ML18100A466
| ML18100A466 | |
| Person / Time | |
|---|---|
| Site: | Salem  |
| Issue date: | 05/31/1993 |
| From: | Swamy S, Yang C WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML18100A462 | List: |
| References | |
| WCAP-13660, NUDOCS 9307130214 | |
| Download: ML18100A466 (80) | |
Text
WCAP-13660 WPF0954A: 1 b/050493 WESTINGHOUSE CLASS 3 TECHNICAL JUSTIFICATION FOR ELIMINATING LARGE PRIMARY LOOP PIPE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE SALEM GENERATING STATION UNITS 1AND2.
VERIFIED:
MAY 1993 D. C. Bhowmick D. E. Prager Work Performed Under Shop Order PDGP-950 WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P. 0. Box 355 Pittsburgh, Pennsylvania 15230-355
© 1993 Westinghouse Electric Corporation All Rights Reserved
TABLE OF CONTENTS Section Title Page EXECUTIVE
SUMMARY
xi
1.0 INTRODUCTION
1-1 1.1 Purpose 1-1 1.2 Scope 1-1 1.3 Objectives 1-1 1.4 Background Information 1-2 1.5 References 1-3 2.0 OPERATION AND STABILITY OF THE REACTOR 2-1 COOLANT SYSTEM 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 3-2 3.3 Loads for Leak Rate Evaluation 3-3 3.4 Load Combination for* Crack Stability 3-3 Analyses 3.5 References 3-4 4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Primary Loop Pipe and Fittings Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties.
4-2 4.4 References 4-3 WPF1080A: 1 b/042893 lV
TABLE OF CONTENTS Section 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 6-3 7.0 FRACTURE MECHANICS EVALUATION 7-1 7.1 Local Failure Mechanism 7-1 7.2 Global Failure Mechanism 7-2 7.3 Results of Crack Stability Evaluation 7-3 7.4 References 7-4 8.0 FATIGUE CRACK GROWTH ANALYSIS 8-1 8.1 References 8-2
9.0 ASSESSMENT
OF MARGINS 9-1
10.0 CONCLUSION
10-1 WPF1080A: 1 b/042893 v
Section APPENDIX A-APPENDIX B -
WPF1080A:lb/042893
- TABLE OF CONTENTS Limit Moment Toughness Criteria for Salem Cast Primary Loop Components Vl A-1 B-1
Table 3-1 3-2 4-1 4-2 4-3
.4-4 4-5 6-1 LIST OF TABLES Dimensions, Nonnal Loads and Nonnal Stresses for Salem Faulted Loads and Stresses for Salem Measured Tensile Properties for Salem Primary Loop Piping Unit 1 Measured Tensile Properties for Salem Primary Loop Piping Unit 2 Measured Room Temperature Tensile Properties for Salem Primary Loop Elbow Fittings Mechanical Properties for Salem Materials at Operating Temperatures Fracture Toughness Properties for Salem Primary Loops for Leak-Before-Break Evaluation at Critical Locations Flaw Sizes Yielding a Leak Rate of 10 gpm at the Governing Locations WPF1080A: lb/042893 Vll 3-5 3-6 4-5 & 4-6 4-7 & 4-8 4-9 &
4-10 4-11 4-12 6-4
LIST OF TABLES (cont)
Table Title Page 7-1 Stability Results for Salem Based on 7-5 Elastic-Plastic I-Integral Evaluations 7-2 Stability Results for Salem Based on 7-6 Limit Load 8-1 Summary of Reactor Vessel Transients 8-3 8-2 Typical Fatigue Crack Growth at [
8-4
]a,c,e (40 Years) 9-1 Leakage Flaw Size, Critical Flaw Sizes and Margins 9-2 for Salem Generating Station Units 1 and 2 B-1 Chemistry and Fracture Toughness Properties of the B-2 Material Heats of Salem Unit 1 B-2 Chemistry and Fracture Toughness Properties of the B-5 Material Heats of Salem Unit 2 WPF1080A: 1 b/042893 Vlli
Figure 3-1 3-2 4-1 4-2 4-3 6-2 6-3 LIST OF FIGURES Title Hot Leg Coolant Pipe Schematic Diagram of Salem Primary Loop Showing Weld Locations Representative Lower Bound True Stress - True Strain Curve for A35 l CF8M at 545°F J vs. Aa for SA35 l CF8M Cast Stainless Steel at 600°F J vs. Aa at Different Temperatures for Aged Material
[
]a,c,e (7 500 Hours at 400°C)
Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures
[
ofUD
]a,c,e Pressure Ratio as a Function Idealized Pressure Drop Profile Through a Postulated Crack WPF1080A: lb/042893 ix Page 3-7 3-8 4-13 4-14 4-15 6-5 6-6 6-7
Figure 7-1 7-2 7-3 7-4 8-1 8-2 8-3 A-1 LIST OF-FIGURES (cont)
[
]a,c,e Stress Distribution Critical Flaw Size Prediction - Hot Leg at Location 1 Critical Flaw Size Prediction - Cross-over Leg at Location 6 Critical Flaw Size Prediction - Cross-over Leg at Location 9 Typical Cross-Section *of [
Reference Fatigue Crack Growth Curves for [
t,c,e Reference Fatigue Crack Growth Law for [
in a Water Environment at 600°F Pipe with a Through-Wall Crack in Bending WPF0954A: 1 b/050493 x
7-7 7-8 7-9 7-10 8-5 8-6 8-7 A-2
EXECUTIVE
SUMMARY
The existing structural design basis for the reactor coolant system of the Salem Generating Station Units 1 and 2 require that the dynamic effects of pipe breaks be evaluated and that protective measures for such breaks be incorporated into the design. However, within the last decade, such breakS have been shown to be highly unlikely and should not be included in the structural design basis of Westinghouse type pressurized water reactors. To eliminate primary loop pipe breaks from the design basis, it must be demonstrated to the satisfaction of the U.S.
Nuclear Regulatory Commission that a leak-before-break situation exists. This report provides such a demonstration for the Salem Generating Station Units 1 and 2.
In this report it is shown that the primary loops are highly resistant to stress corrosion cracking and high and low cycle fatigue. Water hammer is mitigated by system design and operating procedures.
The primary loops were extensively examined. The as-built geometries for the pipe and elbows and loadings were obtained. The materials were evaluated using the Certified aterials Test Reports. Mechanical properties were determined at operating temperatures.
ince the piping systems include cast stainless steel fittings, fracture toughnesses considering thermal aging were 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 found which would cause 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, fatigue crack growth was shown not to be an issue for the primary loops.
It is concluded that dynamic effects of reactor coolant system primary loop pipe breaks need not be considered in the structural design basis of the Salem Generating Stations.
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SECTION
1.0 INTRODUCTION
1.1 Purpose This repon applies to the Salem Generating Station Units 1 and 2 Reactor Coolant System (RCS) primary loop piping. It is intended to demonstrate that for the specific parameters of the Salem Stations, RCS primary loop pipe breaks need not be considered in the structural design basis. The approach taken has been accepted by the Nuclear Regulatory Commission (NRC) (Reference 1-1 ).
1.2 Scope The existing structural design basis for the RCS primary loop requires that dynamic effects of pipe breaks be evaluated. Specifically, as part of the Loss of Coolant Accident (LOC..\\.)
design basis for the Salem Generating Stations the following pipe breaks are postulated in the CS primary loop piping: the six terminal ends in the cold. hot. and crossover legs; a split in e steam generator inlet elbow, and the loop closure weld in the crossover leg. However.
Westinghouse has demonstrated on a generic basis that RCS primary loop pipe breaks are highly unlikely and should not be included in the structural design basis of Westinghouse
~lants (see Reference 1-2). In order to demonstrate this applicability to the Salem Generating Stations. Westinghouse has performed a fracture mechanics evaluation. a determination of leak rates from a through-wall crack. a fatigue crack growth evaluation, and an assessment of margins against crack instability consistent with the leak-before-break (LBB) methodology.
Through this successful application of the LBB methodology, the above eight break locations in the RCS primary loop piping are eliminated from the Salem Generating Stations' structural design basis.
1.3 Objectives In order to validate the elimination of RCS primary loop pipe breaks from the Structural Design Basis for the Salem Stations. the following objectives must be achieved:
- a.
Demonstrate that margin exists between the critical crack size and a postulated crack which yields a detectable leak rate.
- b.
Demonstrate that there is sufficient margin between the leakage through a postulated crack and the leak detection capabHity of the Salem Stations.
WPF0954A: 1 b/050493 1-1 J
- c.
Demonstrate margin on applied load.
- d.
Demonstrate that fatigue crack growth is negligible.
1.4 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 1-3). That topical report employed a dete_rministic fracture mechanics evaluation and a probabilistic analysis to support the elimination of RCS primary loop pipe breaks. 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 1000 pipe breaks. This material was provided to the NRC along with Letter Report
~S-EPR-2519 (Reference 1-4).
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 1-5 and 1-6). The results from the LLNL study_ were_
released at a March 28, 1983, ACRS Subcommittee meeting. These studies which are applicable to all 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 10-12 per reactor year and the mean probability of an indirect LOCA to be 10-7 per reactor y~ar. Thus, the results previously obtained by Westinghouse (Reference 1-3) were confirmed by an independent NRC research study.
Based on the studies by Westinghouse. LLNL. the ACRS, and the AIF, the NRC completed a
- safety review of the Westinghouse reports submitted to address asymmetric blowdown loads that result from a number of discrete break locations on the PWR primary systems. The NRC Staff evaluation (Reference 1-1) concludes that an acceptable technical basis has been provided so that asymmetric blowdown loads need not be considered for those plants mat 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 WPF0954A: lb/050493 1-2
olant loop integrity. In a more formal recognition of LBB methodology applicability for.
WRs, the NRC appropriately modified 10 CFR 50, General Design Criterion 4,
- "Requirements for Protection Against Dynamic Effects for Postulated Pipe Rupture" (Reference 1-7).
This report provides a fracture mechanics demonstration of primary loop integrity for the Salem Generating Stations consistent with the NRC position for exemption from consideration of dynamic effects.
Several computer codes are used in the evaluations. The main-frame computer programs are under Configuration Control which has requirements conforming to NRC' s Standard Review Plan 3.9.1 (Reference 1-8). The fracture mechanics calculations are independently verified (benchmarked).
1.5 References 1-1 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.
1-2 Letter from Westinghouse (E. P. Rahe) to NRC (R. H. Vollmer), NS-EPR-2768, dated May 11, 1983.
1-3 WCAP-9283, "The Integrity of Primary Piping Systems of Westinghouse Nuclear Power Plants During Postulated Seismic Events," March, 1978.
1-4 Letter Report NS-EPR-2519, Westinghouse (E. P. Rahe) to NRC D. G. Eisenhut),
Westinghouse Proprietary Class 2, November 10, 1981.
1-5 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated April 25, 1983.
1-6 Letter from Westinghouse (E. P. Rahe) to NRC (W. V. Johnston) dated July 25, 1983.
WPFI 080A: I b/042893 1-3
1-7 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. 207ffuesday, October 27, 1987/Rules and Regulations, pp. 41288-41295.
1-8 Nuclear Regulatory Commission, Standard Review Plan Section 3.9.1, "Special Topics for Mechanicai Component," NUREG-0800, Revision 2, July 1981.
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SECTION 2.0 OPERATION AND STABILITY OF THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Cracking The Westinghouse reactor coolant system 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., IGSCC, intergranular stress corrosion cracking). This operating history totals over 500 reactor-years, including five plants each having over 18 years of operation and 15 other plants each with over 13 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 established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking jn Pressurized Water Reactors (PWR's). The results of the study performed by the PCSG were esented in NUREG-0531 (Reference 2-1) entitled "Investigation and Evaluation of Stress orrosion 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 IGSC~
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
.rimary coolant systems.
- As stated above, for the Westinghouse plants there is no history of cracking failure in the reactor coolant system loop. The discussion below further qualifies the PCSG' s findings.
WPF1080A: 1 b/042893 2-1
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 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 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 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 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 WPF1080A: lb/042893 2-2
namic loads. To ensure dynamic system stability, reactor coolant parameters are stringently.
controlled. Temperature during normal operation is maintained within a narrow range by control rod position; 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 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. Preoperational testing and operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping are such that no significant water hammer can occur.
2.3 Low Cycle and High Cycle Fatigue Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue u_sage factor evaluation to show compliance with the rules of Section ID of the ASME Code. A further evaluation 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.
- gh 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 Salem. 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.
WPF1080A: 1 b/042893 2-3
SECTION 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 as-built outside diameter and minimum wall thickness of the pipe are 34.00 in. and 2.335 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, dead weight 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 I 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 is postulated to exist at this location which is subjected to both the normal loads and faulted loads to assess leakage and stability, respectively. The ads (developed below) 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 taking into consideration both fracture toughness and stress. The enveloping locations so determined are called toughness critical locations.
Two most critical locations (one from each unit) are identified after the full analysis is completed. Once loads (this section) and fracture toughnesses (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 locations for evaluation are those shown in Figure 3-2.
WPF1080A: lb/042893 3-1
3.2 Calculation of Loads and Stresses The stresses due to axial loads and bending moments are calculated by the following equation:
- where, cr
=
F
=
M
=
A
=
z
=
cr =
stress axial load bending moment pipe cross-sectional area section modulus F
A M
+_
z (3-1)
The bending moments for the desired loading combinations are calculated by the following equation:
- where, M
My Mz
=
=
=
bending moment for required loading Y component of bending moment Z component of bending moment (3-2)
The axial load and bending moments for leak rate predictions and crack stability analyses are computed by the methods to be explained in Sections 3.3 and 3.4.
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Loads for Leak Rate Evaluation The normal operating loads for leak rate predictions are calculated by the following equations:.
F
=
My =
Mz =
The subscripts of the above equations represent the following loading cases:
DW =
TH
=
p
=
dead weight normal thermal expansion load due to internal pressure This method of combining loads is often referred as the algebraic sum* method.
(3-3)
(3-4)
(3-5)
The loads based on this method of combination are provided in Table 3-1 at all the locations entified in Figure 3-2. The as-built dimensions are also given.
3.4 Load Combination for Crack Stability Analyses 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 °"12 to 1.0. The absolute summation of loads are shown in the following equations:
F = I Fnw I + I Fm I + I Fp I + I FssEINERTIA I + I FssEAM I (3-6)
My= I (My)nw I+ I (My)rn I+ I (My)p I+ I (My)ssEINERTIA I+ I (My)ssEAM I (3-7)
Mz = I CMz)nw I + I (Mz)rn I + I (Mz)p I + I (Mz)ssEINERTIA I + I (Mz)ssEAM I
(~-8) where subscripts SSE, INERTIA and AM mean safe shutdown earthquake, inertia and anchor motion, respectively.
Ae loads so determined are used in the fracture mechanics evaluations (Section 7.0) to
~~onstrate the LBB margins at the locations established to be the governing locations.
These loads at all the locations of interest (see Figure 3-2) are given in Table 3-2.
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3.5 References 3-1 USNRC Standard Review Plan 3.6.3, Leak-Before-Break Evaluation Procedures, NUREG-0800.
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Table 3-1 Dimensions, Normai Loads and Normal Stresses for Salem Outside Minimum Axial Bending Total Diameter Thickness Loadh Moment Stress Locationa (in.)
(in.)
(kips)
(in-kips)
(ksi) 1 34.00 2.335 1482 28726 23.06 2
34.00 2.335 1502 8235 11.25 3
34.00 2.335 1340 17114 15.71 4
36.32 2.495 1657
. 2949 7.65 5
36.32 2.495 1658 2647 7.52 6
36.32 2.495 1652 4599 8.42 7
36.32 2.495 1704 1441 7.11 8
36.32 2.495 1704 3008 7.86 9
36.32 2.495 1842 9853 11.64 t::910 32.26 2.215 1426 8633 12.69 11 32.26 2.215 1414 7883 12.12 12 32.26.
2.215 1415 8048 12.24 a See Figure 3-2 b Includes pressure WPF1080A: lb/042893 3-5
a b
c Locationa,b 1
2 3
4 5
6 7
8 9
10 11 12 See Figure 3-2 See Table 3-1 for dimensions Includes pressure WPF1080A: lb/042893 Table 3-2 Faulted Loads and Stresses for Salem Axial Loadc Bending Moment Total Stress (Kips)
(in-Kips)
(ksi) 1674 31146 25.30 1653 9575 12.68 1826 21098 20.11 1863 4332 9.09 1862 3817 8.84 1857 5437 9.59 1791 2329 7.86 1792 4312 8.81 1871 11619 12.59.
1469 9925 13.77
. 1458 8683 12.88 1460 9699 13.58 3-6
I I
L-------------------------------
L Thickness = t F
~
M
- 1.
Crack OD ooa = 34.00 in ta = 2.335 in
~ ormal Loadsa Faulted Loadsb forcec:
1482 kips bending moment: 28726 in-kips forcec:
1674 kips bending moment: 31146 in-kips a See Table 3-1
~ See Table 3-2
~ Includes the force due to a pressure of 2250 psi Figure 3-1 Hot Leg Coolant Pipe WPF0954A: lb/042793 3-7
.1
- , ~-
HOT LEG HOT LEG Reactor Pressure Vessel Temperature 611°F, Pressure: 2250 psi CROSSOVER LEG Temperature 545°F, Pressure: 2250 psi COLD LEG Temperature 545°F, Pressure: 2305 psi COLD LEG REACTOR COOLANT PUMP CROSSOVER LEG
--© Figure 3-2 Schematic Diagram of Salem Primary Loop Showing Weld Locations WPF0954A: l b/042393 3-8
SECTION 4.0 MATERIAL CHARACTERIZATION 4.1 Primary Loop Pipe and Fittings Materials The primary loop pipe materials are A376 TP316, and the elbow fittings are A351 CF8M.
4.2 Tensile Properties The Pipe Certified Materials Test Reports (CMTRs) for Salem were used to establish the tensile properties for the leak-before-break analyses. The CMTRs include tensile properties at room temperature and at 650°F for each of the heats of material. These properties are given in Table 4-1 for Salem Unit 1 and in Table 4-2 for Salem Unit 2.
For the A376 TP316 material, the representative properties at 611°F and 545°F were established from the tensile properties at 650°F temperature given in Tables 4-1 and 4-2 by "lizing Section III of the 1992 ASME Boiler and Pressure Vessel Code. Code tensile properties at 611°F and 545°F were obtained by interpolating between the 500°F, 600°F and 650°F tensile properties. Ratios of the code tensile properties at 611°F and 545°F to the corresponding tensile properties at 650°F temperature were then applied to the 650°F tensile properties given in Tables 4-1 and 4-2 to obtain the plant specific properties for A376 TP316 at 611°F and 545°F.
The Elbow Fittings Certified Materials Test Reports (CMTRs) for Salem were used to establish the tensile properties for the leak-before-break analyses. The CMTRs for elbow fittings include tensile properties at room temperature for each of the heats of material. These properties are given for Salem in Table 4-3.
For the A351 CF8M material, the representative properties at 611°F and 545°F were established from the tensile properties at room temperature given in Table 4-3 by utilizing Section III of the* 1992 ASME Boiler and. Pressure Vessel Code. Code tensile properties at 611 °F and 545°F were established by interpolating between the room temperature and the 650°F tensile properties. Ratios of the code tensile properties at 611°F and 545°F to the corresponding properties at room temperature were then applied to the room temperature
- .operties given in Table 4-3 to obtain the plant specific representative properties for A351 CF8M. at 611°F and 545°F.
WPFlOSOA:lb/042893.
4-1 1
The average* and lower bound yield strengths and ultimate strengths are given in Table 4-4.
The ASME Code moduli of elasticitY are also given, and Poisson's ratio was taken as 0.3.
For leak-before-break fracture evaluations of the toughness critical locations the true stress-true strain curves for A351 CF8M at 545°F must be available. This curve was obtained using the Nuclear Systems Materials Handbook (Reference 4-1). The lower bound true stress-true strain curve is given in Figures 4-1.
4.3 Fracture Toughness Properties The pre-service fracture toughnesses of both forged and cast stainless steels* of interest here have in terms of J1c been found to be very high at 600°F. Typical results for a cast material are given in *Figure 4-2. J1c is observed to be over 2500 in-lbs/in2* Forged materials are* even higher. However, cast stainless steels are subject to thermal aging during service. This thermal aging causes an elevation in the yield strength of the material and a degradation of the fracture toughness, the degree of degradation being somewhat proportional to the level of ferrite in the material.
To determine the effects of thermal aging on piping integrity, a detailed study was carried out in Reference 4-2. In that report, fracture toughness results were presented for a material
[
]a,c,e The effects of the aging process on the end-of-service life fracture toughness are further discussed in Appendix B.
End-of-service life toughnesses for the heats are established using the alternate toughness criteria methodology of Reference 4-5 (Appendix B). By that methodology a heat of material is said to be as good as [
y.c,e if it can be demonstrated that its end-of-service fracture toughnesses equal or exceed those of [
]a,c,e. The worst case fracture toughness values WPF1080A: 1 b/042893 4-2
r Salem Units 1 and 2 loops at critical locations, as taken from Appendix B, are given in.
able 4-5.
Available data on aged stainless steel welds (References 4-2 and 4-3) indicate that Jlc values for the worst case welds are of the same order as the aged material. However, the slope of the J-R curve is steeper, and higher J-values have been obtained from fracture tests (in excess of 3000 in-lb/in2). 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 stress for the weld materials is much higher at the temperaturea. Therefore, weld regions are less limiting than the cast material.
It is thus conservative to choose the end-of-service life toughness properties of [
r,c,e as representative of those of the welds. Also, such pipes and fittings have an end-of-service life calculated room temperature Charpy U-notch energy, (KCU), greater than that of [ *
]a,c,e are also conservatively assumed to have the properties of [
]a,c,e.
In fracture mechanics analyses that follow, the fracture toughness properties given in able 4-5 will be used as the criteria against which the applied fracture toughness values will e compared.
Forged stainless steel piping such as A376 TP316 does not degrade due to thermal aging.
Thus fracture toughness values well in excess of that established for the cast material and welds exist for this material throughout service life and are not limiting.
4.4 References 4-1 Nuclear Systems Materials Handbook, Part I - Structural Materials, Group 1 - High Alloy Steels, Section 2, ERDA Report TID 26666, November, 1975.
4-2 WCAP-10456, "The* Effects of Thermal Aging on the Structural Integrity of Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2, November 1983.
a In the report all the applied J values were conservatively determined by using base.
metal strength properties.
WPF1080A: 1 b/042893 4-3
4-3 Slama, G., Petrequin, P., Masson, S.H., and Mager, T.R., "Effect of Aging on Mechanical Properties of Austenitic Stainless Steel Casting and Welds", presented at Smirt 7 Post Conference Seminar 6 - Assuring Structural Integrity of Steel Reactor.
Pressure Boundary Components, August 29/30, 1983, Monterey, CA.
4-4 Appendix II of Letter from Dominic C. Dilanni, NRC to D. M. Musolf, Northern States Power Company, Docket Nos. 50-282 and 50-306, December 22, 1986.
4-5 Witt, F.J., Kim, C.C., "Toughness Criteria for Thermally Aged Cast Stainless Steel,"
WCAP-10931, Revision 1, Westinghouse Electric Corporation, July 1986, (Westinghouse Proprietary Class 2).
4-6 ASME Boiler and Pressure Vessel Code 1992,Section III.
WPF1080A: lb/042893 4-4
Table 4-1 Measured Tensile Properties for Salem Primary Loop Piping Unit-1 Yield Ultimate Yield Ultimate Strength Strength Strength Strength Heat No.
(psi)
(psi)
(psi)
(psi)
Material Room Temp.
Room Temp.
650°F Temp.
650°F Temp.
F0212 43,500 85,100 23,400 67,400 A376 TP316 43,200 88,000 F0212 38,500 81,600 NIA NIA A376 TP316 42,500 83,600 F0221 44,000 83,900 21,600 65,200 A376 TP316 42,700 86,000 F0213 44,000 86,000 23,700 69,800 A376 TP316 41,000 83,000 F0215 41,000 82,300 21,700 66,800 A376 TP316 46,900 83,700 D8776 35,000 76,700 22,400 63,400 A376 TP316 35,300 77,400 D8776 33,400 75,900 NIA NIA A376 TP316 34,500 77,400 F0225 44,000 86,900 22,100 68,200 A376 TP316 48,500 88,900 D8785 33,400 75,700 22,000 61,000 A376 TP316 33,600 75,300 F0190 42,000 88,800 21,300 58,200 A376 TP316 43,000 85,000 F0222 41,000 82,700 21,500 66,400 A376 TP316 44,500 87,200 F0226 42,200 86,400 21,500 66,200 A376 TP316 45,500 91,400 WPF1080A: lb/042893 4-5
Table 4-1 (cont)
Measured Tensile Properties Salem Primary Loop Piping Unit-1 Yield Ultimate Yield Ultimate Strength Strength Strength Strength Heat No.
(psi)
(psi)
(psi)
(psi)
Material Room Temp.
Room Temp.
650°F Temp.
650°F Temp.
F0382 43,000 85,300 25,300 72,600 A376 TP316 43,700 86,800
. F0369 42,300 89,500 21,700 69,400 A376 TP316 43,700 89,800 F0369 43,000 89,100 26,000 71,400 A376 TP316 47,700 93,700 F0369 46,500 88,400 26,000 71,400 A376 TP316 48,400 91,500 V0630 36,000 78,600 23,400 61,300 A376 TP316 49,400 91,900 F0371 44,300 90,900 25,600 72,300 A376 TP316 48,000 92,700 K2011 33,100 75,200 20,600 56,200 A376 TP316 33,000 75,100 K2011 34,900 75,100 20,600 56,200 A376 TP316 34,300 75,100 WPF1080A: 1 b/042893 4-6
Table 4-2 Measured Tensile Properties for Salem Primary Loop Piping Unit-2 Yield
- Ultimate Yield Ultimate Strength Strength Strength Strength Material Heat No.
(psi)
(psi)
(psi)
(psi)
Room Temp.
Room Temp.
650°F Temp.
650°F Temp.
11681 43,700 82,400 24,600 66,900 A376 TP316 43,700 86,400 11678 43,700 82,400 24,900 67,200 A376 TP316 39,900 84,400 11679 38,700 82,900 24,000 67,400 A376 TP316 42,400 85,600 11676 38,000 80,200 NIA NIA A376 TP316 42,100 84,600 11676 36,400 78,200 21,700 65,800 A376 TP316 36,500 76,500 11677 40,500 84,200 24,900 70,600 A376 TF316 41,700 84,200 11677 38,500 80,600 NIA NIA A376 TP316 44,200 88,200 12337 42,200 83,600 27,100 72,200 A376 TP316 44,900 87,400 K2204 44,700 86,100 29,100 71,800 A376 TP316 52,900 95,100 12337 44,900 85,100 27,100 72,200 A376 TP316 46,300 86,900 K2205 43,700 87,400 25,900 73,400 A376 TP316 48,900*
92,400 WPFl 080A: 1 b/042893 4-7
Table 4-2 (cont)
Measured Tensile Properties for Salem Primary Loop Piping Unit-2 Yield Strength Ultimate Yield Ultimate Strength Strength Strength Heat No.
(psi)
(psi)
(psi)
(psi)
Material Room Temp.
Room Temp.
650°F Temp.
650°F Temp.
K2205 45,900 87,900 NIA NIA A376 TP316 46,200 89,900 K2205 44,400 86,100 25,900 73,400 A376 TP316 47,400 89,900 K2205 43,700 83,400 22,700 70,400 A376 TP316 44,900 83,900 J1682 43,900 86,600 25,600 71,300 A376 TP316 51,100 78,600 J2334 40,000 82,100 22,500 67,000 A376 TP316 43,200 87,900 J2741 43,100 85,900 NIA NIA A376 TP316 45,100 88,400 J2741 44,500 86,900 24;100 66,100 A376 TP316 44,000 86,100 J2741 42,000 84,900 NIA NIA A376 TP316 43,400 87,300 J2741 41,400 84,900 21,300 67,100 A376 TP316 44,900 87,400 WPF1080A: lb/042893 4-8
Table 4-3 Measured Room Temperature Tensile Properties for Salem Primary Loop Elbow Fittings (Unit-1)
Heat No.
Yield Strength (psi)
Ultimate Strength (psi)
Material Room Temperature Room Temperature 33418-3 40200 84100 A351 CF8M 33418-2 42000 85250 A351 CF8M 33049-2 39300 82000 A351 CF8M 32929-2 43500 85000 A351 CF8M 11632-1 37500 75000 A351 CF8M 11708-1 36000 75250 A351 CF8M 12236-1 40500 83500 A351 CF8M 11743-1 43500 85000 A351 CF8M 11086-1 42000 84500 A351 CF8M 11777-4 40500 81500 A351 CF8M r-.12049-2 40500 79000 A351 CF8M 11480-1 45000 85500 A351 CF8M 12511-1 46500 85500 A351 CF8M 12393-:-5 40500 84000 A351 CF8M 12735-1 45000 85500 A351 CF8M 12836-1 45000 81000 A351 CF8M 13133-1 45000 85250 A351 CF8M I
14706-2 42000 77500 A351 CF8M 14619-1 45000 85000 A351 CF8M 11671-1 46500 88500 A351 CF8M 11851-1 40500 83000 A351 CF8M 11829-2 45000 85500 A351 CF8M 12049-1 42000 79250 A351 CF8M 12121-1 45000 84000 A351 CF8M 11086-2 42000 85500 A351 CF8M 11974-2 37000 85500 A351 CF8M 11126-2 44250 86000 A351 CF8M WPFl 080A: 1 b/042893 4-9
Heat No.
36803 37658 38690 39125 40874 39405 37429 39445 39716 36896 38245 52282-4 52162-1 51800-1 63549-1 & 2 50115 49359 49686 50362 63669-1 50594 49165 49083 52870-1 52449-2 52162-2 51800-2 Table 4-3 (cont)
Measured Room Temperature Tensile Properties for Salem Primary Loop Elbow Fittings (Unit-2)
Yield Strength Ultimate Strength (psi)
(psi)
Room Temperature Room Temperature 47000 86500 49500 84250 34900 75000 43200 84900 42500 85400 40200 79900 45660 84480 39250 84300 45300 86800 40600 82700 36500 75000 43800 85800 42600 85750 40200 83250 45300 90600 42200 85950 53500 95500 46250 90050 41100 85250 44250 84750 38350 80300 42800 85600 44900 88000 38350 79800 42750 87500 41500 86550 39300 83450 WPF1080A: I b/042893 4-10 Material A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M
- A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M A351 CF8M
Table 4-4 Mechanical Properties for Salem Materials at Operating Temperatures Material Temperature Average Yield Lower Bound (oF)
Strength (psi)
Yield Stress Ultimate (psi)
Strength (psi) a,c,e Modulus of Elasticity E = 25.245 x 106 psi. at 611°F E = 25.575 x 106 psi. at 545°F Poisson's ratio: 0.3 WPF1080A: 1 b/042893 4-11
[
Table 4-5 Fracture Toughness Properties for Salem Primary Loops for Leak-Before-Break Evaluation at Critical Locations Location<a)
Heat KCU lie Tmat Jmax Comment-No.
(daj/cm2)
(in-Ib/in2)
(non-dim)
(in-lb/in2).
s (a) The locations are shown in Figure 3-2 WPF1080A: lb/042893 4-12 a,c,e J
Figure 4-1 WPF1080A: I b/042893 Representative Lower Bound True Stress - True Strain Curve for A351 CFSM at 545°F 4-13 a,c,e
a,c,e Figure 4-2 J vs. Aa for SA351 CFSM Cast Stainless Steel at 600°F WPF1080A: lb/042893 4-14
a,c,e J,?igure 4-3 J Vs...1a at Different Temperatures for Aged Material
[
]a,c,e (7500 Hours at 400°C)
WPF1080A: 1 b/042893 4-15
SECTION 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 (governing locations). Candidate locations are designated load critical locations or
- toughness critical locations as discussed in Section 3.0. Such locations are established based on the loads (Section 3.0) and the material properties established in Section 4.0. These locations are defined below for Salem. Table 3-2 as well as Figure 3-2 are used for this evaluation.
Load Critical Locations The highest stressed location for the A376 TP316 straight pipes is Location 1 (See Figure 3-2) at the reactor vessel outlet nozzle to pipe weld. Furthermore, since it is on a straight pipe, it is a high toughness location.
-ughness Critical Locations Low toughness locations are at the ends of every elbow. All the elbows except indicated below, exceed the toughness of [
].a.c,e In the case of.Unit 1 low toughness values are at the end of 90 degree elbow in the* cross-over leg. These are locations 6 and 7 (See Figure 3-2 for locations).
Location 6 governs since it has higher stress than Location 7. In the case of Unit 2 low toughness location is at the end of 90 degree elbow with splitter. These are locations 8 and 9. Location 9 governs since it has higher stress than Location 8. It is thus concluded that the enveloping locations are 6 (Unit-1) and 9 (Unit-2). The allowable toughness 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 (i.e. Japp) and tearing modulus (Tapp) are compared are:
(1)
If Japp < Jlc' then the crack is stable; WPFI 080A: 1 b/042893 5-1
(2)
If Japp> he* but, if Tapp< Tmat and Japp < Jmax' then the crack is stable.
These criteria apply to the toughness critical locations. For load critical locations, the limit load method discussed in Section 7.0 is used.
WPF1080A: 1 b/042893 5-2
SECTION 6.0 LEAK RA TE 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 [ t*c,e, both [
6.3 Calculation Method The basic method used in the leak rate calculations is the method developed by [
]
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, Pc, for the primary loop enthalpy condition and an assumed flow. Once Pc was found for a given mass flow, the [
]a,c,e was found from Figure 6-2 (taken from Reference 6-1). For all cases considered, since [
Therefore, this method will yield the two-phase pressure drop due to momentum effects as illustrated in Figure 6-3. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using WPFI 080A: 1 b/042893 6-1
Af>f = [
(6-1) where the friction factor f is determined using the [
]a,c,e The crack relative roughness, e, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [
The frictional pressure drop using equation 6-1 is then calculated for the assumed flow rate and added to the [
]a,c,e to obtain the total pressure drop from the primary system to the atmosphere. That is, for the primary loop Absolute Pressure - 14. 7 = [
(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 turn 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 Reference 6-2 and the leak rates were calculated using the two-phase flow formulation described above. The average material properties of Section 4.0 were used for these calculations.
The flaw sizes to yield a leak rate of 10 gpm were calculated at the governing locations and are given in Table 6-1. The flaw sizes so determined are called leakage flaws.
- The Salem Generating Stations RCS pressure boundary leak detection system meets the intent of Regulatory Guide 1.45. Thus, to satisfy the margin of 10 on the leak rate, the flaw sizes (leakage flaws) are determined which yield a leak rate of 10 gpm.
WPFl 080A: 1 b/042893 6-2
.5 References 6-1
[
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 II-1, NUREG/CR-3464, September 1983.
WPF1080A: 1 b/042893 6-3 i
WPFl 080A: 1 b/042893 Table 6-1 Flaw Sizes Yielding a Leak Rate of 10 gpm at the Governing Locations 6-4 a,c,e
N-* -
- = -
Figure 6-1 WPF0954A: 1 b/042393 a,c,e STAGNATION ENTHALPY I 1o2 Btullbt Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 6-5
a,c,e 0 -
LENGTH/DIAMETER RATIO (L/O, Figure 6-2
. ]a.c.e Pressure Ratio as a Function of LID
- WPF0954A: lb/042393 6-6
a,c,e a,c,e Figure 6-3 Idealized Pressure Drop Profile Through a Postulated Crack WPF0954A: l b/042393 6-7
SECTION 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-tip 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 loaci, the calculated J-integral value is shown to be less.
than the Jlc 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 Tapp= -
da a; applied tearing modulus modulus of elasticity 0.5 (cry + au) (flow stress) crack length yield and ultimate strength of the material, respectively Stability is said to exist when ductile tearing occurs if Tapp is less than Tmat* the experimentally determined tearing modulus. Since a constant Tmat is assumed a further restriction is placed in Japp* Japp must be less than Jmax where Jmax is the maximum value of J for which the experimental Tis greater than or equal to the Tmat used.
As discussed in Section 5.2 the local crack stability will be established by the two-step criteria:
(1) If J
< JI, then the crack is stable.
app c
(2) If Japp > Jlc' but, if Tmat < Tmat and Japp < Jmax* then the crack is stable.
WPF1080A: 1 b/042893 7-1
7.2 Global Failure Mechanism Detennination of the conditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of defonnation accompanying fracture. One method for predicting the failure of ductile material is the plastic instability*
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 fonned. The stress level at which this occurs is tenned 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 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 internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:
[
]a,c,e where:
[
t,c,e crf
=
0.5 (cry + cru) (flow stress), psi
[
[
WPF1080A: 1 b/042893 7-2
e analytical model described above accurately accounts for the piping internal pressure as ell 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 Stability analyses were performed at the critical locations established in Section 5.1. The elastic-plastic fracture mechanics (EPFM) I-integral analyses for through-wall circumferential cracks in a cylinder were performed using the procedure in the EPRI fracture mechanics handbook (Reference 7-2).
The lower-bound material properties of Section 4.0 were applied (see Table 4-4). The fracture toughness properties established in Section 4.3 and the normal plus SSE loads given Table 3-2 were used for the EPFM calculations. Evaluations were performed at the toughness critical locations identified in Section 5.1. The results of the elastic-plastic fracture mechanics I-integral evaluations are given in Table 7-1. The leakage size flaws are presented on the same table.
The critical locations were also identified in Section 5.1. A stability analysis based on limit load was petformed for these locations as described in Section 7.2. The welds at these locations are SMAW or SAW welds. The welds at locations l and 9 are SMAW welds. The weld at Location 6 is SAW weld. The "Z" factor correction for SMAW and SAW were applied (Reference 7-3) as follows:
Z = 1.15 [l.O + 0.013 (00-4)]
z = 1.30 [1.0 + 0.010 (00-4)]
(For SAW weld) where OD is the outer diameter of the pipe in inches.
The Z-factors were calculated for the load critical locations, using the dimensions given in Table 3-1. The Z factors were 1.59, 1.72, and 1.63 for locations 1, 6 and 9 respectively. The
.plied loads were increased by the Z factors 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.
WPF0954A: l b/050493 7-3
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.
Kumar, V., German, M. D. and Shih, C. P., "An Engineering Approach* for Elastic-Plastic Fracture Analysis," EPRI Report NP-1931, Project 1237-1, Electric Power Research Institute, July 1981.
7-3.
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.
WPF1080A: 1 b/042893 7-4
Table 7-1 Stability Results for Salem Based on Elastic-Plastic J-Integral Evaluations Fracture Criteria Calculated Values Flaw Size Jlc Tmat Jmax Location (in)
(in-lb/in2)
(in-lb/in2)
Japp Tapp (in-lb/in2)
WPFl 080A: 1 b/042893 7-5 a,c,e
Table 7-2 Stability Results for Salem Based on Limit Load Location WPF1080A: 1 b/042893 Flaw Size (in.)
7-6 Leakage Flaw Size (in.)
a,c,e
a,c,e
[
]a,c,e Stress Distribution WPF1080A: lb/042893 7-7
OD = 34.00 in t = 2.335 in cr). = 20.86 ksi
<Ju = 56.20 ksi FJ = 1674 kips Mb = 31146 in-kips A376 TP316 Material With SMAW Weld Figure 7-2 Critical Flaw Size Prediction
- Hot Leg at Location 1 WPFO\\i 54A: 1 b/042393 7-8 a,c;e
OD = 36.32 in t = 2.495 in cry = 22.57 ksi (ju = 71.78 ksi Fa= 1857 kips Mb = 5437 in-kips A351 CF8M Material With SAW Weld Figure 7-3 Critical Flaw Size Prediction - Cross-over Leg at Location 6 WPF0954A: lb/042393 7-9 a,c,e
__J
OD = 36.32 in t = 2.495 in cry = 22.57 ksi au= 71.78 ksi Fa= 1871 kips Mb= 11619 in-kips A351 CF8M Material With SMAW Weld Figure 7-4 Critical Flaw Size Prediction - Cross-over Leg at Location 9 WPF0954A: 1 b/042393 7-10 a,c,e
SECTION 8.0 FATIGUE CRACK GROWTH ANALYSIS To detennine the sensitivity of the primary coolant system to the presence of small cracks, a fatigue crack growth analysis was carried out for the [
]a,c,e region of a typical system (see Location [. ]a,c,e of Figure 3-2). This region was selected because crack growth calculated here will be typical of that in the entire primary loop. Crack growths calculated at other locations can be expected to show less than l 0% variation.
A[
]a,c,e of a plant typical in geometry and operational characteristics to any Westinghouse PWR System.
[
]a,c,e All nonnal, upset, and test conditions were considered. A summary of the applied transients is provided in Table 8-1. Circumferentially oriented
- urface flaws were postulated in the region, assuming the flaw was located in three different ations, as shown in Figure 8-1. Specifically, these were:
Cross Section A: [
Cross Section B: [
Cross Section C: [
Fatigue crack growth rate laws were used [
1a,c,e The law for stainless steel was derived from Reference 8-1, with a very conservative correction for the R ratio, which is the ratio of minimum to maximum stress during a transient. For stainless steel, the fatigue crack growth fonnula is:
da = (5.4 x 10-12) K ff4*48 inches/cycle dn e
K
= K (1-R)0.5 eff max R= K. /K mm max WPFl 080A: 1 b/042893 8-1
.1
[
[
where:
[
]a,c,e where AK is the stress intensity factor range.
The calculated fatigue crack growth for semi-elliptic surface flaws of circumferential orientation and various depths is summarized in Table 8-2, and shows that the crack growth is very small, [
]a,c,e 8.1 References 8-1 Bamford, W. H., "Fatigue Crack Growth of Stainless Steel Piping in a Pressurized Water Reactor Environment," Trans. ASME Journal of Pressure Vessel Technology, Vol. 101, Feb. 1979.
8-2
[
]a,c,e 8-3
[
]a,c,e WPFlOSOA:lb/042893.
8-2
Table 8-1 Summary* of Reactor Vessel Transients Number Typical Transient Identification Number of Cycles Normal Conditions 1
Heatup and Cooldown at 100°F/hr 200 (pressurizer cooldown 200°F/hr) 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 106 Upset Conditions Loss of load, without immediate turbine 80 or reactor trip 7
Loss of power (blackout with natural 40 circulation in the Reactor Coolant System) 8 Loss of Flow (partial loss of flow, one 80 pump only) 9 Reactor trip from full power 400 Test Conditions 10 Turbine roll test 10 11 Hydrostatic test conditions Primary side 5
Primary side leak test 50 12 Cold Hydrostatic test 10 WPF1080A: 1 b/042893 8-3
__J
Table 8-2 Typical Fatigue Crack Growth at
[
]a,c,e (40 years)
FINAL FLAW (in.)
Initial Flaw (in.)
[
]a,c,e
[
]a,c,e
[
]a,c,e 0.292 0.31097 0.30107 0.30698 0.300 0.31949 0.30953 0.31626 0.375 0.39940 0.38948 0.40763 0.425 0.45271 0.4435 0.47421 WPF1080A: 1 b/042893 8-4
a,c,e Figure 8-1 Typical Cross-Section of [
WPF1080A:lb/042893.
8-5
Figure 8-2 WPF1080A: lb/042893 Reference Fatigue Crack Growth Curves for [
]a,c,e 8-6 a,c,e
a,c,e Figure 8-3 Reference Fatigue Crack Growth Law for [
]a,c,e in a Water Environment at 600°F WPF1080A: I b/042893 8-7
SECTION
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.
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
- 2.
- 3.
- 10 gpm (the leakage flaw).
Leak Rate - A margin of 10 exists between the calculated leak rate from the leakage flaw and the leak detection capability of 1 gpm.
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 size flaw is stable).
WPF1080A: 1 b/042893 9-1
Table 9-1 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Salem Generating Station Units 1 and 2 Location Leakage Flaw Size Critical Flaw Size Margin a
based on limit load b
based on J integral evaluation WPFl 080A: l b/042893 9-2 a,c,e
SECTION
10.0 CONCLUSION
S This report justifies the elimination of RCS primary. loop pipe breaks from the structural design basis for the Salem Generating Station Units 1 and 2 *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 because of system design, testing, and operational considerations.
- c.
The effects of low and high cycle fatigue on the integrity of the primary piping are negligible.
- d.
Ample margin exists between the leak rate of small stable flaws and the capability of the Salem reactor coolant system pressure boundary Leakage Detection System.
- e.
Ample margin exists between the small stable flaw sizes of item d and larger stable flaws.
- f.
Ample margin exists in the material properties use.d to demonstrate end-of-service life (relative to aging) stability of the critical flaws.
For the critical locations flaws are identified that will be stable because of the ample margins described m d, e, and f above.
Based on the above, it is concluded that dynamic effects of RCS primary loop pipe breaks need not be considered in the structural design basis of the Salem Generating Station Units 1 and 2.
WPF1080A: lb/042893 10-1
[
WPF1080A: lb/042893 APPENDIX A L'1TMOMENT A-1
a,c,e Figure A-1 Pipe with a Through-: Wall Crack in Bending WPFl 080A: lb/042893 A-2
APPENDIX B TOUGHNESS CRITERIA FOR SALEM CAST PRIMARY LOOP COMPONENTS All of the individual cast piping components of the Salem primary loops, do not satisfy the original [
]a,c,e criteria (Reference 4-5). [
WPFl 080A: 1 b/042893 B-1
WPFl 080A: 1 b/042893 Table B-1 Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 1 B-2 a,c,e
WPF1080A: 1 b/042893 Table B-1 (continued)
Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 1 B-3 a,c,e
WPF1080A:lb/042893 Table B-1 (continued)
Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 1 B-4 a,c,e
WPF1080A: 1 b/042893 Table B-2 Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 2 B-5 a,c,e
*-~-* --*
WPFl 080A: 1 b/042893 Table B-2 (continued)
Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 2 B-6 a,c,e
WPFl 080A: 1 b/042893 Table B-2 (continued)
Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 2 B-7 a,c,e
WPF0954A: lb/042793 Table B-2 (continued)
Chemistry and Fracture Toughness Properties of the Material Heats of Salem Unit 2 B-8 a, C, J
a,c,e WPF0954A: 1 b/030893 B-9