ML22132A039

RIL-2022-005, NRC Technical Assessment of Zorita Materials Testing Results, Final May 2022

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RIL 2022-05 NRC TECHNICAL ASSESSMENT OF ZORITA MATERIALS TESTING RESULTS Date Published: May 2022 Prepared by:

M. Hiser P. Purtscher R. Tregoning Research Information Letter Office of Nuclear Regulatory Research

Disclaimer Legally binding regulatory requirements are stated only in laws, NRC regulations, licenses, including technical specifications, or orders; not in Research Information Letters (RILs). A RIL is not regulatory guidance, although NRCs regulatory offices may consider the information in a RIL to determine whether any regulatory actions are warranted.

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TABLE OF CONTENTS LIST OF FIGURES ....................................................................................................................iv LIST OF TABLES ......................................................................................................................vi EXECUTIVE

SUMMARY

..........................................................................................................vii ABBREVIATIONS AND ACRONYMS .......................................................................................ix Chapter 1: Background on Zorita Harvesting and Research .................................................1 1.1 Zorita-Related Research Programs .................................................................................1 Chapter 2: Discussion of Key Results from Zorita Baffle Plate Testing and Characterization ...................................................................................................................4 2.1 Overview of Zorita Plate Materials ...................................................................................4 2.2 Crack Growth Rate Testing Results for Zorita Baffle Plate Materials ...............................6 2.3 Representativeness of High Zorita CGRs ........................................................................9 2.4 Interpreting High CGRs on Zorita Materials ...................................................................11 2.5 Implications of High CGRs from Zorita Plate Materials ..................................................12 2.6 Void Swelling Evaluation from Transmission Electron Microscopy................................. 12 2.7 Implications of Void Swelling Results from Zorita Plate Materials .................................. 14 Chapter 3: Discussion of Key Results from Zorita Core Barrel Weld Testing and Characterization .................................................................................................................16 3.1 Overview of Zorita Weld and Heat-Affected Zone Materials........................................... 16 3.2 Fracture Toughness Testing Results for Zorita Weld and Heat-Affected Zone Materials ...................................................................................................................16 3.3 Implications of Low Fracture Toughness for Zorita Weld Materials ................................ 19 Chapter 4: Conclusions and Recommendations ..................................................................23 References ..............................................................................................................................24 Appendix A: Additional Zorita Plate Testing Results ......................................................... A-1 Appendix B: Additional Zorita Weld and Heat-Affected Zone Testing Results................. B-1 Appendix C: Additional Zorita Plate CGR Testing Tabular Data........................................ C-1 Appendix D: Additional References for Selected Figures.................................................. D-1 iii

LIST OF FIGURES Figure 1: Top: Stress-strain curves for Zorita tensile specimens in air at three fluence levels and two temperatures: 25°C (left) and 320°C (right). Bottom: Yield stress (left) and elongation after fracture (right) as a function of fluence for Zorita tensile specimens in air at three fluence levels and two temperatures (Ref: Figures 4-4, 4-5, 8-1 and 8-2 from MRP-440) ..............................................5 Figure 2: SEM fractography of fracture surfaces from 25 degrees C air tensile tests at 50 dpa (top) and 25 dpa (bottom). (Ref: Figures B-30 and B-47 from MRP-440.) Red boxes indicate locations of more detailed SEM images in MRP-440.......................................................................................................................6 Figure 3: Summary of constant K IASCC CGR data as a function of stress intensity factor in PWR and BWR HWC conditions normalized to 320 degrees C compared to the IASCC CGR model curve from ASME Code Case N-889 for this temperature/chemistry .........................................................................................8 Figure 4: Crack length and stress intensity factor as a function of time during the IASCC CGR test on Specimen B1CT01 during Steps 17a through 21c (Ref: Figure 6-11 from MRP-440) ............................................................................................8 Figure 5: Summary of constant K IASCC CGR data as a function of stress intensity factor in BWR NWC conditions normalized to 288 degrees C compared to the IASCC CGR model curve from ASME Code Case N-889 for this temperature/chemistry .........................................................................................9 Figure 6: Summary of the Zorita CGR data in PWR conditions compared to the IASCC CGR model from ASME Code Case N-889 as a function of fluence................... 11 Figure 7: Swelling observations for highest temperature location from Zorita baffle plate material. Examples of voids or bubbles shown in red circles. (Ref: Figure F-26 from MRP-440) .............................................................................................13 Figure 8: Top: Summary table of Zorita void swelling results (Ref: Table F-5 from MRP-440; bottom: Zorita void swelling results plotted as a function of temperature .......................................................................................................15 Figure 9: Summary of FT test results on Zorita weld and HAZ materials (Ref: Figure 7-12 from MRP-451) ..................................................................................................17 Figure 10: Summary of Zorita weld FT data as a function of dose compared to the data in the literature on irradiated SS weld FT (Ref: Figure 7-14 from MRP-451) .......... 18 Figure 11: J-R curves for selected Zorita HAZ specimens. Left: W1HCT04 at ~0.7 dpa tested in air at 320 degrees C Right: W2HCT06 at ~1.5 dpa tested in PWR primary water at 320 degrees C (Ref: Figures 6-4 and 6-12 from MRP-451)3 .... 18 Figure 12: Summary of Zorita HAZ FT data as a function of dose compared to the data in the literature on irradiated SS HAZ and base metal FT (Ref: Figure 7-15 from MRP-451)3 .................................................................................................19 Figure 13: Summary plot of irradiated SS weld FT data as a function of fluence for Zorita and related materials .........................................................................................20 iv

Figure 14: Side views of specimen W2HCT03 after testing, showing that the crack deviated from the plane defined by the side grooves during the FT test (Ref: Figure C-80 from MRP-451) ......................................................................22 v

LIST OF TABLES Table 1: Summary of Zorita Materials Research Programs .........................................................3 Table 2: Zorita Plate Composition Compared to Type 304 SS Specification ..............................4 Table 3: Chemical Analysis Results from the Zorita Weld Metal [11] ........................................ 16 vi

EXECUTIVE

SUMMARY

The purpose of this report is to review and assess the results from testing of irradiated stainless steel (SS) reactor internals harvested from the José Cabrera Nuclear Power Station (also known as Zorita) in Spain. The goal of this assessment is to briefly summarize the key results from this testing and succinctly identify relevant new information that may impact regulatory decisionmaking related to irradiation-assisted degradation of light-water reactor (LWR) vessel internals.

The Zorita reactor was a single-loop 160 megawatt-electric pressurized-water reactor (PWR) designed by Westinghouse Electric Corporation (Westinghouse) that achieved 26.4 effective full-power years of operation. The Zorita reactor internals were primarily composed of Type 304 austenitic SS. Harvesting from the Zorita internals included several pieces of SS baffle plate ranging from <1 to ~50 displacements per atom (dpa), as well as two pieces of core barrel SS weld ranging from <0.1 to ~2 dpa. The reactor internals were designed by Westinghouse and are very similar to Westinghouse-designed U.S. PWRs, which constitute the majority of the U.S. PWR fleet. These materials are some of the most representatively aged, high-fluence irradiated SS components that have been studied.

A number of research programs have been performed on the Zorita baffle plate and core barrel weld materials at a variety of laboratories in the U.S., Sweden, and Norway. The most significant results from the testing of Zorita baffle plate materials are the repeated observations of high crack growth rate (CGR) during irradiation-assisted stress corrosion cracking (IASCC) CGR testing and the very low amount of observed void swelling. The most significant results from the testing of Zorita weld materials are the very low fracture toughness (FT) values observed in multiple tests.

The Zorita baffle plate void swelling data should be taken as encouraging in that void swelling may not progress as rapidly in light water reactors (LWRs) as previously suggested. However, due to the lower operating temperatures of Zorita, the results cannot conclusively eliminate the potential for significant void swelling, particularly at higher doses and temperatures. The Zorita void swelling data show the strong influence of temperature, consistent with the results from other data in the literature. Industry and regulators should seek to observe additional LWR-irradiated materials at higher doses and temperatures near 360 degrees C to more confidently conclude that void swelling will not pose a significant issue during extended operating periods.

The Zorita baffle plate CGR data suggest that the IASCC CGR model for American Society of Mechanical Engineers (ASME) Code Case N-889 does not sufficiently predict the increased IASCC CGRs at fluences above 20 dpa observed in this material. This deficiency supports the proposed NRC condition on this Code Case, limiting its applicability to materials less than 20 dpa.

Given the small volume of LWR internals exceeding 20 dpa, the practical implications of this condition are likely to be limited in the near term. When assessing the significance of the high CGRs on Zorita materials at high fluence levels, it should also be recognized that these data come from one heat of material irradiated in one reactor. Heat-to-heat variability can lead to significant uncertainty in materials testing, so additional CGR testing of highly irradiated materials should be pursued where practical to augment the Zorita plate testing data and confirm or refute the observations from the Zorita CGR testing.

The Zorita core barrel weld FT data should be carefully considered when assessing irradiated SS weld embrittlement, particularly given the very limited amount of data from in-service welds. The vii

Zorita data should be used to update existing guidance on irradiated SS weld FT as contained in BWRVIP-100, Revision 1, and WCAP-17096. Given the low susceptibility to IASCC, low operating stresses, and flaw-tolerant design of the BWR core shroud and PWR core barrel, it is not expected that these lower weld FT data pose an immediate safety concern. However, these results may necessitate reduced inspection intervals compared to previous guidance to ensure that an acceptable margin to structural integrity exists. Further research on irradiated SS weld materials should prioritize generating additional data at fairly low fluence levels (<2 dpa) and extend data on irradiated SS weld properties up to higher fluences approaching 20-30 dpa.

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ABBREVIATIONS AND ACRONYMS ADAMS Agencywide Documents Access and Management System ANL Argonne National Laboratory ASME American Society of Mechanical Engineers BWR Boiling water reactor BWRVIP Boiling Water Reactor Vessel and Internals Project CGR Crack growth rate EPRI Electric Power Research Institute FT Fracture toughness GB Grain boundaries HAZ Heat-affected zone HWC Hydrogen water chemistry IASCC Irradiation-assisted stress corrosion cracking INL Idaho National Laboratory LWR Light water reactor MRP Materials Reliability Program NEA Nuclear Energy Agency NWC Normal water chemistry OECD Organization for Economic Co-operation and Development PWR Pressurized water reactor PWSCC Primary water stress corrosion cracking RPV Reactor Pressure Vessel SCC Stress corrosion cracking SEM Scanning electron microscopy SMILE Studsvik Materials Integrity for Life Extension SS Stainless steel TEM Transmission electron microscopy WCAP Westinghouse Commercial Atomic Power ZIRP Zorita Internals Research Project ix

CHAPTER 1: BACKGROUND ON ZORITA HARVESTING AND RESEARCH The Zorita reactor operated from 1969 to 2006 with 26.4 effective full-power years of reactor operation. It was a single-loop 160 megawatt-electric pressurized-water reactor (PWR) designed by Westinghouse Electric Corporation (Westinghouse). Harvesting of reactor internals from the Zorita reactor was primarily carried out by the Zorita Internals Research Project (ZIRP),

which was led by the Electric Power Research Institute (EPRI) and included cooperative funding from the NRC and other international organizations [1].

The Zorita reactor internals were primarily composed of Type 304 austenitic SS. Harvesting from the Zorita internals included several pieces of SS baffle plate ranging from <1 to ~50 displacements per atom (dpa), as well as two pieces of core barrel SS weld ranging from <0.1 to ~2 dpa. The reactor internals were designed by Westinghouse and are very similar to Westinghouse-designed U.S. PWRs, which constitute the majority of the U.S. PWR fleet.

Therefore, the materials are highly representative of the SS internals in operating U.S. plants and may be used to validate (or refute) findings from other irradiated materials that may have experienced less representative aging or irradiation conditions, such as accelerated thermal aging or test reactor irradiation conditions with very high neutron fluxes or fast neutron spectrums.

1.1 Zorita-Related Research Programs The initial harvesting of Zorita internals materials, including baffle plate and core barrel weld materials, as well as a three-dimensional fluence and temperature analysis (Materials Reliability Program (MRP) 392, Materials Reliability Program: Zorita Internals Research Project: Radiation and Temperature Analysis of Zorita Baffle Plate and Core Barrel Weld Material [2]), was performed under the ZIRP. EPRI led the multinational ZIRP project team, which included funding or in kind contributions from the NRC, Spanish Nuclear Safety Council (Spanish regulator), Swedish Radiation Safety Authority (Swedish regulator), Mitsubishi Heavy Industries (MHI), Axpo Holding (Swiss utility), and Tractebel (Belgian engineering consultancy). In addition to harvesting, ZIRP performed a thorough testing program of Zorita baffle plate materials at Studsvik (Swedish testing laboratory), covering tensile, irradiation assisted stress corrosion cracking (IASCC) crack initiation, IASCC crack growth, and fracture toughness (FT) (MRP-440, Materials Reliability Program: Zorita Internals Research Project (MRP-440), Testing of Highly-Irradiated Baffle Plate Material [1]). Additional in-kind effort for ZIRP performed by MHI included TEM to assess irradiation damage and void swelling. A follow-on program with a similar scope, funded by the NRC and EPRI at Studsvik, was performed and included optical microscopy, tensile, IASCC crack growth, and FT testing on Zorita core barrel weld materials (MRP-451, Materials Reliability Program: Fluence Effects on Stainless Steel Welds (MRP-451): Crack Growth Rate and Fracture Toughness Testing of Zorita Weld and HAZ Materials [3]).

To supplement the testing on as-received Zorita materials, in 2016, the NRC and EPRI supported the machining and shipment of baffle plate and core barrel specimens from Studsvik to the Halden Reactor in Norway to enable further irradiation of Zorita core barrel weld and heat-affected zone (HAZ) materials to generate data at higher fluences. The Halden reactor permanently shut down unexpectedly in 2018 before significant fluence could be acquired on the Zorita core barrel weld and HAZ samples. However, IASCC crack growth rate (CGR) tests were performed on a small number of Zorita baffle plate samples in the range of 40-50 dpa and core barrel weld and HAZ samples at 1-2 dpa (HWR-1236, Final Report on the BWR Crack 1

Growth Rate Investigation IFA-791 [4] and HWR-1320, Interim Report on the PWR Crack Growth Rate Investigation IFA-817 [5].

Following some observations of higher CGRs in the plate materials and low FT values in the Zorita weld/HAZ materials, the NRC funded the machining and shipping of two sets of additional specimens for independent testing at Argonne National Laboratory (ANL). These Zorita materials have been tested primarily for CGR and FT (results through 2019 are available in ANL 19/45, Crack Growth Rate and Fracture Toughness Tests on Irradiated Ex-Plant Materials [6])

as well as fractography and TEM (ANL-20/50, Irradiated Microstructure of Zorita Materials [7]).

In addition, some of these materials have also been characterized for chemical composition by mass spectroscopy and TEM at Idaho National Laboratory (INL) to confirm they are typical Type 304 SS and assess potential transmutation effects at higher doses (INL/EXT-21-62220, Chemical Compositional Analysis and Microstructural Characterization of Harvested Zorita Reactor Pressure Vessel (RPV) Internals [8]).

In addition to the aforementioned testing programs that included NRC funding in cooperation with the EPRI MRP for PWR internals, EPRI also funded multiple projects, primarily through the Boiling Water Reactor Vessel and Internals Project (BWRVIP). Two key reports documenting BWRVIP testing of the Zorita materials are BWRVIP-294, Revision 2, Fracture Toughness of Zorita RPV Core Internals Applicable to BWRs: Final Report 2019, issued October 2019 [9] and BWRVIP-335, BWR Vessel and Internals Project, Crack Growth Rate Testing of Zorita Core Barrel Materials in BWR Environments, issued August 2020 [10]. The NRC does not have access to these data but has some awareness of a portion of the data through a conference publication [11]. Table 1 below provides a summary of the Zorita materials research programs, including funders, publications, scope, etc.

In the following sections, the authors summarize the key results from these testing programs and assess the major takeaways and relevance of the results for regulatory decisionmaking related to irradiation assisted degradation of LWR internals. Section 3 addresses significant results from the Zorita baffle plate materials, where high-fluence irradiation effects are the major focus. Section 4 addresses key results from the Zorita core barrel weld materials, which received lower fluence but, based on operating experience, represent a region that is more susceptible to cracking. Finally, Appendices A and B summarize the remainder of the results that are of lower importance to regulatory decisions for irradiation assisted degradation of LWR internals.

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Table 1: Summary of Zorita Materials Research Programs BWRVIP ZIRP Zorita Welds Halden ANL/INL Testing NRC, EPRI, NRC and Halden Funders NRC EPRI and others EPRI members ANL-19/45, BWRVIP-HWR-1236 ANL-20/50 and 294, Rev. 2, Publications MRP-440 MRP-451 and INL/EXT and HWR-1320 62220 BWRVIP-335 Testing MHI and Studsvik Halden ANL & INL Studsvik Location Studsvik 40-50 dpa <1-50 dpa plate 10, 25, 50 1, 2 dpa Materials plate, 1-2 (<0.1-1 dpa weld 1-2 dpa weld dpa plate weld/HAZ dpa weld/HAZ also available)

Tensile Testing X X unknown Crack Initiation X unknown CGR Testing X X X X X FT Testing X X X X Optical / SEM X X X X unknown TEM X X X unknown Composition X X 3

CHAPTER 2: DISCUSSION OF KEY RESULTS FROM ZORITA BAFFLE PLATE TESTING AND CHARACTERIZATION The most significant results from the testing of Zorita baffle plate materials are the very low amount of observed void swelling and the repeated observations of high CGR during IASCC CGR testing. This section focuses on these results and a detailed assessment of their relevance to regulatory decisions. Appendix A to this report and MRP-440, ANL-19/45, and HWR-1320 provide a more comprehensive summary of the overall testing of Zorita baffle plate materials by the NRC and EPRI.

2.1 Overview of Zorita Plate Materials Several pieces from the Zorita baffle plate were removed, covering a range of fluence from <1 to 50 dpa. The materials tested in the initial ZIRP program were targeted at 10, 25, and 50 dpa to assess the impact of fluence on microstructure and properties. The baffle plate was fabricated from 28.6-millimeter (mm)-thick plates of Type 304 SS. Documentation, such as certified material test report, on the base and weld metals used was not available. However, chemical analyses of Zorita plate specimens at INL [8] for the Zorita baffle plate materials showed that the composition was within the specifications of Type 304 SS (see Table 2).

Table 2: Zorita Plate Composition Compared to Type 304 SS Specification Carbon Sulfur Nitrogen Chromium Manganese Nickel Silicon Phosphorus Zorita 0.04% 0.02% 0.04% 19.7% 1.48% 9.85% 0.28% 0.02%

Analysis Type 304 SS 0.08% 0.03% 0.10% 18-20% 2% max 8-10.5% 0.75% 0.045% max Specification max max max max The tensile testing results are shown in context with literature data in Appendix A. However, one important observation from the tensile testing was intergranular brittle failure in the room temperature air tensile test at 50 dpa. This is in contrast to the tensile tests performed at lower fluence levels and at higher temperatures, where ductile fracture behavior and fracture surfaces were observed. This intergranular brittle failure at 50 dpa indicates a change in ductility and failure mode due to increased fluence, which is important to highlight as it may help understand the increased IASCC CGR observed at higher fluence.

Figure 1 shows the trend with fluence and temperature for the yield stress and elongation after fracture for the nine air tensile tests across three fluence levels and two temperatures, with the highest dose specimens failing in a brittle fashion at much lower strain than the lower dose specimens at room temperature. Figure 2 shows the fracture surface from the room temperature tensile tests at ~25 and ~50 dpa, respectively. The 50-dpa fracture surface is almost completely intergranular and exhibited little necking, while all other samples showed ductile behavior:

dimple fracture and significant necking.

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Figure 1: Top: Stress-strain curves for Zorita tensile specimens in air at three fluence levels and two temperatures: 25°C (left) and 320°C (right). Bottom: Yield stress (left) and elongation after fracture (right) as a function of fluence for Zorita tensile specimens in air at three fluence levels and two temperatures (Ref: Figures 4-4, 4-5, 8-1 and 8-2 from MRP-440) 5

Figure 2: SEM fractography of fracture surfaces from 25 degrees C air tensile tests at 50 dpa (top) and 25 dpa (bottom). (Ref: Figures B-30 and B-47 from MRP-440.) Red boxes indicate locations of more detailed SEM images in MRP-440.

2.2 Crack Growth Rate Testing Results for Zorita Baffle Plate Materials Researchers at a number of facilities have performed crack growth testing of Zorita baffle plate materials. Initially, the ZIRP project funded several CGR tests on Zorita plate materials ranging from 10 to 50 dpa at Studsvik. Next, a few samples ranging from 40 to 50 dpa were sent to Halden and tested as part of the Halden Reactor Project. Finally, two shipments of Zorita plate specimens (second shipment also included some Zorita weld specimens) were sent to ANL for independent testing funded by the NRC.

These tests have produced a significant amount of CGR data on Zorita plate materials in PWR and boiling water reactor (BWR) hydrogen water chemistry (HWC) environments, which are summarized in Figure 3. All data shown in Figure 3 were collected within valid stress intensity 6

factor (K) conditions 1 and normalized to 320 degrees Celsius (C) in accordance with the IASCC 0F0F CGR model from American Society of Mechanical Engineers (ASME) Code Case N-889, Reference Stress Corrosion Crack Growth Rate Curves for Irradiated Austenitic Stainless Steels in Light Water Reactor Environments [12,13]. In general, a single data point is reported for test segments where testing conditions remain consistent (e.g., environment, stress intensity factor, temperature). The data in Figure 3 show a bimodal CGR behavior in PWR/BWR HWC conditions, in which cracking progresses quite slowly for portions of the test before accelerating very rapidly by multiple orders of magnitude to higher CGRs (generally well beyond the IASCC CGR model in ASME Code Case N-889) during other portions of the test. The data in Figure 3 also shows that the likelihood of high CGR increases at higher fluences. Figure 4 provides an example of one of these higher CGR events, in this case Steps 18 and 19 from an ~50-dpa test.

It should be noted that many of the lower CGR data (<10-7 mm/second) in Figure 3 cover very little crack extension, which may indicate a lack of engagement of the crack, which is a common challenge when performing this type of testing. It should also be noted that in testing at ANL high CGRs were only observed when testing at elevated K levels > 25 MPam.

Testing has also been performed in BWR normal water chemistry (NWC) environments, with the results summarized in Figure 5. High CGRs are observed in BWR NWC conditions, with data above the IASCC CGR model from ASME Code Case N-889 in almost every instance. It should be noted that BWR internals are not expected to see fluences beyond 20-30 dpa [15], and nearly all U.S. BWRs now operate regularly with low-potential HWC conditions. Therefore, the applicability of BWR NWC data to operating plants at these high fluences is limited but nonetheless may be indicative or insightful of the IASCC susceptibility of these highly irradiated Zorita materials.

1. Valid stress intensity factor was determined based on ASTM E399 using an effective yield strength discounting 50% of the irradiation induced hardening as discussed by Jenssen et al. [14].

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Figure 3: Summary of constant K IASCC CGR data as a function of stress intensity factor in PWR and BWR HWC conditions normalized to 320 degrees C compared to the IASCC CGR model curve from ASME Code Case N-889 for this temperature/chemistry Figure 4: Crack length and stress intensity factor as a function of time during the IASCC CGR test on Specimen B1CT01 during Steps 17a through 21c (Ref: Figure 6-11 from MRP-440) 8

Figure 5: Summary of constant K IASCC CGR data as a function of stress intensity factor in BWR NWC conditions normalized to 288 degrees C compared to the IASCC CGR model curve from ASME Code Case N-889 for this temperature/chemistry 2.3 Representativeness of High Zorita CGRs The high CGRs were developed under testing conditions that are representative of material behavior for reactor internals and therefore could occur if similar conditions arise in service.

Based on the testing of the Zorita materials, the conditions that appear to lead to high CGRs are a K > 15 MPam combined with a fluence above 25 dpa. The following bullets summarize some of the factors that were considered in assessing the relevancy of the CGR results to operating plant conditions, while the sub-bullets provide context for what has been observed in the Zorita materials testing:

• Testing transients are representative of operating conditions.

- In some cases, a modest transient such as a planned change in temperature (e.g.,

from 340 degrees C to 290 degrees C) or loading (e.g., transition from partial periodic unloading to constant K) or unplanned change in chemistry preceded an acceleration of the CGR by a few hours or days. However, there is no clear metallurgical explanation for why a modest change in temperature, loading, or chemistry should cause high CGR; changes of this magnitude routinely occur during plant operation.

- These types of modest transients occur in primary water stress corrosion cracking (PWSCC) and IASCC CGR testing regularly without initiating high CGR.

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- There are also many examples in the Zorita CGR testing in which no apparent change in testing conditions precedes the initiation of high CGR. Therefore, the testing conditions are anticipated to be representative, including for the observed high CGR.

• Zorita CGR results are replicated at multiple laboratories and under multiple principal investigators using slightly different testing techniques and loading histories.

- The high CGR behavior has been observed in many Zorita specimens tested at multiple laboratories with extensive experience in this type of testing, so it is not likely to be a result of an artifact of a certain testing approach by a particular investigator at a particular laboratory.

- High CGRs have been observed even more consistently in testing in BWR NWC conditions, as shown in Figure 5, further supporting the idea that this is an inherent behavior of these highly irradiated materials.

- The Zorita CGR tests were run in a similar manner to the way SCC CGR tests have been run historically on SS and nickel-based alloys for IASCC and PWSCC.

• K-validity criteria are met for most high CGR events.

- All data have been analyzed for K validity consistent with accepted practices for irradiated SSs as discussed in Jenssen et al. [14], and most of the high CGR events occurred within K valid portions of the tests. The data presented in Figures 3 and 5 only include the K-valid portions of the tests, which include many high CGR events.

• Testing conditions during high CGR are representative of plant conditions.

- EPRI and its contractors have proposed an explanation that high crack tip strain rates are the cause of the high CGR behavior and are not representative of reactor internals loading [1].

- The EPRI explanation centers around the idea that during testing as CGR increases, insufficient load shedding occurs, leading to an increasing applied K that creates the high CGRs in an unrealistic manner relative to plant operations. It is true that in many instances, insufficient load shedding is observed during these high CGR events.

However, this explanation has two key deficiencies:

(1) The initiation and acceleration of high CGR occurs first without an increasing K (increases in K then happen in some cases because when high CGRs occur, it is challenging to decrease load quickly enough to maintain a constant K). EPRIs proposed explanation for the high CGRs does not explain what causes the initiation and acceleration of high CGR. The most plausible explanation appears to be the inherent stress corrosion cracking (SCC) susceptibility of the material.

(2) In some instances, decreasing K loading has been successfully applied during testing once the high CGRs were observed (see Steps 18 and 19 in Figure 4).

However, decreasing the applied K did not significantly decrease the high CGR.

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2.4 Interpreting High CGRs on Zorita Materials After extensive review of the Zorita CGR testing results from testing programs at Studsvik, Halden, and ANL, the most plausible explanation is that these high CGR data are representative of the IASCC susceptibility of these highly irradiated materials at elevated fluence. The high CGR events have been observed consistently across multiple specimens tested in multiple laboratories in a variety of environmental and loading conditions. Figure 6 shows the difference between observed IASCC CGR data in PWR conditions on Zorita materials and the IASCC CGR model prediction from ASME Code Case N-889 as a function of fluence. This plot clearly shows the increasing propensity for high CGR to occur at higher fluence levels, starting at about 25 dpa. It should be noted that the high CGR data tends to be about 5¬-50 times higher than the IASCC CGR model from ASME Code Case N-889 predicts, which can have a significant impact on how large flaws can grow unless they are detected by periodic inspections.

Figure 6: Summary of the Zorita CGR data in PWR conditions compared to the IASCC CGR model from ASME Code Case N-889 as a function of fluence Researchers have made efforts to understand the microstructure of the Zorita materials to help explain this high CGR behavior [1,7,8]. The IASCC mechanism is understood to generally be driven by changes in microstructure due to radiation hardening and changes in microchemistry due to radiation-induced segregation at the grain boundaries (GBs) causing a decrease in local chromium content, allowing SCC to progress more easily [16]. It is possible that the high CGRs observed in the Zorita materials may be due to further progression of radiation-induced segregation or other aging mechanisms occurring locally at either the GBs or more generally in the material. Microstructural investigations on the Zorita materials have found an increase in 11

void fraction with increasing dose and temperature, although no clear evidence of an accumulation of voids at the GBs [7,8]. Other work on highly irradiated SSs has shown that voids can populate along the GBs and cause a GB weakening effect [17]. This GB weakening mechanism with increasing fluence is also consistent with the observation of intergranular brittle fracture in the 50-dpa air tensile test described previously. The combination of extensive radiation-induced segregation and GB weakening due to increasing voids provides a plausible explanation for the increasing susceptibility to higher CGRs with increasing fluence.

2.5 Implications of High CGRs from Zorita Plate Materials This high CGR data from Zorita suggest that the IASCC CGR model for ASME Code Case N-889 may be insufficient to predict IASCC CGRs at fluences above 20 dpa. CGR data from other irradiated materials at fluence levels above 20 dpa are sparse and in general have been generated from less representative irradiation conditions, such as fast spectrum test reactors.

The higher fluence data from Zorita, therefore, should be given considerable weight when assessing the structural integrity of reactor internals irradiated to high fluence (>20 dpa). Flaw evaluations performed on irradiated materials in this fluence that range above 20 dpa may need an additional safety factor beyond the IASCC CGR model for ASME Code Case N-889 to account for the observed high CGRs from the Zorita materials (which were not included in the development of the IASCC CGR model for ASME Code Case N-889).

Consistent with these observations, NRC staff have proposed three conditions on Code Case N-889 in draft Regulatory Guide DG-1367, Inservice Inspection Code Case Acceptability, ASME Section XI, Division 1. [18] The first condition limits the applicability of the Code Case to fluence levels less than 20 dpa. The remaining two conditions address deficiencies in the Code Case at predicting CGRs at low fluence levels below 0.75 dpa and with predicting irradiated yield stress for cold-worked non-molybdenum bearing SSs.

Other factors to consider are that highly irradiated internals components tend to experience very low operational stresses during normal operation, and irradiation-induced stress relaxation tends to reduce fit up or residual stresses from fabrication/construction. These factors may mitigate the effects of susceptibility to higher CGRs by minimizing the applied stresses that drive crack growth.

It should also be recognized that the data are from probably only one heat of material irradiated in one reactor. Heat to-heat variability is a known phenomenon in materials testing, so additional CGR testing of highly irradiated materials should be pursued where practical to augment the Zorita plate testing data and confirm or refute the observations from the Zorita CGR testing. The OECD/NEA SMILE project, running from 2021 to 2025, plans to harvest and test Type 304 SS internals materials from the Ringhals 2 PWR with fluences up to 50 dpa. The SMILE data should help to provide additional understanding of the IASCC behavior of high-fluence PWR internals.

2.6 Void Swelling Evaluation from Transmission Electron Microscopy As part of the ZIRP project, Mitsubishi Heavy Industries performed detailed TEM on the high dose (50 dpa) sample. This TEM work provided information on various characteristics of irradiation damage, including precipitates, Frank loops, hardness, gas generation, and void swelling. Of particular interest were estimates of void swelling, which has been proposed as a potential life-limiting aging mechanism in irradiated SSs [19]. Figure 7 shows an example TEM 12

image from the highest swelling location. The three TEM images to the right side of Figure 7 show the visual fields analyzed for swelling (voids appear as hazy circles).

Figure 7: Swelling observations for highest temperature location from Zorita baffle plate material. Examples of voids or bubbles shown in red circles. (Ref: Figure F-26 from MRP-440)

The void swelling mechanism in SSs is complex and sensitive to a number of factors, including dose, dose rate, irradiation temperature, and neutron energy spectrum. Past observations of void swelling from fast reactors should be interpreted cautiously because the irradiation conditions can differ in the following important respects from those from LWRs:

• Irradiation temperature: Void swelling is known to increase strongly with increasing temperature [16]. Fast reactor data on void swelling have shown significant swelling at temperatures above 400 degrees C [19]. Peak temperatures in LWR internal components (accounting for gamma heating) are calculated to be about 360 degrees C, with most of the internals at lower temperatures than 360 degrees C [19].

• Neutron energy spectrum: Thermal neutrons tend to generate helium gas in SS during irradiation. Helium has been shown to help voids stabilize and grow more readily, causing an increase in void swelling. Therefore, the neutron energy spectrum (fast only or mixed as in an LWR) influences the generation of helium gas and the rate of swelling.

From a neutron energy spectrum perspective, fast reactor irradiation conditions (which 13

lack a significant thermal neutron flux) would tend to produce less swelling than LWR irradiation conditions [19].

• Dose rate: Higher dose rate generally causes the onset of significant swelling at a higher dose [16,19]. Therefore, for a given dose (fluence), swelling is negatively correlated to dose rate (more swelling with a lower dose rate). Fast reactor irradiations generally create a much higher dose rate than LWR service conditions, so fast reactor irradiations will tend to underpredict swelling at a given dose based on their higher dose rates.

Based on these multiple competing factors that differ between fast reactors and LWRs, the results of void swelling from high-dose LWR-irradiated materials are particularly insightful for understanding how void swelling may progress during extended LWR operation. The top part of Figure 8 shows the results from the TEM analyses of void swelling in the Zorita materials, which indicate very low levels of void swelling (<0.08 percent) in all observed locations, including both the maximum temperature and maximum fluence locations (note the maximum temperature and fluence occur in different locations). The bottom part of Figure 8 shows that the Zorita swelling results correlate much more strongly with temperature than with dose, consistent with the data trends shown in the literature [19].

2.7 Implications of Void Swelling Results from Zorita Plate Materials It should be noted that the Zorita reactor operated at a slightly lower temperature than many U.S. PWRs. Specifically, the maximum temperature was <330 degrees C in the Zorita internals, while many U.S. PWRs are expected to see internal components with peak temperatures near 360 degrees C. These data should therefore be taken as an encouraging sign that void swelling may not progress as rapidly in LWRs as previously suggested [19], but not conclusively to eliminate the potential for significant void swelling, particularly at higher doses and temperatures. Even the Zorita void swelling data show a strong influence of temperature as seen in Figure 8.

After previously expecting void swelling to reach a steady-state rate of 1% per dpa based on data at higher temperatures, more recently [20] has determined that a maximum steady-state swelling rate of 0.07% per dpa (after the incubation period) is more likely for LWR conditions.

This appears to be consistent with the fairly low swelling observed in the Zorita materials.

Nevertheless, industry and regulators should seek to observe additional LWR irradiated materials at higher doses and temperatures near 360 degrees C to more confidently conclude that void swelling will not pose a significant issue during extended operating periods.

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Figure 8: Top: Summary table of Zorita void swelling results (Ref: Table F-5 from MRP-440);

bottom: Zorita void swelling results plotted as a function of temperature 15

CHAPTER 3: DISCUSSION OF KEY RESULTS FROM ZORITA CORE BARREL WELD TESTING AND CHARACTERIZATION The most significant results from the testing of Zorita weld materials are the very low FT values observed in multiple tests. This section details these results and assesses their relevance to regulatory decisions. Appendix B and MRP-451 and BWRVIP-335 provide a more comprehensive summary of the overall testing of Zorita weld and HAZ materials by the NRC and EPRI [3,10].

3.1 Overview of Zorita Weld and Heat-Affected Zone Materials Two separate welds from the Zorita core barrel were removed. Weld 1 was extracted from an axial (vertical) core barrel weld towards the top of active fuel and had a maximum fluence of

~1 dpa (also used in BWRVIP testing). Weld 2 was harvested from an axial (vertical) core barrel weld at the core midplane and includes an intersection with a circumferential (horizontal) weld.

Weld 2 had a maximum fluence of ~2 dpa. The materials tested in the EPRI-NRC program ranged in dose from 0.7 to 1.9 dpa. The core barrel was fabricated from 45-mm-thick plates of Type 304 SS. Metallographic samples from both welds showed that the measured ferrite contents were in the range of 5-7 percent, which is normal for SS welds. Documentation (such as certified material test report) on the base and weld metals used was not available. However, chemical analyses of each of the two welds were typical for Type 308 weld metal, as shown in Table 3 [11].

Table 3: Chemical Analysis Results from the Zorita Weld Metal [11]

Carbon Manganese Sulfur Chromium Iron Nickel Molybdenum Copper Zorita 0.06% 1.47% 0.017% 21.5% 70.5% 9.9% 0.04% 0.13%

Analysis Type 308 SS 0.08% 1.0-2.5% 0.03% 19.5-22.0% Balance 9.0- 0.75% max 0.75%

Specification max max 11.0% max 3.2 Fracture Toughness Testing Results for Zorita Weld and Heat-Affected Zone Materials FT testing of Zorita weld and HAZ materials in air and PWR environments has been performed at Studsvik according to ASTM E1820 through a cooperatively funded project by the NRC and EPRI. Additional FT tests have also been performed at Studsvik under BWR conditions with separate EPRI funding [11]. Figure 9 shows the Zorita weld and HAZ FT testing results as a function of test temperature.

The Zorita weld metal tests displayed low FT such that a stable J-integral vs. crack growth resistance (J-R) curve was not able to be constructed and work per unit of fracture surface area (J) at 1 mm crack extension could not be reported 2 [3]. Unstable crack advance occurred in all 1F1F these specimens; therefore, the linear-elastic fracture mechanics test standard ASTM E399 was used to evaluate the data instead. As seen in Figure 9, there were no significant effects of temperature, dose, or test environment on the Zorita weld metal FT over the range of conditions

2. EPRI has indicated that they are working to develop J-R curves to better analyze these FT tests [21].

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evaluated. Some of the fracture surfaces of low-toughness Zorita weld specimens show fine-scale dimpled features, suggesting that the low toughness may be associated with the size and distribution of particles, such as inclusions that formed on initial solidification of the weld. Figure 10 compares the Zorita weld FT data to that in the literature and shows clearly that the Zorita weld FT data values are much lower than those from the limited data in the literature at similar fluences.

In contrast to the weld metal tests, FT testing of the HAZ resulted in increased resistance to ductile tearing in comparison with the weld material, and J-R curves were constructed. An effect of dose and environment on the measured J at 1 mm was observed. From power law curve fits (i.e., J = C (a)n) to the data, the value of coefficient C decreased from ~300 kilojoules per square meter (kJ/m2) at ~0.7 dpa in air to ~200 kJ/m2 at ~1.5 dpa in a PWR environment as shown in Figure 11. Figure 12 shows that the Zorita HAZ FT was relatively low but, unlike the weld metal data, it fell within the observed range of literature data for HAZ and base metal.

Figure 9: Summary of FT test results on Zorita weld and HAZ materials (Ref: Figure 7-12 from MRP-451) 3 2F2F

3. FT is expressed as Kc, which is calculated the same as KJc in ASTM E1921.

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Zorita data Figure 10: Summary of Zorita weld FT data as a function of dose compared to the data in the literature on irradiated SS weld FT (Ref: Figure 7-14 from MRP-451)3, 4 3F3F Figure 11: J-R curves for selected Zorita HAZ specimens. Left: W1HCT04 at ~0.7 dpa tested in air at 320 degrees C Right: W2HCT06 at ~1.5 dpa tested in PWR primary water at 320 degrees C (Ref: Figures 6-4 and 6-12 from MRP-451)3 4 Please refer to Appendix D for citations shown in this figure.

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Zorita data Figure 12: Summary of Zorita HAZ FT data as a function of dose compared to the data in the literature on irradiated SS HAZ and base metal FT (Ref: Figure 7-15 from MRP-451)3,4 3.3 Implications of Low Fracture Toughness for Zorita Weld Materials Based on Figures 10 and 12, the most significant results from this testing are the very low FT results for the weld metals. There are very limited data on irradiated SS weld materials, and the Zorita welds are representative of the SS welds that are in operating U.S. reactors. Therefore, these new data should be carefully considered to determine whether past assumptions about irradiated SS weld FT are still appropriate. The NRC has approved using the criteria in BWRVIP-100, Revision 1-A, BWR Vessel and Internals Project, Updated Assessment of the Fracture Toughness of Irradiated Stainless Steel for BWR Core Shrouds, issued February 2017

[22] for irradiated SS weld FT to assess flaws in BWR core shrouds. Identical values for FT in irradiated SSs are also contained in Westinghouse Commercial Atomic Power (WCAP)-17096, Revision 2, Reactor Internals Acceptance Criteria Methodology and Data Requirements, issued 2006 [23], which is referenced in MRP-227. Revision 1-A, "Materials Reliability Program:

Pressurized Water Reactor Internals Inspection and Evaluation Guidelines," issued December 2019 [24]. On March 22, 2021, EPRI notified the NRC by letter [25] of a notification under Title 10 of the Code of Federal Regulations Part 21, Reporting of Defects and Noncompliance,

[26] sent to the industry about the potential nonconservatism in BWRVIP-100, Revision 1-A (EPRI, 2017). During a public meeting on May 27, 2021, Westinghouse and the Pressurized Water Reactor Owners Group further indicated that they were assessing if there is an impact on the flaw tolerance methodologies and/or fracture toughness values in WCAP-17096-NP-A, Rev. 2 and WCAP-17096-NP, Rev. 3. [28] Finally, in November 2021, the NRC released publicly an assessment [29], following the process in LIC-504, Revision 5, Integrated Risk-Informed Decisionmaking Process for Emergent Issues, issued March 2020 [30], of the nonconservatism issue in BWRVIP-100, Revision 1-A. The assessment concluded that while there was not an immediate safety issue, this issue should be monitored through the NRC inspection program to confirm industry addresses it appropriately.

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Figure 13 shows a plot of the Zorita weld and HAZ FT data as a function of fluence with the most relevant measured data from the literature, one NRC-approved industry flaw evaluation method (BWRVIP-100, Revision 1-A), and two other proposed lower bound predictions noted in the equations below [31] and the other a modified version of the original MRP-276 equation to fit the latest data). In Figure 12, the BWRVIP-100, Revision 1-A, criteria are clearly not bounding of the data from ANL [32] (i.e., from a cancelled BWR core shroud weld irradiated at Halden) or the Zorita weld data. The results from the Zorita core barrel base metal and HAZ are approximately bounded by BWRVIP-100, Revision 1-A, but the MRP-276 lower bound curve appears to be a better approach that could be modified to bound the measured data for all regions of the weld zone. The modified MRP-276 lower bound curve uses a similar formula, with updated coefficients to ensure it bounds the Zorita data. The original and modified MRP--276 equations are shown below:

Original: = 180 142 (1 )

Modified: = 130 92 (1 )

Figure 13: Summary plot of irradiated SS weld FT data as a function of fluence for Zorita and related materials The nuclear industry makes extensive use of the guidance in BWRVIP-100, Revision 1-A, to determine inspection intervals and evaluate the significance of any flaws detected during in-service inspections of the internals. These low FT data from the Zorita materials suggest the BWRVIP-100, Revision 1-A, and WCAP-17096 guidance are sufficient for FT of irradiated base metal and HAZ, but not for the weld metal. In particular, the fluence threshold for transition from elastic-plastic fracture mechanics to linear elastic fracture mechanics may need to be reduced to appropriately bound these new data. Fortunately, as shown in Appendix B, the weld and HAZ materials from Zorita have shown strong resistance to IASCC crack growth, so there should be 20

less concern with large flaws growing during service into the weld metal. However, these data suggest that embrittlement of the weld metal during service could be a concern for preexisting flaws (from fabrication, for instance) or slowly growing SCC flaws under accident loading conditions if toughness were to continue decreasing at higher fluences. The NRCs LIC 504 assessment [29] indicates the low weld fracture toughness is not an immediate safety issue for BWR core shrouds, based on the operating experience related to flaws found in service and conservative loading assumptions.

Historically, the HAZ has been the region where fatigue cracking and SCC have initiated because of the local stress concentration, high residual stresses, a sensitized microstructure, and lower flow strength than the weld metal [16,33]. In the scenario of a flaw outside of the weld in the HAZ or base metal, the better resistance of the weld to SCC and the higher strength of the weld may prevent the crack in the base metal/HAZ from growing into the weld. Since the service-induced flaws tend to form in the HAZ, most cracks will tend to take the path of least resistance, away from the higher strength, more SCC-resistant, low-toughness weld metal.

Under potential accident conditions in which the load is increased significantly, a flaw that was growing in the HAZ could move into the lower toughness weld metal, as demonstrated during toughness testing of HAZ material (see Figure 14), but only after significant plasticity at the crack tip.

To further explore these results, it would be beneficial to generate additional irradiated SS FT experimental data. The greatest priority would be to generate additional data at these fairly low fluence levels (<2 dpa) on harvested materials that have seen significant thermal aging as well as low-dose irradiation. Ideally, these could be compared to unirradiated thermally aged materials and as- welded properties to differentiate the effects of thermal aging and irradiation on embrittlement. Additional characterization of oxygen content and the original welding process would also be helpful in future work on harvested materials. Finally, higher fluence levels should be explored to understand FT and CGR behavior at higher dose levels that will be seen in both BWRs and PWRs during extended plant operation. The OECD/NEA SMILE project expects to test SS weld materials with up to 7 dpa of exposure from the Ringhals 2 plant, which could help address several of these data needs. Other domestic harvesting opportunities are also being pursued on higher dose SS weld materials.

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Figure 14: Side views of specimen W2HCT03 after testing, showing that the crack deviated from the plane defined by the side grooves during the FT test (Ref: Figure C-80 from MRP-451) 22

CHAPTER 4: CONCLUSIONS AND RECOMMENDATIONS The Zorita materials research programs have investigated highly representative high-fluence irradiated SS components harvested from a commercial PWR. The most significant results from the testing of Zorita baffle plate materials are the repeated observations of high CGR during IASCC CGR testing and the very low amount of observed void swelling. The most significant results from the testing of Zorita weld materials are the very low FT values observed in multiple tests.

The Zorita baffle plate void swelling data should be taken as encouraging in that void swelling may not progress as rapidly in LWRs as previously suggested. However, due to the lower operating temperatures of Zorita, the results cannot conclusively eliminate the potential for significant void swelling, particularly at higher doses and temperatures. The Zorita void swelling data show the strong influence of temperature, consistent with the results from other data in the literature. Industry and regulators should seek to observe additional LWR-irradiated materials at higher doses and temperatures near 360 degrees C to more confidently conclude that void swelling will not pose a significant issue during extended operating periods.

The Zorita baffle plate CGR data suggest that the IASCC CGR model for ASME Code Case N-889 does not sufficiently predict the increased IASCC CGRs at fluences above 20 dpa observed in this material. This deficiency supports the proposed NRC condition on this Code Case, limiting its applicability to materials less than 20 dpa. Given the small volume of LWR internals exceeding 20 dpa, the practical implications of this condition are likely to be limited in the near term. When assessing the significance of the high CGRs on Zorita materials at high fluence levels, it should also be recognized that these data come from one heat of material irradiated in one reactor. Heat-to-heat variability can lead to significant uncertainty in materials testing, so additional CGR testing of highly irradiated materials should be pursued where practical to augment the Zorita plate testing data and confirm or refute the observations from the Zorita CGR testing. The SMILE project will harvest and test Type 304 SS internals materials from the Ringhals 2 PWR with fluences up to 50 dpa. The SMILE data and other future harvesting efforts should help to provide additional understanding of the IASCC behavior of high-fluence PWR internals.

The Zorita core barrel weld FT data should be carefully considered when assessing irradiated SS weld embrittlement, particularly given the very limited amount of data from in-service welds.

The Zorita data should be used to update existing guidance on irradiated SS weld FT as contained in BWRVIP-100, Revision 1, and WCAP-17096. Given the low susceptibility to IASCC, low operating stresses, and flaw-tolerant design of the BWR core shroud and PWR core barrel, it is not expected that these lower weld FT data pose an immediate safety concern.

However, these results may necessitate reduced inspection intervals compared to previous guidance to ensure that an acceptable margin to structural integrity exists. Further research on irradiated SS weld materials should prioritize generating additional data at fairly low fluence levels (<2 dpa) and extend data on irradiated SS weld properties up to higher fluences approaching 20-30 dpa. Further harvesting efforts planned through the SMILE program and at a domestic plant to acquire PWR-irradiated SS weld and HAZ materials at a range of fluences are expected to help address these data needs.

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REFERENCES 5 4F4F 6F6F 1. Electric Power Research Institute, Materials Reliability Program: Zorita Internals Research Project (MRP-440), Testing of Highly-Irradiated Baffle Plate Material, Product ID 3002016015, Palo Alto, CA, October 29, 2019 (proprietary; available for viewing through the NRC Technical Library 6).

5F5F 7F7F 2. Electric Power Research Institute, Materials Reliability Program: Zorita Internals Research Project: Radiation and Temperature Analysis of Zorita Baffle Plate and Core Barrel Weld Material (MRP-392), Product ID 3002003084, Palo Alto, CA, 2015 (proprietary).

8F8F 3. Electric Power Research Institute, Materials Reliability Program: Fluence Effects on Stainless Steel Welds (MRP-451): Crack Growth Rate and Fracture Toughness Testing of Zorita Weld and HAZ Materials, Product ID 3002018250, Palo Alto, CA, July 30, 2020 (proprietary; available for viewing through the NRC Technical Library6).

9F9F 4. Karlsen, T.M., Final Report on the BWR Crack Growth Rate Investigation IFA-791 (HWR-1236), Halden Reactor Project, Halden, Norway, 2018 (proprietary).

10F10F 5. Karlsen, T.M., Interim Report on the PWR Crack Growth Rate Investigation IFA-817 (HWR-1320), Halden Reactor Project, Halden, Norway, 2021 (proprietary).

11F11F 6. Chen, Y., B. Alexandreanu, and K. Natesan, Crack Growth Rate and Fracture Toughness Tests on Irradiated Ex-Plant Materials, ANL-19/45, Argonne National Laboratory, Lemont, IL, July 2020. (ADAMS Accession No. ML20198M503).

12F12F 7. Chen, Y., W-Y. Chen, and B. Alexandreanu, Irradiated Microstructure of Zorita Materials, ANL-20/50, Argonne National Laboratory, Lemont, IL, August 2020. (ADAMS Accession No. ML20269A143).

13F13F 8. Kombaiah, B., C. Judge, J. Charboneau, S. Smith, L. Gimenes Rodrigues Albuquerque, and V. Montes de Oca Carioni, Chemical Compositional Analysis and Microstructural Characterization of Harvested Zorita Reactor Pressure Vessel (RPV) Internals, INL/EXT 62220, Idaho National Laboratory, Idaho Falls, ID, March 2021. (ADAMS Accession No. ML21124A112) 14F14F 9. Electric Power Research Institute, BWRVIP-294, Rev. 2: Fracture Toughness of Zorita RPV Core Internals Applicable to BWRs: Final Report 2019, Product ID 3002015929, Palo Alto, CA, October 2019 (proprietary).

5. Publicly available NRC published documents are available electronically through the NRC Library on the NRCs public Web site at http://www.nrc.gov/reading-rm/doc-collections/. The documents can also be viewed online or printed for a fee in the NRCs Public Document Room (PDR) at 11555 Rockville Pike, Rockville, MD; the mailing address is USNRC PDR, Washington, DC 20555; telephone (301) 415-4737 or (800) 397-4209; fax (301) 415-3548; and e-mail pdr.resource@nrc.gov.

6. The Technical Library, which is located at Two White Flint North, 11545 Rockville Pike, Rockville, Maryland 20852, is open by appointment only. Interested parties may make appointments to examine documents by contacting the NRC Technical Library by email: Library.Resource@nrc.gov between 8:00 a.m. and 4:00 p.m. (EST), Monday through Friday, except Federal holidays.

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15F15F 10. Electric Power Research Institute, BWRVIP-335: BWR Vessel and Internals Project, Crack Growth Rate Testing of Zorita Core Barrel Materials in BWR Environments, Product ID 3002017168, Palo Alto, CA, August 2020 (proprietary).

16F16F 11. Jenssen, A. J. Stjrnster, K. Kese, R. Carter, J. Smith, A. Demma, and M. Hiser, Fracture Toughness Testing of an Irradiated PWR Core Barrel Weld, Fontevraud 9Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 17-20, 2018.

17F17F 12. Eason, E., and R. Pathania, Irradiation-Assisted Stress Corrosion Crack Growth Rates of Austenitic Stainless Steels in Light Water Reactor Environments, 17th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Ottawa, Ontario, Canada, August 9-12, 2015.

18F18F 13. Eason, E. D., and Pathania, R. Disposition Curves for Irradiation-Assisted Stress Corrosion Cracking of Austenitic Stainless Steels in Light Water Reactor Environments, Proc. ASME 2015 Pressure Vessels & Piping Conference, Boston MA. Paper PVP2015-45323, July 19-23, 2015.

19F19F 14. Jenssen, A., J. Stjrnster, C. Topbasi, and P. Chou, Specimen Size Effects on the Crack Growth Rate Response of Highly Irradiated Type 304 Stainless Steel, 19th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Boston, MA, August 18-22, 2019.

20F20F 15. Electric Power Research Institute, BWRVIP-315: BWR Vessel and Internals Project, Reactor Internals Aging Management Evaluation for Extended Operations, Product ID 3002012535, Palo Alto, CA, July 2019 (proprietary).

21F21F 16. Chopra, O.K. Degradation of LWR Core Internal Materials due to Neutron Irradiation, NUREG/CR-7027, ANL-10/11, Argonne National Laboratory, Lemont, IL, December 2010.

(ADAMS Accession No. ML102790482) 22F22F 17. Miura, T., K. Fujii, K. Fukuya, and Y. Kitsunai, Micro-tensile testing for grain boundary fracture in neutron-irradiated stainless steels, Materials in Nuclear Energy Systems, American Nuclear Society, Baltimore, MD, October 6-10, 2019.

23F23F 18. U.S. Nuclear Regulatory Commission. Draft Regulatory Guide DG-1367 Inservice Inspection Code Case Acceptability, ASME Section XI, Division 1, January 2021. (ADAMS Accession No. ML20120A631).

24F24F 19. Garner, F.A. Chapter 10, "Void swelling and irradiation creep in light water reactor (LWR) environments", in Understanding and Mitigating Ageing in Nuclear Power Plants, Ed. P. G.

Tipping, Woodhouse Publishing, 2010, pp. 308-356.

25F25F 20. Garner, F., New data and insights on prediction of void swelling in austenitic pressure vessel internals, Fontevraud 9Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 17-20, 2018.

26F26F 21. Palm, N., B. Carter. Potential Non-Conservatism in EPRI Report, BWRVIP-100, Rev. 1-A, and EPRI Software, BWRVIP-235 (Closed Session), presentation at NRC public meeting on May 27, 2021. (ADAMS Accession No. ML21147A009).

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27F27F 22. Electric Power Research Institute, BWRVIP-100NP, Revision 1-A: BWR Vessel and Internals Project, Updated Assessment of the Fracture Toughness of Irradiated Stainless Steel for BWR Core Shrouds, Product ID 3002008388NP, Palo Alto, CA, February 2017. (ADAMS Accession No. ML17076A228) 28F28F 23. Carpenter, B.T., R.G. Lott, S. Fyfitch, and A. Kulp, Reactor Internals Acceptance Criteria Methodology and Data Requirements, WCAP-17096-NP-A, Revision 2, Westinghouse, Cranberry Township, PA, August 2016. (ADAMS Accession No. ML16279A320) 29F29F 24. Electric Power Research Institute, MRP-227, Revision 1-A: Materials Reliability Program:

Pressurized Water Reactor Internals Inspection and Evaluation Guidelines, Product ID 3002018304, Palo Alto, CA, December 2019 (ADAMS Accession No. ML19339G350).

30F30F 25. Palm, N., Hanley, T. BWRVIP-2021-030: Potential Non-Conservatism in EPRI Report, BWRVIP-100, Rev. 1-A, 3002008388 and Impacted BWRVIP Reports, letter dated March 22, 2021. (ADAMS Accession No. ML21084A164) 31F31F 26. U.S. Nuclear Regulatory Commission. Title 10 of the Code of Federal Regulations Part 21, Reporting of Defects and Noncompliance, 42 FR 28893, June 6, 1977, unless otherwise noted

[77 FR 39905, Jul. 6, 2012; 80 FR 54233, Sep. 9, 2015]. (https://www.nrc.gov/reading-rm/doc-collections/cfr/part021/index.html) 32F32F 27. Radonovich, D.C., E.W. Deemer, R. Hosler, S. Davidsaver, G. Troyer and S. Fyfitch, Reactor Internals Acceptance Criteria Methodology and Data Requirements, WCAP-17096-NP, Revision 3, Westinghouse, Cranberry Township, PA, July 2019. (ADAMS Accession No. ML19218A179) 33F33F 28. D. Radonovich, J.B. Hall. PWROG Evaluation of the Potential Non- Conservative Fracture Toughness in BWRVIP-100, Rev. 1-A, presentation at NRC public meeting on May 27, 2021.

(ADAMS Accession No. ML21147A012).

34F34F 29. Buford, A., Technical Assessment of Nonconservative Fracture Toughness in Boiling Water Reactor Vessel and Internals Project Topical Report, BWRVIP-100, Revision 1-A, November 17, 2021. (ADAMS Accession No. ML21312A543 (package))

35F35F 30. U.S. Nuclear Regulatory Commission. LIC-504, Revision 5 Integrated Risk-Informed Decisionmaking Process for Emergent Issues, March 9, 2020. (ADAMS Accession No. ML19253D401) 36F36F 31. Electric Power Research Institute, Materials Reliability Program: Thermal Aging and Neutron Embrittlement Assessment of Cast Austenitic Stainless Steels and Stainless Steel Welds in PWR Internals (MRP-276), Product ID 1020959, Palo Alto, CA, May 12, 2010.

37F37F 32. Chen,Y., B. Alexandreanu, C. Xu, Y. Yang, K. Natesan, and A. S. Rao, Environmentally Assisted Cracking and Fracture Toughness of an Irradiated Stainless Steel Weld, 19th International Conference on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Boston, MA, August 18-22, 2019.

38F38F 33. Andresen, P.L. Chapter 9, "Stress corrosion cracking (SCC) of austenitic stainless steels in high temperature light water reactor (LWR) environments", in Understanding and Mitigating Ageing in Nuclear Power Plants, Ed. P. G. Tipping, Woodhouse Publishing, 2010, pp. 236-307.

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APPENDIX A: ADDITIONAL ZORITA PLATE TESTING RESULTS A.1 Overview of Zorita Plate Testing Zorita baffle plate testing and characterization was performed at Studsvik, Halden, Argonne National Laboratory, and Idaho National Laboratory under various programs covering tensile, irradiation-assisted stress corrosion cracking (IASCC) crack initiation, IASCC crack growth rate (CGR), fracture toughness (FT), and TEM to assess irradiation damage and void swelling.

The Zorita plate CGR and void swelling results were the most noteworthy and are discussed in detail in Section 3 of the main body of this report, while this appendix covers other results obtained during the test program, including tensile, crack initiation, and FT testing.

A.2 Plate Tensile Testing Figure A-1, from Materials Reliability Program (MRP)-440 [1], summarizes the tensile test results compared to those from other data in the literature. The data are consistent with the general trend for irradiated stainless steels (SSs) and do not offer any significantly new insights into irradiated SS behavior.

A.3 Plate Crack Initiation Testing Figure A-2, from MRP-440, summarizes the IASCC initiation test results on O-ring and uniaxial constant load specimens compared to those from other data in the literature. The data are in line with the general trend for irradiated SSs and tend to confirm the Electric Power Research Institute MRP trendline for IASCC initiation as a function of dose and applied stress.

One interesting aspect of the IASCC initiation data from the Zorita materials was that no O-ring specimens failed (note open blue squares in Figure A-2). This is different from literature data on O-ring samples. However, it should be noted that much of the literature data on O-ring samples used flux thimble tubes with the original surface (including oxide layer) intact, while the Zorita O-ring specimens were machined from a plate and polished. The differing product form and surface condition of the tested Zorita materials may explain the differing behavior compared to O-ring literature data.

A-1

Zorita data Figure A-1: Yield stress (top) and total elongation (bottom) as a function of dose (Ref:

Figures 8-3 and 8-4 from MRP-440)4 A-2

Zorita data Figure A-2: Summary of IASCC crack initiation results on Zorita plate materials compared to literature data (Ref: Figure 8-6 from MRP-440)4 A.4 Plate Fracture Toughness Testing Figure A-3, from MRP-440, summarizes the FT test results compared to other literature data.

The data are consistent with the general trend for irradiated SSs, with the principal new insight being that FT decreases largely saturate at doses above 10 displacements per atom (dpa). This is an important finding as plants age and licensees may find flaws in components at higher doses.

A-3

Zorita data Figure A-3: Summary of FT results on Zorita plate materials compared to literature data (Ref: Figure 8-24 from MRP-440)4 A-4

APPENDIX B: ADDITIONAL ZORITA WELD AND HEAT-AFFECTED ZONE TESTING RESULTS B.1 Overview of Zorita Weld and Heat-Affected Zone Testing The goal of the testing program was to evaluate the effects of irradiation on stainless steel (SS) welds. To this end, tensile testing, microstructural characterization by light optical microscopy, fracture toughness (FT) testing, and stress corrosion crack growth rate (CGR) testing were performed on SS weld and heat-affected zone (HAZ) specimens machined from the Zorita core barrel at doses in the range of 0.7-1.9 displacements per atom (dpa). In addition to testing in air, some tests were conducted in simulated pressurized-water reactor (PWR) primary water conditions and PWR shutdown chemistry conditions at 170 degrees Celsius (C).

The Zorita weld and HAZ FT results were the most noteworthy and are discussed in detail in Section 4 of the main body of this report, while this appendix covers other results obtained during the test program, including tensile and CGR testing.

B.2 Weld and Heat-Affected Zone Tensile Testing Figure B-1 from Materials Reliability Program (MRP)-451 [3] summarizes the tensile test results, in conjunction with literature data on irradiated SS weld/HAZ materials.

For both weld and HAZ materials, the yield stress generally follows the trend of the few other data points available in the literature. However, for ductility, both weld and HAZ materials from Zorita show significantly lower ductility compared to materials at similar doses described in the literature. This interesting observation appears to be consistent with the very low FT also observed in these materials, which is discussed in greater detail in Section 4 of the main body of this report.

B-1

Zorita data Weld YS HAZ YS Zorita data Weld ductility Weld ductility HAZ ductility Zorita data Zorita data Zorita data Figure B-1: Summary of tensile results: weld yield stress (top left), HAZ yield stress (top right), weld ductility (bottom left), HAZ ductility (bottom right) compared to literature data (Ref.: Figures 7-3 through 7-6 from MRP-451)4 B-2

B.3 Weld/Heat-Affected Zone Crack Growth Rate Testing A fairly extensive CGR test matrix was performed on the Zorita weld and HAZ materials as shown below in Table B-1 (Ref: Table 5-1 from MRP-451).

Table B-1: Test Matrix for Crack Growth Rate Testing in PWR Primary Water Specimen ID Material Dose, dpa KI, Test temperature, °C Environment W2WCT03 Weld 1.53 ~30 290, 320, 340 W2WCT06 Weld 1.53 ~15 and 25 290, 320, 340 PWR primary W2HCT03 HAZ 1.48 ~25 290, 320, 340 water*

W2HCT06 HAZ 1.48 ~30 290, 320, 340 W2WCT05 Weld 1.53 ~30 170 Shutdown W2WCT04 Weld 1.53 ~30 170 chemistry*

• PWR primary water simulated by the addition of 1000 ppm B, 2 ppm Li and 30 cm3/kg of hydrogen at the temperatures indicated, and shutdown chemistry simulated by the addition of 2000 ppm B and 15 cm3/kg of hydrogen to water at 170 C Figure B-2, from MRP-451, shows the results from the CGR testing, with temperature correction and comparison to the irradiation-assisted stress corrosion cracking (IASCC) CGR model from American Society of Mechanical Engineers (ASME) Code Case N-889. The key takeaway from the weld/HAZ CGR testing is that all the weld and HAZ specimens tested in this study were very resistant to IASCC in the environments investigated, as it was very difficult to obtain sustained crack growth under constant stress intensity factor (K) conditions. All data fall well below the ASME Code Case N-889 IASCC CGR model curves, even though most of the data were obtained under loading conditions that involved partial periodic unloading and a hold time at maximum load.

B-3

Figure B-2: Summary of constant K and partial periodic unloading (PPU) IASCC CGR data in PWR conditions normalized to 320 degrees C compared to the Eason/Electric Power Research Institute IASCC CGR model curve from ASME Code Case N-889 for this temperature/chemistry (Ref: Figure 7-8 from MRP-451)

B-4

APPENDIX C: ADDITIONAL ZORITA PLATE CGR TESTING TABULAR DATA CGR Data from Studsvik [1]

Table C-1 Crack growth rate data for Specimen B1CT01 (41.5 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 1 0.2 1 320 1.78*10-5 13.0 13.1 27.5 1.089 24 39 50 2 0.6 1 320 8.49*10-6 16.1 16.1 2.5 0.079 30 49 62 3 0.6 0.1 320 8.30*10-7 16.1 16.1 35.2 0.102 30 49 62 4 0.6 0.01 320 1.62*10-7 16.1 16.2 265.5 0.153 30 49 62 5a 0.6 0.001 320 3.52*10-8 16.2 16.3 256.8 0.025 30 49 62 5b 0.6 0.001 320 3.15*10-8 16.3 16.3 152.4 0.017 30 49 62 6a 0.6 100 s hold 320 2.63*10-8 16.3 16.3 25.5 0.007 30 49 62 6b 1 Const. P 320 1.04*10-8 16.3 16.3 76.9 0.003 30 49 62 6c 0.6 100 s hold 320 2.93*10-8 16.3 16.3 238.5 0.022 30 49 62 6d 0.6 1 ks hold 320 2.12*10-8 16.3 16.3 163.4 0.013 30 49 62 6e 0.6 9 ks hold 320 1.41*10-8 16.3 16.3 144.5 0.007 30 49 62 7 1 Const. K 320 2.56*10-8 16.3 16.3 213.5 0.020 30 49 62 8 1 Const. K 340 5.25*10-8 16.3 16.4 262.5 0.048 30 49 62 9 1 Const. K 290 2.37*10-8 16.4 16.4 336.0 0.032 31 50 63 10a 1 Const. K 320 1.02*10-8 16.4 16.5 194.3 0.013 31 50 63 10b 0.6 9 ks hold 320 9.86*10-8 16.5 16.5 44.4 0.016 31 50 63 10c 1 Const. K 320 1.18*10-8 16.5 16.5 265.9 0.016 31 50 63 11 0.2 1 320 9.26*10-5 21.7 22.1 2.5 0.889 40 66 83 12 0.6 1 320 2.33*10-5 26.2 26.2 0.8 0.069 49 79 100 C-1

Table C-1 (continued) Crack growth rate data for Specimen B1CT01 (41.5 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 13 0.6 0.1 320 3.08*10-6 26.2 26.2 9.2 0.101 49 79 100 14 0.6 0.01 320 2.60*10-7 26.2 26.2 113.6 0.108 49 79 100 15 0.6 0.001 320 3.38*10-8 26.2 26.3 185.1 0.026 49 79 100 16a 0.6 100 s hold 320 3.95*10-8 26.3 26.3 170.9 0.022 49 79 100 16b 0.6 1 ks hold 320 2.81*10-8 26.3 26.3 164.2 0.020 49 79 100 16c 0.6 9 ks hold 320 1.49*10-8 26.3 26.3 197.3 0.012 49 80 100 17a 1 Const. K 320 1.07*10-6 26.3 26.5 29.8 0.114 49 80 101 17a1 1 Const. K 320 3.33*10-7 26.3 26.3 14.5 0.015 49 80 101 17a2 1 Const. K 320 1.67*10-6 26.3 26.5 15.3 0.099 49 80 101 17b 1 Const. K 320 1.17*10-5 26.5 27.3 10.8 0.468 49 80 101 17b1 1 Const. K 320 7.56*10-6 26.5 26.8 6.3 0.173 49 80 101 17b2 1 Const. K 320 1.90*10-5 26.8 27.3 4.5 0.295 50 81 103 18 1 dK/da=- 320 1.46*10-5 27.3 22.9 12.0 0.603 52 84 106 16 19 1 dK/da=- 320 1.10*10-5 22.9 21.1 6.7 0.258 45 72 91 16 20 1 Const. K 320 2.97*10-5 21.1 23.1 9.5 0.963 42 68 85 21a 1 dK/da=- 320 2.99*10-5 23.1 21.3 2.4 0.259 48 77 98 16 21b 1 dK/da=- 320 1.27*10-5 20.3 18.9 3.9 0.179 43 70 88 16 7 21c 1 dK/da=- 320 4.54*10-6 18.7 17.9 5.9 0.098 40 65 82 16 21d 1 dK/da=- 320 2.26*10-8 17.7 17.7 95.1 0.007 38 62 78 22 1 Const. K 320 1.73*10-8 17.7 17.7 120.3 0.007 38 62 78 23a7) 0.6 dK/da=+4.7 320 4.59*10-6 17.7 20.4 16.5 0.326 38 62 78 C-2

Table C-1 (continued) Crack growth rate data for Specimen B1CT01 (41.5 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 23b7) 0.6 dK/da=+4.7 320 4.31*10-4 20.4 25.1 0.5 0.714 45 73 92 24a 1 dK/da=+5 320 1.14*10-6 25.1 26.9 36.8 0.256 58 94 119 24b 1 dK/da=+5 320 1.41*10-6 26.9 30.8 32.7 0.546 63 103 130 25 1 Const. K 320 1.61*10-7 30.8 31.3 245.5 0.154 75 122 154 26 1 Const. K 340 2.39*10-7 31.3 32.5 185.9 0.220 77 126 158 27 1 Const. K 290 2.55*10-5 32.5 137.4 22.5 4.166 82 132 167 28 1 Const. P 290 5.19*10-6 3.6 3.6 24.5 0.371 15 24 30

1) Loading condition during testing. A number stands for the frequency in Hz during continuous cyclic load, whereas a number followed by a time and the word hold means the specimen was subjected to partial periodic unloading (PPU) with the hold time at maximum load indicated. Const. K or P means testing was conducted under constant stress intensity factor (K), or load (P), while dK/da=X means K was changed by the nominal rate (in MPam/mm) indicated.

2) Test temperature during the step.

3) Crack growth rate determined by linear regression of the corrected DCPD curve.

4) Stress intensity factor based on the crack length corresponding to the first (KIstart) and last (KIstop) data point of the regression analysis and load data.

5) Difference in test time or crack length between the first and last data point of the regression analysis.

6) K validity according to ASTM E399 and modified criteria based on the same standard. An effective yield strength was used for the modified criteria, i.e., YSeff1=YSunirr + (YSirr-YSunirr)/2 and YSeff2=YSunirr + (YSirr-YSunirr)/3. The actual K during the step was related to the K limit based on the crack length when the step started, i.e., 100% or higher means the criterion was exceeded.

7) Positive dK/da as indicated plus continuous cycling at frequency and R ratio reported.

C-3

Table C-2 Crack growth rate data for Specimen B1CT03 (41.5 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 1 0.2 1 320 3.29*10-5 16.0 16.1 10.8 1.077 30 49 61 2 0.6 1 320 1.48*10-5 20.1 20.1 4.2 0.228 38 61 77 3 0.6 0.1 320 1.93*10-6 20.1 20.1 21.2 0.148 37 61 77 4 0.6 0.01 320 2.16*10-7 20.1 20.1 148.3 0.118 37 61 77 5 0.6 0.001 320 4.13*10-8 20.1 20.1 485.2 0.074 38 61 77 6a 0.6 100 s hold 320 3.97*10-8 20.1 20.2 365.7 0.048 38 61 77 6b 0.6 1 ks hold 320 3.32*10-8 20.2 20.2 353.3 0.045 38 61 77 6c 0.6 9 ks hold 320 2.72*10-8 20.2 20.2 258.5 0.027 38 61 77 6d 0.6 9 ks hold 320 3.11*10-8 20.2 20.3 461.2 0.057 38 61 77 7a 1 Const. K 320 1.09*10-8 20.3 20.1 135.5 0.005 38 61 77 7b 1 Const. K 320 7.22*10-7 20.1 20.1 2.3 0.006 37 61 77 7c 1 Const. K 320 7.61*10-9 20.1 20.3 101.3 0.004 37 61 77 7d 0.6 9 ks hold 320 6.09*10-8 20.3 20.3 67.7 0.014 38 61 78 7e 1 Const. K 320 1.30*10-8 20.3 20.3 197.1 0.007 38 61 78 8 1 Const. K 340 3.17*10-8 20.3 20.4 333.8 0.036 38 62 78 9a 1 Const. K 290 2.92*10-9 20.4 20.4 19.9 0.001 38 62 78 9b 1 Const. K 290 2.60*10-5 23.6 27.3 17.5 1.547 44 72 90 9c 1 Const. P 290 8.91*10-7 14.1 14.2 21.6 0.069 27 44 55 9d 1 Const. P 290 <1*10-9 4.9 1.9 260.6 0.000 9 15 19 10 1 Const. K 290 2.05*10-5 22.7 27.3 28.3 1.978 43 70 89 11 1 dK/da=-16.7 290 1.02*10-5 27.4 22.2 18.7 0.691 57 93 118 12 1 Const. K 290 3.05*10-5 22.2 26.3 10.5 1.212 48 79 99 C-4

Table C-2 (continued) Crack growth rate data for Specimen B1CT03 (41.5 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 13 1 dK/da=- 290 7.87*10-6 26.3 23.1 10.0 0.271 62 100 126 14 1 Const. K 290 5.11*10-5 23.1 31.0 7.2 1.148 56 92 116 15a 1 dK/da=- 290 4.74*10-6 31.0 27.7 15.5 0.269 83 135 170 15b 1 dK/da=- 290 2.48*10-8 27.3 27.4 46.9 0.005 75 122 154 16 16 1 Const. K 290 1.80*10-8 27.4 27.2 341.4 0.021 75 122 154 17 0.6 9 ks hold 290 3.91*10-6 27.2 30.8 30.4 0.509 75 122 154 18 1 Const. K 290 1.81*10-8 30.8 30.8 133.0 0.012 89 145 183 19 1 Const. K 340 1.63*10-8 30.8 31.0 379.1 0.022 89 145 183 20 1 Const. K 320 2.91*10-8 31.0 31.0 12.4 0.002 90 146 184 21 1 Const. P 320 <1*10-9 4.6 4.6 72.9 0.000 13 22 28 22 0.6 9 ks hold 320 3.63*10-6 31.8 34.4 25.0 0.374 93 150 190 23 1 Const. K 320 1.64*10-8 34.4 34.7 435.0 0.028 104 169 214 24 1 Const. K 340 3.28*10-8 34.7 35.5 582.6 0.065 105 171 216 25 1 Const. K 340 1.90*10-5 35.5 60.7 23.3 1.494 108 176 222 26 1 Const. K 340 8.35*10-7 60.7 81.8 178.5 0.576 228 370 467

1) Loading condition during testing. A number stands for the frequency in Hz during continuous cyclic load, whereas a number followed by a time and the word hold means the specimen was subjected to partial periodic unloading (PPU) with the hold time at maximum load indicated. Const. K or P means testing was conducted under constant stress intensity factor (K), or load (P), while dK/da=X means K was changed by the nominal rate (in MPam/mm) indicated.

2) Test temperature during the step.

3) Crack growth rate determined by linear regression of the corrected DCPD curve.

4) Stress intensity factor based on the crack length corresponding to the first (KIstart) and last (KIstop) data point of the regression analysis and load data.

5) Difference in test time or crack length between the first and last data point of the regression analysis.

6) K validity according to ASTM E399 and modified criteria based on the same standard. An effective yield strength was used for the modified criteria, i.e., YSeff1=YSunirr + (YSirr-YSunirr)/2 and YSeff2=YSunirr + (YSirr-YSunirr)/3. The actual K during the step was related to the K limit based on the crack length when the step started, i.e., 100% or higher means the criterion was exceeded.

C-5

Table C-3 Crack growth rate data for Specimen B2CT01 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 1a 0.2 0.9 320 1.08*10-5 12.9 12.9 22.8 0.849 24 39 49 1b 0.2 0.9 320 9.00*10-6 12.9 12.9 4.6 0.164 24 39 49 2 0.6 0.9 320 7.38*10-6 16.0 16.0 3.3 0.087 30 49 61 3 0.6 0.1 320 9.81*10-7 16.0 16.0 28.7 0.102 30 49 61 4 0.6 0.01 320 1.56*10-7 16.0 16.0 176.3 0.099 30 49 61 5 0.6 0.001 320 3.40*10-8 16.0 16.0 320.8 0.043 30 49 61 6a 0.6 100 s hold 320 3.70*10-8 16.0 16.0 264.2 0.036 30 48 61 6b 0.6 1 ks hold 320 3.10*10-8 16.0 16.0 143.8 0.013 30 48 61 6c 0.6 9 ks hold 320 2.27*10-8 16.0 16.1 195.0 0.018 30 49 61 7 1 Const. K 320 4.98*10-8 16.1 16.1 324.4 0.050 30 49 61 8 1 Const. K 340 6.20*10-8 16.1 16.1 503.3 0.119 30 49 61 9 1 Const. K 290 4.60*10-8 16.1 16.2 837.8 0.134 30 49 62 9a 1 Const. K 290 4.61*10-8 16.1 16.1 573.5 0.088 30 49 62 9b 0.6 9 ks hold 290 2.56*10-7 16.2 16.2 24.7 0.028 30 49 62 9c 1 Const. K 290 1.10*10-8 16.2 16.2 239.2 0.020 30 49 62 10 1 Const. K 320 2.08*10-8 16.2 16.2 189.9 0.007 30 49 62 11 0.2 0.9 320 1.12*10-4 21.2 21.4 1.5 0.722 40 64 81 12 0.6 0.9 320 1.84*10-5 25.8 25.8 0.8 0.058 48 78 99 13 0.6 0.1 320 2.12*10-6 25.8 25.8 11.5 0.089 48 78 99 14 0.6 0.01 320 2.38*10-7 25.8 25.8 34.6 0.031 48 78 99 C-6

Table C-3 (continued) Crack growth rate data for Specimen B2CT01 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 15 0.6 0.001 320 5.50*10-8 25.8 25.8 138.9 0.019 48 78 98 16 0.6 9 ks hold 320 1.40*10-8 25.8 25.8 150.0 0.016 48 78 99 17 0.6 0.001 320 3.37*10-8 25.8 25.8 89.9 0.012 48 78 99 18 0.6 0.01 320 2.82*10-7 25.8 25.8 48.1 0.047 48 78 98 19 0.6 0.001 320 4.66*10-8 25.8 25.8 340.0 0.055 48 78 99 20 0.6 100 s hold 320 6.41*10-8 25.8 25.8 168.8 0.039 48 78 99 21 0.6 1 ks hold 320 3.74*10-8 25.8 25.9 170.9 0.029 48 78 99 22 0.6 9 ks hold 320 2.37*10-8 25.9 25.9 359.1 0.031 48 78 99 23 1 Const. K 320 <1*10-9 25.9 25.9 192.6 0.000 48 78 99 24 0.4 9 ks hold 320 5.93*10-8 25.9 25.9 381.5 0.083 48 78 99 25 0.5 9 ks hold 320 1.99*10-8 25.9 26.0 336.2 0.036 48 78 99 26 0.55 9 ks hold 320 <1*10-9 26.0 26.0 334.4 0.008 48 79 99 27 0.40 1 ks hold 320 5.73*10-7 26.0 26.2 86.0 0.167 49 79 99 28 0.50 1 ks hold 320 1.84*10-7 26.2 26.3 145.1 0.102 49 79 100 29 0.60 1 ks hold 320 5.99*10-8 26.3 26.3 148.3 0.023 49 80 101 30 0.60 9 ks hold 320 5.08*10-8 26.3 26.3 94.4 0.016 49 80 101 31 1 Const. K 320 <1*10-9 26.3 26.4 265.2 0.000 49 80 101 32 0.6 9 ks hold 320 <1*10-9 26.4 26.3 310.7 0.008 49 80 101 33 0.4 0.001 320 1.05*10-6 26.3 26.5 38.1 0.155 49 80 101 34 0.5 0.001 320 4.41*10-7 26.5 26.6 33.8 0.088 50 81 102 C-7

Table C-3 (continued) Crack growth rate data for Specimen B2CT01 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 35 0.6 0.001 320 2.52*10-8 26.6 26.7 124.2 0.006 50 82 103 36 0.5 0.001 320 1.65*10-7 26.7 26.8 116.4 0.091 50 82 103 37 0.55 0.001 320 1.46E-7 26.8 26.9 172.2 0.083 51 82 104 38 0.6 0.001 320 <1*10-9 26.9 26.9 186.9 0.000 51 83 105 39 0.4 0.001 320 2.10*10-7 26.9 27.1 158.0 0.129 51 83 105 40 0.5 0.001 320 1.13*10-7 27.1 27.4 358.3 0.139 52 84 106 41 0.5 100 s hold 320 4.47*10-8 27.4 27.4 200.3 0.042 53 86 108 42 0.4 0.001 320 4.86*10-8 27.4 27.4 93.6 0.005 53 86 108 43 0.4 0.1 320 2.11*10-5 27.4 27.9 4.7 0.347 53 86 109 44 0.4 0.01 320 1.32*10-6 27.9 28.0 14.5 0.070 55 89 112 45 0.4 0.001 320 1.91*10-8 28.0 28.1 106.4 0.013 55 90 113 46 0.4 0.01 320 9.81*10-7 28.1 28.4 49.8 0.173 55 90 113 47 0.4 0.003 320 2.40*10-7 28.4 28.5 66.1 0.048 56 92 116 48 0.4 0.001 320 7.46*10-8 28.5 28.6 97.4 0.030 57 92 116 49 0.4 100 s hold 320 1.02*10-7 28.6 28.6 121.9 0.046 57 92 117 50 0.4 1 ks hold 320 2.83*10-8 28.6 28.6 211.8 0.027 57 93 117 51 0.4 9 ks hold 320 <1*10-9 28.6 28.7 99.9 0.000 57 93 117 52 0.3 9 ks hold 320 3.85*10-8 28.7 28.8 147.2 0.018 57 93 118 53 0.3 27 ks hold 320 1.28*10-8 28.8 28.8 195.3 0.008 58 93 118 54a 1 Const. K 320 6.86*10-8 28.8 28.8 82.6 0.018 58 94 118 C-8

Table C-3 (continued) Crack growth rate data for Specimen B2CT01 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 54b1 1 Const. K 320 - 28.8 28.8 17.5 0.000 58 94 118 54b2 1 Const. K 320 2.87*10-6 28.8 29.3 10.0 0.105 58 94 118 54b3 1 Const. K 320 1.59*10-5 29.3 29.9 5.0 0.275 59 96 121 54b4 1 Const. K 320 4.34*10-5 29.9 33.4 7.0 1.082 61 99 125 54c 1 Const. P 320 5.11*10-8 0.7 1.6 172.0 0.039 1 2 3

1) Loading condition during testing. A number stands for the frequency in Hz during continuous cyclic load, whereas a number followed by a time and the word hold means the specimen was subjected to partial periodic unloading (PPU) with the hold time at maximum load indicated. Const. K or P means testing was conducted under constant stress intensity factor (K), or load (P), while dK/da=X means K was changed by the nominal rate (in MPam/mm) indicated.

2) Test temperature during the step.

3) Crack growth rate determined by linear regression of the corrected DCPD curve.

4) Stress intensity factor based on the crack length corresponding to the first (KIstart) and last (KIstop) data point of the regression analysis and load data.

5) Difference in test time or crack length between the first and last data point of the regression analysis.

6) K validity according to ASTM E399 and modified criteria based on the same standard. An effective yield strength was used for the modified criteria, i.e.,

YSeff1=YSunirr + (YSirr-YSunirr)/2 and YSeff2=YSunirr + (YSirr-YSunirr)/3. The actual K during the step was related to the K limit based on the crack length when the step started, i.e., 100% or higher means the criterion was exceeded.

C-9

Table C-4 Crack growth rate data for Specimen B2CT03 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 1 0.2 1 320 3.08*10-5 16.1 16.1 11.5 1.058 30 49 62 2 0.6 1 320 9.83*10-6 20.1 20.1 6.0 0.221 38 61 77 3 0.6 0.1 320 1.38*10-6 20.1 20.0 31.0 0.149 37 61 77 4 0.6 0.01 320 2.05*10-7 20.0 20.1 158.5 0.117 37 61 77 5 0.6 0.001 320 4.93*10-8 20.1 20.1 273.9 0.046 38 61 77 6a 0.6 100 s hold 320 4.40*10-8 20.1 20.1 260.6 0.039 38 61 77 6b 0.6 1 ks hold 320 4.21*10-8 20.1 20.1 240.8 0.055 38 61 77 6c 0.6 9 ks hold 320 2.78*10-8 20.1 20.1 191.8 0.019 38 61 77 7 1 Const. K 320 5.18*10-8 20.1 20.2 481.0 0.072 38 61 77 7a 1 Const. K 320 1.36*10-8 20.1 20.1 165.6 0.008 38 61 77 7b 1 Const. K 320 4.02*10-7 20.1 20.1 4.2 0.004 38 61 77 7c 1 Const. K 320 1.30*10-8 20.1 20.2 71.0 0.004 38 61 77 7d 1 Const. K 320 7.02*10-7 20.2 20.2 12.5 0.035 38 61 77 7e 1 Const. K 320 1.75*10-8 20.2 20.2 221.8 0.016 38 61 77 8 1 Const. K 340 4.53*10-8 20.2 20.2 352.6 0.053 38 61 77 9a 1 Const. K 290 5.21*10-9 20.2 20.3 102.2 0.008 38 61 78 9b 1 Const. K 290 2.19*10-5 20.4 23.6 27.8 2.081 38 62 78 10a 1 Const. K 320 4.45*10-5 23.6 37.6 23.9 3.799 46 75 95 10a1 1 Const. K 320 3.01*10-5 23.6 25.9 8.9 0.992 46 75 95 10b 1 dK/da=-16.7 320 2.79*10-5 37.6 35.6 2.7 0.298 91 148 187 C-10

Table C-4 (continued) Crack growth rate data for Specimen B2CT03 (23.9 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 10c 1 Const. K 320 7.56*10-5 35.6 40.8 2.0 0.536 88 143 181 10d 1 dK/da=-16.7 320 1.94*10-5 40.8 37.2 4.5 0.301 105 171 216 10e 1 Const. K 320 6.27*10-5 37.2 42.2 2.3 0.536 99 160 202 10f 1 dK/da=-16.7 320 1.48*10-5 42.2 39.5 5.4 0.309 117 190 240 10g 1 Const. K 320 6.53*10-5 39.5 46.0 2.3 0.561 113 183 231 10h 1 dK/da=-16.7 320 5.23*10-6 46.0 43.0 14.5 0.293 140 226 286 10k 1 Const. K 320 2.46*10-7 43.0 43.1 11.9 0.012 135 219 276

1) Loading condition during testing. A number stands for the frequency in Hz during continuous cyclic load, whereas a number followed by a time and the word hold means the specimen was subjected to partial periodic unloading (PPU) with the hold time at maximum load indicated. Const. K or P means testing was conducted under constant stress intensity factor (K), or load (P), while dK/da=X means K was changed by the nominal rate (in MPam/mm) indicated.

2) Test temperature during the step.

3) Crack growth rate determined by linear regression of the corrected DCPD curve.

4) Stress intensity factor based on the crack length corresponding to the first (KIstart) and last (KIstop) data point of the regression analysis and load data.

5) Difference in test time or crack length between the first and last data point of the regression analysis.

6) K validity according to ASTM E399 and modified criteria based on the same standard. An effective yield strength was used for the modified criteria, i.e.,

YSeff1=YSunirr + (YSirr-YSunirr)/2 and YSeff2=YSunirr + (YSirr-YSunirr)/3. The actual K during the step was related to the K limit based on the crack length when the step started, i.e., 100% or higher means the criterion was exceeded.

C-11

Table C-5 Crack growth rate data for Specimen B3CT03 (9.4 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 1 0.2 1 320 1.29*10-5 13.0 13.1 26.3 1.075 24 40 50 2 0.4 1 320 2.63*10-5 16.2 16.1 0.9 0.084 30 49 62 3 0.4 0.3 320 6.52*10-6 16.1 16.1 2.1 0.050 30 49 62 4 0.4 0.1 320 2.27*10-6 16.1 16.1 6.5 0.051 30 49 62 5 0.4 0.03 320 7.94*10-7 16.1 16.1 17.3 0.050 30 49 62 6 0.4 0.01 320 3.31*10-7 16.1 16.1 51.2 0.059 30 49 62 7 0.4 0.003 320 1.14*10-7 16.1 16.1 138.7 0.061 30 49 62 8 0.4 0.001 320 5.17*10-8 16.1 16.1 303.3 0.052 30 49 62 9 0.5 0.001 320 4.45*10-8 16.1 16.2 362.4 0.061 30 49 62 10 0.6 0.001 320 2.66*10-8 16.2 16.2 359.9 0.036 30 49 62 11a 0.6 100 s hold 320 2.14*10-8 16.2 16.2 137.3 0.012 30 49 62 11b 0.6 100 s hold 320 2.42*10-8 16.2 16.2 185.0 0.016 30 49 62 12 0.6 1 ks hold 320 1.56*10-8 16.2 16.2 215.6 0.013 30 49 62 13 0.6 9 ks hold 320 7.17*10-9 16.2 16.2 289.0 0.004 30 49 62 14 1 Const. K 320 2.75*10-9 16.2 16.2 527.4 0.007 30 49 62 15 1 Const. K 340 1.50*10-8 16.2 16.2 358.5 0.016 30 49 62 16a 1 Const. K 290 6.78*10-9 16.2 16.2 102.0 0.006 31 49 62 16b 1 Const. K 290 6.14*10-8 16.2 16.2 56.0 0.015 31 49 62 16c 1 Const. K 290 4.17*10-7 16.2 16.3 24.0 0.036 31 50 62 16d 1 Const. K 290 4.87*10-9 16.3 16.3 191.3 0.007 31 50 63 C-12

Table C-5 (continued) Crack growth rate data for Specimen B3CT03 (9.4 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 16e 1 Const. K 290 5.09*10-7 16.3 16.3 20.0 0.032 31 50 63 16f 1 Const. K 290 1.52*10-8 16.3 16.3 66.0 0.009 31 50 63 16g 1 Const. K 290 2.61*10-6 16.3 16.4 18.4 0.165 31 50 63 16h 1 Const. K 290 1.66*10-9 16.4 16.4 651.0 0.007 31 50 63 17a 1 Const. K 320 1.27*10-9 16.4 16.4 166.8 0.003 31 50 63 17b 0.6 9 ks hold 320 2.59*10-7 16.4 16.5 72.1 0.072 31 50 63 17c 1 Const. K 320 1.45*10-9 16.5 16.5 334.9 0.002 31 50 63 187) 0.2 1, dk/da=6 320 6.22*10-5 16.5 22.7 3.7 0.981 31 50 64 197) 0.4 1, dk/da=6 320 3.99*10-5 22.7 23.3 0.7 0.087 43 69 87 207) 0.4 0.3, dk/da=6 320 1.45*10-5 23.3 23.6 1.1 0.055 44 71 90 217) 0.4 0.1, dk/da=6 320 4.45*10-6 23.6 23.9 3.4 0.052 44 72 91 227) 0.4 0.03, 320 1.79*10-6 23.9 24.2 9.0 0.056 45 73 92 dk/d 6 237) 0.4 0.01, 320 5.72*10-7 24.2 24.6 26.8 0.062 46 74 94

/

247) 0.4 0.003,dk/da= 320 1.71*10-7 24.6 24.9 96.0 0.059 47 76 95 6

257) 0.4 0.001,dk/da= 320 7.09*10-8 24.9 25.3 256.8 0.067 47 77 97 267) 0.5 0.001,dk/da= 320 6.43*10-8 25.3 25.7 277.8 0.068 48 78 99 6

277) 0.5 100s h, 320 8.29*10-8 25.7 26.1 227.3 0.072 49 80 100 dk/d 6 28 0.5 1 ks hold 320 5.35*10-8 26.1 26.1 248.0 0.052 50 81 102 29 0.5 9 ks hold 320 1.53*10-8 26.1 26.2 312.8 0.022 50 81 103 30 1 Const. K 320 2.87*10-9 26.2 26.2 197.0 0.003 50 82 103 C-13

Table C-5 (continued) Crack growth rate data for Specimen B3CT03 (9.4 dpa) based on corrected (non-linear) DCPD data R Load1) Temp.2) CGR3) KIstart KIstop t a K validity, %6)

Step ratio condition ºC mm/s MPam4) MPam4) hours5) mm5) 399 YS/2 YS/3 31 1 Const. K 340 8.08*10-9 26.2 26.2 285.4 0.003 50 82 103 32a 1 Const. K 290 1.78*10-8 26.2 26.2 36.5 0.009 50 82 103 32b 1 Const. K 290 5.80*10-6 26.2 28.4 56.6 1.234 50 82 103 32b1 1 Const. K 290 1.59*10-6 26.2 26.4 21.8 0.139 50 82 103 32b2 1 Const. K 290 5.69*10-6 26.4 27.1 20.9 0.446 51 83 105 32b3 1 Const. K 290 1.21*10-5 27.1 28.4 14.2 0.653 54 87 110 32c 1 dk/da=-4 290 2.18*10-5 28.4 28.1 11.3 0.904 58 94 118 32d 1 Const. K 290 2.48*10-5 28.1 30.9 10.7 0.941 60 97 123 32e 1 dk/da=-3 290 2.31*10-5 30.9 32.1 9.8 0.824 70 113 143 33a 1 dk/da=-3 320 4.01*10-5 32.0 37.0 12.2 1.759 76 124 156 33b 1 Const. K 320 6.90*10-5 37.0 41.9 2.5 0.602 101 164 207

1) Loading condition during testing. A number stands for the frequency in Hz during continuous cyclic load, whereas a number followed by a time and the word hold means the specimen was subjected to partial periodic unloading (PPU) with the hold time at maximum load indicated. Const. K or P means testing was conducted under constant stress intensity factor (K), or load (P), while dK/da=X means K was changed by the nominal rate (in MPam/mm) indicated.

2) Test temperature during the step.

3) Crack growth rate determined by linear regression of the corrected DCPD curve.

4) Stress intensity factor based on the crack length corresponding to the first (KIstart) and last (KIstop) data point of the regression analysis and load data.

5) Difference in test time or crack length between the first and last data point of the regression analysis.

6) K validity according to ASTM E399 and modified criteria based on the same standard. An effective yield strength was used for the modified criteria, i.e.,

YSeff1=YSunirr + (YSirr-YSunirr)/2 and YSeff2=YSunirr + (YSirr-YSunirr)/3. The actual K during the step was related to the K limit based on the crack length when the step started, i.e., 100% or higher means the criterion was exceeded.

7) Positive dK/da as indicated plus continuous cycling at frequency and R ratio reported. h stands for hold time.

C-14

CGR Data from Halden (reproduced from Tables 5 and 8 from HWR-1320 [5] and Table 7 from HWR-1236 [4])

Table C-6 Summary of crack growth rate data for CT1 from HWR-1320 (50 dpa)

Time, hrs K validity, %

CGR K, a Step t, hrs Load mm/s MPam mm start stop 399 YS/2 YS/3 1 5 53 48 1.21E-08 13.2 0.002 40 68 87 Constant 2 54 480 426 1.11E-08 14.6 0.017 45 75 97 Constant 3 654 980 326 1.04E-08 15.6 0.012 48 80 103 Constant 4 984 1084 100 2.60E-08 15.5 0.009 47 79 103 Cyclic 5 1090 1220 130 5.94E-09 15.7 0.003 48 81 104 Constant 6 1852 2164 312 3.06E-09 16.2 0.003 49 83 107 Constant 7 2166 2360 194 -3.10E-09 17.6 -0.002 54 90 117 Constant 8 2360 2740 380 1.68E-08 18.6 0.023 57 95 123 PPU 9 2908 3264 356 2.64E-08 20.6 0.034 63 106 136 PPU 10 4420 4776 356 5.55E-09 20.7 0.007 63 106 137 PPU 11 4780 5560 780 1.90E-08 22.9 0.053 70 117 152 PPU 12 5570 5880 310 3.74E-08 24.8 0.042 76 127 164 PPU 13 5892 5920 28 3.28E-07 25.0 0.033 76 128 166 PPU 14 7040 7170 130 1.24E-07 26.6 0.058 81 136 176 PPU 15 7180 7244 64 5.80E-07 27.8 0.134 85 143 184 PPU 16 7560 7800 240 2.38E-07 29.8 0.206 91 153 197 PPU 17 7800 8520 720 1.28E-07 31.6 0.332 96 162 209 PPU 18 8580 8810 230 1.35E-07 33.9 0.112 103 174 225 PPU C-15

Table C-7 Summary of crack growth rate data for CT4 from HWR-1320 (40 dpa)

Time, hrs K validity, %

CGR K, a Step t, hrs Load mm/s MPam mm start stop 399 YS/2 YS/3 1 5 53 48 1.49E-08 13.2 0.003 40 68 87 Constant 2 53 480 427 9.58E-09 14.5 0.015 44 74 96 Constant 3 654 980 326 7.37E-09 15.5 0.009 47 79 103 Constant 4 984 1084 100 2.67E-08 15.4 0.010 47 79 102 Cyclic 5 1090 1220 130 8.45E-09 15.6 0.004 48 80 103 Constant 6 1852 2164 312 8.55E-10 16 0.001 49 82 106 Constant 7 2166 2380 214 -5.22E-09 17.5 -0.004 53 90 116 Constant 8 2380 2740 360 1.74E-08 18.5 0.023 56 95 123 PPU 9 2908 3264 356 7.84E-08 20.6 0.100 63 106 136 PPU 10 4420 4776 356 3.22E-08 21 0.041 64 108 139 PPU 11 4778 4790 12 1.09E-05 23.8 0.471 73 122 158 PPU 12 4790 4798 8 2.48E-05 24.8 0.714 76 127 164 PPU 13 4798 4802 4 3.82E-05 31.4 0.550 96 161 208 PPU 14 4802 4808 6 1.86E-05 29.2 0.402 89 150 193 PPU C-16

Table C-8 Summary of crack growth rate data for CT4 from HWR-1236 (41 dpa)

Time, hrs K validity, %

CGR K, a Step t, hrs mm/s MPam mm start stop 399 YS/2 YS/3 1 362.5 371 8.5 1.72E-05 11.73 0.53 35.8 60.2 77.7 2 371 380 9 9.18E-06 12.6 0.30 38.4 64.6 83.4 3 380 492 112 2.34E-07 13.16 0.09 40.1 67.5 87.2 4 492 560 68 2.64E-06 13.81 0.65 42.1 70.8 91.5 5 560 593 33 5.03E-06 14.47 0.60 44.1 74.2 95.8 6 593 611 18 1.18E-06 14.7 0.08 44.8 75.4 97.4 7 611 711 100 2.06E-07 14.91 0.07 45.5 76.5 98.7 8* 717 876 159 6.02E-08 14.84 0.03 45.2 76.1 98.3 9 1071 1110 39 1.37E-06 15.99 0.19 48.8 82.0 105.9 10 1110 1122 12 4.83E-06 16.8 0.21 51.2 86.2 111.3 11 1122 1140 18 1.64E-05 17.38 1.06 53.0 89.1 115.1 12 1160 1288 128 1.19E-07 11.52 0.05 35.1 59.1 76.3 13 1328 1672 344 5.30E-08 11.89 0.07 36.3 61.0 78.7

• shaded row indicates data generated in hydrogen water chemistry. Remaining data generated in normal water chemistry C-17

CGR Data from ANL (reproduced from Table 9 of ANL-19/45 [6])

Table C-9: SCC CGR test results for the decommissioned Zorita baffle plate materials SCC CGR Sample Dose Test ID. (dpa) Env. K (MPam)a CGR with PPU (x10-8 mm/s) CGR w/o PPU (x10-8 mm/s) 15.5 0.99 -

15.6 - 0.05 21.5 1.51 -

A3CT04 ~0.06 PWR 21.6 - 0.38 27.5 1.77 -

27.5 - 0.20 16.9 2.43 -

Low- 17.2 - 1.47 DO, 21.2 2.86 -

B3CT14 ~8 high- 21.2 - 1.78 purity 27.0 1.71 -

27.0 - 1.27 16.4 1.00 -

Low- 16.4 - 0.96 DO, 20.7 1.35 -

ACT03 ~15 high- 20.7 - 0.59 purity 26.5 3.74 -

26.9 - 2.25 16.7 3.03 -

Low- 17.1 - 1.84 DO, 21.3 3.01 -

B1CT07 ~39 high- 21.1 - 1.82 purity 26.0 3.94 -

26.2 - 1.47 16.4 1.20 -

Low- 16.6 - 0.80 DO, 20.1 1.75 -

B1CT09 ~47 high- 20.2 - 1.24 purity 25.2 2.46 -

25.3 - 1.04 19.8 1.88 -

19.8 - 0.90 23.9 2.19 -

24.4 - 409 24.5 - 49.2 24.3-26.0 - 906 26.0-29.2 - 4290 B1CT08 ~47 PWR 5.5 - 1.5 17 - 0.40 30-39.0 - 4280 32.7-37.0 - 5780 25.3-30.9 - 5350 10.9 - 2.33 66.5-77.6 - 4340 aWhen a rising K condition is present, a K range is provided.

C-18

APPENDIX D: ADDITIONAL REFERENCES FOR SELECTED FIGURES This appendix contains additional references to correlate with the citations shown in Figures 10, 12, A-1, A-2, A-3, and B-1.

For Figures 10, 12, and B-1, the following references from MRP-451 are:

2. Electric Power Research Institute, Materials Reliability Program: Zorita Internals Research Project (MRP-440), Testing of Highly-Irradiated Baffle Plate Material, Product ID 3002016015, Palo Alto, CA, October 29, 2019 (proprietary; available for viewing through the NRC Technical Library6).

6. Jenssen, A. J. Stjrnster, K. Kese, R. Carter, J. Smith, A. Demma, and M. Hiser, Fracture Toughness Testing of an Irradiated PWR Core Barrel Weld, Fontevraud 9Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 17-20, 2018.

8. Electric Power Research Institute, BWRVIP-154, Revision 2: BWR Vessel and Internals Project, Fracture Toughness in High Fluence BWR Materials - Final Report, Product ID 1019077, Palo Alto, CA, 2009.

16. Electric Power Research Institute, BWRVIP-221: BWR Vessel and Internals Project, Crack Growth in High Fluence BWR Materials - Crack Growth Rate Testing of Types 304L and 316L at doses ranging from 3.5 to 13 dpa, Product ID 1019079, Palo Alto, CA, 2009.

18. A. Jenssen, R. Pathania, R. Carter, Crack Growth in Irradiated Austenitic Stainless Steels in BWR Environments, Fontevraud 8 - Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 2014.

30. A. Jenssen, J. Stjrnster, R. Pathania, R. Carter, Crack Growth in Irradiated Stainless Steel Welds in BWR Environments, Fontevraud 9 - Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 2018.

31. Electric Power Research Institute, Materials Reliability Program: A Review of Radiation Embrittlement of Stainless Steels for PWRs (MRP-79) - Revision 1, Product ID 1008204, Palo Alto, CA, 2004.

32. A. Demma, R. Carter, A. Jenssen, T. Torimaru, R. Gamble, Fracture Toughness of Highly Irradiated Stainless Steels in Boiling Water Reactors, 13th International Symposium on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Whistler, BC, Canada, August 2007.

50. U. Ehrnstén, K. Wallin, P. Karjalainen-Roikonen, S. van Dyck and P. Ould, Fracture Toughness of Stainless Steels Irradiated up to ~9 dpa in Commercial BWRs, Fontevraud 6Contribution of Materials Investigations to Improve the Safety and Performance of LWRs, Fontevraud, France, September 2006.

D-1

For Figures A-1, A-2, and A-3, the following references from MRP-440 are:

11. Electric Power Research Institute, BWRVIP-154, Revision 2: BWR Vessel and Internals Project, Fracture Toughness in High Fluence BWR Materials - Final Report, Product ID 1019077, Palo Alto, CA, 2009.

16. A. Jenssen, P. Efsing, B. Forssgren, B. Bengtsson and M. Molin, Examination of Highly Irradiated Stainless Steels from BWR and PWR Reactor Pressure Vessel Internals, Fontevraud 7Contribution of Materials Investigations to Improve the Safety and Performance of LWRs, Avignon, France, September 2010.

20. Electric Power Research Institute, BWRVIP-221: BWR Vessel and Internals Project, Crack Growth in High Fluence BWR Materials - Crack Growth Rate Testing of Types 304L and 316L at doses ranging from 3.5 to 13 dpa, Product ID 1019079, Palo Alto, CA, 2009.

33. A. Jenssen, R. Pathania, R. Carter, Crack Growth in Irradiated Austenitic Stainless Steels in BWR Environments, Fontevraud 8 - Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 2014.

36. Electric Power Research Institute, Models of Irradiation-Assisted Stress Corrosion Cracking of Austenitic Stainless Steels in Light Water Reactor Environments: Volume 2 Disposition Curves Application, Product ID 3002003103, Palo Alto, CA, 2014.

40. A. Jenssen, J. Stjrnster, R. Pathania, R. Carter, Crack Growth in Irradiated Stainless Steel Welds in BWR Environments, Fontevraud 9 - Contribution of Materials Investigations and Operating Experience to LWRs Safety, Performance and Reliability, Avignon, France, September 2018.

41. Electric Power Research Institute, Materials Reliability Program: A Review of Radiation Embrittlement of Stainless Steels for PWRs (MRP-79) - Revision 1, Product ID 1008204, Palo Alto, CA, 2004.

42. A. Demma, R. Carter, A. Jenssen, T. Torimaru, R. Gamble, Fracture Toughness of Highly Irradiated Stainless Steels in Boiling Water Reactors, 13th International Symposium on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Whistler, BC, Canada, August 2007.

43. Grnwall, B, Birath, S, Haag, Y. Ringhals 1, System 216-Hot Cell Examination of a SRM/IRM (dry tube) Guide Tube Manufactured by General Electric - Final Report (in Swedish), STUDSVIK/NS-89/75, Studsvik Nuclear AB, Sweden, 1989.

44. A. Jenssen, V. Grigoriev, R. Jakobsson, P. Efsing, Fracture Resistance Evaluation of a Flux Thimble Irradiated to 65 dpa in a PWR, Fontevraud 6Contribution of Materials Investigations to Improve the Safety and Performance of LWRs, Fontevraud, France, September 2006.

45. Conermann, J, Shogan R. P, Forsyth, D. R, Wilson, I. L, Tang, H. T, Characterization of Baffle-Former Bolts Removed from Service in US PWRs, 10th International Symposium on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Lake Tahoe, NV, August 2001.

D-2

48. R. W. Bosch, M. Vankeerberghen, R. Gérard, F. Somville, Crack initiation testing of thimble tube material under PWR conditions to determine a stress threshold for IASCC, Journal of Nuclear Materials. 461, pp. 112-121, 2015.

49. Electric Power Research Institute, Materials Reliability Program: Effect of Lithium on IASCC Initiation (MRP-413), Product ID 3002008082, Palo Alto, CA, 2016.

51. Electric Power Research Institute, Materials Reliability Program: Characterizations of Type 316 Cold Worked Stainless Steel Highly Irradiated Under PWR Operating Conditions (International IASCC Advisory Committee Phase 3 Program Final Report) (MRP-214),

Product ID 1015332, Palo Alto, CA, 2007.

52. K. Takakura, K. Nakata, M. Ando, K. Fujimoto and E. Wachi, Lifetime Evaluation for IASCC Initiation of Cold Worked 316 Stainless Steels BFB in PWR Primary Water, 13th International Symposium on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Whistler, BC, Canada, August 2007.

53. H. Nishioka, K. Fukuya, K. Fujii, T. Torimaru, IASCC properties and mechanical behavior of stainless steels irradiated up to 73dpa, 13th International Symposium on Environmental Degradation of Materials in Nuclear Power SystemsWater Reactors, Whistler, BC, Canada, August 2007.

54. T. M. Karlsen, M. Espeland, H. Jenssen, Crack Initiation in Irradiated Constant Load Tensile Specimens using Off-line Instrumentation - Results from First and Second Loadings, Institutt for Energiteknikk, Norway, 2007, FC Note 1545. Available on:

Cooperative IASCC Research II Program: Final Comprehensive CIR II CD Version 10.08, Product ID 1021235, EPRI, Palo Alto, CA, 2010.

55. C. Pokor, A. Toivonen, M. Wintergerst, U. Ehrnstén, W. Karlsen, J-P. Massoud, Determination of the Time to Failure Curve as a Function of Stress for a Highly Irradiated AISI 304 Stainless Steel after Constant Load Tests in Simulated PWR Water Environment, Fontevraud 7Contribution of Materials Investigations to Improve the Safety and Performance of LWRs, Avignon, France, September 2010.

63. J. Nakano, T.M. Karlsen, M. Espeland, Summary of Results from the PWR Crack Growth Rate Investigation IFA-670, OECD Halden Reactor Project, Report HWR-843, August 2008.

66. U. Ehrnstén, K. Wallin, P. Karjalainen-Roikonen, S. van Dyck and P. Ould, Fracture Toughness of Stainless Steels Irradiated up to ~9 dpa in Commercial BWRs, Fontevraud 6Contribution of Materials Investigations to Improve the Safety and Performance of LWRs, Fontevraud, France, September 2006.

D-3