ML20235X041
| ML20235X041 | |
| Person / Time | |
|---|---|
| Site: | Vogtle  |
| Issue date: | 07/31/1987 |
| From: | Lee Y, Roarty D, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292H540 | List: |
| References | |
| WCAP-11532, NUDOCS 8707240023 | |
| Download: ML20235X041 (79) | |
Text
WESTINGHOUSE CLASS 3
! WCAP-11532 i
TECHNICAL BASES FOR ELIMINATING PRESSURIZER SURGE LINE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR' V0GTLE UNIT 2 July 1987 S. A. Swamy D. H. Roarty Y. S. Lee A. C. Chan ,
Verified by: k[
F. 4. Witt Approved by: sb$
f 5. 5. AP lus"amy, Manager Struerural Materials Engineering -
Work Performed Under Shop Order GHFJ6502F WESTINGHOUSE ELECTRIC CORPORATION
, Generation Technology Systems Division i P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728 i
3400s:16/971087 ER7 A
MSS $85h5 L l
TABLE OF CONTENTS Section Title Pace
1.0 INTRODUCTION
1-1 l 1.1 Background 1-1 1.2 Scope and Objective 1-1 1.3 References 1-4 l
2.0 FAILURE MECHANISMS FOR FLAWED PIPES 2-1 2.1 General Considerations 2-1 2.2 Global Failure Mechanism 2-1 2.3 Local Failure Mechanism 2-2 2.4 References 2-3 3.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE 3-1 LINE AND THE REACTOR COOLANT SYSTEM l
3.1 Stress Corrosion Cracking 3-1 3.2 Water Hammer 3-3 3.3 Low Cycle and High Cycle Fatigue 3-4 3.4 Summary Evaluation of Surge Line for Potential 3-4 Degradation During Service 3.5 References 3-5 4.0 MATERIAL CHARACTERIZATION 4-1 .
4.1 Pipe, Fittings and Weld Materials 4-1 l 4.2 Tensile Properties 4-2 1
4.3 Fracture Toughness Properties 4-3 4.4 References 4-4 1
5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS 5-1 5.1 Loads for Crack Stability Analysis 5-2 5.2 Loads for Leak Rate Evaluation 5-2 5.3 Summary of Loads Geometry and Materials 5-3 i 5.4 Governing Location 5-3 1
2466s 10/07C787 jj L
TABLE OF CONTENTS (cont.)
Section Title Page .
6.0 FRACTURE MECHANICS EVALUATION 6-1 6.1 Global Failure Mechanism 6-1 6.2 Leak Rate Predictions 6-3 6.3 Local Failure Mechanism 6-6 6.4 Integrity Assessment 6-7 6.5 References 6-8 7.0 FATIGUE CRACK GROWTH EVALUATION 7-1 7.1 Thermal Transient Stress Analysis 7-1 7.2 Fatigue Crack Growth Analysis 7-5 7.3 References 7-10 8.0 ASSESSMENT OF MARGINS 8-1 .
!O CONCLUSIONS 9-1 .
APPENDIX A Limit Moment A-1 7
2480s 10/070787 jjj
LIST OF FIGURES
. Figure Title Page 2-1 Schematic of Generalized Load Deformation Behavior 2-4 4-1 Weld Identification 4-7 l
4-2 Stress-Strain Curvo (minimum properties) 4-8 i 4-3 Stress-Strain Curve (average properties) 4 -9 4-4 Comparison of as-welded SAW J-R Curves from 4 10 Various Size C(T) Specimens and Experiments (from Reference 4-5)
, 5-1 Schematic Layout of Pressurizer Surge Line 5-6 1 , [ la,c e Stress Distribution 6-10 6-2 Loads Acting on the Pipe Model 6-11 6-3 " Critical" Flaw Size Prediction 6-12 6-4 Limit Moment Prediction for 14-Inch Section 6-13 1 6-5 Finite Element Model 6-14 6-6 Boundary Conditions 6-15 6-7 Analytical Predictions of Critical Flow Rates of 6-16 Steam - Water Mixtures 1
6-8 [ ']a,c e Pressure Ratio as a 6-17 Function of L/D i
)
24ses.to/c71os7 jy N'
I LIST OF FIGURES (cont.)
- i
. k Figure Title Page .
6-9 Idealized Pressure Drop Profile Through 6-17 a Postulated Crack 7-1 Comparison of Typical Maximum and Minimum Stress 7-16 Profile Competed by Simplified [
3a,c.e 7-2 Schematic of Surge Line at [ 7-17 ja,c.e 7-3 [' ]a,c.e Maximum and Minimum Stress 7-18 Profile for Transient #4 A-1 Pipe with a Through Wall Crack in Bending A-2 2486s 1C/070787 y I
1 i
i
- LIST OF TABLES l 1
Table No. Title Pace 4-1 Room Temperature Mechanical Properties of the 4-4 Surge Line Materials and Welds of the Vogtle Unit 2 Plant 4-2 Typical Tensile Properties of SA376 TP316, SA351 4-5 CF8A and Welds of Such Material for the Primary Loop 1
1 4-3 Fracture Toughness Properties Typical of t e 4-6 Surge Line 5-1 Summary of Envelope loads 5-4 5-2 Loading Components at Governing Locations 5-5 7-11 Therkal Transients Considered for Fatigue Crack 7-11 Growth Evaluation )
7-2 Stresses for the Minor Transients (ksi) 7-12 7-3 Envelope Normal Loads 7-13 7-4 Through-Wall Stress Profile for Transients Using 7-14 Finite Element Analysis 7-5 Pressurizer Surge Line Fatigue Crack Growth Reruits 7-15 8-1 Comparison of Results vs. Criteria 8-3 24tes toec707a7 yj C_
SECTION 1.0 INTRODUCTION
- t
1.1 Background
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.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 i 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, and thereby eliminate the need for some of the plant hardware.
1.2 Scope 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 juncticn to the pressurizer nozzle junction. Schematic drawing of the piping system is shown in Section 5.0. The recommendation and criteria 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.
2486s 10 470787
- 3) Postulate a surface flaw at the governing location. Determine fatigue crack growth. Show that a through-wall crack will not ,
result.
- 4) Postulate a through-wall flaw at the governing location. The size of the flaw should be large enough so that the leakage is assured of detection with margin using the installed leak detection equioment 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 normal plus SSE loads, demonstrate that there is a margin of at least d between the leakage size flaw and the critical size flaw.
- 6) Review the operating history to ascertain that operating experience has indicated no particular susceptibility to failure from the effects of corrosion, water hammer or low and high cycle fatigue.
- 7) For the base and weld metals actually in the plant provide the material properties including toughness and tensile test data. -
Justify tnat the properties used in the evaluation are representative of the plant specific material. Evaluate long term effects such as thermal aging where applicable.
- 8) Demonstrate margin on applied load.
The flaw stability criteria proposed for the analysis will examine both the global and local stability for a postulated through-wall circumferential flaw. The glebal analysis is carried out using the { Ja,c,e method, based on traditional plastic limit load concepts, but accounting for [ Ja,c.e and taking into account the presence of a flaw. The local stability analysis can be carried out using the method described in NUREG/CR 3464 (1-2). This method is based on linear elastic fracture mechanics and it can be used up to load levels producing small plastic zote size. For higher loads, the local stability analysis is carried out by performing a static I m,wmw 12 _ _ _ _ _ _ _ _ _ _ _ _ - _ _ _ _ __ i
elastic plastic [ ]a,c.e i i of a straight piece of the pressurizer surge line pipe containing a through-wall circumferential flaw. The EPRI elastic plastic fracture handbook *
~ 1 method can also be used for this purpose. -
The leak rate is calculated for ti:e normal operating condition. The leak rate prediction model used in this evaluation is an [ l
]a,c,e The crack opening area i 1
required for calculating the leak rates is obtained by subjecting the ' postulated through-wall flaw to normal operating loads. Surface roughness is ; l accounted for in determining the leak rate through the postulated flaw. l ( i As stated earlier, the evaluations described above considering I circumferentially oriented flaws cover longitudinal cases in pipes and elbows. The likelihood of a split in the elbows is very low because of the ! fact that the elbows are [ Ja c.e and no flaws are actually l l anticipated. The prediction methods for failure in elbows are virtually the same as those for [ . I l Ja,c.e However, the elbows are [ Ja,c,e and, therefore, the 1 l probability of any longitudinal flaw existing in the surge line is much I smaller when compared with the circumferential direction. Based on the above, it is judged that circumferential flaws are more limiting than longitudinal flaws in elbows and throughout the system. The computer cc, des used in this evaluation for leak rate and fracture mechanics calculations have been validated (bench marked), l 1-3 \
1.3 References
~
1-1 Report of the U.S. Nuclear Regulatory Commission Piping Review Committee .
- Evaluation of Potential for Pipe Breaks, NUREG 1061, Volume 3, November ,
1984. 1-2 NUREG/CR-3464, 1983, "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulated Circumferential Through Wall Cracks." 1-3 Begley, J.A., et. al., " Crack Propagation Investigation Related to the Leak-Before-Break Concept for LMFBR Piping" in Proceedings, Conference on Elastic Plastic Fracture, Institution of Mechanical Engineers, London 1978. W l m ,io/onon 14
SECTION 2.0 FAILURE CRITERIA FOR FLAWED PIPES 2.1 General Considerations Active research is being carried out in industry and universities as well as other research organizations to establish fracture criteria for ductile materials. Criteria being investigated include those based on J-integral initiation toughness, equivalent energy, crack opening displacement, crack opening stretch, crack opening angle, r.et-section yield, tearing modulus and void nucleation. Several of these criteria are discussed in an ASTM publication (2-1). A practical approach based on the ability to obtain material properties and to l make calculations using the available tools was used in selecting the criteria for this investigation. The ultimate objective is to show that the pressurizer surge line containing a conservatively assumed circumferential
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l through-wall flaw is stable under the worst combination of postulated faulted and operating condition loads within acceptable engineering accuracy. With this viewpoint, two mechanisms of failure, namely, local and global failure mechanisms are considered. 2.2 Global Failure Mechanism l For a tough ductile material which is notch insensitive the global failure will be governed by plastic collapse. Extensive literature is available on this subject. A PVRC study (2-2), reviews the literature as well as data from several tests on piping components, and discusses the details of analytical I methods, assumptions and methods of correlating experiments and analysis. A schematic description of the plastic behavior and the definition of plastic load is shown-in Figure 2-1. For a given geometry and loading, the plastic load is defined to be the peak load reached in a generalized load versus displacement plot and corresponds to the point of instability. 2486: 10 t70787 2-1 m
1 A simplified version of this criterion, namely, net section yield criterion has been successfully used in the prediction of the load carrying capacity of . pipes containing gross size through-wall flaws (2-3) and was found to correlate well with' experiment. This criterion can be summarized by the - following relationship: Wa < Wn (2-1) where Wa = applied generalized load Wp = calculated generalized plastic lcad Wp represents the load carrying capacity of the cracked structure and it can be obtained by an elastic plastic finite element analysis or by empirical correlation which is based on the material flow properties as discussed in Section 6.1 2.3 Local Failure Mechanism The local mechanism of failure is primarily dominated by the crack tip . behavior in terms of crack-tip blunting, initiation, extension and finally crack instability. The material properties and geometry of the pioe, flaw - size, shape and loadings are parame+ers used in the evaluation of local failure. The stability will be assumed if the crack does not initiate at all. It has been demonstrated that the initiation toughness, measured in terms of JIN from a J-integral resistance curve, is a material parameter defining the crack initiation. If, for a given load, the calculated J-integral value is shown to be less than J of the material, then the crack will not initiate. IN If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation to show stable crack propagation. dJ E (2-2) T,pp = g l l 2486s IC'07C78' 2-2
where T,pp
= cpplied tearing modulus E = modulus of elasticity
=. flow stress = [ Ja,c.e .
of a = crack length c, y u
= yield and ultimate strength of the material respectively.
In summary, the local crack stability will be established by the two-step criteria: J<J Ic, r (2-3) T,pp < Tmat, if J 1 J Ic (2-4) 2.4 References l 2-1 J.D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia, PA 19109, November 1977. l 2-2 J. C. Gerdeen, "A Critical Evaluation of Plastic Behavior Data and a Unified Definition of Plastic Loads for Pressure Components," Welding Research Council Bulletin No. 254. 2-3 Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks, EPRI-NP-192, September 1976. me.4efonos' 2-3 \ _ _ . _ _ _ _ _ _ . _- ___ _ -. _ _ _ _ .-_
3 1
.\
Wp= PLASTIC LOAD l l o 1 3 S N 1 3 1 6 l 6 C l I I - l I AP l GENERALIZED DISPLACEMENT FIGURE 2-1 Schematic of Generalized Load-Deformation Behavior
,,o.mm.: sonne 2-4 4
SECTION 3.0 OPERATION AND STABILITY OF THE PRESSURIZER SURGE LINE AND THE REACTOR COOLANT SYSTEM 3.1 Stress Corrosion Cracking The Westinghouse reactor coolant system primary loop and 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 l 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 l ' 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 (Reference 3-1) entitled " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Light Water Reactor Plants." In that report the PCSG stated: 1 "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 cr 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. anoc*" 3-1 L
Operating experience in PWRs supports this determination. To datc. v stress-corrosion cracking has been reported in the primary piping or safe , ends of any PWR." During 1979, several instances of crccking in PWR feedwater piping led to the establishment of the third PCSG. The investigations of the PCSG reported in NUREG-0691 (Reference 3-2) further confirmed that no occurrences of IGSCC have been reported for PWR primary coolant systems. As stateo ebove, 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 as other material in the system, applicable ASME Code rules, fracture toughness, l welding, fabrication, and processing. The elements of a water environment known to increase the susceptibility of ! 1 austenitic stainless steel tc stress corrosion are: oxygen, fluorides, chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g., sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to operatien and careful control of water chemistry during plant operation are used to prevent the occurren:e 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. un,, wens 32
During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking w.th
^
the major water chemistry control standards being included in the plant operating procedures as a condition for plant operation. For example, during normal power operation, oxygen concentration in the RCS and connecting Class 1 lines is expected to be in the ppb range by controlling charging fltw chem-istry and maintaining hydrogen in the reactor coolant at specified corcentra-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. 3.2 Water 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
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condition 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 roolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow ranje by control rod position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical reactor coolant systems to verify the flow and viorstion characteristics of the system and connecting surge lines. Preoperational testing and operating experience have verified the Westinghouse approach. The operating transients 1 m e,ie m c7s7 3_3 L - - _ _ _ _ _
of the RCS primary piping and connected surge lines are such that no significant water hammer can occur. , 3.3 Low Cycle and High Cycle Fatigue , 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 ill of the ASME Code. A further evaluation of the low cycle fatigue loading is discussed in Chapter 7 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 l reactor coolant loop piping of a number of plants during hot functional j testing. Stresses in the elbow below the RC pump have been found to be very I l i small, , atween 2 and 3 ksi at the highest. 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 i stress intensity factor below the threshold for fatigue crack growth. . j 3.4 Summary Evaluation of Surge Line for Potential Degradation During Service There has never been any service cracking or wall thinning identified in the pressurizer surge lines of Westinghouse PWR design. Sources of such degradation are mitigated by the design, construction, inspection, and cperation 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 l the pressurizer could have loading from water hammer events. However, these loads are effectively filtered by the pressurizer and have a negligible effect on the surge line. 2'"'"*' 3-4
I Wall 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 3 teel, whien is highly resistant to these *
~
degradation mechanisms; Per NbREG-0691, a study of pipe cracking in PWR piping, only two incidents of wall thinning in stainless steel pipe were reported and these were not 'n 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 clee.cly not a mechanism which would affect the surge line. Flow stratification, where low flow conditions permit cold and hot water to separate into distinct layers, can cause significant thermal fatigue i loadings. This was an important issue in PWR feedwater piping where )l temperature differences of 300 F were not uncommon under certain operational l conditions. The pressurizer surge line operates with a maximum potential temperature difference of about 40*F based on hot leg and pressurizer temperatures. This small temperature difference will not produce sufficient j thermal stress to cause any significant fatigue. ; i
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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. The Vcgtle Unit 2 surge line piping and associated fittings are forged product l j forms (see Section 4) which are not susceptible to toughness degradation due to thermal aging. 1
- 3. 5 References -
3-1 Investigation and Evaluation of Stress-Corrosion Cracking in Piping of Light Water Reactor Plants, NUREG-0531, U.S. Nuclear Regulatory Commission, February 1979. 3-2 Investigation and Evaluation of Cracking Incidents in Piping in i Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980. m e " """ 3-5 L
SECTICN 4.0 MATERIAL CHARACTERIZATION 4.1 Ejoe, Fittings and Weld Materials The pipe material of the pressurizer surge line is SA 376-TP316, a wrought product form of the type used for the primary loop piping of several PWR plants. The surge line is connected to the primary loop nozzle. The nozzle material is SA182-F316N. The other end of the surge line is connected to the pressurizer nozzle (carbon steel). The piping layout includes a reducer (SA403-WP304) which connects the 16-inch diameter and the 14-inch diameter piping segments. Note that the surge line system does not include any cast pipe or cast fitting. ' The weld wire used in the shop fabrication is generally of Type 308. The field weld used 308L weld wire. The welding processes used are gas tungsten arc (GTAW), submerged arc (SAW) and shielded metal arc (SMAW). l - Weld locations and corresponding welding processes are identified in Figure 4-1. l a
~
1 In the following sections the tensile and fracture toughness properties of these materials are presented and criteria for use in the leak-before-break analyses are defined. 4.2 Tensile Properties The material certifications for the surge line were used to establish the
- tensile prcperties for the piping, fittings and welds. These properties are given in Table 4-1. The properties in this table are at room temperttare. In the leak-before-break evaluation presented later, the minimum properties at operating temperature are used for the flaw stability evaluation and average properties are used for the leak rate predictions. The viability of using such properties for the surge line is presented below.
As noted in Table 4-1, the specific room temperature properties of the surge line heats compare favorably with the properties of similar material of the a,c.e primary loops (see Table 4-2). [ ] 24Hs 10/070787 4.} E____________. - . -
1 l J 1 i [
)a,c e In brief, the following material properties are the ones used in the analyses set forth in this report.
Minimum Properties for Flaw Stability Analysis . a,c,e , J Average Properties for Leak Rate Calculations a,c e 4.3 Fracture Toughness Properties Series of fracture toughness tests on SA376-TP316 pipe material and E308 welds are reported in References 4-2 and 4-3. These data are summarized in Table 4-3. Additional toughness data for austenitic stainless steel pipe welds is provided in references 4-4, which is based on small size specimen, me, mmer 4.g
l
)
i The maximum value of J in a J-R curve depends upon the planform size of the specimen, if the thickness of the specimen is maintained constant. The J-R curves for austenitic submerged are weld compact specimen of different planform .
, sizes were compared in. Reference 4-5. Also included in the comparison was the J-R curve for a 16-inch diameter pipe. Two important conclusions can be derived i
based on this comparison (See Figure 4-4). First, increased J resulted j max ' , when larger specimen were tested (for the 9.5T specimen J,,x was about 20,000 in-lb/in ,2and for the 16-inch diameter pipe experiment the J max exceeded 25,000 in-lb/in 2). Secondly, the J-R curve for this pipe experiment followed l l the J-R curve pattern for the larger specimen (See Figure 4-4). It is also well f known that the SMAW welds have superior toughness properties as compared to the SAW welds. Therefore, the 9.5T specimen data (from Figure 4-4) can be l conservatively used for leak-before-break application to SMAW welds. From Figure 4-4 the T mat is 76, 65 and 43 corresponding to J values of 6000, 10,000 2 l and 15,000 in-lb/in respectively. The calculated values of Japplied and l T applied in fracture mechanics evaluations will be compared with the above values to demonstrate flaw stability.
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4.4 References 1 4-1 Nuclear Systems Materials Handbook, ERDA Report TID 26666, November 1975, L 1 Part I, Group 1, Section 4. J 4-2 S. S. Palusamy, " Tensile and Toughness Properties of Primary Piping Weld Metal for Use in Mechanistic Fracture Evaluation," WCAP 9787, May,1981 (Westinghouse Proprietary Class 2). 4-3 S. S. Palusamy, et al., " Mechanistic Fracture Evaluation of Reactor Coolant Pipe Containing a Postulated Circumferential Through-Wall Crack," WCAP-9558, Rev. 2, May 1982, (Westinghouse Proprietary Class 2). 4-4 Toughness of Austenitic Stainless Steel Pipe Welds, EPRI NP-4768, Electric Power Research Institute, October 1986. 4-5 G. Wilkowski, et. al., " Analysis of Experiments on Stainless Steel Flux Welds," NUREG/CR-4878, BMI-2151, April 1987. 4-3 _ - eA
n o i t c a u e d r e A 5 5 4 6 0 1 3 2 R A A A . n 2 3 0 8 / / / 4 8 6 3
% i 7 6 7 6 N N N 5 4 7 6 n
o i t a h g c n n o I l 8 5 7 6 0 0 0 0 5 5 E r F e 3 2 6 0 1 6 0 1 9 2 0 O % P 5 5 5 5 4 3 4 4 4 6 5 _ E S H _ E T _ I s 0 0 0 0 0 0 _ T F s 5 0 0 0 0 0 9 R O w e 4, 8, 4, 0, A A A 5, 4, 7 6 E o r 6 _ P S l t 2 1 1 7 / / / 9 1 7 _ O D F S 6 6 6 6 N N N 7 7 5 6 R L T P E N WA L L A D P e h 1 C N t t
- I A 2 a g 0 0 0 0 0 0 0 0 0 0 1
_ 4 N m n 0 0 0 0 0 0 0 0 0 3 8 A S T E H L I i e t r 7, 4, 9, 6, 5, 9, 5, 5, 5, 6, 2, _ L C A N l t 2 1 9 7 2 2 1 1 3 2 0 B E I U U S 8 8 7 8 9 9 9 9 9 8 9 A M R T E E E T L s R A T s _ U M G e T 0 t r A E V e t R N s S E I f 0 0 0 0 0 0 0 0 P L f d 0 0 0 0 0 0 4 2 M E E O l e 2, 2, 9, 4, A A A 5, 3, 9, 9, T G % i 2 2 2 6 / / / 7 9 6 2 R 2 Y 4 4 4 4 N N N 6 4 3 4 M U O S 8 8 O 6 6 6 6 0 8 0 4 N R 1 1 1 1 3 0 3 0 6 _ 3 3 3 3 R 8 3 8 R 3 1 l P P P P E 0 E 0 E P 3 a T T T T - 3 - 3 - W F i - - - - 9 E 4 E 9 - - _ r 6 6 6 6 - - 3 2 e 8 _ 7 7 7 7 5 4 5 4 S 0 A t 3 3 3 3 A A 4 1 a A A A A F S F 5 F A A M S S S S S A S A S S S A H A 4 H 1 7 _ r 6 6 6 0 9 5 0 2 6 7 3 e 1 1 1 7 5 1 4 7 4 0 4 _ t b 0 0 0 5 8 6 9 1 0 1 4 _ a m 4 4 4 6 5 0 1 2 - 4 6 e u J D S C 0 2 3 H N L L L J 7 S I S 1 6 6 e e e e t p p p p r " c i i i i e 4 e u p p p p c 1 l d m d d d d d u x z o r " " l l l l l d " z r o 6" 6" 6 4 e e e e e e 6 o P F 1 1 1 1 W W W h W R 1 N [ j
1 TABLE 4-2 i TYPICAL TENSILE PROPERTIES OF SA376 TP316, SA351 CF8A and WELDS OF
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SUCH MATERIAL FOR THE PRIMARY LOOP Test Temperature Average Tensile Properties J Plant Material (*F) Yield (psi) Ultimate (psi) A SA376 TP316 70 40,900 (48)a 83,200 (48) 650 23,500 (19) 67,900 (19) ) E 308 Weld 70 63,900 (3) 87,600 (3) B SA376 TP316 70 47,100 (40) 88,300 (40) l 650 26,900 (22) 69,100 (25) l ] E 308 Weld 70 59,600 (8) 87,200 (8) 650 31,500 (3) 68,800 (1) C SA376 TP316 70 46,600 (36) 87,300 (36) ' l 650 24,200 (18) 66,800 (19) E 308 Weld 70 61,900 (4) 85,400 (4) D SA351 CF8A . 70 47,300 (14) 84,500 (14) 650 26,000 (4) 70,500 (4) Weld 70 61,200 (31) 84,500 (32)
- a. ( ) indicates the number of test results averaged.
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e, c, a e, c, a
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t a m T
)
2 n i c / I b E J l N - I n L i ( E G _ R U e, S _ ) c, E i e b aj _ H s t 0 0 0 _ T p a 0 0 0 m 5, ( 1 F i , 2, _ O s t 5 0 1 _ e 1 6 6 6 _ L i U _ A t _ C r _ I e _ P p 3 Y o _ - T r 4 P _ S b E E e 0 0 0 _ L I l 0 0 0 B A T R i s d l 7 5, 0, T E n e 1 0 5 P e i 2 2 4 O T Y R P S S E . N p _ H m _ G e ) U T F 0 0 0 O " 0 0 0 T t ( 6 6 6 _ s _ E e R T U T _ C _ A . _ R s _ t F s t 6 e f o _ 6 6 _ 1 1 t _ 3 3 s _ l P P e _ a T T w _ i c o r 6 6 L _ e 7 7 d _ t 3 3 l . _ a A A e b M S S W [
~ -
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Figure 4-2 Stress-Strain Curve (minimum properties) 2480s 10/062$57 4,g 1
' J
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cc
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.'s-1 t-1 Figure 4 3 sness-5trair,icurve (kretage properties)
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Crack Extension, em 8 18 28 38 48 58 E8 78 38 i i e i e i
" j =
. = ,- 5 8
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- 3
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a D agC p 6 a IT C T) 4 o si c - 3 5 . ke o 9.5T C',T) i n 6-inch (152-mm) diometer pipe e 16-inch (406-mm) diamete, pipe i g . . . i l .5 1 1.5 2 2.5 3, , Crack Extension, anches I JD-R curves at 550 F (288 C) Figure 4-4 Comparison of as-welded SAW J-R curves from various size C(T) Specimens and Experiments (from Reference 4-5) e m . m os2sar 4-10 l 1
l SECTION 5.0 LOADS FOR FRACTURE MECHANICS ANALYSIS " l i i Figures 5-1 shows schematic layout of the surge line. I I The stresses due to axial loads and bending moments were calculated by the following equation: o=f+{ (5.1) where, o = stress F = axial load M = bending moment A = metal cross sectional area ; 2 = section modulus i The bending moments for tha desired loading combinations were calculated by the fellowing equation: l M= My 2 g 2 (5.2) l l where, M = bending moment for required loading ; My = Y ccmponent of bending moment M = Z component of bending moment Z l The axial load and bending mcments for crack stability analysis and leak rate predictions were computed by the methods explained in Sections 5.1 and 5.2. ms, auw 53 L-_____--
5.1 Loads for Crack Stability Analysis The faulted loads for the crack stability analysis were calculated by the , following equations: F = I (5.3) lFDW + FTH1 +pF l + IFSSE . N = (5.4) Y l(My)DW + (NY )TH1 I + IINY)SSEl M = l(M7)DW + (NZ)TH1 I + I(N Z)SSEj (5.5) Z Where, the subscripts of the above equations represent the following leading
- cases, DW
- deadwe'.ght TH1 = maximum thermal expansion including applicable thermal anchor motion SSE = SSE loading including seismic anchor motion P = load due to internal pressure 5.2 Loads for Leak Rate Evaluation The normal operating loads for leak rate predictions were calculated by the following equations:
1 F = FDW + FTH2 + F p (5.6) N
- Y (NY)0W + (M Y)TH2 (5.7)
M z
=
(MZ )DW + (NZ )TH2 (5.8) Where, the subscript TH2 represents normal operating thermal expansion loading. All other ' parameters and subscripts are the same as those explained in Section 5.1. e e & 51044M7 5 -2 1
5.3 Summary of Loads, Geometry and Materials The piping layout and the normal plus SSE stresses are shown in Figure 5-1. ' The envelope loads for the 16" section as well as the 14" section are provided i in Table 5-3. The cross-sectional dimensions and materials are also 3 summarized. The loading components for the envelope loads, are provided in ] Table 5-2. 5.4 Governing Location The normal plus SSE axial stresses at various locations along the pressurizer surge line have been provided in Figure 5-1, to enable a comparative j evaluation. The stresses are calculated using minimum wall thickness. The ! maximum stress (30.3 ksi) occurs at the junction of the pressurizer surge line and the primary loop nozzle. The welding process at this location is GTAW and , SMAW (Note that an in-shop weld bead with SPAW procedure was applied to the l pipe end prior to the GTAW field weld). For the 14-inch segment the maximum stress occurs at the pressurizer nozzle junction. The stress at this location is comparatively very low (19.3 ksi). The field weld operation at this location uses the GTAW procedure. The only location on the surge line where SAW welding procedure was used is at node number 2600. The maximum (normal + j SSE) stress at this location is relatively very small (16.3 ksi). l It is noted that the location of highest stress is at the junction of the primary loop nozzle and the pressurizer surge lir.e. Detailed fracture mechanics analyses will be performed at this location. Fracture mechanics assessments will be performed for the remaining two locations identified above. un.ie m n' 5-3 L
0
)
Ms 3 5 7 p 3 8 9 9 i 5 6 0 7 k 5 3 2 9 n i (
)
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)
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- L E s 5E D e V I h EN S c 0 Q 0 0 LE TA n .
B UI i 6 6 4 4 AF OD( 1 1 1 1 TO Y R A L M = A M I 6 6 6 6 U R 6 1 6 1 6 1 6 1 S E 7 3 7 3 7 3 7 3 T 3 P 3 P 3 P 3 P A A T A T A T A T M S - S - S - S - 0 0 6 E 0 0 6 0 0 1 DO 6 6 1 0 0 ON 0 0 9 9 e N 2 2 e 2 2 l l u u d d e e h N g h g c O n c n S I d i S d i T e l t n e l a a t n I t a a t o D l m r o l m r e r i N u r ep i u o p t O a o t a c C F N O c F N O e e S S N n n h O t o h t o c I s i c s i n T e t n e t I A h d a I h d a C g a c g a c 6 O i o o 4 i o o 1 L H L L 1 H L L nA o
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SECTION 6.0 FRACTURE MECHANICS EVALUATION . I 6.1 Global Failure Mechanism Determination of the conditions which lead to failure in stainless steel i should be done with plasti,c fracture methodology because of the large amount l of deformation accompanying fracture. One method for predicting the failure I of ductile material is the plastic instability method, based on traditional plastic limit load concepts, but accounting for strain hardening and taking i into account the presence of a flaw. The flawed pipe is predicted to fail ! when the remaining net section reaches a stress level at which a plastic hinge j is formed. The stress level at which this occurs is termed as the flow stress. The flow stress is generally taken as the average of the yield and ultimate tensile strength of the material at the temperature of interest. This methodology has been shown to be applicable to ductile piping through a large number of experiments and will be used here to pre' dict the critical flaw C site in the surge line. The failure criterion has been obtained by requiring equilibrium of the section containing the flaw (Figure 6-1) when loads are applied. The detailed development is provided in Appendix A for a through-wall circumfarential flaw in a pipe with internal pressure, axial ! force, and imposed bending moments. The limit moment for such a pipe is given by: a,c,e [ ] (6.1) , where: 1 1 [
=
)a,c e l .
uwo.mns 53 L
[ 1 Ja,c.e (6.2) The analytical model described above accurately accounts for the piping internal pressure as well as imposed axial force as they affect the limit moment. Good agreement was found between the analytical predictions and the experimental results (Reference 6-1). A typical segment of the pressurizer surge line under maximum loads of axial force F and bending moment M is schematically illustrated as shown in Figure 6-2. In order to calculate the critical flaw size, a plot of the limit moment versus crack length is generated as shown in Figure 6-3. The critical flaw size corresponds to the intersection of this curve and the maximum load line. . l 6.1.1 Stability Assessment for 16" Section I The " critical" flaw size is calculated to be [ Ja,c.e using the minimum tensile properties for SA376TP316 (wrought) stainless steel. Figure 1 5-1 identifies the location of the critical region. a,c.e Since W > 5533 in-kips for cracks smaller than [ ] and W, = p 5533 in-kips, the global stability criterion of Section 2.2 is satisfied. If IWB 3640 approach is used and if the material strength properties are ; conservatively assumed to be the same as the base metal properties, the critical flaw size for the weld metal would be about [ Ja,c,e , J e me. iome'" 6-2
)
6.1.2 Stability Assessment for 14" Section i The governing location in the 1d" section of the pressurizer surge line is at . the junction of the. surge line and the pressurizer nozzle. The stress at this location is significantly lower than the stress at the governing location in the 16" section. (Therefore, the governing location in the 16" section can be considered as the governing location for the entire surge line). The loads at the junction of the surge line and the pressurizer nozzle are I Fx = 239 kips and Mb = 2098 ir-kips. A plot of the limit moment versus crack length is shown in Figure 6-4. As expected, the " critical" flaw size at i this location is larger (" critical" flaw size = [ Ja,c.e inches). If IWB 3640 approach is used and if the material strength properties are conservatively assumed to be the same as the base metal properties, the critical flaw size for the weld metal would be about [ 3a,c e inches. 6.2 Leak Rate Predictions i 1 Fracture mechanics analysis shows that postulated through-wall cracks in the surge line would remain stable and not cause a gross failure of this component. 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 tSis section is to discuss the method which will be used to predict the flow through such a postulated crack and present the leak rate calculation results for through-wall circumferential cracks. 6.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure causes flashing which can result in choking. For'long channels where the ratio of the channel length L, to hydraulic diameter, Dg , WD ) is g4 greater than [ la,c.e, both [ Ja,c.e must be censidered. In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturatien pressure of the fluid. At this point, the flow begins to flash and choking occurs. Pressure losses due to momentum changes will dominate for [ Ja,c,e 244ts 10/070787 p3 L
However, for large L/D H values, friction pres?"re drop will become.important and must be considered along with the momentum losses due to flashing. 6.2.2 Calculation Methods A. Crack Opening Area The crack opening area for leak rate calculations was obtained by using the Finite Element Method. The finite element model consists of 120 elasto plastic 3-dimensional isoparametric brick elements and 890 nodes. Due to symmetry only one ' half of the circumference, i.e.,180* portion is modeled. The length of the model is 64 inches (4 times the pipe diameter) which is sufficiently long to attenuate the local effect of the crack. The l finite element model used in the analysis is shown in Figure 6-5. The material is assumed to obey von Mises yield criterion and linear isotropic hardening law. The true stress-true strain curve for the material is shown in Figure 4-3. This curve is obtained from the Nuclear Systems Materials Handbook (Reference 6-3). The stress-strain curve is approximated by multilineal representation. The finite element analysis is performed using the ADINA (Reference 6-4) computer code. The boundary conditions are described in Figure 6-6. The pipe is subjected to the internal pressure of 2235 psi. The total axial force resulting from pressure, deadweight and thermal expansion is 310 kips. A bending moment of 3685 in-kips is then superposed while the pressure and the axial load are held constant. Due to non-linear material behavior, the loads are added in incremental steps. B. Leak Rates The basic method used in the leak rate calculations is the method developed by [ 3a ,c e , The ficw rate'through a crack was calculated in the following manner. Figure 6-7 from Reference 6-5 was used to estimate the critical pressure, Pc, for the - surge line enthalpy condition and an assumed flow. Once Pc was found for a
- m,,emm 6-4 l
given mass flow, the [ Ja.c.e was found from Figure 6-8 taken from Reference 6-5. 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 6-9. Now using the assumed flow rate, G, the frictional pressure drop can be calculated using a Pf .= [ Ja,c.e (6.3) where the friction factor f is determined using the [ Ja,c.e The crack relative roughness, c, was obtained from fatigue crack data on stainless steel samples. The relative roughness value used in these calculations was [ la,c,e RMS. The frictioral pressure drop using Equation 6.3 is then calculated for the assumed flow and added to the [ Ja,c,e to obtain the total pressure drop from the primary system to the atmosphere. That is, (6.4) Absolute Pressure - 14.7 = [ Ja,c.e for a given assumed flow G. If the right-hand side of Equation 6.4 does not agree with the pressure difference between the surge line and the atmosphere, then the procedure is repeated until Equation 6.4 is satisfied to within an acceptable tolerance and this results in the flow value through the crack. This calculational procedurt. has been recommended by [ Ja,c.e for this type of [ Ja,c.e calculation. me, cerer:7 g l L
6.2.3 Leak Rate Results Leak rate calculations using finite element method were made for postulated 3 inches and 4 inches long through-wall flaws. The applied loads were normal , operating loads resulting from deadweight, pressure and thermal expansion (F, = 310 kips, Mb = 3685 in-kips). The average stress-strain curve (see Section 4) was used in the analysis. The 3 inches long thrcugh-wall flaw yielded a leak rate of [ Ja,c.e . The 4 inches long flaw yielded a leak rate of [ Ja,c.e Based on these results it was estimated that a[ Ja,c.e long postulated through-wall flaw will yield a leak rate
#' inches long.
of 10 gpm. Thus the " reference flaw" is found to bc[ ] 6.2.4 Leak Detection Capability The Vogtle Unit 2 plant leak detection system inside containment can detect I 1 gpm leak rate as required by Regulatory Guide 1.45. As seen above, a margin of 10 was applied to the leak rate to define the pressurizer surge line leakage size flaw in accordance with NUREG 1061, Volume 3. . 6.3 Local Failure Mechanism , In this section the local stability analysis is performed to show that unstable crack extension will not result for a flaw two times as long as the
" reference" flaw.
At the critical location (the junction of the hotleg nozzle and the pressurizer surge line) the (normal + SSE) outer surface axial stress, o,, is seen to be 30.3 ksi based on the minimum wall thickness. The (normal + SSE) axial force and bending moment are F, = 314 kips and Mb = 5533 in-kips. a,c.e The minimum yield strength for flaw stability analysis is[ 3 ksi (see Section 4). EP'll elastic plastic fracture headbook method is used to calculate the J applied using the normal plus SSE loads. The J applied was calculated for a [ Ja,c.e long postulated through-wall flaw (which is 2 times the reference flaw size) and was found to be[ 3. "' The applied tearing mcdulus Tapplied, was calculated, using the method of un, ,o vem l 6-6 l
I I I i I reference 6-7, corresponding to the above J applied value and was found to be about [ 3a,c.e which is significantly lower than the Tmat f 76 (corresponding to J = 6000 in-lb/in#), as discussed in Section 4. In addition, for a leakage ~ size flaw i.e., the reference flaw of [ la,c,e long, the normal plus ! SSE load was increased by a factor of d The J-T analysis gave an applied J of [ ] a,c.e and a T applied f about 16. For both the above cases the T applied is lower than Tmat (Tmat values are specified in Section 4) and therefore unstable crack propagation will not result. 6.3.1 Crack Extension Considerations l a ,c,e The crat extension corresponding to the maximum calculated J applied 3 in-lb/in , would be about 0.25 in. (see figure 4-4). If the J applied is l calculated for a larger crack length including this crack extension, the ! J applied w uld be less than[
] a,c.e The tearing modulus T applied corresponding to this increased Japplied
- uld be about 17, which is again lower than T mat. Therefore consideration of crack extension, in this case j would not change the crack stability conclusions. l 6.4 Intecrity Assessment I
Local and glebal stability analyses have been performed at the critical location (which is the junction of the pressurizer surge line and the primary i loop nozzle. The weld at this location is SMAW and GTAW (see Section 4). fiowever, crack stability has been demonstrated by conservatively using SAW weld toughness properties. (It is known that the SMAW and GTAW welds exhibit superior toughness properties as compared tc the SAW welds). As stated in Section 5, the only SAW weld location is at node 2600 (see Figure 5-1), where the (normal + SSE) stress is 16.3 ksi. Also it is noteworthy that at the critical location the maximum stress is 30.3 ksi. For this location crack stability demonstration was performed using SAW toughness properties. Thus the analysis for the critical location cover the lower stressed SAW weld location (node 2500, Figure 5-1). 2445s 10/G70787 7 (-
I The maximum stress (19.3 ksi) in the 14-inch segment occurs at the junction of ) the surge line and pressurizer nozzle (node 2900, Figure 5-1). The weld at this location is GTAW. As a matter of fact the governing location in the , 16-inch section can be considered as enveloping this location. As an alternative, an assessment can be performed as follows: The normal operating loads at this location are F x= 233 kips and Mb = 979 in-kips. Based on I available generic evaluations, the " leakage size flaw" would be about 5 inches long. Since the maximum stress at this location is less than the yield strength, the Japplied w uld be relatively small compared to the superior material toughness properties for the GTAW welds. It is judged that the J f r a 10 inches long through-wall flaw subjected to (normal + SSE) applied loads would be lower than J Ic f r GTAW welds. Also the J applied f r the leakage size flaw subjected to /_2 (normal + SSE) loads is expected to be lower than J IC. Because of the high material toughness, the critical flaw size predictions by limit moment [ ]a,c.e and IWB 3640 [ Ja c.e can be used to ascertain that adequate margins exist. Specifically, since the critical flaw size based on IWB-3640 is about 13 inches, a margin of greater than 2 exists with respect to the leakage size flaw. From Figure 6-4, the a,c.e ' limit load corresponding to the leakage size flaw is about[ .] The applied moment being 2097 in-kips, a factor of greater than 3 en applied , maximum moment is evident. 6.5 Reference 6-1 Kanninen, M. F. et al., " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976. 6-2 ASME Section Ill, Division 1-Appendices,1986 Edition, July 1,1986. . 1 1 6-3 Nuclear Systems Matericis Htndbook, ERDA Report TID 26666, November 1975, Part I, Group 1, Section 4.
\
l l l 1 l 2435: 10/t70787 g.g l i
6-4 Klaus-Jurgen Bathe, "ADINA, A Finite Elemant Program for Automatic Dynamic incremental Nonlinear Analysis," Report 92448-1, Acousting and
~
Vibration L'aboratory, Department of Mechanical Engineering M.I.T.,1975, , Revised December 1978. 6-5 [ 3a,c.e I 6-6 [ 3a,c.e , 6-7 Paris, P.C., Tada, H., "The Application of Fracture Proof Design Methods Using Tearing Instability Theory to Nuclear Piping Postulating Circumferential Through-Wall Cracks" Section 1-3, NUREG/CR-3464, September, 1983. 1 I i l i l I l :! m e, imm" 6-9 ., 1 t ..
l l t l1illijlllllJ
- e. .
- c. -
a
+
=
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- _ s i
k D . u A s s t e e t r r t S p e. c, a] .
. [
_ 1 _ i 6 a e 2 r
- u g
i F ( ( _ e . u e o ms e m eo
l
.3 m
k a c 9:' c % 6 .
~ 1 r
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. l
. e _
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. s
. d .
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.. L
_ 2
- e. 6
_ c, r e a u _ ~ . i g F
- i s
p 5 3 2, _ 2 _ =
- P
(<
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I-Illll 1 ll
4 I i a,c.e i l 4 i l l l l i Figure 6-3 " Critical" Flaw Size Prediction i I i
- l l
6-12 ues.maismo ) i 1 1
1 i a,c.e . i l l f l - L _ Figure 6-4 Limit Moment Prediction for 14-inch Section 2at t s /070787.10 6-13 .
a,c.e l l
. \
1 l Figure 6-5 Finite Element Model 2466s/070187.10 b"14 J
i 1 i i 1 i l 1 1 I 8
+a,c..
_ a 4 l l 1 i 1
. i 1
i l l l 1 \ l l l
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Figure 6-6 Boundary Conditions uu.mai.n o 6-15 L
~
.i. a,c.e E
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STAGNATION ENTHALPY (102 Stu/lb) Figure 6-7 Analyt8 cal fredktions of Critical Flow Rates of Steam-Water Mixtures I mwo7aiano 6-16
1
)
i i
\
1 1 i I 1
- a,c.e 1 l
5* Y e z w a w e
.J W
t e , i l 1 i I LENGTH / DIAMETER RATIO (L/D) I i l l Figure 6-8 [ ]8'C'8 Pressure Ratio as'a Function of L/0 u...mmn o 6-17
a,c,e
- ~.
4 a,c,e FI _ j - _ _= L/DH"# \ h L/D H M* ,
= L =
l l
- 1 Figure 6-9 Idealized Pressure Drop Profile Through a Postulated Crack un.mmeno 6-18
'l J
i 1 i f SECTION 7.'0' FATIGUE CRACK GROWTH EVALUATION . 7.1 Thermal Transient Stress Analysis i I i The thermal transient stress analysis was performed for the Vogtle Unit 2 pressurizer surge line to obtain the through wall stress profiles for use in the' fatigue crack growth analysis of Section 7.2. The through wall stress distribution for each transient was calculated for i) 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, deadreight and thermal expansion (at normal operating temperature, 653 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 sicplified methods for all minor transients. l More accurate nonlinear through weil stress distributions were developed for severe transients by [ ja,c.e l , 7.1.1 Critical Location for Fatigue Crack Growth Analysis ; The location selected for the fatigue crack growth analysis was the critical location determined from the fracture analysis. This location had the highest stress as discussed in section 5.4. This point was the weld between the reactor coolant loop hot leg nozzle and j the surge piping. 7.1.2 Design 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 c'onditions in the plant. These 24tts/070747.10 7.}
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 including the operating basis earthquake (OBE), in accordance with design specification (7-1) and the applicable system design criteria document (7-2, 7-3), were considered for this evaluation. All these were included in the final fatigue crack growth analysis. Those transients with a range of applied stress intensity (aK) less than the threshold value of aK would not result in significant fatigue crack growth. l 7.1.3 Simplified Stress Analysis l The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients. [ f ]a,c.e The inside surface stress was calculated by the following equation which is similar to the transient portion of ASME Section III NB3600, Eq. 11:
]a,c.e 54={
where, S$ = inside surface stress
+a,c.e 6
)
1 a ots/C70787.10 y 1
v. I 4 1 TA, TB = parameters defined it. Eq.11 of ASME'!4B 3600. ' AT1, AT2 These are calculated in 3-D thermal tran:,ient analysis. A y negative value of ATI gives a positive tensile stress ] AT2 i
= AT2 at inside surface.
,s For the location evaluated, the weld to the RCC nozzle, the effect of the discontinuity was negligible, i.e. , TA = TB and"a, = ab. The maximum and minimum insidh surface stresses were searched'. front theg S values calculated for J each time step of the transient solution.
The outside surface stresses corresponding to maximum and minimum inside stresses were calculated by the following equations: f I S01 = [ ] (7.5)+a.c,e, l
] (7.6)+a,c.e, S02 = [
I where,
=
S 01 utside surface stress at time t,,x 5 = utside surface istPess'at time t min
- 02
" -+a,c.e, t
All other parameters are as defined previously The material pr;operties for the surge pipe [(SA376 TP316)] and the RCL [ l I l lf h )a,c.e The values of l l 1 E and a, at rocm temperature, provide approximately the same estimation of tFe through wall thermat transient stresses as compared to normal operating temperature properties. The following veltes were conservatively used: l
+a,c.e !
1- l
~
l
'11 ,
.p i
2410s 17*717 10
- 7. 3 m _ _____
1 A comparison was made between the maximum inside and outside surface stresses calculated using simplified and finite element techniques. A detailed discussion of the finite element analysis follows in section 7.1.4. Based on' ,
~
this comparison a factor of 1.2 (20%) was conservatively applied to the stresses from the simplified analysis. The maximum and minimum linear through wall stress distrib'ution for each thermal transient was obtained by [ la,c,e The simplified analysis discussed in this section was performed for all minor thermal transients of [ Ja,c.e Nonlinear through wall stress profiles were developed for the remaining severe transients as explained in Section 7.1.4. The inside and outside surface stresses calculated by simplified methods for the minor transients are shown in Table 7-2. [ ja,c.e This figure shows that the simplified method provides more conservative crack growth. 7.1.4 Nonlinear Stress Distribution for Severe Transients [ . Ja,c.e As mentioned earlier, the surge line section near the RCL hot leg nozzle joint is the critical location for the fracture mechanics analysis and thus was selected for fatigue crack growth analysis. A schematic of the surge line geometry at this location, is shown in Figure 7-2. [ e a,c.e
] -
2'ets 'C7076710 N i_
[ l Ja,c,e Table 7-4 summarizes the through wall stresses for all transients using finite element analysis. 7.1.6 Total Stress for Fatigue crack Grewth The total through wall stress at a section was obtained by superimposing the pressure load stresses and the stresses due to deadweight and thermal expan-sion (normal operating case) on the thermal transient stresses (of Table 7-2 or the nonlinear stress distributions discussed in Section 7.1.4). Thus, the total stress for fatigue crack growth at any point is given by the following equation: Total Thermal Stress Due Stress for Transient to Due to Fatigue = + DW + + Internal (7.7) Crack Growth Thermal Pressure Expansion The envelop thermal expansion, deadweight and pressure loads for calculating the total stresses of Equation 7.7 are summarized in Table 7-3. 7.2 Fatigue Crack Growth Analysis - The fatigue crack growth analysis was performed to determine the effect of the design thermal transients, in Table 7-1. The analysis was performed fer the critical cross section of the model which is identified in Figure 7-2. Y 24mmne 7-5
i i 7.2.1 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 some 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: h=CoAK;" (7.2.1) where "Co" and the exponent "n" are material properties-, 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 minimum to maximum stress for any given transient, as will be discussed later. Fatigue crack growth properties of stainless steel in a pressurized water environment have been used in the analysis. The input required for a fatigue crack growth analysis is basically the information necessary to calculate the parameter AK 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 can be calculated by Equation (7.2.1). This increment of growth is then added to the original crack size, the AK; adjusted, and the analysis proceeds to the next transient. The procedure is continued in this manner until all the transients have been analyzed. 1 The transient sequence, used in this analysis, is as follows: The number of i occurrences of a particular transient are equally spread ever the lifetime so that the crack growth can be determined at intermediate times during the lifetime of the structure. The cycle input data is used to determine a . i
, i 248ts17G787 to J6 l
schedular distribution by dividing each transient into event categories. Specifically, the total cycles of transients are scheduled into five events per year, two events por year, one event per year, one every fourth ' year, and 1 one every eighth year. The crack size is updated by adding the incremental growth due to a transient to the initial crack size. The updated crack size is used as the initial crack size for the next transient loading. The calculations are repeated to account for all the events during a year. The. crack growth results are printed out for each year of plant life. ) l The crack tip stress intensity factors (K )y to be used in the crack growth analysis were calculated using an expression which applies for a semi-elliptic surfcce flaw in a cylindrical geometry (7-6). l The stress intensity factor expression was taken from Reference 7-6 and was calculated usirg 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) = A0+ Ag {+ A2 ({} + A3 ({)3 (7.2.2) The stress intensity factor K y(c) was calculated at the deepest point of the crack using the following expression:
+a,c e i
(7.2.3) 1 i
) '
$ A L- )
I l 1
+a,c.e l,
.l i
i I:
- )
_ Calculation of the fatigue crack growth for each cycle was then carried out j i using the reference fatigue crack growth rate law determined from ! consideration of the available data for stainless steel in a pressurized water l l enviror. ment. This law allows for the effect of mean stress or R ratio (KImin/Elmax) n the growt;l rates. The reference crack growth law for stainless steel in a pressurized water environment was taken from reference 7-7 since no code curvo is available, and . it is defined by the following equation:
. _ a,c.e , i h=, _
(7.2.4) a,c.e where: l l l l l 1 1 and R is the ratio of the minimum K to the maximum K. The environmental ! factor of 2.0 was based on recommendations from an ASME Section XI task group (7-8). , 1
)
i e l 346t s '21C781.10 ]-8
7.2.2 Results Fatigue crack growth analyses were carried out for the critical cross . \ section. The crack growth results for the maximum acceptable ASME Section XI I flaw depth is presented in Table 7-5. The postulated flaw is assumed to be six times as long as it is deep. For the postulated flaw of [ Ja,c,e the result shows that the flaw growth will be less than 60% of the total wall thickness during the 40 year design life of the plant. For smaller flaws, the flaw growth will be lower. j These results also confirm operating plant experience. There have.been no leaks observed in Westinghouse PWR pressurizer surge lines. The worst case transient AK value for the maximum crack depth is [ Ja ,c.e The minimum flow stress for the base metal at 650*F is 44,720 psi (see Section
- 4) which can be used to obtain a conservative estimate of the plastic zone ,
size. The expression for plastic zone size, rp, calculation is: (see Reference ' . 7-9) ; l 2 5 y = (AK ) P flow i l Thus, a conservative upper bound plastic zone size is calculated to be [
]a,c.e . The remaining ligament resulting from the 0.126 in, deep l end-of-fatigue-life flaw is 0.616 in. (i.e 1.415 - 0.799). Thus the plastic l
zone size is much less than the remaining ligament. l l l Based on the above it is concluded that for the Vogtle Plant Unit 2 pressurizer surge line, the fatigue crack growth results during servic'e meet the acceptance criteria. 2415s/070727 10 J.g
l I
7.3 REFERENCES
7-1 [ *
)a,c.e 7-2 [
l 3a,c.e 7-3 [ ja,c,e 7-4 ASME Section III, Division 1-Appendices, 1983 Edition, July 1, 1983. 7-5 WECAN -- Westinghouse Electric Computer Analysis, User's Manual -- Volumes I, II, III and IV, Westinghouse Center, Pittsburgh, PA, Third Edition, 1982. l 7-6 McGowan, J. J. and Raymund, M., " Stress Intensity Factor Solutions for , l Internal Longitudinal Semi-Elliptical Surface Flaws in a Cylinder Under l Arbitrary Loadings", Fracture Mechanics ASTM STP 677, 1979, pp. 365-380. 7-7 James, L. A., and Jones, D.P., " Predictive Capabilities in Environmentally l Assisted Cracking," Special Publication, PVP-Vol. 99, American Society of
. Mechanical Engineers, Nove. 1985.
7-8 " Evaluation of Flaws in Austenitic Steel Piping Section XI Task Group for j Piping Flaw Evaluation, ASME Code, ASME Trans. Journal of Pressure Vessel Technology, Volume 108, Number 3, August 1986. 7-9 Rice, J.R., ASTM STP, 1957, Volume 415, page 247. e un,com ,o 7-10
1 1 i
'l l
TABLE 7-1 { THERMAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION Trans. No. Description No. of Cycles a,c.e 1 i 4 i G G l l l l i
. l l !>
l . 9 s O E7.10 7,
1 TABLE 7-2
- 1 STRESSES F0tt THE MINOR TRANSIENTS (KSI) '
Maximum Corresponding Minimum Corresponding i Transient No. of Inside Outside Inside Outside No. Cycles Stress' Stress Stress Stress _ _a a,c.e l l 1
~
1 eum (a) Pressure transient only (b) Seismic vibration load only l e e un,mnno 7-12
TABLE 7-3 ENVELOPE NORMAL LOADS , F M Condition (kips) (in-kips) Deadweight + Thermal Expansion 21.6 36841 00 = 16.0 inches tmin = 1.415 inches Pressure = 2235 psig O I S m e w crn ia 7 13
TABLE 7-4 THROUGH-WALL STRESS PROFILE FOR TRANSII.NTS USING FINITE ELEMENT ANALYSIS + a,c.e 1 l l l
+does not include 5% margin l ues.5mer io 7 14 l
l
}
1 I TABLE 7-5 PRESSURIZER SURGE LINE FATIGUE CRACK GROWTH RESULTS , d Initial Crack Crack Length After Year Length + (in.) 10 20 30 40 a,c.e ,
- 1
+ Wall thickness = 1.415 inches
)
4 s 2480s/070787.10 7- 5
.i l
M l
+a.c.e i
.i 1
I i l I i
~
\
-i i
i 1
- - l Figure 7-1 Conparison of Typical Maximum and Minimum Stress Profile .
Computed by Simplified [ ]a,c.e j i 248!a 57C 727.10 p
9'.
~
a,c.e l 1 5URGELINE PIPE I
\
Figure 7 2 Schematic of Surge Line at [ Ja,c.e 2486s/071087;10 7,g
r
- a,c.e
_ a,c.e l l
.]
i l l
~
Figure 7.3 [ Ja,c.e Maximum and Minimum ~ 5 tress Profile for Transient #4. mworano 7-18
SECTION 8.0 ASSESSMENT OF MARGINS 2 In the preceding sections, the leak rate calculations, fracture mechanics I analysis and fatigue crack growth evaluations were performed. Margins at the critical location are summarized below: In Section 6.1 the " critical" flaw size using limit load method is calculated , to be [ la,c.e inches long. Using INB-3640 approach, the critical flaw l size at the governing location weld is found to be [ Ja,c,e inches long. In Section 6.3 it is seen that a postulated [ Ja,c,e long through-wall flaw will remain stable when subjected to normal plus SSE loads. Based on the above, the critical flaw size will of course exceed [ ).a,c.e In Section 6.2 it is shown that at the critical location, a flaw of [
]a,c.e would yield a leak rate of 10 gpm. Thus, there is a margin of at least 2 on flaw size.
In Section 6.3 it is shown that the reference flaw [ la,c.e yieldir.g a leak rate of 10 gpm would be stable when subjected to a load equal to /2 (Normal + SSE). In summary, relative to l
- 1. Loads l
The leakage-size crack will rot experience unstable crack extension even if very large loads of / 2 (No mal plus SSE) are applied.
- 2. Flaw Size
, a. A margin of at least 2 exists between the critical flaw and the flaw yielding a leak rate of 10 gpm.
O
"'"'"" 8-1
i l
- 2. If limit load is used as the basis for critical flaw size, larger j margin for global stability would result.
- 3. Leak Rate e
A margin of 10 exists for the reference flaw [ ]a,c.e between calculated leak rate and the criteria of Regulatory Guide 1.45. A summary comparison of criteria and analytical results is given in Table 8-1. The criteria are seen to be met. i l l i I l l l t i usemonno 8-2 l
l i TABLE 8-1 COMPARISON OF RESULTS VS. CRITERIA CRITERION RESULT 3 Met I
- 1. NUREG1061 Volume 3 Section5.2(b)- (Required margin of 2 demonstrated) 1 Margin on Flaw Size i
1
- 2. NUREG1061 Volume 3 Met Section 5.2(i) - (Required margin of /2 demonstrated)
Margin on Load i
- 3. NUREG 1061 Volume 3 Met Section 5.7 - (Margin of 10 on leak rate l Margin on Leak Rate demonstrated)
I
- 4. NRC criteria on alicaable Met fatigue crack growth (af < 60% wall thickness)
(Plastic zone size < remaining ligament) 4 6 mu emna g.3
SECTION 9.0 CONCLUSIONS 2 This report justifies the elimination of pressurizer surge line pipe breaks for Vogtle Unit 2 as follows: -
- a. Stress corrosion cracking is precluded by use of fracture resistant materials in the piping system and controls on reactor coolant chemistry, temperature, pressure, and flow during normal operation.
- b. Water hammer should not occur in the RCS piping (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.
- d. Ample margin exists between the leak rate of small stable flaws and the criterion of Reg. Guide 1.45.
l
- e. Ample marcin exists between the small stable flaw sizes of item d and the critical flaw.
l f. With respect to stability of the reference flaw, ample margin exists l between the maximum postulated loads and the plant specific faulted loads (i.e. Normal + SSE). The pcstulated reference flaw will be stable because of the ample margins in d, e and f and will leak at a detectable rate which will assure a safe plant shutdown. Based on the cbove, it is concluded that pressurizer surge line breaks should not be considered in the structural design basis of Vogtle Plant Unit 2. l l 24sso;7creuo g_y
1 4 N APPENDIX A l 1 4 i LIMIT MOMENT l 1 l i I I l l e i mm.mn.x,"*' A-1 l 1
i I i l l l i APPENDIX A = 4 LIMIT MOMENT 2 1 [ . l I
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i i l l l l l i e
)
a.c.e O A-2
W A i
,a,c e l
l i I i FIGURE A-1 pip ggiH A ,H, ROUGH. WALL CRACX IN BENDING - l A-3 _w}}