ML031910667

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WCAP-16020-NP, Rev 0, Technical Justification for Eliminating 12 Residual Heat Removal (RHR) Lines Rupture as the Structural Design Basis for Callaway Nuclear Power Plant.
ML031910667
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Site: Callaway Ameren icon.png
Issue date: 02/28/2003
From: Bhowmick D, Ching Ng, Petsche G, Swamy S
Westinghouse
To:
Office of Nuclear Reactor Regulation
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ML031910683 List:
References
FOIA/PA-2005-0108 WCAP-16020-NP, Rev 0
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r Westinghouse Non-Proprietary Class 3 WCAP-16020-NP February 2003 Revision 0 Technical Justification for Eliminating 12" Residual Heat Removal (RHR) Lines Rupture as the Structural Design Basis For Callaway Nuclear Power Plant

  • Westinghouse

WESTINGHOUSE NON-PROPRIETARY CLASS 3 WCAP-16020-NP Revision 0 Technical Justification for Eliminating 12" Residual Heat Removal (RHR) Lines Rupture as the Structural Design Basis for Callaway Nuclear Power Plant D. C. Bhowmick C. K. Ng February 2003 Verified: -

F. Petsche Structural Mechanics Technology Approved:

S.AA amy, nager Structural Mechanics Technology Westinghouse Electric Company LLC P.O. Box 355 Pittsburgh, PA 1523D-0355 0 2003 Westinghouse Electric Company LLC All Rights Reserved

TABLE OF CONTENTS LIST OF TABLES .............. v LIST OF FIGURES .. ............ vi 1 INTRODUCTION .. 1-1 1.1 Background .1-1 1.2 Scope and Objective .1-1 1.3 References .1-2 2 OPERATION AND STABILITY OF THE RHR LINES .. 2-1 2.1 Stress Corrosion Cracking .2-1 2.2 Water Hammer .2-2 2.3 Low Cycle and High Cycle Fatigue .2-2 2.4 Other Possible Degradation During Service of the RHR Lines .2-2 3 MATERIAL CHARACTERIZATION .. 3-1 3.1 Pipe Materials And Weld Process...........................................................................3-1 3.2 Material Properties .3-1 3.3 Reference .3-1 4 LOADS FOR FRACTURE MECHANICS ANALYSIS . .4-1 4.1 Nature of the Loads .4-1 4.2 Loads for Crack Stability Analysis .4-2 4.3 Loads for Leak Rate Evaluation .4-2 4.4 Summary of Loads and Geometry for the RHR Lines .4-2 4.5 Governing Locations for the RHR Lines .4-3 5 FRACTURE MECHANICS EVALUATION .. 5-1 February 2003

iv 5.1 Global Failure Mechanism ......................................... 5-1 5.2 Leak Rate Predictions ......................................... 5-2 5.2.1 General Considerations ......................................... 5-2 5.2.2 Calculation Method ......................................... 5-2 5.2.3 Leak Rate Calculations......................................... 5-3 5.3 Stability Evaluation ......................................... 5-4 5.4 References.............................................................................................................5-4 6 ASSESSMENT OF FATIGUE CRACK GROWTH ..................................... 6-1 6.1 Introduction ......................................... 6-1 6.2 Critical Locabon For Fatigue Crack Growth Analysis ......................................... 6-1 6.3 Design Transients ......................................... 6-1 6.4 Stress Analysis ......................................... 6-1 6.5 OBE Loads ......................................... 6-2 6.6 Total Stress For Fatigue Crack Growth ......................................... 6-2 6.7 Fatigue Crack Growth Analysis ......................................... 6-2 6.7.1 Analysis Procedure ......................................... 6-2 6.8 Results ......................................... 6-5 6.9 References ......................................... 6-5 7 ASSESSMENT OF MARGINS .......................................... 7-1 8 CONCLUSIONS ......................................... . 8-1 APPENDIX A - LIMIT MOMENT.............................................................................................. A-1 February 2003

V LIST OF TABLES Table 3-1: Room Temperature Material Properties for the RHR Lines .................................. 3-2 Table 3-2: Representative Tensile Properties for the RHR Lines at Operating Temperatures ................................................................... 3-2 Table 3-3: Modulus of Elasticity (E) for the RHR Lines ........................................................... 3-2 Table 4-1: Summary of Callaway Nuclear Power Plant Piping Geometry and Normal Operating Condition for the Residual Heat Removal Line Loop 1........................ 4-4 Table 4-2: Summary of Callaway Normal Loads and Stresses for Residual Heat Removal Line Loop 1................................................................... 4-5 Table 4-3: Summary of Callaway Nuclear Power Plant Faulted Loads and Stresses for Residual Heat Removal Line Loop 1.................................................................. 4-6 Table 4-4: Summary of Callaway Nuclear Power Plant Piping Geometry and Normal Operating Condition for Residual Heat Removal Line Loop 4 ............................... 4-7 Table 4-5: Summary of Callaway Nuclear Power Plant Normal Loads and Stresses for Residual Heat Removal Line Loop 4 .................................................................. 4-8 Table 4-6: Summary of Callaway Nuclear Power Plant Faulted Loads and Stresses for Residual Heat Removal Line Loop 4 .................................................................. 4-9 Table 5-1 : Leakage Flaw Sizes .................................................................. 5-5 Table 5-2: Summary of Critical Flaw Sizes .................................................................. 5-5 Table 6-1: Design Transients Considered for Fatigue Crack Growth Evaluation .................... 6-6 Table 6-2: RHR Lines Fatigue Crack Growth Results ............................................................ 6-7 Table 7-1 : Leakage Flaw Sizes, Critical Flaw Sizes and Margins ........................................... 7-2 Table 7-2: LBB Conservatism .................................................................. 7-2 February 2003

Vi LIST OF FIGURES Figure 3-1 Callaway Nuclear Power Plant RHR Line Loop 1 Layout ..................................... 3-3 Figure 3-2 Callaway Nuclear Power Plant RHR Line Loop 4 Layout ..................................... 3-4 Figure 4-1 Governing Weld Location for RHR Line Loop 1................................................. 4-10 Figure 4-2 Governing Weld Location for RHR Line Loop 4 ................................................. 4-11 Figure 5-1 Fully Plastic Stress Distribution............................................................................5-6 Figure 5-2 Analytical Predications of Critical Flow Rates of Steam-Water Mixtures .............. 5-7 Figure 5-3 [ ]ace Pressure Ratio as a Function of UD ............................... 5-8 Figure 5-4 Idealized Pressure Drop Profile through a Postulated Crack ............................... 5-9 Figure 5-5 Loads acting on the Model at the Governing Locations ..................................... 5-10 Figure 5-6 Critical Flaw Size Prediction for Node 3020 (RHR Line Loop 4) ........ ................ 5-11 Figure 5-7 Critical Flaw Size Prediction for Node 3285 (RHR Line Loop 1)........................ 5-12 Figure 6-1 Schematic of RHR Line at RCL Hot Leg Nozzle Weld Location ........................... 6-8 Figure 6-2 Reference Crack Growth Curves for Stainless Steel in Air Environments ............ 6-9 Figure A-1 Pipe with A Through-Wall Crack in Bending ................................................... A-2 February 2003

1-1 1 INTRODUCTION

1.1 BACKGROUND

The current structural design basis for the Residual Heat Removal (RHR) lines requires postulating non-mechanistic circumferential and longitudinal pipe breaks. This results in additional plant hardware (e.g., pipe whip restraints and jet shields) which would mitigate the dynamic consequences of the pipe breaks. It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for the RHR lines. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that can be used for establishing that a circumferential type of break will not occur within the RHR lines. The evaluations consider that circumferentially oriented flaws cover longitudinal cases.

1.2 SCOPE AND OBJECTIVE The scope of this report is limited to the high energy Class 1 portion of the RHR lines (primary loop junction to the second isolation valve). A schematic drawing of the piping system is shown in Section 3. The recommendations and criteria proposed in SRP 3.6.3 (Reference 1-2) are used in this evaluation. The criteria and the resulting steps of the evaluation procedure can be briefly summarized as follows:

1. Calculate the applied loads. Identify the location(s) at which the highest faulted stress occurs.
2. Identify the materials and the material properties.
3. Postulate a surface flaw at the governing location. Determine fatigue crack growth.

Show that a through-wall crack will not result.

4. Postulate a through-wall flaw at the governing location(s). The size of the flaw should be large enough so that the leakage is assured of detection with margin using the installed leak detection equipment when the pipe is subjected to normal operating loads.

Demonstrate that there is a margin of 10 between the calculated leak rate and the leak detection capability.

5. Using maximum faulted loads in the stability analysis, demonstrate that there is a margin of 2 between the leakage size flaw and the critical size flaw.
6. Review the operating history to ascertain that operating experience has indicated no particular susceptibility to failure from the effects of corrosion, water hammer or low and high cycle fatigue.
7. For the materials types used in the Plant, provide representative material properties.

Introduction February 2003

1-2 The leak rate is calculated for the normal operating condition. The leak rate prediction model used in this evaluation is an [

ace. The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through-wall flaw to normal operating loads (Reference 1-3). Surface roughness is accounted for in determining the leak rate through the postulated flaw.

It should be noted that the terms Oflaw and 'crack' have the same meaning and are used interchangeably. Goveming location" and Ncritical location" are also used interchangeably throughout the report.

1.3 REFERENCES

1-1 WCAP-7211, Revision 4, "Energy Systems Business Unit Policy and Procedures for Management, Classification, and Release of Information," January 2001.

1-2 Standard Review Plan; public comments solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal RegisterNol. 52, No. 167/Friday, August 28, 1987/Notices, pp.

32626-32633.

1-3 NUREG/CR-3464, 1983, 'The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through Wall Cracks."

IntroductionFeray20 Introduction February 2DO3

2-1 2 OPERATION AND STABILITY OF THE RHR LINES 2.1 STRESS CORROSION CRACKING The Westinghouse Reactor Coolant System (RCS) Class 1 lines have an operating history that demonstrates the inherent operating stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking, IGSCC). This operating history totals over 1100 reactor-years, including 5 plants each having over 30 years of operation, 4 plants each with over 25 years of operation, 12 plants each with over 20 years of operation and 8 plants each with over 15 years of operation.

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

The elements of a water environment known to increase the susceptibility of austenitic stainless steel to stress corrosion are: oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment. Prior to being put into service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled 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 within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the major water chemistry control standards being included in the plant operating procedures as a condition for plant operation. For example, during normal power operation, oxygen concentration in the RCS Class 1 lines is expected to be in the parts per billion (ppb) range by controlling charging flow chemistry and maintaining hydrogen in the reactor coolant at specified concentrations. Maintaining concentrations of chlorides and fluorides within the specified limits also stringently controls halogen concentrations. This is assured by controlling charging flow chemistry. Thus during plant operation, the likelihood of stress corrosion cracking is minimized.

Wall thinning by erosion and erosion-corrosion effects will not occur in the RHR lines due to the low velocity, and the material, austenitic stainless steel, is highly resistant to these degradation mechanisms. Therefore, wall thinning is not a significant concern in the portion of the system being addressed in this evaluation.

Stailityof ines ebruay 200 of th RHRadLines the RHR Operation and Stability Operaion February 2003

2-2 As a result of the recent issue of Primary Water Stress Corrosion Cracking (PWSCC) occurring in V. C. Summer reactor vessel hot leg nozzle, Alloy 82/182 weld is being currently investigated under the EPRI Materials Reliability Project (MRP) Program. It should be noted that the susceptible material under investigation is not found in the RHR lines at the Callaway Nuclear Power Plant.

2.2 WATER HAMMER Overall, there is a low potential for water hammer in the RCS and connecting RHR lines since they are designed and operated to preclude the voiding condition in normally filled lines. The RCS and connecting RHR lines including piping and components are designed for normal, upset, emergency, and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design.

Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by the control rod positions; pressure is controlled also within a narrow range for steady-state conditions by the pressurizer heaters and pressurizer spray. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system and the connecting auxiliary lines. Preoperational testing and operating experience has verified the Westinghouse approach. The operating transients of the RCS primary piping and connected RHR lines are such that no significant water hammer can occur.

2.3 LOW CYCLE AND HIGH CYCLE FATIGUE An assessment of the low cycle fatigue loading is discussed in Section 6 as part of this study in the form of a fatigue crack growth analysis.

Pump vibrations during operation would result in high cycle fatigue loads in the piping system.

During operation, an alarm signals the exceedance of the RC pump shaft vibration limits. Field measurements have been made on the reactor coolant loop piping in a number of plants during hot functional testing. Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest. Field measurements on typical PWR plants indicate vibration amplitudes less than 1 ksi. When translated to the connecting RHR lines, these stresses would be even lower, well below the fatigue endurance limit for the RHR line materials and would result in an applied stress intensity factor below the threshold for fatigue crack growth.

2.4 OTHER POSSIBLE DEGRADATION DURING SERVICE OF THE RHR LINES Thermal stratification occurs when conditions permit hot and cold layers of water to exist simultaneously in a horizontal pipe. This can result in significant thermal loadings due to the Operation and Stability of the RHR Lines O;erationandStability of the RHR Lines February 2003

2-3 high fluid temperature differentials. Changes in the stratification state result in thermal cycling, which can cause fatigue damage. This was an important issue in PWR feedwater line and pressurizer surge line piping, where temperature differentials of 3000 F were not uncommon.

For the RHR piping in the Callaway Nuclear Power Plant, thermal stratification is not a concern during normal plant operation, since the unisolable piping extending from the hot leg to the isolation valve is relatively short. This ensures that turbulant penetration from the hot leg will heat the piping and preclude stratification. In addition, thermal stratification is also not a concern during RHR operation, since the flow rate is sufficiently high to preclude stratification.

The RHR Lines and the associated fittings for Callaway Nuclear Power Plant are forged product forms, which are not susceptible to toughness degradation due to thermal aging.

The maximum normal operating temperature of the RHR piping is about 6190F. This is well below the temperature that would cause any creep damage in stainless steel piping.

of th RHie eray20 Oprto an Operation and Stblt Stability of the RHR Lines February 2003

3-1 3 MATERIAL CHARACTERIZATION 3.1 PIPE MATERIALS AND WELD PROCESS The material types of the RHR lines for the Callaway Nuclear Power Plant are SA376 TP304, SA312 TP304 and SA403 WP304. They are wrought product of the types used for the piping in several PWR plants. The RHR line system does not include any cast pipes or cast fittings.

The welding processes used are Gas Tungsten Arc Weld (GTAW) and Shielded Metal Arc Weld (SMAW) combination or GTAW. Figures 3-1 and 3-2 show the schematic layout of the RHR lines Loop 1 and 4 respectively and also identify the weld location by node points.

In the following sections the tensile properties of the materials are presented for use in the Leak-Before-Break analyses.

3.2 MATERIAL PROPERTIES The room temperature mechanical properties of the Callaway Nuclear Power Plant RHR lines material were obtained from the Certified Materials Test Reports (CMTRs) and are given in Table 3-1. The material properties at temperatures (700F and 619 0F) are required for the leak rate and stability analyses. The minimum and average tensile properties at the temperatures of interest stated above were calculated by using the ratio of the ASME Code Section II (Reference 3-1) properties and those tabulated in Table 3-1. Table 3-2 shows the representative minimum and average tensile properties at various operating temperatures. The modulus of elasticity values was established at various temperatures from the ASME Code Section II (see Table 3-3). In the Leak-Before-Break evaluation, the representative minimum yield and minimum ultimate strengths at operating temperature were used for the flaw stability evaluations and the representative average yield strength properties were used for the leak rate predictions. These properties are summarized in Table 3-2.

3.3 REFERENCE 3-1 ASME Boiler and Pressure Vessel Code Section II, Part D - Material Properties, 2001 Edition, July 1, 2001, ASME Boiler and Pressure Vessel Committee, Subcommittee on Materials.

Characterization February 2003 Material Material Characterization February 2003

3-2 Table 3-1: Room Temperature Material Properties for the RHR Lines Heat No. (SIN)* Material Yield Strength Ultimate Strength (psi) (Psi)

L3384 (348) SA376 TP304 41200 81100 ERLG (441) SA403 WP304 37860 80730 EROD (437) SA403 WP304 37440 82630 J6346 (348) SA376 TP304 42400 84900 ERLG (439) SA403 WP304 37860 80730 6473 (782) SA376/SA312 TP304 40300 84500 D9711 (790) SA403 WP304 36800 81300 D9553 (784) SA403 WP304 40200 84700 5-621 (451) SA312 TP304 43200 89300 J6346 (449) SA376 TP304 42400 84900

  • S/N: Serial Number Table 3-2: Representative Tensile Properties for the RHR Lines at Operating Temperatures Minimum Average Minimum Material Temperature Yield Yield Ultimate l(F) (psi) (psi) (psi)

SA376/SA312 TP304 or SA403 WP304 619 22384 24310 68244 SA376/SA312 TP304 or SA403 WP304 70 36800 39966 80730 Table 3-3: Modulus of Elasticity (E) for the RHR Lines Temperature (OF) E (0t6 psi) 619 25.205 70 28.300 February 2003 Material Characterization Material Characterization February 2003

3-3 3285 3240

'3230 3150 3180 Iot Leg Loop I 3090 3060 Figure 3-1 Callaway Nuclear Power Plant RHR Line Loop 1 Layout Material Characterization February 2003

3-4 3580 3520 Hot Leg Loop 4 23500 r3030

3060 3400 3080 3340 3120 3320 3290 3140 Figure 3-2 Callaway Nuclear Power Plant RHR Line Loop 4 Layout Material Ch r cerz Maera.ebrary.2003 to Characterization February 2003

4-1 4 LOADS FOR FRACTURE MECHANICS ANALYSIS 4.1 NATURE OF THE LOADS Figure 3-1 and Figure 3-2 show schematic layouts of the RHR lines Loop 1 and Loop 4 respectively for Callaway Nuclear Power Plant and identify the weld locations by node points.

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

F M (4-1)

A Z

where, a = Stress F = Axial Load M = Moment A = Metal Cross-Sectional Area Z = Section Modulus The moment for the desired loading combination was calculated by the following equation:

M= VM+M2+Ml (4-2)

where, M = Moment For Required Loading Mx = Torsional Moment My = Y Component of Bending Moment MZ = Z Component of Bending Moment The axial load and moments for crack stability analysis and leak rate predictions are computed by the methods to be explained in Sections 4.2 and 4.3.

Mechanics Analysis Febwary 2003 Loads Fracture Mechanics for Fracture Loads for Analysis February 2003

4-2 4.2 LOADS FOR CRACK STABILITY ANALYSIS In accordance with Standard Review Plan 3.6.3 the absolute sum of loading components can be applied which results in higher magnitude of combined loads. If crack stability is demonstrated using these loads, the LBB margin on loads can be reduced from 42 to 1.0. The faulted loads for the crack stability analysis were calculated by the absolute sum method as follows:

F = IFDWI + IFTHI + IFpI + IFssEI (4-3)

MX= IMxDwI + IMxTHI + IMXSSEI (4-4)

My= 1MyDw + IMYTHI + IMyssEl (4-5)

MZ= IMzDwI + IMZTHI + IMZSSEI (4-6) where DW = Deadweight TH = Normal Thermal Expansion Load P = Load Due To Internal Pressure SSE = Safe Shutdown Earthquake Loading Including Seismic Anchor Motion 4.3 LOADS FOR LEAK RATE EVALUATION The normal operating loads for the leak rate predictions were calculated by the algebraic sum method as follows:

F = FDW + FTH + Fp (4-7)

MX = MX DW + MX TH (4-8)

MY = MYDW + MYTH (4-9)

MZ = MZ DW + MZTH (4-10)

The parameters and subscripts are the same as those explained in Sections 4.1 and 4.2.

4.4

SUMMARY

OF LOADS AND GEOMETRY FOR THE RHR LINES The load combinations were evaluated at the various weld locations. Normal loads were determined using the algebraic sum method whereas the faulted loads were combined using the absolute sum method. The normal operating loadings for the RHR lines are Pressure (P),

Deadweight (DW) and Normal Operating Thermal Expansion (TH) loads. The faulted loadings February 2003 Loads for Loads Mechanics Analysis Fracture Mechanics for Fracture Analysis February 2003

4-3 consist of Normal Operating loads plus Safe Shutdown Earthquake (SSE) loads including the Seismic Anchor Motion.

Tables 4-1 and 4-4 show the piping geometry and normal operating condition for the RHR lines Loop 1 and Loop 4 at the weld locations respectively. The minimum pipe wall thickness at the weld counterbore is used in the analysis. The normal and faulted loads are tabulated in Tables 4-2 and 4-3 respectively at the weld locations for RHR Line Loop 1, while Tables 4-5 and 4-6 show the normal and faulted loads for RHR Line Loop 4.

4.5 GOVERNING LOCATIONS FOR THE RHR LINES All the welds at the RHR lines Loops 1 and 4 are fabricated using the GTAW/SMAW combination or GTAW weld process procedures. The governing locations were established on the basis of the pipe schedules, material type, operating temperature, operating pressure, and the highest faulted stresses at the welds. Figures 4-1 and 4-2 show the schematic layouts of the RHR lines Loop 1 and Loop 4 respectively for the Callaway Nuclear Power Plant and also identify the governing weld locations.

The governing locations enveloping both RHR lines Loop 1 and Loop 4 are found to be:

Node 3285 (RHR Line Loop 1)

Node 3020 (RHR Line Loop 4)

Loads for Fracture Mechanics Analysis Loadsfor Fracture MechanicsAnalysis February 2003

4-4 Table 4-1 : Summary of Callaway Nuclear Power Plant Piping Geometry and Normal Operating Condition for the Residual Heat Removal Line Loop 1 Weld Outer Minimum Normal Operating Location Material Diameter Wall Pressure Temperature Node Type (in) Thickness (psig) ( F) 3020 SA376/SA31(2 TP304 or SA403 WP304 12.750 1.005 2235 619 3030 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 3040 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 3050 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 3060 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 3090 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 3120 SA3761SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3150 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3160 SA3761SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3180 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3190 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3230 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3240 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3280 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3285 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 3300 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 February 2003 Loads for Loads Mechanics Analysis Fracture Mechanics for Fracture Analysis February 2003

4-5 Table 4-2: Summary of Callaway Normal Loads and Stresses for Residual Heat Removal Line Loop 1 Location Axial Force Moment Axial Stress Moment Total Stress Node (Ibs) (in-lbs) (psi) (psi) (psi) 3020 192298 631638 5186 6252 11439 3030 191926 108653 5176 1075 6252 3040 200560 326835 5409 3235 8644 3050 200490 411556 5407 4074 9481 3060 180263 739254 4862 7317 12179 3090 180623 635570 4871 6291 11163 3120 180623 483086 4871 4782 9653 3150 180623 705757 4871 6986 11857 3160 198000 399180 5340 3951 9291 3180 198030 150267 5341 1487 6828 3190 180623 302359 4871 2993 7864 3230 180623 1300822 4871 12876 17748 3240 196572 1020473 5302 10101 15403 3280 196889 1257967 5310 12452 17762 3285 180623 1552056 4871 15363 20234 3300 180623 1406793 4871 13925 18796 Febwary 2003 Loads for Loads Mechanics Analysis Fracture Mechanics for Fracture Analysis February 2003

4-6 Table 4-3: Summary of Callaway Nuclear Power Plant Faulted Loads and Stresses for Residual Heat Removal Line Loop 1 Location Axial Force Moment Axial Stress Moment Total Stress Node (Ibs) (in-Ibs) (psi) (psi) (psi) 3020 225671 1507098 6086 14918 21004 3030 225238 401425 6075 3973 10048 3040 211708 499102 5710 4940 10650 3050 211667 622861 5709 6165 11874 3060 246376 1204405 6645 11922 18567 3090 243579 948680 6569 9390 15960 3120 235313 636467 6346 6300 12646 3150 233470 882540 6297 8736 15032 3160 210206 487527 5669 4826 10495 3180 210206 262627 5669 2600 8269 3190 231833 452128 6253 4475 10728 3230 226147 1454034 6099 14393 20492 3240 209645 1183939 5654 11719 17373 3280 210139 1410188 5667 13959 19626 3285 228015 1751544 6150 17338 23487 3300 228023 1586801 6150 15707 21857 February 2003 Loads Mechanics Analysis Fracture Mechanics Loads for Fracture Analysis February 2003

4-7 I Table 4-4: Summary of Callaway Nuclear Power Plant Piping Geometry and Normal Operating Condition for Residual Heat Removal Line Loop 4 Outer Minimum Normal Operating Material Diameter Wall Pressure Temperature Type (in) Thickness (psig) (OF)

(in)

SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 619 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 SA376/SA312 TP304 or SA403 WP304 12.750 1.005 2235 70 ch nisF acur Mdsfo Lo nayss eb ua y 00 Loads for Fracture Mechanics Analysis February 2003

4-8 Table 4-5: Summary of Callaway Nuclear Power Plant Normal Loads and Stresses for Residual Heat Removal Line Loop 4 Weld Axial Moment Total Stress Location Force Moment Axial Stress Stress Node (Ibs) (In-lbs ) (psi) (psi) 3020 188245 1086916 5077 10759 15836 3030 188176 1010275 5075 10000 15075 3040 182761 943054 4929 9335 14264 3060 182429 1024057 4920 10137 15057 3080 198995 740120 5367 7326 12693 3120 199422 252751 5378 2502 7880 3140 195789 278387 5280 2756 8036 3180 195789 593462 5280 5874 11155 3240 195789 1461963 5280 14471 19752 3290 195789 916018 5280 9067 14348 3320 195789 711832 5280 7046 12326 3340 196080 698683 5288 6916 12204 3400 196080 415925 5288 4117 9405 3420 205577 286575 5544 2837 8381 3500 205725 342087 5548 3386 8935 3520 196080 502244 5288 4971 10260 3580 196080 605231 5288 5991 11279 Febwary 2003 Loads for Mechanics Analysis Fracture Mechanics for Fracture Analysis February 2003

4-9 Table 4-6: Summary of Callaway Nuclear Power Plant Faulted Loads and Stresses for Residual Heat Removal Line Loop 4 Weld Axial Force Moment Axial Stress Moment Total Stress Node (Ibs) (in-Ibs) (psi) (psi) 3020 232496 1826562 6270 18080 24351 3030 232422 1640611 6268 16240 22508 3040 230352 1375521 6213 13616 19828 3060 230064 1251847 6205 12391 18596 3080 225199 1180885 6074 11689 17763 3120 222539 617502 6002 6112 12114 3140 228218 676284 6155 6694 12849 3180 227541 922884 6137 9135 15272 3240 217837 1720972 5875 17035 22910 3290 216718 1106202 5845 10950 16795 3320 213289 780396 5752 7725 13477 3340 211071 787249 5693 7793 13485 3400 210422 515809 5675 5106 10781 3420 207383 375105 5593 3713 9306 3500 210476 558188 5677 5525 11202 3520 212149 706602 5722 6994 12716 3580 212194 766821 5723 7590 13313 February 2003 Loads for Loads for Fracture Mechanics Analysis Fracture Mechanics Analysis February 2003

4-10 Critical Weld Location 3285 3240 3230 3160 3190 3150 3180 Hot Leg Loop I 120 3090 3060 Figure 4-1 Governing Weld Location for RHR Line Loop 1 February 2003 Loads for Fracture Mechanics Analysis Fracture Mechanics Analysis February 2003

4-11 Hot Leg Loop 4

'3500 Critical Weld '3030 Location 3060 3400 3080 3240 3340 3120 3320 3290 3140 Figure 4-2 Governing Weld Location for RHR Line Loop 4 f, Fracture Lo d for Meh nc An.y.

_r ct r Mechanics Analysis Febr. ry.2003 Loads February 2003

5-1 5 FRACTURE MECHANICS EVALUATION 5.1 GLOBAL FAILURE MECHANISM Determination of the conditions, which lead, to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the [ ]aXc~e method based on traditional plastic limit load concepts, but accounting for [ ]ace and taking into account the presence of a flaw. The flawed component is predicted to fail when the remaining net section reaches a stress level at which a plastic hinge is formed. The stress level at which this occurs is termed as the flow stress. [

]ace This methodology has been shown to be applicable to ductile piping through a large number of experiments and is used here to predict the critical flaw size in the RHR lines.

The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 5-1) when loads are applied. The detailed development is provided in Appendix A for a through-wall circumferential flaw in a pipe section with internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:

[ ]a,c,e (5-1) where:

]ace (5-2)

The analytical model described above accurately accounts for the internal pressure as well as the imposed axial force as they affect the limit moment. Good agreement was found between valatio echaics Fracure Febuary200 Fracture Mechanics Evaluation February 2003

5-2 the analytical predictions and the experimental results (Reference 5-1). The flaw stability evaluations using this analytical model are presented in Section 5.3.

5.2 LEAK RATE PREDICTIONS Fracture mechanics analysis shows that postulated through-wall cracks in the RHR lines would remain stable and would not cause a gross failure of this component. However, if such a through-wall crack did exist, it would be desirable to detect the leakage such that the plant could be brought to a safe shutdown condition. The purpose of this section is to discuss the method, which will be used to predict the flow through such a postulated crack and present the leak rate calculation results for through-wall circumferential cracks.

5.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure (causing choking) is taken into account. For long channels where the ratio of the channel length, L, to hydraulic diameter, DH, (L/DH) is greater than [ ]ace, both [ ja,c,e must be considered. In this situation, the flow can be described as being single-phase through the channel until the local pressure equals the saturation pressure of the fluid. At this point, the flow begins to flash and choking occurs. Pressure losses due to momentum changes will dominate for [ ]ace. However, for large L/DH values, the friction pressure drop will become important and must be considered along with the momentum losses due to flashing.

5.2.2 Calculation Method In using the [

la,c,e.

The flow rate through a crack was calculated in the following manner. Figure 5-2 from Reference 5-2 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass flow, the [

]a~ce was found from Figure 5-3 taken from Reference 5-2.

For all cases considered, since [ ] Therefore, this method will yield the two-phase pressure drop due to momentum effects (AP2,) as illustrated in Figure 5-4. Now using the assumed flow rate, G.the frictional pressure drop can be calculated using apf ]ac.e (5-3)

FractureMechaics Evaluatio February 200 Fracture Mechanics Evaluation February 2003

5-3 where the friction factor f was determined using the [ ]a~c~e The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ ]a~c eRMS.

The frictional pressure drop using Equation 5-3 was then calculated for the assumed flow and added to the [ ]axcwe to obtain the total pressure drop from the system under consideration to the atmosphere. Thus, Absolute Pressure - 14.7 = [ ]ace (5-4) for a given assumed flow G. If the right-hand side of Equation 5-4 does not agree with the pressure difference between the piping under consideration and the atmosphere, then the procedure is repeated until Equation 5-4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack.

For the locations at the lower temperatures, the leak rate is calculated by using the simple orifice type flow formula given by [

(5-5)

]axce 5.2.3 Leak Rate Calculations Leak rate calculations were performed as a function of postulated through-wall crack length for the governing locations previously identified. The crack opening area was estimated using the method of Reference 5-4 and the leak rates were calculated using the calculation methods described above. The leak rates were calculated using the normal operating loads at the governing locations identified in Section 4. Average yield strength properties from Table 3-2 were used in the leak rate calculations. The crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for the governing locations at the Callaway Nuclear Power Plant RHR lines are shown in Table 5-1.

Fracture Mechanics Evaluation February 2003

5-4 The Callaway Nuclear Power Plant has an RCS pressure boundary leak detection system which is consistent with the guidelines of Regulatory Guide 1.45 for detecting leakage of 1 gpm in one hour.

5.3 STABILITY EVALUATION A typical segment of the pipe under maximum loads of axial force F and moment M is schematically illustrated in Figure 5-5. In order to calculate the critical flaw size, plots of the limit moment versus crack length are generated as shown in Figures 5-6 to 5-7. The critical flaw size corresponds to the intersection of this curve and the maximum load line. The critical flaw size is calculated using the lower bound base metal tensile properties tabulated in Table 3-2.

The weld process types for all the shop welds and field welds are Gas Tungsten Arc Weld (GTAW) and Shielded Metal Arc Weld (SMAW) combination or GTAW. Although the weld process at the critical weld locations identified in Section 4.5 is GTAW, in order to envelop all the weld process types in the stability evaluation, SMAW weld process type is conservatively assumed. Therefore, the LZ factor correction for the SMAW was applied (Reference 5-5) as follows:

Z = 1.15 [1 + 0.013 (OD - 4)] (for SMAW) (5-6) where OD is the outer diameter in inches. Substituting OD = 12.75 inches, the Z factor was calculated to be 1.28 for SMAW. The Z correction factor for GTAW is 1.0. The applied loads were increased by the Z factors for SMAW and the plots of limit load versus crack length were generated as shown in Figures 5-6 to 5-7 for the critical locations. Table 5-2 shows the summary of critical flaw sizes.

5.4 REFERENCES

5-1 Kanninen, M. F. et al., "Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976.

5-2 [

]a,c,e 5-3 Crane, D.P., "Handbook of Hydraulic Resistance Coefficient."

5-4 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe,"

Section lI-1, NUREG/CR-3464, September 1983.

5-5 Standard Review Plan; Public Comment Solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal RegisterNol. 52, No. 167/Friday, August 28, 1987/Notices, pp. 32626-32633.

Fctr Mehncvlaioeray20 Fracture Mechanics Evaluation February 2003

5-5 Table 5-1: Leakage Flaw Sizes Node Point Temperature Crack Length (in.)

(OF) (for 10 gpm leakage) 3020 619 3.41 3285 1 70 2.97 Table 5-2: Summary of Critical Flaw Sizes Node Point

__________+/-l 3020 Temperature

~(OF) 619 I(in)

Critical Flaw Size 10.25 3285 70 12.99 February 2003 Mechanics Evaluation Fracture Mechanics Evaluation February 2003

5-6 Figure 5-1 Fully Plastic Stress Distribution Evaluation February 2003 Fracture Mechanics Fracture Mechanics Evaluation February 2003

5-7 a,c, e C

I U:

C%

-J 4C STAGNATION ENTHALPY (102 ua/lW Figure 5-2 Analytical Predications of Critical Flow Rates of Steam-Water Mixtures R;acure, Muchanics Evaluation February 2003

5-8 a.ce -

a 0

LU C.,

LENGTH/DIAMETER RATIO (LID Figure 5-3 [ ja,c,e Pressure Ratio as a Function of LJD Fracture Mechanics Evaluation February 2003

5-9 a,c,e ace Figure 5-4 Idealized Pressure Drop Profile through a Postulated Crack Fracture Mechanics Evaluation February 2003

5-10 a

0

  • S
  • S
  • S
  • I
  • S
  • S
  • S
  • I
  • S
  • I
  • S
  • S
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  • S
  • I a a
  • S
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  • U L -4
  • S
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  • a I S

-I S 5

  • S
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  • S S S
  • S
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  • 4
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  • S S S
  • S I-Figure 5-5 Loads acting on the Model at the Governing Locations Fracture Mechanics Evaluation February 2003

5-11 a,c,e OD = 12.75 in ay,= 22.384 ksi F = 232.496 kips t = 1.005 in a. = 68.244 ksi M =1826.562 in-kips SA376/SA312 TP304 or SA403 WP304 with SMAW weld Figure 5-6 Critical Flaw Size Prediction for Node 3020 (RHR Line Loop 4)

February 2003 Fracture Mechanics Evaluation Fracture Mechanics Evaluation February 2003

5-12 a~c,e OD = 12.75 in ay = 36.800 ksi F = 228.015 kips t = 1.005 in o.= 80.730 ksi M = 1751.544 in-kips SA376/SA312 TP304 or SA403 WP304 with SMAW weld Figure 5-7 Critical Flaw Size Prediction for Node 3285 (RHR Line Loop 1)

Evlato February 2003 Fracture Mehnc Frctr Mechanics Evaluation February 2003

6-1 6 ASSESSMENT OF FATIGUE CRACK GROWTH

6.1 INTRODUCTION

The fatigue crack growth of the Callaway Nuclear Power Plant RHR lines was determined by comparison with a generic fatigue crack growth analysis of a similar piping system. The details of the generic fatigue crack growth analysis are presented below. By comparing all parameters critical to the fatigue crack growth analysis between Callaway and the generic analysis, it was concluded that the generic analysis would adequately cover the fatigue crack growth of the Callaway Nuclear Power Plant RHR lines.

Due to similarities in Westinghouse PWR designs, it was possible to perform a representative fatigue crack growth calculation which would be applicable to the Callaway Nuclear Power Plant. A comparison was made of the number of cycles, material, geometry, and types of discontinuities.

6.2 CRITICAL LOCATION FOR FATIGUE CRACK GROWTH ANALYSIS The weld locations at the RCL hot leg nozzles to RHR lines (see Figures 3-1 and 3-2) were determined to be the most critical location for the fatigue crack growth evaluation. The nozzle configuration and weld location is shown in Figure 6-1. The geometry of the pipe was identical between the Callaway Nuclear Power Plant and the generic model (12" Schedule 140). Both analyses used austenitic stainless steel at the critical location.

6.3 DESIGN TRANSIENTS The transient conditions selected for this evaluation are based on conservative estimates of the magnitude and the frequency of the temperature fluctuations documented in various operating plant reports. These are representative of the conditions which are considered to occur during plant operation. The normal operating and upset thermal transients, in accordance with the design specification and the applicable system design criteria document, were considered for this evaluation. Out of these, 20 transients were used in the fatigue crack growth analysis and are listed in Table 6-1. There are some differences between the generic transients used in the fatigue crack growth evaluation and the Callaway Nuclear Power Plant transients but these differences would have insignificant impact on the fatigue crack growth results.

6.4 STRESS ANALYSIS A thermal transient stress analysis was performed for a typical plant similar to the Callaway Nuclear Power Plant to obtain the through-wall stress profiles for use in the fatigue crack growth analysis. The generic RHR line design transients described in Section 6.3 were used.

A simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients. In this method, a 1-D computer program was used to perform the thermal analysis to determine the through wall temperature February 2003 Assessment of Assessment Crack Growth Fatigue Crack of Fatigue Growth February 2003

6-2 gradients as a function of time. The inside surface stress was calculated by using an equation, which is similar to the transient portion of ASME Section III NB 3600, Equation (11). The effect of discontinuity was included in the analysis by performing a separate 1-D thermal analysis for the pipe and nozzle. The maximum and minimum inside surface stresses were then obtained by searching the inside surface stress values calculated for each time step of the transient solution. The outside surface stresses corresponding to the maximum and minimum inside surface stresses were then calculated by a similar method. The maximum and minimum linear through wall stress distribution for each thermal transient was obtained by joining the corresponding inside and outside surface stresses by a straight line. These two stress profiles are called the maximum and minimum through wall stress distributions respectively, for convenience. The stresses due to the generic pressure and the generic moment loading were then superimposed on the through wall cyclical stresses to obtain the total maximum and minimum stress profile for each transient.

6.5 OBE LOADS The stresses due to OBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contribute significantly to crack growth due to the small number of cycles.

6.6 TOTAL STRESS FOR FATIGUE CRACK GROWTH The total through wall stress at a section was obtained by superimposing the generic pressure stress and the generic moment stresses on the thermal transient stresses. Thus, the total stress for fatigue crack growth at any point is given by the following equation:

Total Stress Stress due to For Fatigue Internal + Moment(DW + Thermal Crack Pressure + Thermal Transient Stress Growth Expansion) 6.7 FATIGUE CRACK GROWTH ANALYSIS The fatigue crack growth analysis was performed to determine the effect of the design thermal transients tabulated in Table 6-1. The analysis was performed for the critical cross-section identified in Figure 6-1. A range of crack depths was postulated, and each was subjected to the transients in Table 6-1, which included pressure and moment loads.

6.7.1 Analysis Procedure The fatigue crack growth analyses presented herein were conducted in the same manner as suggested by Section Xl, Appendix A of the ASME Boiler and Pressure Vessel Code (Reference 6-1). The analysis procedure involves assuming an initial flaw exists at some point and predicting the growth of that flaw due to an imposed series of transient stresses. The February 2003 Assessment of Assessment Crack Growth Fatigue Crack of Fatigue Growth February 2003

6-3 growth of a crack per loading cycle is dependent on the range of applied stress intensity factor, AK,, by the following:

da

= COAKI (6-1) where "CO" and the exponent "n" are material properties, and AK, is defined later. For inert environments these material properties are constants, but for some water environments they are dependent on the level of mean stress present during the cycle. This can be accounted for by adjusting the value of "CO" by a function of the ratio of minimum to maximum stress for any given transient, as will be discussed later. Fatigue crack growth properties of stainless steel in a pressurized water environment have been used in the analysis.

The input required for a fatigue crack growth analysis is basically the information necessary to calculate the parameter AK,, which depends on crack and structure geometry and the range of applied stresses in the area where the crack exists. Once AK, is calculated, the growth due to that particular cycle can be calculated by Equation (6-1). This increment of growth is then added to the original crack size, the AK, adjusted, and the analysis proceeds to the next transient. The procedure is continued in this manner until all the transients have been analyzed.

The applied stresses at the flaw locations are resolved into membrane and bending stresses with respect to the wall thickness. Pressure, thermal, and discontinuity stresses are considered in the determination of the Klfactors.

The stress intensity factor at the point of maximum depth is calculated from the membrane and bending stresses using the following equation taken from the ASME Code (Reference 6-1):

K, =+/ [GmMm +ObMb ]

where: am, ab = Membrane and Bending Stress, respectively a = Minor Semi-Axis (flaw depth)

Q = Flaw Shape Parameter Including A Plastic Zone Correction Factor for Plane Strain Condition Q = [ 1 2 -0.212 (aa/as)2 ]

01 = f:I2[l (b 2jad)cos2P dI2 GcS = Yield Strength of the Material February 2003 Assessment of Crack Growth Fatigue Crack of Fatigue Growth February 2003

6-4 a = Or + Cyb b = Major Semi-Axis (Flaw Length/2)

  • = Parametric Angle of the Ellipse Mm = Correction Factor for Membrane Stress Mb = Correction Factor for Bending Stress The appropriate values of Mm and Mb as a function of crack geometry can be found in Reference 6-1. The range of stress intensity factor (AK,) for fluctuation of applied stress is determined by first finding the maximum and minimum stress intensity factor (K,max, K, min) 1 a - K, during a given transient and then calculating the range of stress intensity factor (AK, = Ke min). At times Kemin may go below zero, in these cases, K,min is set equal to zero before AK, is determined.

Calculation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law determined from consideration of the available data for stainless steel in a pressurized water environment. This law allows for the effect of mean stress or R ratio (K, min/K 1 max) on the growth rates.

The reference crack growth law used for the stainless steel RHR pipe system was taken from that developed by the Metal Properties Council - Pressure Vessel Research Committee Task Force In Crack Propagation Technology. The reference curve has the equation:

[ (6-2)

Iacse This equation appears in Appendix C of ASME Section Xl for air environments and its basis is provided in Reference 6-2, and shown in Figure 6-2. For water environments, an environmental factor of [ Iace was used, based on the crack growth tests in PWR environments reported in Reference 6-3.

Assessment of Fatigue Crack Growth February 2003

6-5 6.8 RESULTS Fatigue crack growth analyses were carried out at the critical cross-section. Analysis was completed for a range of postulated flaw sizes oriented circumferentially, and the results are presented in Table 6-2. The postulated flaws are assumed to have an aspect ratio of six to one.

Even for the largest postulated flaw of 0.35 inch, which is about 35 percent of the wall thickness, the result projects that the flaw growth through the wall will not occur during the 40 year design life of the plant. Therefore, fatigue crack growth should not be a concern for the Callaway Nuclear Power Plant RHR Lines.

6.9 REFERENCES

6-1 ASME Boiler and Pressure Vessel Code Section Xl, 2001 Edition, "Rules for Inservice Inspection of Nuclear Power Plant Components" 6-2 James, L. A., and Jones, D. R, "Fatigue Crack Growth Correlations for Austenitic Stainless Steel in Air," in Predictive Capabilities in Environmentally Assisted Cracking."

ASME publication PVP-99, Dec. 1985.

6-3 Bamford, W. H., "Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment," Trans ASME, Journal of Pressure Vessel Technology, Feb.

1979.Engineering Development Labs Report HEDL-TME-76-43, May 1976.

rowth of Fatige Crack Assessmen Febuary 200 Assessment of Fatigue Crack Growth February 2003

6-6 Table 6-1 : Design Transients Considered for Fatigue Crack Growth Evaluation.

Trans. No. Description No. of Occurrences 1 Unit Loading 13,200 2 Unit Unloading 13,200 3 Step Load Increase 2,000 4 Step Load Decrease 2,000 5 Large Step Load Decrease with Steam Dump 200 6 Feedwater Cycling 2000 7 Unit Loading Between 0 and 15% Power 500 8 Unit Unloading Between 0 and 15% Power 500 9 Loss of Load 80 10 Loss of Power 40 11 Partial Loss of Flow-Dead Loop 80 12 Partial Loss of Flow-Active Loop 80 13 Reactor Trip with no Inadvertent Cooldown 230 14 Reactor Trip with Cooldown; No Safety Injection 160 15 Reactor Trip with Cooldown Actuating Safety Injection 10 16 Inadvertent RCS Depressurization 20 17 Control Rod Drop 80 18 Inadvertent Safety Injection 60 19 Turbine Roll Test 20 20 Steady-State and Random Fluctuations 3.2 x 106 February 2003 Assessment of Crack Growth Fatigue Crack of Fatigue Growth February 2003

6-7 Table 6-2: RHR Lines Fatigue Crack Growth Results Initial Crack Depth (in) After a,c,e

_ _ _ _ _ _ _ _ _ I l _ _ _ I _ _ _

4 4 4 I I 4 4. I I 4 4. I Assessment of Fatigue Crack Growth February 2003

6-8 Critical Section for RCL Hot Leg Fatigue Crack Growth Nozzle I I Pipe I

5*37n fl 6.375" R R is the pipe radius Figure 6-1 Schematic of RHR Line at RCL Hot Leg Nozzle Weld Location Assessment of Fatigue Crack Growth Assessme nt of F atig ue Crack Gro wth F ebruary 2 00 3

6-9 3.0X Ur = -- 2Z 7 Solid _lie for 70_

Mulled lines for 5WF/

_ _ ___ _ 7  :~~~~~~7; c CE For other R ratios and

- temperature. see C -3210(b) 2.0 10 0 _ 101 102

__Iks, fIl _

Figure 6-2 Reference Crack Growth Curves for Stainless Steel in Air Environments Assessment of Fatigue Crack Growth Assessment of FatigueCrackGrowth February 2003

7-1 7 ASSESSMENT OF MARGINS In the preceding sections, the leak rates calculations, fracture mechanics analysis and fatigue crack growth assessment was performed.

The results of the leak rates of Section 5.2 and the corresponding stability results of Section 5.3 are used in performing the assessment of margins. Margins are shown in Table 7-1.

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 or more exists between the critical flaw and the flaw having a leak rate of 10 gpm (the leakage flaw).
2. Leak Rate - A margin of 10 exists between the calculated leak rate from the leakage flaw and the leak detection capability of 1 gpm.
3. Loads - At the critical locations, the leakage flaw was shown to be stable using the faulted loads obtained by the absolute sum method (i.e., a flaw twice the leakage flaw size is shown to be stable; hence the leakage flaw size is stable). Therefore a margin on loads of 1.0 (see Section 4.2 for explanation) using the absolute summation of faulted load combinations is satisfied.

All the LBB recommended margins are satisfied.

In this evaluation, the Leak-Before-Break methodology is applied conservatively. The conservatism used in the evaluation is summarized in Table 7-2.

Margins of Margins February 2003 Assessment of Assessment February 2003

7-2 Table 7-1: Leakage Flaw Sizes, Critical Flaw Sizes and Margins Node Critical Flaw Leakage Flaw Margin

_ _ ~~~~Size (in)* Size (in) _

3020 10.25 3.41 3.01 3285 12.99 2.97 4.37 Table 7-2: LBB Conservatism Factor of 10 on Leak Rate Factor of 2 on Leakage Flaw Algebraic Sum of Loads for Leakage Absolute Sum of Loads for Stability Average Material Properties for Leakage Minimum Material Properties for Stability of Margins Assessment Februaty 2003_

Assessment of Margins February 2003

8-1 8 CONCLUSIONS This report justifies the elimination of RHR line pipe breaks as the structural design basis for the Callaway Nuclear Power Plant as follows:

a. Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.
b. Water hammer should not occur in the RCS piping (primary loop and the attached class 1 auxiliary lines) because of system design, testing, and operational considerations.
c. The effects of low and high cycle fatigue on the integrity of the RHR lines were evaluated and shown acceptable.
d. Ample margin exists between the leak rate of small stable flaws and the capability of Callaway Nuclear Power Plant reactor coolant system pressure boundary leakage detection system.
e. Ample margin exists between the small stable flaw sizes of item (d) and the critical flaw size.

The postulated reference flaw will be stable because of the ample margins in items (d) and (e) and will leak at a detectable rate which will assure a safe plant shutdown.

Based on the above, it is concluded that RHR line breaks should not be considered in the structural design basis of the Callaway Nuclear Power Plant.

ConclusionsFeray20 Conclusions February 2003

A-1 APPENDIX A - LIMIT MOMENT

[

Ja,c,e Appendix A - Limit Moment February 2003

A-2

'I I Figure A-1 Pipe with A Through-Wall Crack In Bending Appendix A - Limit Moment Apn.ALmtMmn er ----

February 2003

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