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TLR-RES-DE-2023-14 Thermal Aging of Assw - Final
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TECHNICAL LETTER REPORT TLR-RES/DE-2023-14-Rev 0

Assessment of Thermal Aging Embrittlement of Austenitic Stainless Steel Weld Metals

Prepared by:

J. Poehler, DE/MEB P. Purtscher, DE/MEB

POC:

J. Poehler, DE/MEB

Date:

August 13, 2024

DISCLAIMER: This report does not contain or imply legally binding requirements. Nor does this report establish or modify any regulatory guidance or positions of the U S Nuclear Regulatory Commission and is not binding on the Commission.

ii DISCLAIMER: This report was prepared as an account of work sponsored by an agency of the US Government. Neither the US Government nor any agency thereof, nor any employee, makes any warranty, expressed or implied, or assumes any legal liability or responsibility for any third party's use, or the results of such use, of any information, apparatus, product, or process disclosed in this publication, or represents that its use by such third party complies with applicable law.

iii ABSTRACT

This report summarizes a review by the Nuclear Regulatory Commission (NRC) staff of the current state of knowledge of the effects of thermal aging embrittlement (TE) on austenitic stainless steel weld (ASSW) materials used in light water reactor (LWR) piping, with regard to aging mechanisms and effects relevant to long-term operation. The report focuses on the effects of TE on the mechanical properties, particularly the fracture toughness, of ASSW materials. Consideration of the effects of irradiation embrittlement (IE) on fracture toughness, or the combined effects of TE and IE, is outside th e scope of this report. ASSW materials commonly used in nuclear power plant piping in the US include Type 308, Type 316, and Type 309. ASSW are subject to loss of fracture toughness after exposure to LWR operating temperatures over the operating life of reactors, e.g. 290 °C for 40 or more years. The report also reviews how TE of ASSW is addressed in structural integrity evaluation methods, such as the American Society of Mechanical Engineers (ASME) Code flaw evaluation provisions for piping.The report examines the guidance on aging management of ASSW in NUREG-2191, the Generic Aging Lessons Learned Report for Subsequent License Renewal (GALL-SLR).

Significant findings of this work include that TE results in a reduction in fracture toughness relative to unaged material, with the degree of loss of fracture toughness dependent on the weld process. Submerged arc welds (SAW ) and shielded metal arc welds (SMAW) tend to have lower aged fracture toughness than gas tungsten arc welds (GTAW).The report suggests additional confirmatory testing of the lower bound fracture toughness curve for GTAW.

Additional testing would also be needed to establish the effects of LWR coolant environments on thermally aged fracture toughness, the effect of TE on low -cycle fatigue crack growth rates, and to support development of more refined models for thermally aged fracture toughness of ASSW. The report also concludes that the current ASME Code flaw evaluation provisions do not account for the effect of TE on fracture toughness of ASSW.The report notes that ASSW are not identified as a separate component or material in the GALL-SLR, but are treated the same as the piping base material. Therefore, TE of ASSW is not currently addressed, but cracking of stainless steel piping (including ASSW) is managed by several aging management programs (AMPs) in NRC license renewal guidance.

iv TABLE OF CONTENTS

ABSTRACT................................................................................................................................. iv

LIST OF FIGURES..................................................................................................................... vii

LIST OF TABLES...................................................................................................................... viii

EXECUTIVE

SUMMARY

............................................................................................................. ix

ABBREVIATIONS AND ACRONYMS......................................................................................... x

1 INTRODUCTION...................................................................................................................... 1 1.1 Background....................................................................................................................... 1 1.2 Weld Materials of Interest for Operating Reactors............................................................ 1

2 METALLURGICAL CONSIDERATIONS................................................................................. 3 2.1 Ferrite Limits in ASSW...................................................................................................... 3 2.2 Determining ferrite content in ASSW................................................................................ 3 2.3 Kinetics of Thermal Embrittlement.................................................................................... 3

3 MECHANICAL PROPERTIES of thermally aged austenitic stainless steel welds............ 7 3.1 Impact and Tensile Properties.......................................................................................... 7 3.1.1 Types of Mechanical Property Data.................................................................... 7 3.1.2 Effect of Ferrite Content...................................................................................... 8 3.1.3 Effect of Aging on Tensile Properties.................................................................. 8 3.2 Fracture Toughness.......................................................................................................... 8 3.2.1 General............................................................................................................... 8 3.2.2 NUREG/CR-6428, Revision 1............................................................................. 9 3.2.3 Effect of Test Temperature on Fracture Toughness of Thermally Aged Materials.............................................................................................................. 9 3.2.4 Fracture Toughness of SMAW and SAW (Flux) Welds................................... 11 3.2.4.1 Unaged SMAW and SAW Welds................................................................... 11 3.2.4.2 Aged SMAW and SAW Welds....................................................................... 12 3.2.4.3 Other Fracture Toughness Data for Aged SMAW and SAW Welds............. 14 3.2.4.4 Summary Discussion - SMAW and SAW Welds.......................................... 14 3.2.5 Fracture Toughness of GTAW (non -flux) Welds............................................... 15 3.2.5.1 Unaged GTAW Welds.................................................................................... 15 3.2.5.2 Aged GTAW Welds........................................................................................ 16 3.2.5.3 Other Fracture Toughness Data for Aged GTAW Welds.............................. 17 3.2.5.4 Summary Discussion - GTAW....................................................................... 20 3.2.6 Effects of the Environment on Fracture Toughness of ASSW.......................... 21 3.2.7 Discussion......................................................................................................... 21 3.2.8 Effect of Ferrite Content.................................................................................... 21 3.2.9 Effect of Aging Temperature............................................................................. 22 3.2.10 Summary Discussion - Fracture Toughness.................................................... 24

v 4 CURRENT AGING MANAGEMENT GUIDANCE FOR REDUCTION IN FRACTURE TOUGHNESS OF SS............................................................................................................. 25 4.1 Aging Management Guidance for ASSW........................................................................ 25 4.2 Aging Management Guidance for Reduction of Fracture Toughness of WASS.............. 25 4.3 Operating Experience..................................................................................................... 25 4.4 Potential Changes to Aging Management Guidance for ASSW...................................... 26 4.4.1 Detection of Aging Effects................................................................................. 27 4.4.2 Flaw Evaluation................................................................................................. 27 4.5 Summary......................................................................................................................... 27

5 EVALUATION METHODOLOGIES....................................................................................... 29 5.1 Handling of ASSW Fracture Toughness and Flaw Evaluation in the ASME Code......... 29 5.1.1 ASME Code,Section XI, Nonmandatory Appendix C, Analytical Evaluation of Flaws in Piping............................................................................ 29 5.1.2 Code Case N-906............................................................................................. 31 5.1.3 Conclusions - ASSW Fracture Toughness in the ASME Code,Section XI....................................................................................................................... 32 5.2 Probabilistic Methods...................................................................................................... 32

6 EFFECT OF THERMAL AGING ON OTHER AGING MECHANISMS.................................. 34 6.1 Effect Of Thermal Aging On SCC................................................................................... 34 6.2 Effect of TE on Fatigue................................................................................................... 35

7 CONCLUSIONS..................................................................................................................... 36

8 REFERENCES....................................................................................................................... 38

APPENDIX A Determining Ferrite Content of ASSW........................................................... A-1 A.1 Methods of Determination of Ferrite Content....................................................................... A-1 A.2 Estimation from Chemical Composition................................................................................ A-1 A.3 Measurement using Magnetic Instrument or Feritscope...................................................... A-6 A.4 Estimation of Ferrite Content using Quantitative Metallography.......................................... A-6 A.5 Summary - Determining Ferrite Content.............................................................................. A-6

vi LIST OF FIGURES

Figure 2-1 Fracture toughness of Type 316L GTAW Weld Material with 8% ferrite as a function of aging time, from Hojo, 2014. a. J IC ; b. J6mm.................................... 4 Figure 3-1 Correlations of C and n for J-R curve with CVN energy at 320 ° C (Figs. 9 and 10 from Saillet et. al., 2020)........................................................................... 7 Figure 3-2 Modified J-R Curves for Type 304 SAW weld metal from EPRI NP -4768........... 10 Figure 3-3 Modified J-R curves for Type 316 SAW Weld Metal from EPRI NP -4768........... 10 Figure 3-4 J value at a crack extension of 1mm for unaged SMAW and SAW Welds, from several different sources............................................................................. 11 Figure 3-5 Fracture toughness J-R curves for SMAW weld metal evaluated in NUREG/CR-6428, Rev. 1.................................................................................... 13 Figure 3-6 Comparison of J-R curves for archival Type 308L SMAW weld metal to NUREG/CR-6428, Rev. 1 lower bound curve for SAW and SMAW weld metal.................................................................................................................... 14 Figure 3-7 J value at a crack extension of 1 mm for unaged GTAW welds, from several different sources..................................................................................... 15 Figure 3-8 Fracture toughness J-R curves for austenitic stainless steel GTAW welds from NUREG/CR-6428, Rev. 1............................................................................ 16 Figure 3-9 J-R Curves for as -welded and aged Type 316L weld filler metal. a) Tested at room temperature; b) Tested at 320 ° C........................................................... 18 Figure 3-10 J-R curves for GTAW Welds from Koyama et. al., 1999 Hojo, 2014 and Hong et. al., 2018, compared with the NUREG/CR -6428, Rev. 1 lower bound for GTAW welds....................................................................................... 20 Figure 3-11 Effects of aging at 343 °C on the J -Integral fracture toughness of type 308 stainless steel weld metal. All tests conducted with precracked CVN specimens at 290 °C. (From NRC, 2000b).......................................................... 22 Figure 3-12 J6mm for Type 316L materials with different aging temperatures and the same aging parameter. a) GTAW weld; b) SMAW weld.................................... 23 Figure 5-1 Comparison of J 2.5(J0.1) for CASS with other materials in ASME Code Section XI (From ASME, 2017, Fig. 8)............................................................................. 30 Figure 5-2 Fracture toughness lower bound J -R curves and the data on Type 304, 316L,and CF-3 welds used to develop the ASME Code IWB -3640 analysis. (from NUREG/CR-6428. Rev. 1).......................................................... 31

vii LIST OF TABLES

Table 2-1 Equivalent Aging Times at 400 °C to Represent LWR Piping Temperatures............... 6

viii EXECUTIVE

SUMMARY

Austenitic stainless steel welds (ASSW) are used to join wrought austenitic stainless steel (WASS) piping in operating light water reactors (LWRs). ASSW are duplex alloys containing both austenite and ferrite, similar to cast austenitic stainless steels (CASS). Similar to CASS, ASSW are susceptible to thermal aging embrittlement (TE) which can lead to a reduction of the fracture toughness of material after prolonged service at reactor operating temperatures. The NRC staff reviewed the available mechanical property data for thermally aged ASSW, focusing on fracture toughness data, to determine the potential significance of loss of fracture toughness of these welds. Much of the available fracture toughness data is summarized in Effects of Thermal Aging on Fracture Toughness and Charpy -Impact Strength of Stainless Steel Pipe Welds. NUREG/CR-6428, Revision 1 ( NUREG/CR-6428, Rev. 1). The staff also reviewed flaw evaluation procedures applicable to ASSW in Section XI of the A merican Society of Mechanical Engineers (ASME) Code, to determine if these procedures account for the reduction in fracture toughness due to TE. The staff also reviewed the aging management guidance in the GALL-SLR to determine if TE of ASSW is addressed in this guidance.The conclusions of this report can be used to support potential future research on TE of ASSW, potential future changes to aging management of ASSW in the GALL-SLR, and potential future recommendations for changes to the ASME Code flaw evaluation procedures applicable to ASSW. Consideration of the effects of irradiation embrittlement (IE) on fracture toughness, or the combined effects of TE and IE, is outside the scope of this report.

Significant conclusions of the NRC staffs research include:

Thermal aging of ASSW at LWR operating temperatures causes a significant reduction in fracture toughness, with GTAW welds having higher thermally aged toughness than SAW/SMAW welds.

Additional/confirmatory research is recommended to confirm the lower bound toughness for GTAW, establish environmental effects on thermally aged toughness, and confirm that lower bound toughness curves are appropriate for operation beyond 80 years.

The flaw evaluation methods of the ASME Code, Appendix C, do not account for thermal aging of ASSW, and revision of these methods should be considered.

Thermal aging of ASSW is not explicitly considered in NRC aging management guidance in the GALL-SLR, although aging management programs do generically address cracking of ASSW welds.

ix ABBREVIATIONS AND ACRONYMS

AERM aging effect requiring management ASME American Society of Mechanical Engineers ASSW austenitic stainless steel weld BWR boiling water reactor CASS cast austenitic stainless steel EPRI Electric Power Research Institute FN ferrite number GALL Generic Aging Lessons Learned (report)

GALL-SLR Generic Aging Lessons Learned for Subsequent License Renewal (report)

GMAW Gas metal arc welding GTAW gas tungsten arc welding HAZ heat affected zone HWC Hydrogen Water Chemistry (BWR)

IE irradiation embrittlement IGSCC intergranular stress corrosion cracking L-B lower bound LCF Low-Cycle Fatigue LWR light water reactor MMA Manual metal arc welding NRC Nuclear Regulatory Commission NWC Normal Water Chemistry (BWR)

PWR pressurized water reactor TE thermal aging embrittlement SAW submerged arc welding SCC stress corrosion cracking SMAW shielded metal arc welding TIG Tungsten inert gas welding (synonymous to GTAW)

TLAA time limited aging analysis WASS wrought austenitic stainless steel WRC Welding Research Council

x 1 INTRODUCTION

1.1 Background

Austenitic stainless steel welds (ASSW) are present in both boiling water reactor ( BWR) and pressurized water reactor (PWR) plants. These welds typically join piping made from wrought austenitic stainless steel (WASS) or cast austenitic stainless steel (CASS), and in some cases may be used in dissimilar metal weld joints between low-alloy steels and stainless steels.

Common weld filler metals include Type 308, 308L, 316, and 316L stainless steels. These weld fillers are similar in composition to WASS and CASS. Like CASS, ASSW have a duplex phase composition with some portion being ferrite (typically 5-15%), while most of the microstructure consists of austenite. As a result of containing ferrite, like CASS, ASSW are potentially susceptible to thermal aging embrittlement (TE) during long-term service in light water reactor components that have a service temperature of 250 °C (482 °F), or greater. With regard to TE, the "Generic Aging Lessons Learned for Subsequent License Renewal Report Final Report" Volume 2 (GALL-SLR, V2) states [ it is]:

Also termed thermal aging or thermal embrittlement. At operating temperatures of 260 to 343 °C [500 to 650 °F], CASS exhibit a spinodal decomposition of the ferrite phase into ferrite-rich and chromium-rich phases. This may give rise to significant embrittlement (reduction in fracture toughness), depending on the amount, morphology, and distribution of the ferrite phase and the composition of the steel.

Thermal aging refers to the microstructural changes that occur as a result of exposure of the material to elevated temperatures for a period of time. Thermal aging does not always result in loss of fracture toughness, so the occurrence of thermal aging does not always mean that TE is occurring. There is currently no specific guidance in NUREG -2191, "Generic Aging Lessons Learned for Subsequent License Renewal Report Final Report" (GALL -SLR, V1, GALL-SLR, V2) regarding aging management of TE of ASSW. The purpose of this document is to summarize the state of knowledge with respect to TE of ASSW, review aging management of ASSW in the GALL-SLR, and review the ASME Code flaw evaluation procedures applicable to ASSW. The conclusions of this report can be used to support potential future research on TE of ASSW, potential future changes to aging management of ASSW in the GALL-SLR, and potential future recommendations for changes to the ASME Code flaw evaluation procedures applicable to ASSW.

This technical letter report (TLR ) focuses on changes to the fracture toughness of ASSW caused by TE, and also addresses changes to the stress corrosion cracking (SCC) growth rates caused by TE.Irradiation embrittlement (IE) of ASSW, or the combined effects of TE and IE on ASSW, are outside the scope of this report.

1.2 Weld Materials of Interest for Operating Reactors

In US BWR and PWR plants, the most common stainless steels used for reactor coolant system (and other connected system) piping are Type 304, 304L, 316, and 316L. Matching filler metals for these alloys (respectively) are Type 308, 308L, 316, and 316L.

EPRI NP-4768 (EPRI, 1986) lists the weld materials used in BWR stainless steel welding practice. In BWRs, the gas tungsten arc welding ( GTAW) process (also known as tungsten

1 inert gas (TIG) welding) 1 was used for root passes in piping, with the finish welding being done by either the SAW or SMAW weld processes, except for automatic field welds which used the GTAW process.

Type 309 contains higher chromium and nickel content than Type 308 and is used to make dissimilar metal welds of WASS to low-alloy or carbon steels, with the higher alloying compensating for dilution into the low-alloy /carbon steel base metal. Type 16-8 -2 is a weld filler of interest for high-temperature reactors due to its more advantageous creep properties.

American Society of Mechanical Engineers (ASME) weld material specifications use different designations for weld rods and electrodes used with different processes. For example, for bare stainless steel electrodes used in GTAW, gas metal arc welds (GMAW), and submerged arc (SAW ) welds, Type 308 rods are designated ER308 in specification SFA -5.9/SFA -5.9M, Specification for Bare Stainless Steel Welding Electrodes and Rods. (ASME, 2023a)

Electrodes for SMAW are specified in SFA 5.4/SFA 5.4M, Specification for Stainless Steel Electrodes for Shielded Metal Arc Welding. (ASME, 2023b) Type 308 electrodes are designated E308-XX, with XX having different values depending on the type of welding current and position. In this report, the designation based on the chemical composition is used (e.g.

Type 308) but the ASME designations for the weld rod or electrode are not generally used.

1 In this report, the term GTAW is used. If a reference used the term TIG, it is referred to as GTAW in this report.

The ASME Code,Section XI, Appendix C uses the term nonflux welds, of which GTAW would be a subset.

2 2 METALLURGICAL CONSIDERATIONS

2.1 Ferrite Limits in ASSW

NUREG/CR-6428, Rev. 1 indicates that ferrite in ASSW is considered desirable and beneficial for increasing the tensile strength, for improved resistance to stress corrosion cracking (SCC),

and also for prevention of microfissuring (hot cracking) upon solidification (NRC, 2018). NRC Regulatory Guide 1.31, Control of Ferrite Content in Stainless Steel Weld Metal, (NRC, 2013) specifies a minimum ferrite content of 5% and a maximum ferrite content of 20%. NRC, 2013 states that the [20%] upper limit provides a ferrite content adequate to offset dilution and reduce thermal aging effects.For intergranular stress corrosion cracking (IGSCC) resistance of BWR piping, NUREG-0313 (NRC, 1988) recommends a minimum ferrite content of 7.5%, in conjunction with low-carbon weld filler, for Types 308L, 316L, 309L, and similar grades.The reason for the difference in the minimum recommended ferrite between NRC, 1988 and NRC, 2013 appears to be because the 5% minimum in NRC, 2013 is intended to prevent microfissuring during welding, while the minimum of 7.5% ferrite recommended by NRC, 1988 is for resistance to IGSCC.

2.2 Determining FerriteCcontent in ASSW

Ferrite content in ASSW may be determined in several ways:

Estimate from chemical composition using Hulls factors or other formulae.

Measured using magnetic instrument or ferritescope.

Estimate using quantitative metallography.

Appendix A contains an evaluation of these various methods, including an evaluation of several calculational methods including Hulls.Based on the evaluation in Appendix A, the staff concludes that the Welding Research Council (WRC) WRC -1992 method for estimation of ferrite content from chemical composition is the most accurate method, because

The WRC-1992 method provides balanced, slightly conservative predictions up to 12%

measured ferrite, and has the smallest root mean square deviation ( RMSD).

The WRC-1992 methodology is incorporated into the ASME Code,Section III in Figure NB-2433.1-1 and is referenced in RG 1.31 (NRC, 2013).

With respect to measurement methods, the staff recommends use of a magnetic instrument or ferritescope, rather than quantitative metallography, since the latter method is tedious and time consuming.

2.3 Kinetics of Thermal Embrittlement

Thermal aging in the laboratory is typically conducted at higher temperatures than LWR service temperatures to allow for reasonably limited laboratory aging times. NUREG/CR-6428, Rev. 1 notes that the kinetics of thermal embrittlement in CASS materials are controlled by three processes: spinodal decomposition, precipitation and growth of grain boundary carbides, and precipitation of G ph ase in ferri te. Small changes in the composition may cause the kinetics to vary significantly.

TE is a diffusion-controlled process that tends to slow down as aging time increases, beyond a certain point. This effect is known as saturation. This effect can be seen in Figure 2-6, which shows fracture toughness of Type 316L GTAW weld metal with 8% ferrite as a function of aging time for JIC and J at a crack extension of 6 mm (J6mm). In this case, JIc can be considered to have saturated after 300 hours0.00347 days 0.0833 hours 4.960317e-4 weeks 1.1415e-4 months at 400°C, since it maintains a constant minimum level. By contrast, the J6mm does not appear to have saturated after 60,000 hours0 days 0 hours 0 weeks 0 months at 400°C, since there is still a downward trend with respect to the log of aging time.

aged at 400 deg C

600

500

400

300

200

100

0 1 10 100 1000 10000 100000 Aging time (hours) a.

aged at 400 deg C

1600 1400 1200 1000 800 600 400 200 0

1 10 100 1000 10000 100000 Aging time (hours) b.

Figure 2-1 Fracture toughness of Type 316L GTAW Weld Material with 8% ferriteas a function of aging time, from Hojo, 2014. a. J IC ; b. J6mm

To establish equivalency of the amount of aging that will occur at one temperature versus another, an aging parameter P is defined in NUREG/CR-6428, Rev. 1:

= log() 100019.143 1+273 1673 Eqn. 2-1 4

Where:

= ()

=

= (°)

P provides a method of comparing equivalent aging for different temperature and times normalized to aging at 400 °C. Estimates of activation energies range from 65-230 kJ/mol (NUREG/CR-6428, Rev. 1). Much of the testing for thermal aging uses an aging temperature of 400 °C for 10,000 hours0 days 0 hours 0 weeks 0 months , which has a P value of 4. NUREG/CR-6428, Rev. 1 indicates that studies on low-temperature thermal aging of Types 304L and 316L SS welds containing about 10% ferrite yielded an activation energy of 113 kJ/mol for the thermal embrittlement of Type 304L weld in the range of 335-400°C. An activation energy of 100 kJ/mol has been used in Japan for some studies of the fracture toughness of thermally aged CASS (Kawaguchi et. al.,

2005). Due to the similar duplex microstructures of CASS and ASSW, it is reasonable to use a similar activation energy. Higher activation energies result in a lower P -number for the same aging time and temperature. Aging for 10,000 hours0 days 0 hours 0 weeks 0 months at 400 °C is equivalent to aging for 517,263 hours0.00304 days 0.0731 hours 4.348545e-4 weeks 1.000715e-4 months or 59 effective full power years at 290 ° C, which is representative of PWR cold leg temperature or BWR reactor recirculation temperature. However, for higher temperature components, such as PWR hot leg or pressurizer piping, aging at 400 ° C for 10,000 hours0 days 0 hours 0 weeks 0 months (P =

4) may not represent the full amount of thermal aging that the component would experience.

Further, a higher P-number would be needed to represent operation times greater than 59 EFPY at 290 °C. For example, 72 EFPY at 290 ° C equates to a P-number of 4.086 (using an activation energy of 113 kJ/mol) which would be equal to 12,200 hours0.00231 days 0.0556 hours 3.306878e-4 weeks 7.61e-5 months at 400 ° C. When performing accelerated aging, if saturation can be demonstrated at a lower P -number than the P-number representing the target service time and temperature, shorter aging times can possibly be used.

Table 2-2 shows the P-number for various service temperatures characteristic of LWR piping, and the equivalent P-number for service times corresponding to 54 or 72 EFPY, which would be typical for plants operating out to either 60 or 80 calendar years. Table 2-2 also shows the required accelerated aging time at 400 ° C to achieve the same P -number. Table 2-1 shows that for components with service temperatures higher than 290 ° C, aging for 10,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C will not be adequate.P-numbers for two different activation energies are shown in Table 2-2.

The lower activation energy of 100 kJ/m2 results in higher P-numbers for the same aging time and requires longer accelerated aging times, given the same operating conditions.

The highest service temperature for PWR CASS components is for components in pressurizer surge lines operating at 343 °C. For 72 EFPY, this equates to a P -number of 4.99. The equivalent aging time at 400 °C is 97,325 hours0.00376 days 0.0903 hours 5.373677e-4 weeks 1.236625e-4 months . Hojo et. al., 2014 documents testing of materials aged for 60,000 hours0 days 0 hours 0 weeks 0 months at 400 °C which equates to a P value of 4.8. This would be equivalent to about 36 EFPY at 343 °C, assuming an activation energy of 100 kJ/m 2. Per Table 2-1, an aging time of 97,325 hours0.00376 days 0.0903 hours 5.373677e-4 weeks 1.236625e-4 months at 400 °C would be needed to represent 72 EFPY at 343 °C.

Aging at 450 °C results in more reasonable aging times to achieve a P -number equivalent to 80 years of operation. For example, an aging time of 24,080 hours9.259259e-4 days 0.0222 hours 1.322751e-4 weeks 3.044e-5 months at 450 ° C would result in a P-number of 4.99 equivalent to 72 EFPY of operation at 343 °C.However, aging at 450 °C could potentially result in microstructural changes that do not occur at 400 °C and below so accelerated aging at this temperature may result in overly conservative predictions of toughness.

5 Table 2-1 Equivalent Aging Times at 400 °C to Represent LWR Piping Temperatures

Service Piping EFPY Service Activation P Accelerated Temp. Type Time (hr) Energy(kJ/mol) Aging Time (hr) 290 BWR, 54 473040 100 4.16 14399 PWR CL 320 PWR HL 54 473040 100 4.63 42437 343 PWR 54 473040 100 4.96 90502 PZR 290 BWR, 72 630720 100 4.28 19199 PWR CL 320 PWR HL 72 630720 100 4.75 56583 343 PWR 72 630720 100 5.08 120669 PZR 290 BWR, 54 473040 113 3.96 9145 PWR CL 320 PWR HL 54 473040 113 4.49 31018 343 PWR 54 473040 113 4.86 72994 PZR 290 BWR, 72 630720 113 4.09 12193 PWR CL 320 PWR HL 72 630720 113 4.62 41357 343 PWR 72 630720 113 4.99 97325 PZR

To summarize, longer laboratory aging times than 10,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C will likely be needed to represent the aging that will occur in LWR piping components for extended operation, particularly for operation out to 60 or 80 years.The duration and time of accelerated aging needed are dependent of the activation energy, which varies with both chemical composition and aging temperature.Shorter aging times could be acceptable if saturation can be demonstrated at a lower P-number than that representative of the full service time and temperature.

Another potential source of ASSW materials with long aging times would be piping materials harvested from operating or decommissioning nuclear power plants. These materials would also be aged at temperatures characteristic of in-service exposure in LWR plants, thus any artificial effects of laboratory aging at temperatures higher than normal LWR temperatures would not be present.

6 3 MECHANICAL PROPERTIES OF THERMALLY AGED AUSTENITIC STAINLESS STEEL WELDS

3.1 Impact and Tensile Properties

3.1.1 Types of Mechanical Property Data Several investigators (NUREG/CR-6428, Rev. 1, I-NERI, 2017, and Chen et. al., 2019b) have tested various mechanical properties of both unaged (as -welded) and thermally aged welds.

These include tensile properties (ultimate tensile strength and yield strength), Charpy V -notch (CVN) impact strength, and J-R fracture resistance. These tests have been conducted both at room temperature and elevated temperatures corresponding to LWR operating temperatures.

Fracture toughness (J-R fracture resistance) data is most useful since only fracture toughness data can be used directly in fracture mechanics evaluations of flaws.Therefore, this report focuses on fracture toughness test data for thermally aged ASSW.

Generally, mechanical property testing shows that thermal aging of ASSW causes an indeterminate effect on tensile and yield strength, an increase in the CVN transition temperature, a decrease in CVN upper shelf energy, and a reduction in J -R fracture resistance.

For CASS materials, Saillet et. al., 2020 used the prediction of the absorbed energy at both room temperature (RT) and operating temperature to correlate to the C and n parameters that describe the J-R curve from fracture toughness tests (Figure 3-1). There is some scatter, but a reasonable fit of the J at 1 mm (same as C parameter in power law fit to the J -R curve) can be defined as a function of the CVN absorbed energy when both are measured at elevated temperature. The n parameter in power law fit to the J -R curve is a weak function of the absorbed energy.

Figure 3-1 Correlations of C and n for J-R curve with CVN energy at 320 °C (Figs. 9 and 10 from Saillet et. al., 2020).

7 3.1.2 Effect of Ferrite Content NUREG/CR-6628 ( NRC, 2000b) des cribes CVN testing of Type 308 MMA (SMAW) welds with three different ferrite contents (4, 8, and 12%) as a function of aging time at 343 ° C. The results show that as ferrite increases, the effects of increasing aging on the CVN energy are more pronounced, with larger decreases observed as ferrite increase. Lucas, 2011 presents results of CVN tests of Type 316L weld material with 10% and 14% ferrite.The material was tested at 0, 5000, and 10,000 hours0 days 0 hours 0 weeks 0 months of aging at 300 ° C, 400 ° C, and 430 °C. For elevated temperature (288 °C) tests, the decrease in CVN energy relative to the unaged value for the same aging time was greater for the higher ferrite material, although there was often a slight increase from 5,000 to 10,000 hours0 days 0 hours 0 weeks 0 months . Based on fractography, Lucas noted that while at 5,000 hours0 days 0 hours 0 weeks 0 months spinodal decomposition was complete, at 10,000 h the material is regaining toughness and demonstrated complete ductile failure in both phases throughout the fracture surface. Lucas further noted that this is likely due to the diffusion of interstitial carbon and nitrogen from the ferrite, precipitating as carbides, mitigating the loss of toughness caused by spinodal decomposition. Lucas et. al.,

2016, shows additional results for the same weld materials. At 288 °C, high ferrite (14%)

material aged at 430 °C loses more CVN energy at 10,000 hours0 days 0 hours 0 weeks 0 months than low ferrite (10%) material, and does not see the recovery in CVN energy from 5,000 to 10,000 hours0 days 0 hours 0 weeks 0 months that is seen for material aged at 400 °C.For both high and low ferrite materials, the reduction in CVN energy from aging at 300 °C was considerably less than at higher temperatures.

NUREG/CR-6428, Rev. 1 also discusses the effects of aging on Type 19-9L MMA (SMAW) welds with 5-9% ferrite and Type 308 SMAW welds with 4, 8, and 12% ferrite. These results showed slight increases in tensile strength with aging time and ferrite content (for the Type 308 welds) and no increase in yield strength.

3.1.3 Effect of Agingon Tensile Properties NUREG/CR-6428, Rev. 1 presents tensile and yield strength data (estimated from CVN results) as a function of test temperature, for Type 308 and 316 using SMAW, SAW and GTAW processes, for both aged (10,000 hours0 days 0 hours 0 weeks 0 months at 400 °C) and unaged welds. The results showed no clear effect of aging on these tensile properties.

I-NERI, 2017 showed increases in yield and tensile strengthwith increased aging time up to 10,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C, and decreases in elongation with increased aging time, for Type 316L and Type 347 weld metals.These materials had 11% and 10% ferrite, respectively.

Overall, the effect of thermal aging on tensile properties is not clear. NUREG/CR-6428, Rev. 1 contains much more tensile property data than I-NERI, 2017, and showed no significant effect of aging on tensile properties. I-NERI, 2017 showed increases in ultimate tensile strength and yield strength. As detailed in Section 3.2.5.3 I-NERI, 2017 also documented lower J -R toughness curves for GTAW welds than the bulk of the data used to develop the lower bound J-R curve for GTAW in NUREG/CR-6428, Rev. 1.

3.2 Fracture Toughness

3.2.1 General The ASM Handbook Volume 19: Fatigue and Fracture (ASM, 1996) provides a good summary of the fracture toughness behavior of ASSW, while NUREG/CR-6428, Rev. 1 summarizes the fracture behavior of ASSW based on data in the PIFRAC database. At both ambient and elevated temperatures, fracture toughness is dependent on the weld process, with the order from highest to lowest toughness: 1) GTAW welds (highest), 2) gas metal arc welds (GMAW), 3) 8 SMAW, 4) submerged arc (SAW) welds, and 5) Flux-cored arc weld (FCA) (ASM, 1996). This is also consistent with results from the PIFRAC database which showed GTAW welds have higher toughness than SMAW or SAW welds, although there was no statistical difference between SMAW and SAW welds (NUREG/CR-6428, Rev. 1). The welding processes most commonly used in operating light water reactors are GTAW, SMAW, and SAW welding. ASM, 1996 further indicates that fracture processes are controlled by the density of inclusions rich in manganese and silicon, which is much higher in SMAW and SAW welds than in GTAW welds.

Generally, it has been found that welds with high unaged toughness, such as GTAW welds, experience a greater percentage of toughness loss on aging than welds with low unaged toughness, such as SMAW or SAW welds, although the aged toughness of GTAW welds still remains higher than that of SMAW or SAW welds.

3.2.2 NUREG/CR-6428, Revision 1 At present, the most comprehensive evaluation of the fracture toughness of thermally aged ASSW is contained in NUREG/CR-6428, Revision 1 ( NUREG/CR-6428, Rev. 1), which provides lower bound J-R curves for austenitic stainless steel GTAW, SMAW, and SAW welds, based on available data. The lower bound curve equation is the same for SMAW and SAW welds. The lower bound curves are not specific to delta ferrite content or chemistry. The data covered ferrite contents approximately from 5-11%.The discussion in the following sections provides more detail on the fracture toughness of welds made using the different weld processes from NUREG/CR-6428, Rev. 1, as well as other sources of data.

3.2.3 Effect of Test Temperature on Fracture Toughness of Thermally Aged Materials NUREG/CR-6428 indicates that unaged ASSW tested at reactor temperatures generally have lower fracture toughness than unaged ASSW tested at room temperature. For BWRs, reactor temperature is nominally 282 ° C, and PWR primary coolant piping temperatures range from 288-327 °C (NUREG/CR-6428, Rev. 1). NUREG/CR-6428, Rev. 1 did not contain J-R curves for ASSW tested at room temperature, only at reactor temperatures and higher.

Toughness of Austenitic Stainless Steel Pipe Welds, EPRI NP -4768 (EPRI, 1986b) contains results for ASSW J-R tests of the same material heats tested at both room temperature and reactor operating temperature. These results are contained in Figure 3-2 and Figure 3-3,

showing J-R curves for the same weld materials tested at 75 °F (24 ° C) and 550 ° F (288 °C).

The J2.5 ((J value at a crack extension of 2.5 mm) values from the J-R curves from EPRI, 1986a at 550 ° F are roughly half of the J2.5 values 75 ° F. The J-R tests used to define the lower bound fracture toughness curves in NUREG/CR-6428, Rev. 1, were all conducted at reactor temperatures and higher, so the lower bound curves should be conservative for LWR reactor temperatures.

9 Figure 3-2 Modified J-R Curves for Type 304 SAW weld metal from EPRI NP -4768

Figure 3-3 Modified J-R curves for Type 316 SAW Weld Metal from EPRI NP-4768

10 3.2.4 Fracture Toughness of SMAW and SAW (Flux) Welds 3.2.4.1 Unaged SMAW and SAW Welds

NUREG/CR-6428, Rev. 1 provides lower bound fracture toughness curves for unaged ASSW.

For unaged SAW and SAW welds, the lower bound curve is represented by:

(/2 ) = 1380.45 Eqn. 3-1

Where is in mm.

Equation 3-1 results in a J2.5 value of 208 kJ/m.

Figure 3-4 shows comparison of the J1 values from several different sources.The J 1 value is the J value at a crack extension of 1 mm which is equal to the C value in the general J -R curve power law equation =. The lower bound curve defined in Equation 3-1 is conservative compared to this data, for which the lowest J 1 is about 150 kJ/m2.

J1, Unaged SMAW/SAW Welds at Operating Temperature 600

500

400

300

200

100

0

Figure 3-4 J value at a crack extension of 1mm for unaged SMAW and SAW Welds, from several different sources

Key to Figure 3-4 Number Material Process Reference

[1] 308L SMAW Hojo et. al., 2014

[2] 316L SMAW Hojo et. al., 2014

[3] 308 SMAW Hong et. al., 2018

[4-1] 304 SMAW EPRI, 1986b

[4-2] 304 SAW EPRI, 1986b

[4-3] 316 SAW EPRI, 1986b

[5] 308 SMAW Mills, 1987

[6] 19-9 SMAW Hale et. al., 1990

11 Key to Figure 3-4 Number Material Process Reference

[7 -1] 308, 4% ferrite SMAW NUREG/CR-6628

[7 -2] 308, 8% ferrite SMAW NUREG/CR-6628

[7 -3] 308, 12% ferrite SMAW NUREG/CR-6628

[8] 308 SAW Mills, 1988

[9] CF8 SMAW Gudas et. al., 1981

[10] 316 MMA2 Picker et. al., 1983

[11] 316 MMA2 Garwood, 1984

[12] 304 SMAW Vassilaros et. al., 1985

[13] 308L SMAW Gavenda et. al., 1986

Orange = International data

Blue = Domestic (US) data

Orange shaded = international, not considered in NUREG/CR-6428, Rev. 1 lower bound

3.2.4.2 Aged SMAW and SAW Welds

Figure 3-5 shows the SMAW weld toughness data evaluated in NUREG/CR-6428, Rev. 1. The data covers test temperatures ranging from 288-427 °C, aging temperatures from 300-400 °C, and aging times from 10,000 - 60,000 hours0 days 0 hours 0 weeks 0 months , along with unaged material. The lower bound curve is compared to twenty-two J -R curves, five of which are for unaged material. Four different alloys are represented (308, 308L, 19. 9L, and 316L). The data in Figure 3-5 does not include data for SAW welds. NUREG/CR-6428, Rev. 1 states that the statistical differences in SAW and SMA weld fracture toughness J-R curves have also been evaluated [in NRC, 1995]

and results indicate no difference between SAW and SMA welds J-R curves. NUREG/CR-6428, Rev. 1 also shows a figure with unaged SAW weld J-R curves showing these curves are bounded by the lower bound unaged curve fo SAW /SMAW welds, provided in Eqn. 3-1 above.

2 MMA = Manual Metal Arc welding, equivalent to SMAW.

12

Figure 3-5 Fracture toughness J-R curves for SMAW weld metal evaluated in NUREG/CR-6428, Rev. 1

For thermally aged SAW and SMAW welds, NUREG/CR-6428, Rev. 1, Section 3.3 recommends the following L-B equation for J:

(/2 ) = 117 0.45 Eqn. 3-2 Where is in ().

NUREG/CR-6428, Rev. 1 does not specifically state a temperature for which the above equation is applicable. The J-R testing was conducted at temperatures ranging from 288-427

°C, with most data being from tests between 288-325 °C. However, the lower bound curves can be considered applicable for service temperatures up to the maximum LWR operating temperature for piping of 343 ° C, because the difference in the J -R curve between 325 °C and 343 °C should be insignificant.,

Equation 3-2 results in a J2.5 of 177 kJ/m2.

Considering the data in Figure 3-5, the average decrease in fracture toughness due to thermal aging, based on the difference in the C value 3 between unaged materials and thermally aged materials (or J1) is 21%. For thermally aged materials with a P value greater than or equal to 4, the average decrease is 29%. There is considerable variation in the decrease in C, ranging from 5% to 52%, with t hree materials actually experiencing slight increases in the C value with aging.

3 The C values for the J-R curves represented in Figure 3-5 are contained in Table B.1 of NUREG/CR-6428, Rev. 1.

13 3.2.4.3 Other Fracture Toughness Data for Aged SMAW and SAW Welds

Chen and coworkers performed fracture toughness testing on archival Type 308L weld material from a BWR plant (Grand Gulf, Unit 2) (Chen et. al., 2019b). The welds were made with SMAW joining Type 304L plate material. Fracture toughness tests were performed both on unaged material and material aged for 10,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C. The fracture toughness testing was conducted in high-purity water at a nominal temperature of 315 ° C. The following power law equation was fit to the resulting J -R curve for the aged weld metal:

= 160 0.47 Eqn. 3-3 Where is in.

This equation results in a higher J -R curve than the lower bound equation from NUREG/CR -

6428, Rev. 1 in Eqn. 3 -2, see Figure 3-6. The curve is consistent with the data used in NUREG/CR-6428, Rev. 1 to develop the lower bound curve. As such, the data from Chen et.

al., 2019b serves to confirm the lower bound curve for SMAW recommended in NUREG/CR-6428, Rev. 1. It is also noted that the fracture toughness data used to develop the lower bound curves in NUREG/CR-6428, Rev. 1, was from tests in air. Hong et. al., 2018 reported J -R curves for a Type 308 SMAW aged at 400 ° C for 5000 and 10,000 hours0 days 0 hours 0 weeks 0 months . These curves are bounded by the NUREG/CR-6428, Rev. 1 lower bound curve for SMAW.

400

350

300

250

200 Chen et. al., 2019b

150 NUREG/CR-6428 LB 100 SMAW

50

0 0.00 1.00 2.00 3.00 4.00 5.00 6.00 a, mm

Figure 3-6 Comparison of J-R curves for archival Type 308L SMAW weld metal to NUREG/CR-6428, Rev. 1 lower bound curve for SAW and SMAW weld metal

3.2.4.4 Summary Discussion - SMAW and SAW Welds

A lower bound fracture toughness curve for aged SMAW and SAW welds is provided in NUREG/CR-6428, Rev. 1. The lower bound curve is based on data for SMAW welds only, with applicability to SAW welds established by a statistical study. For SMAW welds, testing of a thermally aged plant archive material Type 308L weld resulted in a toughness curve that was above the proposed lower bound in NUREG/CR-6428, Rev. 1 (Chen et. al., 2019b).

14 3.2.5 Fracture Toughness of GTAW (non-flux) Welds 3.2.5.1 Unaged GTAW Welds

Per NUREG/CR-6428, Rev. 1, unaged GTAW welds can be represented by:

(/2) = 3300.45 Eqn. 3-4

Where is in ().

Figure 3-7 compares the J1 values for various unaged GTAW weld materials. The blue bars in the figure are for domestic (US) data while the orange bars in the figure represent international (non-US) data, with shaded bars indicating data not considered in developing the lower bound curve for unaged GTAW welds in NUREG/CR-6428, Rev. 1. Figure 3-7 shows that there is a lot of variability in the J1 values, but there is not a tendency for either domestic or international data to have higher or lower J1 values, and that the data not reflected in the NUREG/CR-6428, Rev.

1 lower bound curve is bounded by the lower bound curve. The lower bound curve defined by Equation 3-4 is conservative compared to the data shown in the figure, with the exception of the data from Lucas et. al., 2011. NUREG/CR-6428, Rev. 1 did not consider the Lucas et. al., 2011 data in the determination of the l ower bound J-R curves for GTAW, because the authors considered this data to be an outlier.

J1, Unaged G TA W welds at operating temperature

1200

1000

800

600

400

200

0

[1] [2] [3] [4-1] [4-2] [5] [6-1] [6-2] [7] [8] [9-1] [9-2]

Axis Title

Figure 3 J value at a crack extension of 1 mm for unaged GTAW welds, from several different sources

Key to Figure 3-7 Number Material Reference

[1] 316L Hojo, 2014

[2] 316L Hong et. al., 2018

[3] 304 EPRI, 1986b 316L High Lucas et. al., 2011

[4 -1] Ferrite

15 Key to Figure 3-7 Number Material Reference 316L Low Lucas et. al., 2011

[4 -2] Ferrite

[5] 308 Mills, 1987

[6 -1] 16-8-2 Mills, 1988

[6 -2] 16-8-2 Mills, 1988

[7] 316L Faure, 1989

[8] 308L Nakagaki et. al., 1986

[9 -1] 308L Wilkowski et. al., 1987

[9 -2] 308L Wilkowski et. al., 1987

Orange = International data J1, kJ/m2

Blue = Domestic (US) data

Orange shaded = international, not considered in NUREG/CR-6428, Rev. 1 lower bound

3.2.5.2 Aged GTAW Welds

Figure 3-8 shows the GTAW weld toughness data evaluated in NUREG/CR-6428, Rev. 1. The data is from tests conducted at temperatures of 325 ° C or 427 ° C, and the aging temperatures ranged from 300-400 °C and aging times from 10,000-60,000 hours0 days 0 hours 0 weeks 0 months , with some unaged data included as well. Most of the data is for Type 316L weld filler metal, with a few data points for Type 308 material.The lower bound is based on nine J -R curve s, with two of the curves for unaged material and the remaining severn curves for aged material.

Figure 3-8 Fracture toughness J-R curves for austenitic stainless steel GTAW welds from NUREG/CR-6428, Rev. 1 16

For thermally aged GTAW welds, NUREG/CR-6428 Rev. 1, Section 3.3 recommends the following lower bound (L -B) equation for J:

(/2 ) = 270 0.45 Eqn. 3-5 Where is in ().

NUREG/CR-6428, Rev. 1, does not specifically state a temperature for which the above equation is applicable. Since most of the data was from tests conducted at temperatures between 288 ° C and 325 °C. The staff considers the lower bound curves applicable for service temperatures up to the maximum LWR operating temperature for piping of 343 °C, because the difference in the J-R curve s between 325 °C and 343 °C should be insignificant.

Equation 3-5 results in a J2.5 value of 408 kJ/m2 which exceeds the AMP XI.M12 screening criterion for CASS of a J 2.5 value of 255 kJ/m2.

Considering the data in Figure 3-8, and also including the data from I-NERI, 2017, the average decrease in the fracture toughness due to thermal aging, based on the difference in the C value 4 (or J1) between unaged materials and thermally aged materials is 30%. There is considerable variation in the percent decrease in C, ranging from 13% to 56%.

3.2.5.3 Other Fracture Toughness Data for Aged GTAW Welds

In I-NERI, 2017, the changes in mechanical properties and microstructures in ASSWs were investigated after thermally aging at 360 ° C or 400 ° C up to 10,000 hours0 days 0 hours 0 weeks 0 months . The effects of long-term aging on various mechanical properties were measured using tensile and J-R fracture toughness tests. Testing was conducted on three types of weld metal, Type 308L, Type 316L, and Type 347, with all welds being made using the GTAW process. Fracture toughness test results were only reported for material aged at 400 ° C. F or Type 316L weld metal, a significant reduction in the fracture toughness was observed after aging at 400 ° C for 10,000 hours0 days 0 hours 0 weeks 0 months (Figure 3-9 ). By contrast, Type 347 weld filler metal showed little or no reduction of fracture toughness.

Type 347 weld filler is not typically used for primary piping welds in US plants. It is noted that the I-NERI data for Type 316L weld filler shows substantially lower toughness values at 320 ° C to the NUREG/CR-6428, Rev. 1 lower bound equation, with a J2.5 of around 200 kJ/m2 for the I-NERI data, compared to a lower bound J 2.5 value of 408 kJ/m2 from NUREG/CR-6428, Rev. 1.

4 The C values for the J-R curves represented in Figure 3-5 are contained in Table B.1 of NUREG/CR-6428, Rev. 1.

17 a)

b)

Figure 3-9 J-R Curves for as-welded and aged Type 316L weld filler metal. a) Tested at room temperature; b) Tested at 320 °C

18 Hong and coworkers (Hong et. al., 2018) reported on fracture toughness testing on Type 316L and Type 347 GTAW welds. All the welds were aged for 10,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C. The J-R curves for the Type 316L material were not bounded by the NUREG/CR-6428, Rev. 1 lower bound J-R curve for GTAW welds, which was noted by the authors. By contrast, the Type 347 GTAW weld lost relatively little toughness due to aging. The tests of the Type 316L and Type 347 welds reported in Hong et. al., 2018, appear to be for the same materials reported on in I-NERI, 2017, based on identical chemical composition, and the J -R curves appear to be identical.

One factor that could help to understand the low toughness of these welds would be the oxygen content. It has been recognized in the ASME Code that there is a difference in the lower bound toughness for GTAW (no flux) and SMAW/SAW (fluxed welds). GTAW is done with inert shielding gas. If the gas shielding is not effective, the oxygen content may not be as low in all cases, which could explain the difference in toughness for the 316L and 347. However, oxygen is not routinely reported and was not reported in this case.

Koyama et. al., 1999 provides a representative J-R curve for CF-8M with ferrite greater than 20% as well each of the welding processes typically used for Class 1 piping, including TIG (GTAW), SMAW and SAW welds aged 20,000 to 40,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C. These J-R curves are used as the basis for proposed Z-factors for Japanese weld materials developed using similar methodology to the Z-factors in Section XI, Appendix C. The weld material is not identified, but is assumed to be equivalent to Type 316 since the base pipe material joined is equivalent to Type CF8M.The calculated Z-factors for TIG and SMAW are the same (assumed to be good for CF-8M with < 20% ferrite also) while those for SAW and CF-8M with > 20% ferrite are grouped together, but at higher Z-factors (lower toughness) than that for TIG/SMAW. This is different from the Z-factors in ASME appendix C where the Z-factor for the non-flux welds is assumed to be 1 (equivalent to wrought base metal) and the Z-factor for SMAW/SAW (flux) welds are the same.

Hojo, 2014 provides J -R curves for Type 316L GTAW welds with 8% ferrite from the Japanese PWR owners database, with aging temperatures ranging from 300 ° C to 450 ° C, and aging times ranging from 3000 hours0.0347 days 0.833 hours 0.00496 weeks 0.00114 months to 60,000 hours0 days 0 hours 0 weeks 0 months . The materials were all tested at 325 °C. A total of 22 J-R curves were generated for aged materials, with 6 curves with 10,000 hours0 days 0 hours 0 weeks 0 months aging or greater and aging temperatures of 400 ° C or greater. These curves, plus one curve aged at 350

°C for 60,000 hours0 days 0 hours 0 weeks 0 months , all have aging parameter P greater than 4.

Figure 3-10 shows the J-R curve s for non-flux welds from Koyama et. al., 1999 and Hojo, 2014,

the J-R curve from Hong et. al., 2018 for a Type 316L GTAW weld, and the NUREG/CR-6428, Rev. 1 lower bound curve for GTAW welds.The J -R curve shown from Hojo, 2014, was for material aged for 60,000 hours0 days 0 hours 0 weeks 0 months at 400 ° C, representing an aging parameter P of 4.8.

The lowest J-R curve from Hojo, 2014 compares well with the NUREG/CR-6428, Rev. 1, lower bound curve, while the Koyama et. al., 1999 curve is bounded by the NUREG/CR-6428, Rev. 1 curve.

Considering the additional J-R curves for GTAW welds that were not used as a basis for the NUREG/CR-6428, Rev. 1 lower bound J-R curve for GTAW, curves from two out of three sources are either bounded or very close to the NUREG/CR-6428, Rev. 1 lower bound curve.

The data from Hojo, 2014, includes seven J-R curves with aging parameter P greater than 4.

19 Figure 3-10 J-R curves for GTAW Welds from Koyama et. al., 1999 Hojo, 2014 and Hong et. al., 2018, compared with the NUREG/CR-6428, Rev. 1 lower bound for GTAW welds

3.2.5.4 Summary Discussion - GTAW

In NUREG/CR-6428, Rev. 1, a lower bound fracture toughness curve for GTAW welds is proposed based on seven aged and two unaged J-R curves, which is a relatively small number of data. Further, seven of these J-R curves are from one source (MHI). For GTAW, some recent testing has shown J-R toughness curves for Type 316L that are below the lower bound curve proposed in NUREG/CR-6428, Rev. 1 (I-NERI, 2017, Hong et. al., 2018). These references (I-NERI, 2017 and Hong et. al., 2018) do not provide an explanation of why the J -R curve for the Type 316L weld metal falls below the lower bound curve in NUREG/CR-6428, Rev.

1. Since the testing reported in I-NERI, 2017 and Hong et. al., 2018, appear to represent the same test and heat of material, this represents one data point. The material had a low unaged J-R curve, with a J 1.5 value of around 450 kJ/m2 at 320 ° C, compared to a J1.5 of 650 kJ/m2 for the unaged 316L material at 325 ° C in Figure 3-9. Based on the low unaged toughness of the material, as well as the low aged toughness compared to the six other J-R curves for aged 316L material in Figure 3-6, this data appears to be an outlier. Additionally, fracture toughness testing documented in Lucas, 2011 showed even lower J -R curves for both unaged and aged Type 316L weld metal.

Fracture toughness data for aged GTAW welds from Hojo, 2014, is slightly lower than the NUREG/CR-6428, Rev. 1 lower bound curve (Figure 3-10). This data has P of 4.8 which represents SLR conditions. Based on this data, it is suggested to reduce the lower bound GTAW curve C parameter by 10%. This would result in the following recommended equation for the lower bound curve:

(/2 ) = 243 0.45 Eqn. 3-1

20 Considering that the lower bound curve in NUREG/CR-6428, Rev. 1 for GTAW is based on relatively few J-R curves, and does not bound some other J-R data as described above, it would be beneficial to perform more J-R tests on GTAW welds.

3.2.6 Effects of the Environment on Fracture Toughness of ASSW Per NUREG/CR-6428, Rev. 1, fracture toughness in water environments of ASSW may be lower than fracture toughness in air, but there is insufficient data to accurately establish the effects of environment on fracture toughness.NUREG/CR-6428, Rev. 1 cites work by Lucas (Lucas, 2011, and Lucas et. al., 2011), which showed that J-R toughness can be up to 40%

lower when tested in a reactor coolant (BWR) environment (300 ppb oxygen). However, the Type 316L material tested had anomalously low as -welded toughness. In addition, the reduction in toughness in the environment was higher for 10% ferrite material than for 14%

ferrite material. A BWR environment with 300 ppb O2 is probably not representative of the reactor coolant system environment of many US plants since most BWRs operate with hydrogen water chemistry with noble metal addition or online noble metal addition, which results in a much lower oxygen environment, and PWRs operate with very low oxygen, reducing conditions. Therefore, the Luc as 2011 and Lucas et. al., 2011 results may not be representative of the environmental conditions in most US LWRs.

Other J-R testing in a reactor coolant environment cited in NUREG/CR-6428, Rev. 1, and detailed in NRC, 2010 was of neutron-irradiated, wrought Type 304L stainless steel. Other work cited in NUREG/CR-6428, Rev. 1 by Nakajima, et. al., 1986, was also on sensitized, wrought Type 304 stainless steel in a BWR normal water chemistry environment. This testing may not be representative of the potential environmental effects on the toughness of ASSW since it was conducted on WASS, which does not have a duplex microstructure, and was on irradiated materials in one case.

Chen et. al., 2019b documents fracture toughness testing of aged Type 308L material in a high-purity water environment at 315 ° C. The test conditions contained < 10 ppb dissolved oxygen, a 4% hydrogen/nitrogen cover gas, circulation of the coolant at a rate of 20-30 ml /minute, and conductivity < 0.07 S/cm2 throughout the test. As detailed in Section 3.2.4.3, the resulting J-R curve for material aged for 10,000 hours0 days 0 hours 0 weeks 0 months at 400 °C is bounded by the NUREG/CR-6428, Rev. 1, lower bound curve for SAW and SMAW welds.

In summary, there is not enough data on fracture toughness of aged ASSW in reactor coolant environments representative of most US LWRs to make any conclusions regarding whether there is an environmental effect on fracture toughness. One test of SMAW weld material in a high-purity water environment (similar to a PWR environment) has been conducted, which was bounded by the NUREG/CR-6428, Rev. 1 lower bound curve for SAW /SMAW welds.

Additional fracture toughness testing of ASSW in representative reactor coolant environments,

which would include either PWR or BWR HWC environments, would be helpful in establishing whether there is a significant environmental effect on fracture toughness of thermally aged ASSW.

3.2.7 Discussion 3.2.8 Effect of Ferrite Content There is some evidence that thermally aged fracture toughness could be correlated with ferrite content, as is done in the fracture toughness model for thermally aged CASS used as the basis for the screening criteria for loss of fracture toughness used by the NRC in its aging management program for CASS. Examination of the data on ASSW welds in Table B.1 of NUREG/CR-6428, Rev. 1 reveals that the data spans welds with a ferrite content of 4.0% -

21 19.0%, with most data between 5 and 14%. The data covers various types of weld metal, primarily 308, 308L, 316, and 316L. Weld processes represented include SAW, SMAW, and GTAW. There is not enough data with both ferrite content, J -R curve data, and thermal aging, for any given weld filler/process combination to develop a meaningful correlation of fracture toughness to ferrite content and thermal aging time and temperature. One of the few studies tabulated in Table B.1 of NUREG/CR-6428, Rev. 1 that has J-R data and a range of ferrite contents is from NRC, 2000b, for 308 SMAW welds. The material was aged for 50,000 hours0 days 0 hours 0 weeks 0 months at 343 °C. Figure 3-11 shows the J-R curves for the aged welds with the different ferrite contents.

The material toughness actually increased with aging for the 4% ferrite material, and decreased with aging for the 8% and 12% ferrite materials. These results show that loss of toughness increases with increasing ferrite content except at a ferrite content of 4%, at least for this one type of material and process.

Figure 3-11 Effects of aging at 343 °C on the J-Integral fracture toughness of type 308 stainless steel weld metal. All tests conducted with precracked CVN specimens at 290 °C.

(From NRC, 2000b)

3.2.9 Effect of Aging Temperature

Figure 3-12 shows the effect of different aging temperatures on J at a crack extension of 6 mm (J6mm) for two welds with the same P value, based on test results described in Hojo, 2014.

Toughness for material aged at 400° C tends to be roughly the same as results from aging at lower temperatures. Therefore, aging at 400° C would be representative of the behavior after

22 aging at lower temperatures. Overall, there is insufficient evidence that aging at different temperatures results in different amounts of thermal aging given the same P value.

a)

b)

Figure 3-12 J6mm for Type 316L materials with different aging temperatures and the same aging parameter. a) GTAW weld; b) SMAW weld

23 3.2.10 Summary Discussion - Fracture Toughness Section 3 of this report summarizes the existing data on fracture toughness of aged and unaged ASSW. Most of the existing data is captured in NUREG/CR-6428, Rev. 1, which includes lower bound fracture toughness curves for aged and unaged ASSW as a function of weld process, with separate lower bound curves provided for SAW /SMAW and GTAW. Both SAW /SMAW welds and GTAW experience a reduction in fracture toughness due to thermal aging. The aged fracture toughness of SAW /SMAW welds is lower than that of GTAW. However, welds made using the SAW and SMAW process also tend to have a lower initial toughness than GTAW welds. GTAW welds tend to experience a greater percent reduction of the initial weld toughness due to aging, but still retain significantly higher aged toughness than SMAW and SAW welds.

Based on limited data, given the same alloy, higher ferrite content seems to lead to lower fracture toughness, for the same aging time and temperature. The temperature of accelerated aging does not seem to cause a discernable difference in loss of toughness, given the same P value.

Although considerable data on both unaged and aged fracture toughness of ASSW exists, there is insufficient data that has documented chemical composition and ferrite content to enable developing a model correlating chemical composition and/or ferrite content, weld process and aging time to fracture toughness. Additional testing of materials with a range of chemical composition, ferrite content, and weld process could facilitate development of a more accurate fracture toughness model with less uncertainty.

There is also insufficient fracture toughness data on ASSW tested in representative LWR reactor coolant environments to conclusively determine if there is a detrimental effect of the reactor coolant environment on fracture toughness. Additional fracture toughness testing of ASSW in environments representative of those prevalent in the US LWR fleet would be beneficial to help define the environmental effect.

24 4 CURRENT AGING MANAGEMENT GUIDANCE FOR REDUCTION IN FRACTURE TOUGHNESS OF SS

4.1 Aging Management Guidance for ASSW

ASSW in Class 1 piping most likely have a high enough service temperature to make them susceptible to TE. Class 2 systems connected to the reactor coolant system ( RCS) may have some welds that are potentially susceptible to TE due to their operating temperature. The aging effect of loss of fracture toughness due to TE is currently not addressed in the GALL-SLR for ASSW, although the welds are subject to Section XI examinations for other aging mechanisms..

The aging effects requiring mana gement (AERMs) for ASSW in the GALL-SLR are not distinguished from the WASS in the system.

Aging management of Class 1 piping in PWRs for several AERMs is addressed by several line items in GALL-SLR Table IV.C2, Reactor Vessel, Internals, and Reactor Coolant System, Reactor Coolant System and Connected Lines. These line items cover WASS (either clad or unclad), and CASS piping. The aging effects include cracking due to SCC and cumulative fatigue damage. Cracking of pipi ng due to SCC is addressed by one or more of the following AMPs: XI.M1, ASME Section XI Inservice Inspection, Subsections IWB, IWC, and IWD, AMP XI.M2, Water Chemistry, XI.M25, BWR Reactor Water Cleanup System, and AMP XI.M35, ASME Code Class 1 Small -bore Piping, and XI.M7, BWR Stress Corrosion Cracking. AMP XI.M1 generally requires volumetric examinations of larger bore Class 1 and 2 piping welds and immediately adjacent material, and AMP XI.M35 is credited with finding cracking using volumetric or destructive examinations in ASME Code Class 1 small -bore piping not subject to Section XI inservice inspection ( ISI) requirements. AMP XI.M2 is a preventive program for SCC cracking.

ASSW are also present in non-piping components such as reactor vessel internals (RVI).

Several AMPs in the GALL-SLR contain guidance for nondestructive examination of components that include ASSW. These AMPs include XI.M4, BWR Vessel ID Attachment Welds, XI.M8, BWR Penetrations, XI.M9, BWR Vessel Internals, and XI.M16A, PWR Vessel Internals.Some of these AMPs explicitly address loss of fracture toughness due to neutron embrittlement and/or TE.

Cumulative fatigue damage and cracking due to cyclic loading are managed by a time limited aging analysis (TLAA), and in some cases the Fatigue Monitoring AMP (X.M1) may be credited for resolving the TLAA.

CASS piping has the additional AERM of TE, which is managed by AMP XI.M12.

4.2 Aging Management Guidance for Reduction of Fracture Toughness of WASS

The aging effect of loss of fracture toughness due to TE is currently not addressed in the GALL-SLR for WASS. The management of other AERMs for WASS piping is the same as stated in Section 4.1, since wrought piping materials and weld fillers are not called out as separate line items in the GALL-SLR.

4.3 Operating Experience

Failures due to fracture of ASSW in LWR primary systems have been rare or nonexistent.

Similarly, flaws due to in-service degradation have rarely been detected in ASSW, if ever. In BWRs in the 1980s, there was extensive experience with IGSCC in the heat affected zone (HAZ) of welded austenitic stainless steel pipe (EPRI, 2019). IGSCC in the weld metal itself, is not mentioned in EPRI, 2019. The cracking therefore occurred in the vicinity of ASSW, but probably not in the weld metal itself. In PWR primary coolant piping, SCC of stainless steel piping, including the weld has very rarely occurred, and the NRC staff is not aware of any documented incidents of SCC in ASSW. NRC Information Notice 97-19 (NRC, 1997a) documents cracking due to IGSCC in the base metal of a safety injection line just upstream of the check valve adjacent to the unisolable section of the line; however there is no mention of cracking penetrating or originating in the weld metal. NRC Information Notice 97-46 (NRC, 1997b) documents cracking in high-pressure safety injection piping resulting from thermal cycling and flow-induced vibration.Fatigue of reactor coolant system branch connection piping due to thermal cycling has been an ongoing issue in PWRs, and has been addressed by topical report Materials Reliability Program: Management of Thermal Fatigue in Normally Stagnant Non-Isolable Reactor Coolant System Branch Lines, MRP-146 (EPRI, 2016).

Austenitic stainless steel piping in BWRs has a history of cracking due to intergranular stress corrosion cracking (IGSCC)stress corrosion cracking. IGSCC has been mitigated in many BWRs by replacement of piping with piping materials less susceptible to IGSCC, such as low -

carbon grades, or by water chemistry changes. Most commonly this cracking occurs in the HAZ. Therefore, flaw evaluations may need to consider weld fracture toughness properties.

Primary water stress corrosion cracking has occurred extensively in PWRs in nickel -based piping components, particularly in Alloy 600 base material and Alloy 82 and 182 weld material.

However, nickel-based alloy materials are normally welded with nickel -based weld materials such as Alloy 82 and Alloy 182, not ASSW, so this operating experience (OE) is not relevant to ASSW.

4.4 Potential Changes to Aging Management Guidance for ASSW

For TE of CASS, the GALL-SLR, Volume 2 applies a threshold temperature of 250 °C, below which reduction in fracture toughness is not considered significant.This thres hold is also appropriate for ASSW since, for example, based on typical activation energies, the equivalent aging time at 250 °C to service at 290 °C for 72 EFPY would be 456 EFPY.Therefore, for ASSW, aging at 250 °C occurs about six times slower than at typical LWR reactor coolant system temperatures, making this a conservative threshold. Systems that contain stainless steel piping and thus, ASSW with operating temperatures greater than or equal to 250 ° C generally include the reactor coolant system, emergency core cooling systems (ECCS), and auxiliary systems. However, only the reactor coolant system (reactor coolant pressure boundary) consisting of the RCS and connected system piping within the first containment isolation valve, are likely to have temperatures greater than 250 ° C.

Since TE is a potentially significant aging mechanism for ASSW, consideration should be given to adding new line item(s) to the GALL-SLR for austenitic stainless steel welds in Class 1 piping and piping components, for the mechanism should be added as a new component, with the AERM of loss of fracture toughness, if TE is determined to be significant.

For CASS piping components in AMP XI.M12, screening was used to determine which piping components were recommended for augmented nondestructive examination ( NDE) or component-specific flaw tolerance evalua tion. Augmented NDE was recommended because ASME Section XI volumetric examinations only cover welds and the immediately adjacent material, and because historically UT examinations were not effective for CASS due to its grain structure.

For ASSW, most welds are covered by ASME Section XI-required volumetric examinations (typically ultrasonic examination).Section XI UT examinations have better detectability for flaws in ASSW provided Section XI, Appendix VIII-qualified UT methods are applied. The ASME 26 Section XI examinations are performed on a sampling basis (25% of welds). However, many plants have implemented risk-informed ISI programs which allows examination of a lower percentage of welds. Therefore, an assessment of the need for augmented NDE should consider whether a plant that implements a traditional or risk-informed ASME Section XI ISI programs.

With respect to an AMP for loss of fracture toughness for welds, a susceptibility screening approach similar to that for CASS is not possible at present due to the lack of a model for toughness of ASSW as a function of chemical composition and ferrite content. AMP XI.M12 for CASS contains screening criteria in terms of ferrite content and chemical composition that are linked to a fracture toughness J 2.5 value of 255 kJ/m2. The lower bound fracture toughness curve for GTAW welds in NUREG/CR-6428, Rev. 1 has a J 2.5 value that exceeds 255 kJ/m2.

Therefor, pending confirmation of the NUREG/CR -6428, Rev. 1 curve, i t should be possible to screen out welds made with GTAW process with respect to the need for additional aging management activities. However, GTAW welds still experience reduction in fracture toughness due to TE compared to the as -welded state, so the effects of TE should be taken into account if a flaw evaluation is necessary.

If welds are made using more than one weld process, welds should be considered susceptible to TE if part of the weldment is made using either SAW or SMAW. For example, many piping welds use the GTAW process for the weld root, while the bulk of the weld is made using either SMAW or SAW.

4.4.1 Detection of Aging Effects ASSW in pipes greater than or equal to 4 inches in diameter in reactor coolant systems are generally subject to volumetric examination on a sampling basis through GALL-SLR AMP XI.M1, others would be examined through AMP XI.M7, XI.M25, or XI.M35, and AMP XI.M2 will provide preventive measures for SCC cracking and fatigue cracking. Many licensees have implemented risk-informed inservice inspection programs which reduce the percentage of welds examined, or may result in welds with low likelihood of inservice degradation not being inspected. Provided that some percentage of ASSW welds in a system are subject to NDE under AMP XI.M1, then these AMPs should be adequate with respect to detection of aging effects (e.g. cracking). These programs would not detect loss of fracture toughness.

Inspections for cracking should prioritize ASSW that have higher operating temperatures, such as welds in the pressurizer surge line or other lines connected to the pressurizer.For ASSW in RVI, AMPs XI.M4, XI.M8, XI.M9, and XI.M16A call for nondestructive examinations of certain welds, which would detect cracking.

4.4.2 Flaw Evaluation Section 5.1.1 discusses the current ASME Code,Section XI flaw evaluation procedures for ASSW, and concludes that these procedures do not account for TE. 5.1.1 Therefore, the current ASME Code,Section XI flaw evaluation procedures would not be applicable to thermally aged ASSW without modification. Ideally, an AMP for ASSW would reference ASME Code,Section XI procedures for flaw evaluation, if these procedures are updated to account for TE of ASSW. Section 5.1 discusses a possible method for updating the flaw evaluation procedures for ASSW in the ASME Code,Section XI.

4.5 Summary

ASSW in primary piping are covered under the same line items as wrought and CASS base materials in piping in the GALL-SLR.

27

If loss of fracture toughness of ASSW were determined to warrant aging management, new AERM(s) for loss of fracture toughness would need to be added for ASSW welds in the RCS and RCPB.

Appropriate AMPs that could be assigned to manage the aging effect of cracking of ASSW such as AMP XI.M1, ASME Section XI Inservice Inspection, Subsections IWB, IWC, and IWD, AMP XI.M2, Water Chemistry, and AMP XI.M35, ASME Code Class 1 Small-bore Piping.

Pending confirmation of the NUREG/CR-6428, Rev. 1 lower bound fracture toughness curve for GTAW, it may be possible to screen out welds made entirely with the GTAW process from requiring any additional aging management activities..

Plants that are implementing a RI-ISI program may be inspecting a smaller number of ASSW in primary piping than plants that do not have a RI-ISI program.

Current flaw evaluation procedures in the ASME Code,Section XI referenced in the GALL-SLR do not currently take into account TE effects on toughness of ASSW.

28 5 EVALUATION METHODOLOGIES

5.1 Handling of ASSW Fracture Toughness and Flaw Evaluation in the ASME Code

5.1.1 ASME Code,Section XI, Nonmandatory Appendix C, Analytical Evaluation of Flaws in Piping The ASME Code,Section XI, Rules for Inservice Inspection of Nuclear Reactor Facility Components, recognizes differences in the fracture toughness of ASSW made with flux (SMAW, SAW) versus welds made without flux (GTAW). In Nonmandatory Appendix C,

Analytical Evaluation of Flaws in Piping, Figure C-4210-1 provides a flowchart for selection of the flaw analysis method for austenitic piping, which requires different analysis methods depending on whether the analysis is for a wrought material, cast material or weld material.

Beginning with the 2019 edition of Section XI, changes were made in the requirements for analytical procedures for CASS materials to require different analytical methods based on the ferrite content of the material. In Figure C -4210-1, weld materials are differentiated with respect to flux or nonflux welds. For flux welds and higher ferrite cast materials, elastic-plastic f racture mechanics (EPFM) is required using the rules of C-6000. For nonflux welds and cast materials with ferrite content less than or equal to 14%, the limit-load rules of C-5000 are used. When EPFM is performed for flux welds or cast materials, Z-factors are given by C-6330. The Z-factor is essentially a multiplier on the load in the stress ratio of Tables C-5310-1 through C-5310-5 and Table C-5410-1 which provide the allowable end-of-evaluation period flaw depth to pipe thickness ratios for various flaw orientations and loading conditions. So, the Z-factor basically is a correction factor applied to the limit-load allowable flaw sizes which makes the allowable flaw sizes smaller to account for unstable crack extension due to ductile tearing prior to reaching limit load. The Z -factor is given by an equation in C-6330 as a function of pipe size. For CASS materials with ferrite content greater than 25%, C -6330 specifies the use of the same Z-factors that are used for Category 2 ferritic materials in Table C -6330-1, which basically includes ferritic flux welds.

Therefore, in Appendix C, austenitic stainless steel flux welds (SMAW and SAW) and CASS with ferrite content > 14% are recognized as having lower fracture toughness than wrought materials. The required Z-factors for ASSW (SAW and SAW ) were not changed in the 2019 edition of Section XI, and are the same Z-factors now required for CASS with ferrite content greater than 14%, except for CF8M material with ferrite content greater than 25%.

The changes for the analysis methods for CASS in Appendix C were made in Code Record 16-2757 and took effect with the 2019 Edition. The technical basis for Record 16-2757 is contained in PVP2017-66100 (ASME, 2017), which references Journal of Pressure Vessel Technology, 1986, and PVP2003-2026 ( ASME, 2003) with respect to the basis of the Z-factors for SAW and SMAW welds in Appendix C.

Figure 5-1 reproduces a figure from ASME, 2017 showing a constant J2.5 level for SAW and SMAW welds of around 350 kJ/m2, independent of delta ferrite content.

29 Figure 5-1 Comparison of J2.5(J0.1) for CASS with other materials in ASME Code Section XI (From ASME, 2017, Fig. 8)

ASME, 2017 references ASME, 1986, and ASME, 2003 as the source for the SAW/SMAW line in Figure 5-1. Further investigation showed this J2.5 value was based on testing documented in EPRI, 1986b. However, the J 2.5 of around 350 kJ/m2 of the SAW/SMAW line in Figure 5-1 should probably be lower, around 250 kJ/m 2, if data from EPRI, 1986b was used. It is also noted that this line is a bit higher than the J-R curve s for 304 and 316 SAW welds from Landes and McCabe (EPRI, 1986b) shown in Figure 3-2 and Figure 3-3, which have J2.5 values around 250-280 kJ/m2. The development of the current Z-factor methodology in Section XI, Appendix C is described in ASME, 1986 and EPRI, 1986a, which reference EPRI, 1986b for the fracture toughness properties of SMAW and SAW welds. NUREG/CR-6428, Rev. 1 provides a comparison of the new lower bound curves to the previous data from EPRI, 1986b, as reproduced in Figure 5-2. The new lower bound curves are ~35% lower than the EPRI, 1986b curve (lower by around 100 kJ/m2 at a crack extension of 2.5 mm. However, the J2.5 value for Category 2 ferritic material, depicted by the lower red line in Figure 5-1 (from ASME, 2017) is approximately 180 kJ/m2, which is approximately equal to the J 2.5 value from the lower bound curve for SAW and SMAW welds from NUREG/CR-6428, Rev. 1, shown in Figure 5-2.

Therefore, a possible approach for modification of the flaw evaluation rules for SAW and SMAW welds would be to require the use of the Z-factors for Category 2 ferritic material for these welds, as is done for CASS materials having greater than 25% ferrite.

30 Figure 5-2 Fracture toughness lower bound J-R curves and the data on Type 304, 316L,and CF-3 welds used to develop the ASME Code IWB-3640 analysis. (from NUREG/CR-6428. Rev. 1)

In the Appendix C, in Figure C-4210-1, nonflux welds are treated the same as wrought material.

GTAW welds are nonflux welds. However, the fracture toughness of thermally aged GTAW welds is lower than that of wrought materials, although higher than that of thermally aged SAW and SMAW welds.In Figure 5-1, the line for SAW and SMAW welds (unaged) is at around 350 kJ/m2, which was considered an appropriate fracture toughness basis for the Z-factors for CASS with ferrite content greater than 14% (CF3 and CF8) or between 14 and 25% (CF8M). This is the basis for the Z-factors in C -6330(a). The J2.5 value of 408 kJ/m 2 from the lower bound curve for aged GTAW welds from NUREG/CR-6428, Rev. 1 is slightly higher than the SAW and SMAW line in Figure 5-1. Therefore, since the GTAW weld lower bound toughness is bounded by the toughness used for the basis of the Z-factors in C-6330(a), use of the same Z-factors would be appropriate for thermally aged GTAW welds, pending confirmation that J2.5 of 350 kJ/m2 is the correct fracture toughness basis for the SAW/SMAW Z -factors in Appendix C.

It is also noted that the unaged lower bound toughness curves from NUREG/CR-6428, Rev. 1, are not significantly greater than the curves for thermally aged material. For example, for SAW and SMAW welds, the J 2.5 value for unaged material from NUREG/CR-6428, Rev. 1 is 208 kJ/m2 compared to the aged J2.5 of 177 kJ/m2, or only 15% lower for the aged material.For GTAW welds, the respective aged and unaged J 2.5 values are 408 kJ/m2 and 498 kJ/m2, respectively, or 20% lower for the aged material. Therefore, use of the Z-factors recommended above for both thermally aged and unaged weld materials is appropriate.

5.1.2 Code Case N-906 Code Case N-906, Flaw Evaluation Procedure for Cast Austenitic Stainless Steel Piping and Adjacent Fittings,Section XI, Division 1, published in 2020 (ASME, 2020), provides an alternative method of predicting the failure mode and determining Z-factors to that provided in Section XI, Appendix C for CASS materials. The code case is only applicable to circumferential surface flaws. The technical basis for the code case uses a mathematical model which

31 correlates toughness with chemical composition developed from fracture toughness tests of 19 heats of aged CASS material (ASME, 2019). The corresponding Z-factors are then determined using equations based on dimensionless -plastic -zone-parameter (DPZP) analysis.

The DPZP analysis is based on circumferential surface-cracked nuc lear pipe test data to determine when limit-load is applicable and also approximates a correction factor that needs to be applied to the limit-lo ad solution (inverse of the Code Z-factor). By contrast, the Z-factors in Appendix C were developed based on EPFM anal ysis of throughwall cracked pipes under bending, so are more conservative. The Code Case N-906 methodology was validated against full-scale pipe tests (ASME, 2019).

Code Case N-906 provides a method of flaw evaluation for CASS piping that is intended to be more accurate while retaining sufficient conservatism, compared to the method in Section XI, Appendix C. It is possible that a similar approach could be developed for ASSW. The details of the correlation of parameters such as chemistry and/or chemical composition would necessarily differ from the correlation employed in Code Case N -906, and the methodology would also need to account for the weld process.

5.1.3 Conclusions - ASSW Fracture Toughness in the ASME Code,Section XI Based on the above discussion, the staff concludes the following with respect to the current ASME Code,Section XI flaw evaluation procedures for ASSW:

These procedures do not account for thermal aging of ASSW, with respect to the underlying basis for the Z-factors.

These procedures do not contain sufficient conservatism to bound the fracture toughness of thermally aged ASSW.

The Z-factors in Section XI, Appendix C, C-6330(a), which are currently specified for SMAW and SAW welds, Type CF3 or CF8 CASS with ferrite greater than 14%, and Type CF8M CASS with ferrite greater than 14% but less than or equal to 25%, may be appropriate for flaw evaluation of aged ASSW welds made using the GTAW process, pending confirmation of the fracture toughness basis for these Z-factors.

The Z-factors in Section XI, Appendix C for Type 2 ferritic material would be appropriate for flaw evaluation of aged ASSW welds made using the SAW or SMAW process, rather than the Z-factors from C -6330(a).

A possible path for resolution of the issues above would be for the NRC to request that the cognizant ASME Code group (Working Group, Pipe Flaw Evaluation (WGPFE),

Section XI) investigate the basis for the Z-factors applicable to ASSW and take action as necessary to update these factors.

The ASME Code WGPFE should provide clarification/better documentation of the toughness basis for the Z-factors of unaged SAW/SMAW welds, as shown by the line in Figure 5-1 (Figure 8 from ASME, 2027), if these Z-factors will be retained in Section XI.

5.2 Probabilistic Methods

The ASME Code uses deterministic methodologies to evaluate structural integrity of components. In recent years, probabilistic fracture mechanics (PFM) methods have been developed to assess the structural integrity of nuclear power plant components. The xLPR (Extremely Low Probability of Rupture) computer code has the capability to perform PFM analyses on primary piping.The code, which was designed, programmed, and tested under a 32 rigorous software quality assurance program, provides regulators, industry, researchers, and the public with new quantitative capabilities to analyze the risks associated with nuclear power plant piping systems subject to active degradation mechanisms. Core capabilities of the code include modeling fatigue, stress-corrosion cracking, inservice inspection, chemical and mechanical mitigation, leak rates, and seismic effects.The code can output the probability of leakage and probability of rupture. A description of the xLPR code and its technical basis, and verification and validation, can be found in NUREG -2247, Extremely Low Probability of Rupture Version 2 Probabilistic Fracture Mechanics Code, August 2021 (NRC, 2021). In a recent study, xLPR was used to demonstrate that PWR piping systems previously approved for leak -before-break (LBB) continue to exhibit an extremely low probability of rupture consistent with the requirements of Title 10 of the Code of Federal Regulations, Part 50, Appendix A, General Design Criterion (GDC) 4, when subject to the effects of primary water stress -corrosion cracking (PWSCC) (NRC, 2022)

The xLPR code could potentially be used to examine the effects of thermal aging on the toughness distribution for ASSW in piping. Fracture toughness is a user input to xLPR.

However, xLPR does not currently incorporate thermal aging models, or have the capability to dynamically vary fracture toughness with time. Therefore, end-of-life fully aged (saturated) properties could be used, or s ensitivity studies comparing the probabilities of leakage or rupture for materials with different amounts of thermal aging could be performed.

33 6 EFFECT OF THERMAL AGING ON OTHER AGING MECHANISMS

Thermally aged ASSW may have different susceptibility to other aging mechanisms such as SCC. This in turn could affect how aging of ASSW is managed for long-term operation.

6.1 Effect Of Thermal Aging On SCC

Several investigations of SCC growth in thermally aged ASSW have been conducted. Chen et.

al., 2019a found no differences in SCC growth rates in a low corrosion potential LWR environment (< 10 ppb dissolved oxygen, 315-318 °C, conductivity ~ 0.07 S/cm) for aged (400

°C for 10,000 hours0 days 0 hours 0 weeks 0 months ) and unaged Type 308L weld material with 7% ferrite.

Yamada et. al., 2009, performed SCC tests on Type 308L and Type 316L weld metals in both high-potential (8-ppm O2) and low potential (2.7 ppm H2). Both test environments contained boric acid (H2BO3) resulting in 500 ppm boron, and lithium hydroxide resulting in 2 ppm lithium.

The low potential conditions essentially simulate a PWR environment, while the high-potential conditions may have been intended to represent a startup condition in a PWR. The materials were aged at 400 ° C for 10,000 hours0 days 0 hours 0 weeks 0 months . Specimens of the same weld materials with 20% cold work were also tested under the same conditions. Although SCC crack growth rates ( CGRs) were higher in the oxygenated environment compared to the low potential environment for all materials, there was no significant impact on SCC growth rates for aged versus unaged material. Yamada et. a l., 2010 reported similar results for Type 316L weld metal.

Hixon et. al., 2007 found that the SCC CGRs in simulated BWR conditions of Type 316L weld material with a ferrite umber of 13 aged at 400 °C for 5000 hours0.0579 days 1.389 hours 0.00827 weeks 0.0019 months were within the scatter band for wrought material. The aged material CGR was about seven times that of the unaged material.

Lucas et. a l., 2011 conducted SCC tests of Type 316L material under BWR conditions with 300 ppb dissolved oxygen. These test conditions are similar to BWR normal water chemistry (NWC), in which dissolved oxygen content (DO) may typically be around 200 ppb, and other oxidants may also be present (EPRI, 2011). This study found that SCC CGR in BWR conditions for aged Type 316L weld material increased by a factor of four over unaged material for 14%

ferrite material, and a factor of 20 for 10% ferrite material.

Andresen, 2015 reported on SCC CGR testing of Type 308L weld metal with 14.5% ferrite aged at 400 ° C for 10,000 hours0 days 0 hours 0 weeks 0 months . The CGR testing was focused on BWR water, represented by 2000 ppb dissolved oxygen (considerably higher than the DO content under typical BWR NWC conditions), and 63 ppb hydrogen to provide a comparison to lower potential conditions 5. A maximum CGR of 2.4x10 -7 mm/s was reported in the 2000 ppb oxygen conditions, while a maximum CGR of 3 x 10-9 mm/s was observed under the low potential conditions. However, CGR were not reported for unaged material. The CGRs were also reported to be comparable to those for aged stainless steel weld materials in Yamada et. al., 2010 and Lucas et. a l., 2011.

It is noted that for the three studies that conducted SCC tests of ASSW in low potential environments, the average CGR for aged material is 3x10 -9 mm/s versus an average of 4x10-9 mm/s for unaged material, indicating no significant difference. For the three studies that tested

5 The 63 ppb hydrogen condition was intended to be representative of PWR water chemistry, to which 25-50 cc/kg of hydrogen is maintained (equivalent to about 2-4 ppm), or BWR HWC or HWC -NMCA conditions.

34 both aged and unaged ASSW in BWR high-potential environments, the average CGR of aged material was ten times that of the unaged material.

Based on the investigations discussed above, thermal aging at 400 ° C for 5000-10,000 hours0 days 0 hours 0 weeks 0 months accelerates the SCC CGR in high-potential BWR environments by about one order of magnitude, but results in no detectable differences in CGR in low potential PWR environments.

Since the BWR environments tested are high-potential (high dissolved oxygen) environments representative of BWR NWC, these results are not very relevant to US BWRs which generally implement a hydrogen water chemistry (HWC) regime, which has very low dissolved oxygen.

6.2 Effect of TE on Fatigue

The staff performed a literature search for research on the effects of TE on fatigue of ASSW.

Several studies were found (Raske et. al., 1977, James, 1978, Goyal et al., 2009, and Kumar et.

al., 2018), but all were on material aged at 593 °C or above, which is a much higher temperature than the LWR operating temperature range.These studies were focused on high-temperature reactor operating conditions rather than LWR conditions, and therefore are not considered applicable to ASSW aged in LWRs.

35 7 CONCLUSIONS

Thermal aging of ASSW at LWR operating temperatures causes a significant reduction in fracture toughness, as measured by the J -R fracture resistance curve, although there is considerable scatter in the available data for both as -welded (unaged) and thermally aged welds.

ASSW welds made with the GTAW process have higher thermally aged toughness than ASSW welds made using either the SAW or SMAW processes.

The thermally aged fracture toughness of ASSW can be comparable to that of thermally aged high-molybdenum CASS, particularly for flux welds (SAW and SMAW welds).

GTAW welds typically have higher thermally aged fracture toughness than SAW and SMAW welds, but still lower than that measured for wrought materials.

No new data has been found to warrant any change to the lower bound J-R fracture toughness curves in NUREG/CR-6428, Rev. 1 for SAW/SMAW at SLR service conditions.

Limited new data for GTAW suggests that the lower bound toughness curve should be reduced by at least 10% for SLR service conditions. However, additional J -R testing of aged GTAW weld materials would be helpful to confirm the proposed lower bound curve.

Two recent tests of Type 316L GTAW welds found anomalously low fracture toughness for both as-welded and thermally aged material, compared to the bulk of the fracture toughness data for ASSW GTAW welds.

There is not enough data on fracture toughness of aged ASSW in reactor coolant environments representative of most US LWRs to make any conclusions regarding whether there is an environmental effect on fracture toughness.

ASSW harvested from operating or decommissioning plants could be a good source of materials with long aging times which have been exposed to actual conditi ons in operating nuclear power plants, on which fracture toughness testing c ould be conducted.

The Z-factors in the ASME Code,Section XI, Nonmandatory Appendix C, do not account for thermally aged fracture toughness of ASSW.

Based on comparisons with the lower bound toughness curves in NUREG/CR-6428, Rev. 1, o the Z-factors in Section XI, Appendix C, C-6330(a), would be appropriate for flaw evaluation of ASSW welds made using the GTAW process.

o the Z-factors in Section XI, Appendix C for Category 2 ferritic material would be appropriate for flaw evaluation of ASSW welds made using the SAW or SMAW process (flux welds).

Thermal aging of ASSW has a relatively minor effect on SCC growth rate in LWR coolant en vironments, with essentially no effect in PWR environments, and an acceleration by a factor of ten in oxidizing BWR environments (similar to NWC).

With respect to the effect of TE on low -cycle fatigue CGRs, the only studies are of ASSW at much higher temperatures than the LWR service temperature range, so these results are not considered applicable to piping components operating in LWRs.

Therefore, testing on materials aged at temperatures closer to LWR operating

36 temperatures is needed to determine if TE of ASSW affects low -cycle fatigue crack growth for materials in service in an LWR environment.

Based on the evaluation of a limited number of heats, the WRC -1992 diagram is the best method for predicting the ferrite number of ASSW from chemical composition, because it provides balanced, slightly conservative predictions up to 12% measured ferrite, and has the smallest RMSD. The lower bound fracture toughness curves in NUREG/CR-6428, Rev. 1, do not take ferrite content into account. If a model were developed that used ferrite content as an input, an evaluation of the effectiveness of these methods w ith a larger database would be helpful.

Accelerated testing must be for sufficient time and temperature to represent SLR times and temperatures. Aging at 400 ° C for 10,000 hours0 days 0 hours 0 weeks 0 months may not be sufficient to represent SLR aging conditions. It may be possible to demonstrate saturation for shorter aging times. New testing can be done with accelerated aging conditions up to 400 ° C without compromising the integrity of the results when compared to properties for LTO.

There is limited evidence that higher ferrite weld materials will have lower thermally aged toughness, but there is insufficient data on ASSW with both ferrite content and J -R curves to develop a model correlating thermally aged toughness to ferrite content.

37 8 REFERENCES

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43 APPENDIX A DETERMINING FERRITE CONTENT OF ASSW

A.1 Methods of Determination of Ferrite Content

Ferrite content in ASSW may be determined in several ways:

Estimate from chemical composition using Hulls factors or other formulae.

Measured using magnetic instrument or ferritescope.

Estimate using quantitative metallography.

A.2 Estimation from Chemical Composition

NUREG/CR-6428, Rev. 1 recommends that ferrite content be estimated from the chemical composition using Hulls equivalent factors, which provide equations for the Cr and Ni equivalents (essentially the ferrite-stabilizing and austenite-stabilizing elements) (Hull, 1973).

The ferrite content is then calculated as a function of the ratios of the Cr and Ni equivalents.

The ASTM A800/800M methodology incorporates the Schoefer Diagram for the base metal of austenitic iron-chromium -nickel alloy castings. (ASTM, 2020a) ASTM A800/800M also uses a ratio of Cr to Ni equivalents calculated differently than Hulls. ASTM A800/800M references the Schaeffler Diagram and DeLong Diagram for estimation of ferrite content in weld deposits, both of which are also evaluated in this appendix. NUREG/CR-6428, Rev. 1 provides a comparison between the measured and calculated ferrite content for CASS materials for both the Hulls and ASTM A800/800M (ASTM, 2020). Ferrite was measured using a feritscope 6. It was found that the difference between the measured and calculated values using Hulls varied by plus or minus 6%, with the ferrite content being underpredicted for nickel contents 10%. NUREG/CR-6428, Rev. 1 also found that ASTM, 2020 underpredicted the ferrite content compared to the measured values for CASS materials with greater than 15% ferrite. The scope of ASTM, 2020 states that it applies only to CASS materials, not weld materials. The authors of NUREG/CR-6428, Rev. 1 compared the performance of Hulls equivalent factors and ASTM A800/800M with respect to the predictive accuracy for ferrite in CASS, but did not investigate the performance of these methods specifically for ASSW.

The Shaeffler Diagram and Delong Diagram are older methods for estimating the ferrite content of stainless steel weld metals (ASM, 1993). These methods also use a ratio of chromium to nickel equivalents and include multiple lines for different ferrite contents. The Shaeffler Diagram is reasonably accurate for conventional 300 series stainless steel weld deposits from covered electrodes, but is of limited use when less conventional compositions are used and when a high level of nitrogen is present. The DeLong Diagram accounts for the effect of nitrogen as an austenite stabilizer, and predicts the Welding Research Council (WRC) FN within +/- 3% in approximately 90% of the tests of Type 308, 309, 316 and 317 family weld metals (ASM, 1993).

After the standardization of ferrite number determination in 1974 in AWS A4.2 (AWS, 2020), the DeLong Diagram was found to seriously underestimate the ferrite content of weld metals with

6 The Feritscope is manufactured by Helmut Fischer and distributed by Fischer Technologies, Inc. in the US. An older model of the Feritscope may be calibrated using primary standards according to AWS, 2020. The manufacturers brochure for the Feritscope FMP -30 claims adherence to the measurement accuracy speci fied in standard ANSI/AWS A4.2M/A4.2:1997, and also claims traceability to NIST secondary standards.

high manganese content and overestimate the FN of highly alloyed weld metals, such as Type 309. This led to development by the WRC of the WRC-1988 and WRC -1992 diagrams. ASM, 1993 indicates that the Schaeffler Diagram still has some utility because it allows determination of martensite formation, which the WRC-1988 and WRC -1992 diagrams do not because neither diagram includes a manganese term. The WRC -1988 and WRC -1992 diagrams also include lines to indicate compositions for which primary austenite solidification (AF line) or primary ferrite solidification (FA line) occur (WRC, 1989, Kotecki et. al., 1992, ASM, 1993). Whether primary austenite or ferrite solidification occurs is important for prevention of hot cracking, but is not important with respect to thermal aging. The WRC-1992 diagram is also incorporated into the ASME Code,Section III, in Figure NB-2433.1-1, which is specified as the permissible method for delta ferrite determinations for welding materials based on chemical analysis. In addition, the WRC-1992 is called out in RG 1.31 (NRC, 2013) as an example of a constitutional diagram that may be used to estimate ferrite content from chemical composition. The NRC staff performed as assessment comparing the estimated versus measured ferrite content for five methods: Hulls, ASTM A800/800M. Shaeffler Diagram, Delong Diagram, and WRC -1992.

To investigate the accuracy of the ferrite estimation methods specific to ASSW, the staff used the data from NUREG/CR-6428, Rev. 1, Table A.1 for ASSW materials that have both chemical composition and measured ferrite data. There were 19 materials meeting this criterion. The staff plotted the predicted versus measured ferrite content for Hulls, A800/A800M, Schaeffler, DeLong, and WRC-1992 (Figure A-1 t hrough Figure A-5). The staff also determined the RMSD for the predicted versus the measured values for each method. Table A-1 lists the RMSD values for the various methods. The black solid line in each graph represents the one-to-one line (e.g. measured = predicted). Data points falling below the one-to-one line indicate underprediction. The dotted lines are the upper and lower bounds calculated by adding or subtracting the RMSD from the one-to-one line. Both Hulls and ASTM A800/800M underpredict the ferrite content for materials with measured ferrite content greater than 10%.

Table A-1 RMSD for Various Ferrite Estimation Methods

Method RMSD Hulls 4.36 ASTM A800/A800M 5.21 Schaeffler 3.94 DeLong 4.09 WRC-1992 3.95

A-2 Figure A-1 Hulls Equivalent Factors, Predicted vs. Measured Ferrite Content

Figure A-2 ASTM A800/A800M Calculated vs. Measured Ferrite Content for ASSW

A-3 Figure A-3 Schaeffler Diagram, Calculated vs. Measured Ferrite Content for ASSW

Figure A-4 Ferrite Number from DeLong Diagram vs. Measured Ferrite Content

A-4 Figure A-5 Ferrite Number from WRC-1992 Diagram vs. Measured Ferrite.

Figure A-1 and Figure A-2 show that both Hulls and ASTM A800/A800M generally underpredict the measured ferrite content at all ferrite levels from 5% to 20%. Figure A-3 shows that the Schaeffler Diagram overpredicts for measured ferrite up to at least 12%. Both the DeLong diagram (Figure A-4 ) and WRC-1992 diagram ( Figure A-5 ) predict measured ferrite fairly well up to 12%, with less conservative bias than the Schaeffler diagram at low ferrite. Both DeLong and WRC-1992 provide fairly balanced predictions up to 12% ferrite, relatively evenly split between slightly conservative and slightly nonconservative.

Based on the evaluation of a limited number of heats, the WRC -1992 diagram is the best method for predicting the ferrite number of ASSW, because it provides balanced, slightly conservative predictions up to 12% measured ferrite, and has one of the smaller RMSD values (Schaeffler was slightly lower at 3.94 vs. 3.95).

The staff recommends the use of the WRC -1992 for estimating the ferrite content of ASSW based on the following:

It is the most recently developed constitutional diagram correcting several deficiencies in early diagrams;

It is endorsed in RG 1.31 and the ASME Code,Section III;

The results of the staffs assessment in Appendix A show that the WRC-1992 diagram provides balanced, slightly conservative predictions up to 12% measured ferrite, and has the smallest RMSD.

A-5 A.3 Measurement Using Magnetic Instrument or Feritscope

Standard AWS A4.2M: 2020, Standard Procedures for Calibrating Magnetic Instruments to Measure the Delta Ferrite Content of Austenitic and Duplex Ferritic -Austenitic Stainless Steel Weld Metal, (AWS, 2020) provides a standard method for measuring ferrite number (FN) using magnetic instruments. The standard notes that FN is considered to be approximately equivalent to the volume percent ferrite, particularly at low FN values, but that FN may overstate the volume percent ferrite at higher ferrite numbers by a factor on the order of 1.3 to 1.5, which depends to a certain extent upon the actual composition of the alloy in question. Overstatement of FN for higher ferrite content materials may contribute to the apparent underprediction of ferrite content by the methods discussed in Section A-1. NUREG/CR-6428, Rev. 1 also indicates that the accuracy of measurements made with magnetic instruments is affected by the surface roughness of the material (due to the small size of the probe).

Per NUREG/CR-6428, Rev. 1, the feritscope Error! Bookmark not defined. operates on the magneto-induction principle and measures the relative magnetic permeability of the specimen. The manufacturers brochure for the Feritscope indicates that the instrument measures all magnetic portions (phases) of the otherwise non-magnetic structure (matrix) including strain-induced martensite in addition to delta ferrite. The brochure also indicates that the instrument can distinguish sigma phase, a Fe-Cr precipitate, from delta ferrite.

ASTM, 2020 also provides guidance under supplementary requirements for estimation of ferrite content by measurement of magnetic response, and references ASME Standard A799/A799M (ASTM, 2010).

A.4 Estimation of Ferrite Content U sing Quantitative Metallography

Quantitative metallography involves polishing and etching a specimen to reveal the ferrite and austenite phases, superimposing a grid over the image from an optical microscope and determining by point counting the percentage of ferrite in the material. NUREG/CR -6428, Rev.

1 indicates that quantitative metallography provides the most accurate estimate of ferrite content, but that it is tedious and obtaining metallographic samples from various regions of the weld may not be practical. ASTM E562 (ASTM, 2019) provides a standard methodology for quantitative metallography which is referenced in ASTM A800/800M.

A.5 Summary - Determining Ferrite Content

Several methods for measuring and estimating ferrite content in duplex stainless steel materials (including ASSW) have been discussed above. None of the methods are perfect. The staff does not recommend using Hulls factors (as recommended by NUREG/CR-6428, Rev. 1) to predict the ferrite content of ASSW, because it underpredicts the ferrite content for materials with ferrite content greater than 10%. With respect to prediction from chemical composition, the WRC-1992 method appears to be the best because it provides balanced, slightly conservative predictions up to 12% measured ferrite, and has the smallest RMSD. Another factor in support of using the WRC-1992 method is that it is incorporated into the ASME Code,Section III and referenced in RG 1.31.The lower bound fracture toughness curves recommended in NUREG/CR-6428, Rev. 1 do not take into account ferrite content. If a fracture toughness model using ferrite content as an input were developed, an evaluation of the effectiveness of these methods with a larger database would be helpful to confirm the optimal method for ferrite estimation.

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With respect to measuring ferrite content, quantitative metallography is the most accurate, but is tedious and time consuming. Therefore, use of a Feritscope or other magnetic instrument is more practical.Use of the magnetic test method is also endorsed in RG 1.31 (NRC, 2013) for all weld processes.

Since a model correlating thermally aged toughness with ferrite content has not been developed to date, knowledge of the ferrite content of a weld can only be used to determine a general tendency for a lower fully aged toughness with higher ferrite content. Another complicating factor is that the ferrite content predicted from chemical composition may vary significantly from the measured ferrite content due to the welding process.

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