ML20236B412
| ML20236B412 | |
| Person / Time | |
|---|---|
| Site: | Vogtle  |
| Issue date: | 10/31/1987 |
| From: | Chang K, Lee Y, Palusamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19304B619 | List: |
| References | |
| WCAP-11584, NUDOCS 8710260163 | |
| Download: ML20236B412 (104) | |
Text
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[ WESTINGHOUSE PROPRIETARY CLASS 3 l
WCAP- 11584 1
1 TECHNICAL BASES FOR ELIMINATING ACCUMULATOR LINE RUPTURE AS THE STRUCTURAL DESIGN l BASIS FOR V0GTLE UNIT 2 October 1987
)
Y. S. Lee S. A. Swamy D. H. Roarty F. J. Witt Verified by: M'
- K. C.' Chang G ,
Approved by: y .~ [,g M 4
- 5. 5. palusamy, Manager Structural Materials Engineering ,
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Work Performed Under Shop Order GHFJ-6502F i
i i i iO - WESTINGHOUSE ELECTRIC CORPORATION
[ Generation Technology Systems Division l P.O. Box 2728
" Pittsburgh, Pennsylvania 15230-2728 an. inmue 8710260163 871016 5 DR ADOCK 0500 4
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TABLE OF CONTENTS l
Section Title Page .
1.0 INTRODUCTION
1-1 1.1 Background 1-1 1.2 Scope and Objective 1-1 1.3 References 1-4' 2.0- FAILURE CRITERIA FOR FLAWED PIPES 2-1 l 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 ACCUMULATOR LINES 3-1 3.1 Strass 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 Accumulator Line for Potential Degradation During Serivce 3-4 3.5 Assessment of Pipe Degradation or Failure from Indirect Causes 3-5 3.6 References 3-6
'4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Pipe, Fittings and Weld Materials 4-1 4.2 Tensile Properties 4-1 4.3 Fracture Toughness Properties 4-3
- , 4.4 References 4-5 A 5.0' LOADS FOR FRACTURE MECHANICS ANALYSIS 5-1 l 5.1 Loads for Crack Stability Ant. lysis 5-2 5.2 Loads for Leak Rate Evaluation 5-2 5.3 Summary of Loads Geometry and Mattrials 5-2 5.4 Governing Location 5-3 an.mno gj
TABLEOFCONTENTS(cont.)
1 Section' Title h 1
6.0 . FRACTURE MECNANICS EVALUATION 6-1 6.1 Global failure Mechanism 6-1 6.2 Leak Rate Predictions 6-2 6.2.1 General Considerations- 6-2
]
6.2.2 Calculation Method 6-3 J 6.2.3 Leak Rate Calculations 6-4 6.2.4 Leak Detection Capability, Administrative Procedures and Technical Specification Requirements 6-4 l 6.3 Stability Evaluation Using the "Z" Factor Approach 6-6 _!
6.4 Local Stability Analysis 6-7
. 6.4.1 Crack Extension Considerations 6-8 ,
6.5 Integrity Assessment 6-8 6.6 References 6-9 7.0 ASSESSMENT OF FATIGUE CRACK GROWTH 7-1 7.1 Acceptability of Fatigue Crack Growth 7-2 7.2 References 7-3 8.0 ASSESSMENT OF MARGINS 8-1
9.0 CONCLUSION
S 9-1 I
APPENDIX A Limit Moment A-1 j
. J C- APPENDIX B Fatigue Crack Growth Considerations B-1 i B.1 Thermal Transient Stress Analysis B-2 B.1.1 Critical Location for Fatigue Crack B-2 Growth Analysis nn..iomuo ggg
i TABLEOFCONTENTS.(cont.)
1 Section Title Page I
B.1.2 Design Transients B-3 l
B.1.3 Simplified Stress Analysis B-3 1 1
B.1.4 Non-linear Stress Distribution for B-6 1 Severe Transients B.1.5 OBE Loads B-6 B.1.6 Total Stress for Fatigue Crack Growth B-7 B.2 Fatigue Crack Growth Analysis B-7 l B.2.1 Analysis Procedure B-7 B.2.2 Results B-10 B.3 References B-10 4
APPENDIX C Material: Specification,and Fracture Toughness C-1 Properties of the Accumulator Tank, Nozzle, and Safe End C.1 Materials Specification C-1 C.2 Fracture Toughness C-1 l
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L LIST OF FIGURES l
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Figure Title M 1-1 Schematic' Diagram of Accumulator Line, Vogtle Unit 2 1-5 l
l 2-1 Schematic of Generalized Load Deformation Behavior 2-4 j l
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. '4 - 1 True Stress-Strain Curve For SA403-WP316 Stainless 4-14 Steel at 558'F (Minimum Yield Stress, o y = 23.84 ksi) 4-2 True Stress Strain Curve For SA403-WP316 Stainless 4-15 Steel- at 558'F (Average Yield Stress, 26.97 ksi) !
4 Lower Bound J-R Curve for SAW Welds From Reference 4-6 4-15
.. 4-4' Lower Bound J-R Curve for SAW Welds Used in the 4-17 Analysis 5-1 Schematic Layout of Accumulator Line - Loop 1 5-7 1
5-2 Schematic Layout of Accumulator Line - Loop 2 5-8 5-3 Schematic Layout of Accumulator Line - Loop 3 5-9 5-4 Schematic Layout of Accumulator Line - Loop 4 5-10 .;
6-1 ( Ja,c.e Stress Distribution 6-10
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k ' '6-2 Analytical Predictions of Critical Flow Rates of 4 Steam Water Mixtures 6-11 i 1
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j LIST OF FIGURES (cont.) l Figure- Title Page 6-3 Critical or Checked Pressure Ratio As a' Function of L/D 6-12 6-4 Idealized Pressure Drop Profile Through a Postulated Crack 6-13 6 Loads Acting on the Pipe Model at the Governing 6-14 Location 6-6 Critical Flaw Size Prediction for Base Metal Using Limit Load Approach 6-15 6-7 Z-Factor Calculations for SAW Welds to Demonstrate Margin on Flaw Size 6-16 6-8 Z-Factor Calculations for SAW Welds to Demonstrate Margin Loads 6-17 i
A-1 Pipe with a Through Wall Crack in Bending A-2 B-1 Comparison of Typical Maximum and Minimum Stress i Profile Computed by Simplified [
Ja,c.e B-16 ;
i B-2 Schematic of Accumulator Line at [RCL Cold Leg Nozzle i g ']a,c.e B-17 j i B-3 [ .rg ., ) a , c . e and Minimum Stress Profile j for Transient #10 B-18 ]
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i y LIST OF TABLES i
Table'No. ' Title Pm 4-1 Available Room Temperature Mechanical Properties of I l
?the 10 Inch Accumulator Line Materials and Welds ;
of the Vogtle Unit 2 Plant 4-7 4-2 Room Temperature Tensile Properties of the SA35/CF8A Cast 45. Degree Nozzle 4-9 i i
4-3 Typical Tensile Properties of SA376 TP316, SA351 CF8A and Welds of Such Material for the Primary j Loop 4-10 4-4 Comparison of Tensile Properties of the Accumulator ,
- Lines With Those of Typical Wrought Primary Loops and ASME Code Minimum Requirements 4-11 W
4-5 Fracture Toughness Properties Typical Accumulator
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a Line 4-12
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74 - 3 4-6 Chemistry and End of Service Life KCU Toughness
[ for Four 45 Degree Nozzles 4-13 5-1 Summary of Envelope Loads for 10 Inch Pipe 5-4 h
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Sa2 Loading Components at Governing Locations For i M .J 1 ,.
10 Inch Line 5-5 W: 1 .x .,' ?
^ 5-3 Loading Components at the Primary Loop Nozzle Junction 5-6
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s 8-1 Comparison of Results vs. Criteria 8-3
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L L B-1 . Thermal Transients Considered for Fatigue Crack C Growth Evaluation B-12 an.-imori ;. y$$$
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.e E*- LIST OF TABLES (cont.)
. Table No. Title Page
-B-2: Stresses for the Minor Transients (PSI) B-13 B-3 Envelope Normal Loads B-14 B-4 ' Accumulator Line Fatigue Crack Growth Results B-15 l
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W2sTINGHousE PROPRIETARY class 2 SECTION 1.0 INTRODUCTION
1.1 Background
The current structural design basis for the accumulator line recuires
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postulating non-mechanistic circumferential and longitudinal pipe breaks.
This results in additional plant hardware (e.g. pipe whip restraints and jet It shields) which would mitigate the dynamic consequences of the pipe breaks.
h is, therefore, highly desirable to be realistic in the postulation of pipe (
j breaks for these lines and thereby eliminate the need for some of the plant hardware. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that are used for establishing that a circumferentici type break will not occur. The evaluations considering circumferential1y oriented flaws cover longitudinal cases. The scope of the piping covered by the work is shown in Figure 1-1 and includes the entire span of the 10-inch diameter pipe from the primary loop
. junction to the accumulator tank nozzle junction (i.e. anchor to anchor).
1.2 Scope and Objective The general purpose of this investigation is to demonstrate leak-before-break for the accumulator line. Schematic drawings of the piping system are shown in Section 5.0. The recommendation and criteria proposed in NUREG 1061 Volume These criteria and resulting steps of 3 (1-1) are used in this evaluation.
the evaluation procedure can be briefly summarized as follows:
- 1) Calculate the applied loads. Identify the location at which the highest stress occurs.
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- 2) Identify the materials and the associated material properties.
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- l. WESTINGHOUSE PRoFMic!TARY class 2 i
- 3) Postulate a surface flaw at those locations having the least
- favorable combination cf stress and material properties. Determine fatigue crack growth. S 'a 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 :ritalled 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. j i
- 5) Using normal plus SSE loads, demonstrate that there is a margin of at least 2 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 f atigue.
- 7) For the base and weld metals actually in the plant provide the material properties including toughness and tensile test data.
Justify that the properties used in the evaluation are representative of the plant specific mater' .1 Evaluate long term effects such as thermal aging where applice 1a. l l
- 8) Demonstrate margin of at least 1.4 on applied load. j The flaw stability criteria proposed for the analysis examines both the global and local stability for a postulated through-wall circumferential flaw. The global analysis is carried out using the [- Ga,c,e method, l based on traditional plastic limit load concepts, but accounting for [
I; ]a,c.e and taking into account the presence of a flaw. The local stability analysis is 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 zone size. For higher loads, an. iomno 12
the local stability analysis is carried out using the EPRI elastic plastic fracture .han'dbook method. In this application, the latter method was used for local stability analysis. [
The leak ratesis calculated for the normal operating condition. The leak rate prediction'model used in this evaluation is anE
)**C The crack opening area required for calculating the leak rates is obtained by subjecting the postu- i lated through-wall flaw to normal operating loads. Surface roughness is !
accounted for in determining the leak rate through the postulated flaw. ,
.l As stated earlier, the evaluations described above considering circumferen-tially oriented flaws cover longitudinal cases in pipes and elbows. The l 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 anticipated. The prediction methods for failure in elbows are virtually the same as those for
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Ja,c,e However, the' elbows are [ Ja,c.e and, therefore, the probability of any I longitudinal flaw existing in the accumulator line is much 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. !
Several computer codes are used in the evaluations. The main-frame computer programs are under Configuration Control which has requirements conforming to Standard Review Plan 3.9.1. The fracture mechanics calculations are
, independently verified.
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1.3 References
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'* 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.
l 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.
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SECTION 2.0 FAILURE CRITERIA FOR FLAWED PIPES 1
o 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 stratch, crack opening angle, net-section yield, tearing modulus and void nucleation. Several of these criteria are discussed in an ASTM l publication (2-1). j A practical approach based on the ability to obtain material properties and to "
make calculations using' the available tools was used in selecting the criteria for this investigation. The ultimate' objective is to show that the'accumula- j tor line containing a conservatively assumed circumferential through-wall flaw 1
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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. 1 2.2 Global Failure Mechanism 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 methods,_ assumptions and methods of correlating experiments and analysis.
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A schematic description of the plastic-behavior and the definition of plast'ic
-- load is shown in Figure 2-1. For a given geometry and loading, the plastic q load is ' defined to be the peak load reached in a generalized load versus j
-displacement plot and corresponds to the point of instability. !
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 j following relationship:
Wa < Wp (2-1) )'
-where Wa = applied generalized load Wp = calculated generalized plastic load 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 !
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-2.3 Local Failure Wechanism 1 1
The local mechanism of failure is primarily dominated by the crack tip
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behavior in terms of crack-tip blunting, initiation, extension and finally ;
crack instability. The material properties and geometry of the pipe, flaw size, shape and loadings are parameters used in the evaluation of local failure..
The stability will be assumed if the crack does not initiate at all. It has i been demonstrated that the initiation toughness, measured in terms of Jg ,
h, 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 Jg ,' of the material, then the crack will not initiate.
' If. the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation-1 i
'" " "' 2-2
.. e dJ E T,pp: = g g (2-2) l l
l l
where .T,pp
= applied tearing modulus E =: modulus of elasticity of
=
flow stress = (oy + u)/2 a = crack length .]
o,o y u
= yield and ultimate strength of the material respectively. 4 4
In' summary, the local crack stability is established by the two-step criteria:
J<Jge, or (2-3)
T,pp < Tmat, if J g Jyc (2-4) 2.4 References.
. i
~2-l' J.D. Landes, et al., Editors, Elastic-Plastic Fracture, STP-668, ASTM, Philadelphia, PA 19109, November 1977.
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.
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i f-l SECTION 3.0 OPERATION AND STABILITY OF THE ACCUMULATOR LINES 1
3.1 Stress Corrosien Cracking i 1
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 corrosioncracking). This operating history totals over 400 reactor years,
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including five plants each having over 15 years of operation and 15 other 1 plants each with over 10 years of operation.
'In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the j
. 'second Pipe Crack Study Group. _(The 'first Pipe Crack Study Group established I 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: q l
"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
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.- of' producing stress-co.resion~ cracking in the primary systems of PWRs.
,0perating experience in PWRs supports this determination. To date, no stress-corrosion cracking has been reported in-the primary piping or safe ends of any PWR." .l l
During 1979, several instances of cracking in PWR feedwater piping led to the establishment of the third PCSG. The investigations of the_PCSG reported in NUREG-0691L(Reference 3-2) further confirmed that no. occurrences of IGSCC have been reported for PWR 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 pipit;g, the following three conditions must exist simultaneously: high tensile stresses, susceptible material, and a corrosive environment. Since some residual stresses and some l degree of material susceptibility exist in any stainless steel ' piping, the potential for stress corrosion is minimized by properly selecting a material timmune to SCC as well as preventing the occurrence of a corrosive l
- environment. The material specifications consider compatibility with the system's operating environment (both internal and external) as well as other material in the system, applicable ASME Code rules, fracture toughness, welding, fabrication, and processing.
The elements of a water environment known to increase the susceptibility of austenitic stainless steel to stress corrosion are: oxygen, fluorides, ,
chlorides, hydroxides, hydrogen peroxide, and reduced forms of sulfur (e.g.,
sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are f
1 used to prevent the occurrence of a corrosive environment. Prior to being put into service, the piping is cleaned internally and externally. During flushes ,
. and preoperational testing, water chemistry is controlled in accordance with
written specifications. Requirements on chlorides, fluorides, conductivity, J- and'pH are included in the acceptance criteria for the piping.
an.omem 3-2
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During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the major water chemistry control standards being included in the plant I
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 flow chem-istry and maintaining hydrogen in the reactor coolant at specified concentra-tions. Halogen concentrations are also stringently controlled by maintaining concentrations of chlorides and fluorides within the specified limits. Thus during plant operation, the likelihood of stress corrosion cracking is minimized.
3.2 Water Hammer Overall, there is a low potential for water hammer in the RCS and connecting accumulator lines since they are ogsigned and operated to preclude the voiding condition in normally filled lir.es. The RCS and connecting accumulator lines including piping and components, are designed for normal, upset, emergency, and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are stringently controlled. Temperature during normal operation is maintained within a narrow range by control rod position; pressure is controlled by pressurizer heaters and pressurizer spray also within a narrow range for steady-state conditions. The flow characteristics of the system remain constant during a fuel cycle because the only governing parameters, namely system resistance and the reactor coolant pump characteristics are controlled in the design process. Additionally, Westinghouse has instrumented typical ,
- reactor coolant systems to verify the flow and vibration characteristics of the system and connecting accumulator lines. Preoperational testing and
'" " "' 3-3
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l operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping and connected accumulator lines are such l 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 I 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 q fatigue loading is discussed in Chapter 7 as part of this study in the form of a fatigue crack growth analysis. !
High cycle fatigue loads in the system would result primarily from pump r vibrations during operation. During operation, an alarm signals the l' exceedance of the RC pump shaft vibrat bn limits. Field measurements have been made on the reactor coolant loop p uing of a number of plants during hot
.. functional testing. ' Stresses in the elbos below the.RC pump have 'been found i to be very small, between 2 and 3 ksi at tie highest. When translated to the connecting accumulator-lines, these stresses are even lower, well below the - !
j fatigue endurance limit for the accumulator line material and would result in an applied stress intensity factor below the threshold for fatigue crack ]
growth. I 3.4 Summary Evaluation of' Accumulator Line for Potential Degradation During 3 1
Service j 1
There has never been any service cracking or wall thinning identified in the accumulator lines of Westinghouse PWR design. Sources of such degradation are
. mitigated by the design, construction, inspection, and operation of the accumulator lines.
p .
1 There is no mechanism for water hammer in the accumulator piping system.
$ Wall thinning by erosion and erosion-corrosion effects will not occur in the accumulator line due to the low velocity, typically less than 10 ft/see and the material, austenitic stainless steel, which is highly resistance to these mu..umano 3-4
s L
4: degradation mechanisms. Per NUREG-0691' a study of pipe cracking in PWR
, l piping, only two incidents of wall thinning in stainless steel pipe were
'~
r'eported and these were not in the accumulator line. Although it is not clear ,
from the report, the cause of the wall' thinning was related to the high water 3
- velocity and is therefore not a mechanism which would affect the accumulator
_ line. .
Flow stratification, where low flow conditions permit cold and hot water to
! separate into' distinct layers, can cause significant thermal fatigue loadings. This was an important issue in PWR feedwater piping where temperature differences of 300*F were not uncommon under certain operational..
conditions. Stratification is believed to be important where low flow conditions and a temperature differential exist. This is not an issue in_the accumulator line where typically there is no flow during normal plant operation. During RHR operation the flow causes sufficient mixing to i
eliminate stratification.
Finally,, the maximum operating temperature of the accumulator piping, which is ;
about 560*F, is well below the temperature which would cause any creep damage in stainless steel piping.
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3.5 Assessment of Pipe Degradation or Failure From Indirect Causes Appropriate protection against the potential of pipe degradation or failure from indirect causes is provided by plant design features and by the ,
implementation of structure, system, and component design, fabrication and inspection requirements as specified in the design basis. These features and requirements are consistent with those specified in the Standard Review Pian as discussed in the Plant Vogtle FSAR as follows:
- Flood protection is discussed in FSAR section 3.4.1.
. FSAR section 3.5 describes how protection is provided against internally
{ generated missiles both inside and outside the containment.
l mwioouno 3-5
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i Fire protection-is discussed in section 9.5.1 of the FSAR.
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FSAR sections 3.9.B.3 and 5.4 discuss in detail the design and fabrication requirements of class 1, 2, and 3 components and component supports.
]
q In-service inspection and ter. ting of the reactor coolant pressure boundary is j covered in FSAR section 5.2.4 and general discussions of inspection are i provided in FSAR sections 17.1.14 and 17.2.10.
1 It can'be concluded by review of these Sections of the Plant Vogtle FSAR that 1 the required measures are taken to preclude the degradation or failure from outside sources of piping in the plant and that the methods utilized are consistent with those given in the Standard Review Plan. f 3.6 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.
y 4
3-2 Investigation and Evaluation of Cracking Incidents in Piping in Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory Commission, September 1980.
I 4
1 1 '
o an,moune 3-6 )
. SECTION 4.0
- i. MATERIAL CHARACTERIZATION l
4.1 Pipe, Fittings, and Weld Materials The pipe material of the 10-inch accumulator line extending from the RCS cold leg injection point to the first check valve is SA 376-TP316, a wrought product form of the type used for the primary loop piping of several PWR plants. The fittings are SA403-WP316 which is wrought. The pipe materials for the remaining portions of the accumulator lines, which operate at 120'F, are SA376-TP316, SA312-TP316, and SA312-TP304L. For these lines, the fittings are SA403-WP316. The weld wire used in the shop fabrication is generally low carbon 316L. The welding processes used were gas tungsten arc (GTAW),
submerged arc (SAW), gas metal ' arc (GMAW) and shielded metal arc (SMAW). The field welds used 308L weld wire. For. each accumulator line there is a 45 degree nozzle intersecting the cold leg of the primary loop. The material of these nozzles is SA351 CF8A, a cast product form.
In the following section the tensile and fracture toughness properties of the above materials are presented and criteria for use in the leak-before-breat analyses are defined.
Material properties for the accumulator tank, nozzle, and safe end are given in appendix C.
4.2 Tensile Properties The material certifications for the accumulator lines were used to establish
. the tensile properties for the piping, fittings, and welds. Those properties are given in table. 4-1 and the properties of the nozzle material, SA351 CF8A, are given in table 4-2.
sus,nwano 41
l
. The properties in tables 4-l'and 4-2 are those at room temperature. In the leak-before-break evaluation presented later, the minimum properties at l operating temperature are used for the flaw stability evaluation and average ~
properties are used for the leak rate predictions. The viability of using l such properties for the accumulator line is presented below.
[
.)a,c.e
[
i l
l Ja,c.e All the properties presented are seen to exceed the room temperature code minimum properties. Larger margins are noted when comparing ;
the experimental yield stress data with the code minimum properties.
As noted in table 4-1, the specific room temperature properties of the accumu-f later line heats compare favorably with the properties of similar material for the primary loops (see table 4-3). (
)
l l
1 Ja,c.e These properties were used for the crack stability
"'""""' 4-2
l 4
l calculations. The modulus of elasticity of 25.8 x 10 6psi was obtained from the nuclear systems material handbook (reference 4-1) for consistency with the
~
stress strain diagram which was also obtained from that reference. The stress-strain behavior (minimum properties) is shown in figure 4-1. In a similar manner, the average material properties were obtained from the properties of table 4-1. [
ja c.e In brief, the following material properties are the ones used in the analyses set forth in this report. f Minimum Properties for Flaw Stability Analysis-a c.e
\1 Average Properties for Leak Rate Calculations j i
i a,c.e 1
I l
4.3 Fracture Toughness Properties l
[.
- 1
- ').a,c.e As seen from this table, the lowest J l value for the weld. materials observed from the J-R curves was [ Jak,e '
2 in-lb/in . The T met c rresp nding to this value of J Ic was over
[ ) a.c.e The sample yielding the lowest slope of the J-R curve had a an.-i == n o 4-3
Tst f:( ). * ' The corresponding J Ie was found'to be around
[ Ja,c e.in-lb/in .2While the data on welds cited above show superior toughness properties, they were not used in the leak-be' fore-break' evaluation.
~
Instead, lower bo'und toughness data for the SAW welds were used as follows.
The SAW toughness data .from reference 4-6 is seen'to be lower bound toughness-j data. 'This data was obtained using IT specimen. The JIc was 597 in-lb/in 2~
2 and the J,,x was about 2000 in-lb/in . The J-R curve obtained from reference 4-6 is provided in figure 4-3. Similar SAW 1T specimens.were tested at Battelle 2
- Columbus Laboratories (reference 4-7).
The J;c was found to be 547 in-lb/in 2
and the J max was about 6500 in-lb/in . The J-R curve is shown in figure 4-4.
The toughness data from reference 4-6 is also plotted in the same figure 4-4.
Remarkable agreement is observed between these data sets. The lower bo'und fracture toughness from figure 4-4', for SAW welds is thus -
2 !
Jlc = 547 in-lb/in T
mat
= 89 corresponding to J of 1880 in-lb/in 2 1 .. T mat
= 81 corresponding to J of 2100 in-lb/in 2 ,
'T mat
= 76 corresponding to J of 3400 in-lb/in 2 T = 0 corresponding to J of 4550 in-lb/in 2 mat.
These properties were used for the crack stabil.ity evaluations.
Lower bound estimates for the fracture toughness of cast stainless-steel l materials,.taking thermal aging into account, are discussed in reference 4-4. !
(
),a,c,e Forged stainless steel is considered not susceptible to thermal aging for the
. applications at hand; however, thermal aging embrittlement must be considered for the cast 45' nozzle.
ja,c,e I
L I
sur, iciano 44 L
b
3 L
(
is th'elend of. service. life Charpy-U-notch' energy (KCU) followina the procedure of. 4
- b. . ' reference.4-4.- (~ i
,)a,c,e By
"~
- the criteria established in reference 4-5, the fracture toughness is at least as great .as: the toughness of (-' ]a,c.e the benchmark mat'erial of ,
E reference 4-5.
l
(' Ja,c e is the same heat which serves as a lower bound for welds as seen in table 4-5. It is then seen that for the leak-before-break analyses,
[ Ja,c,e The fracture criteria for the cast material are thus a,c,e c
')
For the portions of the accumulator lines operating at 120 F, the JIc toughness, the yield strength and'the ultimate strength will be higher than the corresponding values at 600'F.- Therefore, the properties described above
-serve as:the lower bound properties for fracture mechanics analyses.
4.4 References l
4-1 F. J.'Witt'et al., "Integerity of the Primary Piping System of j Westinghouse Nuclear. Power Plants During Pos'tulated Seismic Events,"' j WCAP-9283, March 1978. j 1
l 4-2 S. S. Palusamy, " Tensile and Toughness Properties of Primary Piping l Weld Metal for Use in Mechanistic Fracture Evaluation," WCAP 9787, May, 1981 (Westinghouse Proprietary Class'2). ;
t 4-3 S. S. Palusamy, et al., " Mechanistic Fracture Evaluation of Reactor j Coolant P_ipe Containing a Postulated Circumferential Through-Wall I Crack," WCAP-9558, Rev. 2. May 1982, (Westinghouse Proprietary j Class 2). !
i 2E:h icus? "
4-5 s i
i 4-4 W. H. Bamford, et al., "The Effects of Thermal Aging on' the Structural Integrity of Cast Stainless Steel Piping for Westinghouse Nuclear Steam Supply Systems," WCAP-10456, November, 1983 i (Westinghouse Proprietary Class 2).
4-5 F. J. Witt and C. C. Kim, " Toughness Criteria for Thermally Agd Cast Stainless Steel," WCAP 10931, Revision 1, July 1986 (Westinghouse Proprietary Class 2). :
1 4-6 Toughness of Austenitic Stainless Steel Pipe Welds, EPRI NP-4768, Electric Powar Research Institute, October 1986.
4-7 G. Wilkowski, et. al. " Analysis of Experiments on Stainless Steel Flux Welds," NUREG/CR-4878, BMI-2151, April,1987.
1 i
W i
l meioomme 4-6
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TABLE 4 i
(
ROOM TEMPERATURE TENSILE PROPERTIES 'i OF THE SA351 CF8A CAST 45 DEGREE N0ZZLE Y! ELD STRESS ULTIMATE STRENGTH ELONGATION REDUCTION 4 i
LOOP -(ksi) '(ksi)- (%) IN AREA (%) l i
1_ 38.050 - 81.950 64 76 2 ~ 38.05 83.2 65 75 3 40.0 87.05 51 60 4L 39.15- 83.9 65 74' 1
.)
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4 , J. . TYPICAL.. TENS 115 FROPERTIES OF SA376 TP315, ,SA351 CF8A and WELDS OF j .
\. sSUCH MATERIAL FOR THE PRIMAP.V LOOP ,
' 7 !
i ,.3 4,
q t[1^
't, p\ - '
{,
g a s n} a s Average Tavsile Properties u.8 '
Test Teaper.sture 7, y : '
it -
,;q o : 3 Plant. Mn.ariad (*F) Yield (psi) ,
U]timate(pql k;
\1 >q
^ . t ,<
SA376 TP315 , 3 'P3,200(48)'
A- -
70 4 40??00 (48)a ,
67,900.(19), 5 '
l'9 650 4.
,2dbd (19) s , <
)
E'308 Weld $0 r03,id0(3) 87,6CM (3) x Yy 1
. . '!t 4
., te l
.B SA376 TP316 70 47,100(40) 88,300 (40) / j.
p "
650 26,900 (22) 69,100(25)- ,M ic n
{} (
4 L
I E 308 Weld 70 59,600(8) 87,200 (ii) '
650 }1,500(1) 68,80Ck(3)- g ,
i
- l3, ,' \ A
'3 '
/J C SA376 TP316 70 ),g,,,600(36) 87,300)(36) ..,
650 '24,200 (18) 66,800 (19) .
- s -
.s .'- '
l -
})\ l V , , ; , l' E'308 Weld 70 61,$00(4) 85,400 (4) j
,\ I 3 'N' ,
i ,
i D SA3510F8A 70 t #4),300 (14) 84,500(14)
- > , lr '
- 23,000 (4)
'% 1 650 70,500 (4)
,7 i
s i
( ,1
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61,200 (31) ' 84,500 (32) s <
s I k
!t 1
4,
- a. ( ) indicates the number .of test results averaged. ,
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TABLE 4-6 CHEMISTRY AND END OF SERVICE LIFE KCU TOUGHNESS FOR FOUR 45' N022LES l
a,c.e
~_
_ )
1 6.
- Calculated and of life toughness KCU-l f
. . 1 I
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I FIGURE 4- 1 TRUE STRESS STRAIN CURVE FOR SA403-WP316 STAINLESS STEEL
- AT558'F(MINIMUMYIELDSTRESS, 23.84 Ksi) 4-14
. 4 m.-- !
_ a,c,e a
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Figure 4-3 LLMER BOUND J-R CURVE FOR SAW WELDS FROM REFERENCE 4-6 ma. mmr io 4-16 w
n
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l Figure 4-4 LOWER BOUND J-R CURVE FOR SAW WELDS USED IN THE ANALYSIS i sen.. woe.msneno 4_17 j
(
-)
SECTION 5.0 l
LOADS FOR FRACTURE MECHANICS ANALYSIS i d
4 Figures 5-1,' 5-2, 5-3, and 5-4 are schematic layouts of the four accumulator lines.
The stresses due to axial loads and bending moments were calculated by the following equation:
o=I+{ (5.1).
where,.
o '= . stress F. =~ axial load- 1 M = bending moment ,
-A' = metal cross-sectional area 2- = section modulus The bending moments for the desired loading combinations were calculated by ;
the following equation:
M=IMy2.,, g 2
(5.2) where,
'N = bending moment for required loading My = Y. component of bending moment N = 2 comp nent of bending moment Z-The axial load and bending moments for crack stability analysis and leak rate 5 predictions were computed by the methods to be explained in Sections 5.1 and 5.2.
maco mue 5-1
f
<. 1 i
5.1 Loads-for Crack Stability Analysis c,-
The faulted. loads for the crack stability analysis were calculated by the
'following equations:
F- =
lF DW I+lF THl + lF lp + IF SSE l (5.3)
M- y
-=
'l(My)DWI+ IIN Y)THI + IIM )SSEl Y
(5.4) l My =- I l(MZ)DWI # !(N Z)TH + IIN )SSEl Z
(5.5)
Where, the. subscripts of the above equations represent the following loading cases, p
DW = deadweight TH = normal thermal expansion 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:
F =
FDW +,FTH + fp (5.6)
N
- Y-(NY)DW'+ INY)TH (5.7)
N 2
(NZ)DW + (N2 )TH (5.8) 5.3 Summary of Loads, Geometry and Materials Table 5-1 provides a summary of envelope loads computed for fracture mechanics
- evaluations in accordance with the methods described in sections 5.1 and 5.2.
The cross-sectional dimen'sions and materials are also summarized. Load data
- are tabulated at the highest stressed location (Node 1042, loop 1), and the
.., second highest str'essed location (Node 1070, loop 1), in table 5-1. The loading components are provided in table 5-2.
- ""'- 5-2
j l i l
5.4 Governing Location l The normal plus SSE axial stresses along the accumulator line starting from
' the primary loop _ junction up to the accumulator tank nozzle junction were compared. _ The maximum stress occurs at location 1042 on loop 1 accumulator line. The welding process at this location is SAW. The second highest stressed location is at node 1070 on loop 1 accumulator line. These locations are identified in figure 5-1. Detailed fracture mechanics analyses were .
performed at the highest stressed location. The accumulator 11ne nozzles at the primary loop are made of cast stainless steel material SA351 CF8A. This ,
material is susceptible to thermal aging degradation. Therefore, the local crack stability analysis is performed at the junction of the primary loop nozzle and the accumulator line accounting for the effects of thermal aging.
The loads for this location are also provided in table 5-1 and the corresponding loading components are provided in table 5-3.
~
e an.mo mmno 5-3
(
)
. s p
Fi 8 1 9' 1 0 6
3 2
k 6 2 6 2
( 1 1 1 1 1 1
)
s E e 9 9 9 9 DAh 5 5 9 9 5 5 II c 9 9 5 5 SD n 9 9 9 9 N i 8 8 I ( 8 8 8 8
)
M - s ULK e 5 5 - 5 5 5 5 MLCh 5 5 5 5 5 5 I AI c 9 9 9 9 9 9 NWH n 8 8 8 8 8 8 I Ti M ( 0 O 0 0 0 0
)
L s ALK e NLCh 0 0 0 0 E IAI c 0 0 0 0 0 0 P MWHn 0 0 0 0 0 0 I O Ti 1 1 P N ( 1 1 1 1 H
C E N L I U D
0 E 1 H 0 0 0 0 0 0 C 4 4 4 4 4 1
4 1
R S 1 1 1 1 O
F 1S )
- D E s 5A D e 5 5 5 5 O I h 5 5 EL S c 7 7 7 7 7 7 L tan 0 0 BE UI i 0 0 0 0 AP OD( 1 1 1 1 1 1
. TO .
L E
V N L E A I
6 6 F R 36 36 61 61 1A 1A O E 01 01 73 73 58 58 T 43 43 3P 3P 3F 3F Y A AP AP AT AT AC AC R M SW SW S- S- S- S -
A W
l P U O S O 1 1 1 L 1 1 1 E 2 2 0 0 0 0 DO 4 4 7 0
7 0
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mr re l
u mr re l
u mr re N op op O a op a a C F NO F NO F NO w N n n d n O t o t o yn o I s i s i raei T e t e t a mpzc l t A hda thda C gac xgac i ozn O i oo eioo roou L HLL NHLL PLNJ ye
t gf n( 4 7 7 i 5 8 2 dZ 2 2, 4, nN -
)
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I
( N 0
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T b A l 2C -
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. AN e 0 5 TR Bt 2 1 E ) n V 1 e O - m G p o o N T o A L d ,)
S a , b T o2 l N l4 -
E 0 t N t1 gf O s n( 3 6 7 P e- i 1 8 1 M h dY 3, 2, 9, O gn nN -
C i o e 1 9 5 Hi Bt - 6 2 G t n N a e I c m D o o A L N O (
L
)
b 0 0 0 2 l 6 7 0 7 l(
a 1
5, 4, 8, ie 6 7 3
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F t
e o l r M t a u d g h m r
s s
+ .
de h ap a i e e E c oy e e h r S n LT DW T P S A w&
TABLE 5-3 LOADING COMPONENTS AT THE PRIMARY LOOP N0ZZLE JUNCTION
~
(Location - 1040, Loop-1)
Load Axial Bending Bending Type Force (1b) Moment MY (ft-lb) Moment M2 (ft-lb)
Dead -390 -415 -2,679 Weight Thermal -13,860 -7,403 -41,770 Pressure ~ 137,400 - -
SSE + '8,270 31,442 13,600 Anch. Mot. I e
s e
4 9
4 -
nu,n n. 5-6
i Node 1042: Highest Load for the 10 in. Pipe o Node 1070: Next Highest Load for the 10 in. Pipe e 1
- k
~
4 toe \0go
$ (*'
o, N
1 i
I
. s !
~
Aco I
I I
l I
FIGURE 5-1 Schematic Layout of Accumulator Line - Loop 1 I A
5-7
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{
L SECTION 6.0 FRACTURE MECHANICS EVALUATION -
6.1 : 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 fer predicting the failure b of: ductile material is the [ Ja c.e method, based on Traditional plastic' l mit i load concepts, but. accounting for (.
Ja c.e and taking into account the presence of a' flaw.. The flawed pipe-is predicted to fail when-the remaining net section reaches a stress level at which a plastic' hinge is formed. The stress level at which this occurs is termed as the flow stress. [
-.: .]a,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 accumulator l'ine. 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 circumferential f. law 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)
( ] .
l where:
. [.
n ja.c e I
... }
l an.nwano 6-1 ,
"M
{-- ,
. -i 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 1 moment. Good agreement was found between the analytical predictions and the experimental results (reference 6-1). Flaw stability' evaluations, using this analytical model, are presented in section 6.3.
~6.2 Leak Rate Predictions The purpose of this section is to discuss the method which vil be used to
- predict the flow through a postulated crack and present the, 1 oak rate calculation results for postulated through-wall circumferential cracks in the accumulator'line.
6.2.1 General Considerations The flow of hot pressurized water through an opening to a lower back pressure (causing choking) is taken'into account. For long channels where the ratio of the channel length, L, to hydraulic diameter, O g , (L/D H
) is greater than
[ 3a,c.e, both [- Ja.c.e must be considered.
In this situation the flow can be described as being single phase through the channel until the local pressure equals the saturation pressure of the fluid.
' At this-point, the flow begins to flash and choking occurs. Pressure losses
. Lduetomomentumchangessilldominatefor[ Ja,c.e However, for l' ;1arge L/DH values., friction pressure drop will become important and must be considered along with the momentum losses due to flashing.
l.
s y ,
6.2.2 Calculation Method.
'In using the isentropic equilibrium model , the basic method used in the leak
- rate calculations is the method developed by (
1 1
l Ja.c.e, j i
The flow rate through a crack was calculated in the following manner. Figure 6-2 fr'om reference 6-2 was used to estimate the critical pressure, Pc,.'for the'. j primary loop enthalpy condition and an assumed flow.. Once Pc'was found for a.
given mass' flow, the [. Ja,c.e- ]
was found from figure 6-3 taken fra reference 6-2. For all cases considered, since [. .Ja.c.e Therefore, this method will yield l the two phase pressure drop due to momentum effects as illustrated in figure .;
6-4. Now using the assumed flow rate, G, the frictional pressure drop can be j
, calculated using
,]a,c.e (6.3)
APf=[ l where'the friction factor f is determined using th'e ( Ja,c.e j The crack relative roughness, e, was obtained from fatigue crack' data on !
stainless steel samples. Tne relative roughness value used in these i calculatie'sn was ( Ja.c.e RMS. l l
The frictional pressure drop using Equation'6.3 is then calculated for the assumed flow and added to the [.
.]a,c.e to obtain-the total pressure drop from the system under consideration'to the atmosphere. Thus, I
, Absolute Pressure - 14.7 = ( ]C(6.4)
I y
i i
1 man.nooseno 6-3 u# .. .. -
3
for a given assumed flow G. If the right-hand side of equation 6.4 does not i agree with the pressure difference between the piping under consideration 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 procedure has been recommended by [ l Ja.c.e for this type of [.
Ja.c.e calculation, i 6.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulated through-wall crack length for the critical location previously identified. The crack opening area was estimated using the method of reference. 6-4 and the leak rates were calculated using the calculation methods described above. The leak rates were calculated using the normal operating loads of axial force F and bending moment M at the governing location identified in section 5.0. These loads are given directly below.
F = 121 kips, M = 884 in-kips The crack length yielding a leak rate of 10 gpm (10 times the leak detection requirement of 1.0 gpm) is found to be ( )"'C long.
Thus reference flaw ~ size of ( Ja,c.e is established.
6.2.4 Leak Detection Capability, Administrative Procedures and Technical Specification Requirements The VEGP leakage detection criterion includes a detected unidentified leak rate of 1.0 gpm and, in accordance with NUREG-1061, Volume 3, a margin of 10 j was applied to the leak rate to define the accumulator line leakage size flaw used in the stability analysis. The basis for the 1.0 gpm leak rate is the i
.' presence (inside containment) of diverse and redundant leakage detection .
)
systems to measure containment gaseous radioactivity, airborne particulate i radioactivity, containment air cooler condensate flow monitor and containment !
O nu nocune 6-4 1
l
sump level. The sensitivity and response time of the detection equipment for unidentified leakage is such that a 1eakage rate, or.its equivalent, of 1 gpm can be detected in approximately one hour as shown in FSAR figure 5.2.5-1. In
= addition, humidity, temperature, and pressure monitoring of the containment
- atmosphere are used for alarms and indirect indication of leakage to the containment. These methods are in compliance with Regulatory Guide 1.45 as discussed in FSAR subsection 5.2.5.
The above methods are supplemented by visual and ultrasonic inspection of the reactor coolant pressure boundary during plant shutdown periods, in accordance with the Inservice Inspection Program (FSAR subsection 5.2.4). ]
' j 1
In addition, technical specification 4.4.6.2.1 requires monitoring of !
containment gaseous or particulate radioactivity --d normal sump inventory and I discharge at least once per 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br />. This section of the technical specification also requires performance of a reactor coolant system inventory balance at least once per 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />.
Rea'etor coolant inventory monitoring provides an indication of system leakage. Operators perform a RCS leakage calculation (VEGP Nuclear Operations Procedure 14905) at least once every 72 hours8.333333e-4 days <br />0.02 hours <br />1.190476e-4 weeks <br />2.7396e-5 months <br />. If the limits of Technical Specification 3.4.6.2 are determined to have been exceeded then the operators are directed to the Abnormal Operating Procedure Reactor Coolant System Leakage (VEGP18004).
The accumulator line is integral to the RCS. Any leakage from the accumulator line would be unidentified leakage until the exact location was determined.
At that time the leak would have to be designated as reactor coolant pressure boundary leakage. Detected leaks will be repaired within the system limiting conditions for operation established in either technical specifications or administrative procedures. When leakage is detected in reactor coolant pressure boundary piping,' Technical Specifications 3.4.6.2 requires that the plant be in hot standby within six hours and in cold shutdown within the next (
l thirty hours. Depair would be required before restart, j i i e I suwtoonstio 6-5 l
1 Experience at other Westinghouse plants indicate a normal background
_ unidentified average leakage rate of between 0.1 gpm and 0.3 spm, and it has
~
also been demonstrated with pressurized pipe tests that leak rates above 0.1 ',
. gpm at'one location.can be readily detected visually. The undefined leakage l rate'at VEGP is expected to be similar to other plants. Experience at similar ]
plants and the results of these tests indicated that a 1.0 gpm leak rate can be reliably detected and located during plant operation.
6.3 Stability Evaluation Using the "2" Factor Approach A typical segment of the pipe under maximum loads of axial force F and bending moment M is schematically illustrated as shown in figure 6-5. In order to calculate the critical flaw size, a plot of the limit moment versus crack length is generated as shown in figures 6-6. 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 4.0. From figure 6-6 the critical flaw size
_is'seentobe[( Ja,c e for the base metal.
k ,
The weld at the location of interest (i.e. the governing location) is an SAW weld. Therefore, a "2" factor correction for SAW welds was applied (references 6-5and6-6)asfollows:
2 = 1.30 (1 + 0.010 (0.0. - 4)) (6.5) where 00 is the outer diameter of the pipe in inches. Substituting'0D = 10.75 inches, the Z factor was calculated to be 1.388. The applied loads were increased by the Z factor and the plot of limit load versus crack length was regenerated as shown in figure 6-7. The lower bound base metal tensile properties (section 4.0) were used for this purpose. From figure 6-7, the ;
critical flaw size is seen to be [ la,c.e long. Noting that the o flaw yielding a leakage of 10 gpm (i.e. leakage size flaw)'was calculated to ,
be ( Ja.c.e long, a factor of 2.7 exists between the leakage size I k
flaw and the critical flaw. Thus, a margin of greater than 2 on flaw size is in evidence. )
l' an.emmue 6-6 i
f
i In order to determine the margin on applied loads (normal plus SSE), the
, applied loads were increased by a factor of 1.963 (i.e. /2 Z) and the plot of limit load versus crack length was generated as shown in figure 6-8. Again the lower bound base metal tensile properties were used for this purpose.
From figure 6-8 the critical flaw size is seen to be ( Ja,c,e long which is larger than [ ]a,c e inches long leakage size flaw. Thus a margin greater than 1.4 on (normal plus SSE) loads is demonstrated.
6.4 Local Stability Analysis In this section the local stability analysis is performed to show that unstable crack extension will not result when postulated through wall flaws 1 i
are subjected to maximum plant loads, i l
I At the critical location identified in section 5.0, the (normal plus SSE) outer surface axial stress; a, is seen to be 23.6 ksi based on the )
minimum wall thickness. The (normal plus SSE) axial force and bending moment
.- are Fx = 168 kips and Mb = 1239 in-kips.
. The minimum yield strength for flaw stability analysis is [ ]a,c e ksi (see section 4). The EPRI elastic plastic fracture handbook method is used to J calculate the J applied using the normal plus SSE loads. The J applied was l calculated for a ( Ja,c,e long postulated through wall flaw (which is 2 times the reference flaw size) and was found to be [ Ja,c.e ,
The applied tearing modulus, T applied was calculated from the basic definition, namely, l
E dJ T
applied ' 2 da (6.6)
,o
. ThehwasobtainedbycalculatingaJcorrespondingtoasmallincrementaa in crack length using elastic plastic load-controlled analysis. The T app}jgg was found to be [ Ja,c.e corresponding to J l applied 0 I 3 l' which is lower than T mat f about 89 (
Ja,c,e In eddition, for a leakage size flaw, i.e. the reference flaw of ( Ja,c,e
- long, the normal plus SSE load was increased by a factor of /2. The J-T m
- .enno 6-7 l
w____ _
F y l 1
.1 l
1 i
analysis gave'an applied J of ( -Ja,c.e and a T,p of
( Ja,c,e ,
,The=Tmat[ ] is J l
- 76. For both the above cases the T applied is lower than T mat and,
" therefore, unstable crack propagation will not result. 1 l '
-6.4.1 Crack Extension Considerations
'The crack extension corresponding to the maximum calculated J applied f
(
Ja,c,e would bs about 0.22 in, at each crack tip (see 1
-figure 4-3). If the J applied is calculated for a larger crack length ,
. allowing for crack extension of 0.22 in. at each crack tip, the Japplied d would be-[' Ja,c e The tearing modulus _ corresponding to this l increased J applied is ( . Ja,c,e which is again lower than the .T I mat
- 70. Therefore, consideration of crack extension, in this case would not j change the crack stability conclusions. .l 6.5. Integrity Assesment of Primary Loop Nozzle Junction e
The loop nozzles connecting the accumulator lines are made of cast stainless
.. steel, SA351 CF8A, which is susceptible to thermal aging. Therefore, fracture mechanics analyses were performed at the junction of the accumulator pipe and the primary loop nozzle. At this location, the leakage size flaw was calculated to be ( Ja,c.e long. The maximum stress at this-
)
location is 17.5 ksi. It is noteworthy that this stress is significantly l lower than the 23.6 ksi stress at the critical location. The (normal plus SSE) axial force and bending moment are Fx = 160 kips and Mb = 839 in-kips. At this location the J a plied was calculated to be [
Ja,c,e for twice the leakage size flaw ( Ja,c,e inches long subjected to normal plus SSE loads. The tearing modulus T applied corresponding to this J applied was calculated using the method described 7 earlier and was found to be about [ Ja,c.e . In addition, the J applied was calculated to be ( la,c,e for the [ ]a,c,e inches long l
. leakage size flaw subjected to / 2 (normal plus SSE) loads. The tearing. modulus corresponding to this J applied was found to be about
(- Ja,c.e . For both the above cases the T ap lied is found to be an, o o """
6-8
I significantly lower than the Tmat of ( ) ' established in section 4.0,
- and.therefore, unstable crack propagation will not result considering the thermal aging effects for the cast nozzle anterial. J
'~
6.6 References I
{
1 6-1 Kanninen, M. F. et al., "Mechanica1' Fracture Predictions for Sensitized Stainless Steel Piping with Circusferentialucracks" EPRI NP-192, September 1976.
{
j 1
6-2 [ '
ja c.e 6-3~[
3a c.e ,
6-4 Tada, H., "The Effects of Shell Qorrectin,tns on Stress Intensity Factors
,, and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983.
l 6-5 NRC letter from M. A. Miller to Georgia Power Company, J. P. O'Reilly, )
dated September 9, 1987. //
6-6 ASME Code Section XI, Winter 1985 Addendum, Article IWB-3640.
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' 5 e F1gure 6-2 Analytical Predictions of Critical Flow Rates t
I 4
of Steara-Water Mixtures \
I r 5-11 ,
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l I
l I
1 1
a,c,e I
a.c.e
/ 7 LQg = 40
- --- L2 g 40y ;
c L :
Figure 6-4 Idealized Pressure Drop Profile Through a Postulated Crack 6-13 99
n ir -' , g
< , s \g u L.
- 9 u g
5 p .=a % =
l l l .
l l I i .
I I I
I I
I I'
i .q I I .
1
- l l
l 1 ;
I I I I l i I I I i V
P = 2285 psi l
F = 168 kips I ' M = 1239 in-kips t .
Figure 6-5 Loads Acting on the Pipe Model at the Governing Location 6-14
l l
l 1 :
k i Setet l
.. NODE -#1042 OD = 10.75 in.
t = 0.8955 in ;
axial force = 168 kips o = 23.94 kai c7 = 67.3 kei i oj=45.57ksi T = 558 F ( l 3
= i i --
l I
FLAW GEOMETRY I
\ s I
\
- l
~
Figure 6-6 Critical Flaw Size Prediction for the Base Metal Using Limit Load Approach -
6-15
J
~
I
~ 1 8,tet I f
J 1
i Node - #1042 - Z Factor Z Factor = 1.388 i OD = 10.75 in j t = 0.8955 in J Axial Force = 168Z = 233.1 kips a = 23.84 ksi g
o7 s 67.3 isi l o}=45.571.si T = 550 F i
FLAW GEOMETRY i
. 1 i
l I
i i
1 l
l CRACK LENGT11 (in) i
, Figure 6-7 ,2-Factor Calculations for SAW Welds to Demonstrate Margin on Flaw Size 6-16
~
ss ft<.sbf 3,C 9 Node - #1042 - W =G Z Factor W =flZ = 1.9625
.. OD = 10.75 in t = 0.8955 in Axial Force = 168W.= 329.7 kips o = 23.84 kai oY = 67.3 ksi o"_=47.57 ksi L l T = 558 F
~~
l FLAW GE0 METRY
. CRACK LENGTH (in) .
. Figure 6-8 Z-Factor Calculations for SAW Welds to Demonstrate
, Margin on Loads 6-17 i
t-
+
SECTION 7.0 ASSESSMENT OF FATIGUE CRACK GROWTH
~
The fatigue crack growth on'the Vogtle Unit 2 accumulator line was determined by comparison with-a generic fatigue crack growth analysis of a similar piping j system. The details of the generic fatigue crack growth analysis are presented in Appendix B. By comparing all parameters critical to the fatigue crack growth analysis, between Vogtle Unit 2 and generic, it was concluded !
.that the generic ~ analysis would envelop the ' fatigue crack growth of the Vogtle
' Unit 2 accumulator line at the governing location identified in section 5.0. ;
Due to similarities in Westinghouse PWR designs it was possible to perform a generic fatigue crack growth calculation which would be applicable to many projects. A comparison was made of stresses and number of cycles, material,
. geometry, and types of discontinuities.-
The following summarizes the parameters which were compared:
'Yogtle Unit 2 Cold Generic Cold Leg Nozzle Leg Nozzle to Pipe Critical Location to Pipe Weld Weld / Weld to Elbow Pipe Outer Diameter 10.75" 10.75" Thickness .895" .895" Material Austenitic Stainless Steel Austenitic Stainless Steel Normal Temperature 550'F 560*F Normal Pressure 2300 psia 2300 psia Normal Operating 10.1 ksi 14.3/19.9 ksi !
Stress (Pres.+ Dwt Thermal Exp.)
Thermal Transients See Appendix B *
- Thermal transient loadings are nearly identical for the two projects. i i
j l
l
- """"' 7- 1 I
This comparison demonstrates the many similarities between the Vogtle Accumulator line and the generic accumulator line evaluation. The only uncertainty in this comparison is the higher level of pressure plus thermal
, expansion stress for Vogtle. Since this is essentially a mean stress (steady state) this difference will have only a minor impact on the fatigue crack growth calculation by increasing the R ratio. It is judged that this would increase the fatigue crack growth by less than [. -) .a . c . e Applying
[- 'Ja,c.e increase to the generic fatigue crack growth data (from appendix B) results in a final flaw depth of approximately [ .Ja.c.e for an initial flaw of [ ').a,c.e These results demonstrate that no significant fatigue crack growth will occur over the 40 year plant design life even for the largest postulated flaw.
7.1 Acceptability Fatigue Crack Growth A detailed discussion pertaining to the fatigue crack growth law used in the analysis described in Appendix B and the data used in definhg the law are provided in Reference (7-1). For the assessment of crack growth
.' acceptability, the crack growth results of the generic analysis presented in Appendix B are used conservatively and are considered applicable to the Vogtle
- Unit 2 Accumulator lines. Detailed discussion in support of this assumption has been provided in the previous se:: tion.
The maximum allowable preservice indication may have a depth of 0.09 in, per IWB-3514.3, Allowable Indication Standard for Austenitic Piping, ASME Code.
Section XI - Division 1,1986 edition. Estimated fatigue crack growth results are given in the previous section of this report. [.
Ja,c.e is conservatively chosen as a basis for examining the acceptability of fatigue crack growth. [t
~
Ja.c.e Thus, the first criterion on flaw depth is ,
satisfied.
i me no 7-2 I l
l J
The worst case transient AK value for the maximum crack depth is [' .)a,c.e L, The flow stress for the base metal'at 560*F is 45.5 ksi. which can be used to l l obtain a conservative estimate of the plastic zone size.
The expression for plastic zone size,pr , calculation is: [. Ja,c.e 5
= g (,AK )2 r
l p l
flow l Thus, the plastic zone size is calculated to be [ Ja.c.e The
- remaining ligament for the 0.224 in, deep end-of-fatigue-life flaw is 0.671
'in. (i.e. 895 - 0.224). Thus, the plastic zone size is less than the remaining ligament.
l Based on the above, it is concluded that for the Vogtle Unit 2 Accumulator l Lines, the fatigue crack growth during service will not be significant.
7.2 References
~
7.1 Bamford, W. H. " Fatigue Crack Growth of Stainless Steel Reactor Coolent Piping in a Pressurized Water Reactor Environment," ASME Trans. Journal of Pressure Vessel Technology, February 1979.
7.2 Rice, J. R., ASTM STP,1967, Volume 415, p. 247.
l l
l l*
e, t i
mano no 73
SECTION 8.0 ASSESSMENT Of MARGINS In the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessment were. performed. Margins at the q critical location are summarized below:
i In Secton 6.3 the " critical" flaw size using limit load method is calculated to be [. .Ja,c.e Icag. Using~the IWB-3640 approach (i.e. "2" factor approach), the critical flaw size'at the governing location weld is found to i be (, Ja c.e long. In section 6.4 it is demonstrated that a !
postulated [ .]a,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 j
In Section 6.2 it is shown that at the critical location, a flaw of [.
Ja,c.e would yield a leak rate of 10 gpm. Thus, there is a nar::in of at'least 2 on flaw size and a margin of 10 with respect to the plant leak
< detection capability of 1 gpm.
In Sections 6.3 and 6.4 it is shown that the reference flaw [ $
,]a,c.e yielding a leak rate of 10 gpm would be stable when subjected to .
a load equal to /2 (normal + SSE). Specifically, using the i IWB-3640 approach (section 6.3) a [ Ja.c.e long through-wall flaw I was shown to be stable when subjected to [2 2 (normal plus SSE) l loads. Also, based on local stability analysis (section 6.4) the leakage size flaw of [ Ja c.e inches was shown to be stable when subjected to
/2 (normal plus SSE) loads, l
l In summary, relative to
. 1. Loads i
The leakage-size crack will not experience unstable crack extension even if very large lo' ads of /2 (Normal plus SSE) are applied.
aan.nasseno 8-1 l
l i
- 2. Flaw Size f
- a. A margin of at least 2 exists between the critical flaw and the flaw
-yielding a leak rate of 10 gpm. j s
- 2. If limit load is used as the basis for critical flaw size, larger margin for global stability would result. ll l
- 3. Leak Rate !
'A margin of 10 exists for the reference flaw [ Ja,c.e between calculated leak rate and the criteria of Regulatory Guide 1.45. -
i A summary comparison of criteria and analytical results is given in Table B-1. The criteria are seen to be met. '
1 l
l l
^
l i
)
i i
j l
t f
^
an.n n. i 8-2 l
f I
TABLE 8-1 1 COMPARISON OF RESULTS VS. CRITERIA CRITERION RESULT _;
- 1. NUREG-1061 Volume 3 Net
- Section 5.2(h) - (Required margin of 2 demonstrated) l Wargin on Flaw Size 4
L I
l 2. NUREG-1061 Volume 3 Met j l Section 5.2(i) - (Required margin of vC2 demonstrated) I Margin on Load
- 3. NUREG-1061 Volume 3 Met ,
3 L Section 5.7 - (Margin of 10 on leak rate ]
Margin.on Leak Rate demonstrated) i 1
- 4. NRC criteria on allowable Net l- fatigue crack growth (af < 60% wall thickness) )
(Plastic zone size < remaining i ligament) 3 I
l 1
e I
a m .noom u n 8-3
_ c, i SECTION 9.0 CONCLUSIONS l
This report justifies the elimination of the accumulator line pipe breaks for l
Vogtle Unit 2 as follows:
4
- a. Stress corrosion cracking is precluded by use of fracture resistant l j
materials in the piping system and controls on reactor coolant
)
chemistry, temperature, pressure, and flow during normal operation. l lr
- 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. i The effects of low and high cycle fatigue on the integrity of the l accumulator line were evaluated and shown acceptable.
l
, d. Ample margin exists between the leak rate of small stable flaws and the criterion of Reg. Guide 1.45.
- e. Ample margin exists between the small stable flaw sizes of item d and the critical flaw.
- f. With respect to stability of the reference flaw, ample margin exists between the maximum postulated loads and the plant specific faulted loads (i.e. Normal + SSE).
The postulated 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 above, it is concluded that the accumulator line breaks should .
not be considered'in the structural design basis of Vogtle Plant Unit 2.
sen.nooseno 91
APPENDIX A LIMIT N0HENT O
S,Cet O
1 I
t ;
)
2Neo-OS2M7.10 ,
A-1 {
i
I 1
I 1
l a ,c .e .
l-I l
< mme I
l
. 1
. I F!alRE A-1 Pipe with a Through-Wall Crack in Sending l m:3. tuu. sun, ..
A-2 1
1
y _- - ;
i l
l l
l
- l t .
APPENDIX B i t
I FATIGUE CRACK GROWTH CONSIDERATIONS I
4 4
2 4
e
)
muaw ,o B-1 1
i
)
l l
l B.1 Thermal Transient Stress Analysis The thermal transient stress analysis was performed for a typical PWR plant to
,- obtain the through wall stress profiles for use in the fatigue crack growth !
analysis of Section B.2. The through wall stress distribution for each t transient was calculated for i) the time corresponding to the maximum inside i
surface stress and, ii) the time corresponding to the minimum inside surface !
stress. These two stress profiles are called the maximum and minimum through wall stress distribution, respectively for convenience. The constant stresses !
due to pressure, deadweight and thermal expansion (at normal operating !
temperature, 550*F) loadings were superimposed on the through wall cyclical stresses to obtain the total maximum and minimum stress profile for each
)
transient. Linear through wall stress distributions were calculated by
{
conservative simplified methods for all minor transients. More accurate nonlinear through wall stress distributions were developed for severe j transients by (. Ja.c.e B.1.1 Critical Location for Fatigue Crack Growth Analysis The accumulator line stress report design thermal transients (Section B.I.2),
1-D analysis data on accumulator line thermal transient stresses (based on ASME Section III NB3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis. (' !
la.c.e This location is selected as the worst location based on the following considerations:
- 1) the fatigue usage factor is highest.
ii) the stress due to thermal expansion is high.
1 l
iii) the effect of discontinuity due to undercut at weld will tend to increase the cyclical thermal transient loads.
iv) the review of data shows that the 1-D thermal transient stresses in the accumulator line piping section are generally higher near the (
,)a,c.e 4
l aan mensno g.2 i
t 1
J
.B.1.2 ~ Design Tra.,sients-The transient conditions selected for this evaluation are based on conservative estimates of.the magnitude and the frequency of the temperature
. fluctuations resulting from various operating conditions in the~ plant. These are representative of the conditions which are considered to occur during i plant operation.. The fatigue evaluation based on these transients provides l confidence that the component is appropriate for its application over the design life of the plant. All the normal operating and upset thermal
)
transients, in accordance-with design specification and-the applicable system l design criteria document (B-1), were considered for this evaluation. Out of.
. these,only.[.
la.c.e These transients were selected on the basis of the following criteria:
i
+a,c e ^
.o
, where, !
. i
+a,c.e B.1.3 Simplified Stress Analysis '
The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients.
[. j
)**C The inside surface stress was calculated by the following .
equation which is similar to the transient portion of ASME Section III NB3600, ;
Eq. 11:
l S$=[' Ja,c.e (B.3) mu,=mne l B-3 i
where, l
_ g :'
Sg = inside surface stress
+a.c.e p.-
i ,
r
, i r f_
, o l I !
.l -
4 q
l
- f. -
1
[' ._
!^
[' .
8 3 '*** The maximum and minimum inside surface stresses were P
. searched from the S4 values calculated for each time step of the transient solution.
The outside surface stresses corresponding to maximum and minimum inside stresses
.were calculated by the following equations:
- ~
S01 = [ ] (B.7)'
S02 = [i )a.c.e (B.8)
W. y mm.mmeno B-4
I where,
{ ;
S =
01 outside surface stress at time t,,x
=
+ S 02 utside surface stress at time t min
+a , c .e ,
a.c.e {
The material properties for the accumulator pipe [ ] and the RCL -]
1
, .]a,c.e The values of E and a, at the normal operating temperature, provide a conservative estimation of the through wall thermal j p- transient stresses as compared to room temperature properties. The following values were conservatively used, which represent the highest of the [
] a,c.e materials:
+a,c.e-The maximum and minimum linear through wall stress distribution for each thermal transient was obtained by [' '
Ja,c.e The simplified analysis discussed in this section was performed for all minor thermal transients of l
( .]a,c.e Nonlinear through wall stress profiles were developed for the remaining severe transients as explained in ;
Section B.1.4. The inside and outside surface stresses calculated by ;
simplified methods for the minor transients are shown in Table B-2. [. !
I. :
.]a c.e This figure shows that the
,- simplified method provides more conservative crack growth.
an,==n o B-5 1 l
-_ - - _ _ - I
- r. t B.1.4 Nonlinear Stress Distribution for Severe Transients l.
, .)a,c.e As mentioned earlier, the accumulator line section near the [ .Ja.c.e is the I worst location for fatigue crack growth analysis. A schematic of the accumulator line geometry at this location, is shown in Figure B-2. ['
l 1
i 1
1 1
~
1
.ja c.e l
B.1.5 OBE Loads Tne stresses due to OBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contribute significantly to I crack growth due to small number of cycles.
O I 9'
nu,mmun B-6 1
B.I.6- Total Stress for Fatigue Crack Growth The total through wall stress at a section was obtained by superimposing the pressure' load stresses and the st'resses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of Table B-2 or the nonlinear stress distributions discussed in Section B.I.4). Thus, the total stress for fatigue crack growth at any point is given by the following equation: '
Total Thermal Stress Due Stress vp for Transient to Due to j Fatigue = + DW + + Internal (B.9) l ll Crack Growth Thermal Pressure
' it Expansion i O
[ The envelope thermal expansion, deadweight and pressure loads for calculating j
- j the total stresses of Equation B.9 are summarized in Table B-3. '
~
B.2 Fatigue Crack Growth Analysis E* The fatigue crack growth analysis was performed to determine the effect of the
., design thermal transients, in Table B-1. The analysis was performed for the
- critical cross section of the model which is identified in Figure B-2. A
$ range of crack depths was postulated, and each was subjected to the transients i I
in Table B-1.
B.2.1 Analysis Procedure l
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 O
g mu. =suo B-7
t
'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" j (B.2.1) where "Co" and the exponent "n" are material ~ properties, and AKg is definedlater,inEquation(B.2.3). 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 "Cc" 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 l
water environment have been used in the analysis.
l The input required for a fatigue crack growth analysis is basically the
' {
^
information necessary to calculate the parameter AK g, which depends on I crack and structure geometry and the range of applied stresses in the area where'the crack exists. Once AK3 is calculated, the growth due to that particular cycle can be calculated by Equation (B.2.1). This increment of growth is then added to the original crack size, the AKg adjusted, and the analysis proceeds to the next transient. The procedure is continued in this l
manner until all the transients have been analyzed.
The crack tip stress intensity factors (K )j to be used in the crack growth analysis were calculated using an expression which applies for a semi elliptic surface flaw in a cylindrical geometry [B-4).
The stress intensity factor expression was taken from Reference 8-1 and was calculated using the actual stress profiles at the critical section. The
!, maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit.by a third order polynomial: -
4 o (x) = A 0+ A l$+A(k)+A({
2 3 (B.2.2) an. - o B-8
i
-The stress _ intensity factor Kj (e) was calculated at the deepest point of
- -]
the crack usingthe following expression: , j 1
+a,c.e ]
Kc)-
i l J
(2.2.3) 1 l
.where h
ll J l
l Bl.
i, f' ' . . !
[ Calculation of the fatigue crack growth for each cycle was then carried out b using the reference fatigue crack growth rate law determined from 0 consideration of the available data for stainless steel in a pressurized water environment. This law allows for the'effect of mean stress or R ratio E
(KImin/Elmax) n the growth rates.
The reference crack growth law for stainless steel in a pressurized water environment was taken from a collection of data (B-5) since no code curve is
., available, and it is defined by the following equation:
] a,c.e h=[ (B.2.4)
[4 mu,- .n. g.g 1
i
- . l; V
l s where K,ff = (Kg ,,x) (1-R)1/2 l=
KI "I" R = Ky ,,x
]
h=crackgrowthrateinmicro-inches / cycle !
1 B.2.2 Results I
. Fatigue crack growth analyses were carried out for the critical cross #
section. Analysis was completed for a range of postulated flaw sizes oriented circumferentially, and the results are presented in Table B-4. The postulated !
[- flaws are assumed to be six times as long as they are deep. Even for the
- largest postulated flaw of [ ;
} Ja c.e the result shows that the flaw growth through the wall will l not occur during the 40 year design life of the plant. For smaller flaws, the I
( flaw growth is significantly lower. . For example, a postulated [ Ja,c.e inch deep. flaw will grow to [ .Ja,c.e which is less than [ Ja.c.e i
the wall thickness. These results also confirm operating plant experience.
B.3 REFERENCES B-1 a,c.e
' B-2 ASME Section III, Division I-Appendices,1983 Edition, July 1,1983.
4 B-3 WECAN -- Westinghouse Electric Computer Analysis, User's Manual -- Volumes 1, II, III and IV, Westinghouse Research and Development Center, i Pittsburgh, PA, Third Edition,1982.
u e
== ie no B-10
f B-4 McGowan, J. J. and Rayrwnd, M., " Stress Intensity Factor Solutions for l
i '
Internal Longitudinal Semi Elliptical Surface flaws in a cylinder Under l
=-
Arbitrary Loadings", Fracture Mechanics ASTN STP 677,1979, pp. 365-380.
~-
B-5 Bamford, W.' H., " Fatigue' Crack Growth of Stainless Steel Reactor Coolant l Piping.in a Pressurized Water Reactor Envir8nment", ASME Trans. Journal ,
of Pressure Vessel Technology, February 1979. !
s
}
i
. )
i i
3
(
i
/
i n
9~
I e s .
.t' ,
t n 2,wuno B-11
.:,r - m
~ , , , .- :
'E
.'"'4/,_ _ . .' ! , . j./,'r 4
,... _ -__ , . h.
m .4 - -
u; > ..- ,a.' ..
r r m'-
1
, l .cn .. .
'9r {' . . 4 @cf., ,,.. ..
- f
.. s y~ ,; ., .
. tin
'g.v,s
,..< C y- jg,p_ -.. <-
, lx 4 rp TABLEB-1l,\'o' p, .,, .
T 1' r .
p!p? .
THERMAL' TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION I ': 4 . ' a
.f Trans. No. of s
. ft,L,.
No. - Description Occurrences o lp .
, ;g.
,, ,/f
+a,c.e
'I Ir A-. ( j. ,
,. . . / ': f . s x '
. [ ' [' -[ , ;. . :. -i I
ij' h rp 6- y
- f r,
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.,. 1 ,
=,-- .
sp )
,L : i,
. ix ,
ywI 1 a
i ' ,l_ l .,
_ < 1 ll - '
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s, vb ,i G$h Q' t e s \
- t. ,,\
.(1'
(,
f( ;
s, -
.e. '
- mu.mmen.
B-12 c _ _ - _ _ _ _ _ _ .__
+
s
. ,\ r, .
,, i S. . /, ,
'k
'r s \,
u.
e g ,t c.
Wya +- q,- ,I
.n
.\ q, f e. W .>
I t (
uJ ;4. ,
L l
7 M c 4-
. \1? t
.s' '
Ww ,
(
- MM '
'r
+ u8 ,.
,8
. ;., .)
i h
M m r
)
W :! -
GC \ l '
H .*
M i s l .\ .
EW l-(. e
- e o ./
' 3 -
m I
/. )/
^
m E
m >
N
- .
- CL, ,i V <
V E H i 2
w .L
- M M LD M ,
E EW 4 ac M H N H m ,
(- '( -
e '.
s .,y Eb E W O MQ i s
.. LLI - H LJ e== i , / i i;s 4 E bt' D
- m' o-o tv: > ' ,
-,' t -
-H( 3 .
g
. y ,
g, r '
.1, ri s
e
, f s ,4 6 j' g A m m l j ,
4 h.
W M
'l.
m W w. 1 ( i 4 . ' %,
W cc n , !.
$4,
- W H s M ~M I I- !)*!
> w i( ).
N W.
1
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t.
j , -
m; [ . '
i_3 1
1 i I
.j(
Y j
.Jy \
/ ) i C >- ,
1
-e' E u
-( i , ~[
i i (
^ e - '
F' :b -Q
, ' I' l, i ,f. , s
.. t v. 'r,' / r, !
D' \
X I -'
W ..
3
- a. $s/; i
- / g ,
4 1
' gM m, i
l i*B
, l c
- ~{ ' .. l 'l } ;
l
-/ l
\
) 1
- ' B-13, >
,, s
.3 \ \ '
i
.e 9
( i .i 'i
- h..in
- p
- nn_ s .. .m ,
l
'f..- Qtj .- i'h(( .g ;
- k. I, f f )
'p '
,N
_ V' 1 ,, l t i r. J .);'ta TABLE B-3 iji
.:-- ; . - [~ .y ENVELOPE NORMAL LOADS
~ ~
CONDITION ," a,c.e
!)
Normal Operating J
s
(
I 1
.o's .
. 1 t
!1 x&
.4 m,
l 4
' _a ' >
l i
I
! 'N s';
L i l 'b Y 4
l 1
4 5:
i
- l 1
4i -
6
'y ;. a l
L'y . .
i M41787,10 1.- B-14
_ _- ________-____x
, TABLE B-4 ACCUMULATOR LIN: FATIGUE CRACK GROWTH RESULTS Wall Thickness = [ 3 +a,c.e INITIAL CRACK LENGTH AFTER YEAR CRACK LENGTH 10 20 30 40 (IN.)
l +8,C,e O
mu.-e.a no B-15 R
1 a
e,s.e
- l l
I i
n a
l.
Il p .,
\
l l
(-
Figure B-1 Comparison of Typical Naximum and Minimum Stress Profile Computed by Simplified [. !+ +a,c.e am.am ,in B-M
e y, *a .c .e
. . 2. -
9 1
-+ +a c.e Accunulater Pipe 4
Figure B-? Schematic of Accumulated Line At I 3 + a ,c .e ame enm esone, is B-17
- a.s.e i
4
. J 1
w l l
l i
1 h
l, ik 10!
i
)
- . 8 i
l e
1 l
J'* ,
+a.c.e Figure B-3 [
] and Minimum Stress l
Profile for Transient #10 1 un.ax nune B-18
- auge
's,c,c A
1
)
e
= .
+a,c,e Figure B-4 ['.
.) Maximum and Minimum Stress Profile for Transient #11 a m ,e m . w o.o in B-19
s.s,e 4
P e
?
i l
t 1
Ih a .
c .
+a c.e Figure B-5 [. .) Waximum and Minimum Stress Profile for Transient #12 i
i a m . w a Sness,is B-20
)
- Setes I
l e
~
8
+a.c.e Figure B-6 [' l Maximum and Minimum Stress Profile for Transient #14 sm.ausuunc B-21
[.
1 APPENDIX C l.
NATERIALS SPECIFICATION AND FRACTURE TOUGHNESS
.. PROPERTIES OF THE ACCUMULATOR TANK, N0Z2LE, AND SAFE END C.1 Naterials Specification i The accumulator nozzle is SA 350 LF-2 with an ultimate tensile strength of between 70 and 95 ksi and a 36 ksi yield strength. The tank material is SA l 264.containing SA 537 C1 1 (carbon steel), ou = 70 ksi (min.),-c =50 y j ksi (min.), and SA 240-TP304 (stainless steel bonded to the carbon steel),
ou=75 ksi, c y=30 ksi. The safe end is made of-SA 312-TP304 (stainless steel pipe): o =75 ksi, o y=30 ksi.
u l
l' .
C.2 Fracture Toughness The certified material test report (CMTR) for the nozzle provides the
- following test results:
LOOP 1 Potential' Test '
Result RT NDT Drop' Weight No Break at O'F (Both heats) -10' :
Charpy Test Energy: 72,60,70,(1stheat) at 60'F. (ft-lb) 55, 51, 50:(2nd heat) 0' LateralExpansion(Mils) 60, 53, 59 (1st heat) 46, 41, 40 (2nd heat)
?' .
Y' 2 .
an. manne C-1 '
LOOP 2
' Potential Test Result , RTNDT
, Drop Weight No Break at O'F (Both heats) -10' Charpy Test Energy (f t-lb) at 60*F 72, 60, 70, (1st heat) [
55, 51, 50 (2nd heat) O' Lateral Expansion (Mils) 60, 53, 59 (1st heat) j 46, 41, 40 (2nd heat) l l
, Drop Weight No Break at O'F (Both heats) -1C' l
Charpy Test Energy (f t-lb) 60*F 72, 60, 70, (1st heat) ;
55, 51, 50 (2nd heat) 0' Lateral Expansion (Mils) 60, 53, 59 (1st heat) 46, 41, 40 (2nd heat) h* ,
N an.-o.a no C-2 4
LOOP 4 Potential
., Te:t' Result RT
. . ~
NOT e Drop Weight No Break at O'F (Both heats) -10' Charpy Test Energy (ft-lb) at 60*F 66, 65, 67, (1st heat) t 136, 130, 123 (2nd heat) O' l
Lateral Expansion (Wils) 61, 60, 63 (1st heat) 93, 91, 89 (2nd heat)
Therefore the RT NDT f the nozzle is O'F. The drop weight test for the tank material was not performed, however, Charpy V-Notch tests at 60'F were made and the test results are as follows:
Potential Test Loop Results RT NDT V Notch Energy 56, 62, 58 (1st heat) .
l Test at (ft-lb) 49, 51, 53 (2nd heat) 60*F 1 Lateral 49,51,53(1stheat) l Expansion 47, 53, 50 (2nd heat) l (mils) '
Energy 56,58,58(1stheat) !
(ft-lb) {'
2 0*F i' Lateral 55, 53, 54 (1st heat)
Expansion (mils)
_ ___ ~ - - -
1 6.,
L .
V v .
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m -
Potential Test Loop Results RT
- NDT qvo U Energy 56, 62, 54 (1st heat)
(ft-lb) 56, 58, 54 (2nd heat) 3 Lateral 49, 51, 53 (1st heat)
Expansion 47, 53, 50 (2nd heat)
(mils)
Energy 63, 66, 62 (1st heat) i (ft-lb) 63, 66, 60 (2nd heat) 1 50. 48. 48 (3rd heat) 4 Lateral 59, 57, 60 1stheat)
Expansion 57, 63, 60 2nd heat)
(mils) 45, 46, 48 3rd heat)
The V-Notch test results of the tank material show very similar results with those of nozzle material. Thus RT NOT f the tank material is considered to be 0*F.
The toughness of ferritic steel is taken from Appendix A,Section XI of the y ASME Code as the lesser of
's Kje = 33.2 + 2.806 Exp (0.02 (T-RTNDT+100)) (C.1) or Kj , = 200 ksi /in. (upper shelf toughness)
Substituting the operating temperature,120' and RT e O'T in equation NDT (C.1), the toughness of ferritic steel at 120*F is found to be 200 ksid.
This is equivalent to Jg , = 1333 in-lb/in2 where Jg , is assumed to be given by J ge = Kg ,2/E where E = 30 x 106 p34,
- s. Tests of actual SA 350 LF-2 forging material have shown J 2 ge values of
[
,,, la.c.e in-1b/in at 50*F and (: )*'C in-lb/in2 at 75'F. The
'6 Ja,c.e RTNDT was [- .
Since the tank and nozzle operate at 120'F with m u-aecuno c.4
a RTNDT of O'F, they operate at or near the upper shelf fracture toughness temperature. Thus the actual Jg , appropriate to the 120'F operating temperature is r. ear the higher test result [ Ja,c.e, l
Thus the accumulator nozzle and the accumulator tank material exhibit a 2
greater toughness level as compared to the JIc = [ -) '
- in-lb/in used as the criterion in the leak-before-break evaluation.
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