NUREG/CR-5314EGG-2562Vol. 3

LieAssmntPoeueLife Assessment Procedures,:for Major LWR Components

Cast Stainless Steel Components

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

Idaho National Engineering LaboratoryEG&G Idaho, Inc.

Prepared forU.S. Nuclear Regulatory Commission

AVAILABIIITY NOTICE

Avallablity d Reference Materials Cited In NRC Publications

Most documents cited In NRC publicatlons wil be available from one of the following sources:

1. The NRC Public Document Room, 2120 L Street, NW, Lower Level. Washington. DC 20555

2. The Superintendent of Documents, U.S. Governrnent Printing Offic. P.O. Box 37082. Washington.DC 20013-7052

3. The National Technical Information Service, Springfield. VA 22181

Although the listing that follows represents the majority of documents cited In NRC publications, It Is notIntended to be exhaustive.

Referenced documents available for Inspection and copyig for a fee from the NRC Pubilo Document RoomInclude NRC correspondence and Internal NRC memoranda; NRC Office of Inspection and Enforcementbulletins. chrculars, Information notices, Inspection and investigation notices; Ucensee Event Reports; ven-dor reports and correspondence; Cornmlsslon papers; and applicant and licensee documents and corrs-spondence.

The following documents In the NUREG series are available for purchase from the GPO Sales Program:formal NRC staff and contractor reports, NRC-sponsored conference proceedings, and NRC booldets andbrochures. Also available are Regulatory Guides. NRC regulations In the Code of Federal Regulations, andNuclear Regulatory Commission Issuances.

Documents available from the National Technical Information Service Include NUREG series reports andtechnical reports prepared by other federal agencies and reports prepared by the Atomic Energy Commis-slon. forerunner agency to the Nuclear Regulatory Comrmsslon.

Documents available from publio and special technical libraries include at open literature Items, such asbooks. Journal and periodcal articles, and transactions. Federal Register notices, federal and state legisla-tion, and congressional reports can usually be obtained from these libraries.

Docurnents such as theses, dissertations, foreign reports and translations, and non-NRC conference pro-ceedings are available for purchase from the organization sponsoring the publication cited.

Single copies of NRC draft reports are available free, to the extent of supply, upon written request to theOffice of Information Resources Management, Distribution Section, U.S. Nuclear Regulatory Commrdsslon.Washington. DC 20555.

Copies of industry codes and standards used In a substantive manner In the NRC regulatory process aremaintained at the NRC Llbrary. 7920 Norfolk Avenue, Bethesda. Maryland. and are available there for refer-ence use by the public. Codes and standards are usually copyrighted and may be purchased from theoriginating organization or, If they are American National Standards, from the American National StandardsInstitute. 1430 Broadway. New York. NY 10018.

DISCLAIMER NOTICE

This report was prepared as an accoult of work sponsored by an agency of the United States GovernmentNeitherthe United States Government norany agency tereoforany oftheir employees, makes any warranty.expresed or Implied, or assumes any legal liability of responsibility for any third pary's use, or the results ofsuch use, of any Information, apparatus, product or process disclosed In this report, or represents that Its useby such third party would not Infringe privately owned rights.

NUREG/CR-5314EGG-2562Vol. 3RM, R9

Life Assessment Proceduresfor Major LWR Components

Cast Stainless Steel Components

Manuscript Completed: September 1990Date Published: October 1990

Prepared byC. E. Jaske, V. N. Shah

G. H. Weidenhamer, NRC Program Manager

Idaho National Engineering LaboratoryManaged by the U.S. Department of E1nergy

EG&G Idaho, Inc.Idaho Fas D) 83415

Prepared forDivision of EngineeringOffice of Nuclear Regulatory ResearchU.S. Nuclear Regulatory CommissionWashington, DC 20555.NRC FIN A6389Under 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 proce-dure accounts for loss of fractur toughness caused by hrmal embrittlement and includesthe folloing: a review ofdesign and fabrication records, nserice inspection recods, andoperating history a fracture mechanics evaluation to determine the required tughness atend-of-life using worstMoads and woast-law indications; current and future toughness es-timates: and criteria regarding continued service, repair, or replacement of the componentbeing evaluated. The report discusses the available Charpy V-notch impact energy. fcturetoughness, tensile strength, fatigue sistance. and ftge-crac k growth data, and presentstwo methods for assessing the degree of thermal embrittlement: metallurgical evaluationand 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 usenuclear power to commercially generate electricityand, therefore, has some of the olde operating com-mrcial reactor As U.S. light water reactors (WRs)have matured, problems associated with timo- or use-dependent degradation (aging) mechanisms such asstress corrosion, radiation embrittlement, fatigue, andother edfcts have occurred and have raised questionsabout the continued safety and viability of older nu-clear plants Some of the recent aging-related prob-lems include primary water sue corrosion crackingof pressurized water reactor (PWR) pressurizer heatersleeves and instrument nozzles and PWR steam genr-ator tube plugs, steam generator tube ruptures causedby high-cycle fatigue and by a faied tube plug, signif-icant wall-thinning of light water reactor metal con-tainments caused by corrosion, fatigue failure ofboiling water reactor (BWR) recirculation pump inter-nals (resulting tn potential damage to reactor pressurevessel core and internals), catastrophic failure of a"nonnuclear" portion of a PWR feedwater line causedby erosion-corrosion, and through-wall thermal-fatigue cracks in high-pressure safety injection linesand a residual heat removal lie.

At the same time, with a continually increasing de-mand for electricity and limited new generating capac-ity under construction, the U.S. electric utilities aremotivated to keep their existing plants operating be-yond the original design liMf at as high a capacity aspossible. The economics of plant life extension areclearly favorable. Studies cosponsored by the U.S.Department of Energy (DOE) and the Electric PowerResearch Institute (EPRI) show that replacing anysingle nuclear plant component can easily be justified,if the life of the plant can be extended for a number ofyears. Extending the life of a 1000-MW plant by20 years is expected to realize a net present worth ofperhaps $1 billion.

Therdore, the potential problems of managing ag-ing in older plants and the resolution of technicalsafety issues in consideration of the development ofappropriate license renewal criteria have become amajor focus for the research sponsored by the U.S.Nuclear Regulatory Commission (USNRC). An im-portant part of the USNRC research effort is theNuclear Plant Aging R ch (NPAR) Pram that isbeing conducted at several national laboratories, in-cluding the Idaho National Engineering Laboratory(INEL). One of the NPAR program tasks at the INELis to develop the appropriate technical criteria for theUSNRC to assess the residual life of the major PWR

and BWR components and structures. Theseassessments will help the USNRC identify and resolvesafety issues associated with LWR aging degradationand develop policies and guidelines for making operat-ing plant license renewal decisions

Most of the effort for this life assessment task isfocused on integrating, evaluating, and updating thetechnical information relevant to aging and license re-newal from current or completed NRC and industryresearch programs. A five-step approach is being pur-suod to accomplish the life assessment task: (1) iden-tify and prioritize major components, (2) identifydegradation sites, mechanisms and stressors, and po-tential failure modes for each component, and thenevaluate the current inservice inspection (1SI) mth-ods, (3) assess advanced inspection, surveillance, andmonitoring methods, (4) develop life assessment mod-els and procedures, and () support the development oftechnical criteria for liese renewal.

A brief discussion of each of these steps follows:

1. Identification and prioritlzatlon of majorcomponuetr Virtually all major equipmentcontained within a nuclear plant complex issubject to some aging degradation and mustbe evaluated in an aging and licensee renewalprogram. From the USNRC's perspective ofensuring the health and safety of the public,the first step in this assessment task: was toIdentify those major components critical tonuclear power plant safety. Components thathelp contain the release of fission productsduring normaL off-normal, or accident con-dhions were selected. The PWR components(in rough order of importance) include thereactor pressure vessel (RPV); the contain-ment and basemat reactor primary coolantpiping, safe ends, and nozzles; steam genera-tors; reactor coolant pump bodies; pressurizerand associated surge and spray lines; controlrod drive mechanisms; cables and connec-tors; emergency diesel generators; RPV inter-nals; RPV supports; and feedwater lines andnozzles. A similar list was developed forBWRs except that the containment wasranked most important. Although some PWRsteam generator and BWR recirculation pip-ing have been replaced, their replacement is amajor, time-consuming operation. In addi-tion, steam generator tubes and recirculationpiping constitute a part of the primary pres-sure boundary, and therefore, aging

iV

I

evaluation of these components is included inthis ask. The lifethme of many of the smaller,less expensive components such as thepumps, valves, sensors, batteries, controls,etc. is often less then the initial license periodof 40 years, and these components are usuallyrepaired, refurbished, or replaced relativelyfrequently. Therefore, these components arebeing studied in other NPAR tasks that arefocusing more on reliability, availability, andmaintainabiity than on life assessments.

2. Identation of the degradation sites, mech-anism and stressors. and the potential fail-mm modes and evaluation of the current 1SImethodr fmTe-oruse pendentdamageoraging can be caused by one or several differ-ent mechanisms active within a component,structure, or material and, if not recognizedand properly managed, may result in sometype of failure or impairment of function. Thedegradation is often the result of interationsbetween design, materials, operationalstressors, and environments. Poor design,improper material selection, severe environ-ments, or inadequate maintenance practicescan accelerate the degradation. Therefore,identification and understanding of the de-sign, materials, stressors, environments, andaging mechanisms is essendaL This step con-sists of identifying and qualitatively evaluat-ing the stressors, potential degradation sitesand mechanisms, probable failure modes, andcurrent ISI methods for each of the selectedmajor coponients. This qualitative analysisof the degradation sites and processes andcurrentlSI methods Isessential to developinga proper understanding of tie impact of agingon safe operation of nuclear power plants andidentifying and prioritizing the unresolvedtechnical issues relevant to nuclear powerplant aging and license renewaL

3. Assessment of advanced Inpection, swoveU-lance, and monitoring methodr Knowledgeof the current damage state of the material isessential for a proper assessment of he resid-nal life of a component or structure. InserviceInspectns are performed to measure the cur-rent state of damage. However, many of thestandard nondestructive examination (NDE)methods employed to satisfy current ISI re-quirnents were developed for the detectionand qualitative assessment of fabrication-related flaws. These methods are not entirely

adequate for residual lifeassessment. Inspec-tions for life assessment generally requiregreater detection reliability and a more quan-itative determination of defects and accumu-lated damage than tradital ISL. Therefore,it is essential that emerging methods for in-section be developed and evaluated and thatthey accurately determine, for example, thesize, shape, location, orientation, and type ofboth surface and internal flaws.

4. Develop (or evaluate) Ife assessment modelsandprocedures: An important feature of thisNPAR task Is the development of models orprocedures to estimate fte remaining usefullife of the major, risk-sIgnificant LWR com-ponents and structures. The residual lifeassessment models and procedures appropri-ate for various structures and componentsmay differ considerably. However, a generalapproach that may work well for many of thereactor primary system pressure boundarycomponents, the containment, and possiblyother major components is as follows: (a)evaluate the present state of the component,(b) estimate the change expected during theplanned operating period, (c) identify appro-priate design requirements, and (d) comparethe estimated condition of the component atthe end of the planned operating period withthe design requirements.

The component state or condition of interestmight be one or more of its material proper-ties. such as fracture toughness, fatigueusage, strength, elasticity, etc., and/or one ormore geometry characteristics, such as flawsize, wall thinning from wear, corrosion, anderosion, tolerances, etc. These characteristicscan be determined by inspection, destructiveexamination, analysis, or other methods.However, note that the use of analysis to esti-matekey matial properties (such as fracturetoughness or elasticity) and component ge-ometry changes from wear, corrosion, ero-sion, flaw growth, etc., generally requires agood knowledge of the initial material condi-ton and tolerances and a very good under-standing of the relationships betweenstressors and states. This understanding mustbe based on appropriate data and physicalmodels.

The changes in the component state or condi-tion during some futuae planned qperating pe-tod (such as a license renewal period) can be

v

estimated based on conservative extrapola-don of previous ISI (for example, wall thick-ness data) and destructive examinationmeasurements or analysis. Again, the analy-sis tools should reflect a good undetandingof the physical relationships between the op-erating stressors and the changes in the com-ponent state or condition.

The component or structural design require-ments should be based on traditional engi-neering analyses (thrmal, hydraulic, stress,neutronic, etc., as appropriate) or testing andreflect appropriate margins between the ex-pected loads and the calculated failure loads.The original design analysis may be suffi-cient for some components. A revised designanalysis may be necessary for other compo-nents subject to additional cycles, previouslyunknown stressors, etc. New informationmay, in some cases, allow a reduction in or re-vised estimate of the margi

The final step, comparing the design require-ments with the expected condition of thecomponent, would be followed, of course, bya decision to replace or repair the componentor leave the component as is. The comparisonof design requirements with a component'sexpected condition might also influence theplanned future operating period (license re-newal period) and the frequency of future ISIor destructive examination activities.

A somewhat different approach is associatedwith components subject to environmentalqualification (EQ) requirements. These arecomponents that must function properly dur-ing a normal operating period of some givenduration and also during selected design basisaccidents (DBA). Here three general ap-proaches might be considered: (a) compareactual operating environments with thepreagingEQ environments, (b) compare esti-mated mechanical properties at the end of theplanned operating period with the preagedEQ sample mechanical properties. and (c) re-move samples from the plant, add additionalaging (to reflect the expected aging during theplanned future operaing period), and subjectthe samples to DBA testing. The first two approachesrequire modeling of the complex re-lationships between environmental stressors,aging degradation (which are often materialproperty changes), and time. ITe third ap-

proach provides results with more certaintybut possibly at the highest cost.

The residual life assessment models or proce-dures will have a strong impact on the devel-opment of technical criteria for licenserenewal. It is important that the limitationsand uncertainties associated with whicheverapproach is chosen be assessed so that confi-dence intervals can be calculated and safetymargins properly assessed,

5. Development of technical criteriafor licenserenewah The results and outputs obtainedfrom the activities discussed above will beused to help the USNRC ident and resolvetechnical safety issues associated with LWRaging degradation and develop policies andguidelines for making license renewal deci-sins The qualitative analysis of the degrada-tion sites and processes and current IS1methods (Step 2) will facilitate the identifica-tion and prioritization of the umesolved tech-nical issues relevant to nuclear power plantaging and license renewal. The qualitativeanalysis will also help provide the USNRCand industry a basis for lifa assessment model(or procedure) development. The assess-ments of the advanced inspection, surveil-lance and monitoring methods could providea technical basis for acceptance or rejectionof any ISI associated with a life assessmentMost important, the use of acceptable lifeassessment models or procedures could formthe technical basis for a revised license re-newal safety analysis report (SAR) as well asa revised NRC standard review plan and reg-

- guides.

Overall project results to date are as follows. Thefirst two steps have been completed, and results arepresented in a two-volume report: Residual LifeAsessmentofMajorLight WaterReactor Components- Overview, NUREG/CR-4731. Some progress hasbeen made on Step 3. the assessment of emerging in-spection, surveillance, and monitoring methods Theemerging inspectio monitoring and material evalua-tion methods are briefly discussed in Volume 1 ofNUREO/CR-4731. An indepth assessment of bothadvanced fatigue monitoring and material evaluationmethods will be published in the near future. Indepthassessments of improved acoustic monitoring, ad-vanced ultrasonic testing, and eddy current techniquesare in progress.

Life assessment procedures will be published soonfor five major components as part of Step 4: PWR

Vi

I

reactor pressure vessels, reinforced concrete contain-ments, cast stainless steel components, PWR stamgenerator tubes, and metal containments. We plan topublish these procedures In this multi-volumeNUREG report, as indicated below.

PWR Stea Generator Tubes - Volume 4Metal Containments - Volume S

Life assessment niodels or procedures for the othermajor components and structures will be presented infuture volumes of this report

PWR Reactor Pressure Vessels - Volume 1Reinforced Concrete Containments - Volume 2Cast Stainless Steel Components - Volume 3

P. E. MacDonaldV. N. ShahSeptember 1990

vii

EXECUTIVE SUMMARY

Many critical pressure boundary components incommercial light waterreactors (LWRs) are composedof cast stainless steels Life assessment procedures areneeded for these components because cast stainlesssteels are subject to thermal embrittlement duringlong-term service exposure at LWR operating temper-atures. The components of concern include pumpbodies, reactor coolant piping and fittings, surge lines(in a few plants). pressurizer spray heads, checkvalves, control rod drive mechanism housings, andcontrol rod assembly housings. These are made ofgrade CF-8, CF-A, or CF-M stainless steel in U.S.LWRs; grade CF-3 stainless steel also is used in someforeign LWRs. The purpose of this project was toreview dte available data on thermal embrittlement ofcast stainless steels and to develop updated proceduresfor life assessment of key LWR cast stainless steelcomponents.

Cast stainless steels have a two-phase microstruc-ture consisting of ferrite islands in an austenite marxWith long-term thermal exposure at LWR operatingtempe e, other phases form in the ferrite phasethat cause it to become hard and brittle, while the aus-tenite remains ductl If the amount of ferrite smalland if it is distributed evenly and finely throughout theaustenite, then the properties of the casting are not sig-nificantly affected by the thmal embrittlement of theferrite. However, as the amount of ferrite and thecoarseness of its distribution increase, the thermal em-brittlement of the ferrite increasingly affects the prop-erte of the casting.

The properties most affected by thtermal embrittle-ment are Charpy V-notch (CVN) impact energy andfracture toughness (JIJ. Both of these properties de-crease as the degree of thermal embrittlement in-creases. If these values become too low, the structuralintegrity of a cast stainless steel component could beseriously impaired The available data indicate thatthermal embrittlement does not have a large effect ontensile, fatigue-crack initiation or low-cycle fatigue-crack growth behavior However, more fatigue-crack-growth data are needed for CF-8 and for all caststainless steels in the high-cycle (low growth rate, nearthreshold) regime. Thus, for life assessment of caststainless steel components, the main concern is loss offracture toughness and impact energy. Data and engi-neering models have been developed to help predictthe degree of embrittlement as a function of thermalexposure history. For reactor internals, there also is aconcern that irradiation may embrittle the austeniteand add to the overall emzbriulementof the component.

No information on this potential problem was foundduring this project, but it is an area that merits futurestudy.

The minimum CVN impact energy aft long-termaging has been found to be proportional to the squareof the fraction of ferrite, the mean fernte spacing, anda chemical-composition parameter. This model shouldbe developed further for application to the assessmentof components. A metparameter can beused to define lower-bound trends to the available im-pact energy values for cast stainless steels as a functionof chemical composition and thermal exposure time,In this project we popoase a model that uses that pa-rameter to predict the impact energy decrease for anyparticular lot of cast stainless steel. This predicted im-pact energy value or the predicted minimum impactenergy value is then used to estimate fracture tough-ness from correlations between impact energy andfracture toughness at both room temperature and2900C This proposed approach should provide a con-servative estimate of frture toughness for use in thestructural integrity assessment of cast stainless steelcomponents.

Insnvice inspection (0SI) is needed to define type,size, and location of defects in cast stainless steel com-ponents so that the structural integrity of these compo-nents can be evaluated. Use of radiography during ISIIs less practical than during fabrication and is inef-

ient. In addition, conventional ultrasonic testing (UI)methods are not reliable to detect flaws in cast stainlesssteel components because coarse grains and differentgrain structures in cast stainless steel result in a lowsignal--to-nise ratio. Advanced UT methods with en-hanced signal-to-noise ratios are being developed.These methods have an improved capability to detectflaws in cast stainless steel components and have beenused to inspect statically cast components in severalPWR plat The effect of grain structure on the propa-gation of ultrasonic waves in cast stainless steel mate-rial has been studied to increase the accuracy oflocating defects. Because of the difficultes with radi-ography and UT methods to detect and size flaws,application of the acoustic emission technique to de-tect crack growth in cast stainless steel needs to beevaluated.

Weld repair procedures are employed to removefabrication defects during the original production ofcast stainless steel components. Itshouldbepossibleto

vii

adapt, modify, and qualify those procedures forperforming weld repairs to components that are inservice in order to extend their useful life.

We developed a procedure for estimating the currentcondition and residual life of key LWR cast stainlesssteel components. The procedure is implemented in-nine major steps. The first three steps involve the col-lection, examination, and storage of records for fabri-cation and construction, inservice inspection, andoperating history. Ike fourth step involves a conserva-tive fatigue and fracture mechanics evaluation to de-termine the worst-case flaw size and the minimumreqWred fracture toughness at the end of the next oper-atng period I the fifth step, the current condition ofthe mateial is assessed using either a proposed analyt-ical model, microstuctural daa or measured proper-ties (or using some combination of those threemethods). In the sixth step, the results of the fourth andfiffh steps are combined to evaluate the structural in-tegrity of the component. The seventh step establisheswhat actions (none, repair, replace, or shut down) areto be taken, and the eighth step establishes the plan forthe next ISL From this point, the steps are repeated (thecomponent is reevaluated) as needed.

Future work is recommended to provide informa-tion for use in implementing the life assessment proce-dure. Guidelines of worst-case thermal emnrittlementdata should be developed. Thermal eonbrittlement sur-veillance programs should be initiated. More impactand fracture toughness data should be developed formaterial aged and tested at maximum LWR operatingtemperature. Ihe aging parameter model should beimproved to better account for the observed mecha-nisms of thermal embrittlement. The work on evalua-tion of the relation of farite amount and spacing to thedegree of embrittlement should be continued. High-cycle faigue-ack-growth tests should be performedon aged and irradiated materials, especially CF-8. Theuse of field replication and magnetic measurements forcharactfizin cast stainless steels should be demon-strated and validated. Techniques involving the use ofsubstandard sized specimens and microhardness toquantify the degree of thermal embrittlement shouldbe developed. Repair welding procedures should bequalified and validated. Fracture-nechanics analysesshould be performed for representative combinationsof components and operating histories to estimate theminimum fracture toughness required for safe opera-don. ISI methods and programs should be modified toinco te procedures and evaluations appropriate tocast stainless steel components.

Lx

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

the US. Nuclear RegulatoryCommission has played in providing programmatic guidanceand review for this report

The authors sincerely appreciate the critical review of this report performed by B. LLanderman. The authors thank 0. K. Chopra and H. M. Chung of Argonne NationalLaboratory and M. B. Lapides of Electnc Power Research Institute for several telephoneconversations and meetings related to aging of cast stainless steel components. The authorsalso thank 0. K. Chopra for providing the data from the Charpy impact tests of aged caststainless steel specimens, and thank K E. Lapides for providing a summary of a draft EPRIreport on aging of cast stainless steel components. The authors thank C. W. Rainger of San-dusky Foundary & Machine Company for reviewing the report.

Finally, the authors thank D. R. Pack and D. L Bramwell of the Idaho National Engineer-ing Laboratory for their technical editing assistance and the significant effort they put forthin coordinating the production of this report, and to K. 0. Roberts and B. L. Thompson fortheir special effot in composing the document.

I

I

CONTENTS

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

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

EXECUTIVE SUMMARY viii

1. . . iACNKN 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 NN THTEREMAMLAELMBR]LEM ....... ................. 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 . . .................... 37.... ............... 37

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 offerriteinanaustenitematrix . ........................................................ 8

5. Effect of thermal aging on the room-tempeaue Charpy V-notch impact energy of CF-3stainless steel .................................................................. 13

6. Effect of thermal aging on the room-temperature Charpy V-notch impact energy of CF-Sstainless steel ...................... 13

7. Effect of thermal aging on the room-temperature Charpy V-notch impact energy of CF-SMstaiesasteel ...................... 14

8. Effect of test temperatmr and thermal aging on the Charpy V-notch impact energy of CF-3stainless 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-8stainless 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-SMstainless 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 materialparameter ' for aged cast stainless steels, adapted from Chopra and Chung (1989a), data suppliedby Chopra (1990c) ................................................................. 21

12. Correlation between room-temperature fracture toughness (Ja and impact energy for cast stainlesssteels . .................................................................. 22

13. Correlation between rm pae tearing modulus (l) and fracture toughness (JIc) for caststainlesssteels .................................................................. 22

14. Correlation between the 2900C fractue toughness (Jic) and onom-tempera impact energy forcast stainless steels ........................ 23

15. Correlation between the 290°C tearing modulus (1) and fracture toughness (Jic) for cast stainlesssteels .......... 23

xiD

I

16. Fatigue crack growth rae of CF-SM stainless steel, in air and in a PWR environment, at a cyclicfrequency of 0.01711z . ............................................................. 25-

17. Predicted lowerbomnd Charpy-mpact values for cast stainless steels as a function of aging timeat 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 fetecontentusing lodel 1 of Aubrey etaL(1982) ........................................... 33

20. Comparison of ferrite content measured using a ferritescope with that determined metallographicallybya 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 FORMAJOR LWR COMPONENTS

VOLUME 3: CAST STAINLESS STEELCOMPONENTS

1. INTRODUCTION

All opeatig commercial nuclear power plants inthe United States are light water reactor (LWR) sys-tems; they employ a variety of cast stainless steel com-ponents in the primary pressure boundary. Caststainless steel is used because of its high resistance tohotcracking, corrosion, and strss-corrosion cracking,and because of its improved mechanical properties(Solomon and Devine, Jr., 1983). In addition, use ofcast stainless steel provides benefits of the castingprocess in manufacturing. Complex shapes, such aspump housings, valve bodies, and fittings are staticallycast; usually the molten metal is poured into a fixedsand mold. Cylindrical shapes, such as pipes are cen-trifugally cast; te molten metal is poured into a ro-tating sand-lined metal casing.

Maintaining the structural integrity of thesecomponents is essential to the safe operation of thepower plants. Thus, any mechanism that maydegrade the mechanical properties of the caststainless steels used in these componentsshould be considered in efforts to extend the operatinglife of an LWR. Thermal embrittlement at LWR oper-ating temperatures is the major degradation mecha-nism of concern for cast stainless steel components.Typical operating temperatures in pressurzed water -reactor (PWR) primary coolant piping varies from 288to 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 waterreactors is about 282?C (S40F) (Shah andMacDonald, 1989). For service exposure in this rela-tively low-temperature range, the embrittlemrent oc-curs at a very slow rate over a period of many years.

As aging progresses, and the degree of embrittle-ment increases, the critical flaw size decreases; thus,even initial defects that do not grow, as well as thosethat may grow subeitically, may in fime become de-fects of critical size. During strucural integrity assess-ments of cast stainless steel components, one mustdetermine if the critical defect size will become toosmall to be reliably detected during inservice inspec-

tion (ISI). Finally, leak-before-break &LB) analysisshould be perfcrmed to deternine if a detectable leakwill develop before a growing crack reaches a criticalsize for unstable ppagation.

The concern about possible thermal embrittlementbecomes progressively more important as cast stain-less steel components enter the second half of the40-yr period of a nuclear power plant's original oper-ating license, and as extending the operating licensefor an additional period of 20 years or more is consid-erel The potential problem of thermal embriulementis aggravated by the fact that the current nondestruc-tive examination (NDE) techniques used for wroughtstainless steels usually are not as effective whenapplied to cast stainless steels.

The objectives of this repo't are to review the avail-able information on thermal embrittlement of caststainless steels, discuss methods of assessing thedegree of embrittlement for inservive cast stainlesssteel components, discuss the possible effects of inser-vicerepair welding on thermal embrittlement, and sug-gest possible approaches for evahmuing the remainingsafe life of cast stainless steel components, approachesthat can be used in efforts to extend the operating li-cense of a light water reactor. To provide backgroundfor these discussions, the following sections of this in-toduction briefly describe (a) the LWR cast stainlesssteel components of interest and (b) the physical met-alluy of cast stainless steels.

1.1 LWR Cast Stainless SteelComponents

Table I lists the eight LWR cast stainless steel com-ponents considered in this project The foLlowing para-graphs give brief descriptions of these components.

Reactor Cool=nPw*pBody. Three types of reactorcoolant pump bodies are used in LWR plants. Ues FandE pumps are used in PWRs, nd Type C pumps areused in BWRs. Pump bodies we manufactured usingCF-8, CF-SA, or CF-8M statically cast stainless steelsections that are subsequently welded. The casting istypically done in more dan one piece because of the

I

Table 1. LWR cast stainless steel components

Component

Reactor CoolantPump Body

CoolantPiping

Coolant PipingFittings

SurgeLines

PressurizerSpray Head

Check Valves

cfpeciCasting

Static

Centzifagal

Static

Centrifugal

Grade

CF-8CF-8MCF-8A

CF-SACF-8M

CF-8A

Numberof Planti Comments

Type F in 61 plantsType } in 15 plantsType C in all BWRs

207

27

44% of Westinghouseplants

All Westinghouseplants

CF-8M SomeCombrustionEngieing Plants

WestinghouseplantsStatic

Static CF-8M Some Westinghouse plants

RecirculationPiping Fittings

Control Rod DriveMechanism Housing

Static

Static

CF-8CF-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 weldedtogether to produce the fmal assembly (Shah andMacDonald, 1989). Recent designs of Type F pumpcasings decreased or eliminated the number of weld-ments as it became practical to manufacture largercasting sections. The pump impellers are made ofCFM- cast stainless steeL Figures 1 through 3 show thepump configurations (ASME Code Section III).

Figure 1 shows a Type F pump. Most Type F pumpsare manufactured by Westinghouse. Their wall thck-ness ranges from 100 to 200 mm (4 to 8 in.). Thesepumps are used in Westinghouse and some Babcock &Wilcox (B&W) plants. Type F pumps are used in atotal of 61 nuclear plants. The Bingham-WillametteCompany also makes Type P pumps that are used insome B&W plants. Tlpe F pumps are manufacturedusing an electroslag welding process with no postweldheat reatment the lack of such treatment can leavesome high residual stresses at levels near yieldstrength. (Section m of the ASME Boiler and Pressure

Vessel Code for Class I components neither requiresnor prohibits postweld heat treatments for austeniticstainless steel welds.)

Figure 2 shows a Type E pump. 7ype E pumps aremanufactured by Byron Jackson Pump Division,Borg-Warner Corporation. These pumps have wallthicknesses 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 withmanual welds followed by a full solution heat treat-ment, which eliminates nearly all the residual stressesif the assembly is slowly cooled from the solutiontreating temperature. However, the weldedcastings arequenched from the solution treatment temperatures[-1038 to 11490C (1900 to 2100F)l to avoid sensiti-zation, which can leave high residual stresses (Shahand MacDonald, 1989). Any improvement providedby solution heat treatment is limited because the sec-tions of th pump body are thick. The optimum im-provement will be at and near the surface, because the

2

I

rate of cooling is important. Tb minimize sensitizationand residual streasses In Type E and C pumps (de-scribed in the next paragraph), aI1 welding performed Dischargeafter the solution heat treatment of the material is lim-ited to a low-heat input no greater than 1900 J/cm nozzle(5000 ) Crot) Horizontalf -4* ross

Figure 3 shows a Type C pump. Type C pumps are Section AAmanufactured by Byron Jackson Pump Division.Borg-Warner Corporation, for use in all GE BWR ump shaftplants. These are smaller than the pmnps used in PWRplants. The smaller size accommodates the lowersystem pressures and space limitations of BWR TypeC pumps also are manufactured with manual weldsfollowed by a full solution heat treatment. In most Closurecases, these welded castings also are quenched from studsthe solution treatment temperature, which can leavehigh residual streCSe.Upe

flangeo a e

Pump casing wall

Discharge casingnozzle

Vertical cross section 9-3044

Weldjoint

Figure 2. PWR Type E coolant pump.

Reactor Coolant Piping. he newer Westinghouseplants use centrifugally cast CF-8A or CF-8M stain-less steel for the main coolant piping (Shah andMacDonald, 1987). CE, B&W, and older

Closure Westinghouse plants did not use cast stainless steel forstuds _main coolant piping. Table 2 gives the names, sizes,

operating age in years, and piping materials for the 27Pump _Westinghouse plants with cast stainless steel primaryshaft Ipiping (gan et al., 1987; USNRC, 1988a). The plants

are listed in chronological order of their operating li-Dscha lcense dates. The typical diameter of te coolant piping

is about 0.8 m (32 in.). Almost no problems have beendetected in centrifugally cast piping to date (Egan

Weld Het al., 1987).joint 4 @ aiPump Reactor Coolant Piping Fittings. The main coolantcasin \ 7 piping fittings, such as elbows, are statically cast

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

a larger population and deeper defects are expectedFigure 1. PWRTypeFcoolantpump. in statically cast components than in welds or

3

I

centrifugally cast components This should be takeninto account when selecting sites for inserviceinspections.

steam space of the pressurizer. The cod-leg coolantsprays through the spray head into the pressurizer toreduce the pressure in the primary coolant system InWestinghouse plants, the spray head is made of caststainless steel and is subjected to an opezating tmper-ature of approximately 3430C (6500F), which is thesanution temperature at the PWR operating pressure.

Check Valves. Check valves with cast stainless steelbodies are installed in LWR safety systems havingstainless steel piping. For example, in Westinghouseemergency core cooling systems (ECCS), there arecast stainless steel check valves to prevent flow fromthe primary loop hot leg or cold leg back into theECCS piping (USNRC, 1988b; Greenstreet et al,1985).

Recirculation Piping Fiaings and Valves. The BWRrecirculation piping fittings, such as elbows and tees,and pump suction and discharge valves, are staticallycast CF-8 and CF-8M stainless steel.

Convrol Rod Drive Mechanism Housing. The con-trol rod drive mechanism (CRDM) housing constitutesa primary pressure boundary. In certain Westinghouseplants, some of the CRDM housings are made of CF-Scast stainless steel (Shah and MacDonald, 1989).

CocImnternals. Core internals made of cast stainlesssteel may be subjected to both thermal and irradiationembrittlement during long-tem service. In CE plants,the control rod assembly shrouds are made of centrifu-gally cast CF- stainless steel. In Westinghouse-typeplants, the lower support structures, cruciform instru-ments guides, and flow mixer plates are made of stain-less steel castings. In GE plants, the orificed fuelsupports are made of CF-8, and parts of the jet-pumpassembly are made of CF-8 and CF-3 stainless steelcastings.

An evaluation of the effects of irdation on stain-less steel castings should be made. This report reviewsthe thernal embrittlement of such castings. The ther-mal embrittlement is caused by embrittlement of theferrite phase in the ferrite/hustenite microstructure thatis inherent in these alloys. The austenite phase is duc-tile and remains so after thermal exposure; thus, whenthe fraction of ferrite is small, the degree of overallembrittlement of the casting is low. However, if thecomponents used in internals are subjected to highlevels of irradion, the austet phase could becomeembrittled as well. In such a case, where both the aus-tenite and ferrite phases are embrittled, the overallstructural integrity of the casting could be in question.

4

FIgure 3. BWRIypeCcoolantpump.

SurgeLine. The surge line connects the reactorcool-ant piping hot leg to the pressurizer. As the tempera-ture of the reactor coolant fluctuates, the volumeincreases or decreases, causing water to flow into orout of the pressurizer through the surge line. Typically,the surge line is made of austenitic stainless steel.However, in some Combustion-Engineering plants thesurge line is made of CF-8M stainless steeL This lineis typically 300-mm (1-in.) diameter Schedule 160pipe (Shah and MacDonald, 1989). The pressurizerend of the surge line has operating temperatures ashigh as -3430C, which is the highest operating temper-ature of the major PWR primary coolant pressureboundary components.

Pressurizer Spray Head. T pressurizer spray headis the termination of the spray line and is located in the

I

Table 2. Castprimary coolantpipingmaterials(Ean ct al., 1987; USNRC, 1988a)

PlantNameKewauneePrairie Island 2cook IFarley ATrojanCook2Beaver Valley INorth Anna INorth Anna 2Farley 2MeGuire 1

McGuire 2Callaway 1Sequoyal 1Sequoyah 2Catawba 1Wolf CreekMillstone 3Citawaba 2

Beaver VAlley 2South TeM ISouth lTxas2Vogtle 2ComanchePeak ICommanche Peak2Watts Bar I

UnitSize

(MWe)

560530

109082g

1130

1054852934788829

11801180115711401140

.11531158115011531113852

125012501113115011501177

Numberof System

224

344

333

34 i4444.4444434 ..4

*4..4 -I. 4

,4 . .

OperatngAge(Years)

141412109

999887555544433221

1'00

.0

PipeMaterial

CF-SMCF-8MCF-8MCF-8ACF-8ACF-SMCF-8MCE-SM

CF-8ACF-8A

CF-SACF-8ACF-8MCF-8MCF-8ACF-8A

CF--ACF-8A

CF-8ACF-8ACF-SACF-8ACF-8ACP-8ACF-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 sincethe operating license dates.

5

1.2 PhysIcal Metallurgy of CastStainless Steels

The required chemical compositions and the re-quired tensile properties of six common grades of caststainless steel are listed in Tables 3 and 4, respectively.TheomnposidonsofCF-3,CF-8,CF-3MandCF-SMare similar to those of the wrought grades Type 304L,Type 304, Type 316L, and Type 316, respectively. Thewrought grades have microstructures that are usuallyfully austenitic, whereas the cast grades typically havemicrtru resconsistingofabout5 to 25% (volume)ferrite and the balance austenite (Peckner andBernstein, 1977). (The ferrite phase in cast stainlesssteels and stainless sted welds is sometimes referred toas the delta ferrite because it forms at high tempera-tures, where the ferrite field of the phase diagram isknown as the delta phase to distinguish it from the low-temperature ferrite phase known as the alpha phase.)Theprmaryfacorcontrollingtheferrite-austenitebal-ance is bulk chemical composition. Chromium, silicon,and molybdenum, and niobium (if present) promote theformation of ferrite, whereas nickel carbon, manga-nese, and nitrogen promote the formation of austenite(Peckner and Bernstein, 1977). The nitrogen is intro-duced during the casting process, but its percentage isnot specified. The typical values of nitrogen in thechemical composition of cast stainless steels may varyfrom 0.03 to 0.08%. Empirical relationships have beendeveloped forestimating the faerite content from chem-ical composition; the use of two such relationships areillustrated later in this report (Aubrey et al., 1982;Bonnet et al, 1990).

The size, distribution, and morphology of the ferritewithin the austenite depends on solidification condi-tions during the casting process. The conditions affect-ing the solidification of metal are complex (Reed-Hill,1973) and a detailed discussion of them is beyond thescope of this reporL Cast stainless steels may solidifywith a columnar or equiaxed grain structure or a mix-ture of both structures. Researchers at ArgonneNational Laboratory (ANL) have observed that insteels of similar chemical compositions, the averagegrain size and spacing between ferrite islands gener-ally tends to increase with larger section sizes and cor-responding slower heat removal rates.' In heavy-wallcastings, there can be considerable variations in micro-structure across the wall of the casting. The complexnature of metal solidification during casting makes itdifficult to predict the finer details of the actual micro-structure 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 ml-crostructural details Is important for the assessment ofcasting integrity and service life, some means of mea-suring them must be employed.

ncreased ferrite content increases the yield and ten-sile strength of the cast stainless steels, so gradesCP3A and CF-8A have controlled ferrite-austeniteratios to increase the minimum strength levels(Peckner and Bernstein, 1977). The minimum strengthlevels are indicated in Table 4. These controlled-ferrite grades have the same chemical requirements asCF-3 and CF-8; the ferrite is controlled by adjustingthe cheidcal composition within the imits shown inTable 3. Experience has shown that castings madefrom the alloys listed in Table 3 may contain up to 30%ferrite when control of the amount of ferrite is not spe-dcfib Although ferrite levels near 30% are not usualfor castings made in accordance with ASTM specifica-dons (e.g., A-743 and A-744), they are possible.

As shown in Tables 1 and 2, LWR components havebeen made from CF-8, CF-8A, and CF-8M stainlesssteeL CF-8 Is used for its genral resistance to oxidiz-ing conditions. Compared with CF-8, the CF-8Malloy has improved resistance to reducing environ-ments as well as high resistance to oxidizing condi-dons. Therefore, the CF-8M alloy is widely used forcorrosion-resistant pumps and valves. CF-8A is usedto provide mechanical strength properties higher thanthose of the CF-8 alloy. The low-carbon versions(CF-3, CF-3A, and CF-3M) may be used in the futureto provide an added margin of resistance to intergran-ular stess-corrosion cracking (IGSCC), especially inBWRs, where many IGSCC problems in recirculationpiping have been encountered in the past (Shah andMacDonald, 1987).

Examples of austenitic-ferritic microstructures ob-served in cast stainless steel are shown in Figure 4.cThese samples were obtained from centrifugal cast-ings. The ferrite phase (darkened networks and/orislands) is contained within a matrix of the austenitephase. 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 ferriteregions is significantly different from that of the aus-nite regions in these alloys. 7h illustrate this point, thelocal 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 con-ducted 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

CFLSACFL8A

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 VesselCode.

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 485MPaU(si) (70) (77) (70) (77) (70) (70)

0.2% offset yield 205 240 205 240 205 205strengthMPa(ksi) (30) (35) (30) (35) (30) (30)

Elongation in 50mm 35.0 35.0 35.0 35.0 30.0 30.0orsn2in.

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

7

p2'~~

- '

61oo90o

6M-2

6M203Glyceregla etch

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 ferritein an austenite matrix.

8

Figure 4 were analyzed using an electron microprobe;the results ar presented in Table 5. Tbe overall bulkcomposition determined by the foundry is comparedwith the local composition within the ferrite and theaustenite regions. Bulk values for silicon, manganese,sulfur, and phosphorus are not shown in Table 5because local values were not measured for these fourelements. The local values were obtained from theaverage of six individual determinations at randomlyselected regions within each particular phase. As ex-pected, the ferrite phase in both samples containedhigher chromium and molybdenum levels and lowernickel levels than the bulk material.

1.3 Benefits of Ferrite In CastStainless Steel

Ferrite in cast stainless steel provides resistance tosensitization, which is caused by precipitation ofchromium-rich carbides MA23C) and the resulting de-pletion of chromium at het grain boundary. Sensitiza-tion describes the susceptibility of a stainless steel tointergranular stress-corrosion cracking (1GSCC) fol-lowing certain heat treatments such as a welding pro-cess or a slow cooling from the solution annealingtreatnent. Duing welding, temperatures are reached

in the heat-affected zone that cause precipitation ofchromium carbides. In cast stainless steel, the chro-mium carbides preferentially precipitate at the ferrite-austenite interfaces, not at the austentite grainboundaries. The faster diffusion rate of chromium inthe ferrite phase results In formation of carbides on theferrite side of the interface. A chromium-depletedzone, which forms on the ferrite side of the interface, isquickly replenished with chromium because of the rel-atively high chromium content of the ferrite phase (seeTable 5) and the high diffusivity of chromium in fer-rite. Industrial experience indicates that componentswith high ferrite levels have more resistance to ISCC(Peckner and Bernstein, 1977)2 than do componentswith low ferrite levels. For example, many duplex (40to 60%) stainless steels have been found to be highlyresistant to IGSCC in aqueous environments in nonnu-clear applications (Peckner and Bernstein, 1977).

The minimum amount of ferrite needed to avoidstress-coroion cracking susceptibility of cast stain-less steel In an LWR environment depends on the car-bon 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

0.053CF-8M Overall'

Austenite

Ferrite

2.3 _b

4.0 _b

a. Overall composition was determined by ladle analysis asrepcrtedby foundry; composition ofausteniteandfer-rite 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 environmentcast stainless steel with a maximum carbon content of0.035% requires a minimum ferrite content of 7.5% forresistance to sensitization during welding (Hazeton,1988). Results of laboratory tests on CP-3 and CP-8materials in an accelerated BWR environment showthat IGSCC is not expected with a ferrite contentgreater than 12% (Hughes et al., 1982). Inadditionthetest results show that no IGSCC effects are expectedwith a maximum carbon content of 0.03% in weld sen-sitized material and of 0.015% in furnace-sensitizedmaterial, regardless of ferrite content. Similar evalua-tions for CF-8M material are needed. The test resultsalso show that in addition to the amount of ferrite, dis-tribution of ferrite or ferrite spacing plays an importantroe in determining stress-corrosion cracking suscepti-bility ofcast stainless steels. A finer distribution of fer-rite provides better resistance to stress-corrosioncracking.

worst case, this is likely to result in a local leak ratherthan in any substantial loss of system integrity, becauseof the heterogeneous nature of the matel. The integ-rity of weldments is important because welding is usedas a joining procedure during abication, a method ofrepairing casting defects during fricatio an instal-ladon procedure, and a method of repairing compo-nents for the purpose of extending their operationallifetimes.

The presence of some ferrite (typically at least 5%)minimizes the occurrence of hot cracking or microfis-suring in castings and weld metal.b In several olderplants, the weld metal in the pump casing and in stain-less steel RCS piping have less than 3% ferrite,whereas in newer plants a minimum of 5% ferrite is re-qired for the reactor coolant pressure boundary com-ponents (USNRC, 1978).

In BWR plants, stress corrsion cracking of the sen-sitized heat-effected zones of cast stainless steel ispossible and has been observed.$ However, in the

a. Private communication with NL E. Lapides,Electric Power Research Istitute, Palo Alto,Cafoia, 1990.

b. PrivatecommunicationwithE.Landermancon-sultant, Pittsburgh, Pennsylvania, September 1988.

10

2. REVIEW OF INFORMATION ON THERMAL EMBRITTLEMENT

Understanding the basic mechanisms that causethernal embrittlement damage to stainless steel corn-ponents is necessary for identifying the aging parame-ters related to design, fabrication, and operation ofthese components. The degree of material embrittle-meat in aged cast stainless steel components is mostoften evaluated by Charpy V-notch (CVN) impact en-ergy tests and, recently, by fracture toughness tests.Tensile and fatigue prp es for the aged componentsalso are needed in evaluations of the integrity of thecast stainless steel components. This section reviewsthe basic mechanisms causing thermal embrittlementdamage, evaluates the currently available Charpy im-pact energy and fracture toughness data, and thenbriefly discusses the low-cycle fatigue behavior of thecast stainless components.

2.1 Basic Mechanisms

The mechanisms of thermal aging that cause em-brittlement in cast stainless steels have been reviewedand evaluated extensively by Chopra and Chung(1985, 1986a, 1986b, 1986c, 1987, 1990a, 1990b),Chung and Chopra (1987), Chung (1989), Chopra andSather (1990), Chopra (1990), Auger et al. (1990),Bonnet et al. (1990), and Brown et al. (1990). Em-brialement or loss of toughness in cast stainless steelsduring elevated-temperature exposure is related (a) tothe precipitation of carbides at the austenite-ferritephase boundaries and (b) to the formation of theCr-rich alpha-prime phase and the Ni-rich andSi-rich 0 phase in the ferrite. The phase-boundarycarbidesplayasignificant role in embrittlement for ex-posures at temperatures greater than 4000C (750M3,but have kss effect on the embrittlement at exposuretemperatures less than 4000C.

For LWR orating temperatures tless than 3500C(660)], the formation of the alpha-prime phase andthe 0 phase in the ferrite are the primary factors in-volved in embrittlement. Also, the kinetics of the for-mation of these phases appear to be different attemperatures ess than 4000C (7500P) than at tempera-tures greater than 4000C. Because of these differencesin formation and precipitation behavior, the results oftests an material subjected to accelerated aging at tem-peratures greater than 4000C should not be extmpo-lated to the lower LWR operating temperatures.Bamford et al. (1987) of Westinghouse describe the re-suits of analyses performed using x-ray energy disper-sive spectroscopy. These results indicate that data from

accelerated aging studies at temperatures greater than4000C (7500F) could be used to predict behavior at op-crating temperatures. Additional, independent re-search is needed to resolve the differences between theANL and Westinghouse studies.

The alpha-prime phase typically forms by theprocess of spinodal decomposition (Chopra andChung, 1987, Sassen et al, 1987). Spinodal decompo-sition refers to the reaction whereby two phases of thesame crystal lattice type, but different compositionsand properties, form because a miscibility gap exists inthe alloy system. In the iron-chromium system, theseimmiscible phases are known as the iron-rich alphaphase and the chromium-rich alpha-prime phase. Thisphase separation process occurs at a very fine scale (onthe order of only a few nanometers) in the ferrite re-gions of cast stainless steel, and use of the atom probefield Ion microscope Is required to resolve the presenceof the alpa-prime, phase (Sassen et al, 1987). Thereare indications that after extensive (many years) agingat LWR operating temperatures the alpha-prime phasecan form also by means of a nucleation and growthprocess, as well as by spinodal decomposition (Chopraand Chung, 1987). Depending on the composition ofthe ferrite and the exposure temperate, either or bothof these processes may be involved in the formation ofalpha-prinmephase.

The 0 phase forms in te ferrite by a nucleation andgrowth process, and its rate of formation increaseswith increased levels of carbon and molybdenum(Chopra and Chung, 1987). Butthe 0 phase has no di-rect effect on the degree of embrittlement, as demon-strated by the following laboratory test. A CF-8stainless steel pump cover was found to be embrittledafter 12 years of service in aBWR (Chopra and Chung,1987). Annealing for one hour at 5500C (10200F)dissolved the alpha-prime phase and restored theCharpy-impact resistance to the level expected forunaged material, but had no effect on the 0 phasepresent. Thus, the G phase had no significant effect onthe degree of embrittlement, and the alpha-primeembrittlement was easily reversed by a short heat treat-ment (one hour) at a moderate temperature [5500C(10203].

The annealing treatment described above wasapplied to verify that embrittlement was caused by thealpha prime. It is not proposed as a method for revers-ing the effects of alpha-prime embrittlement in actualcomponents because of the practical difficulties that

11

would be encountered in applying such a heat treat-ment in the field.

Th pesece of G phase may indirectly affect thedegree of embittlement in cast stainless steel compo-nents; however, its role is not yet well-efined. Itspresence apparently mitigates the degree of emnrittle-ment caused by the alpha-prime phase by slowingdown its formation. On the other hand, the G phase(and perhaps the alpha-prime phase) precipitates morereadily at dislocations (Chopra and Chung, 1987).Thus, cold work or stressing during service, includingfatigue cycling, may increase the degree of embrittle-ment; no data on this potential effect were found in thepresent projecL

Because only the ferrite phase is embrittledby long-term service atLWR operating temperatures, the over-all thermal embrittlement of a cast stainless steelproduct depends on the amount and morphology of theferrite presenL For LWR applications, a current guide-line is that low-temperature embrittlement is a majorconcerm only when the volume fraction of ferrite ex-ceeds a level of approximately 15% (Copeland andGiannuzzi, 1984). The reasoning behind this guidelineis that the ferrite phase tends to form in isolated poolscontained within the austenite when ferrite levels areless than or equal to 15%. In this case, the overalltoughness of the stainless steel casting is not greatlyaffected even if the ferrite were embrittleA However,where ferrite levels are greater than 15%, there is agreater tendency for a continuous path of embrittledmaterial to exist through the thickness of the cast com-ponent, which would greatly reduce its toughness if theferrite regions were embrittled (Chopra and Chung,1986b).

There is recent evidence that thick-walled (typicallygreater than 100 mm) stainless steel castings with fer-rite levels in the range of 10 to 15% also maybe subject to low-temperature embrittlement (Chopraand Chung, 1986b). In heavy-section castings, thegrain size tends to be large and the ferrite spacing(average distance between ferrite islands) is increased.With increasing ferrite spacing at a constant ferritecontent, the tendency for a continuous path of ferrite toexist through the thickness of the cast component in-creases. Bonnet et al. (1990) selectively dissolved theaustentite phase from samples of CF-8M and foundthat the ferrite phase remained continuous at ferritevolume fractions as low as 5%. This result indicatesthat a relatively small amount of ferrite may provide acontinuous path for fracture in aged cast stainlesssteels.

Two different sets of data have recently beenpresented regarding the thermal embrittlement ofstainless steel weldments. One data set indicates thatthe fracture toughness of Type 308 SS weldments wasnot affected by aging at 4270C (800F) for 10,000 h(Mills, 1987). However, the other data set indicatesthat the Charpy V-notch impact energy of Type 308 SSweldments with 8% ferrite may be significantly re-duced by aging at 3430C for 10,000 h: however, corre-sponding fracture toughness data are not presented.More research is needed to ensure that the fracturetoughness of the weldments for cast stainless steelcomponents is not significantly affected by thermalaging at the LWR operating temperatures.

Embrittlement of welds, ifconfirmedby tresuts,could become a major problem because welds gener-ally have lower initial fracture toughness before ther-mal exposure than base metal. Weldments in stainlesssteel castings may contain high residual stresses thatapproach the yield strength (Egan et aL. 1987) becauseSection HI of the ASME Boiler and Pressure VesselCode neither requires nor prohibits postweld heattreatment of these welds. The usual industry practice isto not subject welds in cast stainless steel componentsto postweld heat treatments, because such a heat treat-ment might sensitize base material near the end of theheat-affected zone and make that region susceptible toIGSCC.

2.2 Charpy Impact Energy Data

The degree of low-temperature embrittlement ismost often quantified by the room-temperatureCharpy impact energy after aging at temperatures inthe range of 300 to 4500C (570 to 840MF). Increasedaging temperature is employed to accelerate the rate ofthermal embrittlement compared with that whichoccurs at normal LWR operating temperatures near2880C (550). Charpy V-notch impact energy data,obtained from Chopra (1990), for thermally agedCF-3, CF-8 and CF-SM stainless steel are shown inFigures 5, 6, and 7, respectively. In each figure, theCharpy impact energy is plotted as a function of theaging parameter P. which is defined by the followingrelation:

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

where t is the time, Q is the activation energy, R is thegas constant, and T is the absolute temperature P is de-fined such that it is equivalent to the logarithm (base10) of the number of hours of aging at 4000C (750F).For example, a value of P 4 is equivalent to aging for10,000 h at 4000C.

12

I

104

cg

>, 10ena)cw

E

1020 1 2 3 4 5

Aging Parameter (P)

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

104

UJ

cqa-

10S

0102

0 1 2 . - 3 4 5

Aging Parameter (P)

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

13

I

104

E

(:3

CwUcoa.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 Heatso 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)

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

From multiple correlation analysis of data on theaging of cast stainless steels, the activation energy wasfound to be related to composition by the followingapproximate 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 Chungin their earlier work.

Recently, Chopra and Chur., (1989a and 1989b)have developed new correlations for predicting theactivation energy. They found the following separatecorrelations 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) in-clude additional terms for nickel, manganese, and ni-trogen as well as those for chromium, molybdenum,and silicon. The contribution of molybdenum is posi-tive in Equations (2) and (3) but negative in Equa-tion (4). Also, the contribution of manganese ispositive in Equation (3) but negative in Equation (4).

Chopra and Chung (1989a and 1989b) argue thatEquation (4) is the most appropriate to use in assessing

14

the potential for thermal embrittlement of cast stain-less steel components that are used in U.S. nuclearplants because the compositions included in theFramatome-ANL set of data are more representativeof those used in such components. For this reason,Equation (4) was used, along with Equation (1), incomputing the values of P plotted in Figures 5through 7. The values of Q and P were calculatedfrom the data supplied by Chopra and Sather (1990).Although Equation (4) provides an improvement overEquation (2), there is still room for improvement inthis analytical approach, as is evidenced by the widescatter of the data in Figures 5 through 7. Further-more, Chopra and Sather (1990) reported that the cor-relation represented by Equation (4) was developedfor observed values of Q between 75 and 240 kJAnote.Thus, it should be used only for cases where the com-puted value of Q falls within this range.

The compositions included in the Swiss set of datafall in a relatively narrow range that does not representthose used in cast stainless steel castings produced inthe U.S. Equation (1) and either Equation (2), (3). or(4) should be used to estimate the degree of aging of aspecific heat of material during service. However,because the reasons for the differences among Equa-tions (2), (3), and (4) are not completely understood,additional research concerning the factors contributingto the activation energy for thermal aging of cast stain-less steels is needed.

The recent work by Bonnet et aL (1990) even qucs-tions the validity of using an empirical value of Q cal-culated from the chemical composition. They believethat Q is a function of the aging temperature and ap-proaches a value of approximately 250 kfinole at300 to 325°C, the region of primary interest for LWRapplications. If this belief is trues it implies that valuesof Q calculated using Equation (4) provide overlypessimistic predictions of the aging kinetics for caststainless steels. Unfortunately, their conclusion isbased on the evaluation of data for only two beats ofCF-8M stainless steeL Additional, extensive evalua-

mons would be required before their approach could beapplied to the assessment of aging in LWR cast stain-less steel components

Figures 5 through 7 show lowerbound approxima-tions to the data trends for each alloy. These approxi-mations were developed as part of the work in thecurentprqect. The impact energy decreases as a func-tion of P for values of P less than about 3.5 to 4. Forvalues of P greater than about 3.5 to 4, the impactenergy appears to show no further decrease as a func-tion 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 approxima-tion, the lower-bound CVN-vs.-P lines were assumedto saturate at the CVN value where P was equal to 4 (asshown in Figures 5 through 7).

All of the CVN data for CF-3M fell above thelower-bound approximation, as is shown in Figure S.Some of the data for the KRB plate' fell below thelower4ound approximation for CF-8 (see diamondsymbols in Figure 6); these data were discounted in de-veloping the lower-bound approximation because theyare from European material rather than from U.S. ma-terial. One of the data points for the commercial heats(see square symbol in Figure 6) fell slightly below thelower-bound approximation because it was consideredto be an outlier. Also, for the same reason, one of thedata points for the large experimental heats (see trian-gular symbols in Figure 7) fell below the lowr-boundapproximation for CF-SM The lower-bound approxi-mations 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, impliesthat the maximum degree of thermal embrittlement iscontrolled by factors other than just the chemicalcomposition.

These lower bounds are appropriate for LWR appli-cations because operating temperatures less than400WC are expected and the degree of embrittlement atthese lower temperatures Is not expected to exceed thatobserved at 400C. It would be useful, in future work,to verify this assumption by evaluating the impact be-havior of materials that have been exposed to tempera-turesof 300rC (5700 ) or less for many years. Becausea greater degree of embrittlement is expected from ex-posures at 4000C (7500F) than from exposures at350°C or less, using lower bounds to all of the datashould provide conservative, but overly pessimistic,predictions of thermal embrittlement. Additional ther-mal embrittlement data for the values of aging param-eterP greater than4 are needed to verify that the lowerbound curve levels off, especially at temperatures of350C or less.

The failur of P to fully correlate the changes inimpact energy as a function of aging time andtemperature indicates that the simple expression ofEquation (4) does not accurately model the activationenergy for the complex process of thermal ermbrittle-tnent, which involves the formation of carbide, alpha-prime, and 0-phase precipitates and their interactionsduring aging at temperatures in the range of 300 to450°C (570 to 8405;) (Chopra and Chung, 1986c).

a. From KRB reactor at Gundremmingen, WestGermany.

15

However, until a more accurate model can be devel-opedq the lower bounds shown in Figures 5 through 7provide guidance for LWR applications and indicatean approach that should be developed further.

Most of the data that have been developed to assesthe effects of thermal embrittlement of cast stainlesssteels are room-temperature impact properties, such asthose shown in Figures 5 through 7. For application toLWR components, properties at operatng temperatureare needed Also, fracture toughness and tearing mo-dulus data, rather than Impact energy data, are neededin fracture-mechanics assessments of the safety ofcomponents. Both of these issues are being addressedin 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 impactdata for CF-3, CF-8, and CF-8M, respectively, as afunction of test temperature. Except for Figure Sc, theopen circles represent data for unaged material, thefilled triangles represent data for material aged ap-proximately 10,000 hours at 3500C, and the filledsquares represent data for material aged approximately10,000 hours at 400C. In Figure Sc, the ciles repre-sent data for unaged material, the triangles representdata for material aged 9,980 hours at 350CC, and the

squares represent data for material aged 30,000 hoursat 350C (Figure &c contains no data for material agedat4000 C). TMe filled symbols in Figure 8c are for spec-imens with a L-C orientation (longitudinal specimenwith circumferential crack), and the open symbols inFigure Be are for specimens with a C-L orientation(circumferential specimen with longitudinal crack).

Aging typically reduced the impact resistance of thematerials at all test temperatures, except for the cen-tifugally cast pipe (see Figure 8c), where the aged ma-terial had the same upper-shelf impact resistance asthe unaged material. The centrifugally cast pipe alsohad significantly higher impact resistance than thestatically cast materials, and within the normnd scatterof the data, the specimen orientation aw-C versus C-L)had no sigpificant effect on its impact resistance. Theimpact resistance of the materials aged at 400C wasalways slightly less than that of the materials aged at350CC. For the aged material the room-temperature(25°C) data always provided a conservative estimateof the impact resistance at typical LWR operating tem-peratures (up to approximately 290C. Thus, when nohigh-teupeature Charpy data are available, the lowerbound to the room-temperature Charpy data is likelyto provide a conservative estimate of the lower boundto the high-temperature Charpy data.

5000

4000Cd5

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 stainlesssteel, data from Chopra and Sather (1990).

16

3000

2500

0 Unaged * Aged 9,980 hr at 350°CUnaged (shroud material) a Aged 9,980 hr at 4000C

_o 6>

EU

0.

E

2000

1500

1000

.U.

Heat I of CF-317.1% ferrite

500

0-200 -100 0 100

Test Temperature. °C200 300

b. Cast pump impeller vanes, Heat 1.

8000

7000Heat P2 ol CF-3

15.6% ferriteAged at 3500C

I* Unaged, L-C Orientationo Unaged. CL OrientationA Aged 9,980 hr, L-C Orientationa Aged 9,980 hr. C-L Orientationa Aged 30,000 hr. L-C Orientationb Aged 30,000 hr. C-L Orientation

cml

'U0.

6000

005000

4000

3000

2000

A LY.

an o0* P. I 0

1000 a

I a InL 1.I

- -200 -100 0 100Test 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 stainlessstel, data from Chopra and Sather (1990). (Continued.)

17

5000

4000(UE

0

Cu0.

E

3000

2000

1000

-200 -100 0 100 200

Test Temperature. OC300

Cast 76-mm slab.

Figure 9. Effect of test temperatre and tiermal aging on the Charpy V-notch impact energy of CF-8 stainlesssteel, 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 sinlesssted data from Chopra and Sather (1990).

18

I

(.4E

CA(Ia.

E

5000 -

4000

3000

2000

1000

0- 200 -100 0 . 100 200

Test Temperature, C0 -

300

b. Cast 76-mm slab, Heat 74.

5000

4000C,E

8-

0

C

Caa.E

3000

2000

1000

0 .e=

-200 -100 0 100 . 200 300Test 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 stainlessstel data from Chopra and Sather (1990). (Continued.)

19

I

The curves shown in Figures 8 through 10 are bestfits to the data obtained by means of locally weightedleast-squares regression, using weighting factors(degree of data smoothing) of 66 to 75 percent Thereare no explicit expressions for these curves; they wereobtained to graphically show the trends of the datarather than to provide mathematical expressions forcomputing CVN values In general, the transition re-gion of these curves shifted to higher temperatures andthe upper-shelf energy (USE) decreased with in-creased aging temperature r timem In many cases, theUSE reached a peak at some temperature less than300°C and decreased at temperatures above that peak.

In contrast to the preceding curve fitting approach,Chopra and Chung (1990a) used a hyperbolic tangentfunction of the following form to describe Charpytransition curves

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

where Ko is the lower-shelf energy, T is the test tem-perature, B is the half-distance between the upper-and lower-shelf energy, C is the mid-shelf transitiontemperature in °C, and D is the half-width of the tran-sition region. This approach has the advantage of pro-viding an explicit description of the Charpy transitioncurves, but it does not completely model the observedtrends of tie experimental result. For this reason, theapproach was not employed in the current study.

In their recent work Chopra and Chung (1989a and1989b) have proposed the following parameter for cor-relating the minimum Charpy V-notch impact energyafter long-term aging with the composition and micro-structure of cast stainless steel,

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

+ %Si) L (10 (5)

where 8m is the measured volume fiaction of ferrite inpercent and L is the mean ferrite spacing in pm. Theircorrelation between the parameter 0 and the mini-mum room-temperature Charpy V-notch impactenergy is shown in Figure 11. The power-law equationshown inFigure 11,CVN - 6,270 0<- ,wasob-tained by least-squares regression analysis of the datain 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. Additionalwork should be performed to evaluate the correlation

of this parameter with the minimum high-temperatureChaWyV-otchimpactenegydataandwith fracturo-toughness data at both room and elevated tempera-tures. The 0 parameter could be used, with either(a) measured values of the ferite fraction and spacingor (b) a calculated value of the ferrite fraction and anestimated upper-bound (worstcase) value of the fer-rite spacing, to estimate the minimum room-temperature Charpy impact energy expected afterlong-tem aging of a cast stainless steeL In either case,it would be necessary to know the composition of thestainless steel, including the nitrogen content

Chopra and Chung (1990a) have recommended thata worst-case value for the ferrite spacing of 180 pim beused when a measured value is not available. For ex-ample, 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 ameasured ferrite content of 15%, the value of O iscomputed to be 69.5 for a ferrite spacing of180um. 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-impactenergy expected after long-term aging.

If it were decided that the CVN value based on thewors-case ferrite spacing were lower than desred, theactual ferrite spacing could be measured and used inthe cowrelation. For example, if the ferite spacing forthis 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-impactenergy expected after lng-tm aging.

This 10-parameter approach provides a simple,straightforward method of estimating the minimumroom-temperature Charpy-impact energy expectedfor a cast stainless steel after long-term service.Chopra and Chung (1990a) also have shown that the0 parameter can be used to provide improved corre-lations for aging kinetics. They found improved core-lations between impact energy and the aging parameterP when the materials were grouped by the value of Orather than by alloy type, as in Figures 5,6, and 7. Thisapproach should be evaluated further.

20

I

04 2,000

i. 1,500

P

V 1,000

a

j 500E

0

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

o CF-3, ANL a CF-SM, GFo CF-8, ANL A CF-8M, FRA* CF-SM, ANL * CF-8, EPRI _w a CF-3, GF & CF-3. CEGBr CF-B, GF -All Data _

0< 9 CVN . 6,270 40.6112

I u,. I II 11 uI, IsI II gI IIl I I III------------------0 50 0. l0 150 200 250 300

Material Parameter (0) m 8m2 (C + 0.4O(Cr + Mo + SQ(0.001) L

Figure 11. Correlation between minimum room-temperature Charpy V-notch impact energy and materialparameter 0 for aged cast stainless steels, adapted from Chopra and Chung (1989b), data supplied by Chopra inJune 1990.

2.3 Fracture Toughness Data

The fracture toughness of cast stainless steels hasbeen evaluated by means of J-R curve tests (Bamfordet al., 1985; Chopra and Chung, 1987; Hiser, 1987).'Most of the earlier testing (Bamford et al., 1985). wasperformed at room temperature, but some of the recenttesting (Chopra and Chung, 1987; Hiser, 1987) hasbeen performed at 290°C (555'F). Those invesdga-tions included analyses of test rsults to determine val-ues of JIC. which is a measure of fracture toughness atthe initiaton of crack growth in metallic materials, andtearing modulus, T, which is proportional to the deriv-ative of facture tougness I with respect to crack size,ascoveredby the testproceduresdescribedinASTMDesignadon E 813.

Figures 12 through 15 show fractutoughness datacorrelations for cast stainless steels, as taken from thework of Chopra and Chung (1990a). The data shown inthese 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), ElectricPower Research nstitute (PRI), Battelle (BA!), andKRB. Figure 12 shows the correlation between room-

a. 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 room-temperature impact energy. Ihe solid curve representthe mean trend of the data, whereas the chain-dashedline represents the lower bound to the data In perform-ing a structural integrity evaluation, the lower-bound'relationship can be used to estimate the room-temperature fracture toughness from an estimated ormeasured value of the room-temperature impactenergy for a specific component. Then, the estimatedroom-mperature fractue toughness can be used toestimate a lower-bound value of the room-temperature tearing modulus using the dashed line inFigure 13. As is shown in Figure 14, the lower-bound fracture toughness at room-temperure corre-lation from Figure 12 also provides a conservativeestimate of the fracture toughness at 2900C from avalue of room-temperature Charpy impact energy.Finally, the tearing modulus at 290OC can be estimatedfrom the lower-bound correlaon (dashed line) shownin Figure 15. If fracture toughness could be obtainedfor aged material from the component being evaluatedin a structual integrity assessment, it could be used inplace of values estimated using the prding scheme.

-he value of fiacture toughness at a temperature be-tween the room temperature and 290°C can be esti-mated by linearly interpolating between the datapresented in Figures 12 and 14; similarly, the value oftearing modulus can be obtained from the datapresented in Figures 13 and 15. The values of fracture

21

I

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

1.Y60

z

2

a

ztC

Ul

IL

.Sa0

1000 1500 2000 2500 3000IMPACT ENERGY, KCV (kjIm2 )

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

DEFORMATION, J1I (In.-Ib/In.2 )0 2000 4000 6000 s000 10000

500 - .00

TESTED AT ROOM TEMP.700- AGING TEMP. (C) 700

3 284SYMBOLS9 4OPEN: CF-S , Cf-3 a 350

000- CLOSED: CF-Sm 0 400 800

0 0 427S OO- 0 0450 500

V o 0 UNAGEDcc 0w 400_ 0 V RCANNEALED 400

>0 00

300 o02 0 7300

I-. o0 0 Z~*-

200 00 . o,

100- 1 ° __00

0 500 1000 1500 2000

DEFORMATION JIC (kJ/m2 )

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

22

I

0.U-_ I

RT IMPACT ENERGY, KCV (fttlb)Go 100 160 200

I . . . .I . . . . . . . . .250

ZIIDN n - - * . . . . . . . . . . . . . .

'4

10a040ZI-

0

----

1500-

1000-

.:o

CAST STAINLESS STEL

SYMBOLSOPEN: CF-8 , CF-3CLOSED: CF-SM

AGINGTEMP.(8C) ANL WH FRA

25 0 O350 A4000 o V427-, , 01450 01284

EPRI KRB

m

-'Gao.

a.0*aoooo7C0.

.0. C!.

eo000 Ca.04

'I?

-- a

.I 0

.-A-

00

,-sLOWER BOUND RT Jic

ir

-2000

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

0. . O 1000 -1S 0 20000. 2500 3000 3500 4000RT IMPACT ENERGY, KCV (kjhn2 )

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

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

I1 a I a1

- -- I10000

HnD S .

0

cctu

CSI-

700-

800-Goo-

aoo600

400

3000

200

100

TESTED AT 290-320°C

SYMBOLSOPEN: CF-8 , CF-3CLOSED: CF-GM

0

6A 0

a-

0 e

*O. '

. 0 0 00 0 -a O

L A -A - -

0 *, we

- AGING TEMP. ('C)E 284

A 350

0 400

0 4270 4500 UNAGED

-800

-700

-600

-500

-400

-300

-200

-100

- - - - - u - -

o l 1ooo .50 0 200

DEFORMATION J1c (kJ/m2 )

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

23

I

toughness and tearing modulus at tempeatures lowerthan 290 0C are needed for the pressurized thermalshock analysis of cast stainless steel components

The coteations shown in Figures 12 and 14 maynot be valid for higher values of lIc. because it is lkelythat many of the indicated JIG values do not satisfy thevalidity criteria of ASTM B 813, which requires thatthe thickness, B, and the uncrace ligament, b, of thetest specimen be greater than 25 JQ*. JQ is the condi-tional value of fracture toughness, and sf is the flowstress, which Is the average of the ultimate tensilestrength and het 0.2% offset tensile yield strength Ifthese size conditions, plus the regression-line slopecondition (see Paragraph 9.2 of ASTM B 813), aresatisfied, then JQ equals JI_

Typical tensile properties (Chopra and Chung,1987) of cast stainless steels that were aged for 9,980 hat 350 0C (660 0F) and tested at room tempature andat 288 0C (550 0F) yield sf= 440 MPa at room tempera-tu (Ri) and 300 MPa at 288-C The data shown inFigures 12 and 14 were developed from tests of ITcompact-type specimens, for which typically B = b =25.4 mm. From this information, it is possible to esti-mate the maximum value of JQ for which valid lievalues 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 than447 and 305 WO 2 atRT and 288oC (550o , respecctively, may not sat the validity criteria of ASTM B813. Additional work should be performed to evaluatethe validity of the JIC values plotted in Figures 12and 14.

Based on fracture-mechanics evaluations of theircast stainless steel components, Framatome recom-mends that a lower-bound value of Jk equal to100 kihn 2 be used for assessing those componentThis lower-bound value Is specific to the cases ana-lyzed by Framatome. Comparable lower-bound JICvalues would have to be developed for other specific

cases as they ar evaluated, but it is instructive to usethis value as an example in th present discussion. Forthe lower bound in Figure 12, a Jl, value of 100 kJ/m2

corresponds to a Charpy V-notch impact energy ofabout 735 ktdm 2 at room temperature Only a few ofthe experimental data reviewed in this report fan be-low the Framatome lower-bound value. However, forthe long exposure times (greater than 40 years) asso-ciated with plant-license renewal, there is a concernthat more materials may embrittle to the degree thattheir properties fall below this lower bound or thatthere may be specific cases where a more conservativelower-bound value of impact toughness is necessary.This needs to be determined by case-specific struc-tural integity assessments

2A Tensile and FatigueProperties

In addition to fracture-toughness properties, thetensile, fatigue-crack initiation, and fatigue-ackpropagation properties of cast stainless steels areimportant for use in engineering evaluations of fitnessfor service. Typically, thermal aging increases thetensile ultimate and yield strengths and decreases thetensile ductility (Landerman and Bamford, 1978; andChopera and Chung, 1987).a However, Chopra andChung (1987) also report that a 12-yearreactor-service exposure slightly decreased the tensilestrength of a CF-8 pump cover plate. Overall, thermalaging does cause changes in tensile properties, butsuch changes do not appear to have a significant effecton the design-allowable strengths of cast stainlesssteels.

Landerman and Bamford (1978) reviewed the low-cycle fatiguerack-gwth properties of cast stainlesssteels. They found that the CF-8M stainless steel ex-hibited about the same fatiguo-ack-growth resis-tance as comparable wrought alloys, and that aging at427 0C for 3000 h exhibited no significant effect on thefatigue-cacwk-growth resisac in either air or simu-lated PWR water environments. However, these re-sults are only for aged CF-8M stainless steeL Manycast staless steel components, including reactor cool-ant piping, are made of CF-8 and CF-8A stainlesssteels, and the fatigue-crack growth rates for these ma-terials 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, dependingupon the carbon content. One possible explanation forthis potentially different behavior is that chromiumcarbides may form at the ferrite-austenite interface inCF-8 and CF-8A materials at operating temperatures,whereas the presence of molybdenum prevents suchformaton in CF-SM stainless steeLa This explanationneeds to be validated. If coirsion figue is a probabledegradation 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-operatingtemperatures would not be valid.

Slama et al.P evaluated both the low-cycle fatiguecrack-initiation behavior and fatigue-crack-growthresistance of cast stainless steels at both roomtemperature 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.

10-2daldN = 7.5x 10- 10 AK4

4 I I

J£

I I

(Upper bound designcurve for stainless steels)

1O-3 I-,/

&~ 4

V ' * *

A'a

& a .44k a & aI

*A /

//

I * a

//

/// /

//

/7/

/' / /

E

z

x

J/

10-4 F-

'a & &, & /A'' 4b //

a / /A /

/ O 0° 78 A6

/1'

I/I a78 B4I320 C Air I Virgin

I PWR condition10-5 / /o

/ 0�

/0

/

*78 B7 320 *C Alr I Aged 7500 h& 78 B12 PWR I at 400 *C

== Scatter Band

K10-6I 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 fre-quency of 0.017 Hz (see footnote b).

25

I

thernma aging caused a small (10 to 15%) increase incycli yield strength, and, although one might expectreduced tensile ductility to reduce low-cycle fatigueresistance, they found no significant effect of thermalaging on low-cycle fatigue-crack-initiation resis-tance. Fatigue-crack-growth-rate resistance in airalso was not significantly affected by thermal aging.However, as is illustrated In Figure 16, they found thatfatigue-crak-growth rates for aged CF-8M stainlesssteel increased by as much as a factor of 10 whentested In a simulated PWR water environment, com-pared to testing in aa at 3200C (610P, but did notexceed the upper-bound design fatigue-crack-growth-ate curve for stainless steels.

Additional data for the fatigue crack iniation andgrowth in CF-8, CF-8A, and CFL8M stainless steelmaterials in LWR environments are needed. Data forhigh-cycle fatigue behavior of CFE8M material areneeded because some cast stainless steel surge linesand elbows are subject to thermal stratificatioL. Data

for the guebehaviorof Irmdiatedcaststainless steelmaterials are needed because some reactor internalsare made of these materials No data are presentlyavailable for fatigue behavior of cast stainless steels ina BWR environment. The corrosion-ague behaviorof aged cast stainless steel in a BWVR environmentneeds to be evaluate

In summary, therma aging of cast stainless steelshas been found to produce some increase in tensilestrength and some degradation in known fatigue prop-erties, but it may cause a significant degradation infracture toughness. This degradation could Impair thestructural integrity of cast stainless steel componentsduring long service. Thus, the fracture toughness andimpact energy of thermly embrittled material are themost important items in evaluating the long-termstructural integrity of cast stainless steel components(Bamford et al., 1987). In addition, degraded fatigueand stress-corrosion-cracking resistance, if any,should be taken into account in evaluating the integrityof cast stainless steel components.

26

I

3. INSERVICE INSPECTION OF COMPONENTS

Nondestrutive examination (NDE) procedures areused to detect and characterize the nature of defects incast stainless steel components. Application ofNDE tocast stainless steel components is complicatedbecausethe material microstructure often contains a coarse,variable, and unpredictable combination of columnarand 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 vol-umetric inservice inspection of many LWR compo-nents. However, conventional UT is not sufficientlyreliable for cast stainless steel because coarse grainscause high attenuation and scattering that result in alow signal-to-noise ratio and make defect detectiondifficult. The noise level in cast stainless steel is signif-icantly higher than that in wrought stainless steel. Inaddition, different grain structures in cast stainlesssteel introduce elastic anisotropy, which causes veloc-ity variation and beam skewing and consequently inac-curate detection of the defect. Cast stainless steel iselastically isotopic if composed of equiaxed grains andthe resulting variation in wave velocity with its propa-gation direction is small (2%), whereas cast stainlesssteel is elastically anisotopic if composed of columnargrains and the resulting variation in velocity with itspropagation direction is large (as much as 100% forshear waves) (Kupperman et al.,1987). For these rea-sons, radiography normally is used for the NDE of caststainless steel components, both during fabrication andduring Inservice inspection, especially for thick castsections 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 orwhether the flaw is connected to the surface. For thisreason, radiographic indications must be assumed tobe a flaw of the most detrimental shape and location(Umino and Rao, 1984). Fracture-mechanics analysisindicates that the worst-case flaw will be a surfacecrack of 0.5 aspect (depth/length) ratio. Additionalpenetrant examinations can be performed to determineif the indication is actually a surface flaw. If the indica-tion can be shown to be an imbedded flaw, then thestress intensity factor caused by such an imbeddedflaw can be used in the fraure-mechanics analysis.

Use of radiography during inservice inspection isless practical and is less efficient than during fabrica-tion. In some cases, structural restraints do ao allowsufficient 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 developmentof advanced UT methods to inspect cast stainless steelcomponents. Advanced methods use low-frequency(longer wave length) transducer and short pulses,which enhance the signal-to-noise ratio and Improvethe capability to detect flaws. Two digital signal pro-cessing techniques, time and spatial averaging, canalso provide a high degree of sensitivity and a highsignal-to-noise ratio (Ycong and Ammirato, 1989;Shankar et al., 1988). The time averaging techniqueimproves the presentation of data and makes themclearer and easier to interpret. The spatial averagingtechnique, which is similar to the synthetic aperture fo-cusing technique (SAFI), reduces grain noise indica-tions and, thus, enhances the signal-to-noise ratio forcracks larger than the average grain size. An additionaldifficulty in the inspection of statically cast compo-nents, such as pump casings, comes from rough sur-faces and Irregular surface contours, which cause poortransducer coupling with the component. For reliableexamination, the surface condition needs to be im-proved by mechanical means such as grinding.

The effect of grain structure on the propagation ofultrasonic waves in cast stainless steel material hasbeen studied to increase the accuracy of locating de-fects. Specially designed cast stainless steel specimenswith cracks were tested to experimentally determinethe effects of grain structure on wave propagation.These results are used to identify the grain structure ina cast stainless steel componern, which is in turn usedto compensate for the grain effect in UT examinationto locate defects. Ultrasonic examinations were used tocharacterize grain structures in centrifugally cast stain-less steel piping at the Vogde power plant during pre-serviceinspection.

Ultrasonic examinations were used so inspect weldsand statically cast components, such as pump casingsandelbows,atseveraldomesticPWRplants (eongandAmmirato, 1989). The spatial averaging technique wasused in these inspections to enhance the signal-to-noise ratio. The base metal of the hot-leg elbow at theTrojan plant was inspected because of a concern forhigh stresses. The concern was caused by the fact thatseveral rupture constraints for the primary coolant

a. Private communication with M. S. Good,Batele Pacific Northwest Laboratories, November6, 1986.

27

piping were displaced from their original design posi-tions ASME Section X inservice inspection require-mentsfarcaststainesssteelcomponentdonotincludeinspection of the base metal but include inspection ofwelds. However, the statically cast components arelikely to have a larger number of fabrication defectsthan h weldsw and therefore, may be more susceptibletocmcking.ASME Section XI inspectionrequirementsfor statically cast stainless steel components need to bereevaluated.

Double-probo search units, also called twin-cystalsearch units, were used to inspect PWR Type F pumpcasings (200-mm thick) at the Tihange 1 plant inBelgium. These units employ low-frequency(0.5 MHz) longitudinally polarized waves forenhancement of signal-to-noise ratio but are limitedto a depth range from 10 to 60 mm (Dombret et al,1990). Large-diameter (140-mm) focused transducerswere used to examine down to a depth of 200 mm.

Advanced UT systems also show promise for theNDE of centrifugally cast stainless steel piping, butadditional development and validation are neededbefore they can be used routinely for inservice inspec-tions. An EPRI study (Gregory et al., 1986), usingsamples of centrifugally cast stainless steel pipe withweldments, evaluated the accuracy of the advancedUT methods. The signals appear to be satisfactory forcrack detection. However, because of the type of mi-crostructure, cracks were not always detected or con-sistently sized when the same regions were inspectedfrom both sides of the pipe wall. Field studies at aPWR plant (prior to service) have evaluated the feasi-bility of using advanced UT systems to inspect centrif-ugally cast staidess steel pipe (Avioli, Jr., 1986). Theresults of that work indicate that advanced UT systemscan be used to make measurements successfully in thefield, but that additional development is needed tomake the scanners adaptable to all of the various pip-ing configurations that may be encountered. AnotherEPRI study has evaluated the use of Lamb waves, anultrasonic surface wave technique to detect large-scale

cracks (80% or greater through-wall cracks) in centrif-ugally cast piping (Georgetown University, 1988).Lamb waves are not as severely affected by grainstructure as the conventional ultrasonic waves, and.therefore, experience less attenuation and lower noiselevels. Development of a prototype search unit and acrack sizing capability for this technique are needed sothat its use as an inservice inspection tool can beevaluated.

Use of the acoustic emission technique to detectcrack growth in cast stainless steel components needsto be evaluated because of the difficulties with the ra-diography and UT examinations to detect flaws Crackgrowth through the embrittled ferrite phase releasesthe elastically stored energy in the typical frequencyrange of 0.1 to 0.5 MHz." The higher the degree ofthernal embrittlement, the higher the crack velocitythrough the ferrite phase, and the amplitude of the re-suIting acoustic emission signal will be larger If thecrack growth in the embrittled cast stainless steel isthrough the ferrite phases, it will be composed of sev-eral small steps of brittle crack propagation and is mostlikely to be detected by the acoustic emission tech-nique. Coarse grains of cast stainless steel materialcause little scattering of the acoustic emission signalbecause the signal has a low frequency content (0.1 to0.5 MHz); the corresponding wave length ranges from30 to 6 mm, which exceeds the grain size.

The influence of thermal embrittlement of caststainless steel on acoustic emission signals has notbeen investigated. However, the influence of anotherembrittlement mechanism, i.e., hydrogen embrittle-ment, has been investigated. In hydrogen embrittledmaterial, the crack growth itself generates AE signalswith amplitudes higher than 40 db above electronicnoise. These signals are likely to be detected evenunder plant operational conditions.

a. Private communication withKWU Alliance, May 1990.

P. Jax, Bechtel/

28

I

4. ASSESSMENT OF DEGREE OF MATERIAL EMBRITTLEMENT

For license renewal, one needs to be able to estimatethe amount of thermal embrittlement that has occurredin cast stainless steel LWR components during pastservice and also the amount of embrittlement that isexpected 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) nondestruc-tive testing. These three topics are addressed sepa-rately 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 ofInservice Degradation

The time-temperature aging parameter P, given pre-viously by Equation (1), is the basis for a proposedmodel of inservice degradation caused by thermal em-brittlement. The parameter, through Equation (2), isbased on bulk chemical composition, but it does not di-rectly account for the effects of ferrite volume fractionor ferrite spacing. However, because the ferrite volumefraction is related to composition, P indirectly incorpo-rates the effect of ferrite volume fraction on thermalembrittlement. Also, as was pointed out previously, Pdoes not completely correlate the effects of aging timeand temperature with the measured changes in impactenergy. However, P provides the best model that is cur-rently available. With appropriate conservative as-sumptions, P can be used to estimate the amount ofthermal embrittlement expected in LWR service, andas improved models are developed, those can be incor-porated into the approach suggested below.

Thlie lowcr4ond lines shown in Figures 5 through7 are applicable to exposure temperatures of 4000C(7500F or less and can be described by the followingrelation:

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, exposuretime, and exposure temperature, a lower-bound orminimum Charpy V-notch impact energy can be esti-mated using Equations (4) and (9).

Equation (9) can be used to estimate the range of po-tential CVN values as a function of exposure time at aspecific temperature for CF-SM, CF-8, and CF-3stainless steels. As mentioned, Chopra and Sather(1989) report that the correlation represented by Equa-tion (4) was developed for observed values of Qbetween 75 and 240 kn/mole. For the allowed compo-sitions of these three stainless steels (see Table 3), allthree could have values of Q between 75 and 240kJ/mole. Thus, for purposes of illustration, these wereassumed to be the minimum and maximum values of Qfor a material in service and were used with Equa-tion (9) to estimate lower-bound room-temperatureCharpy values as a function of exposure time at 3000C(570, as presented in Figure 17.

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

and

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

where CVN is the Charpy V-notch Impact energy inkJ/m2 ,PistheagingparameterofEquation(l), andbl,b2, and b3 are material dependent constants with thefollowing values:

Figure 17 shows that CF-3 is expected to be theleast affected by aging, and that CFL8M and C- areexpected to be affected about the same and much morethan Cr-3. For the maximum Q of 240 kJhnole, thesaturation level (P=4) is not expectd to be reachedwithin 100 years. For the minimum Q of 75 kiAnole,the saturation Level (P 4) is expected to be reachedwithin about 11.8 years. For the long-term exposure ofCF-8 or CF-8K the worst-case (minimum QJ room-temperature CVN value is predicted to be about260 Un/m2, whereas the corresponding value for CFL-3is 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 CVN r 735 hU correspondsto J__I- hJ/m 2(sea Figure 12)

.~~r-- I I I I -

1 10 100Time at 3000C. years

Figure 17. Predicted lower-bound Chaipy-impact values for cast stainless steels as a function of aging time at3000C (500MT).

The horizontal, chain-dashed line in Figure 17 at735 kJ/W2 mcrespoids to a lower-bound JIC value of100 kJ/n 2 mentioned by Slama et al of Framame forpossible use in safety assessment of casn stainless steelcomponents. Only the CF-3 alloy is predicted to sat-isfy this limit. Most of the CF-8 and CF-SM composi-tions are predicted to fall below this limit after 10 to20 years of aging.

The wide spread between the maximum and mini-mum lines in Figure 17 further emphasizes the needfor research to better dtfine the value of Q at tempera-ures of 300°C or less discussed previously. If the valueof Q at these lower temperatures is close to 250 J/mole, as was suggested by Bonnet et al. (1990), therate of embritlement would be expected to be close tothe lines for maximum Q in Figure 17.

Figure 17 could be used to estimate the degree ofembritulement of a cast material of unknown composi-tion. However, it is recommended hat the actual com-position be used to compute Q. Then, if the computedQ value is between 75 and 240 kl/mole, a plot such asthat shown in Figure 17 can be prepared for that spe-cific material and the tempeatut(s) of interest

4.2 Metallurgical EvaluationMethods

Metallurgica evaluation methods cmrently providethe best appmach for quantifying thermal embrittle-ment in cast stainless steel components that have beenin servic&e These methods can involve either nonde-structive measurements and examinations or destruc-tive tests and examinations of small samples that havebeen removed from components. When samples aretaken from a component, they must be removed in amanner such that there is no functional damage to thecomponent or such that the component can be accept-ably repaired. Tbe metallurgical evaluation methodseither characterize the microstructure to generate datafor estimating the degree of embrittlement using deg-radation models or characterize mechanical propertiesby testing subsized or miniature samples.

4.2.1 Mlcrostructure Characterization. Mevolume fraction of ferrite in cast stainless stel compo-nents needs to be determined because it is the primarymicrostructural parameter controlling thermal em-brittlement. If the ferite content is below 10%, the po-tential formatia degradation is lilkely to be low. Forferrite levels between 10 and 15%, the potential formaterial degradation is likely to be moderate.

30

I

Significant degradation may take place when the fer-rite level exceeds 15%.

T = temperature of the final heat treat-ment in OC

The amount of ferrite can be estimated from themeasured bulk chemical composition of the material.Procedures for estimating ferrite content have been re-viewed by Leger (1982) and Aubrey et al (1982). Alof thes procedures are based on calculating a chro-mium equivalent value, Crc, and a nickel equivalentvalue, Nrk, and using these equivalent values to com-pute a volume percent ferrite content, %F. he empiri-cal formulas for computing these values have beendeveloped from correlations of experimental data.Aubrey et al. (1982) used linear regression analysis toevaluate four different ferrite estimation formulas foran extensive database of infarmation on CF-3, CF-8,and CF-M stainless steels and found that the follow-ing relationship gave the best overall results:

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

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

+ 20(%C) - 5.18 .

This method is similar to the one defined by Equa-tions (10), (11), and (12), except that it has differentvalues of the coefficients for the terms and includes anadditional term to account for the temperature of the fi-nal beat treatment. Bonnet et al. (1990) report an ac-curacy of +3% ferrite for their method, which is abouttwice as good as the +6% ferrite accuracy cited byAubrey 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 isnot known, and the method of Bonnet et al. cannot be

(10) p applied. When the final heat treating temperaturesare known, the resulting correction factor,[40014[150D -_, could be as large as 25%. For exam-ple, in the work of Aubrey et al the final heat treatingtemperatures ranged fron 1093 to 11770C (2000 to215 V which gives corresponding values of the tern-perature correction factor of 0.98 to 1.24. Thus, it may

(11) be worthwhile in future research to evaluate themethod of Bomnet et al. for typical castings producedin 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 stain-less steel component is known or measured, Equa-tons (10), (11) , and (12) can be used to estimate itsferite content..

Bonnetet at. (1990) have developed the fowing,slightly different method for predicting the amount offerrite from the chemical composition:

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

+ 3.391[400]/[1500 - T1

Leger (1982) found that ferrite content also in-creases as section thickness increases, as is shown bythe data for CF-4M stainless steel in Fgure 18. Theferrite content as a function of thickness appears to sat-rate (reach a constant value) at section thicknesses

greater than approximately 45 mm (1.77 in.). lhere-fore, using the data for the thicker castings would pro-vide conservative (larger than actually expected)estimates of ferrite content for tinner castings Thework of Aubrey et al. (1982) was for castings thickerthan 25 mm (D.98 in.), soEquations (10), (11), and (12)probably provide reasonable predictions for thick-section casngs.

Equations (10), (11), and (12) were used to estimatethe ferrite content of the 45-mm thick samples ofCF-SM that were evaluated by Leger (1982) and of the10-mm (3.94 In.) thick samples of CFLS and CF-SMthat were discussed previously (see Figure 4). Fig-ure 19 duo that there wasagood correlation betweenpredicted and measured ferrite content for these dataThe solid line indicates the mean correlation; thedottedand dashed lines indicate the upper and lower limits,where

31

respectively, which are plus and minus two times thestandard error [estimated at 3.03, as reported byAubrey et al. (1982)] from the mean correlation,respectively.

45 0.63

30

25 071]

. 20 r itFl30

*E 15 -MicrographyU. ,-Ferritescope

10,A- 0.90

5 -

F- ~ 1.04~ 1.09

o 1.240 20 40 60

Thickness (mm)

Figure 18. Ferrite content of CF-8M castings as afunction of section thickness (Leger, 1982).

In order to make a reasmnably conservative estimateof the frrite content, one would add 6.06 to the valuecalculated using Equation (10) far comparison withacceptable ferrite levels. Thus, for the adjusted ferritecontent to not exceed an upper-limit ferrite content of15%, the mean ferrite content calculated using Equa-tion (10) should not exceed 8.94. This value is onlyslightly greater than the minimun ferrite level of 8%that is recommended to avoid sensitization duringwelding (Hazelton, 1988). Furthermore, if one adds6.06 to 8 to account for the expected unceainty, theminimum would be 14.06%. This -5% overlap in de-sirable minimum and maximum levels occurs becauseof the uncertainty associated with calculatng ferritecontent from composition. Therefore, the calculationof fere leveh from chemical composition providesonly approximate guidance for the assessmentof stain-

less steel castings and is of limited use in quantifyingthema embrittlement.

Recent results indicate that ferrit spacing, as wellas ferrite content, may influence the degme of thermalembrittlement (Chopra and Chung, 1986b, 1987b,1989a, 1989b). Procedures for estimating the ferritespacing in a cast stainless steel component are notavailable, but, as discussed previously, Chopra andChung (1990a) recommend that a value of -180 .umbe assumed as a worst-case condition if the actual fer-rite spacing has not been measured. Since the ferritespacing is a significant factor in the 0 parameter,which corelates with the minimum Charpy impact en-ergy expected after long-term exposure, it would beuseful to develop a procedure for estimating the feitespacing in sainless steel castings.

Measuring ferrite content is an alternative to esti-mating it. If a material sample can be removed fromthe component, then quantitative metographic pro-cedures can be used to measure the ferrite content(Leger, 1982; Aubrey et al., 1982). The procedurestypically employ either computerized image analysisor manual point-counting techniques. The average fer-rite spacing also could be measured during metallo-graphic examination. Material sampling must beperformed carefully to ensure that the regions ex-amined represent the bulk material. Because sampleremoval often may be limited to exterior or interior re-gions, quantification of the ferrite volume and spacingwithin the wall of the component is difficult.

The above metallographic measurements can bemade on surface replicas of selected areas of a compo-nent in the field. Metallographic replication has beendeveloped as a viable NDE tool (Neubauer and Wedel,1983); Auerkai, 1983; Viswanathan, 1985; Harth andSherlock, 1985; Henry and Ellis, 1983; Hilton andSteakley, 1986; Janiszewskl et al., 1986; Tlkigawaet al., 1986; Masuyama et al., 1985), and ASTMEmergency Standard ES-12 for field replication hasbeen developed.V Prototype replication equipment hasbeen used to remotely examine the inside surface ofsteam-turbine rotors ldkigawa et al., 1986) and to in-spect steam-generator tubes (Padden, 1983; Paddenand Daniels, 1976, 1981). Replicas must be made atambient temperature during an outage and at surfacelocations that are accessible. As with sample removal,judicious selection of the areas to be replicated isnecessary.

a. Fm apresentationbyJ. DeLong attheASTM-ASME-MPCJ-lWorshapattheASMlWmterAn-nual Meeting, Anaheim, December 93 1986.

32

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

40 - CF-8M

.50 200-

00

o -

~ 0 02405

Measured ferrite content (vol %)

Figure 19. Cmparison of fAre conttpredicted from chemical composition with measured feitecontentus-ing 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 mag-netic than the austenite phase, ferrite content can bemeasured by instruments that measure the magneticproperes of the stainless steel. Field measurements offerrite content, based on magnetc properties, can bemade using a Ferritescope (Leger, 1982; Aubrey et aL,1982). This instnrnent can be used in the field to mea-sure ferrite levels up to 30% and has been demon-strated to produce teliable measurements far sampleswith section sizes ranging from 27 to 152 mm (1.06 to6.0 in.) (Aubrey et al, 1982). Figure 20 shows that agood correlation between ferrite content measured us-ing the Ferrhtscope and using a metallographic point-counting method can be obtained. This methodappears promising, but field validation tests on repre-sentativeLWR cast stainless steel components are nec-essary before it can be used routinely during inserviceinspections. Section m of the ASME Boiler and Pres-sure Vessel Code now permits the use of magneticmeasuring instruments for determining the ferritelevels in welds.

Bonnet et a!. report that the saturation magneticmethod (using a Sigmametre) provides the most reli-able experiental technique for measuring the amountof ferrite. This method is somewhat different than themagnet-induction method on which the Ferritescopeis based. Curtis and Sherwin (1961) discuss using thesaturation magnetic method to measure the feite con-tent of stainless steel welds. Either type of magneticmeasurement method should provide reasonable mea-surements of ferrite amount if it is propemty calNiatedand used by a qualified operator.

Factors other than the bulk chemical compositionand ferrite content and distribution in stainless steelcastings may play an important role in thermal cm-brittlement. For example, as was shown previously(see Table 5), the chemical composition of the feteregions is significantly different from that of the auste-iite regions in these alloys. The partitioning of the

33

1030 0 0

20000

a 0'E .3lo 0

10 20 30Point-counted ferrite (vol %)

Flgure 20. Comparison of ferrite content measured using a ferritescope with that determined metallographicallyby a point-count procedure.

ferrite-promoting elements to the ferrite phase mayaffect the amount of thermal embrittlement. If such aneffect were found in fut studies, a correlation be-tween local composition and the degree of thrmal em-brittlement might be developed. Such a correlationmight be useful in the future development of alloycompositions that would be resistant to partitioningduring casting and to subsequent embrittlement duringservice at LWR operating tperatues.

4.2.2 Testing of Miniature Samples. Mea-suring the mechanical properties of small samples re-moved from stainless steel castings is anotherpotentialmethod for quantifying the degree of thermal em-brittlement caused by service exposure. Two ap-proaches could be used. Fracture touginess could bemeasured directly using subsized specimens. Alterna-tively, impact cne., tensile strength, or Indentationresistance could be measured and used to estimatefracture toughness if appropriate correlations wereavailable. Direct measurement of frac toughnessusually is more costly and normally requires lamersamples than the indirect measurements However, thedirect measurements are generally more reliable thanthe indirect ones

T e has been a gm deal of recent interest in thetesting of subsized specimens to measure mechanicalproperties (Jaske et aL. 1986; Manahan, 1981; Haggaget aL, 1985; Msawa and Hamaguchl, 1986; ASTM,1986; Huang. 1986; Lucas et aL 1986; Corwin and

Hougland, 1986; McConnelletaL, 1986).Muchof thiswork has been motivated by the need to measure theproperties of irradiated materials. For example, theASTM special technical publication (ASTM 1986) in-dludes 21 papers of the ASTM Symposium on the Useof Nonstandard Subsized Specimens for irradiated1bsting. These past studies included the comparativetesting of standard-sized and subsized specimens ofreactor pressure vessel steels and wrought austeniticstainless steels, but they did not include testing of thecast CF-S. C7-SM, and CF-3 stainless steels. Suchtesting is needed to vaLidate the use of subsized speci-mens for these materials. Results of the past evalua-tions of subsized specimens do povide guidance forestimating the minimum specimen size that would beexpected to yield valid measurements of Jk.

Experience indicates that the minimum specimendimension should be at least five to ten times the sig-nificant micostuctural feature (aske et AL, 1986),awhich for the cast stainless steels may be either thegrain size or the ferrite spacing. Experimental studiesare needed to establish what significant microstruc-tural feature controls fracture toughness in cast stain-less stes. Values of grain size and ferr spacing areusually lazier for heavy-section castings than for thin-section castings. If, for example, ferrite spacing is thesignificant feature, the minimum specimen dimensionwould 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 rangefrom 50 to 200 pm (Chopra and Chung, 1987).

As was discussed previously, ASTM E 813 requiresthat the thickness, B, and the uncracled ligament, b, ofa test specimen be greater than 25JcQsf, where JQ is theconditional value of fracture toughness, and sf is theflow stress, which is the average of the ultimate tensilestrength and the 0.2% offset tensil yield strength. Ifthese size conditions, plus the regression-line slopecondition (see Paragraph 9.2 of ASTM E 813), aresatisfied, then JQ equals JIB

Typical tensile properties (Chopra and Chung,1987) of cast stainless stels that were aged for 9,980 hat 3500C (6600F) and tested at rom temperture andat288°C (550) can be used to estimate the minimumvalues of B and b that would be required to obtain validJC values. For example, if Jlj = 250 kjAM2, which Fig-ures 12 and 14 show is a fracture toughness value thatcould be expected for thermally embrittled cast stain-less steel, and sf = 440 MPa (64 ksI) at room tempera-ture and 300 MPa (44 kal) at 2880C (5505F), then

B. b (minimum) 25 (0.25W.44) = 14.2 mm at roomtemperatur

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 stainlesssteels have relatively high fracture toughness, fairlylarge specimens, rather than subsized ones, are neededfor measuring their 1k values

One property that could be measured on small mate-rial samples is microhardness of the ferrite phase.Recent research has shown that the Vickers hardnessnumber of the ferrite phase increased from 260 inunaged material to 377 in aged material from a CF-8casting (Chopra and Clung, 1987). It may be possibleto establish a correlation between the hardness of theferrite phase and the amount of thermal embrittlement.The correlation would have to take into account theeffect of ferrite volume fraction and ferrite spacing, asdiscussed previously. Future research is needed toevaluate the potential for developing this type ofmichardness-embrittlementcorrelation.

4.3 Use of Ultrasonic Testing toCharacterize Materials

If

= 100 kUjm 2, which is the lower bound used byFramatome i

then

B, b (minimum) =25 (0.l01OA4) 5.68 mm at roomtem~perature

a. UnpublishedpaperpresentedbyG.Slamaetal..at SMiR Postconfernce 6 at Monterey, Cal ia,August29-30, 1983.

Use of ultMasonic techniques is being developed toquantify material embrittlement. Both Swanson(1981) and Vary (1976) have found correlations be-tween ultrasonic attenuation and the fiacture tough-ness of wrought low-alloy steels. Their measurementswere iade on well-documented samples under care-fully controlled conditions in the laboratory. Similarstudies have not been performed on cast stainlesssteels, and the work has not been extended to the fieldmeasurement of fracture toughness on components.Thus, extensive research and development work willprobably be necessary before ultrasonic measuentmethods can be used to quantify enbrittlement in caststainless steels.

35

5. WELD-REPAIR PROCEDURES

If inservice NDB reveals that defects ar present ina cast stainless steel component, the possibility ofweld-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 befound In castings of these alloys during componentfabrication. The major concerns with inservicaare that access may be limited and that the em-brittlement may affect weldability. The former iten iscommon to inservice repairs of any alloy, while thelatter item is unique to the cast stainless steels.

The procedures used for weld-repair of staticallycast components during manufaurg have been rieviewed elsewhere (Egan et aL, 1987); a brief review ofthe standard procedures developed for repair weldingduring the fabrication of centrifugally cast stainlesssteel pipe is presented here After the pipe is cast andexcess material is machined from the outside and in-side surfaces, the pipe is fully radiographed to identifymajor and minor defects, in accordance with theASME Code, The locations of all those defects aremapped and recorded for future reference. The defectsare removed or reduced to an acceptable size by grind-ing or thermal gouging, with the final metal being re-moved by grinding using an aluminurm oxide or siliconcarbide wheel. The repair area is dye penetrant in-spected and thoroughly cleaned with organic solventsbefore welding.

Welding procedures and welders are qualified asspecified by the ASME Code (Group P-8). Welding isperformed using the manual shielded metal-arcprocess and Type 316 weld rods that produce at least7% ferrite in the weld metal. The minimum preheat is560C (1330F); no postweld heat treatment isemployed, and the inrpass temperature is keptbelow

a. E G. Kay, Sandusky Foundry & Machine Com-pany, personal commnication of unpublished infor-mation, October 1986.

1770C (350MF) for multiple passes. Welding current ismaintained as low as possible, travel speed is main-tained as high as possible, and stringer beads are used.After completion, all weld repairs are dye-peneftantinspected, and major repairs also are radiographed.

There should be no major problem in developingand qualifying similar weld-repair procedures forfield use on components that have been in service, pro-vided that the defect is local. For example, if a fatiguecrack were found to have developed at a local geomet-ric discontinuit in a component, then it would be fea-sible to consider a repair; but if fatigue damage werefound to be significant throughout a component, a re-pair probably would not be feasible. In repair welding,care must be taken to minimize the introduction of re-sidual stresses. On the other hand, postweld heat treat-ment Is impractical in most cases, and when it is used,it could produce sensitized regions near the ends of theheat-teated area For major repairs, it is likely thatspecific procedures will have to be developed andqualified on a case-by-case basis.

If repair welding is considered to be a viable optionfor license extension, research is needed to evaluateany problems that might be encountered in the weldingof thermally embrittled material. For example, it mightbe found Otat thermally embrittled material is moresusceptible to weld cracking than new material.Because the embrittlement mechanism is different foraging at temperatures greater than 400°C (750TF) thanfor aging at temperatures less than 3506C (6600F), theresearch should concentrate on the welding of materi-als that have been aged for long times at 350VC or less.Studies using samples that have been removed fromactual LWR service would be particularly appropriate.Results of this type of research would provide guid-ance for the development and qualification of weld-repair procedures for cast stainless steel LWRcomponents

36

6. LIFE ASSESSMENT PROCEDURE

There are two basic approaches to maximizing thestructural integrity of LWR components for operationduring 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 orreplace it only when there is cause to do so. This ap-proach is employed widely to critical structural com-ponents whose failure would impact safety and whoseunnecessary replacement wouldbe extremely costly. Asecond approach would be to arbitrarily replace a corn-ponent after some specified service life without at-tempting to determine its current condition. Thismethod is applied to critical structural componentswhose failure would impact safety and whose replace-ment is easier and less costly than evaluating their cur-rent condition. For most cast stainless steel LWRcomponents that are subject to the possibility of ther-mal embrittlement during long-term service, the sec-ond approach is undesirable because of an unfavorablecost/benefit ratio; thus, the first approach is the mostreasonable one.

Figure 21 presents a generic procedure for agingand license renewal evaluations of LWR cast stainlesssteel components. Nine major steps are listed.

The first step is to review all of the design, fabrica-tion, and construction records related to the compo-nents being evaluated. Some of the items listed inStep I may not be available. Those items should beIdentified, and plans to obtain them during the next in-service inspection (151) should be developed, if possi-ble. In cases where it is not possible to obtaincase-specific data for austenitic stainless steel prod-ucts (e g., impact and fracture toughness data are notrequired by Section H1 of the ASME Boiler and Pres-sure Vessel Code), worst-case data should be obtainedfrom handbooks, technical literature, and industrial ex-perience. The information collected in Step 1 shouldbe stored for easy retrieval and reference in the future.

All of the ISI records for the components are re-viewed in the second step. The information on defectindications and locations and maximum undetectedflaw sizes is needed for the fracture mechanics evalua-tion (Step 4). The chemical compositions of the mate-rials should be verified by measurements on thecomponents. Essential information (such as size andlocation of flaws in the base metal) not being obtainedduring the current SI programs should be obtainedduring fuature IS! programs. In many cases, key loca-tions are not accessible to ISI, so the original records

must be used with appropriate factors applied toaccount for uncertainty.

The third step is to review all of the data on operat-ing history. The operating temperature is needed toevaluate the degree of thermal embrittlement, and dataon loadings and cyclic operations are needed to helpestimate the amount of crack propagation that mighthave occurred or might occur in the future. Indicationsof past problems may help point out areas that shouldbe evaluated in Steps 4 and 5.

A fatigue and fractiu-mechanics evaluation is per-formed in the fourth step. The locations and sizes ofthe existing flaws are identified from the ISI records.Simplified, conservative loadings are estimated from areview of the operating history. Stress analyses are per-formed to identify regions of the highest stress, and thestress history at the flaw locations is estimated. If thereview of operating history reveals that certain loadswere not included In the original design, revised calcu-lations for the ASME Section m cumulative fatigueusage fators are made. Crack-propagation analysis isperformed to estimate the maximum possible defectsize at the end of the next operating period (EONOP).Finally, the minimum required fracture toughness cor-responding to the worst-flaw size at the EONOP aredetermined.

The fatigue and fracture-mechanics evaluation per-formed in the fourth step is conservative because it as-sumes that the flaws in the cast stainless steelcomponents are crack-like. This assumption is madebecause the current inservice inspection methods arenot capable of completely characterizing the flaws. Ifan emerging inservice inspection method can posi-tively identify that a given flaw is not crack-like, thena fatigue-crack initiation analysis would precede thecrack propagation analysis. In such a fatigue analysis,an effective stress concentration factor is computed forthe flaw (Hertzberg, 1983), and then the usage factor atEONOP is determined. If the usage factor is greaterthan 1.0, it is assumed that a crack will be initiated atthe given flaw before EONOP; in this case, a crackpropagation analysis would be needed.

In the ffth step, the material condition of the com-ponentis assessed to determine the minimum expectedfracture toughness atEONOP. Three techniques can beemployed to assess the degree of thermal embrittle-ment of a cast stainless steel component, as discussedin 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 procedures1.8 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 Pressure3.7 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 stress4.3.2 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 nextoperating period (EONOP)

5. Assess material condition

5.1 Use simplified analytical model5.1.1 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 CharpyV-notch impact energy at long life

5.1.3 Estimate lower-bound fracture toughness atEONOP from fracture toughness versus impactenergy correlation

5.2 Assess ferrite content

5±-1 Calculate upper bound ferrite content from composition52.2 Measure ferrite content

Metallogaphic sample from componentSurface replication of componentFerritescope

5±3 Use guiddlines to estimate ractre toughnes atEONOP5.3 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 estimatedlower-boundvalue6.1.1 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 criticalflaw 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 remainingsafe 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 afactorof 1.5 to 3

8.3.4 Evaluate selected samples at 5-year intervals to assessRate 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) toestimate the minimum Charpy V-notch impact energyvalueatEONOP,orthe G parameterofEquation (5)to estimate the minimum Charpy V-notch impact en-ergy after long-tem aging. Then, using a Jjc (fracturetoughness) versus CVN corzeation, minimum Ji ises-timated. The second technique is based on eithercalcu-lating or measuring the ferrite content and usingguidelines or an embrittlement model to determine Ifthe ferrite value is acceptably low. The third techniqueis to measure the current Chaipy V-notch impact en-ergy, fracture toughness, or microlardness of samplesthat have been removed from a component; then thefirst technique, i.e; the simplified analytical model, isused to estimate the JIC value at EONOP.

In the sixth step, the minimum required fracturetoughness from Step 4 is compared with the estimated

-actual lower-bound fracture toughness at EONOPfrom Step 5. and the margin of safety is determined.Using the estimated stresses from Step 4 and the esti-mated lower-bound toughness from Step S. criticalcrack size is determined. Then, the detectability of thecritical flaw size is assessed. Finally a leak-before-break (LBB) analysis is performed to deternine themargin between the critical crack length and the leaka-ge-detectable crack length that corresponds to tentimes the detectable leak rate.

Once the component integrity at EONOP isevaluated, a plan of action is developed in the seventhstep. If acceptable safety margins are present, and theASME Section III cumulative fatigue usage factor atEONOP is less than 1.0, continued operation toEONOP may be permitted. Potential need for repair orreplacement must be considered.

The eighth step is to establish a plan for future ISI ofthe component. The plan should define both the extentof inspection required and the interval at which rein-spection should be performed. It may be useful to con-sider developing athermal-mbrittlement svellanceprogram similar to the programs used for evaluatingradiation effects. Such a program would provide valu-able data from realistic exposure conditions.

The ninth step is to reevaluate the component afterfuture inspections or as required by other circum-stances. The revaluation will start at S tep 2, upgraded

by the data from the most recent evaluation, and willp- cedfrom that point.

The major difficulties in implementing this overallprocedure are encountered in Step 5, Assess Compo-nent Condition. The simplified analytical model(Step 5.1) can be applied with a reasonable degree ofconfidence, even though further validation of the cor-relation between CVN and Jic values is needed (seeFigures 12 and 14). However, ferrite assessment(Step 5.2) and the measurement of properties(Step 5.3) have not been fully developed and vali-dated. Quantitative evaluations of possible ferrite mea-surement techniques and subsized specimen testingmethods are needed. Correlations and/or models areneeded to relate ferrite content, spacing, and hardnessto the amount of thermal embrittlement. Also, a modelfor predicting future degradation caused by thermalembrittlement should be developed to remove uncer-tainties associated with the current approach. Thus,considerable work is needed in the area of assessingand forecasting the condition of cast stainless steelcomponents that operate at temperatures near 3000C(570 in light water reactors.

Fracture-mechanics evaluations also could be diffi-cult to conduct in some situations. Framatome has per-formed fracture mechanics evaluations for their caststainless steel components and has developed a mini-mum required value of lO kJ/tm2 for JI, at the end ofoperating life. The establishment of a minimum Jtivalue simplifies the evaluation of standardized compo-nents because there is no need to repeat the fracture-mechanics analysis for each case. However, such asimplified 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-mechanicsanalyses for several different classes of cast stainlesssteel components, so minimum required values of Ji atthe end of operating life could be developed for eachclass. These values then could be applied to the case-by-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 ther-mal embrittlement damage in CF-8, CF-8M, andCF-3 cast stainless steel components subjected to lightwaterreactor (LWR) operating temperatures. Thesa al-loys have an austerite matrix with islands of ferite.For LWR applications, typical volume fractions of theferrite phase are In the range of 8 to 20%. The ferritephase is subject tD embrittlement by the formation ofalpha-prime after long-tem (many years) exposuresat temperatures new 3000C (5709F). The amount, dis-tribution, and degree of embrittlement of the ferritephase deternine the extent to which a component hasbeen degraded by thermal embrittlement. Tis degra-dation 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-growthresistance are not markedly affected by thermal em-britlement, so the primary concern in addressing therenewal or extension of LWR operating licenses is theloss of fracture toughness after long-term service.However, more data on the effect of thermal embritle-ment on fatigue-crack-growth resistance are needed.especially for CF-8 stainless steel and for high-cyclefatigue.

Prom the maximum service temperatme and cheni-cal composition of the alloy, an aging parameter isproposed for estimating the lower-bound room-temperature CVN values for service times up to 100years, or the minimum CVN values at a very long timecan be estimated using the 0 parameter. A CYN im-pact energy versus JD correlation then can be used toestimate a corresponding lower-bound value of thefracture toughness.

Nondestructive examination (NDB) of cast stainlesssteel components yields information on the size, loca-tion, and distribution of defects. This information canbe employed in fractutr-mechanics evaluations ofstructural integrity. However, because of the coarseand variable nature of the microstructure of cast stain-less steels, ultrasonic testing (UO) is limited in its use-fulness as an NDE tool; radiography normally is usedfor the NDE of these materials, especially for thicksections. Research is in progress to develop and vali-date advanced UT methods for inservice inspection ofcast stainless steel components. Howevea, neither cur-rent nor advanced NDE methods characterize the me-

chanical properties or metallugical condition (degreeof embrittlement) of insevice components.

Correlationshavebeen developedfor esmating theferrite content of stainless steel castings from theirchemical composition. However, these correlations donot predict the ferrite spacing, which is a parameterthat influences fracture toughness. Castings that con-tain more than about 10 to 15% ferrite appear to be themost susceptible to thermal embrittlement, but no pre-cise guideline for an acceptable ferrite content hasbeen established, and mateia with even less than 10%ferrite may be degraded by thermal embrittlement.Metallographic examination of small samples re-moved from components or of surface replicas canprovide data on the ferrite content and spacing at spe-cific locations in castings. The feritescope is a mag-netic instrument that can be used in the field tonondestructively measure ferrite contents up to 30%.

Subsized specimens from small samples ofmaterialthat have been removed from service can be used tomeasure fracture toughness. Guidelines for minimumspecimen size ar provided by ASTM B 813. Thoseguidelines are expected to be conservative, but theyhave not been verified directly for the testing of caststainless steels.

Long-term service exposure increases the miro-hardness in the ferrite phase of cast stainless steels, butno correlation between th measured microhardnessand change in impact resistance or fracture toughnesshas been established The correlation also would haveto account for the effects of ferrite content and spacingto provide a quantitative estimate of fracturetoughness.

Weld-repair procedures have been developed forthe removal of casting defects during original fabrica-tion. It shouldbe reasonably easy to adapt these proce-dures to the inservice repair of stainless steel castingsHowever, repair welding has not been fully qualifiedand validated for material that has been thermally em-brittled by long-term service exposures in LWRenvironments.

The recommended approach for evaluating caststainless steel components for LWR license renewals,is to assess the current condition of the component, es-timate its remaining safe life, and decide on continuedas-is use, repair, or replacement of the component The

42

alternative approach of autmatically replacing a con-ponent after a given period of service is not usuallyfeasible.

A nine-step procedure for the aging and license re-newal evaluation of cast stainless steel components hasbeen outlined. Tbe first three steps include the reviewof fabrication and construction records, ISI records,and operating history of the components. In the fourthstep, the data colected from these reviews are used toperform the fracture mechanics evaluation to deter-

mine the minimum required fracture toughness and theworst-flaw size at ihe end of the extended operatingperiod (EONOP). Three different methods are pres-ented in Step 5 to estimate the lowerbond fracturetoughness of cast stainless steel components atEONOP. Then, the minimum required fracture tough-ness at EONOP is compared with the correspondingestimated lower-bound value, and a leak-before-break analysis is performed. The action regarding con-tinued operation, repair, or replacement of acomponent is established in Step 7, and finally, plansfor the next ISI are made.

43

8. RECOMMENDED FUTURE WORK

Improved methods need to be developed for eva-luating the fitness of cast stainless steel LWR compo-nenuts for service during original or extended lcensingperiods and for reducing the conservatism inherent inthe procedures discussed above. Therefore, we recom-mend thai fure work be in the following13 areas:

1. Guidelines of worst-case data should be de-veloped for use when case-specific data arenot available.

2. Thermal-embrittlement surveillance pro-grams similar to the programs used for eva-luating radiation effects should be developed.Such programs would provide valuable datafrom realistic exposure conditions.

3. More CVN and Jxc data on thermally agedmaterial tested at normal LWR operatingtemperatures [about 280 to 29O0C (536 to554F] are needed.

4. An improved aging parameter should be de-veloped that more accurately models theaging mechanisms that cause thermalembrittlement during long-term serviceexposure

5. Work should be undertaken to determine ifthere is some critical combination of ferritelevel and spacing above which ferite distri-bution in cast stainless steel is continuous. Ifsuch a limit exists, then a method of estimat-ing ferrite content and spacing from composi-tion and production history should bedeveloped.

6. Fatigue-cc-growth studies of aged CF-8stainless steel should be performed. Near-threshold, as well as low-cycle, tests areneeded on both aged and irradiated materials.

Data for high-cycle fatigue behavior of agedCF-8M stainless steel are also needed.

7. The use of metallographic surface replica-tion. the feruitescope, and other magnetic-property measurement tools, along witheddy-current methods, should be evaluated,demonstrated, and validated as inserviceNDE tools for evaluating the condition ofstainless steel castings.

8. Experimental studies should be performed todetermine the smallest specimen size forvalid measurement of the fracture toughnessof cast stainless steels.

9. Work should be undertaken to evaluate thecorrelation between the micohbardness of theferrite phase and fracture toughness for caststainless steels.

10. Procedures for repair welding cast stainlesssteels that have been subject to long-erm ser-vice in LWR environments should be quali-fied and validated for use in license renewal.

11. Fracture-mechanics analyses of representa-tive component/load-history combinationsshould be conducted. Such analyses can pro-vide estimates of the required minimum Juvalue.

12. Current work on the development of ad-vanced ultrasonic methods for inspection ofcast stainless steel components should becontinued.

13. Present ISI programs should be modified toobtain essential data that are needed for theevaluation of LWR cast stainless steelcomponents

44

i

9. REFERENCES

Alexander, D. L., et &I. (1990). 'he Effect of Aging at 343°C on Type 308 Stainless Steel Weldments," Fatigue,Degradation, and Fracture-1990, ASME PVP-Vol. 195, MCP-Vol.30, New York, presented at the 1990Pressure Vessels and Piping Coference, Nashville. Tennessee, Jne 17-21 ,1990, pp. 187-192.

ASTM (1986). The Use of Small-Scale Specimensfor Testing Irradiated Material, ASTM STP 888, Philadelphia.

Aubrey, L. S., et al. (1982). "Ferrite Measuremnent and Control in Cast Duplex Stainless Steels," Stainless SteelCastings, ASTM SmP 756, The American Society for Testing and Materials, Philadelphia, pp. 126-164.

Auerkari P. (1983). "Remnant Creep Life Estimation of Old Power Plant Slcam Piping Systens," Advances In LcePrediction Methods, The American Society of Mechanical Engineers, New Yord, pp. 353-356.

Auger, P., et al. (1990). "Atom probe and transmission electron microscopy study of aging of cast duplex stainlesssteels." Materials Science and Technology, 6, pp. 301-313.

Avioli, Jr., M. J. (1986). "Application of Advanced UT Systems to the Preservice Inspection of Centrifugally CastStainless Steel Pipes in Power Plants," Nondestructive Evaluation Program: Progress in 1985. EPRINP-4315-S, pp. 6-1 to 6-8.

Bamford, W. IL, etal. (1985). 'ITermal Aging of Cast Stainless Stel, and Its Impact on Piping Integrity,"Journalof Engineering Materials and Technology, 107, 1, pp. 53-60.

Bamford, W. H., et a. (1987). "Aging Effects on Cast Stainless Steels," Properties of Stainless Steels In ElevatedTemperature Service, UPC-Vol. 26, PVP-Vol. 132, The American Society of Mechanical Engineers,New York,pp.49-63.

Bonnet, S., et al. (1990). "Relationship between evolution of mechanical properties of various cast stainless steelsand metallurgical and aging parameters, outline of current EDF programmes," Materials and Science Tech-nology, 6, pp. 221-229.

Brown, . E., et al. (1990). "he Thermal Aging of Alloys Modelling the Ferrite Phase in Stainless Steel Castings,"Fatigue, Degradation, andFracturc-1990, PVP-Vol. 30, The American Society of Mechanical Engineers,New York, pp. 175-185.

Buchalet, C., Y. Meyaud, and Ph. Thupin (1985). 'Effect of Long-Tam Aging an the Critical Defect Sizes in Pri-mary Piping Cast Stainless Steel Elbows," Proceedings of the Second n ational Conference on Environ-mental Degradation of Materials in Nuclear Power Systems-Water Reactors, Monterey, Calyfornia,September9-12,1985. ANSITMS-AUIE/NACE, pp. 477-484.

Chopra, 0. L (1990). nThermal Aging of Cast Stainless Steels: Mechanisms and Predictions," Fatigue, Degrada-.ton, and Fracture-1990. PVP-Vol. 195, MPC-Vol. 30, Ihe American Society of Mechanical Engineers,New York, pp. 193-214.

Copra, O. K, and HI M. Clung (1985). "Aging of Cast Duplex Stainless Steels in LWR Systems,"NuclearEngl-neering andDesign, 89, pp. 305-318.

Chopra, 0. K., and H. M. Chung (1986a). "Aging of Cast Stainless Steel," Thirteenth Water Reactor SafetyResearch information Meeting, 2,NURE0/CP-072, pp. 387-407.

Chopra, 0. L, and H. M. Chung (1986b). "Investigations of the Mechanism of Thermal Aging of Cast StainlessSteels," presented at the Metal Component Subcommittee ofthe Advsory Committee on Reactor SafeguardsMeetng, Columbus, Ohio, Juy 1-2,1986, transcript available from NRC Public Document Room.

45

Chopra, 0. K., and K K Chung (1986c). Long-Term Embrittlement of Cast Duplex Stainless Steels hi LWRSystems: Semiannual Report October 1983-March 1986, NUREOGCR4744 VoL 1, No. 1, ANL-86-54.

Chopra, OK., and It M. Chung (1987).Long-TermEmbrittlementofCastDuplexStainlessSteeluinLWR Systems:Semiannual Report AprilSeptemberl986. NUREG/CR-4744VoL 1, No.2, ANL;7-16.

Chopra, 0. K., and IL At Chung (1989a). "Aging Of Cast Stainless Se" presented by W. Shack at ACRS MetalComponents Subcommittee Meeting. Battelle, Columbus. Ohio, March 15, 1989, transcript available fromNRC Public Document Room.

Chopra, O. K., and IL M. Chung (1989b). "Initial Assessment of the Processes and Significance of Thermal Agingin Cast Stainless Steels," Sixteenth WaterReactorSafety r1formadonMeetingNationalInstitute of Standardsand Technology. Gaithersburg, Maryland, October 24-27.1988, NUREGIP-097, YoL 3, pp. 519-546.

Chopra. O. K, and IL K Chlung (1990a). "Cast Stainless Steel: Mechanims and Predictions," DTansactions ofthe17th Water Reactor Sofety Research Informadon Meeting, Rockville, Maryland, October 23-25, 1989,NUREG/CP-0105, Vol. 3, pp. 319-344.

Chopra, 0. K., and IL NE Chung (1990b). '1Long-Tebm Embrittlement of Cast Duplex Stainless Steels i LWRSystems Semiannual Repot October 1987-Marh 1988,' NUREGVR-4744 Vol. 3, No. 1, ANL-89=

Chopra, 0. K., and A. Sather(1990). "nitial Assessment of the Mechanisms and Significance ofLow-TempetureEmbrittlerient of Cast Stainless Steels in LWR Systems," NUREGiXR-538S, ANL-89/17.

Chung& IL M.(1989). "'Termal Aging of Decommissioned Reactor Cast Stainless Steel Components and Method-ology for Life Predictions," Lfe Assessment and Lyfe Extensions of Power Plant Componens-1989:Presented at the 1989 ASME Pressure Vessels and Piping Conference-ASME Co-sponsorship, Honolulu,Hawaii.July 23-127, 1989,p. 111.

Copeland, J. P., and A. J. Giannuzzi (1984). Long-Term Integrity of Nuclear Power Plant Components,EPR1-NP-3673-LD, October.

Corwin,W. R., and A. NE Hougland (1986) "Effect of Specimen Size and Material Conditionon the Charpy rmpactProperties of 9Cr-lMo-V-Nb Steel," Te Use of Smal-Scale Specimensfor Testing Irradiated Material,ASTM SIP 888, lbe American Society for Testing and Materials, Philadelphia, pp. 325-338.

Curtis, 0. C., and J. Sherwin (1961). "'Magnetic method for the estimation of ferite in stainless steel welds," BritishJournal ofApplled Physics, 12, pp. 344-345.

Devine, T. 1M (1980), "Influence of Carbon Content and Ferrite Morphology on fte Sensitizatin of Duplex Stain-less Steel American Societyfor Metals, Metallurgical Transaction A, IlA, May. pp. 791 -800.

Dombret, P., P. Caussin, and P. Rorivc (1990), "Developing Ultrasonics forPWR Pump Bowl Inservice Inspection,"Nuclear Engineering International, pp. 42-44.

Egan, 0. R., et al. (1987). Inspection of CentrIgally Cast Stainless Steel Components In PWR, EPRI NP-51 31.

EPRI (1986). InSpectiOn of PWR Cast Pipe," EPRI Journal, 11, 5, pp. 57-58.

Georgetown University (1988).Lamb Wave Inspectionfor Large Cracks in CentriUgaily Cast Stainless Steel, EPRINP-5963.

Greenstret. W. L., et al. (1985). Aging and Service Wear of Check Valves Used In Engineered Safety-FeatureSystems of Nuclear Power Plants. NUREG/CR-4302, Vol. 1, ORNL-6193N1, p. 9.

Gregory, R C., et aL (1986). "Uitrasonic Characterization of Centrifugally Cast Stainless Steel Samples," Nonde-structive Evaluation Program: Progress in 1985, EPRI NP43 15-SR. pp. 3-1 to 3-12.

46

Haggag, F. K, et al. (1985). "The Use of Miniaturized Tests to Predict Flow Properties and Estimate FractureToughness in Deformed StedPlates,"InteatlonalConference andElbWon onFadgue, Corrosion Crack-Ing, Fracture Mechanics, and Failure Analysis, SaltLake Clry Utah, December 1985.

Harth, 0. I, and T. P. Sherlock (1985). -Monitoring the Service Induced Damage in Utility Boiler Pressure Vesselsand Piping Systems,"Residul-LWe Assessment, Nondestructive ExmInaton, andNuclear Heat ExchangerMaterials, PVP-Vo. 98-1, The American Society of Mechanical Engineers, New Yoik, pp. 9-23.

Hazelton.W. S., and W. H Koo (1988). TechnicalReport onMaterial Selection andProcessing GuldelinesforBWRCoolant Prssure Boundary Piping, NUREG-0313, Rev. 2.

Henry, L. F., and .V. Ellis (1983). PlastlcReplicationProceduresforDamageAssessment. progress report on EPRIProject RP2253-1.

Hertzberg, R. W. (1983). Fracture Mechanics of Engineering Materials, Second Edition, New Yori John Wiley& Sons, p. 470.

Hilton, S. O., and M. F. Steakley (1986). "Steam Line inspection Pngam for Tennessee Valley Authority's HighTemperature Steam Piping Systems," ERFJU Conference on Lre Extension and Assessment of Fossil Plants,Washington, D.C. June 1986.

Hiser, A. L. (1987). -knsile and J-R Curve Charac n of Thermally Aged Austenitic Cast Stainless SteeV.Properties of Stainless Steeb in Elevated Temperature Service, MPC-Vol. 26, PVP-VoI. 13Z The AmericanSociety of Mechanical Engineers, New York, 1987, pp. 35-48. (Also, see NUREG/CR-5024, 1988.)

uang, F. I. (1986). 'Use of Subsized Specimens for Evaluating the Fracture Toughness of Irradiated Materials,"The Use of Small-Scale Specimensfor Testing Irradiated Matcrial, ASTM STP 888, Philadelphia,pp. 290-304.

Hughes, N. R. W. L. Clarke, and D. E. Delwiche (1982), !Intergianular Stress-Crrosion Cracking Resistance ofAustenitic Stainless Stee Castings," Stainless Steel Castings, ASTM STP, V. 0. Behal and A. S. Melilli(eds.), pp. 26f47.

Janszewski, J. A., et al. (1986). "A Production Approach to On-Sit Metao pic Analysis," EPNJ Cofrenceon eEtension andAssesmentofFossilPlants, WahingtonD.C.Jlame 1986.

Jaske, C. E., et aL. (1986). "Lfe-Assessment of High-Temperature, Fossi-Plant Components," EPRI Corferenceon Wfe Extension andAssessment of Fossil Plants, Washington, D.C., June 1986.

Jeoog, P., and F. Ammirato (1989). Ultrasonic Exmination of Cast Stainless Steel, EP-W-6299.

Kupperman, D. S., K. J. Reirnann, and J. AhWoLqpez (1987), "Ulasonic NDE of Cast Stainless Steel," NDTInternational, 20,3, pp. 145-152.

Landermra, E. L. and W. H. Bamford (1978). "Fracture Toughness and Fatigue Characteristics of CentrfugallyCast Type 316 Stainless Steel Pipe After Simulated Thermal Service Conditions," Ductility and ToughnessConsiderations In Elevated Temperature Service, MPC-8, The American Society of Mechanical Engineers,New York, 1978, pp. 99-127.

Leger, M. T. (1982). Vftdkctg and Evaluating Ferrite Content in Austenitic Stainless Steel Castings," StainlessSteel Castings, ASTM SIP 756, The American Society for Testing and Materials, Philadelphia, pp. 105-125.

Lucas, 0. E, et aL (1986). "Subsized Bend and Charpy V-Notch Specimens for Iraited Testng," The Use ofSmall-Scale Specimensfor Testing Irraaed Material, ASTM STP 888, PhIladelphia, pp. 305-324.

Manahan,M 1P., etal. (1981). -ThDevelopmentofaMiniaturizedDisc BendTestfor Determination ofPostlrradi-ation Mechanica Properties," Journal ofNuclear Materials, 103-104, pp. 1545-1550.

47

MasyainaP.,etal. (1985). "FindingsonCreep-FatigueDamageinPressuePasof Long-Term Seice-ExposedThermal Powea Plants," Journal of Pressure Vessel Technology, 107. pp. 260-270.

McConnell, P., et aL (1986). "Experience in Subsized Specimen Tisting," The Use of Small-Scale SpecimensforTesting Irradiated Material, ASTM STP 888, Philadelphia, pp. 353-368

Mills, W. J. (1987). "Fracture Toughness of Aged Stainless Steel Primary Piping and Reactor Vessel Materials,"Jownal of Pressure Vessel Technologo 109, November, pp. 440-448.

Misawa, T., and Y. Hamaguchl (1986). "Miniature Specimen Tests for Evaluating Fracture Resistance With Radi-ation and Hydrogen Embrittlements in Peritic Steel," The Mechanism of Fracture, American Society forMetals, Metals Park, Ohio, pp. 521-527.

Neubauer, B., and U Wedel (1983). "Restlife Estimation of Creeping Components by Means of Replicas," Ad-vancesin L4fe Prediction Met hods, The American Society of Mechanica Engineers, New York, pp. 307-313.

Padden, T. R. (1983). Metallurgical Characterization of the Creep Ratcheting Test Article Failure,WAESD-HT-94000-28, Westinghouse Electric Corporation. Madison, Pennsylvania.

Padden, T. R., and C. A. Daniels (1976). In Situ Examination of the Small Scale Steam Generator Model, WARD-110134-1, Westinghouse Electric Corporation, Madison, Pennsylvania.

Padden, T. R. and C. A. Daniels (1981). "Replication Manipulator and Procedure for Tube Bore Surfaces,"presented at fte Annual Meeting of Scanning Electron Microscopy, Inc., Dallas, Texas, April 1981.

Peckner, D., and I. M. Bernstein (1977). Handbook of Stainless Steels, New York: McGraw-Hill Book Company.

Reed-Hill, R. E. (1973). Physical Metallurgy Principles, second edition, New York D. Van Nostrand Company,pp.568-610.

Sassen, J. NE, et al. (1987). "Kinetics of Spinodal Decomposition in the Ferrite Phase of a Duplex Stainless Steel,"Properties ofSainless Steels In Elevated Temperature Service, MPC-VoL 26, PVP-%b. 132, The AmericanSociety of Mechanical Engineers, New York, pp. 65-78.

Shah, V. N., and P. E. MacDonald (eds.), (1987). Residual Life Assessment of Major Light Water ReactorComponents-Overview, VoL 1. NUEGCR-4731, EGO-2469, p. 66.

Shah, V. N., and P. E. MacDonald (eds.), (1989). Residual Lfle Assessment of Major Light Water ReactorComponents-Overview, Vol.2, NUREG/CR-4731, EGG-2469.

ShankarRR G. Selby, and P. Jeong (1988). "Defect Detection and Signal Processing for Cast Stainless Steel, Inter-nadonal Journal ofPressure Vessel & Piping, 35, pp. 57-71.

Solomon, IL D., and T. M. Devine, Jr. (1983). -Keynote Address A Tale of Two Phases," Dqplex Stainless Steels(Conferee Proceedings of the American Society for Metals), R. A. Lula (ed.), Metals Park, Ohio 44073:American Society for Metals, pp. 693-756.

Swanson, 0. D. (1981). Development ofFracture Toughness Measurement Capability by Ultrasound, Fmial ReportBDX-613-2625, Bendix, Kansas City Division.

Taldgawa, T.. at aL (1986). "Life Assessment and Extension of Steam Turine Rotor -Overview of MH's Activi-des," EPRN Coaierence on 4fc Extension and Assessment of Fossil Plants, Washington. D. C.. June, 1986.

Tautwein, A., and W. Gysel (1982). "Influence of Long-Time Aging of CF8 and CF8M Cast Steel at TUmpeaturesBetween 300 and 5000C on Impact Tbughness and Structural Properties:' Stainless Steel Castings, ASTMSTP 756, Philadelphla, pp. 165-189.

48

Umino, D. T., and A. K. Rao (1984).Long-TermlnspectionRequdrementsforPWRPamzp Casings, EPRI NP-3491.

USNRC (1978). "Control of Farite Content in Stainless Steel Weld Metal. Regulatory Guide 131, Revision 3.

USNRC (1988a). Licensed Operating Reactors Status Sunmary Report Data as of 05-31-88, NUREG-0020,Vol. 12, No. 6.

USNRC (1988b). Ofrice of Nuclear Reactor Regulation, Thermal Stresses In Piping Connected To Reactor CoolantSystems, NRC Bulletin No. 88-08, Supplemnent 2, August 4.

Vary, A. (1976). "'The Feasibility of Ranking Material Facte lboughness by Ultaonic Attenuation Measue-ments," Journal of Testing and Evaluation, 4, 4, pp. 25 1-256.

Viswanathan, R. (1985). '"Dissimilar Metal Weld and Boiler Creep Damage Evaluation for Plant Life Extension,"Residual-Life Assessment, Nondestructive Examination, and Nuclear Heat Exchanger Materials,PVP-VoL 98-1, The American Society of Mechanical Engineers, New York, pp. 9-18.

49

NRC FORM 235 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMBER2I1102(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 YEAROctober 1990

Cast Stainless Steel Components 4 FIN OR GRANT NUMBER

A6389S. AUTHOR(S) 6 TYPE OF REPORT

Technical

C.E. Jaske, V.N. Shah 7.PERIODOOVERED fa^#Do:e")

. PEfORMNC ORGANIWZTIN -tME AND ADDRESS 1 Ax ro Ac er w gan and mawvng adat

Idaho National Engineering LaboratoryEG&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 EngineeringOffice of Nuclear Regulatory ResearchU.S. Nuclear Regulatory CommissionWashington.- DC 20555

10. SUPPLEMENTARY NOTES

13. ABSTRACT (200 words or lSss)

This report presents a procedure for estimating the current conditionand residual life of safety-related cast stainless steel components in lightwater reactors (LWRs). The procedure accounts for loss of fractures tough-ness caused by thermal embrittlement and includes the following: a review ofdesign and fabrication records, inservice inspection records, and operatinghistory; a fracture mechanics evaluation to determine the required toughnessestimates; and criteria regarding continued service, repair, or replacementof the component being evaluated. The report discusses the available CharpyV-notch impact energy, fracture toughness, tensile strength, fatigue resis-tance, and fatigue-crack growth data, and presents two methods for assessingthe degree of thermal embrittlement: metallurgical evaluation and analyti-cal modeling of inservice degradation.

12. KEY WOROSIOESCRIPTORS (Lt ods or phases ihat f assis reseathers In ocaflg p 1 unlimitedMENT

LWR cast stainless steel componentsThermal enbrittlement 14. SECURIY CLASSFATM

Charpy impact energy unclassifiedFracture toughness AxLife assessment procedure unclassifiedInservice inspections 1s. NUMBER OF PAGES

Fatigueis. PRICE

NRC FORM 335 (2.89