ML031610768

Rev 0 to WCAP-15837-NP, Technical Justification for Eliminating Large Primary Loop Pipe Rupture as Structural Design Basis for R.E. Ginna Nuclear Power Plant for License Renewal Program

ML031610768
Person / Time
Site:	Ginna
Issue date:	05/31/2003
From:	Bhowmick D, Ching Ng Westinghouse
To:	Office of Nuclear Reactor Regulation
References
FOIA/PA-2005-0108 WCAP-15837-NP, Rev 0
Download: ML031610768 (60)

Text

Westinghouse Non-Proprietary aass 3 WCAP-15837-NP Revision 0 Technical Justification for Eliminating Large Primary Loop Pipe Rupture as the Structural Design Basis for the R. E. Ginna Nuclear Power Plant for the License Renewal Program Westinghouse May 2003

WESTINGHOUSE NON - PROPRIETARY CLASS 3 WCAP-15837 - NP Revision 0 Technical Justification for Eliminating Large Primary Loop Pipe Rupture as the Structural Design Basis for the R. E.

Ginna Nuclear Power Plant for the License Renewal Program D. C. Bhowmick C.K.Ng May 2003 Verifier.a, J. F. Petsche Approved:

A*7 '

S.

wamy, M ager Structural Mechanics Technology Westinghouse Electric Company LLC P.O. Box 355 Pittsburgh, PA 15230-0355 02003 Westinghouse Electric Company LLC All rights Reserved o:\\RGE.doc:1 b-052303 May2003 May2003 o:\\RGE.doc:l b-052303

TABLE OF CONTENTS EXECUTIVE

SUMMARY

vii 1.0 Introduction.........................................................

1-1 1.1 Purpose.........................................................

1-1 1.2 Background Information.........................................................

1-1 1.3 Scope and Objectives.........................................................

1-2 1.4 References..........................................................

1-3 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM..................... 2-1 2.1 Stress Corrosion Cracking.......................................................... 2-1 2.2 Water Hammer.........................................................

2-2 2.3 Low Cycle and High Cycle Fatigue.........................................................

2-3 2.4 References.........................................................

2-3 3.0 PIPE GEOMETRY AND LOADING............

............................................. 3-1 3.1 Introduction to Methodology..........................................................

3-1 3.2 Calculation of Loads and Stresses.........................................................

3-1 3.3 Loads for Leak Rate Evaluation.

......................................................... 32 3.4 Load Combination for Crack Stability Analyses......................................................... 3-3 3.5 References.........................................................

3-3 4.0 MATERIAL CHARACTERIZATION.......................................................... 4-1 4.1 Primary Loop Pipe and Fitings Materials.................................

.4-1 4.2 Tensile Properties.................................

4-1 4.3 Fracture Toughness Properties.................................

4-2 o:\\RGE.doc:1 b-05123/03 May 2003 o:\\RGE.doc:1b 05303 May 2003

.1i, 4.4 References............................................

4-5 5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA.

5-1 5.1 Critical Locations.5-1 5.2 Fracture Criteria.......................

5-1 6.0 LEAK RATE PREDICTIONS.........

6-1 6.1 Introduction.6-1 6.2 General Considerations.................................

6-1 6.3 Calculation Method.................................

6-1 6.4 Leak Rate Calculations.................................

6-2 6.5 References.................................

62 7.0 FRACTURE MECHANICS EVALUATION.................................

7-1 7.1 Local Failure Mechanism.................................

7-1 7.2 Global Failure Mechanism.................................

7-1 7.3 Results of Crack Stability Evaluation

............................... 7-3 7.4 References..................................

7-5 8.0 FATIGUE CRACK GROWTH ANALYSIS.................................

8-1 8.1 References.................................

8-2

9.0 ASSESSMENT

OF MARGINS.................................

9-1 10.0 CONCWSIONS.................................

10-1 APPENDIX A..................................

A-1 LIMIT MOMENT..................................

A-1 o:\\RGE.doc:lb-052303 May2003

V LIST OF TABLES Table 3-1 Dimensions, Normal Loads and Normal Stresses for Ginna Station................. 3-4 Table 3-2 Faulted Loads and Stresses for Ginna Station................................................ 3-5 Table 4-1 Measured Tensile Properties (psi) for the Ginna Station Primary Loop Piping (Material A376 TP316).46 Table 4-2 Measured Tensile Properties (psi) for the Ginna Station Primary Loop Elbow Fittings (Material A351 CF8M).4-7 Table 4-3 Mechanical Properties for Ginna Station Materials at Operating Temperatures 4-8 Table 4-4 Chemistry and Fracture Toughness Properties of the Material Heats of Ginna Station.4-9 Table 4-5 Fracture Toughness Properties for Ginna Station Primary Loops for Leak-Before-Break Evaluation at Critical Locations..........................

4-10 Table 6-1 Flaw Sizes Yielding a Leak Rate of 2.5 gpm at the Governing Locations......... 6-3 Table 7-1 Stability Results for Ginna Station Based on Elastic J-lntegral Evaluations...... 7-6 Table 7-2 Stability Results for Ginna Station Based on Umit Load................................... 7-6 Table 8-1 Summary of Reactor Vessel Transients (60 Years)

......................... 8-3 Table 8-2 Fatigue Crack Growth at [

]J8

( 60 years)................. 8-4 Table 9-1 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Ginna Station.......... 9-2 o:\\RGE.doc:1 b-05/23/03 May 2003 o:\\RGE.doc:1 b3/03 May 2003

vi UST OF FIGURES Figure 3-1 Hot Leg Coolant Pipe.................................................

3-6 Figure 3-2 Schematic Diagram of Ginna Station Primary Loop Showing Weld Locations.. 3-7 Figure 4-1 Pre-Service J vs. Aa for SA351 CF8M Cast Stainless Steel at 600 F......

....... 4-13 Figure 6-1 Analytical Predictions of Critical Rates of Steam-Water Mixtures..................... 6-4 Figure 6-2

[

]a.Ce Pressure Ratio as a Function of UD..........

.................. 6-5 Figure 6-3 Idealized Pressure Drop Profile Through a Postulated Crack........................... 6-6 Figure 7-1

[

]a.c Stress Distribution..............

................................... 7-7 Figure 7-2 Critical Flaw Size Prediction - Hot Leg at Location 1....................

................... 7-8 Figure 7-3 Critical Flaw Size Prediction - Hot Leg at Location 3....................

................... 7-9 Figure 7-4 Critical Flaw Size Prediction - Cold Leg at Location 12.................................. 7-10 Figure 8-1 Typical Cross-Section of [

]ace................................ 8-5 Figure 8-2 Reference Fatigue Crack Growth Curves for Carbon and Low Alloy Ferritic Steels

......................................................................................................................... 8-6 Figure A-1 Pipe with a Through-Wall Crack in Bending.................................................

A-2 o:\\RGE.doc:1 b-052303 May2003 o:ARGE.doc:1b 052303 May 2003

vii EXECUTIVE

SUMMARY

The original structural design basis of the Reactor Coolant System (RCS) for the Rochester Gas and Electric Corporation Ginna Nuclear Power Plant required consideration of dynamic effects resulting from pipe break and that protective measures for such breaks be incorporated into the design. Subsequent to the original Ginna design, additional concem of asymmetric blowdown loads was raised as described in Unresolved Safety Issue A-2 (Asymmetric Blowdown Loads on the Reactor Coolant System) and Generic Letter 84-04 (Reference 1-2).

However, research by the Nuclear Regulatory Commission (NRC) and industry coupled with operating experience determined that safety could be negatively impacted by placement of pipe whip restraints on certain systems.

As a result, 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 technology and material toughness. Generic analyses by Westinghouse for the application of LBB for specific plants were documented in response to the Unresolved Safety Issue A-2 in the NRC letter dated May 6, 1986 (Reference 1-3).

This present WCAP report documents the plant specific geometry, loading, and material properties used in the fracture mechanics evaluation. It also Includes the temperature, pressure and loadings generated as a result of the Ginna Nuclear Power Plant License Renewal Program. Mechanical properties were determined at operating temperatures. Since the piping system includes cast stainless steel fittings, the end of life (60 year) fracture toughness considering thermal aging was determined for each heat of material.

Based on loading, pipe geometry and fracture toughness considerations, enveloping critical locations were determined at which Leak-Before-Break crack stability evaluations were made.

Through-wall flaw sizes were postulated which would cause a leak at a rate of ten times the leakage detection system capability of the plant.

Large margins for such flaw sizes were demonstrated against flaw instability.

Finally, using thermal transient stresses and cycles, fatigue crack growth for 60 years was shown to be acceptable for the primary loops. All the recommended LBB margins (margin on leak rate, margin on flaw size and margin on loads) are satisfied.

It Is therefore concluded that the dynamic effects of RCS primary loop pipe breaks need not be considered in the structural design basis of the Ginna Nuclear Power Plant for the Ucense Renewal Program.

o:\\RGE.doc:1 b-05/23/03 May 2003 o:\\RGE.doc:1 b 05/23103 May 2003

1-1

1.0 INTRODUCTION

1.1 PURPOSE This report applies to the Ginna Station Reactor Coolant System (RCS) primary loop piping. It is intended to demonstrate that for the specific parameters of the Ginna Station, RCS primary loop pipe breaks need not be considered in the structural design basis for the 60 year plant life.

The Nuclear Regulatory Commission (NRC) (Reference 1-2) has accepted the approach taken.

1.2 BACKGROUND

INFORMATION Westinghouse has performed considerable testing and analysis to demonstrate that RCS primary loop pipe breaks can be eliminated from the structural design basis of all Westinghouse plants. The concept of eliminating pipe breaks in the RCS primary loop was first presented to the NRC in 1978 in WCAP-9283 (Reference 1-4). That topical report employed a deterministic fracture mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks from the structural design basis. That approach was then used as a means of addressing Generic Issue A-2 and Asymmetric LOCA Loads.

Westinghouse performed additional testing and analysis to justify the elimination of RCS primary loop pipe breaks from the structural design basis. This material was provided to the NRC along with Letter Report NS-EPR-2519 (Reference 1-5).

The NRC funded research through Lawrence Livermore National Laboratory (LLNL) to address this 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 previously performed by Westinghouse and the research performed by LLNL applied to all Westinghouse plants (References 1-6 and 1-7).

The results from the LLNL study were released at a March 28, 1983, ACRS Subcommittee meeting. These studies, which are applicable to all Westinghouse plants east of the Rocky Mountains, determined the mean probability of a direct LOCA (RCS primary loop pipe break) to be 4.4 x 1012 per reactor year and the mean probability of an indirect LOCA to be 10 7 per reactor year. Thus, the results previously obtained by Westinghouse (Reference 1-4) were confirmed by an independent NRC research study.

Based on the studies by Westinghouse, LLNL, the ACRS, and the AIF, NRC completed a safety review of the Westinghouse reports submitted to address asymmetric blowdown loads that result from a number of discrete break locations on the Pressurized Water Reactor (PWR) primary systems. The NRC Staff evaluation (Reference 1-2) concludes that an acceptable technical basis has been provided so that asymmetric blowdown loads need not be considered for those plants that can demonstrate the applicability of the modeling and conclusions contained in the Westinghouse response or can provide an equivalent fracture mechanics demonstration of the primary coolant loop integrity. In a more formal recognition of Leak-Before-Break (LBB) methodology applicability for PWRs, the NRC appropriately modified Introducton May 2003 o:.RGE.doc:1b-052303

1-2 10CFR 50, General Design Criterion 4, "Requirements for Protection Against Dynamic Effects for Postulated Pipe Rupture" (Reference 1-8).

1.3 SCOPE AND OBJECTIVES The general purpose of this investigation is to demonstrate Leak-Before-Break for the primary loops in the Ginna Station on a plant specific basis for the 60 year plant life.

The recommendations and criteria proposed in Reference 1-9 are used in this evaluation. These criteria and resulting steps of the evaluation procedure can be briefly summarized as follows:

1.

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

2.

Identify the materials and the associated material properties.

3.

Postulate a surface flaw. 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.

A margin of 10 is demonstrated between the calculated leak rate and the leak detection capability.

5.

Using faulted loads, demonstrate that there is a margin of 2 between the leakage flaw size and the critical flaw size.

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 actually used In the plant, provide the properties including toughness and 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 the Ginna Station which is consistent with the NRC position for exemption from consideration of dynamic effects.

The computer codes that are used in this evaluation for leak rate and fracture mechanics calculations have been validated and used for all the LBB applications by Westinghouse.

Introduction May 2003 o:UIGE.doc:1b-052303

1-3

1.4 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 USNRC Generic Letter 84-04, Subject "Safety Evaluation of Westinghouse Topical Reports Dealing with Elimination of Postulated Pipe Breaks in PWR Primary Main Loops," February 1, 1984.

1-3 Nuclear Regulatory Commission Docket #s 50-266 and 50-301 Letter from G. E. Lear, Director PWR Project Directorate #1 Division of PWR Licensing-A, NRC, to C. W. Fay, Vice President Nuclear Power Department Wisconsin Electric Power Company.

1-4 WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events," March 1978.

1-5 Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC (D. G. Eisenhut),

Westinghouse Proprietary Class 2, and November 10, 1981.

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

1-7 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.

1-8 Nuclear Regulatory Commission, 10 CFR 50, Modification of General Design Criteria 4 Requirements for Protection Against Dynamic Effects of Postulated Pipe Ruptures, Final

Rule, Federal RegisterNol.

52, No. 207/Tuesday, October 27, 1987/Rules and Regulations, pp. 41288-41295.

1-9 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.

May 2003 Introduction o:\\RGE.doc:1 b-052303

2-1 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 2.1 STRESS CORROSION CRACKING The Westinghouse RCS primary loops 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.

In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group. (The first Pipe Crack Study Group (PCSG) established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second PCSG was to include a review of the potential for stress corrosion cracking In Pressurized Water Reactors (PWR's).

The results of the study performed by the PCSG were presented in NUREG-0531 (Reference 2-1) entitled "Investigation and Evaluation of Stress-Corrosion Cracking 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 PWR primary system piping is extremely low because the ingredients that produce IGSCC are not all present. The use of hydrazine additives and a hydrogen overpressure limit the oxygen in the coolant to very low levels.

Other impurities that might cause stress-corrosion cracking, such as halides or caustic, are also rigidly controlled. Only for brief periods during reactor shutdown when the coolant is exposed to the air and during the subsequent startup are conditions even marginally capable-of producing stress-corrosion cracking in the primary systems of PWRs. Operating experience in PWRs supports this determination.

To date, no stress corrosion cracking has been reported in the primary piping or safe ends of any PWR."

During 1979, several instances of cracking In PWR feedwater piping led to the establishment of the 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 coolant systems.

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 Material Reliability Project (MRP) Program.

It should be noted that the susceptible material under investigation is not found In the primary loop piping at the Ginna Station.

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 Operation and Stability of the Reactor Coolant System May 2003 o:RGE.doc:1b-052303

2-2 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 external) as well as other material 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 intemally and extemally. 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 is expected to be in the ppb range by controlling charging flow chemistry and maintaining hydrogen in the reactor coolant at specified concentrations. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and fluorides within the specified limits.

Thus during plant operation, the likelihood of stress corrosion cracking is minimized.

2.2 WATER HAMMER Overall, there is a low potential for water hammer in the RCS since It is designed and operated to preclude the voiding condition in normally filled lines. The reactor coolant system, including piping and primary components, is designed for nornal, 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; pressure Is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only goveming 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. Preoperational testing and operating experience has verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer can occur.

Operation and Stability of the Reactor Coolant System May 2003 o:RGE.doc:1 b-052303

2-3 2.3 LOW CYCLE AND HIGH CYCLE FATIGUE An assessment of the low cycle fatigue loadings was carried out as part of this study in the form of 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 are minimized by restrictions placed on shaft vibrations during hot functional testing and operation.

During operation, an alarm signals the exceedence of the vibration limits. Field measurements have been made on a number of plants during hot functional testing, including plants similar to the Ginna Station. Stresses in the elbow below the reactor coolant pump resulting from system vibration have been found to be very small, between 2 and 3 ksi at the highest. These stresses are well below the fatigue endurance limit for the material and would also result in an applied stress intensity factor below the threshold for fatigue crack growth.

2.4 REFERENCES

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

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

Operation and Stability of the Reactor Coolant System o:\\RGE.doc:1b-052303 May 2003

3-1 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 hot leg pipe is shown in Figure 3-1. The outside diameter and minimum wall thickness of the pipe are 33.875 in. and 2.333 in., respectively, as shown in the figure. The normal stresses at the weld locations are from the load combination procedure discussed in Section 3.3 whereas the faulted loads are as described in Section 3.4. The components for normal loads are pressure, deadweight and thermal expansion.

An additional component, Safe Shutdown Earthquake (SSE), is considered for faulted loads. As seen from Table 3-2, the highest stressed location in the entire loop is at Location 1 at the reactor vessel outlet nozzle to pipe weld. This highest stressed location is a load critical location and is one of the locations at which, as an enveloping location, Leak-Before-Break is to be established. Essentially, a circumferential flaw postulated to exist at this location is subjected to both the normal loads and faulted loads to assess leakage and stability, respectively. The loads at this location are also given in Figure 3-1.

Since the elbows are made of cast stainless steel, thermal aging must be considered (Section 4.0). Thermal aging results in lower fracture toughness. Thus, locations other than the load critical locations must be examined by taking into consideration both fracture toughness and stress. The enveloping locations so determined are called touqhness critical locations. Once loads (Section 3.0) and fracture toughness (Section 4.0) are obtained, the load critical and toughness critical locations are determined (Section 5.0). At these locations, leak rate evaluations (Section 6.0) and fracture mechanics evaluations (Section 7.0) are performed per the guidance of Reference 3-1. Fatigue crack growth (Section 8.0) and stability margins are also evaluated (Section 9.0).

The weld locations being evaluated are shown In Figure 3-2.

3.2 CALCULATION OF LOADS AND STRESSES The stresses due to axial loads and moments are calculated by the following equation:

F M A= + M (3-1)

A Z

where, r

=

Stress F

=

Axial Load Pipe Geometry and Loading May2003 o:\\RGE.doc:1b-052303

3-2 M

=

Moment A

=

Pipe Cross-Sectional Area Z

=

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

M=( Mx2+My2+Mz 2)12 (3-2)

where, M

=

Moment for Required Loading Mx

=

X Component of Bending Moment My

=

Y Component of Bending Moment Mz

=

Z Component of Bending Moment NOTE:

X-axis is along the centerline of the pipe.

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

3.3 LOADS FOR LEAK RATE EVALUATION The normal operating loads for leak rate predictions are calculated by the following equations:

F

=

FD + FTH + F (3-3)

MX

=

(MX)DW + (MX)TH (3-4)

MY

=

(MY)DW + (MY)TH (3-5)

MZ

=

(MZ)D + (MZTH (3-6)

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

DW

=

Deadweight TH

=

Normal Thermal Expansion P

=

Load Due To Intemal Pressure This method of combining loads is often referred as the algebraic sum method (Reference 3-1).

The loads based on this method of combination are provided in Table 3-1 at all the weld locations identified in Figure 3-2. The outside diameter and minimum thickness are also given.

Pipe Geometry and Loading May 2003 o:\\RGE.doc:1b-052303

3-3 3.4 LOAD COMBINATION FOR CRACK STABILITYANALYSES In accordance with Standard Review Plan 3.6.3 (Reference 3-1), 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 /2 to 1.0. The absolute summation of loads are shown in the following equations:

F =IFDWI+IFTHI+IFPI+IFSSEINERTAI+IFSSESAMI (3-7)

MX = I (MX)DW I + I (MW)m I + I (MX)SSE INERTIA I + I (MX)SSE SAM I (3-8)

MY = I (MY)DW I + I (MY)TH I + I (MY)SSE INERTIA I + I (MY)SSE SAM I (3-9)

MZ = I (MZ)DW I + I (MZ)TH I + I (MZ)SSE INERTIA I + I (MZ)SSE SAM I (3-10) where subscripts SSE INERTIA and SSE SAM mean Safe Shutdown Earthquake Inertia and Safe Shutdown Earthquake Seismic Anchor Motion, respectively.

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

Table 3-2 shows the enveloped loads from both the loops and Table 3-1 shows the corresponding normal loads.

3.5 REFERENCES

3-1 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.

Pipe Geometry and Loading o:\\RGE.doc:1b-052303 May 2003

3-4 See Figure 3-2 Included Pressure Pipe Geomety and Loading o:\\RGE.doc:1b-052303 Table 3-1 Dimensions, Normal Loads and Normal Stresses for Ginna Station Outside Minimum AilLa"Mmn Location*

Diameter Thickness Axial Loa Momen Total Stress (psi)

(in)

(in)

(b)

(nIs 1

33.875 2.333 1503000 14021826 14720 2

33.875 2.333 1503000 4671201 9240 3

33.875 2.333 1503000 4537361 9162 4

37.188 2.989 1630058 8728777 8507 5

37.188 2.989 1697988 1389966 5834 6

36.188 2.489 1694718 1545528 7176 7

36.188 2.489 1690558 1761887 7264 8

36.188 2.489 1700888 1212286 7039 9

36.188 2.489 1700888 2700010 7755 10 37.188 2.989 1760128 5262643 7550 11 32.125 2.208 1347627 3980436 9235 12 32.125 2.208 1347627 4773674 9780 13 33.063 2.676 1351747 6178548 8728 May 2003

Location*** Axial Load*** (lbs)

Moment (In-lbs)

Total Stress (psi) 1 1753880 16298576 17139 2

1753870 5255054 10668 3

1756940 6610840 11475 4

1960878 13101928 11257 5

1852998 16724794 12345 6

1789808 10914559 12045 7

1785918 4938304 9154 8

1890568 7685554 10873 9

1878728 9524594 11713 10 1803438 14004621 11121 11 1498687 11962249 15456 12 1502957 9259743 13616 13 1485327 12229487 12618 See Figure 3-2 See Table 3-1 for Dimensions Included Pressure May 2003 Pipe Geometry and Loading oARGE.doc:lb-052303 3-5 Table 3-2 Faulted Loads and Stresses for Ginna Station

t

.Fa 7,1 OD = 33.875 in t = 2.333 in Normal Loads Force:

1503 kips Faulted Loads Force:

14022 in-kips Moment:

16299 in-kips See Table 3-1 See Table 3-2 Included the force due to a pressure of 2250 psia Figure 3-1 Hot Leg Coolant Pipe Pipe Geometry and Loading o:\\RGE.doc:1 b-052303 3-6 L

L Moment:

1754 kips May 2003

3-7 CROSSOVER LEG HOT LEG Temperature: 603.90F CROSS-OVER LEG Temperature: 543.1 °F COLD LEG Temperature: 543.1°F Pressure: 2250 psia Pressure: 2250 psia Pressure: 2250 psia Figure 3-2 Schematic Diagram of Ginna Station Primary Loop Showing Weld Locations May 2003 Pipe Geometry and Loading o:\\RGE.doc:1 b-052303 r 1

4-1 4.0 MATERIAL CHARACTERIZATION 4.1 PRIMARY LOOP PIPE AND FITTINGS MATERIALS The material type for the Ginna Station primary loop piping is A376 TP316 and for the elbow fittings is A351 CF8M.

4.2 TENSILE PROPERTIES The piping Certified Materials Test Reports (CMTRs) for the Ginna Station were used to establish the tensile properties for the Leak-Before-Break analysis. The CMTRs include tensile properties at room temperature and/or at 650 F for each of the heats of material.

These properties are given in Table 4-1 for piping and in Table 4-2 for elbows fittings.

For the A376 TP316 material, the representative properties at 603.9 F were established from the tensile properties at 650 F given in Table 4-1 by ublizing Section II of the 2001 ASME Boiler and Pressure Vessel Code (Reference 4-1). Code tensile properties at 603.9'F were obtained by interpolating between the 600OF and 650 F tensile properties. Ratios of the code tensile properties at 603.9 F to the corresponding tensile properties at 650 F were then applied to the 650 F tensile properties given in Table 4-1 to obtain the plant specific properties for A376 TP316 at 603.9F.

The Elbow Fittings Certified Materials Test Reports (CMTRs) for the Ginna Station were used to establish the tensile properties for the Leak-Before-Break analysis. The CMTRs for elbow fittings include tensile properties at room temperature and/or at 650 F for each of the heats of material. These properties are given in Table 4-2.

For the A351 CF8M material, the representative properties at 603.Q9F and 543.1F were established from the tensile properties at 650F given in Table 4-2 by ufilizing section 11 of the 2001 ASME boiler and pressure vessel code. Code tensile properties at 603.9 F and 543.1*F were established by interpolating between the 500 F, 600 F and the 650 F tensile properties.

Ratios of the code tensile properties at 603.9 F and 543.1 F to the corresponding properties at 650*F were then applied to the 650 F tensile properties given in Table 4-2 to obtain the plant specific representative properties for A351 CF8M at 603.9 F and 543.1 F.

The average and lower bound yield strengths and lower bound ultimate strengths are given in Table 4-3. The ASME code Moduli of Elasticity are also given in Table 4-3, and Poisson's Ratio was taken as 0.3.

Material Characterization May 2003 o:ARGE.doc:1b-052303

4-2 4.3 FRACTURE TOUGHNESS PROPERTIES The pre-service fracture toughnesses of cast stainless steels in terms of Jc have been found to be very high at 600 F. [

]ace However, cast stainless steel is susceptible to thermal aging at the reactor operating temperature, that is, about 290 C (5500F). Thermal aging of cast stainless steel results in embrittlement, that is, a decrease in the ductility, impact strength, and fracture toughness of the material. Depending on the material composition, the Charpy impact energy of a cast stainless steel component could decrease to a small fraction of its original value after exposure to reactor temperatures during service.

The susceptibility of the material to thermal aging increases with increasing ferrite contents.

The molybdenum bearing CF8M shows increased susceptibility to thermal aging. The method described below was used to calculate the end of life toughness properties for the cast material.

In 1994, the Argonne National Laboratory (ANL) completed an extensive research program in assessing the extent of thermal aging of cast stainless steel materials. The ANL research program measured mechanical properties of cast stainless steel materials after they have been heated in controlled ovens for long periods of time. ANL compiled a database, both from data within ANL and from intemational sources, of about 85 compositions of cast stainless steel exposed to a temperature range of 290-400 C (550-750*F) for up to 58,000 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> (6.5 years).

From this database, ANL developed correlations for estimating the extent of thermal aging of cast stainless steel (References 4-3 and 4-4).

ANL developed the fracture toughness estimation procedures by correlating data in the database conservatively.

After developing the correlations; ANL validated the estimation procedures by comparing the estimated fracture toughness with the measured value for several cast stainless steel plant components removed from actual plant service. The ANL procedures produced conservative estimates that were about 30 to 50 percent less than actual measured values. The procedure developed by ANL in Reference 4-4 was used to calculate the end of life fracture toughness values for this analysis. ANL research program was sponsored and the procedure was accepted (Reference 4-5) by the NRC.

The chemical compositions are available from CMTRs and are provided in Table 4-4.

Cr(,q)= Chromium Equivalent Ni(eq) = Nickel Equivalent where Cr(eq) and Ni.q are in percent weight

&= Ferrite In Percent Volume Cr(,q), Ni(Q) and ic values obtained from Reference 4-2 are shown in Table 4-4.

Material Characterization May 2003 o:\\RGE.doc:1b-052303

4-3 The following equations are taken from Reference 4-4.

For CF8M steel with <10% Ni, the saturation value of saturation room temperature (RT) impact energy Cvst (J/cm2) is the lower value determined from logoCOvsat = 1.10 + 2.12 exp (-0.041¢)

(4-1) where the material parameter

• is expressed as

• = 5¢ (Ni + Si + Mn)2 (C + O.4N)/5; (4-2) and from log,oCvs, = 7.28 - 0.011&c - 0.185Cr - 0.369Mo - 0.451 Si

-0.007Ni - 4.71 (C + AN)

(4-3)

For CF8M steel with >10% Ni, the saturation value of RT impact energy Cvsat (J/cm2 ) is the lower value determined from logloCvat = 1.10 + 2.64 exp (0.064Q)

(44) where the material parameter is expressed as

• = 8c (Ni + Si + Mn)2 (C + 0.4N)/5 (4-5) and from log 1oCvsw = 7.28 - 0.0118 - 0.1 85Cr- 0.369Mo - 0.451 Si

-0.007Ni - 4.71 (C + 0.4N)

(4-6)

The RT impact energies of the cast stainless steel materials were determined from the chemical compositions available from CMTRs and provided in Table 4-4.

The saturation J-R curve at 290*C(554*F) for static-cast CF8M steel is given by.

Jd = 49 (CVat 0 41 (a)n (4-7) n = 0.23 + 0.06 log,0 (Cvst)

(4-8) where Jd is the "deformation Jr in kJ/m2 and Aa is the crack extension in mm.

Material Characterization May 2003 o:\\RGE.doc:1b-052303

4-4

]a,ce a,c,e The correlations presented in Reference 4-4 are applicable to cast stainless steels used in the U.S. nuclear industry, the steels contain <25% ferrite in almost all cases.

[

Ia,c,e The results from the ANL Research Program indicate that the lower-bound fracture toughness of thermally aged cast stainless steel is similar to that of Submerged Arc Welds (SAWs). The applied value of the J-integral for a flaw in the weld regions will be lower than that in the base metal because the yield strength for the weld materials is much higher at the temperature'.

Therefore, weld regions are less limiting than the cast material.

In the fracture mechanics analyses that follow, the fracture toughness properties given in Table 4-5 will be used as the criteria against which the applied fracture toughness values will be compared.

'In the report, all the applied J values were conservatively determined by using base metal strength properUes.

Material Characterization o:RGE.doc:1b-052303 May 2003

4-5

4.4 REFERENCES

4-1 Boler and Pressure Vessel Code Section II, Part D - Material Properties, 2001 Edition, July 1, 2001, ASME Boiler and Pressure Vessel Committee, Subcommittee on Materials.

4-2 A800/A800M Standard Practice for Steel Casting, Austenitic Alloy, Estimating Ferrite Content Thereof, Section 1 -

Iron and Steel Products, Vol. 01.02, Ferrous Castings; Ferroalloys; Shipbuilding.

4-3

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

4-4

0. K. Chopra, "Estimation of Fracture Toughness of Cast Stainless Steels During Thermal Aging in LWR Systems," NUREG-CR-4513, Revision 1, U. S. Nuclear Regulatory Commission, Washington, DC, August 1994.

4-5

'Flaw Evaluation of Thermally aged Cast Stainless Steel in Ught-Water Reactor Applications," Lee, S.; Kuo, P. T.; Wichman, K.; Chopra, O.; Published in International Joumal of Pressure Vessel and Piping, June 1997.

Material Characterizaton o:\\RGE.doc:1 b-052303 May2003

4-6 Table 4-1 Measured Tensile Properties (psi) for the Ginna Station Primary Loop Piping (Material A376 TP316)

At Room Temperature At 6500 F HEAT SERIAL YIELD ULTIMATE YIELD ULTIMATE NO.

NO.

LOCATION STRENGTH STRENGTH STRENGTH STRENGTH (psi)

(psi)

(psi)

(psi)

D8650 1476 Hot Leg 34500 75300 21000 65100 D8650 1476 Hot Leg 36000 76500 D8650 1477 Hot Leg 36500 76500 D8650 1477 Hot Leg 38000 78500 D8632 1480 Hot Leg 36500 76200 22700 64200 D8632 1480 Hot Leg 37200 78100 D6886 1481 Hot Leg 38000 79100 21100 65700 D6886 1481 Hot Leg 39000 80400 D8548 1157A Cross-over Leg 33000 75900 20800 63100 D8548 11 57A Cross-over Leg 40300 77900 V0246 1996 Cross-over Leg 30000 75800 20600 59700 V0246 1996 Cross-over Leg 30500 75000 D8552 1154 Cross-over Leg 37000 80400 23500 66700 D8552 1154 Cross-over Leg 40000 81400 V0249 1982 Cross-over Leg 31400 75800 21900 63400 V0249 1982 Cross-over Leg 33000 76600 D8647 1471 Cold Leg 36200 75200 21100 66200 D8647 1471 Cold Leg 38500 79000 D8651 1473 Cold Leg 37000 75600 22100 60600 D8651 1473 Cold Leg 40000 80600 Material Characterization May 2003 o:\\RGE.doc:1b-052303

Table 4-2 Measured Tensile Properties (psi) for the Ginna Station Primary Loop Elbow Fittings (Material A351 CF8M)

At Room Temperature At 650°F HEAT YIELD ULTIMATE YIELD ULTIMATE NO.

LOCATION STRENGTH STRENGTH STRENGTH STRENGTH (psi)

(psi)

(psi)

(psi) 5175-1 Hot Leg 42000 90000 01547-1 Hot Leg 49500 91000 5038-1 Cross-over Leg 58500 98000 00024-1 Cross-over Leg 41400 81500 25500 66750 2786-1 Cross-over Leg 54000 94500 03182-1 Cross-over Leg 40500 84500 27000 69800 4693 Cross-over Leg 43500 87000 4312 Cross-over Leg 55500 94000 01493-1 Cross-over Leg 46500 91000 01799-1 B Cross-over Leg 48000 90000 4672 Cross-over Leg 49500 90000 4858 Cross-over Leg 48000 88000 2506-2 Cold Leg 52500 93500 3764-5 Cold Leg 48000 87000 May 2003 Material Characterization o:\\RGE.doc:1 b-52303 4-7

4-8 a,c,e Material Characterization May 2003 o:\\RGE.doc:1b-052303

4-9 a,c,e

• Heats for the Hot Leg; "Heats for the Cold Leg All other heats are in cross-over leg.

N Is assumed as 0.05

'From Equations 4-1 or 4-4 2From Equations 4-3 or 4-6 3Minimum of Cv,at' and Cvm' A-4--f-I ^L-^A 2

1__

oatiGEcl %Ab-teZ3On o.:IRGE.doc:1b-051503 May 2003

4-10 a,c,e Material Characterization May 2003 oAXRGE.dioc:1lb-051 503

4-11 a.c,e Figure 4-1 Pre-Service J vs. Aa for SA351 CF8M Cast Stainless Steel at 600F Matenal Characterizaton o:\\RGE.doc:1b-051503 May 2003 1

5.0 CRITICAL LOCATIONS AND EVALUATION CRITERIA 5.1 CRITICAL LOCATIONS The Leak-Before-Break (LBB) evaluation margins are to be demonstrated for the limiting locations (goveming locations). Candidate locations are designated as load critical location or toughness critical locations as discussed in Section 3.0. Such locations are established based on the loads in Section 3.0 and the material properties established in Section 4.0.

These locations are defined below for the Ginna Station. Table 3-2 as well as Figure 3-2 are used for this evaluation.

Load Critical Locations The highest stressed location for the entire primary loop is at Location 1 (in the Hot Leg)

(See Figure 3-2) at the reactor vessel outlet nozzle to pipe weld. Location I is the critical location for all the weld locations in the primary loop piping. Since it is on a straight pipe with forged material, it is a high toughness location.

Toughness Critical Locations Low toughness locations are at the ends of every elbow. In the case of the hot leg, low toughness is found for Heat No. 01547-1 (see Figure 3-2 for locations). Location 3 has the higher faulted stress than location 4 and will yield a higher Japp. In the case of cross-over leg and cold leg, the low toughness is found for Heat No. 4312. Location 12 has the highest faulted stress in the cross-over leg and cold leg elbow and will yield the highest J.

It is thus concluded that the enveloping locations are 1, 3 and 12. The allowable toughness values for the critical locations are shown in Table 4-5.

5.2 FRACTURE CRITERIA As will be discussed later, fracture mechanics analyses are made based on loads and postulated flaw sizes related to leakage. The stability criteria against which the calculated J and tearing modulus are compared are:

(1)

If Jam < Jic, then the crack will not initiate; (2)

If Japp 2 Jlc, but, if T w < Tmat and J., < Jm,. then the crack Is stable.

where:

Jaw

=

Applied J Jic

=

J at Crack Initiation T.ppm=

Applied Tearing Modulus Critical Locations and Evaluation Criteria May 2003 o:\\RGE.doc:1 b-052303

5-2 Tmat =

Material Tearing Modulus JI

=

Maximum J value of the material For critical locations, the limit load method discussed in Section 7.0 was also used.

Critical Locations and Evaluation Criteria o:RGE.doc:1b-052303 May 2003

6-1 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 through postulated through-wall cracks and present the leak rate calculation results for through-wall circumferential cracks.

6.2 GENERAL CONSIDERATIONS The flow of hot pressurized water through an opening to a lower back pressure causes flashing which can result in choking. For long channels where the ratio of the channel length, L, to hydraulic 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 rate through a crack was calculated in the following manner.

Figure 6-1 from Reference 6-1 was used to estimate the critical pressure, P for the primary loop enthalpy condition and an assumed flow. Once P was found for a given mass flow, the [

Ia,e was found from Figure 6-2 (taken from Reference 6-1). For all cases considered, since [

race Therefore, this method will yield the two-phase pressure drop due to momentum effects (P 24) as illustrated in Figure 6-3, where PO is the operating pressure. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using Apf1

) ac,e (6-1) where the friction factor f is determined using the a."

The crack relative roughness,

, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [

Iaxx The frictional pressure drop using equation 6-1 is then calculated for the assumed flow rate and added to the [

Ia,'

to obtain the total pressure drop from the primary system to the atmosphere. That Is, for the primary loop, Leak Rate Predictions May2003 o:\\RGE.doc:1 b-O052303

6-2 Absolute Pressure - 14.7 = [

]aAe (6-2) for a given assumed flow rate G. If the right-hand side of equation 6-2 does not agree with the pressure difference between the primary loop and the atmosphere, then the procedure is repeated until Equation 6-2 is satisfied to within an acceptable tolerance which in tum leads to correct flow rate value for a given crack size.

6.4 LEAK RATE CALCULATIONS Leak rate calculations were made as a function of crack length at the governing locations previously identified in Section 5.1. The normal operating loads of Table 3-1 were applied in these calculations.

The crack opening areas were estimated using the method of Reference6-2 and the leak rates were calculated using the two-phase flow formulation described above. The average material properties of Section 4.0 (see Table 4-3) were used for these calculations.

The Ginna Station RCS pressure boundary leak detection system of 0.25 gpm meets the intent of Reg. Guide 1.45, which is 1 gpm in 1 hour1.157407e-5 days <br />2.777778e-4 hours <br />1.653439e-6 weeks <br />3.805e-7 months <br /> or less. Thus, to satisfy the margin of 10 on the leak rate, the flaw sizes (leakage flaw sizes) are determined which yield a leak rate of 2.5 gpm.

The flaw sizes to yield a leak rate of 2.5 gpm were calculated at the goveming locations and are given in Table 6-1. The flaw sizes so determined are called leakaae flaw sizes.

6.5 REFERENCES

6-1.

[

] a.c.e.

6-2.

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 11-1, NUREG/CR-3464, September 1983.

May 2003 Leak Rate Predictions o:UGE.doc:1b-0 3/03

6-3 a,c,e Leak Rate Predictions May 2003 o:\\RGE.doc:1 Ob-5/23103

6-4 a,c,e Figure 6-1 Analytical Predictions of Critical Rates of Steam-Water Mixtures Leak Rate Predicions o:\\RGE.doc:1 b-05/23103 May 2003

6-5 a,c,e Figure 6-2 J'1" Pressure Ratio as a Function of LAD Leak Rate Predictions o:\\RGE.doc:1b-05/23/03 May 2003

6-6 F

Figure 6-3 Idealized Pressure Drop Profile Through a Postulated Crack Leak Rate Predictions o:\\RGE.doc:1 b-23/03 a,c,e May2003

7-1 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 of crack-tp blunting, initiation, extension and finally crack instability. The local stability will be assumed if the crack does not initiate at all. It has been accepted that the initiation toughness measured in terms of Jlc from a J-integral resistance curve is a material parameter defining the crack initiation. If, for a given load, the calculated J-integral value is shown to be less than the Jic of the material, then the crack will not initiate. If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:

dJ E da where:

Tapp

=

Applied Tearing Modulus E

=

Modulus of Elasticity rf

=

0.5 (y + c,) = Flow Stress a

=

Crack Length cs,Y aU

=

Yield and Ultimate Strength of the Material, respectively Stability is said to exist when ductile tearing occurs f Tam is less than Tm, the experimentally determined tearing modulus. Since a constant T,, is assumed, a further restriction is placed on J.

J must be less than J,,

where J.

is the maximum value of J for which the experimental T,ma is greater than Tam used.

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

(1) If J,pp < Jic, then the crack will not initiate.

(2)

If Ja

> Jlc, but, if Tax < T,,.t and J., < Jx, then the crack is stable.

7.2 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 plastic instability Fracture Mechanics Evaluation May 2003 o\\RGE.doc:1b-052303

7-2 method, based on traditional plastic limit load concepts, but accounting for strain hardening and taking into account the presence of a flaw. The flawed pipe 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. The flow stress is generally taken as the average of the yield and ultimate tensile strength of the material at the temperature of interest.

This methodology has been shown to be applicable to ductile piping through a large number of experiments and will be used here to predict the critical flaw size in the primary coolant loop piping. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 7-1) when loads are applied.

The detailed development is provided in Appendix A for a through-wall circumferential flaw in a pipe with intemal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:

[

Ia,Ae where:

[

]ae 0

=

0.5 (y + a) = Flow Stress, psi Ia,ce I

Ia,ce Fracture Mechanics Evaluation o:\\RGE.doc:1b-052303 May 2003

7-3 The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect the limit moment.

Good agreement was found between the analytical predictions and the experimental results (Reference 7-1).

For application of the limit load methodology, the material, including consideration of the configuration, must have a sufficient ductility and ductile tearing resistance to sustain the limit load.

7.3 RESULTS OF CRACK STABILITY EVALUATION Local Failure Mechanism jax.e Fracture Mechanics Evaluation o:ARGE.doc:1 b-052303 May 2003

7-4

]ac,e Global failure mechanism The critical locations were also identified in Section 5.1. A stability analysis based on limit load was performed for these locations as described In Section 7.2. The welds at these locations are Gas Tungsten Arc Weld (GTAW) and Manual Metallic Arc Weld (MMAW) which is similar to Shielded Metal Arc Weld (SMAW). The Z' factor correction for GTAW is 1.0, while that for SMAW was applied (Reference 7-5) as follows:

Z = 1.15 [ 1.0 + 0.013(OD-4) where OD is the outer diameter of the pipe in inches.

Fracture Mechanics Evaluation May 2003 o:\\RGE.doc:1b-052303

7-5 The Z-factors for SMAW welds were calculated for the critical locations, using the dimensions given in Table 3-1. The calculated Z factors were 1.60, 1.60 and 1.57 for locations 1, 3 and 12 respectively. The applied loads were increased by the Z factors for SMAW welds and plots of limit load versus crack length were generated as shown in Figures 7-2, 7-3 and 7-4. Table 7-2 summarizes the results of the stability analyses based on limit load. The leakage size flaws are also presented in the same Table.

7.4 REFERENCES

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

7.2 Johnson, W. and Mellor, P. B., Engineering Plasticity, Van Nostrand Reinhold Company, New York, (1973), pp. 83-86.

7.3 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 ll-1, NUREGICR-3464, September 1983.

7.4 Irwin, G R., "Plastic Zone Near a Crack and Fracture Toughness," Proc. 7h Sagamore Conference, R IV-63 (1960).

7.5 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.

F ure Mechanics Evaluaton o:\\RGE.doc:1 b-052303 May 2003

7-6 a,c,e Notes:

• Tn,at is not applicable since Japp < Jc Jpp values were calculated for flaw size(s) of 2 times the 2.5 gpm leakage flaw size(s). As can be seen from Table 7-1, Japp values are much lower than the Jlc allowable values and therefore there are ample margins available for the J calculation.

a,c,e Fracture Mechanics Evaluation o:\\RGE.doc:1b-052303 May 2003

Figure 7-1

]ac.. Stress Distribution Fracture Mechanics Evaluation o:\\RGE.doc:1 b-052303 7-7 May 2003

7-8 a,c,e r

OD = 33.875 in t = 2.333 in ay= 21.011 ksi F = 1753.88 kips cr = 59.700 ksi M = 16298.576 in-kips A376 TP316 wih SMAW weld Figure 7-2 Critical Flaw Sze Prediction - Hot Leg at Location 1 Fracture Mechanics Evaluation o:\\RGE.doc:lb-052303 May 2003

OD = 33.875 in t = 2.333 in a,e,e ay = 26.011 ksi F = 1756.94 kips a,= 66.750 ksi M = 6610.84 in-kips A351 CF8M with SMAW weld Figure 73 Critical Flaw Size Prediction - Hot Leg at Location 3 Fracture Mechanics Evaluation o:XRGE.doc:1b-052303 7-9 May 2003

7-10 a,c,e OD = 32.125 in t = 2.208 in qy= 26.922 ksi F = 1502.957 kips au = 66.750 ksi M = 9259.743 in-kips A351 CF8M with SMAW weld Figure 7-4 Critical Flaw Size Prediction - Cold Leg at Location 12 Fracture Mechanics Evaluation o:ARGE.doc:1 b-052303 May 2003

8-1 8.0 FATIGUE CRACK GROWTH ANALYSIS To determine the sensitivity of the primary coolant system to the presence of small cracks, a fatigue crack growth analysis was carried out for the [

]ace region of a typical system (see Location [

]

of Figure 3-2). This region was selected because crack growth calculated there would be typical of that in the entire primary loop. Crack growths calculated at other locations would be expected to show less than 10% variation.

At [Iace of a plant typical in geometry and operational characteristics to any Westinghouse PWR system.

The normal, upset, and test conditions were considered. A summary of the generic thermal transients used conservatively for the Ginna Station for 60 years is provided in Table 8-1.

Circumferentially oriented surface flaws were postulated in the region.

These flaws were assumed to be located in two different locations, as shown in Figure 8-1. Specifically, these were:

Cross-Section A:

a.ce Cross-Section B: l

] a.c,e Fatigue crack growth rate laws were used [

]ac.e The law for stainless steel was derived from Reference 8-1. A compilation of data for austenitic stainless steel in a PWR water environment was presented in Reference 8-2, and it was found that the effect of the environment on the crack growth rate was very small. From this information t was estimated that the environmental factor should be conservatively set at [

1ac, in the crack growth rate equation from Reference 8-1.

For stainless steel, the fatigue crack growth formula is:

3a.ce The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 8-2. The result shows that the crack growth for 60 Fatigue Crack Growth Analysis May 2003 o:\\RGEdoc:1 b/052303

8-2 years is very small, [

Iaje Therefore fatigue crack growth is not a concern for the Ginna Station primary loop piping.

8.1 REFERENCES

8-1 James, L. A. and Jones, D. P.,

Fatigue Crack Growth Correlations for Austenitic Stainless Steel in Air, Predictive Capabilities in Environmentally Assisted Cracking,"

ASME publication PVP-99, December 1985.

8-2 Bamford, W. H., Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment," Trans. ASME Joumal of Pressure Vessel Technology, Vol. 101, Feb. 1979.

Fatigue Crack Growth Analysis o.\\RGE.doc:1 b-052303 May 2003

8-3 Table 8-1 Summary of Reactor Vessel Transients (60 Years)

Number Number Transient Identification of Cycles Normal Conditions 1

Heatup and Cooldown at 1 00°F/hr 200 2

Load Follow Cycles 18300 (Unit Loading and Unloading at 5 %of full power/min) 3 Step Load Increase and Decrease 2000 4

Large Step Load Decrease, with Steam Dump 200 5

Steady State Fluctuations 1xi 06 Upset Conditions 6

Loss of Load (Without immediate Turbine or Reactor Trip) 80 7

Loss of power (Blackout with natural circulation in the Reactor Coolant System) 40 8

Loss of Flow (Partial Loss of Flow, one pump only) 80 9

Reactor Trip From Full Power 400 Test Conditions 10 Turbine Roll Test 10 11 Hydrostatic Test Conditions - Primary Side Leak Test 50 12 Cold Hydrostatic Test 10 Fatigue Crack Growth Analysis May 2003 o:\\RGE.doc:1 b052303

8-4 Fatigue Crack Growth Analysis o:\\RGE.do:1 b-052303 Table 8-2 Fatigue Crack Growth at [

Ia,"

( 60 years)

Final Flaw (In.)

Initial Flaw (in.)

in

.)

I ]a,"

.I 18lace May 2003

Figure 8-1 Typical Cross-Section of [

Fatigue Crack Growth Analysis o:%RGE.doc:1 bJ052303 8-5 a,c,e

]ace May 2003

8-6 a,c,e Figure 8-2 Reference Fatigue Crack Growth Curves for Carbon and Low Alloy Ferritic Steels Fatigue Crack Growth Anaysis oNRGE.doc:1 b-052303 May 2003

9-1

9.0 ASSESSMENT

OF MARGINS The results of the leak rates of Section 6.4 and the corresponding stability and fracture toughness evaluations of Sections 7.1, 7.2 and 7.3 are used in performing the assessment of margins. Margins are shown in Table 9-1.

In summary, at all the critical locations relative to:

1. Raw 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 2.5 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 0.25 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).

A margin on loads >1 using the absolute summation of faulted load combinations is satisfied.

Assessment of Margins o:\\RGEdoc:1 b052303 May 2003

9-2 a,c,e Based on Limit Load Based on J Integral Evaluation [Note: margin on flaw size(s) is greater than 2 since J values for critical flaw size(s) of 2 times the leakage flaw size(s) are much lower than the Jc values as shown in Table 7-1 ]

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10-1

10.0 CONCLUSION

S This report justifies the elimination of RCS primary loop pipe breaks from the structural design basis for the 60 year plant life of Ginna Nuclear Power Plant as follows:

a.

Stress corrosion cracking is precluded by the use of fracture resistant materials in the piping system and controls on the reactor coolant chemistry, temperature, pressure, and flow during normal operation.

b.

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

c.

The effects of low and high cycle fatigue on the integrity of the RCS piping are negligible.

d.

Ample margin exists between the leak rate of small stable flaws and the capability of the Ginna Nuclear Power Plant Reactor Coolant System Pressure Boundary Leakage Detection System.

e.

Ample margin exists between the small stable leakage flaw sizes of item (d) and larger stable critical flaw sizes.

f.

Ample margin exists in the J-integral material properties used to demonstrate end-of-service life (relative to aging) stability for the critical flaws. Based on the stability analysis results, Cast Austenitic Stainless Steel (CASS) elbow material of Ginna Nuclear Power Plant primary loop piping is not an issue.

For the critical locations, flaws are identified that will be stable because of the ample margins described in tems (d), (e), and (f) above.

Based on the above, the Leak-Before-Break conditions are satisfied for the Ginna Nuclear Power Plant RCS piping. All the recommended margins are satisfied. It is therefore concluded that the dynamic effects of RCS primary loop pipe breaks need not be considered in the structural design basis of the Ginna Nuclear Power Plant for the 60 year plant life as part of the Ucense Renewal Program.

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A-1 APPENDIX A LIMIT MOMENT I

Ia,c.e Appendix A May2003 Appendx A oA438.doc:b-052303 May 2003

A-2 a,c,e Figure A-1 Pipe with a Through-Wall Crack In Bending May 2003 Appendix A o:\\RGE.doc:1 b052303