WCAP-17262-NP, Rev 1, 'Technical Bases for Eliminating ...

LIC-13-0100EnclosureAttachment 5Page 1

"Technical Bases for Eliminating Large Primary Loop Pipe Rupture as theStructural Design Basis for Fort Calhoun Unit 1"

WCAP-17262-NP, Revision 1July 2013

Westinghouse Non-Proprietary Class 3

WCAP-17262-NPRevision 1

Technical Bases forEliminating Large PrimaryLoop Pipe Rupture as theStructural Design Basis forFort Calhoun Unit 1

Westinghouse

July 2013

WESTINGHOUSE NON-PROPRIETARY CLASS 3

WCAP-17262-NPRevision I

Technical Bases for Eliminating Large Primary LoopPipe Rupture as the Structural Design Basis for Fort

Calhoun Unit 1

C. T. Kupper*Piping Analysis and Fracture Mechanics

July 2013

Reviewer: D. C. Bhowmick*Piping Analysis and Fracture Mechanics

A. Udyawar*Piping Analysis and Fracture Mechanics

Approved: S. A. Swamy*, ManagerPiping Analysis and Fracture Mechanics

*Electronically approved records are authenticated in the electronic document management system.

Westinghouse Electric Company LLC1000 Westinghouse Drive

Cranberry Township, PA 16066, USA

© 2013 Westinghouse Electric Company LLCAll Rights Reserved

WESTINGHOUSE NON-PROPRIETARY CLASS 3 iiWESTINGHOUSE NON-PROPRIETARY CLASS 3 ii

Acronym List

ACRS Advisory Committee on Reactor Safeguards

AIF Atomic Industrial ForumANL Argonne National LaboratoryAR Aspect RatioASME American Society of Mechanical EngineersCE Combustion EngineeringCGR Crack Growth RateCMTR Certified Material Test ReportEPFM Elastic-Plastic Fracture MechanicsEPRI Electric Power Research InstituteEPU Extended Power UprateFCG Fatigue Crack Growth

FCS Fort Calhoun StationIGSCC Intergranular Stress Corrosion CrackingLBB Leak-Before-Break

LLNL Lawrence Livermore National LaboratoryLOCA Loss of Coolant AccidentMUR Measurement Uncertainty RecaptureNRC Nuclear Regulatory Commission

OD Outside DiameterOPPD Omaha Public Power District

PCSG Pipe Crack Study GroupPWR Pressurizer Water Reactor

PWSCC Primary Water Stress Corrosion CrackingRCS Reactor Coolant SystemRMS Root Mean SquareRT Room TemperatureSAW Submerged Arc Weld

SAM Seismic Anchor MotionSCC Stress Corrosion CrackingSMAW Shielded Metal Arc Weld

SSE Safe Shutdown Earthquake

WCAP-17262-NP July 2013Revision 1

WESTINGHOUSE NON-PROPRIETARY CLASS 3 iii

TABLE OF CONTENTS

1 .0 In tro d u ctio n .................................................................................................................... 1-11 .1 P u rp o s e .............................................................................................................. 1 -11.2 Background Information ..................................................................................... 1-11.3 Scope and Objectives ........................................................................................ 1-21.4 References ......................................................................................................... 1-3

2.0 Operation and Stability of the Reactor Coolant System ................................................. 2-12.1 Stress Corrosion Cracking .................................................................................. 2-12.2 W ater Hammer ................................................................................................... 2-22.3 Low Cycle and High Cycle Fatigue .................................................................... 2-32.4 W all Thinning, Creep, and Cleavage .................................................................. 2-32.5 References ......................................................................................................... 2-3

3.0 Pipe Geometry and Loading .......................................................................................... 3-13.1 Introduction to Methodology ............................................................................... 3-13.2 Calculation of Loads and Stresses ..................................................................... 3-23.3 Loads for Leak Rate Evaluation ......................................................................... 3-33.4 Load Combination for Crack Stability Analyses .................................................. 3-43.5 References ......................................................................................................... 3-4

4.0 Material Characterization ............................................................................................... 4-14.1 Primary Loop Pipe and Fittings Materials ........................................................... 4-14.2 Tensile Properties ............................................................................................... 4-14.3 Fracture Toughness Properties .......................................................................... 4-24.4 References ......................................................................................................... 4-5

5.0 Critical Location and Evaluation Criteria ........................................................................ 5-15.1 Critical Locations ................................................................................................ 5-15.2 Fracture Criteria ................................................................................................. 5-1

6.0 Leak Rate Predictions .................................................................................................... 6-16.1 Introduction ......................................................................................................... 6-16.2 General Considerations ...................................................................................... 6-16.3 Calculation Method ............................................................................................. 6-16.4 Leak Rate Calculations ...................................................................................... 6-26.5 References ......................................................................................................... 6-2

7.0 Fracture Mechanics Evaluation ...................................................................................... 7-17.1 Local Failure Mechanism ................................................................................... 7-17.2 Global Failure Mechanism .................................................................................. 7-27.3 Results of Crack Stability Evaluation .................................................................. 7-37.4 References ......................................................................................................... 7-4

8.0 Fatigue Crack Growth Analysis ...................................................................................... 8-1

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8.1 References ......................................................................................................... 8-2

9.0 Assessment of Margins .................................................................................................. 9-19.1 References ......................................................................................................... 9-1

10.0 Conclusions .................................................................................................................. 10-1

Appendix A: Limit Moment ................................................................................................... A-1


WESTINGHOUSE NON-PROPRIETARY CLASS 3 V

LIST OF TABLES

Table 3-1 Dimensions, Normal Loads and Stresses for Fort Calhoun Unit 1 .......................... 3-5

Table 3-2 Faulted Loads and Stresses for Fort Calhoun Unit 1 .............................................. 3-6

Table 4-1 Measured Tensile Properties for Fort Calhoun Unit 1 Primary Loop Piping ............. 4-6

Table 4-2 Measured Tensile Properties for Fort Calhoun Unit 1 Primary Loop Elbows ........... 4-7

Table 4-3 Mechanical Properties for Fort Calhoun Unit 1 Materials at Operating Temperatures........................................................................................................................... 4 -8

Table 4-4 Chemistry & Fracture Toughness Piping Properties of the Material Heats of FortC a lh o u n U n it 1 .................................................................................................... 4 -9

Table 4-5 Chemistry & Fracture Toughness Elbow Properties of the Material Heats of FortC a lh o u n U n it 1 .................................................................................................. 4 -10

Table 4-6 Fracture Toughness Properties Used to Evaluate Critical Locations .................... 4-11

Table 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the Governing Locations ................ 6-3

Table 7-1 Stability Results for Fort Calhoun Based on Elastic-Plastic J-Integral Evaluations .7-5

Table 7-2 Stability Results for Fort Calhoun Based on Limit Load .......................................... 7-5

Table 8-1 Reactor Coolant System Operating Transients ....................................................... 8-3

Table 8-2 FCG at Alloy 82/182 Weld (Nozzle to Safe-end Weld) - Outlet Nozzle .................. 8-4

Table 8-3 FCG at Stainless Steel Weld (Safe-end to Pipe Weld) - Outlet Nozzle .................. 8-4

Table 8-4 FCG at Reactor Vessel Nozzle Location - Outlet Nozzle ....................................... 8-4

Table 9-1 Leakage Flaw Sizes, Critical Flaw Sizes by Limit Load and Margins for Fort Calhoun................................. ......................................................................................... 9 -2

Table 9-2 Stability Results for Fort Calhoun Based on Elastic-Plastic J-Integral Evaluations.9-2

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WESTINGHOUSE NON-PROPRIETARY CLASS 3 viWESTINGHOUSE NON-PROPRIETARY CLASS 3 VI

LIST OF FIGURES

Figure 3-1

Figure 3-2

Figure 6-1

Figure 6-2

Figure 6-3

Figure 7-1

Figure 7-2

Figure 7-3

Figure 8-1

Figure 8-2

Figure 8-3

Figure 8-4

Figure A-1

Reactor Coolant System Pipe ........................................................................... 3-7

Schematic Diagram of Fort Calhoun Primary Loop Showing Weld Locations ...3-8

Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures ............. 6-4[ ]a,c,, Pressure Ratio as a Function of L/D ....................................... 6-5

Idealized Pressure Drop Profile Through a Postulated Crack ........................... 6-6

Fully Plastic Stress Distribution ....................................................................... 7-6

Critical Flaw Size Prediction - Location 1 ........................................................ 7-7

Critical Flaw Size Prediction - Location 6 ........................................................ 7-8

Reactor Vessel Outlet Nozzle with Stress Cut Locations ................................. 8-5

Reference Fatigue Crack Growth Curves for Carbon & Low Alloy Ferritic Steels.................. ........................................................................................................ 8 -6

Reference Fatigue Crack Growth Curves for Stainless Steels ......................... 8-7

Reference Fatigue Crack Growth Curves for Alloy 82/182 Welds .................... 8-8

Pipe with a Through-Wall Crack in Bending .................................................... A-2


WESTINGHOUSE NON-PROPRIETARY CLASS 3 vii

EXECUTIVE SUMMARY

The original structural design basis of the reactor coolant system for the Omaha Public PowerDistrict (OPPD) Fort Calhoun Station (FCS) Unit 1 required consideration of dynamic effectsresulting from pipe break and that protective measure for such breaks be incorporated into thedesign. Subsequent to the original Fort Calhoun design, additional concern of asymmetricblowdown loads was raised as described in Unresolved Safety Issue A-2 (AsymmetricBlowdown Loads on the Reactor Coolant System) and Generic Letter 84-04. Fort Calhoun Unit1 was part of the utilities which sponsored Westinghouse to resolve the A-2 issue. Genericanalyses by Westinghouse to resolve the A-2 issue were approved by the NRC anddocumented in Generic Letter 84-04 (Reference 1-1).

Research by the NRC and industry coupled with operating experience determined that safetycould be negatively impacted by placement of pipe whip restraints on certain systems. As aresult, NRC and industry initiatives resulted in demonstrating that leak-before-break (LBB)criteria can be applied to reactor coolant system piping based on fracture mechanics technologyand material toughness.

Subsequently, the NRC modified 10CFR50 General Design Criterion 4, and published inFederal Register (Vol. 52, No. 207) on October 27, 1987 its final rule, "Modification of GeneralDesign Criterion 4 Requirements for Protection Against Dynamic Effects of Postulated PipeRuptures (Reference 1-2)." This change to the rule allows use of leak-before-break technologyfor excluding from the design basis the dynamic effects of postulated ruptures in primary coolantloop piping in pressurized water reactors (PWRs).

This current report (WCAP-17262-NP Revision 1) demonstrates compliance with LBBtechnology for the Fort Calhoun Unit 1 reactor coolant system piping on a plant specificanalysis. Inputs from the license renewal program are used in the LBB analysis. The reportdocuments the plant specific geometry, operating parameters, loading, and material propertiesused in the fracture mechanics evaluation. Mechanical properties were determined at operatingtemperatures. Since the piping systems include cast stainless steel, fracture toughnessconsidering thermal aging was determined for each heat of material for the fully aged condition.

Based on loading, pipe geometry and fracture toughness considerations, enveloping critical(governing) locations were determined at which leak-before-break crack stability evaluationswere made. Through-wall flaw sizes were postulated which would cause a leak at a rate of ten(10) times the leakage detection system capability of the plant. Large margins for such flawsizes were demonstrated against flaw instability. Finally, fatigue crack growth was shown not tobe an issue for the primary loop piping.

The effects due to the license renewal program on the continued applicability of LBB for thereactor coolant loop piping at Fort Calhoun Unit 1 have been evaluated. It is demonstrated thatthe dynamic effects of the pipe rupture resulting from postulated breaks in the reactor coolantprimary loop piping need not be considered in the structural design basis for Fort Calhoun Unit 1for the license renewal period.


WESTINGHOUSE NON-PROPRIETARY CLASS 3 Viii

Revision 0 of this report is still applicable for the Extended Power Uprate (EPU) andMeasurement Uncertainty Recapture (MUR) programs. Revision 1 is applicable for "current"operating conditions where "current" applies to normal operating conditions prior to EPU andMUR.


WESTINGHOUSE NON-PROPRIETARY CLASS 3 1-1

1.0 INTRODUCTION

1.1 PURPOSE

This report applies to the Fort Calhoun Unit 1 Reactor Coolant System (RCS) primary looppiping. It is intended to demonstrate that for the specific parameters of Fort Calhoun, RCSprimary loop pipe breaks need not be considered in the structural design basis. The LBBapproach taken in this report has been accepted by the U.S. Nuclear Regulatory Commission(NRC) (Reference 1-2).

1.2 BACKGROUND INFORMATION

Westinghouse has performed considerable testing and analysis to demonstrate that RCSprimary loop pipe breaks can be eliminated from the structural design basis of all Westinghouseplants. The concept of eliminating pipe breaks in the RCS primary loop was first presented tothe NRC in 1978 in WCAP-9283 (Reference 1-3). That topical report employed a deterministicfracture mechanics evaluation and a probabilistic analysis to support the elimination of RCSprimary loop pipe breaks. That approach was then used as a means of addressing GenericIssue A-2 and Asymmetric LOCA Loads.

Westinghouse performed additional testing and analysis to justify the elimination of RCSprimary loop pipe breaks. This material was provided to the NRC along with Letter ReportNS-EPR-2519 (Reference 1-4).

The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to addressthis same issue using a probabilistic approach. As part of the LLNL research effort,Westinghouse performed extensive evaluations of specific plant loads, material properties,transients, and system geometries to demonstrate that the analysis and testing previouslyperformed by Westinghouse and the research performed by LLNL applied to all Westinghouseplants (References 1-5 and 1-6). The results from the LLNL study were released at a March 28,1983, ACRS Subcommittee meeting. These studies, which are applicable to all Westinghouseplants east of the Rocky Mountains, determined the mean probability of a direct LOCA (RCSprimary loop pipe break) to be 4.4 x 1012 per reactor year and the mean probability of anindirect LOCA to be 107 per reactor year. Although FCS is not a Westinghouse plant, similarprobabilities would be expected for a LOCA at FCS. Thus, the results previously obtained byWestinghouse (Reference 1-3) were confirmed by an independent NRC research study.

Based on the studies by Westinghouse, LLNL, the ACRS, and the AIF, the NRC completed asafety review of the Westinghouse reports submitted to address asymmetric blowdown loadsthat result from a number of discrete break locations on the PWR primary systems. The NRCStaff evaluation (Reference 1-1) concludes that an acceptable technical basis has beenprovided so that asymmetric blowdown loads need not be considered for those plants that candemonstrate the applicability of the modeling and conclusions contained in the Westinghouseresponse or can provide an equivalent fracture mechanics demonstration of the primary coolantloop integrity. In a more formal recognition of leak-before-break (LBB) methodology applicabilityfor PWRs, the NRC appropriately modified 10 CFR 50, General Design Criterion 4,

Introduction July 2013WCAP-1 7262-NP Revision 1


"Requirements for Protection Against Dynamic Effects for Postulated Pipe Rupture"

(Reference 1-2).

1.3 SCOPE AND OBJECTIVES

The general purpose of this investigation is to demonstrate leak-before-break for the primaryloop in Fort Calhoun Unit 1. The recommendations and criteria proposed in References 1-7 and1-8 are used in this evaluation. These criteria and resulting steps of the evaluation procedurecan be briefly summarized as follows:

1. Calculate the applied loads. Identify the locations at which the highest stress occurs.

2. Identify the materials and the associated material properties.

3. Postulate a surface flaw at a governing location. Determine fatigue crack growth. Showthat a through-wall crack will not result.

4. Postulate a through-wall flaw at the critical (governing) locations. The size of the flawshould be large enough so that the leakage is assured of detection with margin using theinstalled leak detection equipment when the pipe is subjected to normal operating loads.Demonstrate a margin of 10 between the calculated leak rate and the leak detectioncapability.

5. Using faulted loads, demonstrate that there is a margin of at least 2 between the leakageflaw size and the critical flaw size.

6. Review the operating history to ascertain that operating experience has indicated noparticular susceptibility to failure from the effects of corrosion, water hammer, low andhigh cycle fatigue, wall thinning, creep, or cleavage.

7. For the materials actually used in the plant provide the properties including toughnessand tensile test data. Evaluate long term effects such as thermal aging.

8. Demonstrate margin on applied load.

This report provides a fracture mechanics demonstration of primary loop integrity for FortCalhoun consistent with the NRC position for exemption from consideration of dynamic effects.

It should be noted that the terms "flaw" and "crack" have the same meaning and are usedinterchangeably. "Governing location" and "critical location" are also used interchangeablythroughout the report.

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1.4 REFERENCES

1-1 USNRC Generic Letter 84-04, Subject: "Safety Evaluation of Westinghouse TopicalReports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary MainLoops," February 1, 1984.

1-2 Nuclear Regulatory Commission, 10 CFR 50, Modification of General Design Criteria 4Requirements for Protection Against Dynamic Effects of Postulated Pipe Ruptures, FinalRule, Federal Register/Vol. 52, No. 207/Tuesday, October 27, 1987/Rules andRegulations, pp. 41288-41295.

1-3 WCAP-9283, "Integrity of the Primary Piping Systems of Westinghouse Nuclear PowerPlants During Postulated Seismic Events," March, 1978.

1-4 Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G. Eisenhut),Westinghouse Proprietary Class 2, November 10, 1981.

1-5 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983.

1-6 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.

1-7 Standard Review Plan: Public Comments Solicited; 3.6.3 Leak-Before-Break EvaluationProcedures; Federal RegisterNol. 52, No. 167/Friday August 28, 1987/Notices,pp. 32626-32633.

1-8 NUREG-0800 Revision 1, March 2007, Standard Review Plan: 3.6.3 Leak-Before-BreakEvaluation Procedures.

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2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT

SYSTEM

2.1 STRESS CORROSION CRACKING

The Westinghouse and Combustion Engineering (CE) designed reactor coolant system (RCS)primary loops have an operating history that demonstrates the inherent operating stabilitycharacteristics of the design. This includes a low susceptibility to cracking failure from theeffects of corrosion (e.g., intergranular stress corrosion cracking (IGSCC)). This operatinghistory totals over 1400 reactor-years, including 16 plants each having over 30 years ofoperation, 10 other plants each with over 25 years of operation, 11 plants each over 20 years ofoperation and 12 plants each over 15 years of operation.

In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second PipeCrack Study Group. (The first Pipe Crack Study Group (PCSG) established in 1975 addressedcracking in boiling water reactors only.) One of the objectives of the second PCSG was toinclude a review of the potential for stress corrosion cracking in Pressurized WaterReactors (PWR's). The results of the study performed by the PCSG were presented inNUREG-0531 (Reference 2-1) entitled "Investigation and Evaluation of Stress CorrosionCracking in Piping of Light Water Reactor Plants." In that report the PCSG stated:

"The PCSG has determined that the potential for stress-corrosion cracking in PWRprimary system piping is extremely low because the ingredients that produce IGSCC arenot all present. The use of hydrazine additives and a hydrogen overpressure limit theoxygen in the coolant to very low levels. Other impurities that might causestress-corrosion cracking, such as halides or caustic, are also rigidly controlled. Only forbrief periods during reactor shutdown when the coolant is exposed to the air and duringthe subsequent startup are conditions even marginally capable of producingstress-corrosion cracking in the primary systems of PWRs. Operating experience inPWRs supports this determination. To date, no stress corrosion cracking has beenreported in the primary piping or safe ends of any PWR."

For stress corrosion cracking (SCC) to occur in piping, the following three conditions must existsimultaneously: high tensile stresses, susceptible material, and a corrosive environment. Sincesome residual stresses and some degree of material susceptibility exist in any stainless steelpiping, the potential for stress corrosion is minimized by properly selecting a material immune toSCC as well as preventing the occurrence of a corrosive environment. The materialspecifications consider compatibility with the system's operating environment (both internal andexternal) as well as other material in the system, applicable ASME Code rules, fracturetoughness, welding, fabrication, and processing.

The elements of a water environment known to increase the susceptibility of austenitic stainlesssteel to stress corrosion are: oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, andreduced forms of sulfur (e.g., sulfides, sulfites, and thionates). Strict pipe cleaning standardsprior to operation and careful control of water chemistry during plant operation are used to

Operation and Stability of the Reactor Coolant System July 2013WCAP-1 7262-NP Revision 1


prevent the occurrence of a corrosive environment. Prior to being put into service, the piping iscleaned internally and externally. During flushes and preoperational testing, water chemistry iscontrolled in accordance with written specifications. Requirements on chlorides, fluorides,conductivity, and pH are included in the acceptance criteria for the piping.

During plant operation, the reactor coolant water chemistry is monitored and maintained withinvery specific limits. Contaminant concentrations are kept below the thresholds known to beconducive to stress corrosion cracking with the major water chemistry control standards beingincluded in the plant operating procedures as a condition for plant operation. For example,during normal power operation, oxygen concentration in the RCS is expected to be in the ppbrange by controlling charging flow chemistry and maintaining hydrogen in the reactor coolant atspecified concentrations. Halogen concentrations are also stringently controlled by maintainingconcentrations of chlorides and fluorides within the specified limits. Thus during plant operation,the likelihood of stress corrosion cracking is minimized.

During 1979, several instances of cracking in PWR feedwater piping led to the establishment ofthe third PCSG. The investigations of the PCSG reported in NUREG-0691 (Reference 2-2)further confirmed that no occurrences of IGSCC have been reported for PWR primary coolantsystems.

Primary Water Stress Corrosion Cracking (PWSCC) occurred in V. C. Summer reactor vesselhot leg nozzle, Alloy 82/182 welds. It should be noted that this susceptible material is found atthe Fort Calhoun Unit 1 Reactor Vessel Inlet and Outlet nozzle locations. Mitigation will beimplemented as required to minimize PWSCC at the Reactor Vessel nozzle locations.

2.2 WATER HAMMER

Overall, there is a low potential for water hammer in the RCS since it is designed and operatedto preclude the voiding condition in normally filled lines. The reactor coolant system, includingpiping and primary components, is designed for normal, upset, emergency, and faulted conditiontransients. The design requirements are conservative relative to both the number of transientsand their severity. Relief valve actuation and the associated hydraulic transients following valveopening are considered in the system design. Other valve and pump actuations are relativelyslow transients with no significant effect on the system dynamic loads. To ensure dynamicsystem stability, reactor coolant parameters are stringently controlled. Temperature duringnormal operation is maintained within a narrow range by control rod position; pressure iscontrolled by pressurizer heaters and pressurizer spray also within a narrow range forsteady-state conditions. The flow characteristics of the system remain constant during a fuelcycle because the only governing parameters, namely system resistance and the reactorcoolant pump characteristics, are controlled in the design process. Additionally, Westinghouseand CE designs have instrumented typical reactor coolant systems to verify the flow andvibration characteristics of the system. Preoperational testing and operating experience haveverified the Westinghouse approach. The operating transients of the RCS primary piping aresuch that no significant water hammer can occur.

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2.3 LOW CYCLE AND HIGH CYCLE FATIGUE

An evaluation of the low cycle fatigue loadings was carried out as part of this study in the formof a fatigue crack growth analysis, as discussed in Section 8.0.

High cycle fatigue loads in the system would result primarily from pump vibrations. These areminimized by restrictions placed on shaft vibrations during hot functional testing and operation.During operation, an alarm signals the exceedance of the vibration limits. Field measurementshave been made on a number of plants during hot functional testing. Stresses in the elbowbelow the reactor coolant pump resulting from system vibration have been found to be verysmall, between 2 and 3 ksi at the highest. These stresses are well below the fatigue endurancelimit for the material and would also result in an applied stress intensity factor below thethreshold for fatigue crack growth. Fort Calhoun RCS configurations are similar and the resultsare concluded to be similar.

2.4 WALL THINNING, CREEP, AND CLEAVAGE

Wall thinning by erosion and erosion-corrosion effects should not occur in the primary looppiping due to the low velocity, typically less than 1.0 ft/sec and the stainless steel material, whichis highly resistant to these degradation mechanisms. The cause of wall thinning is related to thehigh water velocity and is therefore clearly not a mechanism that would affect the primary looppiping.

The maximum operating temperature of the primary loop piping, which is less than 6000F, is wellbelow the temperature that would cause significant mechanical creep damage in stainless steelpiping. Cleavage type failures are not a concern for the operating temperatures and thestainless steel material used in the primary loop piping.

2.5 REFERENCES

2-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of Light WaterReactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979.

2-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized WaterReactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.

Operation and Stability of the Reactor Coolant System July 2013WCAP-17262-NP Revision 1


3.0 PIPE GEOMETRY AND LOADING

3.1 INTRODUCTION TO METHODOLOGY

The general approach is discussed first. As an example a segment of the primary coolant looppipe is shown in Figure 3-1. The as-built outside diameter and minimum thickness of the pipeare, as shown in the figure. The normal stresses at the weld location are from the loadcombination procedure discussed in Section 3.3 whereas the faulted loads are as described inSection 3.4. The components for normal loads are internal pressure, dead weight, pressureexpansion, and normal thermal expansion. An additional component, Safe ShutdownEarthquake (SSE), is considered for faulted loads. As seen from Table 3-2, the highest stressedlocation in the entire loop is Location 1 (Figure 3-2) at the reactor vessel outlet nozzle to pipeweld. The highest stressed location in the cold/cross-over leg piping is Location 6 at the cross-over leg elbow weld. These are the critical locations at which, as enveloping locations,leak-before-break is to be established. Essentially a circumferential flaw is postulated at eachlocation which is subjected to both the normal loads and faulted loads to assess leakage andstability, respectively. The loads (developed below) at the critical locations are also given inFigure 3-1. Loads from Replacement Steam Generator (RSG) that are applicable for currentpower operating conditions are used in the LBB analysis.

Since the piping for Fort Calhoun Unit 1 is made of cast stainless steel, thermal aging effectsmust be considered (Section 4.0). Thermal aging results in lower fracture toughness; thus,locations must be examined taking into consideration both fracture toughness and stress. Onceloads (this section) and fracture toughness (Section 4.0) are obtained, the critical locations areidentified in Section 5.0. At these locations, leak rate evaluations (Section 6.0) and fracturemechanics evaluations (Section 7.0) are performed per the guidance of References 3-1 and 3-2.Fatigue crack growth (Section 8.0) and stability margins are also evaluated (Section 9.0).

The critical locations are determined based on stresses and material properties for Fort CalhounUnit 1. All the weld locations for evaluation are those shown in Figure 3-2. It should be notedthat the Fort Calhoun Primary loop consists of 2 loops with 24 weld locations in each loop. Dueto the similarity in geometry and materials, only one loop with 14 weld locations will be analyzedfor the hot, cold, and crossover legs with the enveloping loads being used at each weld locationfrom both loops. Therefore, all 24 weld locations in each loop are accounted for with the 14weld locations analyzed.

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3.2 CALCULATION OF LOADS AND STRESSES

The stresses due to axial loads and bending moments are calculated by the following equation:

F M (3-1)

A Z

where,

a = stress

F = axial load

M = moment

A = pipe cross-sectional area

Z = section modulus

The moments for the desired loading combinations are calculated by the following equation:

M= Mx +M2+M (3-2)

where,

M) X component of moment, Torsion

my Y component of bending moment

Mz Z component of bending moment

The axial load and moments for leak rate predictions and crack stability analyses are computedby the methods to be explained in Sections 3.3 and 3.4.

Pipe Geometry and Loading July 2013WCAP-17262-NP Revision 1


3.3 LOADS FOR LEAK RATE EVALUATION

The normal operating loads for leak rate predictions are calculated by the following equations:

F = FDw + FTH + Fp + Fpexp (3-3)

Mx= (Mx)Dw + (MX)TH + (Mx)Pexp (3-4)

MY= (My)Dw + (MY)TH + (My)Pexp (3-5)

Mz= (Mz)ow + (MZ)TH + (Mz)Pexp (3-6)

The subscripts of the above equations represent the following loading cases:

DW = deadweight

TH = normal thermal expansion

P = load due to internal pressure

Pexp = pressure expansion

This method of combining loads is often referred to as the algebraic sum method(References 3-1 and 3-2).

The loads based on this method of combination are provided in Table 3-1 at all the weldlocations identified in Figure 3-2. The as-built dimensions are also given in Table 3-1.


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3.4 LOAD COMBINATION FOR CRACK STABILITY ANALYSES

In accordance with Standard Review Plan 3.6.3 (References 3-1 and 3-2), the margin in termsof applied loads needs to be demonstrated by crack stability analysis. Margin on loads of 1.4(4/2) can be demonstrated if normal plus Safe Shutdown Earthquake (SSE) are applied. The 1.4(4I2) margin should be reduced to 1.0 if the deadweight, thermal expansion, internal pressure,pressure expansion, SSEINERTIA and seismic anchor motion (SAM) loads are combined basedon individual absolute values as shown below.

The absolute sum of loading components is used for the FCS LBB analysis which results inhigher magnitude of combined loads and thus satisfies a margin on loads of 1.0. The absolutesummation of loads is shown in the following equations:

F =IFDW I+ I FTH I+ I Fp I+ I FSSEINERTA I+ I FSSEAM I + I Fpexp I

Mx = I (Mx)Dw I+ I (MX)TH I + I (MX)SSEINERTIAI + I (MX)SSEAMI + I (Mx)Pexp I

MY= I (My)Dw I+ I (MY)TH I + I (My)ssEINERTIAI + I (MY)SSEAMI + I (My)Pexp I

Mz= I (Mz)Dw I + I (MZ)TH I + I (MZ)SSEINERTIAI + I (Mz)ssEMI + I (Mz)pexp I

(3-7)

(3-8)

(3-9)

(3-10)

where subscript SSEINERTIA refers to safe shutdown earthquake inertia, SSEAM is safeshutdown earthquake anchor motion.

The loads so determined are used in the fracture mechanics evaluations (Section 7.0) todemonstrate the LBB margins at the locations established to be the governing locations. Theseloads at all the weld locations (see Figure 3-2) are given in Table 3-2.

3.5 REFERENCES

3-1 Standard Review Plan: Public Comments Solicited; 3.6.3 Leak-Before-Break EvaluationProcedures; Federal RegisterNol. 52, No. 167/Friday, August 28, 1987/Notices,pp. 32626-32633.

3-2 NUREG-0800 Revision 1, March 2007, Standard Review Plan: 3.6.3Evaluation Procedures.

Leak-Before-Break


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Table 3-1 Dimensions, Normal Loads and Stresses for Fort Calhoun Unit I

Minimum Axial BendingOutside Diameter Thickness Loadb Moment Total Stress

Locationa (in) (in) (kips) (in-kips) (ksi)

1 38.50 2.688 1822 21287 14.43

2 38.50 2.688 1771 806 6.18

3 38.50 2.688 1771 9396 9.57

4 39.00 2.938 1682 18368 11.63

5 29.25 2.063 1019 2965 8.43

6 29.00 1.938 1019 2732 8.80

7 29.00 1.938 1053 2828 9.10

8 29.00 1.938 1042 1763 8.01

9 29.00 1.938 1025 2178 8.31

10 29.25 2.063 995 3530 8.80

11 29.00 1.938 1044 1345 7.62

12 29.00 1.938 1023 681 6.86

13 29.00 1.938 1039 629 6.91

14 29.00 1.938 1028 1550 7.72

Notes:

a. See Figure 3-2

b. Included Pressure


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Table 3-2 Faulted Loads and Stresses for Fort Calhoun Unit I

Locationa'b Axial Loadc (kips) Bending Moment (in- Total Stress (ksi)kips)

1 1964 23806 15.90

2 1962 2777 7.59

3 1961 11586 11.06

4 2039 21235 13.73

5 1078 4934 10.52

6 1075 4368 10.71

7 1074 4200 10.54

8 1058 2528 8.84

9 1058 3130 9.42

10 1077 4604 10.23

11 1067 2471 8.84

12 1069 1670 8.09

13 1061 1860 8.22

14 1062 3884 10.16

Notes:

a.

b.

C.

See Figure 3-2

See Table 3-1 for dimensions

Included Pressure


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-F>

MP

OD IK

Location 1 OD' = 38.50 in

ta = 2.688

Normal Loadsa Faulted Loadsb

Forcec: 1822 kips Forcec: 1964 kips

Moment: 21287 in-kips Moment: 23806 in-kips

ODa = 29.00 inLocation 6 ta = 1.938

Normal Loadsa Faulted Loadsb

Forcec: 1019 kips Forcec: 1075 kips

Moment: 2732 in-kips Moment: 4368 in-kips

a See Table 3-1b See Table 3-2c Includes the force due to a pressure of 2100 psia

Figure 3-1 Reactor Coolant System Pipe


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WESTINGHOUSE NON-PROPRIETARY CLASS 3 3-8WESTINGHOUSE NON-PROPRIETARY CLASS 3 3-8

Pump

SteamGenerator

* Alloy 82/182 Weld

Cross-Over Leg

HOT LEGTemperature: 5930F*

CROSS-OVER LEGTemperature: 545 0F*

COLD LEGTemperature: 5450F*

Pressure: 2100 psia

Pressure: 2100 psia

Pressure: 2100 psia

* Temperatures applicable to Fort Calhoun current operating conditions (pre-EPU+MUR)

Figure 3-2 Schematic Diagram of Fort Calhoun Primary Loop Showing Weld Locations

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4.0 MATERIAL CHARACTERIZATION

4.1 PRIMARY LOOP PIPE AND FITTINGS MATERIALS

The primary loop pipe materials are A-451 CPF8M, and the elbow fittings are A-351-65 CF8M.

4.2 TENSILE PROPERTIES

The Certified Materials Test Reports (CMTRs) for Fort Calhoun primary loop piping and elbowfittings were used to establish the tensile properties for the leak-before-break analyses. Thetensile properties for the pipe material are provided in Table 4-1; while the tensile properties forthe elbow fittings material are provided in Table 4-2.

For both the A351-65 CF8M and A451 CPF8M materials, the representative properties at FortCalhoun current operating temperatures were established from the tensile properties at roomtemperature by utilizing Section II of the ASME Boiler and Pressure Vessel Code (Reference 4-1). Code tensile properties at the operating temperatures of 5930F for the hot leg and 5450F forthe cold/crossover legs were obtained by interpolation. Ratios of the ASME Code tensileproperties at operating temperature to the corresponding properties at room temperature werethen applied to the room temperature tensile properties obtained from CMTRs (Tables 4-1 and4-2) to obtain the Fort Calhoun specific properties at operating temperatures.

The average and lower bound yield strengths and ultimate strengths for the pipe and elbowfitting materials are tabulated in Table 4-3. The ASME Code modulus of elasticity values arealso given, and Poisson's ratio was taken as 0.3.

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4.3 FRACTURE TOUGHNESS PROPERTIES

The pre-service fracture toughness (J) of cast stainless steels that are of interest are in terms ofJjc (J at Crack Initiation) and have been found to be very high at 6000F. Cast stainless steel issusceptible to thermal aging during service. Thermal aging of cast stainless steel results in adecrease in the ductility, impact strength, and fracture toughness, of the material. Depending onthe material composition, the Charpy impact energy of a cast stainless steel component coulddecrease to a small fraction of its original value after exposure to reactor temperatures duringservice.

The method described below was used to calculate the end of life toughness properties for thecast material of the Fort Calhoun Unit 1 primary coolant loop piping and elbows.

In 1994, the Argonne National Laboratory (ANL) completed an extensive research program inassessing the extent of thermal aging of cast stainless steel materials. The ANL researchprogram measured mechanical properties of cast stainless steel materials after they had beenheated in controlled ovens for long periods of time. ANL compiled a data base, both from datawithin ANL and from international sources, of about 85 compositions of cast stainless steelexposed to a temperature range of 290-4000C (550-7500F) for up to 58,000 hours (6.5 years).From this database, ANL developed correlations for estimating the extent of thermal aging ofcast stainless steel (References 4-2 and 4-3).

ANL developed the fracture toughness estimation procedures by correlating data in thedatabase conservatively. After developing the correlations, ANL validated the estimationprocedures by comparing the estimated fracture toughness with the measured value for severalcast stainless steel plant components removed from actual plant service. The ANL proceduresproduced conservative estimates that were about 30 to 50 percent less than actual measuredvalues. The procedure developed by ANL in Reference 4-3 was used to calculate the end of lifefracture toughness values for this analysis. ANL research program was sponsored and theprocedure was accepted (Reference 4-4) by the NRC.

The chemical compositions of the Fort Calhoun primary loop pipe and elbow fitting material areavailable from CMTRs and are provided in Table 4-4 and Table 4-5 of this report. It should benoted that the Fort Calhoun elbows are CF8M steel whereas the piping segments are CPF8Msteel. The following equations are applicable for both CF8M and CPF8M type materials.

The following equations are taken from Reference 4-3:

Creq = Cr + 1.21 (Mo) + 0.48(Si) - 4.99 (4-1)

Nieq = (Ni) + 0.11(Mn) - 0.0086(Mn) 2 + 18.4(N) + 24.5(C) + 2.77 (4-2)

where Creq = (Chromium equivalent); Nieq = (nickel equivalent);

6c (ferrite content) in percent volume is given by:

Material Characterization July 2013WCAP-1 7262-NP Revision 1


6c=100.3(Creq / Nieq )2-170.72(Creq / Nieq )+74.22 (4-3)

The saturation room temperature (RT) impact energies of the cast stainless steel materials weredetermined from the chemical compositions.

For CF8M steel with < 10% Ni, the saturation value of RT impact energy CVsat (J/cm 2) is the lowervalue determined from

Io9goCVsat = 1.10 + 2.12 exp (-0.0414) (4-4)

where the material parameter 4 is expressed as

4 = 8, (Ni + Si +Mn) 2(C + 0.4N)/5.0 (4-5)

and from

logioCvsat = 7.28 - 0.0116c - 0.185Cr - 0.369Mo - 0.451Si- 0.007Ni - 4.71(C + 0.4N) (4-6)

For CF8M steel with > 10% Ni, the saturation value of RT impact energy CVsat (J/cm 2) is the lower

value determined from

IogioCvsat = 1.10 + 2.64 exp (-0.064¢) (4-7)

where the material parameter 4 is expressed as

S= 5 (Ni + Si +Mn) 2(C + 0.4N)/5.0 (4-8)

and from

IogioCvsat = 7.28 - 0.011 0 - 0.185Cr - 0.369Mo - 0.451Si- 0.007Ni - 4.71(C + 0.4N) (4-9)

The saturation J-R curve at RT, for static-cast CF8M steel is given by

Jd = 16(Cvsat)0 67(Aa)n (4-10)

and for centrifugally cast CF8M steel, by

Jd = 20(Cvsat)067(Aa)n (4-11)

where the exponent n for CF8M steel is expressed as

n = 0.23 + 0.08 loglo (Cvsat) (4-12)

where Jd is the "deformation J" in kJim 2 and Aa is the crack extension in mm.

The saturation J-R curve at 2900C (5540 F), for static-cast CF8M steel is given by

Jd = 49 (Cvsat)041(Aa)n (4-13)

and for centrifugally cast CF8M steel, by

Jd = 57 (Cvsat)0 41(Aa)n (4-14)

where the exponent n for CF8M steel is expressed as

n = 0.23 + 0.06 log1 o (Cvsat) (4-15)

where Jd is the "deformation J" in kJ/m 2 and Aa is the crack extension in mm.

Material Characterization July 2013WCAP-1 7262-NP Revision 1


Note: Conservatively static cast equations are used for J calculations.

[

]a,c,e

The results from the ANL Research Program indicate that the lower-bound fracture toughness ofthermally aged cast stainless steel is similar to that of Submerged Arc Welds (SAWs). Theapplied value of the J-integral for a flaw in the weld regions will be lower than that in the basemetal because the yield strength for the weld materials is much higher at the temperature1 .Therefore, weld regions are less limiting than the cast base material.

In the fracture mechanics analyses that follow, the fracture toughness properties given in Table4-6 will be used as the criteria against which the applied fracture toughness values will becompared.

1 All the applied J values were conservatively determined by using base metal strength properties.


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4.4 REFERENCES

4-1 ASME Boiler and Pressure Vessel Code Section II, Part D. "Materials," 2001 Edition,July 1,2001.

4-2 0. K. Chopra and W. J. Shack, "Assessment of Thermal Embrittlement of Cast StainlessSteels," NUREG/CR-6177, U. S. Nuclear Regulatory Commission, Washington, DC,May 1994.

4-3 0. K. Chopra, "Estimation of Fracture Toughness of Cast Stainless Steels duringThermal Aging in LWR Systems," NUREG/CR-4513, Revision 1, U. S. NuclearRegulatory Commission, Washington, DC, August 1994.

4-4 "Flaw Evaluation of Thermally Aged Cast Stainless Steel in Light-Water ReactorApplications," Lee, S.; Kuo, P. T.; Wichman, K.; Chopra, 0.; Published in InternationalJournal of Pressure Vessel and Piping, June 1997.


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Table 4-1 Measured Tensile Properties for Fort Calhoun Unit I PrimaryLoop Piping

Yield Strength (psi) Ultimate Strength (psi)Heat No. Room Temp. Room Temp.

J-26078901 46000 83600

J-21278901 43110 83200

K-3559012 43100 83700

J-289567890 53100 84700

J-289567890 53100 84700

J-283789012 42500 82100

J-355123456 47800 85500

J-288234567 46600 89220

J-285012345 45600 85700

J-288234567 46600 89220

J-26201234 44460 83920

J-285678901 47600 85700

J-287567890B 46120 85210

J-286234567 53900 85100

J-288901234 47900 83000

J-285678901 47600 85700

J-284456789 51000 84700

J-286901234 49400 78700

J-287567890 46120 85210

J-284456789 51000 84700

Note: The Fort Calhoun Unit 1 Pipe Material is A451-CPF8M


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Table 4-2 Measured Tensile Properties for Fort Calhoun Unit I PrimaryLoop Elbows

Yield Strength (psi) Ultimate Strength (psi)

Heat No. Room Temp. Room Temp.

27129-1 43500 85500

24968-1 31500 86750

32785-1 44300 83800

26895-2 48000 84000

26895-1 41400 82400

33801-2 39800 78300

32515-1 43100 84000

33756-1 42900 85000

33418-1 42000 85250

33867-2 46600 88100

33975-1 45000 85500

33712-1 46000 88300

33676-1 47750 87900

27418-1 39600 78600

26307-2 49100 87500

32515-2 40200 78100

33041-1 45800 86100

36972-3 45000 84000

Note: The Fort Calhoun Unit 1 Elbow Material is A351-CF8M


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Table 4-3 Mechanical Properties for Fort Calhoun Unit I Materials at Operating Temperatures

Lower Bound

Average UltimateYield Modulus of Strength

Temperature Strength Elasticity Yield Stress(OF) (psi) (psi) (psi) (psi)

Hot Leg 593 28716 25.335 x 106 19821 74976

Cold/Crossover 545 29519 25.575 x 106 20375 74976Leg I I I I

Poisson's ratio: 0.3


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WESTINGHOUSE NON-PROPRIETARY CLASS 3H4-9 4-9

7Table 4-4 Chemistry & Fracture Toughness Piping Properties of the Material Heats of Fort Calhoun Unit I

a,c,e


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Tale45 heitr &FWcueSToughOUssElo NO-RPrIpeTrYie CLASSe Maera 4et10otCahuUi

Table 4-5 Chemistry & Fracture Toughness Elbow Properties of the Material Heats of Fort Calhoun Unit I a,c,e

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Table 4-6 Fracture Toughness Properties Used to Evaluate Critical Locations

TraJc (in-lb/in2) Tmat (non- Jax (in-lb/in Heat NumberLocation i dimensional)

a,c,e

t I t


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5.0 CRITICAL LOCATION AND EVALUATION CRITERIA

5.1 CRITICAL LOCATIONS

The leak-before-break (LBB) evaluation margins are to be demonstrated for the critical locations(governing locations). Such locations are established based on the loads (Section 3.0) and thematerial properties established in Section 4.0. These locations are defined below for the FortCalhoun primary loop piping. The faulted loads from Table 3-2 and the weld locations fromFigure 3-2 are used for this evaluation.

Critical Locations

The primary loop is made of cast stainless steel for both pipes and elbows. The higheststressed locations for the primary loop are at Location 1 (in the hot leg) at the reactor vesseloutlet nozzle to pipe weld and Location 6 at the crossover leg elbow weld (for the cross-overand cold leg). Locations 1 and 6 (Figure 3-2) are the critical locations for all the weld locationsin the primary loop piping. Enveloping materials from all the heats (see Section 4) are used forthe LBB fracture mechanics leak rate (see Section 6) and stability (see Section 7) analyses.

5.2 FRACTURE CRITERIA

As will be discussed later, fracture mechanics analyses are made based on loads andpostulated flaw sizes related to leakage. The stability criteria against which the calculated J andtearing modulus are compared are:

(1) If Japp < J1, then the crack will not initiate and the crack is stable;

(2) If Japp > Jjc; and Tapp < Tmat and Japp < Jmax, then the crack is stable.

Where:

Japp = Applied JJ = J at Crack InitiationTapp = Applied Tearing ModulusTmat = Material Tearing ModulusJmax = Maximum J value of the material

For critical locations, the limit load method discussed in Section 7.0 was also used.

Critical Location and Evaluation Criteria July 2013WCAP-17262-NP Revision 1


6.0 LEAK RATE PREDICTIONS

6.1 INTRODUCTION

The purpose of this section is to discuss the method which is used to predict the flow throughpostulated through-wall cracks and present the leak rate calculation results for through-wallcircumferential cracks.

6.2 GENERAL CONSIDERATIONS

The flow of hot pressurized water through an opening to a lower back pressure causes flashingwhich can result in choking. For long channels where the ratio of the channel length, L, tohydraulic diameter, DH, (L/DH) is greater than

a,c,e

6.3 CALCULATION METHOD

The basic method used in the leak rate calculations is the method developed by

]a,c,e

The flow rateReference 6-2condition and

Reference 6-2).

through a crack was calculated in the following manner. Figure 6-1 fromwas used to estimate the critical pressure, P,, for the primary loop enthalpyan assumed flow. Once Pr was found for a given mass flow, the [

]a~c,e was found from Figure 6-2 (taken fromFor all cases considered, since [ ]ac.e

Therefore, this method will yield the two-phase pressure drop due to momentum effects asillustrated in Figure 6-3, where P, is the operating pressure. Now using the assumed flow rate,G, the frictional pressure drop can be calculated using

APf = [ ]a,c,e (6-1)

where the friction factor f is determined using the [c, was obtained from fatigue crack data on stainlessvalue used in these calculations was [ ]a,c,e

]ac'e The crack relative roughness,steel samples. The relative roughness

The frictional pressure drop using equation 6-1 is then calculated for the assumed flow rate andadded to the [ ]ace to obtain thetotal pressure drop from the primary system to the atmosphere.

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That is, for the primary loop:

Absolute Pressure - 14.7 = [ ]a,c,e (6-2)

for a given assumed flow rate G. If the right-hand side of equation 6-2 does not agree with thepressure difference between the primary loop and the atmosphere, then the procedure isrepeated until equation 6-2 is satisfied to within an acceptable tolerance which in turn leads toflow rate value for a given crack size. It should be noted that for FCS, the defined atmosphericpressure is 14.2 psi rather than 14.7 psi. This difference in pressure of 0.5 psi will havenegligible impact on leak-rate calculations.

6.4 LEAK RATE CALCULATIONS

Leak rate calculations were made as a function of crack length at the governing locationspreviously identified in Section 5.1. The normal operating loads of Table 3-1 were applied, inthese calculations. The crack opening areas were estimated using the method ofReference 6-3, and the leak rates were calculated using the two-phase flow formulationdescribed above. The average material properties of Section 4.0 (see Table 4-3) were used forthese calculations.

The flaw sizes to yield a leak rate of 10 gpm were calculated at the governing locations and aregiven in Table 6-1 for Fort Calhoun Unit 1. The flaw sizes so determined are called leakage flawsizes.

The Fort Calhoun Unit 1 RCS pressure boundary leak detection system meets the intent ofRegulatory Guide 1.45, and the plant leak detection capability is 1 gpm. Thus, to satisfy themargin of 10 on the leak rate, the flaw sizes (leakage flaw sizes) are determined which yield aleak rate of 10 gpm.

6.5 REFERENCES

6-1

a,c,e

6-2 M. M, EI-Wakil, "Nuclear Heat Transport, International Textbook Company," New York,N.Y, 1971.

6-3 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the CrackOpening Area of Circumferential and a Longitudinal Through-Crack in a Pipe,"Section I1-1, NUREG/CR-3464, September 1983.

6-4 D. Rudland, R. Wolterman, G. Wilkowski, R. Tregoning, "Impact of PWSCC and CurrentLeak Detection on Leak-Before-Break," proceedings of Conference on Vessel HeadPenetration, Inspection, Cracking, Repairs, Sponsored by USNRC, MarriotWashingtonian Center, Gaithersburg, MD, September 29 to October 2, 2003.

Leak Rate Predictions July 2013WCAP-1 7262-NP Revision 1


Table 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the

Governing Locations

Location Leakage Flaw Size (in)

1 5.28

6 6.64

Reactor Vessel inlet and outlet nozzle locations have Alloy 82/182 welds. These locations arealso analyzed for LBB with Alloy 82/182 material properties. Location 1 has higher faultedstress than location 14; therefore, Location 1 is selected for the LBB evaluations for the Alloy82/182 welds.

II

a,ce


July 2013Revision 1


a, c, e

Ic

STAGNATION ENTHALPY (102 tw/lb

Figure 6-1 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures


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a.-l

w

Ir.15

LEN42THADIAMETER RATIO MIDI

Figure 6-2 [ ]ac~e Pressure Ratio as a Function of LID


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a,c,e

Figure 6-3 Idealized Pressure Drop Profile Through a Postulated Crack

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7.0 FRACTURE MECHANICS EVALUATION

7.1 LOCAL FAILURE MECHANISM

The local mechanism of failure is primarily dominated by the crack tip behavior in terms ofcrack-tip blunting, initiation, extension and final crack instability. The local stability will beassumed if the crack does not initiate at all. It has been accepted that the initiation toughnessmeasured in terms of J1, from a J-integral resistance curve is a material parameter defining thecrack initiation. If, for a given load, the calculated J-integral value is shown to be less than theJic of the material, then the crack will not initiate. If the initiation criterion is not met, one cancalculate the tearing modulus (see equation A-14a of Reference 7-1) as defined by the followingrelation:

dJ ETapp da x

where:

Tapp

E

af

a

('y, (Yu

= applied tearing modulus

= modulus of elasticity

0.5 (oy + (yu) = flow stress

crack length

- yield and ultimate strength of the material, respectively

Stability is said to exist when ductile tearing does not occur if Tapp is less than Tmat, theexperimentally determined tearing modulus. Since a constant Tmat is assumed a furtherrestriction is placed in Japp. Japp must be less than Jmax where Jmax is the maximum value of J forwhich the experimental Tmat is greater than or equal to the Tapp used.

As discussed in Section 5.2 the local crack stability criteria is a two-step process:

(1) If Japp < J1c, then the crack will not initiate and the crack is stable;

(2) If Japp > Jj; and Tapp < Tmat and Japp < Jmax, then the crack is stable.

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7.2 GLOBAL FAILURE MECHANISM

Determination of the conditions which lead to failure in stainless steel should be done withplastic fracture methodology because of the large amount of deformation accompanyingfracture. One method for predicting the failure of ductile material is the plastic instabilitymethod, based on traditional plastic limit load concepts, but accounting for strain hardeningand taking into account the presence of a flaw. The flawed pipe is predicted to fail when theremaining net section reaches a stress level at which a plastic hinge is formed. The stresslevel at which this occurs is termed as the flow stress. The flow stress is generally taken asthe average of the yield and ultimate tensile strength of the material at the temperature ofinterest. This methodology has been shown to be applicable to ductile piping through alarge number of experiments and will be used here to predict the critical flaw size in theprimary coolant piping. The failure criterion has been obtained by requiring equilibrium ofthe section containing the flaw (Figure 7-1) when loads are applied. The detaileddevelopment is provided in Appendix A for a through-wall circumferential flaw in a pipe withinternal pressure, axial force, and imposed bending moments. The limit moment for such apipe is given by:

[ Ia,c,e

where:

a,c,e

0.5 (oy + ca,) = flow stress, psiGf

I

]ace

The analytical model described above accurately accounts for the piping internal pressure aswell as imposed axial force as they affect the limit moment. Good agreement was foundbetween the analytical predictions and the experimental results (Reference 7-2). For applicationof the limit load methodology, the material, including consideration of the configuration, musthave a sufficient ductility and ductile tearing resistance to sustain the limit load.


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7.3 RESULTS OF CRACK STABILITY EVALUATION

J-Integral Analysis:

Stability analyses were performed at the governing locations established in Section 5.1. Theelastic-plastic fracture mechanics (EPFM) J-integral analyses for through-wall circumferentialcracks in a cylinder were performed using the procedure in the EPRI fracture mechanicshandbook (Reference 7-3).

The lower-bound material properties were used. The fracture toughness properties establishedin Section 4.3 and the normal plus seismic loads given in Table 3-2 were used for the EPFMcalculations. The postulated flaw size was 2 times (for flaw size margin of 2) the leakage flawsize established in Section 6.0 (see Table 6-1). Evaluations were performed at the criticallocations identified in Section 5.1. The results of the elastic-plastic fracture mechanics J-integralevaluations are given in Table 7-1.

Limit Load Analysis:

A stability analysis based on limit load was also performed for these locations as described inSection 7.2. The weld process type, at the critical locations 1 and 6, is used as Shielded MetalArc Welding (SMAW). The "Z" correction factor for SMAW (References 7-4 and 7-5) is asfollows:

Z = 1.15 [1.0 + 0.013 (OD-4)] for SMAW

where OD is the outer diameter of the pipe in inches.

The Z-factors were calculated for the critical locations, using the dimensions given in Table 3-1.The applied faulted loads (Table 3-2) were increased by the Z-factors and plots of limit loadversus crack length were generated as shown in Figures 7-2 and 7-3. Lower bound materialproperties were used from Table 4-3. Table 7-2 summarizes the results of the stability analysesbased on limit load. The leakage flaw sizes are also presented on the same table.

The Alloy 82/182 weld has high toughness and it does not degrade due to the thermal aging andtherefore the LIMIT load method with a weld-process 'Z' factor of 1.0 should be used tocalculate the critical flaw sizes. However, a 'Z' factor was conservatively applied as shownbelow to account for the elastic-plastic consideration of the Alloy 82/182 weld material based onReference 7-6. The critical flaw size and the leakage flaw size for the Alloy 82/182 welds areshown at the bottom of Table 7-2.

Z = 0.0000022 x (OD)3 - 0.0002 x (OD)2 + 0.0064 x OD + 1.1355

where OD = Outside Diameter.

Fracture Mechanics Evaluation July 2013WCAP-1 7262-NP Revision I


7.4 REFERENCES

7-1 NUREG-1061 Volume 3, "Report of the U. S. Nuclear Regulatory Commission PipingReview Committee," November 1984.

7-2 Kanninen, M. F., et. al., "Mechanical Fracture Predictions for Sensitized Stainless SteelPiping with Circumferential Cracks," EPRI NP-192, September 1976.

7-3 Kumar, V., German, M. D. and Shih, C. P., "An Engineering Approach for Elastic-PlasticFracture Analysis," EPRI Report NP-1931, Project 1237-1, Electric Power ResearchInstitute, July 1981.

7-4 Standard Review Plan; Public Comment Solicited; 3.6.3 Leak-Before-Break EvaluationProcedures; Federal Registerfol. 52, No. 167/Friday, August 28, 1987/Notices,pp. 32626-32633.

7-5 NUREG-0800 Revision 1, March 2007, Standard Review Plan: 3.6.3 Leak-Before-BreakEvaluation Procedures.

7-6 ASME Pressure Vessel and Piping Division Conference Paper PVP2008-61840,"Technical Basis for Revision to Section XI Appendix C for Alloy 600/82/182/132 FlawEvaluation in Both PWR and BWR Environments," July 28-31, Chicago IL, USA.

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Table 7-1 Stability Results for Fort Calhoun Based on Elastic-Plastic J-Integral

Evaluations

Fracture Criteria Calculated Values

Flaw Jic Jmax Japp

Location Size* (in) (in-lb/in 2) Tmat (in-lb/in2) (in-lb/in2) Tapp a,c,e

KTable 7-2 Stability Results for Fort Calhoun Based on Limit Load

Critical Location Critical Flaw Size (in) Leakage Flaw Size (in)

1 35.27 5.28

6 34.16 6.64

I]a,c,e

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WESTINGHOUSE NON-PROPRIETARY CLASS 3 • v

WESTNGHOSE ON-POPRITAR CLAS 37-6

of

Figure 7-1 Fully Plastic Stress Distribution


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a,c,e

OD = 38.50 in. Uy-min = 19.82 ksi F = 1964 kips

t = 2.688 in. Ou-min = 74.98 ksi M = 23806 in-kips

SA451-CPF8M with SMAW Weld

Note: OD = outer diameter, t = thickness

Figure 7-2 Critical Flaw Size Prediction - Location I


July 2013Revision 1


a,c,e

OD = 29.00 in. Gy-min = 20.38 ksi F = 1075 kips

t = 1.938 in. ou-min = 74.98 ksi M = 4368 in-kips

SA-351 CF8M with SMAW Weld

Note: OD = outer diameter, t = thickness

Figure 7-3 Critical Flaw Size Prediction - Location 6


July 2013Revision 1


8.0 FATIGUE CRACK GROWTH ANALYSIS

To determine the sensitivity of the primary coolant system to the presence of small cracks, afatigue crack growth analysis was carried out at the reactor vessel outlet nozzle safe-end regionfor Fort Calhoun (see Location 1 in Figure 3-2). This region was selected because crack growthcalculated here will be typical of that in the entire primary loop; crack growths calculated at otherlocations can be expected to show less than 10% variation. Fatigue crack growth is the onlycredible crack growth mechanism.

A finite element stress analysis was carried out for the outlet nozzle safe end region withrepresentative nozzle dimensions to determine the stresses resulting from thermal transientsand mechanical loading. These stresses were then combined with conservative weldingresidual stress distributions at the stainless steel and Alloy 82/182 welds to determine thefatigue crack growth for postulated flaws in the various materials at the nozzle safe end region.Table 8-1 summarizes the transients and design cycles applicable for Fort Calhoun Unit 1(Reference 8-2). Extended Power Uprate (EPU) and Measurement Uncertainty Recapture(MUR) transients were not considered in Revision 1 of this report.

Circumferentially oriented surface flaws were postulated at three different locations at the outletnozzle, these locations are: safe end to pipe stainless steel weld, safe end to nozzle Alloy82/182 weld, and the ferritic steel nozzle (Figure 8-1). The total stress at each region was usedto generate crack tip stress intensity factors (K1) for circumferential flaw Aspect Ratios -AR (flawlength/flaw depth) of 2, 6, and 10. Fatigue crack growth analyses were performed for thepostulated flaws with initial flaw depths of 10% through-wall thickness.

Fatigue crack growth rate laws were used from the ASME Code Section Xl (Reference 8-3) forthe ferritic steel and stainless steel as shown in Figures 8-2 and 8-3 respectively. The fatiguecrack growth law for the safe end to nozzle region (Alloy 82/182) was derived from Reference8-1 and illustrated in Figure 8-4. The laws were all structured for applicability to pressurizedwater reactor environments.

The calculated fatigue crack growth results for the semi-elliptic surface flaws of circumferentialorientation and various depths are summarized in Tables 8-2 through 8-4, and the results showthat the crack growth is very small. Therefore, it is concluded that surface flaws will not becomethrough-wall flaws for the end of service life of the plant and also applicable for the period ofextended operation (Section 4.3.1 of Reference 8-4 has indicated that OPPD does not expectthe number of design cycles for the transients that are counted to be exceeded during theperiod of extended operation).

Fatigue Crack Growth Analysis July 2013WCAP-1 7262-NP Revision 1


8.1 REFERENCES

8-1 NUREG/CR-6721, ANL-01/07, "Effects of Alloy Chemistry, Cold Work, and WaterChemistry on Corrosion Fatigue and Stress Corrosion Cracking of Nickel Alloys andWelds," April 2001.

8-2 Fort Calhoun Station Updated Safety Analysis Report (USAR-4.2), Revision 12.

8-3 ASME Boiler and Pressure Vessel Code Section XI, "Rules for Inservice Inspection ofNuclear Power Plant Components," 2001 Edition.

8-4 NUREG-1782, "Safety Evaluation Report Related to the License Renewal of the FortCalhoun Station, Unit 1," Docket No. 50-285.

Fatigue Crack Growth AnalysisWCAP-1 7262-NP

July 2013Revision 1


Table 8-1 Reactor Coolant System Operating Transients

Number Transient Identification DesignCycles

1 Plant Heatup at 100°F/hr 5002 Plant Cooldown at 100°F/hr 5003 Plant Loading at 10% of Full Power/min 150004 Plant Unloading at 10% of Full Power/min 150005 Step Load Increase of 10% of Full Power 20006 Step Load Decrease of 10% of Full Power 20007 Normal Plant Variation* 1068 Reactor Trip 4009 Loss of Reactor Coolant Flow 4010 Abnormal Loss of Load 4011 Loss of Secondary Pressure 512 Hydrostatic Test, 3125 psia 1013 Plant Leak Test, 2250 psia 200

*The RCS average temperature for purpose of design is assumed to increase and decrease a maximum of 6°F in 1 minute. The

corresponding RCS pressure variation is less than 100 psi. This transient has an insignificant effect on fatigue crack growth and was notincluded in this evaluation.


July 2013Revision 1


m

L

Table 8-2 FCG at Alloy 82/182 Weld (Nozzle to Safe-end Weld) - Outlet Nozzle

Thickness = 3.031 in., Inside Radius = 16.219 in.Initial Flaw Size = 10% Through Original Wall Thickness

Initial Flaw Final Flaw Depth at Final Flaw Ratio (alt)Flaw Configuration Depth (in.) End of Service Life at End of Service

(in.) Life

Table 8-3 FCG at Stainless Steel Weld (Safe-end to Pipe Weld) - Outlet Nozzle


Initial Flaw Final Flaw Depth at Final Flaw Ratio (alt)Flaw Configuration Depth (in.) End of Service Life at End of Service

(in.) Life 1ITable 8-4 FCG at Reactor Vessel Nozzle Location - Outlet Nozzle


Initial Flaw Final Flaw Depth at Final Flaw Ratio (a/t)Flaw Configuration Depth (in.) End of Service Life at End of Service Life

(in.)

__ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ 1

a,c,e

a,c,e

[

[,ce

m I

Fatigue Crack Growth AnalysisWCAP-17262-NP

July 2013Revision 1


Figure 8-1 Reactor Vessel Outlet Nozzle with Stress Cut Locations


July 2013Revision 1


1000

700

500

200

S0

U

0

UC0b

0b

0

0U0

c-i

100

70

50

20

10

7

5 *Linear interpolation is recommendedto account for R ratio dependenceof water environment curves, for0.25 < R < 0.65 for steep slope:

da = (1.02 X 10"6) 01 AK5

"9 5

dN

12 5 7 10 20

Stress Intensity Factor Range (AK 1 ksi 41'.)

50 70 100

Figure 8-2 Reference Fatigue Crack Growth Curves for Carbon & Low Alloy FerriticSteels


July 2013Revision 1


3 X 10-4

10o-4

10,5

TC

b•

10.6

10-7

6 10 20

AK (ksi ')

50 100

Note: A Factor of 2.0 is applied to the Air Environment Curve to Represent crack growth rate inPWR Environment

Figure 8-3 Reference Fatigue Crack Growth Curves for Stainless Steels


July 2013Revision 1


10-6

EE

1 -1l I/ , . . J .... ...... 10-111/ " . , . ... j . .... J10.11 10-10 10-9 10-8 10-7 10-6 10-11 10.10 10 .g 10-8

CGRair (MIs) CGRuir (r/s)

10-7 10-5

10-13

L),

:Alloy 600

- C-4.M34e-14+1.6216e-i6 T-1,4896e-1BT 2+t4.3546e-21 T3

o ANL0 James

Ix Amnzallag et al.L Naganto et at

1It U- ~.&..&.I..4.. .I.4..~..

0 100 200 300 400

Temperature (°C)500 600

Figure 8-4 Reference Fatigue Crack Growth Curves for Alloy 821182 Welds


July 2013Revision 1


9.0 ASSESSMENT OF MARGINS

The results of the leak rates of Section 6.4 and the corresponding stability evaluations ofSection 7.3 are used in performing the assessment of margins. Margins are shown in Tables 9-1 and 9-2. All of the LBB recommended margins are satisfied.

In summary, at all the critical locations relative to:

1. Flaw Size - Using faulted loads obtained by the absolute sum method, a margin of 2 ormore exists between the critical flaw and the flaw having a leak rate of 10 gpm (theleakage flaw).

2. Leak Rate - A margin of 10 exists between the calculated leak rate from the leakage flawand the plant leak detection capability of 1 gpm according to FCS Technical Specification2.1.4 (Reference 9-1).

3. Loads - At the critical locations the leakage flaw was shown to be stable using thefaulted loads obtained by the absolute sum method (i.e., a flaw twice the leakage flawsize is shown to be stable; hence the leakage flaw size is stable). A margin of 1 on loadsusing the absolute summation of faulted load combinations is satisfied.

9.1 REFERENCES

9-1 Omaha Public Power District, Fort Calhoun Station Unit 1, Renewed Facility OperatingLicense No. DPR-40, Appendix A, Technical Specification 2.1.4, "Reactor CoolantSystem Leakage Limits."

Assessment of MarginsWCAP-1 7262-NP

July 2013Revision 1


Table 9-1 Leakage Flaw Sizes, Critical Flaw Sizes by Limit Load and

Margins for Fort Calhoun

Critical Critical Flaw Size* Leakage Flaw Size Margin

1 35.27 5.28 6.7

6 34.16 6.64 5.1*Linmit Load Method

I

a,c,e

Table 9-2 Stability Results for Fort Calhoun Based on Elastic-Plastic J-IntegralEvaluations

Fracture Criteria Calculated Values

Critical Flaw Jic Jmax Japp

Location Size* (in) (in-lblin 2) Tmat (in-lb/in2) (in-lb/in 2) Tapp a,c,e.-- II I-

L i-i

Assessment of MarginsWCAP-17262-NP

July 2013Revision I


10.0 CONCLUSIONS

This report justifies the elimination of RCS primary loop pipe breaks from the structural designbasis for the Fort Calhoun Unit 1 as follows:

a. Stress corrosion cracking is precluded by use of fracture resistant materials in thepiping system and controls on reactor coolant chemistry, temperature, pressure,and flow during normal operation.Mitigation measures for PWSCC sensitive Alloy 82/182 welds will be implementedas required on the Fort Calhoun Reactor Vessel nozzles weld locations.

b. Water hammer should not occur in the RCS piping because of system design,testing, and operational considerations.

c. The effects of low and high cycle fatigue on the integrity of the primary piping arenegligible.

d. Ample margin exists between the leak rate of small stable flaws and the capabilityof the Fort Calhoun reactor coolant system pressure boundary Leakage DetectionSystem.

e. Ample margin exists between the small stable flaw sizes of item (d) and largerstable flaws.

f. Ample margin exists in the material properties used to demonstrate end-of-servicelife (fully aged) stability of the critical flaws.

For the critical locations, flaws are identified that will be stable because of the ample marginsdescribed in d, e, and f above. The LBB assessment was performed based on deterministicfracture mechanics methods that have been previously approved by the NRC.

Based on the above, the leak-before-break conditions and margins are satisfied for the FortCalhoun Unit 1 primary loop piping. It is demonstrated that the dynamic effects of the piperupture resulting from postulated breaks in the reactor coolant primary loop piping need not beconsidered in the structural design basis for Fort Calhoun Unit 1 for the license renewal periodwith current power operating conditions.

Conclusions July 2013WCAP-1 7262-NP Revision 1

WESTINGHOUSE NON-PROPRIETARY CLASS 3

APPENDIX A: LIMIT MOMENT

A-1

I

a,c,e

Appendix A: Limit MomentWCAP-17262-NP

July 2013Revision 1

WESTINGHOUSE NON-PROPRIETARY CLASS 3 A-2

Figure A-1 Pipe with a Through-Wall Crack in Bending

Appendix A: Limit Moment July 2013WCAP-1 7262-NP Revision 1

Revise the structural design basis for reactor coolant system (RCS) pipingdescribed in USAR Section 4.3.6.

NRC and industry initiatives using fracture mechanics technology andmaterial toughness demonstrate that leak-before-break (LBB) criteria can beapplied to RCS piping.

Eliminate RCS piping rupture from consideration in the FCS structural

design basis

LBB methodology performed by Westinghouse using the latest LBB criteria.

The analysis was performed using the existing RCS leak detectioncapability and piping stress analysis loads. Thermal aging considerationswere performed using approved Westinghouse methodology.

The analysis is applicable for the period of extended operation (PEO), whichbegins at midnight on August 9, 2013.

No Technical Specification (TS) changes are proposed in this licenseamendment request (LAR).

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