ML20073J989
| ML20073J989 | |
| Person / Time | |
|---|---|
| Site: | Farley  |
| Issue date: | 04/30/1991 |
| From: | Palusamy S, Schmertz J WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19312B431 | List: |
| References | |
| WCAP-12834, NUDOCS 9105090245 | |
| Download: ML20073J989 (87) | |
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Westinghouse Proprietary Class 3 WCAP-12834 TECHNICAL JUSTIFICATION FOR ELIMINATING PRESSURIZER SURGE LINE RUPTURE FROM THE STRUCTURAL DESIGN BASIS FOR FARLEY UNITS 1 AND 2 April 1991 D. C. Showmick S. A. Swamy Y. S. Lee
. D. E. Prager K. R. Hsu Verified: O 84 p'
Sf C. Schmertz Structural Mechanics Technology j ?
n /
Approved: /WA M b'. Y # #*
/ S'. S. Ralusamy, Manager Diagnostics and Monitoring Technology Work Performed Under Shop Order: AKIP-950 WESTINGHOUSE ELECTRIC CORPORATION Nuclear and Advanced Technology Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 e 1991 Westinghouse Electric Corp.
5242s/041991:10
APPENDIX A LIMIT MOMENT 6
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APPENDIX A LIMIT MOMENT
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FOREHORD This document contains Hestinghouse Electric Corporation proprietary information and data which has been identified by brackets. Coding associated with the brackets sets forth the basis on which the information is considered proprietary. These codes are listed with their meanings in HCAP-7211.
The proprietary information and data contained in this report were obtained at considerable Westinghouse expense and its release could seriously affect our competitive position. This information is to be withheld from public disclosure in accordance with the Rules of Practice 10 CFR 2.790 and the information presented herein be safeguarded in accordance with 10 CFR 2.903.
Withholding of this information does not adversely affect the public interest.
This information has been provided for your internal use only and should not be released to outside persons or organizations without the express written approval of Westinghause Electric Corporation. Should it become necessary to
. release this informatlon to such per.ons as part of the review procedure, p'. ease contact Westinghause Electric Corporation, which will make the
, necessary arrangements required to protect the Corporation's proprietary interests.
l 5242s/041991:10 i
.-.,- -..-.. - - ,- .-. . - . ~ ~ _ . . - . - - - - - - - - - .- - - - - -
TABLE OF CONTENTS Section lLtle Eggg
1.0 INTRODUCTION
1-1 1.1 Background 1-1 1.2 Scope and Objective 1*
1.3 References 1-3 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 2-1 2.1 Stress Corrosion Cracking 2-1 2.2 Hater Hammer 2-3 2.3 Low Cycle and High Cycle Fatigue 2-4 2.4 Summary Evaluation of Surge Line for Potential Degradation During Service 2-4
. 2.5 References 2-5 3.0 MATERIAL CHARACTERIZATION 3-1 3.1 Pipe and Held Materials 3-1 3.2 Material Properties 3-1 3.3 References 3-2 4.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 4-1 4.1 Loads for Crack Stability Analysis 4-2
- 4. 2. Loads for Leak Rate Evaluation 4-2 4.3 Loading Condition 4-2 4.4 Summary of Loads and Geometry 4 4.5 Governing Locations 4-5
- 5242s/041991:10 111 l-
TABLE OF CONTENTS (cont.).
Section Title Eage .
5.0 FRACTURE MECHANICS EVALUATION 5-1 5.1 Global failure Mechanism 5-1 5.2 Leak Rate Predictions 5-2 5.3 Stability Evaluation 5-4 5.4 References 5-5 6.0 ASSESSMENT OF FATIGUE CRACK GR0HTH 6-1 6.1 Introduction 6-1 6.2 Initial flaw Size 6-2 6.3 Results of-FCG Analysts 6-2 6.4 References 6-3 7.0 ASSESSMENT OF MARGINS 7-1 ,
8.0 CONCLUSION
S 8-1 ,
APPENDIX A Limit Moment A-1 i
5242s/041991:10 iv
. _ _ _ . _ . - _ ..._.__._.._-___.___.____.___._._.m..___.. . _._.
LIST OF TABLES
~* -
Iakle Tit 1e P_ age 3-1 Room Temperature Mechanical Properties of the Pressurizer Surge Lire Materials and Welds of Farley Unit 1 3-3 l
3-2 Room Temperature Mechanical Proper ties of the Pressurizer Surge Line Materials and Helds of Farley Unit 2 3-4 3-3 Room Temperature ASME Code Minimum Properties 3-5 3-4 Representative Tensile Properties for Farley Unit 1 3-6
. 3-5 Representative Tensile Properties for Farley Unit 2 3-7
- 3-6 Modulus of Elasticity (E) 3-8 4-1 Types of Loadings 4-6 4-2. Normal and Faulted Loading Cases for Leak-Before Break Evaluations 4-7 4-3 Associated Load Cases for Analyses 4-8 4-4 Summary of LBB Loads and Stresses by Case for Farley Unit 1 4-9 l.
4-5 Summary ofELBB Loads and Stresses by Case for Farley Unit 2 4-10 L 5242s/041991:10 V
1 LIST OF TABLES (cont.)
Table I.i tl e Ease -
i 5-1 Leakage Flaw Size Length for farley Unit 1 5-6 l l
l 5-2 Leakage flaw Size Length for Farley Unit 2 5-7 5-3 Summary of Critical Flaw Size for Farley Unit 1 5-8 5-4 Summary of Critical Flaw Size for Farley Unit 2 5-9 6-1 Fatigue Crack Growth Results for 10% of Hall Initial Flaw Size 6-4 7-1 Leakage Flaw Sizes, Critical Flaw Sizes and Margins for Farley Unit 1 7-2 ,
7-2 Leakage Flaw Sizes, Critical Flaw Sizes-and Margins .
.for Farley Unit 2 7-3 7-3 LBB Conservatisms 7-4 1
l l-L
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4 5242s/041991:10 vi
LIST OF FIGURES
. figu_r.ft lille EAggt 5
3-1 Farley Unit 1 Surge Line Layout 3-9 3-2 Farley Unit 2 Surge Line Layout 3-10 '
4-1 Farley Unit 1 Surge Line Showing the Governing Locations 4 11 4-2 Farley Unit 2 Surge Line Showing the 4-12 Governing Locations 5-1 Fully Plastic Stress Distribution 5-10 5-2 Analytical Predictions of Critical Flow Pates of Steam-Water Mixtures 5-11
, 5-3 [ ]a,c.e Pressure Ratio as a function of L/D 5-12 5-4 Idealized Pressure Drop Profile through a _
Postulated Crack 5-13 5-5 Loads Acting on the Model-at the Governing Location 5-14 5-6 Critical Flaw Size Prediction for Farley Unit 1
-Node 2203 Case D 5-15 5-7. Critical Flaw Size Prediction for Farley Unit 1 Node 2233 Case E- 5-16 5-8 Critical Flaw Size Prediction for Farley Unit 1
~
Node 2203 Case F 5-17 5242s/041991:10 vil
LIST OF FIGURES (cont.) l ELqvfte Tit 1e East .
5-9 Critical Flaw Size Prediction for Farley Unit 1 .
Node 2203 Case G 5-ll 5-10 Critical flaw Size Prediction for Farley Unit 1 Node 3214 Case 0 5-19 5-11 Critical Flaw Size Prediction for Farley Unit 1 Node 3214 Case E 5-20 5-12 Critical Flaw Size Prediction for Farley Unit 1 Node 3214 Case F 5-21 5-13 Critical Flaw Size Preiiction for farley Unit 1 Node 3214' Case C 5-22 ,
5-14 Critical Flaw Size Predh' ion for Farley Unit 2 5-23 ,
i Node 2203 Case D 5-15 Critical Flaw Size Prediction for Farley Unit 2 5-24 Node 2203 Case E l
l 5-16 Critical Flaw Size Predictica for Farley Unit 2 5-25 Node 2203 Case F 5-17 Critical Flaw Size Prediction for Farley Unit 2 5-26 Node 2203 Case G 5 Critical Flaw Size Prediction for Farley Unit 2 5-27 Node 2264. Case 0
~
5-19 Critical Flaw Size Prediction for farley Unit 2 5-28 Node 2264 Case E 5242s/041991:10 VIII l
l
LIST OF FIGURES (cont.)
Fiaure .TLue EAge 5-20 Critical flaw Size Prediction for farley Unit 2 5-29 Node 2264 Case F 5-21 Critical- Flaw Size Prediction for Farley Unit 2 5-30 Node 2264 Case G 6-1 Determination of the Effects of Thermal Stratification on Fatigue Crack Growth 6-5 6-2 Fatigue Crack Growth Methodology 6-6 6-3 Fatigue-Crack Growth Rate Curve for Austenitic
. Stainless Steel 6-7 6-4 Fatigue Crack Growth Rate Equation for Austenitic Stainless Steel 6-8 6-5 Fatigue Crack Growth Critical Locations 6-9 _
6-6 Fatigue Crack Growth Controlling Positions at each-Location- 6-10 A-1 Pipe with a Through-Wall Crack in Bending A-3 5242s/041991:10 ie
._._m__ . _ . _ - _ _ _ _ _. _. _ _ . _ _ _ _ _ _ _
l SECTION 1.0
, INTRODUCTION
. 1.1 Bnkg e nd l
The current structural design basis for the pressurizer surge line requires ]
postulating non-mechanistic circumferential and longitudinal pipe breaks.
This results in additional plant hardware (e C. pipe whip restraints and jet shields) which would mitigate the dynamic consequences of the pipe breaks.
It is, therefore, highly desirable to be rea!istic in the postulation of pipe breaks for the surge line. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that can be used for establishing that a circumferential type break will not occur within the pressurizer surge line. The evaluations considering circumferentially oriented flaws cover longitudinal cases. The pressurizer surge line is known to be subjected to thermal stratification and the effects of thermal stratification for Farley surge lines have been evaluated and documented in WCAP-12855. The results of the stratification evaluation as described in HCAP-12855 have been used in the leak-before-break evaluation presented in this report.
1.2 Scang and Objective The general purpose of this investigation is to demonstrate leak-before-break for the pressurizer surge line. The scope of this work covers the entire pressurizer surge line from the primary loop nozzle junction to the pressurizer nozzle junction. A schematic drawing of the piping system is shown in Section 3.0. The recommendations and criteria proposed in NUREG 1061 Volume 3 (1-1) are used in this evaluation. The criteria and the resulting steps of the evaluation procedure can be briefly summarized as follows:
- 1) Calculate the applied loads. Identify the locaticn at which the l* highest stress occurs.
l* 2) Identify the materials and the associated material properties.
l 5242s/041991:10 1-1
- l. -
- 3) Postulate a surface llaw at the governing location. Determine fatigue crack growth. Show that a through-wall crack will not ,
result.
- 4) Postulate a through-wall flaw at the governing location. The size of the flaw should be large enough so that the leakage is assured of detection with margin using the installed leak detection equipment when the pipe is subjected to normal operating loads. A margin of 10 is demonstrated between the calculated leak rate and the leak detection capability.
- 5) Using maximum faulted loads, demonstrate that there is a margin of at least 2 between the leakage size flaw and the critical si:e flaw.
- 6) Review the operating history to ascertain that operating experience has indicated no particular susceptibility to failure from the effects of corrosion, water hammer or low and high cycle fatigue.
- 7) For the materials actually in the plant provide the material h properties to justify that the properties used in the evaluation are ,
representative of the plant specific material.
, The flaw stability analyses is performed using the metnodology cescribed in SRP 3.6.3 (1-2).
The leak rate is calculated for the normal operating condition. The leak rate prediction model used in this evaluation is an (
l a.c.e The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through-wall flaw to normal operating loads (1-3). Surface roughness is accounted for in de'ermining the leak rato through the postulated flaw. -
5242s/042291:10 1-2 1
The computer codes used in this evaluation for leak rate and f*acture
-- mechanics calculations have been validated (bench marked).
1.3 References 1-1 Report of the U.S. Nuclear Regulatory Commission Piping Review Committee
- Evaluation of Potential for Pipe. Breaks, hDREG 1061 Volume 3, November 1984.
1-2 Statidard Review Plan; public comments solicited; 3.6.3 Leak-Before-Break Evaluation Procedures; Federal Register /Vol. 52, No.167/ Friday, August 28, 1987/ Notices, pp. 32626-32633.
'-3 NUREG/CR-3464, 1983, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks."
1-4 WCAP-12855, Structural Evaluation of the'Farley Units 1 & 2 Pressurizer Surge Lines, Considering the Effects of Thermal Stratification 5242s/041991:10 1 i
j SECTION 2.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 2.1 Stress Corrosion Cracking The Westinghouse reactor coolant system primary loop and connecting Class 1 )
lines have an operating history that demonstrates the inherent operating stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosion cracking). This operating history totals over 400 reactor-years, including five plants each having over 15 years of operation and 15 other plants each with over 10 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 in Pressurized Water Reactors (PWR's). The results of the stu'.y performed by the PCSG were presented in NUREG-0531 (Reference 2-1) entitlad " 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-corrosten cracking in the primary systems of PWRs.
5242s/041991:10 2-1
. . ~ _
Operating experience in PHRs supports this determination. To date, no stress-corrosion cracking has been reported in the primary piping or safe ,,
ends of any PHR."
During 1979, several instances of cracking in PHR 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 PHR primary coolant systems.
As stated above, for the Westinghouse 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 environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well ss other material in the system, applicable ASHE 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 redeced 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 l written specifications. Requirements on chlorides. fluorides, conductivity.
- L .and pH are included in the acceptance criteria for the piping.
l l
l l 5242s/041991:10 2-2 i
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During plant operation, the rer.d. ' c 5 W V a sh u u ry is monitored and
, maintained within very speci t i fr % F + <n nt concentrations are kept below the thresholds known +o a a nfur^ '9 d r ss corrosion cracking with
- the major water chemistry cont (Cl h saras owing included in the plant operating procedures as a condition for plant operation. For example, during normal power operation, oxygen concentration in the RCS and connecting Class I lines is expected to be in the ppb range by controlling charging flow chem-istry and maintaining hydrogen in the reactor coolant at specifled concentra-tions. Halogen concentrations are also stringently controlled by maintaining !
concentrations of chlorides and fluorides within the specified limits. This is assured by controlling charging flow chemistry. Thus during plant opera-tion, the likelihood of stress corrosion cracking is minimized.
2.2 Mater Hammer Overall, there is a low potential for water hammer in the RCS and connecting surge lines since they are designed and operated to preclude the voiding condttion in normally filled lines. The RCS and connecting surge line including piping and components, are designed for normal, upset, emergency,
. and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by 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
- .onstant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and vibration characteristics of the system and connecting surge lines. -Prooperational testing and operating experience have verified the Westinghouse approach. The operating transients 5242s/041991
- 10 2-3
1 of the RCS primary piping and connected surge lines are such that no significant water hammer can occur. ,
2.3 Low Cycle and High Cycle Qtigut ,
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 III of the ASME Code. A further evaluation of the low cycle fatigue loading is discussed in Section 6.0 as part of this study in the form of a fatigue crack growth analysis.
Pump vibrations during operation would result in high cycle fatigue loads in the piping system. During operation, an alarm signals the exceedance of the RC pump shaft vibration limits. Field measurements have been made on the reactor coolant loop piping 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. Recent field measurements on typical PHR' plants indicate vibration amplitudes less than 1 ksi. When translated to the connecting surge line, these stresses would be even lower, well below the fatigue endurance limit for the surge line material and would ,
result in an applied stress intensity tactor below the threshold for fatigue crack growth.
2.4 -Summary Evaluat_ ion of Surge Line for Potential Degradation Durina Servict There has never been any service cracking or wall thinning identified in the pressurizer surge lines of Westinghouse PHR design. Sources of such degradation are mitigated by the design, construction, inspection, and operation of the pressurizer surge piping.
There is no mechanism for water hammer in the pressurizer / surge system. The pressurizer safety and relief piping system which is connected to the top of the pressurizer could have loading from water hammer events. However, these -
loads are effectively mitigated by the pressurizer and have a negligible effect on the surge line. -
5242s/041991:10 2-4
Hall thinning by erosion and erosion-corrosion effects will not occur in the surge line due to the low velocity, typically less than 1.0 ft/sec and the material, austenitic stainless steel, which is highly resistant to these degradation mechanisms. Per NUREG-0691, a study of pipe cracking in PHR piping, only two incidents of wall thinning in stainless steel pipe were reported and these were not in the surge line. Although it is not clear from the report, the cause of the wall thinning was related to the high water velocity and is therefore clearly not a mechanism which wculd affect the surge line.
It is well known that the pressurizer surge lines are subjected to thermal stratification and the effects of stratification are particularly significant during certain modes of heatup and cooldown operation. The effects of stratification have been evaluated for the Farley plant surge lines and the loads, have been derived in HCAP-12855. These loads have been used in the leak-before-break evaluation described in this report.
- The Farley Units 1 & 2 surge line piping and associated fittings are forged product forms (see-Section 3) which are not susceptible to toughness degradation due to thermal aging.
Finally, the maximum operating temperature of the pressurizer surge piping, which is about 650'F, 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-Hater Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory _
Commiss'on February _1979.
2-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Hater Reactors,-NUREG-0691, U.S. Nuclear Regulatory Commission; September 1980.
5242s/041991:10 2-5=
I SECTION 3.0,,
. MATERIAL CHARACTERIZATION 3.1 Pioe and He_1d Materials i
The pipe materials of the pressurizer surge line for the farley Unit 1 are j SA376/TP316 and SA403/WP316, and the pipe materials for Unit 2 are SA3'i6/TP304 ,
and SA403/HP316. These are a wrought product form of the type used for the primary loop piping of several PHR plants. The surge line is connected to the primary loop nozzle at one end and the other end of the surge line is connected to the pressurizer nozzle. The surge line system does not include any cast pipe or cast fitting. The welding processes used are shielded metal arc (SMAH) and submerged arc (SAW). Held locations are identified in Figures 3-1 and 3-2. ,
In the following section the tensile properties of the materials are presented for use in the leak-before-break analyses.
3.2 Material Procerties The room temperature mechanical properties of the farley Units 1 & 2 surge line material, were obtained from the Certified Materials Test Reports and are given in Table 3-1 and 3-2. The room temperature ASHE Code minimum properties are given in Table 3-3. It is seen that the measured properties well exceed those-of the Code. The representative minimum and average tensile properties were established (see Tables 3-4 and 3-5). The material properties at temperatures (135'F, 205'F, 300*F, 330'F and and 653'F) are required for the leak rate and stability analyses discussed later. The minimum and average tensile properties were calculated by using the ratio of the ASHE Code Section III properties at the temperatures of interest stated above. Tables 3-4 and 3-S show the tensile properties at various temperatures for the Farley Units 1
& 2. The modulus of elasticity values were established at various temperatures from the ASME Code Section III (Table 3-6). In the leak-before-break evaluation, the representative minimum properties at temperature are used for 5242s/041991:10 1
I the flaw stability evaluations and the representative average properties are used for the leak rate predictions. The minimum ultimate stresses are used ,
for stability analyses. These properties are summarized in Tables 3-4 and 3-5.
3.3 Referencet 3-1 ASME Boiler and Pressure Vessel Code Section III, Division 1. Appendices July 1, 1989.
I 5242s/041991:10 3-2
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1 1ABLE 3-1
\
Room Temperr.ture Hechanical Properties of the Pressurizer Surge Line
- - Haterials and Helds of Farley Unit 1 YIELO ULT! HATE
< 10 HEAT NO./ SEA 1ALla HATERIAL 11gMGIB Sig B GTJL T [LD E g/A (psi) (psi) (%) (%)
1 SA376/TP316 J3536/9106 42,400 83,600 55.0 70.8 41,200 85,000 55.2 71.1 2 SA376/TP316 J2619/7214 42,400 86,100 52.7 68.2 42,200 86,900 54.6 67.7 1 3 SA376/TP316 J3536/9106 42,400 83,600 55.0 70.8 41,200 85,000 55.2 71.1 !
i 4 SA403/HP316 56060/LR3003 38,300 82,500 77.5 l
. i 5 SA376/TP316 J3536/9106 42,400 83,600 .. 70.8 41,200 85,000 55.2 71.1 i
t i Shop Held-(SW) - Fabricated by SHAH and SAW combination with GTAH for the root Field Held (FH) - Fabricated by GTAH (insert) and SHAH combination '
b
+ ,
I r
5242s/041991:10 3-3
.. - - .... - .. _~ . . - - - . ... . - . - ...-... - - -. - .. - -- -
h L
TABLE 3-2 Room Temperature Hechanical Properties of the Pressurizer Surge Line
' Materials and Helds of Farley Unit 2 -
YIELD ULTIMATE i 1D HEAT NO./ SERIAL NO. MATERIAL. SJRLMIB STRENGTH- ELONL RL4 (psi) (psi) (%) (%) !
1 SA376/TP304 55540/13689 45,700 88,200 57.5 72.6 42,900 83,800 59.0 73.0 t
2 SA376/TP304 55540/13689 45,700 88,200 57.5 72.6 42,900 83,800 59.0 73.0 <
3 SA376/TP304 55540/13694Y 44,100 85,400 60.2 73.4
-45,700 88,600 58.0 72.1 i
4 SA403/WP316 EMPR 34,370 79,000 61.5 75.7 5 SA376/TP304 55540 44,100 85,400 60.2 73.4 45,700 88,600 58.0- 72.1
'l Shop Weld (SW) - Fabricated by GTAW (insert) and SMAH combination t
Field Held-(FW) - Fabricated by GTAH (insert) and SMAH combination _
l l 5242s/041991:10 3-4 C _ ___. . _ . ~ . -
.._.___....__..,.,_...-_.._._______.~.._.,__...[
i l i
i l I TABLE 3-3 1.
l 1
. Room Temperature ASME Code Minimum Properties 'I
.i 3
- 1 1
Material Yield Sinn UltiN te Streli j j (psi) (psi) l l <
i i !
1
$A376/TP304 30,000 75,000 3 4 :
i J
i t f I
i 4
h i- :
t I
i 4
i l.*-
i t
5242s/041991:10 3 I I I
TABLE 3-4 Representative Tensile Properties for farley Unit 1 Minimum Tempera *ure Minimum Average Ultimate ti1111111 (*f_1 U.Old_IDSD Y1Ridlnil lRIM SA376/TP316 100 38,300 41,520 82,500 and 135 36,420 39,480 82,500 SA403/WP316 205 32,780 35,540 82,410 300 29,750 32.250 80,740 330 29,020 31,460 80.210 653 23,590 25,580 78,980 5242s/042291:10 3-6
i TABLE 3-5 c.
Representative Tensile Properties for farley Unit 2 f Minimum ,
Temperature Minimum Average Ultimate Material l'F) Yield (esi) yleid (osi) Jgti).
SA376/TP304 100 42,900 44,600 83,800 135 40,390 41,990 82,230 205 35,580 36,980 79.050 653- 25,580 26,590 70,950 SA403/HP316 100 34,370 34,370 79,000 135 32,680 32,680 79,000
. 300 26,690 26,690 77,310 330 26.040 26,040 76,810-
. 653- 21,170 21,170 75,630 r
5242s/041991:10 3-7 i
TABLE 3-6 Modulus of Elasticity (E)
Ihnenture Llkill
(*F) 100 28,138 135 27,950 205 27,600 300 27,000 330 26,850 653 25.035 5242s/041991:10 3-8
b ty. Field weld o sy. Shop yeid 4
os 44 n m' 0 \
5
\
Q % g )) C ~
/ 1 v
s r
e e
h re 3*1 par),y Unit i Surge Line Layout 3-9
5
.i FW Field W'Id
$w. shop Weld '
P 5
0$
E
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- T. ~
N
/ k v
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O w
Figure 3-2 Farley Unit 2 Surge Line Layout 3-10
SECTION 4.0 LOADS FOR FRACTURE HECHANICS ANALYSIS figures 3-1 and 3-2 show Jchematic layouts of the surge lines for farley Units )
I & 2 and ident My the weld locations. l The stresses due to axial loads and bending moments were calculated by the following equation:
o f+{ (4-1) where, ;
o - stress f - axial load H - bending moment A = metal cross-sectional area
. Z - section modulus-The bending moments for the desired loading combinations were calculated by the following equation:
Hg - (M y 2 + H 2 (4-2) where, Mg - bending moment for required loading My -- Y component of bending moment My - Z component of bending moment The axial load and bending moments for crack stability analysis and leak rate predictions are computed by the methods to be explained in Sections 4.1 and i^ 4.2 which follow. '
- a i
- 5242s/041991:10 4-1
. _ . , - . _ - . _ . _ . _ . _ _ _ _ _ . . . _ - _ _ _ . _ _ _ _ . . . . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ ~ _ _ - . _ _ _ _ . _ . _ _
4.1 Loads _for Crack Stability Anajy111 The faulted loads for the crack stability analysis were calculated by the absolute sum method as follows: .
r .
Ir ggl + lr 7g l Ir,I + lr33gl m 3)
My .
lH YDW I+IN Y TH I+IN Y SSE I (4~4)
H z
=
lHzogl+lMZ TH I+IN2SSEl (4-5)
DH = Deadweight TH = Applicable thermal load (normal or stratified)
P . Load due to internal pressure l SSE - SSE loading including seismic anchor motion 4.2 Loads for leak Rate Evahl h The normal operating loads for leak rate predictions were calculated by the algebraic sum method as follows: ,
F
=FDH + FTH + I p I4-0) -
My -
(My )DH + (My)TH I4~7}
Mg -
(MZ )DH + IN Z)TH (4-8) 1- The parameters and subscripts-are the same as those explained in Section 4.1, 1
4.3 Loadina Cond_itions Because thermal-stratification can cause large stresses at heatup and cooldown temperatures 'in the range of 455'F of the RCS fluid, a review of stresses was used to identify the worst situations for LBB applications. The loading states so identified are given in Table 4-1.
5242s/042291:10 4-2
Seven loading cases were identified for LBB evaluation as given in Table 4-2.
. Cases A, B, C are cases for leak rate calculations with the remaining cases being the corresponding faulted situations for stability evaluations.
The cases postulated for leak-before-break are summarized in Table 4-3. The cases of primary interest are the postulation of a detectable leak at normal power conditions ((case A or B) and the determination of pipe stability during the subsequent cooldown to detect and repair the leak. The combination B/F or A/F, depending on whether stratification is assumed at normal power, depicts this scenario. Case C/G has a large stratification 4T (320*F) and includes SSE. This case was postulated assuming the plant could remain in mode 5 condition with a bubble in the pressurizer (RCS < 200*F) for an extended period of time (days) for certain types of plant activities such as secondary side maintenance. Under this condition, it was postulated that the leak could be detected (Case C with large stratification AT) and that an SSE could occur coincidentally with the large stratification. This case is an extreme postulation with irdeed a very low probability of occurrence.Ja .c.e The combination [
j a .C,0 5242s/041991:10 4-3
The more realistic cases (
ja.c.e
(_
Ja .c.e The logic for this AT ( Ja .c.e is based on the following:
Actual practice, based on experience of other plants with this type of situation, indicates _that the plant operators complete the cooldown as quickly as possible once a leak in the primary system is detected. Technical
- Specifications may require cold shutdown within 36 hours4.166667e-4 days <br />0.01 hours <br />5.952381e-5 weeks <br />1.3698e-5 months <br /> but actual practico ,
is that the plant depressurizes the system as soon as possible once a primary system leak is detected. Therefore, the hot leg is generally on the w&rmer ,
side of the limits ()200'F) when the pressurizer bubble is quenched. Once the bubble is quenched, the pressurizer is cooled down fairly quickly reducing the AT in the system.
4.4 Summarv of Loads and Geometry
- The load combinations were evaluated at the various weld locations. Normal loads were determined using the algebraic sum method whereas faulted loads ,
were combined using the absolute sum method. ;
l l 5242s/041991:10 4-4 pg-. - -
.scw,., p.-e. -
l l
1 4.5 Governing Latations l
l All the welds at farley Units 1 and 2 surgelines are fabricated using the SMAH I
. and SAW procedure. The following governing locations were established for the welds.
SMAH Weld .
Node 2203 for farley Unit 1. !
Node 2203 and Node 2264 for Farley Unit 2.
$AH Held Node 3214 for farley Unit 1.
The loads and stresses at these governing locations for all the loading i combinations are shown in Tables 4-4 and 4-5.
Figare 4-1 shows the governing locations for farley Unit 1.
Figure 4-2 shows the governing Iccations for farley Unit 2.
t 4
^
5242s/041991:10 4-5
TABLE 4-1 Types of Loadings Pressure (P)
Dead Height (OH)
Normal Operating Thermal Expansion (TH)
Safe Shutdown Earthquake and Seismic Anchor Motion (SSE)"
_ _ a.C e a
SSE is used to refer to the absolute sum of these loadings, i
l l t 4-6
, TABLE 4-2
. Normal and faulted Leading Cases for Leak-Before-Break Evaluations CASE A: This is the normal operating case at an RCS temperature of 653'T consisting of the algebraic sum of the loading components due to P, DW and TH.
a,c.e CASE B:
CASE C:
CASE D: This is the faulted operating case at an RCS temperature of 653*F consisting of the absolute sum (every component load is taken as positive) of P, DW, TH and SSE, CASE E: a,c.e CASE F:
CASE G:
<m "
4-7
_ _ _ ..__..____--_____.____.__....._.-_m._____._______- -
TABLE 4-3 Associated Load Cases for Analyses ,
A/D This is heretofore standard leak-before-break evaluation.
a.c.e A/F B/E B/F a
B/G C/G" 8
These are judged to be low probability events. .
l l
4-8
TABLE 4-4 Sumary of LBB Loads and Stresses by Case for Farley Unit 1 Node Case FX (Ibs) SX IESI) NB (in-lb) Sg (psi) ST IESI) l 2203 A 214493 4281 2107831 14359 18640 2203
- - a,c.e 2203 ~
2203 D 259753 5185 2577009 17555 22739
- -- a,c.e l 2203 2203 2203 _ ,
3214 A 219513 4381 327918 2234 6615 3214
- a,c.e 3214 3214 0 253733 5065 567508 3866 8930 3214 a c.e 3214 3214 e
4 4
t 4-9
l l
\
TABLE 4-5 .
Sumary of LBB Leats and Stresses by Case for farley Unit 2 .
1 i
Node Case FX (ibs) SX (PSI) N6 (in-lb) Sg (psi) ST IP8i) 2203 A 214493 4281 2111199 14381 18663 .
~ ~
2203 e,c.e 2203 2203 0 259753 5185 2580940 17581- 22766
~ ~
2203*
a,c.e 2203 2203 2264 A 232623 46f.3 292002 1989 6632 ,
2264 a,c.e 2264 -
2264 D 243623 4863 806603 5495 10357 a,c.e 2264 2264 2264 A
4 4
4-10
r T '
i ;
?
~
PZR O
C b
s E b
,_ 3214 r O w
g Highest Stressed i Weld Location
.Z (.iAW) a 2203 h
5 s
4o
/
\ Highest Stressed
[ Weld Location (SMAW)
- <o i
j 3 I
l \!
n i
o 4
t 6 a) 2 .
cW 2 oA LM S
d(
l ed Wez ry el h a t n OA d2 sn eo ri t t Sa c
t o sL) e W hdA glM i eS liW(
R Z [
3 P 0 2
2 oO
^
/
O
& y
@g& t
- 2, @ I~ : * ?X~? E E3s g$5e r8% 8.
a
, l illlllIl : ,
SECTION 5.0
.- FRACTURE MECHANICS EVALUATION 5.1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture. One method for predicting the failure of ductile material is the ( 3a .c.e method, based on traditional plastic limit load concepts, but accounting for [ 3a ,c.e and taking into account the presence of a flaw. The flawed component is predicted to fail when the remaining
- net section reaches a stress level at which a plastic hinge is formed. The stress level at which this occurs is termed as the flow stress. [
c 3a ,c.e This methodology has been shown to be applicable to ductile piping through a large number of experiments and is used here to predict the critical flaw size in the pressurizer surge line. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 5-1) when loads are applied. The detailed development is provided in Appendix A for a through-wall circumferential flaw in a pipe section with internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by: _
( 3a ,c.e (5-1) where:
[
3a ,c.e ,
5242s/041991:10 5-1
[
Ja,c.e (5-2)
The analytical model described above accurately accounts for the internal pressure as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental rer.ults (reference 5-1). Flaw stability evaluations, using this analytical moael, are presented-in section 5.3.
5.2 Leak Rate Predictions Fracture mechanics analysis shows in general that postulated through-wall cracks in the surge line would remain stable and do not cause a gross failure of this component. However, if such a through-wall crack did exist, it would .
be desirable to detect the leakage such that the plant could be brought to a safe shutdown condition. The purpose of this section is to discuss the method which will be used to predict the flow through such a postulatec crack and present the leak rate calculation results for through-wall circumferential j cracks.
5.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure l
(causing choking) is taken into account. For long channels where the ratio of the channel length, L, to hydraulic diameter, DH, (L/D )g is greater than a a
[ J ,c.e, both [ J 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.
5242s/041991:10 5-2
At this point, the flow begins to flash and choking occurs. Pressure losses
- due_to momentum changes will dominate for [ Ja .c.e However, for large L/0Hvalues, the friction pressure drop will become important and must be considered along with the momentum losses due to flashing.
5.2.2 Calculational Method In using the [
3a .c.e ,
The flow rate through a crack was calculated in the following manner. Figure ,
5-2 from reference 5-2 was useti 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 5-3 taken from reference 5-2. For all cases considered,
, since [ Ja .c.e Therefore, this method will yield the two-phase pressure drop due to momentum effects as illustrated in figure
. 5-4. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using
[ "]c.e (5-3) where the friction factor f is determined using the [ l a.c.e The crack relative roughness, c, was-obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these a
calculations was [ J .c.e RMS.
The frictional pressure drop using Equation 5-3 is then calculated for the assumed-flow and.added to the [
Ja.c e to obtain the total pressure drop from the system under
[
consideration to the atmosphere. Thus, l'
Absolute Pressure - 14.7 = [ ]a c.e (5-4) ,
5242s/041991:10 5-3
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 receated until equation 5-4 is satisfied to within an acceptable tolerance and this results in the value of flow ,
through the crack.
5.2.3 Leak Rate Calculations Leak rate calculations were performed as a function of postulated through-wall crack length for the critical locations previously identified. The crack opening area was estimated using the method of reference 5-3 and the leak rates were calculated using the calculational methods described above. The leak rt'es were calculated using the norrnal 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 critical locations at the Farley Units 1 & 2 pressurizer surge lines are shown in Tables 5-1 and 5-2, The Farley plant RCS pressure boundary leak detection system, as documented in FSAR Section 5.2.7 and the NRC Safety Evaluation Report Section 5.6, 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.
5.3 Stability Evaluatj_qQ A typical segment of the 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 to 5-21. The critical flaw size corresponds to the intersection of this curve and the maximum load line. The critical flaw size is calculated using the lower bound base metal tensile properties established in section 3.0. -
5242s/041991:10 5-4
The welds at the location of interest (i.e. the governing locations) are SAW o and SMAH. -Therefore, "Z" factor correction for SMAH_and SAH welds were applied (references 5-5 and 5-6) as follows:
2 = 1.15 [1 + 0.013 (0.D. - 4)] (for SMAH) (5-5)
Z = 1.30 [1 + 0.010 (0,0. - 4)) (for SAH) (5-6) where OD is the outer diameter in inches. Substituting 00 - 14.00 inches, the Z factor was calculated to be 1.30 for SMAH and 1,43 for SAH. The applied loads were increased by the Z factors and the plots of limit load versus crack length were generated as shown in figure 5-6 to 5-21. Tables 5-3 and 5-4 show the summary of critical flaw sizes for farley Units 1 & 2.
5,4 References 5-1 Kanninen, M. F. et al., " Mechanical fracture Predictions for Sensitized
. Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976.
5-2 [
3a ,c.e I 5-3 Tada, H., "The Effects of Shell Corrections on Stress Intensity Factors and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section II-1,. NUREG/CR-3464, September 1983.
5-4 NRC letter from M. A. Miller to Georgia Power Company, J. P. O'Reilly, dated September 9, 1987.
5-5 ASME Code Section XI, Winter 1985 Addendum, Article IHB-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.
5242s/041991:10 5-5 l
TABLE 5-1 Leakage Flaw Size for Farley Unit 1 Node Point Load Case Temoerature Crack Length (in.
('F) (for 10 gpm leakage) a.C.e 2203 3214 e
D 5-6
TABLE 5-2 Leakage Flaw Size for Farley Unit 2 Ende Point Load Case TemoeralqLe Crack Lenath (in.)
('F) (for 10 gpm leakage) a,c.e 2203 2264 W
9 e
e 5-7
l TABLE 5-3 Summary of Critical Flaw Size for farley Unit 1 Critical Node Point Load Case Temoerature Flaw Size (in)
('F) a,c.e
~
I"*
2203 3214 .
Wu I
i i 5-8 1
TABL. 5-4 Sumary of Critical Flaw Size for Farley Unit 2 Critical Node Point Load Case Temneratur_e flaw Size (in)
(*F) a c.e 2203
-2264 e man m
e e
5-9
a,c.e 1
l l
Figure 5-1 Fully Plastic Stress Distribution )
5-10
- a,c,e
=
c.
E.-
2_
2 8
s W
U 3
~
STAGNATION ENTHALPY (102 Stu/lb) l l _
Figure 5-2 Analytical Predictions of Critical Flow Rates of Steam-Water Mixtures 5-11
- a :,e
~
o 8
e o-in w
e 3
M w
a b
.4 4
9 t:
m .
v LENGTH /DI AMETER RATIO (L/D)
Figure 5-3 [ Ja ,c.e Pressure Ratio as a function of L/0 ,
i 5-12
_, .. . . . - - . . . - . - - . = = . - ~ _ . -- . _ . _ _. .. _ .
. _ heCet a c.e
- l 1
Figure 5-4. Idealized Pressure Drop Profile Through a Postulated Crack 5-13
(T ~ %
b _ {l s
i b.
~
z 9
I I 1 1 1 l 1
1 I
l l
^
k I
I I
l
\ \
l I
i i
i !
e '
Figure 5-5. Loads Acting on the Model at the Governing Location 5-14
t d
8Ce M
.S
?
oE v
/
=
o 5 A
i, --
PI PE 0D=14. GG . T: 251 SICY:23.6 SICU=79.9 Fa=260. M=.258E+04 0
Figure 5-6. Critical Flaw Size Prediction for farley Unit 1 Node 2203 Case 0 l I
5-15 1
_1
- . _ . - . . - . - - . . . _ . . . . . . ~ - . . - . . . - . . . - - . . - . . . ,~ - . - - . - . . . , . - - . .. .
D a.C e m
% 8
.N v
5 E '
D
=
3 4 > o PIPE OD=14.99 T=1.251 SIGY=23.6 SIGu=79.e Fa=259. M=.255E+94 l
1
- l
. Figure 5-7. Critical Flaw Size Prediction for Farley Unit 1 Node 2203 Case E 5-16
a.C,e
. I
?
.5 v
Z e
t:
5A i J < .
PIPE OD=14.OO T=1.250 SICY=32.0 SIGU=82.4 Fa=53.7 M:.271E+04 9
Figure 5-8 Critical Flaw Size Prediction for Farley Unit 1 Node 2203 Case F 5-17
- . - ._ . . - ... ~_. . . . _...- -- - . - . . . . . . . . ... . . ~ . . . . . -
l l
~
~
8,C.e
..D
?
.5 w
E E -
- g. .
3 J ..
PIPE OD=14.90 T=1.250 SICY=36.4 SIGU=82.5 Fa=59.8 M=.372E+04 l
Figure 5-9 Critical Flaw Size Prediction for Farley Unit 1 l Node 2203 Case G
{
! 5-18
.- - - - . . . - _ - . _ - - . - - . - ~ . _ . - - . - . . . . - . . . . . . . - -.- --. ._.- ,
I I
l l
B,C,e
. D f
v i
=
5a PIPE OD=14.99 T=1.250 SICY=23.6 SIGU=79.9 Fa=254. M=568.
Figure 5-10 Critical Flaw Size Prediction for Farley Unit 1 Node 3214 Case D 5-10
-r- e .y --
.* _ .w-,
8.C,e
.S
?=
v f
I:
M l
PIPE OD=14.99 T=1.259 SICY=23.6 SIGU=79.9 Fa=253. M:624.
9 i
Figure 5-11 Critical Flaw Size Prediction for Farley Unit 1 l
Node 3214 Case E 5-20
- 4,C,e
'e 3I
.N-
-v 1
o t:
ac
-g.
d i . .
PIPE OD=14.99 T=1.250 SICY=29.9 SIGU=89.2 Fa=51.7 M=.167E+94 l
i .,
l l.
Figure 5-12 Critical Flaw Size Prediction for Farley Unit 1 Node 3214 Case F 5-21 1
. . _ . _ _ _ _ _ .. _ __. ~ - ._. _ _ . _ _ _ _ - _ . . - - _
l i
1
- a,C e i
S
.t
?
.c v
E E -
2: '
J , , ., ,
PIPE OD=14.90 T=1.250 SICY=29.8 L n'.d=89.7 Fa=56.8 M=.236E+04 Figure 5-13 Critical Flaw Size Prediction for farley Unit 1 Node 3214 Case G 5-22
l 7
a,C.e
.S
?
.=
w E
B t:
, E ;3
!ARE UN;;" 2 N@E 2203(SMAW) CASE D
' PIPE OD=14.90 T=1.250 SICY=25.6 SIGU=70.9 Fa=260. M=.258E+04 i
Figure 5-14 Critical Flaw Size Prediction for Farley Unit 2 Node 2203 Case D 5-23
. _ . _ . . . . . . _ _ . ~ _ . _ _ _ _ . . . . . . . . _ . . _ . _ _ _ - . . _ _ _ . . _ _ _ , _ _ _ _ . ._ _
a,C,e
.S
?
.=
v E
s 2:
W I
2 3
~
FARJ lET 2 N@E 2203(SMAW)
PIPE OD=14.99 T=1.250 SICY=25.6 SICU=79.9 CASE E Fa=259. M=.254E+94 Figure 5-15 Critical Flaw Size Prediction for Farley Unit 2 Node 2203 Case E 5-24
t
- A.Cie M
.S
?
.5 v
E g
=
1 D
=
3 PIPE OD=14.99 T=1.250 SICY=35.6 SIGU=79.1 Fa=53.7 M=.238E+04 t
O l
Figure 5-16 Critical Flaw Size Prediction for Farley Unit 2 Node 2203 Case F 5-25
i i
l 9
f I
~ 8,C,e I
l 9
b c
H 3
l l
PIPE OD=14.99 T=1.250 SICY=40.4 SIGU=82.2 Fa=59.8 M=.331E+04 l
Figure 5-17 Critical Flaw Size Prediction for farley Unit 2 Node 2203 Case G l 5-26 N _ . _ _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _
- 8.c,0
.I.'
?
.5 w
E
=
D
, 5 A
b, s ,< -
PIPE OD=14.99 T=1.250 SICY=21.2 SIGU=75.6 Fa=244. M:807.
Figure 5-18 Critical Flaw Size Prediction for Farley Unit 2 Node 2264 Case D l
5-27
~ a sCoe in
.D
?
.5 v
8
=
m 3
PIPE OD=14.99 T=1.25s SICY=21.2 SIGU=75.4 Fa=242. M=586.
Figure 5-19 Critical Flaw Size Prediction for Farley Unit 2 Node 2264 Case E 5-28 l
.. _ _ _ . _ .. - . _ . . _. _ _ = - . _ _ . . _ . _ _ _ _ _ _ _ _ _ . ._ _ _ _ _ . . _ _ _ . . _ . _ . _ _ . - . _ . . . . _ .
a.c.e '
M
.h
?
.=
v 1 6*
E:
M PIPE OD=14.99 T=1.250 SICY=26.9 SICu.?6 8 Fa=51.2 M=.156E+04 Figure 5-20 Critical Flaw Size Prediction for Farley Unit 2 Node 2264 Case F 5-29
1 l
l l
a,Cse m
-k
'.E b
c i
H f
E 3
PIPE OD=14.99-T=1.250 SICY=26.7 SICU=77.3 Fa=57.7 M=.249E+94 w
Figure 5-21 Critical Flaw Size Prediction for farley Unit 2 Node 2264 Case G 5-30
SECTION 6.0
, ASSESSMENT OF FATIGUE CRACK GROWTH
. 6.1 Introdntion To determine the sensitivity of the pressurizer surge line to the presenc3 of small cracks when subjected to the transients discussed in WCAP-12855, fatigue crack growth analyses were performed. This section summarizes the analyses I and results.
Figure 6-1 presents a general flow diagram of the overall process. The methodology consists of seven basic steps as shown in figure 6-2. Steps 1 through 4 are discussed in WCAP-12855. Steps 5 through 7 are specific to fatigue crack growth and are discussed in this section.
There is presently no fatigue crack growth rate curve in the ASME Code for 4
austenitic stainless steels in a water environment. However, a great deal of work has been done recently which supports the development of such a curve.
An extensive study was performed by the Materials Property Council Working Group on Reference Fatigue Crack Growth concerning the crack growth behavior of these steels in air environments, published in reference 6-1. A reference curve for stainless steels in air environments, based on this work. is in the 1989 Edition of Section XI of the ASME Code. This curve is shown in figure 6-3.
A compilation of data for austenttic stainless steels in a PWR water environment was made by Bamford (reference 6-2), and it was found that the effect of the environment on the crack growth rate was very small, for this reason it was estimated that the environmental factor should be set at 1.0 in the crack growth rate equation from reference 6-1. Based on these works (references 6-1 and 6-2) the fatigue crack growth law used in the analyses is as shown in figure 6-4.
1 r
5242s/041991:10 6-1 ,
,i 6.2 Initial flaw Size i ,
Various initial surface flaws were assumed to exist. The flaws were assumed to be semi-elliptical with a six-to-one aspect ratio. The largest initial .
flaw assumed to exist was one with c depth equal to 10% of the nominal wall thickness, the maximum flaw size that could be found acceptable by Section XI of the ASME Code.
6.3 Results of FCG Analysis All five locations, representing all cross sections of the surge line where thermal stratification could occur, were evaluated for fatigue crack growth.
Figure 6-5 identifies the five locations. Figure 6-6 shows the position at each location where crack growths were calculated.
l Results of the fatigue crack growth analysis are presented in table 6-1 for an initial flaw of 10% nominal wall thickness.
Conservatisms existing in the fatigue crack growth analysis are listed below.
- 1. Plant operational transient data has shown that the conventional design transients contain significant conservatisms
[
i ya ,c.e
- 4. FCG neglects subsurface fatigue usage prior to crack initiation 5242s/041991:10 6-2
. _ _ - . _ _ _ . - - - . - - _ _ - . - . . . ~ _ - . _ = - . . - _... . . - .- - - .- .___- - _
6.4 Referrates 6-1. James, L. A. and Jones, D. P., " fatigue Crack Growth Correlations for Austenitic Stainless Steel in Air," in Er.cdic11yg_C.opabilitin_10 Environ. mentally Assit1ELCatling ASME publication PVP-99, December 1985.
6-2. Bamford H. H., " fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment," 6.5MLTnnsi Journal of Pressure Vessel Technology, Feb. 1979.
l l
i 4
52425/041991:10 6-3
l TABLE 6-1 FATIGUE CRACK GROHTH RESULTS FOR 10% HALL INITIAL FLAH DEPTH e
Initial Initial final (40 yr) Final Flaw Location Position Size (in) (% Hall) Size (in) (% Hall) a,c e '
4 O
. 6-4 l
l
t Of TE AMINAfl0N 07 THE (FFICT5 0F THE RM AL $TR AT!FICAfl0N
_ a, ,u 4
4 Figure 6-1 Determination of the Effects of Thermal Stratification on Fatigue Crack Growth 6-5
l a.c.e a' I
i 1
l Figure 6-2 Fatigue Crack Growth Methodology 1
6-6
. . . n . _- - - - . _ ~ ~ _ . - _ . . , _ _ ._ _ _ . - - -. - _ - . - - - . .-- .-._~_.-._
e 4
30atd
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'/
- . .. ,e
. . . . . ,i ,y
,e \\\< > > t /
l l'
/ ///
, lfl 2 I i+ I / . ' ~{ > I ,
\_ 1 / /) // i
! !/ / /// i l / / / /// I
.i. . . / //
,,. l . W- f *,'
i
>I ll !1lI!
Ii //
l ,
,1 // !/ r/ i s ! / / // /
l t l ////
tod
, ii
'l
/ I,L w ....
, , ,,5f "l .z ll l
[ 'I V 1 'I
- i > / /I
/ / f I I/
ll / I) /
/ / /// !
/ j fj (
, ,, / / h/ ,,,
t... ,,,
..~ m Figure 6-3 Fatigue Crack Growth Rate Curve for Austenitic Stainless Steel 6-7
. _ _ _ . _ _ _ _ ~ - _ . . _ . - - - . - - - - _ _ _ _ _ - - - - _ . - ----
r l 3
h - C F S E AK .30 where i
h-CrackGrowthRateininches/ cycle l C - 2.42 x 10-20 ;
F - Frequency factor (F - 1.0 for temperature below 800'F)
S - R ratio correction (S - 1.0 for'R - 0; S l i 1.8R for 0 < R <
.8; and S - -43.35 + 57.97R for R > 0.8) ,
E - Environmental Factor (E - 1.0 for_PWR) , [
i AK - Range of stress intensity factor, in psi
- in ,
R -
The_ ratio of the minimum Kg (K'Imin) to-the maximum Kg (KImax}'
s 4
Figure 6-4.- Fatigue Crack Growth Equation for Austenitic Stainless Steel l.
5242s/042291:10 6-8 e-- r ,,. , , , , , .--,,.~r. , , , ,--.,.,.n,,.,,,,,,._,,w.-,,,.. mm, .,,-._..,,,,nn,,.,,,-..,,.,,,,w,__.,,.,,_,,-.--,,,,,.,s,,,,,~,,,,
. _ _ _ - - _ . . __ . .._.______________._____m- _ _ _ ._ . _ _ _ . _ _ . . . _ _ _ _ _ . _ _
. LOC 4
\.0C . 5
- LOC 3
, t.0C 2 LOC 1
+ p /
e h
figure 6-5. Fatigue Crack Growth Critical locations 6-9
- " '"-""W T
- 4'
i
-- a . c . e i m
P e
Figure 6-6. Fatigue Crack Growth Controlling Positions at Each Location ,
6-10
SECTION 7.0 i ASSESSMENT Of MARGINS In the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessment were performed. Margins at the critical location are summarized below:
In Secton 5,3 using the IHB-3640 approach (i.e. "Z" factor approach), the
" critical" flaw sizes at the governing locations are calculated. In Section I 5.2 the crack lengths yielding a leak rate of 10 gpm (10 times the 1eP.9 detection capability of 1.0 gpm) for the critical locations are calculated.
The leakage size flaws, the instability fiaws, and margins are given in Tables 7-1 and 7-2. The margins are the ratio of instability flaw to leakage flaw. l Tha margins for analysis combination cases A/0, ( Ja ,c.e well exceed the factor of 2. The margin for the extremely low probability event defined by [ Ja.c.e has also exceeded the factor of 2. As stated in Section 4.3, the probability of simultaneous occurrence of SSE and maximum stratification due to shutdown because of leakage is estimated to be '
very low.
. In this evaluation, the leak-before-break methodology is applied conservatively. The conservatisms used in the evaluation are summarized in Table 7-3.
9 l
i l
l 5242s/041991:10 7-1
-_.-_~_.,_.m.._.. . _ _ _ _ . , . . . , , . . . . _ _ _ . _ . . . . , _ . . . _ . , _ _ _ _ _ _ . _ . _ . . . , , _ . . . . . , _ , , . _ . _ _ , , . . , _ , - . , . _ , _
l l
TABLE 7.)
Leakage Flaw Sizes, Critical flaw Sizes and Margins for farley Unit 1 Load Critical Flaw Leakage Flaw Rods Cut __ Size (in) __ Size (in)- Mrsin 2203 A/D 13.16 3.30 3.98 a,c.e l
l 3214 _A/D 18.76 6.70 2.80 -
a.c.e
- J These are judged to be low probability events 7-2 1
TABLE 7-2 9
Leakage flaw Sizes. Critical flaw Sizes and Margins for farley Unit 2 i
Load Critical Flaw Leakage flaw Rode Cut Size (in) Size (in) - B10910 2203 A/D 12,53 3.40 3.60
~~ ~
a,c.e 2264 _ A/C 18.35 6.40 2.86 ,
8,C,9 i
l a
These are judged to be low probability events 7-3
i SECTION 8.0
, CONCLUSIONS This report justifies the elimination of pressurizer surge line pipe breaks as the structural design basis for farley Units 1 and 2 as follows: l l
- a. Stress cor osion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation,
- b. Water hammer should not occur in the RCS piping (primary loop and the attached class 1 auxiliary lines) because of system design, testing, and operational considerations. .
- c. The effects of low and high cycle fatigue on the integrity of the surge line were evaluated and shown acceptable. The effects of thermal stratification were evaluated and shown acceptable.
- d. Adequate margin exists between the leak rate of small stable flaws and the capability of Farley Units 1 and 2 reactor coolant system pressure boundary. leakage detecticn system.
- e. Ample margin exists between the small stable flaw sizes of item d and the critical flaw size.
The postulated reference flaw will be stable because of the ample margins in d and e and will leak at a detectable ratt "hich will assure a safe plant shutdown.
Based on tha above, it is concluded that pressurizer surge line breaks should not be considered in the structural design basis of Farley Units 1 & 2.
5242s/041991:10 8-1
-~ _ . _