L-2015-272, 1301103.401, Revision 0, Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L, May 2015

1301103.401, Revision 0, Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L, May 2015

ML15314A161
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Site:	Saint Lucie
Issue date:	05/08/2015
From:	Alleshwaram A Structural Integrity Associates
To:	Florida Power & Light Co, Office of Nuclear Reactor Regulation
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ML15314A218	List: L-2015-272, 1301103.401, Revision 0, Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L, May 2015 ML15314A160
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FPL Letter L-2015-272 ATTACHMENT 5 SIA non-proprietary Report No. 1301103.401, Rev. 0, "Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L," dated May 2015 Next 82 pages

Report No. 1301103.401 Revision 0 Project No. 1301103 May 2015 Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L Preparedfor."

St. Lucie Nuclear Station, Units 1 & 2 Florida Power & Light Florida City, FL Contract Number 02322154 Preparedby:

Structural Integrity Associates, Inc.

San Jose, California Preparedby: Date: 5/08/2015 Apamna Alleshwaramn Reviewed by: Date: 5/08/2015 Nathaniel Cofie Approved by: Date: 5/08/2015 Nathaniel Cofie S*ctur*! tn*rgOl A*ociates, Inc."

REVISION CONTROL SHEET Document Number: 1301103.401

Title:

Flaw Tolerance Evaluation of St. Lucie Surge Line Welds Using ASME Code Section XI, Appendix L Client: Florida Power & Light SI Project Number: 1301103 Quality Program: [] Nuclear D] Commercial Section Pages Revision [ Date Comments 1.0 1-1 2 0 5/08/2015 Initial Issue 2.0 2-1 2 3.0 3 3-4 4.0 4-1 8 5.0 5-1 15 6.0 6-1 - 6-22 7.0 7-1 19 8.0 8-1 9.0 9-1 - 9-3

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Table of Contents Section Pg

1.0 INTRODUCTION

............................................................................ 1-1 2.0 TECHNICAL APPROACH ................................................................ 2-1 3.0 ENVIRONMENTALLY ASSISTED FATIGUE SCREENING ...................... 3-1 3.1 EAF Screening and Detenrmination of Critical Locations................................ 3-1 3.1.1 Screening of PressurizerSurge Nozzle ...................................................... 3-1 3.1.2 Screening of Hot Leg Surge Nozzle .......................................................... 3-3 3.1.3 Screening of Surge Line Piping.............................................................. 3-3 4.0 PIPING INTERFACE LOADS AND THERMAL TRANSIENTS................... 4-1 4.1.1 Piping Loads.................................................................................... 4-1 4.1.2 Thermal Transients............................................................................ 4-1 4.1.3 ThermalAnchor Movenients................................................................... 4-2 4.1.4 StratificationHeight........................................................................... 4-2 5.0 STRESS ANALYSIS ........................................................................ 5-1 5.1 Finite Element Model ...................................................................... 5-1 5.2 Material Properties for Stress Analysis ................................................... 5-2 5.3 Thermal and Stratification Stress Analyses Including Pressure and TAM ............. 5-2 5.3.1 DeadweightAnalysis........................................................................... 5-3 5.3.2 Mechanical Boundaiy Conditions............................................................ 5-3 5.4 Stress Analysis Results..................................................................... 5-4 6.0 ALLOWABLE FLAW SIZE EVALUATION ........................................... 6-1 6.1 Interface Loads.............................................................................. 6-2 6.2 Load Combination.......................................................................... 6-3 6.3 Material Properties for Allowable Flaw Size Determination ........................... 6-4 6.4 Welding Process ............................................................................ 6-4 6.4.1 Z-factor.......................................................................................... 6-5 Report No. 1301103.40 1.R0 iii

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6.5 Thermal Transient Loads................................................................... 6-5 6.6 Operating Conditions....................................................................... 6-6 6.7 Allowable Circumferential Part Through-Wall Flaw.................................... 6-6 6.7.1 Base Metal.................................................................................... 6-6 6.7.2 Weld Metal ................................................................................... 6-7 6.8 Allowable Flaw Size Determination Results ............................................. 6-9 7.0 CRACK GROWTH EVALUATION...................................................... 7-1 7.1 Design Inputs ............................................................................... 7-1 7.1.1 Geometry and Operating Conditions....................................................... 7-1 7.1.2 Loads.......................................................................................... 7-2 7.1.3 Thermal Transients and Thermal StratificationEvents ................................... 7-3 7.1.4 Residual Stresses............................................................................. 7-4 7.1.5 Crack Growth Laws.......................................................................... 7-5 7.1.6 PostulatedInitial Surface Flaw............................................................. 7-9 7.2 Calculations................................................................................ 7-10 7.2.1 Stress Intensity Factors..................................................................... 7-10 7.2.2 Crack Growth............................................................................... 7-12 7.3 Crack Growth Analysis Results.......................................................... 7-13 7.4 Inspection Interval ........................................................................ 7-14 8.0

SUMMARY

AND CONCLUSIONS ...................................................... 8-1

9.0 REFERENCES

............................................................................... 9-1 Report No. 1301103.401 .R0 iv

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List of Tables Table Pg Table 4-1: Deadweight and Seismic Loads at the Surge Line Ends............................... 4-3 Table 4-2: Surge Line Transients..................................................................... 4-4 Table 4-3: Stratification Transients for Surge Line [10] ........................................... 4-6 Table 4-4: Thermal and Seismic Anchor Movements.............................................. 4-7 Table 4-5: Stratification Height ...................................................................... 4-8 Table 5-1: Material Properties........................................................................ 5-6 Table 5-2: Hanger Spring Constants and Preloads ................................................. 5-6 Table 5-3: Linearized Stress Paths and Nodes ...................................................... 5-7 Table 6-1: Geometry and Operating Conditions................................................... 6-10 Table 6-2: Thermal Expansion Loads .............................................................. 6-11 Table 6-3: Interface Loads .......................................................................... 6-13 Table 6-4: Material Properties of Bounding Surge Line Elbow.................................. 6-15 Table 6-5: Loads for CASSPAR Base Metal Analyses ........................................... 6-17 Table 6-6: Inputs for CASSPAR Base Metal Analyses ........................................... 6-18 Table 6-7: Allowable Part Through-Wall Circumferential Flaw Sizes in the Weld Metal for Unit 1......................................................................................... 6-19 Table 6-8: Allowable Part Through-Wall Circumferential Flaw Sizes in the Weld Metal for Unit 2......................................................................................... 6-20 Table 6-9: Allowable Part Through-Wall Circumferential Flaw Sizes in the Base Metal for Unit 1......................................................................................... 6-21 Table 6-10: Allowable Part Through-Wall Circumferential Flaw Sizes in the Base Metal for Unit 2......................................................................................... 6-22 Table 7-1: Thermal Transients and Thermal Stratification Events............................... 7-15 Table 7-2: Seismic Loads............................................................................ 7-16 Table 7-3: Circumferential Allowable Flaw Size ................................................. 7-17 Table 7-4: Circumferential Crack Growth Results ................................................ 7-17

List of Figures Figure Page Figure 1-1: Schematic of Surge Line at St. Lucie Units 1 and 2................................... 1-2 Figure 4-1: PIPESTRESS St. Lucie Units 1 and 2 Surge Line Model ............................ 4-3 Figure 4-2: Surge Line High and Low Pressure Stratification During Heatup ................... 4-7 Figure 5-1: ANSYS Finite Element Model of the Surge Line ..................................... 5-8 Figure 5-2: Applied Mechanical Boundary Conditions for Deadweight Case.................... 5-9 Figure 5-3: Applied Temperatures for the Stratification Low Pressure Event at A320°F ..... 5-10 Figure 5-4: Stress Path Definitions for the Elbow Location at the Hot Leg End ............... 5-11!

Figure 5-5: Stress Path Definitions for the Middle Elbow Location............................. 5-12 Figure 5-6: Stress Intensity Contour Plot for a Typical Thermal Transient ..................... 5-13 Figure 5-7: Stress Intensity Contour Plot for the Stratification Low Pressure Event atA32O0°F........................................................................................ 5-14 Figure 5-8: Stress Intensity Factors of Axial and Circumferential Flaws for Stress Path 4 Hot Standby Stratification Event ................................................................... 5-15 Figure 7-1: Semi-Elliptical Circumferential Flaw on the Inside Surface of a Cylinder ........ 7-18 Figure 7-2: Through-Wall Residual Stress Distribution .......................................... 7-19 Report No. 1301103.401 .R0 vi I Ty Asoiats, nc

1.0 INTRODUCTION

A flaw tolerance evaluation in accordance with ASME Code,Section XI, Appendix L [1] has been performed to manage fatigue at critical locations of the St. Lucie Units 1 & 2 pressurizer surge line locations by inspection and flaw tolerance evaluations. Specifically, the critical locations of concern are where the calculated fatigue cumulative usage factor (CUE) was determined to exceed the allowable usage factor when environmentally assisted fatigue (EAF) is considered. When this occurs, an alternative ASME Code,Section XI Appendix L flaw tolerance evaluation can be performed. The ultimate objective of the evaluation was to determine if the provisions of Appendix L can be met and use that as a basis for determining the required successive examination interval for the locations where the CUE exceeds 1.0 when BAF is considered.

The pressurizer surge lines at St. Lucie Units 1 & 2, connect to the pressurizer with a nozzle that is oriented in the vertical direction as shown in Figure 1-1. They then have a 900 elbow and run horizontally, with expansion loops to the hot leg. At the hot leg there is a vertical drop to the hot leg surge nozzle. The surge lines in both units are 12 inch Sch. 160 stainless steel piping.

The flaw tolerance evaluation was performed to determine the acceptability of all the locations of concern with respect to the requirements of ASME Code Section XI Appendix L. The evaluation was performed for the bounding location on the surge line. Based on a comparison of geometry, material properties and applicable loads, the results of the detailed evaluation of the bounding location are applicable to the other critical locations of concern on the surge line.

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Pr-essurizer

'Hot :Leg Nozzle Figure 1-1: Schematic of Surge Line at St. Lucie Units 1 and 2 Report No. 1301103.401 .R0 1-2 ID*l tOUFYA/C~WIc

2.0 TECHNICAL APPROACH The evaluation was performed in accordance with the requirements of ASME Code,Section XI, Appendix L. Consistent with the procedure in ASME Code Section XI, Appendix L the methodology used to determine the successive inspection schedule consists of the following principal tasks:

• Determine the loads and stresses at the critical locations of the surge line. These are addressed in Sections 4.0 and 5.0.

• Use the stresses at the critical location to determine the allowable flaw depths for various service levels. This is addressed in Section 6.0.

• Postulate hypothetical flaws at the critical location. Select appropriate crack models to simulate the postulated flaws. This is discussed in Section 7.0. As discussed below, circumferential flaws are considered in the crack growth evaluation since they are bounding.

• 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 detenrmine 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. Details are provided in Section 7.0.

• Determine the required successive inspection schedule by comparing the crack size of the end of the evaluation period to the allowable flaw size. This is discussed in Section 7.0.

For the St. Lucie surge line geometry, significant axial stresses are expected due to the relatively short section of pipe between the hot leg nozzle and first elbow on the hot leg end, as shown in Figure 1-1. The short pipe section makes the surge line more rigid, thereby restricting thermal expansion of the horizontal sections of the surge line and increasing axial stresses, particularly at the elbow. From the stress analyses using the finite element model, as discussed in Section Report No. 1301103.401 .R0 2-1

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5.04, axial stresses are much higher than hoop stresses, especially for thermal stratification events due to the resultant bending moment from the differential thermal expansion between the top and bottom of the horizontal section of the surge line. As such, the crack growth for a circumferential flaw with the higher axial stresses is expected to be significantly greater than the crack growth for an axial flaw with the lower hoop stresses. Hence, only crack growth of a circumferential flaw is considered in this evaluation, and the crack growth of an axial flaw is bounded by the analysis of a circumferential flaw for St. Lucie.

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3.0 ENVIRONMENTALLY ASSISTED FATIGUE SCREENING 3.1 EAF Screening and Determination of Critical Locations EAF screening calculations, based on previous fatigue analyses, were performed on all the critical locations, including the pressurizer and hot leg surge nozzles, to provide the basis for evaluating only the surge piping as the bounding location and not the pressurizer and hot leg surge nozzles.

In accordance with NIJREG- 1801, Revision 2 (or the Generic Aging Lessons Learned report)

[32], EAF may be evaluated using the fatigue life reduction factor (Fan) formulations in NUJREG/CR-5704 [14] for austenitic stainless steel components or in NU7REG/CR-6583 [33] for carbon or low alloy steel components. Fen values are multiplied by the fatigue usage factors computed using the air fatigue curve to obtain total cumulative environmental usage (Uen) factors. If the Uen for a location remains less than 1.0, then the component is acceptable for the period of operation and does not require flaw tolerance analyses to address EAF.

The limiting value for Fen for stainless steel material, assuming a temperature above 200°C and a zero (slowest possible) strain rate, is 15.35. It is conservative to use stainless steel Fen values for nickel alloy materials.

3.1.1 Screening of Pressurizer Surge Nozzle a) Unitl1:

i. The Unit 1 pressurizer surge nozzle and attached elbow were replaced during refueling outage SL1-20, which occurred in the Fall of 2005 [6, Section 1.4].

Since the components were replaced in kind, it was concluded through an equivalency evaluation that the existing design basis stress calculations are unaffected by the replacement [7].

ii. The existing design basis fatigue analysis of the Unit 1 pressurizer surge nozzle reported a cumulative fatigue usage factor (CUF) of 0.9672 [6, p. 3-10]. Note that this fatigue analysis was performed utilizing a Seal Shell type of analysis [6],

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which is a very conservative, simplified method of analysis typically used in the 1960s.

iii. Subsequently, the Unit 2 surge nozzle (before application of weld overlay and with similar or identical design to that of Unit 1) had a fatigue analysis performed using more modem FEA methods. The Unit 2 pressurizer surge nozzle CUF for normal fatigue was reported to be 0.1073, which indicates a factor of nine reduction in CUE by using the modem FEA methods. If the Unit 1 nozzle was reanalyzed with similar modem FEA methods, more realistic and reduced stress estimates would be expected, similar to the Unit 2 results [6].

b) Unit 2:

1. In addition to the normal fatigue of the pressurizer surge nozzle for Unit 2 being 0.1073, since the new nozzle has not been fatigued by the approximately first 30 years of plant operation, it is clearly bounded by the limiting location in the surge piping. The Uen at bounding location of the surge piping was computed to be 13.544 at the elbow above the hot leg [5, p. 8], using projected numbers of cycles (less severe than design numbers of cycles assumptions). Therefore environmental CUE would be expected to be less than allowable for the PSL Unit 1 surge nozzle, if current analysis techniques were utilized.

ii. A preemptive WOL was applied to the pressurizer surge nozzle of Unit 2 by ARE VA during the SL2-17 refueling outage. The results of the WOL fatigue evaluation are documented in AREVA Document No: 32-9041583-000 [4].

According to that document, the CUE at NW1[ ]The corresponding EAF usage when multiplied by the Fen of[- ]

Similarly, the CUE at PI-]and the corresponding EAF usage i4 ]and the CUF at W[]and the corresponding EAF usage is{ ]

c) Summary of Screening for Pressurizer Surge Nozzle: Since all values of Uen are less than 1.0 and are therefore acceptable for the period of extended operation, the Unit 1 and Unit 2 Pressurizer surge nozzle locations are screened out with respect to EAF.

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3.1.2 Screening of Hot Leg Surge Nozzle a) Unit 1:

1. Hot leg surge nozzle WOL cumulative fatigue usage for plant operation and without environmental effects was previously calculated and the maximum CUE reported as 0.103 [5, Table 2]. However, this result is located at an outside surface and, therefore, not susceptible to EAF.

ii. The inside surfaces, exposed to the water environment, need to include BAY effects. The highest normal fatigue usage on an inside surface, for the stainless steel components, is 0.023 [5, Table 2]. When factoring in the maximum Fen for stainless steel of 15.35 [14] for all the inside surface stainless steel locations, the maximum occurring environmentally-assisted fatigue usage (EAF) is 0.358 [5, Table 5].

b) Unit 2:

i. For the hot leg surge nozzle, a preemptive WOL was applied by AREVA during the SL2-17 refueling outage. The results of the WOL fatigue evaluation are documented in AREVA Document No: 32-904 1586, Revision 1 [3]. According to that document, the CUE at NW_18(f Jindth corresponding BAY usage is[ JAlso, the CUF at PWA_I80[ ]

[ ]and the corresponding BAY usage is[ ]

c) Summary of Screenin2 for Hot leg Surgie Nozzle: Since all values of Uen are less than 1.0 and are therefore acceptable for the period of extended operation, the Unit 1 and Unit 2 hot leg surge nozzle locations are screened out with respect to BAY.

3.1.3 Screening of Surge Line Piping a) Unit 1: The cumulative fatigue usage of the surge line piping, for plant operation and without environmental effects is 0.883 for Unit 1 [5, Table 1]. The maximum occurring Report No. 1301103.401 .R0 3-3 *j *,ku*uaIitgliyAsocaes I

environmentally-assisted fatigue usage (EAF), based on this analysis is 13.554 for Unit 1

[5, Table 4].

b) Unit 2: The cumulative fatigue usage of the surge line piping, for plant operation and without environmental effects is 0.654 for Unit 2 [5, Table 1]. The maximum occurring environmentally-assisted fatigue usage (EAF), based on this analysis is 10.039 for Unit 2

[5, Table 4].

c) Summary of Screeninjg for Surge Line Piping: All three elbows of the surge piping, for both units, fail to meet the allowable CUE of 1.0 when effects of EAF are included [5, Section 6.0].

Based on the information provided in Sections 3.1.1 through 3.1.3, the surge line piping (shown in Figure 4-1) is considered to be the bounding location and will need to be re-evaluated using modem FEA methods (performed herein) in order to reduce conservatism for the subsequent flaw tolerance evaluation.

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4.0 PIPING INTERFACE LOADS AND THERMAL TRANSIENTS Piping interface loads and thermal transients developed herein will be used to perform stress and flaw tolerance evaluations.

4.1.1 Piping Loads The piping interface loads at the hot leg surge nozzle and pressurizer surge nozzle ends of the surge line were determined in a piping analysis in Reference 8, and select loads are reproduced in Table 4-1 below. In the piping analysis, the hot leg and pressurizer surge nozzle safe end to pipe interfaces are Point 30 and 200, respectively, as shown in Figure 4-1. Break loads are not considered since the main RCS loop is qualified for leak-before-break [9], and loads in the surge line resulting from branch line breaks elsewhere are expected to be much less than maximum seismic loads.

4.1.2 Thermal Transients Table 4-2 lists the transients from References 10 and 11 to be analyzed, along with the number of cycles during a 40-year plant life, the time history for pressure and temperature, and the calculated heat transfer coefficients. Transients for emergency and faulted conditions are not included since the resulting thermal transient stresses are intended to support the crack growth during normal plant operation of the ASME Code, Appendix L flaw tolerance evaluation [1].

The thermal stratification events for the surge line are listed in Table 4-3. Note that the stratification events are evaluated in a finite element analysis (FEA) as steady state thermal events. It is assumed that before stratification occurs, the surge line has a uniform temperature of 653°F-ATstrat and 440°F-ATstrat for high and low pressure stratification transients, respectively, as shown in Figure 4-2, where AT is the temperature difference of the top and bottom fluids during stratification.

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4.1.3 Thermal Anchor Movements The thermal and seismic anchor movements at the hot leg surge nozzle safe end are given in Reference 12 for St. Lucie Units 1 and 2 and shown in Table 4-4 below.

The hot leg thermal anchor movements (TAMs) given are for the normal operating temperature of 604°F. They are scaled for other temperature values using the following scaling factor:

(TmE -- 70)/(604 -- 70)

These TAMs are applied at the surge line elbow connecting the hot leg surge nozzle safe end.

4.1.4 Stratification Height Table 4-5 shows the calculation of stratification height. For heatup and cooldown, only five representative cases are shown in Table 4-5. Stratification height H, which is the height of the flowing fluid, can be calculated based on the H/d values in Table 4-5. The maximum heatup/cooldown stratification height is 0.3506*10.126 = 3.55". The design specification does not state that stratification occurs during transients other than heatup and cooldown. Hence it is concluded that stratification occurs only during heatup and cooldown. The stratification height is used as input in to the thermal stratification stress analysis discussed in Section 5.3.

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Table 4-1: Deadweight and Seismic Loads at the Surge Line Ends FX FY FZ MIX MY MZ Location Description lb lb lb ft-lb ft-lb ft-lb Dead Weight 17 1386 34 721 640 4898 Point 30 ____

UNIT OBEo) ~ -621 -414 -659 -1967 -1995 -3037 1 Point Dead Weight -17 646 -34 -3023 97 4449 200 OBE -559 -412 -686 -4189 -675 -4108 Dead Weight 16 1465 32 395 595 4409 Point 30 UNIT OBE -635 -423 -673 -2008 -2043 -3103 2 Point Dead Weight -16 742 -32 -2723 88 4059 200 OBE -573 -421 -702 -4282 -692 -4198 Notes:

(1) SSE loads are twice the OBE loads [10, 11]

(2) Table reproduced from Table 11 of Reference 8 IOOB 200 -

Y A - on 150B z

30 150 Source: Reference [8]

Figure 4-1: PIPESTRESS St. Lucie Units 1 and 2 Surge Line Model Report No. 1301103.401 .R0 4-3 *Sucira /citgriYAscats

Table 4-2: Surge Line Transients Time TPZR Tm.L TBULK P Cycles h

Description sec 0 F 0 F 0 F psia Btu/hr-ft2-0 F Plant Heatup 0.0 70 70 70 15 500 36

[10, 11, Figure 2] 20880 653 540 653 2250 126 Plant Cooldown 0.0 653 540 653 2250 500 126

[10, 11, Figure 2] 20880 70 70 70 15 36 Plant Loading (A) 0.0 653 540 653 2250 15000 456

[10, 11, Figure 3] 50.0 653 540 653 2250 456 50.1 653 544 544 2310 383 1120.0 653 604 604 2310 400 1120.1 653 604 653 2310 456 0.0 653 604 653 2250 15000 456 Plant Loading (B)

[10, 11, Figure 3] 0.1 653 604 604 2250 400 Plant Unloading (A) 0.0 653 604 653 2240 15000 664

[10, 11, Figure 3] 0.1 653 604 604 2240 583 30.0 653 604 604 2250 583 30.1 653 604 653 2250 664 200.0 653 604 653 1980 664 1100.0 653 540 653 2200 664 1100.1 653 540 540 2200 556 1200.0 653 540 540 2320 556 0.0 653 540 540 2250 15480 4794 Plant Unloading (t3) (I)

[10, 11, Figure 3] 0.1 653 540 653 2250 5726 10% Step Increase 0.0 653 595 653 2250 2000 591

[10, 11, Figure 4] 10.0 653 595 653 2250 591 10.1 653 595 595 2250 512 140.0 653 604 604 2340 519 140.1 653 604 653 2340 591 350.0 653 604 653 2250 591 350.1 653 604 604 2250 519 10% Step Decrease 0.0 653 604 653 2260 2000 712

[10, 11, Figure 4] 0.1 653 604 604 2278 625 10.0 653 604 604 2279 625 10.1 653 604 653 2280 712 150.0 653 595 648 2180 675 Report No. 1301103.401.R0 4-4

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Time TPZR Tiu. TBtJLK P h Description scCycles seF ° F psia Btu/hr-ft2 -0 F 150.1 653 595 595 2180 617 350.0 653 595 595 2180 617 350.1 653 595 653 2280 712 Reactor Trip, Loss of Flow, 0.0 653 604 653 2250 480 5726 Loss of Load 2.0 653 604 653 2250 5726

[10, 11, Figures 5 and 91 2.1 653 604 604 2250 5027 13.0 653 604 604 2400 5027 13.1 653 604 653 2400 5726 60.0 616 543 616 1760 5027 120.0 616 543 616 1720 5027 120.1 616 543 540 1720 4794 Hydrostatic Test, 3125 psia 100 15 10 46

[10, 11, Figure 6] 400 3125 98 100 15 46 Leak Test, 2250 psia, Up 0.0 100 400 200 46

[10, 11, Figure 7] 10800.0 400 2250 98 Leak Test, 2250 psia, Down 0.0 400 2250 200 98

[10, 11, Figure 7] 10800.0 100 400 46 Note:

(1) Also bounds end of reactor trip and loss of flow or load.

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Table 4-3: Stratification Transients for Surge Line [10]

Transient Btu/hr-ft2 -°F Cycles Heatup, AT =320°F, Low Pressure 203 75 Heatup, AT = 320 0F, High Pressure 254 75 Heatup, AT = 250°F, Low Pressure 202 375 H-eatup, AT =250°F, High Pressure 244 375 0

Heatup, AT =200° F, Low Pressure 198 400 Heatup, AT = 200°F, High Pressure 234 400 Heatup, AT =150°F, Low Pressure 190 500 Heatup, AT = 150°F, High Pressure 219 500 Heatup, AT = 90°F, Hot Standby 193 87,710 Cooldown, AT =320°F, Low Pressure 203 75 Cooldown, AT = 320°F, High Pressure 254 75 0

Cooldown, AT =250 F, Low Pressure 202 375 Cooldown, AT =2500 F, High Pressure 244 375 0

Cooldown, AT = 200 F, Low Pressure 198 400 Cooldown, AT = 2000 F, High Pressure 234 400 Cooldown, AT = 150°F, Low Pressure 190 500 Cooldown, AT = 150°F, High Pressure 219 500 Cooldown, AT = 90°F, Hot Standby 193 87,710 Note:

(1) Heatup and cooldown stratification transients can be evaluated together because the stratification conditions between them are identical.

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K*3" 600 4-4Q.. . . . .. . . .. . . . .. .. . . . .

tLl 400 a--

I-l 200 0

0 1 2 3 4 5 TIME (HOURS)

Figure 4-2: Surge Line High and Low Pressure Stratification During Heatup

[

]

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Table 4-5: Stratification Height Low P Low P Low P Low P High P High P High P High P Hot AT=320 0F AT=250 0 F AT=200 0F AT=150 0 F AT=320 0F AT=2500 F AT=200 0F AT=I5O0 F Standby ydd 0.2297 0.2389 0.2494 0.2636 0.1784 0.1852 0.1i924 0.2015 0.2236 H/d 1( ) 0.3055 0.3177 0.3317 0.3506 0.2373 0.2463 0.2559 0.2680 0.2974 H, in 3.0935 3.2174 3.3588 3.5501 2.4026 2.4942 2.5912 2.7137 3.0114 a2 ,,rad 1.9993 2.0427 2,0916 2.1567 1.7443 1.7795 1.8163 1.8621 1.9702 1

c N,rad 2.3425 2.3953 2.4550 2.5347 2.0351 2.0773 2.1215 2.1766 2.3071 A:./d 2 0.1362 0.1440 0.1530 0.1654 0.0949 0.1002 0.1058 0.1130 0.1311 A/ 0.2032 0.2146 0.2276 0.2455 0.1426 0.1504 0.1587 0.1693 0.1958 Wr/d 0.8413 0.8528 0.8653 0.8812 0.7657 0.7769 0.7884 0.8022 0.8333 Ri 13.7479 13.1463 12.5129 .11.7349 18.2186 17.4855 16.7652 15.9275 14.1735 d, ft 0.8438 0.8438 0.8438 0.8438 0.8438 0.8438 0.8438 0.8438 0.8438 2

A:,, ft 0.0970 0.1025 0.1090 0.1178 0.0676 0.0713 0.0753 0.0805 0.0934 Wy, ft 0.7099 0.7196 0.7302 0.7436 0.6461 0.6556 0.6653 0.6770 0.7032 AH, ft2 0.1447 0.1528 0.1621 0.1748 0.1016 0.1071 0.1130 0.1206 0.1394 2

g, ft/sec 32.174 32.174 32.174 32.174 32.174 32.174 32.174 32.174 32.174 T*F120 190 240 290 333 403 453 503 563 r,~.g, lbnmJt 61.7 60.3 58.98 57.58 56.079 53.462 51.16 48.80 44.84 T50o,, 0F 440 440 440 440 653 653 653 653 653 r, lbm/it3 51.76 51.76 51.76 51.76 36.95 36.95 36.95 36.95 36.95 Dr, lbm/ft3 9.94 8.54 7.22 5.82 19.12 16.51 14.21 11.85 7.89 2

g', ft/sec 6.18 5.31 4.49 3.62 16.65 14.37 12.37 10.31 6.87 Q, ft3/sec 0.0891 0.0892 0.0892 0.0892 0.0892 0.0892 0.0892 0.0891 0.08914 Q, gpm (2) 40 40 40 40 40 40 40 40 40 Notes:

(1) H/d is the height of the flowing fluid (H, in) divided by the inside diameter of the pipe (d, in).

(2) A bounding value of 40 GPM is assumed based on Reference 13.

Report No. 1301103.401 .RO 4-8 IllorlyAsocaes Ic

5.0 STRESS ANALYSIS A three-dimensional (3-D) finite element model for the surge line piping is developed using the ANSYS Finite element analysis (FEA) software package [2] which is also used to perform the stress analysis. The following is the order in which the FEA were conducted:

• Develop a finite element model of the surge line piping as shown in Figure 4-1.

• 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 loading. 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 with the highest total stress intensities, for locations in the base metal and the weld.

• Extract component stresses at critical base metal and weld locations and store them in computer files.

5.1 Finite Element Model The hot leg surge line piping is 12" schedule 160 [10, pp.82, 93, 94], with an outside diameter of 12.75 inches and thickness of 1.3 12 inches. The dimensions and elevations of the surge line piping are provided in the FPL engineering specifications [10, pp.52, 82, 93, 94].

The three dimensional (3-D) model includes the entire surge line from the elbow following the safe end of the hot leg nozzle up to and including the safe end of the pressurizer surge nozzle.

The Unit 1 and 2 surge lines have identical geometric layout, but slightly different placement and stiffness of spring hangers. Since the environmentally-assisted fatigue analysis showed Unit 1 to

be bounding [5], the finite element model developed herein was based on Unit 1 inputs, but its analysis results will serve as a conservative envelope for the surge line of both units. The finite element model is shown in Figure 5-1.

The 3-D model is constructed using the ANSYS 8-node structural solid elements, SOLID45 for stress and SOLJD70 for thermal. The spring hangers attached to the surge pipe are modeled using the ANSYS longitudinal spring-damper structural element, COMBIN14. The spring constants for all the three hangers are listed in Table 5-2 [17], with zero damping, because no dynamic effects are explicitly analyzed.

5.2 Material Properties for Stress Analysis The surge line is fabricated from cast austenitic stainless steel (CASS), SA-3 51 Grade CF8M

[11, p. 3 1]. Material properties for the surge line piping are obtained from the 2001 ASME Boiler and Pressure Vessel Code with 2003 Addenda [15], and tabulated in Table 5-1. The ASMIE Code specifications for the materials used in this analysis are identical to those in the ASTM specification for the purposes of this analysis.

Note that during the Unit 1 pressurizer replacement, the existing vertical surge line spool piece to the 90 degree elbow was replaced with a one piece 90 degree elbow made of SA-403 WP347

[16, p.26]. As far as this linear elastic FEA is concerned, the SA-403 WP347 material has the same thermal and mechanical properties as SA-35 1 CF8M [11, p.3 1]. Therefore, the entire surge line is modeled using the same material properties in the FEA.

5.3 Thermal and Stratification Stress Analyses Including Pressure and TAM The thermal and stratification stress analyses were performed using the temperature distributions computed in the thermal analyses for various time steps of each transient. The thermal transient stress analyses include the scaled TAMs and internal pressure.

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The stratification stress analyses include the scaled TAMs. The applied temperatures for the Stratification Low Pressure Event at A320°F are shown in Figure 5-3. In this figure, the stratification height discussed in Section 4.1.4 is the boundary between the hot and cold fluids.

The high pressure cases and the hot standby case use an upper bound pressure of 2235 psi and the low pressure cases use a lower bound pressure of zero.

5.3.1 Deadweight Analysis For the deadweight analysis, the weight to be considered includes the weight of steel, water in the pipe, and the external insulation. The combined weight is evenly distributed within the pipe material as an equivalent weight density, which is calculated to be 0.393 lb/in 3.

Spring stiffness [17] and preloads [8], as listed in Table 5-2, are also included in this load case.

The preloads are the minimum preloads among Units 1 and 2, which are conservative and bounding since lower spring preloads will yield greater reaction forces and moments for the piping (i.e., the lower preloads compensate lesser of the deadweight).

A vertical acceleration of lg is applied in the Y-direction to simulate the equivalent gravitational force.

5.3.2 Mechanical Boundary Conditions For all analyses the pilot node attached to the pressurizer surge nozzle end is constrained in all translational and rotational degrees-of-freedom (DOF), while the surge line piping itself is allowed to expand in the radial direction.

For the thermal transient and stratification analyses, the pilot node attached to the hot leg side elbow end is the location where scaled TAMs are applied.

For the Deadweight case, the pilot node attached to the hot leg side elbow is constrained in all DOF. The mechanical boundary conditions for this analysis are shown in Figure 5-2.

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5.4 Stress Analysis Results A total of twenty stress paths are chosen at high stress locations for the purpose of reporting the stress analysis results. As can be seen from Figures 5-6 and 5-7 regions of high stress were observed around the elbow above the hot leg for the thermal stratification transient and another elbow at an intermediate portion (between Nodes 100 and 100B in Figure 4-1) for the normal thermal transient of the surge piping. Comparing the stresses in Figures 5-6 and 5-7, it can be seen that Node 30 is the bounding location since the stresses are much larger for the thermal stratification transient. This location will therefore be used for the subsequent flaw tolerance evaluation in Section 6. Paths are chosen at both weld and base metal locations for subsequent flaw tolerance analyses at both elbows. Weld locations are limiting with respect to EAF, because they have fatigue strength reduction factors that may contribute to any postulated fatigue crack initiation effects. Base metal locations are selected for subsequent evaluation of thermal aging of CASS material components.

Four (4) stress paths are chosen in the base metal portion of the elbow at the hot leg surge nozzle end of the surge line and are based on the inspection of results from the thermal and mechanical stress analyses. Four (4) paths are also chosen at the weld attaching the elbow to the horizontal portion of surge pipe to capture the high stresses at the weld location.

Similarly, four (4) stress paths are chosen in the base metal portion of the elbow at the middle of the surge line and are based on the inspection of results from the thermal and mechanical stress analyses. Eight (8) stress paths are chosen at the welds attaching the elbow to the surge pipe to capture the high stresses at the weld location.

Table 5-3 is a summary of the stress paths and nodes. The paths are shown in Figures 5-4 and 5-5. Component stresses for the thenrmal transients and stratification transients are extracted for these paths in a local Toroidal coordinate system with stress components identified as Radial (SX), Axial (SY), and Hoop (SZ). The representative stress intensity contour plots for the Heatup transient and stratification low pressure event at A320°F are shown in Figures 5-6 and 5-7 respectively. The transient stresses are used in Section 7.0 for the crack growth analyses.

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Figure 5-78 shows the stress intensity factors for axial and hoop flaws in Stress Path 4 for Hot Standby Stratification at A90°F (Transient 21). Per the crack growth analysis in Section 7.0, Hot Standby Stratification at A900 F is the transient with the greatest contribution to crack growth, and Stress Path 4 is the most limiting stress path in the bounding elbow. The results in Figure 5-8 show that the stress intensity factors for circumferential flaws are higher than the stress intensity factors for axial flaws. As such, circumferential flaws are expected to grow faster than axial flaws. This validates the technical approach and assumption that crack growth of an axial flaw is bounded by the circumferential flaw for the St. Lucie surge line geometry, as outlined in Section 2.0.

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Table 5-1: Material Properties Mean Coefficient ougs of Specific Heat, Density Dcrpin Temperature Modulus, Conductivity, k Diffusivity,Cp(bn)

(0 F) E x 106 (BTU/hr.ft-°F)(l) d (ft~lhr) (TUbmF (s) Expansion, (T/b-F a x106 (1/ 0F) 70 28.3 8.5 8.2 0.139 0.121 100 28.1 8.6 8.3 0.140 0.121 150 27.9 8.8 8.6 0.142 0.124 200 27.6 8.9 8.8 0.145 0.124 250 27.3 9.1 9.1 0.147 0.127 300 27.0 9.2 9.3 0.150 0.127 A-351 Grade 0.283(2) 350 26.8 9.3 9.5 0.152 0.128 CF8M (6rlN-M) 400 26.5 9.5 9.8 0.155 0.129 033' 450 26.2 9.6 10.0 0.157 0.130 500 25.8 9.7 10.2 0.160 0.130 550 25.6 9.8 10.5 0.162 0.133 600 25.3 9.8 10.7 0.165 0.133 650 25.1 9.9 10.9 0.167 0.133 700 24.8 10.0 11.2 0.170 0.135 0

Notes 1. Convert to BTU/sec-in- F for input to ANSYS.

2. Density of Steel only. Used in all analyses except Deadweight case.

3. Density of Steel + Water + Insulation used in Deadweight case.

Table 5-2: Hanger Spring Constants and Preloads Stiffness Spring Spring Hanger Constant, Preloads (ib)

No. 117] K (lbs/in) [8]

1171 H-1B 4532 2382 H-2B 4532 2294 H-3B 4008 2930 Report No. 1301103.401.R0 5-6

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Table 5-3: Linearized Stress Paths and Nodes Path Number Inside Node Outside Node Elbow Location P1 7776 8144 P2 10080 10448 Hot Leg Elbow P3 12368 12608 P4 5344 5728 P5 42366 42674 Intermediate P6 44640 45008 Elbow P7 46928 47168 P8 39904 40288 Weld Location P9 7656 8024 P10 9960 10328 Hot Leg Elbow _______

P11 12248 12488 P12 5209 5608 P13 38064 38156 P14 38640 38732 P15 39212 39272 Intermediate P16 37453 37552 Elbow P17 42216 42584 P18 44520 44888 P19 46808 47048 P20 39769 40168 Report No. 1301103.401.R0 5-7

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Figure 5-I: ANSYS Finite Element Model of the Surge Line Report No. 1301103.401 .R0 5-8 Sf " '~tr'/ * ,"'c I

Figure 5-2: Applied Mechanical Boundary Conditions for Deadweight Case 5-9

~j~SkucftuW lutegrity Asso~s, k~

Report No. 1301103.401 .R0 -

TYPE NUN alT; a~JL 12 11 ii 22 22333. 333 404. 444 155. 556 226. 667 297. 778 368. 889 440 frSLSU1RGLL1N Finite Ulcmnt Mr:k .. ..

Figure 5-3: Applied Temperatures for the Stratification Low Pressure Event at A320 0 F Rieport No. 1301103.401 .R0 5-10 IidwgrltyAsoitsI

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• ~ .......... .. Pe.. nselected r*\\stratlon np ur po ups Flt1-..... Stress path De1tniatiolS for the Elbow O~ ' t b o e

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S._. 55 TTh~F*,-4980 1929.87 10759.7 19589.5 248419 323 4.37249.2 6344.78 15174.6 44.324.4141 Note: At Time, T =24880.0 seconds of Plant Heatup Transient, units in terms of psi Figure 5-6: Stress Intensity Contour Plot for a Typical Thermal Transient R~eport No. 1301103.401 R0 5-13 hl~

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384.434 23969.7 4/55 1140.2 975 12177.1 35762.3 59347.6 82932.9 106518 0

Figure 5-7: Stress Intensity Contour Plot for the Stratification Low Pressure Event at A329 F report No. 1301103.401 .R0 5-14

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80 A Circumferential Flaw A 60 A

50 A

• "40 A A

• 30 A

20 A 10 A

0 0 0.2 0.4 0.8 1 1.2 Depth (in)

Figure 5-8: Stress Intensity Factors of Axial and Circumferential Flaws for Stress Path 4 Hot Standby Stratification Event 5-15 ~cIwwI I#iwgdly Msociate~ Inc~

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6.0 ALLOWABLE FLAW SIZE EVALUATION 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, the flaw evaluation procedures of IWB-3640 are used when applicable in the determination of the allowable flaw sizes. As discussed in Section 5.4, axial stresses are bounding so only circumferential flaws will be considered in the allowable flaw size evaluation.

The base metal for the surge line piping and elbows is cast austenitic stainless steel (CASS), and the weld metal is stainless steel. As such, the base metal and weld metal must be evaluated separately in the flaw tolerance evaluation. Flaws will be postulated in the base metal and weld metal at the critical surge line location (per Reference [5], Node Point 30 in Figure 4-1), which is the elbow directly attached to the hot leg surge nozzle.

For the weld metal, guidance for calculation of allowable flaw sizes is provided by ASME Code Section XI, Appendix L (L-3 000) [1]. 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 [1], which contains the screening criteria procedure to determine the applicable failure mode and evaluation for allowable flaw sizes with appropriate structural factors.

For the CASS base metal, there is no guidance in the ASME Section XI, IWIB-3640 for determination of allowable flaw sizes for materials with ferrite levels greater than or equal to 20%. As such, a probabilistic fracture mechanics (PFM) analysis is used tO determine allowable flaw depths as a function of flaw length for several failure probabilities. The PFM analysis is performed using the CAS SPAR software. Reference 25 provides the details of the fracture mechanics and probabilistic models that are programmed into the CASSPAR software. Using component-specific material property inputs, CASSPAR determines allowable flaw depths, a, for Report No. 1301103.401 .R0 6-1 *3SIIiIIIr' lt g~ril soits n

a part through-wall circumferential flaw for several flaw lengths, (represented by the flaw angle 20), for several failure probabilities. The critical depth for tearing instability is determined where the J-integral tearing modulus (J-T) curve due to the applied loads intersect the material J-T curve.

A subsequent crack growth evaluation presented in Section 7.0 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, 2001 Edition with Addenda through 2003 [1] is used for the determination of the allowable flaw sizes when applicable.

6.1 Interface Loads The following loads at the critical location (Node Point 30 in Figure 4-1) are evaluated:

• Deadweight (DW): Deadweight is obtained from Table 4-1.

• Operating Basis Earthquake (OBE): The OBE load is obtained from Table 4-1.

• Safe Shutdown Earthquake (SSE): Per the design specifications [10, 11], SSE loads are twice the OBE loads.

• Thermal Expansion (TE): The thermal expansion loads from thermal transients and thennal stratification events are obtained from Tables 4-2 and 4-3. The highest thermal expansion stresses from the thermal loads are shown in Table 6-3 as secondary bending stress. These loads include the thermal and seismic anchor movements from Table 4-4.

• Pressure: The pressure during thermal transients or thermal stratification events is obtained from the design specifications [10, 11]. The thermal expansion loads from Section 4.2 include the operating pressure during thermal transients or thermal stratification events. Hence that load has a combination of secondary thermal expansion stress plus pressure stress.

For the weld metal analyses, the membrane stress due to pressure and the bending stress from thermal expansion are required as separate entries for the evaluation per ASME Code Section XI, Report No. 1301103.40 1.R0 6-2 *gI* Asow",s Ic

Appendix C [1] procedures. Knowing the operating pressure, the membrane stress due to pressure can be calculated. The resulting pressure stress can then be subtracted from the stresses due to the thermal expansion plus pressure load provided in Section 4.0 to obtain the bending stress due to thermal expansion only.

For the base metal analyses, the above procedure cannot be applied since the forces and moments have to be directly input into the GAS SPAR software tool that is being used to calculate the allowable flaw size for the CASS material. Hence the piping loads from the combination of thermal expansion plus pressure loads from Section 4.0 are conservatively entered as the bending loads due to thermal expansion. In addition, the operating pressure is entered to determine the membrane stress. Hence the pressure stress is doubly applied and will lead to conservative estimate of the allowable flaw sizes.

The interface loads are summarized in Table 6-3. The loads input to CASSPAR are shown in Tables 6-5 and 6-6.

6.2 Load Combination The load combinations represent the normal conditions (Service Level A), upset conditions (Service Level B), and emergency/faulted conditions (Service Levels C and D). Load combinations are discussed in the extended power uprate documents [10, 11] and described below. Individual loads that define the load combinations are provided in Table 5-3.

Service Level A This load combination includes the internal pressure under normal operating conditions, deadweight, and highest thermal expansion load, whether due to Service Level A thermal transient or thermal stratification event.

Service Level B This load combination includes internal pressure under upset conditions, deadweight, OBE, and highest Service Level B thermal transient load.

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Service Levels C and D These load combinations include internal pressure under emergency/faulted conditions, deadweight, SSE, and highest Service Level C/D thermal transient load.

6.3 Material Properties for Allowable Flaw Size Determination Per References 10 (page 32) and 11 (page 28), the base metal for the surge line including elbows is ASTM SA-35 1 Grade CF8M (centrifugally cast) as shown in Table 6-4. The material property inputs for CASSPAR including the percent chemical compositions of nickel (Ni), silicon (Si),

carbon (C), nitrogen (N), manganese (Mn), and the value of delta ferrite of the CASS base metal are obtained from the Tables 2-1 .3a and 2-1 .3b of Reference 22 and shown in Table 6-6.

Between the ladle and check analysis, the chemical composition with the highest delta ferrite is used to yield more conservative allowable flaw sizes and bound the allowable flaw sizes of compositions with lower delta ferrite. A delta ferrite standard deviation of 2.5% is used for all analyses.

Material strength data for the base metal (i.e., Heat A2562 for Unit 1 and Heat S-250 for Unit 2, per Reference 22) at room temperature is obtained from the Certified Material Test Reports (CMTR) [19], and linear extrapolation based on data from ASME Code Section II, Part D [15] is used to obtain values at higher operating temperatures. The yield stress, oy, and ultimate stress, Gu, at the operating temperatures are shown in Table 6-4. As discussed in the previous paragraph, the calculated material properties are applicable to both the base metal and weld metal.

For the weld metal, the supplied references including the design specifications do not specify the weld material for the bounding elbow. SA-2 13, TP304 was assumed for the weld material, which is consistent with previous analyses [29, Table 1].

6.4 Welding Process The reference materials including the design specifications [10, 11] and welding data [20] do not specify the welding process for the bounding elbow. In this evaluation, shielded metal arc Report No. 1301103.401 .R0 6-4

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(SMAW) or submerged arc (SAW) welding processes are conservatively assumed since they have relatively low toughness compared to other welding processes such as gas tungsten arc welding (GTAW). The failure mode for these welds is ductile tearing. Per ASME Code Section XI, Appendix C (C-6330), load factors (Z-factors) are introduced to account for the assumption of limit load as a failure criterion for the low toughness welds fabricated using SMAW or SAW welding processes in the determination of the allowable flaw size.

6,4.1 Z-factor Per ASMVE Code Section XI, Appendix C (C-6330) [1], Z-factors for SMAW or SAW welding processes in austenitic weld materials are calculated as follows:

Z --1.30 [1 +0.010 (NPS -4)] for NPS > 4 (6-1) where, NPS = Nominal pipe size (in)

For the 12-inch NPS surge line piping, the calculated Z-factor is 1.404.

6.5 Thermal Transient Loads The thermal loads for thermaal transients and thermal stratification events are obtained from the Section 4.0. Twenty-six thermal transients (Loads 1-26) and nine thermal stratification events (Loads 27-3 5) are analyzed. Service levels for each transient are defined in the extended power uprate documents [10, 11]. The loads for each transient, derived from Table 4-2 are shown in Table 6-2. Thermal expansion stresses are calculated from moments of each load, and the highest stresses for each service level are selected for the flaw tolerance evaluation. For Service Level A, High Pressure Stratification at A320°F (Load 27) results in the highest stresses for Units 1 and 2. For Service Level B, Reactor Trip/Loss of Flow/Load 2 (Load 20) results in the highest stresses for Units 1 and 2. There are no thermal transients classified as Service Levels C or D, per the extended power uprate documents [10, 11]. Hence for Service Levels C and D, the thermal load is the load under normal operating conditions (Load 39).

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6.6 Operating Conditions For Service Level A, the temperature at the bounding elbow is taken as the mean temperature in the surge line during the thermal stratification event. The highest stresses for thermal stratification events occur during High Pressure Stratification at A320°F. Per the design specifications [10, 11], the normal operating temperature at the pressurizer end of the surge line is 653 0F. The mean temperature in the surge line and hence, the bounding elbow is 653 - 320/2

=493 0F for Service Level A. For Service Level B, C, and D, the temperature at the bounding elbow is taken as the normal operating temperature at the hot leg end of the surge line, as shown in Table 6-1. Per Section 6.3, the temperatures are used to obtain the yield stress, a~y and ultimate stress, au, for the base metal and weld metal at different service levels, as shown in Table 6-4.

For Service Levels A and B, the internal pressure is the maximum pressure during the thermal load. For Service Levels C and D, the internal pressure is the normal operating pressure.

6.7 Allowable Circumferential Part Through-Wall Flaw The allowable circumferential part through-wall flaw is evaluated separately for the base metal and the weld metal.

6. 7.1 Base Metal The allowable part through-wall flaw in the base metal is evaluated using the software, CASSPAR [25], which employs elastic-plastic fracture mechanics methodology. The CAS SPAR inputs for pipe geometry, CASS material composition and loading are shown in Table 6-6. All CAS SPAR runs are performed using 5,000,000 trials with no correlation between toughness and tensile properties.

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Per Section 6.6, High Pressure Stratification at A320*F (Load 27) and Reactor Trip/Loss of Flow/Load 2 (Load 20) result in the highest stresses for Service Levels A and B, respectively.

The thermal load for normal operating conditions (Load 39) is evaluated for Service Level C/D.

Similar to the procedure in ASME Code Section XI, IWB-3 640/Appendix C, a fraction of the secondary thermal expansion stress is used in the load combination for materials that fail by ductile tearing [26]. The thermal load is divided by the safety factor in Appendix C of ASME Code Section XI. For Service Level C/D, the safety factor for Service Level D is conservatively used, as it yields higher forces and moments. Both Unit 1 and Unit 2 are evaluated separately.

Hence, there are six CASSPAR runs. The total forces and moments are combined using square root of the sum of the squares of the vector components.

CAS SPAR outputs the allowable flaw depths as a ratio of the wall thickness, a/t, and the flaw length is expressed as a ratio of the pipe circumference, e/nt, for various failure probabilities. Per Reference 23, the failure probabilities corresponding to the service levels are 10-6 for Service Level A, 10-5 for Service Level B, and 10-4 for Service Level C/D.

An initial flaw size is taken as 1/4 t and 1:6 for the aspect ratio [23]. For a pipe thickness of 1.312 inches and OD of 12.75 inches, the length of this postulated flaw is 1.968 inches or 6/it of 0.055.

The flaw length will get longer as a result of crack growth. The allowable flaw depth to thickness ratio (a/t) will be determined after crack growth in a subsequent analysis in Section 7.0.

6. 7.2 Weld Metal For flux welds such as SMAW/SAW, as assumed in Section 6.4, elastic-plastic fracture mechanics (EPFM) methodology described in ASME Code Section XI, Appendix C per the screening criteria should be applied [1]. The technical approach consists of determining the allowable flaw size (circumferential extent and through-wall depth) in the pipe that will cause the flawed pipe to fracture by ductile crack extension.

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

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• .'* '  :*:(6-2) and the stress ratio for membrane stress is calculated as:

Stress Ratio = ZSFmcrm (6-3) c-f where, c-rm = 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 for Service Level B, and deadweight + SSE for Service Levels C/D o-e = Secondary bending stress (ksi) from thermal expansion stresses for Service Level A (thermal transients or thennal stratification events) and for Service Levels B, C, and D (thermal transients only).

o-: = Flow stress (ksi), which is equal to the average of the yield strength, ay, and the ultimate tensile strength, cmu, (Table 6-4).

S = Safety factor for bending stress, depending on service level [1, Table C-2621].

SF,n = Safety factor for membrane stress depending on service level [1, Table C-2621].

Z =Z-factor for flux welds (Section 6.4.1).

Based on the calculated stress ratios in Equation (6-2), the allowable flaw depth-to-thickness ratio under stress due to combined loading for Service Levels A, B, C, and D is obtained from Table C-5310-1 through Table C-5310-4 of ASMVE Code Section XI, Appendix C [1]. Stress ratios in Equation (6-3) are used in Table C-5310-5 [1] to determine the allowable flaw depth-to-thickness ratio for membrane loading only. The allowable flaw sizes for combined loading and for only membrane loading are compared, and the smaller value is reported as the allowable flaw size for the weld:

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6.8 Allowable Flaw Size Determination Results The calculated Unit 1 and Unit 2 allowable flaw depths as a function of crack length in the weld metal are shown in Table 6-7 and Table 6-8, respectively, for all service levels. The postulated flaw length after crack growth will be used to determine the allowable a/t from these tables. The maximum allowable a/t in these tables is limited to 75% of pipe wall thickness.

The calculated Unit 1 and Unit 2 allowable flaw depths as a function of crack length in the base metal are shown in Table 6-9 and Table 6-10, respectively. The maximum allowable a/t in these tables is limited to 75% of pipe wall thickness consistent with ASME Code Section XI, Appendix C. The postulated flaw length after crack growth will be used to determine the allowable a/t from these tables.

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Table 6-1: Geometry and Operating Conditions Property Value Reference Pipe Specification 12-inch NPS [10, Figure 19]

Schedule 160 [11, Figure 19]

Pipe Thickness 1.312 in Standard Pipe Schedule Outer Diameter 12.75 in Standard Pipe Schedule Inner Diameter 10.126 in Standard Pipe Schedule Z-factor 1.404 Section 6.4.1 Normal Operating Pressure j 2250 psia [10, 11]

Normal Operating Temperature 604 0F [10, page 12]

at Hot Leg End of Surge Line [11, page 101 0

Normal Operating Temperature 653 F [ 10, page 12]

at Pressurizer End of Surge Line [ 11, page 10]

Maximum Temperature Difference in Surge Line during A 320 0F [10, page 12]

Thermal Stratification Event0') [11, page 10]

0 Mean Temperature in Surge Line during Thermal 493 F Section 6.6 Stratification Event0')

Note:

(1) Per Section 6.5, the highest stress for Service Level A occurs during High Pressure Stratification at A320*F (Load 27) for both Unit 1 and 2. Per Section 6.6, the calculated temperature during that thermal stratification event is 493 °F for Service Level A.

6-10 SIwc&LraI Iiilogry Associates, Inc.

Report No. 1301103.401.R0 -1

Table 6-2: Thermal Expansion Loads Load Service Level Maximum 0

Description [0 1 eprtr )[0 l Case [0 ]Tmeaue(F[0 1 1 Plant Heatup A 540 2 Plant Cooldown A 540 3 Plant Loading 1 A 604 4 Plant Loading 2 A 604 5 Plant Loading 3 A 604 6 Plant Loading 4 A 604 7 Plant Unloading 1 A 540 8 Plant Unloading 2 A 540 9 Plant Unloading 3 A 540 10 Plant Unloading 4 A 540 11 10% Step Upl1 A 604 12 10% Step Up 2 A 604 13 10% Step Up 3 A 604 14 10% Step Up 4 A 604 15 10% Step Down 1 A 604 16 10% Step Down 2 A 604 17 10% Step Down 3 A 604 18 10% Step Down 4 A 604 19 Reactor Trip, Loss of Flow, Loss of Load 1 B 604 20 Reactor Trip, Loss of Flow, Loss of Load 2 B 604 21 Reactor Trip, Loss of Flow, Loss of Load 3 B 604 22 Reactor Trip, Loss of Flow, Loss of Load 4 B 604 23 Hydrostatic Test Up A 604 24 Hydrostatic Test Down A 604 25 Leak Test Up A 604 26 Leak Test Down A 604 27 Low Pressure Stratification at A320°F A 440 28 High Pressure Stratification at A320°F A 653 29 Low Pressure Stratification at A250°F A 440 30 High Pressure Stratification at A250°F A 653 31 Low Pressure Stratification at A200OF A 440 32 High Pressure Stratification at A200°F A 653 Report No. 1301103.401.R0 6-11

-1 W c tualuorily Associates, InoP

33 Low Pressure Stratification at A150°F A 440 34 - High Pressure Stratification at A150°F A 653 35 - Hot Standby Stratification at A90°F A 653 39 Normal Operating Conditions A, B, C, D 604 Report No. 1301103.401.R0 6-12

-1 8*ucTd lr,*'*Associates, inc.0

Table 6-3: Interface Loads Primary Primary Secondary Maximum Moment Membrane Bending Bending 4 Pressure Load Case( ) Reference (ft-lbf) Stress Stress Stress (psia) (ksi (ksi) (ksi)

P Mi1 My Mz Gm Gb Ge Deadweight Table 4-1 - 721 640 4898 -0.489-OBE Table 4-1 - -1967 -1995 -3037 -0.405-SSE0)~ Section 6.1 -- - -- 0.8 10-High Pressure Stratification Section 6.5 2250 339297 89984 109252 4.34 -31.66 Unit 2 at A320°F( )

1 Reactor Trip/Loss of Section 6.5 2400 -7267 106427 -20708 4.63 -6.01 Flow!Load 2(3 Normal Section 6.5 2250 10031 124407 -1709 4.34 -7.88 Operating(4 )

Deadweight Table 4-1 - 395 595 4409 -0.437-OBE Table 4-I - -2008 -2043 -3103 -0.414-SSE01 ) Section 6.1 -- - -- 0.828-High Pressure Stratification Section 6.5 2250 339297 89984 109252 4.34 -31.66 Unit 2 at A320°F( )

2 Reactor Trip/Loss of Section 6.5 2400 -7267 106427 -20708 4.63 -6.01 Flow/Load 2(3 Normal Section 6.5 2250 10031 124407 -1709 4.34 -7.88 Operating( 4 )

Notes:

(1) Per the design specifications [10, 11], safe shutdown earthquake (SSE) loads are twice the operating basis earthquake (OBE) loads.

Report No. 1301103.401 .R0 6-13 ISECrIYAsoiteIfc

(2) Per Section 6.5, High Pressure Stratification at A320°F (Load 27) results in the highest stresses for Service Level A for both Unit 1 and 2. The maximum pressure for the thermal load is from the design specifications [10, 11].

(3) Per Section 6.5, Reactor Trip/Loss of Flow/Load 2 (Load 20) results in the highest stresses for Service Level B for both Unit 1 and 2. The maximum pressure for the thermal load is from the design specifications [10, 11].

(4) Per Section 6.5, there are no thermal transients classified as Service Levels C or D, and therefore, the thermal load is the load under normal operating conditions (Load 39).

Report No. 1301103.401.R0 6-14 * /l

Tal P-:Mtrialprpertty foudig LineeElbow lure)

Property VaIue~3 ) Reference Base Metal Material SA-351 Grade Design Specifications CF8M [10, page 32]

(Centrifugally Cast)

Piece Number 505-03-1 Saturated Fracture Toughness Determination [22]

Heat A2562 Saturated Fracture Toughness Unit 1 Determination [22]

1 cyu @ 493°F (Service Level A)( ) 75.0 ksi CMTR [19]

Base Gy @ 493°F (Service Level A) (1) 34.4 ksi CMTR [19]

Metal c*u @ 604°F (Service Levels B, C, and D)( 2) 75.6 ksi CMTR [19]

oy @ 604 0F (Service Levels B, C, and D) (2) 33.7 ksl CMTR [19]

(and Weld Metal(4)) Base Metal Material SA-35 1 Grade Design Specifications CF8M [11, page 28]

(Centrifugally Cast)

Piece Number 751-107 Saturated Fracture Toughness Determination [22]

Heat S-250 Saturated Fracture Toughness Unit 2 Determination [22]

ou @ 493°F (Service Level A)(') 79.7 ksi CMTR [19]

ovy@ 493°F (Service Level A)(') 37.8 ksi CMTR [19]

cyu @ 604"F (Service Levels B, C, and D) (2) 80.3 ksi CMTR [19]

Gy @ 604"F (Service Levels B, C, and D) (2) 37.1 ksi CMTR [ 19]

Notes:

(1) Per Section 6.5, the highest stresses for Service Level A occur during High Pressure Stratification at A320°F (Load 27) for both Unit 1 and 2. Per Section 6.6, the calculated temperature for Service Level A during that thermal stratification event is 493 0F.

(2) Per Section 6.6, the temperature for Service Level B, C, and D is taken as the nornal operating temperature at the hot leg end of the surge line, 604 0F.

(3) Values for material properties are scaled from ASME code data for yield and ultimate strength using CMTR values as the minimums.

Report No. 1301103.401.R0 6-15 "*

(4) Per Section 6.3, the weld metal is assumed to have the same properties as the metal in the elbow. Per ASME Code Section IX, QW-153 [18], the weld metal must be stronger than the base metal of the adjacent components.

Report No. 1301103.401 .RO 6-16 kitr ft gr socae,/n.

Table 6-5: Loads for CASSPAR Base Metal Analyses Service Level SRF Pipestress Analysis CASSAR Input Thra odwt odForces (Ibs) -Moments (ft-lbs) Force Moment Highest Stresses FOrce Moment FX Fy F, M, My M, Load Combo (kips) (in-kips)

PSL1 A TE(3 2.7 2.3 -2131 -4805 14369 339297 89984 109252 TE +DW 5.43 1946.16 High Pressure A320°F DW - - 17 1386 34 721 640 4898 (Load 27)(2)

PSL1 B TE( 3) 2.4 2.0 -2671 2609 173281 -7267 106427 -20708 TTE+DW+OBE 8.60 680.41 Reactor Trip/Loss of Flow/Load 2 DW - - 17 1386 34 721 640 4898 (Load 20)/2) OBE - - -621 -414 -6591 -1967 -1995 -3037 PSL1C/D TE 1.3 1.4 1-1200 773 16246~ 10031 124407 -1709 TE+DW+SSE 14.29 1136.87 Normal Operating DW - - 17 1386 34 721 640 4898 (Load 39)(2) SSE(" - - 1242 -828 -1318j -3934 73990 -6074 PS2A TE( 3) 2.7 2.3 1-2131 -4805 143691339297 89984 109252 TE +DW 5.42 1940.56 High Pressure A320°F DW - - 16 1465 32 39 59 40 (Load 27)(2)j________j__________

P2B TE(3 2.4 2.0 1-2671 2609 173281-7267 106427 -20708 TE+DW+OBE 8.64 681.87 Reactor Trip/Loss of Flow/Load 2 DW - - j16 1465 32 39 59 40 (Load 201)2) OBE - - -635 -423 -673 -2008 -2043 -3103 PSL2C/D TE 1.3 1.4 1-1200 773 16246110031 124407 -1709 TE+DW+SSE 14.34 1136.68 Normal Opearting DW - - 16 1465 32 39 95 40 (Load 39)(2) SSE(" - - 1270 -846 -13461 -4016 -4086 -6206 Notes:

(I) Safe shutdown earthquake (SSE) loads are twice the operating basis earthquake (OBE) loads rio , 1].

(2) Per Section 6.5, High Pressure Stratification at A320°F (Load 27) and Reactor Trip/Loss of Flow/Load 2 (Load 20) result in the highest stresses for Service Levels A and B respectively. There are no thermal transients classified as Service Levels C or D, and therefore, the thermal load is the load during normal operating conditions (Load 39).

(3) Per Section 6.1, the pressure is double counted for the base metal analyses.

Report No. 1301103.401 .R0 6-17

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Table 6-6: Inputs for CASSPAR Base Metal Analyses I ~Force Moment IPressure( 1 ) A Ferrdte~2 )

Case R4 (inch) t (inch) I Std Dev Ni Si Mn C N CASSPAR Filename

...... _ _____ (kips) (in-kips) 4. ksia) (%)L.. 4 PSL1, Level A, High Pressure 6320°F 5.063 1.312 5.43 1946 16 ]. 2.25 25.64 2.5 9.06 0.72 0.63 0.04 0.055 PSL1 Strat A~xis PSL1, Level B, ReactorTrip/Loss of Flow/Load 2 5.063 1.312 8.60 680.41 2.40 25.64 2.5 9.06 0.72 0.63 0.04 0.055 PSL1ITrip B.xls PSL1, Level C/D, Normal Operating* 5.063 1.312 14.29 1136.87 j 2.25 25.64 2.5 9.06 0.72 0.63 0.04 0.055 PSL1LNoneCD~xls PSL2, Level A, High Pressure A320°F P512, Level B, Reactor Trip/Loss of Flow/Load 2 J5.063 5.063 1.312 1.312 5.42 8.64 1940.56 681.87 4 2.25 2.40 20.45 20.45 2.5 9.27 0.85 0.69 0.06 0.040 P5L2 Strat A.xls I5.063 2.5 9.27 0.85 0.69 0.06 0.040 PSL2_Trip B.xls PSL2, Level C/D, Normal Operating 1.312 14.34 1136.68 2.25 20.45 2.5 9.27 0.85 0.69 0.06 0.040 P5L2 None CD.xls Notes:

(1) Per Section 6.1, the pressure is double counted in the base metal analyses.

(2) The delta ferrite level is limited to 25% in Reference 23. However, slight extrapolation to ferrite levels greater than 25% is acceptable because of the small bandwidth between calculated and measured delta ferrite in Figure 5 of Reference 27.

Report No. 1301103.401I.R0 6-18

• rASOCI*.Sral.

Table 6-7: Allowable Part Through-Wall Circumferential Flaw Sizes in the Weld Metal for Unit 1 St. Lucie Unit 1 Allowable Flaw Size Calculation flOut data Dimensions

• NPS~

Ro Ri J tniom Z LA~J 6.38 j 5.036.... 31 12_2.5 L~i Loads Service P MX MY MZ a.~ aY, a*

Load Level (psi) llb-ft) (lb-ft) - (Ib-ft) -(ksi) (ksi) (ksi)

Deadweight A,8, CO -- 721 4 49 - 0.489 -

OBE 8 - -1967 -1995 -07 - 0.405 -

SSE C,Dg ...... -3934 *-399 *-6074 -- 0.809~ -

Stratification (Load 27) A 2250 339297 894 109252 4.341 - 31-663 Reactor Trip (Load 20) B 2400 -7267 106427 -20708 4.631 - 6.011 Normal Operating (Load 39) C, D 2250 10031 -124407 - 1709 4.341 ____- 7.883 Stress Ratios Service afm a*b ____e S - SJ op S Su af Stress Ratio Level (ki (ksi) (ksi) ____,F)__ ~ (ksi) (ksi) (ksi) Conto Merit A 4.341 0.489 31.663 - 2.7 - 2.3 493 34.4 750 54.7 0.477 0.30 B 4.631 0.894 6.011 2.4 2.0 604 33.7 '75.6 54.7 0.21_9 0.29 C 4.341 1.298 - 7.883 1.8 1.6 604 33.7 75.6 54.7 0.271 0.20 D 4.341 1.298 7.883 1.3 1.4 604 33,7 75.6 54.7 0.290 0.14 Results Allowable Flaw Depth-to-Thickness Ratio Ratio of Flaw Length to Pipe Circumference, IliD 0

o___________ 0.1 0.2 0.3 0.4 0.5 0.6 0.75 Service Flaw Length, 1.'(degree)

Level 0 ] 36 72 108 144 180 216 270 A 0.75 l 0.75 0.54 0.38 0.30 0.275 0.26 0.25 B 0.75 I 0.75 0.75 0.75 0.75 0.74 0.71 0.66 C 0.75 I 0.75 0.75 0.75 0.75 0.71 0.67 0.63 0 0.75 0.75 0.75 0.75 0.75 0.71 0.6S 0.60 Report No. 1301103.401 .R0 6-19 * :IFICZ*~ I*e~~ SOC~WIc

Table 6-8: Allowable Part Through-Wall Circumferential Flaw Sizes in the Weld Metal for Unit 2 St. Lucie Unit 2 Allowable Flaw Size Calculation input data Dimensions Ro Ri tnom ] Z FI~Z1 (in) 6.38 j (in) 5.06 J I (in) 1.31 j (i n 122.5j Lu~i L~i Loads Service p MX MY MZ ts,, *b

• Load Level (psi) (lb-ft) -(Ib-ft) (Ib-ft) (ksi) (kai) (ksi)

Deadweig~ht A, B,C, 0D 395 595 4409 - 0.437 --

OBE B - 20 2043 -3103 - 0.414 --

SSE C, D - -4016 -408 -620 - 0.827 --

Stratification (Load 27) A 2250 339297 894 10925 4.341 - 31.663 Reactor Trip (Load 20) B 2400 - -7267 106427 -20708 4.631 - 6.011 Normal Operating (Load 39) C, D 2250 10031 124407 -1709 4....4341

-- - 7.883 Stress Ratios Service ($ b ' SFm, SFb Top SI Su 5, f Stress Ratio Level (ksi) (ksi) (ksi) -_ __ ___ ('F) (ksi) (ksi (ksi) Combt Membt A 4.341 0.437 31.663 2.7 2.3 493 j 37.8 79:7 8. 0.443 j 0.2 B 4.631 0.851 6.011 2.4 2.0 64 37.1 80.3 58.7 0..203 0,.27 C 4.341 1.265 7.883 1,8 1.6 604 37.1 80.3 5j 0.252 0.* J D 4.341 1.265 7.883 1.3 1.4 604 37.1 80.3 58 0.6 01 Results Allowable Flaw Depth-to-Thickness Ratio Ratio of Flaw Length to Pipe Circumference, Lht D 0 0.1 o__________ 0.2 0.3 I 0.4 I 0.5 I 0.6 I0.75 Service Flaw Length, 1, (degree)

Level 0 36 72 108 144 180 216 270 A 0.75 0.75 0.66 0.48 0.38 0.342 0.32 0.31 B 0.75 0.75 0.75 0.75 0.75 0.75 0.74 0.68 C 0.75 0.75 0.75 0.75 0.75 0.72 0.69 0.66 D 0.75 0./5 0.75 0.75 0.75 0.72 0.67 0.63 Report No. 1301103.401 .R0 6-20 /ocilte,/IncY Assl*

Table 6-9: Allowable Part Through-Wall Circumferential Flaw Sizes in the Base Metal for Unit 1 Service Level A - Failure Probability of 106:

0.7500 0.7500 0.7500 0.6700 0.6000 0.5700 Service Level B - Failure Probability of 10s:

0.3 0.7500 U.5 0.7500 0,71 0.7500 Service Level C/D - Failure Probability of I0-4:

1.E-04 ] 0.7500 0.7500 0.7500 0.7500 0.7500 0.7500 6-21

-2

~j5SbvcturW lutogrilyAssociats~ k~

ReportNo. 1301 103.401.R0

Table 6-10: Allowable Part Through-Wall Circumferential Flaw Sizes in the Base Metal for Unit 2 Service Level A - Failure Probability of 10-6:

Service Level B - Failure Probability of 10-5:

0.5 0.7500 0,2 0.50 2 0.7500 Service Level CID - Failure Probability of 104~:

0.50 0.7500i 0-.7500ioo ul io ou:o

~j~SfrvobiW INto grHy Msociato~ Inc~

Report No. 1301 103.401 .R0 6-22

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7.0 CRACK GROWTH EVALUATION The crack growth evaluation is performed for both the base metal and the weld for both units at the elbow location, which is considered bounding for the surge line. The crack growth evaluation will be used in conjunction with calculated allowable flaw sizes from Section 6.0 to determine the required inspection interval for a postulated flaw in the surge line at the bounding elbow location. As discussed in Section 5.4, axial stresses are bounding so only circumferential flaws will be considered in the crack growth evaluation.

Fatigue crack growth is computed using linear elastic fracture mechanics (LEFM) techniques.

Representative fracture mechanics models are used to determine stress intensity factors (K) for each loading. The stress intensity factors for each type of load are computed as a function of crack depth and superimposed for the various operating states. Stresses that contribute to fatigue crack growth are compiled from stress analyses discussed in Section 5.0.

The weld metal and base metal are evaluated separately. For a postulated initial flaw of 25% of the original thickness, crack growth is simulated until the crack has reached the allowable flaw depth or a sufficiently long operating period, whichever comes first. The allowable operating time after crack growth is reported as the required successive inspection schedule based on the ASME Code,Section XI, Appendix L [1] procedures.

7.1 Design Inputs The following design inputs are used in this evaluation:

7.1.1 Geometry and Operating Conditions For both units, the bounding location for the surge line is the elbow directly attached to the hot leg surge nozzle. Per the design specifications [10, 11], the surge piping is 12-inch NPS schedule 160. Detailed geometrical information is provided in Table 6-1.

Report No. 1301103.401 .R0 7-1 *o

The maximum temperature and operating pressure range for the thermal transients and thermal stratification events are determined from Tables 4-2 and 4-3 and shown in Table 7-1.

Shown in Table 7-1, the projected number of cycles for the thermal transients and thermal stratification events over 60 years of operation are taken from the thermal cycle evaluation [30],

and the annual number of cycles are calculated for crack growth analysis.

7.1.2 Loads The loads are described in the following subsections.

7.1.2.1 Crack Face Pressure Crack face pressure is applied to all load combinations. A unit load of 1 ksi is scaled to the minimum and maximum operating pressures of each thermal transient or thermal stratification event.

• For thermal transients, the crack face pressures are set to the pressure range of the transient.

• The high pressure stratification events occur only at the normal operating pressure with no pressure fluctuations. Therefore, the minimum and maximum crack face pressures are both set to the normal operating pressure.

• The pressure for the low pressure stratification events are not specified in the design specifications [10, 11 ]. Therefore, the bounding pressure range is assumed, and the minimum and maximum crack face pressures are set to zero and normal operating pressure, respectively.

Report No. 1301103.401 .R0 7-2 * /I1C

7.1.2.2 Deadweight and Seismic Loads Deadweight stresses are taken from the finite element model discussed in Section 5.0. The deadweight is considered bounding for both units. The stress intensity factors for deadweight are extracted using the software package SI-TIFFANY [24], as discussed further in Section 7.2.1.

Operating basis earthquake (OBE) forces and moments are obtained from the design specifications [10, 11] and shown in Table 7-2. Safe shutdown earthquake (SSE) loads are twice the OBE loads, per the design specifications. The primary bending stresses for both units are calculated in Section 6.0, and the higher bending stresses for Unit 2 are bounding for both units and used in the crack growth analysis.

7.1.3 Thermal Transients and Thermal Stratification Events Thermal stress analyses of thermal transients and thermal stratification events have been performed using the finite element model given in Section 5.0. The stress intensity factors for thermal transients and thermal stratification events are extracted using the software package SI-TIFFANY [24], as discussed further in Section 7.2.1. Since all thermal transients and thermal stratification events have been analyzed individually in the finite element analysis (Section 5.0) and not scaled to representative cases, the stress scale factors for all cases are equal to 1. As shown in Table 7-1, the maximum temperature for thermal transients and thermal stratification events are taken from the design specifications [10, 11]. From the thermal cycle evaluation [30],

the 60-year projected cycles for Unit 1 are higher than those of Unit 2 and are therefore, considered bounding for both units.

Thermal stresses for normaal operating conditions are extracted at the start of thermal transients (e.g., time =0) that being under normal operating conditions.

The maximum rise time from all transients except heatup and cooldown is estimated to be 1400 seconds. The rise time for heatup and cooldown transients is 17,000 - 20,000 seconds. Although the rise time for heatup and cooldown is longer than 10,000 seconds, the AK is smaller for Report No. 1301 103.401 .R0 7-3 3IicriafiiUrlvsscite, nc

heatup and cooldown than the other transients, and hence, the contribution to crack growth from heatup and cooldown is expected to be insignificant compared to the other transients. To simplify the crack growth equation for calculations and to allow for comparative analysis of crack growth contributions between all thermal transients, 10,000 seconds was chosen as the bounding maximum rise time for all thermal transients except for Hot Standby Stratification (Case 21).

7.1.4 Residual Stresses Reproduced from NUREG 0313 [31, Figure 3 in Appendix A] and shown in Figure 7-2, the through-wall residual stress distribution in the axial direction is given as:

o" =r [1-6.91~ +8.69t~9 -0.48K -2.03~j (7-1) where, c-, = inside surface stress, assumed as yield stress (ksi) x = distance into pipe wall from inside diameter (in) t = wall thickness (in)

Exact material strength data for the weld metal is not available. However, per ASME Code,Section IX, QW-1 53, the weld metal must be selected to be stronger than adjacent components, namely the base metal. Therefore, the material strength data for the base metal from the Certified Material Test Reports (CMTR) [19] are applicable to the weld metal. Per Table 6-4, the bounding yield stress for the CASS base metal is 37.8 ksi. This is also used for the weld metal as explained above.

The pc-CRACK software [21] which is used for the crack growth evaluation, limits stress inputs to third order polynomials. Equation 7-1 is plotted and fitted to a third order polynomial using a Report No. 1301 103.401 .R0 7-4 Wruorrl lloril soi/eIc

thickness of 1.3 12 inches (Section 7.1.1) and the bounding yield stress. The axial residual stresses (in ksi) as a function of the distance from the inside surface (in inches) in third order polynomial form are:

G*(x) = 38.5 -214 x +/- 247 x2 - 76.6 x3 (7-2)

The residual stresses are only applied to the crack growth in the weld metal (Stress Paths 9-12).

The crack growth for the base metal (Stress Paths 1-4) excludes residual stresses.

7.1.5 Crack Growth Laws The CASS base metal and weld metal are analyzed separately using different crack growth laws, as discussed below.

7.1.5.1 Base Metal Crack growth in CASS, Grade CF8M is similar to wrought Type 316 stainless steel. As such the crack growth is calculated using the reference fatigue crack growth rate curves for wrought austenitic stainless steels in deaerated water [28]:

(da/dN) = Co(AK)" (7-3) where, da/dN = growth rate, in/cycle Co = material parameter for loading rate and environment AK = stress intensify factor range, ksi ina" Discussed further in Section 7.2.1, the stress intensity factor range is defined as the difference between the maximum applied stress intensify factor, Kmax, and the minimum applied stress intensity factor, Kmj,, for a load combination:

Report No. 1301103.401 .R0 7-5

• S g* ~l

AK =Kmax -Kmin (7-4)

The material parameters for Type 316 stainless steel are calculated as:

CO C SR ST SENV (7-5) where, C 4.43x10o-7 for AK > AKth

= 0 for AK < AKth n = 2.25 SR - l+/-e8 02

. (R- 0.748) for0_<R< 1.0

= 1 for R <0

= e-2516/TK for 300 *F_<T*_ 650 0F ST

= 3.39 x 105e(-2516rrK 0"03 01TK) for 70 0 F _<T _<300 0F SENV R = Kmin / Kmax TK = [(T -32)/1.8 +273.15], K TR = Rise Time, s T - Temperature, 0F A~th = 1.0 ksi-in'*'

The crack growth is summed on a yearly basis, irrespective of the sequence of crack growth contributions of the individual load combinations. The crack growth parameters for Low Pressure Stratification Events (Cases 13, 15, 17, and 19) and Hot Standby Stratification (Case

21) are examined separately in the subsequent sections. For all other transients (Cases 1 - 12) and High Pressure Stratification Events (Cases 14, 16, 18, and 20), the bounding values are calculated as follows. The maximum rise time, TR, is 10,000 seconds. The highest temperature, T, for all thermal transients and thermal stratification events is 653 0F. The maximum R-ratio is Report No. 1301103.401 .R0 7-6

• d11r0.

0.9 for a conservatively higher crack growth rate. Conservatively, the material parameters corresponding to the highest crack growth rate are calculated and applied as bounding for all load combinations except Low Pressure Stratification Events and Hot Standby Stratification:

TK = [(653 - 32)/1.8 + 273.15] =618 SR= 1+e8 02. (0.9 -0 7.48 ) = 4.38 ST= e-2516 /6 18 =0.0171 SINV = (10000)0.3 = 15.8 Co = (4.43 x 10-7) (4.38) (0.0171) (15.8) = 5.24 x 10-7 This calculated coefficient Co is the base case, which is used to scale the multipliers for the remaining stratification events.

7.1.5.1.1 Low Pressure Stratification Events Per Table 7-1, the Low Pressure Stratification Events (Cases 13, 15, 17, and 19) occur at a temperature of 440 0F, which is lower than the temperatures of the other transients. From initial analysis of the stresses, Low Pressure Stratification Events have a lower R-ratio, as well, and hence, a bounding R-ratio of 0.5 is used. The maximum rise time, TR, is 10,000 seconds.

Hence, the material parameters for Low Pressure Stratification Events are:

TK = (440 - 32)/1.8 + 273.15 = 500 SR = l+e 8 °2 (°5 -0 748

. ) 1.13 ST =e-2516 /500 =0.0065 SEN~V= (10000)0.3 = 15.8 Co =(4.43 x 10-7) (1.13) (0.0065) (15.8) --5.20 x 1-Report No. 1301103.401 .R0 7-7 * * ,O

A ratio between the coefficient for Low Pressure Stratification Events and the coefficient for the base case is calculated and accounted for by applying a multiplier for all stress intensity factors in the Low Pressure Stratification Events load combination.

7.1.5.1.2 Hot Standby Stratification Hot Standby Stratification (Case 21) is a static event that occurs over a period of time without any changes in the temperature or pressure. Hence, the maximum rise time, TR, is closer to the lower limit of 1400 seconds, as stated in Section 7.1.3, than the bounding 10,000 seconds used for other transients. The temperature is 653 °F, and the maximum R-ratio is 0.9. The material parameters for Hot Standby Stratification are:

TK = (653 - 32)/1.8 +273.15 =618 SR = 1+e8 °2 (°9 0-748

. ) = 4.38 ST e-2516 /618 = 0.0171 Si~v = (1400)0.3 = 8.79 Co (4.43 x 10-7) (4.38) (0.0171) (8.79) = 2.91 x 10-7 A ratio between the coefficient for Hot Standby Stratification and the coefficient for the base case is calculated and accounted for by applying a multiplier for all stress intensity factors in the Hot Standby Stratification load combination.

7.1.5.2 Weld Metal Per Section 6.3, the weld metal is assumed to be SA-213 TP304. Crack growth in Type 304 stainless steel is calculated using the ASME austenitic steel fatigue crack growth law in air, which is built into pc-CRACK [21] and based on equations in ASMiE Section XI, Appendix C (C-3210) [1]:

Report No. 1301103.401 .R0 7-8 * /~t'*' ' ~ ~

da o (X) (7-6) where, log 10 C = -10.009+/-+8.12x10-4T -l1.13x10-6T 2 ++/-1.02xl10 9 T3 S=I R<0

=l+l.8R 0 <R <0.79

= -43.35 + 57.97R R < 0.79 R= Kminl/Kmax AK = Kmax - Kmin n= 2.25 To account for the PWR water environment, a multiplier of two is applied to Equation 7-6 as suggested in Reference 26. In this calculation, the factor of two is accounted for by doubling the number of cycles for each load combination.

7.1.6 Postulated Initial Suiface Flaw The postulated initial flaw is a semi-elliptical circumferential flaw on the inside surface with an aspect ratio of 6-to-i, per ASME Section XI, Appendix L (L-3210) [1]. Per the EPRI technical basis document for CASS flaw tolerance evaluation [23, Section 4.0], the postulated surface flaw depth is 25%, as shown in Figure 7-1. For the 1.312 inch thick surge line, this is equivalent to an initial flaw that is 0.328 inches deep and 1.968 inches long. The postulated flaw characteristics are used in crack growth evaluation for both the base metal and the weld metal.

Report No. 1301103.401.R0 7-9 V *nw'curaf IIIIOUFII Ass*[o'ite, n

7.2 Calculations A two-step software approach is used to first calculate the maximum and minimum stress intensity factors using SI-TIFFANY [24] and then calculate the crack growth using pc-CRACK

[21].

7.2.1 Stress Intensity Factors The stress intensity factors (K) for the individual loads are determined using the model shown in Figure 7-1 which is incorporated in to both SI-TIFFANY and pc-CRACK. The individual stress intensity factors that contribute the nominal maximum applied stress intensity factor, Kmax, and the minimum applies stress intensity factor, Kmin, are summarized in the tabulations below. The stress intensity factor range, AK, is computed by taking the difference of the summed Kmax and Kmin. Kresidual and Kdeadweight are from constant loads and do not contribute to the AK range but affect the R-ratio (Kmin/Kmax), which accounts for mean stress effects. For the thermal transients, Kmax and Kmin are taken as the maximum and minimum over the duration of the transient. For the thermal stratification events, the Kmax includes thermal stratification event and the "Kmin" is without any stratification. Note that not all terms are included depending on the stress path (e.g., residual stresses are only included for the stress paths in the weld metal) and the service level (e.g., thermal stratification only occurs for Service Level A).

For the thermal transients (Cases 1-12), Kmax and Kmin are taken as the maximum and minimum over the duration of the transient, as shown. Note that residual stresses are only included for the stress paths in the weld metal (Stress Paths 9-12).

KmaxK in..

Kdeadweight Kdeadweight KUcrack face pressure Kcrack face pressure Ktrnsient max, thermal expansion + pressure Ktransient mai, thermal expansion + pressure Kresidual (Stress Paths 9-12 only) Uiresidual (Stress Paths 9-12 only)

Report No. 1301103.401 .R0 7-10 IIII*dI AsowaesInc

For the high pressure thermal stratification events (Cases 14, 16, 18, and 20) and hot standby stratification (Case 21), the Kmax contribution is the thermal stratification event, and the Kmin contribution is normal operating conditions without any stratification. Note that residual stresses are only included for the stress paths in the weld metal (Stress Paths 9-12).

Km ax Kmin Kdeadweight Kdeadweight K crack face pressure K crack face pressure Kstratification, thermal expansion + pressure Knormai operating conditions, thermal expansion + pressure Kresidual (Stress Paths 9-12 only) Kresidua (Stress Paths 9-12 only)

For the low pressure thermal stratification events (Cases 13, 15, 17, and 19), the low pressure is not specified. Therefore, the Kmax contribution is the thermal stratification event, and the "Kmin" contribution is set to zero, which conservatively assumes cold conditions. Note that residual stresses are only included for the stress paths in the weld metal (Stress Paths 9-12).

Kmax Kin .... ....

Kdieadweight Kdeadweight Kcrack face pressure Kstratification, thermal expansion + pressure Kresidual (Stress Paths 9-12 only) Kresiduaj (Stress Paths 9-12 only)

For the OBE and SSE, Kmax and Kmin are taken as the positive and negative of the seismic loads, respectively, and the crack face pressure is at normal operating pressure. Note that residual stresses are only included for the stress paths in the weld metal (Stress Paths 9-12).

Kinax Kmin 7-11

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Report No. 1301103.401 .R0 -1

Kdeadweight Kdeadweight Kcrack face pressure Kcrack face pressure KOBE/SSE, positive KOBE/SSE, negative Kresidual (Stress Paths 9-12 only) Kresidual (Stress Paths 9-12 only)

The software package SI-TIFFANY [24] is used to obtain tables of the Kmax and Kmin for various flaw depths and aspect ratios for the thermal transients, thermal stratification events, and deadweight. The stresses from the finite element analyses in Section 5.0 include the operating pressure and thermal expansion stresses, and hence, the pressure and thermal expansion stresses for SI-TIFFANY are input as zero, to avoid double counting. The pipe diameter and wall thickness are taken from Table 6-1.

As shown in Table 7-1, twelve thermal transients (Cases 1 - 12), and nine thennal stratification events (Cases 13 - 21) are analyzed, from the finite element analysis discussed in Section 5.0.

An additional deadweight case (Case 22) is also analyzed in Section 5.0. For the bounding elbow, Stress Paths 1 - 4 are in the base metal, and Stress Paths 9 - 12 are in the weld metal, as shown in Figure 5-5.

7.2.2 Crack Growth Per Table 6-1, pipe dimensions are:

Inside Radius: 5.063 in Pipe Thickness: 1.3 12 in The aspect ratio of the flaw is allowed to vary. Per Section 7.1.6, the dimensions of the postulated initial flaw are:

Crack Depth: 0.328 in Half Crack Length: 0.984 in Report No. 1301103.401 .R0 7-12 S/*O

For the base metal, the crack growth law in Equation 7-3 through 7-5 are implemented in pc-CRACK together with the CASS-specific crack growth parameters in Section 7.1.5.1. For the weld metal, the crack growth law represented by Equation 7-6 is used in pc-CRACK with all the material parameters already built into pc-CRACK. Per Section 7.1.5.2, the number of cycles for each load combination is doubled to account for the PWR water environment.

Table 7-land Table 7-2 show the design inputs for the thermal and seismic loads, respectively, from Section 7.1.2. Since each load is been analyzed separately in the stress analysis discussed in Section 5.0, the scale factors for all thermal and seismic loads are equal to 1. Scale factors are used to scale the 1-ksi unit pressure to the maximum operating pressures during the thermal loads. The residual stress distribution in Section 7.1.4 is added for crack growth in the weld metal only (Stress Paths 9-12).

The crack growth is performed for 60 years, and the time for the postulated flaw to grow beyond the allowable flaw size, if it occurs, is reported.

7.3 Crack Growth Analysis Results Table 7-3 reproduces the maximum allowable flaw length and depth for all service levels from the allowable flaw size determination (Tables 6-7 thru 6-10). The final flaw length-to-circumference ratio (0/rc) was obtained after crack growth by multiplying the final crack depth after crack growth in Table 7-4 by 6 to maintain the 6:1 aspect ratio of the postulated flaw.

Table 7-4 summarizes the crack growth results for a postulated circumferential flaw. For Stress Paths 1-4 in the base metal, a postulated circumferential flaw will grow to the allowable flaw size after the operating periods shown with the limiting operation period of 252 months (21 years) for Stress Path 4. For Stress Paths 9-12 in the weld metal, the crack growth of a postulated circumferential flaw will not grow beyond the allowable flaw size for at least 720 months (60 years) of operation. Therefore, the base metal is limiting, and the allowable operating period is at least 21 years. As noted in the technical approach in Section 2.0 and Report No. 1301103.401 .R0 7-13

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discussed in Section 5.4, crack growth of an axial flaw is bounded by the circumferential crack growth analysis results.

7.4 Inspection Interval Fatigue crack growth analysis of a postulated flaw in the weld metal and base metal at the bounding surge line location, namely the elbow directly attached to the hot leg nozzle, shows that the allowable operating period is at least 21 years. This analysis is bounding for both Unit 1 and 2.

Using the allowable operating period of 21 years, the successive inspection schedule for the surge line at St. Lucie, Unit 1 and 2 is ten years per guidelines of Table L-3420-1 of Appendix L and IWIB-2410 of ASME Code,Section XI [1].

Report No. 1301103.401 .R0 7-14 *

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Table 7-1: Thermal Transients and Thermal Stratification Events Case Description Maximum Pressure( 2 ) Projected Annual Temperature( 2 ) (ksi) Cycles(3 ) Cyclest4 )

(0 F) Min. Max. (60 years) (per year) 1 Plant Heatup 540 0 2.25 143 2.38 2 Plant Cooldown 540 0 2.25 143 2.38 3 Plant Loading A 604 2.25 2.31 900 15 4 Plant Loading B 604 2.25 2.31 900 15 5 Plant Unloading A 604 1.98 2.32 900 15 6 Plant Unloading B 540 1.98 2.32 900 15 7 10% Step Increase 604 2.25 2.34 390 6.5 8 10% Step Decrease 604 2.18 2.28 390 6.5 9 Reactor Trip, Loss of Flow, Loss of 604 1.72 2.40 115 + 1 + 8 2.07 Load 10 Hydrostatic Test 604 0 3.125 13 0.22 11 Leak Test Up 604 0.4 2.25 113 +3 1.93 12 Leak Test Down 604 0.4 2.25 113 +3 1.93 13 Low Pressure Stratification at A320°F 440 0 2.25 21 + 21(1) 0.7 14 High Pressure Stratification at A320°F 653 2.25 2.25 21 + 21(') 0.7 15 Low Pressure Stratification at A250°F 440 0 2.25 107 + 107(1) 3.57 16 High Pressure Stratification at A250°F 653 2.25 2.25 107 + 107C'O 3.57 17 Low Pressure Stratification at A200°F 440 0 2.25 114 + 114(1) 3.8 18 High Pressure Stratification at LA200°F 653 2.25 2.25 114 + 114(1) 3.8 19 Low Pressure Stratification at A150°F 440 0 2.25 143 + 143(1) 4.77 20 High Pressure Stratification at A150°F 653 2.25 2.25 143 + 143(1) 4.77 21 Hot Standby Stratification at A90°F 653 2.25 2.25 25085 + 836 25085(1)

(1) Thermal stratification events occur during both heatup and cooldown. Hence, the cycles are doubled to account for occurrence during both heatup and cooldown.

(2) The temperatures and pressures for thermal transients and thermal stratification events are taken from the design specifications [10, 11].

(3) The projected cycles for thennal transients and thermal stratification events are taken from the thermal cycle evaluation [5]. The projected cycles for Unit 1 are higher than those of Unit 2 and are therefore, considered bounding for both units.

(4) For the base metal, the crack growth analysis uses the annual number of cycles. For the weld metal, the annual number of cycles is doubled to account for the PWR water environment, as described in Section 7.1.5.2.

Report No. 1301103.401.R0 7-15  ;: id'/ soat,

Table 7-2: Seismic Loads Primary Projected Annual SrieOperating Moment Bending Cycles( 3 ) Cycles( 4 )

Load Case Ref. Pressure (lbf-ft) Stress(2 ) (per 60 (per Level (ksi) (ksi) years) year)

M1 My Mz 6 Operating Basis B [10, 11] 2.25 2008 2043 3103 0.414 40 0.67 Earthquake Safe Shutdown Eatqae) C,D [10, 11] 2.25 4016 4086 6206 0.827 40 0.67 (1) Per the design specifications [10, 11], safe shutdown earthquake (SSE) loads are twice the operating basis earthquake (OBE) loads.

(2) Per Section 7.1.2.2, the higher bending stresses from Unit 2 are considered bounding for both units.

(3) The projected cycles for OBE events are taken from the thermal cycle evaluation [30]. Projected SSE cycles are assumed to be the same in number as projected OBE cycles.

(4) For the base metal, the crack growth analysis uses the annual number of cycles. For the weld metal, the annual number of cycles is doubled to account for the PWR water environment, as described in Section 7.1.5.2.

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Table 7-3: Circumferential Allowable Flaw Size Final Half Aloae Locaion Stress Flaw Fa Loaton Path Length Fa 1 0.163 0.75 Base 2 0.159 0.75 Metal 3 0. 160 0. 75 4 0.162 0.75 9 0.062 0.75 Weld 10 0.057 0.75 Metal 11 0.072 0.75 12 0.061 0.75 Note:

(1) Allowable flaw depths determined from Tables 6-7 through 6-10.

Table 7-4: Circumferential Crack Growth Results Crack Growth Results Allowable Loain Stress Fia lwFinal Half Flaw Length Operating Path Dph(8/u) Period S [a/t] in in (O/n) months 1 0.7481 0.9815 2.9445 0.163 432 Base 2 0.7241 0.9500 2.8500 0.159 624 Metal 3 0.7327 0.9613 2.8839 0.160 384

______ 4 0.7394 0.9701 2.9103 0.162 252 9 0.2808 0.3684 1.1052 0.062 720 Weld 10 0.2608 0.3422 1.0266 0.057 720 Metal 11 0.3285 0.4311 1.2933 0.072 720 12 0.2807 0.3682 1.1046 0.061 720 Report No. 1301103.401.R0 7-17

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t a

x Initia a = 0.25t1

.niti aa2c=. 1/6 Figure 7-1: Semi-Elliptical Circumferential Flaw on the Inside Surface of a Cylinder Report No. 1301103.401.R0 7-18 *g/g

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8.0

SUMMARY

AND CONCLUSIONS The flaw tolerance of the bounding surge line location, namely the elbow directly attached to the hot leg nozzle at St. Lucie Units 1 and 2 has been evaluated and the required successive inspection schedule has been determined for a postulated 25% deep flaw with a length six times the depth per the requirements of ASME Code,Section XI, Appendix L.

The flaw tolerance evaluation consisted of determining the loads at the bounding location and performing a finite element stress analyses to determine stresses at the base metal and weld locations. The stresses are used to determine the allowable flaw sizes and perform a crack growth evaluation to determine how long it will take the postulated flaw to reach the allowable flaw sizes. It was established that axial stresses are bounding so only flaws in the circumferential directions were considered in the evaluation. The analyses indicated that for a postulated circumferential flaw in the base metal, it will take 21 years to reach the allowable flaw size. For a postulated circumferential flaw in the weld metal, it will take at least 60 years before the allowable flaw size is reached. As noted in the technical approach in Section 2.0, crack growth of an axial flaw is bounded by the circumferential crack growth analysis results.

Based on a comparison of geometry, material properties and applicable loads, the results of the detailed evaluation of the bounding location are applicable to the other weld locations on the surge line.

Report No. 1301103.401 .R0 8-1

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9.0 REFERENCES

1. ASME Boiler & Pressure Vessel Code,Section XI, 2001 Edition, with Addenda through 2003.

2. ANSYS Mechanical APDL and PrepPost, Release 14.5 (w/Service Pack 1), ANSYS Inc.,

September 2012.

3. AREVA Proprietary Document No. 32-9041586-001, "St. Lucie Unit 2 Hot Leg Surge Nozzle Weld Overlay Analysis," SI File No. 1301 103.204P.

4. AREVA Proprietary Document No. 32-9041583-000, "St. Lucie Unit 2 Pressurizer Surge Nozzle Weld Overlay Analysis," SI File No. 1301 103.204P.

5. SI Calculation Package No. 1301346.306, Rev. 0, "Surge Line Environmentally-Assisted Fatigue Analysis."

6. SI Report No. SIR-01-086, Rev. 4, "Position Document to Address GSI-190 Issues Related to Fatigue Evaluation for St. Lucie Units 1 and 2," SI File No. 13 00246.402.

7. FPL Letter NP-EPU-09-1635 to Gary Stevens (SI),

Subject:

"St. Lucie Plant - Unit 1 Extended Power Uprate Project, Transmittal of AREVA Letter AREVA-09-03261 ,"

August 4, 2009, SI File No. 090035 8.202.

8. SI Calculation Package No. 1301346.303, Rev. 1, "Surge Line Piping Normal Fatigue Analysis (Simplified Elastic-Plastic)".

9. Email attachment from Satyan-Sharma Tirumani (FPL) to Aparna Alleshwaram (SI), dated 12/13/2013, "NRC SER on EPU Section 2 2 2 RCS.docx," SI File No. 1301103.202.

10. FPL Document No. 19367-3 1-5 Rev. 14, Engineering Specification for Reactor Coolant Pipe and Fittings for FPL, SI File No. 1301103.207.

11. FPL Document No. 13172-31-5 Rev. 10, Design Specification for Reactor Coolant Pipe and Fittingsfor St Lucie Unit 2, SI File No. 1301103.207.

12. Email attachment from Satyan-Sharma Tirumani (FPL) to Aparna Alleshwaram (SI), dated 1/17/2014, "FPL- 14-4.pdf," PROPRIETARY, SI File No. 1301 103.205P.

13. CEOG Report CEN-3 87-NP, Revision 1-NP, PressurizerSurge Line Flow Strat~cation Evaluation, SI File No. 1301103.208.

14. NUREG/CR-5704 (ALNL-98/3 1), "Effects of LWR Coolant Environments on Fatigue Design Curves of Austenitic Stainless Steels," April 1999.

15. ASMIE Boiler and Pressure Vessel Code, Sections II - Part D, 2001 Edition with 2003 Addenda.

16. FPL Specification No. 04150, Rev. 0, "Replacement Pressurizer Installation for Unit 1 Pressurizer Replacement Project," SI File No. 1301103.201.

17. Spring Hangar Drawings, SI File No. 1301346.208.

Report No. 1301103.401 .R0 9-1 * * 'o

4a. Bergen-Paterson Pipe Support Corp., Unit 1 Pressurizer Surge Drawings

i. 8770-Hi-B, Rev. 2 ii. 8770-H2-B, Rev. 2 iii. 8770-H3-B, Rev. 4 4b. B.F. Shaw Co. Ebasco Services Inc., Unit 2 Pressurizer Surge Drawings

i. H-1B, Rev. 3 ii. H-2B, Rev. 2 iii. H-3B, Rev. 3

18. ASME Boiler and Pressure Vessel Code, Sections IX, 2001 Edition with 2003 Addenda.

19. FPL Transmittal CSI-14-07, "CMTRs of the Cast Austenitic Stainless Steel (CASS) Pipes and Piping Components for the Development of a CASS Aging Management Program for St. Lucie Units 1 and 2," from Soraya Benitez (FPL) to Aparna Alleshwaram (SI), May 21, 2014. SI File No. 1301079.201.

20. FPL Transmittal ENG-SPSL-14-0003, Transmittal of PSL - PZR Surge Line Piping Weld Data," from Paul Atkinson (FPL) to Aparna Alleshwaram (SI), January 28, 2014. SI File No. 1301103.207.

21. pc-CRACK 4.1 CS, Version Control No. 4.1.0.0, 2013.

22. SI Calculation No. 130 1079.401, Rev. 0, "Saturated Fracture Toughness Determination."

23. MaterialsReliability Program:Technical Basisfor ASME Section XI Code Case on Flaw Tolerance Evaluation of Cast Austenitic Stainless Steel (CASS) Piping (MRP-362). EPRI, Palo Alto, CA: 2013. 3002000672.

24. SI Report No. DEV1310.402, Rev. 1, "SI-TIFFANY 2.01 Users' Manual," September 2014.

25. SI Report No. 1000883.406, Rev. 1, "Probabilistic Reliability Model for Thermally Aged Cast Austenitic Stainless Steel Piping."

26. ASME Section XI Task Group for Piping Flaw Evaluation, ASME Code, "Evaluation of Flaws in Austenitic Steel Piping," Journal of Pressure Vessel Technology, Vol. 108, August 1986.

27. NUREG/CR-4513, Rev. 1, "Estimation of Fracture Toughness of Cast Stainless Steels During Thermal Aging of LWR Systems," May 1994.

28. W. J. Mills, "Critical Review of Fatigue Crack Growth Rates for Stainless Steel in Deaerated Water - Parts 1 and 2," EPRI MRP-2010 Conference and Exhibition: Materials Reliability in PWR Nuclear Power Plants, Colorado Springs, CO, June 28 - July 01, 2010.

29. SI Calculation No. 0800024.30 1, Rev. 2, "Material Properties for St. Lucie, Unit 1 Hot Leg Surge / Hot Leg Shutdown Cooling Nozzles Preemptive Weld Overlays and Hot Leg Drain Replacement."

30. SI Report No. 1301346.403, Rev. 3, "Thermal Cycle Evaluation for St. Lucie Units 1 and 2" (formerly SI Report No. SIR-01-102).

Report No. 1301103.401 .R0 9-2 "*'tr'"'

31. NUJREG-03 13, Rev.2, "Technical Report on Material Selection and Processing Guidelines for BWR Coolant Pressure Boundary Piping," U.S. Nuclear Regulatory Commission, January 1988.

32. NUIREG-1801, Revision 2, "Generic Aging Lessons Learned (GALL) Report," U. S.

Nuclear Regulatory Commission, December 2010.

33. NUJREG/CR-6583 (ANL-97/18), "Effects of LWR Coolant Environments on Fatigue Design Curves of Carbon and Low-Alloy Steels," March 1998.

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