ML15147A617

Palisades - Enclosure 1, Relief Request Number RR 4-21 Proposed Alternative, in Accordance with 10 CFR 50.55a(z)(2), Hardship Without a Compensating Increase in Level of Quality and Safety

ML15147A617
Person / Time
Site:	Palisades
Issue date:	05/22/2015
From:	Hardy J A Entergy Nuclear Operations
To:	Document Control Desk, Office of Nuclear Reactor Regulation
Shared Package
ML15147A613	List: ML15147A616 ML15147A617 ML15147A618 ML15147A619 PNP 2015-037, Appendix a, Computer Files Listing, File No. 1200895.306, Revision 1
References
PNP 2015-037
Download: ML15147A617 (105)

Text

ENCLOSURE1 ENTERGY NUCLEAR OPERATIONS, INC.PALISADES NUCLEAR PLANTRELIEF REQUEST NUMBER RR 4-21 PROPOSED ALTERNATIVE in Accordance with 10 CFR 50.55a(z)(2)

Hardship Without a Compensating Increase in Level of Quality and Safety1. ASME CODE COMPONENT(S)

AFFECTED

/ APPLICABLE CODE EDITIONComponents

/ Numbers:Code of Record:See Enclosure Table 1Pressure Retaining Dissimilar Metal Piping Butt WeldsContaining Alloy 82/182American Society of Mechanical Engineers (ASME) Section Xl,2001 Edition through 2003 Addenda as amended by10 CFR 50.55aASME Code Case N-770-1, "Alternative Examination Requirements and Acceptance Standards for Class 1 PWRPiping and Vessel Nozzle Butt Welds Fabricated with UNSN06082 or UNS W86182 Weld Filler Material With or WithoutApplication of Listed Mitigation Activities, Section Xl, Division 1"N-770-1 Inspection Item: A-2 and BDescription:

Unit / Inspection Interval:

Class 1 Pressurized Water Reactor (PWR) pressure retaining Dissimilar Metal Piping and Vessel Nozzle Butt Welds containing Alloy 82/182Palisades Nuclear Plant (PNP) / Fourth 10-Year IntervalDecember 13, 2006 through December 12, 20152. APPLICABLE CODE REQUIREMENTS The ASME Boiler and Pressure Vessel Code, Rules for Inservice Inspection of NuclearPower Plant Components, Section Xl, 2001 Edition through 2003 Addenda, as amended by10 CFR 50.55a.With the issuance of a revised 10 CFR 50.55a in June 2011, the Nuclear Regulatory Commission (NRC) staff incorporated, by reference, Code Case N-770-1.

Specificimplementing requirements are documented in 10 CFR 50.55a(g)(6)(ii)(F) and are listedbelow:A. Regulation 10 CFR 50.55a(g)(6)(ii)(F)(1) states "Licensees of existing, operating pressurized water reactors as of July 21, 2011 must implement the requirements ofASME Code Case N-770-1, subject to the conditions specified in paragraphs 1 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE (g)(6)(ii)(F)(2) through (g)(6)(ii)(F)(10) of this section, by the first refueling outageafter August 22, 2011."B. Regulation 10 CFR 50.55a(g)(6)(ii)(F)(3) states that baseline examinations forwelds in Code Case N-770-1, Table 1, Inspection Items A-1, A-2, and B, must becompleted by the end of the next refueling outage after January 20, 2012.The welds covered by this proposed alternative would be classified as Inspection Items A-2and B (described below) for which visual and essentially 100 percent volumetric examination, as amended by 10 CFR 50.55a(g)(6)(ii)(F)(4),

in part, are required, per NRCinterpretation.

ASME Code Case N-770-1, Table 1, Examination Categories, as amended by 10 CFR50.55a(g)(6)(ii)(F)

CLASS 1 PWR Pressure Retaining Dissimilar Metal Piping and Vessel Nozzle Butt WeldsContaining Alloy 82/182Parts Examined Insp Extent and Frequency of Examination ItemBare metal visual examination each refueling outage.Unmitigated butt weld Essentially 100% volumetric examination for axial andat Hot Leg operating A-2 circumferential flaws in accordance with the applicable temperature

(-2410) < requirements of ASME Section Xl, Appendix VIII, every five625-F (3290C) years. Baseline examinations shall be completed by the end ofthe next refueling outage after January 20, 2012.Bare metal visual examination once per interval.

Unmitigated butt weld Essentially 100% volumetric examination for axial andtemperature

(-2410)>

B circumferential flaws in accordance with the applicable tempere (-2410) ad B requirements of ASME Section Xl, Appendix VIII, every second525°F (274°C) and < inspection period not to exceed 7 years. Baseline examinations 580°F (304-C) shall be completed by the end of the next refueling outage afterJanuary 20, 2012.ASME Section Xl, Appendix VIII, Supplement 10, "Qualification Requirements for Dissimilar Metal Piping Welds," is applicable to dissimilar metal (DM) welds without cast materials.

3. REASON FOR REQUESTExaminations of the DM welds listed in Enclosure Table 1 of this request could not beperformed as required by ASME Code Case N-770-1, as conditioned by10 CFR 50.55a(g)(6)(ii)(F).

These DM welds are nominal pipe size (NPS) 2 inches and greater full penetration branchconnection welds installed in primary coolant loop piping. See the Enclosure for typicalconfiguration.

2 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE The relevant conditions for this proposed alternative are ASME Section XI Code CaseN-770-1, and 10 CFR 50.55a(g)(6)(ii)(F) items (1) and (3), which address performing therequired baseline examinations.

Regulation 10 CFR 50.55a(g)(6)(ii)(F)(1) requires that licensees implement the requirements of ASME Code Case N-770-1, subject to the conditions specified in paragraphs (g)(6)(ii)(F)(2) through (g)(6)(ii)(F)(10) of this section, by the first refueling outage after August 22, 2011.Regulation 10 CFR 50.55a(g)(6)(ii)(F)(3) requires that baseline examinations for welds inCode Case N-770-1 Table 1, Inspection Items A-1, A-2, and B be completed by the end of thenext refueling outage after January 20, 2012.Relief is requested from 10 CFR 50.55a(g)(6)(ii)(F) items (1) and (3) for performance ofrequired baseline volumetric examinations of the eight cold leg welds and one hot leg weldlisted in the Enclosure Table 1.HardshipThe PNP welds in question are outside of the current Performance Demonstration Initiative (PDI) demonstrated joint configurations.

That is, there currently are no PDI demonstrated volumetric techniques for this PNP weld joint configuration.

Volumetric examination techniques for this complex weld configuration (see Attachment 1, Figures 1 and 2) requiredevelopment and demonstration through the PDI qualification program.

Under the PDIprogram, representative mock-ups would need to be fabricated in accordance with the PDIspecimen fabrication program in order to develop examination techniques.

Qualification ofprocedures and personnel would be needed in order to reliably perform qualified examinations of these welds. It is estimated that these activities would take a minimum of approximately 18months.Attempting "best efforf' phased array ultrasonic examinations on these welds for which relief isrequested without the needed technique development and demonstrations would not producereliable results.

Even though qualified procedures exist for the weld thickness and diameter, the geometry would negatively impact sound path calibration, search unit focusing, properinside diameter impingement angles, and cause mis-orientation angles. As a result, thisgeometric complexity of the configuration would challenge the capability of current procedures to reliably characterize and size indications identified during the examinations and could leadto false-positive indications, as well as unnecessary increased radiation exposure toexamination personnel.

Additional information concerning the hardships associated with the use of a manual phasedarray ultrasonic search unit, an eddy current or ultrasonic inspection from the inner diameter ofthe component, and a high-angle ultrasonic inspection method, and dose estimates associated with these methods, is provided in the ENO response to question RAI-2.2 inReference 3.A design mitigation strategy has not been approved for this weld configuration.

Creating amitigation strategy would require significant time for development of tooling, qualification ofprocedures and personnel, and mockup verification, in order to develop an approved methodthat would also minimize radiation exposure to personnel.

3 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Due to the location of the welds, performing unplanned manual phased array ultrasonic testexaminations of the nine welds would involve significant radiation exposure to personnel.

Total dose incurred by examination, radiation protection, and supervisory personnel duringultrasonic testing of the nine weld locations is estimated to be at least 37 rem. This totalincludes preparation activities, and credits dose reduction controls and measures such asshielding, decontamination of components, high efficiency particulate air filter ventilation units,cameras, and remote telemetry.

Dose rates vary, on contact, from 120 mrem/hour to 15000mrem/hour for the weld locations.

General area dose rates in the vicinity of the weld locations vary from 25 mrem/hour to 200 mrem/hour.

The ENO response to question RAI-2.1 in Reference 3 provides detailed radiological doseestimates associated with performing unplanned volumetric inspections and dose estimates associated with performing the volumetric inspections with appropriate planning and foresight.

The total dose for performing unplanned volumetric inspections was estimated to beapproximately 37 Rem. The estimated total dose with appropriate planning and foresight wasestimated to be 14.5 Rem. Reference 3 provides a breakdown of dose for each weld forerecting and removing scaffolding, conducting

surveys, removing insulation, etc. The timesfor each activity along with the general area dose rates and dose rates at 30 cm from thenozzles are provided in the reference.

Development of a qualified procedure for the volumetric inspections would includeidentification of examination techniques that would minimize radiological exposure toexamination personnel.

4. PROPOSED ALTERNATIVE AND BASIS FOR USEProposed Alternative

1) Perform periodic system leakage tests in accordance with ASME Section Xl Examination Category B-P, Table IWB-2500-1.

2) Perform visual examinations (per Code Case N-722-1) and dye penetrant surfaceexaminations (per ASME Section X1 Examination Category B-J, Table IWB-2500-1) ofthe welds in accordance with ASME requirements.

3) Perform a volumetric examination, using ASME Code, Section Xl, Appendix VIII,Supplement 10 qualified procedures, equipment and personnel, on each of the ninesubject welds of this alternative during the next scheduled refueling outage (1 R24).4) Until the next scheduled refueling outage, if unidentified PCS leakage increases by 0.15gpm above the WCAP-16465NP baseline mean, and is sustained for 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />, ENO willtake action to be in Mode 3 within 6 hours6.944444e-5 days <br />0.00167 hours <br />9.920635e-6 weeks <br />2.283e-6 months <br /> and Mode 5 within 36 hours4.166667e-4 days <br />0.01 hours <br />5.952381e-5 weeks <br />1.3698e-5 months <br />, and perform baremetal visual examinations of the nine subject welds of this alternative, unless it can beconfirmed that the leakage is not from these welds.Regulation 10 CFR 50.55a(g)(6)(ii)(F)(1) states "Licensees of existing, operating pressurized water reactors as of July 21, 2011 must implement the requirements of ASME Code CaseN-770-1, subject to the conditions specified in paragraphs (g)(6)(ii)(F)(2) through(g)(6)(ii)(F)(10) of this section, by the first refueling outage after August 22, 2011."4 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Regulation 10 CFR 50.55a(g)(6)(ii)(F)(3) states that baseline examinations for welds in CodeCase N-770-1, Table 1, Inspection Items A-1, A-2, and B, must be completed by the end ofthe next refueling outage after January 20, 2012.Pursuant to 10 CFR 50.55a(z)(2),

ENO will comply with 10 CFR 50.55a(g)(6)(ii)(F) bycompleting the baseline volumetric examinations as required by Code Case N-770-1, for eachof the nine subject welds of this alternative, prior to the end of the next scheduled refueling outage (1 R24), currently scheduled to start in fall 2015.Basis for UseTable 1 in Attachment 1 of this enclosure describes the eight cold leg welds and one hot legweld that have not been examined in accordance with Code Case N-770-1 examination requirements, as required by 10 CFR 50.55a(g)(6)(ii)(F) items (1) and (3).Examination HistoryTable 1 in Attachment 1 also provides examination history information for the nine weldlocations for which relief is requested.

No evidence of through-wall cracking for thesecomponents has been identified during these inspections.

Moreover, for the three weldlocations that were not subject to surface or visual examinations during the 1 R23 refueling outage (i.e., weld no.'s 3, 6, and 7 in Table 1), maintenance activities in the vicinity of theweld locations during the 1 R23 refueling outage did not identify observations of leakagefrom the welds.Structural Evaluation

Background

ENO submitted a proposed alternative (relief request number RR 4-18) concerning volumetric examinations of the subject branch connection welds on February 25, 2014(Reference 1), and supplemental information was submitted on March 1, March 4, March 6,March 9, and March 11, 2014 (References 2 through 7).ENO provided the following three calculations in the March 6, 2014 supplemental submittal (Reference 4):1. Structural Integrity Associates, Inc, Calculation 1200895.306, "Hot Leg Drain NozzleWeld Residual Stress Analysis and Circumferential Crack Stress Intensity FactorDetermination,"

Revision 0, dated March 6, 2014.2. Structural Integrity Associates, Inc., Calculation 1200895.307, "Hot Leg Drain NozzleCrack Growth Analyses,"

Revision 0, March 6, 2014.3. Structural Integrity Associates, Inc., Calculation 1200895.308, "Hot Leg Drain NozzleLimit Load Analyses for Flawed Nozzle-to-Hot Leg Weld," Revision 0, March 6, 2014.The SmartCrack software program was used to analyze crack propagation in the branchconnection welds. SmartCrack uses a closed-form solution approach to calculate stressintensity factors for a crack with a two-dimensional residual stress field, and requiresassumptions to be made concerning the model geometry, crack dimensions, boundary5 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE conditions, and stress field. This closed-form solution approach allowed for expeditiously analyzing crack growth under the time constraints imposed during the refueling outage in2014.On March 12, 2014 (Reference 8), the NRC staff verbally authorized the use of reliefrequest number RR 4-18 at PNP until the next scheduled refueling outage, scheduled in thefall of 2015 (1 R24). A NRC safety evaluation detailing the technical basis for the verbalauthorization was subsequently issued on September 4, 2014 (Reference 9). The safetyevaluation credited the ENO flaw and structural evaluation, in addition to the identified hardship and conditions of relief identified in the proposed alternative, in concluding thatENO had provided sufficient information to demonstrate reasonable assurance of thestructural integrity of the nine subject welds for one cycle of operation without performing avolumetric examination.

On February 27, 2015, SI verbally notified ENO of a discrepancy discovered in a calculation that was submitted on March 6, 2014 (Reference

4) in support of relief request number RR4-18. In a letter to ENO dated March 6, 2015, SI stated that the discrepancy wasdiscovered in the residual stress calculation (File No. 1200895.306) within the loadapplication for the normal operating conditions (NOC) evaluation.

Specifically, the insidediameter surface elements adjacent to the circumferential free end and the axial free end ofthe hot leg pipe did not receive pressure loading (i.e., the effective pressure was 0 psig),while the remaining internal surfaces were pressurized to 2,122 psig for the NOC cycles.The corrected calculations and updates of the associated SI memorandums are provided inEnclosure

2. These calculations and memorandums are provided for informational purposes only, and do not provide the technical basis for the enclosed proposed alternative.

Structural Evaluation Using Finite Element Analysis Methodology The technical basis for the proposed alternative is provided in Attachment 2 of thisenclosure, which includes the following report and calculations:

SI Report No. 1400669.401

.RO, "Evaluation of the Palisades Nuclear Plant Branch LineNozzles for Primary Water Stress Corrosion Cracking,"

Revision 0, dated May 14, 2015.SI Calculation, File No. 1400669.313, "Crack Growth Analysis of the Hot Leg DrainNozzle,"

Revision 0, dated May 11, 2015.SI Calculation, File No. 1400669.323, "Crack Growth Analysis of the Cold Leg BoundingNozzle,"

Revision 0, dated May 11, 2015.SI Calculation, File No. 1400669.310, "Finite Element Model for Hot Leg Drain Nozzle,"Revision 0, dated March 9, 2015.SI Calculation, File No. 1400669.320, "Finite Element Model Development for the ColdLeg Drain, Spray, and Charging Nozzles,"

Revision 0, dated April 3, 2015.SI Calculation, File No. 1400669.312, "Hot Leg Drain Nozzle Weld Residual StressAnalysis,"

Revision 0, dated May 5, 2015.6 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE SI Calculation, File No. 1400669.322, "Cold Leg Bounding Nozzle Weld Residual StressAnalysis,"

Revision 0, dated May 5, 2015.In these calculations, SI used a finite element analysis (FEA) approach to evaluatepostulated flaws in the hot leg and cold leg nozzles.

These models were used to performweld residual stress evaluations and calculations of stress intensity factors in the DM welds.Utilizing these new stress intensity factor distributions for postulated circumferential andaxial flaws in the DM welds, crack growth due to PWSCC was evaluated for both the hot legand cold leg configurations.

Crack growth durations were then plotted on charts to show theservice life of the hot leg and cold leg configurations based on crack growth from anassumed initial flaw depth of 0.025 inch. It should be noted that PWSCC was the only crackgrowth mechanism considered in this evaluation (i.e., PWSCC growth of a postulated axialand circumferential flaw in the weld).The FEA approach is a more sophisticated approach for evaluating a postulated flaw thanthe SmartCrack closed-form solution approach used to support relief request number RR4-18. Using a FEA approach, such as the ANSYS FEA software, enables a more realistic and accurate representation of the cracks in a nozzle as it grows. The FEA approach is adetailed numerical method that uses the actual residual stresses directly within the model,without having to make assumptions concerning model geometry and without transitioning the stress inputs into a separate

program, like the assumptions made in SmartCrack.

Using the FEA approach, the calculations determined that, for the hot leg drain nozzle, thetime for an initial 0.025 inch deep flaw to grow to 75% through-wall was calculated to be30.5 years for the bounding axial flaw (36.7 years to go 95% through-wall) and 33.9 yearsfor the circumferential flaw (42.1 years to go 95% through-wall).

For the cold leg drain nozzle, the time for an initial 0.025 inch deep flaw to grow to 75%through-wall is 64.5 years for the bounding axial flaw (77 years to go 95% through-wall) and55.6 years for the circumferential flaw (66.2 years to go 95% through-wall).

By the 1 R24 refueling outage, PNP will have operated for 27.4 effective full power years(Reference 21).The SI report in Attachment 2 discusses a benefit in crack growth rate when Alloy 182 weldmetal underwent post weld heat treatment (PWHT). This benefit ranged from a factor of twoto four. Dominion Engineering recommended that crack growth rates be reduced by aconservative factor of two based on this data. Using this factor for the nozzles at PNP willincrease the crack growth times previously discussed.

For comparison, only the boundingcrack growth life will be discussed here. This is the axial flaw for the hot leg drain nozzle,and the circumferential flaw for the cold leg nozzles.

A factor of two will increase the limitinghot leg drain nozzle duration to 61.0 years, and the bounding cold leg nozzle to 111.2 years,for flaws to grow 75% through wall. Similarly, the durations for 95% through-wall flawswould increase to 73.4 years for the hot leg drain nozzle and 132.4 years for the cold legnozzle.Attachment A of the SI report in Attachment 2 includes a calculation of the initiation time forthe hot leg drain nozzle and cold leg bounding nozzle using one of the models in the xLPRprogram that was developed by Electric Power Research Institute (EPRI). The calculation ofinitiation time based on the PNP results shows that the time to initiation for the hot leg drainnozzle is approximately 130 years. If the crack initiation time is combined with the PWHT7 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE reduction in crack growth rate, the time for a flaw to grow to 75% through-wall wouldapproach 200 years for the hot leg drain nozzle and would exceed 200 years for 95%through-wall.

Using the crack initiation value for the cold leg nozzles, the time for a flaw togrow to 75% through-wall would exceed 600 years. For both the hot leg and cold legnozzles, the time for a flaw to grow to 75% through-wall is well beyond the life of the plant.Additional Information Post Weld Heat Treatment All welds referenced in this relief request were post weld heat treated at the vendor's facilityduring original fabrication.

Per Combustion Engineering detail weld procedure (MA-41,Revision 0), heat treatment of the hot leg nozzle weld (weld number 5-675) consisted of anintermediate post weld heat treatment at 11000F, -00F/+ 500F, for 15 minutes.

Final heattreatment at the hot leg nozzle consisted of 11 500F, +/- 250F, for one hour per inchthickness of weld. No heat treatment was performed at the site for the subject welds. SeeReference 2.The post weld heat treatment was applied to the nozzle welds by heating the affectednozzle/weld as part of the hot leg pipe assembly in a furnace.

The hot leg drain nozzle(piece number 675-03) was installed into the hot leg pipe (pipe assembly 673-04) by cuttinga hole into the hot leg piping and then fitting and welding the hot leg drain nozzle in placeusing an intermediate heat treat of 11 00°F + 50°F for 15 minutes.

The backing ring for theweld was removed, the weld was dye penetrant tested, and then back welded. The backweld was dye penetrant tested, and any indications were repaired, and then a final dyepenetrant test was performed.

The weld was then back clad and dye penetrant tested. Anyindications were removed and dye penetrant tested again. The hot leg drain nozzle weldwas radiographed and any defects were removed, weld repaired, and then dye penetrant tested and radiographed.

The pipe assembly with the hot leg drain nozzle installed waspost weld heat treated in a furnace using a post weld heat treat of 1150°F +/- 250F, holdingthis temperature for one hour/inch thickness of weld. See Reference 5.Weld Repair HistoryThe manufacturing/quality plan provided in the specification for the PCS piping providesinstructions for performing weld repairs based on the results of NDE testing.

Any defectsidentified in the nozzle welds would have been removed prior to final furnace heat treatment of the assembly.

PNP is unable to locate post-fabrication documentation other than theweld radiographs taken after the final furnace heat treatment.

These radiographs represent the condition of the subject welds at the time of installation at the site. A search of PNPrecords did not identify any repairs performed on the subject welds since installation.

SeeReference 2.8 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Weld and Cladding GeometryThe hot leg weld is 1/22 inch wide at the ID and then increases in width at a 7-/2' (+2°/-0°)

angle to the OD of the hot leg pipe. The thickness of the hot leg pipe is 3-34 inches.See Reference 2.The thickness of the alloy 182/82 cladding on the inside surface of the weld is 1/4 inchnominal thickness with a minimum thickness of 5/32 inch. The inner radius at the hot legdrain is 21.34 inches measured from the inside surface of the cladding to the centerline ofthe hot leg pipe. See Reference 5.Detailed information concerning the weld geometry is provided in References 2 and 5.Basis for Assuming No Weld RepairThe presence of an initial weld repair from plant construction (e.g., extending 50% of thewall thickness from the inside diameter (ID)) is often assumed when modeling Alloy 82/182piping butt welds. Often for piping butt welds, the residual stress calculated for the ID is asmall tensile value, or even compressive, in the absence of an assumed weld repair. In suchcases, the possibility of a significant weld repair being present on the weld ID can have arelatively large effect on the calculated

stresses, especially on and near the ID surface.However, for the Alloy 82/182 branch connection welds at PNP, there are two reasons whyit is not necessary to include a weld repair assumption in the analysis.

First, the design forthis weldment specifies a 3600 backweld on the ID surfaces of the pipe that is about 0.25inch thick. This design feature results in elevated residual stress levels at the ID surfaceprior to the PWHT being applied.

The residual stress levels at the inside surface due to thepresence of the backweld are similar to what would be expected due to the presence of aweld repair on the ID surface.Second, any weld repairs would have been made prior to PWHT being applied, and wouldbe expected to extend over a relatively limited circumferential portion of the original weld.Similar to the situation for the elevated residual stresses due to the presence of thebackweld, the PWHT would relax the residual stresses in the weld repair area, including thesubstantial relaxation expected at the surface exposed to primary coolant.

Moreover, in theunlikely case that initiation occurred in the area of a weld repair, the weld repair would be anadditional source of non-axisymmetric crack loading that would tend to drive crack growththrough-wall over a relatively local circumferential region, ultimately resulting in detection ofleakage prior to the possibility of unstable pipe rupture.See Reference 2.Basis for Five-Cycle Shakedown Assumption Operational cycles are frequently included in welding residual stress calculations as a part ofdetermining the operating stress condition.

In particular, the standard modeling practiceadopted by the xLPR (Extremely Low Probability of Rupture) welding residual stress team,which includes the NRC, national laboratories, and industry participants, specifies that thewelded configuration should be cycled between operating conditions and residual conditions to shake down the nonlinear material hardening behavior.

Typically, three to five cycles areused to shake down the material's behavior.

9 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Since the primary interest from the residual stress analysis is to provide residual stresses forcalculating stress corrosion crack growth under normal operating conditions, it is desirable todetermine a stabilized residual stress state that will not change under normal operating cycles. The as-welded residual stresses usually contain localized peak stresses at somenodal locations.

Applying a few operating cycles will stabilize the stress peaks and valleysdue to the slight stress redistribution at elevated temperatures.

It has been determined through experience that the residual stresses will stabilize after three to five cycles. Fivecycles was used for conservatism.

See Reference 2.Operatinq Conditions The operating temperature of a component is a primary factor influencing the initiation ofPWSCC. Research by EPRI (Reference

17) indicates that the difference in the operating temperature between hot leg locations and cold leg locations is sufficient to significantly influence the time to initiation of PWSCC, with the susceptibility increasing with temperature.

The research reports PWSCC is less likely to occur in cold leg temperature penetrations.

All but one of the welds covered by this relief are found in lower temperature regions of thesystem, typically at temperatures near to Tco0d, which is approximately 5370F. This means,for these welds, there is a lower probability of crack initiation, and a slower crack growth rate(Reference 18).Leakage Detection Capabilities The leak detection methodology presently used by industry is very sensitive.

After a numberof recent operating events, the industry imposed an NEI 03-08, "Guideline for theManagement of Materials Issues,"

requirement to improve leak detection capability.

As aresult, virtually all pressurized water reactors (PWRs) in the United States, including PNP,have a leak detection capability of less than or equal to 0.1 gpm (Reference 16). All plants,including PNP, also monitor seven-day moving averages of reactor coolant system leakrates.Action response times following a detected primary coolant system leak vary, based on theaction level exceeded and whether containment entry is required to identify the source of theleak.Action levels have been standardized for all PWRs, and are based on deviations from:* the seven day rolling average,* specific values, and* the baseline mean.Leak rate action levels are identified in Pressurized Water Reactor Owners Group (PWROG)report, WCAP-1 6465, and are stated below:Each PWR utility is required to implement the following standard action levels for reactorcoolant system (RCS) inventory balance in their RCS leakage monitoring program.10 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Action levels on the absolute value of unidentified RCS inventory balance (fromsurveillance data):Level 1 -One seven day rolling average of unidentified RCS inventory balance valuesgreater than 0.1 gpm.Level 2 -Two consecutive unidentified RCS inventory balance values greater than0.15 gpm.Level 3 -One unidentified RCS inventory balance value greater than 0.3 gpm.Note: Calculation of the absolute RCS inventory balance values must include therules for the treatment of negative values and missing observations.

Action levels on the deviation from the baseline mean:Level 1 -Nine consecutive unidentified RCS inventory balance values greater thanthe baseline mean [p] value.Level 2 -Two of three consecutive unidentified RCS inventory balance valuesgreater than [p + 2a], where a is the baseline standard deviation.

Level 3 -One unidentified RCS inventory balance value greater than [p +3G].These action levels have been incorporated into PNP procedures.

A small steam leak from a weld flaw would, over time, result in a rise in containment sumplevel rate of increase.

Containment sump level is continually monitored, and if a rise in therate of containment sump level increase is observed, plant procedures direct plant operators to identify the source of the leakage.

Operators may also be alerted to a leak from a flaw bycontainment radiation monitoring instrumentation.

This instrumentation, required by theTechnical Specifications, is capable of detecting a 100 cm3/min leak in 45 minutes, based on1% failed fuel. Periodic system leakage tests are performed in accordance with ASMESection XI. Operator walkdowns of containment are periodically performed during poweroperations at lower levels of containment to detect leakage.Therefore, with the periodic system leakage tests, the visual and surface examinations performed during 2012 and 2014 refueling

outages, the results of the SI evaluation, containment monitoring activities, and the testing, examination, and PCS leakagerequirements in the proposed alternative, an acceptable level of quality and safety isprovided for identifying degradation from PWSCC prior to a safety-significant flawdeveloping.

5. DURATION OF PROPOSED ALTERNATIVE The duration of the proposed alternative is until refueling outage 1 R24, which is currently scheduled to begin in fall 2015.11 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE

6. REFERENCES

1. Entergy Nuclear Operations, Inc. letter PNP 2014-015, "Relief Request Number RR4-18 Proposed Alternative, Use of Alternate ASME Code Case N-770-1 BaselineExamination,"

dated February 25, 2014 (ADAMS Accession No. ML14056A533).

2. Entergy Nuclear Operations, Inc. letter PNP 2014-021, "Response to Request forAdditional Information dated February 26, 2014, for Relief Request Number RR 4-18 -Proposed Alternative, Use of Alternate ASME Code Case N-770-1 BaselineExamination,"

dated March 1,2014 (ADAMS Accession No. ML14072A361).

3. Entergy Nuclear Operations, Inc. letter PNP 2014-022, "Response to Second Requestfor Additional Information dated February 27, 2014, for Relief Request Number RR4-18 -Proposed Alternative, Use of Alternate ASME Code Case N-770-1 BaselineExamination,"

dated March 4, 2014 (ADAMS Accession No. ML14063A089).

4. Entergy Nuclear Operations, Inc. letter PNP 2014-028, "Supplemental Response toRequest for Additional Information dated February 26, 2014, for Relief RequestNumber RR 4-18 -Proposed Alternative, Use of Alternate ASME Code Case N-770-1Baseline Examination,"

dated March 6, 2014 (ADAMS Accession No. ML14066A409).

5. Entergy Nuclear Operations, Inc. letter PNP 2014-029, "Response to Third Request forAdditional Information dated March 5, 2014, for Relief Request Number RR 4-18 -Proposed Alternative, Use of Alternate ASME Code Case N-770-1 BaselineExamination,"

dated March 6, 2014 (ADAMS Accession No. ML14070A182).

6. Entergy Nuclear Operations, Inc. letter PNP 2014-030, "Second Supplemental Response to Request for Additional Information dated February 26, 2014, for ReliefRequest Number RR 4-18 -Proposed Alternative, Use of Alternate ASME Code CaseN-770-1 Baseline Examination,"

dated March 9, 2014 (ADAMS Accession No.ML1 4069A004).

7. Entergy Nuclear Operations, Inc. letter PNP 2014-031, "Response to Fourth Requestfor Additional Information dated March 11,2014, for Relief Request Number RR 4-18 -Proposed Alternative, Use of Alternate ASME Code Case N-770-1 BaselineExamination,"

dated March 11, 2014 (ADAMS Accession No. ML14070A477).

8. NRC Electronic Mail, "Palisades Nuclear Plant -Verbal Authorization for ReliefRequest RR 4-18 -MF3508,"

March 13, 2014 (ADAMS Accession No. ML14073A274).

9. NRC letter, "Palisades Nuclear Plant -Proposed Alternative, Use of Alternate ASMECode Case N-770-1 Baseline Examination (TAC No. MF3508),"

dated September 4,2014 (ADAMS Accession No. ML14223B226).

10. 10 CFR 50.55a, "Codes and standards,"

December 11, 2014.11. ASME Section Xl, "Rules For Inservice Inspection of Nuclear Power PlantComponents,"

2001 Edition with Addenda through 2003.12. ASME Section Xl, Division 1, Code Case N-460, "Alternative Examination Coveragefor Class 1 and Class 2 Welds,Section XI, Division 1."12 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE

13. Material Reliability Program:

Primary System Piping Butt Weld Inspection andEvaluation Guideline (MRP-139),

Revision 1, EPRI, Palo Alto, CA, 2008 (ADAMSAccession No. ML1009700671).

14. Nondestructive Evaluation:

Procedure for Manual Phased Array Ultrasonic Examination of Dissimilar Metal Welds, EPRI-DMW-PA-1, Revision 3,1016645, EPRIPalo Alto, CA, 2008.15. "Changing the Frequency of Inspections for PWSCC Susceptible Welds at Cold LegTemperatures",

in Proceedings of 2011 ASME Pressure Vessels and PipingConference, July 17-21, 2011, Baltimore, MD.16. WCAP-1 6465-NP, Rev. 0, "Pressurized Water Reactor Owners Group Standard RCSLeakage Action Levels and Response Guidelines for Pressurized Water Reactors,"

Westinghouse Electric Co., September 2006 (ADAMS Accession No. ML070310082).

17. Electric Power Research Institute:

PWSCC of Alloy 600 Materials in PWR PrimarySystem Penetrations, EPRI, Palo Alto, CA, 1994, TR-103696 (ADAMS Accession No.ML013110446).

18. Materials Reliability Program Crack Growth Rates for Evaluating Primary Water StressCorrosion Cracking (PWSCC) of Alloy 82,182, and 132 Welds (MRP-1 15), EPRI, PaloAlto, CA, 2004,1006696 (ADAMS Accession No. ML051100204).

19. PNP Technical Specification Surveillance Procedure RT-71A, "Primary CoolantSystem, Class 1 System Leakage Test," Revision 18.20. Pressurized Water Reactor (PWR) Owner's Group Letter OG-12-89, "Transmittal of'Final Relief Request Framework' under Relief Request for Large Diameter Cold LegLocations with Obstructions (PA-MSC-0934),"

March 8, 2012.21. WCAP-15353-Supplement 1-NP, Palisades Reactor Pressure Vessel FluenceEvaluation, Revision 0, May 2010 (ADAMS Accession Number ML1 10060692).

7. ATTACHMENTS Attachment 1Table 1 -Weld Examination HistoryFigure 1 -Nozzle Assembly Materials Figure 2 -Hot Leg Drain Nozzle Configuration (Representative) 13 of 14 ENCLOSURE 1PROPOSED ALTERNATIVE Attachment

21. Structural Integrity Associates, Inc. Report No. 1400669.401

.RO, "Evaluation of thePalisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion Cracking,"

Revision 0, dated May 14, 2015.2. Structural Integrity Associates, Inc. Calculation, File No. 1400669.313, "CrackGrowth Analysis of the Hot Leg Drain Nozzle,"

Revision 0, dated May 11, 2015.3. Structural Integrity Associates, Inc. Calculation, File No. 1400669.323, "CrackGrowth Analysis of the Cold Leg Bounding Nozzle,"

Revision 0, dated May 11,2015.4. Structural Integrity Associates, Inc. Calculation, File No. 1400669.310, "FiniteElement Model for Hot Leg Drain Nozzle,"

Revision 0, dated March 9, 2015.5. Structural Integrity Associates, Inc. Calculation, File No. 1400669.320, "FiniteElement Model Development for the Cold Leg Drain, Spray, and ChargingNozzles,"

Revision 0, dated April 3, 2015.6. Structural Integrity Associates, Inc. Calculation, File No. 1400669.312, "Hot LegDrain Nozzle Weld Residual Stress Analysis,"

Revision 0, dated May 5, 2015.7. Structural Integrity Associates, Inc. Calculation, File No. 1400669.322, "Cold LegBounding Nozzle Weld Residual Stress Analysis,"

Revision 0, dated May 5, 2015.14 of 14 ENCLOSURE 1ATTACHMENT 1Table 1Weld Examination HistoryNo. Description ISI Weld ID Location 1R19 1R20 1 R21 1 R22 1 R23Examinations Examinations Examinations Examinations Examinations Visual Surface1 2 inch Cold Leg PCS-30-RCL-1A-P-50A Discharge (Report#

4046 (Report#

1R23-Charging Nozzle 11/2 Leg Exam number PT- 14-025)06-26) P-405Visual2. 2 inch Cold Leg PCS-30-RCL-1A-P-50A Suction (Report#

4047 SurfaceDrain Nozzle 5/2 Leg Exam number (Report#

1R23-07-2.1)PT-i14-031) 3 inch Cold Leg PCS-30-RCL-1B-P-50B Discharge Visual Surface3. Pressurizer Spray 10/3 Leg (Report#

VT (Report#

1 R22-Nozzle 069) PT-12-039) 2 inch Cold Leg PCS-30-RCL-1 B- P-50B Suction Visual SurfaceDrain Nozzle 5/2 Leg (Report#

VT (Report #1 R23-2 inch Cold Leg PCS-30-RCL-2A-P-5OC Discharge Visual Surface5. 2aing Nozzle 11/2 Leg (Report#

VT (Report#

1 R23-Charging Nozzle_11/2_Leg083)

PT-14-019) 3 inch Cold Leg PCS3RCL2A P-SC Discharge Visual Surface6. Pressurizer Spray 11/3 Leg (Report#

VT (Report#

1 R22-Nozzle 11/3 Leg_035)

PT-12-032) 2 inch Cold Leg PCS-30-RCL-2A-P-50C Suction VisualDrain Nozzle 5/2 Leg 038)2 inch Cold Leg PCS-30-RCL-2B-P-50D Suction Visual Surface8. Drain and 5/2 Leg (Report#

VT (Report#

1R23-Letdown Nozzle 071) PT-14-020)

Visual Visual Visual VisualDrinNzze

/ (eor# 07 Rpot#V Visual Vsa9. 2 inch Hot Leg PCS-42-RCL-1 H- A Hot Leg (Report#

4047 (Report#

VT (Report#

VT (Report#

1 R22- (Report#

1R23-Drain Nozzle 3/2 Exam number 062) 022) VT-12-076)

VT-14-059) 107-23.1) 1)1 of 3 ENCLOSURE 1ATTACHMENT 1Figure 1Nozzle Assembly Materials Alloy 82/182Stainless Steel Cladding ICarbon Steel Cold Leg,--IFull Penetration DMAlloy 82/182 WeldINot to Scale2 of 3 ENCLOSURE 1ATTACHMENT 1Figure 2Hot Leg Drain Nozzle Configuration (Representative)

(excerpt from PNP vendor drawing VEN-M1-D Sheet 108, Revision 8)PIOC 4(IYRP'FJ7R WE"L.D/,NVG REAMOL ro SOM 444 f "LO 617 OQAIN AIOZZLAX4LE 61.103 of 3 ENCLOSURE 1ATTACHMENT 2The attached Structural Integrity Associates, Inc. report and calculations provide the technical basis for the proposed alternative.

1. Structural Integrity Associates, Inc. Report No. 1400669.401

.RO, "Evaluation of thePalisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion Cracking,"

Revision 0, dated May 14, 2015.2. Structural Integrity Associates, Inc. Calculation, File No. 1400669.313, "CrackGrowth Analysis of the Hot Leg Drain Nozzle,"

Revision 0, dated May 11, 2015.3. Structural Integrity Associates, Inc. Calculation, File No. 1400669.323, "CrackGrowth Analysis of the Cold Leg Bounding Nozzle,"

Revision 0, dated May 11,2015.4. Structural Integrity Associates, Inc. Calculation, File No. 1400669.310, "FiniteElement Model for Hot Leg Drain Nozzle,"

Revision 0, dated March 9, 2015.5. Structural Integrity Associates, Inc. Calculation, File No. 1400669.320, "FiniteElement Model Development for the Cold Leg Drain, Spray, and ChargingNozzles,"

Revision 0, dated April 3, 2015.6. Structural Integrity Associates, Inc. Calculation, File No. 1400669.312, "Hot LegDrain Nozzle Weld Residual Stress Analysis,"

Revision 0, dated May 5, 2015.7. Structural Integrity Associates, Inc. Calculation, File No. 1400669.322, "Cold LegDrain Nozzle Weld Residual Stress Analysis,"

Revision 0, dated May 5, 2015.213 pages follow 5215 Hellyer Ave.Suite 210San Jose, CA 95138-1025 Phone: 408-978-8200 Fax: 408-978-8964 www.structint.com clohse@structint.com May 14, 2015Report No. 1400669.401.RO Quality Program:

ED Nuclear (Safety-related)

D Commercial Mr. Keith SmithEntergy Nuclear Operations, Inc.Palisades Nuclear Plant27780 Blue Star Memorial HighwayCovert, MI 49043

Subject:

Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary WaterStress Corrosion Cracking

Dear Keith:

In Relief Request RR 4-18 [1], Entergy Nuclear Operations (Entergy) committed tovolumetrically examine a total of nine (9) pressure retaining dissimilar metal welds (DMWs)containing Alloy 600 nozzles and Alloy 182 weld material at the Palisades Nuclear Plant(Palisades) during the upcoming 1R24 refueling outage (Fall of 2015). Of the nine weldlocations, eight are cold leg nozzles (three 2-inch diameter drain nozzles, one 2-inch diameterletdown nozzle, two 2-inch diameter charging

nozzles, and two 3-inch diameter spray nozzles)and one 2-inch diameter hot leg drain nozzle.On December 1, 2014, Entergy contracted with Structural Integrity Associates, Inc. (SI) [2] tosupport a flaw readiness evaluation for the 1 R24 examinations of the nine reactor coolant system(RCS) branch connection Alloy 600 nozzle boss welds at Palisades.

SI began working in earneston this project in early January 2015. The base work scope involved performing flawevaluations for the hot leg and cold leg branch line DMWs. SI's work was interrupted in mid-February 2015 to support a possible relief request to defer inspections for the upcoming outage.To support Entergy's new Relief Request, SI was requested to provide a technical basis fordeferral of inspection of the hot leg drain nozzle and eight cold leg branch line nozzle DMWs.In order to perform the necessary flaw evaluations, finite element models were developed for thehot leg and cold leg nozzles.

These models were used to perform weld residual stressevaluations and calculations of stress intensity factors in the DMWs. Utilizing these new stressintensity factor distributions for postulated circumferential and axial flaws in the DMWs, crackgrowth due to primary water stress corrosion cracking (PWSCC) was evaluated for both the hotToll-Free 877-474-7693 Akron, OH Austin, TX Chaulotte, iC Chattanooga, TN330-M999753 512-533-9191 704-597-5554 423-553-1180

ChIcago, IL Denver, CO San Diego, CA San Jose, CA State College, PA Toronto, Canada877-474-7693 303-792-0077 858-455-6350 408-978-8200 814-954-7776 905-829-9817 Mr. Keith Smith May 14, 2015Report No. 1400669.401

.RO Page 2 of 29Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion Crackingleg and cold leg configurations.

Crack growth durations were then plotted on charts to show theservice life of the hot leg and cold leg configurations based on crack growth from an assumedinitial flaw depth of 0.025 inch.This report summarizes the evaluation results.

It should be noted that PWSCC was the onlycrack growth mechanism considered in this evaluation (i.e., PWSCC growth of a postulated axialand circumferential flaw in the weld). The overall goal of SI's evaluation was to show that ittakes considerable time to grow postulated flaws to allowable flaw sizes in both the hot leg drainnozzle and the cold leg nozzle configurations, and that the crack growth time for these locations is quite significant.

Finite Element AnalysesFor the hot leg drain nozzle, a three-dimensional (3-D) finite element model was developed using the ANSYS software

[3]. The area of interest was the nozzle-to-hot leg DMW. The modelused elastic-plastic material properties intended for weld residual stress analysis.

The modelincluded a local portion of the hot leg pipe and cladding, the drain nozzle, and the nozzle-to-hot leg DMW, including the inside diameter (ID) patch weld, as shown in Figure 1. As depicted inthe figure, a single 900 quadrant of the drain nozzle penetration is modeled due to geometric symmetry.

The included portion of the hot leg piping measured 36 inches longitudinally and 180degrees circumferentially.

The mesh of the finite element model is shown in Figure 2.For the cold leg nozzles, one bounding 3-D finite element model was developed using theANSYS software

[4]. All three nozzle configurations (i.e., drain/letdown,

charging, and spray)are of similar size (diameter) near the forging boss area (within 1/16 inch). Therefore, the largestID and smallest outside diameter (OD) of the three nozzle types was chosen for the boundingmodel. The spray and drain/letdown nozzles have identical nozzle and boss OD dimensions of 4-9/16 inches and 6-3/16 inches, respectively, which are slightly smaller than the charging nozzleOD dimensions of 4-5/8 inches and 6-1/4 inches. For the nozzle ID, the charging nozzle wasbored out to 2-5/8 inches in the first 1.5 inches to accommodate a thermal sleeve. Forconservatism, it was assumed that the entire nozzle ID is 2-5/8 inches. Therefore, the ID of thebored out charging nozzle and the OD of the spray/drain/letdown nozzles were used for thebounding cold leg nozzle configuration.

For the cold leg nozzles, the area of interest was the nozzle-to-cold leg DMW. The model usedelastic-plastic material properties intended for weld residual stress analysis.

The model includeda local portion of the cold leg pipe and cladding, the nozzle, and the nozzle-to-cold leg DMW,including the ID patch weld, as shown in Figure 3. As depicted in the figure, a single 90'quadrant of the nozzle penetration is modeled due to geometric symmetry.

The included portionof the cold leg piping measured 36 inches longitudinally and 180 degrees circumferentially.

Themesh of the finite element model is shown in Figure 4.Structural Integrity Associates, Inc.

Mr. Keith Smith May 14, 2015Report No. 1400669.401.RO Page 3 of 29Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion CrackingWeld Residual Stress AnalysesWeld residual stress analyses were performed for the hot leg drain nozzle [5] and bounding coldleg nozzle [6], and the following steps of initial construction were simulated in the analyses:

1. Deposit cladding on hot leg/cold leg pipe inside surface.2. Install nozzle, backing ring, and deposit boss weld.3. Remove backing ring and deposit ID patch weld.4. Post weld heat treatment (PWHT), including creep effects based upon experimental data.5. Subject the configuration to a hydrostatic test.6. Impose five cycles of "shakedown" at normal operating temperature and pressure tostabilize the residual stress fluctuations due to stress redistribution.

For the hot leg drain nozzle, the boss weld connects the drain nozzle to the hot leg piping. Asshown in Figure 5, the weld is composed of 40 nuggets deposited in 20 weld layers. As depictedin Figure 6, the ID patch weld is composed of 6 nuggets deposited in 2 layers. Stabilized stressresults normal to a circumferential flaw are shown in Figure 7, and those normal to an axial flaware shown in Figure 8. They represent the operating conditions of the hot leg, and include theresidual stresses due to welding plus the operating pressure of 2085 psig and steady-state temperature of 583'F.For the cold leg nozzles, the boss weld connects the bounding nozzle to the cold leg piping. Asshown in Figure 9, the weld is composed of 40 nuggets deposited in 20 weld layers. As depictedin Figure 10, the ID patch weld is composed of 6 nuggets deposited in 2 layers. Stabilized stressresults normal to a circumferential flaw are shown in Figure 11, and those normal to an axialflaw are shown in Figure 12. They represent the operating conditions of the cold leg, andinclude the residual stresses due to welding plus the operating pressure of 2085 psig and steady-state temperature of 5370F.Crack Growth Evaluations Crack growth evaluations were performed for the hot leg drain nozzle [7] and the bounding coldleg nozzle [8]. For both the hot leg drain nozzle and bounding cold leg nozzle, stress intensity factors (Ks) at five depths for a 3600 inside surface connected, part-through wall circumferential flaw as well as two axial thumbnail flaws, at the 0- and 90-degree azimuthal locations of thenozzle, were calculated using finite element analysis techniques.

For the circumferential flaw,the maximum K values around the nozzle circumference for each flaw depth were extracted andused as input for performing the PWSCC growth analyses.

For the axial flaws, the K at thedeepest point of the thumbnail shape was used as input for performing the PWSCC growthanalyses.

For the crack growth analyses, a small initial flaw size of 0.025 inch was chosen. This is basedon the expected flaw size that would be possible since the crack growth evaluation is assumed toV Structural Integrity Associates, Inc.

Mr. Keith Smith May 14, 2015Report No. 1400669.401

.R0 Page 4 of 29Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion Crackingstart on day one of plant operation.

The final flaw size for these analyses is 75% of the wallthickness.

This final depth is chosen as it is the maximum allowable flaw depth per Section XIof the ASME Code for pipe flaw evaluations.

The 75% requirement is driven by the tables andequations in Appendix C, Article C-5000 of Section XI of the ASME Code. The equations andtables are limited to 75% of the pipe thickness.

The values for a through-wall flaw (95% depth)are also calculated to explore the time for a flaw to grow through-wall.

In response to NuclearRegulatory Commission RAI's [ 12] pertaining to Relief Request RR 4-18 [ 1 ], SI performed limitload analyses to show that the hot leg/cold leg/nozzle structures remained structurally stable forboth circumferential and axial through-wall flaws.The key parameters used in the crack growth calculations included:

• Initial flaw depth = 0.025" for both hot and cold leg configurations

• Temperature

= 583'F (hot leg) and 537'F (cold leg)* Wall thickness

= 4" (hot leg pipe thickness) and 3" (cold leg pipe thickness)

An example of the 3600 circumferential flaw is shown in Figure 13 for the hot leg nozzle andFigure 18 for the bounding cold leg nozzle. A total of five circumferential flaw depths weremodeled.

Similarly, five axial flaw depths were modeled at the 0- and 90-degree azimuths.

Figure 14 shows the axial flaws modeled at the 0-degree azimuth for the hot leg nozzle, andFigure 19 shows the comparable axial flaws modeled for the bounding cold leg nozzle.Stress intensity

factors, Ks, were calculated for the 3600 circumferential flaws as well as theaxial flaws at the 00 and 900 locations.

The stress intensity factors were calculated using thestress distributions for residual stress plus normal operating conditions shown in Figures 7 and 8for the hot leg drain nozzle, and Figures 11 and 12 for the bounding cold leg nozzle. In addition, a far field in-plane bending moment is applied to the free end of the hot leg and cold leg runpiping to account for piping moments in the main loop piping. This combined loading is usedfor the determination of the stress intensity factors for both the circumferential and axial flaws.Figures 15 and 16 show the calculated stress intensity factors for the hot leg drain nozzle for,thecircumferential and axial flaws, respectively.

Figures 20 and 21 show the calculated stressintensity factors for the bounding cold leg nozzle for the circumferential and axial flaws,respectively.

Crack growth evaluations were performed for all three flaws. For the hot leg drain nozzle,Figure 17 shows that the time for an initial 0.025" deep flaw to grow to 75% through wall is 30.5years for the bounding axial flaw (on the 0' plane) and 33.9 years for the circumferential flaw.For the 100% through wall flaw (growth up to 95%) the bounding axial flaw takes 36.7 years (onthe 00 plane) and the circumferential flaw takes 42.1 years as shown in Figure 17.For the bounding cold leg nozzle, Figure 22 shows that the time for an initial 0.025" deep flaw togrow to 75% through wall is 64.5 years for the bounding axial flaw (on the 00 plane) and 55.6years for the circumferential flaw. For the 100% through wall flaw (growth up to 95%) theV Structural Integrity Associates, Inc.

Mr. Keith Smith May 14, 2015Report No. 1400669.401.RO Page 5 of 29Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion Crackingbounding axial flaw takes 77.0 years (on the 00 plane) and the circumferential flaw takes 66.2years as shown in Figure 22.Potential Additional Margin Due to PWHTDominion Engineering documented in Reference 9 that industry papers [10, 11] note a benefit infatigue crack growth rate when Alloy 182 weld metal underwent PWHT. This benefit rangedfrom a factor of two to four. There was a recommendation that crack growth rates be reduced bya factor of two based on this data. If such a factor were used for the nozzles at Palisades, itwould increase the crack growth times previously discussed.

For comparison, only the boundingcrack growth life will be discussed here. This is the axial flaw for the hot leg drain nozzle, andthe circumferential flaw for the cold leg nozzles.

A factor of two would increase the limiting hotleg drain nozzle duration to 61.0 years, and the bounding cold leg nozzle to 111.2 years for flawsto grow 75% through wall. Similarly, the durations for 100% through wall flaws (growth up to95%) would increase to 73.4 years for the hot leg drain nozzle and 132.4 years for the cold legnozzle.Crack Initiation TimeThe original relief request [1] did not include any consideration for crack initiation.

Attachment A of this report includes a calculation of the initiation time for the hot leg drain nozzle and coldleg bounding nozzle using one of the models in the xLPR program that was developed by EPRI.The calculation of initiation time based on the Palisades results shows that the time to initiation for the hot leg drain nozzle is approximately 130 years. Attachment A contains additional detailsregarding the calculation.

Combining crack initiation time with the PWHT reduction in crackgrowth rate, the time for a flaw to grow to 75% through wall will approach 200 years for the hotleg drain nozzle. Using the crack initiation value for the cold leg nozzles, the time for a flaw togrow to 75% through wall would exceed 600 years.Conclusions Based on the assessments and calculations documented herein, it is concluded that these weldsdo not experience fast growing flaws due to PWSCC. The hot leg drain nozzle limiting crackgrowth life, not considering the benefits of PWHT or crack initiation time, is 30.5 years for aflaw to grow to 75% through wall, and 36.7 years for a flaw to grow to 100% through wall (95%through wall). Similarly, the cold leg nozzles take 55.6 years for the limiting flaw to grow to75% through wall, and 66.2 years for 100% through wall (95% through wall). The cold legnozzles lag the hot leg nozzle by roughly a factor of two.It should be noted that the resulting crack growth times are specified as the duration duringwhich the reactor is at operating pressure and temperature and not simply years since the reactorwas licensed to operate.

Per Reference 13, the total Effective Full Power Years (EFPY) at theend of the following refueling outage (1R25) is expected to be 28.8, which is less than the 75%through-wall crack growth time for both the hot leg nozzle and bounding cold leg nozzle.1 Structural Integrity Associates, Inc.

Mr. Keith Smith May 14, 2015Report No. 1400669.40l.RO Page 6 of 29Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion CrackingThese crack growth lives do not consider the potential additional margin on PWSCC growth ratefor PWHT or crack initiation time. Taking into account these margins would increase the timefor a crack to grow to 75% or 100% through wall to almost 200 years or more. Based on thisinformation, deferring inspections on these locations for one more plant operating cycle will notrepresent a decrease in nuclear safety.References

1. Relief Request Number RR 4-18, "Proposed Alternative, Use of Alternate ASME Code CaseN-770-1 Baseline Examination,"

PNP 2014-015, February 25, 2014.2. Entergy Nuclear Operations, Inc. Contract Order No. 10426669, Effective Date January 1,2015.3. Structural Integrity Associates Calculation 1400669.3 10, Rev. 0, "Finite Element Model forHot Leg Drain Nozzle."4. Structural Integrity Associates Calculation 1400669.320, Rev. 0, "Finite Element ModelDevelopment for the Cold Leg Drain, Spray, and Charging Nozzles."

5. Structural Integrity Associates Calculation 1400669.312, Rev. 0, "Hot Leg Drain NozzleWeld Residual Stress Analysis."

6. Structural Integrity Associates Calculation 1400669.322, Rev. 0, "Cold Leg BoundingNozzle Weld Residual Stress Analysis."

7. Structural Integrity Associates Calculation 1400669.313, Rev. 0, "Crack Growth Analysis ofthe Hot Leg Drain Nozzle."8. Structural Integrity Associates Calculation 1400669.323, Rev. 0, "Crack Growth Analysis ofthe Cold Leg Bounding Nozzle."9. Letter from G. White (Dominion Engineering, Inc.) to W. Sims (Entergy),

"Effect of Post-Weld Heat Treatment Applied to Alloy 82/182 Full-Penetration Branch Pipe Connection Welds at Palisades,"

DEI Letter L-4199-00-01, Rev. 0, dated February 25, 2014.10. T. Cassagne, D. Caron, J. Daret, and Y. Lef~vre, "Stress Corrosion Crack Growth RateMeasurements in Alloys 600 and 182 in Primary Water Loops under Constant Load,"Proceedings of 9,4nternational Symposium on Environmental Degradation of Materials inNuclear Systems-Water

Reactors, The Minerals, Metals & Materials

Society, Warrendale, PA, 1999, pp. 217-224.11. S. Le Hong, J. M. Boursier, C. Amzallag, and J. Daret, "Measurements of Stress Corrosion Cracking Growth Rates in Weld Alloy 182 in Primary Water of PWR," Proceedings of 1O"International Conference on Environmental Degradation of Materials in Nuclear PowerSystems--Water

Reactors, NACE International, 2002.12. Entergy Nuclear Operation, Inc. letter PNP 2014-028, "Supplemental Response to Requestfor Additional Information, dated February 26, 2014, for Relief Request Number RR 4-18,Proposed Alternative, Use of Alternate ASME Code Case N-770-1 Baseline Examination,"

dated March 6, 2014.13. Email from Daniel Depuydt (Entergy) to Norman Eng (SI), dated May 13, 2015, "

Subject:

Palisades EFPY at 1R25," includes attached file "WCAP-15353-Suppl 1 Excerpt.pdf,"

SI File No. 1400669.205.Structural Integrity Associates, Inc.

Mr. Keith Smith MReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water StressCorrosion Crackingay 14, 2015?age 7 of 29Very truly yours,Preparer:

Chris S. Lohse, P.E.Senior Consultant 5/14/2015 DateReviewer:

Reviewer:

Dilip DedhiaSenior Associate 5/14/2015 DateNorman EngAssociate 5/14/2015 DateApprover:

Richard A. Mattson, P.E.Senior Associate 5/14/2015 DateAttachment A -Crack Initiation Studycc: R. BaxM. FongC. FourcadeF. KuG. MukhimW. WongV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401

.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 8 of 29mANbYbR14.5Figure 1. Components Included in the Hot Leg Drain Nozzle Finite Element ModelV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 9 of 29Figure 2. Isometric View of the Hot Leg Drain Nozzle Finite Element ModelNote: Nozzle weld detail is shown in the bottom right comer.V Stnuctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401

.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 10 of 29Figure 3. Components Included in the Bounding Cold Leg Nozzle Finite Element ModelV Stnuctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.40 1.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 11 of 29Figure 4. Isometric View of the Bounding Cold Leg Nozzle Finite Element ModelNote: Nozzle weld detail is shown in the bottom left comer.r Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 12 of 29Figure 5. Weld Nugget Definitions for the Hot Leg Drain Nozzle Boss WeldV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 13 of 29Figure 6. Weld Nugget Definitions for the Hot Leg Drain Nozzle ID Patch WeldV Stnuctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.40 1.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 14 of 29Figure 7. Radial Stresses at Operating Conditions for the Hot Leg Drain NozzleNote: Radial stresses are shown in the nozzle axis radial direction and units are in ksi.V Sinuctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401

.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 15 of 29Figure 8. Circumferential Stresses at Operating Conditions for the Hot Leg Drain NozzleNote: Circumferential stresses are shown in the nozzle axis circumferential direction and units are in ksi.V Stnuclurul Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 16 of 29Figure 9. Weld Nugget Definitions for the Bounding Cold Leg Nozzle Boss WeldV Stnuctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401

.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 17 of 29Figure 10. Weld Nugget Definitions for the Bounding Cold Leg Nozzle ID Patch Weld!I Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.40 1.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 18 of 29Figure 11. Radial Stresses at Operating Conditions for the Bounding Cold Leg NozzleNote: Radial stresses are shown in the nozzle axis radial direction and units are in ksi.Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.40 1.R0Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 19 of 29Figure 12. Circumferential Stresses at Operating Conditions for the Bounding Cold Leg NozzleNote: Circumferential stresses are shown in the nozzle axis circumferential direction and units are in ksi.V Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 20 of 29Figure 13. Circumferential Flaw with Crack Tip Elements Inserted for the Hot Leg Drain NozzleNote: Deepest circumferential flaw is shown as an example.IV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 21 of 29Figure 14. Axial Flaws on the 0' Face with Crack Tip Elements Inserted for the Hot Leg Drain NozzleNote: Only the 0' axial flaws are shown. The 900 flaws are similar.V Stirctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 22 of 29Ci-r40353025201510500.00.20.4 0.60.81.0Depth (alt)Figure 15. Stress Intensity Factors as a Function of Depth for the Hot Leg Drain Nozzle Circumferential FlawsV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 23 of 292520In-r15U-I-______

_____-u-Axial_0

_ -Axial 9010500.00.2OA0.60.81.0Depth (a/t)Figure 16. Stress Intensity Factors as a Function of Depth for the Hot Leg Drain Nozzle Axial FlawsV Stctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.R0 Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 24 of 291.0/0.90.80.7'0.6(0 0.50 0.40.20.10.0//iI0iI- .-I -Circ. CrackAxial Crack_0-*- Axial Crack_90so01020 30Time (years)40GOFigure 17. Crack Growth for All Flaw Types in the Hot Leg Drain Nozzle (Initial Flaw of 0.025")V Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14,2015Page 25 of 29Figure 18. Circumferential Flaw with Crack Tip Elements Inserted for the Bounding Cold Leg NozzleNote: Deepest circumferential flaw is shown as an example.V Sructural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 26 of 29Figure 19. Axial Flaws on the 0° Face with Crack Tip Elements Inserted for the Bounding Cold Leg NozzleNote: Only the 00 axial flaws are shown. The 900 flaws are similar.!StV ctural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401.RO Evaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 27 of 296050n 40~30x2 201000.000.200.400.600.801.00Depth (a/t)Figure 20. Stress Intensity Factors as a Function of Depth for the Bounding Cold Leg Nozzle Circumferential FlawsV Structural Integrity Associates, Inc.

Mr. Keith SmithReport No. 1400669.401

.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 28 of 29353025= 20v15105-4Axial_0

-U-Axial_90 0 .......0.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00Depth (a/t)Figure 21. Stress Intensity Factors as a Function of Depth for the Bounding Cold Leg Nozzle Axial FlawsV Structural Integrily Associates, Inc.

Mr. Keith SmithReport No. 1400669.40 1.ROEvaluation of the Palisades Nuclear Plant Branch Line Nozzles for Primary Water Stress Corrosion CrackingMay 14, 2015Page 29 of 291.00.90.80.7"0.60-0.5_ 0.40.30.20.10.0--4-100.10e-." --0 .0 e 0-e .--"-.00-Circumferential

---Axial_0--- -Axial 900 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80Time (yrs)Figure 22. Crack Growth for All Flaw Types in the Bounding Cold Leg Nozzle (Initial Flaw of 0.025")V Structural Integrity Associates, Inc.

Attachment ACrack Initiation StudyIntroduction The crack growth evaluations performed for relief request RR 4-18 did not include anycalculation of initiation times for the hot leg drain nozzle. The evaluations only included a crackgrowth evaluation starting from day one of operation of the plant. This is conservative as sometime is required for cracks to initiate prior to PWSCC growth occurring.

This section calculates initiation time for the hot leg drain nozzle and the bounding cold leg nozzle using the moduledeveloped for the xLPR program.

The xLPR program is a joint venture between the industry andthe NRC to address the extreme low probability of rupture for Alloy 82/182 piping welds inPWR plants.Initiation ModelxLPR reviewed a number of initiation models as part of the program.

The xLPR Version 1.0code used one model that was benchmarked for use on Alloy 82/182 weld material.

The modelused was named Direct Model 2 and was developed by EPRI in Report No. 1019032 [2]. It wasbenchmarked for use for Alloy 82/182 material and the inputs for use are in the xLPR Version1.0 report by EPRI [1]. The model calculates the initiation time based on the material properties (yield and ultimate strength),

temperature, and surface stress. The model also has two threshold stresses.

The first is the stress initiation threshold (Uth) where if the surface stress is less, noinitiation will be predicted.

The equations below calculate this initiation threshold.

Allequations presented in this section are rearranged from those provided in Reference 2 for ease ofuse and presentation.

071h = Z07Ysz=z1 +z2, n()'; = 'T',--071.where-.s = local yield stress (input)C,,4, = local ultimate stress (input)z1,z2= CW-SCC threshold parameters (input)Additionally, there is a maximum stress (Omax) where initiation would occur instantly.

This valueis calculated using the following equations.

Attachment A to 1400669.40 1.RO A-1 Structural Integrity Associates, Inc.

0*max D = ve"'071'O'swherev, w = cold work microcracking resistance parameters (input)For the evaluation documented herein, the stresses are between the minimum and maximumthreshold values. Therefore, the time to initiation is calculated using the following equations.

t ,',"o°n = BGeQ1RT n D-zwhereG = m-q ln(D)lv-z)D = v.- e"`':z =z, +/-z2ln(O)n = k )a(4- -1)" ( 0Cv'sandtlNI,nomT= initiation time under fixed set of conditions (output)temperature (input)V Structural Integrity Associates, Inc.Attachment A to 1400669.40 1.ROA-2 a surface stress (input)E = local elastic modulus (input)Q = activation energy for initiation (input)R = universal gas constant (input)B = proportionality constant (input)q = environment-cold-work exponent (input)a,b,c,k = general cold work parameters (input)The inputs for the Palisades specific configurations are given below. Note that approximate values are taken for some of the physical properties of the material (E, yield strength, andultimate strength).

The actual values are close, and the differences here lead to insignificant changes in the results.

The B value is a distribution in the Reference 1 report as xLPR is aprobabilistic program.

For this deterministic

analysis, a conservative value was taken by usingthe 95% lower bound based on the inputs in Reference 1 (Appendix D).The B value is a log normal distribution with values given below [1, Appendix D]:Mean = 1.20e-9 hours -- 1.37e-13 years (is equal to eA-29.62)

Standard deviation (heat to heat) = 1.607Standard deviation (within heat) 0.555Overall standard deviation

= 1.6072 + 0.5552 = 1.70The 95% percentile is 1.65 standard deviations away from the mean.95th = -29.62 -(1.65*1.70)

= -32.425 -- e^-32.425

= 8.3e-15 (in units of years)T = 583°F (hot leg) (from body of report)T = 537°F (cold leg) (from body of report)Hot Leg Circumferential Surface Stress = -41 ksi (see Figure 8 in the body of this report)Cold Leg Circumferential Surface Stress = -45 ksi (see Figure 12 in the body of this report)E = 29e6 psi [3, at approximately 583°F (Hot Leg) and 537°F (Cold Leg)] the actual value at theoperating temperatures are slightly below 29e6 psi. The difference will have an insignificant impact on the results.Q/R = 22,000 K [1, Appendix D]B = 8.3e-15 years (calculated above)q = 0.375 [2, Table 5-4]a = 0.25 [2, Table 5-4]b -0.75 [2, Table 5-4]c -0.25 [2, Table 5-4]k 10 [2, Table 5-4]z= 0.35 [2, Table 5-4]Z2 = 0.333 [2, Table 5-4]v -0.66 [2, Table 5-4]w 0.5 [2, Table 5-4]Yield Strength

= 32,000 psi used (30,000 psi is the approximate value at 583°F (Hot Leg) [3] and30,200 psi is the approximate value at 537'F (Cold Leg) [3]). Using 32,000 psi is conservative as it reduces the time to initiation.

Ultimate Strength

= 80,000 psi [3, at approximately 583°F (Hot Leg) and 537°F (Cold Leg)]Attachment A to 1400669.401

.RO A-3 V Structural Integrity Associates, Inc.

Calculations The model was used with Palisades' specific information to calculate the time to initiation.

Forthe hot leg drain nozzle, the surface stress of 41 ksi (shown in Figure 8, page 15 of this report)and the hot leg temperature are used in the model to calculate the time to initiation.

Figure A-Ishows the plot for this case. The plot shows the initiation time for a given surface stress. For the41 ksi stress of the hot leg drain nozzle, the initiation time is roughly 130 years.For the bounding cold leg nozzle, the surface stress of 45 ksi (shown in Figure 12, page 19 ofthis report) and the cold leg temperature are used in the model to calculate the time to initiation.

Figure A-2 shows the plot for this case. The plot shows the initiation time for a given surfacestress. For the 45 ksi stress of the bounding cold leg nozzle, the initiation time is roughly 600years, 4.6 times longer than the hot leg drain nozzle despite the slightly higher surface stress.For comparison, various different temperatures are considered to put the hot leg drain nozzle andbounding cold leg nozzle in perspective with other typical PWR plant components that have seenactual cracking in the industry.

The Palisades hot leg temperature is fairly low when compared to other PWR plants in theindustry.

Other PWR plants typically operate at a hot leg temperature above 600'F. Forcomparison, a hot leg temperature of 604'F is evaluated and compared to the hot leg drain nozzleat Palisades.

The results of this comparison are presented in Figure A-3. Additionally, thehottest plant components are more susceptible to cracking.

For PWR plants, susceptible locations exist in the pressurizer that operates at around 653°F. Figure A-4 shows the plot forinitiation at the pressurizer temperature.

It should also be noted that most locations that have experienced cracking in the industry did notreceive a PWHT. Therefore, the residual stresses for these other locations would have higherresidual

stresses, on the order of-50 -60 ksi. Looking at these higher stresses from the hot legor pressurizer initiation plots shows that the time to initiation for these locations would beroughly 30 and 5 years for the hot leg and pressurizer, respectively (approximate values taken at55 ksi surface stress).Using the Palisades hot leg drain nozzle surface stress (41 ksi) for comparison, the higher hot legtemperature of 604'F would reduce the time to initiation by a factor of over 2 when compared tothe Palisades temperature

(-130 years vs. -60 years). Comparing the Palisades hot leg drainnozzle results to a pressurizer temperature shows a factor of-10 for time to initiation

(-130years vs. -12 years). This shows the extreme temperature dependence related to initiation.

Forthe Palisades hot leg drain nozzle, the lower temperature gives the location at least a factor of 2over other hot leg locations in PWR plants.Attachment A to 1400669.401

.RO A-4 Structural Integrity Associates, Inc.

References

1. Materials Reliability Program:

Models and Inputs Selected for Use in the xLPR PilotStudy (MRP-302),

EPRI, Palo Alto, CA: 2010, 1022528.2. Stress Corrosion Cracking Initiation Modelfor Stainless Steel and Nickel Alloys: Effectsof Cold Work, EPRI, Palo Alto, CA: 2009, 1019032.3. ASME Boiler and Pressure Vessel Code,Section II, Part D, "Material Properties,"

2004Editions with no Addenda.Attachment A to 1400669.401

.RO A-5 Structural Integrity Associates, Inc.

Time to Initiation 2502006.m150o 100a'E500-Initiation 30 35 40 45 50 55Stress, ksi60 65 70 75 80Figure A-1.Initiation Time for Palisades Hot Leg Drain Nozzle at the Hot LegTemperature of 583°F)669.401.RO A-6 structural Integrity Associates, Inc.Attachment A to 140(

Time to Initiation 14001200~800CS6004-E-1002000-- Initiation 30 35 40 45 50 55 60 65 70 75 80Stress, ksiFigure A-2. Initiation Time for Palisades Bounding Cold Leg Nozzle at the Cold LegTemperature of 537°FAttachment A to 1400669.40 L.RO A-7 Structural Integrity Associates, Inc.

Time to Initiation

EI-.120100806040200-Initiation 30 35 40 45 50 55 60 65 70 75 80Stress, ksiFigure A-3. Initiation Time for Typical Hot Leg Temperature of 604'Fnt A to 1400669.40 1.RO A-8 v Structural Integrity Associates, Inc.Attachme Time to Initiation 252015C4UEa,50-S0-Initiation 30 35 40 45 50 55 60 65 70 75 80Stress, ksiFigure A-4. Initiation Time for Typical Pressurizer Temperature of 653°FAttachment A to 1400669.401

.R0 A-9 Structuri Integrity Associates, Inc.

!Structural Integrity Associates, Inc.CALCULATION PACKAGEFile No.: 1400669.313 Project No.: 1400669Quality Program Type: Z Nuclear D Commercial PROJECT NAME:Palisades Flaw Readiness Program for 6R24 NDE Inspection CONTRACT NO.:10426669CLIENT: PLANT:Entergy Nuclear Operations, Inc. Palisades Nuclear PlantCALCULATION TITLE:Crack Growth Analysis of the Hot Leg Drain NozzleDocument Affected Project Manager Preparer(s)

&Revision Pages Revision Description Approval Checker(s)

Signature

& Date Signatures

& Date01 -23 Initial Issue Preparer:

A-I -A-3Computer Files ,,,,Norman EngNE 5/11/15 Minji FongMF 5/11/15Checkers:

Wilson WongWW 5/11/15Gole MukhimGSM 5/11/15Page 1 of 23F0306-01 R2 Vj tuaturI ltgry Assocaes, IncGTable of Contents1.0 O B JE C T IV E .........................................................................................................

42.0 D E SIG N IN PU T S ......................................................................................................

42.1 Piping Interface Loads .................................................................................

42.2 Residual Stresses at Normal Operating Temperature and Pressure

..............

52.3 Mechanical Load Boundary Conditions

......................................................

52.4 C rack G row th R ate ........................................................................................

53.0 A SSU M PT IO N S ......................................................................................................

64.0 DETERMINATION OF STRESS INTENSITY FACTOR ......................................

64.1 Crack Face Pressure Application

.................................................................

64.2 K Calculation for Circumferential Flaws .....................................................

74.2.1 Finite Element Model with Circumferential Flaws ......................................

74.2.2 Stress Intensity Factor Results ......................................................................

84.3 K Calculation for Axial Flaws ......................................................................

84.3.1 Finite Element Model with Axial Flaws ........................................................

84.3.2 Stress Intensity Factor Results ......................................................................

85.0 CRACK GROWTH CALCULATION

....................................................................

96.0 C O N C LU SIO N S .....................................................................................................

97.0 R E F E R E N C E S ............................................................................................................

11APPENDIX A COMPUTER FILES LISTING ...............................................................

A-1File No.: 1400669.313 Page 2 of 23Revision:

0F0306-01 R2 Stwabural NIg"ely AssociaMes, Inc."List of TablesTable 1: Stress Intensity Factors for Circumferential Flaws ............................................

12Table 2: Stress Intensity Factors for Axial Flaws ...............................................................

12Table 3: Crack Growth Time to 75% Through-Wall

........................................................

12Table 4: Crack Growth Time to 95% Through-Wall

........................................................

12Table 5: A llowable Detected Flaw Size ............................................................................

13List of FiguresFigure 1. Base Finite Element M odel M esh .....................................................................

14Figure 2. Applied Mechanical Load Boundary Conditions

...............................................

15Figure 3. Circumferential Flaw with Crack Tip Elements Inserted

....................................

16Figure 4. Transferred Residual Stress + NOC + Pressure Stress for Circumferential Flaws. 17Figure 5. Stress Intensity Factors as a Function of Depth for Circumferential Flaws ..... 18Figure 6. Axial Flaws with Crack Tip Elements Inserted

.................................................

19Figure 7. Transferred Residual Stress + NOC + Pressure Stress for Axial Flaws ........

20Figure 8. Stress Intensity Factors as a Function of Depth for Axial Flaws .......................

21Figure 9. Crack Growth for All Flaw Types with 0.025" Initial Flaw Size .....................

22Figure 10. Crack Growth for All Flaw Types with 0.1" Initial Flaw Size ........................

23File No.: 1400669.313 Revision:

0Page 3 of 23F0306-01R2 jjS eforul agrfly Associates, Inc:1.0 OBJECTIVE The objective of this calculation package is to determine maximum allowable flaw sizes for 18 and 36months of continued operation based on crack growth analyses for a series of postulated flaws in the hotleg drain nozzle boss weld in support of a Primary Water Stress Corrosion Cracking (PWSCC)susceptibility study at the Palisades Nuclear Plant (Palisades).

The stresses due to the hot leg pipeinterface loads which are determined in this calculation, and residual stresses extracted from a previousanalysis

[1] are used to calculate stress intensity factors (K) which are used to perform crack growthanalyses.

The PWSCC crack growth analyses are performed using the pc-CRACK

[2] program for bothcircumferential and axial flaws. The allowable detected flaw sizes are determined by back-calculating the predicted growth time to a maximum flaw size of 75% through wall thickness per ASME CodeSection XI, IWB-3643.

2.0 DESIGN INPUTSThe finite element model shown in Figure 1 was developed in Reference

[3] and is used for thedetermination of stress intensity factors.2.1 Piping Interface LoadsReference 4 (PDF file, page 88) indicates that, for the hot leg, the bounding thermal transient stress of1.0 10 ksi is due to case Thermal 002, the deadweight (DW) stress is 0.096 ksi and the friction stress is1.056 ksi. The hot leg loads are applied as an equivalent bending moment to the axial free end of themodeled hot leg. The equivalent bending moment is based on the combined stress which is assumed tooccur at the outside surface of the hot leg. The maximum combined bending stress is:DW + Friction

+ Thermal = 0.096 + 1.056 + 1.010 = 2.162 ksiThe moment based on the bending stress is calculated as:M -. n. (24.81254

-20.81254)

.2.162 =13099in--kips OR 4 24.8125where,M = Moment (in-kips)

G = Stress on hot leg pipe free end (ksi)I = Moment of Inertia -(7/4)(OR4 -IRa) (in4)IR = Inside radius of hot leg pipe (in) = 20.8125"

[3]OR = Outside radius of hot leg pipe (in) = 24.8125"

[3]Since half the hot leg pipe is modeled, the equivalent moment applied to the model is 6549.5 in-kips(= 13099 in-kips /2). The moment is applied to the axial end of the hot leg run piping by means of a pilotFile No.: 1400669.313 Page 4 of 23Revision:

0F0306-01 R2 V Smnflrrl Mgily Associates, IncWnode pair to transfer the loading.

The pilot node pair is composed of a target node at the center of thepipe (ANSYS TARGE 170 element) and a set of surface contact elements on the axial end of the pipe(ANSYS CONTA 174 element).

The surface elements are bonded to the pilot node in a slave/master coupling relationship, so that the moment load applied to the pilot node is transferred to the end of thepipe. The hot leg drain nozzle piping loads are considered to have negligible effects on the resulting K'sfor the boss weld, and are therefore not considered.

2.2 Residual Stresses at Normal Operating Temperature and PressureResidual stresses at the fifth operating condition cycle (at time = 2484 minutes) are taken fromReference

1. These stresses include the effects of normal operating temperature of 583'F and pressure of2085 psig [1].2.3 Mechanical Load Boundary Conditions The mechanical load boundary conditions for the stress analysis are symmetric boundary conditions atthe symmetry planes of the model, axial displacement restraint at the end of the nozzle, and axialdisplacement restraint on the pilot node, as shown in Figure 2. In the case where axial flaws are modeledon the symmetry planes, the boundary conditions are released at the nodes where the flaw exists.2.4 Crack Growth RateThe default PWSCC crack growth rate in pc-CRACK

[2] is used. This relation is based on expressions in Reference

[5, Section 4.3] and the resulting equation for the crack growth rate is as follows:da _____1 T+460 _1__--d= Cexp!-Q I r +460 (K-Ki)P for K > Kthcit re60 ,f +6)For times (t) in hours, temperatures (T and Tef) in 'F, crack length (a) in inches and K in ksi-/in, thefollowing reference values are used:Tref = 617°FC = 2.47x 10-7 (constant) 13 = 1.6 (constant)

Q = 28181.8°R (constant)

Kth = 0 (threshold stress intensity factor below which there is no crack growth)T = operating temperature at location of crackFile No.: 1400669.313 Page 5 of 23Revision:

0F0306-01 R2 V0SbvfkwI hwgyAsodWV~kn

3.0 ASSUMPTIONS

The following assumptions are used in this analyses:

" The hot leg drain nozzle piping loads are not considered in calculating stress intensity factorssince loads on the drain nozzle do not produce Mode I crack opening stress intensity factors thatcontribute to crack growth in the boss weld.* The maximum combined stress on the hot leg piping is assumed to occur at the outside surface ofthe hot leg.4.0DETERMINATION OF STRESS INTENSITY FACTORThe stresses described in this section are used with a modified version of the finite element model(FEM) developed previously in Reference

[3] to determine stress intensity factors.

The modification ofthe FEM consists of adding crack tip elements as addressed in Section 4.2 and 4.3. The stress intensity factors (Ks) are calculated using the KCALC feature in ANSYS [6] which is based on linear elasticfracture mechanics (LEFM) principles.

For the LEFM evaluations, only the elastic properties are used inthe FEA.4.1 Crack Face Pressure Application In order to determine the Ks for the circumferential and axial flaws due to residual

stresses, the stresseson the boss weld-to-nozzle interface, at the fifth operating condition (at time = 2484 minutes in theresidual stress analysis

[1]), are extracted from the residual stress analysis and reapplied on the crackface as surface pressure loading.This approach is based on the load superposition principle

[7], which is utilized to transfer the stressesfrom the weld residual stress finite element model onto the fracture mechanics finite element model thatcontains crack tip elements.

The superposition technique is based on the principle that, in the linearelastic regime, stress intensity factors of the same mode, which are due to different loads, are additive(similar to stress components in the same direction).

The superposition method can be summarized with the following sketches

[7, page 66]:-~~~~~~r

-rN- ý P()- ý -P(a)(b) (c) (d)File No.: 1400669.313 Revision:

0Page 6 of 23F0306-01 R2 jjsflhrw latIegrli Associad s, Inc@A load p(x) on an uncracked body, Sketch (a), produces a normal stress distribution p(x) on Plane A-B.The superposition principle is illustrated by Sketches (b), (c), and (d) of the same body with a crack atPlane A-B. The stress intensity factors resulting from these loading cases are such that:Ki(b) = Ki(c) + Ki(d)Thus, Ki(d) = 0 because the crack is closed, and:Ki(b) = Ki(c)This means that the stress intensity factor obtained from subjecting the cracked body to a nominal loadp(x) is equal to the stress intensity factor resulting from loading the crack faces with the same stressdistribution p(x) at the same crack location in the uncracked body.4.2 K Calculation for Circumferential Flaws4.2.1 Finite Element Model with Circumferential FlawsThe stress intensity factors for full circumferential flaws in the nozzle boss weld are determined by finiteelement analysis using deterministic linear elastic fracture mechanics (LEFM) principles.

As a result,five fracture mechanics finite element models are derived to include "collapsed" crack meshing thatrepresent full (3600) circumferential flaws surrounding the nozzle at various depths within the bossweld.The circumferential flaws align with the interface between the boss weld and the nozzle. The modeledcrack depths are: 0.13", 1.17", 2.05", 3.13", and 3.97" as measured at the 0' axial side of the hot legpipe.The modeling of the flaws, or cracks, involves splitting the crack plane and then inserting "collapsed" mesh around the crack tips followed by concentrated mesh refinements that surround the "collapsed" mesh, and are referred to as "crack tip elements".

This step is implemented on a source finite elementmodel without the cracks (the FEM developed in Reference

3) where the crack tip elements are insertedby an in-house developed ANSYS macro.For the fracture mechanics models, 20-node quadratic solid elements (ANSYS SOLID95) are used in thecrack tip region, while 8-node solid elements (ANSYS SOLID 185) are used everywhere else in themodel. The mid-side nodes for the SOLID95 elements around the crack tips are shifted to the "quarterpoint" locations to properly capture the singularities at the crack tips, consistent with ANSYSrecommendations.

The finite element model for the 3.97" deep circumferential flaw, with the crack tipmesh, is shown in Figure 3 as an example; the crack tip mesh for the other crack depths follows the samepattern.File No.: 1400669.313 Page 7 of 23Revision:

0F0306-01R2 C OMsM-, IN9rM Associs, kmiThe quarter point mid-side nodes combined with the extra layers of concentrated elements around thecrack tips provide sufficient mesh refinement to determine the stress intensity factors for the fracturemechanics analyses.

4,2.2 Stress Intensity Factor ResultsThe radial stresses (radial to the nozzle axis) on the weld/nozzle interface are transferred to thecircumferential flaws as crack face pressure per the superposition principle described in Section 4.1.Figure 4 depicts, as an example, the transferred radial stresses as crack face pressure for the 3.97"circumferential flaw depth. During the crack face pressure

transfer, the operating pressure of2085 psi [1] is added to the crack face pressure to account for the internal pressure acting on the crackface due to cracking.

A far field in-plane bending moment per Section 2.1 is also applied to the free endof the hot leg run piping to account for piping moments in the main loop piping.Each crack model is analyzed as a steady state stress pass at the operating and reference temperature of583'F [1] in order to use the material properties at the operating temperature, but without inducingadditional thermal stresses.

At the completion of each analysis, the ANSYS KCALC post-processing is performed to extract the K'sat each crack tip node around the nozzle. The maximum K results are summarized in Table 1 for variouscrack depth ratios "a/t". Since the crack tip location is same in the circumferential flaw, the maximum Kfrom all locations at each crack size is conservatively used for the K vs. a profile.

The "K vs. a/t" trendsare then plotted in Figure 5.4.3 K Calculation for Axial Flaws4.3.1 Finite Element Model with Axial FlawsThe stress intensity factors for axial flaws are determined using the same methodology as thecircumferential flaws. However, the mesh of weld nuggets was removed to insert thumbnail shape flawsin the model. Also, the orientation and shape of the flaws allow all crack depths at the 00 and 900 facesof the symmetric hot leg pipe model to be inserted simultaneously.

Figure 6 shows the five modeledcrack depths (0.25", 1.06", 1.90", 2.91", and 3.85") on the 00 face (hot leg axial face) and the 900 facewith crack tip elements inserted.

The modeling of the axial flaws uses the same crack tip elements as described in Section 4.2.1. Thecrack tip mesh is the same pattern used in the circumferential flaws and is shown in Figure 6 for theaxial flaws at the 0' and 90' faces.4.3.2 Stress Intensity Factor ResultsSimilar to the circumferential flaw analyses, the crack opening residual stresses and additional operating pressure are transferred to the axial flaws as crack face pressure.

Figure 7 depicts, as an example, theFile No.: 1400669.313 Page 8 of 23Revision:

0F0306-01R2 Sanwbur Iu griy Assocats, IncWtransferred hoop stresses as crack face pressure for the axial flaws. In addition, a far field in-planebending moment per Section 2.1 is applied to the free end of the hot leg run piping to account for pipingmoments in the main loop piping. The K results at the deepest point of the flaws are summarized inTable 2 for various crack depth ratios "a/t" and plotted in Figure 8. Since the deepest point of thepostulated axial flaws has the smallest remaining wall thickness, the K at the deepest point is used forthe K vs. a profile.5.0 CRACK GROWTH CALCULATION Stress intensity factors (Ks) at four depths for 3600 inside surface connected, part-through-wall circumferential flaws as well as two axial thumbnail flaws at the 0-and 90-degree azimuthal locations ofthe nozzle, are calculated using finite element analysis (FEA). For the circumferential flaw, themaximum K values around the nozzle circumference for each flaw depth are extracted and used as inputinto pc-CRACK to perform the PWSCC crack growth analyses.

For the axial flaws, the K at the deeppoint of the thumbnail shape is used as input for performing the PWSCC crack growth analyses.

Sincethe K vs. a profile is used as input, the shape of the component is not relevant.

For the crack growth analyses, two initial flaw sizes are chosen. These are based on expectedengineering flaw sizes that could be present for a flaw that would then grow by PWSCC. The final flawsize for these analyses is 75% of the wall thickness.

This final depth is chosen as it is the maximumallowable flaw depth per Section XI of the ASME Code for pipe flaw evaluations.

Additionally, a finalflaw size of 95% of the wall thickness is also considered in this calculation.

The key parameters used in the crack growth calculations included:

Two initial crack depths = 0.025" and 0.1" (assumed)

Temperature

= 583°F (operating temperature

[1])Wall thickness

= 4.0" (hot leg pipe thickness

[3])The resulting crack depths for the circumferential and axial flaws, as a function of time, as calculated bypc-CRACK are shown in Figure 9 for the 0.025" initial flaw size, and Figure 10 for the 0.1" initial flawsize. The time for a flaw to grow from the initial flaw size to 75% and 95% through-wall is tabulated inTable 3 and Table 4 for both circumferential and axial flaw types, respectively.

Table 5 shows themaximum allowable detected flaw sizes for the postulated flaws if continued operation for 18 and 36months is considered.

6.0 CONCLUSION

S Stress intensity factors were calculated for the 3600 circumferential flaws as well as the axial flaws at the00 and 900 locations.

The stress intensity factors were calculated using residual stress distributions forresidual stress plus normal operating conditions.

In addition, a far field in-plane bending moment isFile No.: 1400669.313 Page 9 of 23Revision:

0F0306-01 R2 Vj ufurb l lategrity Associates inc."applied to the free end of the hot leg run piping to account for piping moments in the main loop piping.This combined loading is used for the determination of the stress intensity factors for both thecircumferential and axial flaws. Figure 5 and Figure 8, as well as Table 1 and Table 2, show thecalculated stress intensity factors for the circumferential and axial flaws.Crack growth evaluations were performed for circumferential and axial flaw configurations using twodifferent initial flaw sizes. As shown in Figure 9 and Table 3, the shortest time for an initial 0.025" deepflaw to grow to 75% through-wall in all cases is 30.5 years for an axial flaw on the 00 plane. Figure 10and Table 3 show that the shortest time for an initial 0. 1" deep flaw to grow to 75% through-wall in allcases is 29.7 years for an axial flaw on the 0' plane. Table 5 shows the maximum allowable detectedflaw sizes for 18 and 36 months continued operation.

File No.: 1400669.313 Revision:

0Page 10 of 23F0306-01 R2 Vjsb7aJhI late rit y Associates, InoG

7.0 REFERENCES

1. SI Calculation No. 1400669.312, Rev. 0, "Hot Leg Drain Nozzle Weld Residual StressAnalysis."

2. pc-CRACK 4.1, Version 4.1 CS, Structural Integrity Associates, December 2013.3. SI Calculation No. 1400669.310, Rev. 0, "Finite Element Model for Hot Leg Drain Nozzle."4. Palisades

Document, Report No. CENC- 1115, "Analytical Report for Consumers Power Piping,"SI File No. 1300086.204.

5. Materials Reliability Program:

Crack Growth Rates for Evaluating Primary Water StressCorrosion Cracking (PWSCC) ofAlloy 82, 182 and 132 Welds (MRP-1 15), EPRI, Palo Alto, CA:2004, 1006696.6. ANSYS Mechanical APDL and PrepPost, Release 14.5 (w/ Service Pack 1), ANSYS, Inc.,September 2012.7. Anderson, T. L., "Fracture Mechanics Fundamentals and Applications,"

Second Edition, CRCPress, 1995.File No.: 1400669.313 Revision:

0Page 11 of 23F0306-01R2 ntJ rulugriti Associates, IncOTable 1: Stress Intensity Factors for Circumferential FlawsCrack Depth aft Max. K(in) (ksi-in°)

0.13 0.03 19.611.12 0.28 22.762.04 0.51 14.783.15 0.79 13.033.97 0.99 35.15Table 2: Stress Intensity Factors for Axial FlawsHL Axial Plane 00 HL Circ. Plane 90°Crack Depth K at Deep Pt Crack Depth K at Deep Pt(in) a/t (ksi-in°'

5) (in) a/t (ksi-in°'5) 0.25 0.06 18.50 0.25 0.06 14.161.06 0.26 18.64 1.10 0.27 14.431.90 0.47 17.60 1.96 0.49 14.772.91 0.73 19.34 2.95 0.74 16.983.85 0.96 23.44 3.81 0.95 21.41Table 3: Crack Growth Time to 75% Through-Wall Table 4: Crack Growth Time to 95% Through-Wall Initial Flaw Size Axial Crack Axial Crack Circ. Crack(in) (00 plane) (900 plane) (years)(years) (years) (years)0.025 36.7 50.3 42.10.100 36.0 49.2 41.4File No.: 1400669.313 Revision:

0Page 12 of 23F0306-01 R2 SfncktraI hlfgrnly AssociMes, c."Table 5: Allowable Detected Flaw SizeAllowable Detected Flaw Size (a/t)Hot Leg Thickness, t = 4.00"Months of Axial Flaw Axial Flaw Circumferential Continued at 00 plane at 900 plane FlawOperation a/t a (in) a/t a (in) a/t a (in)18 0.7070 2.83 0.7165 2.87 0.7260 2.9036 0.6690 2.68 0.6785 2.71 0.6975 2.79File No.: 1400669.313 Revision:

0Page 13 of 23F0306-01 R2 Cjbabuu"

&o* wadEYANSYSR14.5Figure 1. Base Finite Element Model MeshFile No.: 1400669.313 Revision:

0Page 14 of 23F0306-OIR2 UCoupled to pilot nodeSymmetry boundary conditions Figure 2. Applied Mechanical Load Boundary Conditions File No.: 1400669.313 Revision:

0Page 15 of 23F0306-01R2 Van" rMW AbSOGWM Aso ftes.Figure 3. Circumferential Flaw with Crack Tip Elements Inserted(Note: Deepest circumferential flaw shown for example)File No.: 1400669.313 Revision:

0Page 16 of 23F0306-01 R2 Van" MWW AMOWN, WNFigure 4. Transferred Residual Stress + NOC + Pressure Stress for Circumferential Flaws(Note: Deepest circumferential flaw shown for example)File No.: 1400669.313 Revision:

0Page 17 of 23F0306-01 R2 VjbObiW ,atgrl Awsdaf WIn"rI40353025201510500.0 0.2 0.4 0.6 0.8 1.0Depth (a/t)Figure 5. Stress Intensity Factors as a Function of Depth for Circumferential FlawsFile No.: 1400669.313 Revision:

0Page 18 of 23F0306-01R2

!VftWM1MAMW~fnc 0°Figure 6. Axial Flaws with Crack Tip Elements InsertedFile No.: 1400669.313 Revision:

0Page 19 of 23F0306-01R2 Can"A~pfy AdMMkicP-KCALC analy,) 0...... -- -1.96433

---

6.993 ..KCALC analysis usincq invorted crack faoe stre~sOAoFigure 7. Transferred Residual Stress + NOC + Pressure Stress for Axial FlawsFile No.: 1400669.313 Revision:

0Page 20 of 23F0306-01 R2 Van" MW* Awad0s, knO-12520rbi1510501L-I-0.0 0.2 0.4 0.6 0.8 1.0Depth (a/t)Figure 8. Stress Intensity Factors as a Function of Depth for Axial FlawsFile No.: 1400669.313 Revision:

0Page 21 of 23F0306-01R2 Van" MobW*hwf Assocites, Wc1.00.90.80.70.640.5cv 0.40.30.20.10.0IICirc. Crack---Axial Crack_0-.- Axial Crack_90500 10 20 30 40Time (years)60Figure 9. Crack Growth for All Flaw Types with 0.025" Initial Flaw SizeFile No.: 1400669.313 Revision:

0Page 22 of 23F0306-01 R2I V3SbvtW Mlatgnfy

Associat, Ina10.90.80.7~0.65r0.40.30.2///Circ. Crack---Axial Crack_0---Axial Crack_90500.1 AV .0010 2030Time (years)4060Figure 10. Crack Growth for All Flaw Types with 0.1" Initial Flaw SizeFile No.: 1400669.313 Revision:

0Page 23 of 23F0306-01R2 5StunwlraIbgrfty

Assocaes, inc.APPENDIX ACOMPUTER FILES LISTINGFile No.: 1400669.313 Revision:

0Page A- I of A-3F0306-01R2 10 OSV19brw.

litgrit y Assocates, kIWOFile Name Description Palisades HL Drain.INP Input file to create base geometry model [3]MPropMISO.INP Elastic-plastic material properties inputs [3]BCNODES.INP Component setting for the boundary conditions.

Crack tip node inputs for fracture mechanics model conversion forNODESC##.INP circumferential flaws. ##=03, 30, 50, 75, and 95 with 03 = 0.13", 30 =1.17", 50 = 2.05, 75 = 3.13", 95 = 3.97" of flaw sizeFMHLC##.INP Geometry input files to create circumferential flaw at specified depth.##=03, 30, 50, 75, and 95FM HL C## COORD.INP Input files to determine circumferential crack face element centroidFMHL- Ccoordinates.

1. 1. =03, 30, 50, 75, and 95FMHLC## COORDl.txt Circumferential crack face element centroid coordinate outputs.##=03, 30, 50, 75, and 95FM HL C## GETSTR.INP Input files to extract circumferential crack face stresses from residualF H C_ T RNstress analysis.

1. 1. =03, 30, 50, 75, and 95Extracted circumferential crack face stresses from residual stressSTR_FieldOperC##1.txt analysis.

1. 1. =03, 30, 50, 75, and 95Input files to transfer stresses into circumferential crack face pressureFMHLC##_IMPORT.INP (plus operating pressure on crack face and applied pipe moment).##=03, 30, 50, 75, and 95Formatted K result outputs for circumferential flaws. ##=03, 30, 50,FMHLC##_IMPORTK.CSV 7,ad9.... 75, and 95HLAXIAL.INP Input file to modify base mesh for axial crack tip insertion Axial* *-Nodes.INP Crack tip node inputs for fracture mechanics model conversion foraxial flaws. **=00 and 90 with 00 = 0' plane and 90 = 900 planeFMHLAXL**.[NP Geometry input files to create axial flaws on the plane. **-= 00 and 90FMHLAXL**_COORD.INP Input files to determine axial crack face element centroid coordinates.

F -**=00, and 90FMHLAXL**_COORDI

.txt Axial crack face element centroid coordinate outputs.

• • =00, and 90FMHLAXL**

GETSTR.INP Input files to extract axial crack face stresses from residual stressanalysis.

• • =00, and 90Extracted axial crack face stresses from residual stress analysis.

STR_FieldOperAXL**1.txt

• • =00, and 90Input files to transfer stresses into axial crack face pressure (plusFMHL_AXL**_IMPORT.INP operating pressure on crack face and applied pipe moment).

• • =00,and 90FMHLAXL**_IMPORTK.CSV Formatted K result outputs for axial flaws. **=00, and 90AnTip8 I_KCALC.INP KCALC post-processing input filepc-CRACK PWSCC growth input file for circumferential flaw.CirFlaw $$$$.pcf

$$$$=0.025 and 0.1 with 0.025 = 0.025" and 0.1 = 0. 1" initial flawsizeFile No.: 1400669.313 Revision:

0Page A-2 of A-3F0306-01 R2 CjSbiIaUrIN Iate r Assocates, IreFile Name Description AxialFlaw 0_$$$$.pcf pc-CRACK PWSCC growth input file for axial flaw on 0' plane.$$$$=0.025 and 0.1AxialFlaw 90_$$$$.pcf pc-CRACK PWSCC growth input file for axial flaw on 900 plane.$$$$=0.025 and 0.1CirFlaw_$$$$.rpt pc-CRACK PWSCC growth output file for circumferential flaw.$$$$=0.025 and 0.1AxialFlaw 0 $$$$.rpt pc-CRACK PWSCC growth output file for axial flaw on 00 plane.$$$$=0.025 and 0.1AxialFlaw 90_$$$$.rpt pc-CRACK PWSCC growth output file for axial flaw on 900 plane.$$$$=0.025 and 0.1File No.: 1400669.313 Revision:

0Page A-3 of A-3F0306-01 R2

Integrity Associates, Incy File No.: 1400669.323 v t Project No.: 1400669CALCULATION PACKAGE Quality Program Type: Z Nuclear EL Commercial PROJECT NAME:Palisades Flaw Readiness Program for 1 R24 NDE Inspection CONTRACT NO.:10426669CLIENT: PLANT:Entergy Nuclear Operations, Inc. Palisades Nuclear PlantCALCULATION TITLE:Crack Growth Analysis of the Cold Leg Bounding NozzleDocument Affected Project Manager Preparer(s)

&Revision Pages Revision Description Approval Checker(s)

Signature

& Date Signatures

& Date0 1-23 Initial Issue Preparer:

A-i -A-2Computer FilesNorman Eng Wilson WongNE 5/11/15 WW 5/11/15Checkers:

Minji FongMF 5/11/15Gole MukhimGSM 5/11/15Page 1 of 23F0306-01 R2 VjSbvoirui laMgdfy Associaes, Inc.Table of Contents1.0 O B JE C T IV E .........................................................................................................

42.0 D E SIG N IN PU T S ......................................................................................................

42.1 Piping Interface Loads .................................................................................

42.2 Residual Stresses at Normal Operating Temperature and Pressure

..............

52.3 Mechanical Load Boundary Conditions

......................................................

52.4 C rack G row th R ate ........................................................................................

53.0 A SSU M PT IO N S ......................................................................................................

64.0 DETERMINATION OF STRESS INTENSITY FACTOR .....................................

64.1 Crack Face Pressure Application

..................................................................

64.2 K Calculation for Circumferential Flaws .....................................................

74.2.1 Finite Element Model with Circumferential Flaws ......................................

74.2.2 Stress Intensity Factor Results ......................................................................

84.3 K Calculation for Axial Flaws ......................................................................

84.3.1 Finite Element Model with Axial Flaws ........................................................

84.3.2 Stress Intensity Factor Results ......................................................................

85.0 CRACK GROWTH CALCULATION

....................................................................

96.0 C O N C L U SIO N S .....................................................................................................

97.0 R E FE R E N C E S .......................................................................................................

11.... 1APPENDIX A COMPUTER FILE LISTING ......................................................................

A-IFile No.: 1400669.323 Page 2 of 23Revision:

0F0306-01R2 jS"ncfurWa gritry ASSocts, Wnc;List of TablesTable 1: Stress Intensity Factors for Circumferential Flaws ............................................

12Table 2: Stress Intensity Factors for Axial Flaws .............................................................

12Table 3: Crack Growth Time to 75% Through-Wall

........................................................

12Table 4: Crack Growth Time to 95% Through-Wall

........................................................

12Table 5: Allowable Detected Flaw Size ............................................................................

13List of FiguresFigure 1. Base Finite Elem ent M odel M esh ........................................................................

14Figure 2. Applied Mechanical Load Boundary Conditions

...............................................

15Figure 3. Circumferential Flaw with Crack Tip Elements Inserted

....................................

16Figure 4. Transferred Residual Stress + NOC + Pressure Stress for Circumferential Flaws.. 17Figure 5. Stress Intensity Factors as a Function of Depth for Circumferential Flaws ...... 18Figure 6. Axial Flaws with Crack Tip Elements Inserted

...................................................

19Figure 7. Transferred Residual Stress + NOC + Pressure Stress for Axial Flaws .............

20Figure 8. Stress Intensity Factors as a Function of Depth for Axial Flaws ........................

21Figure 9. Crack Growth for All Flaw Types with 0.025" Initial Flaw Size ......................

22Figure 10. Crack Growth for All Flaw Types with 0.1" Initial Flaw Size .........................

23File No.: 1400669.323 Revision:

0Page 3 of 23F0306-OIR2 Vjshncbunw hit gnfiy Assocates, liec1.0 OBJECTIVE The objective of this calculation package is to determine maximum allowable flaw sizes for 18 and 36months of continued operation based on crack growth analyses for a series of postulated flaws in thecold leg bounding nozzle boss weld in support of a Primary Water Stress Corrosion Cracking (PWSCC)susceptibility study at the Palisades Nuclear Plant (Palisades).

The stresses due to the cold leg pipeinterface loads which are determined in this calculation, and residual stresses extracted from a previousanalysis

[1], are used to calculate the stress intensity factors (K) which are used to perform crack growthanalyses.

The PWSCC crack growth analyses are performed using the pc-CRACK

[2] program for bothcircumferential and axial flaws. The allowable detected flaw sizes are determined by back-calculating the predicted growth time to a maximum flaw size of 75% through wall thickness per ASME CodeSection XI, IWB-3643.

2.0 DESIGN INPUTSThe finite element model shown in Figure 1 was developed in Reference

[3] and is used for thedetermination of stress intensity factors.2.1 Piping Interface LoadsReference 4 [PDF file page 88] indicates that, for the cold leg, the bounding thermal transient stress is7.307 ksi due to case Thermal 009, the deadweight (DW) stress is 0.459 ksi and the friction stress is0.429 ksi. The cold leg loads are applied as an equivalent bending moment to the axial free end of themodeled cold leg. The equivalent bending moment is based on the combined stress which is assumed tooccur at the outside surface of the cold leg. The maximum combined bending stress is:DW + Friction

+ Thermal = 0.459 + 0.429 + 7.307 = 8.195 ksiThe moment based on the bending stress is calculated as:M .-I ;r (17.843754

-14.843754).8.195 M=......

19056 in-kipsOR 4 17.84375where,M = moment applied to the free end of the cold lega= stress on the cold leg pipeI = Moment of Inertia -(it/4)(OR 4 -IR4)IR = Inside radius of nozzle (in) = 14.84375"

[3]OR = Outside radius of nozzle (in) = 17.84375

[3]File No.: 1400669.323 Page 4 of 23Revision:

0F0306-01R2 VjtI ntwokrh laterfiy Associats, IncPSince half the cold leg pipe is modeled, the equivalent moment applied to the model is 9528 in-kips(= 19056 in-kips /2). The moment is applied to the axial free end of the cold leg run piping by means ofa pilot node pair to transfer the loading.

The pilot node pair is composed of a target node at the center ofthe pipe (ANSYS TARGE 170 element) and a set of surface contact elements on the axial end of the pipe(ANSYS CONTA 174 element).

The surface elements are bonded to the pilot node in a slave/master coupling relationship, so that the moment load applied to the pilot node is transferred to the end of thepipe. The cold leg bounding nozzle piping loads are considered to have negligible effects on theresulting K's for the boss weld, and are therefore not considered.

2.2 Residual Stresses at Normal Operating Temperature and PressureResidual stresses at the fifth operating condition cycle (at time = 2106 minutes) are taken fromReference

[1]. These stresses include the effects of normal operating temperature of 5370F and pressureof 2085 psig [1].2.3 Mechanical Load Boundary Conditions The mechanical load boundary conditions for the stress analyses are symmetric boundary conditions atthe symmetry planes of the model, axial displacement restraint at the end of the nozzle, and axialdisplacement restraint on the pilot node, as shown in Figure 2. In the case where axial flaws are modeledon the symmetry planes, the boundary conditions are released at the nodes where the flaw exists.2.4 Crack Growth RateThe default PWSCC growth rate in pc-CRACK

[2] is used. This relation is based on expressions inReference

[5, Section 4.3] and the resulting equation for the crack growth rate is as follows:=I(for Ke> Kh'dt= T+460 T,.,f +460)(-])P frK tFor times (t) in hours, temperatures (T and Trej) in 'F, crack length (a) in inches and K in ksi4in, thefollowing reference values are used:Tref 617'FC = 2.47x 1 0-7 (constant)

= 1.6 (constant)

Q = 28181.8°R (constant)

Kth = 0 (threshold stress intensity factor below which there is no crack growth)T = operating temperature at location of crackFile No.: 1400669.323 Page 5 of 23Revision:

0F0306-OIR2 Can"kv hME Assodaf fte3.0 ASSUMPTIONS The following assumptions are used in this analyses:

" The cold leg bounding nozzle piping loads are not considered in calculating stress intensity factors since loads on the nozzle do not produce Mode I crack opening stress intensity factorsthat contribute to crack growth in the boss weld." The maximum combined stress on the cold leg piping is assumed to occur at the outside surfaceof the cold leg.4.0 DETERMINATION OF STRESS INTENSITY FACTORThe stresses described in this section are used with a modified version of the finite element model(FEM) developed previously in Reference

[3] to determine stress intensity factors.

The modification ofthe FEM consists of adding crack tip elements as addressed in Section 4.2 and 4.3. The stress intensity factors (Ks) are calculated using the KCALC feature in ANSYS [6] which is based on the linear elasticfracture mechanics (LEFM) principles.

For the LEFM evaluations, only the elastic properties are used inthe FEA.4.1 Crack Face Pressure Application In order to determine the Ks for the circumferential and axial flaws due to residual

stresses, the stresseson the boss weld-to-nozzle interface, at the fifth operating condition (at time = 2106 minutes in theresidual stress analysis

[1 ]), are extracted from the residual stress analysis and reapplied on the crackface as surface pressure loading.This approach is based on the load superposition principle

[7], which is utilized to transfer the stressesfrom the weld residual stress finite element model onto the fracture mechanics finite element model thatcontains crack tip elements.

The superposition technique is based on the principle that, in the linearelastic regime, stress intensity factors of the same mode, which are due to different loads, are additive(similar to stress components in the same direction).

The superposition method can be summarized with the following sketches

[7, page 66]:P(x) 7(x) P7()y(a) (b) (c) (d)File No.: 1400669.323 Page 6 of 23Revision:

0F0306-01R2 V Stru lua/ tfrit, Associates, IncOA load p(x) on an uncracked body, Sketch (a), produces a normal stress distribution p(x) on Plane A-B.The superposition principle is illustrated by Sketches (b), (c), and (d) of the same body with a crack atPlane A-B. The stress intensity factors resulting from these loading cases are such that:Ki(b) = Ki(c) + Ki(d)Thus, Ki(d) = 0 because the crack is closed, and:Ki(b) = Ki(c)This means that the stress intensity factor obtained from subjecting the cracked body to a nominal loadp(x) is equal to the stress intensity factor resulting from loading the crack faces with the same stressdistribution p(x) at the same crack location in the uncracked body.4.2 K Calculation for Circumferential Flaws4.2.1 Finite Element Model with Circumferential FlawsThe stress intensity factors for full circumferential flaws in the nozzle boss weld are determined by finiteelement analysis using deterministic linear elastic fracture mechanics (LEFM) principles.

As a result,five fracture mechanics finite element models are derived to include "collapsed" crack meshing thatrepresent full (3600) circumferential flaws surrounding the nozzle at various depths within the bossweld.The circumferential flaws align with the interface between the boss weld and the nozzle. The modeledflaw depths are: 0.13", 0.88", 1.45", 2.32", and 2.99" as measured at the 0' axial side of the cold legpipe.The modeling of the flaws, or cracks, involves splitting the crack plane and then inserting "collapsed" mesh around the crack tips followed by concentrated mesh refinements that surround the "collapsed" mesh, and are referred to as "crack tip elements".

This step is implemented on a source finite elementmodel without the cracks (the FEM developed in Reference

3) where crack tip elements are inserted byan in-house developed ANSYS macro.For the fracture mechanics models, 20-node quadratic solid elements (ANSYS SOLID95) are used in thecrack tip region, while 8-node solid elements (ANSYS SOLID 185) are used everywhere else in themodel. The mid-side nodes for the SOLID95 elements around the crack tips are shifted to the "quarterpoint" locations to properly capture the singularities at the crack tips, consistent with ANSYSrecommendations.

The finite element model for the 2.99" deep circumferential flaw, with the crack tipmesh, is shown in Figure 3 as an example; the crack tip mesh for the other crack depths follows the samepattern.File No.: 1400669.323 Page 7 of 23Revision:

0F0306-01R2 CjjS"nbrI Jut NW*t Associates, klceThe quarter point mid-side nodes combined with the extra layers of concentrated elements around thecrack tips provide sufficient mesh refinement to determine the stress intensity factors for the fracturemechanics analyses.

4.2.2 Stress Intensity Factor ResultsThe radial stresses (radial to the nozzle axis) on the weld/nozzle interface are transferred to thecircumferential flaws as crack face pressure per the superposition principle described in Section 4.1.Figure 4 depicts, as an example, the transferred radial stresses as crack face pressure for the 2.99"circumferential crack depth. During the crack face pressure

transfer, the operating pressure of 2085 psi isadded to the crack face pressure to account for the internal pressure acting on the crack face due tocracking.

A far field in-plane bending moment per Section 2.1 is also applied to the free end of the coldleg run piping to account for the pipe moment in the main loop piping.Each crack model is analyzed as a steady state stress pass at the operating and reference temperature of537'F [1] in order to use the material properties at the operating temperature, but without inducingadditional thermal stresses.

At the completion of each analysis, the ANSYS KCALC post-processing is performed to extract the K'sat each crack tip node around the nozzle. The maximum K results are summarized in Table I for variouscrack depth ratios "a/t". Since the crack tip location is same in the circumferential flaw, the maximum Kfrom all locations at each crack size is conservatively used for the K vs. a profile.

The "K vs. a/t" trendsare then plotted in Figure 5.4.3 K Calculation for Axial Flaws4.3.1 Finite Element Model with Axial FlawsThe stress intensity factors for axial flaws are determined using the same methodology as thecircumferential flaws. However, the mesh of weld nuggets was removed to insert thumbnail shape flawsin the model. Also, the orientation and shape of the flaws allow all crack depths at the 00 and 900 facesof the symmetric cold leg pipe model to be inserted simultaneously.

Figure 6 shows the five modeledcrack depths (0.25", 0.78", 1.37", 2.16", and 2.90") on the 0' face (cold leg axial face) and 90' face withcrack tip elements inserted.

The modeling of the axial flaws uses the same crack tip elements as described in Section 4.2.1. Thecrack tip mesh is the same pattern used in the circumferential flaws and is shown in Figure 6 for theaxial flaws at the 0' and 90' faces.4.3.2 Stress Intensity Factor ResultsSimilar to the circumferential flaw analyses, the crack opening residual stresses and additional operating pressure are transferred to the axial flaws as crack face pressure.

Figure 7 depicts, as an example, theFile No.: 1400669.323 Page 8 of 23Revision:

0F0306-OIR2 Vjm$n fcbruI lategrfy Associaes, IWO~transferred hoop stresses as crack face pressure for the axial flaws. In addition, a far field in-planebending moment per Section 2.1 is also applied to the free end of the cold leg run piping to account forthe pipe moment in the main loop piping. The K results at the deepest point of the flaws are summarized in Table 2 for various crack depth ratios "a/t" and plotted in Figure 8. Since the deepest point of thepostulated axial flaws has the smallest remaining wall thickness, the K at the deepest point is used forthe K vs a profile.5.0 CRACK GROWTH CALCULATION Stress intensity factors (Ks) at four depths for 3600 inside surface connected, part-through-wall circumferential flaws as well as two axial thumbnail flaws at the 0-and 90-degree azimuthal locations ofthe nozzle, are calculated using finite element analysis (FEA). For the circumferential flaw, themaximum K values around the nozzle circumference for each flaw depth are extracted and used as inputinto pc-CRACK to perform the PWSCC crack growth analyses.

For the axial flaws, the K at the deeppoint of the thumbnail shape is used as input for performing the PWSCC crack growth analyses.

Sincethe K vs. a profile is used as input, the shape of the component is not relevant.

For the crack growth analyses, two initial flaw sizes were chosen. These are based on expectedengineering flaw sizes that could be present for a crack that would then grow by PWSCC. The final flawsize for these analyses is 75% of the wall thickness.

This final depth is chosen as it is the maximumallowable flaw depth per Section XI of the ASME Code for pipe flaw evaluations.

Additionally, a finalflaw size of 95% of the wall thickness is also considered in this calculation.

The following are the additional parameters needed for the crack growth calculations:

Two initial crack depths = 0.025" and 0.1" (assumed)

Temperature

= 537°F (operating temperature

[1])Wall thickness

= 3" (Cold Leg thickness

[3])The resulting crack depths for the circumferential and axial flaws, as a function of time, as calculated bypc-CRACK are shown in Figure 9 for the 0.025" initial flaw size and Figure 10 for the 0.1" initial flawsize. The time for a flaw to grow from the initial flaw size to 75% and 95% though-wall is tabulated inTable 3 and Table 4 for both circumferential and axial flaw types, respectively.

Table 5 shows theallowable detected flaw sizes for the postulated flaws if continued operation for 18 and 36 months isconsidered.

6.0 CONCLUSION

S Stress intensity factors were calculated for the 360' circumferential flaws as well as the axial flaws at the0' and 90' locations.

The stress intensity factors were calculated using residual stress distributions forresidual stress plus normal operating conditions.

In addition, a far field in-plane bending moment isFile No.: 1400669.323 Page 9 of 23Revision:

0F0306-01R2 dslate gfrtgr r Associates, Inc.applied to the free end of the cold leg run piping to account for piping moments in the main loop piping.This combined loading is used for the determination of the stress intensity factors for both thecircumferential and axial flaws. Figure 5 and Figure 8 as well as Table I and Table 2, show thecalculated stress intensity factors for the circumferential and axial flaws.Crack growth evaluations were performed for circumferential and axial flaw configurations using twodifferent initial flaw sizes. As shown in Figure 9 and Table 3, the shortest time for an initial 0.025" deepflaw to grow to 75% through-wall in all cases is 55.6 years for a circumferential flaw. Figure 10 andTable 3 show that the shortest time for an initial 0.1" deep flaw to grow to 75% through-wall in all casesis 53.5 years for a circumferential flaw. Table 5 shows the maximum allowable detected flaw sizes for18 and 36 months of continued operation.

File No.: 1400669.323 Revision:

0Page 10 of 23F0306-01 R2 jSlcbul lat y Associates, Inc;

7.0 REFERENCES

1. SI Calculation No. 1400669.322, Rev. 0, "Cold Leg Bounding Nozzle Weld Residual StressAnalysis."

2. pc-CRACK 4.1, Version 4.1 CS, Structural Integrity Associates, December 2013.3. S1 Calculation No. 1400669.320, Rev. 0, "Finite Element Model Development for the Cold LegDrain, Spray, and Charging Nozzles."

4. Palisades

Document, Report No. CENC-1 115, "Analytical Report for Consumers Power Piping,"SI File No. 1300086.204.

5. Materials Reliability Program:

Crack Growth Rates for Evaluating Primary Water StressCorrosion cracking (PWSCC) ofAlloy 82, 182 and 132 Welds (MARP-115),

EPRI, Palo Alto, CA:2004, 1006696.6. ANSYS Mechanical APDL and PrepPost, Release 14.5 (w/ Service Pack 1), ANSYS, Inc.,September 2012.7. Anderson, T. L., "Fracture Mechanics Fundamentals and Applications,"

Second Edition, CRCPress, 1995.File No.: 1400669.323 Revision:

0Page 11 of 23F0306-01R2 Vj~sinatirsi mfgrlfy Associaes, kIcOTable 1: Stress Intensity Factors for Circumferential FlawsDepth a/t Max K(in) (ksi-in^0.5) 0.13 0.04 22.070.83 0.28 28.501.42 0.47 20.842.31 0.77 20.972.99 1.00 48.27Table 2: Stress Intensity Factors for Axial FlawsCL Axial Plane 00 CL Circ. Plane 900Crack Depth aft K at Deep Pt Crack Depth aft K at Deep Pt(in) (ksi-in0'5) (in) (ksi-in°'

5)0.25 0.08 21.26 0.24 0.08 18.430.78 0.26 20.95 0.83 0.28 19.001.37 0.46 19.58 1.43 0.48 20.822.16 0.72 22.89 2.19 0.73 25.672.90 0.97 28.34 2.85 0.95 30.93Table 3: Crack Growth Time to 75% Through-Wall Axial Crack Axial CrackInitial Flaw Size Circ. Crack(in) (00 plane) (90* plane) (years)(years) (years) (years)0.025 64.5 67.2 55.60.100 62.2 64.4 53.5Table 4: Crack Growth Time to 95% Through-Wall Initial Flaw Size Axial Crack Axial Crack Circ. Crack(in) (00 plane) (900 plane) (years)(years) (years) (years_0.025 77.0 77.9 66.20.100 74.6 75.0 64.0File No.: 1400669.323 Revision:

0Page 12 of 23F0306-01R2 anwbrui liftrily Associates, ic"Table 5: Allowable Detected Flaw SizeAllowable Detected Flaw Size (a/t)Cold Leg Thickness, t = 3.00"Months of Axial Flaw Axial Flaw Circumferential Continued at 00 plane at 900 plane FlawOperation a/t a (in) a/t a (in) a/t a (in)18 0.7238 2.1715 0.7202 2.1605 0.7269 2.180736 0.6871 2.0613 0.6788 2.0363 0.6273 1.8818File No.: 1400669.323 Revision:

0Page 13 of 23F0306-OIR2 Can" bt* AwaIft WFigure 1. Base Finite Element Model MeshFile No.: 1400669.323 Revision:

0Page 14 of 23F0306-OIR2

!UDVcbw MIM AWQGWK knGW4T ,,4Pilot Nodeand AxialRestraint Axial displacement restraint Symmetry boundaryconditions Fracture mechanics analysis using imported crack face stressFigure 2. Applied Mechanical Load Boundary Conditions File No.: 1400669.323 Revision:

0Page 15 of 23F0306-01 R2

~SVoytd" higr Associai knFigure 3. Circumferential Flaw with Crack Tip Elements Inserted(Note: Deepest circumferential flaw shown for example)File No.: 1400669.323 Revision:

0Page 16 of 23F0306-01R2

~jSVan"u ftIY Aswdafte kncPFigure 4. Transferred Residual Stress + NOC + Pressure Stress for Circumferential Flaws(Note: Deepest circumferential flaw shown for example)File No.: 1400669.323 Revision:

0Page 17 of 23F0306-01 R2 V a~Sn"iu NOfWdl Assacefts W.96050-40C630xS201000.000.200.400.600.801.00Depth (a/t)Figure 5. Stress Intensity Factors as a Function of Depth for Circumferential FlawsFile No.: 1400669.323 Revision:

0Page 18 of 23F0306-01R2 V aud"bJ, hbuI AssDOW kei0' plane90 planeFigure 6. Axial Flaws with Crack Tip Elements InsertedFile No.: 1400669.323 Revision:

0Page 19 of 23F0306-01 R2

~SVaud"W kfty Assodae6 km0' plane900 planeFigure 7. Transferred Residual Stress + NOC + Pressure Stress for Axial FlawsFile No.: 1400669.323 Revision:

0Page 20 of 23F0306-01 R2 353025= 201Jev 15X105[ Axial-j00.00 0.10 0.20 0.30 0.40 0.50 0.60 0.70 0.80 0.90 1.00Depth (a/t)Figure 8. Stress Intensity Factors as a Function of Depth for Axial FlawsFile No.: 1400669.323 Revision:

0Page 21 of 23F0306-01R2

~jSVaftr W hry Assoclels, We1.00.90.80.7-'0.60.40.30.20.10.0-'I-.-- -A-- A-" A-A-A.o-, --Circumferential

---Axial_0-.-Axial 9060 65 70 75 800 5 10 15 20 25 30 35 40 45 50 55Time (yrs)Figure 9. Crack Growth for All Flaw Types with 0.025" Initial Flaw SizeFile No.: 1400669.323 Revision:

0Page 22 of 23F0306-01 R2 C an"w~ii his rM Assoclites, hicP10.90.80.7-0.6Q-0.50.4U.0.30.20.10-4-e-- ---'j-Circumferential

---Axial_0-.-Axial_900 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80Time (yrs)Figure 10. Crack Growth for All Flaw Types with 0.1" Initial Flaw SizeFile No.: 1400669.323 Revision:

0Page 23 of 23F0306-01R2