ML20207R835
| ML20207R835 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 03/31/1987 |
| From: | Lee Y, Schmertz J, Swamy S WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| Shared Package | |
| ML19292G928 | List: |
| References | |
| WCAP-11352, NUDOCS 8703180184 | |
| Download: ML20207R835 (90) | |
Text
- _ _ _ _ _ - ______________
W ctingh usa Cicco 3 WCAP-11352 i
TECHNICAL BASES FOR ELIMINATING ACCUMULATOR CLASS 1 LOW PRESSURE LINE RUPTURE AS THE STRUCTURAL DESIGN BASIS FOR SOUTH TEXAS PROJECT UNITS 1 AND 2 March 1987 Y. S. Lee S. A. Swamy j
J. C. Schmertz Verified by:
D. H. Roarty Approved-C S3.~Pguaimy, Manager -
StnacMiral Materials Engineering Work Performed Under Shop Order HPLJ-2028 WESTINGHOUSE ELECTRIC CORPORATION Generation Technology Systems Division P.O. Box 2728 Pittsburgh, Pennsylvania 15230-2728
!DO IDb!K b50$oS9s A
TABLE OF CONTENTS 4
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 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 AND THE REACTOR COOLANT SYSTEM 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 References 3-4 4.0 MATERIAL CHARACTERIZATION 4-1 4.1 Pipe, Fittings and Weld Materials 4-1 4.2 Tensile Properties 4-2 4.3 Fracture Toughness Properties 4-2 4.4 References 4-3 5.0 LOADS FOR FRACTURE HECHANICS 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 5.4 References 5-3 11 j
l
TABLE OF CONTENTS (cont.)
l t.
Section Title Page 6.0 FRACTURE MECHANICS EVALUATION 6-1 6.1 Global Failure Mechanism 6-1 6.2 Leak Rate Predictions 6-3 6.2.1 General Considerations 6-3 l
6.2.2 Calculation Method 6-4 6.2.3 Leak Rate Calculations 6-4 6.3 Local Failure Mechanism 6-5 6.4 References 6-7
7.0 ASSESSMENT
OF FATIGUE CRACK GROWTH 7-1 7.1 Acceptability of Fatigue Crack Growth 7-1 7.2 References 7-1
8.0 ASSESSMENT
OF MARGINS 8-1
9.0 CONCLUSION
S 9-1 APPENDIX A Limit Moment A-1 APPENDIX B Fatigue Crack Growth Considerations B-1 B.1 Thermal Transient Stress Analysis B-2 B.2 Fatigue Crack Growth Analysis B-7 B.3 References B-10 iii
LIST OF FIGURES 3
l
{
Figure Title g
2-1 Schematic of Generalized Load Deformation Behavior 2-4 5-1 Schematic Layout of Accumulator Line Loop 1 5-10 5-2 Schematic Layout of Accumulator'Line Loop 2 5-11 5-3 Schematic Layout of Accumulator Line Loop 3 5-12 f
6-1
(
Ja.c.e Stress Distribution 6-9 l
6-2 Loads Acting on the Pipe Model 6-10 6-3 Critical Flaw Size Prediction of 12 Inch Accumulator 6-11 Line Operating at 120*F 6-4 Critical Flaw Size Prediction of 10 Inch Accumulator 6-12 Line Operating at 120*F 6-5 Critical Flaw Size Prediction of 8 Inch Accumulator 6-13 Line Operating at 120*F 6-6 Leak Rate Versus Crack Length For 12 Inch 6-14 Accumulator Line Operating at 120'F b-7 Leak Rate Versus Crack Length For 10 Inch 6-15 Accumulator Line Operating at 120*F 6-8 Leak Rate Versus Crack Length For 8 Inch 6-16 Accumulator Line Operating at 120'F iv
LIST OF FIGURES (cont.)
L Figure Title M
A-1 Pipe with a Through Wall Crack in Bending A-3 8-1 Comparison of Typical Maximum and Minimum Stress B-15 Profile Computed by Simplified (
ja.c.e B-2 Schematic of Accumulator Line at [
B-16 3a,c.e B-3
[
la.c.e and Minimum Stress Profile B-17 for Transient #10 B-4
(
Ja,c.e Maximum and Minimum Stress B-18 Profile for Transient #11 5-5
(
Ja.c.e Maximum and Minimum B-19 Stress Profile for Transient #12 B-6
[
Ja.c.e Maximum and Minimum B-20 Stress Profile for Transient #14 V
LIST OF TABLES i
Table No.
Title M
4-1 Room Temperature Mechanical Properties of the 4-4 12 Inch Low Pressure Accumulator Line Materials and Welds of the South Texas Project Unit 1 Plant 4-2 Room Temperature Mechanical Properties of the 4-5 10 Inch and 8 Inch Low Pressure Acetmulator Line Materials and Welds of the South Texas Project Unit 1 Plant 4-3 RoomTemperatu{eMechanicalPropertiesofthe 4-7 12 Inch Low Pressure Accumulator Line Waterials and Welds of the South Texas Project Unit 2 Plant 4-4 Room Temperature Mechanical Propprties of the 4-8 10 Inch and 8 Inch Low Pressure Accumulator Line Waterials and Welds of the South Texas Project Unit 2 Plant 4-5 Typical tensile properties of SA376 TP316 and 4-9 Welds of Such Waterial for the Primary Loop 4-6 Comparison of Room Temperature Tensile Properties 4-10 of the 12, 10, and 8 Inch Accumulator Lines Operating at 120*F With Those of Typical Wrought Primary Loops and ASME Code Minimum Requirements 4-7 Fracture Toughness Properties Typical of Accumulator 4-11 Line vi
LIST OF TABLES (cont.)
)
3 Table No.
Title Pajte 5-1 Summary of Envelope Loads for 12 Inch Low Pressure 5-4 Pipe 5-2 Summary of Envelope Loads for 10 Inch Pipe 5-5 5-3 Summary of Envelope Loads for 8 Inch Pipe 5-6 5-4 Loading Components at Governing Locations For 5-7 12 Inch Low Pressure Line 5-5 Loading Components at Governing Locations For 5-8 10 Inch Low Pressure Line 5-6 Loading Components at Governing Locations For 5-9 8 Inch Low Pressure Line 8-1 Comparison of Results vs. Criteria 8-5 B-1 Thermal Transients Considered for Fatigue Crack B-12 Growth Evt.luation B-2 Stresses for the Minor Transients (PSI)
B-13 B-3 Envelope Normal Loads B-13a q
B-4 Accumulator Line Fatigue Crack Growth Results B-14 vii
SECTION
1.0 INTRODUCTION
i
1.1 Background
l The current structural design basis for the accumulator 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 breaks for the accumulator line. Presented in this report are the descriptions of a mechanistic pipe break evaluation method and the analytical results that are used for establishing that a circumferential type break will not occur within the Class 1 portions of the accumulator lines. The evaluations considering circumferential1y 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 accumulator line. Schematic drawings of the piping system are 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.
i
- The scope of this work covers the Class 1 portion, operating at 120*F, of the Emergency Core Cooling System. The rest of the Class I line is analyzed in WCAP-11383, (7-1).
1-1
d f
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 j
of the flaw should be large enough so that the leakage is assured of detection with margin using the installed leak detection equipment when the pipe is subjected to normal operating loads. The low pressure system for this plant is capable of detecting a leakage of 0.5 gpm which is a factor of two less than the amount required by Regulatory Guide 1.45.
If a margin of 10 is retained between the calculated leak rate and the leak detection capability, the postulated flaw is required to yield at least 5 gallons per minute i
l leakage when subjected to normal operating loads.
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.
l l
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 that 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 global analysis is carried out using the [
la c.e method, based on traditional plastic limit load concepts, but accounting for
[
la.c.e and taking into account the presence of a flaw. The 1-2
(.
local stability analysis can_be carried out using'?,he method described in NUREG/CR 3464(1-2). This method is based on linear elastic fracture mechanics and it ca'n be used up to load levels producing stresses near the i
yield point. For higher loads, the local stability analysis is carried out by performing a static elastic plastic [-
]a,c.e of a straight pieca'of the accumulator line pipe containing a through-wall circumferential flaw.
The leak rate is calculate.d for the normal operating condition. The leak rate prediction model.used in this evaluation-is a single phase flow in a tigh't crack including friction losses associated with crack ocening geometries. The crack opening area required for calculating the leak rates is obtained by subjecting the postulated through-wall flaw to normal eperating loads.
Surface roughness is accounted for in determining the leak rate through the-postulated flaw.
As stated earlier..tna evaluations described above considering circumferen-tially 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 anticipated. The prediction methods for failure in elbows are virtually the same as those for
[
Ja.c.e However, the elbows are [
Ja,c.e and, therafore, the probabil.ity of any longitudinal flaw existing in the accumulator line is much smaller when compared with the circumferential dirrction.
Based on the"above, it is juaged that circumferential fisys are more limiting than longituoinal flaws in elbows and throughout the system.
The computer codes used in this evaluation have been validated (bench marked) as described in References (1-4) and (1-5).
7 G
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 l
Circumferential Through Wall Cracks."
1-3 Begley, J.A., et. al., " Crack Propagation Investigation Related to the Leak-Before-Break Cor. cept for LMFBR Piping" in Proceedings, Conference on Elastic Plastic Fracture, Institution of Mechanical Engineers, London 1978.
1-4 Swamy, S.A., et. al., " Additional Information in Support of the Elimination of Postulated Pipe Ruptures in the Pressurizer Surge Lines of South Texas Project Units 1 and 2" WCAP-11256, September 1986, Westinghouse Proprietary Class 2.
1-5 Swamy, S. A., et. al., " Technical Basis for Eliminating Pressurizer Surge Line Ruptures as the Structural Design Basis for South Texas Project Units 1 and 2," WCAP-11256 Supplement 1, November 1986, Westinghouse Proprietary Class 2..
8 1-4
[
SECTION 2.0 FAILURE CRITERIA FOR FLAWED PIPES 2.1 General Considerations j
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 i
initiation toughness, equivalent energy, crack opening displacement, crack opening stretch, crack opening angle, net-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 make calculations using the available tools was used in selecting the criteria for this investigation. The ultimate objective is to show that the accumula-tor line containing a conservatively assumed circumferential 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 For a tough ductile material which is notch insensitive the global failure will be governed by plastic collapse.
E::twisive 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.
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.
i l
l l
2-1
A simplified version of this criterion, namely, not 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 < 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 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 pipe, flaw l
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 been demonstrated that the initiation toughness, measured in terms of JIc 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 f the material, then the crack will not initiate.
Ic If the initiation criterion is not met, one can calculate the tearing modulus as defined by the following relation:
dJ E
(~)
- Ha Q 8 app 2-2
m "e-where T
applied tearing modulus
=
app E
modulus of elasticity
=
'f flow stress = [
Ja.c.e
=
crack length
=
yield and ultimate strength of the material c.o
=
y u
respectively.
In summary, the local crack stability will be established by the two step criteria:
J<J r
Ic, (2-3)
T,pp < Tmat, M J g JIc (2-4)
An additional supplementary criterion is that J < J,,, where J,,x is the maximum value of J obtained from J tests for the material in question.
In this report, however, J has been chosen as a stand-alone criterion.
gc 2.4 References 2-1 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.
4 2'-3
j w,= PLASTIC LOAD 8
1 I
t S
I Se A
l 5
l 5
i o
I i
I l
l AP GENERAll2ED DISPLACEMENT e
l l
FIGURE 2-1 Schematic of Generalized Load-Deformation Behavior 2-4
SECTION 3.0 L
OPERATION AND STABILITY OF THE ACCUMULATOR LINES AND THE REACTOR COOLANT SYSTEM 3.1 Stress Corrosion Cracking
.The Westinghouse reactor coolant system primary loop and connecting Class 1 lines have an operating history that demonstrates the inherent operating stability characteristics of the design. This includes a low susceptibility to cracking failure from the effects of corrosion (e.g., intergranular stress corrosioncracking). This operating history totals over 400 reactor years, including five plants each having over 15 years of operation and 15 other plants each with over 10 years of operation.
In 1978, the United States Nuclear Regulatory Commission (USNRC) formed the second Pipe Crack Study Group. (The first Pipe Crack Study Group established in 1975 addressed cracking in boiling water reactors only.) One of the objectives of the second Pipe Crack Study Group (PCSG) was to include a review of the potential for stress corrosion cracking in Pr narized Water Reactors I
(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:
"The PCSG has determined that the potential for stress-corrosion cracking in PWR primary system piping is extremely low because the ingredients j
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 3-1
,1 4
r----
s:
me
l l
l of producing stress-corrosion cracking in the primary systems of PWRs.
Operating experience in PWRs supports this determination. To date, no stress-corrosion cracking has been reported in the primary piping or safe ends of any PWR."
During 1979, several instances of cracking in PWR feedwater piping led to the l
establishment of the third PCSG. The investigations of the PCSG reported in l
NUREG-0691 (Reference 3-2) further confirmed that no occurrences of IGSCC have been reported for PWR primary coolant systems.
As stated above,for the Westinghouss 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, 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.,
l sulfides, sulfites, and thionates). Strict pipe cleaning standards prior to operation and careful control of water chemistry during plant operation are used to prevent the occurrence of a corrosive environment. Prior to being put into service, the piping is cleaned internally and externally. During flushes and preoperational testing, water chemistry is controlled in accordance with written specifications.
Requirements on chlorides, fluorides, conductivity, and pH are included in the acceptance criteria for the piping.
3-2
During plant operation, the reactor coolant water chemistry is monitored and maintained within very specific limits. Contaminant concentrations are kept below the thresholds known to be conducive to stress corrosion cracking with the' major water chemistry control standards being included in the plant operating procedures as.a condition for plant oporttion. 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 j
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.
1 3.2 Water Hammer Overall, there is a low potential for water hammer in the RCS and connecting accumulator lines since they are designed and operated to preclude the voiding condition in normally filled lines. The RCS and connecting accumulator lines including piping and components, are designed for normal, upset, emergency, and faulted condition transients. The design requirements are conservative relative to both the number of transients and their severity. Relief valve actuation and the associated hydraulic transients following valve opening are considered in the system design. Other valve and pump actuations are relatively slow transients with no significant effect on the system dynamic loads. To ensure dynamic system stability, reactor coolant parameters are l
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 l
C 3-3
operating experience have verified the Westinghouse approach. The operating transients of the RCS primary piping and connected accumulator lines are such that no significant water hammer can occur.
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 III 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.
High cycle fatigue loads in the system would result primarily from pump vibrations during operation. During operation, an alarm signals the exceadance of the RC pump shaft vibration limits. Field measurements have been made on the reactor coolant loop piping of a number of plants during hot functional testing. Stresses in the elbow below the RC pump have been found to be very small, between 2 and 3 ksi at the highest. When translated to the connecting accumulator lines, these stresses are even lower, well below the i
fatigue endurance limit for the accumulator line material and would result in an applied stress intensity factor below the threshold for fatigue crack growth.
3.4 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.
1 3-2 Investigation and Evaluation of Cracking Incidents in Piping in q
Pressurized Water Reactors, NUREG-0691, U.S. Nuclear Regulatory
{
Commission, September 1980.
3-4
y SECTION 4.0 MATERIAL CHARACTERIZATION l
1 4.1 Pipe, Fittings, and Weld Materials The pipe materials of the 12,10, and 8 inch lines operating at 120'F are SA 376-TP316, SA312-TP316L, and SA312-TP304L. The fittings are SA403-WP316, SA182-F316L, and SA403-TP316L. The weld wire used in the shop fabrication is l
generally low carbon 316L; in some instances the weld wire has low carbon with 1
high silicon (316LSi). The welding processes used were gas tungsten arc (GTAW), submerged arc (SAW), gas metal arc (GMAW) and shielded metal arc (SMAW).
In the following section 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 12, 10, and 8 inch portions of the accumulator lines operating at 120*F were used to establish the tensile properties for the piping, fittings and welds. The properties are given in Tables 4-1 and 4-2 for Unit 1, and Tables 4-3 and 4-4 for Unit 2.
The properties in Table 4-1 through 4-4 are those at room temperature.
In the leak-before-break evaluations presented later, the ASME Section III code minimum properties at operating temperatures are used. The viability of using such properties for the 12,10, and 8 inch low pressure accumulator lines, is presented below.
[
ja,c.e 4-1
E
\\
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.
l i
t Based on this discussion it is concluded that the use of ASME Section III code minimum properties is jus'tified.
[
ja c e 4.3 Fracture Toughness Properties I
The low pressure accumulator lines are operating at 120*F and the toughness at 120*F will be hig'her than the toughness at 600*F for which extensive toughness data are available.
[
[
) a,c,e,
The toughness at 120*F will be higher than the toughness at 600'F.
In general, for materials exhibiting ductile fracture over the range of tempera-tures of interest, such as the stainless steel product forms under considera-tion, the initiation toughness, J is less at reactor operating temperature Ie (600*F) than at the low temperature of concern (120*F). T usually mat increases somewhat with temperature however.
Therefore, it is conservative to use the higher temperature J as instability criterion for operation at Ic 120'F. Furthermore, crack instability of the low pressure accumulator lines is determined based on Jyc = [
3a,c.e including thermal aging, even though the thermal aging does not occur at the 120'F operating
[
temperature.
Therefore, JIc = [
] * is a very low value compared to the J values of the actual materials.
yc l
4-2 4
._,.-__,__.,,_,,-o
4.4 References 4-1 F. J. Witt et al., " Integrity of the Primary Piping System of Westinghouse Nuclear Power Plants During Postulated Seismic Events,"
WCAP-9283, March 1978.
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).
l 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 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 (Westinghouse Proprietary Class 2).
4-3 i mimmii su irie imi
TABLE 4 ROOM TEMPERATURE MECHANICAL PROPERTIES OF THE 12 INCH LOW PRESSURE ACCUMULATOR LINE MATERIALS AND WELOS OF THE SOUTH TEXAS PROJECT UNIT 1 PLANT 0.2% OFFSET ULTIMATE FLOW LOOP PRODUCT HEAT YIELO STRESS STRENGTH STRESS ELONGATION REDUCTION NO.
FORM NUMBER MATERIAL (psi)
(psi)
(psi)
PER INCH IN AREA-1 Pipe HT L5093 SA376-TP316 42,700 88,200 65,500 53.2 64.1 1
Fitting HT 53919 SA403-WP316 42,000 77,500 60,000 62.0 75.5 1
Weld HT 17138 SFA5.9-ER316L 67,200 83,600 75,400 50.0-65.5 2
Pipe HT L5091 SA376-TP316
-38,100 82,600 60,400 60.5 71.2 2
Pipe HT L5093 SA376-TP316 42,700 88,200 65,500 53.2 64.1 2
Pipe HT L5091 SA376-TP316 38,100 82,600 60,400 60.5 71.2 2
Fitting HT 53919 SA403-WP316 42,000 77,500 59,800 62.0 75.5 2-Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 43.0 65.9 4.
3 Pipe HT L5093 SA376-TP316 42,700 88,200 65,500 53.2 64.1 l-3 Pipe HT P8611 SA376-TP316 46,700 92,600 69,700 50.5 63.1 3
Fitting HT 55894 SA403-WP316 50,500 79,000 64,800 45.0 70.5 3
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.0 3
Weld HT 00575 SFAS.4-E316L 57,400 80,400 68,900 45.0 63.8 3
Pipe HT L5093 SA376-TP316 42,700 88,200 65,500 53.2 64.1 3
Fitting HT 55895 SA403-WP316 51,000 81,500 66,300 56.0 71.5 3
Fitting HT 53919 SA403-WP316 42,000 77,500 59,800 62.0 75.5 3
Weld HT 17138 SFA5.9 ER316L 65,400 88,700 77,100 48.0 65.9 3
Weld HT 0683A SFA5.4-E316L 58,600 79,700 69,200 40.0 66.2 3
Pipe HT 11-226 SA312-TP316L 41,700 82,100 61,900 54.0 N/A.
3 Weld HT 306402 SFA5.9-ER316L 67,000 90,000 78,500 30.0 48.0
+.... ~
4 TABLE 4-2 ROON TEMPERATURE MECHANICAL PROPERTIES OF THE 10 AND 8 INCH LOW PRESSURE ACCUNULATOR LINE NATERIALS AND WELDS OF THE SOUTH TEXAS PROJECT UNIT 1 PLANT 0.2% OFFSET ULTINATE FLOW LOOP PRODUCT HEAT YIELD STRESS STRENGTH STRESS ELONGATION REDUCTION NO.
FORN [1]
NUMBER MATERIAL (psi)
(psi)
(psi)
PER INCH IN AREA 1
Pipe HT 1081-5-1 SA376-TP316 42,050 82,500 63,200 61.0 N/A 1
Pipe HT 1081-10-2-1 SA376-TP316 44,600 85,100 64,900 53.0 N/A 1
Fitting HT 55707 SA403-WP316 51,000 81,500 66,300 56.0 74.5 1
Fitting HT 53896 SA403-WP316 49,000 81,500 65,300 62.0 75.5 1
Weld HT 17138 SFAS.9-ER316L 65,400 88,700 77,100 48.0 65.9 1
Weld HT X4329 SFA5.9-ER316Lsi 70,600 92,500 81,200 38.0 59.6 1
Pipe 10"-8" HT 10891-5-2-2 SA376-TP316 49,100 81,200 65,200 60.0 N/A 1
Pipe HT 1081-1-1 SA376-TP316 41,150 80,900 61,000 64.0 N/A r
a.
1 Fitting HT 55709 SA403-WP316 48,500 81,000 64,800 60.0 73.5-th 1
Fitting HT 53920 SA403-WP316 42,000 77,500 59,800 62.0 75.5 1
Fitting HT 53918 SA403-WP316 41,400 85,900 63,700 52.8 N/A 2
Pipe 8" HT 10891-5-5-2 SA376-TP316 49,100 81,200 65,200 60.0 N/A 2
Pipe HT 1081-1-1 SA376-TP316 41,150 80,900 61,000 64.0 N/A 2
Fitting HT 55708 SA403-WP316 48,500 81,500 65,000 56.5 74.5 2
Fitting HT 53920 SA403-WP316 42,000 77,500 59,800 62.0 75.5 2
Fitting HT 53918 SA403-WP316 41,400 85,900 63,700 52.8 N/A 2
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.9 2
Pipe HT 1081-2-2 SA376-TP316 42,100 82,900 62,500 63.0 N/A 2
Fitting HT 55709 SA403-WP316 48,500 81,000 64,800 60.0 73.5 2
Fitting HT 53896 SA403-WP316 49,000 81,500 65,300 62.0 75.5 2
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.9 2
Weld HT X4329 SFA5.9-ER316Lsi 70,600 92,500 81,600 38.0 59.6 3
Pipe HT 1081-2-3-2-1 SA376-TP316 51,400 91,050 71,200 51.0 N/A 3
Pipe HT 1081-5-1 SA376-TP316 42,050 82,500 62,300 61.0 N/A 3
Fitting HT 55708 SA403-WP316 48,500 81,500 65,000 56.5 74.5 3
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.9 3
Weld HT 0683A SFAS.4-E316L 58,600 79,700 69,200 40.0 66.2 3
Pipe HT 1081-1-1 SA376-TP316 41,150 80,900 61,000 64.0 N/A i
TABLE 4-2 (Cont'd.)
ROON TEMPERATURE NECHANICAL PROPERTIES OF THE 10 l
AND 8 INCH LOW PRESSURE ACCUMULATOR LINE MATERIALS l
AND WELOS OF THE SOUTH TEXAS PROJECT UNIT 1 PLANT 0.2% OFFSET ULTINATE FLOW i
LOOP PRODUCT HEAT YIELO STRESS STRENGTH STRESS ELONGATION RE00CTION NO.
FORN [1]
NUMBER WATERIAL (psi)
(psi)
(psi)
PER INCH IN AREA 3
Pipe 8" HT 0891-5-2-2 SA376-TP316 49,100 81,200 65,200 60.0 N/A i
3 Fitting HT 53920 SA403-WP316 42,000 77,500 59,800 62.0 75.5 3
Fitting HT 53918 SA403-WP316 41,400 85,900 63,700 52.8 N/A
[1] 10" line, unless indicated as 8".
1 2.
01 l
e i
i i
l j
TABLE 4-3 ROON TEMPERATURE NECHANICAL PROPERTIES OF THE 12 INCH LOW PRESSURE ACCUNULATOR LINE NATERIALS AND WELDS OF THE SOUTH TEXAS PROJECT UNIT 2 PLANT 1
0.2% OFFSET ULTINATE FLOW i
LOOP PRODUCT HEAT YIELD STRESS STRENGTH STRESS ELONGATION REDUCTION NO.
FORM NUMBER NATERIAL (psi)
(psi)
(psi)
PER INCH IN AREA I
1 Pipe HT L5093 SA376-TP316
.40,100 83,000 61,600 59.0 67.9 1
Fitting HT 53919 SA403-WP316 42,000 77,500 59,800 62.0 75.5 I
Weld HT 306402 SFA5.9-ER316L 67,000 90,000 78,500 31.0 48.0 j
1 Weld HT 19759 SFA5.9-ER316L 61,700 81,200 71,700 30.0 47.0 2
Pipe HT L5093 SA376-TP316 40,100 83,000 61,600 59.0 67.9 4
2 Fitting HT 53919 SA403-WP316 42,000 77,500 59,800 62.0 75.5 2
Weld HT 306402 SFA5.9-ER316L 67,000 90,000 78,500 31.0 48.0 m-2 Weld HT 19759 SFA5.9-ER316L 61,700 81,700 71,700 30.0 47.0 24 3
Pipe HT L5093 SA376-TP316 40,100 83,000 61,600 59.0 67.9 3
Pipe HT L5092 SA376-TP316 41,100 98,400 69,800 50.3 63.1 3
Fitting HT 55896 SA403-WP316 50,000 82,000 66,000 56.0 73.5 3
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.9 3
Weld HT 0575 SFA5.4-E316L 57,400 80,400 68,900 45.0 63.8 3
Weld HT 0683A SFA5.4-E316L 58,600 79,700 69,200 40.0 66.2 3
Pipe HT L5092 SA376-TP316 39,300 84,200 61,800 59.0 72.6 3
Fitting HT 55895 SA403-WP316 51,000 81,500 66,300 56.0 71.5 3
Fitting HT 53919 SA403-WP316 42,000 77,500 59,800 62.0 75.5 3
Weld HT 306402 SFA5.9-E316L 67,000 90,000 78,500 30.0 48.0 3
Weld HT 19759 SFA5.9-E316L 64,400 86,400 75,400 35.0 41.6 l
i TA8LE 4-4 i
ROOM TEMPERATURE MECHANICAL PROPERTIES OF THE 10 AND 8 INCH INCH LOW PRESSURE ACCUMULATOR LINE MATERIALS AND WELOS OF THE SOUTH TEXAS PROJECT UNIT 2 PLANT' i
0.2% OFFSET ULTIMATE FLOW l
LOOP PRODUCT HEAT YIELD STRESS STRENGTH STRESS ELONGATION REDUCTION NO.
FORM [1]
NUM8ER MATERIAL (psi)
(psi)
(psi)
PER INCH IN AREA 1
Pipe HT 1081-8-1-1 SA376-TP316 41,150 80,900 61,000 64.0 N/A 1
Pipe HT 1081-8-1-1 SA376-TP316 45,150 86,600 65,900 54.0 N/A 1
Pipe HT 1081-9-2 SA376-TP316 41,050 79,600 60,300-63.0 N/A i
1 Fitting HT 55707 SA403-WP316 51,000 81,500 66,300 56.0 74.5 1
Fitting HT 53836 SA403-WP316 49,000-81,500 65,300 62.0-75.5 1
Weld HT 17138 SFAS.9-ER316L 65,400 88,700 77,100 48.0 65.9 1
Weld HT X4329 SFA5.9-ER316Lsi 70,600 92,500 81,600 38.0 59.6 1
Pipe 8" HT M7864 SA312-TP304L 38,100 82,700 60,400 57.0 N/A 9
1 Fitting HT E9968 SA403-WP316L 36,558 75,779 56,200 60.0 N/A 1
on 1
Fitting HT F0359 SA403-WP316L 42,500 79,250 60,900 50.0 54.0 1
Fitting HT F0024 SA403-WP316L 32,420 77,070 54,700 65.0 N/A 1
Fitting HT 1498N SA182-F316L 38,127 75,506 56,800 56.5 77.0 1
Weld HT 40704 SFA5.9-ER308L 62,900 89,900 76,400 37.0 71.4 1
Weld HT 19759 SFA5.9-ER316L 64,400 86,400 75,400 35.0 41.6 1
Weld HT 306402 SFA5.9-ER316L 67,000 90,000 78,500 30.0 48.0 2
Pipe HT 1081-8-1-1 SA376-TP316 41,150 80,900 61,000 64.0 N/A-2 Pipe HT 1081-9-2 SA376-TP316 41,050 79,600 60,300 63.0 N/A 2
Fitting HT 55708 SA403-WP316 48,500 81,500 65,000 56.5
'74.5 i
i 2
Fitting HT 53896 SA403-WP316 49,000 81,500 65,300 62.0 75.5 2
Weld HT 17138 SFA5.9-ER316L 65,400 88,700 77,100 48.0 65.9 2
Weld HT X4329 SFA5.9-ER316Lsi 70,600 92,500 81,600 38.0
.59.6 l
3 Pipe HT 1081-15-1 SA376-TP316 41,650 78,550 60,100 57.0 N/A 3
Pipe HT 1081-9-2 SA376-TP316 41,050 79,600 60,300 63.0 N/A 3
Fitting HT 55708 SA403-WP316 48,500 81,500 65,000-
- 56. 5.-
74.5-t 3
Weld HT 17138 SFAS.4-ER316L 65,400 88,700 77,100 48.0 65.0 3
Weld HT 00575 SFA5.4-E316L 57,400 80,400 68,900 45.0 63.8 i
3 Weld HT X4329 SFA5.9-ER316LSI 70,600 92,500 81,600 38.0 59.6 i.
[1] 10" line unless indicated as 8".
TABLE 4-5 TYPICAL TENSILE PROPERTIES OF SA376 TP316 AND WELDS OF SUCH MATERIAL FOR THE PRIMARY LOOP I
1 Test Temperature Average Tensile Properties Plant Material
(*F)
Yield (psi)
Ultimate (psi)
A SA376 TP316 70 40,900(48)*
83,200 (48)
E308 Weld 70 63,900(3) 87,600'(3) l l
B SA376 TP316 70 47,100(40) 88,300 (40)
E308 Weld 70 59,600(8) 87,200 (8)
C SA376 TP316 70 46,600(36) 87,300 (36)
E308 Weld 70 61,900(4) 85,400 (4)
- a. (
) indicates the number of test results averaged.
4-9
TABLE 4-6 COMPARISON OF ROOM TEMPERATURE TENSILE PROPERTIES OF THE 12,10, AND 8 INCH ACCUMULATOR LINES OPERATING AT 120*F WITH THOSE OF TYPICAL WROUGHT PRIMARY LOOPS AND ASME CODE MINIMUM REQUIREMENTS Properties (ksi)
Line/ Component Material Yield Ultimate 8
a Acc/ Pipe SA376-TP316 38.1 to 50.3 81.4 to 89.0 Acc/ Pipe SA312-TP304L 38.1 82.7 b
b Loop / Pipe SA376-TP316 40.9 to 47.l 83.2 to 88.3 8
a Acc/ Fitting SA403-WP316L 32.4 to 42.5 75.8 to 79.3 a
8 Acc/ Fitting SA403-WP316 37.5 to 51.0 77.5 to 83.8 Acc/ Fitting SA182-F316L 38.1 75.5 8
8 Acc/ Weld E316L,E316 LSi 57.4 to 67.2 79.7 to 90 Acc/ Weld ER308L 62.9 89.9 ASME Code Minimum Requirements
~
Pipe SA376 TP316 30.0 75.0 Pipg SA312 TP304L 25.0 70.0 Fittings SA403-WP316 30.0 75.0 Fittings SA403-WP316L 25 70 Fittings SA182-F316L 25 65 Welds E316L, E316LSi 70.0 Welds ER308L 75 a.
Range of material certification data b.
Range of averages or average 4-10
m TABLE 4-7 FRACTURE TOUGHNESS PROPERTIES TYPICAL OF THE ACCUNULATOR LINE Test Temp.
Tensile Properties (psi)
J Ic 2
Material
(*F)
Yield Ultimate (in-lb/in )
Tut a,c.e SA376 TP316 600 21,700 65,500 SA376 TP316 600 20,500 60,100 b
b Weld (E316) 600 45,000 61,200 d
Weld 600
[
ja,c.e
- b. Lowest of 6 tests.
a,c.e m>
SECTION 5.0 l
l l
LOADS FOR FRACTURE MECHANICS ANALYSIS h
Figures 5-1, 5-2 and 5-3 are schematic layouts of the three accumulator lines. Note that the high pressure region is not included in the scope of this work, because it already addressed in reference 5-1.
i The stresses due to axial loads and bending moments were calculated by the following equation:
o=k+$
(5.1)
- where, o
=
stress axial load F
=
bending moment M
=
A metal cross-sectional area
=
2 section modulus
=
i The bending moments for the desired loading combinations were calculated by l
the following equation:
M=
M 2+ M y
Z (5.2)
- where, M
bending moment for required loading
=
Y component of bending moment M
=
y 2 component of bending moment M
=
7 The axial load and bending moments for crack stability analysis and leak rate predictions are computed by the methods to be explained in Sections 5.1 and 5.2.
5-1
.4-...--~~.---------.--.,-+-,--,e.,---
,---.,,s---
y,--
w-
5.1 Loads for Crack Stability Analysis
}
The faulted loads for the crack stability analysis were calculated by the following equations:
lFDW + FTH1 + F l + lFSSE F
l (5.3)
=
p IIN )DW + IN )TH1 I + IIN )SSEI (5.4)
N Y
Y Y
Y IIN )DW
- IN )TH1 I + IIN )SSEl (5.5)
M Z
2 2
Z Where, the subscripts of the above equations represent the following loading
- cases, DW deadweight
=
TH1 maximum thermal expansion including applicable thermal anchor
=
motion SSE loading including seismic anchor motion SSE
=
load due to internal pressure P
=
5.2 Loads for Leak Rate Evaluation The normal operating loads for leak rate predictions were calculated by the following equations:
F FDW + FTH2 + Fp (5.6)
=
Y
@Y)DW + (N )TH2 (5.7)
N Y
(N )DW + (N IZ TH2 (5.8)
N Z
Z j
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.
5-2
5.3 Summary of Loads, Geometry and Materials Tables 5-1, 5-2, and 5-3 provide a summary of envelope loads computed for
)
fracture mechanics evaluations in accordance with the methods described in j
Section 5.1, and 5.2.
The cross sectional dimensions and materials are also summarized. Load data are tabulated at the highest stressed location (node 785M, loop 3), and the second highest stressed location (Node 80AE, loop 3),
for the 12 in. low pressure pipe in Table 5-1.
Table 5-2 gives load data for the 10 in. pipe for its highest stressed location (node 2998, loop 2), and its second highest stress location (node 1258, loop 1). Table 5-3 gives load data for the 8 in. pipe for its highest stressed location (node 284, loop 2) and itsnexthigheststressedlocation(node 141, loop 1).
The loading components are provided in tables 5-4, 5-5, and 5-6.
The critical locations are noted in figures 5-1, 5-2, and 5-3.
5.4 References 5-1 Swamy, S. A.,
et al.,
" Technical Bases for Eliminating Class 1 Accumulator Line Ruptures as the Structural Design Basis for South Texas Project Units 1 and 2," WCAP-11383, January 1987 (Westinghouse Proprietary Class 2.)
5-3
. ~..
TABLE 5-1 SIM4ARY OF ENVELOPE LOADS FOR 12 INCH LOW PRESSURE PIPE l
l OUTSIDE MINIMUN N
N0DE DIA.
WALL [a]
F (IN-LOCATION CONDITION NO.
LOOP MATERIAL (INCHES)
SCHEDULE (INCHES) (KIPS) KIPS)
I I
Highest Faulted 785M/TT9 3
SA376-TP316 12.75 140 1.005 48 1134 Load j
Location Normal Operating 785M/TT9 3
SA376-TP316 12.75 140 1.005 45 950 i
T Next Faulted 80AE/800 3
SA376-TP316 12.75 140 1.005 46 1032 Highest i
Load i
location Normal Operating 80AE/800 3
SA376-TP316 12.75 140 1.005.
38 836 I
i
[a] Taken at weld counterbore 8
n
j TABLE 5-2 SlM4ARY OF ENVELOPE LOADS FOR 10 INCH PIPE l
OUTSIDE MINIMUN N
NODE DIA.
WALL [a]
F (IN-LOCATION CONDITION NO.
LOOP MATERIAL (INCHES)
SCHEDULE (INCHES) (KIPS) KIPS)
Highest Faulted 299B 2
SA376-TP316 10.75 140 0.8955 40 290 Load Location Normal Operating 2998 2
SA376-TP316 10.75 140 0.8955 38 205 Next Faulted 125B 1
SA376-TP316 10.75 140 0.8955 41 238 T
Highest Load Location Normal Operating 1258 1
SA376-TP316 10.75 140 0.8955 39 158 4
1 i
[a] Taken at weld counterbore I
i l
I i
l I
l l
TABLE 5-3
SUMMARY
OF ENVELOPE LOADS FOR 8 INCH PIPE l
OUTSIDE MINIMUM N
NODE DIA.
WALL [a]
F (IN-LOCATION CONDITION NO.
LOOP MATERIAL (INCHES)
SCHEDULE (INCHES) (KIPS) KIPS)
Highest faulted 284 2
SA376-TP316 8.625 160 0.8135 26.7 328 Load Location Normal Operating 284 2
SA376-TP316 8.625 160 0.8135 24 214 i
Next Faulted 141 1
SA376-TP316 8.625 160 0.8135 27 302 Highest Load Location Normal Operating 141 1
SA376-TP316 8.625 160 0.8135 24 177
[a] Taken at weld counterbore
TABLE 5-4 LOADING COMPONENTS AT GOVERNING LOCATIONS FOR 12 INCH LOW PRESSURE LINE
. Highest Load Next Highest Load t
(Location - 785M/TT9, Loop-3)
(Location - 80AE/800, Loop-3)
Load Axial Bending 8ending Axial Bending Bending Type Force (1b)
Moment MY (ft-lb) Moment NZ (ft-lb)
Force (lb)
Moment NY (ft-lb) Noment NZ (ft-lb)
Dead
-2929 588 Weight 3529
-10903 489 457 Thermal
-9767
-5148 75536
-8593 49007 48613 Pressure 57589 57589 E+
3479 7658 14679 8254 4813 17577 I
TABLE 5-5 LOADING COMPONENTS AT GOVERNING LOCATIONS FOR 10 INCH LOW PRESSURE LINE l
Highest Load Next Highest load (Location - 2998, Loop-2)
(Location - 1258, Loop-1)
Load Axial Bending Bending Axial Bending Bending Type Force (Ib)
Noment NY (ft-lb) Noment NZ (ft-lb)
Force (Ib)
Noment NY (ft-lb) Noment NZ (ft-lb)
Dead
-636 2149 1999
-92
-315 1504 I
Weight Thermal
-1482 4288 13804
-1142
-3940 10960 1
Pressure 39986 39986 1
l
~
i T
SSE +
2326 2396 6648 2305 2788
6075 I
Anch. Not.
t
'M
's ik i g
kN c
,3
~ _
s t
6 b
~'
~
n_,
TABLE 5-6 LOADING COMPONENTS AT GOVERNING LOCATIONS FOR 8 INCil LOW PRESSURE LINE Highest Load Next Highest Load (Location - 284, Loop-2)
(Location - 141, Loop-1)
Load Axial Bending Bending Axial Bending Bending Type Force (1b)
Moment NY (ft-Ib) Noment NZ (ft-lb)
Force (Ib)
Noment NY (ft-Ib) Noment NZ (ft-lb)
Dead
-65
-4690
-1633 48 371
-4159 Weight Thermal
-205
-11689
-5420
-29 11483
-4559 Pressure 24246 24246 i
SSE +
2715 8580.1 4099.3 2732 7937 6847 Anch. Mot.
l e
9
\\
1 i
Node: 141.Next Highest Load for 8 in. Pipe Node: 1253 Next' Highest Load for 10 in. Pipe I
l b/
S h
141 125B l
l
's 1
P = 665 psi T = 120'F FIGURE 5-1.
Schematic Layout of Accumulator Line Loop 1 G
5-10 l
l Node 284: Highest Load for 8 in. Pipe
)
Node 299B: Eighest Load for 10 in. Pipe M
~
1 Z
'"" ?
284 T
l P = 555 p,,
T = 1 goop FIGURE S.g Schematic Layout of Accumulator Line Loop g 5-11
/
C Node 785N/TT9: Highest Load for 12 in. Pipe Node 80AE/800-: Next Highest Load for 12 in. Pipe 80AE/800 Y
785M/TT9 P = 665 psi T = 120*F s
FIGURE 5-3.
Schematic Layout of Accumulator Line Loop 3 5-12
5.
SECTION 6.0 FRACTURE MECHANICS EVALUATION 6.1 Global Failure Mechanism
{.
Determination of the ccnditions which lead to failure in stainless steel i
should be done with plastic fracture methodology because of the large amount of deformation accompanying fracture.
One method for predicting the failure of ductile material is the (
la,c.e method, based on traditional plastic limit load concepts, but accounting for [
]a,c.e and taking into account the presence of a flaw. The flawed pipe is predicted to fail when the remaining not 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.
The flow stress is generally taken as the average of the yield and ultimate tensile strength of the mat'erial at the temperature of interest. This methodology has been shown to be applicable to ductile piping through a large number of experiments and will be used here to predict the critical flaw size in the accumulator line. The failure criterion has been obtained by requiring equilibrium of the section containing t.he flaw (Figure 6-1) when loads are applied.
The detailed development is provided in i
Appendix A for a through-wall circumferential flaw in a pipe with internal pressure, axial force, and imposed bending moments. The limit moment for such a pipe is given by:
a,c.e
[
]
(6.1) where:
ja.c.e 6-1
---g%w.-t'---*---
--r-7-'w--w m--ww-ww r'www-,----wv.-ww i--v
- =-
-v---------
-- + - -- - - --- -
i S
i r
i Ja,c.e (6.2) i The analytical model described above accurately accounts for the piping iriernal 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 accumulator pipe 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 figures 6-3, 6-4, and 6-5 for 12 in.,10 i
in., and 8 in. accumulator lines operating at 120*F, respectively.
The.
critical flaw size corresponds to the intersection of this curve and the maximum load line.
The critical flaw sizes of [
~
Ja,c e pipe operating l
at 120*F and 665 psig are calculated using ASME Code (reference 6-2) minimum tensile properties for SA376 TP316 (wrought) stainless steel. Figures 5-1, 5-2, and-5-3 identify the locations of the critical-regions.
If the limit moments are larger than the maximum applied moments for cracks l
smaller than the critical flaw sizes corresponding to the maximum applied moments, then the global stability criteria of section 2.2 are satisfied.
- Figures 6-3 through 6-5 show the critical flaw size predictions. The critical flaw sizes corresponding to the limit moments are shown below for the 12 in.,
10 in., and 8 in. pipes:
6-2
-. -. =
.--,--.we-------..
...-m...
r ir--
-,w-
--.rv-,
~,..---,-,-.--...,----e-,
- - - = -,, - - *, - - - - - - - - -, _ - - -
~
12 inch pipe a,c.e I
10 inch pipe 8 inch pipe If the IWB 3640 approach, reference 6-9, is used and if the material strength properties are conservatively assumed to be the same as the base metal properties, tha critical flaw sizes for the weld metal for 12 in.,10 in., and 8 in. pipes operating at 120*F are as follows:
- i 12 inch pipe a,c.e 10 inch pipe 8 inch pipe 6.2 Leak Rate Predictions Fracture mechanics analysis shows that postulated through-wall cracks in the accumulator 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 this section is to discuss the method which will be used to predict the flow through such a postulated crack and present the leaf rate calculation results for through-wall circumferential cracks.
6.2.1 General Considerations l
' The pressure in the 12 in.,10 in., and 8 in. pipes is 665 psig at 120*F.
Therefore, the fluid inside the pipe is subcooled liquid. Since the saturation temperature at atmospheric conditions is 212'F, subcooled liquid cannot evaporate when the fluid expands ~to atmospheric pressure. Hence the leak rate calculation should be performed based on single phase (liquid only) flow.
6-3
6.2.2 Calculation Method The leak rate for this case is obtained by using the simple orifice type flow formula given by [-
ja.c.e 6.2.3 Leak Rate Calculations Leak rate calculations were made as a function of postulated through-wall crack length for the critical locations previously identified. The crack opening area was estimated using the method of Reference 6-5 and the leak rate was calculated using equation 6.3.
The leak rates are calculated using the-normal operating loads of axial force F and bending moment M for 12 in.,10 in., and 8 in, accumulator lines operating at 120*F. These are given directly below.
6-4 a
um a
i m
F=
45 kips, M = 950.36 in-kips for 12 in. pipe F=
38 kips, M = 205.0 in-kips for 10 in. pipe F=
24 kips, M = 214 in-kips for 8 in, pipe i
The leak rates for various postulated crack lengths for the 12 in.,10 in.,
and 8 in. pipe lines are shown in figures 6-6 through 6-8 respectively.
In these figures, the crack length yielding a leak rate of 5 gpm (10 times the leak detection capability of 0.5 gpm detection capability for the plant.are found to be as follows:
a,c.e Thus " reference" flaw sizes of (4.3 in., 6.5 in., and 4.8 in.Ja,c.e for respective pipe diameters 12 in., 10 in., and 8 in. are established.
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 locations, the outer surface axial stresses, (o,), the circumferential stresses, ( c), and radial stresses, (o ), f r 12 in.,
r 10 in., and 8 in. pipes operating at 120*F and internal pressure of 665 psi l
are calculated to be as follows, using formulas from reference 6-6:
o, = 12.5 ksi
= 3.25 ksi o = 0 ksi (12 in. pipe) c r
6.03 ksi 3.024 ksi 0 ksi (10 in. pipe) 10.53 ksi 2.562 ksi 0 ksi (8 in. pipe)
The Von Mises effective stress, o,ff, (see reference 6-7) is given by "eff '
II r) *I r) *I "c)
~
a c
a l
un, um no 6-5
and the o,ff for 12 in.,10 in., and 8 in. pipe operating at 120*F are calculated to be as follows:
12 in. pipe o,ff = 11.2 ksi 10 in. pipe o,ff = 5.22 ksi 8 in. pipe o,ff = 9.51 ksi Thus the effective stress is less than the yield stress and by the Von Mises plasticity theory yielding does not occur. Hence, linear elastic fracture mechanics is applicable for analyzing the pipes with hypothesized flaws. The analytical method used for the local stability evaluation at this location is summarized below.
The stress intensity factors corresponding to tension and bending are expressed, respectively, by (see reference 6-5.)
Kt " 't
/ wa FI")
t Kb * 'b
/ wa F(*)
b where F (a) and F (a) are stress intensity calibration factors corresponding t
b to tension and bending, respectively, a is the half-crack length, a is the half-crack angle, e is the remote uniform tensile stress, and is the t
b remote fiber stress due to pure bending. Data for F (a) and F (")
8"'
t b
given in Reference 6-5.
The effect of the yielding near the crack tip can be incorporated by Irwin's plastic zone correction method (see Reference 6-8) in which the half-crack length, a, in these formulas is replaced by the effective crack length, a,ff, defined by 2
1 K
a,ff = a + - q 2n o,
6-6
i for plane stress plastic corrections, where o is the yield strength of y
l the material and K is the total stress intensity factor due to combined l
tensile and bending loads (i.e., K = K +K).
Finally, the J,pp value b
E is determined by the relation J,pp = K /E, where E is Young's Modulus.
J,pp was calculated using normal plus SSE loads for a flaw two times as long as the " reference" flaw for the three pipe sizes and the resulting numerical values are given below:
12 in. pipe flaw size =
a,c.e
~
~
a,c.e 10 in. pipe flaw size =
8 in. pipe flaw size =
In addition, for the leakage size flaws i.e., the reference flaws for the three pipes, the normal plus SSE load was increased by d for all the pipe lines. Calculations for J using these increased loads yield the following:
12 in. pipe, a,c.e 10 in. pipe, i
8 in. pipe, Clearly, the applied J is lower than [
Ja,c.e for both the above cases and therefore unstable crack propagation will not result.
t 6.4 References 6-1 Kanninen, M. F. et al., "Nechanical Fracture Predictions for Sensitized q
Stainless Steel Piping with Circumferential Cracks" EPRI NP-192, September 1976.
6-2 ASME Section III, Division 1-Appendices,1986 Edition, July 1,1986.
i 6-3 Crane, D. P., " Handbook of Hydraulic Resistance Coefficient."
6-7
6-4 Dillio, C. C., Thermal Engineering Int. Textbook Co. Scranton, PA p.
273, 1969.
6-5 Tada, H.,
"The Effects of Shell Corrections on Stress Intensity Factors
(
and the Crack Opening Area of Circumferential and a Longitudinal Through-Crack in a Pipe," Section 11-1, NUREG/CR-3464, September 1983.
6-6 Durelli, A. J., et. at., Introduction to the Theoretical and Experimental Analysis of Stress and Strain, McGraw Hill Book Company, New York, (1958),pp.233-236.
6-7 Johnson, W. and Mellor, P. B., Engineering Plasticity, Van Nostrand Reimhold Company, New York, (1973), pp. 83-86.
6-8 Irwin, G. R., " Plastic Zone Near a Crack and Fracture Toughness," Proc.
7theSagamoreConference,P.IV-63(1960).
6-9 " Evaluation of Flaws in Austenitic Steel Piping" - EPRI Final Report, April 1986, Prepared by Section XI Task Group for Piping Flaw Evaluation.
6-8
2
.'s,--
,m 24, J
-ma a
s -
a--
- W O
l 1
i i
U.
O+
I i
e U
m E+
C
.O e-M l
j
.a
=
Aa 4
M
=a.
m E
~
c-,
M e
l a
w a
a f
O I
I G
}
W b
J N
i l
6 i
J e
J t
i i
l a
G 6-9
l l
.. j T*'( + a
.9
. 1.
=
I o
8, l
l l
I l
I l
t I
r l
I 1
s L.q.
l J
g i
l 4
l l
l 3
l l
I I
1 1
I c
1 1
I i
V V
a
-.1 E.E.!!
$f S 14. M 6-10
_ a,c.e I
nam accu 00 = 12.75" t = 1.005" P = 665 p.sig F = 48 kips e = 29.16 ksi y
o = 75.0 ksi u
op = 52.08 ksi 0
Temp = 120 F l
l l
Figure 6-3 " Critical" Flaw Size Prediction for 12 Inch Accumulator Line Operating at 120*F 6-11
=____-_-.
a,c.e 1
L i
k l
l FLtm oterm t
OD = 10.75 t = 0.8955 P = 665 psig F = 40 kips o = 29.16 ksi y
o,= 75.0 ksi op = 52.08 ksi 0
Temp = 120 F i
(
l l
l i
l Figure 6-4 Critical Flaw Size Prediction for 10 Inch Accumulator Line Operating at 120'F 6-12
l i
l a,c.e t
I
/
I k
i PLAsf ESETRT r
00 = 8.625" t = 0.8135" P = 665 psig F = 27 kips o = 29.16 y
o = 75.0 ksi y
op = 52.08 ksi 0
Temp = 120 F i
l i
l Figure 6-5 Critical Flaw Size Prediction for 8 Inch Accumulator Line Operating at 120'F 6-13
_ a.c.e Figure 6-6 Leak Rate Versus Crack Length for the 12 Inch Accumulator Line Operating at 120*F 6-14
~
..a,c.e l
Figure 6 7 Leak Rate Versus Crack Length for the 10 Inch Accumulator Line Operating at 120*F 6-15
(
a,c.e e
Figure 6-8 Leak Rate Versus Crack Length for the 8 Inch Accumulator Line Operating at 120*F 6-16 9
a
\\
SECTION 7.0 i
ASSESSMENT OF FATIGUE CRACK GROWTH 7.1 Acceptability of Fatigue Crack Growth The purpose of the fatigue crack growth (FCG) analysis is to demonstrate that a postulated flaw will not grow through the wall under all design and operational loadings. An assessment of the FCG for the South Texas 12, 10, and 8 inch low pressure accumulator pipe was done by comparison with a similar analysis.
For the high pressure portion of the accumulator line a generic FCG analysis was applied (reference 1). Loadings were reviewed on the low pressure piping and it was determined that the generic evaluation would envelop this piping also. In fact the loadings are so small on the low pressure piping that the generic FCC analysis would grossly overpredict the crack growth for the low pressure piping.
In addition to loadings, the material and geometries (wall thickness) of the low pressure piping were reviewed and were found to be compatible with the generic analysis. Appendix B discusses the generic FCG analysis in detail.
In conclusion, the generic fatigue crack growth analysis will envelop the fatigue crack growth on the low pressure portions of the South Texas-accumulator lines.
l The actual crack growth numbers are summarized in Table B-4 of Appendix B.
l 7.2 References 7-1 Swamy, S. A., et. al., " Technical Basis for Eliminating Class 1 Accumulator Line Rupture as the Structural Design Basis for South Texas Project Units 1 and 2" (WCAP-11383 January 1987, Westinghouse ProprietaryClass2).
l 7-1
. u
SECTION 8.0 ASSESSMENT OF MARGINS In the preceding sections, the leak rate calculations, fracture mechanics analysis and fatigue crack growth assessments were performed. Margins are discussed below.
The maximum nominal axial stresses at the outside surface of the pipes for 12 in.,10 in. and 8 in. presented in section 6.3 are compared with the yield stress, o = 29.2 ksi at operating temperature. The results are given y
below:
12 inch pipe:
a, = 12.5 ksi, T = o,/o = 0.43 j
10 inch pipe:
o, = 6.03 ksi, T = o,/c = 0.21 y
8 inch pipe:
a, = 10.53 ksi, T = o,/o = 0.36 j
Thus, the maximum faulted condition axial stress including the deadweight, thermal, pressure and SSE loads is less than half of the yield strength of the material of the pipes.
In section 6.1, the critical flaw sizes using limit load methods for the above three different pipe sizes are calculated. Using the IWB-3640 approach, the critical flaw sizes for the three different pipe sizes are as calculated in section 6.1.
Those results are as follows:
12 inch pipe Critical flaw size using limit moment:
[
la,c.e Critical flaw size using IWB-3640:
(
Ja,c.e 10 inch pipe Critical flaw size using limit moment:
(
Ja,c.e Critical flaw size using IWB-3640:
(
Ja,c.e 8-1
8 inch pipe i
Critical flaw size using limit moment: (
Ja,c.e Critical flaw size using IWB-3640:
[
Ja.c.e In section 6.3 it is seen that the J values at maximum load for the three different pipe sizes are as given below:
l l
~
12 inch pipe:
J,pp =
j a,c.e 10 inch pipe:
J,pp =
8 inch pipe:
J,pp =
These J values are lower than the lower bound J value of (
Ja,c.e Ic considering thermal aging effects. The flaw sizes yielding a J value of (
Ja c.e for the three different pipe sizes are:
"'Cd
~
~
12 inch pipe:
10 inch pipe:
8 inch pipe:
Based on the above data, the critical flaw sizes will exceed [
Ja,c.e foi the 12 in.,10 in., and 8 in, pipes respectively.
In section 6.2 it is seen that at the critical locations the references flaw sizes (Cf) for the three different pipe dimensions are as follows:
a,c.e for 12 inch pipe for 10 inch pipe for 8 inch pipe These yield a leak rate of 5 gpm. Thus, there is a margin of at least, a,c.e
._ = 3.7 for the 12 inch pipe
= 2.4 for the 10 inch pipe
= 2.1 for the 8 inch pipe 8-2 e
,---__m__
y
,-,,_____,________,..____.-----,_~__.,,.-._m__.
--- m....
on flaw sizes. Furthermore, thermal aging cannot take place at the 120'F operating temperature and the J values for all three pipe sizes should be e
higher than [
)"'C' Therefore, the margin shown should also have a larger value.
In section 6.3 it was shown that the reference flaws yielding a leak rate of 5 gpm would be stable when subjected to a load equal to /2 (normal plusSSE).-
In summary, relative to
~
1.
Loads Maximum stresses at the critical locations in the 12 in.,10 in.,
a.
and 8 in. pipes are less than 43 percent of the ASME Code minimum yield strength at temperature, 120*F.
b.
The leakage-size crack will not experience unstable crack extension even if larger loads of /2 (normal plus SSE) are applied.
2.
Flaw size a.
A minimum margin of 2.1 exists in the 12 in.,10 in., and 8 in, pipes between the critical flaw and the flaw yielding a leak rate of 5 gpm.
b.
If the limit load is used as the basis for critical flaw size, the minimum margin for global stability in the 12 in.,10 in., and 8 in.
pipes is 2.7 (based on the 8 in, pipe.)
3.
Leak Rate A margin of 10 exists for the reference flaws; (
Ja.c.e for 12 inch pipe, [
Ja c.e for 10 inch pipe and [
]a,c.e for 8 inch pipe between the calculated leak rate and the leak detection capability of the plant.
8-3
A summary comparison of criteria and analytical results is given in Table 8-1.
The criteria are seen to be met for all three different pipe. sizes.
l i
I d
8-4
TABLE 8-1 COMPARISON OF RESULTS VS. CRITERIA CRITERION RESULT 1.
NUREG1061 Volume 3 Net Section 5.2(h) -
(Required margin of 2 demonstrated)
Margin on Flaw Size 2.
NUREG1061 Volume 3 Net Section 5.2(i) -
(Required margin of /2 demonstrated)
Margin on Load 3.
NUREG 1061 Volume 3 Net Section 5.7 -
(Nargin of 10 on leak rate Margin on Leak Rate demonstrated) 4.
NRC criteria on allowable Net fatigue crack growth
(
Ja,c.e (af < 60% wall thickness) 5.
NRC criteria on allowable Net fatigue crack growth a,c.e (Plastic zone size < remaining ligament) 8-5 9
e
.,----,a---,----e-
--,--,,-,,,-, _ a r-- - - - - -, --,,,-- - - - -- - - -, - - -,, - - - -
l SECTION
9.0 CONCLUSION
S h
This report justifies the elimination of accumulator line pipe breaks (low pressure segment) for the South Texas Project Units 1 and 2 as follows:
l a.
Stress corrosion cracking is precluded by use of fracture resistant l
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 accumulator line piping are negligible.
d.
Ample margin exists between the leak rate of small stable flaws and the leak detection capability of the plant.
e.
Ample margin exists between the small stable flaw sizes of item d l
and the critical flaw.
f.
Ample margin exists in the material properties used to demonstrate end-of-service life (relative to aging) stability of the critical flaws.
g.
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).
9-1
- - ~.
y
'l
- /
=
4 g
/
3,,
- g. '
f The reference flaw will be stable throughout reacterTifs htcause of the ample margins in d, e, f and g and will. leak at a detectable rate which will assure
.s t
a safe plant shutdown.
.~.
i Based on the aboves it is concluded that dynamic effects resulting from accumulator line (low pressure Class);l~ portion) pipe breaks"should'not-be considered in the ' tructural desibn insi pf South Te(as Project Units 1 and 2.
k s
f 1
1 1
s..
3 9
i 4
t I
- e 9-2 i
. _ _ - ~. _ _ _ _. _ _ _. _ _,.
). a O
P l
e APPENDIX A LIMIT MONENT
)
e f
e A-1 7
m_
4 APPENDIX A LIMIT M0 MENT
[
e 4,C,9
)
O S
e A-2
(
w t
a,c.e n
FIGURE A-1 Pipe with a Through-Wall Crack in Bending A. 3
l b
h APPENDIX B FATIGUE CRACK GROWTH CONSIDERATIONS B-1
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 transient was calculated for i) the time corresponding to the maximum inside surface stress and,11) the time corresponding to the minimum inside surface stress. These two stress profiles are called the maximum and minimum through wall stress distribution, respectively for convenience. The constant stresses due to pressure, deadweight and thermal expansion (at normal operating temperature, 550*F) loadings were superimposed on the through wall cyclical stresses to obtain the total maximum and minimum stress profile for each transient. Linear through wall stress distributions were calculated by conservative simplified methods for all minor transients. More accurate nonlinear through wall stress distributions were developed for severe transients by [
3a,c.e B.I.1 Critical Location for Fatigue Crack Growth Analysis The accumulator line stress report design thermal transients (Section B.1.2),
1-D analysis data on accumulator line thermal transient stresses (based on ASME Section III NB3600 rules) and the geometry were reviewed to select the worst location for the fatigue crack growth analysis.
[
l Ja c.e This location is selected as the worst location based on the following considerations:
i) the fatigue usage factor is highest.
l ii) the stress due to thermal expansion is high.
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 (
3a,c.e B-2
B.I.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 conditions in the plant. These are representative of the conditions which are considered to occur during plant operation. The fatigue evaluation based on these transients provides confidence that the component is appropriate for its application over the design life of the plant. All the normal operating and upset thermal transients, in accordance with design specification and the applicable system design criteria document (B-1), were considered for this evaluation. Out of these, only [
Ja,c.e These transients were selected on the basis of the following criteria:
+a,c.e (B.1)
(B.2)
- where,
+a,c.e B.I.3 Simplified Stress Aaa?ysis The simplified analysis method was used to develop conservative maximum and minimum linear through wall stress distributions due to thermal transients.
l
[
]a,c.e The inside surface stress was calculated by the following l
equation which is similar to the transient portion of ASME Section III NB3600, Eq. 11:
Sg=[
la,c.e (B.3)
B-3
- where,
+a,c.e<
l l
l
[
Ja,c.e The maximum and minimum inside surface stresses were searched from the S$ 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)+a,c.e, 502 = [
]
(B.8)+a,c.e, l
B-4 l
l n -,-
I where.
l l
+a,c.e,
(
The material properties for the accumulator pipe [
l
)*'C
The values of E and e, at the normal operating i
temperature, provide a conservative estimation of the through wall thermal transient stresses as compared to room temperature. properties. The following values were conservatively used, which represent the highest of the [
Ja.c.e materials:
~~
+a c.e 1
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
[
Ja.c.e Nonlinear through wall stress profiles were developed for the remaining severe transients as explained in l
Section B.I.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 I
simplified method provides more conservative crack growth.
B-5
a 8.1.4 NonlineshStressDistributionforSevereTransients
[
]
As mentionei earlier, the accumulator line section near the [
]a,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.
[
t l
l 3a,c.e B.1.5 OBE Loads The stresses due to OBE loads were neglected in the fatigue crack growth analysis since these loads are not expected to contribute significantly to crack growth due to small number of cycles.
B-6
k-
~ '
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 stresses due to deadweight and thermal expansion (normal operating case) on the thermal transient stresses (of Table B-2 or the nonlinear stress distributions discussed in Section B.1.4). Thus, the total stress for fatigue crack growth at any point is given by the following equation:
Total Thermal Stress Due Stress for Transient to Due to Fatigue
=-
+
DW +
+
Internal (8.9)
Crack Growth Thermal Pressure Expansion The envelope thermal expansion, deadweight and pressure loads for calculating the total stresses of Equation B.9 are summarized in Table B-3.
B.2 Fatigue Crack Growth Analysis 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 in Table B-1.
B.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 c
l B-7
~,
1 I
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 AK, by the following i
g relation:
h=CoAK" (B.2.1) g where "Co" and the exponent "n" are material properties, and AK is g
defined later, in Equation (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 "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, which depends on g
crack and structure geometry and the range of applied stresses in the area y is calculated, the growth due to that where the crack exists. Once AK particular cycle can be calculated by Equation (B.2.1). This increment of growth is then added to the original crack size, the AKy adjusted, and the analysis proceeds to the next transient. The procedure is continued in this manner until all the transiants have been analyzed.
The crack tip stress intensity factors (K ) to be used in the crack growth g
analysis were calculated using an expression which applies for a semi elliptic surface flaw in a cylindrical geometry (B-4).
The stress intensity factor expression was taken from Reference B-4 and was.
calculated using the actual stress profiles at the critical section. The maximum and minimum stress profiles corresponding to each transient were input, and each profile was fit by a third order polynomial:
c (x) = A + Ay j+ A ({} + A ({)
(B.2.2) 0 2
3 l
l B-8
The stress intensity factor K (e) was calculated at the deepest point of y
the crack using the following expression:
+a,c.e (B.2.3) l Calculation of the fatigue crack growth for each cycle was then carried out using the reference fatigue crack growth rate law determined from consideration of the available data for stainless steel in a pressurized water environment. This law allows for the effect of mean stress or R ratio (K! min /Elu x) on 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:
h"I 3 "#
(B.2.4) p l
l l
B-9
\\
where K,ff = (K,,,) (1-R)1/2 g
1 K "i" I
R=KImax h=crackgrowthrateinmicro-inches / cycle B.2.2 Results-Fatigue crack growth analyses were carried out for the critical cross section. Analysis was completed for a range of postulated flaw sizes oriented circumferential1y,'and the results 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 (0.300 inch which is 33 percent of the wall thickness,]a,c.e the result shows that the flaw grewth through the wall will not occur during the 40 year design life of the plant. For smaller flaws, the flaw growth is significantly lower. For example, a postulated (.100]a,c.e inch deep flaw will grow to [.132"]a,c.e which is less than (3/16]a,c.e the wall thickness. These results also confirm operating plant experience.
There have been no leaks observed in Westinghouse PWR accumulator lines in over 400 reactor years of operation.
B'. 3 REFERENCES a,c.e B-1 B-2 ASME Section III, Division 1-Appendices, 1983 Edition, July 1, 1983.
B-3 WECAN -- Westinghouse Electric Computer Analysis, User's Manual -- Volumes 1, II, III and IV, Westinghouse Center, Pittsburgh, PA, Third Edition, 1982.
l B-10
B-4 McGowan,lJ. J. and Raymund, M., " Stress Intensity Factor Solutions for Internal Longitudinal Semi-Elliptical Surface Flaws in a Cylinder Under Arbitrary Loadings" Fracture Mechanics ASTM STP 677, 1979, pp. 365-380.
B-5 Bamford, W. H., " Fatigue Crack Growth of Stainless Steel Reactor Coolant Piping in a Pressurized Water Reactor Environment", ASME Trans. Journal of Pressure Vessel Technology, February 1979.
)
e i
e B-11
6 TABLE B-1 THERNAL TRANSIENTS CONSIDERED FOR FATIGUE CRACK GROWTH EVALUATION Trans.
No. of No.
Description Occurrences
+a c.e
~
l b
i I.
l i
I e
e O
l l
l B-12 l
s u
m a
W a
=N 3
I m5 a
W ra 8*8 v
m m
W sm b
mb m
W EU to m
44 5
W y
gG
.m
'W
=
= m E
E 8 ~8 u
,=>
5 m
m to M
C O
g "*
m x w 12m5 80
.6 EG EW E
a I
i s.c B-13
s.
TABLE B-3 ENVELOPE NORMAL LOADS CONDITION a,c.e Nonnal Operating e
6 e
l l
8-13a e
TABLE 8-4
?
ACCUMULATOR LINE FATIGUE CRACK GROWTH RESULTS Wall Thickness = [
]+
+a,c.e INITIAL CRACK LENGTH AFTER YEAR CRACK LENGTH.
10 20 30 40 (IN.)
M M
+4,C,e m
D
)'
s e
B-14 m
l i
]
+a.c.e l
l Figure 8-1 Comparison of Typical Maximum and Minimum Stress Profile Computed by Simplified [
]
+a,c.e B-15
+a c.e 2.
+a,c.e Accumulator Pipe
+a,c.e Figure B-2 Schematic of Accumulator Line At
[
]
B-16 m-v
~
i t
+a.c.e
)
i 1
)
i
+a,c.e Figure B-3 [
] Maximum and Minimum Stress Profile for Transient #10 B-17
i
+a,c.e
~
1 g
+a,c.e Figure B-4 [
] Maximum and Minimum Stress Profile for. Transient #11 8-18
~
+s.c.e i
I
(
+a.c.e Figure B-5 [
] Maximum and Minimum Stress Profile for Transient #12 B-19 l
I 4
+Ee8oe 3
}
+a,c.e Figure B-6 [
] Maximum and Minimum Stress Profile.for Transient #14 1
B-20
--r
-- - - - ' ' ' - - -