 1987; USNRC, 1988a)

PlantName

Kewaunee

Prairie Island 2

cook I

Farley A

Trojan

Cook2

Beaver Valley I

North Anna I

North Anna 2

Farley 2

MeGuire 1

McGuire 2

Callaway 1

Sequoyal 1

Sequoyah 2

Catawba 1

Wolf Creek

Millstone 3

Citawaba 2

Beaver VAlley 2

South TeM I

South lTxas2

Vogtle 2

ComanchePeak I

Commanche Peak2

Watts Bar I

Unit

Size

(MWe)

560

530

1090

82g

1130

1054

852

934

788

829

1180

1180

1157

1140

1140

.1153

1158

1150

1153

1113

852

1250

1250

1113

1150

1150

1177

Number

of System

2

2

4

3

4

4

3

3

3

3

4 i

4

4

4

4.

4

4

4

4

4

3

4 ..

4

*4.. 4 -

I. 4

,4 . .

Operatng

Age

(Years)

14

14

12

10

9

9

9

9

8

8

7

5

5

5

5

4

4

4

3

3

2

2

1

1'

0

0

.0

Pipe Material

CF-SM

CF-8M

CF-8M

CF-8A

CF-8A

CF-SM

CF-8M

CE-SM

CF-8A

CF-8A

CF-SA

CF-8A

CF-8M

CF-8M

CF-8A

CF-8A

CF--A CF-8A

CF-8A

CF-8A

CF-SA

CF-8A

CF-8A

CP-8A

CF-8A

a. These values were derived from United States Nuclear Regulatory Commission operating reports

(NUREG-0020) for plant service up to y 1990. The values represent the number of years that have elapsed since

the operating license dates.

5
1.2 PhysIcal Metallurgy of Cast

Stainless Steels

The required chemical compositions and the required tensile properties of six common grades of cast

stainless steel are listed in Tables 3 and 4, respectively.

TheomnposidonsofCF-3,CF-8,CF-3MandCF-SM

are similar to those of the wrought grades Type 304L,

Type 304, Type 316L, and Type 316, respectively. The

wrought grades have microstructures that are usually

fully austenitic, whereas the cast grades typically have

micrtru resconsistingofabout5 to 25% (volume)

ferrite and the balance austenite (Peckner and

Bernstein, 1977). (The ferrite phase in cast stainless

steels and stainless sted welds is sometimes referred to

as the delta ferrite because it forms at high temperatures, where the ferrite field of the phase diagram is

known as the delta phase to distinguish it from the lowtemperature ferrite phase known as the alpha phase.)

Theprmaryfacorcontrollingtheferrite-austenitebalance is bulk chemical composition. Chromium, silicon,

and molybdenum, and niobium (if present) promote the

formation of ferrite, whereas nickel carbon, manganese, and nitrogen promote the formation of austenite

(Peckner and Bernstein, 1977). The nitrogen is introduced during the casting process, but its percentage is

not specified. The typical values of nitrogen in the

chemical composition of cast stainless steels may vary

from 0.03 to 0.08%. Empirical relationships have been

developed forestimating the faerite content from chemical composition; the use of two such relationships are

illustrated later in this report (Aubrey et al., 1982;

Bonnet et al, 1990).

The size, distribution, and morphology of the ferrite

within the austenite depends on solidification conditions during the casting process. The conditions affecting the solidification of metal are complex (Reed-Hill,

1973) and a detailed discussion of them is beyond the

scope of this reporL Cast stainless steels may solidify

with a columnar or equiaxed grain structure or a mixture of both structures. Researchers at Argonne

National Laboratory (ANL) have observed that in

steels of similar chemical compositions, the average grain size and spacing between ferrite islands generally tends to increase with larger section sizes and corresponding slower heat removal rates.' In heavy-wall

castings, there can be considerable variations in microstructure across the wall of the casting. The complex

nature of metal solidification during casting makes it

difficult to predict the finer details of the actual microstructure in the cast stainless steel components of

a. Private communication with Dr. T. Kassner,

Argonne National Laboratory, May 1988.

interest in this project. Thus, if knowledge of these mlcrostructural details Is important for the assessment of

casting integrity and service life, some means of measuring them must be employed.

ncreased ferrite content increases the yield and tensile strength of the cast stainless steels, so grades

CP3A and CF-8A have controlled ferrite-austenite

ratios to increase the minimum strength levels

(Peckner and Bernstein, 1977). The minimum strength

levels are indicated in Table 4. These controlledferrite grades have the same chemical requirements as

CF-3 and CF-8; the ferrite is controlled by adjusting the cheidcal composition within the imits shown in

Table 3. Experience has shown that castings made

from the alloys listed in Table 3 may contain up to 30%

ferrite when control of the amount of ferrite is not spedcfib Although ferrite levels near 30% are not usual

for castings made in accordance with ASTM specificadons (e.g., A-743 and A-744), they are possible.

As shown in Tables 1 and 2, LWR components have

been made from CF-8, CF-8A, and CF-8M stainless

steeL CF-8 Is used for its genral resistance to oxidizing conditions. Compared with CF-8, the CF-8M

alloy has improved resistance to reducing environments as well as high resistance to oxidizing condidons. Therefore, the CF-8M alloy is widely used for

corrosion-resistant pumps and valves. CF-8A is used

to provide mechanical strength properties higher than

those of the CF-8 alloy. The low-carbon versions

(CF-3, CF-3A, and CF-3M) may be used in the future

to provide an added margin of resistance to intergranular stess-corrosion cracking (IGSCC), especially in

BWRs, where many IGSCC problems in recirculation

piping have been encountered in the past (Shah and

MacDonald, 1987).

Examples of austenitic-ferritic microstructures observed in cast stainless steel are shown in Figure 4.c

These samples were obtained from centrifugal castings. The ferrite phase (darkened networks and/or

islands) is contained within a matrix of the austenite

phase. The sample shown in Figure 4a had about 16%

ferrite; the sample shown in Figure 4b had about 28%

fer The local chemical composition of the ferrite

regions is significantly different from that of the ausnite regions in these alloys. 7h illustrate this point, the

local chemical compositions of the samples shown in

b. Private communication with D. 3. Roach,

Battelle Columbus, Ohio, August 22, 1986.

c. C. IL Jaske, unpublished results of studies conducted at Battelle, Columbus, Ohio, 1987.

6

I
Table 3. Required chemical compositions of cast stainless steelse

* Chemical Composition

(percent by weight)

Element

Carbon

Managanese

Sio

Sulu

Phosphorus

Chromium

Nickel

Molybdenum

CF-3A

0.03

150

2.00

0.040

0.040

17.0 to 21.0

8.0 to 12.0

0.50

CFLSA CFL8A

0.08

1.50 -

2.00

0.040

0.040 .- -

18.0 to 21.0

8.0to 11.0

0.50

CF-3M

0.03

1.50

1.50

0.040

0.040

17.0 to 21.0

8.0to 12.0

2.0-3.0

CFE;M

0.08

1.50

1.50

0.040

0.040

18.0 to 21.0

8.0 to 11.0

2.0-3.0

a Frm Secdon II-Mat Specifications,PartA-F=rous Materials, 1983 ASME BoilerandPressure Vessel

Code.

b. Maximua except where range is indicated. -

Table 4. Required room-temperaturc tensile pperties of cast sainless stee

Propert y

(minimum) CF-3 CF-3A . CF-8 CF-8A CF-3M CF-aM

Ihisile stength, 485 530 485 530 485 485

MPaU(si) (70) (77) (70) (77) (70) (70)

0.2% offset yield 205 240 205 240 205 205

strengthMPa(ksi) (30) (35) (30) (35) (30) (30)

Elongation in 50mm 35.0 35.0 35.0 35.0 30.0 30.0

orsn2in.

a. Rom Section U - Material Specifications, Part A - Frous Materials, 1983 ASME Boiler and Pressure Vessel

Code

7
p2'~~

- '

61oo90o

6M-2

Glyceregla etch 6M203

a. CF-8 stainless steel

1L - W - V -N -. 

10oX Glyceregla etch 6M202

b. CF-8M stainless steel

Figure 4. TI'pical microsntruces of centrifugafly cast anstelii-fenitic stainless steels, with islands of ferrite

in an austenite matrix.

8
Figure 4 were analyzed using an electron microprobe;

the results ar presented in Table 5. Tbe overall bulk

composition determined by the foundry is compared

with the local composition within the ferrite and the

austenite regions. Bulk values for silicon, manganese,

sulfur, and phosphorus are not shown in Table 5

because local values were not measured for these four

elements. The local values were obtained from the

average of six individual determinations at randomly

selected regions within each particular phase. As expected, the ferrite phase in both samples contained

higher chromium and molybdenum levels and lower

nickel levels than the bulk material.

1.3 Benefits of Ferrite In Cast

Stainless Steel

Ferrite in cast stainless steel provides resistance to

sensitization, which is caused by precipitation of

chromium-rich carbides MA23C) and the resulting depletion of chromium at het grain boundary. Sensitization describes the susceptibility of a stainless steel to

intergranular stress-corrosion cracking (1GSCC) following certain heat treatments such as a welding process or a slow cooling from the solution annealing

treatnent. Duing welding, temperatures are reached

in the heat-affected zone that cause precipitation of

chromium carbides. In cast stainless steel, the chromium carbides preferentially precipitate at the ferriteaustenite interfaces, not at the austentite grain

boundaries. The faster diffusion rate of chromium in

the ferrite phase results In formation of carbides on the

ferrite side of the interface. A chromium-depleted

zone, which forms on the ferrite side of the interface, is

quickly replenished with chromium because of the relatively high chromium content of the ferrite phase (see

Table 5) and the high diffusivity of chromium in ferrite. Industrial experience indicates that components

with high ferrite levels have more resistance to ISCC

(Peckner and Bernstein, 1977)2 than do components

with low ferrite levels. For example, many duplex (40

to 60%) stainless steels have been found to be highly

resistant to IGSCC in aqueous environments in nonnuclear applications (Peckner and Bernstein, 1977).

The minimum amount of ferrite needed to avoid

stress-coroion cracking susceptibility of cast stainless steel In an LWR environment depends on the carbon content in the cast stainless steel; higher carbon

a Private communication with D. B. Roach,

Battele Columbus, Ohio, August 22, 1986.

Table 5. Results of electron microprobe analysis of two centrifugally cast stainless steel samples

Chemical Compositiona

(percent by weight)

CF-8 Overall8

Austenite

Ferrite

Fe

69.2

69.6

66.7

63.7

63.6

62.0

Cr

203

19.9

27.7

21A

213

25.7

Ni

8A0

8A

3.7

9.70

10.6

6.0

Mo

0.11

0.03

0.06

2.81

C

0.044

0.06

0.03

CF-8M Overall' 0.053

Austenite

Ferrite

2.3 _b

4.0 _b

a. Overall composition was determined by ladle analysis asrepcrtedby foundry; composition ofausteniteandferrite regions was determined by electron microprobe measurements.

b. Value was not determined.

9
I

content requires a higher minimum amount of ferrite

(Devine, 1980) For example, in a BWR environment

cast stainless steel with a maximum carbon content of

0.035% requires a minimum ferrite content of 7.5% for

resistance to sensitization during welding (Hazeton,

1988). Results of laboratory tests on CP-3 and CP-8

materials in an accelerated BWR environment show

that IGSCC is not expected with a ferrite content

greater than 12% (Hughes et al., 1982). Inadditionthe

test results show that no IGSCC effects are expected

with a maximum carbon content of 0.03% in weld sensitized material and of 0.015% in furnace-sensitized

material, regardless of ferrite content. Similar evaluations for CF-8M material are needed. The test results

also show that in addition to the amount of ferrite, distribution of ferrite or ferrite spacing plays an important

roe in determining stress-corrosion cracking susceptibility ofcast stainless steels. A finer distribution of ferrite provides better resistance to stress-corrosion

cracking.

worst case, this is likely to result in a local leak rather

than in any substantial loss of system integrity, because

of the heterogeneous nature of the matel. The integrity of weldments is important because welding is used

as a joining procedure during abication, a method of

repairing casting defects during fricatio an installadon procedure, and a method of repairing components for the purpose of extending their operational

lifetimes.

The presence of some ferrite (typically at least 5%)

minimizes the occurrence of hot cracking or microfissuring in castings and weld metal.b In several older

plants, the weld metal in the pump casing and in stainless steel RCS piping have less than 3% ferrite,

whereas in newer plants a minimum of 5% ferrite is reqired for the reactor coolant pressure boundary components (USNRC, 1978).

In BWR plants, stress corrsion cracking of the sensitized heat-effected zones of cast stainless steel is

possible and has been observed.$ However, in the

a. Private communication with NL E. Lapides,

Electric Power Research Istitute, Palo Alto,

Cafoia, 1990.

b. PrivatecommunicationwithE.Landermanconsultant, Pittsburgh, Pennsylvania, September 1988.

10
2. REVIEW OF INFORMATION ON THERMAL EMBRITTLEMENT

Understanding the basic mechanisms that cause

thernal embrittlement damage to stainless steel cornponents is necessary for identifying the aging parameters related to design, fabrication, and operation of

these components. The degree of material embrittlemeat in aged cast stainless steel components is most

often evaluated by Charpy V-notch (CVN) impact energy tests and, recently, by fracture toughness tests.

Tensile and fatigue prp es for the aged components

also are needed in evaluations of the integrity of the

cast stainless steel components. This section reviews

the basic mechanisms causing thermal embrittlement

damage, evaluates the currently available Charpy impact energy and fracture toughness data, and then

briefly discusses the low-cycle fatigue behavior of the

cast stainless components.

2.1 Basic Mechanisms

The mechanisms of thermal aging that cause embrittlement in cast stainless steels have been reviewed

and evaluated extensively by Chopra and Chung

(1985, 1986a, 1986b, 1986c, 1987, 1990a, 1990b),

Chung and Chopra (1987), Chung (1989), Chopra and

Sather (1990), Chopra (1990), Auger et al. (1990),

Bonnet et al. (1990), and Brown et al. (1990). Embrialement or loss of toughness in cast stainless steels

during elevated-temperature exposure is related (a) to

the precipitation of carbides at the austenite-ferrite

phase boundaries and (b) to the formation of the

Cr-rich alpha-prime phase and the Ni-rich and

Si-rich 0 phase in the ferrite. The phase-boundary

carbidesplayasignificant role in embrittlement for exposures at temperatures greater than 4000C (750M3,

but have kss effect on the embrittlement at exposure

temperatures less than 4000C.

For LWR orating temperatures tless than 3500C

(660)], the formation of the alpha-prime phase and

the 0 phase in the ferrite are the primary factors involved in embrittlement. Also, the kinetics of the formation of these phases appear to be different at

temperatures ess than 4000C (7500P) than at temperatures greater than 4000C. Because of these differences

in formation and precipitation behavior, the results of

tests an material subjected to accelerated aging at temperatures greater than 4000C should not be extmpolated to the lower LWR operating temperatures.

Bamford et al. (1987) of Westinghouse describe the resuits of analyses performed using x-ray energy dispersive spectroscopy. These results indicate that data from

accelerated aging studies at temperatures greater than

4000C (7500F) could be used to predict behavior at opcrating temperatures. Additional, independent research is needed to resolve the differences between the

ANL and Westinghouse studies.

The alpha-prime phase typically forms by the

process of spinodal decomposition (Chopra and

Chung, 1987, Sassen et al, 1987). Spinodal decomposition refers to the reaction whereby two phases of the

same crystal lattice type, but different compositions

and properties, form because a miscibility gap exists in

the alloy system. In the iron-chromium system, these

immiscible phases are known as the iron-rich alpha

phase and the chromium-rich alpha-prime phase. This

phase separation process occurs at a very fine scale (on

the order of only a few nanometers) in the ferrite regions of cast stainless steel, and use of the atom probe

field Ion microscope Is required to resolve the presence

of the alpa-prime, phase (Sassen et al, 1987). There

are indications that after extensive (many years) aging

at LWR operating temperatures the alpha-prime phase

can form also by means of a nucleation and growth

process, as well as by spinodal decomposition (Chopra

and Chung, 1987). Depending on the composition of

the ferrite and the exposure temperate, either or both

of these processes may be involved in the formation of

alpha-prinmephase.

The 0 phase forms in te ferrite by a nucleation and

growth process, and its rate of formation increases

with increased levels of carbon and molybdenum

(Chopra and Chung, 1987). Butthe 0 phase has no direct effect on the degree of embrittlement, as demonstrated by the following laboratory test. A CF-8

stainless steel pump cover was found to be embrittled

after 12 years of service in aBWR (Chopra and Chung,

1987). Annealing for one hour at 5500C (10200F)

dissolved the alpha-prime phase and restored the

Charpy-impact resistance to the level expected for

unaged material, but had no effect on the 0 phase

present. Thus, the G phase had no significant effect on

the degree of embrittlement, and the alpha-prime

embrittlement was easily reversed by a short heat treatment (one hour) at a moderate temperature [5500

C

(10203].

The annealing treatment described above was

applied to verify that embrittlement was caused by the

alpha prime. It is not proposed as a method for reversing the effects of alpha-prime embrittlement in actual

components because of the practical difficulties that

11
would be encountered in applying such a heat treatment in the field.

Th pesece of G phase may indirectly affect the

degree of embittlement in cast stainless steel components; however, its role is not yet well-efined. Its

presence apparently mitigates the degree of emnrittlement caused by the alpha-prime phase by slowing

down its formation. On the other hand, the G phase

(and perhaps the alpha-prime phase) precipitates more

readily at dislocations (Chopra and Chung, 1987).

Thus, cold work or stressing during service, including

fatigue cycling, may increase the degree of embrittlement; no data on this potential effect were found in the

present projecL

Because only the ferrite phase is embrittledby longterm service atLWR operating temperatures, the overall thermal embrittlement of a cast stainless steel

product depends on the amount and morphology of the

ferrite presenL For LWR applications, a current guideline is that low-temperature embrittlement is a major

concerm only when the volume fraction of ferrite exceeds a level of approximately 15% (Copeland and

Giannuzzi, 1984). The reasoning behind this guideline

is that the ferrite phase tends to form in isolated pools

contained within the austenite when ferrite levels are

less than or equal to 15%. In this case, the overall

toughness of the stainless steel casting is not greatly

affected even if the ferrite were embrittleA However,

where ferrite levels are greater than 15%, there is a

greater tendency for a continuous path of embrittled

material to exist through the thickness of the cast component, which would greatly reduce its toughness if the

ferrite regions were embrittled (Chopra and Chung,

1986b).

There is recent evidence that thick-walled (typically

greater than 100 mm) stainless steel castings with ferrite levels in the range of 10 to 15% also may

be subject to low-temperature embrittlement (Chopra

and Chung, 1986b). In heavy-section castings, the

grain size tends to be large and the ferrite spacing

(average distance between ferrite islands) is increased.

With increasing ferrite spacing at a constant ferrite

content, the tendency for a continuous path of ferrite to

exist through the thickness of the cast component increases. Bonnet et al. (1990) selectively dissolved the

austentite phase from samples of CF-8M and found

that the ferrite phase remained continuous at ferrite

volume fractions as low as 5%. This result indicates

that a relatively small amount of ferrite may provide a

continuous path for fracture in aged cast stainless

steels.

Two different sets of data have recently been

presented regarding the thermal embrittlement of

stainless steel weldments. One data set indicates that

the fracture toughness of Type 308 SS weldments was

not affected by aging at 4270C (800F) for 10,000 h

(Mills, 1987). However, the other data set indicates

that the Charpy V-notch impact energy of Type 308 SS

weldments with 8% ferrite may be significantly reduced by aging at 3430C for 10,000 h: however, corresponding fracture toughness data are not presented.

More research is needed to ensure that the fracture

toughness of the weldments for cast stainless steel

components is not significantly affected by thermal

aging at the LWR operating temperatures.

Embrittlement of welds, ifconfirmedby tresuts,

could become a major problem because welds generally have lower initial fracture toughness before thermal exposure than base metal. Weldments in stainless

steel castings may contain high residual stresses that

approach the yield strength (Egan et aL. 1987) because

Section HI of the ASME Boiler and Pressure Vessel

Code neither requires nor prohibits postweld heat

treatment of these welds. The usual industry practice is

to not subject welds in cast stainless steel components

to postweld heat treatments, because such a heat treatment might sensitize base material near the end of the

heat-affected zone and make that region susceptible to

IGSCC.

2.2 Charpy Impact Energy Data

The degree of low-temperature embrittlement is

most often quantified by the room-temperature

Charpy impact energy after aging at temperatures in

the range of 300 to 4500C (570 to 840MF). Increased

aging temperature is employed to accelerate the rate of

thermal embrittlement compared with that which

occurs at normal LWR operating temperatures near

2880C (550). Charpy V-notch impact energy data,

obtained from Chopra (1990), for thermally aged

CF-3, CF-8 and CF-SM stainless steel are shown in

Figures 5, 6, and 7, respectively. In each figure, the

Charpy impact energy is plotted as a function of the

aging parameter P. which is defined by the following

relation:

P - logo (t/cp[(Q/R)(1/T - 1/673)]) (1)

where t is the time, Q is the activation energy, R is the

gas constant, and T is the absolute temperature P is defined such that it is equivalent to the logarithm (base

10) of the number of hours of aging at 4000C (750F).

For example, a value of P 4 is equivalent to aging for

10,000 h at 4000C.

12

I
104

cg

>, 10

en

a)

c

w

E

102

0 1 2 3 4 5

Aging Parameter (P)

Flgure 5. Effectofd ienalaging on theroom-temnperathreChaipy y-notch impactenergyofCF-3 stainless steel.

104

UJ

cqa10S

0102

0 1 2 . - 3 4 5

Aging Parameter (P)

Figure6. Effectofthermal agingontheroom-tperatureCharpyV-notchimpactenergyofCF-- stainless steel.

13
I

104

E

(:3

C

w

U

coa. E

103

logjO(CVN) - 3.412 - .245 P. P S 4 _

log1O(CVN) - 2.432, P > 4

0 _ 0____

00 ; * ; 0 a 18 __. ._

__. __ . _ ...._.._..... E=-r._ ...

_.=_::-;- .-..-. -0--------

. __..._.___..__...................... .,._ _ ._ ..__.. ............

* CF-BM Small Experimental Heats

o CF-8M Commercial Heat

& CF-8M Large Experimental Heats

_ CF_8M Lower Bound

-CF-8M Lower BoundI__ ______ 102

0 I 2 3 4 5

Aging Parameter (P)

Figure 7. Effect of thermal aging on the room-temperature Charpy V-notch impact energy of CF-8M stainless

steel.

From multiple correlation analysis of data on the

aging of cast stainless steels, the activation energy was

found to be related to composition by the following

approximate relation:'

Q (ki/mole) = - 182.6 + 19.9(%Si)

+ 11.08(%Cr) + 14.4(%Mo) (2)

where the percentages are in terms of weight.

Equations (1) and (2) were used by Chopra and Chung

in their earlier work.

Recently, Chopra and Chur., (1989a and 1989b)

have developed new correlations for predicting the

activation energy. They found the following separate

correlations for the Swiss data (ftautwein and Gysel,

1982) and the Framatomea-ANL (Chopra and Chung,

1989a and 1989b) data, respectively.

a. Unpublished paper presented by G. Slama ct aL,

at SMiRT Posconference Seminar 6 at Monterey,

California, August 29-30,1983.

Q(kJ/mole) = -66.70 + 6.91(%Cr) - 5.44(%Ni)

+ 8.08(%Mo) + 17.2(%Si)

+ 44.1(%Mn) + 297(%N)

(Swiss data) (3)

Q(kJ/mole) = 90.60 + 9.63(%Cr) - 8.12(%Ni)

- 7.54(%Mo) + 20.6(%Si)

- 123(%Mn) + 318(%N)

(Framatome - ANL data) (4)

Compared with Equation (2), Equations (3) and (4) include additional terms for nickel, manganese, and nitrogen as well as those for chromium, molybdenum,

and silicon. The contribution of molybdenum is positive in Equations (2) and (3) but negative in Equation (4). Also, the contribution of manganese is

positive in Equation (3) but negative in Equation (4).

Chopra and Chung (1989a and 1989b) argue that

Equation (4) is the most appropriate to use in assessing

14
the potential for thermal embrittlement of cast stainless steel components that are used in U.S. nuclear

plants because the compositions included in the

Framatome-ANL set of data are more representative

of those used in such components. For this reason,

Equation (4) was used, along with Equation (1), in

computing the values of P plotted in Figures 5

through 7. The values of Q and P were calculated

from the data supplied by Chopra and Sather (1990).

Although Equation (4) provides an improvement over

Equation (2), there is still room for improvement in

this analytical approach, as is evidenced by the wide

scatter of the data in Figures 5 through 7. Furthermore, Chopra and Sather (1990) reported that the correlation represented by Equation (4) was developed

for observed values of Q between 75 and 240 kJAnote.

Thus, it should be used only for cases where the computed value of Q falls within this range.

The compositions included in the Swiss set of data

fall in a relatively narrow range that does not represent

those used in cast stainless steel castings produced in

the U.S. Equation (1) and either Equation (2), (3). or

(4) should be used to estimate the degree of aging of a

specific heat of material during service. However,

because the reasons for the differences among Equations (2), (3), and (4) are not completely understood,

additional research concerning the factors contributing

to the activation energy for thermal aging of cast stainless steels is needed.

The recent work by Bonnet et aL (1990) even qucstions the validity of using an empirical value of Q calculated from the chemical composition. They believe

that Q is a function of the aging temperature and approaches a value of approximately 250 kfinole at

300 to 325°C, the region of primary interest for LWR

applications. If this belief is trues it implies that values

of Q calculated using Equation (4) provide overly pessimistic predictions of the aging kinetics for cast

stainless steels. Unfortunately, their conclusion is

based on the evaluation of data for only two beats of

CF-8M stainless steeL Additional, extensive evaluamons would be required before their approach could be

applied to the assessment of aging in LWR cast stainless steel components

Figures 5 through 7 show lowerbound approximations to the data trends for each alloy. These approximations were developed as part of the work in the

curentprqect. The impact energy decreases as a function of P for values of P less than about 3.5 to 4. For

values of P greater than about 3.5 to 4, the impact

energy appears to show no further decrease as a function of P. In facts it even shows a slight increase. Tbus,

the effect of aging on Charpy-impact energy appears

to satuate at long times. As a conservative approximation, the lower-bound CVN-vs.-P lines were assumed

to saturate at the CVN value where P was equal to 4 (as

shown in Figures 5 through 7).

All of the CVN data for CF-3M fell above the

lower-bound approximation, as is shown in Figure S.

Some of the data for the KRB plate' fell below the

lower4ound approximation for CF-8 (see diamond

symbols in Figure 6); these data were discounted in developing the lower-bound approximation because they

are from European material rather than from U.S. material. One of the data points for the commercial heats

(see square symbol in Figure 6) fell slightly below the

lower-bound approximation because it was considered

to be an outlier. Also, for the same reason, one of the

data points for the large experimental heats (see triangular symbols in Figure 7) fell below the lowr-bound

approximation for CF-SM The lower-bound approximations for CF-8 and CF- M were almost the same,

whereas that for CF-3 was higher than the other two.

This result, as is discussed later in this rport, implies

that the maximum degree of thermal embrittlement is

controlled by factors other than just the chemical

composition.

These lower bounds are appropriate for LWR applications because operating temperatures less than

400WC are expected and the degree of embrittlement at

these lower temperatures Is not expected to exceed that

observed at 400C. It would be useful, in future work,

to verify this assumption by evaluating the impact behavior of materials that have been exposed to temperaturesof 300rC (5700 ) or less for many years. Because

a greater degree of embrittlement is expected from exposures at 4000C (7500F) than from exposures at

350°C or less, using lower bounds to all of the data

should provide conservative, but overly pessimistic,

predictions of thermal embrittlement. Additional thermal embrittlement data for the values of aging parameterP greater than4 are needed to verify that the lower

bound curve levels off, especially at temperatures of

350C or less.

The failur of P to fully correlate the changes in

impact energy as a function of aging time and

temperature indicates that the simple expression of

Equation (4) does not accurately model the activation

energy for the complex process of thermal ermbrittletnent, which involves the formation of carbide, alphaprime, and 0-phase precipitates and their interactions

during aging at temperatures in the range of 300 to

450°C (570 to 8405;) (Chopra and Chung, 1986c).

a. From KRB reactor at Gundremmingen, West

Germany.

15
However, until a more accurate model can be developedq the lower bounds shown in Figures 5 through 7

provide guidance for LWR applications and indicate

an approach that should be developed further.

Most of the data that have been developed to asses

the effects of thermal embrittlement of cast stainless

steels are room-temperature impact properties, such as

those shown in Figures 5 through 7. For application to

LWR components, properties at operatng temperature

are needed Also, fracture toughness and tearing modulus data, rather than Impact energy data, are needed

in fracture-mechanics assessments of the safety of

components. Both of these issues are being addressed

in recent work (Landerman and Bamford, 1978;

Buchalet et al., 1985; Chopra and Chung, 1989b;

Hiser, 1987: Bamford et aL, 1987).

Figures 8, 9, and 10 show Charpy V-notch impact

data for CF-3, CF-8, and CF-8M, respectively, as a

function of test temperature. Except for Figure Sc, the

open circles represent data for unaged material, the

filled triangles represent data for material aged approximately 10,000 hours at 3500C, and the filled

squares represent data for material aged approximately

10,000 hours at 400C. In Figure Sc, the ciles represent data for unaged material, the triangles represent

data for material aged 9,980 hours at 350CC, and the

squares represent data for material aged 30,000 hours

at 350C (Figure &c contains no data for material aged

at4000C). TMe filled symbols in Figure 8c are for spec- imens with a L-C orientation (longitudinal specimen

with circumferential crack), and the open symbols in

Figure Be are for specimens with a C-L orientation

(circumferential specimen with longitudinal crack).

Aging typically reduced the impact resistance of the

materials at all test temperatures, except for the centifugally cast pipe (see Figure 8c), where the aged material had the same upper-shelf impact resistance as

the unaged material. The centrifugally cast pipe also

had significantly higher impact resistance than the

statically cast materials, and within the normnd scatter

of the data, the specimen orientation aw-C versus C-L)

had no sigpificant effect on its impact resistance. The

impact resistance of the materials aged at 400C was

always slightly less than that of the materials aged at

350CC. For the aged material the room-temperature

(25°C) data always provided a conservative estimate

of the impact resistance at typical LWR operating temperatures (up to approximately 290C. Thus, when no

high-teupeature Charpy data are available, the lower

bound to the room-temperature Charpy data is likely

to provide a conservative estimate of the lower bound

to the high-temperature Charpy data.

5000

4000

Cd5

2.

0

LU

Cu

0.

3000

2000

1000

0

-200 -100 0 100 200 300

Test Temperature, OC

a. Cast76-mm siab,Heat69.

Figure 8. Effect of test temperature and thermal aging on the Charpy V-notch impact energy of CF-3 stainless

steel, data from Chopra and Sather (1990).

16
3000

2500

0 Unaged * Aged 9,980 hr at 350°C

Unaged (shroud material) a Aged 9,980 hr at 4000C

> _o 6

E

U

0.

E

2000

1500

1000

.U.

Heat I of CF-3

17.1% ferrite

500

0

-200 -100 0 100

Test Temperature. °C

200 300

b. Cast pump impeller vanes, Heat 1.

8000

7000

Heat P2 ol CF-3

15.6% ferrite

Aged at 3500C

I * Unaged, L-C Orientation

o Unaged. CL Orientation

A Aged 9,980 hr, L-C Orientation

a Aged 9,980 hr. C-L Orientation

a Aged 30,000 hr. L-C Orientation

b Aged 30,000 hr. C-L Orientation

cml

'U

0.

6000

0

5000 0

4000

3000

2000

A LY.

an o0

* P. I 0

1000 a

I a I nL 1.I

--200 -100 0 100

Test Temperature, 'C

200 300

c. Centifugy cast pipe, Heat P2.

Figure 8. Effect of test temperture and thermal aging on te Charpy V-notch Impact energy of CF-3 stainless

stel, data from Chopra and Sather (1990). (Continued.)

17
5000

4000

(UE

0

Cu

0.

E

3000

2000

1000

-200 -100 0 100 200

Test Temperature. OC

300

Cast 76-mm slab.

Figure 9. Effect of test temperatre and tiermal aging on the Charpy V-notch impact energy of CF-8 stainless

steel, Heat 68 data from Chopla (1990).

5000

4000

E

0' z

Zu

0.

E

3000

2000

1000

0

-200 -100 0 100 200 300

Test Temperature, °C

a. Cast 76-mn slab, Heat 70.

FIgur 10. Effec of test tempeate and thermal aging on t ChaVpy V-ntch impact engy of C78hM sinless

sted data from Chopra and Sather (1990).

18

I
(.4E

C

A

(I

a.

E

5000 -

4000

3000

2000

1000

0

- 200 -100 0 . 100 200

Test Temperature, C0 -

300

b. Cast 76-mm slab, Heat 74.

5000

4000

C,E

8-

0

C

Ca

a. E

3000

2000

1000

0 .e=

-200 -100 0 100 . 200 300

Test Temperature. OC

c. Cast 76-mm slab, Heat 75.

Figure 10. Effectof test temperature and thermal aging on the Charpy V-notch impact energy of CF-8M stainless

stel data from Chopra and Sather (1990). (Continued.)

19
I

The curves shown in Figures 8 through 10 are best

fits to the data obtained by means of locally weighted

least-squares regression, using weighting factors

(degree of data smoothing) of 66 to 75 percent There

are no explicit expressions for these curves; they were

obtained to graphically show the trends of the data

rather than to provide mathematical expressions for

computing CVN values In general, the transition region of these curves shifted to higher temperatures and

the upper-shelf energy (USE) decreased with increased aging temperature r timem In many cases, the

USE reached a peak at some temperature less than

300°C and decreased at temperatures above that peak.

In contrast to the preceding curve fitting approach,

Chopra and Chung (1990a) used a hyperbolic tangent

function of the following form to describe Charpy

transition curves

CVN - Ko + B{l + tanh[cr-C)/D4)

where Ko is the lower-shelf energy, T is the test temperature, B is the half-distance between the upperand lower-shelf energy, C is the mid-shelf transition

temperature in °C, and D is the half-width of the transition region. This approach has the advantage of providing an explicit description of the Charpy transition

curves, but it does not completely model the observed

trends of tie experimental result. For this reason, the

approach was not employed in the current study.

In their recent work Chopra and Chung (1989a and

1989b) have proposed the following parameter for correlating the minimum Charpy V-notch impact energy

after long-term aging with the composition and microstructure of cast stainless steel,

= [%C + 0.4(%N)1 (%CR + %Mo

+ %Si) L (10 (5)

where 8m is the measured volume fiaction of ferrite in

percent and L is the mean ferrite spacing in pm. Their

correlation between the parameter 0 and the minimum room-temperature Charpy V-notch impact

energy is shown in Figure 11. The power-law equation

shown inFigure 11,CVN - 6,270 0<- ,wasobtained by least-squares regression analysis of the data

in that figure as part of the work in the current study.

Except for the Framatome data point (filled triangle),

the correlation appears to be reasonable. Additional

work should be performed to evaluate the correlation

of this parameter with the minimum high-temperature

ChaWyV-otchimpactenegydataandwith fracturotoughness data at both room and elevated temperatures. The 0 parameter could be used, with either

(a) measured values of the ferite fraction and spacing

or (b) a calculated value of the ferrite fraction and an

estimated upper-bound (worstcase) value of the ferrite spacing, to estimate the minimum roomtemperature Charpy impact energy expected after

long-tem aging of a cast stainless steeL In either case,

it would be necessary to know the composition of the

stainless steel, including the nitrogen content

Chopra and Chung (1990a) have recommended that

a worst-case value for the ferrite spacing of 180 pim be

used when a measured value is not available. For example, for a CF-8 stainless steel with 0.05%C,

0.07%N, 20.6%Cr, 8.3%Ni, 0.3%Mo, 1.1%Si and a

measured ferrite content of 15%, the value of O is

computed to be 69.5 for a ferrite spacing of

180um. Then, the equation shown in Figure 11 gives

CVN = 6,270 IV" = 6,270 (69.5)Y 2

= 469 klY/mr

for the minimum room-temperature Charpy-impact

energy expected after long-term aging.

If it were decided that the CVN value based on the

wors-case ferrite spacing were lower than desred, the

actual ferrite spacing could be measured and used in

the cowrelation. For example, if the ferite spacing for

this CF-8 staidess steel were measured to be 100pm,

the value of 0 would be 38.6, giving

CVN = 6,270 -` = 6,270 (38 6)0*111

= 672 klJme

for the minimum room-temperature Charpy-impact

energy expected after lng-tm aging.

This 10-parameter approach provides a simple,

straightforward method of estimating the minimum

room-temperature Charpy-impact energy expected

for a cast stainless steel after long-term service.

Chopra and Chung (1990a) also have shown that the

0 parameter can be used to provide improved correlations for aging kinetics. They found improved corelations between impact energy and the aging parameter

P when the materials were grouped by the value of O

rather than by alloy type, as in Figures 5,6, and 7. This

approach should be evaluated further.

20

I
04 2,000

i. 1,500

P

V 1,000

a

j 500

E

0

- - - - - - - - - - - - --- - - - - - - - *A T l nll I 1 16 11 1& lin ii I

o CF-3, ANL a CF-SM, GF

o CF-8, ANL A CF-8M, FRA

* CF-SM, ANL * CF-8, EPRI _

a w CF-3, GF & CF-3. CEGB

r CF-B, GF -All Data _

0< 9 CVN . 6,270 40.6112

III I u,. 11 II uI, IsI II gI IIl I I I ------------------ 0 50 l00. 150 200 250 300

Material Parameter (0) 8mm 2 (C + 0.4O(Cr + Mo + SQ(0.001) L

Figure 11. Correlation between minimum room-temperature Charpy V-notch impact energy and material

parameter 0 for aged cast stainless steels, adapted from Chopra and Chung (1989b), data supplied by Chopra in

June 1990.

2.3 Fracture Toughness Data

The fracture toughness of cast stainless steels has

been evaluated by means of J-R curve tests (Bamford

et al., 1985; Chopra and Chung, 1987; Hiser, 1987).'

Most of the earlier testing (Bamford et al., 1985). was

performed at room temperature, but some of the recent

testing (Chopra and Chung, 1987; Hiser, 1987) has

been performed at 290°C (555'F). Those invesdgations included analyses of test rsults to determine values of JIC. which is a measure of fracture toughness at

the initiaton of crack growth in metallic materials, and

tearing modulus, T, which is proportional to the derivative of facture tougness I with respect to crack size,

ascoveredby the testproceduresdescribedinASTM

Designadon E 813.

Figures 12 through 15 show fractutoughness data

correlations for cast stainless steels, as taken from the

work of Chopra and Chung (1990a). The data shown in

these figures are from Argonne National Laboratory

(ANL) (Chopra and Chung, 1989a and 1989b),

Westinghouse (WH) (Landerman and Bamford, 1978;

Bamford et al., 1985), Framatomea (FRA), Electric

Power Research nstitute (PRI), Battelle (BA!), and

KRB. Figure 12 shows the correlation between rooma. Unpublished paper presented by G. Slama et al.

at SMiRT Posconfrece Seminar 6 at Monterey,

California, August 29-30,1983.

temperature fracture toughness, JIl and roomtemperature impact energy. Ihe solid curve represent

the mean trend of the data, whereas the chain-dashed

line represents the lower bound to the data In performing a structural integrity evaluation, the lower-bound

'relationship can be used to estimate the roomtemperature fracture toughness from an estimated or

measured value of the room-temperature impact

energy for a specific component. Then, the estimated

room-mperature fractue toughness can be used to

estimate a lower-bound value of the roomtemperature tearing modulus using the dashed line in

Figure 13. As is shown in Figure 14, the lowerbound fracture toughness at room-temperure correlation from Figure 12 also provides a conservative

estimate of the fracture toughness at 2900C from a

value of room-temperature Charpy impact energy.

Finally, the tearing modulus at 290OC can be estimated

from the lower-bound correlaon (dashed line) shown

in Figure 15. If fracture toughness could be obtained

for aged material from the component being evaluated

in a structual integrity assessment, it could be used in

place of values estimated using the prding scheme.

-he value of fiacture toughness at a temperature between the room temperature and 290°C can be estimated by linearly interpolating between the data

presented in Figures 12 and 14; similarly, the value of

tearing modulus can be obtained from the data

presented in Figures 13 and 15. The values of fracture

21
I

IMPACT ENERGY. KCV (ft.Ib) 100 150 204

1.Y6

0

z

2

a

z

tC

Ul

IL

.Sa0

1000 1500 2000 2500 3000

IMPACT ENERGY, KCV (kjIm2 )

FIgure 12. Coelation betwen room-tempete fractur toughness (JIc) and impact energy for cast stainless

stee&

DEFORMATION, JI 1 (In.-Ib/In.2 )

0 2000 4000 6000 s000 10000

500 - .00

TESTED AT ROOM TEMP.

700- AGING TEMP. (C) 700

3 284 SYMBOLS9 4

OPEN: CF-S , Cf-3 a 350 000- CLOSED: CF-Sm 0 400 800

0 0 427

S OO- 0 0450 500

V o 0 UNAGED

cc 0 w 400_ 0 V RCANNEALED 400

>0 00

300 o02 0 7300

I-. o0 0 Z~*-

200 00 . o,

100- ° __00 1 

0 500 1000 1500 2000

DEFORMATION JIC (kJ/m2 )

Figure 13 Cowelation between room-tempeature tearing modulus (T) and fractue toughness (ic) for cast

stainless steels.

22

I
0

.U-_ I

RT IMPACT ENERGY, KCV (fttlb) Go 100 160 200 I . . . .I . . . . . . . . . 250 ZIIDN n - - * . . . . . . . . . . . . . .

'4

10

a

04

0Z

I0

----

1500-

1000-

.:o

CAST STAINLESS STEL

SYMBOLS

OPEN: CF-8 , CF-3

CLOSED: CF-SM

AGING

TEMP.(8C) ANL WH FRA

25 0 O

350 A

4000 o V

427-, , 0

1450 01

284

EPRI KRB

m

-'Gao.

*aoooo7a.0C

0.

.0

. C!.

eo000 C

a. 04

'I?

-- a

.I 0

.-A00

,-sLOWER BOUND RT Jic

i

r

-2000

........................ ,......... . U. -

0. . O 1000 -1S 0 20000. 2500 3000 3500 4000

RT IMPACT ENERGY, KCV (kjhn2 )

FIgure 14. Corrlaton between the 2900C fracture toughness (JI0 and room-temperature impact energy for cast

stainless swls.

-1 I DEFORMATION, J1c (In.-lb/ln.2) I000 4000 0ooD o000 I1 a I a

1 - -- I 10000

HnD S .

0

cc

tu

CS

I700-

800- Gooaoo

600

400

3000

200

100

TESTED AT 290-320°C

SYMBOLS

OPEN: CF-8 , CF-3

CLOSED: CF-GM

0

6

A 0

a0 e

*O. '

. 0 0 0

0 0 -a O

L A - A - -

0 *

, we

- AGING TEMP. ('C)

E 284

A 350

0 400

0 427

0 450

0 UNAGED

-800

-700

-600

-500

-400

-300

-200

-100

- - - - - u - -

o l .50 1ooo 0 200

DEFORMATION J1c (kJ/m2 )

Figure 15. Coelation between fte 290 0C ering modlus (1) and frature toughness (jc) for cast stainless

stels.

23
I

toughness and tearing modulus at tempeatures lower

than 290 0C are needed for the pressurized thermal

shock analysis of cast stainless steel components

The coteations shown in Figures 12 and 14 may

not be valid for higher values of lIc. because it is lkely

that many of the indicated JIG values do not satisfy the

validity criteria of ASTM B 813, which requires that

the thickness, B, and the uncrace ligament, b, of the

test specimen be greater than 25 JQ*. JQ is the conditional value of fracture toughness, and sf is the flow

stress, which Is the average of the ultimate tensile

strength and het 0.2% offset tensile yield strength If

these size conditions, plus the regression-line slope

condition (see Paragraph 9.2 of ASTM B 813), are

satisfied, then JQ equals JI_

Typical tensile properties (Chopra and Chung,

1987) of cast stainless steels that were aged for 9,980 h

at 350 0C (660 0F) and tested at room tempature and

at 288 0C (550 0F) yield sf= 440 MPa at room temperatu (Ri) and 300 MPa at 288-C The data shown in

Figures 12 and 14 were developed from tests of IT

compact-type specimens, for which typically B = b =

25.4 mm. From this information, it is possible to estimate the maximum value of JQ for which valid lie

values can be obtained, as follows:

JQ(maximum) = 440 x 25.4/25

= 447 kJ/m at RT (6)

and

JQ(maximum) = 300 x 25.4/25

= 305 kJ/m at 288C . (7)

Jjc values in Figure 12 and 14 that are greater than

447 and 305 WO 2 atRT and 288oC (550o , respecc

tively, may not sat the validity criteria of ASTM B

813. Additional work should be performed to evaluate

the validity of the JIC values plotted in Figures 12

and 14.

Based on fracture-mechanics evaluations of their

cast stainless steel components, Framatome recommends that a lower-bound value of Jk equal to

100 kihn 2 be used for assessing those component

This lower-bound value Is specific to the cases analyzed by Framatome. Comparable lower-bound JIC

values would have to be developed for other specific

cases as they ar evaluated, but it is instructive to use

this value as an example in th present discussion. For

the lower bound in Figure 12, a Jl, value of 100 kJ/m2

corresponds to a Charpy V-notch impact energy of

about 735 ktdm 2 at room temperature Only a few of

the experimental data reviewed in this report fan below the Framatome lower-bound value. However, for

the long exposure times (greater than 40 years) associated with plant-license renewal, there is a concern

that more materials may embrittle to the degree that

their properties fall below this lower bound or that

there may be specific cases where a more conservative

lower-bound value of impact toughness is necessary.

This needs to be determined by case-specific structural integity assessments

2A Tensile and Fatigue

Properties

In addition to fracture-toughness properties, the

tensile, fatigue-crack initiation, and fatigue-ack

propagation properties of cast stainless steels are

important for use in engineering evaluations of fitness

for service. Typically, thermal aging increases the

tensile ultimate and yield strengths and decreases the

tensile ductility (Landerman and Bamford, 1978; and

Chopera and Chung, 1987).a However, Chopra and

Chung (1987) also report that a 12-year

reactor-service exposure slightly decreased the tensile

strength of a CF-8 pump cover plate. Overall, thermal

aging does cause changes in tensile properties, but

such changes do not appear to have a significant effect

on the design-allowable strengths of cast stainless

steels.

Landerman and Bamford (1978) reviewed the lowcycle fatiguerack-gwth properties of cast stainless

steels. They found that the CF-8M stainless steel exhibited about the same fatiguo-ack-growth resistance as comparable wrought alloys, and that aging at

427 0C for 3000 h exhibited no significant effect on the

fatigue-cacwk-growth resisac in either air or simulated PWR water environments. However, these results are only for aged CF-8M stainless steeL Many

cast staless steel components, including reactor coolant piping, are made of CF-8 and CF-8A stainless

steels, and the fatigue-crack growth rates for these materials in a PWR environment may be significantly

a. Unpublished paper presented by 0. Slama et al.

at SMIRT Postconference Seminar 6 at Monterey,

California, August 29-30,1983.

24
different than that for CF-SM material, depending

upon the carbon content. One possible explanation for

this potentially different behavior is that chromium

carbides may form at the ferrite-austenite interface in

CF-8 and CF-8A materials at operating temperatures,

whereas the presence of molybdenum prevents such

formaton in CF-SM stainless steeLa This explanation

needs to be validated. If coirsion figue is a probable

degradation mechanism for CF-S and CF-8A stainless

a. Private communication with a M. Chung,

Argonne National Laboratory, April, 1990.

steels, accelerated aging at higher-than-operating

temperatures would not be valid.

Slama et al.P evaluated both the low-cycle fatigue

crack-initiation behavior and fatigue-crack-growth

resistance of cast stainless steels at both room

temperature and 3200 C (6100F). They found that

b. Unpublished paper presented by 0. Slama et al.

at SMiRT Postconference Seminar 6 at Monterey,

California, August 29-30,1983.

10-2

daldN = 7.5x 10- 10 AK4

4 I I

J£

I I

(Upper bound design

curve for stainless steels)

1O-3 I-

,/

&~ 4

V ' * *

A'a

& a .44 k a & a

I

*A /

//

I * a

/

/

//

/ /

/ /

/

7/ /

' / /

E

z

x

J/

10-4 F-

'a & &, & /

A'' 4b //

a / /

A /

/ O 0° 78 A6

/

1'

I/I a78 B4

C Air I320 I Virgin

I PWR condition

10-5 / /o

/ 0

/0

/

*78 B7 320 *C Alr I Aged 7500 h

& 78 B12 PWR I at 400 *C

== Scatter Band

10-6K I I i i I I I f

20 30 40 50 60 70 80 90

.1K (MPavffi)

Figure 16. Fatigue crack growth rate of CF-SM stainless steel, in air and in PWR environment, at a cyclic frequency of 0.017 Hz (see footnote b).

25
I

thernma aging caused a small (10 to 15%) increase in

cycli yield strength, and, although one might expect

reduced tensile ductility to reduce low-cycle fatigue

resistance, they found no significant effect of thermal

aging on low-cycle fatigue-crack-initiation resistance. Fatigue-crack-growth-rate resistance in air

also was not significantly affected by thermal aging.

However, as is illustrated In Figure 16, they found that

fatigue-crak-growth rates for aged CF-8M stainless

steel increased by as much as a factor of 10 when

tested In a simulated PWR water environment, compared to testing in aa at 3200C (610P, but did not

exceed the upper-bound design fatigue-crackgrowth-ate curve for stainless steels.

Additional data for the fatigue crack iniation and

growth in CF-8, CF-8A, and CFL8M stainless steel

materials in LWR environments are needed. Data for

high-cycle fatigue behavior of CFE8M material are

needed because some cast stainless steel surge lines

and elbows are subject to thermal stratificatioL. Data

for the guebehaviorof Irmdiatedcaststainless steel

materials are needed because some reactor internals

are made of these materials No data are presently

available for fatigue behavior of cast stainless steels in

a BWR environment. The corrosion-ague behavior

of aged cast stainless steel in a BWVR environment

needs to be evaluate

In summary, therma aging of cast stainless steels

has been found to produce some increase in tensile

strength and some degradation in known fatigue properties, but it may cause a significant degradation in

fracture toughness. This degradation could Impair the

structural integrity of cast stainless steel components

during long service. Thus, the fracture toughness and

impact energy of thermly embrittled material are the

most important items in evaluating the long-term

structural integrity of cast stainless steel components (Bamford et al., 1987). In addition, degraded fatigue

and stress-corrosion-cracking resistance, if any,

should be taken into account in evaluating the integrity

of cast stainless steel components.

26

I
3. INSERVICE INSPECTION OF COMPONENTS

Nondestrutive examination (NDE) procedures are

used to detect and characterize the nature of defects in

cast stainless steel components. Application ofNDE to

cast stainless steel components is complicatedbecause

the material microstructure often contains a coarse,

variable, and unpredictable combination of columnar

and equiaxed grains (Egan et al., 1987; EPRI, 1986;

Avioli, Jr., 1986; Gregory et al., 1986; Umino and Rao,

1984). Ultrasonic testing U is preferred for the volumetric inservice inspection of many LWR components. However, conventional UT is not sufficiently

reliable for cast stainless steel because coarse grains

cause high attenuation and scattering that result in a

low signal-to-noise ratio and make defect detection

difficult. The noise level in cast stainless steel is significantly higher than that in wrought stainless steel. In

addition, different grain structures in cast stainless

steel introduce elastic anisotropy, which causes velocity variation and beam skewing and consequently inaccurate detection of the defect. Cast stainless steel is

elastically isotopic if composed of equiaxed grains and

the resulting variation in wave velocity with its propagation direction is small (2%), whereas cast stainless

steel is elastically anisotopic if composed of columnar

grains and the resulting variation in velocity with its

propagation direction is large (as much as 100% for

shear waves) (Kupperman et al.,1987). For these reasons, radiography normally is used for the NDE of cast

stainless steel components, both during fabrication and

during Inservice inspection, especially for thick cast

sections such as those used in pump bodies.

Radiography indicates the presence of a flaw, but,

used alone, it cannot determine the depth of the flaw or

whether the flaw is connected to the surface. For this

reason, radiographic indications must be assumed to

be a flaw of the most detrimental shape and location

(Umino and Rao, 1984). Fracture-mechanics analysis

indicates that the worst-case flaw will be a surface

crack of 0.5 aspect (depth/length) ratio. Additional

penetrant examinations can be performed to determine

if the indication is actually a surface flaw. If the indication can be shown to be an imbedded flaw, then the

stress intensity factor caused by such an imbedded

flaw can be used in the fraure-mechanics analysis.

Use of radiography during inservice inspection is

less practical and is less efficient than during fabrication. In some cases, structural restraints do ao allow

sufficient access for radiography. Therefore, both the

NRCa and the nuclear industry (especially EPRI)

(Avioli, Jr., 1986; Gregory et al., 1986; EPRI, 1986)

have initiated research programs for the development

of advanced UT methods to inspect cast stainless steel

components. Advanced methods use low-frequency

(longer wave length) transducer and short pulses,

which enhance the signal-to-noise ratio and Improve

the capability to detect flaws. Two digital signal processing techniques, time and spatial averaging, can

also provide a high degree of sensitivity and a high

signal-to-noise ratio (Ycong and Ammirato, 1989;

Shankar et al., 1988). The time averaging technique

improves the presentation of data and makes them

clearer and easier to interpret. The spatial averaging

technique, which is similar to the synthetic aperture focusing technique (SAFI), reduces grain noise indications and, thus, enhances the signal-to-noise ratio for

cracks larger than the average grain size. An additional

difficulty in the inspection of statically cast components, such as pump casings, comes from rough surfaces and Irregular surface contours, which cause poor

transducer coupling with the component. For reliable

examination, the surface condition needs to be improved by mechanical means such as grinding.

The effect of grain structure on the propagation of

ultrasonic waves in cast stainless steel material has

been studied to increase the accuracy of locating defects. Specially designed cast stainless steel specimens

with cracks were tested to experimentally determine

the effects of grain structure on wave propagation.

These results are used to identify the grain structure in

a cast stainless steel componern, which is in turn used

to compensate for the grain effect in UT examination

to locate defects. Ultrasonic examinations were used to

characterize grain structures in centrifugally cast stainless steel piping at the Vogde power plant during preserviceinspection.

Ultrasonic examinations were used so inspect welds

and statically cast components, such as pump casings

andelbows,atseveraldomesticPWRplants (eongand

Ammirato, 1989). The spatial averaging technique was

used in these inspections to enhance the signal-tonoise ratio. The base metal of the hot-leg elbow at the

Trojan plant was inspected because of a concern for

high stresses. The concern was caused by the fact that

several rupture constraints for the primary coolant

a. Private communication with M. S. Good,

Batele Pacific Northwest Laboratories, November

6, 1986.

27
piping were displaced from their original design positions ASME Section X inservice inspection requirementsfarcaststainesssteelcomponentdonotinclude

inspection of the base metal but include inspection of

welds. However, the statically cast components are

likely to have a larger number of fabrication defects

than h weldsw and therefore, may be more susceptible

tocmcking.ASME Section XI inspectionrequirements

for statically cast stainless steel components need to be

reevaluated.

Double-probo search units, also called twin-cystal

search units, were used to inspect PWR Type F pump

casings (200-mm thick) at the Tihange 1 plant in

Belgium. These units employ low-frequency

(0.5 MHz) longitudinally polarized waves for

enhancement of signal-to-noise ratio but are limited

to a depth range from 10 to 60 mm (Dombret et al,

1990). Large-diameter (140-mm) focused transducers

were used to examine down to a depth of 200 mm.

Advanced UT systems also show promise for the

NDE of centrifugally cast stainless steel piping, but

additional development and validation are needed

before they can be used routinely for inservice inspections. An EPRI study (Gregory et al., 1986), using

samples of centrifugally cast stainless steel pipe with

weldments, evaluated the accuracy of the advanced

UT methods. The signals appear to be satisfactory for

crack detection. However, because of the type of microstructure, cracks were not always detected or consistently sized when the same regions were inspected

from both sides of the pipe wall. Field studies at a

PWR plant (prior to service) have evaluated the feasibility of using advanced UT systems to inspect centrifugally cast staidess steel pipe (Avioli, Jr., 1986). The

results of that work indicate that advanced UT systems

can be used to make measurements successfully in the

field, but that additional development is needed to

make the scanners adaptable to all of the various piping configurations that may be encountered. Another

EPRI study has evaluated the use of Lamb waves, an

ultrasonic surface wave technique to detect large-scale

cracks (80% or greater through-wall cracks) in centrifugally cast piping (Georgetown University, 1988).

Lamb waves are not as severely affected by grain

structure as the conventional ultrasonic waves, and.

therefore, experience less attenuation and lower noise

levels. Development of a prototype search unit and a

crack sizing capability for this technique are needed so

that its use as an inservice inspection tool can be

evaluated.

Use of the acoustic emission technique to detect

crack growth in cast stainless steel components needs

to be evaluated because of the difficulties with the radiography and UT examinations to detect flaws Crack

growth through the embrittled ferrite phase releases

the elastically stored energy in the typical frequency

range of 0.1 to 0.5 MHz." The higher the degree of

thernal embrittlement, the higher the crack velocity

through the ferrite phase, and the amplitude of the resuIting acoustic emission signal will be larger If the

crack growth in the embrittled cast stainless steel is

through the ferrite phases, it will be composed of several small steps of brittle crack propagation and is most

likely to be detected by the acoustic emission technique. Coarse grains of cast stainless steel material

cause little scattering of the acoustic emission signal because the signal has a low frequency content (0.1 to

0.5 MHz); the corresponding wave length ranges from

30 to 6 mm, which exceeds the grain size.

The influence of thermal embrittlement of cast

stainless steel on acoustic emission signals has not

been investigated. However, the influence of another

embrittlement mechanism, i.e., hydrogen embrittlement, has been investigated. In hydrogen embrittled

material, the crack growth itself generates AE signals

with amplitudes higher than 40 db above electronic

noise. These signals are likely to be detected even

under plant operational conditions.

a. Private communication with

KWU Alliance, May 1990.

P. Jax, Bechtel/

28

I
4. ASSESSMENT OF DEGREE OF MATERIAL EMBRITTLEMENT

For license renewal, one needs to be able to estimate

the amount of thermal embrittlement that has occurred

in cast stainless steel LWR components during past

service and also the amount of embrittlement that is

expected to occur during the license renewal period.

The estimation techniques involve a combination of

(a) analytical modeling of inservice degradation, (b)

metallurgical evaluation methods, and (c) nondestructive testing. These three topics are addressed separately in the following subsections of this report.

Alloy bi _b2 b3

CF-8M 3.412 -0245 2.432

CF-8

CF-3

3.30 -0.220 2.42

3.310 -0.0954 2.928

.Substituting Equation (1)

following expression:

into Equation (8) gives the

4.1 Analytical Modeling of

Inservice Degradation

The time-temperature aging parameter P, given previously by Equation (1), is the basis for a proposed

model of inservice degradation caused by thermal embrittlement. The parameter, through Equation (2), is

based on bulk chemical composition, but it does not directly account for the effects of ferrite volume fraction

or ferrite spacing. However, because the ferrite volume

fraction is related to composition, P indirectly incorporates the effect of ferrite volume fraction on thermal

embrittlement. Also, as was pointed out previously, P

does not completely correlate the effects of aging time

and temperature with the measured changes in impact

energy. However, P provides the best model that is currently available. With appropriate conservative assumptions, P can be used to estimate the amount of

thermal embrittlement expected in LWR service, and

as improved models are developed, those can be incorporated into the approach suggested below.

Thlie lowcr4ond lines shown in Figures 5 through

7 are applicable to exposure temperatures of 4000C

(7500F or less and can be described by the following

relation:

logic (CVN) = bi + bk log10 I t/exp[(Q/R) (9a)

(lI/T - 1/673)]1 forP 5 4

logo (CVN) = b for P > 4 . (9b)

For a given alloy, chemical composition, exposure

time, and exposure temperature, a lower-bound or

minimum Charpy V-notch impact energy can be estimated using Equations (4) and (9).

Equation (9) can be used to estimate the range of potential CVN values as a function of exposure time at a

specific temperature for CF-SM, CF-8, and CF-3

stainless steels. As mentioned, Chopra and Sather

(1989) report that the correlation represented by Equation (4) was developed for observed values of Q

between 75 and 240 kn/mole. For the allowed compositions of these three stainless steels (see Table 3), all

three could have values of Q between 75 and 240kJ/

mole. Thus, for purposes of illustration, these were

assumed to be the minimum and maximum values of Q

for a material in service and were used with Equation (9) to estimate lower-bound room-temperature

Charpy values as a function of exposure time at 3000C

(570, as presented in Figure 17.

-log c 1(CVN) - b + be Pfor P 5 4 (8a)

and

1ogi 0 (CVN)= b for P > 4 (Sb)

where CVN is the Charpy V-notch Impact energy in

kJ/m2 ,PistheagingparameterofEquation(l), andbl,

b2, and b3 are material dependent constants with the

following values:

Figure 17 shows that CF-3 is expected to be the

least affected by aging, and that CFL8M and C- are

expected to be affected about the same and much more

than Cr-3. For the maximum Q of 240 kJhnole, the

saturation level (P=4) is not expectd to be reached

within 100 years. For the minimum Q of 75 kiAnole,

the saturation Level (P 4) is expected to be reached

within about 11.8 years. For the long-term exposure of

CF-8 or CF-8K the worst-case (minimum QJ roomtemperature CVN value is predicted to be about

260 Un/m2, whereas the corresponding value for CFL-3

is predicted to be high, equal to about 850 k1m 2.

29
I

-CF-8M Max. 0

-CF-SM Min. 0

- - -CF-S Max. 0 --- CF-3 Max. 0 -' -Framatome Limit

-- -CF-8 Min. a . -. CF-3 Min.

W~E

El 103

a.

I. 102

Maximum activation energy (Max. 0) equals 240 kh/mole and -

minimum activation energy (Min. 0) equals 75 Wi/moe.

Limit of r CVN 735 hU corresponds

to J__I- hJ/m 2(sea Figure 12)

.~~r-- I I I I -

1 10 100

Time at 3000C. years

Figure 17. Predicted lower-bound Chaipy-impact values for cast stainless steels as a function of aging time at

3000C (500MT).

The horizontal, chain-dashed line in Figure 17 at

735 kJ/W2 mcrespoids to a lower-bound JIC value of

100 kJ/n 2 mentioned by Slama et al of Framame for

possible use in safety assessment of casn stainless steel

components. Only the CF-3 alloy is predicted to satisfy this limit. Most of the CF-8 and CF-SM compositions are predicted to fall below this limit after 10 to

20 years of aging.

The wide spread between the maximum and minimum lines in Figure 17 further emphasizes the need

for research to better dtfine the value of Q at temperaures of 300°C or less discussed previously. If the value

of Q at these lower temperatures is close to 250 J/

mole, as was suggested by Bonnet et al. (1990), the

rate of embritlement would be expected to be close to

the lines for maximum Q in Figure 17.

Figure 17 could be used to estimate the degree of

embritulement of a cast material of unknown composition. However, it is recommended hat the actual composition be used to compute Q. Then, if the computed

Q value is between 75 and 240 kl/mole, a plot such as

that shown in Figure 17 can be prepared for that specific material and the tempeatut(s) of interest

4.2 Metallurgical Evaluation

Methods

Metallurgica evaluation methods cmrently provide

the best appmach for quantifying thermal embrittlement in cast stainless steel components that have been

in servic&e These methods can involve either nondestructive measurements and examinations or destructive tests and examinations of small samples that have

been removed from components. When samples are

taken from a component, they must be removed in a

manner such that there is no functional damage to the

component or such that the component can be acceptably repaired. Tbe metallurgical evaluation methods

either characterize the microstructure to generate data

for estimating the degree of embrittlement using degradation models or characterize mechanical properties

by testing subsized or miniature samples.

4.2.1 Mlcrostructure Characterization. Me

volume fraction of ferrite in cast stainless stel components needs to be determined because it is the primary

microstructural parameter controlling thermal embrittlement. If the ferite content is below 10%, the potential formatia degradation is lilkely to be low. For

ferrite levels between 10 and 15%, the potential for

material degradation is likely to be moderate.

30

I
Significant degradation may take place when the ferrite level exceeds 15%.

T = temperature of the final heat treatment in OC

The amount of ferrite can be estimated from the

measured bulk chemical composition of the material.

Procedures for estimating ferrite content have been reviewed by Leger (1982) and Aubrey et al (1982). Al

of thes procedures are based on calculating a chromium equivalent value, Crc, and a nickel equivalent

value, Nrk, and using these equivalent values to compute a volume percent ferrite content, %F. he empirical formulas for computing these values have been

developed from correlations of experimental data.

Aubrey et al. (1982) used linear regression analysis to

evaluate four different ferrite estimation formulas for

an extensive database of infarmation on CF-3, CF-8,

and CF-M stainless steels and found that the following relationship gave the best overall results:

Cr , %Or + %Mo + 0.65 (%Si) - 17.6

and

N c %Ni + 0.08 (%Mn) + 83 (%N)

+ 20(%C) - 5.18 .

This method is similar to the one defined by Equations (10), (11), and (12), except that it has different

values of the coefficients for the terms and includes an

additional term to account for the temperature of the final beat treatment. Bonnet et al. (1990) report an accuracy of +3% ferrite for their method, which is about

twice as good as the +6% ferrite accuracy cited by

Aubrey et al. for their method.

%F = 100.30 (Cr./W - 1772 (Cr./

+74.22

where

CR = %COr + 121(%Mo)+ 0.48(%Si)

-4.99

and

In many cases, the final beat treating temperature is

not known, and the method of Bonnet et al. cannot be

(10) p applied. When the final heat treating temperatures are known, the resulting correction factor,

[40014[150D -_, could be as large as 25%. For example, in the work of Aubrey et al the final heat treating

temperatures ranged fron 1093 to 11770

C (2000 to

215 V which gives corresponding values of the ternperature correction factor of 0.98 to 1.24. Thus, it may

(11) be worthwhile in future research to evaluate the

method of Bomnet et al. for typical castings produced

in the United States.

Ni, - Ni + 0.11(%Mn) - 0.0086(%Mn) 2

+ 18A (%N) + 24.5(%.) + 2.77.

+ 2.77. (12)

Thus, if the bulk chemical composition of a cast stainless steel component is known or measured, Equatons (10), (11) , and (12) can be used to estimate its

ferite content..

Bonnetet at. (1990) have developed the fowing,

slightly different method for predicting the amount of

ferrite from the chemical composition:

%F -[2l.80(Wr/ - VCr.Ni)

+ 3.391[400]/[1500 - T1

Leger (1982) found that ferrite content also increases as section thickness increases, as is shown by

the data for CF-4M stainless steel in Fgure 18. The

ferrite content as a function of thickness appears to satrate (reach a constant value) at section thicknesses

greater than approximately 45 mm (1.77 in.). lherefore, using the data for the thicker castings would provide conservative (larger than actually expected)

estimates of ferrite content for tinner castings The

work of Aubrey et al. (1982) was for castings thicker

than 25 mm (D.98 in.), soEquations (10), (11), and (12)

probably provide reasonable predictions for thicksection casngs.

Equations (10), (11), and (12) were used to estimate

the ferrite content of the 45-mm thick samples of

CF-SM that were evaluated by Leger (1982) and of the

10-mm (3.94 In.) thick samples of CFLS and CF-SM

that were discussed previously (see Figure 4). Figure 19 duo that there wasagood correlation between

predicted and measured ferrite content for these data

The solid line indicates the mean correlation; thedotted

where and dashed lines indicate the upper and lower limits,

31
respectively, which are plus and minus two times the

standard error [estimated at 3.03, as reported by

Aubrey et al. (1982)] from the mean correlation,

respectively.

45 0.63

30

25 071]

. 20 itFl3 r 

0

*E 15 -Micrography

U. ,-Ferritescope

10

,A- 0.90

5 -

F- ~ 1.04

~ 1.09

o 1.24

0 20 40 60

Thickness (mm)

Figure 18. Ferrite content of CF-8M castings as a

function of section thickness (Leger, 1982).

In order to make a reasmnably conservative estimate

of the frrite content, one would add 6.06 to the value

calculated using Equation (10) far comparison with

acceptable ferrite levels. Thus, for the adjusted ferrite

content to not exceed an upper-limit ferrite content of

15%, the mean ferrite content calculated using Equation (10) should not exceed 8.94. This value is only

slightly greater than the minimun ferrite level of 8%

that is recommended to avoid sensitization during

welding (Hazelton, 1988). Furthermore, if one adds

6.06 to 8 to account for the expected unceainty, the

minimum would be 14.06%. This -5% overlap in desirable minimum and maximum levels occurs because

of the uncertainty associated with calculatng ferrite

content from composition. Therefore, the calculation

of fere leveh from chemical composition provides

only approximate guidance for the assessmentof stainless steel castings and is of limited use in quantifying

thema embrittlement.

Recent results indicate that ferrit spacing, as well

as ferrite content, may influence the degme of thermal

embrittlement (Chopra and Chung, 1986b, 1987b,

1989a, 1989b). Procedures for estimating the ferrite

spacing in a cast stainless steel component are not

available, but, as discussed previously, Chopra and

Chung (1990a) recommend that a value of -180 .um

be assumed as a worst-case condition if the actual ferrite spacing has not been measured. Since the ferrite

spacing is a significant factor in the 0 parameter,

which corelates with the minimum Charpy impact energy expected after long-term exposure, it would be

useful to develop a procedure for estimating the feite

spacing in sainless steel castings.

Measuring ferrite content is an alternative to estimating it. If a material sample can be removed from

the component, then quantitative metographic procedures can be used to measure the ferrite content

(Leger, 1982; Aubrey et al., 1982). The procedures

typically employ either computerized image analysis

or manual point-counting techniques. The average ferrite spacing also could be measured during metallographic examination. Material sampling must be

performed carefully to ensure that the regions examined represent the bulk material. Because sample

removal often may be limited to exterior or interior regions, quantification of the ferrite volume and spacing

within the wall of the component is difficult.

The above metallographic measurements can be

made on surface replicas of selected areas of a component in the field. Metallographic replication has been

developed as a viable NDE tool (Neubauer and Wedel,

1983); Auerkai, 1983; Viswanathan, 1985; Harth and

Sherlock, 1985; Henry and Ellis, 1983; Hilton and

Steakley, 1986; Janiszewskl et al., 1986; Tlkigawa

et al., 1986; Masuyama et al., 1985), and ASTM

Emergency Standard ES-12 for field replication has

been developed.V Prototype replication equipment has

been used to remotely examine the inside surface of

steam-turbine rotors ldkigawa et al., 1986) and to inspect steam-generator tubes (Padden, 1983; Padden

and Daniels, 1976, 1981). Replicas must be made at

ambient temperature during an outage and at surface

locations that are accessible. As with sample removal,

judicious selection of the areas to be replicated is

necessary.

a. Fm apresentationbyJ. DeLong attheASTMASME-MPCJ-lWorshapattheASMlWmterAnnual Meeting, Anaheim, December 93 1986.

32
at 40 - O0 CF-8M . : .00

40 - CF-8M

.5

0 200-

00

o -

~ 0 02405

Measured ferrite content (vol %)

Figure 19. Cmparison of fAre conttpredicted from chemical composition with measured feitecontentusing Model 1 of Aubrey et aL (1982) (CF-8M data from Lger, 1982; other data from Aubrey etal., 1982).

Because the feritephase is signifiantly more magnetic than the austenite phase, ferrite content can be

measured by instruments that measure the magnetic

properes of the stainless steel. Field measurements of

ferrite content, based on magnetc properties, can be

made using a Ferritescope (Leger, 1982; Aubrey et aL,

1982). This instnrnent can be used in the field to measure ferrite levels up to 30% and has been demonstrated to produce teliable measurements far samples

with section sizes ranging from 27 to 152 mm (1.06 to

6.0 in.) (Aubrey et al, 1982). Figure 20 shows that a

good correlation between ferrite content measured using the Ferrhtscope and using a metallographic pointcounting method can be obtained. This method

appears promising, but field validation tests on representativeLWR cast stainless steel components are necessary before it can be used routinely during inservice

inspections. Section m of the ASME Boiler and Pressure Vessel Code now permits the use of magnetic

measuring instruments for determining the ferrite

levels in welds.

Bonnet et a!. report that the saturation magnetic

method (using a Sigmametre) provides the most reliable experiental technique for measuring the amount

of ferrite. This method is somewhat different than the

magnet-induction method on which the Ferritescope

is based. Curtis and Sherwin (1961) discuss using the

saturation magnetic method to measure the feite content of stainless steel welds. Either type of magnetic

measurement method should provide reasonable measurements of ferrite amount if it is propemty calNiated

and used by a qualified operator.

Factors other than the bulk chemical composition

and ferrite content and distribution in stainless steel

castings may play an important role in thermal cmbrittlement. For example, as was shown previously

(see Table 5), the chemical composition of the fete

regions is significantly different from that of the austeiite regions in these alloys. The partitioning of the

33
1030 0 0

200

00

a 0

'E .3lo 0

10 20 30

Point-counted ferrite (vol %)

Flgure 20. Comparison of ferrite content measured using a ferritescope with that determined metallographically

by a point-count procedure.

ferrite-promoting elements to the ferrite phase may affect the amount of thermal embrittlement. If such an

effect were found in fut studies, a correlation between local composition and the degree of thrmal embrittlement might be developed. Such a correlation

might be useful in the future development of alloy

compositions that would be resistant to partitioning

during casting and to subsequent embrittlement during

service at LWR operating tperatues.

4.2.2 Testing of Miniature Samples. Measuring the mechanical properties of small samples removed from stainless steel castings is anotherpotential

method for quantifying the degree of thermal embrittlement caused by service exposure. Two ap- proaches could be used. Fracture touginess could be

measured directly using subsized specimens. Alternatively, impact cne., tensile strength, or Indentation

resistance could be measured and used to estimate

fracture toughness if appropriate correlations were

available. Direct measurement of frac toughness

usually is more costly and normally requires lamer

samples than the indirect measurements However, the

direct measurements are generally more reliable than

the indirect ones

T e has been a gm deal of recent interest in the

testing of subsized specimens to measure mechanical

properties (Jaske et aL. 1986; Manahan, 1981; Haggag

et aL, 1985; Msawa and Hamaguchl, 1986; ASTM,

1986; Huang. 1986; Lucas et aL 1986; Corwin and

Hougland, 1986; McConnelletaL, 1986).Muchof this

work has been motivated by the need to measure the

properties of irradiated materials. For example, the

ASTM special technical publication (ASTM 1986) indludes 21 papers of the ASTM Symposium on the Use

of Nonstandard Subsized Specimens for irradiated

1bsting. These past studies included the comparative

testing of standard-sized and subsized specimens of

reactor pressure vessel steels and wrought austenitic

stainless steels, but they did not include testing of the

cast CF-S. C7-SM, and CF-3 stainless steels. Such

testing is needed to vaLidate the use of subsized specimens for these materials. Results of the past evaluations of subsized specimens do povide guidance for

estimating the minimum specimen size that would be

expected to yield valid measurements of Jk.

Experience indicates that the minimum specimen

dimension should be at least five to ten times the significant micostuctural feature (aske et AL, 1986),a

which for the cast stainless steels may be either the

grain size or the ferrite spacing. Experimental studies

are needed to establish what significant microstructural feature controls fracture toughness in cast stainless stes. Values of grain size and ferr spacing are

usually lazier for heavy-section castings than for thinsection castings. If, for example, ferrite spacing is the

significant feature, the minimum specimen dimension

would range from about 0.25 to 2 mm, which

a. A. R. Rosenfield, unpublished research results,

Battelle, Columbus, Ohio, 1985.

34
corresponds to typical ferrite spacing values that range

from 50 to 200 pm (Chopra and Chung, 1987).

As was discussed previously, ASTM E 813 requires

that the thickness, B, and the uncracled ligament, b, of

a test specimen be greater than 25JcQsf, where JQ is the

conditional value of fracture toughness, and sf is the

flow stress, which is the average of the ultimate tensile

strength and the 0.2% offset tensil yield strength. If

these size conditions, plus the regression-line slope

condition (see Paragraph 9.2 of ASTM E 813), are

satisfied, then JQ equals JIB

Typical tensile properties (Chopra and Chung,

1987) of cast stainless stels that were aged for 9,980 h

at 3500C (6600F) and tested at rom temperture and

at288°C (550) can be used to estimate the minimum

values of B and b that would be required to obtain valid

JC values. For example, if Jlj = 250 kjAM2, which Figures 12 and 14 show is a fracture toughness value that

could be expected for thermally embrittled cast stainless steel, and sf = 440 MPa (64 ksI) at room temperature and 300 MPa (44 kal) at 2880C (5505F), then

B. b (minimum) 25 (0.25W.44) = 14.2 mm at room

temperatur

and

B. b (minimum) - 25 (0.25)D.30) = 20.8 umm at 288C.

and

B, b (minimum) = 15 (0.10/0.30) = 8.33 mm at 288C.

Thus, because even severely embrittled cast stainless

steels have relatively high fracture toughness, fairly

large specimens, rather than subsized ones, are needed

for measuring their 1k values

One property that could be measured on small material samples is microhardness of the ferrite phase.

Recent research has shown that the Vickers hardness

number of the ferrite phase increased from 260 in

unaged material to 377 in aged material from a CF-8

casting (Chopra and Clung, 1987). It may be possible

to establish a correlation between the hardness of the

ferrite phase and the amount of thermal embrittlement.

The correlation would have to take into account the

effect of ferrite volume fraction and ferrite spacing, as

discussed previously. Future research is needed to

evaluate the potential for developing this type of

michardness-embrittlementcorrelation.

4.3 Use of Ultrasonic Testing to

Characterize Materials

If

= 100 kUjm 2, which is the lower bound used by

Framatome i

then

B, b (minimum) =25 (0.l01OA4) 5.68 mm at room

tem~perature

a. UnpublishedpaperpresentedbyG.Slamaetal..

at SMiR Postconfernce 6 at Monterey, Cal ia,

August29-30, 1983.

Use of ultMasonic techniques is being developed to

quantify material embrittlement. Both Swanson

(1981) and Vary (1976) have found correlations between ultrasonic attenuation and the fiacture toughness of wrought low-alloy steels. Their measurements

were iade on well-documented samples under carefully controlled conditions in the laboratory. Similar

studies have not been performed on cast stainless

steels, and the work has not been extended to the field

measurement of fracture toughness on components.

Thus, extensive research and development work will

probably be necessary before ultrasonic measuent

methods can be used to quantify enbrittlement in cast

stainless steels.

35
5. WELD-REPAIR PROCEDURES

If inservice NDB reveals that defects ar present in

a cast stainless steel component, the possibility of

weld-repairing that defect may be considered. CF-S,

CFSM, and CF-3 stainless steels are readily welded.

Welding is routiney used to repair defects that may be

found In castings of these alloys during component

fabrication. The major concerns with inservica

are that access may be limited and that the embrittlement may affect weldability. The former iten is

common to inservice repairs of any alloy, while the

latter item is unique to the cast stainless steels.

The procedures used for weld-repair of statically

cast components during manufaurg have been rie

viewed elsewhere (Egan et aL, 1987); a brief review of

the standard procedures developed for repair welding

during the fabrication of centrifugally cast stainless

steel pipe is presented here After the pipe is cast and

excess material is machined from the outside and inside surfaces, the pipe is fully radiographed to identify

major and minor defects, in accordance with the

ASME Code, The locations of all those defects are

mapped and recorded for future reference. The defects

are removed or reduced to an acceptable size by grinding or thermal gouging, with the final metal being removed by grinding using an aluminurm oxide or silicon

carbide wheel. The repair area is dye penetrant inspected and thoroughly cleaned with organic solvents

before welding.

Welding procedures and welders are qualified as

specified by the ASME Code (Group P-8). Welding is

performed using the manual shielded metal-arc

process and Type 316 weld rods that produce at least

7% ferrite in the weld metal. The minimum preheat is

560C (1330F); no postweld heat treatment is

employed, and the inrpass temperature is keptbelow

a. E G. Kay, Sandusky Foundry & Machine Company, personal commnication of unpublished information, October 1986.

1770C (350MF) for multiple passes. Welding current is

maintained as low as possible, travel speed is maintained as high as possible, and stringer beads are used.

After completion, all weld repairs are dye-peneftant

inspected, and major repairs also are radiographed.

There should be no major problem in developing

and qualifying similar weld-repair procedures for

field use on components that have been in service, provided that the defect is local. For example, if a fatigue

crack were found to have developed at a local geometric discontinuit in a component, then it would be feasible to consider a repair; but if fatigue damage were

found to be significant throughout a component, a repair probably would not be feasible. In repair welding,

care must be taken to minimize the introduction of residual stresses. On the other hand, postweld heat treatment Is impractical in most cases, and when it is used,

it could produce sensitized regions near the ends of the

heat-teated area For major repairs, it is likely that

specific procedures will have to be developed and

qualified on a case-by-case basis.

If repair welding is considered to be a viable option

for license extension, research is needed to evaluate

any problems that might be encountered in the welding

of thermally embrittled material. For example, it might

be found Otat thermally embrittled material is more

susceptible to weld cracking than new material.

Because the embrittlement mechanism is different for

aging at temperatures greater than 400°C (750TF) than

for aging at temperatures less than 3506C (6600F), the

research should concentrate on the welding of materials that have been aged for long times at 350VC or less.

Studies using samples that have been removed from

actual LWR service would be particularly appropriate.

Results of this type of research would provide guidance for the development and qualification of weldrepair procedures for cast stainless steel LWR

components

36
6. LIFE ASSESSMENT PROCEDURE

There are two basic approaches to maximizing the

structural integrity of LWR components for operation

during and beyond the original license period. F ,

one can evaluate the current condition of a component,

estimate its remaining safe IMe, and decide to repair or

replace it only when there is cause to do so. This approach is employed widely to critical structural components whose failure would impact safety and whose

unnecessary replacement wouldbe extremely costly. A

second approach would be to arbitrarily replace a cornponent after some specified service life without attempting to determine its current condition. This

method is applied to critical structural components

whose failure would impact safety and whose replacement is easier and less costly than evaluating their current condition. For most cast stainless steel LWR

components that are subject to the possibility of thermal embrittlement during long-term service, the second approach is undesirable because of an unfavorable

cost/benefit ratio; thus, the first approach is the most

reasonable one.

Figure 21 presents a generic procedure for aging

and license renewal evaluations of LWR cast stainless

steel components. Nine major steps are listed.

The first step is to review all of the design, fabrication, and construction records related to the components being evaluated. Some of the items listed in

Step I may not be available. Those items should be

Identified, and plans to obtain them during the next inservice inspection (151) should be developed, if possible. In cases where it is not possible to obtain

case-specific data for austenitic stainless steel products (e g., impact and fracture toughness data are not

required by Section H1 of the ASME Boiler and Pressure Vessel Code), worst-case data should be obtained

from handbooks, technical literature, and industrial experience. The information collected in Step 1 should

be stored for easy retrieval and reference in the future.

All of the ISI records for the components are reviewed in the second step. The information on defect

indications and locations and maximum undetected

flaw sizes is needed for the fracture mechanics evaluation (Step 4). The chemical compositions of the materials should be verified by measurements on the

components. Essential information (such as size and

location of flaws in the base metal) not being obtained

during the current SI programs should be obtained

during fuature IS! programs. In many cases, key locations are not accessible to ISI, so the original records

must be used with appropriate factors applied to

account for uncertainty.

The third step is to review all of the data on operating history. The operating temperature is needed to

evaluate the degree of thermal embrittlement, and data

on loadings and cyclic operations are needed to help

estimate the amount of crack propagation that might

have occurred or might occur in the future. Indications

of past problems may help point out areas that should

be evaluated in Steps 4 and 5.

A fatigue and fractiu-mechanics evaluation is performed in the fourth step. The locations and sizes of

the existing flaws are identified from the ISI records.

Simplified, conservative loadings are estimated from a

review of the operating history. Stress analyses are performed to identify regions of the highest stress, and the

stress history at the flaw locations is estimated. If the

review of operating history reveals that certain loads

were not included In the original design, revised calculations for the ASME Section m cumulative fatigue

usage fators are made. Crack-propagation analysis is

performed to estimate the maximum possible defect

size at the end of the next operating period (EONOP).

Finally, the minimum required fracture toughness corresponding to the worst-flaw size at the EONOP are

determined.

The fatigue and fracture-mechanics evaluation performed in the fourth step is conservative because it assumes that the flaws in the cast stainless steel

components are crack-like. This assumption is made

because the current inservice inspection methods are

not capable of completely characterizing the flaws. If

an emerging inservice inspection method can positively identify that a given flaw is not crack-like, then

a fatigue-crack initiation analysis would precede the

crack propagation analysis. In such a fatigue analysis,

an effective stress concentration factor is computed for

the flaw (Hertzberg, 1983), and then the usage factor at

EONOP is determined. If the usage factor is greater

than 1.0, it is assumed that a crack will be initiated at

the given flaw before EONOP; in this case, a crack

propagation analysis would be needed.

In the ffth step, the material condition of the componentis assessed to determine the minimum expected

fracture toughness atEONOP. Three techniques can be

employed to assess the degree of thermal embrittlement of a cast stainless steel component, as discussed

in Section 4. The first technique uses the simplified

37
Figure 21. Generic procedure for the evaluation of LWR cast stainless steel components.

1. Collect and review design, fabrication, and construction records

1.1 Design drawings and specifications

1.2 Measure chemical composition of material

13 All available mechanical propties for original material

1.3.1 Tens1le ultimate, yield, elongation, reduction of area

1.3.2 Impact: energy versus temperature, fracture appearance

1.3.3 Fracture toughness

1.3.4 Fatigue strength

1.3.5 Fatigue-crack propagation behavior

1.4 ASME Section M cumulative fatigue usage factors

1.5 As-built drawings and dimensions

1.6 Name, location, and dat of fabrication

1.7 Field installation procedures

1.8 Type, location, and repair procedure for all casting defects

1.9 Number, location, type, and heat treatment of welds

2. Review inservice inspection (ISI) records

2.1 Location of inspected regions

2.1.1 Were all welds inspected?

2.1.2 Was base metal inspected?

2.1.3 Were regions of high incidence of weld-repair during fabrication inspected?

2.2 Nature, size, distribution, and disposition of indications (if any)

2.3 Estimate maximum undetected flaw size based on ISI procedures

2A Measurements of chenical composition (if any)

2.5 Other information

3. Review operating history

3.1 Hours of operation

3.2 Number of stastop cycles

3.3 Number of major load swings

3A Number of upsets

3.5 Temperature

3.6 Pressure

3.7 Records of any past problems related to the component

4. Perform fatigue and fracture mechanics evaluation

4.1 Identify size and locations of existing flaws from ISI records

38
Figure 21. Generic procedure for the evaluation of LWR cast stainless steel components. (Continued.)

4.2 Estimate simplified, conservative loadings fom review of oeating history

4.3 Perform stzess analysis of component

4.3.1 Identify regions of highest stress

4.3.2 Estimate stress history at potential locations of flaws

4.4 Update ASME Section m cumulatv fatigue usage factors, if design basis violated

4.5 Compute possible flaw growth (if any) during next operting period

4.6 Compute the minimum required fract toughness and worst-case flaw size at dhe end of next

operating period (EONOP)

5. Assess material condition

5.1 Use simplified analytical model

5.1.1 Use composition and maximum operating temperature to estimate lower-bound

Charpy V-notch impact energy value at EONOP

5.12 Or use composition, ferrite content, and ferrite spacing to estimate minimun Charpy

V-notch impact energy at long life

5.1.3 Estimate lower-bound fracture toughness atEONOP from fracture toughness versus impact

energy correlation

5.2 Assess ferrite content

5±-1 Calculate upper bound ferrite content from composition

52.2 Measure ferrite content

Metallogaphic sample from component

Surface replication of component

Ferritescope

5±3 Use guiddlines to estimate ractre toughnes atEONOP

5.3 Measure properties of small Samples from component

5.3.1 Frcture toughness testing

5.32 Charpy testing

5.3.3 Microhardnesscorrelation

5.3A Use simplified analytical model to estimate faCtre toughness at EONOP

6. Evaluate component integrity at EONOP

6.1 Compare the minimnum required fracture toughness atEONOP with the corresponding estimated

lower-boundvalue

6.1.1 Is the minimum required fracture toughness smaller han the estimated lower-bound value

by an acceptable margin?

6.1.2 Estimate critical flaw size atEONOP

6.1.3 Is critical flaw size safely detectable?

6.2 Evaluate the potential for leak-before-break (LB) to occur.

39
Figure 21. Teneric procedure for the evaluation of LWR cast stainless steel components. (Continued.)

6.2.1 Is critical flaw size less than the wall thickness?

6.2.2 Will a through-wall crack produce a safely detectable leak before it reaches a critical size?

7. Establish action to btaken

7.1 Continued operation to EONOP if acceptable safety margins ar present

7.2 Determine if repair procedures are needed

7.3 Determine if replacement is needed.

8. Establish plan for next ISI

8.1 Determine extent of inspection required

8.1.1 Consider areas of high stress

8.1.2 Consider areas where defects are most likely to exist

8.1.3 Consider areas where weld repairs have been made

8.1.4 Consider areas where temperature is the highest

8.1.5 Consider areas where cumulative fatigue usage factors are high

8.1.6 Consider areas wher a flaw may grow by fatigue loading

8.1.7 Consider areas where IGSCC may occur: welds with less than 8% ferrite

8.2 Determine next inspection interval

8.2.1 Use information on worscase flaw size and crack-growth rate from Step 4 and on critical

flaw size from Step 6

8.2 Compute expected remaining safe life

8.13 Set interval for next inspection at lesser of predetermined fraction (e.g., 1/4) of remaining

safe life, or 10 years

8.3 Develop a thermal-embrittlement surveillance program

8.3.1 Make Charpy V-notch or fracture-toughness specimens of archival material

8.3.2 Use both base metal and weld metal

8.3.3 Expose the specimens at temperatures that lead the embrittlement rate of components by a

factorof 1.5 to 3

8.3.4 Evaluate selected samples at 5-year intervals to assess

Rate of embrittlement (change in impact energy or fracture toughness)

Degree of emnbritement (value of impact energy or fracture toughness)

9. Re-evaluatecomponent(s)

9.1 Perform as required

9.2 Return to Step 2 of procedure

40
analytical model expressed by Equations (4) and (9) to

estimate the minimum Charpy V-notch impact energy

valueatEONOP,orthe G parameterofEquation (5)

to estimate the minimum Charpy V-notch impact energy after long-tem aging. Then, using a Jjc (fracture

toughness) versus CVN corzeation, minimum Ji isestimated. The second technique is based on eithercalculating or measuring the ferrite content and using

guidelines or an embrittlement model to determine If

the ferrite value is acceptably low. The third technique

is to measure the current Chaipy V-notch impact energy, fracture toughness, or microlardness of samples

that have been removed from a component; then the

first technique, i.e; the simplified analytical model, is

used to estimate the JIC value at EONOP.

In the sixth step, the minimum required fracture

toughness from Step 4 is compared with the estimated

-actual lower-bound fracture toughness at EONOP

from Step 5. and the margin of safety is determined.

Using the estimated stresses from Step 4 and the estimated lower-bound toughness from Step S. critical

crack size is determined. Then, the detectability of the

critical flaw size is assessed. Finally a leak-beforebreak (LBB) analysis is performed to deternine the

margin between the critical crack length and the leakage-detectable crack length that corresponds to ten

times the detectable leak rate.

Once the component integrity at EONOP is

evaluated, a plan of action is developed in the seventh

step. If acceptable safety margins are present, and the

ASME Section III cumulative fatigue usage factor at

EONOP is less than 1.0, continued operation to

EONOP may be permitted. Potential need for repair or

replacement must be considered.

The eighth step is to establish a plan for future ISI of

the component. The plan should define both the extent

of inspection required and the interval at which reinspection should be performed. It may be useful to consider developing athermal-mbrittlement svellance

program similar to the programs used for evaluating

radiation effects. Such a program would provide valuable data from realistic exposure conditions.

The ninth step is to reevaluate the component after

future inspections or as required by other circumstances. The revaluation will start at S tep 2, upgraded

by the data from the most recent evaluation, and will

p- cedfrom that point.

The major difficulties in implementing this overall

procedure are encountered in Step 5, Assess Component Condition. The simplified analytical model

(Step 5.1) can be applied with a reasonable degree of

confidence, even though further validation of the correlation between CVN and Jic values is needed (see

Figures 12 and 14). However, ferrite assessment

(Step 5.2) and the measurement of properties

(Step 5.3) have not been fully developed and validated. Quantitative evaluations of possible ferrite measurement techniques and subsized specimen testing

methods are needed. Correlations and/or models are

needed to relate ferrite content, spacing, and hardness

to the amount of thermal embrittlement. Also, a model

for predicting future degradation caused by thermal

embrittlement should be developed to remove uncertainties associated with the current approach. Thus,

considerable work is needed in the area of assessing

and forecasting the condition of cast stainless steel

components that operate at temperatures near 3000C

(570 in light water reactors.

Fracture-mechanics evaluations also could be difficult to conduct in some situations. Framatome has performed fracture mechanics evaluations for their cast

stainless steel components and has developed a minimum required value of lO kJ/tm2 for JI, at the end of

operating life. The establishment of a minimum Jti

value simplifies the evaluation of standardized components because there is no need to repeat the fracturemechanics analysis for each case. However, such a

simplified approach may not be applicable to all U.S.

LWRs because of the variation in component deign,

fabrication, and operation from one reactor to another.

It would be useful to perform fracture-mechanics

analyses for several different classes of cast stainless

steel components, so minimum required values of Ji at

the end of operating life could be developed for each

class. These values then could be applied to the caseby-case evaluation of components for license renewal.

a. UnpublishedpaperpresentedbyG.Slamaetal.,

at SMiRT Postconference Seminar 6 at Monterey,

California, August 29-30,1983.

41
I

7. SUMMARY

This report has reviewed the quantification of thermal embrittlement damage in CF-8, CF-8M, and

CF-3 cast stainless steel components subjected to light

waterreactor (LWR) operating temperatures. Thesa alloys have an austerite matrix with islands of ferite.

For LWR applications, typical volume fractions of the

ferrite phase are In the range of 8 to 20%. The ferrite

phase is subject tD embrittlement by the formation of

alpha-prime after long-tem (many years) exposures

at temperatures new 3000C (5709F). The amount, distribution, and degree of embrittlement of the ferrite

phase deternine the extent to which a component has

been degraded by thermal embrittlement. Tis degradation is manifested by a loss of Charpy V-notch

(CVN) impact energy and fracture toughness (Jx).

Other mechanical properties, such as tensile strength,

fatigue strength, and low-cycle fatgue-crack-growth

resistance are not markedly affected by thermal embritlement, so the primary concern in addressing the

renewal or extension of LWR operating licenses is the

loss of fracture toughness after long-term service.

However, more data on the effect of thermal embritlement on fatigue-crack-growth resistance are needed.

especially for CF-8 stainless steel and for high-cycle fatigue.

Prom the maximum service temperatme and chenical composition of the alloy, an aging parameter is

proposed for estimating the lower-bound roomtemperature CVN values for service times up to 100

years, or the minimum CVN values at a very long time

can be estimated using the 0 parameter. A CYN impact energy versus JD correlation then can be used to

estimate a corresponding lower-bound value of the

fracture toughness.

Nondestructive examination (NDB) of cast stainless

steel components yields information on the size, location, and distribution of defects. This information can

be employed in fractutr-mechanics evaluations of

structural integrity. However, because of the coarse

and variable nature of the microstructure of cast stainless steels, ultrasonic testing (UO) is limited in its usefulness as an NDE tool; radiography normally is used

for the NDE of these materials, especially for thick

sections. Research is in progress to develop and validate advanced UT methods for inservice inspection of

cast stainless steel components. Howevea, neither current nor advanced NDE methods characterize the mechanical properties or metallugical condition (degree

of embrittlement) of insevice components.

Correlationshavebeen developedfor esmating the

ferrite content of stainless steel castings from their

chemical composition. However, these correlations do

not predict the ferrite spacing, which is a parameter

that influences fracture toughness. Castings that contain more than about 10 to 15% ferrite appear to be the

most susceptible to thermal embrittlement, but no precise guideline for an acceptable ferrite content has

been established, and mateia with even less than 10%

ferrite may be degraded by thermal embrittlement.

Metallographic examination of small samples removed from components or of surface replicas can

provide data on the ferrite content and spacing at specific locations in castings. The feritescope is a mag- netic instrument that can be used in the field to

nondestructively measure ferrite contents up to 30%.

Subsized specimens from small samples ofmaterial

that have been removed from service can be used to

measure fracture toughness. Guidelines for minimum

specimen size ar provided by ASTM B 813. Those

guidelines are expected to be conservative, but they have not been verified directly for the testing of cast

stainless steels.

Long-term service exposure increases the mirohardness in the ferrite phase of cast stainless steels, but

no correlation between th measured microhardness

and change in impact resistance or fracture toughness has been established The correlation also would have

to account for the effects of ferrite content and spacing

to provide a quantitative estimate of fracture

toughness.

Weld-repair procedures have been developed for

the removal of casting defects during original fabrication. It shouldbe reasonably easy to adapt these proce- dures to the inservice repair of stainless steel castings

However, repair welding has not been fully qualified

and validated for material that has been thermally embrittled by long-term service exposures in LWR

environments.

The recommended approach for evaluating cast

stainless steel components for LWR license renewals,

is to assess the current condition of the component, estimate its remaining safe life, and decide on continued

as-is use, repair, or replacement of the component The

42
alternative approach of autmatically replacing a conponent after a given period of service is not usually

feasible.

A nine-step procedure for the aging and license renewal evaluation of cast stainless steel components has

been outlined. Tbe first three steps include the review

of fabrication and construction records, ISI records,

and operating history of the components. In the fourth

step, the data colected from these reviews are used to

perform the fracture mechanics evaluation to determine the minimum required fracture toughness and the

worst-flaw size at ihe end of the extended operating

period (EONOP). Three different methods are presented in Step 5 to estimate the lowerbond fracture

toughness of cast stainless steel components at

EONOP. Then, the minimum required fracture toughness at EONOP is compared with the corresponding

estimated lower-bound value, and a leak-beforebreak analysis is performed. The action regarding continued operation, repair, or replacement of a

component is established in Step 7, and finally, plans

for the next ISI are made.

43
8. RECOMMENDED FUTURE WORK

Improved methods need to be developed for evaluating the fitness of cast stainless steel LWR componenuts for service during original or extended lcensing

periods and for reducing the conservatism inherent in

the procedures discussed above. Therefore, we recommend thai fure work be in the following

13 areas:

1. Guidelines of worst-case data should be developed for use when case-specific data are

not available.

2. Thermal-embrittlement surveillance programs similar to the programs used for evaluating radiation effects should be developed.

Such programs would provide valuable data

from realistic exposure conditions.

3. More CVN and Jxc data on thermally aged

material tested at normal LWR operating

temperatures [about 280 to 29O0C (536 to

554F] are needed.

4. An improved aging parameter should be developed that more accurately models the

aging mechanisms that cause thermal

embrittlement during long-term service

exposure

5. Work should be undertaken to determine if

there is some critical combination of ferrite

level and spacing above which ferite distribution in cast stainless steel is continuous. If

such a limit exists, then a method of estimating ferrite content and spacing from composition and production history should be

developed.

6. Fatigue-cc-growth studies of aged CF-8

stainless steel should be performed. Nearthreshold, as well as low-cycle, tests are

needed on both aged and irradiated materials.

Data for high-cycle fatigue behavior of aged

CF-8M stainless steel are also needed.

7. The use of metallographic surface replication. the feruitescope, and other magneticproperty measurement tools, along with

eddy-current methods, should be evaluated,

demonstrated, and validated as inservice

NDE tools for evaluating the condition of

stainless steel castings.

8. Experimental studies should be performed to

determine the smallest specimen size for

valid measurement of the fracture toughness

of cast stainless steels.

9. Work should be undertaken to evaluate the

correlation between the micohbardness of the

ferrite phase and fracture toughness for cast

stainless steels.

10. Procedures for repair welding cast stainless

steels that have been subject to long-erm service in LWR environments should be qualified and validated for use in license renewal.

11. Fracture-mechanics analyses of representative component/load-history combinations

should be conducted. Such analyses can provide estimates of the required minimum Ju

value.

12. Current work on the development of advanced ultrasonic methods for inspection of

cast stainless steel components should be

continued.

13. Present ISI programs should be modified to

obtain essential data that are needed for the

evaluation of LWR cast stainless steel

components

44

i
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49
NRC FORM 235 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER

2I1102(Assigned by NRC Add VoL. SWP. Rev..

=2= BIBLIOGRAPHIC DATA SHEET *AddN,.) M u oNUREG/CR-5314, Vol. 3

2. ii.E AND SUBTIE EGG-2562

Life Assessment Procedures for Major LWR REPORT PUBLIEUE

Components I YEAR

October 1990

Cast Stainless Steel Components 4 FIN OR GRANT NUMBER

A6389

S. AUTHOR(S) 6 TYPE OF REPORT

Technical

C.E. Jaske, V.N. Shah 7.PERIODOOVERED fa^#Do:e")

. PEfORMNC ORGANIWZTIN -tME AND ADDRESS 1 ro Ac er Ax w g an and mawvng adat

Idaho National Engineering Laboratory

EG&G Idaho, Inc.

Idaho Falls, Idaho 83415

L SPONOPNGORG1N-NAME AND AIDDRESS iWNlFI, ope osubovw';#crmctrroNA$DWuO, fsrJinL r~v^&kem RBu~y Coomn,

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington.- DC 20555

10. SUPPLEMENTARY NOTES

13. ABSTRACT (200 words or lSss)

This report presents a procedure for estimating the current condition

and residual life of safety-related cast stainless steel components in light

water reactors (LWRs). The procedure accounts for loss of fractures toughness caused by thermal embrittlement and includes the following: a review of

design and fabrication records, inservice inspection records, and operating

history; a fracture mechanics evaluation to determine the required toughness

estimates; and criteria regarding continued service, repair, or replacement

of the component being evaluated. The report discusses the available Charpy

V-notch impact energy, fracture toughness, tensile strength, fatigue resistance, and fatigue-crack growth data, and presents two methods for assessing

the degree of thermal embrittlement: metallurgical evaluation and analytical modeling of inservice degradation.

12. KEY WOROSIOESCRIPTORS (Lt ods or phases ihat f assis reseathers In ocaflg p 1 unlimitedMENT

LWR cast stainless steel components

Thermal enbrittlement 14. SECURIY CLASSFATM

Charpy impact energy unclassified

Fracture toughness Ax

Life assessment procedure unclassified

Inservice inspections 1s. NUMBER OF PAGES

Fatigue

is. PRICE

NRC FORM 335 (2.89