L-21-282, Proposed Inservice Inspection Alternative IR-063

Proposed Inservice Inspection Alternative IR-063

ML22006A167
Person / Time
Site:	Perry
Issue date:	01/05/2022
From:	Penfield R Energy Harbor Nuclear Corp
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
L-21-282
Download: ML22006A167 (86)

Text

energy Energy Harbor Nuclear Corp.

Perry Nuclear Power Plant IO Center Road harbor P.O. Box 97 Perry, Ohio 44081

Rod L. Penfield 440-280-5382 Site Vice President, Perry Nuclear

January 5, 2022 L-21-282 10 CFR 50.55a

ATTN: Document Control Desk U.S. Nuclear Regulatory Commission Washington, DC 20555-0001

Subject:

Perry Nuclear Power Plant Docket No. 50-440, License No. NPF-58 Proposed lnservice Inspection Alternative IR-063

In accordance with 10 CFR 50.55a(z)(1 ), Energy Harbor Nuclear Corp. hereby requests Nuclear Regulatory Commission (NRC) staff approval of a proposed inservice inspection alternative to the American Society of Mechanical EngineersSection XI, Table IWB-2500-1, "Examination Category B-D, Full Penetration Welded Nozzles in Vessels," for use at Perry Nuclear Power Plant. The proposed alternative is enclosed and identifies the affected components, applicable code requirement, and a description and basis for the proposed alternative.

NRC staff review and approval of the proposed alternative is respectfully requested by January 31, 2023 to allow for application of the alternative during the spring 2023 refueling outage.

There are no regulatory commitments contained in this submittal. If there are any questions or if additional information is required, please contact Mr. Phil H. Lashley, Manager - Fleet Licensing, at (330) 696-7208.

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. / DJ~t/4~ L1*;?1. //JOI/,

Rod L. Penfier

Enclosure:

10 CFR 50.55a Request IR-063 Perry Nuclear Power Plant L-21-282 Page 2

cc: NRC Region III Administrator NRC Resident Inspector NRC Project Manager

Enclosure L-21-282

10 CFR 50.55a Request IR-063

(7 pages follow)

10 CFR 50.55a Request IR-063

Proposed Alternative in Accordance with 10 CFR 50.55a(z)(1)

-- Alternative Provides Acceptable Level of Quality and Safety --

Page 1 of 7

1. ASME Code Components Affected

Code Class: Class 1

Description:

Reactor Pressure Vessel (RPV) Reactor Feedwater (RFW) Nozzles Examination Category: B-D Item Numbers: B3.90 B3.100 Component IDs:

Table 1 ASME Component ID Component Description Item No.

1B13-N4A-KA Feedwater Nozzle N4A to Vessel B3.90 1B13-N4A-IR Feedwater Nozzle N4A Inner Radius B3.100 1B13-N4B-KA Feedwater Nozzle N4B to Vessel B3.90 1B13-N4B-IR Feedwater Nozzle N4B Inner Radius B3.100 1B13-N4C-KA Feedwater Nozzle N4C to Vessel B3.90 1B13-N4C-IR Feedwater Nozzle N4C Inner Radius B3.100 1B13-N4D-KA Feedwater Nozzle N4D to Vessel B3.90 1B13-N4D-IR Feedwater Nozzle N4D Inner Radius B3.100 1B13-N4E-KA Feedwater Nozzle N4E to Vessel B3.90 1B13-N4E-IR Feedwater Nozzle N4E Inner Radius B3.100 1B13-N4F-KA Feedwater Nozzle N4F to Vessel B3.90 1B13-N4F-IR Feedwater Nozzle N4F Inner Radius B3.100

2. Applicable Code Edition and Addenda

American Society of Mechanical Engineers (ASME) Boiler and Pressure Vessel Code,Section XI, Rules for Inservice Inspection of Nuclear Power Plant Components, 2013 Edition (no Addenda).

3. Applicable Code Requirement

The applicable Code requirement is contained in ASME Section XI, Division 1, Subsection IWB, Table IWB-2500-1, Examination Category B-D, Full Penetration Welded Nozzles in Vessels. Class 1 nozzle-to-vessel welds and nozzle inner radii examination requirements are delineated in Item Numbers B3.90, Nozzle-to-Vessel Welds, and B3.100, Nozzle Inside Radius Section. The method of

10 CFR 50.55a Request IR-63 Page 2 of 7

examination is volumetric. With respect to the extent of examination, all nozzles with full penetration welds to the vessel shell (or head) and integrally cast nozzles must be examined each interval.

4. Reason for Request

The proposed alternative provides an acceptable level of quality and safety based on a plant-specific evaluation using a probability fracture mechanics (PFM) methodology endorsed by the Nuclear Regulatory Commission (NRC) in Electrical Power Research Institute (EPRI) Technical Reports BWRVIP-108-A, Technical Basis for the Reduction of Inspection Requirements fo r the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Inner Radii, [Reference 1] and BWRVIP-241-A, Probabilistic Fracture Mechanics Evaluation for the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Blend Radii, [Reference 2].

5. Proposed Alternative and Basis for Use

In lieu of performing examinations on 100 percent of the population of RFW nozzle assemblies, Energy Harbor Nuclear Corp. proposes to examine a minimum of 25 percent of the population of the nozzle-to-vessel welds and a minimum of 25 percent of the population of the inner radii using volumetric inspection methods performed in accordance with ASME Section XI, Appendix VIII, Performance Demonstration for Ultrasonic Examination Systems, as modified by 10 CFR 50.55a.

Basis for Use

EPRI Technical Reports BWRVIP-108-A [Reference 1] and BWRVIP-241-A

[Reference 2] contain the technical basis supporting ASME Boiler and Pressure Vessel Code Case N-702, Alternative Requirements for Boiling Water Reactor Nozzle Inner Radius and Nozzle-to-Shell Welds,Section XI, Division 1

[Reference 3] for reducing the inspection of RPV nozzle-to-vessel welds and nozzle inner radius regions from 100 percent of the population of the nozzles to 25 percent of the population of nozzles for each nozzle type during each 10-year ISI interval.

However, the reports and Code Case N-702 explicitly exclude reactor feedwater nozzles stating that these nozzles are managed under a separate mandated program directed by NUREG-0619, BWR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking: Resolution of Generic Technical Activity A-10 (Technical Report) [Reference 4]. This request proposes that the RFW nozzle-to-vessel welds and inner radii examinations mandated under NUREG-0619 can be subsumed by the current ASME Section XI requirements and the number of inspections reduced based on this relief request.

Discussion on NUREG-0619

The Perry Nuclear Power Plant (Perry) complies with the inspection requirements of NUREG-0619 by implementation of the NRC approved alternative 10 CFR 50.55a Request IR-63 Page 3 of 7

GE-NE-523-A71-0594-A, [Reference 5], which stipulates inspections for the RFW nozzle inner radii, nozzle inner bore regions, and the spargers. The nozzle inner radii inspections are performed in compliance with ASME Section XI, as modified by the Performance Demonstrated Initiative (PDI) Program description and meets the established criteria of Reference 5. Section 6.3 of Reference 5 states that after compliance with ASME Section XI, Appendix VIII, the examination frequency will be the ASME Section XI examination frequency for non-interference fit plants such as Perry.

NUREG-2221, Technical Bases for Changes in the Subsequent License Renewal Documents NUREG-2191 and NUREG-2192, December 2017, states that the recommendation for condition monitoring of the RFW nozzles be performed under ASME Section XI ISI inspections for the condition monitoring basis for managing cracking in boiling water reactor (BWR) feedwater nozzles induced by cyclical loading mechanism. This is based in part on improvements in ASME Section XI directed volumetric examination techniques, which now meet the requirements proposed in Reference 5. Thus, there is a general acknowledgement that the RFW nozzle examinations need no longer go beyond ASME Section XI guidelines and a separate mandated program is not required.

Justification for Reduction to 25 Percent

One of the criteria for demonstrating plant-specific applicability of BWRVIP-108-A

[Reference 1] and BWRVIP-241-A [Reference 2] is to show that the maximum RPV heatup and cooldown rate for the unit is limited to less than 115°F per hour. Perry Technical Specification 3.4.11, RCS Pressure and Temperature (P/T) Limits, limits the maximum reactor vessel operational heatup and cooldown rates to less than or equal to 100°F per hour, which is within the 115°F per hour limit. Therefore, this criterion is satisfied.

Using the same analytical methodology as employed in BWRVIP-108-A and BWRVIP-241-A [References 1 and 2], it can be shown that Perrys RFW nozzle failure probabilities due to a low temperature over pressure (LTOP) event at the nozzle radius region and the nozzle-to-vessel weld are very low and meet the NRC acceptance criteria in NUREG-1806, Technical Basis for Revision of Pressurized Thermal Shock (PTS) Screen Limit in the PTS Rule (10 CFR 50.61). Based on the results of the plant-specific evaluation, and industry and internal operating experience, the inspection of 25 percent of the population of RFW nozzles is considered technically justified.

The plant-specific PFM analysis employed a Monte Carlo simulation using Structural Integrity Associates, Inc. proprietary software VIPERNOZ, which was developed for RPV nozzle weld inspection with BWRVIP-108-A. Perrys nozzle stresses are used with probabilistic distributions from BWRVIP-108-A and BWRVIP-241-A [References 1 and 2] to evaluate the plant specific probabilities. Stress results for Perry are provided in Attachment 1 of this request. The PFM evaluation is provided in 10 CFR 50.55a Request IR-63 Page 4 of 7

Attachment 2. The attached plant-specific analysis conservatively extends to 60 years. The PFM evaluation conclusions are summarized in Section 5.0 of Attachment 2.

For normal operation, the results for Stress Path 1 at the nozzle blend radius bounds the other stress paths. For Stress Path 1, five (5) failures occurred during normal operation in 1 million simulations for 60 years of plant operation. The probability of failure (PoF) for normal operation for Stress Path 1 is calculated to be 5 failures / 1 million simulations / 60 years = 8.33 x 10-8 per year. The calculated PoF for normal operation for Stress Path 1 is less than the allowable PoF of 5.0 x 10-6 per year from NUREG-1806.

For LTOP events, the results for Stress Path 1 at the nozzle blend radius bounds the other stress paths. For Stress Path 1, 52 LTOP failures occurred in 1 million simulations for 60 years of plant operation. The conditional PoF for LTOP events for Stress Path 1 is calculated to be 52 failures / 1 million simulations / 60 years =

8.67 x 10-7 per year. Accounting for an LTOP event occurrence of 1.0 x 10-3 per year, the calculated PoF for LTOP events for Stress Path 1 is 8.67 x 10-10 per year, which is less than the allowable PoF of 5.0 x 10-6 per year from NUREG-1806.

The PFM results are also confirmed by a deterministic fracture mechanics (DFM) evaluation (Attachment 2, Appendix A), using the methodology in Section 6 of BWRVIP-108-A.

Thus, using the PFM methodology in BWRVIP-108-A and BWRVIP-241-A, which is the technical basis for ASME Code Case N-702, the Perry feedwater nozzles are qualified for inspection relief from 100 percent to 25 percent of the feedwater nozzle population every 10 years from the fourth 10-year ISI interval for up to 60 years of plant operation.

Code Case N-702 is listed in Regulatory Guide (RG) 1.147, Revision 19, Table 2, Conditionally Acceptable Section XI Code Cases. The required condition associated with Code Case N-702 is as follows:

The applicability of Code Case N-702 for the first 40 years of operation must be demonstrated by satisfying the criteria in Section 5.0 of NRC Safety Evaluation regarding BWRVIP-108 dated December 18, 2007 (ML073600374) or Section 5.0 of NRC Safety Evaluation regarding BWRVIP-241 dated April 19, 2013 (ML13071A240).

The condition is intended to verify the applicability of the technical basis documents

[References 1 and 2] to the licensees design. Because the technical basis documents exclude the RFW nozzles, the Code Case is not applicable; however, Energy Harbor Nuclear Corp. considers the intent of the Code Case, the condition, and acceptance criteria are demonstrated by the plant-specific analysis for Perry.

10 CFR 50.55a Request IR-63 Page 5 of 7

Perry Operating Experience

A detailed evaluation of the historical problems of the RFW nozzle and sparger is presented in NEDE-21821, BWR Feedwater Nozzle / Sparger Final Report, March 1978. The solution of the feedwater nozzle and sparger cracking problems involved several elements, including material selection and processing, nozzle clad elimination, and thermal sleeve and sparger redesign. Perry implemented these changes during construction including clad elimination around the nozzle and a welded thermal sleeve and safe end design.

As documented in Perrys Updated Safety Analysis Report Section 5.3.3.3, The shell and vessel head were made from formed low alloy steel plates, and the flanges and nozzles from low alloy steel forgings. The section goes on to state, Post weld heat treatment of 1,100°F minimum was applied to all low alloy steel welds. This heat treatment should reduce residual stresses from any repairs such that they would not be a dominant force requiring consideration in the analysis.

Weld residual stress (after post weld heat treatment) was included in the probabilistic fracture mechanics design input as described in Attachment 2.

A review of the most recent examination results for each component listed in Table 1 identified all ultrasonic examination results were acceptable to the requirements of ASME Boiler and Pressure Vessel Code Section XI, 2001 Edition with the 2003 Addenda. All of the examinations were volumetric, using ultrasonic testing (UT) methodology performed in accordance with ASME Section XI, 2001 Edition and 2003 Addenda, and as modified by the PDI program and 10 CFR 50.55a requirements. All of the inner radius welds in Table 1 had 100 percent exam coverage. All of the nozzle-to-vessel welds had greater than or equal to 82.7 percent exam coverage. Table 2 describes the coverage, exam method, and when the most recent exam occurred for each weld. The reduced exam coverage for the nozzle-to-vessel welds was due to nozzle forging configurations, for which relief was granted in the NRC safety evaluation dated September 18, 2020 (Accession No. ML20252A026). Only the volumetric examinations required by ASME Section XI, Table IWB-2500-1 Category B-D are discussed in this request.

10 CFR 50.55a Request IR-63 Page 6 of 7

Table 2 Component Description Year of Exam Percent ID Last Exam Method Coverage Performed 1B13-N4A-KA Feedwater Nozzle N4A to Vessel 2013 UT 83.2%

1B13-N4A-IR Feedwater Nozzle N4A Inner Radius 2013 UT 100%

1B13-N4B-KA Feedwater Nozzle N4B to Vessel 2013 UT 83.2%

1B13-N4B-IR Feedwater Nozzle N4B Inner Radius 2013 UT 100%

1B13-N4C-KA Feedwater Nozzle N4C to Vessel 2013 UT 83.2%

1B13-N4C-IR Feedwater Nozzle N4C Inner Radius 2013 UT 100%

1B13-N4D-KA Feedwater Nozzle N4D to Vessel 2013 UT 83.2%

1B13-N4D-IR Feedwater Nozzle N4D Inner Radius 2013 UT 100%

1B13-N4E-KA Feedwater Nozzle N4E to Vessel 2013 UT 82.7%

1B13-N4E-IR Feedwater Nozzle N4E Inner Radius 2013 UT 100%

1B13-N4F-KA Feedwater Nozzle N4F to Vessel 2013 UT 83.2%

1B13-N4F-IR Feedwater Nozzle N4F Inner Radius 2013 UT 100%

Industry Operating Experience

NUREG-0619 was published in November 1980 to address instances of thermal fatigue cracking of the RFW and control rod drive return line nozzle. The guidance gave design, operating, and inspection recommendations to address these concerns. As identified in a July 25, 2006 BWRVIP letter to the NRC (Accession No. ML062080159), a survey of all U.S. BWRs shows the majority of RPV nozzles had no reportable indications in the nozzle-to-vessel weld or their inner radii. A few nozzles contained subsurface indications, which were determined to be acceptable.

The data in the letter indicates that the inspections performed as of that date, using reliable techniques, have shown that there are no active degradation mechanisms for the nozzle-to-vessel welds and inner radii regions. A search of the Institute of Nuclear Power Operations (INPO) database performed in December of 2021 found that this trend appears to continue with no reports of indications in RPV nozzle-to-vessel welds or inner radii regions found.

Proposed Alternative

Energy Harbor Nuclear Corp. proposes to inspect 25 percent of the population of RFW nozzle-to-vessel welds and inner radius regions each 10-year ISI inspection interval at Perry using volumetric inspection methods that meet or exceed the requirements of ASME Section XI, 2013 Edition (no addenda), Appendix VIII, Performance Demonstration for Ultrasonic Examination Systems, as modified by 10 CFR 50.55a. The plant-specific evaluation performed for Perrys RFW nozzles shows that for a 60-year plant life, the failure probabilities due to a LTOP event at the nozzle inner radius region and the nozzle-to-vessel weld meet the acceptance criterion in NUREG-1806. Hence, the alternative provides an acceptable level of quality and safety pursuant to 10 CFR 50.55a(z)(1) for the RFW nozzle-to-vessel welds and nozzle inner radii sections identified in Table 1.

10 CFR 50.55a Request IR-63 Page 7 of 7

This proposed alternative does not seek relief from any other aspect of the ASME Section XI Code. Therefore, if an indication is detected that exceeds ASME inspection criteria, scope expansion (extent of condition) will still be performed in accordance with ASME Section XI, subsection IWB-2430, Additional Examinations, for the Code of Record in place at the time of discovery.

6. Duration of Proposed Alternative

The proposed alternative is requested for the remainder of the fourth 10-year ISI inspection interval that ends on May 17, 2029, recognizing that the existing 40-year license expires November 7, 2026.

7. Precedent

Relief to inspect 25 percent of the RFW nozzle-to-vessel welds and inner radius regions each 10-year ISI interval was authorized for Columbia Generating Station, Docket No. 50-397, by NRC safety evaluation dated April 14, 2021 (ADAMS Accession No. ML21096A048). The relief was approved for the remainder of the 60-year plant life, including the period of extended operation.

8. References

1. BWRVIP-108-A, Technical Basis for the Reduction of Inspection Requirements for the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Blend Radii, EPRI, Palo Alto, CA: 2018, 3002013092.

2. BWRVIP-241-A, Probabilistic Fracture Mechanics Evaluation for the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Blend Radii, EPRI, Palo Alto, CA: 2018, 3002013093.

3. ASME Boiler and Pressure Vessel Code, Code Case N-702, Alternative Requirements for Boiling Water Reactor (BWR) Nozzle Inner Radius and Nozzle-to-Shell Welds,Section XI, Division 1, February 20, 2004.

4. NUREG-0619, BWR Feedwater Nozzle and Control Rod Drive Return Line Nozzle Cracking: Resolution of Generic Technical Activity A-10 (Technical Report), Revision 1, November 1980.

5. GE-NE-523-A71-0594-A, Alternate BWR Feedwater Nozzle Inspection Requirements, Revision 1, May 2000

Attachment 1 10 CFR 50.55a Request IR-063

Feedwater Nozzle Loads, Finite Element Model, and Stress Analysis

(44 pages follow)

File No.: 2001178.301 Project No.: 2001178 Quality Program Type: Nuclear Commercial

CALCULATION PACKAGE

PROJECT NAME:

Perry Feedwater Nozzle PFM Evaluation and Inspection Relief Request

CONTRACT NO.:

48309846

CLIENT: PLANT:

Energy Harbor Perry Nuclear Power Plant

CALCULATION TITLE:

Feedwater Nozzle Loads, Finite Element Model, and Stress Analys is

Document Affected Project Manager Preparer(s) &

Revision Pages Revision Description Approval Checker(s)

Signature & Date Signatures & Date 0 1 - 28 Reissue with A A-2 typographical changes in B B-13 red for Table 3 by Kevin Wong (KLW) Kevin Wong Garivalde S. Dominguez 12/22/21 12/17/21 12/17/21

Kevin Wong 12/17/21

Charles Fourcade 12/17/21

Table of Contents

1.0 OBJECTIVE................................................................................................................ 4 2.0 METHODOLOGY........................................................................................................ 4 3.0 ASSUMPTIONS.......................................................................................................... 4 4.0 DESIGN INPUTS AND LOADS................................................................................... 4 4.1 Thermal Transients......................................................................................... 4 4.2 Bounding Transients....................................................................................... 5 4.3 Mechanical Loads........................................................................................... 6 5.0 FINITE ELEMENT MODEL......................................................................................... 6 5.1 RPV Cladding.................................................................................................. 6 5.2 Removal of Weld Overlay................................................................................ 7 5.3 Mesh................................................................................................................ 7 5.4 Mesh Sensitivity Study.................................................................................... 7 6.0 STRESS ANALYSIS................................................................................................... 8 6.1 Unit Pressure Stress Analysis......................................................................... 8 6.2 Thermal Transient Stress Analysis.................................................................. 8 6.3 Piping Interface Loads Stress Analysis........................................................... 9 6.4 Critical Stress Paths........................................................................................ 9

7.0 CONCLUSION

............................................................................................................ 9

8.0 REFERENCES

.......................................................................................................... 10 COMPUTER FILES....................................................................................... A-1 HEAT TRANSFER CALCULATION.............................................................. B-1 B.1 Applied Fluid Temperatures..................................................................................... B-2 B.2 Heat Transfer Coefficients for Forced and Natural Convection............................... B-3 B.3 Steam Condensation............................................................................................... B-5

File No.: 2001178.301 Page 2 of 29 Revision: 0 F0306-01R4

List of Tables

Table 1: Pre-and Post-Power Uprate Temperatures and Flow Rate................................... 11 Table 2: Thermal Transient Definitions................................................................................ 12 Table 3: Thermal Transient Definitions................................................................................ 13 Table 4: Piping Load Summary [5]....................................................................................... 19

List of Figures

Figure 1. FW Nozzle as Modeled Dimensions [2]................................................................ 20 Figure 2. FW Nozzle Material Identifications....................................................................... 21 Figure 3. ANSYS Finite Element Model............................................................................... 22 Figure 4. Pressure Load Applied to FEM (in psi)................................................................. 23 Figure 5. Mechanical Boundary Conditions (BCs).............................................................. 24 Figure 6. Original FEM Mesh Density Pressure Stress Results (i n psi), Blend Radius........ 25 Figure 7. Refined FEM Mesh Density Pressure Stress Results (in psi), Blend Radius........ 26 Figure 8: Typical Applied Heat Transfer Coefficient and Temper ature Profile..................... 27 Figure 9: Piping Load Orientation........................................................................................ 28 Figure 10: Piping Load Application and Boundary Conditions............................................. 28 Figure 11: Stress Paths........................................................................................................ 29

File No.: 2001178.301 Page 3 of 29 Revision: 0 F0306-01R4

1.0 OBJECTIVE The objective of this calculation is to establish the design in puts and methodology, modify a previously developed finite element model (FEM) and perform stress analysis to be used in the probabilistic fracture mechanics (PFM) evaluation of reactor pressure vessel (RPV) feedwater (FW) nozzle at Perry Nuclear Power Plant (Perry).

2.0 METHODOLOGY Stress distributions due to pressure, thermal transient, and me chanical loads will be used in the subsequent PFM evaluation. A three step methodology is performed herein:

i. Loads: Establish design inputs and methodology to be used in th e FEM. A thermal transient analysis will be performed based on system thermal cy cling that occurs at the FW nozzle. Concurrent with the thermal transients, pressure an d piping interface loads are also considered. Power uprate affects the FW temperature an d rated flow per nozzle as shown in Table 1 [1].

ii. FEM: Modify a previously developed three-dimensional (3-D) fini te element model (FEM) of the Perry feedwater nozzle [2] to be used in the stress anal yses. Additional methodology for the FEM development was established in a previous loads calculation [3]

and are noted herein.

iii. Stress Analyses: Perform stress evaluation by applying loads to the FEM and obtain stress distributions for the RPV FW nozzle at the nozzle blend radius and nozzle-to-shell weld.

3.0 ASSUMPTIONS All assumptions except Assumption 5 (i.e., the assumption regar ding the RPV cladding) associated with the previous FEM calculation [2, Section 2.0] are applica ble to this calculation. All assumptions associated with the previous loads calculation [3, Section 3.0] are applicable to this calculation.

The following are additional assumptions within this calculatio n:

a. For Transient 4b (Turbine Roll), 4140 seconds is calculated by assuming a decrease rate of 100°F/hr.

b. For Transient 22 (Reactor Overpressure), a conservative approac h is assumed by omitting the steady-state temperature after 2760 seconds with indefinite time in the transient.

4.0 DESIGN INPUTS AND LOADS

4.1 Thermal Transients The thermal cycle diagrams for the FW nozzle [6] and RPV [7] were used to define the thermal transients in Table 3; in addition, power uprate conditions [1] in Table 1 are applied.

Table 2 lists the projected 60-year cycles [4.a, Tables 8 & 9] with the following notes:

For the following transients, the 60-year projected cycles are a summation of the normal and alternate transients:

o Transient 3 (Startup): Normal Startup 3-A and Alternate Startup 3-B.

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o Transient 4 (Turbine Roll): Normal Turbine Roll 4-A and Alternate Turbine Roll 4-B.

o Transient 14 (Hot Standby): Normal Shutdown - Hot Standby 14-A and Alternate Shutdown - Hot Standby 14-B.

o Transient 15 (Shutdown): Normal Shutdown - Blowdown to Condense r 15-A and Alternate Shutdown - Blowdown to Condenser 15-B.

o For Transient 16 (Shutdown, Vessel Flooding): Normal Shutdown - Vessel Flood 16-A and Alternate Shutdown - Vessel Flood 16-B.

It is assumed that there are 40 internal cycles for each OBE ev ent [4.b]. Therefore, 2 OBE events 40 internal cycles = 80 OBE cycles are evaluated for 60 years of plant operation.

4.2 Bounding Transients The transients are then screened to obtain the limiting events that bound the other transients. The limiting events are shaded in grey in Table 3. The criteria for selecting the worst cases are (a) transients that have a large temperature difference and (b) tra nsients that have a drastic rate of change in temperature. The screening method is calculated in the supp orting spreadsheet 2001178.301-FW_Bounding_Transient_Screening.xlsm, and the following transients are selected as bounding:

(1) Transient 3b (Start-up): For the FW temperature side (T FW), the event has a step-down temperature change from 400°F to 185°F (see Table 3). For the R PV temperature side (T A): the event has temperature range from 100°F to 552°F, with temperatu re rate of change of 100°F/hr.

(2) Transient 4b (Turbine Roll): For the T FW side, the event has a temperature change from 70°F to 325°F in 2 minutes duration (see Table 3).

(3) Transient 6 (Daily Reduction, 50% Power): This transient is included to segregate the high cycle yet less severe events such as Daily reduction to 50% power a nd Rod pattern change from the other thermal transients that generates high stresses.

(4) Transient 8 (Turbine Trip): For the T FW side, the event has a temperature change from 427.8°F to 90°F in 1.5 minutes duration (see Table 3).

(5) Transient 10 (Turbine Generator Trip): For the T FW side, the event has a of temperature change from 427.8°F to 275°F in 1 minute duration. For the T A side, the event has temperature change from 552°F to 350°F both in increasing and decreasing direction (see Table 3).

(6) Transient 15b (Shutdown): For the T FW side, the event has a step-down temperature change from 425°F to 200°F (see Table 3). For the T A side, the event has temperature range from 502°F to 250°F, with temperature rate of change of 100°F/hr.

(7) Transient 20 (Loss of FW pumps): For the T FW side, the event has a step-down temperature change from 485°F to 100°F (see Table 3). For the T A side, the event has temperature range from 350°F to 582°F.

(8) Transient 21 (Safety Relief Valve (SRV) Blow Down): For the T FW side, the event has a step-down temperature change from 427.8°F to 275°F and from 275°F to 100°F (see Table 3). For the TA side, the event has temperature range from 552°F to 375°F in 10 minutes duration.

(9) Transient 22 (Reactor Overpressure): For the T FW side, the event has a step-down temperature change from 427.8°F to 275°F (in 1 minute) and from 275°F to 10 0°F (in 15 minutes) (see Table 3). For the TA side, the event has temperature range from 561°F to 400°F in 15 minutes duration.

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4.3 Mechanical Loads The mechanical loads [5] are shown in Table 4. Mechanical for ces and moments are listed in kips and inch-kips, respectively. The thermal sleeve mechanical loadings [5] are small and are considered to be negligible; thus the thermal sleeve mechanical loadings shall n ot be evaluated.

5.0 FINITE ELEMENT MODEL The details of the previously developed FEM are provided in the previous FEM [2] and loads [3]

calculations:

- Geometry: Figure 1 shows the dimensions of the FEM based on nozzle drawings referenced in the previous FEM calculation [2, Section 3.1] with additional m odifications described in the subsequent sections.

- Materials: Figure 2 shows the materials used in the FEM, and th e material properties are tabulated in the previous FEM calculation [2, Table 1]. The material and material properties of the added cladding are described in the next section.

- Applied Fluid Temperatures: The calculation of applied fluid temperatures follows the methodology in the previous loads calculation [3, Section 4.5] for the thermal transients in Table

3. Detailed calculations of the applied fluid temperatures are provided in Appendix B and documented in the supporting files.

- Heat Transfer Coefficients: The calculation of heat transfer co efficients by forced convection, natural convection, and steam condensation as a function of feedwater flow and location follows the methodology in the previous loads calculation [3, Sections 4.6 and 4.7]. Detailed calculations of the heat transfer coefficients are provided in Appendix B and documented in the supporting files.

- Assumptions: All assumptions except Assumption 5 (i.e., the ass umption regarding the RPV cladding) associated with the previous FEM calculation [2, Section 2.0] are applicable to this calculation. All assumptions associated with the previous load s calculation [3, Section 3.0] are applicable to this calculation.

The FEM is a quarter model which consists of two axes of symmet ry in the longitudinal and circumferential directions of the vessel shell (Figure 5). Hen ce, a quarter model (0° to 90°) with the appropriate boundary conditions is adequate for the stress analysis of loads that were axisymmetric and resulted in the same stresses as a 360° model. For mechanical piping load which are not axisymmetric, a separate full 3-D FEM (360° model) is constructed from the quarter size model by reflecting about the symmetry planes (Figure 10). Piping inter face moments were applied at one free end of the pipe.

Two minor modifications described in the subsequent sections ar e made to the previously developed FEM for this stress analysis. The modified FEM is documented in the supporting files in Appendix A.

5.1 RPV Cladding The previously developed FEM did not model the RPV cladding [3, Section 2.0, Assumption 5], since the location of interest for the previous analysis was sufficie ntly far away from the clad regions. For this stress analysis and the subsequent PFM evaluation of the nozzle blend radius and the nozzle-to-shell weld, the RPV cladding of a nominal thickness of 3/16 inch [8, Sheet 1 of 23] has been added to the FEM at the RPV inner surface, as shown in Figure 2. The m odeling of the nominal thickness is File No.: 2001178.301 Page 6 of 29 Revision: 0 F0306-01R4

conservative with regards to stresses compared to the minimum s pecified thickness. The RPV cladding is modeled as 309/308L Austenitic Stainless Steel with material properties consistent with the Reference [2, Table 1]

5.2 Removal of Weld Overlay One Perry feedwater nozzle has a weld overlay (WOL) [9] applied at the nozzle-to-safe end weld N4C-KB. The other Perry feedwater nozzles had a mechanical stress improvement process (MSIP) applied at the nozzle-to-safe end weld [10]. The stress effects of th e WOL and MSIP are localized to the nozzle-to-safe end weld and the adjacent base metal. The regio n of interest at the nozzle blend radius and nozzle-to-shell weld for this evaluation is sufficiently far away (i.e., greater than one nozzle diameter) from the nozzle-to-safe end weld such that the effect s of the WOL or MSIP at the region of interest are negligible, based on previous WOL stress evaluations. As such, the stresses at the region of interest are essentially the same for all the Perry feedwater nozzles irrespective of the WOL or MSIP.

The previously developed FEM included the WOL at the nozzle-to-safe end weld. Since the effects of the WOL are negligible at the nozzle blend radius and the nozzle-to-shell weld, the WOL has been removed from the FEM.

5.3 Mesh A 3-D linear-elastic FEM of the FW nozzle is built using the fi nite element software ANSYS 14.5 [11].

The FEM consists of eight node SOLID45 structural elements (see Figure 3). These are converted to eight node SOLID70 thermal elements for thermal analyses. The A NSYS input files FWgeom.INP and FWgeom-2.INP can be read into ANSYS to create the FEM. ANSYS 14.5 [11] was used from Reference [2] and the same version was used to modify the FEM. For the stress analysis such as pressure stress analysis (see Section 5.4) and thermal transient and mechanical piping stress analysis (see Section 6.2 and 6.3), ANSYS 18.1 [12] is used.

5.4 Mesh Sensitivity Study In order to verify that the mesh density of the FEM is adequate to capture stresses at the blend radius and safe end, a mesh sensitivity study is performed. A unit pressure analysis is performed by applying a 1000 psi pressure to the inside surfaces of the vessel and nozzle. Cap loads are applied to the top of the vessel and attached pipe end using the equation:

P R2 P (1) i cap RR22 o i

where Ri is the inner radius of the vessel or attached piping and R o is the outer radius of the vessel or attached piping.

These cap loads are given a negative sign in order to exert tension on the model. Nodes are also coupled in the axial direction for both the top of the vessel a nd the pipe section to simulate the attached vessel and piping, respectively. Symmetry boundary conditions are applied to the vertical and horizontal cut planes to simulate parts of the vessel that ar e not modeled. Figure 4 shows the pressure load applied to the FEM and Figure 5 shows the boundar y conditions applied to the FEM.

The ANSYS input file Pressure.INP contains the loading and boundary conditions for the pressure loading.

Figure 6 shows the mesh and pressure stress results at the blen d radius for the original FEM. The original FEM has a maximum stress intensity of 50,097 psi. Figure 7 shows the mesh and pressure File No.: 2001178.301 Page 7 of 29 Revision: 0 F0306-01R4

stress results at the blend radius for the refined FEM. The re fined FEM has a maximum stress intensity of 50,513 psi.

Since the stress intensity differences for nodes at the blend radius is less than 1% for the original and refined FEMs, the original mesh (as shown in Figure 3) is suffi cient for future analyses.

The refined FEM ANSYS input files are saved in the supporting files as FWgeom-Refined.INP and Pressure-Refined.INP (see Appendix A)

6.0 STRESS ANALYSIS The developed finite element model (FEM) of the Feedwater nozzl e is used to perform thermal, pressure and mechanical piping stress analyses using the ANSYS FEA software [12]. A thermal analysis is performed for each transient defined in Section 4.2. Unit pressure and piping interface load analyses are performed separately from the thermal transient an alyses. The details of the evaluations are discussed below.

6.1 Unit Pressure Stress Analysis A 1000 psi unit pressure stress analysis is performed in Section 5.4.

6.2 Thermal Transient Stress Analysis The FEM developed in Section 5.0 is used as input for the thermal transient stress analysis using ANSYS [12]. ANSYS SOLID45 struct ural elements are switched to SOLID70 thermal elements for thermal temperature distribution analysis. For thermal stress analysis, SOLID45 structural elements are used. The boundary conditions used for the thermal stress analysis are the same as those used for the pressure stress analysis, described in Section 5.4 and show n in Figure 5.

Previously defined bounding thermal transients in Section 4.2 a re evaluated, applying heat transfer coefficients and fluid temperatures which are shown in Figure 8.

All ANSYS input files used for the thermal transient stress ana lysis, as listed below, are saved in the project computer files:

HTBC.INP: Thermal boundary regions and heat transfer application FWnoz_T#.INP:

FWnoz_T#_mntr.INP: Thermal and stress analysis input files FWnoz_S#.INP: Where # indicates the transient number

Transient Definitions:

Transient 3b: Startup Transient 4b: Turbine Roll Transient 6: Weekly Reduction, 50% Power Transient 8: Turbine Trip Transient 10: Turbine Generator Trip Transient 15b: Shutdown Transient 20: Loss of FW Pump Transient 21: SRV Blowdown Transient 22: Reactor Overpressure

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6.3 Piping Interface Loads Stress Analysis For the attached piping interface loads stress analysis, a full 3-D FEM is used rather than the 1/4 model used for the other analyses. The original geometry is modified to represent the full 3-D model.

Mechanical piping interface loads are applied to the piping face of the FEM. The mechanical loading is outlined in Table 4. Deadweight, seismic primary, seismic restrained free end and thermal restrained free end loads are applied. Design mechanical loads are for primary stress checks and are not considered for fatigue. Since seismic primary and seismic restrained free loads are identical, a single seismic load case is analyzed. The ANSYS input file FWnoz_Mechanical.inp contains the seismic and thermal piping loads and boundary conditions for th e piping stress analysis respectively.

Figure 9 shows the local coordinate system for the design forces and moments. These coordinates are modified to account for global ANSYS coordinates for use with the FEM. In order to account for the difference in applied location, M is corrected for by the equation:

M LL, appliedMHL

where, ML = Applied moment, Figure 9 H = Applied force, Figure 9 L = Length from design location R (see Figure 9) to applied location

= (162.8025-151) = 11.8025 inch

where 162.8025 inches is obtained from the finite element model : from the center of the RPV to the FW piping end and 151 inch is obtained from Reference [5, Sheet 5].

Loads are applied using the ANSYS TARGE170 target element type to create a pilot node at the free end of the nozzle. The ANSYS CONTA175 contact element type is u sed to create a contact surface at the free end of the nozzle. The pilot node and surface are bond ed together, so that the moment applied to the pilot node is transferred to the nozzle free end.

6.4 Critical Stress Paths The critical stress paths are selected for two locations on the feedwater nozzle. Two paths (P1 and P2) are chosen at the blend radius, and two paths (P3 and P4) are chosen at the nozzle-to-shell weld.

Detailed nodal location of paths are as follows:

Path 1, P1: nozzle-to-vessel shell blend radius, x-z plane, from node 113476 to node 114562 Path 2, P2: nozzle-to-vessel shell blend radius, y-z plane, from node 113521 to node 115123 Path 3, P3: nozzle-to-vessel shell weld, x-z plane, from node 1 31478 to node 135428 Path 4, P4: nozzle-to-vessel shell weld, y-z plane, from node 1 31469 to node 135395

The stress extraction is generat ed under RSYS = 0, where Path 1 and 3 results have the hoop stress in Y-direction and Path 2 and 4 results have the hoop stress X-dir ection.

7.0 CONCLUSION

The FW nozzle finite element model is developed, and pressure, thermal transients, and mechanical piping loads are applied to perform stress analyses. The stress results for each of the defined stress paths will be used in the subsequent PFM evaluation.

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8.0 REFERENCES

1. GE Design Specification No. 26A5308, Revision 3, Reactor Vessel - Power Uprate, SI File No.

2001178.206.

2. SI Calculation No. 1300310.304, Revision 0, Perry Feedwater Nozzle Finite Element Model.

3. SI Calculation No. 1300310.305, Revision 0, Feedwater Nozzle Design Loads Calculation for use in Environmental Fatigue Analysis.

4. Projected 60-Year Cycles

a. Structural Integrity Associates Calculation No. 2001140.301, Re vision 1, Fatigue Update for Perry Nuclear Power Plant from 10/1/2016 to 10/31/2021.

b. Email from J. Zbiegien (Energy Harbor) to K. Wong (SI),

Subject:

RE: [EXTERNAL] RE:

Projected Cycles, December 3, 2021, 7:44AM, SI File No. 200117 8.207.

5. GE Design Specification Data Sheet No. 22A5536AF, Revision 1, Safe End, Feedwater Nozzle, SI File No. 0800184.202.

6. Feedwater Thermal Cycle Diagrams

a. Energy Harbor Drawing No. 306-0081-00000, Revision E, Feedwate r Temperature/Pressure Cycles, SI File No. 2001178.204.

b. Energy Harbor Drawing No. 306-0082-00000, Revision D, Feedwate r Temperature/Pressure Cycles, SI File No. 2001178.204.

7. Reactor Thermal Cycle Diagrams

a. Energy Harbor Drawing No. 08-0037-00001 (General Electric Drawing No. 762E458, Revision 7, Sheet 1, Reactor Cycles), SI File No. 20001178.20 3.

b. Energy Harbor Drawing No. 08-0037-00002 (General Electric Drawing No. 762E458, Revision 7, Sheet 2, Reactor Cycles), SI File No. 2001178.203.

c. Energy Harbor Drawing Update Notice No. 08-0596-001-001, Revision 0, Reactor Cycles, SI File No. 2001178.203.

d. Energy Harbor Drawing Update Notice No. 08-0596-001-003, Revision 0, Reactor Cycles, SI File No. 2001178.203.

8. CBI Nuclear Company Section No. 238-D11.3, Revision 5, Perry 1 - 238 BWR 6 Vessel Contract 73-C108, Stress Report (Code), Water Lev. Instr. Nozzle Design Report - Section D 11.3, SI File No. 2001178.212.

9. Energy Harbor IDCN No. 305-006-108-995013, Revision 1, RV. Fee dwater Nozzle Weld Arrangement, SI File No. 2001178.211.

10. Energy Harbor Calculation No. EA-0264, Revision 1, Flaw Evaluation, Feedwater Nozzle B13-N4E-KB, SI File No. 2001178.210.

11. ANSYS Mechanical APDL and PrepPost, Release 14.5 (w/Service Pack 1), September 2012.

12. ANSYS Mechanical APDL (UP20170403) and Workbench (March 31, 201 7), Release 18.1, SAP IP, Inc.

File No.: 2001178.301 Page 10 of 29 Revision: 0 F0306-01R4

Table 1: Pre-and Post-Power Uprate Temperatures and Flow Rate

Parameter Pre-Power Uprate Post Power Uprate

100% Rated FW Flow/Nozzle (gpm) 7410* [1, sheet 35 of T4] 6555 [1]

FW Temperature (°F) 420 [1] 427.8 [1]

Note:

• This is the as-analyzed flow rate; the actual flow rate before power uprate was lower than the power uprate value of 6555 gpm.

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Table 2: Thermal Transient Definitions Event Description 60-Year Projected Cycles

[4.a, Tables 8 & 9]

1 Boltup 37 2 Design hydrotest 47 3 Startup 166 + 10 (Note 1) 4 Turbine roll 167 +1 (Note 1) 5 Daily reduction 685 6 Weekly reduction 436 7 Rod pattern change 685 8 Turbine trip 22 9 Partial feedwater heater bypass 159 10 Turbine generator trip 13 11 Other scrams 87 13 Reduction to 0% power 209 14 Hot standby 127 + 4 (Note 1) 15 Shutdown 165 + 4 (Note 1) 16 Shut down, vessel flooding 160 + 10 (Note 1) 17 Shutdown 183 18 Unbolt 36 19 Refueling 33 20 Composite Loss of Feedwater 20 21 SRV blowdown 2 22 Reactor Overpressure 1 23 Automatic Blowdown 1 24 Improper Start of Cold Recirc Loop 1 25 Sudden Start of Pump in Cold Recirc Loop 4 27 Pipe Rupture 1

- OBE 2 x 40 (Note 2)

Notes:

1. For the noted transients, the 60-year projected cycles are a su mmation of the normal and alternate transients:

a. Transient 3 (Startup): Normal Startup 3-A and Alternate Startup 3-B.

b. Transient 4 (Turbine Roll): Normal Turbine Roll 4-A and Alternate Turbine Roll 4-B.

c. Transient 14 (Hot Standby): Normal Shutdown - Hot Standby 14-A and Alternate Shutdown - Hot Standby 14-B.

d. Transient 15 (Shutdown): Normal Shutdown - Blowdown to Condenser 15-A and Alternate Shutdown - Blowdown to Condenser 15-B.

e. For Transient 16 (Shutdown, Vessel Flooding): Normal Shutdown - Vessel Flood 16-A and Alternate Shutdown - Vessel Flood 16-B.

2. It is assumed that there are 40 internal cycles for each OBE event [4.b]. Therefore, 2 OBE events x 40 internal cycles = 80 OBE cycles are evaluated for 60 years of p lant operation.

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Table 3: Thermal Transient Definitions

Event Event Time, FW Temp RPV Temp RPV Time FW Number** Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 70 70 0 0 1 Bolt-up 1080 100 70 0 1080 0 2880 100 70 0 1800* 0 0 100 100 0 0 1 180 180 1250 1 0 2 Design Hydrotest 1801 180 180 1250 1800* 0 1802 100 100 20 1 0 3602 100 100 20 1800* 0 Startup 0 70 100 20 0 3a (CU heated 1080 100 100 20 1080 0 after 17352 100 552 1051 16272 2 initiated) 19152 100 552 1051 1800* 2 0 70 100 20 0 1080 100 100 20 1080 0 Startup 2880 100 100 20 1800* 0 3b** (CU heated 13680 400 400 1051 10800 0 before 13681 185 400.03 1051 1 2 initiated) 19152 185 552 1051 5471 2 20952 185 552 1051 1800* 2 0 100 552 1051 2 Turbine 1800 100 552 1053 1800 33 Roll (CU 1920 325 552 1055 120 33 4a heated 2520 325 552 1066 600 33 after initiated) 4320 427.8 552 1100 1800 100 6120 427.8 552 1100 1800* 100 0 185 552 1051 2 Turbine 1800 185 552 1051 1800* 2 Roll (CU 5940 70 552 1051 4140 33 4b** heated 6060 325 552 1053 120 33 before 6660 325 552 1065 600 33 initiated) 8460 427.8 552 1100 1800 100

10260 427.8 552 1100 1800* 100

• Assumed for stress analysis

• • Selected as bounding transient

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Table 3: Thermal Transient Definitions (continued)

Event Event Time, FW Temp RPV Temp RPV Time FW Number Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 427.8 552 1100 100 Weekly 1800 427.8 552 1100 1800* 100 6** Reduction, 1950 360 552 1063 150 50 50% Power 3750 360 552 1063 1800* 50 3900 427.8 552 1100 150 100 5700 427.8 552 1100 1800* 100 0 427.8 552 1100 100 Rod 1800 427.8 552 1100 1800* 100 7 Pattern 2700 394 552 1078 900 75 Change 4500 394 552 1078 1800* 75 5400 427.8 552 1100 900 100 7200 427.8 552 1100 1800* 100 0 427.8 552 1100 100 1800 427.8 552 1100 1800* 100 1890 90 552 1100 90 100 8** Turbine 2490 90 552 1100 600 100 Trip 2610 325 552 1100 120 100 3210 325 552 1100 600 100 5010 427.8 552 1100 1800 100 6810 427.8 552 1100 1800* 100 0 427.8 552 1100 100 Partial 1800 427.8 552 1100 1800* 100 9 FWH 1890 265 552 1100 90 100 Bypass 3690 265 552 1100 1800* 100 3870 427.8 552 1100 180 100 5670 427.8 552 1100 1800* 100

• Assumed for stress analysis

• • Selected as bounding transient

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Table 3: Thermal Transient Definitions (continued)

Event Event Time, FW Temp RPV Temp RPV Time FW Number Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 427.8 552 1100 100 1800 427.8 552 1100 1800* 110 1810 396 565 1175 10 110 1815 384 565 1175 5 110 1830 347 538 991 15 110 1860 275 537 987 30 110 Turbine 2760 185 512 871 900 2 10** Generator 8583 185 350 120 5823 2 Trip 12183 185 450 585 3600 2 15753 185 549 1035 3570 33 15855 255 552 1050 102 33 15873 325 552 1050 18 33 16473 325 552 1053 600 33 18273 427.8 552 1100 1800 100 20073 427.8 552 1100 1800* 100 0 427.8 552 1100 100 1800 427.8 552 1100 1800* 110 1815 384 538 991 15 110 1860 275 539 985 45 110 2760 185 564 869 900 2 8568 185 350 120 5808 2 11 SCRAM All Others 12183 185 450 585 3615 2 15753 185 549 1035 3570 33 15855 255 552 1050 102 33 15873 325 552 1050 18 33 16473 325 552 1053 600 33 18273 427.8 552 1100 1800 100 20073 427.8 552 1100 1800* 100 Reduction 0 400 552 1100 100 13 to 0% 1800 200 552 1053 1800 2 Power 3600 200 552 1050 1800* 2 0 200 552 1050 2 14a Hot 1800 200 552 1050 1800* 2 Standby 1860 100 552 1050 60 2 3660 100 552 1050 1800* 2

• Assumed for stress analysis

• • Selected as bounding transient

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Table 3: Thermal Transient Definitions (continued)

Event Event Time, FW Temp RPV Temp RPV Time FW Number Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 200 552 1050 2 1800 200 552 1050 1800* 2 1920 435 552 1050 120 2 14b Hot 2760 435 552 1050 840 0 Standby 2761 200 552 1050 1 2 4561 200 552 1050 1800* 2 4681 435 552 1050 120 0 KLW 12/22/21 6481 435 552 1050 1800* 0 0 100 552 1050 2 15a Shutdown 1800 100 552 1050 1800* 2 12672 100 250 33 10872 2 14472 100 250 33 1800* 2 0 435 552 1050 0 1800 435 552 1050 1800* 0 2160 425 542.0 783.3 360 0 2161 200 542.0 782.6 1 3 15b** Shutdown 3121 200 515.3 243.7 960 3 3241 395 512.0 188.4 120 0 3601 385 502.0 34.0 360 0 3602 200 502 33 1 3 12672 200 250 33 1800* 3 0 200 250 33 3 1800 200 250 33 1800* 3 1801 40 250 33 1 0 2311 40 236 33 510 0 16b Vessel 2312 125 236 33 1 0 Flooding 2822 125 222 33 510 0 2823 100 222 33 1 25 3423 100 205 33 600 25 3424 125 205 33 1 0 5224 125 205 33 1800* 0 0 125 205 33 0 1800 125 205 33 1800* 0 17 Shutdown 2700 100 179 33 900 0 5580 100 100 33 2880 0 7380 100 100 33 1800* 0

• Assumed for stress analysis, **Selected as bounding transient File No.: 2001178.301 Page 16 of 29 Revision: 0 F0306-01R4

Table 3: Thermal Transient Definitions (continued)

Event Event Time, FW Temp RPV Temp RPV Time FW Number Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 100 70 33 0 18 Unbolting 1800 100 70 33 1800* 0 2880 70 70 33 1080 0 4680 70 70 33 1800* 0 0 427.8 552 1100 100 1800 427.8 552 1100 1800* 100 1803 424 582 1335 3 0 1813 438 561 1125 10 0 1815 440 561 1125 2 0 1920 582 554 1067 105 0 2340 525 525 833 420 0 2700 573 567 1180 360 0 2710 562 561 1125 10 0 2820 561 561 1125 110 0 3600 561 561 1125 780 0 20** LOFP 4020 490 490 607 420 0 4500 573 567 1180 480 0 4510 572 561 1125 10 0 4620 561 561 1125 110 0 5400 561 561 1125 780 2 5580 530 538 961 180 2 6000 485 485 579 420 2 6001 100 350 120 1 33 9601 100 518 898 3600 33 9721 325 524 923 120 44 KLW 10321 427.8 552 1053 600 100 12/22/21 12121 427.8 552 1100 1800* 100 0 427.8 552 1100 100 1800 427.8 552 1100 1800* 110 21** SRV 1860 275 534 1093 60 110 Blowdown 2400 170 375 1035 540 3 2760 100 364 996 360 3 11760 100 100 20 1800* 3

• Assumed for stress analysis

• • Selected as bounding transient

File No.: 2001178.301 Page 17 of 29 Revision: 0 F0306-01R4

Table 3: Thermal Transient Definitions (continued)

Event Time, FW Temp RPV Temp RPV Time FW Number Event Name seconds (H to B), (D to G), Pressure, Increment, Flow

°F °F psig seconds (%)

0 427.8 552 1100 100 1800 427.8 552 1100 1800* 110 1802 422.7 594 1510 2 110 1832 346.3 561 1175 30 110 22** Reactor Overpressure 1860 275 561 1175 28 2 2760 100 400 240 900 2 2880 325 552 1050 120 33 3480 325 552 1053 600 33 5280 325 552 1100 1800* 100 0 427.8 552 1100 100 1800 427.8 552 1100 1800 110 23 Automatic Blowdown 1860 275 375 1053 60 110 2760 100 100 709 900 3 4560 100 100 20 1800* 3 Improper 0 394 552 1100 75 24 Start of Cold 1800 394 552 1100 1800* 75 Recirculation 2700 427.8 552 1100 900 100 Loop 4500 427.8 552 1100 1800* 100 Sudden Start 0 394 552 1100 75 of Pump in 1800 394 552 1100 1800* 75 25 Cold 2700 427.8 552 1100 900 100 Recirculation Loop 4500 427.8 552 1100 1800* 100 Pipe Rupture 0 427.8 552 1100 100 27 and 1800 427.8 552 1100 1800* 100 KLW Blowdown 1815 259 259 1091 15 0 12/22/21 3630 259 259 20 1800* 0

--- Leak Test 0 100 100 0 0 1 100 100 400 1 0

• Assumed for stress analysis

• • Selected as bounding transient

File No.: 2001178.301 Page 18 of 29 Revision: 0 F0306-01R4

Table 4: Piping Load Summary [5]

Nozzle End Loads [5, Sheet 5]

H M Radius Load (kips) (inch- (inch) kips)

Deadweight 12.7 353 151

Primary Seismic 27.4 494 151

Thermal Restrained Free 58.8 1470 151

File No.: 2001178.301 Page 19 of 29 Revision: 0 F0306-01R4

Notes: 1. This dimension is to the center of the weld.

2. As modeled dimension.

3. The length of straight pipe is 2.70 at the OD and 2.69 at the ID. 2.70 is modeled for both.

Figure 1. FW Nozzle as Modeled Dimensions [2]

File No.: 2001178.301 Page 20 of 29 Revision: 0 F0306-01R4

Cladding (Type 309/308L)

Vessel (SA-533, Gr. B, Cl. 1)

Nozzle Forging (SA-508, Gr. 2, Cl. 2)

Safe End Weld (Ni-Cr-Fe, Inconel 182)

Safe End (SA-508, Gr. 1)]

Cladding (Type 309/308L)

FW Piping (SA-508, Gr. 1)

Figure 2. FW Nozzle Material Identifications

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Figure 3. ANSYS Finite Element Model

File No.: 2001178.301 Page 22 of 29 Revision: 0 F0306-01R4

RPV End Cap Load (blue)

Internal Pressure (red)

Piping End Cap Load (yellow)

Figure 4. Pressure Load Applied to FEM (in psi)

File No.: 2001178.301 Page 23 of 29 Revision: 0 F0306-01R4

Symmetry BC

Couple

Symmetry BC

Symmetry BC

Couple

Figure 5. Mechanical Boundary Conditions (BCs)

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Figure 6. Original FEM Mesh Density Pressure Stress Results (i n psi), Blend Radius

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Figure 7. Refined FEM Mesh Density Pressure Stress Results (in psi), Blend Radius

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Figure 8: Typical Applied Heat Transfer Coefficient and Temper ature Profile

(Top figure shows heat transfer coefficient in BTU/(sec-in2-°F), lower figure shows applied temperature in °F.)

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Figure 9: Piping Load Orientation

Figure 10: Piping Load Application and Boundary Conditions

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P1 Z P2

P3 P4

Y X

Figure 11: Stress Paths

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COMPUTER FILES

File No.: 2001178.301 Page A-1 of A-2 Revision: 0 F0306-01R4

This document contains Structural Integrity Associates; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

LOADS CALCULATION

Filename Description

21001178.301-HTCs.xlsm Applied fluid temperature and heat transfer coefficients calculation in spreadsheet 21001178.301-Bounding_Transient Screening.xlsm Bounding Transient screening in spreadsheet

FEM AND STRESS ANALYSIS

Filename Description FWGEOM.INP FWGEOM-2.INP FW Nozzle finite element model FWGEOM-Refined.INP FWGEOM-2-Refined.INP FW Nozzle finite element model using refined mesh Pressure.INP FW Nozzle model under internal pressure Pressure-Refined.INP FW Nozzle model under internal pressure using refined mesh FWnoz_Mechanical.INP FW Nozzle model under mechanical piping load FWnoz_T#.INP APDL input file to perform thermal transient analysis FWnoz_T#_mntr.INP APDL input file to define load steps for thermal stress analysi s FWnoz_S#.INP APDL input file to perform thermal stress analysis HTBC.INP APDL input file to apply heat transfer coefficients for thermal transient analysis

BC.inp Structural boundary condition for thermal transient anal ysis

GenStress.mac Macro that generates polynomial coefficients of stress for all paths.

GetPath.mac Macro that generates path in ANSYS APDL GETPATH.TXT Text File that defines stress paths Stress outputs*

1. = Pressure (Unit Pressure run)

FWNOZ_#_COE _%.CSV = Mechanical (Nozzle Unit Piping Load)

FWNOZ_#_MAP _%.CSV = S3a, S4b, S6, S8, S10, S15b, S20, S21, S22 (Thermal Transient run)

% = (Paths) P1, P2, P3, P4

File No.: 2001178.301 Page A-2 of A-2 Revision: 0 F0306-01R4

This document contains Structural Integrity Associates; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

HEAT TRANSFER CALCULATION

File No.: 2001178.301 Page B-1 of B-13 Revision: 0 F0306-01R4

B.0 OBJECTIVE The objective of this appendix is to implement the methodology of Reference [3] in calculating the applied fluid temperatures and heat transfer coefficients for each region of the FEM.

B.1 Applied Fluid Temperatures Fluid temperature boundaries are defined [3] using the lettered nodes in Figure B-1 as guidance. The following methodology is used in defining temperatures within the following regions in Figure B-1:

H-M: T = TFW, FW fluid temperature as defined on the FW nozzle thermal cycl e diagram, °F

M-A & A-B: )T (BTFW T( FWC5)(TA-1) where:

TA = RPV Region A fluid temperature from the RPV thermal cycle diagram, °F C5 = 0

Therefore, for these regions, T = T FW

B-C: This region is assumed to be insulated

Point C: )T (BTFW T( 2C1 FW')(C)(TA-2) where:

C2 = 1.0 C1 = 0.45 for 20% - 100% FW flow and 1.0 for 0% FW flow (linearly interpolate for flows in between 0% and 20%)

Point D: T = TA

C-D: For model simplicity, the temperature between point C and D is assumed to be the average of the feedwater flow and vessel temperature in Region A. For transients with less than 20% rated feedwater flow, the temperature is assumed uniform at the vessel Region A temperature. The average temperature approach is slig htly conservative when compared to the specified linear temperature transition because a larger thermal gradient will be applied to both ends of the C-D region wheneve r the temperatures at Point C and D differ. Therefore, Table 3 applies the average t emperature in this region.

If the stress analyst desires, a linear transition is also acce ptable.

D-G: Same as point D

N-O: T = 100°F

File No.: 2001178.301 Page B-2 of B-13 Revision: 0 F0306-01R4

B.2 Heat Transfer Coefficients for Forced and Natural Convectio n Heat transfer coefficients (HTCs) are defined for regions of the model based on the thermal boundaries identified in Figure B-1. HTCs for forced flow conditions are calculated with power uprate modifications applied as appropriate. The HTCs are calculated for a range of temperatures and flow rates. Table B-3 gives the results for varying temperatures and flow rates. Maximum HTCs for each region are printed in bold. Bounding natural convection heat transfer coefficients are also presented.

The following methodology is used when defining the HTCs for ea ch of the FW nozzles regions:

Forced Convection:

H-I: h values based on cross-sectional properties and curve values, which are a function of temperature:

h (Bf (T) -3)

DA 0.8 DF

where:

f(T) = Temperature dependent factor D = Hydraulic Diameter, ft A = Cross-sectional flow area, ft 2 F = Volumetric flow rate, gpm

The HTC at 100% flow is given as h/(% flow/100)0.8 for various temperatures.

Therefore, h values at other flow rates can be calculated as:

h = (% flow /100)0.8 [ h/(% flow/100)0.8] (B-4)

These h values are corrected for surface roughness of the pipe accordi ng to the following formula:

frough h ' (Bh -5) f smooth

The ratio (frough/fsmooth) ranges from 1.1 to 1.3 depending on temperature and flow rate.

The computed h values conservatively have a roughness factor of 1.3 applied to all flows at all temperatures. After the roughness factor is appli ed to the original h values, hfinal is computed as follows:

0.8 hh gpm ' 100 7410% 6555 (B-6) final Flow gpm

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I-J: It is assumed that the HTCs for this region are the average bet ween Region H-I and Region J-K [3, Assumption 3.5].

J-K & L-M: The HTC is in the same format shown in Equation B-4. These re gions were analyzed for a leakage flow of 15 gpm at 25% pressure drop (100% flow). The reduced FW flow rate is assumed to have negligible impact on the leakage flow rate [3, Assumption 3.4].

Therefore, the original heat transfer formula based on percent flow is valid for power uprate flow rates. Thus, hfinal is computed as follows:

hfinal = (% flow/100)0.8 [ h/(% flow/100)0.8] (B-7)

K-L: This region is defined by the same h values as Regions J-K and L-M, as modified by the following equation for the insulating effect of the primary sea l:

h 1 (B-8) effective 1L K h seal

Where:

L = Seal thickness = 0.06 in = 0.005 ft K = Thermal conductivity of seal = 9.5 BTU/hr-ft-°F

M-A: It is assumed that the HTCs for this region are the average be tween Region L-M and Region A-B [3, Assumption 3.5].

A-B & C: Power uprate flow rates are used in this computation. h' is 900 Btu/(hr-ft 2-F) for Region A-B and C. The maximum of h1 and h2 below are used to then calculate the HTC for this region.

0.8 Z @temp _ of _ annulus %Flow 6555gpm h o 1 h (B'-9)

Z @100 F 100 7410gpm

Z @temp _ of _ annulus '

h o 2 100 (B-10) 0 h.28*

Z @ F

1 KP 3 Z (B-11) r

0.8 Where

K = Thermal Conductivity, Btu/hr-ft-°F Pr = Prandtl Number

= Kinematic Viscosity = Dy namic Viscosity / Density, ft 2/hr

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Due to the high evaluated leakage rate, the inner annulus temperature is assumed to be equivalent to the Feedwater flow temperature [3, Assumption 3.12]).

B-C: This region is assumed to be insulated.

C-E: It is assumed that the HTCs for this region are the average between Point C and Region E-G [3, Assumption 3.6]

E-G: Similar to Region A-B, the HTCs for this region are computed f rom the same equations but using an h value of 3500 Btu/hr-ft 2-F.

O-N: This region has a constant HTC of 0.2 BTU/hr-ft 2-°F.

Natural Convection:

The maximum possible natural convection coefficient is 1,000 BTU/ hr-ft 2-°F. This HTC shall be used for all FW flows at 0% of rated flow, except as noted below.

K-L: This region has HTC of 655 BTU/hr-ft 2-°F.

A-B & C: The HTCs for this region and point are computed using a constant flow rate of 20% rated flow.

B-C: This region is assumed to be insulated.

C-E: It is assumed that the HTCs for this region are the average between Point C and Region E-G [3, Assumption 3.6]

E-G: Similar to Region A-B, the HTCs for this region are computed f rom the same equations but using an h' value of 3500 Btu/hr-ft 2-F.

O-N: This region has a constant HTC of 0.2 BTU/hr-ft 2-°F.

B.3 Steam Condensation During the loss of FW pumps transient, there are times when the flow is zero and vessel water level drops below the FW nozzle, exposing the nozzle to steam. Since the nozzle and thermal sleeve are initially below the saturation temperature, laminar film conden sation heat transfer occurs on the nozzle and thermal sleeve surfaces. As condensation occurs in the ann ular gap between the thermal sleeve and the nozzle, the condensate w ill flow downward and out of the gap. Additional steam will flow into the gap to replace the condensate. Therefore, the thermal slee ve can be assumed to have no effect on the heat transfer at the nozzle surface behind the gap. The av erage HTC for film condensation in horizontal tubes for steam condensation (known as the Kern Corr elation) is given as:

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3 ' 1/ 4 h D0.612g (B-12) l(lv)klhfg (T T )

l g w

Where:

g = Acceleration due to gravity, 4.17 x 10 8 ft/hr2

= Mass density of liquid at average temperature, lbm/ft 3

= Mass density of vapor, lbm/ft 3 kl = Conductivity of liquid at average temperature, Btu/hr-ft-F h'fg = hfg + 0.68cp(Tg - Tw), Btu/lbm hfg = Heat of condensation at vapor temperature, Btu/lbm cp = Specific heat of liquid at average temperature, Btu/lb-F Tg = Saturated vapor temperature = T final, F Tw = Wall temperature = T initial, F

= Dynamic viscosity of liquid at average temperature, lbm/ft-hr D = Inner diameter of horizontal pipe, ft

Table B-1 and Table B-2 list selected properties of saturated w ater and saturated steam. Steam properties are interpolated at T g, and water properties are interpolated at T f, which is taken as the average of Tg and Tw. Table B-4 shows the h value calculation for steam condensation at different surface and steam temperatures for a nozzle inner diameter of 12.00 inches. The maximum h value for steam condensation is 1024 BTU/hr-ft 2-°F. Therefore, for time steps during the zero flow conditions of Transient 20 where the FW nozzle is filled with s team, a conservative HTC of 1050 BTU/hr-ft2-°F is applied to all interior surfaces from Points H to G.

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Table B-1: Thermo-Physical Properties for Saturated Water [3]

Water Property Value at Fluid Temperature, T

Temperature,

°F 40 50 70 100 200 300 400 500 550 600 Thermal Conductivity, 0.325 0.332 0.347 0.364 0.394 0.395 0.381 0.349 0.325 0.292 k,

Btu/hr-ft-°F Specific Heat, cp, 1.00 1.00 0.998 0.998 1.00 1.03 1.08 1.19 1.31 1.51 Btu/lbm-°F Density,

, 62.4 62.4 62.3 62.0 60.1 57.3 53.6 49.0 45.9 42.4 lbm/ft3 Volumetric Rate of 2.0 4.9 12 20 40 60 80 100 110 120 Expansion,

, 10-5 Dynamic Viscosity, 104 88 65.8 45.8 20.5 12.6 9.1 7.1 6.4 5.8 ft3/ft3-°F, 10-5 Prandtl Number, 11.60 9.55 6.82 4.52 1.88 1.18 0.927 0.87 0.93 1.09 Pr

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Table B-2: Thermal Properties for Saturated Steam [3]

Temperature Specific Density Enthalpy Tg, °F Volume g, lbm/ft3 hfg, Btu/lbm vg, ft3/lbm 500 0.6761 1.4791 714.8 510 0.6153 1.6252 701.3 520 0.5605 1.7841 687.3 530 0.5108 1.9577 672.7 540 0.4658 2.1468 657.5 550 0.4249 2.3535 641.6 560 0.3877 2.5793 625.0 570 0.3537 2.8273 607.6 580 0.3225 3.1008 589.3 590 0.2940 3.4014 570.1

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Table B-3: Forced Convection Heat Transfer Coefficients, BTU/h r-ft2-°F [3]

Thermal Fluid % Feedwater Flow Boundary Temperature

°F 5 25 50 75 100 110 40 192 696 1212 1676 2110 2277 50 192 696 1212 1676 2110 2277 100 254 921 1604 2219 2793 3014 H to I 200 358 1299 2261 3127 3936 4248 300 424 1536 2674 3698 4655 5024 400 455 1648 2870 3970 4997 5393 500 461 1672 2911 4026 5068 5469 552 455 1648 2870 3970 4997 5393 40 584 2116 3684 5096 6415 6923 50 584 2116 3684 5096 6415 6923 100 772 2798 4871 6738 8482 9154 I to J 200 1089 3948 6874 9508 11968 12916 300 1288 4667 8126 11239 14148 15269 400 1382 5007 8718 12058 15179 16381 500 1403 5083 8850 12241 15409 16630 552 1382 5007 8718 12058 15179 16381 40 976 3536 6157 8516 10720 11569 50 976 3536 6157 8516 10720 11569 J to K 100 1290 4674 8139 11257 14170 15293 and L to 200 1821 6598 11487 15888 20000 21585 M 300 2152 7798 13578 18780 23640 25513 400 2308 8366 14565 20146 25360 27369 500 2344 8494 14789 20456 25750 27790 552 2308 8366 14565 20146 25360 27369 40 645 1236 1452 1553 1614 1632 50 645 1236 1452 1553 1614 1632 100 768 1351 1540 1626 1675 1690 K to L 200 930 1475 1630 1697 1735 1746 300 1009 1528 1667 1725 1759 1768 400 1042 1548 1681 1736 1768 1777 500 1049 1553 1684 1739 1769 1778 552 1042 1548 1681 1736 1768 1777 40 576 1862 3242 4484 5645 6092 50 576 1862 3242 4484 5645 6092 100 771 2472 4304 5953 7493 8087 M to A 200 1099 3501 6095 8430 10612 11452 300 1306 4145 7217 9983 12566 13562 400 1406 4452 7752 10722 13497 14566 500 1429 4521 7872 10888 13706 14792 552 1404 4450 7748 10717 13490 14559

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Table B-3: Forced Convection Heat Transfer Coefficients, BTU/hr-ft2-°F [3] (continued)

Thermal Fluid % Feedwater Flow Boundary Temperature

°F 5 25 50 75 100 110 40 176 188 327 452 569 614 50 176 188 327 452 569 614 100 252 269 469 648 816 881 A to B 200 378 404 703 972 1223 1320 and Point 300 461 492 857 1185 1492 1610 C 400 504 539 938 1297 1633 1762 500 513 548 955 1321 1662 1794 552 500 535 931 1287 1620 1749 600 485 518 902 1247 1570 1695 B to C All Insulated 50 430 459 799 1105 1392 1502 100 616 658 1146 1584 1994 2152 200 924 986 1717 2375 2990 3227 C to E 300 1126 1203 2095 2897 3647 3936 400 1233 1317 2293 3171 3992 4308 500 1255 1341 2334 3228 4064 4386 600 1193 1266 2204 3049 3838 4142 50 684 730 1271 1759 2214 2389 100 980 1047 1822 2521 3173 3424 200 1469 1569 2732 3779 4757 5134 E to G 300 1792 1914 3332 4609 5802 6261 400 1961 2095 3648 5045 6351 6854 500 1997 2133 3713 5136 6465 6977 600 1886 2014 3507 4851 6106 6590 O to N All 0.2 0.2 0.2 0.2 0.2 0.2

Table B-4: Steam Condensation Heat Transfer Coefficients, BTU/ hr-ft2-°F [3]

h g 4.17E+08= 0.612* {l (l - v )g kl 3 h'fg /[d(Tg - Tw )]}1/4 [29]

h'fg = h fg + 0.68cp (Tg - Tw ), [29]

D (in )= 12.00 Tg v, lb m/ ft3 h fg, Btu/lbm Tf, °F l, lb m/ ft3 cp, Btu/lbm-°F l, lbm/ft-hr kl, Bt u / h r-ft -°F Tw,°F h 'fg, Bt u / lb m h, Bt u / h r-ft ² -°F 561 2.6041 623.3 494.4 49.3 1.184 0.260 0.351 427.8 730.5 592 561 2.6041 623.3 525.5 47.3 1.271 0.244 0.334 490 684.6 654 561 2.6041 623.3 556.5 45.3 1.370 0.229 0.316 552 631.6 1024

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Table B-5: Bounding Transients Heat Transfer at Each Time Step

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Table B-5: Bounding Transients at Each Time Step (continued)

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Figure B-1: FW Nozzle Thermal Boundaries [3]

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Attachment 2 10 CFR 50.55a Request IR-063

Probabilistic Fracture Mechanics Evaluation for Perry Feedwater Nozzle

(30 pages follow)

File No.: 2001178.302 Project No.: 2001178 Quality Program Type: Nuclear Commercial

CALCULATION PACKAGE

PROJECT NAME:

Perry Feedwater Nozzle PFM Evaluation and Inspection Relief Request

CONTRACT NO.:

48309846

CLIENT: PLANT:

Energy Harbor Perry Nuclear Power Plant

CALCULATION TITLE:

Probabilistic Fracture Mechanics Evaluation for Perry Feedwater Nozzle

Document Affected Project Manager Preparer(s) &

Revision Pages Revision Description Approval Checker(s)

Signature & Date Signatures & Date 0 1 - 24 Reissue with A A-4 typographical changes B B-2 in red on page 11 by Kevin Wong (KLW) 12/22/21 Kevin Wong 12/17/21 Kevin Wong 12/17/21

Garivalde Dominguez 12/17/21

Table of Contents

1.0 INTRODUCTION

......................................................................................................... 4 2.0 METHODOLOGY........................................................................................................ 4 2.1 Assumptions.................................................................................................... 5 3.0 DESIGN INPUTS........................................................................................................ 6 3.1 Stress Analysis................................................................................................ 6 3.2 Deterministic Parameters................................................................................ 6 3.2.1 RPV Dimensions............................................................................................. 6 3.2.1 Normal Operating Conditions and LTOP event conditions.............................. 6 3.2.2 Transients and Projected 60-Year Cycles....................................................... 6 3.2.3 Piping Loads.................................................................................................... 7 3.2.4 In-Service Inspection....................................................................................... 7 3.2.5 Weld Residual Stresses.................................................................................. 8 3.3 Random Variables........................................................................................... 8 3.3.1 Materials, Material Chemistry, and Fracture Toughness................................. 8 3.3.2 Fluence............................................................................................................ 8 3.3.3 Stress Corrosion Cracking Initiation................................................................ 8 3.3.4 Fatigue Crack Growth..................................................................................... 9 4.0 PROBABILISTIC FRACTURE MECHANICS EVALUATION.................................... 10 5.0 RESULTS AND CONCLUSIONS.............................................................................. 11

6.0 REFERENCES

.......................................................................................................... 12 DETERMINISTIC FRACTURE MECHANICS EVALUATION........................ A-1 SUPPORTING FILES.................................................................................... B-1

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List of Tables

Table 1: Bounding Thermal Transients Cycles.................................................................... 15 Table 2: Random Variables Parameter Summary................................................................ 16 Table 3: Deterministic Parameter Summary........................................................................ 17 Table 4: Probability of Detection (PoD) Distribution [20]....................................................... 18 Table 5: Probability of Failure for 60 Years of Operation..................................................... 19

List of Figures

Figure 1: Stress Paths for Feedwater Nozzle...................................................................... 20 Figure 2: Through-wall Stress Distributions, Unit Pressure................................................. 21 Figure 3: Through-wall Stress Distributions, Nozzle Unit Momen t Load.............................. 22 Figure 4: Through-wall Stress Distributions, Three Most Severe Thermal Transients......... 23 Figure 5: Weld Residual Stress Distribution for Stress Paths 3 and 4 at Nozzle-to-Shell Weld

............................................................................................................................ 24

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1.0 INTRODUCTION

ASME Code Case N-702 [1] allows for the reduction of in-service inspection population from 100% to 25% of nozzle blend radii and nozzle-to-shell welds every 10 ye ars for boiling water reactor (BWR) reactor pressure vessel (RPV) nozzles. Feedwater nozzles were excluded from the scope of ASME Code Case N-702 due to the cracking issues that occurred in BWR feedwater nozzle inner radii during the 1970s.

Feedwater nozzle cracking issues in BWRs were addressed by vari ous items in NUREG-0619 [2].

Since implementation of these items the BWR fleet has operated many years and demonstrated that the cracking issues have been managed effectively. Therefore, there is now a potential to seek relief for the feedwater nozzle exam coverage. BWRVIP-108-A [3] is the techni cal basis for ASME Code Case N-702 for the reduction of inspection requirements for the boiling water reactor nozzle-to-vessel shell welds and nozzle inner radius. That work was supplemented by p robabilistic fracture mechanics (PFM) evaluation for the BWR nozzle-to-vessel shell welds and nozzle blend radii (BWRVIP-241-A) [4].

Hence, the evaluation for the Perry Nuclear Power Plant (Perry) feedwater nozzles will follow the methodology used in BWRVIP-108-A and BWRVIP-241-A for establis hing the technical basis.

Additional guidance for NRC PFM submittals is provided in EPRI Letter 2019-016 [5] and proposed NRC Regulatory Guide 1.245 [6].

A relief request [7.a] for the Columbia feedwater nozzles using the BWRVIP-108-A and BWRVIP-241-A methodology was recently approved by the NRC [7.b]. The Perry feedwater nozzle evaluation follows the methodology used in the Columbia PFM evaluation and guidance from the safety evaluation report

[7.b] for the Columbia relief request. Appendix A performs a supplemental deterministic fracture mechanics evaluation, following the methodology in BWRIP-108-A.

A plant specific PFM evaluation is performed to support a relie f request for reduction of inspection population from 100% to 25% every 10 years for in-service inspe ctions (ISI) of the nozzle blend radii and nozzle-to-shell welds for the Perry feedwater nozzles. The evaluation will include 25% inspection of the feedwater nozzle population for every ten years from the fourth 10-year ISI interval to 60 years of plant operation for the first three 10-year ISI intervals.

2.0 METHODOLOGY Monte Carlo simulations are performed with the SI-proprietary s oftware VIPERNOZ [8], which was developed for RPV nozzle weld inspections with BWRVIP-108-A [3] and is the successor to the EPRI software VIPER for RPV shell weld inspections with BWRVIP-05 [9 ]. BWRVIP-108-A and BWRVIP-241-A use a Monte Carlo method to determine the probability of failure. The Perry feedwater nozzle stresses are used with probabilistic distributions from BWRVIP-108-A and BWRVIP-241-A to evaluate the plant specific probabilities.

The acceptance criterion from NUREG-1806 [21] limits the differ ence in probability of failure per year due to the low temperature over pressure (LTOP) event to be no more than 5x10 -6 per year when changing from full (100%) in-service inspection (ISI) to 25% in spection of the feedwater nozzle population every ten years. Thus, the acceptance criterion is also met if the probability of failure for 25% inspection every ten years is less than 5x10 -6 per year.

This analysis evaluates 25% inspection of the feedwater nozzle population every ten years from the fourth 10-year ISI interval to 60 years of plant operation with 100% inspection of the feedwater nozzle population with nozzle-specific coverage for the first three 10 -year ISI intervals. If the resulting File No.: 2001178.302 Page 4 of 24 Revision: 0 F0306-01R4

probability of failure (PoF) per year due to a postulated LTOP event (including 1x10 -3 probability of LTOP event occurrence per year [3, pg. 5-13]) is less than the allowable PoF of 5x10 -6 per year from NUREG-1806 [21], then there is a technical basis for inspection reduction based on BWRVIP-108-A and BWRVIP-241-A. The probability of failure for normal operation is also evaluated relative to the acceptance criterion.

2.1 Assumptions The following PFM assumptions used in the evaluation are based on previous BWRVIP development projects:

1. Flaws are assumed to be aligned parallel with the weld directio n as justified in BWRVIP-05P [9].

2. One stress corrosion crack initiation and 0.1 fabrication flaws are assumed per nozzle blend radius as justified in BWRVIP-108-A SER [3.b].

3. One stress corrosion crack initiation and 1.0 fabrication flaw are assumed per nozzle/shell weld as justified in BWRVIP-108-A [3].

4. The NRC Pressure Vessel Research Users Facility (PVRUF) flaw size distribution is assumed to apply as justified in EPRI Report W-EPRI-180-302 [20].

5. The weld residual stress distribution at the nozzle/shell weld is assumed to be a cosine distribution through the wall thickness with 8 ksi mean amplitude and 5 ksi standard deviation from BWRVIP-108-A [3].

6. Upper shelf fracture toughness is set to 200 ksi in with a standard deviation of 0 ksi in for un-irradiated material consistent with BWRVIP-108-A [3].

7. Standard deviation of the mean K IC is set to 15 percent of the mean value of the K IC as justified in BWRVIP-108-A SER [3.b].

The following are additional assumptions associated with this c alculation:

8. It is assumed that there are 40 internal cycles for each OBE event [13.b].

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3.0 DESIGN INPUTS

The stress analysis performed in a separate calculation [11] as design input to the PFM evaluation is described in Section 3.1. Section 3.2 describes the design inputs modeled deterministically as constants, and Section 3.3 describes the probabilistic inputs c onsidered to be random variables.

3.1 Stress Analysis Coefficients of linearized stresses for unit pressure (1000 psi internal pressure), deadweight, thermal moment, seismic, and the bounding thermal transients in Table 1 were extracted from the stress analysis calculation [11] using a finite element model of the Perry feedwater nozzle. Figure 1 shows the four extracted stress paths for evaluation [11, Figure 11 and Section 6.4]:

Stress Path 1 at the nozzle blend radius (Y-direction for hoop stress)

Stress Path 2 at the nozzle blend radius (X-direction for hoop stress)

Stress Path 3 at the nozzle-to-shell weld (Y-direction for hoop stress)

Stress Path 4 at the nozzle-to-shell weld (X-direction for hoop stress)

Figure 2 and Figure 3 show the through-wall stress distribution s for all stress paths for the unit pressure and thermal moment, respectively. Figure 4 shows the through-wall stress distributions for all stress paths for the three most severe thermal transients. Through-wall stress distribution data and plots for all thermal transients in the stress analysis are provided in t he supporting files. The stress analysis results are used as deterministic inputs described in the next sections.

3.2 Deterministic Parameters Details of the deterministic parameters are described in the fo llowing sections.

3.2.1 RPV Dimensions Table 3 summarizes the relevant Perry RPV dimensions [15] for t he PFM evaluation.

3.2.1 Normal Operating Conditions and LTOP event conditions.

Table 3 summarizes the normal operating conditions from the feedwater thermal cycle diagrams [17]

and the LTOP event pressure and temperature, which are specifie d in the BWRVIP-05P Safety Evaluation Report [9.b].

The stresses due to normal operating pressure and LTOP pressure for the PFM evaluation are evaluated using unit pressure stresses from the stress analysis calculation, as described in Section 3.1.

The unit pressure stresses are also scaled to the minimum and maximum pressures during each bounding transient [11, Table 3] for fatigue crack growth.

3.2.2 Transients and Projected 60-Year Cycles The thermal transients for the Perry feedwater nozzle are obtai ned from the thermal cycle diagrams for the RPV and feedwater nozzle [17] with consideration for extend ed power uprate (EPU) [18]. The stress analysis calculation tabulated the transient definitions with temperatures and pressures [11, Table 3] and selected nine bounding transients. The bounding t ransient criteria are (a) transients that have a large temperature difference and (b) transients that have a drastic rate of change in temperature

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[11, Section 4.2]. Table 1 summarizes the thermal transients with corresponding 60-year projected cycles [13.a, Tables 8 and 9], the bounding transients with gro uped cycles for evaluation, and the following notes:

For the following transients, the 60-year projected cycles are a summation of the normal and alternate transients:

a. Transient 3 (Startup): Normal Startup 3-A and Alternate Startup 3-B.

b. Transient 4 (Turbine Roll): Normal Turbine Roll 4-A and Alternate Turbine Roll 4-B.

c. Transient 14 (Hot Standby): Normal Shutdown - Hot Standby 14-A and Alternate Shutdown - Hot Standby 14-B.

d. Transient 15 (Shutdown): Normal Shutdown - Blowdown to Condense r 15-A and Alternate Shutdown - Blowdown to Condenser 15-B.

e. For Transient 16 (Shutdown, Vessel Flooding): Normal Shutdown - Vessel Flood 16-A and Alternate Shutdown - Vessel Flood 16-B.

It is assumed that there are 40 internal cycles for each OBE ev ent [13.b]. Therefore, 2 OBE events [13.a, Table 8] x 40 internal cycles = 80 OBE cycles are evaluated for 60 years of plant operation.

The stresses due to thermal transients are from the stress analysis calculation, as described in Section 3.1.

3.2.3 Piping Loads Stresses due to piping loads [12] have been analyzed in the pre vious stress analysis calculation [11, Section 6.3], as described in Section 3.1. The thermal moment stresses are scaled to the minimum and maximum temperatures during each bounding transient [11, Table 3], using the following scaling T factor:

0 (3-1) 7 70 where,

= RPV normal operating temperature (°F) = 552 °F [17.a]

= maximum or minimum operating temperature (°F)

The PFM evaluation also conservatively included stresses due to deadweight and seismic OBE, which were excluded as negligible in the Columbia feedwater nozzle PFM evaluation [7.b, Section 3.2.4]. The stresses from the seismic piping load are conservatively evalua ted at the minimum and maximum temperatures for the bounding transient 3B (Startup) and the OBE cycles in Table 1.

3.2.4 In-Service Inspection For the first 30 years of plant operation, 100% of the feedwate r nozzle population has been inspected, and the minimum feedwater nozzle inspection coverages were 100% and 82.7% for the nozzle blend radius and nozzle-to-shell weld, respectively [26].

From the fourth 10-year ISI interval onward, 25% inspection is used for 40 to 60 years of plant operation.

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The probability of detection (PoD) distribution function associated with inspection method is shown in Table 4 [20] and is an independent input from the ISI inspectio n interval, population, and coverage above. The POD curve is consistent with the POD used in BWRVIP -108-A [3.a, Figure 2-1], which included data for both the nozzle blend radius and nozzle-to-sh ell weld [3.b, Section 4.2].

3.2.5 Weld Residual Stresses Consistent with BWRVIP 108-A [3], the weld residual stresses (W RS) are assumed present in the nozzle-to-shell welds for Stress Paths 3 and 4. The WRS distribution at the nozzle/shell weld is assumed to be a cosine distribution through the wall thickness with 8 ksi mean amplitude and 5 ksi standard deviation (see Figure 5). No WRS is present in the nozzle blend radius region for Stress Paths 1 and 2.

3.3 Random Variables Random variables used in the PFM evaluation are summarized in Table 2 with details in the following sections.

3.3.1 Materials, Material Chemistry, and Fracture Toughness Table 2 includes the weld chemistries (%Cu and %Ni) and initial RT NDT along with the standard deviation and distributions for the nozzle forging (SA-508, Gr. 2, Cl. 2) and the nozzle-to-shell welds (SA-508, Gr.

09L, CL. 2) [19]. The mean values for initial RT NDT are plant-specific [19], and the remaining values and distributions are used from BWRVIP-108-A and the associated saf ety evaluation report [3].

Consistent with BWRVIP-108-A, the upper shelf fracture toughnes s of 200 ksiin for unirradiated materials is used. The feedwater nozzles are not in the RPV be ltline region and operate at high temperatures for the upper shelf fracture toughness to be appli cable.

3.3.2 Fluence The Perry feedwater nozzle (N4) experiences fluence of less tha n 1.00 x 1017 n/cm2 for 54 EFPY for 60 years of plant operation [22]. As such, a bounded fluence of 1.00 x 1017 n/cm2 is used for the evaluation of the feedwater nozzles.

3.3.3 Stress Corrosion Cracking Initiation The stress corrosion cracking (SCC) initiation model in the VIPERNOZ program is a power law relationship. The cast stainless steel SCC data in a BWR enviro nment is used as specified in BWRVIP-05 [9, Section 8.2.2.2], and used in BWRVIP-108-A [3] and BWRVI P-241-A [4]. Although the Perry feedwater nozzles are not cast stainless steel, sensitivity studies in BWRVIP-108-A Safety Evaluation Report concluded that use of cast stainless steel data is conse rvative [3.b, Section 4.5, RAI 2-10]. This model has the form;

84.21 (0. 4-1)

where: T = time, hours

= applied stress, ksi

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The residual plot shows that a lognormal distribution produces the best fit for the data. The lognormal residual plot with the linear fit of the data is shown below:

0.9248 0.0003 (4-2)

where: = (x - ) /

= data mean

= data standard deviation x = ln (Tactual/Tpredicted)

3.3.3.1 Stress Corrosion Cracking Growth The SCC growth model in VIPERNOZ program is a power law relatio nship from NUREG/CR-6923 [23].

The relationship used is:

6.8 (210 4-3)

where: da/dt = stress corrosion crack growth rate, in/hr K = sustained crack tip stress intensity factor, ksi in

The residual plot shows that a Weibull distribution produces th e best fit for the data. The Weibull residual plot with the linear fit of the data is shown below:

0.9085 0.3389 (4-4)

where: Y = ln (ln (1/ (1-F) ))

F = cumulative distribution from 0 to 1 x = ln ((da/dt) actual / (da/dt) predicted)

3.3.4 Fatigue Crack Growth Consistent with BWRVIP-108-A and BWRVIP-241-A, the fatigue cra ck growth data for SA-533 Grade B Class 1 and SA-508 Class 2 (carbon-molybdenum steels) in a reac tor water environment are reported in Reference [24] for weld metal testing at an R-ratio (algebraic ratio of Kmin/Kmax, R) of 0.2 and 0.7. To produce a fatigue crack growth law and distribution for the VIPERNOZ software, the data for R= 0.7 was fitted into the form of Paris Law. The R= 0.7 fatigue crack growth law was chosen for conservatism.

The curve fit result of the mean fatigue crack growth law is pr esented with the Paris law shown as follows:

3.817 10. (4-5)

where a = crack depth, in n=cycles File No.: 2001178.302 Page 9 of 24 Revision: 0 F0306-01R4

K = Kmax - Kmin, ksi-in0.5

A comparison to the ASME Section XI fatigue crack growth law in a reactor water environment is documented in Reference [20] and it shows a reasonable comparis on where the Section XI law is more conservative on growth rate at high K.

Using the rank ordered residual plot, it is shown that a Weibull distribution is representative for the data.

The Weibull residual plot with the linear curve fit of the data is shown below:

4.1500 0.3712 (4-6)

where y = ln( ln( 1 / (1 - F) )

x=ln((da/dn) actual / (da/dn)mean)

F=cumulativepr obability distribution

4.0 PROBABILISTIC FRACTURE MECHANICS EVALUATION The probabilistic fracture mechanics evaluation is performed fo r 25% inspection of the feedwater nozzle population every ten years from the fourth 10-year ISI interval onward for up to 60 years of plant operation, after 100% inspection was performed for the first 30 years of plant operation.

For the nozzle blend radius region, a nozzle blend radius crack model [10] is used in the probabilistic fracture mechanics evaluation. For this location and crack mod el, the applicable stress is the stress perpendicular to a path defined 90 degrees from the tangent drawn at the blend radius.

For the nozzle-to-shell weld, either a circumferential or an axial crack, depending on weld orientation, can initiate due to either component fabrication (i.e. consider ing only welding process) or stress corrosion cracking. The probability of failure for a circumfer ential crack is less than an axial crack due to the difference in the stress (hoop versus axial) and the inf luence on the crack model. However, this probabilistic fracture mechanics evaluation for the nozzle and vessel shell weld considers both circumferential and axial cracks (depending on weld orientation).

An axial elliptical crack model with a crack aspect ratio of a/ l = 0.5 is used in the evaluation for the nozzle-to-shell weld. The inspection probability of detection (PoD) curve in Table 4 from Reference

[20, Figure 8-10] is utilized with a ten-year inspection interval. The calculation of stress intensity factor is at the deepest point of the crack.

Crack growth calculations include stress corrosion cracking and fatigue crack growth, which includes loads due to deadweight, internal pressure, nozzle moment, seis mic OBE, and thermal transients.

The approach used for this evaluation is consistent with the me thodology presented in BWRVIP-05

[9.a]. A Monte Carlo simulation is performed using a variant of the VIPER program [14] and follows the flow diagram in BWRVIP-05 [9.a, Figure 8-11]. The Monte Carlo method can be used to solve probabilistic problems using deter ministic computation. A mean value, standard deviation, and distribution curve as defined in the random variables summary ( Table 2) define a set of possible inputs and their probabilities of occurring. Using this domain of pos sible inputs, a set of inputs are generated

File No.: 2001178.302 Page 10 of 24 Revision: 0 F0306-01R4

for use in determining whether the nozzle will fail using conve ntional deterministic fracture mechanics methodology. The number of simulations in which the nozzle is determined to fail divided by the number of simulations run (i.e., realizations) gives the probability of failure. The failure criteria are defined as:

Rupture: Kapplied > KIc, where Kapplied is the applied stress intensity factor, and K Ic is the fracture toughness

Leakage: Crack depth > 80% of vessel wall thickness

The VIPER program was developed as part of the BWRVIP-05 [9.a] effort for Boiling Water Reactor (BWR) reactor pressure vessel (RPV) shell weld inspection recommendations. The software was modified into a separate version in BWRVIP-108-A [3], identifie d as VIPER-NOZ, for RPV nozzle weld inspections, which was used in this project. The detailed description of the methodology incorporated in the VIPER and VIPER-NOZ program is documented in BWRVIP-05 [9] and BWRVIP-108-A [3],

respectively. The modified software is identified as VIPER-NOZ to distinguish from the original VIPER software and is verified on a project specific basis [8] to ens ure the modifications made to the VIPER software are fully quality assured.

5.0 RESULTS AND CONCLUSIONS

Table 5 presents the results of the probabilistic fracture mechanics evaluation for normal operation and LTOP events.

For normal operation, the results for Stress Path 1 at the nozz le blend radius bounds the other stress paths. For Stress Path 1, 5 failures occurred during normal o peration in 1 million simulations for 60 years of plant operation. The probably of failure (PoF) for n ormal operation for Stress Path 1 is calculated to be 5 failures / 1 million simulations / 60 years = 8.33 x 10-8 per year. The calculated PoF for normal operation for Stress Path 1 is less than the allowab le PoF of 5 x 10 -6 per year from NUREG-KLW 12/22/21 1806 [21].

For LTOP events, the results for Stress Path 1 at the nozzle blend radius bounds the other stress paths.

For Stress Path 1, 52 LTOP failures occurred in 1 million simulations for 60 years of plant operation.

The conditional PoF for LTOP events for Stress Path 1 is calcul ated to be 52 failures / 1 million simulations / 60 years = 8.67 x 10 -7 per year. Accounting for an LTOP event occurrence of 1 x 10 -3 per year [3, pg 5-13], the calculated PoF for LTOP events for Stres s Path 1 is 8.67 x 10 -10 per year, which is less than the allowable PoF of 5 x 10 -6 per year from NUREG-1806 [21]. KLW 12/22/21 The PFM results are also confirmed by the deterministic fractur e mechanics evaluation in Appendix A, using the methodology in BWRVIP-108-A [3.a, Section 6].

Thus, using PFM methodology in BWRVIP-108-A and BWRVIP-241-A, w hich is the technical basis for ASME Code Case N-702, the Perry feedwater nozzles are qualified for inspection relief from 100% to 25% of the feedwater nozzle population every 10 years from the fourth 10-year ISI interval for up to sixty years of plant operation.

File No.: 2001178.302 Page 11 of 24 Revision: 0 F0306-01R4

6.0 REFERENCES

1. ASME Code Case N-702, Alternative Requirements for Boiling Wat er Reactor (BWR) Nozzle Inner Radius and Nozzle-to-Shell Welds,Section XI, Division 1, Febr uary 20, 2004.

2. U.S. NRC Report, Revision 1: BWR Feedwater Nozzle and Control R od Drive Return Line Nozzle Cracking: Resolution of Generic Technical Activity A-10, NUREG- 0619, Nuclear Regulatory Commission (NRC), November 1980.

3. EPRI BWRVIP-108-A

a. BWRVIP-108-A: BWR Vessel and Internals Project, Technical Basis for the Reduction of Inspection Requirements for the Boiling Water Reactor Nozzle to Vessel Shell Welds and Nozzle Blend Radii, EPRI, Palo Alto, CA 2007, CA. EPRI PROPRIETARY.

b. U.S. NRC Report, Safety Evaluation by the Office of Nuclear Re actor Regulation, BWRVIP-108, BWR Vessel and Internals Project Technical Basis for the Reduction of Inspection Requirements for the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Inner Radii, December 19, 2007, ADAMS Accession No. ML073600374.

4. BWRVIP-241-A: BWR Vessel Internal Project, Probabilistic Fracture Mechanics Evaluation for the Boiling Water Reactor Nozzle-to-Vessel Shell Welds and Nozzle Blend Radii, EPRI, Palo Alto, CA.

1021005. EPRI PROPRIETARY.

5. BWRVIP Letter 2019-016, from N. Palm (EPRI) to D. Rudland and P. Raynaud (U.S. NRC), White Paper on Suggested Content for PFM Submittals to the NRC, Febru ary 27, 2019, ADAMS Accession No. ML19241A545.

6. U.S. NRC DRAFT Regulatory Guide 1.245

a. U.S. NRC DRAFT Regulatory Guide DG-1382, Proposed New Regulator y Guide 1.245, Preparing Probabilistic Fracture Mechanics Submittals, September 2021, ADAMS Accession No. ML21034A328.

b. U.S. NRC DRAFT Report for Comment NUREG/CR-7278, Technical Bas is for the use of Probabilistic Fracture Mechanics f or Regulatory Applications, September 2021, ADAMS Accession No. ML21257A237.

7. Columbia Feedwater Nozzle Relief Request 4ISI-09

a. Energy Northwest Letter G02-20-048, Columbia Generating Statio n, Docket No. 50-397 Fourth Ten-Year Interval Inservice Inspection (ISI) Program Rel ief Request 4ISI-09, April 22, 2020, ADAMS Accession Number ML20114E235.

b. U.S. NRC Report, Safety Evaluation by the Office of Nuclear Re actor Regulation, Relief Request 4ISI-09, Alternate Examination of Reactor Vessel Feedwater Nozzles and Nozzle-to-Shell Welds, Columbia Generating Station, Energy Northwest, Docket No. 50-397, April 14, 2021, ADAMS Accession Number ML21096A048.

8. SI Calculation 2001178.303, Revision 0, Verification of Software VIPERNOZ Version 1.3.

File No.: 2001178.302 Page 12 of 24 Revision: 0 F0306-01R4

9. EPRI BWRVIP-05

a. BWRVIP Report, BWR Reactor Pressure Vessel Shell Weld Inspection Recommendations (BWRVIP-05P), Electric Power Research Institute TR-105697, Sep tember 1995. EPRI PROPRIETARY.

b. U.S. NRC Report, Final Safety Evaluation of the BWR Vessel Int ernals Project BWRVIP-05 Report, TAC No. M93925, Division of Engineering Office of Nucl ear Reactor Regulation, Nuclear Regulatory Commission, July 28, 1998.

10. Delvin, S. A., Riccardella, P. C., Fracture mechanics analysis of JAERI model pressure vessel test, ASME PVP Conference, Paper 78-PVP-91, 1978.

11. SI Calculation Package No. 2001178.301, Revision 0, Feedwater Nozzle Loads, Finite Element Model, and Stress Analysis.

12. CBI Nuclear Company, Charge No. 238-D4,

Subject:

238 Dia. BWR 6 Feedwater Nozzle, Section D4, Rev. 14, SI File No. 2001178.209.

13. Projected 60-Year Cycles

a. Structural Integrity Associates Calculation No. 2001140.301, Re vision 1, Fatigue Update for Perry Nuclear Power Plant from 10/1/2016 to 10/31/2021.

b. Email from J. Zbiegien (Energy Harbor) to K. Wong (SI),

Subject:

RE: [EXTERNAL] RE:

Projected Cycles, December 3, 2021 7:44AM, SI File No. 2001178.207.

14. VIPER, Vessel Inspection Program Evaluation for Reliability, Version 1.1 (April 2007), Structural Integrity Associates.

15. Energy Harbor Drawing No. 08_0021-00000, Revision 4 (CBI Nuclear Company Drawing, Contract No. 73-C108/14, Dwg. 59, Revision 3, N4 Nozzle Forging (Feedwa ter)), SI File No. 2001178.205.

16. CBI Nuclear Company Section No. 238-D11.3, Revision 5, Perry 1 - 238 BWR 6 Vessel Contract 73-C108, Stress Report (Code), Water Lev. Instr. Nozzle Design Report - Section D 11.3, SI File No. 2001178.212.

17. Thermal Cycle Diagrams

a. Energy Harbor Drawing No. 08-0037-00001 (General Electric Drawing No. 762E458, Revision 7, Sheet 1, Reactor Cycles), SI File No. 20001178.20 3.

b. Energy Harbor Drawing No. 08-0037-00002 (General Electric Drawing No. 762E458, Revision 7, Sheet 2, Reactor Cycles), SI File No. 2001178.203.

c. Energy Harbor Drawing Update Notice No. 08-0596-001-001, Revision 0, Reactor Cycles, SI File No. 2001178.203.

d. Energy Harbor Drawing Update Notice No. 08-0596-001-003, Revision 0, Reactor Cycles, SI File No. 2001178.203.

e. Energy Harbor Drawing No. 306-0081-00000, Revision E, Feedwate r Temperature/Pressure Cycles, SI File No. 2001178.204.

f. Energy Harbor Drawing No. 306-0082-00000, Revision D, Feedwate r Temperature/Pressure Cycles, SI File No. 2001178.204.

g. Energy Harbor Engineering Change Package N o. 08-0596-000, Revision 0, SI File No.

2001178.204.

h. Energy Harbor Engineering Change Package No. 08-0596-001, Revision 0, SI File No.

2001178.204.

File No.: 2001178.302 Page 13 of 24 Revision: 0 F0306-01R4

18. General Electric Document No. 26A5308, Revision 3, Reactor Ves sel - Power Uprate, SI File No.

2001178.206.

19. Email from J. Zbiegien (Energy Harbor) to K. Wong (SI), Subj: PY FW Nozzle RR Design Inputs with attachment referencing Energy Harbor Document EA-0246, Revision 0 A-02, 11/17/21 9:21AM, SI File No. 2001178.213.

20. SI Report No. W-EPRI-180-302, Revision 0, Evaluation of effect of inspection on the probability of failure for BWR Nozzle-to-Shell-Welds and Nozzle Blend Radii Re gion.

21. Technical Basis for Revision of Pressurized Thermal Shock (PTS) Screening Limit in the PTS Rule (10 CFR 50.61), NUREG-1806, Vol. 1, August 2007.

22. Email from J. Zbiegien (Energy Harbor) to K. Wong (SI), Subj: RE: [External] RE: PY FW Nozzle RR Design Inputs with attachment referencing Energy Harbor Doc ument DI-EA-0235, 11/18/21 12:40PM, SI File No. 2001178.214.

23. NUREG/CR-6923, Appendix B.8, Expert Panel Report on Proactive Materials Degradation Assessment, Published February 2007.

24. Bamford, W. H., Application of corrosion fatigue crack growth rate data to integrity analyses of nuclear reactor vessels, Journal of Engineering Materials and Technology, Vol. 101, 1979, SI File No. 1300341.213.

25. EPRI Letter 2012-138, BWRVIP Support of ASME Code Case N-702 I nservice Inspection Relief, August 31, 2012, SI File No. 1300341.213.

26. Email from J. Zbiegien (Energy Harbor) to K. Wong (SI), Subj: Exam Percent Coverage, referencing Perry ISI Final Report P1R014, 12/15/21 11:36AM, SI File No. 2001178.216.

27. pc-CRACK, Version 5.0 CS, Structural Integrity Associates, Inc., December 30, 2020.

28. API 579-1/ASME FFS-1, Fitness-For-Service, June 2016.

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Table 1: Bounding Thermal Transients Cycles

Perry Feedwater Transients [11, Table 2] Bounding Analysis 60-Year Projected Bounding New Grouped Event Description Cycles Group Group 60-Year to 11/7/2046 Event ID Cycles

[13.a, Tables 8 & 9]

1 Boltup 37 2 Design hydrotest 47 Startup 3b 260 3 Startup 166 + 10 (Note 1) 4 Turbine roll 167 +1 (Note 1) Turbine Roll 4b 168 5 Daily reduction 685 Weekly 6 Weekly reduction 436 Reduction 6 1806 7 Rod pattern change 685 8 Turbine trip 22 Turbine Trip 8 181 9 Partial feedwater heater bypass 159 10 Turbine generator trip 13 Turbine 10 100 11 Other scrams 87 Generator Trip 13 Reduction to 0% power 209 14 Hot standby 127 + 4 (Note 1) 15 Shutdown 165 + 4 (Note 1) 16 Shutdown, vessel flooding 160 + 10 (Note 1) Shutdown 15b 931 17 Shutdown 183 18 Unbolt 36 19 Refueling 33 20 Composite Loss of Feedwater Pumps 20 LOFP 20 20 21 SRV blowdown 2 SRV Blowdown 21 2 22 Reactor Overpressure 1 23 Automatic Blowdown 1 Reactor 24 Improper Start of Cold Recirc Loop 1 Overpressure 22 8 25 Sudden Start of Pump in Cold Recirc Loop 4 27 Pipe Rupture 1

- OBE 2 x 40 (Note 2) OBE OBE 80

Notes:

1. For the noted transients, the 60-year projected cycles are a su mmation of the normal and alternate transients:

a. Transient 3 (Startup): Normal Startup 3-A and Alternate Startup 3-B.

b. Transient 4 (Turbine Roll): Normal Turbine Roll 4-A and Alterna te Turbine Roll 4-B.

c. Transient 14 (Hot Standby): Normal Shutdown - Hot Standby 14-A and Alternate Shutdown - Hot Standby 14-B.

d. Transient 15 (Shutdown): Normal Shutdown - Blowdown to Condense r 15-A and Alternate Shutdown -

Blowdown to Condenser 15-B.

e. For Transient 16 (Shutdown, Vessel Flooding): Normal Shutdown - Vessel Flood 16-A and Alternate Shutdown - Vessel Flood 16-B.

2. It is assumed that there are 40 internal cycles for each OBE event [13.b]. Therefore, 2 OBE events x 40 internal cycles = 80 OBE cycles are evaluated for 60 years of p lant operation.

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Table 2: Random Variables Parameter Summary

Source: BWRVIP-108-A [3] unless otherwise noted.

Random Parameter Mean Standard Deviation Distribution Reference Flaw density, nozzle/shell weld (fabrication) 1 per weld Mean Poisson [3, 3, 3]

Flaw density, nozzle and nozzle/shell weld (SCC 1 per weld Mean Poisson [3, 3, 3]

initiation)

Flaw density, nozzle blend radius (fabrication) 0.1 per nozzle Mean Poisson [3, 3, 3]

Flaw size (fabrication) n/a n/a PVRUF [3]

Weld residual stress, 8 through-wall (ksi) inside surface 5 Normal [3, 3, 3]

cosine distribution N4 nozzle % Cu 0.26 0.045 Normal [3, 3, 3]

to shell % Ni 1.2 0.0165 Normal [3, 3, 3]

weld Initial RTNDT (F) -20 13 Normal [19, 3, 3]

% Cu 0.09189 0.04407 Normal [3, 3, 3]

N4 nozzle % Ni 0.78 0.068 Normal [3, 3, 3]

forging Initial RTNDT (F) -20 26.48 Normal [19, 3, 3]

Fast neutron fluence (n/cm2) 1.00 x 1017 0.2 (20%) n/a [22, 3]

KIC upper shelf (ksiin) 200 0 Normal [3, 25, 3]

Residual SCC initiation time (hr) T = 84.2x1018()-10.5 y=0.9248x-Lognormal [3, 3, 3]

0.0003 K dependent Residual da/dt = 6.82x10 -12(K)4 y=0.9085x-Weibull [23, 3, 3]

SCCG (in/hr) K >50 ksiin 0.3389 K independent da/dt = 2.8x10-6, n/a n/a [23]

K <50 ksiin SCC threshold (ksiin) 10 2 Normal [3, 3, 3]

Fatigue crack growth (FCG) da/dn=3.82 Residual (in/cycle) x10-9(dK)2.927 y=4.155x-Weibull [3, 3, 3]

0.3712 FCG threshold (ksiin) 0 0 Normal [3, 3, 3]

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Table 3: Deterministic Parameter Summary

Parameter Value Reference

Dimensions Vessel Thickness (excludes cladding) 6.375 inches Nozzle Drawing Vessel Radius (to vessel surface) 120.1875 inches [15]

Vessel Clad Thickness 0.1875 inches Nozzle Drawing [16]

Operating Conditions - Post-EPU [18]

Normal Operating Temperature 552 °F RPV Thermal Cycle Diagram [17.a]

Feedwater Thermal Normal Operating Pressure 1100 psig (Note 1) Cycle Diagram

[17.e]

LTOP Event Temperature 100 °F BWRVIP-05 SER LTOP Event Pressure 1200 psig [9.b, Section 2.6.2.2]

Note:

1. The higher operating pressure of the feedwater nozzle is conservatively used compared to the lower RPV operating pressure of 1050 psig [17.a].

File No.: 2001178.302 Page 17 of 24 Revision: 0 F0306-01R4

Table 4: Probability of Detection (PoD) Distribution [20]

Flaw Size, in. Cumulative PoD 0.00 0.20 0.05 0.32 0.10 0.46 0.15 0.61 0.20 0.75 0.25 0.85 0.30 0.91 0.35 0.95 0.40 0.96 0.45 0.97 0.50 0.98 0.55 0.99 0.60 1.00

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Table 5: Probability of Failure for 60 Years of Operation

Probability of Failure (PoF)

Location Stress 60 Years of Plant Operation Allowable PoF per year Path PoF per year due to PoF per year due to NUREG-1806 [21]

Normal Operation LTOP Event (Note 1)

Nozzle Blend Radius 1 8.33 x 10-8 8.67 x 10-10 2 < 1.67 x 10-8 (Note 2) < 1.67 x 10-11 (Note 3) 5.0 x 10-6 Nozzle-to-Shell Weld 3 < 1.67 x 10-8 (Note 2) < 1.67 x 10-11 (Note 3) 4 < 1.67 x 10-8 (Note 2) < 1.67 x 10-11 (Note 3)

Notes:

(1) The LTOP PoF accounts for a 1x10-3 probability of LTOP event occurrence per year

[3, pg 5-13].

(2) No failures occurred during normal operation. As such, the PoF for normal operation is calculated as less than 1 failure / 1 million simulations / 60 years = 1.67 x 10-8 per year.

(3) No LTOP failures occurred. As such, the LTOP PoF is calculated as less than 1 failure / 1 million simulations / 60 years

• 1 x 10-3 probability of LTOP event occurrence = 1.67 x 10 -11 per year.

File No.: 2001178.302 Page 19 of 24 Revision: 0 F0306-01R4

P1 P2

Z

P3 P4

Y X

Figure 1: Stress Paths for Feedwater Nozzle

Source: Stress Analysis Calculation [11, Figure 11]

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Figure 2: Through-wall Stress Distributions, Unit Pressure

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Figure 3: Through-wall Stress Distributions, Nozzle Unit Moment Load

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Figure 4: Through-wall Stress Distributions, Three Most Severe Thermal Transients

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Figure 5: Weld Residual Stress Distribution for Stress Paths 3 and 4 at Nozzle-to-Shell Weld

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DETERMINISTIC FRACTURE MECHANICS EVALUATION

File No.: 2001178.302 Page A-1 of A-4 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

Appendix A performs a deterministic fracture mechanics evaluati on of the Perry feedwater nozzles, similar to the supplemental analysis performed for the Columbia feedwater nozzle [7.b, Section 3.2.6].

Each VIPER-NOZ simulation in the PFM evaluation consists of a series of deterministic fracture mechanics (DFM) realizations using the sampled probabilistic inputs in Table 2. A confirmatory DFM evaluation for the Perry feedwater nozzles was performed, using the methodology in BWRVIP-108-A

[3.a, Section 6] and the nuclear QA fracture mechanics software, pc-CRACK 5.0 [27]. The four stress paths in the PFM evaluation were evaluated for crack growth in the DFM evaluation.

Initial Flaw Size Consistent with the BWRVIP-108-A DFM evaluation, an initial flaw depth of 0.15 inch and an aspect ratio of 6:1 for the crack length to crack depth is used for both axial and circumferential cracks.

Stresses The stresses for deadweight, unit pressure, unit moment, bounding thermal transients, and seismic OBE in Section 3.1 of the PFM evaluation were used for the DFM evaluation. Cyclical stresses are applied according to the grouped 60-year cycles in Table 1.

The weld residual stresses in Section 3.2.5 of the PFM evaluation were used for Stress Paths 3 and 4 at the nozzle-to-shell welds.

Fracture Mechanics and Crack Growth Models The DFM evaluation used following mean values of PFM inputs in Table 2:

Stress Corrosion Cracking: Crack growth equation and threshold value of 10 ksiin Fatigue Crack Growth: Crack growth equation and threshold value of 0 ksiin Material Fracture Toughness, K IC, of 200 ksiin

In the PFM analysis, the same mean parameters are used in the e valuation but with an applied standard deviations and distribution to vary for each set of re alization. The fracture mechanics software pc-CRACK 5.0 includes crack models from API-579-1 [28]. For S tress Paths 1 and 2 in the nozzle blend radius, the crack model is semi-circular nozzle corner cr ack with a given initial depth. The cracks are modeled to propagate from the inside blend radius at the smallest distance between the inner and outer blend radius.

Figure A-1: Nozzle Corner Crack Model for Stress Paths 1 and 2 File No.: 2001178.302 Page A-2 of A-4 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

For the nozzle-to-shell weld, the crack-models are a semi-ellip tical axial crack for Stress Path 3 and a semi-elliptical circumferential crack for Stress Path 4. They a re developed from the inside surface of a cylinder with a given initial depth and length. The cracks are propagated through the cylinder wall thickness from the inside radius to outside radius in depth and length with a fixed aspect ratio.

Figure A-2: Semi-Elliptical Axial Crack in a Cylinder Model for Stress Path 3

Figure A-3: Semi-Elliptical Circumferential Crack in a Cylinde r Model for Stress Path 4

File No.: 2001178.302 Page A-3 of A-4 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

DFM Results The results of the DFM evaluation after 60 years of plant opera tion are in terms of the final flaw depth, ar, the final depth-to-thickness ratio, a r/t, and the applied stress intensity factor, K i, of the final flaw size as compared to the fracture toughness of the material, K IC. For all stress paths after 60 years of crack growth, the final flaw depth is less than 10% of the thickness, and the final applied stress intensity factor is less than 50% of the material fracture toughness. Results of the DFM evaluation using average values show that fracture for all four stress paths is not expe cted since the crack growth is not significant if the average growth rate is used, consistent with DFM evaluations in BWRVIP-108-A [3.a, Section 6] and the Columbia supplemental analysis [7.b, Section 3.2.6].

DFM Crack Growth Material Location Stress 60 Years of Operation Fracture Path Final Flaw Depth Final Ki Toughness af (inch) af/t (ksiin) KIC (ksiin)

Nozzle Blend Radius 1 0.60 0.065 91.8 2 0.16 0.018 26.0 200 Nozzle-to-Shell Weld 3 0.35 0.055 46.1 4 0.32 0.050 43.7

File No.: 2001178.302 Page A-4 of A-4 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

SUPPORTING FILES

File No.: 2001178.302 Page B-1 of B-2 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.

Stress Analysis Supporting Files

Filename Description

Stress coefficients and through-wall stress plots for unit pressure, deadweight, 2001178.302 - Stresses.xlsx unit moment, seismic, and thermal transients from the stress analysis calculation

[11].

Probabilistic Fracture Mechanics Supporting Files

PFM / File Name Description Path1.INP Path1.OUT VIPERNOZ input and output files for Path 1 at nozzle blend radius.

Path2.INP Path2.OUT VIPERNOZ input and output files for Path 2 at nozzle blend radius.

Path3.INP Path3.OUT VIPERNOZ input and output files for Path 3 at nozzle-to-shell weld.

Path4.INP Path4.OUT VIPERNOZ input and output files for Path 4 at nozzle-to-shell weld.

VIPERNOZ1p3.EXE VIPERNOZ executable program

ISPCTPOD.INP VIPERNOZ probability of detection curve input file

FLWDSTRB.INP VIPERNOZ flaw size distribution curve input file

Deterministic Fracture Mechanics Supporting Files

DFM / File Name Description Path1.INP or Path 1 at nozzle blend radius.

Path1.OUT pc-CRACK input and output files f Path2.INP or Path 2 at nozzle blend radius.

Path2.OUT pc-CRACK input and output files f Path3.INP Path3.OUT pc-CRACK input and output files for Path 3 at nozzle-to-shell weld Path4.INP Path4.OUT pc-CRACK input and output files for Path 4 at nozzle-to-shell weld

File No.: 2001178.302 Page B-2 of B-2 Revision: 0 F0306-01R4 This document contains Structural Integrity; client; or supplier proprietary information. This document may not be disclosed, wholly or in part, to any third parties without the prior written consent of Structural Integrity Associates, Inc.