ML20205R408
| ML20205R408 | |
| Person / Time | |
|---|---|
| Site: | South Texas  |
| Issue date: | 08/31/1988 |
| From: | Manley R, White R ANAMET LABORATORIES, INC. |
| To: | |
| Shared Package | |
| ML20205R372 | List: |
| References | |
| 8804-06, 8804-07, 8804-6, 8804-7, NUDOCS 8811100020 | |
| Download: ML20205R408 (59) | |
Text
_ _ _ _ _ ._.
__e _ a ATTACHMENT ^
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-GT HbAEr27' PAGE l 0F FAILURE ANALYSIS AND STRUCTURAL INTEGRITY EVALUATION OF LEAKING SMALL BORE ALUMINUM-BRONZE CAST VALVE BODIES AND FITTINGS IN THE ECW SYSTEM 4
Report for Houston Lighting and Power Company South Texas Project Prepared for K. G. Hes s a
By /
'R. V White Matterials and Corrosion Group Manager l
I Approved MM - A '/
i R. A. Manley
- g Manager v i
s Materials and Quality Services Research and Development BECHTEL NATIONAL. INC.
SAN FRANCISCO I
Job No. 19280-001-042 Tech. Report No. 8804-06 FA Log No. 0318157 BLN No. 8804-07 1071m August 1988 S Revision 3 0011100020 681101 8 F;DR ADOCK O'3>0
u AT TACHMENT ST.HL.AE 47 PAGE n OF ,
q ABSTRACT ..
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t' This taport presents the metallurgical f y;1ure analysis and the attJCtural f integrity evealuation of the leaking small bore (2-inch diameter and smaller) d aluminum-bronse castings in the Essential Cooling Water system (ECW) at the y South Texas Project (STP) Unit No.1. J c
The failure analysis of two valves (one leaker, one non-leakar) and one fitting O consisted of metallurgical cross sections, heat treatment studies, chemical and X-ray diffraction studies, microstructure analysis and the verification that 1 the field replication procedure is satisfactory on aluminum 'oronse. The failure analysis revealed that the leaks are caused by dealloying corrosion. I j The plain ca et aluminum-bronse alloys in use (952 and 954) are susceptible to k this form of corrodion. l i
The structural integrity evaluation consisted of cross sectioning and mapping i the dealloying in 32 additional casti' go, performing miniature tensile tests on {
j eastings with known dealloying, perfo ning proof tests on three fittings, two
! containing through wall leakage, and a statistical analysis of the extent of :
the dealloying plus fracture and limit load analysis.
- The metallurgial and structural integrity analysis has demonstrated that the 9 i material has significant margin against failure for the design loading l conditions. Proof tests of actual components with des 11oying demonstrated load ;
5 capacity on the orde; of 50 times the design pressure. ;
j a
i Key Words
! dealloying i a luminum-b ronse
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5 071m u - _ . _ _ . - , _
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ATTACHMEN,T 2 ^
ST HbAE<2 PAGE 92 OF rd"lh FAILURE ANALYSIS AND STRUCTURAL INTEGRITY OF LEAKING SMALL BORE ALUMINUM-BRONZE CAST VALVE BODIES AND FITTINGS IN THE ECW SYSTEM CONTENTS
.P,agg ADSTRACT ................................................................ i CONTENTS................................................................ 11 L I S T O F I LLU ST RAT ION S . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . iii 1.0 FAILURE ANALYSIS OF FIRST THREE SAMP LES . . . . . . . . . . . . . . . . . . . . . . 1.1 1.1 In t ro d u c t i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.1 1.2 Co nclusions o f the Firs t Diagnostic Phase . . . . . . . . . . . . . . 1.1 1.3 Materials .............................................. 1.2 1.4 Ev a lua t i o n Me thod s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.2 1.5 Dis cu s s io n o f Re sul t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.3 2.0 STRUCTURAL INTEGRITY ..................... ................... 2.1 2.1 In t r od u c t i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.1 2.2 Statistical Analysis ................................... 2.1 2.3 S t r e s s Ev a l u a t i o n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.3 2.4 Fracture Mechanics Integrity Analysis................... 2.5 2.5 P r oo f Te s t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.8 1
2.6 Dealloying Rate Studies................................. 2.9 APPENDIX Crack-Tip Opening Displacement Testing of Valve-Body Material / !(
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. ATTACHME ST44L Me*j],jil -
PAGE V OF69' j
LIST OF ILLUSTRATIONS l
TABLES 18.1 1 Table 1.1 Analysis of Cross Sections ................................ 1.5 Table 1.2 Otemical Analyses ......................................... 1.7 Table 2.1 Statistical Sampling of ECW System Fittings . . . . . . . . . . . . . . . 2.11 Tahle 2.2 Design and Operating Data ................................. 2.13 Table 2.3 Summary of Fitting Dimensions and Highest Stress Levels 2.14 Table 2.4 Stress values.............................................. 2.15 Table 2.5 Tensile Tes t Re sult s on De *11oyed Sample s . . . . . . . . . . . . . . . . . 2.23 FICURES Figure 1.1 As - r e c e iv e d p a r t s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.8 Figure 1.2 Ty p i c a l m i c r o s t ru c t u r e s . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.9 Figure 1.3 Lo ca tion o f c ros s s ec t i ons , valve EW2 69. . . . . . . . . . . . . . . . . . . . 1.11 Figure 1.4 Loc a tion o f c ro s s sec tions , valve EW315 . . . . . . . . . . . . . . . . . . . 1.12 Figure 1.5 Lo ca t ion o f c ro s s s e c tions , elbow . . . . . . . . . . . . . . . . . . . . . . . . . 1.13 Figure 1.6 Comparison of photomicrographs of a specimen a nd a r e pl i c a o f t he s pe c ime n . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1.14 Figure 2.1 Bending Stress capacity for Dealloying 2-inch Fitting...... 2.16 Based on LEFM Figure 2.2 Conditions for Net-Section Plastic Collapse for 360'. ...... 2.17 Part-Through Dealloying Fig:xe 2.3 Conditions for Net-Section Plastic Collapse for. . . . . . . . . . . . 2.18 Through-Wall Dealloying Figure 2.4 Aluminum Bronze Fitting Stress Ratio to Allowable.. . . . . . . . . 2.19 Figure 2.5 Dis t ribution o f Sampled Fitt ings. . . . . . . . . . . . . . . . . . . . . . . . . . . 2.20 Figure 2.6 De alloying o f Aluminum B rons e . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2.21 Figure 2.7 Aluminum Bronze Fitting Stress Ratio to p Allowable ( As Designed) ................................... 2.22 6 t
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-iii-1071m I
____________________.-._-_._____J
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r MTACHMENT ST HL AE MY PAGE 6 OF5 1.0 FAILURE ANALYSIS OF FIRST THREE SAMPLES 1.1 Introduction On April 1, 1988, weeping type leaks were found in valves and fittings in the 2 inch and smaller piping in the Essential Cooling Water System (ECW). Ric was in unit number 1 of the Scuth Texas Project (STP). As of May 1,1988, 90 of the approximately 800 valves and fittings had signs of weeping. h e weeping is limited to the valves and fittings, which are all aluminum-bronze eastings. he aluminum-bronze piping and welds show no sign of weeping. h is is consistent with the literature that shows that alloys with less than 8 percent aluminum (wrought pipe and weld metal) are resistant to dealloying and alloys with greater than 8 percent aluminum (valve and fitting castings) are subject to dealloying under certain conditions.
Bechtel's Materials and Quality Services was initially requested to perform a failure analysis of a weeping valve and a fitting. he intent of
' this initial analysis was to identify the mechanism that is causing the weeps and the underlying metallurgical, design and environmental factors that led to the failures.
1.2 Conclusions of the First Diagnostic Phase A. De leaks are caused by dealloyir.g corrosion going through the wall of the cast valve or caat fitting.
B. Only castings have any eign of dealloying.
C. Bere is no sign of dealloying corrosion in either the weld metal or tha wrought Al-bronze pipe.
D. he greatest amount of through-wall corrosion occurs in the region of the crevice and gap between the pipe and valve o pipe and fitting (socket end).
E. h e metallurgical microstructure of the valves and fittings is a mix of three phases alpha, beta and gama2 It to the aluminum in the beta and gama2 phases that is preferentially corrodei. he microstructures examined to date have a complete network of beta or gama2 so that there is a path through the castf ug for dealloying to cause leaks.
F. Forced air cocling as specified by the supplier's procedure does not prevent dealloying.
C. %e chemistry of one valve body was check anclyzed and was in conformance with the smcification, ASME SB-148 CA 954.
H. h e corrosion products are as expected. Rey contain aluminum, iron, calcium, and copper. here is some evidence of Callionella (an iron 1071m 1.1
-- n o. =
AT TACHf AENT.:2 ST HL AlA~?llR PAGE ? 0F fit bacterium) in the corrosion products. Conversely, Nalco (June 9, 1988 report) found no iron bacteria and little microbiological activity. As of the date of this report, we have found no evidence that Gallionella or other bacteria have any ef fect on Al-bronze.
h I. Metallurgical plastic replicas were taken from a laboratory specimen and directly from an elbow. A comparison of the photomicrographs taken at 480X of the lab specimen and the two replicas shows them to be essentially the same. Field replicas can be used to nondestructively determine the microstructure of Al-bronze castings.
J. ne pH of the residual water inside valve EW-269 (serial No.61-382) was 6.0.
K. Et fitt!ngs and valves are experiencing dealloying corrosion because:
- 1. The microstructure contains the beta and gansna2 phases which are susceptible to dealloying.
h
- 2. The pH in the crevices due to the electrochemistry of metal corrosion in the crevice is acidic, which promotes dealloying.
L. D e biocides and other water treatment chemicals do not appear to have promoted the dealloying.
k 1.3 Materials A. We valve bodies are specified to be cast aluminuar-bronze to specification SB-148 CA 954 (10 - 11.5 percent aluminum).
B. We fittings are specified to be cast aluminuar-bronze to i specification SB 148 CA 952 (8.5 - 9.5 percent aluminum).
C. ne pipe material conforms to SB 169 C61400 (6-8 percent aluminum).
1.4 Evaluation Methods A. h e as-received pieces were photographed and are shown in Figure 1.1.
B. In order to properly interpret tha microstructure of Al-bronze, two actions were taken. De supplier of the raw castings was asked to send us photomicrographs of the typical microstructures of Al-b ronze . Figure 1.2a is the information from them. Second, three small samples from Valve EW-269 (serial No.61-382) were heated above j 1100*F for one hour. One sample was cooled in water, one in air and one was lef t in the furnace and slow cooled by turning of f the power. Figure 1.2b shows the microstructure of the three specimens.
C. Cross sections were taken through the valve that leaked, the fitting (elbow) that ?.eaked and the valve that did not leak. Table 1.1 compiles all the information obtained from this work and references the applicable figures that show the microstructures and macrostructures.
1071m 1.2
ATTACHMENT 3 ST HL AE 3712 PAGE ~/ OF W D. The chemistry of Valve EW 269 was determined. R e results are listed in Table 1.2.
E. h e residue in the socket weld crevice in the elbow was analyzed by EDXA for elements and by X-ray dif fraction for compounds. The residue was also analyzed for bacteria that a.e known to be active in microbiological 1y influenced corrosion. We results are in Table 1.2.
F. When Valve EW 269 was unpacked, it was found to have been sealed up by plastic bags and tape. There was still a slight amount of moisture trapped in the valve. he pH was measured with litmus test paper and was pH 6.0.
C. It is cocenon practice to use a technique called replication at a field location, when a metallurgist wants to know the microstructure of a piece of equipment but can' t cut out a sample of material. his e.ethod is well understood for carbon and stainless steels. In case it was needed in the field, a test was made to bS sure that the techniques also work for Al-bronze. We standard field replication techniques for grinding, polishing and etching work quite well. ,
Replicas were taken from a laboratory specimen and from an elbow l polished and etched as if it was being done in the field. The photomicrographs of the two replicas and of the actual piece of material are almost identical (see Figure 1.6).
H. The chemistry as well as the pH of a saturated solution of welding flux were taken to determine if the flux might have been the source of the low pH.
M Discussion of Results A. The study to determine the visual appearance of the three phases (alpha, beta and gamma 2) was completed in a satisfactory manner. A comparison of sample photographs from the supplier (Figure 1.2a) and I the photomicrographs prepared by Bechtel were used to identify the three phases.
B. The microstructure of a fitting that was tempered and forced air cooled by Bechtel was identical to the as-received material. Thi s confirms that forced air cooling does not prevent dealloying, which k
is the corrosion mechanism that is causing the leaks.
C. We chemistry of the base metal is as expected, as was the analysis of the corrosion products. The influence of HIC on this problem is unknown. Fragments (not active) of Gallionella were found in uoist residue in the crevice of the elbow.
D. he pH of the sample of residual water found in Valve EW 269 was slightly low. The site records show a pH of 8 to 9 for the system.
However, it was reported that a pH of 4 was measured in the water weeping from a leaking valve. This low pH is most likely a result of the electrochemical effect of the crevice.
1071m 1.3
ATTACHMENT '
ST HL AE b7 f PAGE 9 OF 9 l F. . EDXA analysis of the welding flux reveals significant amounts of fluorine and sodium and lesser amounts of aluminum and silicon. Wet l chemical analysis revealed 1.01 percent boron. The pH of a saturated solution both of as-received and of high temperature baked flux dissolved in distilled water (pH 6) was pH 7.2, i.e., the flux is basic.
i 1071m 1.4
TMLE 1.1 MW.YSIS OF 00sS SELTINS OF 1xmu. smus SEMAI.10. VM.VE OR TYPE OF WultSr CA!E ILCAI. IEiAUDCICAL ILCATION FtD11 M 010CRARI MOREfr m0m.r mum, aoss ammT10N OF smucIm E OF ooss m.
ID ED. IEAT NO., SECTIN DEA 11NING !ECTEDN 61-382 w1ve dramafererre 27% Algha Fig. 1.3 Omtimous netserk ')D40 100 to 200 and beta of beta degrees 0.12 in. 507/50E dealloying md W 269 115174-44 leakirg ends minor pits {
61-382 w1ve Iagitudinal 100E Algha and Fig. 1.3 Ihalloying is V' 7041
@ 180 degrees plus beta shaped. Contim- f 0.025 in. on ous .etwork of
," the OD a:s! beta in valve.
Sa m 269 H5174-44 ID 50/50t leaking end Pipe and weld no beta I
61-382 w1ve Ia gitudinal Itme Algha and Fig. 1.3 No dealloyin . 7042 0180 degrees beta Continuous thru brunee md small grey netierk of valve and dots in both beta carbon steel ghanes pipe (drain 50/50 It> leak W 269 115174-44 comection) MD MNY.o Fig. 1.4 Ihalloyirg Q>r,n g:
r 466-185 valve In uitudinal SE Alpu and 7043 0180 degrees Some uc e beta. Omtimous QA7' Bronze valve to 2SE Samil grey dots beta g to bronce in both pianes tEtwork <
Q pipe SE/53 m 315 115174-36 No leak l
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RtOE"r FITTIE, ID C3, IEAT NO. SECTIM N N Obow 1mgitudinst 10m Alp.cand Fig. 1,5 Beta rietwork 7044 g 180 degrees at socket beta easi-contiru n=-
Bronr.e elbow wld gap. Ihalloyi:ig at to Inunse gap PIPe 16 528-31 92/10% fa'irig end i
Blow dmanfermee 6M kgha and Ihalloying on 7.A5 100 to 200 beta ID on the ert. ire degrees specimen.
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Elbow Iagitudinal 4E Algha arul Deallah on tie 7066 90 degmes aumy beta ID amud t:r gap frtas leak and and den tev:mi back frne gap weld 8 5 alplu/
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I FICL'RE 1.6 Comparison of replica micrograph with laboratory micrographs Valve EW 269
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TABLE 1.2 CHEMICAL ANALYSES
- 1. Base metal of Valve EW-269 by emission spectrographic analysis.
Element Cu Al Fe Mn Ni
% Weight 84.9 10.6 4.25 0.11 0.06
- 2. Corrosion product analysis by energy dispersive X-ray analysis (EDYA)
Element _A1 Fe S Ca C1 g g
% Screen Height 100 100 20 20 15 10 8
- 3. Corrosion product compound analysis by x-ray diffraction Iron oxide hydrate and calcium carbonate were the only compounds with strong patterns. Since the EDKA results show a strong peak for aluminum e A the only aluminum compound that has a very weak crystal pattern is O1 A1(OH)3 This compound is most commonly the result of the corrosion of aluminum.
- 4. Bacterial analysis of the corrosion product revealed evidence of callionella, which is one of the bacteria known to influence the corrosion of steel and stainless steel.
i 1.7 1071m
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ST HL AEO79
! PAGE /9 0F S 2.0 STRUCTURAL INTEGRITY l
l 2.1 Introduction During the period prior to completion of replacement, STP asked Bechtel to do an analysis of the system in the current and future corroded condition on the basis of statistical probability of the condition of the system based on analysis of samples of leaking and non-leaking joints. n e statisticci analysis was done by Bechtel Western Power Company's mechanical staff.
Bechtel was also asked to analyse the system from an ASME Code standpoint and from a fracture mechanics standpoint. At STP's request, Aptech Engineering Services was retained to do the fracture mechanics analysis. As part of the structural integrity analysis, Bechtel had proof tests conducted to determine the margins.
2.2 Statistical Analysis in order to obtain data for statistical analysis, 65 joints were sectioned circumferential1y and longitudinally. Longitudinal cross sections were used to confirm the observation that the worst dealloying areas were the socket veld crevices. Additional circumferential cross sections were nsd to determine the maximum amount of dealloying. h is was accomplished by making circumferential cross sections of the crevice area and polishing until the A great +ist amount of dealloying was found. A sumary of the data obtained is Ol contained in Table 2.1.
D e data in Table 2.1 were used to perform a statistical analysis. The procedure used and results obtained are described in the following sections.
2.2.1 Classification of Samples The circumferential cross section for each end of the sampled fitting and valves was classified from observations as a leaker or a non-leaker. nus ,
the extent of dealloying for the leaker ends sampled represents the condition
! of any valve or fitting socket that was currently found leaking on the ECW l system. ne extent of dealloying for the non-leaking ends sampled represents the extent of dealloying of the current ends af ter the leakers were removed.
l The results of the samples, presented in Table 2.1, are also shown in Figure 2. 5.
2.2.2 Confidence Level It is necessary to insure that the sampled fittings adequately represent j the whole population of fittings. herefore, the degree of confidence has been estimated.
I 1073m 2.1 L
.c ____
AT TACHMENT 4 ST41L AE 27 yd PAGE3 csOF 69 From the sample, a probability p of finding a certain fraction deelloyed in a socket end is determined. This probability is then used to estimate the probability P of finding a certain f raction dealloyed in the population. The standard error of proportion is determined using the following relationt p(1 p) N-n (equatien 2-1)
,p .
n N-1
) 3 wheret N is the population size, n is the sample size, and p is the probability of finding a certain fraction dealloyed in the gf ( ;
sample. ,
The error of estimation E for a 95% confidence level is found by using: t i
i E = 1.9 6
- s p (equation 2-2)
I Therefore, the probability P of finding a certain fraction dealloyed in the population of fittings is given by:
P=p1E (equation 2-3) e f
From equation 2-1 it is seen that as either the sample size approaches I the population size or rhe probability approaches sero, or one, the error band l will approach zero. To assure a 95% confidence level for a 95% probability we i vill solve equation 2-1 such that p + E = 0.05. The proportion of non-leaking socket ende exceeding 55% dealloying is estimated to f all between 0% and 5% t (1.6% 13.4%) while the proportion of leaking socket ends exceeding 92%
dealloying is estimated to fall between 0% and 5% (1.2% + 3.8%).
2.2.3 Sample Distribution The data presented in Table 2.1 were used to generate Figure 2.5. In this figure, the percent of exceeding a certain fraction of the cross section being dealloyed is plotted for all samples taken. The leskers and non-leakers are plotted separately. As shown on this figure, the leaking ends have a higher percent of the cross section dealloyed than the non-leakers. Thue , it ,
can be deduced that replacement of the leaking socket ends will remove the number of fittings with high dealloying and the remaining population will be characterised by the non-leaking data.
2.2.4 Leakage as a Function of Dealloyed Cross Section As shown above, the leaking socket ends have correspondingly more of the cross section dealloyed than the non-leaking socket ends. Therefore, an attempt to determine the fraction of leaking socket ends as a function of 1073m 2.2
_m_..
lA PAGE31 0F dealloyed cross-sectional area is made. For this purpose, the probability distribution function of socket ends as a function of dealloyed cross section area needs to be determined.
Some assumptions must be made concerning the possible shape of this probability distribution function. Originally, the socket ends had little or no dealloying, thus the distribution function was narrow near reto. As time went on and corrosion set in, more and more socket ends were being dealloyed.
his process can take the shape of an expor.ential or Weibull distribution function. As more and more fittings are dealloyed, tb3 median shif ts toward s higher values and the shape changes towards a normal or loF-normal distribution. From the data obtained from the sample, it con be seen that the non-leaking socket ends can be characterized by a Weibull distribution, while the leaking socket ends approximate a normal distribution. I n e above distribution functions were then used to determine the percentage of leaking socket ends that can be observed as a function of cross section dealloying as shown in Figure 2.5.
2.3 Stress Evaluation To establish the margins existing in the original design a review of the stress analysis of the small bore ECW system was performed.
In addition to this review further calculations were performed to determine the margins that exist when the stresses are compared with the minimum properties of the dealloyed material as determined by testing.
2.3.1 Objectives The specific objectives of this task were as follows:
- 1) to review the original design calculations and establish the original design margins.
r
- 2) to demonstrate existing margins for dealloyed fittings using ASME ,
Section III methods and the minimum properties of dealloyed '
materials as e nablished by tests.
2.3.2 Analysis Assumptions n e original design bases for the small bore Al-bronze piping are established through ASME Section III Subsections NC 3600. In this particular analysis method stress multipliers or stress intensification factors (SIF) are used to account for geometric discontinuities. All st resses in the original design analysis are referenced to the section properties of the piping material. For purposes of the stress review, the results of the original analysis for equations 9 and 11 of ASME Section III were reviewed and compared to allowable design stresses. [
For equation 9B, the primary stress equation due to pressure, weight, and Operating Basis Earthquake (OBE), the allowable stress is 21.6 kai and for equation 11, the ef fect of pressure, weight, other sustained load and thermal i expansion, the allowable stress is 45.0 ksi.
i 1073m 2.3
~ - 1 p y yAcM SnHL AE- _PAGE M OF G To evaluate the margins that exist between design loads and actual eviterial properties, tensile tests (see Table 2.5) were performed on material that was up to 100 percent dealloyed. h e average ultimate tensile strength for the 100 percent dealloyed material was 30 kei. his may be compared with b 75 kai before dealloying. In the evaluation of actual stresses that exist in these fittings a value for SIF has been spplied to the crevice location. h e values for SIF f have been established from the Reference 2 and Refetence 3. In all cases the larger value (conservative case) from either reference was used. Finally for the stress evaluation, ASME III Appendix III was used to establish a value for Sh, a value of the quarter of the ultimate tensile strength of 100 percent dealloyed material (7.5 ksi) was used. h e ASME Section III (subsection NC-3650) equations with stress multipliers to account for geometry discontinuity due to dealloyed material are b ! outlined below: PD "A * "B 1.2S (95) EQ9: C p + 0.75 (Cg) 4 h P 2.4S (9D) 4t f Z f h EQ11: C p
+ 0.75 (Cg) + Cg < Sh*8A f f f Pd : Design Pressure = 120 psi Pp : Peak Pressure = 40 psi tg: Minimum Fitting Wall Thickness Per ANSI B16.11 Zg: Minimum Section Modulus of Fitting = nR 2 tg D: Outside Diameter of Fitting C : Stress Multiplier for Pressure Stress (= 2.74) Per Reference 2 Ck:StressMultiplierforPressureHomentStress(=3.42)UsingCreater Value From Reference 2 or Reference 3 SA tallowable stress range for expansion stress Sh :1/4 (30.0 kei) = 7.5 kai per ASME Section III Appendix III MA : Resultant Moment Due to Weight Mg: Resultant Moment Due to Earthquake and Occasional Loads MC P.ange of Resultant Moments Due to hermal Expansion 2.3.3 Analysis and Results h e original ASME Section III analysis was reviewed and stress values compared to allowable values for the original material. he stress ratios 2 presented in the Figure 2.7 illustrate the fraction of the total allowable stress that exists in each fitting under the design conditions of sustained plus secondary loads.
It can be seen from these data that there are significant margins existing in the original design. he highest stress in the ECW small bore 1073m 2.4
I F A77AC11M ITO - AE-srap*.13 N 7gh' piping is about 77 percent of the code allowable stress. his margin was further enhanced by replacement of components with less than 70 percent margin 2 in the design wall thickness. An alternative design criterion was established with which to compare the calculation results of equations 9 and 11. In this case the allowable stress was based on the average ultimate tensile strength of 100 percent dealloyed material. nis value is 30 kai. Consistent with ASME Section III Appendix III, a value for Sg was established as one quarter of this value or 7.5 kei. L e stress allowables for equation 9 using ASME III Sections NC 3600 are 1.2 Sh or 9 kai and for equation 11 the value is 18.75 kai. Evaluating the system with this design basis gives the results shown in Figure 2.4. In this figure the calculated stress values have been compared to the average ultimate tensile strength as defined by tests. Again it may be seen from Figure 2.4 that in all cases significant margins exist between the sustained plus secondary loads and the available material tensile prorrties. 2.3.4 Conclusions Based on the margins that exist in the original design condition, the reevaluation of stresses using fitting section properties and tensile properties from dealloyed matet .1 tests, all fittings are acceptable from the stress point of view. his is the case even when the worst case Stress Intensification Factors (SIF) values for fittings are used from the published literature. Bis result stems from the initial design margins illustrated in Table 2.4, the ef fect of the increased section properties of the fitting (as compared to the pipe properties that were used in the original design analysis) I and the low operating pressures of the ECW system. 2.4 Fracture Mechanics Integrity Analysis 2.4.1 Introduction The ASME Code design rules for piping provide margins against failure for loading conditions encountered during normal service as well as postulated conditions such as seismic loads and abnormal events. In assessing the structural integrity of partially dealloyed aluminum bronze piping components, a conservative evaluation has been performed to assure that adequate margins still remain. Wie was accomplished by evaluating the condition where the dealloyed region is assumed to have lost its load carrying capacity and will behave like a crack-like flaw. Unde r these conditions , flaw evaluation procedures similar to Section XI of the ASME Code have been applied. 2.4.2 Linear Elastic Fracture Mechanics Linear elastic fracture mechanics techniques (LEFM) were used to establish the load carrying capacity of partially dealloyed fittings when the dealloyed region is treated as a crack-like flaw. When LEFM principles are applied, the flaw tolerance of the component can be quantified in terms of applied stress, flaw sise and shape, and the fracture toughness of the material. By defining any two of these parameters, the third parameter can be quantified by fracture mechanics relationships. Selection of appropriate values for fracture toughness, and a conservative representation of the site and extent of dealloying as a flaw, will give a determination of the of tuctural g 1073m 2.5
- - - - __-.w------
ST.HL AE. PAGE39 4/A , capacity for a partially dealloyed fitting. In terms of LETM, crack instability (or propagation) is predicted if the following condition is satisfiedt i Kg ) Kge - where Kg = applied stress intensity factor (a function of stress and l flaw and part geometry) j and Kre = critical stress intensity factor (or fracture toughness) Kge is a material property which can be determined through testing, l like a yield strength in a tensi1.e test. In this case, K ; toughness data (such as Otarpy V-notch impact properties)re for data or any aluminum other bronze were not readily available. B erefore, crack tip opening displacement (CT M tests were run on two valve bodies. We valve body material (CA954 alloy) was ! chosen because it was of sufficient size (i.e., the fittings were too small) to ; obtain reasonable size CTOD specimens. Furthermore, it represents the material with the lowest toughness of the cast aluminum bronze used at STP. We CTOD Rese are equivalent to Krc values of 3 . values 63.$ and were 95.13.2 mils 27nd ksi-int 9.5the using mils. conservative relationship Ete2 2 y
= c
( E (1 v) [ wheret E = Young's modulus A Cy = the yield strength /_1\ 6 e = critical crack tip opening displacement i v = Poisson's ratio See the Appendix for the CTOD test details. (l ne 952 alloy used for fittings is expected to have better toughness on A ! average than the 954 alloy because it should have less beta phase because of Q3 the lower specified aluminum content. - For purposes of illustration, the toughness of cast aluminum bronse was taken - as 65 kei-int /2 discussions withbasedontherelationshipbetweenCTODandKg$andon several sources. We test data (Reference 4 for aluminum , bronze pipe and welds in 10- and 30-inch diameter cooling lines have excellent 1 i toughness, with Kge values estimated to be in the 150 - 200 kai-inl/2 range. j Kg can be expressed, in simplified terins, as follows: Kg = OF (n a )1/2 ( where: o = nominal stress ' [' a = crack size F = functional relationship that accounts for flaw shape,
- body geometry and type of loading.
I I 1073m 2.6
hlh. 3 PAGEOS OF 59 i h e relationship between critical bending stress and dealloyed area was determined for the case of two limiting conditions: a complete 360 degree circumferential dealloyed region originating from the inside surf ace, and a through-vall dealloyed region extending part way around the circumference. In directly applying LEFM to these conditions, the dealloyed regions were conservatively modelled as flaws, debonded and without any load carrying capacity. n e methods contained in Reference 5 and 6 were used to establish the stress intensity factors for these flaw geometries. We capacity of the remaining uncracked section for the two limiting cases is shown in Figure 2.1 as a function of the extent of dealloying (cracked cross-section) and two racture toughness. Even for the lowest value of appropriate Kge used (i.e. 65 ksi-inl/ estimates of (2), significant load-carrying capacity remains for through-vall flaws exceeding 50% of the circumference or 4.3 inches in surface flaw length in 2-inch pipe. h For the summary stresses given in Table 2.3, the computed through-wall flaw lengths that can be tolerated in the system are as follows: Nominal Pipe Site S oc ke t Tolerable Site I (inc h) Diam. Circumference Pe rc en t Degrees , 1 1.75 5.50 65 234 2 2.75 8.64 59 212 It is reasonable to assume, given the conservative nature of the 1.ETM evaluation assumptions, that non-ductile failure of partially dealloyed material is not an issue of concern. It is therefore concluded that adequate toughness exists to prevent fracture of ECW system fittings and plastic ( collapse is the govtrn#.ag f ailure mode. l f 2.4.3 Limit Load Analysis l Because sufficient fracture resistance exists in partially dealloyed fittings to allow for utilisation of the inherent tensile strength margins in ; t the ECW system design, the structural integrity for a net-section plastic collapse failure mode was evaluated. We two previous limiting cases of , part-through and through-vall partially dealloying were again analysed by . assuming no load carrying capacity of fully-dealloyed regions. W e bending I stress for net-section plastic collapse was established from Reference 7. Reference 7 provides the technical basis for the flaw evaluation procedures of ; Paragraph IWB-3640 and Appendix C for austenitic piping given in Section XI of the ASME Code. For a material flow stress defined as the average of the specified l minimum yield strength an>* ltimate tensile strength, the tolerable levels of i part-through dealloying unt.ormly distributed around the circumference are ; shown in Figure 2.2. Ample bending capacity is observed up to dealloy depths in excess of 80 percent even when the residual strength of 100 percent dealloyed material has been neglected. When the strength of dealloyed material is used: the bending load capacity of the fitting is computed to be appr.eximately five times the worst case design loads. his is shown by the horizontal line in Figure 2.2. 1073m 2.7
ATTACHMENT 2
,, . y -- ~
he through-wall dealloyed model, as shown in Figure 2.3, gives a lower prediction of bending stresses to cause plastic collapse than the 360 degree part-through model; however, significant margins are still calculated for through-wall dealloying extending 180 degrees around the circumference. Again, because the strength of the fully dealloyed material is significant, the available bending load capacity will be greater than these conservative ' calculations and will not fall below the case of the fully dealloyed condition (Sce again Figure 2.3). 2.4.4 Summary Based on LEFM analysis under the conservative assumption that dealloyed regions are debonded and will behave as cracks, very large amounts of i dealloying can be tolerated in small bore fittings without a concern for non-ductile failure. Similarly, on a plastic collapse basis there is adequate strength even in the fully dealloyed condition, as described in Figure 2.2, to support the intended design loads with sufficient safety margins. It is reasonable to conclude that, based on the above analyses, the components will have significant integrity and margin against failure for the design loading conditions. 2.5 Proof Tests A leaking coupling, a leaking tee and a non-leaking elbow were selected for tests to failure. Steel plates were welded onto the open ends and then i they were subjected to increasing water pressure. h e two leakers failed by ' leaking f aster than the pump (200 cubic em/ minute) could keep up with it rather ! than by breaking. h e non-leaker failed by the end plate blowing off. The , pressure to cause failure was as follows: ! l Percent of Cross Section ' Dealloyed at Area of Item condition Maximum Attack Area Failure Pressure (percent) (psi) Coupling Leake r 38 8950 Tee Leaker 52 5900 , Elbow Non-Le ake r 0 6500 l The f ailure of the tee was at the veld we made to seal the casting. The welder noticed the casting material was difficalt to weld (presumably because l l it was dealloyed at that point) and indicated he felt it would fail at that closure. Therefore, we feel the coupling failure where the closure welds were made on wrought pipe is most representative of the pressure carrying capacity of the leaking dealloyed Al-bronze. For a pressure of 8950 psi in the 3,03 diameter coupling, the stress in the 0.255 inch wall to cause failure is 49.54 kai. Bis is three times the design allowable stress of 15.7 kei. A conservative reference burst pressure was calculated by the following equation from Reference 8: P3=c u Feyt in W 1073m 2.8 i
AT I AUNMLNT ST HL AE 77 l PAGE 37 OF . __ wheret O u' = nominal tensile strength Feyt = correction fact':. for cylinders (0.250/eu + 0.2 27) I W = vall ratio (outside diameter /inwide diameter) eu = true strain at maximum load i n e calculated burst pressure was 6800 psi. nerefore, the "proof" strength of the coupling exceeded the reference burst strength of a sound casting. I 2.6 Dealloying Rate Studies , he dealloying rates reported by Upton in Reference 9 are as follows: t Twelve mont.h reate ' He at Treatment Maximum Corrosion Rate (rnpy) Water quenched 0.02 general - 0.03 in crevices As-received bar stock 0.03 general l 0.15 in crevices l l Furnace cooled 2.0 1 ) Ferrara and Caton report in Reference 10 a maximum dealloying rate of l 30 mpy af ter one year and 22.5 mpy af ter two years. Both Upton's and l'errara's i results are lower than the 44 mpy average and 83 mpy maximurs estimated at STP. We maximum rates are based on the assumption that dealloying occurred in 3 years to 100 percent of the wall and the average rate assumes 50 percent. ne S7P socket data are plotted in Figi 'e 2.6 based on the cocunonly used expression for corrosion ratet V=Ki t*K2 where V = corrosion rate in spy K2 " I/2 Xg = experimentally determined - 1.36 x 10 4 for STP socketa t = time in hours The data from Ferrara and Caton and from Upton were used to validate the form of the equation. 1073m 2.9
m, ry g MgNT T LGE D 0
".EFERENCES-
- 1) ASM loiter and Pressure Vessel Code Section III 1974 with Winter. l 1975 Addenda f
- 2) "Stress indices for girth welded joints, including radial weld !
shrinkage, mismatch and tapered w't.1 transitions." NUREG/CR-371 l (ORNL/SUR-2913) by E. C. Rodabaute and S. E. Moore. [ [
- 3) "Stress indices for ANSI Standard 516.11 socket welding fLttings." i ORN!.-TM-4929 by E. C. Rodabaugh and S. E. Moore.
- 4) "Structural Integrity Analysis for Essential Cooling Water Lines."
July 1983 by G. R. Egan and S. R. Paterson. , l 5) "Elastic-Flastic Fracture Analysis of through-wall and Surface Flaws l in cylinder" EPRI Report NP-5596 January 1988 by V. Runar and , M. D . Ge rman. l
- 6) "Evaluation of Flaws in Carbon Steel Piping," EPRI RP 1757-65. Draft !
Final Report. March 1988. [ l t l
- 7) "Evaluation of Flaws in Austenitic Steel Piping." EPRI Report l
EP-4690-SR (April 1986) S) C. P. Royer and S. T. Rolfe, Ef fect of Strain-Hardening Expontat and } Strain Concentrations on the Rurating Behavior of Pressure Vessels. l ASME Faper 74-MAT-1 (1974). [ i
- 9) 3. Upton Corrosion (June 1963) 204t. 1 I
- 10) R.J. Ferrara and T.F. Caton Materials Performance (February 1982) 30. l l
l b I L [ t I i I [ l 1073e 2.10 l i
[jf/ h W 4,T
.. ~
y a bll ~ ., i- S HLA 74A Table 2.1 PAGE 39 0F M s , Statistical Sampling of ECW Systeia Fittings
- Percent of Cross-Section Area 1.ost to Dealloying laskers
~
No n-1.e ake r s Valves EW102 A 43 5 47 EW103 A 24 ,
- B o EW104 A 54
'8 10 EW105 A 30
-i s 8 to EW335 A 14 8 37 EW337 A 59 8 21 EW338 A 10
*B
- 22 EW369 A 24 8 3 EW116 A 9 B 6 EW126 A 52 3 82 EW415 A 8 B 29 EW323 A 54 B 13 EW235 A 46 B 12 EWO 31 A 35 B 37 EW332 A 5 5 17 EW214 A 88 8 20 EW115 A 59 3 38 E%017 A 61 5 18 EW114 A 0 B 0 A 1 EW3 51
(.. , 8 11 EW215 A 68 5 3 A8 A 0 B 0 DA5 A 80 8 19 e 1073m 2.11
7 7XCHM d . ST ML AE- 'If
, PAGE 30 Sr Table 2.1 (continued)
Statistical Sampling of ECW System Socket Ends P6rcent of Cros a-Section Area Lost to Dealloying-Leakers ~ Non-Leakers Valves (continued) 102 A 19 , B 36 MA13 A 57 8 44
- ) . A2 A 62 B 64 LO582 A 29 5 28 ;
Tees
't337 A 0 5 0 ,
C 0 T338 A 47 B 37 i C 32 Berst A 52
*ie s t B 38 C ,
48 , Coupling s ' C-6 A 38 MA12 A 50 ; i I Total Sampled 24 41 dean Percent Dealloyed 48.79 19.39 Standard Deviation 21.02 17.39 , i I i r i i i l i 1073m 2.12
- ^ ' - -
l i FATTACHMEN j' ST HL AE T798 ;- PAGETl OF 69 i: Tame 2.2 -l l Design and Operational Data for the Aluminum i Bronze Cooling Water Piping i' I (1) Materials: Pipe CA61400 (SB 169 or 150) , Couplings CA95200 (SB 148 or 271) Valve Bodies CA95400 (SB 148 or 271) (2) Pipe size: 2-inch NPS and smaller, Schedule 40 (3) Temperatures: 150'F maximum design; 160'F operating (4) Pressure: 120 psi maximum design; 41 psi operating (5) Pluid: Brackish Water (6) Design Code: ASME Section III, Class 3 k i l, f 1073m 2.13
TABLE 2.3 i
SUMMARY
OF FITTING DIMENSIONS AND HIGHEST STRESS LEVEIS NOMINAL DESIGN OPERATING DIMENSIONS FITTING SIZE PRESSURE PRESSURE Do t FITTING STRESSES * (ksi) (PSIG) (IW) (IN) Pm Pb Pe (IN) (PSIG) . l 40 1.223 0.184 0.20 0.72 -- l 1/2 120 _ . .A 120 40 1.778 0.224 0.24 4.27 1.16 1 5 40 2.415 0.250 0.29 1.96 2.11 1-1/2 120 2.952 0.273 0.32 4.26 1.81 2 120 40 l { 1.~
- NOTE: Fitting stresses listed are the highest (worst case) stresses calculated without distinction among service level (i.e., normal, upset, emergency or faulted). The l
I stresses are summarized in categories of primary membrano (Pm), primary bending (Pb), and thermal expansion (Pe) with SIFs removed (i.e. unconcentrated) to be I consistent with evaluation definitions of IWB-3640 and proposed IWB-3650. l ns umm
~
&w 2i$
4' , A7 l l - 2 l l
TABLE 2.4 STRESS EVALUATION TOTAL NUMBER TOTAL NUMBER _OF FITTINGS STRESS RATIO: FIT TING OF MEET MAX' STRESS SIZES FliTINGS ALLOWABLES *
- 30 KSI
, EO.9 .074 EO.11 .039 g., g, EO.9 .289 56 EO.11 .19 2 f* 56 m 60 EO.9 .181
~ EO.11 .283 G M/2** M 15 4 EO.9 .326 EO.11 .276 2" 15 4 TOTM. 283 - 283 NOTES: . I $g
- 1. " ONE OF THESE FITTINGS IS %" . 4. EOUATION 11 ALLOWABLE IS (S A - SH) 0.25Sn)
IM o n i ' SA - (1.25S c
- 2. ""Su -% (30 as ) 5. TOTAL NUMBER OF FITTINGS ABOVC b
REPRESENTS STRESS DATA PONTS EXCLUDNC b
- 3. EQUATION 9 ALLOWABLE IS 1.2 Sn THOSE FITTINGS AND VALVES REPLACED AND/OR DELETED l
NROS29.DGN
@$h hi
. b_ - - - _
~ m .
H F i H i E L S - K 0 n o i a \
\
g
\ l 9
g d e i W M N i e s a A - L s 0 R
%x'N B i s
g 1 F a g
' 8 O t
n i o i i t e 0/ t 3 i 0 i F I 7 D i E h Y c i i S n i H u .
- s,i 0
6 A E 2 I S D K n d e y o i 5 6 s\ *~~ i i 0 sa l 1 - 5 l a c w D e i hs K i A N r o sK l 0 O I f I 4 T y C t 0 E i c 0 i S a 1 S . p W 0 O R xc a i 3 C I A. C s s R i F e ' . O _ t r U A 0 S i T l e 2 N g n 2 =HW- E C i d n
' gW G U
i E R B e O I 0 P l G- M 1 1 c T - 2 ' KM i e r u g _- 0 i F 0 0 0 y 0 oi o 7 m 5 4 2 n SM g =$g' Q o
~.5 i l 1
- !i! i ' !
4 i Figure 2.2 Conditions for Net-Section Plastic Collapse for 360 Degree Part-Through Dealloying i 70 g a g i g i g i g i g a g i g i i g i
~
~
360" PART-THt00GH WALL DEALLOYNG SO - x
- pm - 1.u4 KS '~ ~
i g m-100 X DEM.LOY { 50 - (Ug - 30 KS) l - CJ-0 % DEALLOY ! o _ (7f - 45 KSD
, ~q Neglects The Strength j
$ - of Dealloyed Region s
s
' Based Upon 30 -
s 100% Dealloyed N Strength Q s _
's m
20 - 's s - s _
\ '
j$>.; g, _ N' s RM
$RE
' ' ' l ' l ' l ' l ' l ' l %5 J
0 3 20 30 40 50 60 70 80 90 100 i 10 . i PERCENT DEX LOY. */t x 100 h) ) 1
Figure 2.3 Conditions for Net-Section Plastic Collapse for Through-Wall Dealloying i ) 70 , , , , , , , , , , l60 "O
- TMt00GH WALL DEALLOYNG Pm - 104 KSI 50 -
-100 X DEALLOY -
(5r - 30 KSn 4 l
-0 X DEALLOY l m q _
s (Uf - 45 KSI) g I
- A Based Upon -
100% Dealloyed 30 - Neglects The ' Strength Q Strength of s Dealloyed Region N - 20 - 's ' i s mm s
~. Ev>m S s M
~
s _ %E o i l i i i i i i i l i I 'i ~l- i I i li DE sh 0 10 20 30 40 50 60 70 80 90 100 PERCENT DEALLOY, 8/TT x 100 (X)
l FIGURE 2.4 l l I SOUTH TEXAS PROJECT - UNIT 1 i ALUMINUM BRONZE FITTING STRESS RATIO 1 i 1.0 - - - - - -- - am l Mi! EQ. 9 STRE33 tJuiT = 30.0 M3s EQ. 11 STRE3S IJuli - 30.0 K3s "
.S * "
l .7 -~ ! l
.e- -
l_
- . \
g en .3 -*
.4--
.3 --
") l N
O m .30 1.0 FITTING SIZE 1.3 2.0
1 j ,' miEgM sD = 5 0 -- 90 1 J 0 5 - 9 9 3 _ 5 8 809
)
0 5 1 8 8 5 0 1 7 8 3 0 7 5 7 E D
) Y i!
5 0 O S 6 7 L R . 3 L E - 5 S 1 0 5 EA K) A S 1 6 6 D 1 _ 2 WG ,) E4 L( I C N E I 5 0 N - P T ii! 5 6 O N - U I O G E T ,) T N I Z F 0 C F N ! 5 5 E : O ED , 3 5 S g - R L 5 - 8
/ P MM I111 ,3 4 0S 5 S O
A 0 5 R UNI S 8 4 4 C S MF , 3 R : U. O 5 0OF E K )4 Iii11 3 A( 2 _ MN ,3 4 T E 1 O L T I 1 1 30 5 N E 2 TI U 3 C NB 3 R g - U TR I I11 5 1 2 0E 3 P P T S 3 0 5 S D I 11 ,' 2 2 ll11 50 1 2 . ll1I 0 151 d 1 Iii 1l l i 5 z
'5-
;IIi ggIlllI 0
1 1 01 9 " 7 6 5 ^ N Uua' "
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~ , ' il) i:ll !!
e 9
-m.e e o ATTACHN1Lt
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- Cf
_a
-S
=M M
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~n
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=P m
O I l l l l a $ R H a (SD) IN31X3 CMA0Th00 2.21
i i FIGURE 2.7 SOUTH TEXAS PROJECT - UNIT 1 ALUMINUM BRONZE FITTING STRESS RATIO TO ALLOWABLE (As Designed) i 3.0
,9 -- EcuArio.4 90. ALLOW. - 21 ti K 5s
, EcuATION I1. ALLOW. - 43 nSa l .5-- I
.7--
I h I i
" .6--
g l U *3 ' x
, y .3 --
l ~ l .4 og9 I i
.s -- Ed m$g l 555
.2 - -
ri i .1 -- O- --
.50 1.0 1.5 2.0 L'
- PIPE SIZE mess .3 cucumn o u-.o e ..c .. ci.n . .o. .. ..s
__ . _ . _ _ _ . .- . ._ . . . , _ ~ -
g , l r)y ' 1 FTf"1 g .
. u
,PAGE F57 TABLE 2.5 Tensile Test Results on Dealloyed Samples of CA 954 Material from Fittings Sample Percent Dealloyed* Tensile (psi) 1 100 31,700 2 85 31,000 3 34 56,500 4 11 76,800 5 100 28,900
- Based on SCM of tensile fracture surf ace 1073m 2.23 1
' ATTACHMENTS ,
IGE 9 APPENDIX Lab. No. 588.043 June 23, 1988 CRACK-TIP OPENING OISPLACEMENT TESTING OF VALVE-BODY MATERIAL 9 Anamet Laboratories, Inc.
LABORATORY CERTIFICATE I}hA PAGE43 FSF 74B Anamet Laboratories, Inc. 3400 INVESTMENT BOULEVARD . HAYWARD, C ALIFORNIA 9454$ 3811 . (415) 887 8811 August 25, 1988 LABORATORY NUMBER: 588.043 P. O. OS9-1-06995, Am. 7 Revision 1 Req. BLN: 8804-07
SUBJECT:
Crack-Tip Opening Displacement Testing of Valve-Body Material MARK: Hills McAnna 2 150 229-0351 979-15 972 CA 954 5902 DATE SUBMITTED: May 5, ' 88 REPORT TO: Bechtel Group, Inc. ATTN: Mr. Yun Chung P. O. Box 3965 San Francisco, CA 94119 INTRODUCTION Two valve-body castings were submitted by Bechtel Group, Inc. for crack-tip opening displacement (CTOD) testing. The castings,were reportedly made from the aluminum-bronze alloy 095400 and will be identified as Valve Body 1 and Valve Body 2 in this report. Testing was required to be performed at room tem-perature in accordance with the British Standards Institution method BS5762:1979, using standard three-point bend specimens. The specimens for testing were required to be taken from near the flat side containing studs. A photograph furnished by Bechtel showing the location of one specimen is in Appendix A. The second specimen was from a similar location directly across the opening of the velve body. TESTING - Two standard three-point bend specimens meeting the require-ments of BS5762:1979 were machined from each valve body. Figure 1 shows a sketch of the test specimen. The specimens had nominal thicknesses (B) and widths (W) of 0.2375 in, and 0.475 in., l'
LABOR AToRY CERTIFICATE , Anomet Laboratories, Inc. .PAGE W H AYWARD, C AUFomNIA Lab. No. 588.043 respectively, and were 2.2 in. long. The loading span was 4W or 1.9 in. as required by the specification. A chevron notch with the notch-tip at about 0.15 in. from one edge was machined at mid-length in each specimen by electrical discharge machining. The measured thicknesses and widths of the test specimens are listed in Table 1. The specimens were fatigue precracked at the tip of the machined notch by cyclic loading in three-point bending at room temperature. The ratio of minimum to maximum load, R, was 0.1. The specification BS5762:1979 requires that the maximum l stress intensity factor used for fatigue pr'ecracking be no greater than 0.63 o y B 1/ 2 where o y is the 0.2% offset yield strength (Note: As defined earlier, B is the specimen thickness). Tensile tests performed on specimens taken from the two valves showed a yield strength of 35,500 psi for valve Body 1 and 26,800 for Valve Body 2 (Table 2). When precracking was attempted at maximum stress intensities calculated using these yield strengths, crack growth was too slow to be practical. All subsequent fatigue precracking was performed by adjusting the maximum loads with an objective of completing the precracking in less than 500,000 cycles. The general procedure followed was to start with a conservatively low maximum load and then to increase the load gradually in small steps when crack growth was observed to be slow or non-exis tent. Following fatigue precracking, 0.070-in. thick knife edges for supporting the crack-tip opening displacement gage were bonded to the specimen edge on either side of the notch opening. The knife edges were 0.15-in apart. Specimens were then frac-tured in three-point bending at room temperature (70*P). Plots of load vs. gage displacement obtained from the tests are included in Appendix B'.' , An examination. on the fatigue-crack fronts in the f ractured test specimens sh' owed that the crack length (i.e., the length of the machined notch piu's the length of the fatigue crack) was 2
-w LA8 ORATORY CERTIFICATE , f --
j PAGE45 0F Anamet Laboratories, Inc, _ E[vwamo. c Auroama Lab. No. 588.043 minimum at one of the two specimen surfaces and was maximum near mid-thickness. The crack length was measured at each specimen surface and at the location of maximum crack length. Additional crack-length measurements were performed at 25% B, 50% B, and 75% B f rom one specimen surf ace. These latter measurements (labeled ai, a2 and a3) were used to calculate the average crack length (a). Table 3 shows the three measurements, maximum crack longth, minimum crack length, and average crack leng;h for each specimen. The individual and average crack lengths met the specified re-quirements. The average crack lengths were used in calculating the. crack-tip opening displacements and fatigue-precracking stress intensities. The load-displacement plots for all four specimens showed that considerable stable, slow crack growth occurred before the maximum load was reached in the specimens. As a result, only the l crack-tip opening displacements at maximum load could be calcu-lated for the specimens. These calculations were made using the equation: K2 (1-v2) 0.4(W-a)V p Om" 2o E + 0.4W + 0.6 a + s l Where 6, = crack-tip opening displacement it maximum load K = stress intensity factor at maximum load v = Poisson's ratio a = 0.2% offset yield strength E = Young's modulus of elasticity V p
= plastic portion of the gage displacement at maximum load z = distance of clip-gage location from specimen surface (0.070 in.)
3
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/ i LABOR AToRY CERTIFICATE IPAG Anamet Laboratories, Inc.
Lab. No. 588.043 H AYWARD. C ALIFORNIA Revision 1 The stress intensity factor was calculated from the equation:
=
YP K BW 1/ 2 where Y = compliance function P = ma::i=u= load prior to specitaan f ractitre The Poisson's ratio and the Young's modulus of elasticity were assumed to be 0 316 and 16 x 10 6 psi, respectively, for the valve-body material (Ref. Alloy 095400, Metals Handbook, Vol. 2, Ninth Edition, American Society for Metals). The crack-tip opening displacements calculated as described above are listed in Table 4. The maximum loads and stress intensities used in the final stage of fatigue precracking are compared in Table 5 with the maximum loads and stress intensities prior to specimen fracture during CTOD testing. CONCLUSIONS In the present investigation, CTOD tests were performed on 0.237-in. thick standard three-point bend specimens taken from two valve-body castings. The test results showed that slow, stable crack growth occurred in all specimens prior to fracture. The CTOD values calculated at maximum load prior to fracture were as follows: O l l l [ l t 1 44
ATTACHMENT 2 ST.HL AE Y1t/d LABOR AToRY CEQTIFICATE PAGE 4? OF6v Anamet Laboratories, Inc. H AYWARD. C AllFORNI A Revision 1 f ( Valve Body No. Specimen No. CTOD at Maximum Load (in.) I 1 ITR 0.0038 iTL 0.0025 Average 0.0032 2 2TR 0.0101 2TL 0.0089 Average 0.0095 Submitted by: Approved by: I, I h MA [ . /3 . M e w M. Dilip Bhandarkar E. B. Flora Principal Metallurgist President j a /mh/ 23'i 5'
LA8t R ATORY CERTIFIC ATE
~
Annmet Laboratories, Inc. i PAGE 4 R O H AYWARD, C AUPOR NI A Lab. No. 588.043 TABLE 1 THICKNESSES AND WIDTHS OF THREE-POINT BEND CTOD SPECIMENS Valve Body No. 1 Valve Body'No. 2 specimen specimen specimen specimen iTR ITL 2TR 2TL Thickness (B), in. 0.237 0.237 0.237 0.236 Width (W), in. 0.475 0.476 0.475 0.475 TABLE 2 RESULTS OF ROOM-TEMPERATURE TENSILE TESTS Valve Body 1_ Valve Body 2 Property _ Ultimate Tensile 85,000 90,900 Strength, psi 0.2% offset 35.500 26,800 Yield Strength, pai 24% Elongation in 1-in. Gage 10% 6
- 1 LABORATORY CERTIFICATE by#
T Anamet Laboratories, Inc. y g HAY! AMD. CAUPORNIA Lab. No. 588.043 TABLE 3 CRACK LENGTH MEASUREMENTS FOR CTOD SPECIMENS Valve Body No. 1 Valve Body No. 2 specimen specimen specimen specimen Crack Length (in.) 1TR ITL 2TR 2TL At 25% B (at) 0.240 0.255 0.240 0.235 0.245 0.263 0.245 0.243 At 50% B (a2) 0.245 0.250 0.235 0.235 At 75% B (a3) Maximum 0.245 0.265 0.245 0.245 Minimum 0.210 0.225 0.207 0.208 Average 0.243 0.256 0.240 0.238 ag + a2 + ag flot e : Average Crack Length, a= 3 9 4 7
f477Atl&f NT LABORATORY CERTIFICATE ST-HL AE-Anamet ~-~' Laboratories. Inc. Lab. No. 588.043 HAYWARD. C ALIFORNIA Revision 1 TABLE 4 RESULTS OF CTOD TESTING Valve Body No. 1 Valve Body No. 2 Specimen Specimen Specimen Specimen 1TR ITL 2TR 2TL Average Crack Length, a (in.) 0.243 0.256 0.240 0.238 Ratio of Average Crack Length to Width, a/W 0 512 0 538 0 505 0 501 Compliance Function, Y 11.03 12.04 10.78 10.65 Plastic Component of OageDigplacement, Vp 13 55 8.75 37 55 32.20 (10- in.) Maximum Load, P (1b) 452 381 545 546 Crack-Tip Opening Displacement 6, (in.) 0.0038 0.0025 0.0101 0.0089 8'
m. I ATTAcRMENT'4 , LAs RAToRY CERTIFICATE ST HL AEgOf I p , Annmet Laboratories, Inc. tf. AGE 51 HAYWARD, C ALIPORNIA Lab. No. 588.043 i TABLE 5 LOADS AND STRESS INTENSITIES FOR FINAL STAGES OF FATIGUE PRECRACKING OF CTOD SPECIMENS Valve Body No. 1 Valve Body No. 2 specimen specimen specimen specimen 1TR ITL 2TR 2TL Maximum for Final Stage- ' of Fatig0e Precracking Load (Lb) 180 160 180 175 Stress (ksi-in {/gensity r)
) 12.2 11.8 11 9 11 5 Maximum Prior to Specimen Fracture 381 545 546 Load (ib) 452 Stress Tritiensity (ksi-ini/ 4) 10 5 28.1 36.0 35 8 l 9
m: r- . . - _ ATTACHM NT4 ST HL AE 7 # J PAGE 42 F<R Lab. No. 588.043 12 0'
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l 1 3 Figure 1 Sketch of the three-point bend specimen used for CTOD testing. 10
A1T/sCHME 4TR ST HL AE. 749 LABORATCRY CERTIFICATE PAGE 47 O 59 Anamet Laboratories, Inc. H AYWARD, C ALtPORNI A Lab. No. 588.043 m i i~ i s i APPENDIX A , i i CTOD SPECIMEN LOCATION i 1 i 1 i l I l I i t 11 l l P
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i i f i f m w < ta . .. . .___ ...__ _ __ _ _ _ m Figure A-1 Photograph of the valve-hody casting, showing the location of one of the two CTOD specimens machined from the casting. The other specimen was taken from the location directly across from the first.
t.Asam ArcMY canitricATE ' d}T H - 3 PAGE g6 0F 59 Anamet Laboratories, Inc. M AYWARD, C AllPORNI A Lab. No. 538.043 i APPENDIX B PLOTS OF LOAD VS. CLIP GAGE DISPLACEMr:'IT I i i
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