ML20092J357
| ML20092J357 | |
| Person / Time | |
|---|---|
| Site: | Comanche Peak  |
| Issue date: | 01/31/1992 |
| From: | Adamonis D, Lee Y, Witt F WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19344C380 | List: |
| References | |
| TXX-92074, WCAP-13168, NUDOCS 9202240193 | |
| Download: ML20092J357 (84) | |
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t W stinghouse Proprietary Class 3 WCAP-13168 i *, 1 TECHNICAL JUSTIFICATION FOR ELIMINATING 10" ACCUMULATOR LINES RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR THE COMANCHE PEAK NUCLEAR PLANT UNIT 2 January 1997 Y. S. Lee J. C. Schmertz
, S. A. Swamy
- Verified by: /r4' F#~J. Witt
'~
')
Approved by - 1 D. C. Adamonis, Manager Materials, Mechanics and Diagnostic Technology Work Performed Under Shop Order: JJTP-5600 WESTINGHOUSE ELECTRIC CORPORATION
. Nuclear and Advanced Technology Division P.O. Box 2728
- Pittsburgh, Pennsylvania 15230-2728 mm.
l' e 1991 Westinghouse Electric Corp. All Rights Reserved WPF1082J/012492:10 l
TABLE OF CONTENTS 1 Section Title EiLqe
1.0 INTRODUCTION
1.1 Background 1-1 1.2 Scope and Objective 1-1 1.3 References 1-3 l 2.0 OPERATION AND STABILITY OF THE ACCUMULATOR LINE AND THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Cracking 2-1 2.2 Water Hammer 2-3 2.3 Low Cycle and High Cycle Fatigue 2-3 2.4 Potential Degradation During Service 2-4 2.5 References 2-4 3.0 MATERIAL CHARACTERIZATION 3.1 Pipe and Weld Materials 3-1 3.2 Material Properties of the Accumulator Lines 3-1 3.3 Tensile Properties of the injection Nozzles 3-2 3.4 Fracture Toughness Properties of the Injection 3-2 Nozzles 3.5 References 3-3 4.0 LOADS FOR FRACTURE' MECHANICS ANALYSIS 4.1 Nature of Loads 4-1
. 4.2 Loads for Crack Stability Analysis 4-2 4.3 Loads for Leak Rate Evaluation 4-2 4.4 Summary of Loads and Geometry 4-3 4.5 Governing Locations 4-3 5.0 FRACTURE MECHANICS EVALUATION 5.1 Failure Mechariism 5-1
', 5.2 Leak Rate Predictions 5-2 5.2.1 General Considerations 5-2 5.2.2 Calculation Method 5-2 WPF1082J/012492:10 iii
TABLE OF CONTENTS (Cr..t'd.) Section Ij_tle EAq.e 5.2.3 Leak Rate Calculations 5-4 5.2.4 Leak Detection Capability 5-4 5.3 Stability Evaluation of the Accumulator-Lines 4 5.4 Local Stability Evaluation of the Injection 5-5 Nozzles 5.5 References 5-6 6.0 ASSESSMENT OF FATIGUE CRACK GROWTH 6-1 7.0 ASSESSMENT OF MARGINS 7-1
8.0 CONCLUSION
S 8 > APPENDIX A Limit Moment A-1 APPENDIX B Fatigue Crack Growth Considerations B-1 B.1 Thermal Transient Stress Analysis B1 B.I 1 Critical location for Fatigue Crack Growth Analysis B-1 B.I.2 Design Transients B2 B.l.3 Simplified Stress Analysis B-2 B.I 4 Non linear Stress Distribution for Severe Transients B-5
. B.1.5 Total Stress for Fatigue Crack Growth B-6 B.2 Fatigue Crack Growth Analysis B-6 B.2.1 Analysis Procedure B-6 B.2.2 Results B-9 B.3 References B-10 9
J e WPF1082J/012492:10 iV I
LIST OF TABLES Table Title hgg 3-1 Room Temperature Mechanical Properties of the Accumulator Line Materials for Loop 1 of the Comanche Peak Unit 2 Nuclear Power Plant - Unit 2 3-5 s 3-2 Room Temperature Mechanical Properties of the Accumulator Line Materials for Loop 2 of the Comanche Peak Unit 2 Nuclear Power Plant - Unit 2 3-6 3-3 Room Temperature Mechanical Properties of the Accumulator Line Materials for Loop 3 of the Comanche Peak Unit 2 Nuclear Power Plant - Unit 2 3-7 3-4 Room Temperature Mechanical Properties of the Accumulator Line Materials for Loop 4 of the Comanche Peak Unit 2 Nuclear Power Plant - Unit 2 38 3-5 Room Temperature ASME Code Minimum Properties 3-9 3-6 Representative Tensile Properties for the Comanche Peak Unit 2 Nuclear Power Plant 10" Accumulator Lines 3-10 3-7 Modulus of Elasticity (E) 3-11 3 - 8 .. Available Mechanical Properties of the Accumulator Injection Nozzles at 650'F 3-12 3-9 . Chemistry and End of Service Life KCU Toughness
~
for the Accumulator Injection Nozzles 3-13
'. 4-1 Summary of Normal and Faulted Loads and Stresses at Governing Locations 4-4 WPF1082J/012492:10 v
I LIST OF TABLES (Cont'd.) Table ILtle flag 5-1 Leak Rate Crack lengths for the Governing Locations of the Comanche Peak Unit 210" Accumulator Lines 5-7 5-2 Summary of Critical Flaw Sizes for the Gove.ning Locations of the Comanche Peak Unit 2 10" Accumulator Lines 5-8 7-1 Leakage Flaw Sizes, Critical Flaw Sizes, and Margins 7-2 7-2 LBB Conservatisms 7-3 B-1 Thermal Transients Considered for Fatigue Crack B-10 Growth Evaluation B-2 Transient Stresses for Accumulator Line B-ll i B-3 Envelope Normal Loads 8-12 B-4 Accumulator Line Fatigue Crack Growth Results B 13 l l I I O m WPF1082J/012492:10 Vi
LIST Of FIGURES Fioure Title Eagg 3-1 Layout of the Accumulator Line for Loop 1 3-14 3-2 Layout of the Accumulator Line for Loop 2 3-15 3-3 Layout of the Accumulator Line for Loop 3 3-16 i 3-4 Layout of the Accumulator Line for Loop 4 3-17 3-5 True Stress Strain Curve for SA 351-CF8A Stainless Steel at 550'F 3-18 5-1 Fully Plastic Stress Distribution 5-9 5-2 Analytical Predictions of Critical Flow Rates of 5-10 Steam-Water Mixtures r 5-3 [ ]a,c.e Pressure Ratio as a 5-11 Function of L/D 5-4 Idealized Pressure Drop Profile through a Postulated Crack 5-12 5-5 Loads Acting on the Model at the Governing Location 5-13 6 Critical Flaw Size Prediction for the Comanche Peak ' Nuclear Power Plant, Node 2041-SAW 5-14 5-7 Critical Flaw Size Prediction for the Comanche Peak ~ Nuclear Power Plant, Node 1040-SAW 5-15 '. 5-8 Critical Flaw Size Prediction for the Comanche Peak Nuclear Power Plant, Node 2332-SAW 5-16 WPF1082J/012492:10 Vil
LIST OF FIGURES (Cont'd.) Fioure Title EAgg 5-9 Critical flaw Size Prediction for the Comanche Peak Nuclear Power Plant, Node 2520-SAW 5-17 5-10 Critical Flaw Size Prediction for the Comanche Peak Nuclear Power Plant, Node 2900-SAW 5-18 A-1 Pipe with a Through Wall Crack in Bending A-2 B-1 Comparison of Typical Maximum and Minimum Stress B-14 Profile Computed by Simplified and [ j a,c e B-2 Typical. Schematic of Accumulator Line at [ B-15 j ,c.e a B-3 [ ]a,c.e Maximum and Minimum Stress B-16 Profiles for Transient #10 B-4 [ ]a,c e Maximum and Minimum Stress 8-17 Profiles for Transient #11 B-5 [ ]a,c e Maximum and Minimum Stress B-18 Profiles for Transient #12 B-6 [ ]a,c,e Maximum and Minimum Stress B-19 Profiles for Transient #14 8 4
SECTION 1.0
'. INTRODUCTION ~~
1.1 Eackground The current structural design basis for the 10"_ accumulator lines requires postulating non-mechanistic circumferential and longitudinal pipe breaks. This results in additional plant hardware (e.g. pipe whip restraints and jet shields) which would mitigate the dynamic consequences of the pipe breaks.- It is, therefore, highly desirable to be realistic in the postulation of pipe breaks for these lines and thereby eliminate the need for some of the plant hardware. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that are used for establishing that a circumferential type break will not occur. The method applied is the leak-before-break procedure. The evaluations considering circumferentially oriented flaws envelop longitudinal cases. _. , 1.2 Scone and Ob.iective -i The purpose of this investigation is to demonstrate leak-before-break for the 10" accumulator lines. The scope includes the entire accumulator lines, from the cold leg anchor point to the accumulator tank anchor point. Schematic drawings of the piping system are shown in section 3.0. The recommendations and criter.ia proposed in NUREG 1061 Volume 3 (1-1)* 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 location at which the highest stress occurs.
- 2) Identify the materials and the associated material properties.
-
- Numbers in parentheses refer to the references given at the end of the section.
WPF1032J/012492:10 1-1
- 3) Postulate a surface flaw. Determine fatigue crack growth. Show that a through-wall crack will not result.
;{
Postulate a through-wall flaw at the governing location with the 4) least favorable combination of stress and material properties. The size of the flaw-should be large enough so that the leakage is l assured of detection with margin using the installed leak detection ' equipment when the pipe is subjected to normal operating loads.
- 5) Using maximum faulted loads, demonstrate that there is a margin of at least 2 between the leakage size flaw and the critical size flaw, i
- 6) Review the operating history to ascertain that operating experien. ,
has indicated no particular susceptibility to failure from the effects of corrosion, water hammer, or low and high cycle f atigue.
- 7) Justify that the material properties used in the evaluation are representative of the plant specific material. Evaluate long term ,
effects such as thermal aging where applicable. . e . The flaw stability analysis is performed using the methodology described in SRP 3.6.3 (1-2). The leak rates AN alculated for the normal operating condition loads. The leak rate prediction mdel used in this evaluation is an [
]a,c e The crack opening area required for calculating the ieak rates is obtained by subjecting the postulated through-wall flaw to normal operating loads (1-3). Surface roughness is accounted for in determining the leak rate through the postulated fl aw.
The computer codes used in this evaluation for leak rate and fracture , mechanics calculations have been validated (bench marked). s l WF1082J/012492:10 1-2
1.3 Referencu 11 Report of the U.S. Nuclear Regulatory Comission Piping Review Comittee - Evaluation of Potential for Pipe Breaks, NUREG 1061, Volume 3, November 1984. 12 Standard Review Plant public comments solicited 3.6.3 Leak Before-Dreak Evaluation Procedurest Federal Register /Vol. 52, No. 167/ Friday, August 28,1987/ Notices,pp. 32626 32633. 1-3 NUREG/CR 3464, 1983, "The Application of fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks." k 1 \ l w inazs/o12492:1o 1-3
- . _ . . _ . _ _ - .._ . ~ _ _ _ , - , _ _ _ _ _ _ _ . . _ . . ._.
SECTION 2.0
. OPERATION AND STABILITY OF liiE ACCUMULATOR LINES AND THE REACTOR COOLANT SYSTEM t
2.1 Stress Corrosion Crackina The Westinghouse type reactor coolant system primary loop anc connecting Class I lines have an operating history that demonstrates the inherent operating , stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress ! corrosion cracking). This operating history totals over 450 reactor years, including five plants each having over 17 years of operation and 15 other t plants each with over 12 years of operation, r in 1976, 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 in Pressurized Water Reactors . (PWR's). The results of the study performed by the PCSG were presented in NUREG 0531 (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."
l l WN1082J/012692:10 21
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 (2-2) further confirmed that no occurrences of IGSCC have been reported for PWR primary coolant systems. As stated above, for the Westinghouse type plants there is no history of cracking failure in the reactor coolant system loop or connecting Class 1 piping. The discussion below further qualifies the PCSG's findings. 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 environe:ent. The material specifications consider compatibility with the system's operating envirenment (both internal and external) as well as other material in the system, applicable ASMC Code rules, fracture toughness, welding, fabrication, , and processing, The elements of a water envirenu nt 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, sulphites, 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. OLring 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 I operating procedures as a condition for plant operation. For example, during normal power operation, oxygen concentration in the RCS and connecting Class 1 WPf1082J/012492:10 2-2
lines 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 witnin the specified limits. Thus during plant operation, the likelihood of stress corrosion cracking is minimized.
2.2 Hater HammE Overall, there is a low potential for water hammer in the RCS and connecting accumulator lines since they are designed and operated to preclude the voiding condition in normally filled lines. The RCS and connecting accumulator lines including piping and components, are designed for normal, upset, emergency, and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design. Other valve and pump actuations are
, relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are a stringently controlled. Temperature during normal operation is maintained within a narrnw range by control rod position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady state :onditions. 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 characteristi(+. cre controlled in the design proce.,s. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system and connecting accumulator lines. Preoperational testing and operating experienca have verified the Westinghouse approach. The operating transients of the RCS primary piping and connected accumulator lines are such that no significant water hammer can occur. ~ , 2.3 Low Cycle and Hiah Cycle Fatiaut ,
Low cycle fatigue considerations are accounted for in the design of the piping system through the fatigue usage factor evaluation to show compliance with the rules of Section 111 of the ASME Code. A further evaluation of the low cycle WPF1082J/012492:10 23
i
- I e fatigue loading is discussed in Section 6.0 as part of this study in the form i of a fatigue crack growth analysis. .-
High cycle fatigue loads in the system would result primarily from pump vibrations during operation. During operation, an alarm signals the exceedance of the RC pump shaft vibration limits. Field measurements have been made on the reactor coolant loop piping of a number of plants during hot , functional testing. Stresses in the elbow below the RC pump have been found , to be very small, between 2 and 3 ksi at the highest. When translated to the connecting accumulator lines, these stresses are even lower, well below the fatigue endurance limit for the accumulator line material and would result in an applied stress intensity factor below the threshold for fatigue crack growth. Vibratory fatigue loads are monitored for the 10 inch accumulator line during the hot-functional testing of the plant and are well below the high cycle fatigue allowables. 2,4 Potential Deoradation Durina Service Wall thinning by erosion and erosion corrosion effects will not occur in the 10" accumulator lin: due to the icw velocity, typically less than 10 ft/sec and the material, austenitic stainless steel, which is highly resistant to these degradation mechanisms. The Comanche Peak Unit 2 accumulator lines nozzles are forged product forms which are not susceptible to toughness degradation due to thermal aging. Finally, the maximum operating temperature of the accumulator lines piping, whi:h is about 560*F or below, is well below the temperature which would cause any creep damage in stainless steel piping. 2.5 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, i WPF1082J/012492:10 2-4
. _ . . . . _ _ _ _. . _ _ _ . ~ - . _ _ _ . _ - _ _ _ . _ _ . _ _ _ _ _ _ .
22 Investigation and Evaluation of Cracking Incidents in Piping in
. Pressurized Water Reactors, NUREG 0691, U.S. Nuclear Regulatory Commission, September 1980.
i h 0 e 9 1 W11082J/012492:10 2-5
i F SECTION 3.0 MATERIAL CHARACTERIZATION 3.1 Pine and_ Weld Materials The materials of the accumulator lines are A376/TP316, A403/WP316, and L A403/WP304. The injection nozzle material in SA351 CF8A, a cast product form of the type used for primary loop piping of several PWR plants. The accumulator line is connected to the primary loop at one end, and the other end is connected to the accumulator tank. The welding processes used are gas tungsten arc weld (GTAW) shielded metal arc weld (SMAW), and submerged arc weld (SAW). The normal operating pressure and temperature before the first valve from the cold leg are 2250 psia and 550'F respectively. The pressure and temperatare between the first and third valve are 2250 psia and 120'F. The pressure and temperature after the third valve are 700 psia and 120'F. Weld locations and governing locations are identified in Figures 3 1 through
, 3 4 with the pipe geometries. . In the following sections the tensile properties of the materials are presented and criteria for use in the leak-before break analyses are defined.
3.2 Material Pronerties of the Accumulator Lines The room temperature mechanical properties of the Comanche Peak linit 2 Nuclear Power Plant accumulator line materials were obtained from the Certified Materials Test Reports and are provided in Tables 3-1 through 3 4. The room temperature ASME Code-(31) minimum properties are given in Table 3-5. It is seen that the measured properties well exceed those of the Code. The representative minimum and average tensile properties were established from the results given in Tabler. 3 1 through 3-4. The material properties at temperatures of 120'F and 550'F are required for the leak rate and stability
', ana,1yses discussed later. The minimum and average tensile properties were calculated by using the ratio of the ASME Section 111 properties at the temperatures of interest stated above. Table 3 6 shows the tensile properties i at the two temperatures of interest. The modulus of elasticity values were established at various temperatures from the ASME Section 111 (Table 3-7), in WPF1052J/012492:10 3-I
l the leak before break evaluation, the representative minimum properties at temperature are used for the flaw stability evaluations and the representative .- average properties are used for the leak rate predictions. Those properties ' are summarized in Table 3 6. 3.3 Tensile Properties of the In_iection Noztigi - The material certifications for the injection nozzles were used to establish the tensile properties. These properties are given in Table 3-8 at room temperature. From Table 3 8 the average yield strength value of SA351 CF8A (
]'dd The modulus of elasticity was obtained from the Nuclear Systems Materials Handbook (reference 3 2) for consistency with the stress strain diagram which was also obtained from that reference. Thestressstraincurve(minimumproperties)isshownin Figure 3-5. This curve is used in the crack stability analyses.
l 3.4 Eritture Touchness Properties of the In.iection Nozzles Because the accumulator injection nozzle is a cast stainless steel product form operating at 550'F, thermal aging toughness degradation can take place. . l , [ ,
]^* the end of service life Charpv U-notch energy (KCU)- following .'
l the procedure of reference (3-4). [ ] a,c.e 1 WPF1082J/012492:10 3-2
(
. ). By the criteria established in reference (3-3), the fracture toughness of the SA351 Cf8A is at least as great as the toughnessof[ ) , d the benchmark material of reference (3 3).
Available data on aged stainless steel wolds (reference (3 5)) indicate the J,, 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/in ). The applied value of the J-integral for a flaw in the weld a regions will be lower than that in the base metal because the yield stress for the weld materials is much higher at temperature. Therefore, weld regions are less limiting than the cast material. Therefore, the toughness values for LBB evaluation are established as those of [ )*,c.e: 7 y . c. . 3.5 References 3-1 ASME Boiler and Pressure Vessel Code Section 111, Division 1 Appendices July 1, 1989. 3-2 Nuclear Systems Material Handbook, ERDA Report TID 26666, November 1975, part 1. Group 1, Section 4. 3-3 f. J. Witt and C. C. Kim, " Toughness Criteria for Thermally Aged Cast Stainless Steel," WCAP 10931, Revision 1, July 1986 (Westinghouse Proprietary Class 2).
. 3-4 Slama, G., Potrequin, P. , Masson, S. H., and Mager, T. R., "Effect of Aging on Hechanical Properties of Austenitic Stainless Steel Casting . and Welds," presented at SMIRT 7 Post Conference Seminar 5 - Assuring
!* Structural Integrity of Steel Reactor Pressure Boundary Components, August 29/30, 1983, Monterey, CA. wricaza/oir492 to 33 L _ - - . _ _ .___ _ _ _-- __ _ . .- ..-__ ._.
35 WCAP-10456, "The Effects of Thermal Aging on the Structural Integrity l of Cast Stainless Steel Piping for W NSSS," W Proprietary Class 2, .- l l November 1983. l 9 1 l WPF1082J/012492:10 3-4 l
ROOM TEMPERATURE PROPERTIES OF THE ACCUMULATOR LINE MATERIALS FOR LOOP 10F THE COMANCHE PEAK NUCLEAR POWER PLANT - UNIT 2 Material / Type Yield Strength Ultimate Strength Elongation Area Red. ID Heat No./ Serial No. (psi) (psi) (%) (%) SA403/WP316 78,200 38,400 53.5 N/A ** A* D5660 44,400 57.0 N/A 8 3085-6-2 SA376/TP316 86,000 80,000 51,000 61.5 N/A C 49201 SA403/WP316 86,000 44,400 57.0 N/A D 3085-6-2 SA376/IP316 53.5 78,200 38,400 N/A E D5660 SA403/WP316 83,050 43,050 57.0 N/A F 1081-21-1 SA376/TP316 61.0 76.5 SA403/WP316 80,000 49,500 G 53893 57.0 83,050 43,050 N/A H 1081-21-1 SA376/TP316 77.5 79,500 43,000 62.0 I 49203 SA403/WP316 83,050 43,050 57.0 N/A J 1081-21-1 SA376/TP316 76.5 80,000 49,500 61.0 K 53893 SA403/WP316 83,050 49,500 57.0 N/A L 1081-21-1 SA376/TP316 1 78,350 40,050 56.0 N/A M 1081-17-2 SA376/TF316 38,400 53.5 N/A 0 D4583 SA403/WP316 78,200 Y 79,600 41,050 63.0 N/A
" P 1081-9-1 SA376/TP316 82,500 52,500 56.0 73 Q 49199 SA403/WP316 79,600 41,050 63.0 N/A R 1081-9-1 SA376/TP316 80,000 42,350 60.0 N/A 5 1081-18-1 SA375/TP316 59.0 86,000 44,100 N/A T 53755 SA403/WP316 60.0 N/A SA376/TP316 80,000 42,350 U 1081-18-1 65.0 SA403/WF316 87,900 48,800 N/A V 53000
- As shown in Figure 3-1
** Not available
ROOM TEMPERATURE FROPERTIES OF THE ACCUMULATOR LINE MATERIALS FOR LOOP 2 0F THE COMANCHE PEAK NUCLEAR POWER PLANT - UNIT 2 ID Heat No./ Serial No. Material / Type Yield Strength Ultimate Strength Elongation Area Red. (psi) (Psi) (%) (%) t A* 05660 SA403/WP3'6 78,200 38,400 53.5 N/A ** B 3085-6-2 SA376/TP3:6 86,000 44,400 57.0 N/A C 05660 SA403/WP346 78,200 38,400 53.5 N/A . D 3085-6-2 SA376/TP316 86,000 44,400 57.0 N/A E 05660 SA403/WP316 78,200 38,400 53.5 N/A F 3085-6-2 SA376/TP316 86,000 /4,000 57.0 N/A G 3085-4-2-2 SA316/TP316 83,200 44,500 66.0 75.5 H 55705 SA403/WP316 80,000 44,000 62.5 N/A I 3085-4-2-2 SA376/TP316 83,200 44,500 66.0 i,/A J 86,000 37,000 56.0 76.0 54029 SA403/WP316 u K 3085-4-2-2 SA376/TP316 83,200 44,500 66.0 N/A 5 L ' 55705 SA403/WP316 80,000 44,000 67.5 75.5 M 3085-4-2-2 SA375/TP316 83,200 44,500 66.0 N/A N 1081-2-1 SA376/TP316 82,900 42,100 63.0 N/A 0 05712 SA403/WP316 78,200 38,400 53.5 N/A P 1081-18-1 SA376/TP316 80,000 42,350 60.0 N/A 0 04583 SA403/WP316 78,200 38,400 53.5 73 R 1081-19-1 SA376/TP316 81,750 42,500 61.0 N/A S D4583 SA403/WP316 78,200 38,400 53.5 N/A T 1081-19-1 SA316/TP316 81,750 42,500 61.0 N/A U 1081-19-1 SA376/TP316 81,750 42,500 61.0 N/A V 53000 SA403/TP304 87,900 48,800 65.0 N/A l
- As shown in Figure 3-2
** Not available -
4 e e O d q
ROOM TEMPERATURE PROPERTIES OF THE ACCUMULATOR LINE MATERIALS FOR LOOP 3 0F THE COMANCHE PEAK NUCLEAR POWER PLANT - UNIT 2 ID Heat No./5erial No. Material / Type Yield St ength Ultimate Strength Elongation Area Red. (psis (psi) (%) (%) A* 49201-4 SA403/WP316 80,000 51,000 61.5 N/A ** B 3085 4-2-2 SA376/TP316 83,200 44,500 66.0 N/A C 49201-15 SA403/WP316 80,000 51,000 61.5 N/A D 3085-4-2-2 5A376/TP316 83,200 44,500 66.0 N/A E 49201-9 SA403/WP316 80,000 51,000 61.5 N/A F 3085-4-2-2 SA376/TP316 83,200 44,500 66.0 N/A G 3085-6-2 SA376/LP316 86,000 44,400 57.0 N/A H 55706-1 SA403/WP316 77,500 42,000 62.0 75.5 86,000 44,000 1 3085-6-2 SA376/lP316 57.0 N/A J 54029-1 SA403/WP316 86,000 37,000 56.0 76.0 L K 3085-6-2 SA376/TP316 86,000 44,000 57.0 N/A L .55706-2 SA403/WP316 77,500 42,000 62.0 75.5 M 3085-6-2 SA376/TP316 86,000 44,000 57.0 N/A N 1081-14-2 SA376/TP316 87,250 49,000 47.0 N/A 0 D-5712 SA403/WP316 78,200 38,400 53.5 N/A P 1081-14-? SA376/TP316 82,750 49,000 47.0 N/A Q 52975 SA403/WP316 80,000 52,500 61.0 74.5 R 1081-18-1 SA376/TP316 80,000 42,350 60.0 N/A S D-4583 SA403/WP316 78,200 38,400 53.5 N/A T 3085-4-1 SA376/TP316 87,000 45,500 57.0 N/A U 1081-14-1 SA376/TP316 81,150 44,150 64.0 N/A V 53000 SA403/WP304 87,900 48,000 65.0 N/A As shown in Fi9ure 3-2
** Not available
ROOM TEMPERATURE PROPERTIES OF THE ACCUMULATOR LINE MATERIALS FOR LOOP 4 0F THE COMANCHE PEAX NUCLEAR POWER PLANT - UNIT 2 ID Heat No./ Serial No. Materfal/ Type Yield Strength Ultimate Strength Elongation Area Reduction (psi) (psi) (%) (%) A* 55705-2 SA403/WP316 80,000 44,000 62.5 75.5 i 8 3085-6-2 SA376/TP316 86,000 44,400 57.0 N/A ** l C 55705-1 SA403/WP316 80,000 44,000 62.5 75.5 0 3085-6-2 SA376/TP316 86,000 44,400 57.0 N/A , E 55705-5 SA403/WP316 80,000 44,000 62.5 75.5 ! F 3085-6-2 SA376/TP316 86,000 44,400 57.0 N/A j G 1081-2-2-1 SA376/TP316 82,900 42,100 63.0 76.5 H 53893-2 SA403/WP316 80,000 49,500 61.0 N/A 1 1081-2-2-1 SA376/TP316 82,900 42,100 63.0 77.5 J 49203 SA403/WP316 79,500 43,000 62.0 N/A K 1081-21-1 SA376/TP316 83,050 43,050 57.0 N/A L 53893-4 SA403/WP316 80,000 49,500 61.0 N/A Y M 1081-2-2-1 SA376/TP316 82,900 42,100 63.0 N/A N 1081-19-2 SA376/TP316 79,600 41,050 63.0 N/A 0 D4583 SA403/WP316 78,200 38,400 53.5 N/A P 1081-2-1 SA376/TP316 82,900 42,100 63.0 N/A Q 53755 SA403/WP316 86,000 44,100 59.0 N/A R 1081-2-1 SA376/TP316 82,900 42,100 63.0 N/A S 1081-9-1 SA376/TP316 79,600 41,050 63.0 N/A T 49778 SA403/WP304 85,600 44,900 61.0 N/A
- As shown in Figure 3-4
** Not available
r TABLE 3 5 Room Temperature ASME Code Minimum _ Properties Material Yield Stress Ultimate Stre n (psi) (psi) A403/WP304 30,000 75,000 A376/TP316 30,000 75,000 and A403/WP316 e e
?
I 4 i 1. WPf1082J/012492:10 3-9 _. _ .,..._.a . _ . _ . . _ - _ _ _ . _ _ _ , _ . . _ _ _ . _ - _ . . _ _ _ _ = , _ . _ _ _ . . _ . , _ _ _ _ _ _ _ _ _ . , _ . . _, , . . _ . . . _ _ _ _ _ . . _ _ . , . , _ _ . _ _
b 1 TABLE 3 6 . Tensile Properties for the Comanche Peak Unit 2 Nuclear Power Plant 10 Accumulator Lines " l' Minimum Temperature Minimum Average Ultimate KLterial ('F) Yield fosi) Yield f osi) _(osil A403/WP304 120 43,417 46,052 85,285 A376/TP316 120 38,727 42,997 78,062 550 23,027 25,567 60,413 A403/WP316 120 35,778 40,827 77,214 550 23,538 26,860 63,734 SA351/CF8A 550 22,178 25,318 70,465 4 e 9 s 9 i WPF1082#012492:10 3-10 l
TABLE 3 7 I Modulus of Elasticity (E) Temperature E ('F) (ksi) l 120 28.031 550 25,550
?
i k j. l e 9 WPF1082J/012492:10 3-11
AVAILABLE MECHANICAL PROPERTIES OF THE 4 ACCtMJLATOR INJECTION N0ZZLES AT ROOM TEMPERATURE 0.2% offset Ultimate % % toop Product Heat Yield Stress Strength Elongation Reduction i No. Form Number Material fosil fosil _ per Inch in Area
! 1 Nozzle 3-3659/0763 SA351-CFBA 35124 82085 62.0 N/A 2 Nozzle 3-3698/1154 SA351-CF8A 41398 85640 59.0 N/A 3 Nozzle 3-3719/3333 SA351-CF8A 42180 89269 59.6 N/A 4 Nozzle 3-3695/0762 SA351-CF8A 41682 88458 56.2 N/A u
h N/A - Not available i i j 1 4 e , $ @ 8 . e
i, t TABLE 3 9 i CHEMISTRY AND END Of SERVICE Life KCU TOUGilNESS j TOR ACCUMULATOR INJECTION N0ZZLES t i Cr 51 Ni Ho C Mn N Cb KCU Heat 1(h $_ 5 5_ 5._ L $_ 5_ 5_ daJ/cm'
- - a,c.e :
i 3 3659/0763 ! i 3 3698/1154 : 33719/3333 { 33695/0762 M e O I 1 I e 4-t e WPf1082J/012492 10 3 13
. . - . . - ..-...._.--...-...-.L. , . - - __,- - --,. - ..-. -.... - . . - _ - - , - . , - ,
Westinghwse Proprietory C16sn ! Ptpe Outstee pianeter 10,75 in. Mlatewe Wall thicLaens 0,476 ta. . Accumulator isnt heasle 14fe (not
- Ninteue W411 Thltteess 0.3tst in.
$W
- thop Wela .
rv . n.is v.ie Act1M.LAim TNM I N (d) R y D W y ,
,g
,. .' W m W U W .P 0 N
- 5) -
. g g
.y 0 ;
-W ~ Wp K ry W
\
N W- - w. . i W Figure 3 1. Layout of the Accumulator Line for Loop 1 wro981uo10392:10 3-14
Wettinghouse Prctriotary (1666 2 P'De Owlltee Otometer 10.75 in. Nintaus Well Thttkhess 0.818 in. 4tmleter f ant hergle late Ind: Mlaimoe Well Thicknell 0.3205 in. 5W Shop tels e tv - rieia voie Q) x N ' \sw '
,- W ACCLH TNH 2 VN msr xs CV T, , sw (U)
(f) W SW x, (R) x-- H sw (b (i)' x* (b g
~
, sw 3
kg g
,(3)' 3 N
sw sw s
)
SW g_ _
\ \ (T)
% M (0) ($[
x:+ w t ,1 (L1 CCLD LEG Figure 3 2. Layout of the Accumulator Line for Loop 2 WP F 0981J/010392:10 3-15
watinghoves Proprietary Ctest 2 PfPe Outside Otoseter 10,76 in, flinieue Well Thttkness 0.878 in. Attwelator feat Nortle Safe (nel Iltateue Wall fhttaness 0.Itti in.
$W * $h6p Weld rv . ii.ie Wei, G
I h (S E W W NN p
*/()
f w A00LKLAim TNA 3
, /, ' w- -
m (Et N-* x __ y;/* _ m
'0)
~
(H) w yy , - r u toet(sAw)g g
- q). %$ ~ g w
U) /
. L -
&N aio 'to y
Figure 3 3. Layout of the Accumulator Line for Loop 3 wro9eistoio392:w 3 16
Westir$ house Proprietary Citse i . I t
.. rip. ovisie. oi ier is.n $a. !
mint u Wall 1hicknest 0.875 in. l au m i i.e t..a n.: i. s.t. tavi
*iaiu v.ii inico.is o nu in. !
5W Shop Wele in . ri.\e we\e , 4 (. tC) w , w //~ l w A
- -y D
fv r w_ _' OttD LEG , t W FM W' / 0 W' ACnMA.ATOR T#4
@], w y TCX AIATAT-04
- w/ '
w g W W/ i R W yh N N ^< i u , A s o w o ,v Q.. - w I pm)
,, g . 0) l-i i-e l*
l-. Figure-3 4. Layout of the Accumulator Line for Loop 4 wProes1J/oto392:10 3-17 e-- e---- .,v...--.,-e---m-- .w m , . . . , - . , , , +--.-- ..- 7 -,gw.r- ,. - ----rwe.n.-vr- r 7
---,r----g.-..e ,- m - y= = ~ +-
a .c .< . I l r i i 1 . 4 i _. m-+ -( i Figure 3 5. True Stress Strain Curve for SA351 CF84 Stainless Steel at 550'F . Wo9etJ/ottonito 3_t3
SECTION 4.0 i
., LOADS FOR FRACTURE MECHANICS ANALYSIS ! - 4 4,1 Nature of the loads ,
4 : Under normal operating conditions, the accumulator lines are subjected to axial and bending loads which arise from deadweight, pressure, and thermal expansion. Under faulted conditions, the loads caused by Safe Shutdown , Earthquake (SSE) are superimposed on these normal operating loads. The stresses due to axial loads and bending moments were calculated by the ! following equation: I o={+j (4 1) where, o - stress ;
. F = axial load M - bending moment .
A - pipe cross-sectional area
^
Z = section a,odulus
- The x direction is the axial direction of the pipe with y and z denoting the remair.ing orthogonal directions. The bending moments for the desired loading combinations were calculated by the following equation:
M = (MhMj) ua (4-2) where. 1 M - bending moment for required loading My 4 y component of bending moment H - component of bending moment Z WPF1032J/012492:10 4-1
. . . . _- -- .. - _. = _
The axial load-and bending moments for crack stability analysis and leak rate predictions were computed by the methods explained in Sections 4.1 and 4.2 ., which follow. . 4.2 L21ds for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the absolute sum method as follows: F - lFDWl+ lFTHl+ lFpl + lFSSE! I4'3) My = l(M)DWI+IIN)TH!+IIN)SSE! y Y Y (4'4) My = l(NIZ DW !+l(N)THl+l(M)SSEl Z Z (4-5) Where, the subscripts of the above equations represent the following loading cases: DW- = deadweight TH - normal thermal expansion . SSE = SSE loading including seismic anchor motion P - ioad due to internal pressure . 4,3 Loads for leak Rate Evaluation The normal operating loads for leak rate predictions were calculated by the algebraic sum method as follows: F = FDW + FTH + Fp (4-6) N " Y (NY )DW + (N IY TH (4-7) N Z (NZ)DW + (NZ)TH (4-8) The parameters and the subscripts are the same as those explained in Section 4.2. . wf1082u012692:10 42
t 4.4 $mrimary of Lotds and GestnfirX l The load combinations were evaluated at the various weld locations. Normal : loads were determined using the algebraic sem method whereas faulted loads J were combined using the absolute sum method as discussed above. 4.5 Governing _ Locations l The governing locations were established on the basis of the pipe schedules, - types of material, operating temperature, material properties, the highest faulted stresses for the welds, and the types of welds. The shop welds (SW) were SAW and the- field welds-(FW) were rnade by the combination of GTAW and SMAW. Maximum faulted loads of the node point in the neighborhood of the : nozzles were used for the loads of the injection nozzles. This node was identified as node 2041 in loop 3. All four loops were investigated and the following governing locations were identified: ' Material Temperature Node ('F) , I SA351/Cf8A 550 2041/ Loop 3 SA403/TP316 550 2041/ Loop 3 120, 2332/ Loop 4 SA376/TP316 550 1040/ Loop 2 120 2520/ Loop 4 SA403/TP304 120 2900/ Loop 3 l The loads and stresses for the governing locations are shown in Table 4-1. In developing these tables the appropriate wall thicknesses were used. The governing locations have been indicated in the layout sketches of figures
, 3 2, 3-3 and 3 4.
WPF1082J/012492:10 4-3
TABLE 4 1 i Summary of hrmal and faulted loads and Stresses at Governing Locations Node load Axial Force
- Axial Bending 8ending Total
& Case Stress Homent Stress Stress Loop (lbs) (psi) (inlbs) (psi) (psi) 2041/3 Normal 129,137 4,757 361,541 5,827 10,584 t faulted 147,316 5,427 950,338 15,317 20.744 1040/2 Normal 127.385 4,693 392,791 6,331 11,024 Faulted 147,362 5,429 618,966 9,976 15,404 ,
2332/4 Normal 139,243 5,130 342,416 5,519 10,648 ' Faulted 151,284 5,573 744,449 11,999 17,572 2520/4 Normal 136,521 5,029 488,785 7,878 12,907 - Faulted 139,630 5,144 1,066,744 17,193 22,337 2900/3 Normal 55,955 5,202 50,155 1,844 7,047 Faulted 58,185 5,410 233,841 8,600 14,010
- Pressure included.
t I
+
b e b WPF1082J/01249h10 4-4 i
SECTION 5.0
. FRACTURE MECHANICS EVALUATION 5.1 failure Mechanism Determination of the conditions which lead to failure should be done with plastic fracture methodology because of the large 4 mount of deformation accompanying fracture. One method for predicting the failure of ductile material is the [ Ja,c.e method, based on traditional plastic limit load concepts, but accounting for [ ]a c.e and taking into account tht presence of a flaw. Tha flawed pipe is predicted to fail when the remaining ,1et section reaches a r. tress level at which a plastic hinge is formed. The scress level at which this occurs is called the flow stress. [
j a,c.e This methodology has been shown to be applicable to ductile piping through a large number o' experiments and is used hern to predict the critical flaw sizes in the accumulator lines. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 5-1) when
, loads are applied. The detailed development is provided in Apr dix 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.. (5-1) with-
[ ] (5-2) [ WPF1082J/012492:10 5-1 l 1
{ a,c.e .- The analytical model described above accurately accounts for the piping internal pressure as well as imposed axia*. force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental results (5-1). Flaw stability evaluations using this an4lytical model, are presented in section 5.3. : 5.2 Leak Rate Predictions The purpose of this section i. .s discuss tne method which will be used to predict the flow through a postulated crack and present the leak rate calculation results for postulated through wall circuMerential cracks in the accumulator lines. 5.2,1 General Consideratio u 4 The flow of hot pressurized water through an opening to a lower back pressure (causing choking) is taken into account. Fnr long channels where the ratio of . the channel length, i., to hydraulic diameter, Dg , (L/Dg ) is greater than [ ]a,c,e, both [ ]a,c,e must be considered. In this situation the flow can be described as being single-phase through the channel until the local pressure equals the saturation pressure of the fluid. > At this point, the flow begins to flash and choking occurs. Pressure losses
- . due to momentum changes will dominate for [ ]a,c e However, for large L/Dy values, friction pressure drop will become important and must be considered along with the momentum icsses due to flashing.
5.2.2 [.alculation Method Using an [ j ,c.e a , WPF1082J/012492:10 5-2
The flow rate through a crack was calculated in the following manner. Figure
- - 5-2 from reference 5-2 was used to estimate the critical pressure, Pc, for the primary loop enthalpy condition and an' assumed flow. Once Pc was found for a
~ a given mass flow, the [ J .c.e was found from Figure 5-3 taken from reference 5-2. For all cases considered, since [ la,c,e Therefore, this method will yield the two-phase pressure drop due to momentum effects as illustrated in Figure
- 54. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using APf= ( )* 'd (5 3) 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 caiculations was
[ ,
]a,c,e pq3, The frictional pressure drop using equation 5-3 is then calculated for the . assumed flow and added to the momentum pressure drop calculated using the Fauske model to obtain the total pressure drop from the primary system to the atmosphere. Thus, Absolute Pressure - 14.7 = [ ]a,c e(5-4) for a given assumed flow G. If the right-hand side of equation 5-4 does not agree with the pressure difference between the piping under consideration and the atmosphere, then the procedure is repeated until equation 5-4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack.
For the locations at the lower temperature, single phase calculations for the
. leak rate in gallons per minute (GPM) were performed, using an equation from reference 5-3 as follows:
-,c.e
( ) (5-5) WPF1082J/012492:10 5-3 l
l [
]*.e.e 5.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulr.ed through-wall crack length for the five critical locations p 'viously identified. The crack opening areas were estimated using the method of reference 5-4 and the leak rates were calculated using t!ie calculational methods described above. The leak rates were calculated using the normal operating loads at the governing nodes identified in section 4.0. The crack lengths yielding a leak rate of 10 gpm (10 times the leak detection capability of 1.0 gpm) for the critical locations at the Comanche Peak Nuclear Power Plant Unit 2 are shown in Table ,
5-1. 5.2.4 Leak Detection Canability The Comanche Peak Unit 2 Nuclear Power Plant leak detection system inside the containment can detect 1 gpm leak rates as required by Regulatory Guide 1.45. As seen above, a margin of 10 was applied to the leak rate to define the accumulator line leakage size flaws in accordance with NUREG 1061, Volume 3. 5.3 Stability Evaluation of Accumulation Lines A typical segment of a pipe under maximum loads of axial force F and bending moment M is schematically illustrated as shown in Figure 5-5. In order to calculate the critical flaw size, plots of the limit moment versus crack length are generated as shown in Figures 5-6 through 5-10. As mentioned in . 4 I WPf1082J/012492:10 5-4 1
Section 4.0, shop welds were performed by SAW and field welds were performed
. by the combination of GTAW and SMAW. Therefore field weld locations are conservatively considered to be SMAW. The nodes of maximum load under faulted condition are found to be all shop weld (SW) and therefore the leak-before-break concept is demonstrated for SAW welding procedures. The critical flaw size corresponds to the intersection of the limit moment curve and the maximum moment load line. The critical flaw size is calculated using the lower bound base metal tensile properties established in section 3.0.
The "Z" factor correction for SAW weld was applied (5-5 and 5-6) as follows: Z - 1.30 [1 + 0.010 (0D - 4)) (5-6) where OD is the auter diameter in inches. Substituting 00-10.75 inches, the Z factor was calculated to be 1.39 for SAW. The applied loads at the SAW locations were increased by the Z factors and the plots of limit load versus crack length were generated as shown in Figure 5-6 to 5-10. Table 5-2 shows the summary of critical flaw sizes for the Comanche Peak Unit 2 nuclear power plant 10" accumulator lines. S.4 lacal Stability Analysis of the in.iection Nozzles In this section the local stability analysis is performed to show that unstable crack extension will not occur when postulated through wall flaws in the cast injection nozzles are subjected to maximum loads. At the critical nozzle identified in Section 3.0, the (normal plus SSE) outer surface axial stress, o,, is seen to be 20.7 ksi based on the minimum wall thickness (see Table 4-1 of Section 4.0). The (normal plus SSE) axial force and bending moment are Fx = 147 kips and3M - 950.3 in-kips, i The minimum yield <trength for flaw stability analysis is 22.2 ksi (see
', Section 3.0). The EPRI elastic plastic fracture handbcok method was used to calculate the J,, using the normal plus SSE loads. The J, was calculated for a [ ] long postulated through wall ' law (which is 2. times the reference flaw size) and was found to be [ ]. Since the J ,
value is greater than the Ji , value of [ ] the tearing modulus WPF1082J/012492:10 5-5 L_____-_-_-_-__-_
was evaluated. The applied tearing modulus, Ty , was found to be [ ]'dd. Both J,, and Ty , are below the allowables of ( ..
]'d** respectively, given in Section 3.0. Therefore, unstable crack propagation will not result. -
7 5.5 References 5-1 Kanninen H. F. et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976. 5-2 ( ja,c e 5-3 " Thermal Engineering," C. C. 01111o and E.P. Nye, International Text Company, pp. 270-273, 1969. 5-4 Tada, H., "The Effects _of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal , Through-Crack in a Pipe," Section II-1, NVREG/CR-3464, September 1983. 5 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640, 5-6 Standard Review Plan; Public Comment Solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal Register /Vol. 52, No.167/ Friday, August 28, 1987/ Notices, pp. 32626-32633. i 1 . e . l l WPF1082J/012492:10 5-6 i
TABLE 5-1
.. 1 Leak Rate Crack Lengths for the Governing Locations of Comanche Peak Unit 2 Accumulator Lines 1
- Ende Point Material Location Temperature Crack Lenath fin.) ,
(*F) (for 10 gpm leakage) i q ,c,. i e
. '"~ .Before the first valve from RCS (t - 0.875 in.)
** After the first valve from RCS (t - 0.875 in.)
*** AccL.iulator tank nozzle junction (t - 0.3285 in.)
9 9 0, WPF1082J/012492:10 5-7 I _ _ _ _ _ _ _ __
TABLE 5-2 Summary of Critical Flaw Sizes.for the Governing Locations of the Comanche Peak Unit 2 10" Accumulator Lines Temperature Critical Flaw Size (in.) Node Point tittfr_tal Locat100 (*F) A
$_A3
_ a,c.e I l i l i WPF1082J/012492:10 5-8 l
d.c.e i 4 Figure 5-1. Fully Plastic Stress Distribution 5-9 WPF1058J/011492:10
t
- a,c.e m
I S g
~ ~
STAGNATNNd ENTHALPY (10 2 temW Figure 5-2. Analytical Predictions of Critical Flow Rates of Steam-Water $ Mixtures wriossuo11t92:10 5-10
- - a,c e "o
O i 9 l. I 3 5 4 , LENGTM/Df AMETER MATIO (L/Di .t , . Figure 5-3. [ ]a,c,e Pressure Ratio as a Function of L/D 4 L woriosa;/011492:io 5-11
i i
-4,:,e r-4.c.e
/ f l
l= l 4 Figure 5-4. Idealized Pressure-Drop Profile Through a Postulated Crack . WPF1050J/011492:10 5-12
1 R g y -. - - e 2 l
' -F' q 1 y wL -
o t.; i e 3
. 'E E I l
l l l l , l I I I i 1 l '
"_. _a l I :
I I l I I I i I i i I sp e Figure 5-5. Loads Acting on a Pipe Model wn tossuot u92:10 5-13
F a.c.e 1
)
l i [ rigure 5-6. Critical Flaw Size Prediction for the Comanche Peak Nuclear I Power Plant Node 2041 - (AW wnossuo11492:10 5-14
a,c.e
- 7 y
=. Figure 5-7. Critical Flaw Size Prediction for the Comanche Peak Nuclear Power Plant Node 1040 - SAW WPF1058J/011492:10 5-15
.;:] -
,~
\
- a.c,e I
i i t L l l i i
- r i
- k-l-
~c Figure 5-8. Critical Flaw Size Prediction for the Comanche Peak Nuclear -
Power Plant Node 2332 - SAW 5-16 wro9sufo11392:ta
- 4 dCe i i
L
- : et . _
l Figure 5-9. Critical Flaw Size Prediction for the Comanche Peak Nuclear Power-Plant Node 2520 - SAW w n o981af01139221o 5-17
.u .. . _ _ , - . . _ _ . _ . - - . _ , _ . . _
+
a,c.e i i i .
.J .
t Figure 5-10. Critical Flaw Size Prediction for the Comanche Peak Nuclear Power Plant Node 2900 - SAW wpro981a/011392:10 . - 5-18
-]
SECTION 6.0
.. ASSESSMENT OF FATIGUE CRACK GROWTH The purpose of the fatigue crack growth (FCG) analysis is to demonstrate that a postulated flaw will not grow through the wall under all. operational loadings.
The fatigue crack growth in the Comanche Peak Unit 2 Nuclear Power Plant 10" accumulator lines was determined by comparison with a generic fatigue crack growth analysis of a similar piping system. The accumulator lines extending from the RCS cold leg injection points to the tank compare reasonably well with the generic analysis, having essentially the same geometry, materials, and fatigue crack growth rate. Based on comparing all parameters critical to the fatigue _ crack growth analysis, it was concluded that the generic analysis would envelop their fatigue crack growth, lhe details of the generic fatigue crack growth analysis are presented in appendix B. Fatigue crack growth results are summarized in table B-4 of appendix B. Due to similarities in Westinghouse PWR designs it was possible to perform a generic fatigue crack growth calculation which would be applicable to many
, plants. A comparison was made of stresses, number of cycles, materials, and geometry.
The following summarizes the parameters which were compared: Generic Cold Leg Nozzle Comanche Peak Unit 2 Cold Critical Location To Pine Weld Leo Nozzle to Pipe Weld Pipe Outer Diameter 10.75" 10.75" Thickness 0.895" .875" Material Austenitic Stainless Steel Austenitic Stainless Steel-Normal Temperature 550'F 550'F Normal Pressure 2235 psig 2235 psig Normal Operating Stress (Press, DW, 10.1 ksi 10.6 ksi Thermal Exp.)
', Thermal Transients See Appendix B
- Thermal transient loadings are nearly identical for this comparison.
l l WPF0981J/011792:10 6-1 l 1
The maximum allowable preservice indication may have a depth of about 0.1 in, per IWB-3514.3, Allowable Indication Standard or Austenitic Piping, ASME Code,. ' Section XI Division 1, 1985 edition. Typical fatigue crac!; growth results for various initial flaw depths are given in Table B-4 in Appendix B. From the table an initial crack 0,10 inch deep is calculated to grow a depth of 0,132 in, at end. of life. Similarly a crack having an initial depth of 0,1S in, grows to 0.186 in, In conclusion, the fatigue crack growths calculated for the generic case, as summarized in section B,2.2, are applicable to the Comanche Peak Unit 2 Nuclear Power Plant accumulator lines. These results demonstrate that no significant fatigue crack growth will occur over the 40 year plant design life. t WPF0981J/011792:10 6-2
SECTION 7.0 .. ASSESSMENT OF MARGINS
- In the preceding sections, the leak rate calculations, fracture mechanics analyses and fatigue crack growth assessment were performed. Margins at the critical locations are summarized in Table 7-1.
In summary, relative to
- 1. Flaw Size A margin of at leart 2 exists between the critical flaws and the flaws yielding a leik rate of 10 gpm.
- 2. Leat Rate For the reference flaw sizes a margin of 10 exists between the calculated leak rate and the 1 gpm leak detection criteria of o
Regulatory Guide 1.45. In the evaluation, the leak-before-brenk methodology is applied conserva-tively. The conservatisms used in the evaluation are summarized in Table 7-2. WPF1082J/012492:10 7-1
TABLE 7-1 LEAKAGE FLAW SIZES, CRITICAL FLAW SI7ES AND MARGINS e e,c.e i s I (*) For the cast injection nozzles, J , was less than J , and T,, was less than T , for a flaw having a length of ( Ja,c.e,
. e e
WPF1082J/012492:10 7-2 l
TABLE 7-2 LBB CONSERVATISMS
- l o Factor of 10 on Leak Rate l
o Factor of 2 on Leakage Flaw for all cases ) o Algebraic Sum of Loads for Leakage o Absolute Sum of Loads for Stability o Average Material Strengths for Leakage o Minimum Material Strengths for Stability o WPF1082J/012492:10 7-3
SECTION 8.0
. CONCLUSIONS
- This report justifies the elimination of 10" accumulator lines pipe breaks from the structural design basis for the Comanche Peak Unit 2 Nuclear Power Plant as follows:
- a. Stress corrosion cracking is precluded by use of fracture resistant materials in the pipe system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.
- b. Water hammer should not occur in the RCS piping (primary loop and the attached auxiliary lines) because of system design, testing, and operational considerations,
- c. The effects of low and high cycle fatigue on the integrity of the accumulator line piping are negligible.
e
, d. Adequate margin exists between the leak rate of small stable flaws and the capability of the Comanche Peak Unit 2 plant's reactor coolant system pressure boundary leakage detection system.
- e. Ample margin exists between the small stable flaw sizes of item d and the critical flaws.
The postulated reference flaws will be stable because of the ample margins in d and e and will leak at detectable rates which will assure a safe plant shutdown. Based on the above, it is concluded that pipe breaks in the 10" accumulator lines need not be considered in the structural design basis of the Comanche i ', Peak Unit 2 Nuclear Power Plant. , WPf1082J/012492:10 8-1
APPENDIX A LIMIT HOMENT t-The internal stress system at the crack plane has to be in equilibrium with the applied loading, i.e., the hydrostatic pressure P, axial force F, and the bending momentb M . The angle 8 which identifies the point of stress inversion follows from the equilibrium of horizontal forces (see Figure A-1). That is:
-._ a ,c.e
._7 l
t f I i i 4 i i J l 9 4 e WPF1082J/012492:10 A-1 l
m , .. I P 4 t 8 Figure A-1. Pipe with a Through-Wall Crack in Bending l WDF1058J/011492:10 A-2 l l
. APPENDIX B FATIGUE CRACK GROWTH CONSIDERATIONS 8.1- Thermal Transient Stress Analysis The thermal transient stress analysis was performed for a typical PWR plant to obtain the through wall stress profiles for use in the fatigue crack growth analysis of Section B.2. The through wall stress distribution for each transient was calculated for 1) the time corresponding to the maximum inside surface stress and, ii) the time corresponding to the minimum inside surface stress. These two stress profiles are called the maximum and minimum through wall stress distribution, respectively for convenience. The constant stresses due to pressure, deadweight and thermal expansion (at normal operating temperature, 550*F) loadings were superimposed on the through wall cyclical stresses to obtain the total maximum and minimum stress profile for each transient. Linear through wall stress distributions were calculated by conservative simplified methods for all minor transients. More accurate . nonlinear through wall stress distributions were developed for severe transients by [ ]a,c,e B.I.1 Critical location for Fatiaue Crack Growth Analysis The accumulator line design thermal transients (Section B.1.2),1-D analysis data on accumulator line thermal transient stresses (based on ASME Section III NB3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis. [
Ja ,c,e This location is selected as the worst location based on the following considerations:
', , i) the fatigue usage factor is highest. , ii) the stress due to thermal expansion is high.
iii) the effect of discontinuity due to the undercut at the weld will tend to increase the cyclical thermal transient loads. ( WPF1082J/012492:10 B-1
9 iv) the review of data shows that the 1-0 thermal transient stresses in the accumulator line pipin section are generally higher near the [RCL ** cold leg nozzle compared to rest of the accumulator line.]a,c,e 4 B.I.2 Desian Transients The transient conditions selected for this evaluation are based on conservative estimates of the magnitude and the frequency of the temperature fluctuations resulting from various operating conditions in the plant. These are representative of the conditions which are considered to occur during plant operation. The fatigue evaluation based on these transients provides confidence that the component is appropriate for its application over the design life of the plant. All the normal operating and upset thermal transients, in accordance with design specification and the applicable system , design criteria document (B-1), were considered for this evaluation- Out of these, only [
]a,c.e These transients were selected on the basis of the following criteria: ,
[ (B.1) - (B.2)
) a,c,e B.1.3 Simolified Stress Analysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients. .
f . l
. ] d The inside surface stress was calculated by the following equation which is similar to the transient portion of ASME Section III NB3600, Eq. 11:
wncs2J/012492:10 B-2
'. Sg-( Ja c.e (B.3) i where, - a,c.e i
I
?
i i I I i i 4
)
J [ la.c.e The maximum and minimum inside surface stresses were searched from the S g values calculated for each time step of the transient solution. The outside surface stresses corresponding to maximum and minimum insidt stresses were calculated by the following equations: S01 - ( P' ' (B,7) 502 = [ P' ' (B.8) 4 4 WoF1082J/Dt?&92 10 B-3
where, p - ..c.. ,. l l b_. - The material properties for the accumulator pire ((SA376/P316)) and the RCL [ a J.c.e The values of E and a, at the normal operating temperature, provide a conservative estimation of the through wall thermal transient stresses as compared to room temperature properties. The following values were conservatively used, which represent the highest of the [ Ja,c.e materials: a,c.e
- J The maximum and minimum linear through wall stress distribution for each thermal transient was obtained by [
] a,c.e The simplified analysis discussed in this section was performed for all minor thermal transients of a
[ J.c.e Nonlinear through wal' stress profiles were developed for the remainit.g severe transients as explained in Section B.1.4. The inside and outside surface stresses calculated by simplified methods for the minor transients are shown in lable B-2. [ la,c.e This figure shows that the , simplified method provides more conservative crack growth. , WPF1082J/012492:10 B-4 I
B.l.4- Ronlinear Stress Distribution fer severe Transients e. [
}a,c.e As mentioned earlier, the accumulator line sectiois near the [ ]a,c.e is the worst location for fatigue crack growth analysis. A schematic of the accumulator line geometry at this location, is shown in figure B 2, [
0 O m 8,C,e
-9 0
4 4 wrioara/012592:10 B5
4 B.1.5 Total Stress for Fatique Crack Growth The total through wall stress at a section was obtained by superimposing the pressure load stresses and the stresses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of Table B 2 or the nonlinear stress distributions discussed in Section B.1.4). Thus, the total stress for fatigue crack growth at any point is given by the following equation: Total Stress Thermal Stress Due Stress for Transient to Due to Fatigue = Stress + DW + + internal (B.9) Crack Growth Thermal Pressure Expansion The envelope thermal expansion, deadweight and pressare loads for calculating the total stresses of Equation B.9 are summarized in Table B-3. o B.2 .Fatioue Crack Growth Analysf1 The fatigue crack growth analysis was performed to determine the effect of the design thermal transients given in Table B-1. The analysis was performed for the critical cross section of the model which is identified in Figure B 2. A range of crack depths was postulated, and each was subjected to the transients in Table B-1, 1 B.2.) Analysis Procedure The fatigue crack growth analyses presented herein were conducted in the same manner as suggested by Section XI, Appendix A of the ASME Boiler and Pressure Vessel Code. The analysis procedure involves assuming an initial flaw exists at soma point and predicting the growth of that flaw due to an imposed series of stress transients. The growth of a crack per loading cycle is dependent on , the range of applied stress intensity factor AKg , by the following relation: , jhh=courf y (B 10) wnos2not2492:10 G-6
where "Co" and the exponent "n" are material properties, and AK g is defined later, in equation (B-10). For inert environments these material properties are constants, but for some water environments they are dependent on the level of mean stress present during the cycle. This can be accounted for by adjusting the value of "Co" and "n" by a function of the ratio of m W am to maximum stress for any given transient, as will be discussed 1.ter, fatigue crack growth properties of stainless steel in a pressurized water environment have been used in the analysis, i The input required for a fatigue crack growth analysis is basically the information necessary to calculate the parameter AK ,g which depends on crack and structure geometry and the range of applied stresses in the area where the crack exists. Once AK; is calculated, the growth due to that particular cycle i can be calculated by Equation (B.10). This increment of growth is then added to the original crack size, the AK; adjusted, and the analysis proceeds to the next transient. lhe procedure is continued in this manner until all the transients have been analyzed. 4 The crack tip stress intensity factors (K ) gto be used in the crack growth
. analysis were calculated using an expression which applies for a semi-elliptic surface flaw in a cylindrical geometry (B 4).
The stress intensity factor expression was taken from reference B 4 and was calculated using the actual stress profiles at the critical section. The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial: o (x) =A 3
+Ad +A ( * )+A ( * ) 2 3 3 (B 11)
The stress intensity factor K;(d) was calculated at the deepest point of the crack using the following expression: p.c.e (g.12) wnics2n012497:10 B7
a,C,e i e i I Calculation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law determined from consideration of the available data for stainless steel in a pressurized water environment. This law allows for the effect of mean stress or R ratio 7 (Kimin/Kimax) on the growth rates. The reference crack growth law for stainless steel in a pressuria.ed water environment was taken from a collection of data (B 5) since no code curve is available, and it is defined by the following equation:
-=I 1 * * ** (B-13) where K,fg =(Kg ,3x) (1-R)l/2 R=
Kym
= Crack growth rate in micro-inches / cycle ,
WPf1082J/012492:10 B-8
8.2.2 Results s fatigue crack growth analyses were carried out for the critical cross section. ' Analysis was completed for a range of postulated flaw sizes oriented circumferential1y, and the results ue presented in Table B 4. The postulated flaws are assumed to be six times as long as they are deep. Even for the largest postulated flaw of ( Ja .c.e the result shows that the flaw growth through the wall will - not occur during the 40 year design life of the plant. For smaller flaws, the flaw growth is significantly lower. For example, a postulated [ Ja,c.e inch , deep flaw will grow to [ Ja ,c.e which is less than [ Ja.c.e the wall thickness. These results also confirm operating plant experience. There have been no leaks observed in Westinghouse PWR accumulator lines. B.3 REFERENCES B-1 { Westinghouse System Standard Design Criteria 1.3, " Nuclear Steam Supply System Design Transients," Revision 2, April 15, 1974.]a,c o r)
, B-2 ASME Section III, Division I-Appendices, 1983 Edition, July 1, 1983.
B3 WECAN -- Westinghouse Electric Computer Analysis, User's Marual -- Volumes 1, !!, 111 and IV, Westinghouse Center, Pittsburgh, PA, Third Edition, 1982. B4 McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for Internal Longitudinal Semi-Elliptical Surface flaws in a Cylinder Under 2 Arbitrary Loadings", Fracture Mechaniqi ASTM STP 677, 1979, pp. 365-380. B5 Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment", ASME Tranh Journal of Pressure Vessel Technology, February 1979. O wnos2no12492:10 B9
.._... _ _ _- _ _. _ .-._._ _ _, _ _a...._ _ , . - _ _ _ _ _ - . .
t TABLE B 1 THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATI@ - , 1-
- a,c.e s ;
i e o. i i l 4 wrtossuo11492 to B 10 1
. . . _ . _ . _ _ , . . _ . . . , - . . . _ _ _ _ _ . . . , _ . .. . _c..___. . _ _ _ . _ _ . _ . - . _ . . _ . . _ . _ .
eae b [ h e h e S
- M j
~
L5 W l ZW m 4 va RW mo Wm j C4 M b WW O W W M W
~ CC
}2 W KW m
ZW ~ mO E-m Q Q,. w N E W r m Z (S D et ZW g.g M Cr J & O& EW o W
; mO J .Z Wm
)- E WW tt >
- 8. ; W O9 E V W
CC O La. W W M W W W W W 4 E CE W 7& M EW M MW Q m Z m bN OW J eW O> ZW e W Z W m . MC e gm e 1 I W
TABLE B 3
+
[LiVELOPE NORMAL LOADS a.C,0 4 W l f l i 4 I WPF1058J/011492:10 B.12
TABLE B 4 O ACCUMULATOR LINE FATIGUE CRACK GROWTH RESULTS Wall Thickness - [0.895 in.] a,c.e l l l l 1 a f 4 Q 9 4 4 WPF1058J/011492:10 B-13
~ -
l l
+
l i-1 7 a.c.c i I i
)'
i 4
.i ,
I { i 8 I h k s Figure B-1 Comparison of Typical Maximum and Minimum Stress Prgg1gs Computed by Simplified [ ] witoazuotr492:10 B-14
i. a
- a,c.e 2.
a c.c
~
a s
. Figure B 2 TypicalSchgagigof_AccumulatorLineAt[
WPF1082J/012492:10 B-15 ho usi is - . - ..
i ( F l 9 a.c.e ) i f 6 4 r
\
Note: The minimuin stress is zero through the thickness. Figure B 3 [ ]a,c.e Maximum and Minimum Stress Profiles . for Transient #10 , WPF1082J/012492:10 8-}6
I l t. 6 4.c.e e s e i i
~
J s'
, Figure B 4 [ Ja,c.e Maximum and Minimum Stress Profiles for Transient #11 w9710e2J/012492:10 8-17
e a,c.e 1 4 l Figure B-5 ( Ja,c.e Maximum and Minimum Stress Profiles for Transient #12 wioszuo12492:10 B 18
9
/
4
~
q ace n i
?
4 I I i
.~
l *' i,
$ a figure B-6 ( J .c.e Maximum and Minimum Stress Profiles for Transient #14 l g.}g WPF1082J/012497:10
. , _ _ _ _ _ . _ . . . . _ _ _ _ _ _. _ . . . . _ . , . . . . .}}