ML20090B397

Revised License Renewal Commitment Pressurizer Surge Line Weld Inspection Frequency

ML20090B397
Person / Time
Site:	North Anna
Issue date:	03/26/2020
From:	Mark D. Sartain Virginia Electric & Power Co (VEPCO)
To:	Document Control Desk, Office of Nuclear Reactor Regulation
References
20-049
Download: ML20090B397 (93)

Text

VIRGINIA ELECTRIC AND POWER COMPANY RICHMOND, VIRGINIA 23261 March 26, 2020 United States Nuclear Regulatory Commission Serial No.: 20-049 Attention: Document Control Desk NRA/GDM: R1 Washington, D. C. 20555 Docket Nos.: 50-338/339 License Nos.: NPF-4/7 VIRGINIA ELECTRIC AND POWER COMPANY NORTH ANNA POWER STATION UNITS 1 AND 2 REVISED LICENSE RENEWAL COMMITMENT PRESSURIZER SURGE LINE WELD INSPECTION FREQUENCY Virginia Electric and Power Company (Dominion Energy Virginia) made a specific commitment associated with the renewed operating licenses for North Anna Power Station (North Anna) Units 1 and 2 to provide for NRG review and approval a summary of the inspection details for the pressurizer surge line weld at the connection of the surge line to the Reactor Coolant System (RCS) hot-leg piping. (Reference Updated Final Safety Analysis Report (UFSAR) Section 18.3.2.4, "Environmentally Assisted Fatigue,"

and Table 18-1, Item 24). To fulfill this commitment, Dominion Energy Virginia provided the inspection plan summary for the pressurizer surge line weld to the NRG by letter dated July 1, 2014 (Serial No.14-285, ADAMS Accession No. ML14189A129). As part of the inspection plan summary, we informed the NRG that we would revise the North Anna Augmented lnservice Inspection Plan to require baseline volumetric and surface examinations of two welds associated with the pressurizer surge line during the period of extended operation at a frequency of once per inspection period (i.e., once every 40 months during each of the fifth and sixth inservice inspection (ISi) intervals.) Details of the inspection plan are outlined in UFSAR Section 18.2.1, "Augmented Inspection Activities."

The "once every 40 months" weld inspection schedule was selected because no guidance existed for determination of the surge line weld inspection frequency at the time decisions were being made for the first North Anna license renewal period. Subsequent to this effort, Florida Power and Light (FPL) proposed to use ASME Code Section XI, Appendix L, "Operating Plant Fatigue Assessment," from the 2001 ASME Code edition, with the 2002 and 2003 Addenda, for establishing the inspection frequency of the pressurizer surge line at Turkey Point Power Station. Effective August 17, 2017, the latest ASME Code edition approved by the NRG is the 2013 Edition, which includesSection XI, Appendix L. In addition, ASME Code Case N-809, "Reference Fatigue Crack Growth Curves for Austenitic Stainless Steels in Pressurized Water Reactor Environments,"

includes the latest crack growth data and has been approved by ASME. Although the NRG has not generically endorsed Code Case N-809 in Regulatory Guide 1.147, Revision 18, "lnservice Inspection Code Case Acceptability, ASME Section XI, Division 1," at this time, the staff has reviewed and approved precedent license renewal commitments

Serial No.20-049 Docket Nos. 50-338/339 Revised License Renewal Commitment Page 2 of 4 pertaining to fatigue of surge line welds for Turkey Point (ADAMS Accession Nos.:

Submittal: ML12152A156 and Approval: ML 1314A595), St. Lucie (ADAMS Accession Nos.: Submittal: ML15314A160 and Approval: ML16235A138) and Surry Power Stations (ADAMS Accession Nos.: Submittal: ML17339A170 and Approval: ML18166A329.)

Based on the developments and precedents noted above, Dominion Energy Virginia has performed a flaw evaluation for the North Anna pressurizer surge line welds based on the use of the ASME 2013 Edition of Section XI, Appendix L. The ASME Section XI, Appendix L, flaw tolerance evaluation for North Anna, which is based upon the use of bounding stress in the vicinity of the hot leg, shows the crack growth for a postulated circumferential flaw will reach the allowable flaw size after an operating period of 60 years.

Thus, a proposed inspection frequency of once per 10-year inspection interval for surface and volumetric examinations for the pressurizer surge line weld is conservative and justified. This inspection frequency is also consistent with the frequency used in ASME Section XI for examination of reactor coolant piping. The ASME Section XI, Appendix L, flaw tolerance evaluation for the North Anna Units 1 and 2 pressurizer surge line welds, which was prepared by Structural Integrity Associates, Inc., is provided in the attachment.

During development of the Appendix L evaluation of the pressurizer surge line for NAPS Units 1 and 2, it was confirmed that the weld that connects the hot leg nozzle to the reactor coolant piping will not have a concern due to environmental assisted fatigue as this weld location is protected by a thermal sleeve. However, the weld that connects the hot leg nozzle to the pressurizer surge line is not protected by a thermal sleeve and will therefore require management for environmental assisted fatigue concerns. Dominion Energy Virginia proposes to update the inspection plan to inspect one weld instead of two during each 10-year inspection interval. Specifically, the weld that connects the hot leg nozzle to the reactor coolant piping that is protected by a thermal sleeve is being removed from the augmented inspection plan outlined in UFSAR Section 18.2.1. The weld that connects the surge line piping to the hot leg nozzle will continue to be inspected for management of environmental assisted fatigue at a once per 10-year inspection interval frequency instead of a once per inspection period frequency (once every 40 months).

Therefore, based upon the industry precedents noted above, the inclusion of Appendix Lin ASME Section XI, the ASME approved Code Case N-809, and the plant specific North Anna technical evaluation included in the attachment, Dominion Energy Virginia requests NRC approval to revise the North Anna Units 1 and 2 pressurizer surge line weld inspection plan as follows:

1. Reduce the number of welds in the license renewal commitment to be inspected from two welds to one weld that attaches the hot leg nozzle to the pressurizer surge line, and

Serial No.20-049 Docket Nos. 50-338/339 Revised License Renewal Commitment Page 3 of 4

2. Reduce the inspection frequency of the license renewal commitment from once every inspection period (once every 40 months) to once per 10-year inspection interval for the fifth and sixth ISi intervals.

Dominion Energy requests NRG approval by August 31, 2020 to coincide with the planned North Anna Units 1 and 2 Subsequent License Renewal submittal. Upon receipt of NRG approval, the North Anna UFSAR will be revised to reflect the new inspection scope and frequency for the North Anna Units 1 and 2 pressurizer surge line weld based upon the date of completion of the last weld inspection not to exceed 10 years. The proposed inspection scope and frequency for the North Anna Units 1 and 2 pressurizer surge line weld is applicable to both the first license renewal period as well as subsequent license renewal.

If you have any questions or require additional information, please contact Mr. Gary D.

Miller at (804) 273-2771.

Respectfully, Mark D. Sartain Vice President - Nuclear Engineering and Fleet Support Commitment made in this letter:

Upon receipt of NRG approval, the North Anna UFSAR will be revised to reflect the new inspection scope and frequency for the North Anna Units 1 and 2 pressurizer surge line weld based upon the date of completion of the last weld inspection not to exceed 10 years.

Attachment:

Structural Integrity Associates, Inc., Report No. 1700553.402, Rev. 2, December 2019, "Flaw Tolerance Evaluation of the North Anna Unit 1 and 2 Hot Leg Surge Line Nozzles using ASME Code Section XI, Appendix L"

Serial No.20-049 Docket Nos. 50-338/339 Revised License Renewal Commitment Page 4 of 4 cc: U . S. Nuclear Regulatory Commission - Region II Marquis One Tower 245 Peachtree Center Ave., NE Suite 1200 Atlanta, GA 30303-1257 NRC Senior Resident Inspector North Anna Power Station Mr. G. Edward Miller NRC Senior Project Manager - North Anna U.S. Nuclear Regulatory Commission One White Flint North Mail Stop 09 E-3 11555 Rockville Pike Rockville, MD 20852-2738 Mr. Vaughn Thomas NRC Project Manager - Surry U.S. Nuclear Regulatory Commission One White Flint North Mail Stop 04 F-12 11555 Rockville Pike Rockville, MD 20852-2738

Serial No.20-049 Docket Nos. 50-338/339 Attachment Structural Integrity Associates, Inc., Report No. 1700553.402, Rev. 2, December 2019, "Flaw Tolerance Evaluation of the North Anna Unit 1 and 2 Hot Leg Surge Line Nozzles using ASME Code Section XI, Appendix L" Virginia Electric and Power Company (Dominion Energy Virginia)

North Anna Power Station Units 1 and 2

/axline@structlntcom 5215 Hellyer Avenue, Suite 210, San Jose, CA 951381 (408) 833-7316 REPORT NO. 1700553.402 REVISION: 2 PROJECT NO. 1700553.00 December 2019 Flaw Tolerance Evaluation of the North Anna Unit 1 and 2 Hot Leg Surge Line Nozzles using ASME Code Section XI, Appendix L Prepared For:

North Anna Power Station, Units 1 & 2 Dominion Energy 5000 Dominion Blvd Glen Allen, VA Contract Number 70330436, Revision 0, with PO CO 002 & 003 Structural Integrity Associates, Inc.

San Jose, California info@structint.com m 1-877-45!-POWER e structint. com@

REVISION CONTROL SHEET Report Number: 1700553.402 Flaw Tolerance Evaluation of the North Anna Unit 1 and 2

Title:

Hot Leq Surqe Line Nozzles usinq ASME Code Section XI, Appendix L Client: Dominion Energy, Virginia SI Project Number: 1700553.00 Quality Program: ~ Nuclear D Commercial SECTION PAGES REVISION DATE COMMENTS All All 0 10/29/2019 Initial Issue All All 1 12/13/2019 Revised to incorporate Client comments All All 2 12/14/2019 Revised to clarify Non-Proprietary status of the Report.

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TABLE OF CONTENTS

1.0 INTRODUCTION

.... .............. ......... ....... ....... .......... ............ .... .......... .. ...... .... .. ........... ....... 1-1 2.0 TECHNICAL APPROACH .. ........ ... ........ ......... ........... ..... .................... ..... .. .......... ....... .... . 2-1 2.1 Heat Transfer Coefficients ... ........ .. ... .... ..... ..... ........... ..... .. .................. ... ....... ... .. .. ....... 2-2 2.1.1 Swirl Penetration ... .. .... ................... .. .. .. .. ... ........ ... ...... ............ ...... ..... ........ .. ....... 2-2 2.1.2 Internal Flow without Stratification ..... .. .... ..... ... ... .. .. .............. .. .. ...... ...... .... .. ........ 2-3 2.1.3 Stratified Flow and Natural Convection ....... ... .. .... .. .. .. ................................. .. ..... 2-4 2.1.4 Effective Heat Transfer Coefficient for Region of Thermal Sleeve .. .. .. .. .. ...... ..... 2-4 2.2 Flow Rates .... .. .... ... ... .... ... .. ...... ... ... .. ....... .. ..... .. .. .. .. .. .. .. ... ... .. .. .. ....... .. .. .. .. ... ..... .. .... .. ... 2-6 2.2.1 Heatup/Cooldown lnsurge/Outsurge Stratification Sub-Events ........ .. .. .............. 2-6 2.2.2 Other Events ...... ...... .. .... ......... .... .... .. .. .. ....... ..... .. ..... ... .. .. .. .. .. .. .... ... .. ..... .... ... .... ... 2-7 2.3 Stratification .. ... .... .. .. .... .. ...... ... ... .... .. ................ .. .... ........ ..... ... .... .... ... ... ............ .... ...... 2-7 2.4 Thermal Zones .... ... ... .. .. .. .. .. ... ........ .. ........ .. ............ .. .. ........ .. .. .. .. .. ........... ...... .... .... .. ... 2-9 2.5 Heat Transfer Coefficient Results .. .. ....... ...... .... .. .. .. .. .. ..... .. .. ................ .. .. .... ..... .... ... 2-10 3.0 DESIGN INPUTS AND ASSUMPTIONS ..... .... .......... ...... ....... .. ... ....... .... ...... .. ............. ..... 3-1 3.1 Piping Dimensions and Materials and Water Properties .. .... ............. ..... .................... 3-1 3.2 Operating Conditions ... .... ........ ... .... .... ............ ....... .... ... .. ...... .......... .. ..... .......... .. ......... 3-2 3.3 Piping Loads ... .. ... .......... .. ........................ .. ... .... .... ....... ........ ....... .. .. ... .... .. .. ....... ... .. .... . 3-2 3.4 Thermal Transients for Fatigue Crack Growth Analysis .. .. .. .... .. .... .. .. .. .. .... .. ... ............ 3-3

3. 5 Analysis Parameters and Approaches .......... .. ........... .. .. ...... ............ .. .... .. .. .. .. .. ... .. .. ... 3-4 4.0 STRESS ANALYSIS .......... ...................... ... .... ....................... ................ ........ ... ............ ... 4-1 4.1 Finite Element Model. ......................... ...... ..... .. .. .. ..... ........ ..... ... .... .. .... .. .... .... .. ..... ....... 4-2 4.2 Thermal Anchor Movements ..... ........... .... .. ........ .... ......... ..... .. .. .. .. .... .... .... ..... ..... ....... . 4-4 4.3 Materi al Properti es .............. ... .................. .. .... ................ .. .. .. ... ... .. .. ... .. .. .. .... .. .. ... .... .... 4-4 4.4 Thermal/Mechanical Stress Analysis .............. ......... .. .. ...... ...... .. ..... ..... ........ ....... .. ..... 4-4 4 .5 Deadweight Analysis .. .. .. .. .... .. ..... .. ......... .. .. .... .. ............ ... .... ... .... .. .. ........... .. ...... .... ..... 4-5 4.6 Mechanical Boundary Conditions ... ....... .... .... ... ...... .. ........ .. ........... ... .. .. ... ... ..... .... ..... .. 4-5 4.7 Stress Analysis Results ... .. ...... .. ... ... .. .. .. .. ..... .. .... ... ... ................ .. .... .... ... .... .. .. .. .. .. .. .. ... 4-6 5.0 ALLOWABLE FLAW SIZE EVALUATION ...... .... .. ..... ...... ........................ .......... ........ ....... 5-1 5.1 Allowable Flaw Size Determination .......... .. ........ .... ........ .. .... ........... ...... .... .. .............. . 5-1 5.2 Interface Loads ....... .............. .. ...... .. ..... .. .. ...... ..... ................ ....... ... .. .. .......... .. .. ... ......... 5-1 5.3 Load Combinations ... .. .. ............................. .. .. ................ ... ... .. ... ... ..... .. ......... ... .. .. ...... . 5-1 5.4 Material Properties fo r Allowable Flaw Size Determination .... .... ........ .. .... ........ ..... ..... 5-2 5.5 Welding Process .... ........... .... ........ ... .. ... ... .... ... ... ........ .. .... .... .... .... .. .. .. ... .. .. ... ... ........... 5-3 5.6 Z-Factor ... ........ .... .. ....... ... ......... ..... .. .: ........ .. ........ .. ..... .. .. .. ... ..... .. ... .. ...... .. .. .. .. .... ... ...... 5-3 5.7 Allowable Circumferential Part Through-Wall Flaw .. ........ ... .. .... .. .. .. .... .. .... .... .. ...... ..... 5-3 Report No. 1700553.402 R2 Page I iii SJ Structural Integrity info@structint.com m 1-877-4SI-POWER e structint.com ~

Associates. Inc.

5.8 Allowable Axial Part Through-Wall Flaw .................... ...... ................ .. ........ ......... .... .. . 5-5 6.0 CRACK GROWTH EVALUATION .................................................................................... 6-1 6.1 Loads ...... .. ............................ ............................. ... .. ......... ...... .. .................. ... ... .. .. ... .. .. 6-1 6.2 Thermal Transients for Crack Growth Analysis .......... .. ............................... ............... 6-2 6.3 Weld Residual Stress .... .. .... .... .. ............ .. .. ......... ...... ........ .... .. .. ........ .. .. .. ..... .. ........ .... . 6-2 6.4 Stress Intensity Factors ....... ...... .......... .. ..... .. ....... ........... ............................... .. .. ......... 6-3 6.5 Postulated Initial Surface Flaw ................. ............. .. ........... ...... ..... ....... .. ................... . 6-4 6.6 Stainless Steel Fatigue Crack Growth Law .......... .... .... .. ...... .. .. ........ .. ........................ 6-6 6.7 Crack Growth Analysis .. .. .... ........ ... .... ....... ..... .... ... ... ........ ... ... ......... .......... ... .............. 6-7 6.8 Crack Growth Results .. ....... ........... .. ...... .. ..... ......................... .. .... .. .... .. ......... .... .. .. ..... 6-9 7.0

SUMMARY

AND CONCLUSIONS ................................................................................... 7-1

8.0 REFERENCES

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Associa/es. Inc.

LIST OF TABLES Table 3-1. Deadweight and Seismic Piping Interface Loads ....... ... ... .. ...... .. ........ .. ...... ....... .. ..... 3-6 Table 3-2. Thermal Expansion Piping Interface Loads ... .... ... .. ...... .. ..... .. ..... ............ ...... .......... . 3-7 Table 3-3. Number of Thermal Transient Cycles ...... ...... ... .. ..... ............ ..... ......... ... .. .. .. .. .. .. ....... 3-8 Table 3-4 . Thermal Transients .... ... .... .... .. .. ................ ....... ....... ... ... ... .... .... .. .. .. ..... .. ....... .. ... ... .... 3-9 Table 3-5. Equivalent Conductivity of Water Gap (ke) ... ...... .. .. .. .. ............. .. ......... ........ ....... .... . 3-14 Table 3-6. Total Added Thermal Resistance, Thermal Sleeve and Water Gap .. ... .... ..... ........ 3-14 Table 4-1. Piping Stress Values - Unit 1 versus Unit 2 ........ ... ... .. ... ........... .. .. ...... ....... ......... ... .. .4-2 Table 4-2. Thermal Anchor Movements for Design Conditions .. .......... ............ ............... ...... .... 4-7 Table 4-3. Material Properties used in the Finite Element Model {TP316 Stainless Steel) .. ... ..4-7 Table 4-4. Thermal Transients Evaluated .................. ...... ....... .... ... ... .. ..... .... ............ .. .... ........ ... 4-8 Table 5-1: Geometry, Operating Conditions, and Material Properties ..... .. ... ........ ... .............. ... 5-6 Table 5-2. Allowable Part Through-Wall Circumferential Flaw Sizes for Unit 1 .... .......... .... ...... 5-7 Table 5-3. Allowable Part Through-Wall Circumferential Flaw Sizes for Unit 2 ...................... .. 5-8 Table 5-4 . Allowable Part Through-Wall Axial Flaw Sizes for Unit 1 and Unit 2 ..... .. .... .. ....... ... 5-9 Table 6-1: Thermal Transients for Crack Growth Analysis ... ................ .............. ... ....... ... ...... .. 6-10 Table 6-2. Initial Flaw Sizes for Circumferential Flaw .......... ................................................... 6-1 1 Table 6-3 . Initial Flaw Sizes for Axial Flaw .. ................. .. ...... ......... .. .... .... .. .. ..... ................ ....... 6-12 Table 6-4 . Crack Growth Results for Circumferential Flaw - Lower Stratification Loading ..... . 6-13 Table 6-5. Crack Growth Results for Circumferential Flaw - Upper Stratification Loading ...... 6-14 Table 6-6. Crack Growth Results for Axial Flaw - Lower Stratification Loading ... .. ................. 6-15 Table 6-7. Crack Growth Results for Axial Flaw- Upper Stratification Loading ....... .. ... .. .. .. .. .. 6-16 Report No. 1700553.402 R2 Page I v e

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LIST OF FIGURES Figure 2-1. Unit 1 Surge Line Nodes ....................................................................................... 2-11 Figure 2-2. Unit 2 Surge Line Nodes ............ .. ............... ........ ..... .......... ............. ...................... 2-12 Figure 2-3. Stratification Geometry ............. .. .... ........................................... .. ...... .. .. ....... .. .. .... 2-13 Figure 2-4 . lnsurge/Outsurge Stratification Modeling ...................... ...... ......................... ......... 2-13 Figure 2-5 . Swirl Flow Thermal Zone ...................................................................................... 2-14 Figure 2-6 . Stratified Lower Region Thermal Zone ................................. .. ........... .. .. .. ............. 2-15 Figure 2-7. Stratified Upper Region Thermal Zone ................ .. .... .... ..................... .. ... ........... .. 2-16 Figure 4-1. Unit 2 Surge Line Dimensions (Global Coordinates) ...... ....................................... .4-9 Figure 4-2. Components Included in the 3-D Finite Element Model ....................................... .4-10 Figure 4-3. Applied Boundary Conditions (Deadweight) ......... .. ...... .. ........ .. .. .. .. .... .... .............. 4-11 Figure 4-4. Applied Thermal Boundary Conditions for Thermal/Mechanical Transient Analysis (Lower Line Stratified) .. ... .................... ...... .... .. ....... ...... .. .. .. ............ ... .. ... 4-12 Figure 4-5 . Applied Thermal Boundary Conditions for Thermal/Mechanical Transient Analyses (Upper Line Stratified) ............. .. .... .. ................ .. ........ .. ................ .. ....... .4-13 Figure 4-6. Temperature Contour During Transient MOP320 (Time= 9,920 seconds) ...... .... 4-14 Figure 4-7. Applied Mechanical Boundary Conditions for Thermal/Mechanical Stress Analyses .. ........ .................... ... .. .. .. ... ... .. .. .. .. .... ........... .... ................. ....... .. ........... .. 4-15 Figure 4-8, Total Stress Intensity Due to MOP320 Transient (Lower Line Stratified) ............ .4-16 Figure 4-9. Total Stress Intensity Due to MOP320 Transient (Upper Line Stratified) ............ . 4-17 Figure 4-10. Stress Path Locations ... .... .. ... .. .. ..... .... ..... .. ... .......... ..... .... ... .. .. .... ... .... ................ .4-18 Figure 6-1. Through-Wall Residual Stress as a Function of Depth .. .............................. ......... 6-17 Figure 6-2. Semi-Elliptical Flaws on the Inside Surface of a Cylinder ...... .. ............................. 6-18 Report No. 1700553.402 R2 Page I vi

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1.0 INTRODUCTION

North Anna Power Station (NAPS) Units 1 and 2 has committed to inspecting the weld between the surge line nozzle and the hot leg (a NUREG/CR-6260 location [1]) every forty months as part of the 60-year license renewal commitment, as specified in the NAPS UFSAR

[31, Section 8.3.2.4]. NAPS has requested a flaw tolerance evaluation of the weld to justify an extended reinspection interval of ten years. In addition, for Subsequent License renewal to 80 years, it is necessary to address this same location , plus others in the surge line (if any) that may be more limiting, with respect to fatigue crack growth.

A flaw tolerance evaluation in accordance with the 2013 Edition of the ASME Code,Section XI ,

Nonmandatory Appendix L [2] is performed in this project to manage environmentally-assisted fatigue at the critical locations of the NAPS Units 1 and 2 pressurizer surge line.

As part of initial License Renewal (LR), selected welds have been included in the NAPS augmented in-service inspection (ISi) program . Specifically, the pressurizer surge line locations chosen by NAPS for inspection were the NAPS Units 1 and 2 Pressurizer Surge Line Nozzle Hot Leg component full penetration nozzle-to RCS pipe welds (SW-31 (U 1) and SW-36 (U2)), and the full penetration nozzle-to-surge line butt welds (No. 37 (U1) and 3 (U2)). The NRC Safety Evaluation Report (SER) for initial LR as documented in NUREG-1766 [41] states that the "surge-line-to-hot-leg pipe connection is a limiting location from a fatigue perspective when considering reactor water environmental effects. NAPS has committed to additional actions regarding this location during the period of extended operation . The staff considers the surge line a bounding example to represent the effects of the reactor water environment on the fatigue life of pressurizer components during the period of extended operation."

As a result of this and subsequent evaluation of environmentally assisted fatigue (EAF) at North Anna Power Station (NAPS) Units 1 & 2 for Subsequent License Renewal (SLR), the full penetration nozzle-to-RCS pipe weld locations (SW-31 (U1) and SW-36 (U2)) were screened out. There is a thermal sleeve in the RCS hot leg nozzle that connects to the surge piping [5] .

This thermal sleeve also protects RCS pipe welds (SW-31 (U1) and SW-36 (U2)). Taken together, the lower Cumulative Usage Factor (CUF) values and existence of a thermal sleeve has resulted in screening out of RCS pipe welds SW-31 (U1) and SW-36 (U2) for EAF considerations, and they do not require management for EAF during SLR per Reference 38 .

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The Appendix L flaw tolerance evaluation is contained in References 5 through 9, and focuses on the surge line piping alone, and not the RCS piping , as the RCS pipe welds (SW-31 (U 1) and SW-36 (U2) have been screened out, as documented in Reference 38. The limiting Appendix L locations are the hot leg nozzle-to-surge line butt welds (No. 37 (U1) and 3 (U2)).

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2.0 TECHNICAL APPROACH The evaluation was performed in accordance with the requirements of the 2013 Edition of the ASME Code,Section XI, Appendix L [2] . Effective 8/17/2017, the latest ASME Code edition approved by the NRC is the 2013 Edition, which includesSection XI , Appendix L. Code Case N-809 [32], which includes the latest crack growth data, has been approved by ASME. Although Code Case N-809 has not been officially endorsed by the NRC, the NRC has reviewed and approved the use of CC N-809 in precedent license renewal commitments pertaining to fatigue of surge line welds for Turkey Point (Submittal: ML12152A156 and Approval: ML 1314A595) and St. Lucie (Submittal: ML15314A160 and Approval: ML16235A138). Both evaluations used the same crack growth models and data , which were published at an EPRI conference prior to ASME approval of the Code Case.

One location is evaluated on Unit 1 (node170/ 171) and one location is evaluated for Unit 2 (nodes 170/171) for EAF considerations for license renewal and subsequent license renewal.

While not required for initial license renewal or subsequent license renewal, Dominion also contracted SIA to evaluate additional locations for Unit 1 (nodes 175, 190, and 198 per Reference [6]) and Unit 2 (nodes 173, 187, and 201 per Reference [6]) for asset management considerations.

The methodology used to determine the successive inspection schedule consists of the following principal tasks:

• Determine the loads [5] and stresses [6, 8] at the critical locations in the surge line (Sections 3.0 and 4.0), which has been identified as the butt weld No. 37 (U1) and 3 (U2) between the surge piping and the nozzle off the Reactor Coolant System (RCS) hot leg

[38] .

• Use the stresses at the critical locations to determine the allowable flaw depths for various service levels [7] (Section 5.0).

• Postulate hypothetical flaws at the critical locations. Select appropriate crack models to simulate the postulated flaws [9]. Both axial and circumferential flaws will be considered.

(Section 6.0).

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• Use the stresses determined at the critical locations and the selected crack models to compute stress intensity factors for all the applicable normal and upset condition loads.

Perform fatigue crack growth analyses with the resulting stress intensity factors to determine the end-of-evaluation-period flaw size and/or determine the time (allowable operating period) necessary for the postulated initial flaw to grow to the allowable flaw depth [9] (Section 6.0) .

• The steps above are repeated for the additional asset management locations.

The following Sections 2.1 through 2.5 provide a high level technical overview of the heat transfer, fluid dynamics, transient definitions used in the development of the loads for assessment of the various locations.

2.1 Heat Transfer Coefficients As described in the following sections , heat transfer coefficients for specific locations in the surge line are calculated based on surge flow during thermal transients. These coefficients are needed for the thermal transient analysis (Section 4.0).

2. 1. 1 Swirl Penetration Even when there is no insurge or outsurge flow in the surge line, ([lnsurge] equals surge line net flow into the pressurizer, and [Outsurge] equals net flow out of the pressurizer) , there is swirl penetration from the reactor coolant system (RCS) hot leg into the surge line such that forced convection will exist, and the RCS nozzle-to-surge line piping butt weld will remain at the hot leg temperature, unless there is a large outsurge of flow from the surge line. Guidance is taken from the proprietary EPRI Materials Reliability Program (MRP) document MRP-132 [1 OJ.

Adjustments to MRP-132 are taken from a supplemental guidance report MRP-146S [11] to ensure that the heat transfer coefficient formulas are conservative, as described in MRP Letter 2007-013 [12] .

The methodology outlined was used to model the swirl penetration due to hot leg flow when there is no flow in the surge line, similar to that used in the Surry SLR application [43].

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2. 1.2 lntemal Flow without Stratification For surge line insurge and outsurge, the following expression is defined for turbulent flow in a uniformly filled cylinder:

h *d [13, Eq. 6-4a]

- - = Nu= 0.023

• Re 08

• Pr 04 k

where:

h = heat transfer coefficient (Btu/hr*ft2* °F) k = thermal conductivity of the fluid (Btu/hr*ft*°F) d = pipe inner diameter (ft)

Nu = Nusselt number (dimensionless)

V = kinematic viscosity of the fluid (ft2/sec)

Re = Reynolds number of fluid (dimensionless)

Pr = Prandtl number of fluid (dimensionless)

The Reynolds number, Re, for turbulent flow is defined as:

Re= V*d = 4

• Q V Jf

• V

• d where:

V = bulk velocity of the flow (ft/sec)

Q = volumetric flow rate (ft 3/sec)

The above heat transfer coefficient equation is valid for Reynolds numbers greater than 2500

[13]. Water properties are calculated at the bulk temperature of the fluid. When the Reynolds number is less than 2500, the flow is in a transition zone between turbulent and laminar conditions. For laminar flow, the following may be used to estimate heat transfer coefficients.

h *d

- - = Nu = 4.364 [13, Eq. 5-107]:

k The turbulent flow formula produces a significantly conservative (much higher) heat transfer coefficients, h, in all cases. The expression for turbulent flow is therefore conservatively used for all laminar flow conditions.

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2. 1. 3 Stratified Flow and Natural Convection Stratification is expected to occur in the surge line, when the flow rate is insufficiently high to fill the surge piping. For natural convection (zero flow) and stratified flow conditions, it is recommended to apply the following heat transfer coefficient to the entire pipe inner surface [14,

p. 3.7-1]:

h= 0.56*k[,B*g*d 3 .t,,,T -v_ 2

• Pr]o2s d

where:

h = heat transfer coefficient (Btu/hr*ft2- ° F) k = conductivity of the fluid (Btu/hr*ft* °F)

~ = temperature coefficient of volume expansion (fluid) (1/°F) g = acceleration of gravity (ft/sec 2

)

d = pipe inner diameter (ft) t,,,T = ITo - T(O)l/2 (°F)

V = kinematic viscosity of the fluid (ft2/sec)

Pr = Prandtl number of fluid (dimensionless)

To = temperature of ambient fluid (°F)

T(O) = initial fluid temperature (°F) k, ~. v, and Pr are calculated at temperature T = (To+ T(0))/2. The ambient fluid temperature, To ,

is modeled to be 120°F, which is typical of containment temperatures . The initial fluid temperature, T(O), is set as the warmer T1op temperature (Top of Surge Line), which conservatively yields a greater ~ T.

2. 1.4 Effective Heat Transfer Coefficient for Region of Thennal Sleeve There is a thermal sleeve in the RCS hot leg nozzle that connects to the surge piping [5] and the sleeve and the water gap between the sleeve and the nozzle wall provide additional thermal resistance. The following formula gives the heat transfer of fluid in enclosures.

[13, Eq. 7-64]

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where:

k = thermal conductivity (Btu/hr-ft-°F) ke = equivalent conductivity (Btu/hr-ft-°F)

C = 0.11 (dimensionless) for 6000 < Gr8 Pr < 106

= 0.40 (dimensionless) for Gr8 Pr> 106 [13, Table 7-3]

Gro = Grashof number= gl3(T1 - T2)83/v 2 (dimensionless)

Pr = Prandtl number (dimensionless) n = 0.29 (dimensionless) for 6000 :< Gro Pr< 10 6

= 0.20 (dimensionless) for Gr8 Pr> 106 [13, Table 7-3]

= 0.0 for a horizontal annulus [13, Table 7-3]

= annulus length (ft) , not used since m = 0

= annulus width (ft)

= acceleration due to gravity= 32.174 ft/sec 2

= temperature coefficient of volume expansion (fluid) (11°F)

= temperature difference across annulus (°F)

= kinematic viscosity of the fluid (ft2/sec)

Fluid properties are taken at the average temperature of the fluid. The thermal resistance due to the thermal sleeve itself is given by t/kTs, where tis the thermal sleeve thickness and kTs is the thermal sleeve conductivity. The total added resistance, R, is:

R = t/kTs + 8/ke (hr-ft2 -°F/BTU)

The effective heat transfer coefficient, hett, for the nozzle in the region of the thermal sleeve is, therefore:

hett = 1 / (1/hsiv + R) (BTU/hr-ft2-°F) where:

hs1v = heat transfer coefficient at thermal sleeve inside surface (BTU/hr-ft2-°F)

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2.2 Flow Rates The surge line flow rates are described in the following sections.

2.2. 1 Heatup/Coo/down lnsurge/Outsurge Stratification Sub-Events For Heatup/Cooldown (HU/CD) lnsurge/Outsurge (1/0) stratification sub-events, flow rates are not specifically available. The initial flow rate was therefore selected to be 70 gpm [5], which was calculated to be the approximate upper limit of flow during which stratification can still occur. This flow rate also produces a stratified interface of approximately mid-height, which produces the most conservative (largest) pipe bending stresses. For the insurge and outsurge portions of the event where stratification disappears, the flow rate was selected to be 200 gpm

[5], consistent with similar, previously performed calculations.

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2.2.2 Other Events For all other events, constant outsurge flow rates are based on a spray bypass flow of 2 gpm, except as follows.

200 gpm outsurge was conservatively chosen for Heatup and Cooldown (not including the 1/0 stratification sub-events), such that the entire surge line is uniformly filled with fluid at the pressurizer temperature.

When flow rate is specified to be zero or approximately zero, 2 gpm outsurge (approximate spray bypass flow rate) is selected for analysis purposes. This flow rate was selected to ensure numerical stability in the stratification calculations and produces a conservative or negligible effect on the resulting thermal profile.

2.3 Stratification Thermal stratification can potentially occur during each transient if the required flow rate and temperature conditions are mutually satisfied. If the flow rate is high enough, then an insurge or outsurge will sweep the line, preventing thermal stratification . If the flow rate operates in a sufficiently low flow regime and there is a temperature difference between the pressurizer and RCS hot leg , then a thermally stratified interface may develop due to the buoyancy caused by the density differences of the fluid above and below the fluid flow boundary layer. When the low flow rate changes but remains within the flow regime necessary to maintain stratification conditions, the stratified interface may move up or down, resulting in local stress perturbations.

For each transient definition, the surge flow rate history, in addition to pressurizer water and hot leg temperatures are provided. It is desirable to use these parameters to determine if stratification exists at each point in time during the transient, and if so to also determine the stratified interface level and the top and bottom surge line temperatures.

EPRI Thermal Stratification, Cycling, and Striping (TASCS) Report [14] is used to determine the steady state stratification interface height in the surge line.

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Figure 2-3 shows the geometric parameters used [14, p. 3.1-9]. Interface level H and critical depth Ye are measured from the top if the flowing water is on the top or the bottom if the flowing water is on the bottom and are located at the outlet of flow.

The stratification level at any point in time is applied uniformly (no surge line slope, resulting in stratification in all horizontal portions).

The pipe Richardson number, Rip, is calculated to determine whether the flow will remain stratified , based on the input temperatures and flow rates~

For stratification and insurge/outsurge (1/0) cycling, which occurs during heatup and cooldown events , the maximum system I:),, T is conservatively used to define the stratification in the surge line. The top fluid temperature is conservatively determined to be equal to the pressurizer water temperature (T PzR ) , and the bottom temperature is conservatively determined to be the pressurizer temperature minus the maximum stratification I:),, Ts1ra1 specified. When outs urge is large enough , the entire fluid temperature becomes T PZR - When insurge is large enough , the entire fluid temperature becomes (T PZR - I:),, Ts1ra1).

The temperature ramp rate during insurges and outsurges for the heatup and cooldown stratification 1/0 sub-events was taken as 2400°F/hr, consistent with similar, previously performed evaluations. Minimum1000 second hold times are conservatively added between each change in flow rate , to achieve thermal equilibrium (quasi-steady state conditions) .

Validation of thermal equilibrium is performed as part of the stress analyses.

Figure 2-4 illustrates the modeling of the postulated heatup/cooldown sub-events. The sequence of events is given as follows:

1. The surge line starts at a stratified condition.

2. An insurge of flow in the surge line occurs , filling the pipe with RCS fluid .

3. The surge line returns to stratified conditions again.

4. An outsurge of flow in the surge line occurs , filling the pipe with pressurizer fluid.

5. The surge line returns to its stratified condition.

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2.4 Thermal Zones During normal conditions, even when significant surge flow exists in the surge line, swirl flow from the RCS creates a generally uniform temperature around the RCS hot leg surge nozzle and a portion of the adjacent piping upstream of the surge nozzle to piping weld. Using the Unit 2 bounding surge line finite element model (See Section 4.0 for description and justification) ,

Figure 2-5 shows the swirl flow thermal zone near the hot leg nozzle. During outsurges with large flow rate, the swirl flow may be swept away, producing an essentially uniform temperature in the surge line and surge nozzle.

The surge line has two, separate horizontal sections , each of which has the potential to stratify.

Therefore, for conservatism, stratification is analyzed separately for each of the two horizontal segments to determine the most conservative condition . Using the surge line finite element model (See Section 4.0) , Figure 2-6 and Figure 2-7 show the lower and upper stratified thermal regions, respectively. A total of four thermal zones are used in the analysis and are described in Section 3.4.

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2.5 Heat Transfer Coefficient Results Using the methodology described in Section 2.1 through Section 2.4, calculated heat transfer coefficients [5] in Table 3-4 are to be applied in three regions.

• hsrg is for the region of the surge piping,

• hs1v is for the region of the nozzle upstream of the thermal sleeve, and

• hett is the effective heat transfer coefficient for the nozzle in the region behind the thermal sleeve.

For the RCS hot leg, the hot leg temperature is applied directly to the inside surface, because of the high flow rate.

Table 3-5 shows the calculation and results for the added resistance to heat transfer resulting from the thermal sleeve and water gap. For locations with both thermal sleeve and water gap, Table 3-6 shows the total added resistance, based on the thermal sleeve thickness, and an estimated value of kTs of 10 Btu/hr-ft-° F for the thermal sleeve stainless steel material over a range of temperatures [6].

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+ y

. t/Ul>"T.#

~~x

'*l!S

+ z.

1 t: Z:*llO ,

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,_t8* \

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Figure 2-1. Unit 1 Surge Line Nodes Source: Reference 3 Report No. 1700553.402 R2 PAGE I 2-11 fr Structural Integrity Associates. Inc info@structint.com m 1-877-45!-POWER e structint.com@

w

,, j

, /'1 -RC- £.{/D-;l5'otR:git!

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Figure 2-2. Unit 2 Surge Line Nodes Source: Reference 4 Report No. 1700553.402 R2 PAGE 12-12 SJ Structural Integrity Associates, Inc i nfo@structint.com m 1* 877-4SI-POWER e structint.com @)

Figure 2-3. Stratification Geometry HU/CD 1/0 Stratification Event Modeling 500.0 Initial= 70gpm outsurge TPzR 450.0 .

I I

I 400.0 ,

I I I

350.0

• I I I I

2400 °F/ hr I I E 300.a I

• I

': I

.3~ I 250.0 a,

I - -Ttop Cl. I I E

~ 200.0 ,I I I

- -

• Tbot II

• I I 150.0

---

1 PZR - 11 I strat I I

""- I 100.0 200gpm insu rge - 70gpm outs urge I 200gpm outsurge I 70gpm outsurge I 50.0 0.0 1000 2000 3000 4000 SOOD 6000 7000 Time (seconds)

Figure 2-4. lnsurge/Outsurge Stratification Modeling Report No. 1700553.402 R2 PAGE 12-13 13 Structural Integrity Associates. Inc info@structint.com m 1-877-451-POWER e structint.com @)

1 ANSYS RIU N:DAL SOI.DTIC1\J S'IEP=l APR 11 2018 SOB =l 14:47: 37 TIME=l Plill ID . 1 BFETEMP (AVG)

CMX =2.35215 SM.\l =620 SMX =653 Swirl flow thermal zone Tnoz - heff I ' \

I I

I I

I I

\

\

\

\,,,

620 627 . 333 634 .667 642 649 . 333 623 . 667 631 638 . 333 645 . 667 653 CASEl - GENERAL THERMAL EXPANSIC1\J Figure 2-5. Swirl Flow Thermal Zone (Units of the plots are F) 0 Report No. 1700553.402 R2 PAGE 12-14 SJ Structural Integrity Associates. /11c info@structint.com m 1-877-451-POWER e structint.com @)

1 ANSYS Rl4.5 llilJAL SOI.DTIGJ S'IEP=2 APR 11 2018 SOB =1 14: 49:04 TIME=2 Piill ID . 1 BFETEMP (A"\X;)

CMX =4 .48603 SM\J =110 SMX =450 Stratified lower region thermal zone

\

I

,I

/

/

/

110 185. 556 261 .111 336 . 667 412 . 222 147.778 223.333 298.889 374 .444 450 CASE2 - I..GiER REGIGJ STRATIFIED Figure 2-6. Stratified Lower Region Thermal Zone 0

(Units of the plots are F)

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l ANSYS Rl 4.5 N:DAL SOIDTICN STEP=3 APR 11 2018 SUB =1 14: 49 :18 TIME=3 PLOI' NJ . 1 BEETEMP (A'ifu)

CMX =2 . 40789 SM'-J =110 SMX =450 Stratified upper region thermal zone 110 185 . 556 261.111 . 336 . 667 . 412 . 222 147.778 223 . 333 298 .889 374 . 444 450 CASE3 - UPPER REGICN STRATIFIED Figure 2-7. Stratified Upper Region Thermal Zone 0

(Units of the plots are F)

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3.0 DESIGN INPUTS AND ASSUMPTIONS The following design inputs have been used to perform stress and flaw tolerance evaluations .

3.1 Piping Dimensions and Materials and Water Properties Reactor Coolant System (RCS) Hot Leg Piping [24, Figure 21 Inside Diameter: 29.035 inches Wall Thickness: 2.445 inches Surge Line Piping (14-inch Schedule 160) [3, page 371 [4, page 351 Outside Diameter: 14.000 inches Wall Thickness: 1.406 inches Inside Diameter: 11.188 inches Material: A376 TP316 Insulation: 4-inch thick, type Class X Hot Leg Surge Nozzle [3, page 371 [4, page 351 Outside Diameter: 14inches Wall Thickness: 1.250 inches Inside Diameter 11.5 inches Material: A182 F316 Thermal Sleeve Presence of Thermal Sleeve [161 Inside Diameter: 11.188" [modeled same as pipe]

Water properties are taken from Reference [39 , Table 331.

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3.2 Operating Conditions The operating conditions for the surge line are:

Hot Leg Temperature (0% power no load): 547°F [40]

Hot Leg Temperature (100 % power): 612 °F [40]

Pressurizer Temperature (100% power) : 652°F [40]

RCS Flow Rate: 103,667 gpm/loop [40]

3.3 Piping Loads Piping loads are evaluated at four limiting locations in the Unit 1 surge line, as shown in Figure 2-1, and at four limiting locations in the Unit 2 surge line, as shown in Figure 2-2.

Since the entire surge line piping system is modeled using FEA, this will automatically account for piping interface loads (moments and forces) induced by thermal expansion during thermal stratification. However, it is necessary to also account for dead weight, thermal expansion ,

thermal anchor movements, and seismic loads.

Table 3-1 summarizes the bounding piping forces and moments for deadweight and seismic.

Thermal anchor movements are addressed in Section 4.2.

Table 3-2 summarizes the bounding piping forces and moments for thermal expansion.

These loadings are used for both the allowable flaw size determination (Section 5.0) and the fatigue crack growth (Section 6.0) .

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3.4 Thermal Transients for Fatigue Crack Growth Analysis The most significant transients occur during heatups and cooldowns . NAPS has implemented a Modified Operating Procedure (MOP) intended to mitigate insurge/outsurge cycling . The method, known as Modified Steam Bubble MOP, energizes the backup heaters during heatups and cooldowns, which has the effect of producing constant outflow in the surge line. For conservatism , and while not observed during plant operation , 2 contingency transients per Heatup or Cooldown are postulated to account for any future cycling that may potentially occur, despite the MOP.

The surge line is also subjected to other normal and occasional transients that occur in the RCS.

The reference [19] White Paper, identifies transient loads relevant to fatigue crack growth in surge line piping by reviewing the results of similar 3-loop Westinghouse plants previously analyzed for flaw tolerance with environmental effects. It determined that the crack growth contribution from transients not associated with insurge/outsurge stratification events during heatups and cooldowns , is small , when accounting for more realistic numbers of occurrences ,

combined with a very conservative number of heatups and cooldowns postulated to occur over a 10-year period. The relevant RCS transients identified include Plant Heatup, Plant Cooldown, Large Step Load Decrease, Reactor Trip, and Operating Basis Earthquake.

Transient data come from multiple sources of information. For the insurge/outsurge stratification cycling during Heatups and Cooldowns, conservative transients designed to bound the problem were devised, as described in Section 2.3. Contingency stratification cycling is postulated to occur at a pressurizer temperature of 450°F (440 psig pressure) when the system temperature differential is maximized, as shown on Figure 2-4. For Heatup, a 100°F/hr ramp rate is used ,

which is the same ramp rate as the pressurizer. For Cooldown, a 200°F/hr ramp rate is used ,

wh ich is the same ramp rate as the pressurizer. For other relevant transients, guidance from MRP-393 [18] were used to adequately characterize realistic transient behavior.

Table 3-3 lists the number of occurrences for thermal transients includ ing MOP HUs and CDs during a 10-year interval. For the contingency insurge/outsurge stratification cycles, the numbers of cycles for a 10-year period were prorated based on 50 Heatup and Cooldown cycles (or 200 over 40 years). Transients for emergency and faulted conditions are not included since the resulting thermal transient stresses are intended to support the ASME Code, Appendix L Report No. 1700553.402 R2 PAGE I 3-3 SJ Structural Integrity Associates. Inc info@structint.com m 1-877-451-POWER G structint.com @)

flaw tolerance evaluation of crack growth during normal plant operation. As more fully described in Section 5.3, the loadings of concern are those related to plant specific loading cycles consistent with the plant design and operating practices, or actual plant operating data.

Table 3-4 shows the thermal transients to be analyzed. Temperatures are applied in four thermal zones, as described in Section 2.4. (1) T1op is the top temperature of the stratified portion of surge piping, (2) Tbot is the bottom temperature of the stratified portion of piping , (3)

Tnoz is the temperature in the region of swirl inside the RCS hot leg nozzle and surge piping , and (4) THL is the temperature in RCS hot leg piping.

The thermal transients are to be considered as concurrent with the scaled anchor displacements, as described in Section 4.2. In addition, OBE occurs simultaneously with normal operation , up to the total number of OBE events.

3.5 Analysis Parameters and Approaches The following are analysis parameters and approaches used in this report:

1. The ambient containment temperature is set to be 120°F. This is conservatively selected as a lower bound value of the outer pipe wall temperature during stratified conditions.

2. Spray bypass flow rate is selected to be 2 gpm (Section 2.2). This is a typical value consistent with other Westinghouse-designed plants.

3. Since the piping system is insulated, the outside surfaces are conservatively modeled to be adiabatic, such that the component is treated as perfectly insulated.

4. Seismic OBE loads are input in the analysis as applied at normal operating conditions . This is consistent with the design basis of other Westinghouse-designed plants.

5. The temperature difference across the thermal sleeve annulus is modeled to be 50°F, which is judged to be conservative (higher than expected/actual) since this is a narrow gap filled with a good conductor (water). The impact of this temperature parameter is not sensitive to the temperature difference across the thermal sleeve gap.

6. For conservatism , 2 contingency stratification cycling transients are postulated to account for any future cycling that may potentially occur, despite the Modified Operating Procedures.

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7. For Plant Heatup, pressurizer temperature ramp rates were set at 100°F/hr. For Plant Cooldown , ramp rates were set at 200°F/hr (see Section 3.4). These are universally consistent with technical specification limits for Pressurized Water Reactors.

8. The stress-free reference temperature for thermal stress calculation is modeled to be 70°F, which is judged to be a reasonable temperature at installation and is used for thermal strain calculations.

9. The surge line welds are conservatively evaluated as shielded metal arc welds (SMAW) since the selection of SMAW (a flux welding process) is conservative as it has more stringent requirements for allowable flaw size determination.

10. The crack face pressure loading is modeled as a far-field loading, which is imposed in pc-CRACK [34] as a uniform membrane stress. This is the same approach used in APl-579-1[36].

11 . Full structural weld overlays were applied at the pressurizer surge nozzle-to-pipe welds [37] .

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Table 3-1. Deadweight and Seismic Piping Interface Loads Node Piping Load Service Mx (ft-lbs} Mv (ft-lbs}

Mz (ft-lbs}

MsRss Ref a*. -*

L!evel ~

r(ft-lbs)

Unit 1 Dea<llweight A,13,C,D 2,&f86 Y90 5,784 6,512 [3]

170/171 QB ~ 6 19 ' 995- 6,880 _ 18,]47 27,865 [3]

ss~ 25,357 22 *- 35012 C, 1!) 7,387 ~ ' 984 [3]

Deadweight A,B,C,D 1,784 516 2,550 3,155 [3) 175 OBE B 6,333 5,958 12,901 15,558 {3]

SSE C, D 7,401 6,370 14,556 17,528 [3)

Qeaciweieifii ~.B,C,D -6,037 ~.839 -1,772 6,555 13)

!I 190 OB'lt B ' 27,802 7,3'76 - - 1 9,306 J4,642 [3]

S£1i C~ D 32,6578 '" Y,773 22,666 40,522 [3)

Deadweight A , B,C , D 3,431 -1,926 2,598 4,715 [3]

198 OBE B 36,108 11,359 135,884 141,058 [3]

SSE C,D 42,294 13,858 159,932 166,009 [3]

Unit2 Dead!Yelght A,B,!; , D _ 18,776 -3,010 4,_084 19,449 [4) _,

17Cil/171 OEIE -

B- -

24,125 12,281 21,575 34,617 '[4] ..

SSE G, [i)  ; 28,308 15~019 27,636 42.~1.6 [41 Deadweight A, B, C, D 9,763 -2,577 3,921 10,832 [4]

173 OBE B 15,298 15,531 21,777 30,814 [4]

• SSE C,D 17,731 19,424 27,915 38,353 [4] '

Deadweight A,B ,C, D -588 -1, 107 1,5Z3 1,972 [4] l 1, - -

II 187 06!; B 22,874 10,>46 18,030 ~

31,010 [4] *

£§E C, [i) 29,522 12,647 22,541 39,238 [4]

Deadweight A , B, C,D 6,315 -2,523 801 6,847 [4]

201 OBE B 39,374 24,340 12,322 47,902 1[.11!]

SSE C, D 51,264 29,409 13,178 60,552 raJ Notes:

1) X, Y, and Z directions are shown in Figure 2-1 and Figure 2-2.

2) OBE and SSE are reversible .

3) Note that OBE includes both inertial and anchor movement loads, and SSE only includes inertial loads , since anchor movement loads were not included in the analysis for SSE in the stress analysis outputs. The magnitude of the seismic anchor movement (SAM) is very small(< 0.0625"), and as a result there is no significant impact on results. [3, 4]

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Table 3-2. Thermal Expansion Piping Interface Loads Node Thermal Serwice -*

Mx (ft-lbs) Mv (ft-lbs) Mz (ft-lbs)

MsRss Ref Level (ft-lbs)

Unit 1 Strat. (M ode 2) A(1) -228,35Y 16,200 -54,540 ~8,996 [3)

~

-~*

170/1¥1 Normal 0 11). C, D -164,480 125,934 -38,784 210,754 [3)

~

Max. Op. B -182,992 140,107 -43,149_ 234,474 [3)

Strat. (Mode 2) A(1) -207,381 12,665 -46,479 212,903 [3) 175 Normal Op. C, D 23,677 92,048 25,987 98,533 [3)

Max. Op . B 26,342 102,408 28,912 109,623 [3)

Nj:A.(1) I Strat. (M@ele ~) 11.67,32~ 20,196 -1,235 ll.6~J i47 [31 I ll.90 Norma l OJ:). A(ll, C:, D i 8,34ll 2110,275  %,325 215,097 [3l .I

- - -- ~

Max. Qlg. B 31,531 233,941 39,301 239,305 [3)

Strat. (Mode 2) N/A(1l 145,004 -14,179 34,181 149,651 [3) 198 Normal Op. A(1l, C, D -4,879 -190,133 -34,973 193,384 [3)

Max. Op. B -5,049 -196,757 -36,191 200,121 [3J Unit 2 II " "'

Strat. (Meae i ) ;A.(ll 84,J.98 -ll.34,631 lll7,248 197,388 [4J 170/171 Nonmal 011). C, D ' 1,923 151,271 26,147 il.53,526 [4j Max. Gp . B 2,139 168,296 29,090 170,805 [3)

Strat . (Mode 2) A(ll 79,711 -111,104 116,574 179,687 [4) 173 Norma l Op. C, D -9,121 120,641 25,982 123,744 [4)

Ma x. Op. B -10,148 134,219 28,906 137,671 ..

[3]

Strat. (Mode 2) A(1) 60,628 21,352 155,117 167,908 [4]

I 187 I Nmmal Clfl. C, D :l,fili. - -~Z§§.. ~.ML 87,746 [41 Max. QID:.. B -2,976 -95,419 -20,406 97,622 [3)

Strat. (Mode 2) A(1) -82, 537 -36,024 -230,382 247,358 [4) 201 Normal Op. C, D -40,744 -68,117 -81,743 113,938 [4)

Max. Op. B -42,163 -70,490 -84,591 117,908 [3)

Notes:

(1) The higher thermal expansion load (M sRss value), whether due to normal operating condition or thermal stratification, is used for Service Level A.

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Table 3-3. Number of Thermal Transient Cycles Cycles per Transient Notes 10-Years Plant Heatup 50 Plant Cooldown 50 Large Step Load Decrease 2 Reactor Trip 50 MOPHU320 5 (1)

MOPHU290 5 (1)

MOPHU270 13 ( 1)

MOPHU260 23 ( 1)

MOPHU250 38 ( 1)

MOPHU240 18 ( 1)

MOPCD320 10 (2)

MOPCD300 10 (2)

MOPCD290 10 (2)

MOPCD280 10 (2)

MOPCD270 20 (2)

MOPCD260 5 (2)

MOPCD250 18 (2)

MOPCD240 18 (2)

OBE (Event) 1 (3)

Notes:

(1) 2 contingency transients are postulated for each heatup.

(2) 2 contingency transients are postulated for each cooldown.

(3) OBE occurs at normal operation, 20 internal cycles per event.

(4) NAPS operates on an 18 month operating cycle.

(5) Stratification transients are defined in Ref 17 Report No. 1700553.402 R2 PAGE I 3-8 fJ Structural Integrity Associates. Inc info@structint.com m 1-877-45I-POWER e structint.com ~

Table 3-4. Thermal Transients Time p TPZR THL Hbot Ttop Tbot hsrg Tnoz hslv hett Transient (sec) (psig) ("F) ("F) (in) ("F) ("F) BTU/hr*ft * °F 2

("F) BTU/hr*ft2*°F BTU/hr*ft2 *°F Plant Heatup 0 0 70 70.0 0.000 70.0 70.0 119.0 70.0 119.0 61.3 20952 2235 652 612.0 3.082 652.0 612.0 261.3 652.0 252.9 84.3 SS(1l 2235 652 612.0 3.079 652.0 612.0 261.3 652.0 252.9 84.3 Plant Cooldown 0 2235 652 612.0 3.079 652.0 612.0 261.3 652.0 252.9 84.3 10476 0 70 70.0 0.000 70.0 70.0 119.0 70.0 119.0 61.3 ss 0 70 70.0 0.000 70.0 70.0 119.0 70.0 119.0 61.3 Large Step 0 2235 652 612.0 10.409 652.0 612.0 261.3 612.0 1621.3 11 7.4 Load Decrease 10 2042 652 608.1 10.423 652.0 608.1 261.3 608.1 1624.4 11 7.4 15 1977 652 602.6 10.440 652.0 602.6 261.3 602.6 1628.4 117.4 20 1955 652 596.8 10.455 652.0 596.8 261 .3 596.8 1629.4 11 7.4 25 1945 652 590.5 10.469 652.0 590.5 261 .3 590.5 1629.9 117. 4 30 1948 652 584.6 10.481 652.0 584.6 261 .3 584.6 1630.0 117.4 35 1951 652 579.3 10.489 652.0 579.3 261 .3 579.3 1638.6 117.5 40 1955 652 574.9 10.496 652.0 574.9 26 1.3 574.9 1645.9 117.5 45 1960 652 570.7 10.502 652.0 570.7 261 .3 570.7 1652.8 117.5 50 1963 652 567.6 10.506 652.0 567.6 261 .3 567.6 1657.7 117.6 55 1967 652 564.5 10.510 652.0 564.5 261.3 564.5 1661.4 117.6 60 1968 652 561.9 10.513 652.0 561.9 261 .3 561.9 1663.8 117.6 65 1969 652 560.0 10.515 652.0 560.0 261.3 560.0 1665.5 117.6 70 1969 652 558.4 10.517 652.0 558.4 261 .3 558.4 1667.0 117. 6 80 1969 652 555.6 10.520 652.0 555.6 261 .3 555.6 1669.4 117.6 90 1968 652 553.8 10.522 652.0 553.8 261.3 553.8 1671.0 117.6 100 1968 652 551 .6 10.525 652.0 551 .6 261.3 551.6 1672.9 117.6 120 1967 652 549.6 10.527 652.0 549.6 261.3 549.6 1674.6 117.6 140 1968 652 547.9 10.528 652.0 547.9 261.3 547.9 1676.2 117.7 Report No. 1700553.402 R2 PAGE 13-9 t1 Structural Integrity info@structint. com m 1-877-45!-POWER e structint.com @)

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Table 3-4. Thermal Transients Time p TPZR THL Hbot Ttop Tbot hsrg Tnoz hslv hett Transient (sec) (psig) (°F) (°F) (in) (°F) (* F) BTU/hr-ft2* °F (OF) BTU/hr*ft2*°F BTU/hr-ft2*°F 160 1970 652 546.3 10.530 652 .0 546 .3 261.3 546.3 1678.3 117.7 190 1977 652 544.6 10.531 652 .0 544.6 261.3 544 .6 1680.5 117.7 450 2116 652 546.5 10.530 652.0 546.5 261 .3 546 .5 1678.1 117.7 530 2138 652 548.3 10.528 652.0 548.3 261 .3 548.3 1675. 7 117.7 615 2154 652 550.0 10.526 652 .0 550.0 26 1.3 550.0 1674.3 117.6 695 2162 652 551.6 10.525 652.0 551 .6 261 .3 551.6 1672.9 117.6 785 2169 652 553.3 10.523 652.0 553.3 261 .3 553.3 1671 .5 117.6 875 2168 652 554.9 10.521 652 .0 554.9 261 .3 554 .9 1670 .0 117.6 1885 2235 652 554.9 10.521 652 .0 554.9 261 .3 554.9 1670 .0 117.6 ss 2235 652 560.9 10.514 652 .0 560.9 261 .3 560.9 1664.7 117.6 Reactor Trip 0 2235 652 612.0 10.409 652 .0 612.0 261 .3 612.0 1621.3 117.4 12 2235 652 612.0 10.409 652.0 612.0 261.3 612.0 1621.3 117.4 12 2055 652 612.0 10.409 652.0 612.0 261 .3 612.0 1621 .3 117.4 22 2038 652 59 7.6 10.453 652 .0 597.6 261 .3 59 7.6 1629.3 11 7.4 43 2046 652 581. 7 10.485 652 .0 581 .7 261.3 581. 7 1634.5 117.4 63 2052 652 572.0 10.500 652.0 572.0 261 .3 572.0 1650.7 117.5 74 2059 652 568.6 10.505 652 .0 568.6 261.3 568.6 1656.1 117.6 84 2066 652 566.0 10.508 652.0 566.0 261 .3 566.0 1660.0 117.6 94 2072 652 564.9 10.510 652.0 564.9 261 .3 564 .9 1661 .1 117.6 104 2077 652 563.3 10.511 652.0 563.3 261.3 563 .3 1662 .5 117.6 125 2084 652 560.8 10.514 652 .0 560.8 261 .3 560.8 1664.8 117.6 145 2091 652 559.7 10.516 652 .0 559. 7 261 .3 559 .7 1665 .8 117.6 176 2097 652 558.9 10.51 7 652.0 558.9 261.3 558 .9 1666.5 117.6 21 7 2103 652 557.4 10.518 652.0 557.4 261 .3 557.4 1667. 8 117.6 258 2110 652 556.9 10.519 652.0 556.9 261 .3 556.9 1668 .3 117.6 299 2118 652 556.5 10.519 652 .0 556.5 261 .3 556.5 1668.6 117.6 331 2124 652 556.3 10.520 652.0 556.3 261 .3 556 .3 1668.9 117.6 371 2131 652 555.9 10.520 652 .0 555.9 261 .3 555 .9 1669.2 117.6 412 2137 652 555 .6 10.520 652.0 555.6 261.3 555 .6 1669.4 11 7.6 Report No. 1700553.402 R2 PAGE 13-10 e

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Table 3-4. Thermal Transients Time p T PZR T HL Hbot Ttop Tbot hsrg Tnoz hslv hett Transient (sec) (psig) (*F) (*F) (in) (*F) (*F) BTU/hr*ft2* °F (*F) BTU/hr*ft2*°F BTU/hr*ft2 *°F 454 2144 652 555.3 10.521 652.0 555.3 261 .3 555.3 1669.7 117.6 494 21 50 652 555.0 10.521 652.0 555.0 261.3 555.0 1669.9 117.6 535 2155 652 554.7 10.521 652.0 554.7 261.3 554.7 1670.2 117.6 576 2161 652 554.3 10.522 652.0 554.3 261 .3 554.3 1670.6 117.6 617 2166 652 553.9 10.522 652.0 553.9 261 .3 553.9 1670.9 117.6 658 2172 652 553.5 10.523 652.0 553.5 261.3 553.5 1671.3 117.6 701 2178 652 553.1 10.523 652.0 553.1 261.3 553.1 1671.6 117.6 732 2184 652 552.8 10.523 652.0 552.8 261 .3 552.8 1671.9 117.6 762 219 1 652 552.9 10.523 652.0 552.9 261.3 552.9 1671 .8 117.6 783 2196 652 552.9 10.523 652.0 552.9 261 .3 552.9 1671.8 117.6 804 2204 652 553.0 10.523 652.0 553.0 261.3 553.0 1671.7 117.6 824 2212 652 553.0 10.523 652.0 553.0 261 .3 553.0 1671.7 117.6 845 2219 652 553.1 10.523 652.0 553.1 261 .3 553.1 1671.6 117.6 865 2226 652 553.1 10.523 652.0 553.1 261.3 553.1 1671.6 117.6 886 2232 652 553.2 10.523 652.0 553.2 261.3 553.2 1671.5 117.6 1504 2227 652 555.4 10.521 652.0 555.4 261.3 555.4 1669.7 117.6 1860 2220 652 555.4 10.521 652.0 555.4 261.3 555.4 1669.7 117.6 2351 2227 652 555.4 10.521 652.0 555.4 261 .3 555.4 1669.7 117.6 2515 2234 652 555.4 10.521 652.0 555.4 261 .3 555.4 1669.7 117.6 2576 2240 652 555.4 10.521 652.0 555.4 261 .3 555.4 1669.7 117.6 3600 2240 652 555.3 10.521 652.0 555.3 261 .3 555.3 1669.7 117.6 Norm. Op. for QBE ss 2235 652 612.0 10.232 652.0 612.0 261.3 612.0 1621.3 117.4 MOPHUCD 0 440 450 130.0 7.011 450.0 130.0 206.5 130.0 927.1 111.3 320 480 440 450 130.0 11.188 130.0 130.0 159.7 130.0 927. 1 11 1.3 1480 440 450 130.0 11 .188 130.0 130.0 159.7 130.0 927.1 111 .3 1960 440 450 130.0 7.011 450.0 130.0 206.5 130.0 927. 1 11 1.3 2960 440 450 130.0 7.011 450.0 130.0 206.5 130.0 927.1 111 .3 3440 440 450 130.0 0.000 450.0 450.0 260.3 450.0 260.3 85. 1 4440 440 450 130.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4920 440 450 130.0 7.011 450.0 130.0 206.5 130.0 927.1 11 1.3 Report No. 1700553.402 R2 PAGE I 3-11 SJ Structural Integrity i nfo@structint.com m 1-877-45!-POWER " structint.com ~

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Table 3-4. Thermal Transients Time p TPZR THL Hbot Ttop Tbot hsrg Tnoz hslv hett Transient (sec) (psig) (*F) (*F) (in) (*F) (*F) BTU/hr*ft2*°F (*F) BTU/hr*ft2*°F BTU/hr*ft2 *°F 5920 440 450 130.0 7.011 450.0 130.0 206.5 130.0 927. 1 11 1.3 MOPHUCD 0 440 450 150.0 6.980 450.0 150.0 206.5 150.0 1018.2 11 2.5 300 450 440 450 150.0 11 .188 150.0 150.0 172.1 150.0 1018.2 112.5 1450 440 450 150.0 11 .188 150.0 150.0 172.1 150.0 1018.2 11 2.5 1900 440 450 150.0 6.980 450.0 150.0 206.5 150.0 1018.2 112.5 2900 440 450 150.0 6.980 450.0 150.0 206.5 150.0 1018.2 112.5 3350 440 450 150.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4350 440 450 150.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4800 440 450 150.0 6.980 450.0 150.0 206.5 150.0 1018.2 11 2.5 5800 440 450 150.0 6.980 450.0 150.0 206.5 150.0 1018.2 112.5 MOPHUCD 0 440 450 160.0 6.962 450.0 160.0 206.5 160.0 1062.0 113.1 290 435 440 450 160.0 11.188 160.0 160.0 177.9 160.0 1062.0 113.1 1435 440 450 160.0 11 .188 160.0 160.0 177.9 160.0 1062.0 113.1 1870 440 450 160.0 6.962 450.0 160.0 206.5 160.0 1062.0 113.1 2870 440 450 160.0 6.962 450.0 160.0 206.5 160.0 1062.0 11 3.1 3305 440 450 160.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4305 440 450 160.0 0.000 450.0 450.0 260.3 450.0 260.3 85. 1 4740 440 450 160.0 6.962 450.0 160.0 206.5 160.0 1062.0 11 3.1 5740 440 450 160.0 6.962 450.0 160.0 206.5 160.0 1062.0 113.1 MOPHUCD 0 440 450 170.0 6.943 450.0 170.0 206.5 170.0 1107.9 11 3.6 280 420 440 450 170.0 11.188 170.0 170.0 184.0 170.0 1107.9 113.6 1420 440 450 170.0 11 .188 170.0 170.0 184.0 170.0 1107.9 113.6 1840 440 450 170.0 6.943 450.0 170.0 206.5 170.0 1107.9 113.6 2840 440 450 170.0 6.943 450.0 170.0 206.5 170.0 1107.9 113.6 3260 440 450 170.0 0.000 450.0 450.0 260.3 450.0 260.3 85. 1 4260 440 450 170.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4680 440 450 170.0 6.943 450.0 170.0 206.5 170.0 1107.9 113.6 5680 440 450 170.0 6.943 450.0 170.0 206.5 170.0 1107.9 113. 6 MOPHUCD 0 440 450 180.0 6.920 450.0 180.0 206.5 180.0 1145.8 114.0 270 405 440 450 180.0 11.188 180.0 180.0 189.0 180.0 1145.8 114.0 Report No. 1700553.402 R2 PAGE 13-12 e

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Table 3-4. Thermal Transients Time p TPZR THL H bat Ttop Tbat hsrg Tnoz hslv he11 Transient (sec) (psig) (°F) (OF) (in) (°F) (°F) BTU/hr*ft *°F 2

(°F) BTU/hr*ft *°F 2

BTU/hr*ft2 *°F 1405 440 450 180.0 11 .1 88 180.0 180.0 189.0 180.0 1145.8 11 4.0 1810 440 450 180.0 6.920 450.0 180. 0 206.5 180.0 1145.8 11 4.0 2810 440 450 180.0 6.920 450. 0 180.0 206.5 180.0 1145.8 11 4.0 3215 440 450 180.0 0.000 450.0 450.0 260.3 450.0 260.3 85. 1 4215 440 450 180.0 0.000 450.0 450.0 260. 3 450.0 260.3 85. 1 4620 440 450 180.0 6.920 450.0 180.0 206.5 180.0 1145.8 114.0 5620 440 450 180.0 6.920 450.0 180.0 206.5 180.0 11 45.8 114.0 MOPHUCD 0 440 450 190.0 6.896 450.0 190.0 206.5 190.0 1185.7 114.3 260 390 440 450 190.0 11 .188 190.0 190.0 194.3 190.0 11 85.7 11 4.3 1390 440 450 190.0 11 .188 190.0 190.0 194.3 190.0 11 85.7 114.3 1780 440 450 190.0 6.896 450.0 190.0 206.5 190.0 1185.7 114.3 2780 440 450 190. 0 6.896 450.0 190.0 206.5 190.0 1185.7 114.3 3170 440 450 190.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4170 440 450 190.0 0.000 450.0 450.0 260.3 450.0 260. 3 85.1 4560 440 450 190.0 6.896 450.0 190.0 206.5 190.0 1185.7 114.3 5560 440 450 190.0 6.896 450.0 190.0 206.5 190.0 1185.7 114.3 MOPHUCD 0 440 450 200. 0 6.871 450.0 200.0 206.5 200.0 1224.6 114.7 250 375 440 450 200.0 11.188 200.0 200.0 199.4 200.0 1224.6 11 4.7 1375 440 450 200.0 11.188 200.0 200.0 199.4 200.0 1224.6 114.7 1750 440 450 200.0 6.871 450.0 200.0 206.5 200.0 1224.6 114.7 2750 440 450 200.0 6.871 450.0 200.0 206.5 200. 0 1224. 6 114.7 3125 440 450 200.0 0.000 450.0 450.0 260.3 450.0 260.3 85. 1 41 25 440 450 200.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4500 440 450 200.0 6.871 450.0 200.0 206.5 200.0 1224.6 114.7 5500 440 450 200.0 6.871 450.0 200.0 206.5 200.0 1224.6 114.7 MOPHUCD 0 440 450 210.0 6.845 450.0 210.0 206.5 210.0 1263.6 115.0 240 360 440 450 210.0 11 .188 210.0 210.0 204.4 210.0 1263.6 115.0 1360 440 450 210.0 11.188 210.0 210.0 204.4 210.0 1263.6 115.0 1720 440 450 210.0 6.845 450.0 210.0 206.5 210.0 1263.6 115.0 2720 440 450 210.0 6.845 450.0 210.0 206.5 210.0 1263.6 11 5.0 Report No. 1700553.402 R2 PAGE I 3-1 3 I)

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Table 3-4. Thermal Transients Time p TPZR THL Hbot Ttop Tbot hsrg Tnoz hs1v h ett Transient (sec) (psig) (°F) (°F) (in) (°F) (°F) BTU/hr*ft2* °F (°F) 2 BTU/hr*ft *°F BTU/hr*ft2 *°F 3080 440 450 210.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4080 440 450 210.0 0.000 450.0 450.0 260.3 450.0 260.3 85.1 4440 440 450 210.0 6.845 450.0 210.0 206.5 210.0 1263.6 11 5.0 5440 440 450 210.0 6.845 450.0 210.0 206.5 210.0 1263.6 115.0 Table Notes:

1. SS = Steady-State Table 3-5. Equivalent Conductivity of Water Gap (ke) k, ke Btu/hr- Btu/hr-T1 T2 T, °F v, ft2/sec ft-°F Pr ~. 1/-F Gro Gro Pr kefk ft-°F 125 75 100 7.42E-06 0.364 4.540 2.23E-04 7.38E+03 3.35E+04 2.26 0.82 225 175 200 3.37E-06 0.391 1.877 3.92E-04 6.28E+04 1.18E+05 3.25 1.27 325 275 300 2.14E-06 0.397 1.146 5.61 E-04 2.22E+05 2.54E+05 4.06 1.61 425 375 400 1.64E-06 0.382 0.897 7.53E-04 5.11E+05 4.59E+05 4.82 1.84 525 475 500 1.42E-06 0.349 0.853 1.12E-03 1.02E+06 8.69E+05 5.80 2.02 625 575 600 1.34E-06 0.298 1.042 1.99E-03 2.01E+06 2.09E+06 7.35 2. 19 Avg = 1.63 Table 3-6. Total Added Thermal Resistance, Thermal Sleeve and Water Gap kTS, tfkTS, tfkTs+8/ke, Btu/hr-ft-°F hr-ft2 -°F/Btu hr-ft2 -°F/Btu 10 0.00150 0.0079 Report No. 1700553.402 R2 PAGE 13-1 4 I)

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4.0 STRESS ANALYSIS A three-dimensional (3-D) finite element model (FEM) for the surge line piping [6] is developed using the ANSYS Finite element analysis (FEA) software package [20] to perform the stress analysis [8] . The following is the order in which the FEA was conducted:

• Develop a finite element model (FEM) of the surge line piping, as shown in Figure 4-2.

• Perform thermal transient analyses to obtain the temperature history for the applicable normal plant transients.

• Perform thermal transient analyses to obtain the steady-state temperature for the applicable stratification events.

• Perform stress analyses using temperature results from thermal transient definitions [5]. These stress analyses include the appropriate internal pressure and thermal anchor movements (TAMs) at the corresponding temperature time steps.

• Perform stress analysis for the Deadweight case.

• Review stress results and select stress extraction paths at locations, for locations in the base metal and the weld and store them in computer files.

The following Sections 4.1 through 4.7 provide a high level technical overview of the finite element modeling, displacement loadings, materials, and analyses performed for assessment of the various locations.

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4.1 Finite Element Model The pressurizer surge line hot leg nozzle is located in a horizontal section of piping for Unit 2 [4].

The equivalent location in Unit 1 is shown to be located in a vertically oriented section of piping

[3]. Due to the differences between the two units, a comparison of piping stress equations (Equations 10 & 12) is made to determine which Unit to model. The piping stress equations values in the surge line of Unit 2 [4] are shown in Figure 4-1 to bound those of the Unit 1 surge line [3], as follows:

a. The Eq 10 & Eq 12 piping stresses for the top five nodes are listed and they demonstrate that the Unit 2 piping system is bounding.

Table 4-1. Piping Stress Values - Unit 1 versus Unit 2 UNIT 1 [3] UNIT 2 [4]

Eq . 10 Eq . 10 Eq . 12 Eq . 12 Allowable Allowable Node Stress Allow Node Stress Allow Stress Stress (ksi (ksi) (ksi) (ksi) 170 59.5 3Sm 25.2 49.5 170 ru 3Sm ~ 49.5 175 36 .1 3Sm N/A 49.5 173 ~ 3Sm N/A 49.5 180 47.6 3Sm N/A 49 .5 N/A <1l N/A N/A N/A N/A 190 47.3 3Sm N/A 49 .5 187 ru

~

3Sm ~ 54.1 198 82 .2 <2 ) 3Sm 52.6 52.6 201 30 .8 <2) 3Sm N/A 49.5 250 53.1 3Sm 31 .5 49.5 250 ru 3Sm ~ 49.5 Table Notes:

1. No equivalent node 180 (lug) on Unit 2

2. Node 198 is located at the top of the last elbow into the PZR, and Node 201 is located on the bottom of the last elbow. Stresses at Node 198 would be expected to be greater as it is closer to the piping anchor point at the Pressurizer surge nozzle.

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b. Thus, it is reasonable to use the Unit 2 piping geometry to bound Unit 1 based on the following:

• A comparison of piping stress equations values between Unit 1 and Unit 2 shows that Unit 2 has more limiting values.

• The transient loading that contributes to the formation and growth of flaws due to cyclic fatigue is similar between Unit 1 and Unit 2, as is the general pipe size and material properties of the surge line piping. Due to geometric layout differences, the Unit 2 surge line piping produces higher stresses, increasing the probabil ity for the formation of flaws and fatigue crack growth that is bounding for both Units.

Therefore, the Unit 2 surge piping configuration shown in Figure 4-1 is chosen to construct the finite element model for analysis.

The dimensions of the pressurizer surge line [4] are shown in Figure 4-1.

As shown in Figure 4-2 , the finite element model includes:

• The surge line piping

• The hot leg surge nozzle

• The support hangers

• A portion of the hot leg piping A three-dimensional (3-D) model is constructed using 8-node structural solid ANSYS SOLID45 elements. The thermal equivalent element type for the thermal transient analyses is SOLID70.

The spring hangers are simulated using the ANSYS COMBIN14 element type. The spring hanger SH-01 located at the upper elevation (see Figure 4-1 and Figure 4-2) is modeled using a preload of 4,659 lbs and spring stiffness of 1,080 lbs/inch [4]. The hanger SH-02 located in the lower horizontal run is modeled with preload of 4,304 lbs and spring stiffness of 900 lbs/inch [4] .

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4.2 Thermal Anchor Movements For the development of the finite element model, Unit 2 is bounding for Unit 1, based on comparison of piping stress values for the two units (See Section 4.1 for additional discussion) .

As such, Table 4-2 tabulates the thermal anchor movements Load Case No. 1 at Node 7 (hot leg to surge piping intersection point) and Node 250 (pressurizer surge nozzle safe end to piping) at the Design temperature 680°F for Unit 2 [5] . Therefore, anchor movements for each point should be separately scaled to the indicated temperature with a thermal scale factor (TSF) as follows .

For Point 7 (hot leg to surge piping intersection point): TSF = (THOT-70)/(680-70)

For Point 250 (pressurizer surge nozzle safe end to piping): TSF = (TPZR-70)/(680-70)

Where THOT is the hot leg temperature, and TPZR is the pressurizer temperature.

4.3 Material Properties The surge line and hot leg FEM are constructed of SA-376 TP316 stainless steel [4] . The hot leg surge nozzle FEM is constructed of SA-182 F316 stain less steel [4]. Per ASME Section II ,

Part D, the material properties are the same for both (16Cr-12Ni-2Mo) , and therefore the temperature-dependent linear-elastic material properties shown in Table 4-3 are included in the finite element model input file . The ASME Code specifications for the materials modeled in the analysis (SA- material) are identical to those in the ASTM specification (A- material) for the purposes of the analysis.

4.4 Thermal/Mechanical Stress Analysis Stress analyses are performed for thermal transients, mechanical piping loads, and internal pressure. For thermal transients, thermal analyses are performed to determine the transient temperature distribution fo r each transient in Section 3.0. The temperature distribution is then used as input to perform a stress analysis for each transient. Each transient stress analysis is performed concurrently with the specified pressure and thermal anchor movements (TAM 's) time history for that transient, appl ied concurrently included.

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Mapped through-wall stresses for the individual transients and deadweight are extracted along several paths at each weld location identified and are saved for use in crack growth evaluations.

Stratification is possible in either the bottom horizontal section of piping or the top horizontal section of piping. Therefore, two separate sets of thermal/mechanical evaluations are performed, with the only difference being which section of horizontal piping (lower or upper) is stratified, as shown in Figure 4-4 through Figure 4-6.

Table 3-4 tabulates the temperature and pressure time histories for the thirteen bounding thermal transients. Two separate sets of analyses are performed, in which either the lower or upper horizontal section of piping is evaluated as stratified. Table 4-4 lists the transients evaluated along with the ANSYS event naming convention used.

4.5 Deadweight Analysis Elastic stresses due to deadweight are evaluated separately from the thermal/mechanical analyses. A static acceleration of 1 G is applied in the vertical direction to simulate the deadweight load. The density of the piping is increased from the ASME Code value of 0.290 lbs/in 3 (Table 4-3) to 0.383 lbs/in 3 , which is based on the total weight of the piping including water (255.85 lbs/ft) [3, 4].

The stresses due to deadweight loading will be combined with the individual transient stresses in the crack growth calculation (Section 6.0), which will define the load combinations evaluated.

4.6 Mechanical Boundary Conditions Symmetric boundary conditions are applied to the symmetry face of the hot leg pipe as well as to one free end of the hot leg. The other end of the hot leg pipe is coupled in the axial direction to simulate the unmodeled section of piping. Symmetry boundaries are also applied to the top of the surge line at the pressurizer end.

The applied load and boundary conditions for the deadweight case are shown in Figure 4-3 and the boundary conditions for the thermal/mechanical analysis are shown in Figure 4-7.

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4.7 Stress Analysis Results Figure 4-4 and Figure 4-5 show representative plots of the thermal loads applied for the MOP320 transient, for example, for the lower stratified line and for the upper stratified line, respectively. The resulting thermal gradient contour plots are provided in Figure 4-6.

Two separate sets of thermal transient analyses are performed [8], in which either the lower or upper horizontal section of piping is evaluated as stratified. Example stress intensity contour plots are shown in Figure 4-8 and Figure 4-9 for the lower stratified line, and for the upper stratified line, respectively.

Mapped through-wall stresses were extracted along several stress paths in Figure 4-10

[8]. The stress paths are chosen based on close examination of both axial and hoop stress history at each weld location, throughout the various transient histories. The transient stress results are extracted and saved to support the ASME Section XI , Appendix L crack growth analyses in Section 6.0 .

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Table 4-2. Thermal Anchor Movements for Design Conditions DEFLECTIONS (IN) ROTATIONS (RAD)

Point No.

DX DY DZ RX RY RZ 7 Run -1.205 0.052 0.112 0.0002 -0.0001 0.0015 250 Taper -0.092 -0.024 -0.043 0.0001 0.0001 0.0000 Table 4-3. Material Properties used in the Finite Element Model (TP316 Stainless Steel)(3H4 l Mean Thermal Temperature Young's Modulus Thermal Conductivity Specific Heat(2l Expansion (OF) (x10 6 psi) (Btu/hr-ft-°F) (Btu/lb-°F)

(x1 Q-6 in/in/°F) 70 28 .3 8.5 8.2 0.118 100 28 .1(1) 8.6 8.3 0.118 150 27 .8(1) 8.8 8.6 0.121 200 27.5 8.9 8.8 0.121 250 27.3(1) 9.1 9.1 0.124 300 27 .0 9.2 9.3 0.124 350 26.7(1) 9.4 9.5 0.1 25 400 26.4 9.5 9.8 0.1 26 450 26 .2(1) 9.6 10.0 0.127 500 25.9 9.7 10.2 0.127 550 25 .6(1) 9.8 10.5 0.129 600 25 .3 9.8 10.7 0.129 650 25 .1(1) 9.9 10.9 0.130 700 24.8 10.0 11.2 0.131 3

Density (p) = 0.290 lb/in [25], modeled as temperature independent.

Poisson's Ratio (u) = 0.31 [25], modeled as temperature independent.

Notes:

1. Interpolated.

2. Specific Heat values are derived from the equation shown in General Note (a) of Table TCD [25], Specific Heat= TC/ (TD x density).

3. Material Composition-16Cr-12Ni-2Mo [25].

4. Values per Reference [25].

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Table 4-4. Thermal Transients Evaluated Transient Event ANSYS Event Name Plant Heatup HEATUP Plant Cooldown T2DW Large Step Load Decrease LSLD Reactor Trip TRIP Normal Operation NOP MOPHUCD 320 MOP320 MOPHUCD 300 MOP300 MOPHUCD 290 MOP290 MOPHUCD 280 MOP280 MOPHUCD 270 MOP270 MOPHUCD 260 MOP260 MOPHUCD 250 MOP250 MOPHUCD 240 MOP240 Report No. 1700553.402 R2 PAGE 14-8 13 Structural Integrity Associates, Inc. info@structint. com m 1-877-4SI-POWER e structint.com@)

a f!"~6-r!S ~ F"T) ...tw ff?oM To A'/. A'( Ar.

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,. P. <Et.G V t"Aq-<: I~

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Figure 4-1. Unit 2 Surge Line Dimensions (Global Coordinates)

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1 ANSYS R1'4.~

Spring

....- Hanger SH-Ol Pressurizer /

Surge Line Piping Hot Leg Spring Pipe Hot Leg Hanaer I 1\ou le SH-~2 "'--

Figure 4-2. Components Included in the 3-D Finite Element Model (Bounding FEM)

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l=

Figure 4-3. Applied Boundary Conditions (Deadweight}

(Bounding FEM)

(Symmetric boundary conditions are applied to the symmetry face of the hot leg pipe as well as to one free end of the hot leg (Blue constraints). The other end of the hot leg pipe is coupled in the axial direction to simulate the unmodeled section ofpiping (Green constraints). Symmetry boundaries are also applied to the top of the surge line at the pressurizer end)

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130 201.111 212 222 343 333 414 444 165.556 36 . 667

• 307 . 778

• 378 . 889

• 450 Figure 4-4. Applied Thermal Boundary Conditions for Thermal/Mechanical Transient Analysis (Lower Line Stratified)

Transient MOP320 shown, loads applied at time = 9,920 seconds.

(Units for HTC is BTU/sec-in 2 -°F, TBULK is °F)

(Bounding FEM)

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036 .001195 Heat Transfer Coefficient 130 165.556 201.111 236 667 272.222 307.778 343 . 333 414 . 444 7 . 0 450 Bulk Temperature Figure 4-5. Applied Thermal Boundary Conditions for Thermal/Mechanical Transient Analyses (Upper Line Stratified)

Transient MOP320 shown, loads applied at time = 9,920 seconds.

0 (Units for HTC is BTU/sec-in2- F, TBULK is °F) (Bounding FEM)

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NJJAL 9:V.JI'I~

ST£P-H7

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~ ... 129 . 31 SH:{ - 452.23 129.31 47 416 .35 a) Lower Line Stratified t.O:W. s::I.UI'I~

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D<< ..2 . 75585 St-N .. 129.996 SH:{ -450 . 003 129.996 "65.552 201.109 236.665 272 . 221 307.778 343 . 334 378.89 414 . 447 450.003 b) Upper Line Stratified Figure 4-6. Temperature Contour During Transient MOP320 (Time= 9,920 seconds) 0 (Units for temperature is F)

(Bounding FEM)

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l EUMNI'S TYPE !01 II

=

PRFS--N:llM 440 Figure 4-7. Applied Mechanical Boundary Conditions for Thermal/Mechanical Stress Analyses (Units for pressure is psi. Applied displacements and rotations in inches and radians.)

(Bounding FEM)

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~ L s::uJTICN SU,_,

S'If".P-119 TI!-£-9920 SINT (AVG) i,sy-,-o IM< -4.7063 S?-N --696.253

rt{ *52671.9 696.253 *7 a) Full Model tlllAL s::uJ'I'ICN S'I'£2"-U9 SllB - 1 Tn-£"ro9920 SIN!' {AVG)

RSYS-0 D<< --4 .70!53 SM'1 *!596 .253 SM<

• 5267 l. 9 696.253 6471.3 12246.4 18 21.5 23796.6 29571.6 35346.7 1 .8 46896.9 52671.9 b) Bottom View of Hot Leg/Nozzle Figure 4-8, Total Stress Intensity Due to MOP320 Transient (Lower Line Stratified)

Transient MOP320 shown, loads applied at time= 9,920 seconds, OD= 19.355 inch.

(Units for stress is psi. Results include thermal, pressure, and TAM's) (Bounding FEM)

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NnU. 9:l.U!ICN S'I3?-ll6 S03 -1 TD£F9'920 SINT (Av;;)

RSYS-0 a<< -2 .75555 9-N -367 . 943

~ - 52270 . 5

~ :B-57914.3 367 . 943 7668.8 23435 . 7 g 02. 7 34969 . 6 073 . 6 46503 . 5 70 . 5 5

a) Full Model N::I:lN.. s::IJ.ll'IQi S'!t:f-,116 SU, -1 TD-fl-9920 SlNI' (A~)

PSY,;-0

~ - 2 .75585 g.t,1 -357 . 943 9-!X-5mo .5 367 . 943 6134.89 11901.8 17668 . 8 23435 . 7 29202 . 7 34969 . 6 40736 . 6 46503.5 52270 . 5 b) Top View of Hot Leg/Nozzle Figure 4-9. Total Stress Intensity Due to MOP320 Transient (Upper Line Stratified)

Transient MOP320 shown, loads applied at time =9,920 seconds.

(Units for stress is psi. Results include thermal, pressure, and TAM's)

(Bounding FEM)

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Station V Paths 21-22, 24,25 Station IV Station I Station II

• Stress paths 1-5 correspond to Weld location 170/171 (Station I).

• Stress paths 6-10 correspond to location 173 (Station 11).

• Stress paths 11-14 correspond to location 187 (Station Ill).

• Stress paths 15-20 correspond to location 201 (Station IV).

• Stress paths 21 , 22, 24, 25 correspond to weld location 250 (Station V).

Figure 4-10. Stress Path Locations (Bounding FEM)

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5.0 ALLOWABLE FLAW SIZE EVALUATION Allowable flaw sizes are calculated for circumferential and axial part through-wall flaws in the following sections.

5.1 Allowable Flaw Size Determination One important aspect of a flaw tolerance evaluation is the determination of the allowable flaw sizes. These are the flaw sizes that cannot be exceeded when the Code structural factors are applied under the applied loads. As required by Appendix L of the ASME Code Section XI [2],

the flaw evaluation procedures of IWB-3640 are used when applicable in the determination of the allowable flaw sizes.

For the weld metal and adjacent base metal, guidance for calculation of allowable flaw sizes at the hot leg surge nozzle-to-pipe weld is provided by ASME Code Section XI , Appendix L (L-3000) [2]. The allowable flaw size (circumferential depth/length) for the weld will be determined based on the rules in ASME Code Section XI, Subsections IWB-3640 and Appendix C [2], which contains the screening criteria procedure to determine the applicable failure mode and evaluation for allowable flaw sizes with appropriate structural factors.

A crack growth evaluation presented in Section 6.0 will determine the allowable operating periods based on postulated initial flaw sizes and the allowable flaw sizes calculated herein .

The ASME Boiler & Pressure Vessel Code ,Section XI, 2013 Edition [2] is used for the determination of the allowable flaw sizes when applicable .

5.2 Interface Loads The piping interface loads are listed in Section 3.0.

5.3 Load Combinations The load combinations represent the normal conditions (Service Level A) , upset conditions (Service Level B), and emergency/faulted conditions (Service Levels C and D).

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Service Level A This load combination includes the internal pressure under normal operating conditions ,

deadweight, and the highest thermal expansion load at normal operating temperature or due to thermal stratification. All load cases are absolute summed.

Service Level B This load combination includes internal pressure under maximum operating conditions ,

deadweight, OBE, and the thermal expansion load at the maximum operating temperature. All load cases are absolute summed.

Service Levels C and D These load combinations include internal pressure under emergency/faulted conditions, deadweight, SSE, and the thermal expansion load at normal operating temperature. No Service Level C or D transients are included . All load cases are absolute summed.

Section XI, Appendix L, Paragraph L-2210 [2] states : The loadings in the Design Specification, plant specific loading cycles consistent with the plant design and operating practices, or actual plant operating data, shall be used , as appropriate. Dominion and SIA agreed that the appropriate transients are those from Section 3.4 (Table 3-4 and Table 4-4).

5.4 Material Properties for Allowable Flaw Size Determination Material strength data for the base metal (both Units 1 and 2, per Reference 7) was obtained from the Certified Material Test Reports (CMTR). The minimum material strength data is used to calculate the allowable axial flaw size for both Units 1 and 2 for all locations, and to calculate the allowable circumferential flaw at Node 198 (Unit 1). All other circumferential flaw locations use ASME Code minimum properties [25] . The yield stress, cry, and ultimate stress , cru, at the operating temperatures are shown in Table 5-1. The material properties are applicable to both the base metal and weld metal.

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5.5 Welding Process As stated earlier, the use of a flux process (SMAW or SAW) is conservatively selected for the weldments. Per ASME Code Section XI, Appendix C (C-6330) [2] , Z-factors (load factors) are calculated to account for the use of limit load as a failure criterion for the lower toughness welds fabricated using SMAW or SAW welding processes (flux processes) in the determination of the allowable flaw size for austenitic weldments:

5.6 Z-Factor Per ASME Code Section XI, Appendix C (C-6330) [2], Z-factors (load factors) are calculated to account for the use of limit load as a failure criterion for the lower toughness welds fabricated using SMAW or SAW welding processes (flux processes) in the determination of the allowable flaw size for austenitic weldments :

Z = 1.30 [1 + 0.010 (NPS -4)] for NPS > 4 (5-1) where ,

NPS = Nominal pipe size (in)

The calculated Z-factor for the 14-inch NPS surge line piping is 1.430.

5.7 Allowable Circumferential Part Through-Wall Flaw For flux welds such as SMAW/SAW, the elastic-plastic fracture mechanics (EPFM) methodology in the ASME Code ,Section XI, Appendix C [2] , should be used (per the screening criteria in Figure C-6200-1 ). The technical approach consists of determining the allowable flaw size (circumferential extent and through-wall depth) in the pipe that will not result in fracture by crack extension , including the necessary safety factors (SFs).

For circumferential flaws , the stress ratio for combined loading is calculated as:

Stress Ratio = _I_ (CYm + CYb + CYe ) (5-2)

CY! SF;,

and the stress ratio for membrane stress only is calculated as:

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. ZSFu Stress Ratzo = m m (5-3) uf where, O' m = Primary membrane stress (ksi) due to pressure for all service levels O'b = Primary bending stresses (ksi) from:

• deadweight only for Service Level A,

• deadweight + OBE (Inertial + SAM) for Service Level B, and

• deadweight + SSE (Inertial) for Service Levels C/D

~ = Secondary bending stress (ksi) from thermal expansion stresses for:

• Service Level A (thermal transients or thermal stratification events) and

• Service Levels B, C, and D (thermal transients only).

Cit = Flow stress (ksi), which is equal to the average of the yield strength, Oy, and the ultimate tensile strength, Ou.

S~ = Structural factor for bending stress, depending on service level

[2, Table C-2621].

SF,,, = Structural factor for membrane stress depending on service level

[2, Table C-2621] .

Z = Z-factor for flux welds Based on the calculated stress ratios in Equation 5-2, the allowable flaw depth-to-thickness ratio due to combined loading for Service Levels A, B, C, and D are obtained from Table C-5310-1 through Table C-5310-4 of ASME Code Section XI, Appendix C [2]. Stress ratios in Equation 5-3 are used in Table C-5310-5 [2] to determine the allowable flaw depth-to-thickness ratio for membrane loading only. The allowable flaw sizes for combined loading and for membrane only loading are compared , and the smaller value is reported as the allowable flaw size for the weld. ASME Code [25] material strength properties are used for all locations except Node 198 in Unit 1, which uses component specific CMTR material strength (Section 5.4).

The calculated allowable circumferential flaw depths for Units 1 and 2 as a function of crack length for all Service Levels are shown in Table 5-2 and Table 5-3 , respectively.

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5.8 Allowable Axial Part Through-Wall Flaw The allowable axial flaw size is also determined in accordance with ASME Code,Section XI, Appendix C [2]. The allowable flaw depth is determined using Table C-5410-1 [2].

The stress ratio is calculated as follows:

Z *SF a-StressRatio = "' h a- I where, a-h = PRm/t, hoop stress (ksi)

P = Internal pressure (ksi)

Rm = Mean radius (in) t = Wall thickness (in)

C5f = Average of yield and tensile stresses, flow stress (ksi)

Z = Z-factor for flux welds SF,n = Structural factor for membrane stress depending on Service Level

[2, Table C-2621].

The allowable flaw length for stability of a through-wall flaw is calculated as:

z ]1/2 lallow = 1.58(Rmt) 1 / 2 [ (;;h) -1 where, a-h = PRm/t, hoop stress (ksi)

P = Internal pressure (ksi)

Rm = Mean radius (in) t = Wall thickness (in)

C5f = Average of yield and tensile stresses , flow stress (ksi)

Z = Z-factor for flux welds The end-of-evaluation-period flaw length shall be limited to less than the allowable flaw length.

A review of the CMTR material strength properties for all surge pipe sections at U1 & U2 was performed and the minimum yield stress and ultimate stress values, that bound all CMTR data

[7], are used in calculating the allowable axial flaw size for all locations on the surge line.

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The calculated allowable axial flaw depths for both Units 1 and 2 as a function of crack length for all Service Levels are shown in Table 5-4.

Table 5-1: Geometry, Operating Conditions, and Material Properties Property Value Reference 14-inch NPS Pipe Specification [3,4]

Schedule 160 Thickness at Hot Leg Surge Nozzle-to-Pipe Weld 1.406 in [3,4]

Outer Diameter at Hot Leg Surge Nozzle-to-Pipe Weld 14.00 in [3,4]

Inner Diameter at Hot Leg Surge Nozzle-to-Pipe Weld 11 .188 in [3,4]

Z-factor 1.430 Section 5.6 Normal Operating Pressure 2235 psig [40]

Normal Operating Temperature 612 °F [40]

at Hot Leg End of Surge Line - 100% power Normal Operating Temperature 652 °F [40]

at Pressurizer End of Surge Line - 100% power Temperature in Surge Line during Thermal Stratification 450°F Section 2.3 Event Ou @ Circumferential Flaw Material Property - (Node 198)(2) 82.725 ksi CMTR [7]

oy @ Circumferential Flaw Material Property - (Node 198)(2) 39 .325 ksi CMTR [7]

Ou @ Axial Flaw Material Property- Unit 1 & 2 82.0 ksi CMTR [7]

Oy @ Axial Flaw Material Property- Unit 1 & 2 33 .5 ksi CMTR [7]

Notes for Table 5-1 :

(1) Per Section 5.4, the weld metal properties are the same as the metal in the surge line piping . Per ASME Code Section IX, QW-153 [42], the weld metal must be stronger than the base metal of the adjacent components.

(2) All other locations utilize ASME Code properties [25] for allowable circumferential flaw determination.

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Table 5-2. Allowable Part Through-Wall Circumferential Flaw Sizes for Unit 1 Ratio of Flaw Length to Pipe Circumference, lt I n D Service 0 0.1 0.2 0.3 0.4 0.5 0.6 0.75 Node Level Flaw Length, 7.; {degree) 0 36 72 108 144 180 216 270 A 0.75 0.75 0.75 0.59 0.48 0.42 0 .39 0.38 B 0.75 0.75 0.61 0.43 0.35 0.31 0.29 0.28 170/171 C 0.75 0.75 0.74 0.60 0.48 0.42 0.39 0.37 D 0.75 0.75 0.74 0.53 0.42 0.36 0.34 0.32 A 0.75 0.75 0.75 0.68 0.54 0.48 0.44 0.43 B 0.75 0 .75 0.75 0.75 0.69 0.60 0.55 0.53 175 C 0.75 0.75 0.75 0.75 0.73 0.66 0.61 0.56 D 0.75 0.75 0.75 0 .75 0.72 0.64 0 .59 0.54 A 0.75 0.75 0.75 0.65 0.52 0.46 0.42 0.41 B 0.75 0.75 0.53 0.37 0.30 0.27 0.25 0.25 190 C 0.75 0.75 0.74 0.57 0.46 0.39 0.37 0.35 D 0.75 0.75 0.70 0.49 0.39 0.34 0.32 0.30 A 0.75 0.75 0.75 0.75 0.70 0.61 0.56 0.53 B 0.75 0.45 0.23 0.16 0.14 0.13 0.13 0.13 198 C 0.75 0.75 0.49 0.34 0.27 0.24 0.22 0.22 D 0.75 0.64 0.39 0.28 0.22 0.20 0.18 0.18 Note: The limiting (smaller) allowable flaw size from either Unit 1 or Unit 2 will be used in the evaluation.

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Table 5-3. Allowable Part Through-Wall Circumferential Flaw Sizes for Unit 2 Ratio of Flaw Length to Pipe Circumference, ].; I nD Service 0 0.1 0.2 0.3 0.4 0.5 0.6 0.75 Node Level Flaw Length, ].; (degree) 0 36 72 108 144 180 216 270 A 0.75 0.75 0.75 0.62 0.50 0.44 0.41 0.40 B 0.75 0.75 0.70 0.50 0.40 0.35 0.33 0.32 170/171 C 0.75 0.75 0.75 0.66 0.53 0.46 0.42 0.40 D 0.75 0.75 0.75 0.61 0.49 0.42 0.39 0.37 A 0.75 0.75 0.75 0.73 0.58 0.51 0.47 0.46 B 0.75 0.75 0.75 0.66 0.53 0.46 0.43 0.42 173 C 0.75 0.75 0.75 0.73 0.63 0.55 0.50 0.47 D 0.75 0.75 0.75 0.71 0.60 0.52 0.48 0.45 A 0.75 0.75 0.75 0.74 0.65 0.57 0.52 0.50 B 0.75 0.75 0.75 0.75 0.65 0.57 0.53 0.51 187 C 0.75 0.75 0.75 0.75 0.71 0.63 0.58 0.54 D 0.75 0.75 0.75 0.75 0.70 0.61 0.56 0.52 A 0.75 0.75 0.72 0.56 0.45 0.40 0.37 0.36 B 0.75 0 .75 0.75 0.64 0.52 0.45 0.42 0.41 201 C 0.75 0.75 0.75 0.71 0.59 0.51 0.47 0.45 D 0.75 0.75 0.75 0.69 0.56 0.49 0.45 0.42 Note: The limiting (smal ler) allowable flaw size from either Unit 1 or Unit 2 will be used in the evaluation.

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Table 5-4. Allowable Part Through-Wall Axial Flaw Sizes for Unit 1 and Unit 2 Nondimensional Flaw Length, l, I v(Rmt) Al lowable Thru-wall 0 0.5 1.0 2.0 3.0 4.0 5.0 6.0 Flaw Length, Node Service Flaw Length, ].; (inch) ia.V.ow Leve l (inch) 0.00 1.49 2.98 5.95 8.93 11.90 14.88 17.85 A 0.75 0.75 0.66 0.45 0.37 0.33 0.31 0.29 16.01 B 0.75 0.75 0.63 0.41 0.33 0.29 0.27 0.26 13.40 All Nodes Un its 1 & 2 C 0.75 0.75 0.75 0.71 0.64 0.60 0.58 0.56 16.01 D 0.75 0.75 0.75 0.75 0.74 0.72 0.70 0.70 16.01 Report No. 1700553.402 R2 PAGE 15-9 tJ Structural Integrity Associates. Inc i nfo@structint.com m 1-877-45I-POWER e structint.com@)

6.0 CRACK GROWTH EVALUATION The crack growth evaluation is performed for the bounding surge line weld locations, which were identified in the stress analysis. Fatigue crack growth is computed using linear elastic fracture mechanics (LEFM) techniques . The crack growth evaluation is used in conjunction with calculated allowable flaw sizes to determine the required inspection interval for a postulated flaw in the surge line at the bounding weld locations.

For a postulated initial flaw, crack growth is simulated until the flaw has reached the allowable flaw size based on the ASME Code,Section XI, Appendix L [2] procedures. The required inspection period is equal to the time to reach the allowable flaw size.

Crack growth evaluation uses two software packages, SI-TIFFANY [35] for calculating the stress intensity factors and pc-CRACK [34] for calculating the crack growth.

6.1 Loads The following loads are applied to the crack growth evaluation:

• Deadweight - The stresses due to deadweight are extracted from the finite element model in Section 4.0 .

• Seismic - The SRSS of moments due to OBE is used to calculate a primary bending stress. Since crack growth only considers Service Level A and B loads, SSE is not included in the evaluation.

• Weld Residual Stresses - The weld residual stresses are calculated as a function of flaw depth, as discussed further in Section 6.3 .

• Thermal Transients - The temperatures and pressures due to the thermal transients in Table 6-1 are extracted from the finite element model in Section 4.0.

• Crack Face Pressure - The thermal transient loads include the pressure history during the transients and do not account for the crack face pressure. The stress analyses in Section 4.0 used an un-cracked structure to develop the stress results that are used in the closed form solution of the crack model. Initiation of the crack introduces a new surface for Report No. 1700553.402 R2 PAGE 16-1

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pressure loading. This new crack surface pressure loading is modeled to be a far field loading (surface pressure with no end effects). A unit load of 1 ksi is scaled to the minimum and maximum operating pressures of each thermal transient in Table 6-1 , and then applied as an additional membrane stress.

Stress intensity factors for deadweight and thermal transients are calculated from the finite element stresses using SI-TIFFANY [35], as discussed further in Section 6.4.

6.2 Thermal Transients for Crack Growth Analysis The maximum temperature and operating pressure range for the thermal transients are determined from Table 3-4 and shown in Table 6-1. The number of annual cycles is calculated from the 10-year cycles in Table 3-3. Since the MOP Heatup and Cooldown transients are defined together for each temperature (MOPHUCD) , the cycle counts are summed for each combined MOP Heatup (MOPHU) and Cooldown (MOPCD) pairing.

6.3 Weld Residual Stress The weld residual stresses are calculated as a function of flaw depth . The distribution is reproduced from NUREG-0313 [33, Figure 3 in Appendix A] and shown in Figure 6-1. The through-wall residual stress distribution in the axial direction obtained from laboratory-measured data is given as:

2 3 4 (JG)= (Ji [ 1- 6.91 G) + 8.69 G) - o.48 G) - 2.03 G) ] (6-1) where, O"i = inside surface stress, given as the yield stress of the weld metal (ksi) x = distance into pipe wall from inside diameter (in) t = wall thickness (in)

The yield strength of 30 .0 ksi for Type 316 stainless steel, obtained from the ASME Code ,

Section 11 [25] is used for the inside surface stress. Equation 6-1 is plotted and fitted to a third order polynomial using the wall thickness of 1.406 inches (Section 3.1) and the minimum yield Report No. 1700553.402 R2 PAGE 16-2 t1 Structural Integrity Associates. Inc info@structint.com m 1-877-45!-POWER e structint.com @

stress of 30.0 ksi at the inside surface, x = 0. The axial residual stresses (in ksi) as a function of the distance from the inside surface (in inches) in third order polynomial form are:

O"(x) = 30.632 - 158.96x + 170.88x 2 - 49.003x 3 (6-2) 6.4 Stress Intensity Factors Stress analyses of deadweight and the thermal transients have been performed in the stress analysis [8]. SI-TIFFANY [35], generates tables of the maximum and minimum stress intensity factors, Kmax and Kmin, respectively, for various flaw depths and aspect ratios for deadweight and thermal transients.

The stress intensity factors for each deadweight and thermal transient case are determined using the representative fracture mechanics models in Figure 6-2 , which are incorporated into SI-TIFFANY.

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6.5 Postulated Initial Surface Flaw Per ASME Section XI, Appendix L (L-3210) [2], the postulated initial flaw for austenitic piping is a semi-elliptical circumferential or axial flaw on the inside surface. The initial flaw depth of the postulated flaw is determined from the applicable inservice inspection acceptance standard in Table IWB-3410-1 [2] using a flaw aspect ratio of 0.167 per ASME Section XI , Appendix L-3212

[2] .

The flaw depth ratio of 0.14408 alt is determined, using as inputs, the surge pipe thickness of 1.406 inches, and the prescribed aspect ratio of 0.167. Using linear interpolation in Table IWB-3514-1 [2] for austenitic piping, the postulated initial flaw depth (a =0.14408 x t =0.14408 x 1.406" = 0.202 inches) is calculated .

The next step per Appendix L guidance is to calculate the initial flaw aspect ratio. The aspect ratio for the semi-elliptical surface flaw is determined from Table L-3210-2 [2]. This Table uses the component thickness, and a calculated parameter, the membrane-to-gradient cyclic stress ratio. The membrane-to-gradient cyclic stress ratio is defined in Tables L-3210-2 [2] as follows:

Table L-3210-2 Austenltic Piping Postulated Equivalent Single Crack Aspect Ratios (a/ e)

IJ.<Tm//J.<Tg Nominal Wall Thickness, t, in. (mm)

[Note (1)] S0.218 (5.5) 0.344 (8.7) o. 719 (18.3) 1.125 (28.6) .1!2.125 (54) 0 0.0105 0.0107 0.0081 0.0082 0.0088 0.1 0.0280 0.0253 0.0265 0.0289 0.0362 0.25 0.0410 0.0446 0.0654 0.0807 0.1667 1 0.0556 0.0833 0.1351 0.1639 0.1667 3 0.0588 0.1031 O.i539 0.1667 0.1667 IX> 0.1667 0.1667 0.1667 0.1667 0.1667 GENERAL NOTE: Linear interpolation is permissible.

NOTE:

(1) The lnembra*ne-tb-gradlent cyclic stress ratio 1s stated as follows:

Lla,n Llag

=L ~

Dcota1 x(Mm)

Llag i n; = (Liam + LIC1gr x N1 Drota1 = In; I

Summation is over all types of transient loading conditions.

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Extracted Note 1 from Table L-3210-2 [2]

where ,

!::,,.am = cyclic membrane stress i::,,.a9 = cyclic linear and nonlinear gradient stress Di = (LiO"m + LiCT9 )~ X Ni D total = L i Di Ni= number of cycles for i1h load pair or transient loading condition n = fatigue crack growth rate exponent (slope of the log (da/dN) versus log (I::,,.K) curve)= 2.25 [32]

The cyclic membrane stress and the cyclic linear and nonlinear gradient stress for each transient are obtained from the stress analysis. The stress results are extracted at each path and the calculated membrane stress value obtained for all time points in the transient. The remainder of the stress path results (total stress - membrane stress) are considered to be "linear and nonlinear gradient" stresses and are again obtained throughout the transient. A maximum and minimum value of both membrane and "linear and non-linear gradient" stresses are obtained , and from these four values a membrane stress range value 11am (maximum -

minimum) is determined , along with a "linear and non-linear gradient" range value !1a9 . These are combined to obtain a Mm value for each transient. These individual transient values are tl.ag then used to obtain a location specific Mm value, considering all transients. Using this value of tl.ag Mm, shown in Table 6-2 and Table 6-3 , Table L-3210-2 is entered and a value for the aspect tl.ag ratio a/I is obtained for a thickness of 1.406". The calculated aspect ratio is applied to the initial flaw depth determined previously and then both are used in the crack growth calculation .

Table 6-2 shows the initial flaw sizes for a circumferential flaw with stratification transients occurring in both the lower horizontal piping and the upper horizontal piping regions, Report No. 1700553.402 R2 PAGE 16-5 tr Structural Integrity Associates. Inc info@structint.com m 1-877-45!-POWER e structint.com ~

respectively. Table 6-3 shows the initial flaw sizes for an axial flaw with stratification transients occurring in both the lower horizontal piping and the upper horizontal piping reg ions, respectively.

6.6 Stainless Steel Fatigue Crack Growth Law The reference fatigue crack growth rate curves are taken from Code Case N-809 [32] for Type 304 and Type 316 stainless steels and associated weld metals in a PWR environment are:

da/dN = Co*liKn, units of inch/cycle (6-3) where:

Co = scaling parameter that accounts for the effect of loading rate and environment on fatigue crack growth rate

=CST SR SENv n = slope of the log (da/dN) versus log (.6.K) curve= 2.25 C = nominal fatigue crack growth rate constant

= 4.43 x 10-7 for .6.K ;;::: .6.Kth

= 0 for .6.K < .6.Kth

.6.K = stress intensity factor range, ksi-i in

.6.Kth = 1.00 ksi-i in ST = parameter defining effect of temperature on FCG rate

= e-2s15rrK for 300°F :s; T :5 650 °F

= 3.39x10 5 e[(-2516ff K)-o.o 3on Kl for 70°F :5 T < 300°F T = metal temperature, °F SR = parameter defining the effect of R-ratio on FCG rate

= 1.0 for R < 0

= 1 + es.02(R-0.748) for o :s; R < 1.0 R = Kmin/Kmax = R ratio SENv = parameter defining the environmental effects on FCG rate

= T R0 .3 TR = rise time, sec TK = [(T-32)/1.8+273.15], (Temperature in Kelvin, K)

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The bounding maximum rise time of all transients in Table 3-4 is 20,952 seconds for Plant Heatup. The maximum temperature of 652°F is applied to calculate the crack growth rate. The use of the longest rise time among all thermal transients conservatively yields increased crack growth.

The referenced fatigue crack growth rate curves are calculated and input into pc-CRACK [34].

6.7 Crack Growth Analysis The crack growth law in Section 6.6 is available within pc-CRACK. The material for the surge line welds is modeled as being equivalent to A-376 TP 316 Stainless Steel, which is the base metal of the surge line. Crack growth analysis use the representative fracture mechanics models in Figure 6-2, which are incorporated into pc-CRACK.

Per Section 3.1, the piping dimensions are:

Inside Radius: 5.594 in Pipe Thickness: 1.406 in The dimensions of the postulated initial flaws are listed in Table 6-2 and Table 6-3. The aspect ratio of the flaw is allowed to vary as the flaw grows.

The sequence of events for fatigue crack growth is computed on a yearly basis. Thus, the cycles shown in Table 6-1 are used to create a number of cycles/year, and thus a yearly crack growth calculation . In fatigue crack growth, there is no requirement for load pairing between transients per ASME Code,Section XI [2]. Therefore, each thermal transient is analyzed in an arbitrary sequential sequence. This approach is consistent with Subsubarticle C-3210 of the ASME Code,Section XI [2].

The stress intensity factors for deadweight and the thermal transients are calculated using SI-TIFFANY [35] and are computed as a function of crack depth and can thus be superimposed for the various operating states.

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For normal operation (Normal or Service Level A) conditions , the individual stress intensity factors that contribute to the nominal maximum applied stress intensity factor, Kmax , and the minimum applied stress intensity factor, K mi n, are summarized in the tabulation below. The stress intensity factor range, liK , is computed by taking the difference of the summed K max and K min, K residua1 and K deadweig ht are from constant loads and do not contribute to the liK range but affect the R-ratio (Kmi n/Kmax) , which accounts for mean stress effects. The crack face pressures, K crack tace pressure max and K cracktace pressure min, are the min imum and maximum operating pressure during the transient in Table 6-1.

Km ax Kmin Kcieadweight Kd eadwe ight Kresidual Kresidual K crack face pressure max K crack face press ure min Kth ermal transient max Kthermal transient min For the OBE event (Upset or Service Level B) condition , Kmax and Kmin are taken as the positive and negative seismic bending stress and the crack face pressure is taken at normal operating pressure.

Kmax Km in Kdea dweig ht K deadweight Kresidual Kresidual K crack face pressure (normal operating pressure) K crack face pressure (normal operating pressure)

K norm al operation Knormal operation K o sE , positive K o sE , negative The crack growth is performed for an evaluation period of eighty (80) years . The time for the postulated flaw to grow beyond the allowable flaw sizes in Section 5.0 is reported in Table 6-4 through Table 6-7 . If the flaw does not grow beyond the allowable flaw sizes, the final flaw size at the end of evaluation period (80 years) is noted.

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6.8 Crack Growth Results Table 6-4 and Table 6-5 summarize the crack growth results for a postulated circumferential flaw subjected to stratification transients occurring in both the lower horizontal piping and the upper horizontal piping regions , respectively. Table 6-6 and Table 6-7 summarize the crack growth results for a postulated axial flaw subjected to stratification transients occurring in both the lower horizontal piping and the upper horizontal piping regions, respectively. The path locations in the surge line are shown in Figure 4-10.

These results utilize the bounding Unit 2 geometry and bounding loads from Unit 1 and 2, the initial flaw size corresponding to either the upper or lower stratification loading, and therefore the determined crack growth durations are applicable for both Unit 1 and 2.

The allowable flaw size for each path is linearly interpolated using the final flaw length . The most limiting crack growth location is stress paths P1 through P3, which are the surge pipe-to-RCS nozzle butt weld connection.

For an axial flaw, the most limiting crack growth location shows greater than 80 years of allowable flaw growth.

It takes 60 years for the postulated circumferential flaw at P3 to reach the allowable flaw size .

Therefore, the bounding allowable operating period is at least 60 years for the entire surge line.

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Table 6-1: Thermal Transients for Crack Growth Analysis Maximum Pressure Range (ksig) Annual Transient Temperature (°F)

Cycles PZR HL Max. Min.

Plant Heatup 652 606 2.235 0 5 Plant Cooldown 652 606 2.235 0 5 Large Step Load 652 606 2.235 1.945 0.2 Decrease Reactor Trip 652 606 2.24 2.038 5 OBE (Normal 2(2) 652 606 2.235 2.235 Operation)

MOPHUCD 320 450 130 0.44 0.44 0.5 + 1(1)

MOPHUCD 300 450 150 0.44 0.44 1(1)

MOPHUCD 290 450 160 0.44 0.44 0.5 + 1(1)

MOPHUCD 280 450 170 0.44 0.44 1(1)

MOPHUCD 270 450 180 0.44 0.44 1.3 + 2( 1)

MOPHUCD 260 450 190 0.44 0.44 2.3 + 0.5( 1)

MOPHUCD 250 450 200 0.44 0.44 3.8 + 1.8(1)

MOPHUCD 240 450 210 0.44 0.44 1.8 + 1.8(1)

Notes:

(1) The MOP Heatup and Cooldown transients are defined together for each temperature (MOPHUCD), and as such, the cycle counts from Table 3-3 are summed for each combined MOP Heatup (MOPHU) and Cooldown (MOPCD) pairing.

(2) OBE has 20 internal cycles per 1 event. 1 event per 10 years with 20 cycles per event equals 2 cycles per year.

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Table 6-2. Initial Flaw Sizes for Circumferential Flaw Path Locations Station No. (1) lia m//J.JJg a/I alt,% a, in I, in UPPER Stratification Loading I P1 - P5 0.114 0.0378 14.41 0.20258 5.36 II P6 - P10 0.256 0.1053 14.41 0.20258 1.92 II I P11 - P14 0.166 0.0633 14.41 0.20258 3.21 IV P15 - P20 0.100 0.0309 14.41 0.20258 6.56 P21 , P22, V 0.280 0.1073 14.41 0.20258 1.89 P24, P25 LOWER Stratification Loading I P1 - P5 0.164 0.063 14.41 0.20258 3.22 II P6 - P10 0.356 0.113 14.41 0.20258 1.80 Ill P11 - P14 0.120 0.0409 14.41 0.20258 4.95 IV P15 - P20 0.194 0.077 14.41 0.20258 2.64 P21, P22, V 0.176 0.0682 14.41 0.20258 2.97 P24, P25 Note:

(1) Path locations are shown in Figure 4-10.

(2) Since the aspect ratio (a/c) is smaller than 1/8 (0.125), a full 360° circumferential flaw model is used to calculate the stress intensity factors.

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Table 6-3. Initial Flaw Sizes for Axial Flaw Path Locations Station No. (1) l:::..amll:::..a9 a/I alt,% a, in I, in UPPER Stratification Loading I P1 - P5 0.122 0.0420 14.41 0.20258 4.82 II P6 - P10 0.185 0.0726 14.41 0.20258 2.79 Ill P11 - P14 0.392 0.1162 14.41 0.20258 1.74 IV P15 - P20 0.418 0.1183 14.41 0.20258 1.71 P21, P22, V 0.467 0.1222 14.41 0.20258 1.66 P24, P25 LOWER Stratification Loading I P1 - P5 0.141 0.051 14.41 0.20258 3.98 II P6 - P10 0.264 0.106 14.41 0.20258 1.92 Ill P11 - P14 0.436 0.120 14.41 0.20258 1.69 IV P15 - P20 0.520 0.126 14.41 0.20258 1.61 P21 , P22, V 0.390 0.1160 14.41 0.20258 1.75 P24, P25 Note:

(1) Path locations are shown in Figure 4-10.

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Table 6-4. Crack Growth Results for Circumferential Flaw - Lower Stratification Loading Final Flaw Size STATION Path allowable Year a, in C, in(S) I, deg(5l alt alt P1 80 0.4880 1.7598 36.05 0.7498 0.3471 P2 80 0.4909 1.7589 36.03 0.7499 0.3491 I P3 80 0.4720 1.7484 35.82 0.7500 0.3357 P4 80 0.3873 1.6803 34.42 0.7500 0.2755 P5 80 0.3838 1.6794 34.40 0.7500 0.2730 P6 80 0.3134 0.9513 19.49 0.7500 0.2229 P7 80 0.3673 1.0085 20.66 0.7500 0.2612 II P8 80 0.3520 0.9927 20.34 0.7500 0.2504 pg 80 0.2677 0.9299 19.05 0.7500 0.1904 P10 80 0.3520 0.9959 20.40 0.7500 0.2504 P11 (1l 80 0.371 4 4.5403 93.01 0.4266 0.2642 P12( 1l 80 0.3991 4 .8790 99.94 0.3958 0.2839 Ill P13(1l 80 0.3774 4.6137 94.51 0.4200 0.2684 P14(1l 80 0.3727 4.5562 93.33 0.4252 0.2651 P15 80 0.3264 1.3552 27.76 0.7500 0.2321 P16 80 0.3431 1.3627 27.91 0.7500 0.2440 P17 80 0.3562 1.3694 28 .05 0.7500 0.2533 IV P18 80 0.3590 1.3711 28.09 0.7500 0.2553 P19 80 0.3261 1.3550 27.76 0.7500 0.2319 P20 80 0.3327 1.3579 27.82 0. 7500 0.2366 P21 80 0.3370 1.5283 31.31 0.4891 (2) 0.2397 P22 80 0.3398 1.5295 31.33 0.4889 (2) 0.2417 V(2)

P24 80 0.3343 1.5267 31.27 0.4894 ( 2) 0.2378 P25 80 0.3395 1.5294 31.33 0.4889 (2) 0.2415 Notes:

(1) Per the crack growth model [34] the crack grows with fixed aspect ratio. The final flaw length is ca lculated using the flaw depth and the same initial aspect ratio.

(2) The allowable flaw size utilized CMTR data [7].

(3) The limiting allowable flaw sizes from either Unit 1 or Unit 2 are used.

(4) Bounding Paths are shown in BOLD for each axial pipe location (Figure 4-10).

(5) c = half final crack length, I = arc length of full flaw (6) Initial Flaw sizes for Lower Stratification loading used.

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Table 6-5. Crack Growth Results for Circumferential Flaw - Upper Stratification Loading Final Flaw Size STATION Path allowable Year a, in C, in<5l I, deg<5l alt alt P1 (1l 65 0.5022 6.6429 136.08 0.3576 0.3572 P2(1l 61 0.5020 6.6402 136.02 0.3577 0.3570 I p3(1) 60 0.5013 6.6310 135.83 0.3581 0.3565 p4(1) 80 0.4617 6.1071 125.10 0.3820 0.3284 p5(1) 80 0.4648 6.1481 125.94 0.3801 0.3306 P6 80 0.3759 1.0730 21.98 0.7500 0.2674 P7 80 0.3491 1.0534 21 .58 0.7500 0.2483 II P8 80 0.3503 1.0541 21 .59 0.7500 0.2491 pg 80 0.3513 1.0559 21.63 0.7500 0.2499 P10 80 0.3450 1.0502 21.51 0.7500 0.2454 P11 80 0.3615 1.6513 33.83 0.7500 0.2571 P1 2 80 0.3567 1.6494 33.79 0.7500 0.2537 111 P13 80 0.3555 1.6493 33 .79 0.7500 0.2528 P14 80 0.3599 1.6501 33.80 0.7500 0.2560 P15(1l 80 0.3683 5.9499 121.88 0.5176 0.2619 P16(1l 80 0.3836 6.1971 126.95 0.5021 0.2728 P17(1l 80 0.3661 5.9144 121 .15 0.5198 0.2604 IV P18(1l 80 0.4251 6.8675 140.68 0.4601 0.3023 P19<1l 80 0.4147 6.6995 137.24 0.4707 0.2950 P20(1l 80 0.4038 6.5234 133.63 0.4817 0.2872 P21 80 0.3228 1.0127 20.74 0.3754 (2 ) 0.2296 P22 80 0.3153 1.0074 20.64 0.3774 (2) 0.2243 v c2i P24 80 0.3 155 1.0072 20.63 0.3775 (2 ) 0.2244 P25 80 0.3421 1.0253 21.00 0.3708 (2) 0.2433 Notes:

(1) Per the crack growth model [34] the crack grows with fixed aspect ratio. The final flaw length is calculated using th e flaw depth and the same initial aspect ratio.

(2) The allowable flaw size utilized CMTR data [7] .

(3) The limiting allowable flaw sizes from either Unit 1 or Unit 2 are used .

(4) Bounding Paths are shown in BOLD for each axial pipe location (Figure 4-10).

(5) c = half final crack length, I = arc length of full flaw.

(6) Initial Flaw sizes for Upper Stratification loading used .

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Table 6-6. Crack Growth Results for Axial Flaw - Lower Stratification Loading Final Flaw Size STATION Path allowable Year a, in C, in<4 l I, in<4 l alt alt <1l< 2l P1 80 0.5550 2.1347 4.2694 0.5343 0.3947 P2 80 0.5387 2.1210 4.2420 0.5363 0.3831 I P3 80 0.5537 2.1340 4.2680 0.5344 0.3938 P4 80 0.3908 2.0199 4.0398 0.3720 0.2780 P5 80 0.4003 2.0237 4.0474 0.3715 0.2847 P6 80 0.2989 1.0096 2.0192 0.6071 0.2126 P7 80 0.3762 1.0744 2.1488 0.5845 0.2676 II P8 80 0.3473 1.0451 2.0902 0.5947 0.2470 pg 80 0.3050 1.0133 2.0266 0.6058 0.2169 P10 80 0.3705 1.0681 2.1362 0.5867 0.2635 P11 80 0.3375 0.9352 1.8704 0.6331 0.2400 P12 80 0.3245 0.9263 1.8526 0.6363 0.2308 Ill P13 80 0.3091 0.9145 1.8290 0.6404 0.2198 P14 80 0.3434 0.9389 1.8778 0.6318 0.2442 P15 80 0.3432 0.9014 1.8028 0.6450 0.2441 P16 80 0.3046 0.8696 1.7392 0.6561 0.2166 P17 80 0.3155 0.8788 1.7576 0.6529 0.2244 IV P18 80 0.3238 0.8854 1.7708 0.6506 0.2303 P19 80 0.3044 0.8700 1.7400 0.6559 0.2165 P20 80 0.3240 0.8834 1.7668 0.6513 0.2304 P21 80 0.3283 0.9531 1.9062 0.6269 0.2335 P22 80 0.3201 0.9471 1.8942 0.6290 0.2277 V

P24 80 0.3202 0.9460 1.8920 0.6294 0.2277 P25 80 0.3163 0.9444 1.8888 0.6299 0.2250 Note:

(1) The allowable flaw sizes from Table 5-4 are used.

(2) The allowable flaw size utilized CMTR data [7].

(3) Bounding Paths are shown in BOLD for each axial pipe location (Figure 4-10) .

(4) c = half final crack length, I= length of full flaw.

(5) Initial Flaw sizes for Lower Stratification loading used.

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Table 6-7. Crack Growth Results for Axial Flaw - Upper Stratification Loading Final Flaw Size STATION Path allowable Year a, in C, in(4) I, in(4> alt alt (1H2>

P1 80 0.5783 2.5573 5.1146 0.4718 0.4113 P2 80 0.5679 2.5477 5.0954 0.4733 0.4039 I P3 80 0.6055 2.5720 5.1440 0.4697 0.4307 P4 80 0.4079 2.4542 4.9084 0.3166 0.2901 PS 80 0.4233 2.4589 4.9178 0.3160 0.3011 P6 80 0.3949 1.4811 2.9622 0.4422 0.2809 P7 80 0.3807 1.4724 2.9448 0.4453 0.2708 II PB 80 0.4048 1.4864 2.9728 0.4404 0.2879 pg 80 0.3836 1.4735 2.9470 0.4449 0.2728 P10 80 0.3917 1.4789 2.9578 0.4430 0.2786 P11 80 0.3399 0.9616 1.9232 0.6239 0.2417 P12 80 0.3362 0.9605 1.9210 0.6243 0.2391 Ill P13 80 0.3340 0.9601 1.9202 0.6244 0.2376 P14 80 0.3423 0.9627 1.9254 0.6235 0.2435 P15 80 0.2913 0.9069 1.8138 0.6430 0.2072 P16 80 0.2703 0.8944 1.7888 0.6474 0.1922 P17 80 0.2566 0.8874 1.7748 0.6499 0.1825 IV P18 80 0.2743 0.8972 1.7944 0.6464 0.1951 P19 80 0.3246 0.9386 1.8772 0.6319 0.2309 P20 80 0.3620 0.9674 1.9348 0.6219 0.2575 P21 80 0.3148 0.9028 1.8056 0.6445 0.2239 P22 80 0.3142 0.9027 1.8054 0.6445 0.2235 V

P24 80 0.3326 0.9172 1.8344 0.6394 0.2366 P25 80 0.3186 0.9064 1.8128 0.6432 0.2266 Note:

(1) The allowable flaw sizes from Table 5-4 are used.

(2) The allowable flaw size utilized CMTR data [7].

(3) Bounding Paths are shown in BOLD for each axial pipe location (Figure 4-10).

(4) c = half final crack length , I= length of full flaw.

(5) Initial Flaw sizes for Upper Stratification loading used .

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INSIDE WALL OUTSIDE WALL 50~--~..--------,,-------.~......----~......---,

o GE26 o GE 26 (4 azimuths)

A ANL 26 (2 azimuths)

<> ANL 26 (IN-SERVICE FROM KRB) t> ANL 20

- oo

.at:

( /)

(,/)

w 10 0 ,-- o _____ . 0 a.._-'>-~-

0 O $

ex:: 0 o

oc51 a

a ea B

(/)

-10 0 O 0 Oo

-20 0

-30 6A e e e

-40 0 0.2 0.4 0.6 0.8 1.0 alt Figure 6-1. Through-Wall Residual Stress as a Function of Depth Report No. 1700553.402 R2 PAGE I 6-17

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(a) Circumferential Flaw t

(b) Axial Flaw Figure 6-2. Semi-Elliptical Flaws on the Inside Surface of a Cylinder Report No. 1700553.402 R2 PAGE I 6-18 e

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7.0

SUMMARY

AND CONCLUSIONS The flaw tolerance of the bounding surge line location, namely the RCS hot leg surge nozzle-to-pipe weld (No. 37 (U1) and 3 (U2)), at NAPS Units 1 and 2 has been evaluated and the required successive inspection schedule has been determined for a postulated flaw per the requirements of ASME Code,Section X I, Appendix L. Stress analysis resu lts in Section 4.0 confirm that the hot leg surge nozzle-to-pipe weld is the limiting location in the surge line [38].

The flaw tolerance evaluation consisted of determining the loads at the bounding location and performing a finite element stress analyses to determine stresses due to thermal transients.

The stresses are used to determine the allowable flaw sizes and perform a crack growth evaluation to determine the allowable operating period based on crack growth of a postulated flaw compared to allowable flaw sizes.

The allowable operating period for the bounding surge line location (Welds SW-31 (U 1) and SW-37 (U2)) is at least 60 years. Therefore, per the guidelines of Table L-3420-1 of ASME Code,Section XI, Appendix Land IWB-2410 of ASME Code,Section XI , the successive inspection schedule for the surge line (Welds No. 37 (U1) and 3 (U2)) is ten years.

While not required for initial license renewal or subsequent license renewal , Dominion also contracted SIA to evaluate additional locations for Unit 1 (nodes 175, 190, and 198 per Reference [6]) and Unit 2 (nodes 173, 187, and 201 per Reference [6]) for asset management considerations. The allowable operating period for these locations is at least 60 years.

Therefore, per the guidelines of Table L-3420-1 of ASME Code,Section XI, Appendix Land IWB-2410 of ASME Code,Section XI, the successive inspection schedule for these asset management locations on the surge line is sixty years.

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8.0 REFERENCES

1. U.S. NRC NUREG/CR-6260, "Application of NUREG/CR-5999 Interim Fatigue .

Curves to Selected Nuclear Power Plant Components," February 1995.

2. ASME Boiler & Pressure Vessel Code,Section XI, 2013 Edition.

3. Unit 1, Dominion Calculation No. 14938.56-NP(B)-005-XC, Revision 2, "Pipe Stress Analysis - Effects of Thermal Stratification and Thermal Striping on the Pressurizer Surge Line" including NUPIPE-SW ME-110 stress analysis output from 3/11/89, SI File No. 1700553.202.

4. Unit 2, Dominion Calculation No. 14938.56-NP(B)-004-XC, Revision 3, "Pipe Stress Analysis - Effects of Thermal Stratification and Thermal Striping on the Pressurizer Surge Line" including NUPIPE-SW ME-110 stress analysis output from 3/08/89, SI File No. 1700553.202.

5. SI Calculation No. 1700553.301 , "Loads for North Anna Surge Line," (For revision number refer to SI Project Revision Log, 1700553, latest revision).

6. SI Calculation No. 1700553.302, "Finite Element Model Development of the Pressurizer Surge Line," (For revision number refer to SI Project Revision Log, 1700553, latest revision).

7. SI Calculation No. 1700553.303, "Allowable Flaw Size Determination for North Anna Units 1 and 2 Surge Line," (For revision number refer to SI Project Revision Log , 1700553, latest revision).

8. SI Calculation No. 1700553.304, "Finite Element Stress Analysis of the Pressurizer Surge Line-Appendix L Flaw Growth Evaluation - Subsequent License Renewal," (For revision number refer to SI Project Revision Log, 1700553, latest revision) .

9. SI Calculation No. 1700553.305, "Crack Growth Evaluation for North Anna Surge Line," (For revision number refer to SI Project Revision Log, 1700553, latest revision) .

10. Materials Reliability Program: Thermal Cycling Screening and Evaluation Model for Normally Stagnant Non-lsolable Reactor Coolant Branch Line Piping with a Generic Application Assessment (MRP- 132), EPR I, Palo Alto, CA 2004.

1009552. (1 ).

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11. Materials Reliability Program: Management of Thermal Fatigue in Normally Stagnant Non-lsolable Reactor Coolant System Branch Lines - Supplemental Guidance (MRP- 146S). EPRI, Palo Alto, CA: 2009. 1018330.

12. Materials Reliability Program Letter MRP 2007-013, "

Subject:

MRP-132/146/170 Guidance Related to Heat Transfer Coefficients for Use in Stress Analysis," April 2, 2007, SI File No. 1700553.212.

13. Holman, J.P., Heat Transfer, Ninth Edition, McGraw-Hill, 2002.

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15. Dominion Document No. DOM-SLR-17-007, Subject* Summary of Transients for Flaw Tolerance Evaluation of the SPS Pressurizer Surge Line for SLR, May 17, 2017, SI File No. 1601007.205.

16. Dominion Drawing No. 12050-FM-093A, Revision 42, FlowNalve Operating Numbers Diagram, Reactor Coolant System - Loop 3, North Anna Power Station Unit 2, Virginia Power, SI File No. 1700553.213.

17. Design Inputs for NAPS SLR Surge Line Appendix L Evaluation , SI File No.

1700553.208.

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1700553.401.

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ANSYS , Inc., September 2012.

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22. Dominion Energy Purchase Order 70329627, Provide Environmentally Assisted Fatigue Screening Activities for North Anna Unit 1 and Unit 2, SI File No.

1700553.220.

23. Email from Charles A. Tomes (Dominion) to Tim Gilman, Dated March 13, 2018 4:14 AM, "Re: nozzle thickness ," SI File No. 1700553.204.

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24. Email from Tim Gilman (SI) to Charles A. Tomes (Dominion), Dated March 12, 2018 4:59 PM, "Re: nozzle thickness," with nozzle drawing attachment "03-12 66.pdf," SI File No. 1700553.204.

25.ASME Boiler and Pressure Vessel Code, Section 11, Part D - Material Properties, 2007 Edition with Addenda through 2008.

26. Nuclear Power Piping, American Society of Mechanical Engineers, USAS 831.7, 1968.

27.ASME Boiler & Pressure Vessel Code, Section Ill , 1986 Edition with 1987 Addenda.

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Subject:

Weld Process for hot leg surge nozzle to pipe weld for SPS unit 1," dated June 15, 2017, SI File No. 1700553.218.

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1700553.215.

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September 201 5.

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

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37. Email from Charles A. Tomes (Dominion) to Tim Gilman (SI) , "

Subject:

More Information on NAPS Surge Line," dated April 6, 2018 with attachment "Framatome-18-00822 NP Letter for NAPS PZR WOL Material and CUF-DRAFT (2).pdf," SI File No. 1700553.216 .

38. Dominion Letter, DOM-SLR-19-005, Dated September 16, 2019, "Review of CUF values for NAPS Units 1 and 2 Pressurizer Surge Line RCS Hot Leg Location ... ",

SI File No 1700553.235.

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40. North Anna Operating Conditions, E-mail from Charles A. Tomes (Dominion) to Jim Axline (SI),

Subject:

"FW: RE: 1700553.301 , Rev A - Loads for North Anna Surge Line Flaw Growth Evaluation", Dated Friday, January 11 , 2019 8:41 A M, SI File No. 1700553.222.

41. NUREG-1766, Safety Evaluation Report Related to the License Renewal of North Anna Power Station, Units 1 and 2, and Surry Power Station , Units 1 and 2, December 2002.

42.ASME Boiler & Pressure Vessel Code,Section IX, 20 13 Edition.

43. Surry Power Station , Units 1 and 2, Appl ication for Subsequent License Renewal ,

Oct 2018, ADAMS No. ML18291A828.

Note: 1) This document is a supporting technical reference for the work performed, however this report does not contain proprietary information from the indicated reference document.

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