ML25357A197

Docket No. Stn 50-528, Renewed Operating License No. NPF-41, Relief Request 77 - Pressurizer Lower Instrument Nozzle

ML25357A197
Person / Time
Site:	Palo Verde
Issue date:	12/19/2025
From:	Spina J Arizona Public Service Co
To:	Office of Nuclear Reactor Regulation, Document Control Desk
Shared Package
ML25357A196	List: ML25357A197
References
102-09020-JLS/MDD
Download: ML25357A197 (0)

Text

10 CFR 50.55a A member of the STARS Alliance, LLC Callaway

• Diablo Canyon

• Palo Verde

• Wolf Creek Attachments 5 and 6 transmitted in this Enclosure contains PROPRIETARY information.

When Attachments 5 and 6 are separated, this transmittal is decontrolled.

Jennifer Spina Vice President Nuclear Regulatory & Oversight Palo Verde Nuclear Generating Station P.O. Box 52034 Phoenix, AZ 85072 Mail Station 7602 Tel: 623.393.4621 102-09020-JLS/MDD December 1, 2025 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington, DC 20555-0001

Subject:

Palo Verde Nuclear Generating Station (PVNGS) Unit 1 Docket No. STN 50-528 Renewed Operating License No. NPF-41 Relief Request 77 - Pressurizer Lower Instrument Nozzle Pursuant to Title 10 of the Code of Federal Regulations (10 CFR) 50.55a, Codes and Standards, paragraph (z)(1), Arizona Public Service Company (APS) requests Nuclear Regulatory Commission (NRC) staff authorization of Relief Request 77, on the basis that the proposed alternative provides an acceptable level of quality and safety.

Through a modification, four Unit 1 upper Pressurizer instrument nozzles and two lower nozzles were replaced during Unit 1 refueling outage 25 (1R25). These nozzle replacements were a preemptive mitigation for potential Primary Water Stress Corrosion Cracking (PWSCC) identified on the TE-101 thermowell nozzle during Unit 1 refueling outage 24 (1R24), which resulted in pressure boundary leakage [Relief Requests 70 and 73, Agencywide Documents Access and Management System (ADAMS) Accession Numbers ML24197A199 and ML25104A042, respectively]. No age-related flaws were known to exist on the six instrument nozzle welds prior to modification implementation.

No pressure boundary leakage was identified. A corrosion evaluation and an ASME Section III Analysis were previously submitted to the NRC as part of commitments made in Palo Verde Relief Request 76 (ADAMS accession No. ML25142A405) and remain relevant to this relief request.

Relief Request 77 applies to modification of the lower instrument nozzle RC-023 (V208).

It is being submitted to provide appropriate analyses and justify continued use of the nozzle modification for the remainder of the plant operating life.

Authorization is requested for continued operation through the duration of the initial license extension (June 1, 2045). Approval is requested prior to startup from the next Unit 1 refueling outage (1R26), in November 2026, as the current one cycle relief request will expire.

APS makes no new commitments to the NRC by this letter.

102-09020-JLS/MDD ATTN: Document Control Desk U.S. Nuclear Regulatory Commission Relief Request 77 - Pressurizer Lower Instrument Nozzle Page 2 Attachments 5 and 6 transmitted in this Enclosure contains PROPRIETARY information. When Attachments 5 and 6 are separated, this transmittal is decontrolled.

Should you need further information regarding this letter, please contact Michael. D.

DiLorenzo, Licensing Department Leader, at (623) 393-3495.

Sincerely, JS/MDD/cr

Enclosure:

Arizona Public Service Company, Palo Verde Nuclear Generating Station, Unit 1, Relief Request Number 77 cc:

J. D. Monninger NRC Region IV Regional Administrator W. T. Orders NRC NRR Project Manager for PVNGS E. R. Lantz NRC Senior Resident Inspector for PVNGS Spina, Jennifer (Z08962)

Digitally signed by Spina, Jennifer (Z08962)

Date: 2025.12.19 09:36:45

-07'00'

Enclosure ARIZONA PUBLIC SERVICE COMPANY PALO VERDE NUCLEAR GENERATING STATION, UNIT 1 Relief Request Number 77

10 CFR 50.55a Attachments 5 and 6 transmitted in this Enclosure contain PROPRIETARY information. When Attachments 5 and 6 are separated, this transmittal is decontrolled.

i Contents Page 1.0 ASME CODE COMPONENT AFFECTED................................................................. 1 2.0 APPLICABLE CODE EDITION AND ADDENDA....................................................... 1 3.0 APPLICABLE CODE REQUIREMENTS.................................................................. 1 4.0 REASON FOR REQUEST................................................................................... 2 5.0 PROPOSED ALTERNATIVE AND BASIS FOR USE.................................................. 7 6.0 DURATION OF PROPOSED ALTERNATIVE.......................................................... 12 7.0 PRECEDENTS................................................................................................ 13

8.0 REFERENCES

................................................................................................ 14 List of Figures Figure 4-1 Pressurizer Lower Instrument Nozzle - Pre-1R25 Configuration..................... 5 Figure 4-2 Pressurizer Lower Instrument Nozzle - Modified Configuration....................... 6 : Non-Proprietary Version Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds - Framatome Non-Proprietary, Document Number 32-9399132-001 : Non-Proprietary Version Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Modification -

Non-Proprietary, Document Number 32-9398249-000 : Affidavit Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds - Framatome Proprietary, Document Number 32-9399131-001 : Affidavit Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Modification, Document Number 32-9398289-000 : Proprietary Version Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds - Framatome Proprietary, Document Number 32-9399131-001 : Proprietary Version Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Modification, Proprietary Document Number 32-9398289-000

Enclosure Relief Request Number 77 ii Nomenclature Acronym Definition ALJGW As-Left J-Groove Weld APS Arizona Public Service Company ASME American Society of Mechanical Engineers ATTB Ambient Temperature Temper Bead CEA Control Element Assembly EPFM Elastic Plastic Fracture Mechanics GTAW Gas Tungsten Arc Welding HAZ Heat Affected Zone HIC Hydrogen Induced Cracking ISI Inservice Inspection JGW J-Groove Weld LAS Low Alloy Steel LEFM Linear Elastic Fracture Mechanics NDE Nondestructive Examination NRC Nuclear Regulatory Commission OCJ One-Cycle Justification OD Outside Diameter PDI Performance Demonstration Initiative PT (Liquid) Penetrant Testing PVNGS Palo Verde Nuclear Generating Station PWHT Post Weld Heat Treatment PWSCC Primary Water Stress Corrosion Cracking PZR Pressurizer RCS Reactor Coolant System RFO Refueling Outage RG Regulatory Guide RTNDT Reference Temperature Nil Ductility Transition UT Ultrasonic Testing

Enclosure Relief Request Number 77 1

1.0 ASME CODE COMPONENT AFFECTED Component:

Pressurizer Lower Instrument Nozzle Code Class:

1 Examination Category:

B-P, American Society of Mechanical Engineers (ASME) Code Section XI Item No.

B15.10, Table IWB-2500-1 (B-P)

==

Description:==

Pressurizer Lower Instrument Nozzle RC-023 (V208),

Nominal Pipe Size 1-inch Interval:

Fourth (4th) (June 1, 2019, to July 17, 2028) 2.0 APPLICABLE CODE EDITION AND ADDENDA The current edition for the fourth Inservice Inspection (ISI) interval is the American Society of Mechanical Engineers (ASME) Code,Section XI, 2013 Edition.

The Code of Construction for the Pressurizer is the ASME Code Section III, 1971 Edition with Addenda through Winter 1973.

The modification installation Code of Construction is the ASME Code,Section III, 1974 Edition with Addenda through Winter 1975.

The fourth ISI interval for PVNGS Unit 1 began on June 1, 2019, and is currently scheduled to end on July 17, 2028.

3.0 APPLICABLE CODE REQUIREMENTS ASME Code,Section XI, 2013 Edition Flaw Removal IWA-4412 states Defect removal shall be accomplished in accordance with the requirements of IWA-4420.

IWA-4421 states Defects shall be removed or mitigated in accordance with the following requirements:

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Flaw Evaluation IWA-3300(b) states, in part, Flaws shall be characterized in accordance with IWA-3310 through IWA-3390, as applicable.

IWB-3420 states Each detected flaw or group of flaws shall be characterized by the rules of IWA-3300 to establish the dimensions of the flaws. These dimensions shall be used in conjunction with the acceptance standards of IWB-3500.

IWB-3610(b) states, in part, For purposes of evaluation by analysis, the depth of flaws in clad components shall be defined in accordance with Figure IWB-3610-1....

Successive Examinations IWB-2420(a) states, in part, The sequence of component examinations which was established during the first inspection interval shall be repeated IWB-2420(b) states, in part, If a component is accepted for continued service in accordance with IWB-3132.3 or IWB-3142.4, the areas containing flaws or relevant conditions shall be reexamined ASME Code Case N-722-1, Additional Examinations for PWR Pressure Retaining Welds in Class 1 Components Fabricated With Alloy 600/82/182 Materials Item No. B15.180, Instrumentation Connections, require visual examination each refueling outage with IWB-3522 acceptance standards.

Welding Code Case N-638-10, Similar and Dissimilar Metal Welding Using Ambient Temperature Machine GTAW Temper Bead Technique Case N-638-10 provides requirements for automatic or machine gas tungsten arc welding (GTAW) of Class 1 components without the use of preheat or post-weld heat treatment.

Paragraph 4(a)(2) of Code Case N-638-10 requires the completed weld to be nondestructively examined after the three tempering layers have been in place for at least 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br />.

4.0 REASON FOR REQUEST Palo Verde Nuclear Generating Station (PVNGS) is in the second period of the fourth 10-year ISI interval. Modification activities to the Unit 1 pressurizer instrument nozzles were performed during the 1R25 refueling outage. The modifications activities performed replaced the primary water stress corrosion cracking (PWSCC) susceptible Alloy 82 pressure retaining weld material with Alloy 52M weld material. This relief request applies to modification of the lower instrument nozzle RC-023 (V208). The pressurizer instrument nozzle is Item No. B15.10 in Table IWB-2500-1, ASME Section XI, and Item No. B15.180 in Table 1, Code Case N-722-1.

The original Alloy 600 pressurizer lower instrument nozzle was pre-emptively replaced in 1992 with an Alloy 690 nozzle, an Alloy 690 outer sleeve, an Alloy 82 weld pad, and an Alloy 82 nozzle-to-weld pad J-groove weld. The current modification consisted of

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removal the existing Alloy 690 nozzle, Alloy 82 J-groove weld, and Alloy 82 weld pad.

ASME Section XI, Code Case N-638-10, and ASME Section III were used to apply a new weld pad and J-groove weld on the outer surface of the pressurizer bottom head surface using Alloy 52M filler metal. Figure 4-1 depicts the nozzle configuration prior to RFO 1R25. Figure 4-2 provides a sketch of the pressurizer bottom head instrument nozzle modification. The new weld pad was welded to the outer surface of the pressurizer bottom head using machine Gas Tungsten Arc Welding (GTAW) Ambient Temperature Temper Bead (ATTB) welding, with inert shielding gas. The modification attached an Alloy 690 nozzle to the Alloy 52M weld pad with a partial penetration weld using a manual GTAW welding technique and Alloy 52M filler metal.

During implementation of the modification, it was observed that there was moisture between the Alloy 690 outer sleeve and the low alloy steel bottom head. The presence of moisture in the annulus indicated a possible leak path through the autogenous weld of the corrosion liner and the original J-groove weld. Subsequently, APS performed a borescope inspection of the autogenous outer sleeve weld to confirm the presence of a crack in the autogenous weld. The visual examination documented an approximate 1/2 to 3/4-inch linear indication in the autogenous weld and an approximate 3/16 to 1/4-inch linear indication on the toe of the autogenous weld.

The linear indications along with the presence of moisture between the corrosion liner and the bottom head indicated a crack in the autogenous weld or a possible crack in the original J-groove weld remnant. Due to the possibility of a crack in the original J-groove weld, a crack growth evaluation was performed of a postulated worst-case (largest) crack in the J-groove weld which bounds indications found in the autogenous weld.

A flaw evaluation was performed that demonstrated the acceptability of leaving the original partial penetration J-groove attachment weld, with a maximum postulated flaw, in place for the life of the pressurizer (see Basis for Flaw Analytical Evaluation below).

IWA-4412 contains requirements for the removal of, or the reduction in size of defects.

The postulated flaw in the original J-groove weld and autogenous weld will not be removed; therefore, an alternative is proposed for these requirements.

IWA-4412 requires defect removal in accordance with IWA-4420. IWA-3300 requires flaws detected during inservice examinations to be sized. IWB-2420 requires successive examinations of flaws accepted for continued service. IWB-3400 and IWB-3600 were written with the expectation that volumetric NDE techniques such as Ultrasonic Testing (UT) would be used to determine the flaw size and shape. In support of the flaw evaluation, the ASME Code paragraphs IWB-3420 and IWB-3610(b) require characterization of the flaw. There is no Performance Demonstration Initiative (PDI) qualified technique to perform NDE of the configuration of the partial penetration J-groove weld and autogenous weld of the Alloy 690 sleeve at the ID surface of the pressurizer that can be used to accurately characterize the location, orientation, or size of a potential flaw in the original J-groove weld. Therefore, a postulated flaw was demonstrated to be acceptable for the life of the modification and an alternative is proposed for IWA-4412, IWA-3300, and IWB-2420.

NB-4620 requires all welds to be post-weld heat treated except as otherwise permitted in NB-4622.7. APS previously proposed the relief request for one cycle of operation to install a weld pad using ATTB welding in accordance with ASME Case N-638-10. The NRC has approved Case N-638-10 to allow ATTB welding with austenitic filler materials within 1/8-inch of or on ferritic materials without the requirement for preheat or post-weld heat treatment. The current version of this Case requires that the three tempering

Enclosure Relief Request Number 77 4

weld layers be in place for at least 48-hours prior to performance of surface and volumetric NDE. Liquid penetrant and ultrasonic acceptance examinations were performed before the 48-hour period ended. Technical justification for austenitic filler materials has been developed to allow NDE methods to be performed after completion of the weld repair, without waiting for the 48-hour hold time as described in Section 5.C below.

As discussed, APS submitted a relief request for one cycle operation, which began after 1R25. This relief request is to continue operation at PVNGS Unit 1 through the initial license extension (June 1, 2045). Although this relief request is limited to the initial license extension (60 years of plant operation), APS has determined that the modification performed to pressurizer lower instrument nozzle RC-023 (V208) utilizing the alternatives specified in this relief request will provide an acceptable level of quality and safety through 80-years of plant operation.

The proposed alternative is provided in accordance with 10 CFR 50.55a(z)(1).

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Figure 4-1 Pressurizer Lower Instrument Nozzle - Pre-1R25 Configuration

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Figure 4-2 Pressurizer Lower Instrument Nozzle - Modified Configuration

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5.0 PROPOSED ALTERNATIVE AND BASIS FOR USE A. Proposed Alternatives In accordance with 10 CFR 50.55a, Codes and Standards, paragraph (z)(1), APS proposes specific alternatives to the requirements specified in Section 3 above on the basis that performing the alternatives stated below provide an acceptable level of quality and safety.

A design analysis (Reference 8.8) was performed in accordance with ASME Code Section III, Subsection NB, 2013 Edition. The analysis demonstrates that the pressurizer lower instrument nozzle replacement meets all the applicable requirements of Section III. All primary stresses, primary plus secondary stresses, fatigue criteria and sizing requirements are satisfied. Thus, the new nozzle will not eject from the pressurizer under design and all applicable service level conditions throughout the life of the modification. The fatigue life of the replacement lower instrument nozzle is acceptable for an 80-year service life.

In addition to the ASME Code rules the following alternatives are proposed:

Flaw Removal and Flaw Evaluation As an alternative to flaw removal or reduction in size, of the original J-groove weld on the inner surface of the pressurizer bottom head, to meet the applicable acceptance standards per IWA-4412, APS implemented the lower instrument nozzle modification which utilized an OD weld pad and replacement full nozzle as described in the previous relief request for one cycle of operation.

The as-left J-groove weld (ALJGW) analysis report (Reference 8.11) performed a crack growth analysis for a postulated radial-axial corner flaw through the entire J-groove weld and buttering to demonstrate that the postulated flaw is acceptable for 80 years of operation.

As an alternative to performing the NDE required to characterize a flaw under IWB-3420 and IWB-3610(b) in the pressurizer lower instrument nozzle penetration, APS proposes analyzing a maximum postulated flaw that bounds the range of flaw sizes that could exist in the original J-groove weld and buttering. See "Basis for Flaw Analytical Evaluation below.

Welding As an alternative to NB-4620, APS proposed in the previous relief request for one cycle of operation, to install a weld pad using ATTB welding in accordance with ASME Case N-638-10. The NRC has approved ASME Case N-638-10 in Reg. Guide 1.147, Revision 20, to allow ATTB welding of dissimilar materials.

Examination (liquid penetrant surface and UT volumetric) of the completed weld pad was performed in accordance with ASME Section III acceptance criteria after the weld pad has been prepared for NDE and dimensionally inspected. The examinations of the weld pad were acceptable.

Pursuant to 10 CFR 50.55a(z)(1), APS proposes an alternative to ASME Section XI and ASME Case N-638-10. An alternative was proposed to the requirements of N-638-10, Paragraph 4(a)(2), that requires a 48-hour hold time prior to performing

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NDE. APS performed the weld modification with austenitic filler material in accordance with the ATTB welding technique of Code Case N-638-10, with one exception. As an alternative to performing the required NDE at least 48-hours after the three tempering layers reached ambient temperature, APS proposed to perform the NDE methods after completion of the weld modification. The bases for these alternatives are provided below.

APS performed the weld modification to the pressurizer bottom head by using austenitic Nickel-Alloy 52M (SFA-5.14, ERNiCrFe-7A) filler material over the modification area. The weld modification included a minimum of three (3) layers per the temper bead rules in Case N-638-10. The ATTB technique of ASME Section XI Case N-638-10 was approved by the NRC in Reg. Guide 1.147, Revision 20.

B. Basis for Flaw Analytical Evaluation The assumptions of IWB-3600 of ASME Section XI for analytical flaw evaluation are that cracks are fully characterized in accordance with IWB-3420 to compare the calculated parameters to the acceptable parameters addressed in IWB-3500. There are no qualified UT examination techniques for examining the original nozzle-to-Pressurizer (PZR) bottom head J-groove weld. Therefore, since it is impractical to characterize the flaw geometry that may exist therein, it is conservatively assumed that the as-left condition of the remaining J-groove weld includes flaws extending through the entire Alloy 82/Alloy 182 J-groove weld and buttering.

Since uphill and downhill hoop stresses in the J-groove weld at the spherical head are the higher stressed location at the nozzle penetration, the preferential direction for cracking is radial relative to the PZR head. Therefore, a radial-axial flaw (radial with respect to the nozzle axis) in the Alloy 82/Alloy 182 J-groove weld and buttering is postulated and would propagate by PWSCC through the weld and buttering to the interface with the low alloy steel PZR material. Any growth of the postulated as-left flaw into the PWSCC resistant low alloy steel would be by fatigue crack growth under cyclic loading conditions.

Based on a combination of linear elastic fracture mechanics (LEFM) and elastic-plastic fracture mechanics (EPFM) evaluation, the Life of Repair analyses (Reference 8.11) demonstrated that the postulated flaw is acceptable for 80 years of operation after the repair, which exceeds the remaining number of years of the current extended plant design life (60-years). In addition, the primary stress criteria specified in IWB-3610(d)(2) as well as 3.1(c) and 3.2(a)(3) of Code Case N-749 are satisfied since the limit load analysis shows that the structure does not collapse at a pressure above 150% of the design pressure. The process for the Life of Repair analyses was as follows:

1.

A three-dimensional finite element model of the pressurizer lower instrument nozzle was generated. This analysis includes weld simulation of the original J-groove weld (JGW) attaching the original instrument nozzle to the pressurizer bottom head, and simulation of the modifications involving an outer diameter weld pad and JGW attaching the replacement nozzle to the weld pad and autogenous weld attaching the outer sleeve to the JGW. The state of WRS at the end of the final welding step after shakedown, as determined by the ANSYS finite element analysis, is summarized to support the subsequent flaw evaluation of the original ALJGW.

2. The transients applicable for the as-left J-groove weld flaw evaluation are all

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specified design transients (normal/upset, emergency/faulted and test conditions). Using the EPFM acceptance criteria outlined in Sections 3.1 and 3.2 of the ASME Code Case N-749 and in accordance with temperature requirements specified in NRC Regulatory Guide 1.147, the results demonstrated that the value of Japplied, including appropriate structural factors, is below the J-integral of the material toughness, at a ductile crack extension of 0.1-inch for flaws at the size of the end-of-evaluation period.

3. Fatigue crack growth for cyclic loading conditions using operational stresses from pressure and thermal loads and crack growth rates from ASME Section XI, Non-mandatory Appendix A, Sub-article A-4300 for ferritic material in a primary water environment was calculated. Based on the results of LEFM analysis only or a combination of LEFM and EPFM analyses, a postulated flaw remaining in the original Alloy 82/Alloy 182 J-groove weld and buttering for the modified pressurizer lower instrument nozzle was shown to be acceptable.

In addition, evaluation of primary plus secondary stresses and metal fatigue usage were performed for the new weld pad, new J-groove weld, replacement nozzle, and pipe weld which demonstrated compliance with the Design Specification and the applicable ASME Criteria for reinforcement requirements, weld sizes, and nozzle thickness.

Relief is requested from flaw characterization specified in IWB-3420.

C. Basis for Elimination of the ambient 48-Hour Hold Time Elimination of the 48-hour hold was based on the white paper developed for the proposed change to ASME Code Case N-888-1. The white paper was previously provided in APS letter dated April 23, 2025 (Agencywide Documents and Access Management System (ADAMS) Accession No. ML25113A296) which was based on a white paper PVP 2023-107489, Elimination of the 48-hour Hold for Ambient Temperature Temper Bead Welding with Austenitic Weld Metal. Although this ASME Case is not approved in Reg. Guide 1.147, it has been approved by the ASME Section XI Standards Committee. Since Code Case N-888 is the culmination of temper bead code cases that have been produced over the years, combining requirements from N-638, N-839, and Appendix I in cases such as N-740 and N-754, etc., the justification is also applicable to the planned use of Code Case N-638-10 at PVNGS Unit 1.

D. Corrosion Evaluation The pressurizer lower instrument nozzle modification left a small portion of low alloy steel in the pressurizer bottom head exposed to primary coolant. An evaluation was performed (Reference 8.9) for the potential corrosion concerns at the pressurizer low alloy steel wetted surface. Galvanic corrosion, hydrogen embrittlement, stress corrosion cracking (SCC), and crevice corrosion are not expected to be a concern for the exposed low alloy steel base metal. General corrosion of the exposed low alloy steel base metal will occur in the region of the pressurizer bottom head between the weld pad and the nozzle outer sleeve. Due to the depletion of oxygen, tight geometry, and lack of Reactor Coolant System (RCS) flow at the exposed low alloy steel, corrosion products are expected to fill the gap created by the repair and general corrosion rates will significantly decrease after a period of time. However, a sustained corrosion rate, considering the time periods of plant start up, plant operating, and plant shutdown, was applied and the resultant increase in bore

Enclosure Relief Request Number 77 10 diameter was considered in the reinforcement calculation (per NB-3330) as part of the ASME Section III analysis. The corrosion evaluation (Reference 8.9) and the ASME Section III analysis (Reference 8.8) have been previously submitted to the NRC as part of commitments made in Palo Verde Relief Request 76 (ADAMS Accession No. ML25142A405).

The low alloy steel behind the outer sleeve was likely already exposed to primary coolant by the linear indications that were discovered during the modification process. It is conservatively assumed that these indications (and exposed low alloy steel) could have existed since 1992, when the prior nozzle replacement occurred.

The corrosion evaluation (Reference 8.9) concludes that the corrosion considerations are the same for any existing, prior exposure as they are for the wetted surface created by the modification. The same, sustained corrosion rate is applicable from 1992 for the life of the repair for this location.

Additionally, the modification utilized a replacement Alloy 690 nozzle and new Alloy 52M weld pad and J-groove weld to attach the Alloy 690 nozzle. Alloy 52M is also used in a weld configuration to attach the replacement Alloy 690 nozzle to the Type 304 replacement instrument piping. The aforementioned corrosion evaluation (Reference 8.9) also considers corrosion of the replacement piping at the attachment point to the new nozzles.

General Corrosion As previously mentioned, a sustained corrosion rate, considering the time periods of plant start up, plant operating, and plant shutdown, was applied for the locations of the exposed low alloy steel and the resultant increase in bore diameter was considered in the reinforcement calculation as part of the ASME Section III analysis (Reference 8.8).

Although general corrosion of the replacement Alloy 690 may be present, the corrosion rate for this material is small and not expected to be of concern for this modification.

Galvanic Corrosion The results of the NRCs boric acid corrosion program have shown that the electrochemical potential difference between SA-533 Grade B, Alloy 600, and Type 308 stainless steel (nominal chemistry of the pressurizer cladding) is not significant enough to consider galvanic corrosion as a strong contributor to the overall boric acid corrosion process (NUREG-1823). Therefore, it was judged that galvanic corrosion between the exposed pressurizer low alloy steel, Alloy 690, or their weld metals is not a concern for this repair configuration. This is supported by several laboratory studies documented in EPRI Report 1000975 in which low alloy steel specimens were coupled and uncoupled to stainless steel exposed to a borated water environment at various temperatures. The corrosion rates for the coupled and uncoupled conditions were determined to be similar. Additionally, galvanic corrosion rate of carbon steel coupled to stainless steel in a boric acid solution in the absence of oxygen is about equal to the general corrosion rate. The results of this study are also applicable to nickel-based alloys such as Alloy 690, since austenitic stainless steels have a similar electrochemical potential as nickel-based alloys.

Enclosure Relief Request Number 77 11 Hydrogen Embrittlement Hydrogen embrittlement occurs when the ductility of a material is degraded due to the diffusion of hydrogen into the metal lattice. This type of damage usually occurs in combination with an acting stress and is observed most often in plastically deformed metals of high pressure hydrogen environments. Concentrations of hydrogen in the reactor coolant system and hydrogen generated from corrosion are deemed insufficient to induce hydrogen cracking in the low alloy steel of the pressurizer. Therefore, it was determined that hydrogen embrittlement is not expected to be a concern for the exposed pressurizer low alloy steel in the repaired configuration.

Stress Corrosion Cracking Under normal PWR conditions (deaerated with hydrogen overpressure and low conductivity), primary water is not a particularly aggressive environment for SCC of low alloy steel unless a departure from normal operating conditions occurs.

Additionally, this environment does not generally support localized corrosion of low alloy steel, which precludes the formation of a pit or notch that would constitute a stress concentrator or SCC initiation site. Finally, a review of relevant laboratory work and field experience has been previously performed by the Combustion Engineering Owners Group (CE NPSD-1198-NP, Revision 00), with the conclusion being that low alloy and carbon steels will not be subject to SCC in an RCS environment. Therefore, it has been determined that stress corrosion cracking of the low alloy steel of the pressurizer is not a concern for this repair configuration.

Alloy 690 and Alloy 52M are concluded to have a low susceptibility to primary water stress corrosion cracking (PWSCC) based on laboratory studies and are expected to have superior PWSCC performance to the existing configuration, which utilized Alloy 82 weld material. Therefore, PWSCC of Alloy 690 and Alloy 52M is not expected to be a concern.

For the replacement instrument piping, SCC susceptibility for the replacement instrument piping adjacent to the new socket weld is expected to be the same as or less than that of the existing configuration, primarily due to the use of replacement Type 304 piping in lieu of reusing the existing piping. Practices consistent with Regulatory Guide 1.44 used in the welding specification, to control sensitization, also help to reduce susceptibility to SCC.

Crevice Corrosion The geometry of the gap between the pressurizer low alloy steel and the nozzle could create the conditions necessary for crevice corrosion. However, operating experience for PWRs shows that crevice corrosion is not a common concern for low alloy steels in the RCS, and laboratory testing has found that crevice corrosion rates of low alloy steel in aerated and deaerated primary water is less than the respective general corrosion rate. Furthermore, operating experience suggests that the crevice formed by the repair will fill with general corrosion products over time, isolating the surface of the low alloy steel from the RCS and decreasing the rate of corrosion within the crevice. Therefore, it was determined that crevice corrosion of the low alloy steel is not expected to be a concern. Crevice concern is not expected to be a concern for Alloy 690 or Alloy 52M due to these materials excellent resistance to general and crevice corrosion under typical PWR conditions.

Enclosure Relief Request Number 77 12 Low Temperature Crack Propagation (LTCP)

This corrosion concern is applicable to the nickel-based alloys and is considered a form of hydrogen embrittlement. While RCS temperatures in the shutdown and startup conditions are low enough to induce LTCP in very high-pressure hydrogen environments, such environments are not typical of PWR systems. Therefore, LTCP is not expected to be a concern for Alloy 690 or Alloy 52M.

E. Loose Parts Evaluation Given the original pressurizer lower instrument nozzle J-groove weld will not be removed, APS completed a loose parts evaluation to assess the potential for J-groove weld fragments, or the outer sleeve entering the pressurizer during power operation (Reference 8.10). The most probable scenario is that any fragments of the weld or pieces of the sleeve will remain in the pressurizer lower head.

F. Conclusion The modification to the pressurizer lower instrument nozzle produced an effective modification that restored and maintained the pressure boundary integrity of the penetration in the pressurizer. This alternative provides improved structural integrity and reduced the likelihood of leakage for the primary system.

The modified configuration meets the requirements of ASME Section III, and the cumulative usage factors at the critical locations are less than 1.0 for the number of design cycles specified for 80 years of plant operation. A postulated flaw in the original partial penetration J-groove weld was demonstrated to be acceptable for 80 years of operation. Corrosion mechanisms (galvanic corrosion, hydrogen embrittlement, stress corrosion cracking, crevice corrosion, and general corrosion) are not expected to be a concern for the life of the modification.

The temper bead technique is an effective tool for performing repairs on carbon and low alloy steel (P-No. 1 and P-No. 3) materials. Case N-638-10 provisions allow for ambient temperature temper bead welding with no post weld heat treatment. The data and testing performed shows that when austenitic weld metal is used the level of diffusible hydrogen content in the ferritic base metal heat affected area (HAZ) is too low to promote hydrogen-induced cracking (HIC). Therefore, the 48-hour hold requirement in Code Case N-638-10 is not necessary prior to examination of the weld as HIC is not considered credible.

In accordance with 10 CFR 50.55a(z)(1), APS has concluded that the proposed alternatives for flaw evaluation and eliminating the 48-hour hold provide an acceptable level of quality and safety as discussed above.

6.0 DURATION OF PROPOSED ALTERNATIVE The proposed alternatives are provided for the remainder of the PVNGS, Unit 1 initial license extension which expires June 1, 2045.

Enclosure Relief Request Number 77 13 7.0 PRECEDENTS The following relief request was previously approved to eliminate the 48-hour hold time specified in Case N-638-10:

NRC verbal authorization on May 9, 2023 [Agencywide Documents Access and Management System (ADAMS) Accession No. ML23129A312] for Beaver Valley, Unit 2 relief request 2_TYP-4-RV-06 (ADAMS Accession No. ML23118A381).

Letter from David Gudger (Constellation Energy Generation, LLC) to U.S. NRC, "Submittal of Emergency Relief Request I5R-11 Concerning the Installation of a Weld Overlay on Reactor Pressure Vessel Recirculation Inlet Nozzle N2E Safe End-to-Nozzle Dissimilar Metal Weld (32-WD-208)," dated March 24, 2023, (ADAMS Accession No. ML23083B991).

The following relief requests were previously approved for the flaw analytical evaluation:

NRC approval via verbal authorization on October 27, 2023 (ADAMS Accession No. ML23303A011) for Palo Verde Nuclear Generating Station, Unit 1. The NRC Safety Evaluation was subsequently issued on August 26, 2025 (ML25230A047).

NRC approval via verbal authorization on November 6, 2020 (ADAMS Accession No. ML20314A028) for Peach Bottom Atomic Power Station, Unit 2. The NRC Safety Evaluation was subsequently issued on April 23, 2021 (ADAMS Accession No. ML21110A680).

NRC verbal authorization on April 16, 2012, for Quad Cities, Unit 2 (ADAMS Accession No. ML12107A472). The NRC Safety Evaluation was subsequently issued on January 30, 2013 (ADAMS Accession No. ML13016A454).

NRC approval via a verbal authorization on May 17, 2017, for Limerick, Unit 2 (ADAMS Accession No. ML17137A307). The NRC Safety Evaluation was subsequently issued on August 14, 2017 (ADAMS Accession No. ML17208A090).

Enclosure Relief Request Number 77 14

8.0 REFERENCES

8.1 ASME Code,Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components," 2013 Edition.

8.2 ASME Code, Case N-638-10, Similar and Dissimilar Metal Welding Using Ambient Temperature Machine GTAW Temper Bead Technique,Section XI, Division 1, dated May 6, 2019.

8.3 Code Case N-749, Alternative Acceptance Criteria for Flaws in Ferritic Steel Components Operating in the Upper Shelf Temperature Range,Section XI, Division 1, dated March 16, 2012.

8.4 ASME Code,Section III, Nuclear Power Plant Components, 1971 Edition including Addenda through Winter 1973.

8.5 ASME Code,Section III, Nuclear Power Plant Components, 1974 Edition including Addenda through Winter 1975.

8.6 ASME Code,Section III, Rules for Construction of Nuclear Facility Components, 2013 Edition.

8.7 NRC Letter to APS, NRC Verbal Authorization to Arizona Public Service Accepting the Provisions of Relief Request 76 for one Cycle of Operation dated April 27, 2025, ADAMS Accession Number ML25118A063 8.8 Palo Verde Unit 1 Pressurizer Lower Instrument Nozzle Replacement Section III Qualification, Document Number 32-9388449-002. The evaluation was provided in APS letter dated May 22, 2025 (ADAMS Accession No. ML25142A406) 8.9 Corrosion Evaluation for Palo Verde Unit 1 Pressurizer Upper and Lower Instrument Nozzle Modification, Document Number 51-9384346-002. The evaluation was provided in APS letter dated May 22, 2025 (ADAMS Accession No. ML25142A406) 8.10 Palo Verde Unit 1 Pressurizer Upper and Lower Instrument Nozzle Modification Loose Parts Evaluation, Document Number 51-9385873-003. The evaluation was provided in APS letter dated May 22, 2025 (ADAMS Accession No. ML25142A406) 8.11 Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds, Document Number 32-9399131-001. [Attachments 1 (non-Proprietary, Document Number 32-9399132-001) and 3 (Proprietary)]

Enclosure Relief Request Number 77

Non-Proprietary Version Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds - Framatome Non-Proprietary, Document Number 32-9399132-001

Page 1 of 38 CALCULATION PACKAGE File No.: 2551740.303NP Project No.: 2551740 Quality Program Type:

Nuclear Commercial Framatome Document No.: 32-9399132-001 PROJECT NAME:

Framatome Palo Verde Repair of Pressurizer Lower Instrument Nozzles CONTRACT NO.:

1024054690 CO 002 Line 30, CO 003 Line 40 CLIENT:

Framatome Inc.

PLANT:

Palo Verde Nuclear Generating Station, Unit 1 CALCULATION TITLE:

Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds

- Framatome Non-Proprietary Document Revision Affected Pages Revision Description Project Manager Approval Signature & Date Preparer(s) &

Checker(s)

Signatures & Date 0

1 - 20 A A-2 B B-16 Computer Files Initial Issue Richard Mattson 10/8/2025 Preparer:

Charles Fourcade 10/8/2025 Reviewers:

Adam Butzin 10/8/2025 Dilip Dedhia 10/8/2025 Nathaniel Cofie 10/8/2025



Controlled Document Framatome Inc.

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 2 of 38 F0306-01R4 Document Revision Affected Pages Revision Description Project Manager Approval Signature & Date Preparer(s) &

Checker(s)

Signatures & Date 1

15 Corrected Typographical Error Richard Mattson 10/27/2025 Preparer:

Richard Mattson 10/27/2025 Reviewer:

Charles Fourcade 10/27/2025 Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 3 of 38 F0306-01R4 Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 4 of 38 F0306-01R4 Table of Contents 1.0 OBJECTIVE................................................................................................................6 2.0 TECHNICAL APPROACH........................................................................................... 6 2.1 Crack Growth Rate.......................................................................................... 7 2.1.1 Carbon Steel Fatigue Crack Growth................................................................ 7 3.0 ASSUMPTIONS..........................................................................................................8 4.0 DESIGN INPUTS........................................................................................................9 5.0 FRACTURE MECHANICS ANALYSIS........................................................................ 9 5.1 Allowable Stress Intensity Factors................................................................ 10 6.0 CRACK GROWTH CALCULATION.......................................................................... 11

7.0 CONCLUSION

S........................................................................................................12

8.0 REFERENCES

..........................................................................................................13 COMPUTER FILES.......................................................................................A-1 FLAW STABILITY EVALUATION..................................................................B-1 Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 5 of 38 F0306-01R4 List of Tables Table 1. Bounding Thermal Transients Analyzed [9]........................................................... 14 Table 2. Insurge/Outsurge Transients [10]........................................................................... 15 Table 3. Stress Intensity Factors used for Crack Growth Evaluation................................... 16 Table 4. Stress Intensity Factors used for Flaw Stability Evaluation.................................... 16 Table 5. Final Crack Growth Results at 80 Years................................................................ 17 List of Figures Figure 1. Flaw Sizes Evaluated............................................................................................ 18 Figure 2. Crack Tip Numbers............................................................................................... 19 Figure 3. Illustration of Allowable KIc Distribution as a Function of Temperature for the Leak Test Transient............................................................................................. 20 Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 6 of 38 F0306-01R4 1.0 OBJECTIVE The lower instrument nozzles on the Palo Verde Nuclear Generating Station, Unit 1, pressurizer were installed using Alloy 82/182 weld material. As a result, these locations are susceptible to Primary Water Stress Corrosion Cracking (PWSCC). In response, Arizona Public Service has decided to develop a pad-based repair for these locations.

The mitigation/repair for these nozzles involves machining the original nozzles to a point that is inside the pressurizer vessel outside surface, and the installation of a welded pad on the outside surface of the vessel, followed by the installation of a new nozzle on the welded pad. The resulting repair reestablishes the pressure boundary at the outside surface of the vessel and leaves the original J-groove weld in place. The designs have been previously prepared by Framatome, Inc. [1].

The objective of this calculation package is to perform a fracture mechanics evaluation, including crack growth for the remainder of plant life and life extension. The crack growth due to fatigue crack growth (FCG) is determined for a series of postulated flaws, with the initial flaw encompassing the entire J-groove weld and butter, and subsequent self-similar flaws extending into the vessel base metal, subjected to stresses caused by bounding thermal transients, weld residual stress, internal pressure, and piping loads. This postulated initial flaw ensures that any pre-existing indications in the J-groove weld are bounded. An initial check of the linear elastic fracture mechanics (LEFM) allowable is performed, followed by a more detailed elastic-plastic fracture mechanics check of the final flaw size. A fatigue crack growth evaluation is performed, using the pc-CRACK [4] program, to establish the maximum growth into the vessel material.

2.0 TECHNICAL APPROACH The technical approach includes the following steps:

1.

Assume conservatively that the Alloy 82/182 J-groove weld to be completely cracked since it is susceptible to PWSCC. This is then the initial postulated flaw.

2.

Perform fatigue crack growth into the low alloy ferritic vessel steel material starting with the initial postulated flaw.

3.

Determine the allowable flaw size in the ferritic vessel steel material.

4.

Determine the safe operating period from Steps 2 and 3 (how long it would take the initial postulated flaw to reach the allowable flaw size by fatigue crack growth).

ASME Code,Section XI, Nonmandatory Appendix A, Subarticle A-5200 [5] specifies requirements for flaw growth. The following discussion describes how the specific flaw growth analysis was performed, including crack modeling, growth durations, and final flaw acceptance criteria in the ferritic pressurizer base metal.

The stress intensity factor (K) distributions due to the various applied loads are computed using the finite element model developed in Reference [7], using the ANSYS finite element software [8]. Ks are developed as functions of postulated crack depth and linearly superimposed from the various operating states to obtain the total applied K. The stress intensity factors for the thermal transients, pressure, and weld residual stress are used in this calculation. Likewise, the stress intensity factors for Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 7 of 38 F0306-01R4 weld residual stress are determined using the stresses developed in Reference [3]. The K values are then used for the crack growth and flaw stability analyses.

An initial check of flaw acceptance using LEFM is performed based on the criteria in ASME Code,Section XI, Subarticle IWB-3600 [5] and the calculated stress intensity factors. The combined stress intensity factors are compared with the acceptance criteria discussed in Section 5.1. As the LEFM analysis did not show acceptability, an elastic-plastic fracture mechanics (EPFM) analysis was performed as documented in Appendix B.

The stress intensity factor range, K, for each transient is obtained to calculate the crack growth due to FCG. As self-similar crack growth is assumed (see Figure 1), it is appropriate to use an average stress intensity factor for each crack front. Additionally, flaw stability is performed to satisfy Section XI allowable crack size requirements.

The crack growth laws used for the crack growth calculation are discussed in Section 2.1.

22.1 Crack Growth Rate Crack growth is due to fatigue crack growth only, as the entire Alloy 82/182 weld is assumed cracked, and therefore, crack growth is limited to only the ferritic vessel material, as described below.

2.1.1 Carbon Steel Fatigue Crack Growth The FCG rate in ppc-CRACK [4] is based on ASME Code,Section XI, Nonmandatory Appendix A, Subarticle A-4300 [5], and the resulting equation for the crack growth rate is as follows:

n I

K C

dN da

)

(

0 '

(1) where:

n

WKHVORSHRIWKHORJGDG1YHUVXVORJ.I)

C0

= a scaling constant

IRU.I .th

.th WKUHVKROG.I

= 5.0 for R < 0

= 5.0(1-5IRU5

R

= R ratio (Kmin/Kmax)

The limitations in 10CFR50.55(a) [6] on R ratios are not applicable for this configuration, as there are no negative R ratios.

Reference fatigue crack growth behavior of material exposed to light-water reactor environments is JLYHQE\(TXDWLRQXVLQJ.I = Kmax - Kmin. If Kmin is equal to or less than zero, R = 0 is used.

(a) )RUORZ.I values, n = 5.95 and C0 = 1.02 x 10-12 S where S is given by S = 1.0

5

= 26.9 R - 5.725 (0.25 < R < 0.65)

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 8 of 38 F0306-01R4

= 11.76

5

(b) )RUKLJK.I values, n = 1.95 and C0 = 1.01 x 10-7 S where S is given by S = 1.0

5

= 3.75 R + 0.06 (0.25 < R < 0.65)

= 2.5

5

The applicable set of material parameters n and C0 LVGHWHUPLQHGE\FDOFXODWLQJWKH.I at which the two curves intersect. This is given by:

.I = 17.74

5

.I = 17.74 [(3.75R + 0.06) / (26.9R - 5.725)]0.25 (0.25 < R < 0.65)

.I = 12.04

5

,IWKHUDQJHRIDSSOLHGVWUHVVLQWHQVLW\IDFWRULVORZHUWKDQWKLVYDOXHWKHORZ.I parameters apply; RWKHUZLVHWKHKLJK.I parameters are used. The rate constant, C0, produces fatigue crack growth UDWHVLQXQLWVRILQFUDFNJURZWKF\FOHZKHQ.I is in units of NVL¥LQFor the purpose of crack growth into the vessel base metal, the flaw is postulated to initiate at the weld butter/vessel interface which is at a depth of Flaw 1 [

] shown in Figure 1.

3.0 AASSUMPTIONS x

The location of the flaw in the original J-groove weld is assumed to align with the axial symmetry plane of the vessel shell as the hoop stresses from internal pressure will be dominant along this direction. The weld residual stresses and the thermal transient stresses will be fairly uniform around the circumference of the J-groove weld, based on prior experience of similarly configured nozzles. The flaws in the base metal are therefore modeled on the axial symmetry plane of the vessel shell.

x Thermal couples are applied across the narrow gaps between the vessel shell and the outside surfaces of the nozzles. This is a typical method of simulating the heat transfer between two surfaces in close proximity which do not experience any significant flow, thereby allowing the temperature to transfer seamlessly between the components.

x The initial flaw size for the crack growth evaluation is assumed equal to Flaw 1, which extends slightly beyond the entire J-groove weld and butter.

x Due to the explicit modeling of the flaws in ANSYS, initial review of the crack growth results shows that the flaw growth is essentially self-similar, e.g., growth similar in all directions, and therefore average stress intensity factors along each crack front are calculated and used for the crack growth evaluation.

x For the fatigue crack growth, the transients are assumed to occur in the order given in Table 1.

Use of the order in the table has a negligible effect on the final results as the analysis occurs over a large number of years and the crack growth due to each transient is small.

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 9 of 38 F0306-01R4 x

The instrument nozzle piping loads [1] were determined to be negligible (producing maximum K values less than 0.1 ksi-¥in), and are therefore neglected from the analyses (the total stress and K results are several orders of magnitude larger) as they contribute insignificantly to the conclusions of both the crack growth and the flaw stability evaluations.

4.0 DDESIGN INPUTS The following design inputs are used in the evaluation:

x The geometry/model, including the crack tips, was developed in Reference [7].

x The transients are listed in References [9] and [10], and the heat transfer coefficients to be applied are specified in Reference [2]. The transients are defined in Table 1 and Table 2. The transients are defined as shown below:

x The weld residual stress was determined in Reference [3] and is used to calculate the K values herein.

5.0 FFRACTURE MECHANICS ANALYSIS The stress intensity factors (Ks) for the thermal transients and pressure loads are determined herein, using the finite element model created in Reference [7], along with the transients defined in Reference

[9] and Reference [10]. The Ks due to weld residual stress were determined using the model and Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 10 of 38 F0306-01R4 stress results contained in Reference [3]. The flaw dimensions are shown in Figure 1, with crack tip numbering provided in Figure 2. For the crack growth calculations, the thermal transient Ks are added to the weld residual stress Ks and pressure Ks for all transient times, and then averaged along each crack front. Then, the minimum and maximum average K values are determined from all the transient time steps. The averaged Ks that are used in the crack growth evaluation are shown in Table 3.

Per ASME Code,Section XI, Appendix A, Subarticle A-5200, crack growth evaluations need not include Emergency/Faulted Conditions; however, flaw stability must be evaluated for Emergency/Faulted Conditions in addition to Normal/Upset Conditions per ASME Code,Section XI, Subsubarticle IWB-3610. For each thermal transient, the maximum total K (thermal transient plus pressure) is determined for all times evaluated. The pressure is scaled according to the time points shown in Table 1 and Table 2. The stress intensity factors used in the flaw stability evaluation are tabulated in Table 4.

Table 5 includes the total stress intensity factors that result from adding the maximum and minimum averaged thermal transient and residual stress K values to appropriately factored pressure values, at the end of the 80-year evaluation period.

5.1 AAllowable Stress Intensity Factors The material fracture toughness (KIa and KIc), given in Section XI, Nonmandatory Appendix A, Subarticle A-4200 of the ASME Code [5] for low alloy steels, is used to determine the allowable stress intensity factor. The equations for low alloy steel are used here for the SA-533, Grade B, Class 1 vessel material. The allowable stress intensity factors are calculated by taking KIc /2 for emergency/faulted conditions, and conditions where pressure is less than 20% of the Design Pressure, and KIc /10 for normal and upset conditions, per Section XI, Subsection IWB-3610 of the ASME Code [5].

The material fracture toughness, KIc is provided in ASME Code,Section XI, Nonmandatory Appendix A, Subarticle A-4200 [5], and given by:

KIc = 33.2 +20.734*EXP[0.02*(T-RTNDT)]

where T is the temperature in the flaw region at the current time in the analysis and RTNDT is the reference nil ductility temperature of the vessel material.

The RTNDT is specified as [

] for the Palo Verde pressurizer plate materials [12]. Using an upper bound RTNDT value of [

] and the equation for KIc, the allowable stress intensity factors are determined to be [

] (at temperatures exceeding approximately

[

] ) for normal/upset and emergency/faulted conditions (and conditions for which pressure is less than 20% of Design Pressure), respectively. Note that the KIc value is set to a maximum value of 200 ksi-¥LQ (slightly lower than the maximum value of 220 ksi-¥in shown in Figure A-4200-1 [5]) for the evaluations performed herein. Therefore, the allowable K decreases below [

] however it should be noted that once the pressure drops below 20% of Design Pressure (approximately 500 psi),

the allowable increases once again.

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 11 of 38 F0306-01R4 The determined allowable fracture toughness, KIc, is compared with the calculated stress intensity factors in Table 4.

The transient nature of the allowable fracture toughness is illustrated in Figure 3 for the leak test.

The normal/upset allowable K is shown in Figure 3, using an RTndt value of [

] as a function of temperature, for the leak test transient, which exhibits the allowable K transition point at low temperatures and pressure below ~500 psi (20% Design Pressure), which occurs around [

]

Therefore, the total applied K (weld residual+transient+pressure) should be examined at or near this point to determine the LEFM criteria is satisfied. The evaluation is discussed in detail in Appendix B of the calculation. As shown in Figure 3, the allowable at 70°F is approximately [

]

increases to [

] at approximately [

] (and ~500 psi), at which point the allowable immediately drops to [

] (due to the increase in safety factor from ¥2 to ¥10),

increasing to [

] at approximately [

] As shown in Table 4, there are numerous K values that exceed the allowable upper shelf LEFM KIc value of [

] for transient conditions at elevated temperatures (i.e, above upper shelf material properties), and thus an EPFM evaluation is performed in Appendix B to demonstrate flaw stability in the vessel base metal, for which upper shelf material properties may be justified, at elevated temperatures. Additionally, the emergency/faulted transient (Secondary Loss of Pressure) is evaluated, as shown in Appendix B.

The EPFM evaluation is shown to be acceptable, and the flaw is stable at the maximum flaw evaluated at the end of the 80-year evaluation period. Included in Appendix B is an evaluation to verify LEFM criteria is satisfied at low temperature (i.e., below the upper shelf), for the bounding low temperature Leak Test conditions occurring near a low of [

]

6.0 CCRACK GROWTH CALCULATION Stress intensity factors (Ks) at four flaw depths in the vessel base metal were calculated using finite element analysis (FEA) to develop a K vs. a (K as a function of crack depth) trend. The average K at each crack front is used as input for performing the FCG analyses. Since the K vs. a profile is used as input to pc--CRACK [4], the shape of the component is not relevant. The K vs. a values that are used in the crack growth calculations are shown in Table 3.

The FCG laws used in this calculation are discussed in Section 2.1, and this portion of the evaluation uses the Carbon Steel in LWR crack growth law for assessing growth in the vessel base metal. The sequence of events for FCG is shown in Table 1 and Table 2, and is computed on a yearly basis.

Although in FCG the transients should be considered in approximate chronological order per the ASME Code,Section XI, Nonmandatory Appendix C, Subarticle C-3000 [5], it is acceptable to analyze them in numerical order because the analysis occurs over a large number of years and the crack growth due to each transient is small. All of the Normal/Upset thermal transients are analyzed sequentially per Table 1 and Table 2, to determine K values for these transients. This approach is consistent with Subsubarticle C-3210 of the ASME Code,Section XI [5]. A unit 1,000 psi internal pressure evaluation is likewise evaluated for each of the four flaws, and it includes crack face pressure to the flaw surface, including part of the nozzle remnant as an additional conservatism.

Additionally, the weld residual stress at the cold condition (70°F) is extracted from Reference [3] and is applied to the crack faces to solve the Ks for each of the four flaws.

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 12 of 38 F0306-01R4 In FCG, the individual terms that constitute nominal Kmax and Kmin values are summarized in the following tabulations.

Kmax Kmin Kresidual@70 Kresidual@70 Kpressure max Kpressure min Kthermal transient max Kthermal transient min The individual Ks for nominal Kmax and Kmin are combined (summed) with all appropriate scale factors applied. The stress intensity factors for the transient total Ks (weld residual stress plus pressure plus thermal transient) are listed in Table 3.

Likewise, the maximum and minimum stress intensity factors for each thermal transient are also summarized in Table 3. The.LVFRPSXWHGE\WDNLQJWKHGLIIHUHQFHRIWKHUHVXOWLQJWRWDO.max and Kmin values. Note that Kresidual@70 LVDFRQVWDQWDQGWKHUHIRUHGRHVQRWFRQWULEXWHWRWKH.UDQJH

However, Kresidual@70 affects the value of the R-ratio, and is therefore included. All the K values are entered into pc--CRACK to calculate FCG.

Table 5 provides the last load block summary at the end of the evaluation period (at year 80). As shown in the table, the final crack depth is [

] equal to approximately [

] of growth after 80 years. The demonstration of flaw stability in the ferritic low alloy steel vessel material is contained in Appendix B. The evaluation documents that the EPFM evaluation is acceptable, and the flaw is stable at the final flaw depth resulting from the crack growth evaluation, at the end of an 80-year operating interval.

The supporting files are listed in Appendix A.

7.0 CCONCLUSIONS As required by ASME Code,Section XI, Subsubarticle IWB-3610 and Nonmandatory Appendix A, Article A-5000, fatigue crack growth evaluations were performed, using an initial flaw size equivalent to the full size of the Alloy 82/182 J-groove weld and weld butter. The crack growth duration obtained from the pc--CRACK analysis is documented in Appendix A computer files. As indicated in Table 5, the final flaw depth due to FCG is [

] equal to approximately [

] of growth after 80 years. This final flaw depth is slightly larger than that of Flaw 3 shown in Figure 1.

Stress intensity factors were calculated for the postulated J-groove weld flaws for the applicable transients, internal pressure, and weld residual stresses. As shown in Table 4, some stress intensity factors exceed ASME Code allowable limits using linear elastic fracture mechanics methods.

Therefore, an elastic-plastic fracture mechanics analysis is performed per Code Case N-749 [11] to take advantage of the ductility of the vessel shell material at the upper shelf temperature.

Furthermore, Code Case N-749 provides different safety factors on primary stress (internal pressure and mechanical piping loads, if applicable) and secondary stresses (thermal transient, and weld residual stress, if applicable). This is documented in Appendix B, which is supplemental to the main body of this calculation and shows that the postulated flaws in the vessel shell material remain stable Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 13 of 38 F0306-01R4 (will not suffer non-ductile failure), and are therefore deemed acceptable for the remainder of the design life (including the license renewal period) or an 80-year operating interval.

88.0 REFERENCES

1.

Framatome Design Specification No. 08-9383300-004, 04/24/2025, Palo Verde Unit 1 Pressurizer Upper and Lower Instrument Nozzle Modification, SI File No. 2451268.202P, Framatome Inc. Proprietary.

2.

Framatome Specification No. 51-9395124-000, Fracture Mechanics Analysis Inputs Transmittal for Framatome Document 32-9388449-000, SI File No. 2551740.203P.

3.

SI Calculation No. 2551740.302P, Rev. 0, Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Weld Repair Modification.

4.

pc-CRACK 5.0.0.1, Version Control No. 5.0.0.1, Structural Integrity Associates, Inc.,

July 11, 2024.

5.

ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 2013 Edition.

6.

Code of Federal Regulations, 10CFR50.55a, (xxviii) Section XI condition: Analysis of flaws, April 4th, 2022.

7.

SI Calculation No. 2551740.301P, Rev. 0, Finite Element Model Development of the Pressurizer Lower Instrument Nozzle Weld Modification.

8.

ANSYS Mechanical APDL (UP20170403) and Workbench (March 31, 2017), Release 18.1, SAS IP, Inc.

9.

SI Calculation No. 2000645.301, Rev. 2, Transient Loads Definition for Updated Surge Line Analyses.

10. SI Calculation No. 0900426.302, Rev. 0, Computation of Heat Transfer Loads in Pressurizer during Insurge/Outsurge Transients.

11. ASME Boiler and Pressure Vessel Code, Code Case N-749, Alternative Acceptance Criteria for Flaws in Ferritic Steel Components Operating in the Upper Shelf Temperature Range,Section XI, Division 1 and PVP2012-78190, Alternative Acceptance Criteria for Flaws in Ferritic Steel Components Operating in the Upper Shelf Range, July 2012.

12. Palo Verde Generating Station Design Input Transmittal, contains Document ID N001-0604-01052, Certified Material Test Reports for the Pressurizer Upper and Lower Heads, SI File No.

2451268.201.

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 14 of 38 F0306-01R4 Table 11. Bounding Thermal Transients Analyzedd [9))

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 15 of 38 F0306-01R4 Table 22. IInsurge/Outsurge Transients [10))

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 16 of 38 F0306-01R4 Table 33. SStress Intensity Factorss used for Crack Growth Evaluation Table 44. Stress Intensity Factorss used for Flaw SStability Evaluation Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 17 of 38 F0306-01R4 Table 55. FFinal Crackk Growthh Results at 880 Years Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 18 of 38 F0306-01R4 Figure 11. FFlaw Siizes Evaluated Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 19 of 38 F0306-01R4 Figure 22. CCrackk Tip Numbers

[

]

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page 20 of 38 F0306-01R4 Figure 33. IIllustration of Allowable KIc Distribution as a Function of Temperaturee for the Leak Test TTransient Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page A-1 of 38 F0306-01R4 COMPUTER FILES Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page A-2 of 38 F0306-01R4 File Name Description Mechanical and Thermal Analyses Files Inst_noz.INP ANSYS input file to create the base model [7].

FM_Inst_noz_PRESS.INP PRESS.INP PRESS_COMPONENTS.INP ANSYS input files to run the fracture mechanics evaluations for the unit 1,000 psi pressure analyses.

FM_Inst_noz.INP Crack_Nodes.inp ANSYS input files to create the fracture mechanics finite element model for the transient analyses.

MOMENT.INP FM_MOMENT_GEOM$.INP Crack_Nodes$.inp ANSYS input files to create fracture mechanics models and runs for the unit 1,000 in-lb moment analyses ($ = 1-4 for flaws 1 to 4).

RESID_FLAW$.INP

_GETSTR.INP STRESS2D.TXT ANSYS input files to extract the weld residual stress from the STRESS2D.RST file (Reference [3]), and to apply the loads, and solve the fracture mechanics evaluations for weld residual stress ($ = 1-4 for flaws 1 to 4).

TRANSIENTS.INP TRANS.MAC THM_COMPONENTS2.INP ANSYS input files to run the thermal transient fracture mechanics evaluations.

Faulted.inp ANSYS input files to perform the Loss of Pressure transient and extract K results (to support emergency/faulted EPFM evaluation).

AnTip.mac / AnTip.exe Macro to insert crack tip elements into the fracture models.

Post Processing and Results Files AnTip_KCALC.INP KCALC post-processing input file.

STR_#_FLAW$_KvT.CSV Formatted K result output, # = transient name; $ = 1-4 for flaws 1 to 4.

PRESS_FLAW$_KvT.CSV Formatted K result output for 1,000 psi pressure runs;

$ = 1-4 for flaws 1 to 4.

MOM_FLAW$_KvT.CSV Formatted K result output for 1,000 in-lb moment runs;

$ = 1-4 for flaws 1 to 4.

RESID_FLAW$_KvT.CSV Formatted K result output for weld residual runs;

$ = 1-4 for flaws 1 to 4.

STR_LOP_FLAW$_KvT.CSV Formatted K result output for Loss of Pressure transient;

$ = 1-4 for flaws 1 to 4.

FCG.pcf ppc-CRACK input file for crack growth.

FCG.rpt ppc-CRACK output file for crack growth.

K_Total_bottom_Noz.xlsm Excel spreadsheet containing summary of K results.

K_Total_Bottom_Noz_EPFM.xlsm Excel spreadsheet containing EPFM K summary.

2551740.303.N749.xlsx Exel spreadsheet containing EPFM evaluation.

PRESS_TEST.INP Unit pressure test for LEFM test.

Getpath.txt, GETPATH_PRESS.TXT,GETPATH_RESID.TXT Stress path definition for LEFM test.

Resid_K_test.pcf Resid_K_test.rpt ppc-CRACK input/out file for LEFM K checking.

PRESS_TEST_MAP_P$.CSV; STRESS2D_MAP_P$.CSV Stress results for LEFM test.

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-1 of 38 F0306-01R4 FLAW STABILITY EVALUATION Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-2 of 38 F0306-01R4 B.1 INTRODUCTION The analyses performed in the main body of this calculation addresses the fatigue crack growth (FCG) for a series of postulated flaws in the pressurizer Alloy 82/182 J-groove welds and the adjacent ferritic pressurizer vessel, at the Palo Verde Nuclear Generating Station. Linear elastic fracture mechanics (LEFM) analyses were performed to determine flaw acceptance for normal/upset and emergency/faulted conditions.

Multiple crack depths were considered, as modeled in the supporting calculation [B-4]. The depth and length for the initial postulated crack was chosen as the full extent of the susceptible Alloy 82/182 material of the J-groove weld and weld butter, along with a portion of the original nozzle remnant. The subsequent crack depths are near-evenly distributed and exhibit a shape that matches the first crack.

For the LEFM evaluation in the main body of this calculation, tensile residual stresses are present in the original J-groove weld, resulting from the original welding processes. Thus, in accordance with the requirements of ASME Code,Section XI, residual stresses, operating pressure, and thermal transients were included in the LEFM evaluation. This approach has led to difficulty in demonstrating acceptability of the postulated J-groove weld cracking into the low alloy steel vessel base material, as demonstrated by the LEFM analyses.

LEFM techniques for nuclear plants were developed primarily for the irradiated RPV beltline region and low temperature applications of carbon and low alloy steel, in which the material exhibits little or no ductility.

However, the J-groove weld cracking occurs predominantly at operating temperatures, where considerable ductility exists. It is therefore the objective of this Appendix to perform an elastic-plastic fracture mechanics (EPFM) analysis to demonstrate acceptability of the postulated flaws, taking credit for the ductility that exists. All postulated flaw depths are evaluated to ensure that the EPFM margins are achieved during the evaluation period considered herein.

ASME Code,Section XI, Subsection IWB-3612 [B-1] provides the LEFM acceptance criteria for ferritic steel components 4 inches or greater in thickness, which is applicable to the pressurizer shell evaluated herein.

In lieu of the acceptance criteria provided in IWB-3612, this evaluation will use ASME Code Case N-749

[B-5], which provides acceptance criteria for flaws in ferritic steel components/vessels operating in the upper shelf temperature range. Code Case N-749 provides different safety factors on primary stress (internal pressure and mechanical piping loads, if applicable) and secondary stresses (thermal transient, and weld residual stress, if applicable). Details of the use of Code Case N-749 are described in Section B.3.

B.2 ASSUMPTIONS The following assumption was made in this calculation:

x The CVN values cited in Reference [B-8] and Reference [B-9] were actually said to be applicable to the pressurizer bottom head, and considered representative for the upper shell as well. As a further conservatism herein, the CVN values will be assumed to be 50 ft-lbs (a lower bound value required for all plate materials), therefore providing additional margin in the stability evaluation.

B.3 METHODOLOGY The selected Code path is the use of ASME Code Case N-749 [B-5]. This Code Case addresses alternate acceptance criteria for flaws in ferritic steel components operating in the Upper Shelf temperature range.

This Code Case has been Conditionally approved in Regulatory Guide 1.147 [B-6], and the conditions do apply herein.

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-3 of 38 F0306-01R4 Code Case N-749 lays out the acceptance criteria as follows:

Section 2(a) [B-5]

Applicability of these acceptance criteria is limited to ferritic steel components on the upper shelf of the Charpy energy curve when the metal temperature exceeds the upper shelf transition temperature, Tc.* Tc is defined as follows:

Tc = RTNDT + 105°F This was conditioned by the USNRC in their approval of the Code Case in Regulatory Guide 1.147, Rev. 19 [B-6]:

Section 3.1(a) [B-5]

Code Case N-749 provides different safety factors on primary stress (internal pressure and mechanical piping loads, if applicable) and secondary stresses (thermal transient, and weld residual stress, if applicable). For normal and upset conditions, J shall be evaluated at loads equal to 2.0 times the primary loads and 1.0 times the secondary loads, including thermal stresses. The applied J shall be less than or equal to the J-integral of the material at a ductile crack extension of 0.10 in.

(2.5 mm).

Section 3.1(b) [B-5]

For emergency and faulted conditions, J shall be evaluated at loads equal to 1.5 times the primary loads and 1.0 times the secondary loads, including thermal. The applied J shall be less than or equal to the J-integral of the material at a ductile crack extension of 0.10 in. (2.5 mm).

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-4 of 38 F0306-01R4 Section 3.1(c) [B-5]

The primary stress limits of NB-3000 shall be satisfied, assuming a local area reduction of the pressure retaining membrane equal to the area of the end-of-evaluation-period flaw, as determined using the flaw characterization rules of IWA-3000. For emergency and faulted conditions, J shall be evaluated at loads equal to 1.5 times the primary loads.

If these four conditions are met, no further qualification is required.

BB.4 MATERIAL PROPERTIES The material properties required for the Code Case N-749 evaluation are the basic material properties and the J-R curve.

B.4.1 Basic Material Properties The vessel shell is fabricated from SA-533, Grade B, Class 1 material [B-4]. The following material properties are required for the EPFM analysis:

Youngs modulus, E = 25,800 ksi [B-2] E' = E/(1-v2) [B-1, K-1300] = 28,352 ksi Poissons ratio, v = 0.3 [B-2]

Yield stress, YS = 41.5 ksi [B-2]

Ultimate tensile stress, US = 80.0 ksi [B-2]

Flow stress, Vf = 60.75 ksi (YS+US)/2 The values for the Youngs modulus and yield stress are taken at the operating temperature of 653qF from the ASME Code,Section II, Part D [B-2]. Note that the flow stress specified in ASME Code,Section XI, Nonmandatory Appendix K (85 ksi) [B-1,K-1300] is conservatively used, as this produces a conservative allowable Jmat value.

B.4.2 J-R Curve Figure B-1, obtained from Reference [B-7], presents J-T materials curves for irradiated and unirradiated nuclear vessel steels at various upper shelf Charpy V-notch (CVN) energy levels (in joules). The results show a rough correlation, in that higher J-T curves are generally exhibited for higher Charpy V-notch energy levels. A correlation curve has been developed in Reference [B-7] (see Figure B-2 and Figure B-3) between Charpy V-notch energy and the parameters C and m of a J-R curve power law fit of the following form:

- &Dm [B-3]

The resulting Jmat allowable is calculated, using D = 0.1 crack extension, and the C and m terms derived using the tests of the Palo Verde pressurizer base material, conducted in the transverse orientation. These exhibited Charpy V-notch energy levels ranging from a minimum of 98 ft-lbs up to a maximum of 119 ft-lbs at or near the upper shelf temperature [B-9] (measured at +50°F; the upper shelf temperature is most probably higher than this, and actual upper shelf energies applicable at plant operating temperature are thus expected to be much higher) [B-8]. Thus, the 98 ft-lbs value was used previously (in separate fracture mechanics evaluations outside of this one) as a conservatively low estimate of the Charpy V-notch upper shelf energy (CVN) level for the Palo Verde pressurizer base material. As a further conservatism herein, the Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-5 of 38 F0306-01R4 CVN values will be assumed to be 50 ft-lbs (a lower bound value required for all plate materials), therefore providing additional margin in the stability evaluation.

For the parameters C and m, which are dependent upon the CVN value and flow stress (85 ksi), Figure B-2 and Figure B-3 are used to determine values of the coefficient "C" and the exponent "m" for the power law J-R curve fit. Using a lower bound CVN of 50 ft-lbs results in C = 1.34 and m = 0.268, with a corresponding Jmat = 0.724 in-kips/in2 at 0.1 crack extension.

B.5 APPLIEDD J CALCULATIONS B.5.1 Bounding Loading Conditions Analyses for J applied are performed in accordance with the ASME Code,Section XI, Code Case N-749 [B-5]. This allows EPFM J-Integral estimates to be developed from LEFM K calculations, which were already calculated in the main body of this calculation. As seen in the main body of the calculation, several transient conditions do not meet the LEFM allowable criteria. Referring to Table 4, it can be seen that the

[

] transient is bounding, having a maximum K total in excess of [

] (corresponding to a maximum pressure of [

] as shown in Table 1 of the calculation). Therefore, the K values for the Trip transient are used to perform the Service Level A/B stability calculation.

The Service Level C/D stability calculation is performed for the Secondary Loss of Pressure transient, per Reference [B-10]. The transient is defined as a drop in temperature from [

]

followed by an upward ramp back to [

] The pressure drops from [

] in the first [

] followed by a return to [

]

B.5.2 Stress Intensity Factor Distributions The linear elastic fracture mechanics analyses described in the main body of the calculation include stress intensity factors (K) due to thermal transients, operating pressures, and weld residual stress. These are extracted along the crack fronts for each flaw configuration, crack depth (a), and load type. The crack depth (a) is obtained from the finite element model. The maximum K across the crack front due to residual stress, pressure, and thermal transient stress is obtained from the base calculation (see Table 3 of the main body of the calculation). The values for K for unit pressure are scaled separately to the bounding values for Service Levels A/B (Normal/Upset), and Service Levels C/D (Emergency/Faulted) conditions. The Ks due to primary stress (KIp) include maximum operating pressure alone, whereas the Ks due to secondary stress (KIt) include thermal transient stresses alone (since weld residual stress is not required per Code Case N-749). The stress intensity factors are combined in 2551740.303.N749.XLSX and shown in Table B-1.

B.5.3 J Integral Calculation The Code Case N-749 procedure involves the calculation of a plastic zone corrected crack size for small scale yielding from elastically calculated K-values, in accordance with the following:

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-6 of 38 F0306-01R4

=

[B-5, Basis document, Page 9]

Per Code Case N-749, Section 4.1, no structural factors are applied when calculating the plastic zone correction factor.

The applied J integral is then calculated from the effective crack-tip stress intensity factor with a plastic-zone correction (Kleff) as follows:

= 1 +

/

[B-5, Basis document, Page 9]

where,

=

[B-5, Basis document, Page 2]

=

[B-5, Basis document, Page 9]

=

[B-5, Basis document, Page 9]

where J is in in-lb/in2, K is in ksi-in, and E' is in ksi.

K'total = K'Ip + K'It B.6 RESULTS AND DISCUSSION 1st Condition: The initial RTNDT for the SA-533, Grade B, Class 1 base material is shown to be [

] or lower for the various Palo Verde pressurizer top head plates [B-9]. As the evaluated pressurizer vessel is not subject to radiation embrittlement, there is no need to consider any adjusted reference temperature (ART), and therefore the RTNDT is used directly.

Taking direction from the Conditional approval [B-6] with the applicable RTNDT of [

]

Tc = 154.8°F + 0.82 x RTNDT

= 154.8°F + 0.82 x [

]

= [

]

The operating temperature for the pressurizer vessel is 653°F. The limiting upper shelf EPFM conditions which have been evaluated herein included the Trip transient (ranges between [

] for Normal/Upset Conditions. The bounding Emergency/Faulted Conditions evaluated herein include the Loss of Secondary Pressure transient [

] [B-10]. However, the Heatup, Cooldown, Design Hydro Test, and Leak Test transients all have temperature ranges which fall below the Tc value, and Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-7 of 38 F0306-01R4 therefore require additional evaluation. Hence, the secondary requirement of the Conditional approval specifies the minimum temperature below which LEFM must be applied.

Tc1 = 95.36°F + 0.703 x RTNDT

= 95.36°F + 0.703 x [

]

= [

]

The Heatup, Cooldown, and Design Hydro Test transients all have temperature ranges which fall below the Tc1 value of [

]

The heatup and cooldown transient conditions do not bound since the temperature is approximately

[

] at 20% Design Pressure, or 500 psi approximately (at which point the safety factor switches from 10 to 2, as discussed in Section 5.1 of the calculation).

The Hydro/Leak Test temperature is only [

] and therefore becomes the bounding LEFM condition, for which the cold normal/upset allowable is KIc/10 at or above approximately [

]

The total K values are evaluated for all flaw sizes at or below this temperature to check that the LEFM allowable criteria is satisfied. Due to the large K contribution from weld residual stress, some of the total K values exceed the allowable K based on LEFM brittle fracture criteria.

To address the brittle fracture concern at low temperature, a simplified evaluation is performed. Since the majority of the K at low temperature is due to the weld residual stresses, which are decreasing away from the J-groove weld region, the total K (pressure + thermal transient + piping loads + weld residual stress) decreases as the flaw grows deeper through the base metal, towards the OD/pad surface. To evaluate this analytically, a simplified representative flaw model is used, with mapped stresses applied to determine the K values for the limiting cold brittle fracture conditions. The J-groove flaw is best represented using the LEFM Nozzle Corner Crack flaw model (Model 601), available within the pc--CRACK [B-11] fracture mechanics software, as shown below in Figure B-4.

The through-wall hoop stress is extracted from the unit pressure analysis and the weld residual stress analysis (at cold conditions), and is aligned approximately with the stress path direction shown in Figure B-4. For the Leak Test thermal transient, the bounding thermal transient K value of [

] at

[

] into the transient is used (shown in Figure B-5), corresponding to approximately

[

] and a pressure of approximately [

] which is the inflection point between the lower and the upper bound allowable K value (at approximately 20% Design Pressure) as shown in Figure B-6. The total K is plotted as a function of through-wall crack depth, along with allowable K in Figure B-6. As shown in the figure, the applied total K drops below the allowable K (determined at [

] with a structural factor of 10 (i.e., KIc/10) at approximately [

] depth. Therefore, the crack would be expected to arrest prior to growing through the full vessel wall thickness.

Therefore, the LEFM criteria is met as the flaw grows through the vessel wall (i.e., the flaw arrests, even considering an assumed brittle fracture material behavior at low temperature).

Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-8 of 38 F0306-01R4 2nd and 3rd Conditions: Using the equations listed above, the plastic zone size adjusted K and J integral values are calculated in 2551740.303.N749.XLSX and provided in Table B-1. The applied J values with appropriate structural factors discussed in Section B.3 are less than the value of J @ 0.1 extension, thus showing sufficient margin to instability.

4th Condition: The primary stress limits of ASME Code,Section III, NB-3000 shall be satisfied, as demonstrated in the ASME Code qualification (performed by Framatome), which shall verify that the design of the weld repair pad meets the Code requirements without consideration for the flawed components.

The J-R curve used for the allowable Jmat at 0.1 crack extension is obtained using the lower bound 50 ft-lbs CVN, and the general derivations per Reference [B-3]. Bounding transient conditions were selected based on the most limiting pressure and thermal stress and K results. The bounding Normal/Upset Service Level A/B transient evaluated is the Surge 7 transient (at [

] pressure) for Flaws 1-3, and the Trip 1 transient (evaluated at [

] pressure) for Flaw 4. The Secondary Loss of Pressure transient was evaluated for the Emergency/Faulted transient [B-10]. All events are evaluated using the limiting J value

[

] calculated in Section B.4. As shown in the tables, the EPFM evaluation shows acceptability for all flaw depths, using the most limiting J criteria.

The final results are provided in Table B-1, which shows that the EPFM criteria are satisfied for all flaw sizes evaluated.

B.7 CONCLUSIONS Based on the EPFM analyses presented herein, it is concluded that potential cracks in the low alloy steel pressurizer are acceptable for an 80-year operating interval. The analyses assume worst case flaws encompassing the entire remaining J-groove weld and weld butter and extending well into the vessel base metal. EPFM analyses are performed using structural factors from Code Case N-749 (see Section B.2).

The results indicate that the value of J applied, including appropriate structural factors, is below the J-integral of the material toughness, at a ductile crack extension of 0.1 (per Reference [B-3]) for flaws up to the evaluated 80-year crack size, in accordance with the flaw evaluation principles of ASME Code Case N-749.

Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-9 of 38 F0306-01R4 B.8 REFERENCES B-1.

ASME Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 2013 Edition.

B-2.

ASME Boiler and Pressure Vessel Code,Section II, Part DProperties, 2013 Edition.

B-3.

F. Loss, F. J., Menke, B. H., Hiser, A. L., and Watson, H. E., J-R Curve Characterization of Irradiated Low-Shelf Nuclear Vessel Steels, Elastic-Plastic Fracture: Second Symposium, Volume II--Fracture Resistance Curves and Engineering Applications, ASTM STP 803, C. F. Shih and J. P.

Gudas, Eds., American Society for Testing and Materials, 1983, pp. II-777-II-795.

B-4.

SI Calculation No. 2551740.301P, Revision 0, Finite Element Model Development of the Pressurizer Lower Instrument Nozzle Weld Modification.

B-5.

ASME Boiler and Pressure Vessel Code, Code Case N-749, Alternative Acceptance Criteria for Flaws in Ferritic Steel Components Operating in the Upper Shelf Temperature Range,Section XI, Division 1 and PVP2012-78190, Alternative Acceptance Criteria for Flaws in Ferritic Steel Components Operating in the Upper Shelf Range, July 2012.

B-6.

Regulatory Guide 1.147, Revision 19, INSERVICE INSPECTION CODE CASE ACCEPTABILITY, ASME SECTION XI, DIVISION 1.

B-7.

NUREG-0744, Vol. 1 & 2, Rev. 1, Resolution of the Task A-11 Reactor Vessel Materials Toughness Safety Issue, Appendices B-K, Division of Safety Technology, Office of Nuclear Reactor Regulation, U.S. Nuclear Regulatory Commission, Washington, D.C. 20555, October, 1982.

B-8.

E-mail Transmission from Ram Indap (APS) to Pete Riccardella (SI), Valve SI 56 and the Bottom Head Charpy Values, dated May 04, 2004, SI File No. PV-04Q-218.

B-9.

Palo Verde Generating Station Design Input Transmittal, contains Document ID N001-0604-01052, Certified Material Test Reports for the Pressurizer Upper and Lower Heads, SI File No.

2451268.201P.

B-10.

APS SDOC 13-MN725-A00945, Rev. 008, Design Specification No. 0900426.403, Revision 5, January 2013, Palo Verde Nuclear Generating Station Units 1, 2 and 3 Pressurizer Assembly, SI File No. 0900426.403.

B-11.

pc-CRACK 5.0.0.1, Version Control No. 5.0.0.1, Structural Integrity Associates, Inc.,

July 11, 2024.

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-10 of 38 F0306-01R4 Figure B-1: J-T Diagram for Several Reactor Vessel Steels and Welds Showing Rough Correlation with Charpy V-notch Upper Shelf Energy (From Reference B-3, page D-21)

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-11 of 38 F0306-01R4 Figure B-2: Correlation of Coefficient C of Power Law J R-curve Representation with Charpy V-notch Upper Shelf Energy (From Reference B-3, page D-24)

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-12 of 38 F0306-01R4 Figure B-3: Correlation of Exponent m of Power Law J R-curve Representation with

&RHIILFLHQW&DQG)ORZ6WUHVVo (From Reference B-3, page D-24)

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-13 of 38 F0306-01R4 Figure B-4: Nozzle Corner Crack Flaw Model and Extracted Stress Path Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-14 of 38 F0306-01R4 Figure B--5:: LEFM Stress Intensity Factors for Leak Test (Thermal Stress Only)

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File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-15 of 38 F0306-01R4 Figure B--6:: LEFM Stress Intensityy Factor Results for Leak Test Controlled Document

File No.: 2551740.303NP Revision: 1 Framatome Document No.: 32-9399132-001 Page B-16 of 38 F0306-01R4 Table B--1:: Stress Intensity Factor Summary Controlled Document

Enclosure Relief Request Number 77

Non-Proprietary Version Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Modification, Non-Proprietary Document Number 32-9398249-000

Page 1 of 42 CALCULATION PACKAGE File No.: 2551740.302NP Project No.: 2551740 Quality Program Type:

Nuclear Commercial Framatome Document No.: 32-9398249-000 PROJECT NAME:

Framatome PV Repair of Pressurizer Lower Instrument Nozzles CONTRACT NO.:

P.O. 1024054690 CLIENT:

Framatome Inc.

PLANT:

Palo Verde Nuclear Generating Station, Unit 1 CALCULATION TITLE:

Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Weld Modification - Non-Proprietary Document Revision Affected Pages Revision Description Project Manager Approval Signature & Date Preparer(s) &

Checker(s)

Signatures & Date 0

1 - 40 A A-2 Initial Issue Richard Mattson 9/18/25 Ryan Keller 9/18/25 Richard Bax 9/18/25 Richard Mattson



Controlled Document Framatome Inc.

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 2 of 42 F0306-01R4

&RQWUROOHG'RFXPHQW

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 3 of 42 F0306-01R4 Table of Contents 1.0 OBJECTIVE................................................................................................................ 5 2.0 DESIGN INPUTS........................................................................................................ 5 2.1 Finite Element Model....................................................................................... 5 2.2 Material Properties.......................................................................................... 5 2.3 Normal Operating Conditions.......................................................................... 5 3.0 ASSUMPTIONS.......................................................................................................... 5 4.0 TECHNICAL APPROACH........................................................................................... 7 4.1 Welding Heat Input Simulation........................................................................ 8 4.2 Weld Nugget Simulation.................................................................................. 9 4.3 Post Weld Heat Treatment.............................................................................. 9 4.3.1 Creep Properties........................................................................................... 10 5.0 RESIDUAL STRESS SIMULATION PROCESS....................................................... 11 5.1 Cladding........................................................................................................11 5.2 Weld Butter.................................................................................................... 11 5.3 Post Weld Heat Treatment............................................................................ 11 5.4 Original J-Groove Weld................................................................................. 11 5.5 Shakedown Evaluation No. 1........................................................................ 11 5.6 1992 Weld Pad Repair.................................................................................. 12 5.7 1992 Outer Sleeve Fillet Weld....................................................................... 12 5.8 1992 Nozzle-to-Pad Weld............................................................................. 12 5.9 Shakedown Evaluation No. 2........................................................................ 12 5.10 Current Repair Weld Pad.............................................................................. 13 5.11 Shakedown Evaluation No. 3........................................................................ 13 6.0 CALCULATIONS.......................................................................................................13 7.0 RESULTS OF ANALYSIS......................................................................................... 14

8.0 CONCLUSION

S........................................................................................................14

9.0 REFERENCES

..........................................................................................................15 COMPUTER FILES....................................................................................... A-1

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 4 of 42 F0306-01R4 List of Tables Table 1. Component Materials............................................................................................. 16 Table 2. Creep Properties.................................................................................................... 16 List of Figures Figure 1. Components Included in the Finite Element Model.............................................. 17 Figure 2. ANSYS 2-D Finite Element Model of Instrument Nozzle / Repair Pad for Weld Residual Stress Analysis (with Minimum Dimensions)........................................ 18 Figure 3. Weld Bead Distribution......................................................................................... 19 Figure 4. Applied Pressure Loading..................................................................................... 20 Figure 5. Example Applied Mechanical Boundary Conditions............................................. 21 Figure 6. Predicted Fusion Boundary for the Cladding........................................................ 22 Figure 7. Predicted Fusion Boundary for the Weld Butter.................................................... 23 Figure 8. Predicted Fusion Boundary for the Original J-Groove Weld................................. 24 Figure 9. Predicted Fusion Boundary for the 1992 Weld Pad Repair.................................. 25 Figure 10. Predicted Fusion Boundary for the Outer Sleeve Fillet Weld.............................. 26 Figure 11. Predicted Fusion Boundary for the Nozzle-to-Pad Weld..................................... 27 Figure 12. Predicted Fusion Boundary for the Weld Pad Repair......................................... 28 Figure 13. Post Cladding Hoop Stress at 70°F.................................................................... 29 Figure 14. Post Buttering Hoop Stress at 70°F.................................................................... 30 Figure 15. Hoop Residual Stress after PWHT at 70°F.......................................................... 31 Figure 16. Post Original J-Groove Weld Hoop Stress at 70°F.............................................. 32 Figure 17. Post Initial Operation Hoop Stress at 70°F......................................................... 33 Figure 18. Post 1992 Weld Repair Pad Hoop Stress at 70°F.............................................. 34 Figure 19. Post Outer Sleeve Fillet Weld Hoop Stress at 70°F............................................ 35 Figure 20. Post 1992 Nozzle-to-Pad Hoop Stress at 70°F................................................... 36 Figure 21. Post Secondary Operation Hoop Stress at 70°F................................................. 37 Figure 22. Post Weld Pad Repair Hoop Stress at 70°F....................................................... 38 Figure 23. Final Weld Pad Repair with New Nozzle Hoop Stress at 70°F........................... 39 Figure 24. Final Weld Pad Repair with New Nozzle Hoop Stress at NOC........................... 40

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 5 of 42 F0306-01R4 1.0 OBJECTIVE Framatome proposes to modify the pressurizer lower instrument nozzles at Palo Verde Nuclear Generating Station, Unit 1 with a repair that consists of removing the existing nozzle, applying an outside diameter weld pad, inserting a new nozzle, and applying a J-groove/fillet weld between the new nozzle and the weld pad to establish a new pressure boundary.

The objective of this calculation package is to perform a component specific weld residual stress analysis, using the ANYS finite element software [1], of the pad-based repair for the lower pressurizer instrument nozzles. The weld residual stress analysis includes the pressurizer shell cladding, the original J-groove weld butter, the original nozzle-to-pressurizer J-groove weld, the 1992 welded pad repair components, the 1992 repair outer sleeve, the 1992 nozzle-to-pad weld, installation of the new welded pad, and a simplified final configuration that includes the new nozzle.

The stress results from this analysis will be used in a separate fatigue crack growth calculation.

2.0 DESIGN INPUTS 2.1 Finite Element Model Figure 1 identifies the different components of the model. A bounding finite element model for the lower pressurizer instrument nozzles, including material properties, is obtained from Reference [2] (input file INST_NOZ_RES.INP). The resulting finite element model is shown in Figure 2. This finite element model utilizes the minimum dimensions (length and width) for the repair.

The Reference [2] model only includes the final configuration of the nozzle (i.e., the planned new weld pad and nozzle). The lower pressurizer instrument nozzles were previously modified in 1992 as part of an Alloy 600 PWSCC mitigation campaign. As part of the 1992 repair, the original Alloy 600 nozzle was removed, and the original borehole was oversized, removing portions of the original J-groove weld, weld butter, and pressurizer bottom head wall [12, Section 1.5]. The pre-repair geometry is reconstructed using the original design drawings [13, 14] to better capture the complete weld residual stress effects throughout the lifetime of the component.

2.2 Material Properties The material designations of the various components of the model are listed in Table 1. The temperature-dependent nonlinear material property values are also obtained from Reference [2] (or from Reference [15], in the case of Alloy 600). This analysis applies the multilinear isotropic (MISO) hardening material behavior available within the ANSYS finite element program, with modified stress-strain curves specifically developed for weld residual stress analysis. The MISO hardening material behavior has been shown to be suitable for weld residual stress analysis [3].

2.3 Normal Operating Conditions The full load operating pressure and maximum operating temperature for the pressurizer are 2,250 psia (2,235 psig) and 653°F, respectively [8, Table 5].

3.0 ASSUMPTIONS The following assumptions are used in the residual stress evaluation:

x Assumptions in Reference [2] related to the development of the finite element model (FEM) are applicable in this calculation.

x All welding simulations are assumed to be in an air environment. A convection heat transfer coefficient of 5.0 BTU/hr-ft2-°F at 70°F bulk ambient temperature is applied to simulate an air

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 6 of 42 F0306-01R4 environment on the inside and outside surfaces during the application of all welding processes.

The weld residual stress analysis assumes that the welding is done in the plant shutdown condition. Thus, it assumes uninsulated piping in stagnant air (on both the ID and OD surfaces) and uses recommended heat transfer values from the SI weld residual macro verification and validation package [3].

x A maximum interpass temperature of 350°F between the depositions of lumped weld bead nuggets is assumed for all welding processes. This is a typical interpass limit for production field welding, and specifically the repair pad welding.

x Since the welding parameters for the original nozzle-to-pressurizer J-groove weld and the subsequent 1992 weld pad and repair nozzle-to-weld pad J-groove weld are not available, a typical heat input of 28 kJ/inch, and a typical heat efficiency of 0.8, is assumed for these welds.

A typical heat input of 35 kJ/in is used for the new weld pad repair. These parameters are consistent with the analysis process in Reference [3].

x For a 2-D model, each individual weld bead is effectively treated as a complete ring of weld metal, which will conservatively increase the weld residual stress.

x The cladding is assumed to be fully deposited in a three-layer, single pass process. This will produce conservatively high weld residual stresses due to the shrinkage of three large single nuggets rather than the multiple weld passes of the actual cladding process. Note that the weld residual stresses in the pressurizer base metal are largely mitigated by the PWHT process (compare Figure 14 and Figure 15).

x The cladding is assumed to be applied with the nozzle bore and the weld butter/J-groove weld preparation already in place, which is not consistent with the typical nozzle installation process (e.g., the bore hole and weld preparation would be applied after the cladding process.)

However, the weld residual stresses due to the application of cladding before or after the nozzle bore and original J-groove weld preparation should yield the same results. This is based on the fact that the original J-groove weld preparation is produced by machining away material, including the local cladding, which would remove the local weld residual stresses induced by the cladding that was removed. In addition, the application of the PWHT will further reduce variation due to the reduction and smoothing of the weld residual stresses.

x The weld butter is simplified as a three-layer multi-pass process, as shown in Figure 3. This approach is conservative as the actual weld usually consists of multiple weld beads and layers, and simplifying them into larger weld passes would yield larger regions of tensile residual stresses. Note that the application of the PWHT and hydrostatic pressure would further reduce variation due to the reduction and "smoothing" of the weld residual stresses.

x The original nozzle-to-pressurizer J-groove weld would have been involved in at least one hydrostatic test and a period of normal operation prior to the installation of the 1992 weld repair pad and the current weld repair pad. For this evaluation, only normal operation will be simulated prior to the installation of the 1992 weld repair pad and the current weld repair pad. The hydrostatic test following the original nozzle-to-pressurizer J-groove weld will tend to reduce weld residual stresses, thus excluding it is conservative.

x The 1992 weld pad repair and outer sleeve fillet weld would have also been involved in at least a period of normal operation prior to the installation of the current weld repair pad. It is considered unlikely that a hydrostatic test would have been performed following the 1992 repair.

Therefore, only normal operation will be simulated prior to the installation of the welded pad.

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The sacrificial plug is not modeled because the sacrificial plug material is removed for the installation of the replacement nozzle. The plug itself will be installed with only a limited number of tack welds, which would generate only very localized weld residual stresses. The plug is then removed by machining, further removing any residual stresses from the plug installation.

Therefore, the sacrificial plug should have minimal effect on the final weld residual stress results.

x There is evidence of a 1987 repair, seen in the starting configuration of Reference [14]. It appears that the original nozzle was truncated, and a thinner nozzle inserted through the original nozzle and welded in place. However, since none of the components directly affect the J-groove weld, and new welds are distant, the effects of this repair on the J-groove weld are assumed to be negligible, so the 1987 repair has not been modeled in this analysis.

x As part of the new weld pad repair, the outer sleeve, installed in the 1992 repair, will be truncated and then receive a roll expansion. The effects of roll expansion are not expected to be significant and have not previously been considered at Palo Verde. These effects will not be considered in this analysis as well.

x The new nozzle-to-pad weld and the new nozzle-to-piping weld are not specifically simulated to save computation time, as the effects of these welds at the original nozzle-to-pressurizer J-groove weld, with regard to a fatigue crack growth evaluation, will be insignificant.

44.0 TECHNICAL APPROACH The residual stresses due to welding are controlled by various welding parameters, thermal transients due to the application of materials in the welding process, temperature dependent material properties, and elastic-plastic stress reversals. The analytical technique uses finite element analysis to simulate the multi-pass weld processes.

A weld residual stress (WRS) evaluation process developed in an internal Structural Integrity Associates calculation package [3] is used in this calculation. The WRS analysis process documented in Reference [3] follows the guidelines discussed in MRP-316, Revision 1 [4] and MRP-317, Revision 1 [5], and is validated by comparisons of numerical results with accepted measured residual stress data. The analysis process, verified in Reference [3], was refined into an automated WRS analysis module for ANSYS [6]. This module is used for the WRS evaluations documented in the current calculation and briefly described as follows.

Weld-induced residual stress analyses are nonlinear, path-dependent problems that result from the cumulative stress-strain cycling history inherent with the heating and cooling of materials during the welding process. Each simulation is performed as a continuous analysis so that the temperatures and stress histories from different welds are accounted for, e.g., the residual stresses and strains caused by the previous weld pass are used as initial conditions for the next weld pass.

The weld bead depositions are simulated using the element birth and death feature in ANSYS. The element birth and death feature in ANSYS allows for the deactivation (death) and reactivation (birth) of the elements stiffness contribution when necessary. It is used such that elements that have no contribution to a specific phase of the weld simulation process are deactivated (via the ANSYS EKILL command) because they have not been deposited. The deactivated elements have near-zero conductivity and stiffness contribution to the structure. When those elements are required in a later welding phase, they are then reactivated (via the ANSYS EALIVE command).

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 8 of 42 F0306-01R4 4.1 Welding Heat Input Simulation The analyses consist of a thermal pass to determine the temperature distribution due to the welding process, and an elastic-plastic stress pass to calculate the residual stresses through the thermal history induced by each weld pass. Appropriate weld heat efficiency, along with sufficient cooling time, are utilized in the thermal pass to ensure sufficient melting of weld beads, adequate heat penetration into the base metal, and a reasonable size prediction of the heat affected zone (HAZ).

Analytically, the deposition of the weld metal is simulated by imposing a delta heat generation function on the elements representing the active weld bead for each weld. In ANSYS, the heat generation applied to the weld beads is a volumetric energy rate, as energy per volume per time. Two-dimensional (2-D) axisymmetric PLANE55 elements are used in the thermal analysis, while 2-D axisymmetric PLANE182 elements are used in the stress analysis.

For the 2-D analysis, the heat generation rate () to be input into ANSYS is the heat efficiency ()

times the calculated total heat input (Q ) divided by the weld bead area (Abead) and the ramp time (tramp):

=

The heat input application is consistent with the validated process described in Section 3.2.4 of Reference [3]. Note that because the evaluation is performed using 2-D axisymmetric elements, the ANSYS program automatically accounts for volume behavior via projected revolution. Therefore, the heat generation rate can be simplified down to an area basis rather than a volume basis.

Per Reference [3], the heat generation rate is initially applied as a triangular heat generation function on the active weld elements. However, the triangular function is just an initial input curve. The weld nugget temperature is checked at each 5% heat input increment and, if the nugget elements have reached the melting temperature, the heat input will be halted, and cooling will start. The heat input curve can be suspended at any 5% increment interval because no additional energy is necessary when the nugget elements have melted.

x Since the welding parameters for the cladding, buttering, original J-groove weld, the 1992 outer sleeve fillet weld and the 1992 weld pad repair/nozzle-to-pad weld are not available, a typical heat input of 28 kJ/inch, with an overall heat efficiency of 0.8, and a weld bead area of 0.025 in² are assumed for the cladding, buttering, original J-groove weld, the 1992 outer sleeve fillet weld and the 1992 weld pad repair/nozzle-to-pad weld.

x Since the welding parameters for the weld repair pad and replacement nozzle-to-repair pad attachment weld are not available, a typical heat input of 35 kJ/inch, with an overall heat efficiency of 0.8, and a weld bead area of 0.025 in² are assumed for the weld repair pad.

The reasonableness of the above assumptions is confirmed by temperature fusion plots for each welding process (see Section 6.0). The fusion plots show the maximum occurring temperature at any point in the simulated welding process to show if the weldment material reaches the melting temperature, and that the heat penetration depth, where temperatures are above 1,300°F, is similar in size to the typical heat affected zone (HAZ) of roughly between 1/8 and 1/4.

Note that the heat efficiency value represents a composite value reflecting the concepts of arc efficiency, melting efficiency, etc., and is an optimum value to produce reasonable heat penetration in the analysis.

As the model is 2-D, weld head travel speed cannot be specifically modeled. Instead, a head input cycle (i.e., the width of the triangular heat generation function) is specified. In this evaluation, a Q

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 9 of 42 F0306-01R4 3 second heat input cycle was selected as representative for welding on 1.050 diameter [13.b]

components. The 3 second cycle is consistent with values developed during the verification and validation [3] for similar diameter piping systems where evaluated residual stresses were compared to actual measured results. This value, along with the weld bead size, heat input, and welding efficiency, generates the final temperature response in the evaluation, which are shown in a series of weld fusions plots (see Section 7.0). These plots can show over or under heating behavior which can then be corrected by revising the above variables.

The WRS module automatically calculates the appropriate time intervals for the thermal pass to ensure that sufficient heat penetration is achieved, the required interpass temperature between weld passes is met, and a reasonable overall temperature distribution within the finite element model is achieved. The resulting temperature time history is then imported into the stress pass in order to calculate the residual stresses due to thermal cycling of the weld elements using nonlinear, elastic-plastic load/unload stress reversal relations.

44.2 Weld Nugget Simulation Design drawings for the weld pad repairs [7] provide a list of the installation order of the weld pad repair. Logically, the cladding, butter, and J-groove weld are performed before the 1992 interim repair and the current weld pad repair. A breakdown of the weld order and number of beads for each weld are summarized as follows (see Figure 3 for a colored bead plot):

1.

The cladding is performed as a weld in 3 layers, using 3 weld nuggets.

2.

The weld butter is applied in 3 layers. A total of 21 nuggets are defined for this weld.

3.

The original J-groove weld is performed in 3 layers. A total of 14 nuggets are defined for this weld.

4.

The 1992 weld repair pad is performed in 5 layers. A total of 42 nuggets are defined for this weld.

5.

The 1992 outer sleeve fillet weld is performed in 1 layer, using a single weld nugget.

6.

The 1992 nozzle-to-pad weld is performed in 5 layers, using 5 weld nuggets.

7.

The new weld repair pad is applied in 9 layers. A total of 111 nuggets are defined for this weld.

4.3 Post Weld Heat Treatment Per the ASME Code requirements in the original code of construction [9], the cladded and buttered pressurizer was put through a post weld heat treatment (PWHT) before the J-groove weld was installed. Since no information is available on the actual PWHT details of the pressurizer, the finite element analysis is performed based on the minimum requirements in the ASME Code, 1971 Edition through Winter 1973 Addenda [9]. The ASME Code PWHT guidelines [9, NB-4600] state that:

x Above 600°F, the rate of heating shall not exceed 400°F per hour divided by the maximum metal thickness of the shell in inches, but in no case more than 400°F/hour [9, NB-4623.2]. For the 4.16 inch thick vessel shell (including cladding), the heating rate is 400/4.16= 96°F/hour.

Therefore, for the purposes of this analysis, the heating rate from 70°F to the PWHT holding temperature is conservatively evaluated at a higher value of 100°F/hour.

x The minimum PWHT holding temperature for the P-3 vessel material is 1100°F [9, Table NB-4622.1-1].

x The minimum hold time is one hour per inch of weld thickness, with a minimum hold time of one hour [9, Table NB-4622.1-1]. Note that for the purposes of this analysis, the hold time is 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.

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The difference between the highest and lowest temperature points in the vessel being heated shall not exceed 100°F during the holding period [9, NB-4623.3].

x Above 600°F, the cooling rate shall not exceed 500°F per hour divided by the maximum metal thickness of the shell [9, NB-4623.5]. Therefore, the cooling rate is 500/4.16 = 120°F/hour. Note that for the purposes of this analysis, a cooling rate of 120°F/hour is used for temperatures above 600°F.

x From 600°F, the vessel is cooled in still air [9, NB-4623.5].

x An additional steady state load step at 70°F is imposed at the end of the PWHT process.

4.3.1 Creep Properties One purpose of PWHT is to relieve the residual stresses from welding, and the major factor for stress relief is creep at high temperatures. In general, creep becomes significant at temperatures above 800°F; thus, creep behavior under 800°F will not be considered in this analysis.

There are two main categories of creep: primary and secondary. The primary creep addresses the creep characteristics for a short duration at the early stages of the creep regime, while the secondary creep accounts for the creep behavior for a long duration - usually more than 10,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br />. Based on this definition, the PWHT falls within the primary creep characteristics. However, primary creep rates for materials are difficult to obtain, so conservative secondary creep rates are used since primary creep rates are typically an order of magnitude higher than those for secondary creep.

In general, the primary creep rate for the materials is governed by the following equation:

=

The creep data for the SA-533, Grade B, Class 1 pressurizer material is based on the 1/2Mo material provided in Reference [10], which has a similar chemical composition. The creep data for the Alloy 82/182 is not available, so the creep properties for their base metal is used in this analysis instead. The creep data for Alloy 600 (for Alloy 82/182) are provided in Reference [11]. All the creep strengths are provided at two creep rates [10, 11] for each temperature point.

When creep strength is provided at two creep rates at the same temperature point, as listed in Table 2, then and can be calculated as follows:

=

=

=

ln

= ln

=

ln

ln

=

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 11 of 42 F0306-01R4 The resulting creep properties for the appropriate materials are tabulated in Table 2.

5.0 RRESIDUAL STRESS SIMULATION PROCESSS 5.1 CCladding The first weld simulated is the cladding, which is applied as three layers each of which is a single weld nugget. This approach is very conservative, but the application of a post weld heat treatment tempers the effects of this methodology. Only the pressurizer and the cladding are involved. After welding, the model is cooled down to a uniform ambient temperature of 70°F. No other components are attached to these components during this welding process.

5.2 W

Weld Butter The second weld simulated is the J-groove weld buttering. Following the application of the weld butter, the model is cooled down to 70°F 5.3 PPost Weld Heat Treatment The PWHT is simulated as detailed in Section 4.3. The heat up is applied at a rate of 100°F/hour and held at the PWHT temperature of 1100°F for 5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br />. Then, the model is cooled at the rate of 120°F/hour to 600°F, and then air cooled to the ambient temperature of 70°F.

5.4 OOriginal J-GGroove Weld The third weld simulated is the original fabrication J-groove weld. Only the pressurizer, cladding, weld butter, J-groove weld, and original nozzle are involved. ANSYS thermal couples are used to transfer heat across the small annulus between the vessel and the nozzle. These couples are left in place until the original nozzle is removed to ensure proper heat transfer during subsequent welds. After welding, the model is cooled down to a uniform ambient temperature of 70°F.

5.5 SShakedown Evaluation No. 1 Using guidance from MRP-317, page 5-8, Shakedown Evaluation [5], the current as-welded condition is cycled five (5) times between no internal pressure and a uniform temperature of 70°F to normal operating conditions (NOC) pressure and temperature.

The inclusion of these 5 cycles can generate some changes to the WRS results and an overall smoothing of the stress contours. Typically, the variation in residual stress stabilizes after 2 cycles, but 5 cycles are included for completeness. This is intended to capture the operation that would have occurred between the installation of the DMW / ID repair and the application of the weld repair pad. The time duration of the cycle (zero state/NOC/zero state) does not affect the shakedown behavior.

The NOC loads consist of a uniform normal operating temperature of 653°F, and an internal operating pressure of 2,250 psia (2,235 psig) [8]. The operating temperature is treated as a uniform temperature applied to the entirety of the structure. The operating pressure is applied as an internal pressure on the inside surfaces. A pressure induced end-cap load is applied to the modeled free end of the attached original nozzle in the form of tensile axial pressure, and the value is calculated as follows:

=

=

where, Pnozz

= End cap pressure at modeled original nozzle free end (psi)

P

= Internal pressure (psi)

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 12 of 42 F0306-01R4 IDnozz

= Inside diameter of the original nozzle free end (in) [13.b]

ODnozz = Outside diameter of the original nozzle free end (in) [13.b]

Symmetric boundary conditions are applied to the circumferential free end of the pressurizer wall (see Figure 5). As this is a 2-D axisymmetric model, the only applied boundary conditions are perpendicular to the cut plane through the pressurizer wall/cladding thickness. In addition, axial degree-of-freedom couples are applied to the free end of the original instrumentation nozzle. Figure 4 shows the applied internal pressure loading, which includes the end cap line loads that are simulated as axial tensile pressures applied to the free end of the original instrument nozzle.

5.6 11992 Weld Pad Repair The fourth weld simulated is the 1992 weld pad repair. Prior to this weld, most of the original instrument nozzle is removed by machining (modeled using the EKILL feature), leaving only a short stub near the original J-groove weld. The machining of these components is assumed to be instantaneous and occurs at 70°F. Thermal couples are used to bridge the annular region between the original instrument nozzle remnant and the vessel. After the welding of the pad repair, the model is cooled down to 70°F.

5.7 11992 Outer Sleeve Fillet Weld The fifth weld simulated is the 1992 outer sleeve fillet weld. Before this weld is applied, the rest of the original nozzle and a portion of the original J-groove weld, butter, and vessel wall are machined out (modeled using the EKILL feature). The machining of these components is assumed to be instantaneous and occurs at 70°F. The outer sleeve is then inserted (using the EALIVE feature), and the outer sleeve fillet weld applied. It should be noted that this weld is autogenous, however, there is no impact on this analysis method. ANSYS thermal and mechanical contact elements (CONTA172 and TARGE169) are used to transfer heat across the small annulus between the vessel and the sleeve.

These contacts are left in place for the remainder of the analysis to ensure proper heat transfer during subsequent welds. Following the application of the fillet weld, the model is cooled down to 70°F.

5.8 11992 Nozzle-tto-PPad Weld The sixth weld simulated is the 1992 nozzle-to-pad weld. The 1992 nozzle is inserted (using the EALIVE feature), and thermal couples are used to bridge the narrow gap between the outer sleeve and the new nozzle. It should be noted that this weld also captures the end of the outer sleeve introduced in the previous welding step. After welding, the model is cooled down to 70°F.

5.9 SShakedown Evaluation No. 2 After the completion of the J-groove weld repair, another shakedown evaluation is run in accordance with MRP-317, page 5-8, Shakedown Evaluation [5]. The post 1992 weld repair pad welded condition is cycled five (5) times between no internal pressure and a uniform room temperature of 70°F to representative normal operating conditions pressure and temperature. This cycling is consistent with that discussed in Section 5.5, with the exception that the nozzle end cap pressure has changed to reflect the new nozzle free end dimensions, as calculated below:

=

=

where, Pnozz

= End cap pressure at modeled 1992 nozzle free end (psi)

P

= Internal pressure (psi)

IDnozz

= Inside diameter of the 1992 nozzle free end (in) [7]

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 13 of 42 F0306-01R4 ODnozz = Outside diameter of the 1992 nozzle free end (in) [7]

5.10 CCurrent Repair Weld Pad The final weld simulated is the current repair weld pad. The 1992 nozzle and a portion of the vessel are removed and the original 1992 weld pad repair is machined down flush with the OD of the vessel (both are modeled using the EKILL feature). After welding, the model is cooled down to a uniform ambient temperature of 70°F.

5.11 SShakedown Evaluation No. 3 After the completion of the repair weld pad, the size of the bore hole is expanded, which removes a portion of the current weld repair and a portion of the outer sleeve (both are modeled using the EKILL feature). A new nozzle (with a portion of the attached safe end) is added to the pad (using the EALIVE feature), though the actual new nozzle-to-repair pad weld is not simulated. ANSYS thermal contact elements (CONTA172 and TARGE169) are used to transfer heat between the vessel and the sleeve remnant.

Per MRP-317, page 5-8, Shakedown Evaluation [5], the post weld repair pad welded condition is cycled five (5) times between no internal pressure and a uniform room temperature of 70°F to representative normal operating conditions pressure and temperature. This cycling is consistent with that discussed in Section 5.5, with the exception that the nozzle/safe-end end cap pressure has changed to reflect the new nozzle/safe-end free-end dimensions, as calculated below:

=

=

where, Pnozz

= End cap pressure at modeled new nozzle/safe-end free end (psi)

P

= Internal pressure (psi)

IDnozz

= Inside diameter of the new nozzle/safe-end free end (in)

ODnozz = Outside diameter of the new nozzle/safe-end free end (in)

Symmetric boundary conditions are applied to the circumferential free end of the pressurizer wall. As this is a 2-D axisymmetric model, the only applied boundary conditions are perpendicular to the cut plane through the pressurizer wall/cladding thickness. In addition, axial degree-of-freedom couples are applied to the free end of the new instrumentation nozzle. Figure 4 shows the applied internal pressure loading, which includes the end cap line loads that are simulated as axial tensile pressures applied to the free end of the original instrument nozzle.

6.0 CCALCULATIONSS The analytical procedure described in Section 4.0 is implemented using the WRS module in ANSYS [6].

Observing the fusion boundary prediction after the weld simulation provides reasonable assurance that appropriate heat input has been applied. Figure 6 through Figure 12 show the predicted fusion boundaries for all the welding processes as generated by ANSYS for the specific welding sequences.

The fusion boundaries represent the predicted maximum temperature contour mapping that the weld bead elements will reach during each welding process, with purple representing temperature above the melting point.

Note that the figures are composites showing the maximum temperature among all beads of each weld.

This is made possible by an ANSYS macro (MapTemp.mac) that reads in the maximum predicted temperature across the different weld bead elements during the welding process and displays them as a maximum temperature contour plot.

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 14 of 42 F0306-01R4 The figures show that all weld elements have reached the melting temperature. They also show that the heat penetration depth, where temperatures are above 1,300°F, is similar in size to the typical heat affected zone of roughly between 1/8 and 1/4".

The ANSYS input and output files for the analysis are listed in Appendix A.

77.0 RESULTS OF ANALYSIS Figure 13 through Figure 22 depict the hoop residual stress distribution for the post cladding, buttering, PWHT, original J-groove weld, Shakedown #1 cycles, 1992 weld pad repair, 1992 outer sleeve fillet weld, the 1992 nozzle-to-pad weld, Shakedown #2 cycles, and finally, the weld pad repair, respectively.

The hoop direction is with respect to a local cylindrical coordinate system about the original instrument nozzle.

Figure 23 depicts the hoop residual stress distributions at 70°F after the application of the new instrument nozzle and following 5 cycles of normal operating conditions (i.e., Shakedown #3).

Figure 24 depicts the resultant residual plus normal operating condition hoop stress distribution following the 5 cycles of normal operating conditions. Hoop stress for these plots is relative to the new instrument nozzles axis of orientation.

The weld residual stresses generated in this evaluation will be applied to a crack growth model in a later fracture mechanics calculation.

8.0 CONCLUSION

S A component-specific weld residual stress analysis has been performed for the pad-based repair of the pressurizer lower instrument nozzles to be installed at the Palo Verde Nuclear Generating Station, Unit 1. The resulting weld residual stresses will be used in a later crack growth calculation.

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 15 of 42 F0306-01R4

9.0 REFERENCES

1.

ANSYS Mechanical APDL (UP20170403) and Workbench (March 31, 2017), Release 18.1, SAS IP, Inc.

2.

SI Calculation No. 2551740.301P, Revision 0, Finite Element Model Development of the Pressurizer Lower Instrument Nozzle Weld Modification, SI PROPRIETARY.

3.

SI Software V&V Package No. 0800777.312, Revision 0, Black Box Testing and Verification of Weld Residual Stress Finite Element Analysis Module, Version 4.0, SI PROPRIETARY.

4.

Materials Reliability Program: Finite-Element Model Validation for Dissimilar Metal Butt-Welds (MRP-316, Revision 1): Volumes 1 and 2, EPRI, Palo Alto, CA: 2015. 3002005498.

5.

Materials Reliability Program: Welding Residual Stress Dissimilar Metal Butt-Weld Finite Element Modeling Handbook (MRP-317, Revision 1), EPRI, Palo Alto, CA: 2015. 3002005499.

6.

Weld Residual Stress Analysis Module for ANSYS, Version 4.00, Structural Integrity Associates, Inc., March 2016, SI File No. 0800777.218, SI PROPRIETARY.

7.

Framatome Drawing No. 02-8164311-E, Revision 004, Palo Verde Unit 1 Pressurizer Lower Instrument Nozzle Modification, SI File No. 2551740.201P, Framatome Inc. Proprietary.

8.

SI Calculation No. 2000645.301, Revision 2, Transient Loads Definitions for Updated Surge Line Analyses.

9.

ASME Boiler and Pressure Vessel Code,Section III, Nuclear Power Plant Components. 1971 Edition with Addenda through Winter 1973.

10. Pipe and Tubes for Elevated Temperature Service, United States Steel Co., 1949.

11. Publication SMC-027, Inconel Alloy 600, Special Metals Corp., 2004, SI File 0800777.211.

12. Framatome Design Specification No. 08-9383300-004, April 24, 2025, Palo Verde Unit 1 Pressurizer Upper and Lower Instrument Nozzle Modification, SI File No. 2451268.202P, Framatome Inc. Proprietary.

13. Combustion Engineering Drawings, PV-04Q-202.

a.

Drawing No. E-78373-641-001, Revision 02, Lower Vessel Assembly and Heater Holes, Arizona Public Service I, 96 ID Pressurizer.

b.

Drawing No. E-78373-684-001, Revision 02, Instrument Nozzles and Heater Sleeves, Arizona Public Service I, 96 ID Pressurizer.

c.

Drawing No. E-78373-651-001, Revision 02, Bottom Head Assembly, Arizona Public Service I, 96 ID Pressurizer.

14. BWXT Drawing No. 1214122E, Revision 02, Lower Instrument Nozzle Installation, SI File No, 2451268.206P, Framatome Inc. Proprietary.

15. Structural Integrity Calculation No. 0800777.307, Revision 8, Material Properties for Residual Stress Analyses, Including MISO Properties Up to Material Flow Stress.

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 16 of 42 F0306-01R4 Table 1. Component Materials Component Material Pressurizer Base Metal SA-533, Grade B, Class 1 Pressurizer Cladding Inconel (Treated as Alloy 82)

Original J-groove Weld Butter Alloy 182 Original J-groove Weld Alloy 182 1992 Sleeve/Fillet Weld Alloy 690/52M 1992 Repair Pad Alloy 82 Original Nozzle Alloy 600 1992 Interim Nozzle Alloy 690 Replacement Nozzle Alloy 690 Repair Pad Alloy 52M Note:

1.

This table is based on Table 1 of Reference [2].

Table 2. Creep Properties Material Temperature

(°F)

Creep Strength (ksi))

A (ksi/hr) n 0.0001%/hr 0.00001%/hr SA-533, Grade B, Class 1 (Based on 1/2Mo steel) 800 30.2 21.3 1.73E-16 6.60 900 20.5 13.0 2.34E-13 5.06 1000 10.8 6.3 3.85E-11 4.27 1100 3.0 2.0 1.95E-09 5.68 Alloy 182 (Based on Alloy 600) 800 40.0 30.0 1.50E-19 8.00 900 28.0 18.0 2.87E-14 5.21 1000 12.5 6.1 3.02E-10 3.21 1100 6.8 3.4 1.72E-09 3.32 Note: Data for SA-533, Grade B, Class 1 from Reference [10], data for Alloy 182 from Reference [11].

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 17 of 42 F0306-01R4 Figure 11. CComponents Included in the Finite Element Model (Modified from Reference [2]. See Section 2.1.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 18 of 42 F0306-01R4 Figure 22. ANSYS 22-DD Finite Element Model of Instrum ment Nozzle / Repair Pad for Weld Residual SStress Analysis (with Minimum Dimensionss)

(Modified from Reference [2]. See Section 2.1.)

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 19 of 42 F0306-01R4 Figure 33. Weld Bead Distribution

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 20 of 42 F0306-01R4 Figure 44. Applied Pressuree Loading (Units for pressure in ksi.)

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 21 of 42 F0306-01R4 Figure 55. Examplee Applied M Mechanicall Boundary Conditions

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 22 of 42 F0306-01R4 Figure 66. Predicted Fusion Boundaryy for tthe Cladding (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 23 of 42 F0306-01R4 Figure 77. Predicted Fusion Boundaryy for tthe Weld Butter (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 24 of 42 F0306-01R4 Figure 88. Predicted Fusion Boundaryy for the OOriginal J--Groove Weld (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 25 of 42 F0306-01R4 Figure 99. Predicted Fusion Boundaryy for the 1992 Weld Pad RRepair (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 26 of 42 F0306-01R4 Figure 110. Predicted Fusion Boundary for the OOuter Sleeve Fillet Weld (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 27 of 42 F0306-01R4 Figure 111. Predicted Fusion Boundary for the NNozzle-tto-PPad Weld (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 28 of 42 F0306-01R4 Figure 112. Predicted Fusion Boundary for the Weld Pad Repair (Units for temperature in °F. The purple region represents melting temperature in excess of 2500°F.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 29 of 42 F0306-01R4 Figure 113. Post CCladding Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 30 of 42 F0306-01R4 Figure 114. Post BButtering Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 31 of 42 F0306-01R4 Figure 115. Hoop Residual Stress after PWHT at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 32 of 42 F0306-01R4 Figure 116. Post OOriginal J--Groove Weld Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 33 of 42 F0306-01R4 Figure 117. PPost Initiall Operation HHoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

(Results include 5 NOC cycles.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 34 of 42 F0306-01R4 Figure 118. Post 1992 Weld Repair Pad Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 35 of 42 F0306-01R4 Figure 119. Post Outerr Sleeve Fillet Weld Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 36 of 42 F0306-01R4 Figure 220. Post 1992 NNozzle-tto-PPad Hoop Stress at 70°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 37 of 42 F0306-01R4 Figure 221. Post SSecondary Operation Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

(Results include 5 NOC cycles.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 38 of 42 F0306-01R4 Figure 222. Post Weld Pad Repair Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 39 of 42 F0306-01R4 Figure 223. Final Weld Pad Repair with New Nozzle Hoop Stress at 70°°F (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

(Results include 5 NOC cycles.)

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page 40 of 42 F0306-01R4 Figure 224. FFinal Weld Pad Repairr with New Nozzle Hoop Stress at NOC (Cylindrical coordinate system of the original nozzle used. Units for stress in psi.)

(Results include 5 NOC cycles.)

File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page A-1 of 42 F0306-01R4 COMPUTER FILES

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File No.: 2551740.302NP Revision: 0 Framatome Document No.: 32-9398249-000 Page A-2 of 42 F0306-01R4 Filename Description INST_NOZ_RES.INP Finite element model input file generated in Reference [2]. Modified to include the intermediate configurations.

BCNUGGET2D.INP ANSYS input file for weld bead and boundary area definition.

THERMAL2D.INP ANSYS input file to perform the thermal pass.

STRESS2D.INP ANSYS input file to perform the stress pass.

WELD#_mntr.inp ANSYS input file generated by thermal pass.

Contained load steps for stress pass, # = 1 - 7.

• .mac WRS analysis macro files required for analysis,

• = various file names THERMAL2D.TXT Text file of parameter for thermal pass STRESS2D.TXT Test file of parameter input file for stress pass

Enclosure Relief Request Number 77

Affidavit Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds - Framatome Proprietary, Document Number 32-9399131-001

A F F I D A V I T

1.

My name is Ronda M. Lane. I am Manager, Product Licensing for Framatome (formally known as AREVA Inc.), and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by Framatome to determine whether certain Framatome information is proprietary. I am familiar with the policies established by Framatome to ensure the proper application of these criteria.

3.

I am familiar with the Framatome information contained in Enclosure 2, Fracture Mechanics and Crack Growth Analysis of the Pressurizer Lower Instrument Nozzle J-Groove Welds marked as Framatome Proprietary Information herein referred to as Document.

Information contained in this Document has been classified by Framatome as proprietary in accordance with the policies established by Framatome for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by Framatome and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) Trade secrets and commercial or financial information.

6.

The following criteria are customarily applied by Framatome to determine whether information should be classified as proprietary:

(a)

The information reveals details of Framatomes research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c)

The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for Framatome.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for Framatome in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by Framatome, would be helpful to competitors to Framatome, and would likely cause substantial harm to the competitive position of Framatome.

The information in this Document is considered proprietary for the reasons set forth in criteria 6(b), (c), and (e).

7.

In accordance with Framatomes policies governing the protection and control of information, proprietary information contained in this Document has been made available, on a limited basis, to others outside Framatome only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8.

Framatome policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

I declare under penalty of perjury that the foregoing is true and correct.

Executed on October 29th, 2025.

Ronda M. Lane Ronda M. Lane

Enclosure Relief Request Number 77

Affidavit Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Modification, Proprietary Document Number 32-9398289-000

A F F I D A V I T

1.

My name is Philip A. Opsal. I am Manager, Product Licensing for Framatome Inc. (formally known as AREVA Inc.), and as such I am authorized to execute this Affidavit.

2.

I am familiar with the criteria applied by Framatome to determine whether certain Framatome information is proprietary. I am familiar with the policies established by Framatome to ensure the proper application of these criteria.

3.

I am familiar with the Framatome information contained in Framatome Document No. 32-9398289-000, Weld Residual Stress Analysis of the Pressurizer Lower Instrument Nozzle Weld Modification, referred to herein as this Document.

Information contained in this Document has been classified by Framatome as proprietary in accordance with the policies established by Framatome for the control and protection of proprietary and confidential information.

4.

This Document contains information of a proprietary and confidential nature and is of the type customarily held in confidence by Framatome and not made available to the public. Based on my experience, I am aware that other companies regard information of the kind contained in this Document as proprietary and confidential.

5.

This Document has been made available to the U.S. Nuclear Regulatory Commission in confidence with the request that the information contained in this Document be withheld from public disclosure. The request for withholding of proprietary information is made in accordance with 10 CFR 2.390. The information for which withholding from disclosure is requested qualifies under 10 CFR 2.390(a)(4) Trade secrets and commercial or financial information.

6.

The following criteria are customarily applied by Framatome to determine whether information should be classified as proprietary:

(a)

The information reveals details of Framatomes research and development plans and programs or their results.

(b)

Use of the information by a competitor would permit the competitor to significantly reduce its expenditures, in time or resources, to design, produce, or market a similar product or service.

(c)

The information includes test data or analytical techniques concerning a process, methodology, or component, the application of which results in a competitive advantage for Framatome.

(d)

The information reveals certain distinguishing aspects of a process, methodology, or component, the exclusive use of which provides a competitive advantage for Framatome in product optimization or marketability.

(e)

The information is vital to a competitive advantage held by Framatome, would be helpful to competitors to Framatome, and would likely cause substantial harm to the competitive position of Framatome.

The information in this Document is considered proprietary for the reasons set forth in paragraphs 6(b), 6(d), and 6(e) above.

7.

In accordance with Framatomes policies governing the protection and control of information, proprietary information contained in this Document has been made available, on a limited basis, to others outside Framatome only as required and under suitable agreement providing for nondisclosure and limited use of the information.

8.

Framatome policy requires that proprietary information be kept in a secured file or area and distributed on a need-to-know basis.

9.

The foregoing statements are true and correct to the best of my knowledge, information, and belief.

I declare under penalty of perjury that the foregoing is true and correct.

Executed on September 24, 2025.

Philip A. Opsal Manager, Product Licensing, Framatome Inc.

Philip A. Opsal