ML20154Q132
| ML20154Q132 | |
| Person / Time | |
|---|---|
| Site: | Browns Ferry  |
| Issue date: | 06/14/1985 |
| From: | Kuo A, Riccardella P, Tang S STRUCTURAL INTEGRITY ASSOCIATES, INC. |
| To: | |
| Shared Package | |
| ML18030B197 | List: |
| References | |
| SIR-85-008, SIR-85-008-R00, SIR-85-8, SIR-85-8-R, TVA-06, TVA-6, NUDOCS 8603210047 | |
| Download: ML20154Q132 (110) | |
Text
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Fracture Mechanics Evaluation of Recirculation System Piping Welds in Browns Ferry Unit 2 Nuclear Power Plant
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Report No.: SIR-85-008 Revision 0 SI Project No: TVA-06 -
June 14, 1985 I .
Fracture Mechanics Evaluation of Recirculation System Piping Welds in Browns Ferry Unit 2 Nuclear Power Plant 0 _
Prepared by:
M Structural Integrity Associates IU San. Jose, California t
Prepared for:
Tennessee Valley Authority
, Prepared,by: ._
A. Y. KtV 4()
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Date:
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W/V.!-- Date: N0lW Reviewed and /
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Approved by: d 4,4 A
'P. C. Riccataclla~' V "
Date: 6/I3/
Project Manager h ^ STRUCTURAL
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_ REVISI0N ,
CONTROL SHEET -
SECTION PARAGRAPH (S) DATE REVfSION REMARKS All
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All 6/14/85 0- Initial Issue
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TABLE OF CONTENTS Page -
LIST OF TABLES .. . . . . . . . . . . . . . . . iii LIST OF FIGURES ..
iv
1.0 INTRODUCTION
. . . . . . . . . . . . . . . . . 1-1 2.0
SUMMARY
OF INSPECTION RESULTS . . . . .. . . . . . 2-1 3.0 EVALUATION METHODOLOGY . . . . . . . . . . . . . 3-1
[ 3.1 Applied Stresses '
j 3.1.1 Stresses Due to Operational Loadings 3-1 3-2
~ 3.1.2 Residual Stresses . .. . . . . . . . . .
3.2 Stress Intensity Factors . . . . . . . . . . 3-3
. 3.3 Crack Growth . . . . . . . . . . . . . . 3-5 3.4 Allowable Flaw Size (IWB-3640) . . . . . . . . 3-6 3.5 Allowable Flaw Size (EPFM) . . . . . . . . . . 3-7 3.5.1 3600 Part-Through-Wall Cracks in Tension . . . 3-7 3.5.2 Through-Wall Cracks in Tension . . . . . . 3-9 3.5.3 Limitations for J-Controlled Growth . . . . 3-10 3.5.4 Stress-Strain Laws Considered . . . . . . 3-11
' ' 3.5.5 Weld Metal Toughness Data . . . . . . . . 3-14 3.5.6 Critical Flaw Size Determination . . . . . 4-1
[\ 4.0 EVALUATIONS AND RESULTS . . . . . . . . . . . . 4-1 4.1 Weld KR-2-14 . . . . . . . . . . . . . . 4-1 4.2 Weld KR-2-36 . . . . . . . . . . . . . . 42 4.3 Weld KR-2-41 . . . . . . . . . . . . . . 4-4 4.4 Weld KR-2-37 . . . . . . . . . . . . . . 4-5 .
5.0 DISCUSSION AND CONCLUSIONS 5-1
6.0 REFERENCES
.. . . . . . . . . . . . . . . . 6-1 l- l
_
- 8 STRUCTURAL INTEGRITY 3
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_ - r . - - , -
LIST OF TABLES Table - - - -
Page 2-1 Upper Bound Crack Sizes and Worst Case Crack Locations. . . 2-2 3-1 Summary of Applied Stress at Weld KR-2-14 _. . . . . . 3-16 3-2 Summary of Applied Stress at Weld KR-2-36 . . . . . . . 3-17 3-3 Summary of Applied Stress at Weld KR-2-41. '
. . . . . . 3-18 3-4 Allowable End-of-Evaluation Period Flaw Dep'th '
to Thickness Ratio for Circumferential Flaws -
I I - Normal Operating (Including Upset and Test)
Conditions . . . . . . . . . . . . . . . . . 3-19 3-5 F for a Circumferential1y Cracked Cylinder in i
Tension (t/Rj = 1/10) . . . . . . . . . . . . . 3-20 3-6 h1 for a Circumferentially Cracked Cylinder
. . in Tension (t/Ri = 1/10) . . . . . . . . . . . . 3-20
~~ 3-7 Ff6raCircumferential,Through-WallC[ack
[j in a Cylinder of t/R = 1/10 in Tension .
. . . . . . . . 3-21 3-8 h1 for a Circumferential Through-Wall Crack in
) a Cylinder of t/R = 1/10 in Tension .
4
. . . . . . . . 3-21 3-9 Material Stress-Strain Properties of Base and Weldment Materials Used in Analysis . . . . . . . . . 3-22 3-10 Three Commonly used Ramberg-Osgood Constants for Weldment Materials . . . . . . . . . . . . . . 3-23 3-11 Welding Processes For Stainless Steel Pipe . . . . . . 3-24 4
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LIST OF FIGURES l
Figure . Page 2-1 Flbw Tndications in Weld KR-2-14 . . . . . . . . . 2-3 2-2 Flaw Indications in Weld KR-2-36 . . . . . . . . . 2-4 2-3 Flaw Indications in Weld KR-2-41 . . . . . . . . . 2-5 i
2-4 Flaw Indications in Weld KR-2-37 . . . .". . . . . 2-6 3-1 Comparison of Measured and Computed Residual Axial Stresses Along Inner Surface of a Welded, IHSI Treated 12-Inch Sweepolet . . . . . . . . .
3-27
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3-2 Comparison of Measured and Computed Residual Circumferential Stresses Along Inner Surface of a Welded, IHSI Treated 12-Inch Sweepolet . . . . .
3-28 I 3-3 Computed Residual Stresses at Weld Centerline for i IHSI Treatment of a 12-Inch Sweepolet . . . . . . . 3-29 I 3-4 Computed Residual Stresses 0.75 inch (1.9 cm)
'j from Weld Centerline, 0.25 inch (0.64 cm) from
__ Co_il Centerline, for IHSI Treatment of..a 12 Inch Sweepolet . . . . . . . . . . . . . . . . 3-30 1
h 3-5 Post-IHSI Residual Stress Distribution . . . . . . . 3-31 3-6 Analytical Model for Post-IHSI Residual Stress Calculation . . . . . . . . . . . . . . . . 3-32
' 3-7 Post-IHSI Residual Stress Distribution . . . . . . . 3-33
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3-8 Magnification Factors of Circumferential Crack in a Cylinder (a/t = 0.1) . . . . . . . . . . . 3-34 3-9 Stress Corrosion Crack Growth Data for Sensitized Stainless Steel in BWR Environment (Ref. 7) . . . . .
3-35 3-10 Common Assumptions Used to Estimate Circumferential Crack Growth . . . . . . . . . . . . . . .
3-36
. 3-11 Average Effective Circumferential Crack Growth Rate As a function of Operation Periods Used in Calculation of Time Between Inspections . . . . .
3-37 3-12 Tearing Modulus Concept for Stable Crack Growth . . .
3-38 3-13 Circumferentially Cracked Cylinder in Tension . . . . 3-39 STRUCTURE
. . iv INTEGRITY ASSOCIATESINC 4
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LIST OF FIGURES (continued)
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3 .4 Through-Wall Flawed Cylinder Under Remc e Tension 1- . . . 3-40
_- 3-15 Ramberg-Osgood Characterization Stress-Striin Curves . .
3-41 3-16 Compilation of Material Toughnest J-T Curves (from Data of Refs.17 to 21) .
3-42
" 3-17 Lower Bounds of J-T Data for Wrought Stainless Steel Base Metal and for Stainless Steel Weld Metal from TIG, SMAW and SAW Welding Processes .- .
3-43 3 '.3 Effect of Ernst Correction on Lower Bound Weld and Base Plate J-T Curves . . . . . . . . . . .
3-44 2 .)
- Lower Bound J-T Reference Curves for use in Elastic-Plastic Fracture Mechanics Analysis of Austenitic Stainless Steel Pipes . . . . . . . .
3-45
- ?0 Material J-R Curve Derived from Lower Bound J-T Diagram for SAW/SMAW Weldment Material 3-46
-21 Material J-R Curve from Lower Bound J-T Diagram for SAW/SMAW Weldment Material (Expanded Scale) . . . . 3-47 fI '
4-1 Stress Intensity Factor Versus Crack Depth for Weld KR-2-14 . . . . . .
48
-2 Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - Weld KR-2-14 . . .
49 4-3 Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld KR-2-14 . . . . . .
4-10 4-4 Stress Intensity Factor Versus Crack Depth for Weld KR-2-36 . . . . . . . . .
4 11 4-5 Predicted Stress Corrosion Crack Growth for
- Observed Ultrasonic Flaw Indication - Weld KR-2-36 . . .
4-12 4-6 Comparison of Predicted Crack Growth with Allowable
} Flaw Size Limits - Weld KR-2-36 . . . . .
- ~
1 . . . .
4 13 4-7 Stress Intensity Factor Versus Crack Depth for Weld KR-2-41 . . . . . . . .
4-14 4-8 .
Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - Weld KR-2-41 . . .
4-15 i ~
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4-- Comparison of Predicted Crack Growth witE i Allowable Flaw Size Limits - Weld KR-2-41 . . . . . . 4-16 4-10. Stress Intensity Factor Versus Crack Depth '
l for Weld KR-2-37 . . . . . . . . . . . . . . 4-17 6
r 4 - ?1 Predicted Stress Corrosion Crack Growth for !
Observed Ultrasonic Flaw Indication - Weld KR-2-37 . . . 4-18 4-12 Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld KR-2-37 . .
. . . . 4-19 3 .
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1.0 INTR 0gUCTION During th_e 1984/85 outage at the Browns Ferry Unit 2 Nuclear Power Plant, ultrasonic (UT) examination of the recirculation system piping produced indications at four weld joints which are believed to result from inter-granular stress corrosion cracking (IGSCC). Similar indications have been observed at a number of other Boiling Water Reactors (BWRs) in the U.S. and overseas.
r These four welds have been evaluated to demonstrate their acceptability in accordance with ASME Section XI requirements, supplemented by the recom-6
' mendations of NRC Generic Letter 84-11. The welds were also analyzed using Elastic Plastic Fracture Mechanics Tearing Instability methodology to account for possible effects of low toughness weld metal. All of the welds I
were treated by induction heating stress improvement (IHSI) to inhibit
- further IGSCC propagation.
5
. Structural Integrity Associates (SI) was contracted by the Tennessee Valley Authority (TVA) to perform the evaluations of the four weld joints. This report documents the results of the analyses, which demonstrate that design basis safety margins are maintained in these welds, considering worst case interpretation of the UT indications.
Section 2 of this report summarizes the inspection results. Section ,
3 describes the flaw evaluation methodology used to evaluate the welds, and Section 4 presents the evaluation results. Section 5 presents the conclusion of the evaluation regarding the continued, safe operation of the plant.
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- 2.0
SUMMARY
OF INSPECTION RESULTS After a thorough in-service inspection of the recirculation and associated stainless ~ steel piping systems, IGSCC-like indications were found in three
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ring header-to-sweepolet welds and one ring header-to-end cap weld. Figures
" 2-1 to 2-4 provide a weld-by-weld summary of these indications, including indication sizes detected after IHSI treatment *. All the indications are circumferentially oriented, and have been consdrvatively assumed to be cracks or crack-like for purposes of this evaluation.
Upper bound crack dimensions and worst case positions with respect to the applied stresses were used in the crack growth calculations and are tabulated
. in Table 2-1.
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- Some changes in indication sizes occured between the pre- and post- o IHSI inspections of these welds, but they were not significant.
2-1 STRUCTURAL INTEGRITY ASSOCIATESINC
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_ TABLE 2-1 Upper Bound Crack Sizes and Worst Ca:e Crack Locations ,
r Crack Depth Crack Length Weld No.
Worst Applied Stress l (% Wall Thickness) (inches) location (decrees)*
KR-2-14
{ 12.6 i
2.1 80**
KR-2-36 25 2.2 80**
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KR-2-41 19 4 60**
KR-2-37 12 5 any position I l g
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the highest stress location (see Section 3.1).
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3.0 EVALUATION METHODOLOGY .
3.1 Applied Stresses Two major types of stresses are considered in this evaluation, stresses due
, to operational loadings and residual stresses. The operational stresses include pressure, deadweight, shrinkage due to weld overlay repair of weld GR-2-15, thermal, and seismic. Residual stresses were evaluated considering the beneficial effects of the IHSI treatment which was performed on these welds. These stresses are described in more detail in the following tections.
- 3.1.1 Stresses Due to Operational Loadings 1
The applied moments on sweep-o-lets KR-2-14, KR-2-36 and KR-2-41 were provided by TVA (Ref.1). Due to the complexity of the sweep-o-let geometry, the stress due to these moments varies in a non-linear fashion along the azimuthal location of the weld between the ring header and the sweep-o-let.
! Also, the stress varies with distance from the crotch region of the sweep-o-let (Ref. 2). The highest stress location irea sweep-o-let is at the crotch
] region, and the stress decays rapidly as one moves away from that region.
I Since the weld seam between the sweep-o-let and the ring header is somewhat removed from the crotch region,.it does not see the full stress concentration attributable to the crotch region, but the stresses are still higher than the
, nominal bending stresses caused by the pipe applied moments. A stress concentration methodology for such a sweep-o-let weldment location was developed in Reference 2, and is used in this evaluation to obtain the appropriate stresses.
Table 31 to 3-3 present the applied stresses due to the various applied j
loadings for welds KR-2-14, KR-2-36 and KR-2-41 calculated in accordance with the Reference 2 methodology. As azimuthal angle increases from 00 (longitudinal section) to 900 (transverse sections), the stress concen-tration factor on bending moment increases from 1 to 3. The corresponding .
j stress concentration factor for pressure incr6ases from 0.6 to 1. These
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inside surf ace concentration factors also include local through-wall bending e stress effects on the moment terms. At the weld location, the magnitude of j
. 3-1 ^STRUCTURAI INTEGRITY ASSOCIATESINCl
the out' side surface stress is approximately two-thirds of the inside surface
. stress and is compressive (Ref. 2).
Tables 3-1 to 3-3 also give the resultant through-wall membrane and bending -
stresses, and the ASME Code stress ratio (Pm + Pb )/Sm for normal and faulted conditions.
The maximum normal condition membran_e stress is seen to be 15.4 ksi for all welds and locations, which corresponds to a maximum stress ratio of 0.91. Note that this stress ratio conservatively includes thermal expansion and weld overlay shrinkage stresses as,a primary stress term. The maximum faulted condition stress ratio is only negligibly higher (0.925) and, therefore, normal conditions will govern the allowable flaw size calculations.
t 3.1.2 Residual Stresses Residual stresses are known to play a significant role in IGSCC. A favorable f residual stress pattern can arrest further crack growth, while an unfavor-able one can accelerate crack growth. A general survey of available
' analytical and experimental results was performed to establish the most appropriate residual stress profile for use"in the subsequent crack growth
[\ analysis.
- In evaluating the indications two representative post-IHSI residual stress distributions, one for the sweep-o-let welds and one for the j end cap weld, are considered.
, Rybicki, et al I (Ref. 3) have presented extensive analytical results on
(_ induction heating of welded stainless steel pipes. These analytical results cover a wide range of piping welds and fittings. Also Ishikawajima-Harima
- Heavy Industries (IHI) Company in Japan did an in-depth study to qualify and verify the IHS1 process for boiling water reactor piping (Ref. 4).
Figures 3-1 to 3-4 present computed and experimentally measured residual i
~ stresses for a sweep-o-let weld with IHSI treatment. Figures 3-1 and 3-2 present the inner surface stresses, circumferential and longitudinal, versus distance from the weld centerline. The measured stresses compare very favorably to the analytical results. As shown in the figures, the surface residual stress in a 12-inch sweep-o-let is about 20 to 40 ksi compressive.
Figures 3-3 and 3-4 present through-wall analytical results, from two I,
different finite element models. They give about the same m nitud 3-2 $T INTEGRITY N m ASSOCIATESINC
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surface stress, as compared to each other and to those in figures 3-1 and 3-2.
The two finite element models also give similar through-wall residual stress patterns. '
f 3
Finally, Figure 3-5 presents test data on 12 inch sweep-o-let weld to a 22 inch pipe, from Reference 5.
The test data give only inside and outside surface stresses, but at three different azimuthal angles; 00, 450, 900 Since no through-wall test data are available from that test, linear through-wall stress profiles are assumed. Of the three angles examined, the inside surface stress at 00 had the least compressive stress, but all three through-wall stress profiles are similar. Also, the surface stresses agree
. reasonably with the previous analytical and experimental results. Thus, for conservatism, the 00 residual stress distribution of Figure 3-5 was used for
.i sweep-o-lets in this evaluation.
Figures 3-6 and 3-7 present an analytical model and IHSI residual stress results for a 16-inch end cap (Ref. 4). No results, either analytical or
}
experimental, are available for a 22 inch end cap. Therefore, the 16 inch end cap analytical IHSI residual stress is assumed for the 22 inch end cap in the g recirculation system. ~
Figure 3-6 presents the finite element model identifying all the dimensions, boundary conditions and length of the heating coil. Figure 3-7 presents the inside and outside surface stresses as a function of the distance .from the weld centerline. Near the weld centerline, the compressive surface stresses are on the order of 30 ksi.
Since no through-wall data are available, a linear through-wall stress profile is assumed for the subsequent end cap crack growth analysis.
3.2 Stress Intensity Factors l
~
Pipe dimensions used in this analysis are as follows (Ref. 1):
I '
~~
22 Inch Pipe 12 Inch Pipe j, Outside Diameter (in.)
22 12.75 Inside Diameter (in.) 19.75 11.592 Pipe Wall Thickness (in.) 1.125 i 0.579
. 3-3 ^ STRUCTURAL INTEGRITY ASSOCIATESINC
l An analytical model of a 3600 circumferential crack in a cylinder of radius to thickness ratio of 10:1 (Ref. 6) was used for the fracture mechanics crack growth evaluation.
The applied loading consists of piping loads due to shrinkage, dead weight, pressure, thermal expansion, and seismic and the residual stress distributions discussed in Section 3.2. For the piping loads, the loading consists of the piping stresses tabulated in Tables 3-1 to 3-3 for the sweep-o-lets, or just internal pressure stress for the 22 inch end cap (5.662 ksi axial, 11.324 ksi circumferential).
The post-IHSI residual stress distributions are given in Figures 3-5 and 3-7. :-
For purposes of the fracture mechanics analysis, the axial stress distribu-o tions from these loading cases have been expressed in terms of third degree polynomials of the form:
1 g a = Ao + Ai x + Apx2 + A3x3 (1) f' where eis axial stress in the units of ksi, x is the distance from the inside s
surface, and A0-A3 are the coefficients resulting from the curvefit.
i The stress intensity factor for a circumferential crack in a cylinder of radius to thickness ratio of 10:1 can be expressed as follows (Ref. 6):
K1 = 6 (A FO1+ AFI2+ AF 23+ha3 AF) 34 (2) where F 1 , F 2 , F 3, and F 1 4 are magnification factors and a is crack depth as s_ shown in Figure 3-8.
For the linear elastic fracture mechanics portion of j
the analysis, the stress intensity f actors can be calculated independently for piping stress and post-IHSI residual stress distributions, and the resultant stress intensity factor is the superposition of the two loading l-cases.
I i -
I 3-4 STRUCTURAL INTEGRITY ASSOCIATESINC
3.3 Crack Growth A large body of laboratory data exist on stress corrosion crack growth ratas for sensitized stainless steels in simulated BWR environments. These data are surmiarized in Figure 3-9, taken from Reference 7. These data were obtained using fracture mechanics type specimens with different crack si2rs
_ and loadings, which can be characterized by the crack tip stress intensity factor K. The data represent a wide variation in material sensitization, as well as levels of dissolved oxygen in the water. While subject to sone criticism because the simulated water chemistry in these tests did mit
.,- . , contain levels of impurities (chlorides, sulfates, etc.) that could exist ih j operating BWRs, the "best estimate" curve of Figure 3-9 is widely believed ta 1
- provide a reasonably conservative bound of stress corrosion crack growch rate for weld sensitized 304 stainless steel in BWR environments. This cure can be described by a power law representation of the form:
da/dt = 2.27 x 10-8(g)2.26 (3) where a is the crack depth in units of inches, t is time in units of hours, and K is the stress intensity factor in units of Ksi Vin.
t Crack growth analyses typically make use of one of the two assumpticns
-illustrated in Figure 3-10 regarding crack length extension, self-similar crack growth or constant aspect ratio crack growth. The former assumes that the incremental crack extension is the same at all points on the crack frord,.
while the latter assumes that the ratio of depth to length remains constant during crack extension. Considering field and laboratory experience with circumferential crack extension, it appears that the self-similar assumption may underpredict crack length versus time, while the constant aspect ratio assumption overpredicts.
Recent work by Gerber (Ref. 8) under contract to EPRI provides a new approach
!_ for addressing circumferential crack extension which is more technicall'y ;
defensible than the above self-similar or constant aspect ratio approaches. f
.- This approach utilizes data generated in a laboratory stress corrosion test ,
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ASSOCIATESINC
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~
Laboratories (Ref. 9). IGSCC was induced in this pipe through loading to a high applied stress in a simulated BWR environment, which was accelerated by -
the use of graphite wool to create an artificial crevice. Crack growth occurred and was monitored both during operation and at several scheduled shutdown intervals for the test. A number of small' cracks initiated early in the test, the length of which was periodically measured and the initiation of new cracks was noted and their lengths subsequently tracked as well. At the completion of the test, there were a total of 63 cracks with a combined length of 32.57 inches.
,_ ~
L The average effective circumferential crack extension observed in this test is presented in Figure 3-11. This rate includes both growth of existing cracks as well as new defects initiating and contributing to the effective crack growth rate in each inspection interval. Examination of Figure 3-11 suggests that an average effective circumferential crack growth rate of 0.5 mils / hour should give a reasonably conservative estimate. Thus, 0.5 mils / hours was used as the crack length growth rate in this report. It should
]'
__ be pointe _d out, however, that although this is_an average effective rate, it j is based on a laboratory test in which the local environment, load and cycles l
were all intentionally modified to accelerate IGSCC relative to actual plant
, conditions. Test and analytical data (Reference 4) have also shown that the
[ IHSI will suppress not only crack initiation but also crack propagation for
. cracks in both the length and depth directions.
3.4 Allowable Flaw Size (IWB-3640)
Based on detailed calculations presented in References 10 and 11, allowable flaw sizes for various levels of primary and applied loading (Pm + Pb) have
_ been specified in ASME Section XI, IWB-3640 (Ref. 12). A tabulation of I '
allowable flaw sizes as. a function of applied load is given in Table 3-4, l
_ which is taken directly from Section XI, IWB-3640. Note that this table permits very large defects in some cases (as great as 75% of pipe wall) and l[ does not include consideration of any stress other than primary, notably -
secondary and peak stresses from the design stress report as well as any weld l residual stresses or misalignment / fit-up stresses which might exist from .
construction. The argument for this exclusion is that, given the extremely
~
i l-3-6 STRUCTURAL
~
INTEGRITY ASSOCIATESINC
high ductility of austenitic stainless steel, these strain controlled
~
effects will self-relieve after a small amount of plastic deformation and/or stable cracLextension, and will have little or no impact on the loads and flaw sizes needed to cause unstable crack propagation or pipe rupture.
I i
However, some recent fracture toughness data may invalidate the above argument, at least for some classes of austenitic weld metal (Ref.13). To account for possibility of low ductility weld metal, secondary stresses from stress report were also included in the IW8-3640 allowable flaw size determinations in this report, although it is not required by the ASME Code.
h
' It is important to note that the very low measured toughness occurred only in a small percentage of the materials addressed in Reference 13, and may be of
( only limited concern from a probabilistic viewpoint. Indeed, most IGSCC l
observed to date has been restricted to weld heat affected zones, which should exhibit the high toughness attributed to base material. Also, the low l toughness data to date has been limited to flux types of weldments (submerged
~~
~
arc or shielded metal arc), which are not used in current construction practice nor in weld overlay repairs of pipe cacks. Nevertheless, tc address these possible concerns, the analysis procedure used in this report includes j
thermal expansion effects as a primary stress condition in determining allowable flaw size from Table 3-4.
3.5 Allowable Flaw Size (EPFM)
Methodologies from References 14 and 15 are also used in this report to calculate applied J and T values for circumferential through-wall or part-
~
~
through-wall cracks in pipes as functions of applied loading. Details of the
~
methodology used are provided below. These computed, applied J and T values
~
l .
are then compared to a d/T material curve on a J/T stability diagram (as in Figure 3-12) to provide a second means of determining allowable flaw size.
3.5.1 3600 Part-Through-Wall Cracks in Tension As shown in Figure 3-13, consider a cylinder with an inner radius Ri , outer radius Ro, and wall thick. ness t =
Ro -
Rj, containing an internal t
.^ STRUCTURAL i 3-7 INTEGRITY ASSOCIATES.INC
. , ,_ - _ + _
i -
l axisymmetric part-through crack of depth a. a" denotes the far field uniform
, te sile stress and c = t-a the uncracked ligament. A radius to thickness r dio (R i/t) of 10 is used in this report, which corresonds approximately to tne Schedule-80 piping used in service. The elastic-plastic formulae for J-applied in this case have been obtained from Reference 14 and are as follows:
. Jappl = Je + Jp C
=fi (a e,Rj/Ro)(P2 /E') + a co co C (a/t) h1 '(P/Po )"*I (4)
I
-here: .
aF 2 f 1(ae, Rj/Ro) =
[
r (RoA - Rj )2 I
I E' = E/(1-v2)
E = Young's Modulus v = Poisson's Ratio "o= Yield Stress co= Yield Strain
. 3 a,n= Material constants of Ramberg-Osgood Model i
a e = a + [1 + (P/P o
)2]'I [(n" 1) v o5 )/(2r)]
K = c* gra F Po = 2/4 3 eo [Ro2 - (Ri + a) ]
P= 3" r(Ro 2 - Rj2)
F = function given in Table 3-5 h1 = function given in Table 3-6 For materials with n values between 1 and 10 but not exactly as provided in Table 3-6, the corresponding h1 values can be calculated by interpolation.
The non-dimensional tearing modulus, T appl is calculated by:
E l .
Tappi * ( co2 ) (d Jappj/da) t (5) e 3-8 STRUCTURAL INTEGRITY ASSOCIATESINC
~ - - - , . ..y . ,
where Tappl is the applied tearing modulus from loading and all the other c mtities are defined in the same manner as those in Equation (4).
~
e o
Ine applied -tearing modulus can be determin d numerically by applying a finite difference scheme on the above definition.of Tappl e.9.,
Tappl * ( E' } J(a+da) - J(a) e2 da where da is a small crack length increment.
, -.2 Through-Wall Cracks in Tension a second case, consider a cylinder containing a circumferential through-11 crack of length 2a, and subjected to remote uniform tension as shown in gure 3-14. In this figure, R denotes the mean radius, t the wall thickness,
,' the total angle span of the crack, and 2a = 2RY the total length of the "ack. 2b = 2rR is the pipe circumference, and P is the applied load. Load
's applied by a uniform stress field at its. ends given by e" = P/(2rRt) (7)
,3in, a radius to thickness ratio of 10 is used to approximate service ping.
For a Ramberg-Osgood material, the elastic-plastic J-integral
- estimation has been obtained from Reference 15 and is given as follows:
Jappl = Je + Jp Jappl
- fl (a e, ( ) (P /E) 2 + ac co o C (a/b) h1 (P/P o)"*1 (8) where:
f(a,f)=aF/(4:Rt) 1 e 2' 22
, E = Young's modulus l co = Yield stress to = Yield strain I
l_. a,n = material constants of Ramberg-Osgood Model i -
l~
^ STRUCTURAL INTEGRITY 3-9 . .
ASSOCIATES INC I4ti
' ~
ae = a + [1 + (P/Po ) ]~ [(" )( )2]/(S:-)
K = cr"V r a F i -
', Po = 2 cro Rt [r 2 sin-1(1/2 sin Y)] -
F = function given in Table 3-7 h1 = function given in Table 3-8 Interpolation is again used for materials with n values between 1 and 7, but r .t exactly as provided in Table 3-8.
The non-dimensional tearing modulus Tappl can be evaluated by differen-tiation of Jappl in the same manner as described in Section 3.5.1 for part-
/ ,
t : ugh-wall cracks. ~
- l. ~. .3
- Limitations for J-Controlled Growth I . . order for the above tearing modulus stability concept to be valid, certain 7.itations on the theory must be checked. These limitations are necessary
{ ' ensure that the incremental crack growth and non-proportional loading I
Sne in the immediate vicinity of the crack-tip are sufficiently small to g ,astify use of the J-integral in the analysis of crack growth, a condition II which is defined in Reference 16 as "J-controlled growth". These conditions
~
l e generally satisfied, in the large scale yielding range, if the uncracked i
,; ament of the cracked cross-section (b) is sufficiently large to satisfy
- following criteria:
co = h h >> 1 (9) t
- I and e
_ P == b cr) >> 1 (10)
While there are no generally accepted rules for how much greater than 1 these pl.;ameters must be ?.o ensure J-controlled growth, a value of (o =5 to 10
' i; suggested as anquate for Equation (9) and a value of P=25 has been used in a number of sources for Equation (10). These parameters are thus
[I- alculated in the J-T analyses which fcilow, and compared to the above values
- to provide an assessment of the validity of the calculations.
~
3-10 STRUCTURAL INTEGRITY ASSOCIATEStINC
1
, 3.5.4 Stress-Strain Laws Considered 1
~
. Thc cimary material stress-strain law used in this report is based on test data for stainless steel weldments and base metal reported by Westinghouse, (Ref.17). Figure 3-15 illustrates these stress-strain data at operating
~ ~
temperature and their Ramberg-Osgood representations. A complete tabulation 0: ;.aterial tensile properties and corresponding Ramberg-Osgood parameters fc these materials are listed in Table 3-9 at both 750F and 5500F. The weldment material properties at 5500F are used in this report as the primary basis for the J-T analysis results. -
Hc ;ver, since studies have shown J-T analyses to oe extremely sensitive to t' :pecific material stress-strain law characteristics used in the anal-
~ '
y: ., Ramberg-Osgood constants have been obtained for a number of different I
- i. nless steel weldment materials. Table 3-10 lists three comonly used I
c sets for weldment material at 5500F. As a parametric study, allowable
_ : $ size results have been calculated using all three data sets.
{
4 .
2 5.5 Weld Metal Toughness Data b
~
i ta on the elastic-plastic toughness properties of austenitic stainless j~ s 21 welds are presented in References 17 to 21 in the form of J-resistance
~
c .es, J-T values, and/or tabulated J cI and J-Resistance curve slope values. These data have been used to determine lower bound J-T material toughness curves for comparison with applied values to assess crack instability and allowable flaw size by the J-T method. The effects and results of welding process have been considered in establishing lower bound toughness properties for such stability evaluations.
A compilation of applicable material toughness J-T curves from the above references is shown in Figure 3-16. These curves represent wrought stainless st2el base metal toughness, along with the toughness of stainless steel weld l m-[31 representing submerged arc welding (SAW), stick or shielded metal arc welding (SMAW), and gas tungsten arc welding (GTAW) or tungsten inert gas
,! JIG) welding processes. With the exception of the data from Reference 20,
.ne J-T curves in Figure 3-16 were derived from J vs. crack extension (Aa)
, curves (J-R curves) and the following equation.
~
STRUCTURAL
- 3-11 INTEGRITY ASSCCIATES,1NC I
T= dJ
- 2 ,
{33) eg da where T ~= tearing modulus E' = E/(1 -v2) -
E =
elastic modulus = 30,000,000 psi V =
Poisson's ratio = 0.3
. of = material flow stress = 50,000 psi' dJ/da = slope of J-R curve at specific da
, E :ause of the absence of tensile properties in many cases, the above values or . flow stress and elastic modulus were assumed throughout. J-R curves and
~
. sile properties were not available from Reference 20, so J-T curves were t ad directly from this reference.
I l .
pre 3-17 presents lower bounds of the toughness data for the various j -
degories of material in Figure 3-16. It can be seen that there are distinct I ~
ifferences in lower bound material toughness _between the wrought base metal 2nd the various weld metals. The base metal is the toughest material 8
(largest values of J and T), with the SAW and SMAW weld metals being the least ough, and the TIG weld metal having intermediate toughness.
i,_
[ shown in the previous figures, SAW and SMAW welds possess lower tough
~ than TIG weld metal. Differences in the welding processes, primarily heat input differences and the use of flux versus inert gas shielding, can be used to explain such toughness differences. Key features of the TIG (GTAW), SMAW and SAW processes (Ref. 22) are given in Table 3-11, along with the
} relationship to weld metal toughness. Essentially, SAW is a high heat input, L flux-shielded process which can result in relatively coarse microstructures and relatively heavy, slag /non-metallic inclusion contents. Such micro-i structures would tend to give reduced toughness (Ref. 23). In comparison, TIG
, i:,a low heat input process with inert gas shielding rather than flux.
{ T 2refore, TIG welds should be of superior toughness. SMAW has intermediate neat input and shielding by gas and molten flux from the electrode covering.
- hus, SMAW welds are expected to have intermediate toughness - lower than TIG
~ , ,
and slightly higher than SAW. Figure 3-16 generally illustrates this expected trend. -
~
STRUCTURK 3-12 INTEGRITY ASSOCIATESIM a
e-
~
~
~
Lcwer bound J-T toughness curves for use in this analysis were derived from dr,a of Figure 3-16 and 3-17. Essentially, the data were divided into three -
. c igories based on the preceding toughness discussions: base material, TIG j- w ld metal, a'dn SAW and SMAW weld metal. For each of these categories, the
~ 1:gr bound curves at low J values were corrected for specimen size effects,
- -1 merged with the lower bound curves at higher J values from Reference 20
- a conservative manner.
To account for the effects of specimen size and geometry in the small s ecimen data of References 17, 18, 19 and 21, a modified J approach, known
- a Jm, was used along with a modified T, Tm (Ref. 24 and 25). Reference 24
~
' 59ws that Jm and Tm can correlate data for test situations in which the
' :itions for J-controlled crack growth described in Section 3.5.3 are e
ssly violated. This approach is applied here to adjust the lower bound j r .;a of Figure 3-17.
I j -
'h Reference 23, Jm is computed for the comp,ct tension specimen as I
311cws: _
Jm = J + y JP .da a[o b
(12)
~ ~
7 = (1 + 0.76 ( }) (13)
L where Jp is the nonlinear part of the deformation theory J, b is the remaining ligament, W is the specimen width, ao and a are the initial and extended crack lengths respectively, and y is as defined above. Also, from Reference 24, Tm is evaluated as:
l Tm = TTl+ k,7 Jp i b (14)
.l _
I l ~ the preceding equations for Jm and Tm, J and Jp must be defined as functions of crack extension for each material evaluated. Such definitions jave been obtained from the power law curve-fits of the J-R data of the lower 1
- f
'- . D STRUCTURAL 3-13 INTEGRITY ASSOCIATESINC-
..a
~ bou ' materials in References 17,18,19 and 21. The resulting values of Jm-
~
Tm $r the lower bound curves for each material category of Figure 3-17 are -
il. atrated in. Figure 3-18. In each case the Jm-Tm curve branches upward i
from the respective J-T curve at a prescribed poin.t on the curve.
F'ylly,lowerbound.1.n - Tm curves of Figure 3-18 are then faired into the h' 1 J, large Aa data of Reference 20 to obtain tt el lower bound J-T curves s own in Figure 3-19 for each material. Again, it can be seen in Figure 3-19 that three distinct levels of material toughness exist: the toughest-base
[
m "irial, the intermediate toughness - TIG welds, and the lowest toughness SA- and SMAW welds. These lower bound curves of Figure 3-19 are employed in
[ t' report to determine predicted fracture stresses for the subject pipe I -
w 3 using elastic-plastic fracture mechanics analyses.
I i !
brder to make crack growth corrections to applied J-integral values, a r
.arence J-R curve was derived from the lower bound J-T curve in Figure 3-I j 2, Since the tearing modulus (T) is a function of the slope of the J-R curve i
- /da), the reference J-R curve in Figure 3-20 was obtained by integrating
.e lower bound J-T curve in Figure 3-19. The lower data points at small da
- ~ Figure 3-20 represent the raw data from the unmodified J-R curve for the j i ar bound material toughness. Figure 3-21 shows an expanded scale of this I
i
~
? aa regime comparing the raw data (FUC-9) with the extrapolated Jm data.
I: can be seen that, in validating the raw dat 1, Jm gives significant toughness advantages over the deformation J Unw !ata).
3.5.6 Critical Flaw Size Determination The above J-integral estimation methods and material data are then used to establish allowable flaw sizes for the subject welds for comparison to the allowable flaw sizes for' these welds based on ASME Section XI, IWB-3640 methodology discussed previously (Section 3.4). The basic technique is i
! 11'ustrated in Figure 3-12. The intersection of the applied and material J-T curves in this figure yields a critical value of J for predicted nstability of the weld. This
- critical value of J uniquely defind a
_ critical stress for a given flaw size, or conversely, a critical flaw size '
for a given stress level. Th,e later definition is used here.
L 3-14 ^ STRUCTURAL INTEGRITY ASSOCIATESINC
~
At this point, it must be noted that the IWB-3640 tables for permissible flaw s- ~2s were developed based on an inherent safety factor of 2.773 on stress or .,
l .1 to net section collapse of the cracked cross-section (Ref.11). Thus, in order to provide a consistent basis for comparison, the applied loading on each pipe weld to be evaluated (Ref.1) is multiplied by a factor of 2.773 hefore applying the J-T critical flaw size determination described above.
.;1owable sizes for 3600 part-through-wall cracks and finite length,
- rirough-wall crace.s, defined in this manner, are thus used as end-points to prescribe a secor.d allowable flaw size locus for the subject welds, with the same safety margins, but under the assumption of lower bound, flux-type
( , - 'erial toughness from Figure 3-18. The new allowable flaw size loci are i nstructed by drawing a smooth curve parallel to the corresponding
~
awable locus from IWB-3640 between the two end points. Since such a
'tical flaw size determination potentially reflects less than full limit
.ad ductility in the pipe cross-section, it is also appropriate to include bal secondary stress terms (such as thermal expansion) in the above
$liedloading.
m .
- ritical flaw size loci have been determin[d in this manner for the four i ,
selds with UT indications in Section 4.0 of this report. They are then
- ompared to the
~ IWB-3640 based allowable flaw sizes, as well as to the
' .;erved flaw sizes, plus predicted IGSCC crack propagation during subse-lantoperation.
L
. i..
^ STRUCTURAL 3-15 MT M TY ASSOCWESINC
. a.s
TABLE 3-1
, Summary o4 A: plied Stress at held G-2-14
, l@CH MOMENTE LDie Mvy Mzz O
CASES (FT-LBS) (FT-LIS) 12' PIM Cd -2222 6560 PRESSIAE= 1150 PSI
~iE W L 16!19 -47473 TNK= 0.579 IN
- C9E-IY 2441 2467 OD= 12.75 IN EE-ZY 3190 2 06 Z= 64.5 INit!
SSE-IY 4134 4115 SW 16950 PSI -
'!SE-ZY 46B3 5173 i
. * ' NAEE 5241 -3835
~
~-
~
STESES 70RML TD kELD DLE TD ERAtCH IOOTS (PSI)
AW5LE (CE5REE) 0 10 20 30 40 50 60 70 80 90 SCF 3 M0 RENT 1 1 1.05 1.17 1.24 1.42 1. 9 2.!: 2.69
- 3
-K S PRES 0.6 0.61 0.64 0.67 0.69 0.75 0.86 0.29 0.95 1 6
IE!Ei 45SSLEE 7597.150 7723.769 210 .626 B483.484 8736.722 9496.437108S9.24 !!269.1012022.8212661.
tu 1220. 465 1273. 700 1352. 664 1478. 472 1482. 813 1563. 674 1742. 838 1708. 4a3 1665.2:5 1240.1 SL9 FACE TEF.% E822.186 9218.75 979.47910703.5510779.93 !!!23.7812622.7612378.2212070.07 8996 MIRASE712.4883 B71.9679 ;054.1501293.271454.9241711.906 2162.1:2 2459.E212916.369 2925.209 CSE 1111.251276.2?: 1472.6761732.E41 1890.595 215!.900 2633.220 2S02.722 32?4.388 142.82:
, SSE 1723 1986.595 2294.070 2710.513 2949.888 !!61.615 4112.237 4:20.E07 5152.735 4921.!!6 N351;RE 7597.150 7723.769 2103.626 B48:.484 6736.722 9496.437100S?.2411:19.1012029.8212661.91 CUIS1;( ;w
-813.643 -849.139 -901.776 -985.643 -992.542 -1042.44 -1161.89 -Il38.*S -1110.15 -E26.790 SURFACE THERML -5288.12 -6145.83 -6527.65 -71:5.70 -7186.62 -7549.19 -8415.64 -8252.19 -9' 951NrASE-475.658 CBE
'B1.311 -702.767 -662.238 -969.949 -!!41.27 -1441.41 -163?.58 -1944.24 -1 SE -740.837 -850.861 -991.784 -!!59.22 -1260.39 -1435.93 -1755.48 -1922.52 -2196.25 -20 5.25
-1152 -1324.39 -1529.38 -1907.00 -1965.02 -2241.07 -2741.52 -3013.57 -3C.15 -3280.74
~
PEES'IE + CW + TFEFM. + Sm!Pt.A5E les.est 9391.506 96i7.sii 20 36.6710729.:a iiO24.00 ii929.66 33644.03 iiO23.s7 i480..io 14a55.5, BENDIN5
.. 8971.782 9470.360 10165.24 !!229.49 !!436.39 12166.14 13773.93 13788.82 !!876.40 1096 PFESS'EE+ Du + THER'AL + 02E +SHRINXt6E f0'fmE 1Pr.+P!)IS3 9576. 7169830. 556 1082.12 ! ! 019.19 ! !!39.10 12238. 64 14082. 90 14509. 00 15E!.16 15:7 0.564999 0.579973 0.612514 0.650099 0.668973 0.724994 0.830B49 0.255998 0.9057910.907
- PFESSJE 't + THE?NL + SSE +SHRivAGE
! !E.9 5 IPP+ci 2 5.1 9679.506 9948.940 10519.02 11191.13 !!515.48 12439.9: 14:29.41 14780.!3 15662.99 156'5.77 0.571062 0.596958 0.6205910.65965: 0.679379 0.7:6869 0.845 9 0.671996 0.924064 0.924D
. +
s 3-16 STRCTURAL l DiTLr;RITY l isSSCGATES1NC 1
TABLE 3-2
. Scar.arv of A: plied Stress at held KR 6 -
~
l
. SPANCtfM0MENTS c LDft Myy M2 i CASES (FT-i.S$1 (FT-LBS)
I
. 12' PIFT DW 1922 -3U9 PPISSSE: 1150 PSI ,
MF7!AL 12!35 38810 THK 0.579 IN i 08E-IY 5097 599o OD= 12.75 IN TE-ZY 5050 4797 != 64.5 Ill413 .
SSE-IY 7ES B454 SM=. 16'50 PSI 7E-ZY 7524 7042 ST ' AASE 325 146 STRESSES NORML TO WELD [E'TO BRANCH M2ENTS (PSI)
AN5LE (DEEREI)
[ 0 10 20 30 40 50 60 70 80 90 t W PD LOOT 1 1 1.05 1.17 1.24 1.42 1.8 2.12 2.69 3
? . MES PRES 0.6 0.61 0.64 0.67 0.69 0.75 0.66 0.89 0.95 1
} -
'tSSSE 7597.!!0 772!.769 8102.626 64E;.434 8736.722 9496.43710889.2411269.1012028.2212661.91 IN5IDE CW 658.41E6 710.5091778.0617 876.;275 910.4412 999.9481 !!49.990 !!E?.7631254.8361072.744 h
i SUFFICE M2'AL 7220.465 510.E87 795!.750 6664.083 8695.2519096.99410090.29 9825.714 9476.872 6912.52 M !nA5E27.16279 !7.24978 48.5:520 62.59478 7;.995?2 90.56599 !!B.7022 140.1507 172.8682 181.!?5:
- ' 05E 2008.55E 2005.959 2659.E2 3139.542 3412.614 3856.860 4750.51: 5:17.176 5939.294 5663.441 SSE 2S82.976 2 19.S68 23 2.681 4540.564 4944.926 5642.647 6909.854 7605.0 0 6679.970 8 % 4.553
- ??SURE M97.153 7723.769 8103.626 8482.484 87:6.722 9496.43710289.24 !!;69.1012029.2212661.91 OLITSitt CW
-438.945 -473.672 -518.707 -584.218 -606.960 -659.965 -766.660 -793.175 -636.557 -715.162
' SURFICE TIERMt. -481!.64 -5007.25 -5301.16 -5776.05 -5796.B: -6064.66 -6726.85 -650.47 -6317.91 -4602.37 SRIVA5E-18.1085 -24.83:1 -32.!434 -41.9299 -49.206 -60.3773 -79.1548 -93.4338 -115.245 -120.930 03E -!!39.03 -1 2 7.23 -177!.16 -2093.02 -2275.07 -2591.24 -3167.00 -3478.!! -3959.!2 -3775.62 SSE -1921.99 -2213.24 -2559.12 -3027.04 -3296.61 -3761.76 -4606.56 -5070.02 -5786.64 -526.~7 PESS'F1 + Cs
- TE%L + SHRIEf4E E?.SPME 8914.824 9100.210 9566.681 10014.03 10!50.00 11192.68 12782.41 13!29.!7 13846.25 14022.03 PDOINS 658S.372 6882.205 7:15.272 8002. 5 4 8066.407 8481.257 9465.818 9296.!57 9087.147 6905.591 PESStFE+ tw + THERML + CBE + SHRINKAGE ,
E'2PME '249. 5849484. 520 10009. 97 10607. 29 10918. 77 !! E 40. 49 13574.16 13997. 90 148;6.13 14966. 94 (PM+PBl/59 0.545698 0.55?$58 0.59059 0.625799 0.644175 0.692554 0.9008:5 0.82585 0.8752S8 0.08D05
~
PESStFi := + THEP.ML + SSE +5HR!*AM ES/W 9395. 20 9653.52! 10206.4610840.79 !!!?4.15121M.!! !!934.0514395.8815297.91 15407.!
(FmF E I.9 0.554296 0.569:29 0.6021:10.63o574 0.659242 0.715818 0.E22068 0.849!!4 0.902:36 0.902975 i
~
3-17 STRUCTURAL INTEGRITY 1 ASSOCIATES.INC l
l .~
. t .
I '
,~ .
TABLE 3-3
.1 Suar.arv of Acplied Stress at held KR-2-41 ~
~
J.
^
~ BRANCH C ENTE 2M My Mt CASE! (FT-LBS) (FT-LIS)
- p pipt N -33 1 2995 PREES'$E= !!50 PS:
TEWL -6752 42Q81 THX: 0.579 IN OBE-If 6591 4514 OD: 12.7* IN EE-IY 2745 2853 Z= 64.5 INlt! -
SSE-IY E216 5786 SPF 16950 PSI
, CE-ZY 3999 4155 5 ;NKASE 40 19 STFISSES NORR 10 kE5D DLE TO BPJWCH FOOTS (PSil ANSLE (CESKEE)
] 0 10 20 30 40 50 60 70 s 80 90 ste 2AD M0ENT 1 1 1.05 1.17 1. 4 I -
jMS PFIS 0.6 0.6! 0.64 0.67 1.42 1.8 2.12 0.69 :
0.69 0.75 0.86 0.89 0.95 1 _ _
TSSLF.E lEICE N 7597.153 77:!.769 810:.C6 S48!.454 8736.7:2 9496.4!7~10889.24 !! 69.1012008.82126 iT SLTitC TErNL '55.34!8 654.50'.2 770.5049 925.2437 1021.475 !!81.001 1465.863 16:7.049 1901.1:4 185 7829.02; 7928.217 8175.842 B667.624 8438.002 E512.464 9004.221817?.!?B 69S4.846 3768.558 M IMA5E3.534253 4.77:
446 6.160!!8 7.9:5208 9.289353 11. 2162 14.78:10 17.388:7 21.26566 2:.22*.58 1
OBE 12 1370.604 1651.396 1976.!!2 2404.568 2686.362 3140.434 3941.150 4454.024 5:
41.587 5:10.790 1649.488 221'.693 26 40.003 202. 2 1 !566.694 4158.162 *204.206 5864.599 6379.000 6 OUTSIDE "ISS'SE N 7597.150 7723.769 8102.C6 648!.484 8734.722 9496.4:710889.24 !!269.!012028.8
-370. :: -436.350 -513.669 -616.829 -680.983 -787.347 -977.242 -1091.49 -1:67.42 -12!9.
SURFACE TEDX -5219.34 -5:25.47 -5450.56 -5778.41 -56:5.33 -5674.97 -6002.83 -5452.7 SGIRA5E-2.56!S USE -3.18:09 -4.10687 -5.2901: 6.19290 -7.54774 -9.85473 -11.5922 -14. 43
-9: .7 6 -1100.93 -1317.40 -1603.24 -1790.90 -209!.62 -2627.4; -2969.34 -3494.!9 -347 .86 SSE
' -12! .99 -1477.12 -1760.20 -21:4.90 -2377.79 -2772.10 -3469.47 -3909.7; -4586.:0 -4541.:9
~ ~ ~~~ ~
PPi!STE + W + TE7e'A + SHR!*A5E ERBRA'tE BENDINS 8N5.134 915.022 595.711 1008!.61 10314.85 !!!!3.*J 126:6.7412?08.071E13.2713603 6?B9.922 7156.263 7460.422 0000.669 7890.639 8087.'39 8737.472 8194,663 74:2.788 4700. 7:
FFiSSJE+ N + THEFNL + GE +SHRlWr4E T.$ PA E (PM+PBUE5 922 .549 94:0.254 99 7 063 10484.42 10752.57 !!637.:1 !!:93.60 13650.41 14386.97 1 0.!44163 0. 5 6 57 0.585549 0.618 5 0 0.634960 0.686567 0.754 S 0.00'!34 0.848789 0.8 PFISSSE- + THEF.% + SSE +SWRING6E YnERM .
(FMd : S1 9303.!a2 9'24.3041005.7610617.2410909.2911806.9! 1504.11 !!835.51 8.94 14659.92147 i
0.548872 0.56190' 0.592TO 0.6263910.643616 0.696574 0.796702 0.819204 0.669554 0.66489 l
i -
i l
3-18 STRUCTURAL INTEGRITY ASSOCIATES INC
.m
. Y _
1 .
3 I
TABLE 3-4 ALLOWABLE END.0F. EVALUATION PERIOD FLAW DEPTH 1 TO THICKNESS RATIO FOR CIRCUMFERENTIAL FLAWS - NORMAL OPERATING (INCLUDING UPSET AND TEST) CONDITIONS f + P, Rate of Flaw Length, /,. to Poe Carcumference [ Note O))
I. '
[ k e (2)]
J.
0.0 0.5 0.1 0.2 03 0.4 er More 1
.5 (4) (4) (4) (4) (4) (4)
. . '. 4 0.75 0 40 0.21 0.15 (43 (4)
.3 0.75 0.75 0.39 0.27 0.22 0 19
.2 0.75 0.75 0 56 0 40 0.32 0 27 f 1.1 0.75 0.75 0 73 j 1.0 0.75 0.75 0.75 0.51 0 63 0 42 0 51 0 34 0 41 09 0.75 0.75 0.75 0.73 0 59 0 47
.
- 0.8 0.75 0 75 0.75 0.75 0.68
- 0 53 0.7 0.75 0.75 0.75 0 75 0 75 0 58 4 .T 0.6 0 75 0.75 0.75 0.75 0.75 0 63 NC~ 5 .
11: an depth == a.for a surface f'a* -
24 for a sats. arf ace Ea*
! = nominal thickness
. var inteWatson is permiss.bie-e
. (2 ' : enmary membrane strest nmary t'ending stress l - a'loaanie design stress inteas ty On accordance with Section !!D (31 .-f erence cased on nomena's ce o ameter.
(4) f* 3 3514.3 shall be used.
I l
i
! STRUCTUREL 3-19 ,
INTEGRITY ASSOCIATESINC
~
g- .
.I .
TABLE 3-5 F for a Circumferentially Cracked Cylinder in Tension (t/Ri = 1/10),
, a/t 1/8 1/4 1/2 , 3/4
, i F 1.19 1.32 1.82 2.49 l
4 t -
1
[ TABLE 3-6 y
, , h1 for a Circumferentially Cracked Cylinder j
in Tension (t/Ri = 1/10) b.
a/t n=1 n=2 n=3 n=5 n=7 n=10
)
1/8 4.00 5.13 6.09 7.69 9.'09 11.1 1/4 4.17 5.35 6.09 6.93 7.30 7.41 1/2 5.40 5.90 5.63 4.51
! 3.49 2.47
_ 3/4 5.18 3.78 2.57 1.59 1.31 1.10 i l t
3-20 STRUCTURAL INTEGRITY ASSOCIATESINC
---,--+--+m- ,,-----m ,,w'm+ -- w-w- w *w wm"'- t*Y' r w-' 2 e
~
c
. TABLE 3-7 1
F for a Circumferential, Through-Wall Crack
, in a Cylinder of t/R = 1/10 -
in Tension l'
~
l a/b 1/16 1/8 1/4 -
1/2 k
. F 1.077 1.259 1.802 4.208 TABLE 3-8 l h1 for a Circumferential Through-Wall Crack
~
in a Cylinder of t/R = 1/10 in Tension a/b n=1 n=2 n=3 n=5 n=7 1/16 2.979 3.967 4.655 5.576 6.104 1/8 3.221 4.157 4.708 5.163 5.102 1/4 3.677 4.159 4.032 3.238 2.605 1/2 3.091 2.220 1.713 1.137 0.816
~
l l .
7 3-21 STRUCTURAL INTEGRITY AssoCIIIIF. SINC
, TABLE 3-9 Material Stress-Strain Properties of Base and Weldment Materials Used in Analysis -
4 .
i 304 750F 304 5500F TIG 750F TIG 5500F
. .e strain at Pm 0.546* 0.347 0.299 0.103 s ess at Pmax 149,380 88,650 121,890 70,100
" 30.7 17.3 13.63 2.83
~
~
n 1.92 2.49 4.00 11.84 "o 38,200 24,800 68,900 53,900 F ieofFit 0.166-0.888 0.04-0.888 0.114 .299 0.022-0.114 b
YS+ 43,000 24,800 68,900 53,900 TS 86,000 62,600 90,500 63,400 Elong. % 80.3 45 55 O "
28
% RA 81 70 69 69 o
Diametral gage
+ Cross-head measure,ments L o 0.4" Gage length
,E = 30 x 106 pst V = 0.3 i -
3-22 P STRUCTURK
' INTEGRITY ASSOCIATESINC
~
\~ .
~
) .
~
Table 3-10 nree Commonly used Ramberg-Osgood Constants for Weldment Materials ,
ro a n Primary Curve
, . 53.9 - 2.83 11.84 3
Alternate Curve A 44.8 3.39- 6.89
. Alternate Curve B 49.4 9.0 9.8 6
e m
em i
. l N
1 i
p .
l l
~
3-23 STRUCTURAL
, INTEGRITY i ASSOLMINC l
- , ,, , - re
~
'j .
~
TABLE 3-11
~
~
WELDING PROCESSES
}, _
FOR STAINLESS STEEL PIPE A
it SUBMERGED ARC WELD (SAW)
. AUTOMATIC PROCESS
_ . ARC BETWEEN BARE METAL CONSLNABLE .
ELECTRODE (WIRE)ANDWORKPIECE
. ARC SHIELDED BY GRANULAR AND FUSIBLE .
FLUX WHICH BLANKETS MOLTEN WELD METAL
. HIGH WELD DEPOSITION RATE AND SPEED l . DISADVANTAGES ~
J -
. . SLAG MUST BE REMOVED AFTER EACH l PASS TO AVOID ENTRAPMENT IN WELD I METAL J . HIGH HEAT INPUT CAN GIVE SLOW
} . COOLING RATES AND C0 ARSE, LOW TOUGHNESS MICROSTRUCTURE b
PICKUP FROM THE FLUX CAN CHANGE COMPOSITION OF DEPOSIT
. RISK OF MICROFISSURING USED FOR MOST SHOP WELDS - NOT IN FIELD RELATION TO WELD METAL TOUGHNESS RELATIVELY HEAVY SLAG / INCLUSION CONTENT C0 ARSE MICROSTRUCTURE HIGH HEAT INPUT CAN GIVE HIGHER FERRITE CONTENTS THE AB0VE CAN LEAD TO REDUCED TOUGHNESS - PROBABLY THE LOWEST FOR THE WELD PROCESSES CONSIDERED HERE l
3-24
^ STRUCTURAL INTEGRITY ASSOCIATESINC-
~
~
TABLE 3-11 (Continued)
~
_. SHIELDED METAL ARC WELD (SMAW)
I ,
. MANUAL PROCESS
~
. ARC BETWEEN FLUX-C0VERED CONSUMABLE ELECTRODE AND WORKPIECE
. SHIELDING BY GASE0US SHIELD AND MOLTEN FLUX OR SLAG
~
FROM ELECTRODE COVERING
. MOST VERSATILE PROCESS - POSITIONS, ETC.
. DISADVANTAGES
. SLAG BLANKET - SOURCE OF INCLUSIONS I
_ . VISIBILITY IMPAIRED BY SLAG SLAG REMOVAL BETWEEN PASSES IS NECESSARY l . MOISTURE PICKUP IN ELECTRODES
. LOW DEPOSITION EFFICIENCY j ,
USED FOR REPAIRS AND FOR CERTAIN PORTIONS OF FIELD AND SHOP WELDS --
RELATION TO WELD METAL TOUGHNESS INTERMEDIATE TO HEAVY INCLUSION
[ CONTENT INTERMEDIATE HEAT INPUT AND DILUTION EXPECT INTERMEDIATE TOUGHNESS I
i STRUCTURAL
' 3-25 INTEGRITY ASSOCIATESINC
~
, TABLE 3-11
~ (Concluded)
~
GAS TUNGSTEN ARC WELD (GTAW), OR TUNGSTEN INERT GAS (TIG)
AUTOMATIC OR MANUAL PROCESS ARC SETWEEN NONCONSINABLE ELECTRODE (TUNGSTEN) AND
[ WORKPIECE - FILLER METAL (WELD WIRE) CAN BE ADDED T0 i
WELD POOL - SHIELDED BY INERT GAS (ARGON OR HELIUM) r .
MULTI-POSITION, HIGH QUALITY WELD, BUT LOW DEPOSITION 1
=
RATES
. NO FLUX USED - NO SLAG -
3 .
INSIGNIFICANT CHANGES IN FILLER COMPOSITION DURING DEPOSIT - LOW PICKUP OF CONTAMINANTS USED MOSTLY FOR FIELD WELDS, SOME SHOP WELDS, AND ALL WELD OVERLAYS t
i RELATION TO WELD METAL TOUGHNESS
' REDUCED HEAT INPUT THROUGH PULSING GIVES FINER,
{ .
1 TOUGHER MICROSTRUCTURE AND POTENTIALLY LOWER FERRITE NO SLAG-METAL REACTIONS At[5 RESULTANT NONMETALLIC s.
INCLUSIONS
. INSIGNIFICANT PICKUP OF CONTAMINANTS THE ABCVE CAN LEAD TO THE HIGHEST TOUGHNESS FOR THE WELD PROCESSES CONSIDERED HERE 1
l
~
I_
-l .
i -
3 3-26 STRUCTURAL INTEGRITY ASSCCIATESINC
. T '
I -
~
~
J.
e a
'O COM PUTED 6"--* If P1PE %
, STRESSES 0.7, H 2* -- '
MEASURED STRESSES I -- -- / WELD o O* AZIMUTH o + C J . L 0 90* AZIMUTH T NA2lMUTH PE
~
20 -
- R = 6.375" 10 0 1~'
Cm 3 2 -1 1 2 3
[* O , .' ,
, O
. -I .0 -0.5 O O.5 1.0 I $ DISTANCE FROM WELD CENTERLINE -
10 0
{ I"*"
2 t-
-20 -
o .
-200 o Q O U a g-40 -
g . 3co g -o e O o O
1 l..o _
0 4*
- V -
-50 0 J. . .go . FUStoN LINES W -
-600 PIPE S10E SWEEPOLET S10E Figure 3-1. CoTparison of Measured and Computed Residual Asial Stresses o Along Inner Surface of a Welded, lHS! Treated I?-Inch Sweepolet l
3-27
^ STRUCTURAL INTEGRITY ASSOCIATESINC
' ~ *~ ,
F 8 8 i i -
- y-- **--- * - - - . - % % .3 m,
, , t (15 24cm) -
- 6 in -
WELDS AXISYMMETRIC 2 in(5 08cml-e i
b SAVFEM SHELL h 8
~
o 7 INTEGRATION POINTS o 5 INTEGRATION POtHTS !
o 3 INTEGRATION POINTS 0.7in(178cm)
-R = 6.375 in(6619cm) a MPo MPo
-300 -200 .aco o too 200 300 -300 -200 -100 0 10 0 200 300 8 ' OUTER SURFACE g
ro O n-- g 0.e -- _ ,o q
as
--Z D =-- C
'" ~ * ~ '
c ,, 06- 06--
N scL
~~ ~'
i2 -
82 0.4 - 2 04-om Os _
_ _u .
_ oe o gg : .
a 0 02--
04 -
02-- ( 'd -
04 8 --Ea 8 -- .
0, ^
- , , , INNER SURFACE i o i , ,
o
-40 -20 0 20 40 -40 -20 0 20 40 REstDUAL AXIAL STRESS, nel RLSIOUAL CIRCUMFERENTIAL STRESS, kol l
t s .
Figure 3-3. Cornputed Residual Stresses at Weld Centerline for IH51 Treatment of a 12-Inch SweepOlet
>'* ft)
EW3 53to
, MC c- O n
- 4 9 84 -
3 (3 p .
t .
,- , y - - - W 9 **9 ~1 M ""' l i
t .
- i (15 24cml 6m -
2 in(S 08 cm)- - -
WELDS AxlSYMMETRIC l O 75 en ti 91cm)+ +
SAVFEM SHELL ---
o 7 INTEGRATION POINTS
~ !
a 5 INTEGRATION POINTS a 3 INTEGRATION POINTS 0 7in(1.78 cm)
-R =6.375 in(1619cm)
MPa 300 -200 -eno o 40 0 200 300 MPo
-300 -200 -100 o t00 200 300
'8 ' 3 OUTER SURFACE ' '
N o 6--
a O a a 05-- % '
gg .
"' W O
o E o ,
o et E (0 3* 4:
"Os g
.'.mt .
i nm E C5 w
os "
0 og - a 02--y 0 gg o 02-- e
~-d -
04 8 "
g, 8 ~~
-40
-20
, , INNER SURFACE , .
p, 0 20 40
-40 -20 0 20 40 RESIDUAL AXtAL STRESS. hel RESIDUAL CIRCUMFERENTIAL STRESS, kol we Figure 3-4.
Computed Residual Stresses 0.75 Inch (1.9 cm) from Weld Centerline. 0.25 Inch 2>m3 (0.134 r.m) from Coil Lenterline, for tit.t treatment of a 12-inch Sweepolet
, . >3td Md n
>3 M! -
~. .
.j -
40 C o _ 50 6 900 30 ~
o 0 450 j_ 4g c:
00 -
20 " D450 900
[
- 3U
~
A
/ // 20 l j/ _
lo _
p IU i / // -
y c ,
'2
, 4 '/
' */t E
-_ g . .
. 6_ '. s 1 $
h 3 /
I p y/ / _
-10
$-10L / /
c/ / l _-20
//
-20 -
/ _.-30
/,
-30 50
-40 10 00 Figure 3-5. Post-lHS1 Residual Stress Distribution STRUCTURAL 3-31 INTEGRITY
. ASSOCIA'IFSINC f
g l
l I
)
D D I D O D I O T
T l
A T A T r A e S A t S nl ,
S S S S E S E S C'c" R E R E t' d
9 T S
R T
T S
R T
l L
S L
S M n
, A P A P r r
I O I O r 3
X O X O f ,
A H A H
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- n q i o
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i i
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"t e
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l a
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u
,. .*-e 0 I i y O ? 5 5
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i C -
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(
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A o, "* -
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'4 0
0
\
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1
,
- 7
3
' g, g e N- r
- u
'o i F
g 0
~0-e .
3
- ~
[ - - - -
0 0 0 o 4 3 2 2 0 0 a 0 0 1 3 4
- - 2_ - -
- * , ;n a:.
. Sgu g .
g yw . M9mcf 3
- '2aMO k ;
3g
- t c A
.D*
i -
)
e -4 ,3 i 8 ' i o . 4, . . g ,2 ,
1 3 -
4 i R a F1
, k
- z. -
r2
- ab 5 r
13 8
- -t --+-
0 #4 l l -
c
. c 1 A "i */ **
't % * '2 'h8 ^i
- t 's hl $ .4(P
)
h, e
o .
0 .2 .4 , i: ,5 Fra c ie:al Dia:a.,e e T.,;c:gh ,.*a:1 (a /t Figure 3-8. Magnification Factors of Circ:derential Crack in a Cylinder (a/t = 0.1) l 1
3-34 STEUCTURAL I
INTEGRITY ASSDC!rbT. SINC
J
- 3 30-3 -
I
~
pUpper Bound (Furnace Sensitizi d)
, y da/dt = 5.65x10-9(K)3 07
' 10*4
- a g @ Best Estimate (Weld Sensitize <)
,A j da/dt = 2.27x10-8's)2.26
,4 , %....a.,,... , . . .
Sy latafesP:4 ect . T ets *e
/ /
/ / d 8oth s 'stt 3 a' s t u'.s 3, t y .
g sata' cas: ct ties -
30-$ -
[M7 D 8f 6 52 8 3 W' 2. =
8 2 a.r= Og:Gt *'."T e s s's e
[Y
-- "- T -Q $4=litsFtDlf=IatL' C2ap C 3 g 8G4 8P1332 ; ag s m ile h T O sei sitize o at iisc*e. a. e s.-
1 y v S O ca .T,ist
$ ma=C guant.ct%to
@ SotCw3*. Cit *D .
10-6 -
g asasacs.. wita C=. a g i
,, - st anc= Las
,, - @ tant . anCD%=t osa? Las engs a Me
- O stest.taf f 0 Sv witDihC LT8 A 337'f.7e 6 8 asa 03t&a ag a s 3,Ts 3o.7 e i i e e i 0 30 E N 40 so 6o 70
_ Figure 3-9. Stress Corrosion Crack Growth Data for Sensitized Stainless Steel in BWR Environment (Ref. 7 )
3-35 STRUCTURAL
- - INTEGRITY ASSOCIATESINC r-,, ..-v. -e- - , , n - --.
! i i i :
1 i - -
- .
- i - -- ;
% 9 M 9 i ~i I
1.0 - i .
.9 -
.e _ ,
.7 ..
x 3 l U3 .6 -
h Th up .5 -
I)
LC 4
- ., 0 e
w w" .
d, 3. " .3 -
2
- .2 - O' t
.1 - '
t I I I I l I I 2 3 4 5 6 1 Number of Operatson Personis Inr Intled in the Average Crack Growth Rats t.nlaulation I I I I l 1
I i
?,0(10 4,000 6,000 11,000 10,000
- TJ) 12,000 14,000 H
D3 Approximate Time Between inspection (Hours)
! O Figure 3-11. Average Effective Circumferential Crack Growth Rate As a function h of Operation Periods Used in Calculation of Time Between Inspections EM -
7 .
J J
l INS"ASIUTY r I lr - uarut
- > var (SLCPE ,
i I J,g f
\
l APPUED
< l aa =
r .. ,: i og
)
Figure 3-12. Tearing Modulus Concept for Stable
. Crack Growth e
I, 3-38 ^ STRUCTURAL INTEGRITY !
ASSOCIATESINC L -
l
qD ado d Dooh R;
=
Ro -
k I -
I
_c lC_
l I
d yp o y 8 yyyp Figure 3-13. Circumferential1y Cracked Cylinder in Tension i
1-i l 3 39 STRUCTURAL t INTEGRITY ASSOCIATESINC g s',
- . -_ ~.- .~1- -,--__ ,._. - - _ _ _ . , _ _ , _ . _ . . , _ . - . _
y , _4_._ _ o . . _ . _ . - _ , . . _ . , . , . _ , _ _ . , , ,
s CB o . - .
7 O
e 1
s[,
/ -
. v.,, .
u._.-
! . {i ,
t I
2 F
i h
Figure 3-14.
Through-k'all flawed Cy1'inder Under Remote Tension I
i-
-i -
- STRUCTURAL
, 34O INTEGRITY ASSOCIATESINC
-- -- . - - - , ,.- - -- - . . ,-----n-- .
n -- - a - - - - . , . . - . , . , , , - - , - - - - - - - - ~ . . - , --,,.-r- , , , - - , ,
1 i
\; .
\ .
s
\ . - -
t
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)
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n J e io '
9 i '
H c n
s U [j
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=
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- i. -
c .s
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r .
U b i . y q i -
' m v t . \ 20 "
a t
\ e Cs N
E g 1 w . 2 ai C C j * \ -
. g
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-- Z
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e, ,
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i OO 6 i i .
. 1
. o O O O O O O to O 4 6 N Om O ON O@ Oc O4 O6 O CO N O -
O
('s yT s p' 'a .- g
' ^ STRUCTURAL
- 3 41 INTEGRITY ASSOCIATESINC
, -- - ,- m ,- .-- ._
.-y .__.,
m
. . _ - , -- , m --*1 i 1 i .
JT DATA BASE 20 m - - - - - - . - -- - - --
19 - -
f- --- -
c-------------------------------
/
I 18 -Y k - -- ,
, y - - - - -- - -----
17 --
J-f-
/ './ ./
,/-b . '- - --- - -
16 --
f ~ - --
- -. ,/- - - - - - - - -
15 --
4- - - ---------
14 - -- -
~- ------ --
/ - - - -
'g-13 -
f- -
1*- --
0
- s. 33 _ __ / _ , _ _ _- ____.
?>
a a
.- jo. ___.__'1- .
~ .e ,~ ,
2 9_ ____ ,/_ _,<_ _
$ 8- - V--/-4-- _____. _ _ _ _ _ _ _ _ _ _ _ _ _ _
\
, 7 -
,r - y'- - - - - - - - - - - - - - ---
-'\
6-
= -
V- - -- -- ---- - - - - - --
~ -
5- ----- -
4-- - - - - - - - - - - -
i 3-- ------
- - ---- --i ~N 2 _ ._ ______ f
/ - - - - - - - -
- ~ ~ - - - - - - -
- m. . . _ --
3__ _ __-. _________. . . . . _ _ _ _ _ _
,y/
__-/ W. . , .. y ~
_-,,__,__.,7 , , - _ _ - - -
0 - -
wtn 28 0 --------------- - - - - -
HM 200 -
400 Mc '
600 g00 T b>3 Figure 3-16.
- Compilation (from data ofof Material Toughness J-T Curves Refs. 17 to 21) ,
y -
I i
_ _ _ . w .
=
sI__ - _ - _ -
i -
- = _ _- - _ - . __ w_ .__ 0
_ - _ iw!.
- 0 m.
- .s _
6 G
- l T e
I e m y___ ___
- t o S r s
f sl e a
- = - _
- l t n e q.,s. - -
_ _ - _ - - ia M
- - t d S l t
S-t W h
e
_s . D- _-
_ _ _ . - . gl N- -
0 u e o e U_ - - -
-s_ _- . .
l 0 r t W S s O- _ . 4
.s B- -
_% _=_
r s o s s se
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' Figure 3-20. Material J-R Curve Derived from Lower Bound J-T Diagram for SAW/SMAW Weldment Material ,
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', for SAW/SMAW Weldment Material (Expanded Scale)
1
~
,- 4.0 EhlUt.TIONSANDRESULTS 4.1 Weld KR-2-14 1
Input to the flaw evaluation for this weld was as'follows:
o l
Indication t.ength = 2.1 inches Indication Depth = 0.1418 inch '
Pipe 0.D. = 12.75 inches /22 inches '
(Riser circumference used in normalizing crack length for conservatism)
Pipe Wall Thickness = 1.125 inches
, Applied Stresses (Table 3-1) l Pressure + DW + Thermal + Shrinkage = 14.80 ksi membrane,13.88 ksi bending Pressure + DW + Thermal + Shrinkage + Seismic = 15.38 ksi membrane Residual Stresses (Figure 3-5) ..
I i
Figure 4-1 provides applied stress intensity factor versus crack depth data e
for the two load cases used in the evaluation (piping loads and IHS! residual stress). Assuming the indication to be IGSCC, these stress intensity curves were used to perform post-IHSI IGSCC crack growth estimates and the resulting crack growth predictions are illustrated in Figure 4-2. The analysis results in no predicted crack growth for the balance of plant life.
~
The allowable end-of-cycle flaw size was determined in accordance with ASME Section XI, Article IWB-3640, and using the J-T procedure described in
- Section 3.5. The results are illustrated in Figure 4-3 in terms of allowable l flaw depth versus length.
Note that, although not required by IWB-3640, thermal expansion stresses have been included in the evaluation to account for the possible effects of low toughness weldment material. Also, in
![ accordance with the recommendations of NRC Generic Letter 84-11, a maximum .
)
allowable flaw size of 2/3 of the IWB-3640 limit (shown as a dashed line in
-. figure 4-3) is used to allow for uncertainty in flaw depth sizing. .
. l 4-1 ^ STRUCTURAL
!. ' INTEGRITY ASSOCIMESINC
j ..
Also shown in Figure 4-3 are allowable flaw size curves calculated by elastic-plastic fracture mechanics (EPFM) for the three different sets of -
Ramberg-Osgood constants of Table 3-10.
It is seen that the EPFM results yield somewhat more conservative allowable flaw size, but compare favorably
.. to the 2/3 of IWB-3640 limit.
Referring to Figure 4-3,
~
it is seen that the 2/3 of IWB-3640 limit is satisfied indefinitely by the analysis, since 'no crack propagation is predicted.
To add further assurance, the IGSCC crack growth analysis has been repeated assuming various initial flaw sizes ranging upward from the observed UT depth. No crack propagation is predicted in the post-IHSI condition for initial crack depths up to 0.414 inch, or 37% of the pipe wall.
It is also noteworthy that, given the relatively short length of the observed indication (5.3% of circumference), it would not lead to rupture of the pipe joint even if the above crack growth or initial flaw size estimates are significantly in error. Leak-before-break is clearly the appropriate hypothetical failure mode for this indication.
} On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication and the IHS1 treatment which has been applied, will not lead to a reduction in plant safety margins, or a plant operational concern.
4.2 Weld KR-2-36 Input to the flaw evaluation for this weld was as follows:
Indication Length = 2.2 inches Indication Depth = 0.2813 inch
^
Pipe 0.D. = 12.75 inches /22 inches (Riser circumference used in normalizing crack length for conservatism) -
Pipe Wall Thickness = 1.125 inches '
l 4-2 STRUCTURAL INTEGRITY
. ASSOCIATESINC g
Ahl'iedStresses(Table 3-2)
Pressure + DW + Thermal + Shrinkage = 13.85 ksi membrane, 9.09 ksi bending '
f Pressure + DW + Thermal + Shrinkage + Seismic = 14.97 ksi membrane Residual Stresses (Figure 3-5)
Figure 4-4 provides applied stress intensity factor versus crack depth data for the two load cases used in the evaluation (piping loads and IHSI residual stress). Assuming the indication to be IGSCC, these stress intensity curves were used to perform post-IHSI IGSCC crack growth estimates and the resultirng crack growth predictions are illustrated in Figure 4-5. The analysis results
, in no predicted crack growth for the balance of plant life.
The allowable end-of-cycle flaw size was determined in accordance with ASME Section XI, Article IWB-3640, and using the J-T precedure described in Section 3.5. The results are illustrated in Figure 4-6 in terms of allowable A flaw depth versus length. Note that, although not required by IWB-3640, thermal expansion stresses have been included in the evaluation to account
.h' for the possible effects of low toughness weldment material. Also, in accordance with the recomendations of NRC Generic Letter 84-11, a maximum allowable flaw size of 2/3 of the IWB-3640 limit (shown as a dashed line in Figure 4-6) is used to allow for uncertainty in flaw depth sizing.
Also shown in Figure 4-6 are allowable flaw size curves calculated hy-elastic-plastic fracture mechanics (EPFM) for the three different sets of Ramberg-Osgood constants of Table 3-10. It is seen that the EPFM results yield somewhat more conservative allowable flaw sizes, but compare favorably to the 2/3 of IWB-3640 limit.
Referring to Figure 4-6, it is seen that the 2/3 of IWB-3640 limit is
~
prcdicted to be satisfied indefinitely by the analysis, since no crack propagation is predicted. To add f urther assurance, the IGSCC crack growth 1 analysis has been repeated assuming various initial flaw sizes ranging upward from the observed UT depth. No crack propagation is predicted in the I post-IHS! condition for initial crack depths up to 0.612 inch, or 54% of the STRUCTUHAL 4-3 INTEGRITY i
ASSOCIATESINC l
pipe wall.
It is also noteworthy that, given the relatively short length of l the observed indication (5.5% of circumference), it would not lead to rupture of the pipe joint even if the above crack growth or initial flaw size estimates.-are significantly in error. Leak-oefore-break is clearly the appropriate hypothetical f ailure mode for this indication.
On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication and the IHS1 treatment which has been applied, will not lead to a reduction in plant safety margins, or a plant operational concern.
4.3 Weld KR-2-41 -
Input to the flaw evaluation for this weld was as follows:
Indication Length = 4 inches Indication Depth = 0.2138 inch
~
~
~
Pipe 0.D. = 12.75 inches /22 inches (Riser circumference used iri normalizing crack length for conservatism)
Pipe Wall Thickness = 1.125 inches Applied Stresses (Table 3-3)
Pressure + DW + Thermal + Shrinkage = 12.64 ksi membrane, 8.74 ksi bendin Pressure + DW + Thermal + Shrinkage + Seismic = 14.47 ksi membrane
' Residual Stresses Zero and Post-lHSI (Figure 3-5)
~ ~
Ficure 4-7 provides applied stress intensity factor versus crack depth data
~ for the two load cases used in the evaluation (piping loads and IHSI residual stress). Assuming the indication to be IGSCC, these stress intensity curves were used to perform post-IHSI IGSCC crack growth estimates and the resulting crack growth predictions are illustrated in Figure 4-8. The analysis results in no predicted crack growth for the balance of plant life.
I t
4-4 STRUCTURAL INTEGRITY ASSOCIATESINC
~
The allowable end-of-cycle flaw size was determined in accordance with ASME Section XI, Article IWB-3640 and using the J-T procedure described in Section 3.5. The results are illustrated in Figure 4-6 in terms of allowable flaw ,
depth versus-length. Note that, although not required by IWB-3640, thermal expansion stresses have been included in the evaluation to account for the a
I possible effects of low toughness weld material. Also, in accordance with the recommendations of NRC Generic Letter 84-11, a maximum allowable crack size of 2/3 of the IWB-3640 limit is used to allow for uncertainty in crack depth sizing.
Also shown in Figure 4-9 are allowable flaw size curves calculated by elastic-plastic fracture mechanics (EPFM), for the three different sets of
. Ramberg-Osgood constants of Table 3-10.
It is seen that the EPFM results yield somewhat more conservative allowable flaw sizes, but compare favorably to the 2/3 of IWB-3640 limit.
g Referring to Figure 4-9, it is seen that the flaw is predicted to remain at x 3 its present size indefinitely, and thus satisfy the allowable flaw size limit by a large margin for the balance of plant life. To add further assurance, the IGSCC crack growth analysis has been repeated assuming various initial flaw sizes ranging upward from the observed UT depth. No crack propagation is predicted in the post-IHS1 condition for initial crack depths up to 0.6S4
, inch or 61% of the pipe wall.
'~
On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication, will not lead to a reduction in plant safety margins, or a plant operational concern.
4.4 Weld KR-2-37 Input to the flaw evaluation for this weld was as follows:
~
Indication Length = 5 inches Indication Depth = 0.135 inch 4-5 STRUCTURAL INTEGRITY ASSOCIATESINC w a
Pipe 0.D. = 22 inches Pipe I.D. = 19.75 inches
_ Pipe Wall Thickness = 1.125 inches Applied Stresses -
Pressure = 5.622 ksi Residual Stresses (Figure 3-7) '
Figure 4-10 provides applied stress intensity factor versus crack depth data i
for the two load cases used in the evaluation (pressure and IHSI residual stresses). Assuming the indication to be IGSCC, these stress intensity curves were used to perform IGSCC crack growth, estimates for both cases, and the resulting crack growth predictions are illustrated in Figure 4-1. The analysis case results in no predicted crack growth for the balance of plant life.
The allowable end-of-life flaw size was determined in accordance with ASME g Section XI, Article IWB-3640, and using the J-T procedure described in Section 3.5. The results are illustrated in Figure 4-12 in terms of allowable flaw depth versus length. Also, in accordance with the recom-mendations of NRC Generic Letter 84-11, a maximum allowable crack size of 2/3 of the IWB-3640 limit is used to allow for uncertainty in crack depth sizing.
Also shown in Figure 4-12 are allowable flaw size curves calculated by elastic-plastic fracture mechanics (EPFM) for the three different sets of Ramberg-Osgood constants of Table 3-10.
It is seen that the EPFM results yield less conservative allowable flaw sizes in this weld.
Referring to Figure 4-12, it is seen that the flaw is predicted to remain at its present size indefinitely, and thus satisfy the allowable flaw size limit
' by a large margin for the balance of plant life. To add further assurance,
~ the IGSCC crack growth analysis has been repeated assuming various initial flaw sizes ranging upward from the observed UT depth. No crack propagation j
is predicted in the post-IHS! condition for crack depths up to 0.81 inches, or 72% of the pipe wall. ,
I
' 4-6 STRUCTURAL INTEGRITY -
ASSOCIATESINC
t On the basis of the above evaluation, it is concluded that continued operation of the plant with this weld, considering the observed indication, ,
[
,- will not lead to a reduction in plant safety maagins, or a plant operational ;
j concern.- --
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, ga lELDHH-2-14 sm Mc O Figure 4-1. Stress Intensity Factor Versus Crack Depth for Weld KR-2-14 Ng -
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/ H0llRS IIHE b
lg$pgg gg
>9 HELDHR-2-14,POSMHS1
- M Figure 4-2. Predicted Stress Corrosion Crack Growth for Observed Ultrasonic Flaw Indication - ,
Wold kR-7-14
~
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)
~
- ^
1.0 IWB-3640 f _ _ _ _ _ 2/3 of IWB-3640 g
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\ n EPFM (Alternate Curve A) 1 75% ,\ "
EPFM (Alternate Curve B)
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initial Flaw Size
- O & Post-lHS! Flaw Growth Prediction 0 t i- 0 I ' '
0.2 0.4 0.6 0.8 1.0 f/ Circumference Fiqure 4-3.
' Comparison of Predicted Crack Growth with l Allowable Flaw Size Limits - Weld KR-2-14
- _ s
~
4-10 QSTRUCTURAL JETENTY ASSOCIATESINC
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' 4-12 INTEGRITY ASSOCIATESINC
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en 1.0
. IWB-3640 .
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, EPFM (Primary cr-c Curve) 0.8 ~ i i "
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_ 0 0.2 0.4 0.6 0.8 1.0 f/ Circumference 1
g- Figure 4-6. Comparison of Predicted Crack Growth with 1 4 Allowable flaw Size Limits - Weld KR-2-36 I
l 4-13 STRUCTURAL j
' INTEGRITY ASSOCIATESINC
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Q Figure 4-7. Stress Intensity Factor Versus Crack Depth for Weld KR-7-41
0.214 1:POSTIHSI .
C : : : : :
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u s1 IlELDHR-2-41,POSI-IHSI
!. EM -
! O E I Figure 4-8. Predicted Stress Corrosion Crack Growth for Observed Ultransonic Flaw Indication -
F
- 1.0 O
g -IWB-3640 1
2/3 of IW8-3640 3 EPFM (Primary e-c Curve)
. 0.8 -
75% \ '\
g O EPFM (Alternate Curve A)
\ s v EPFM (Alternate Curve B) i, 0.6 _
s
, i N
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's a
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9 5____.-.O v
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g Initial Flaw Size
& Post-IHSI Flaw Growth Prediction l?
j- -
0 i ; i i 0
0.2 0.4 0.6 0.8 1.0 f/ Circumf erence i l
Figure 4-9. Comparison of Predicted Crack Growth with Alloweble Flaw Size Limits - Weld KR-2-41 e 4-16 STRUCTURAL INTEGRITY ASSOCIATESINC
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INCHES hg C C"3 CRACHDEPfil
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llELDKR-2-37 N'7< Figure 4-10. Stress intensity Factor Versins Crack Denth for LJolrf KR-7 ~47
O
- 4. ...................................................
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, 4-18 INTEGRITY ASSOCUEEE 1
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& Post-IHSI Flaw Grcwth Prediction O
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. f/ Circumference t
l' figure 4-12. Comparison of Predicted Crack Growth with Allowable Flaw Size Limits - Weld KR-2-37 ,
I i~
4-19 STRUCTURAL
- MTEGMTY
_ ASSOCIATESIhC
5.0 D'ISCUSSION AND CONCLUSIONS
- This report presents fracture mechanics flaw e.;aluations for four welds in '
the Browns _Eerry_ Unit 2 recirculation piping system (three sweep-o-let to ring header welds, and one ring header to end cap. weld).
The four welds contained relatively small, crack-like indications. These welds, along with the other, uncracked welds in the plant, were treated by Induction Heating Stress Improvement (IHSI) to produce a favorable residual stress pattern and thus reduce their susceptability to IGSCC degradation.
The flaw evaluations were based on the post-IHSI indication sizes, which differed somewhat from the pre-IHSI inspections, but not significantly.
The evaluations presented in this report were performed in accordance with ASME Section XI, IWB-3640 and the recommendations of NRC Generic Letter 84-11.
These conventional approaches were also supplemented by Elastic Plastic Fracture Mechanics Tearing Instability analyses to account for the possible
,, effects of low toughness weld metal. The results of the analyses for all four
- ~
\
welds' indicate that design basis safety margiris are maintained in the welds, by a large margin, considering the worst case effects of the observed flaws; and that these margins are maintained indefinitely during the life of the plant, due to the beneficial effects of the IHSI treatment, which is expected "to inhibit further IGSCC propagation. It is also noteworthy that all of the indications had circumferential lengths less than 10% of pipe circumference. l Thus, even in the event cf large uncertaintics in UT depth sizing or crack growth predictions, the governing failure mode would still be leak-before-break.
~
On the basis of these factors, it is concluded that the inspection results j and corrective actions taken should not result in any reduction in design basis safety margins or increase in the probability of a pipe rupture at the I plant.
One final point of significance is that the IHSI treatments, which were '
l j.O performed on a large percentage of the remaining uncracked welds, should I
I
. 5-1 STRUCTURAL INTEGMTY ASSOCIATESINC
- x. .
O
~
4 greatiyreducetheprobabilityoffutureIGSCCinthesewelds. Thus, it is
.; )I reasonable to expect that the plant will operate for a long period of time
- with no further degradation due to IGSCC, and mo reduction in leak-before-break margins relative to plants with piping not susceptable to IGSCC.
m 4
e 1
l e
?.
l l
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e g I_ .
e T '
5- 2 STRUCTURAL
- INTEGRITY ASSOCIATESINC
. . - - - - -- , ,- - -..,e. . -,. , -- % __e -y-- , , . . - , - , , , .
. ~ . ,
6.0 REFEREriCES 1.
Transmittals from E. Wilson, TVA, Jan. 29, 1985 and May 3, 1985.
2.
SI Report, " Design Report for Recirculation. Piping Sweep-o-lets Repair and Flaw Et aluation, Browns Ferry Nuclear Power Plant, Unit 1", SIR 006, Sept. 1984.
3.
EPRI Report NP-2662-LD, " Computational Residual Stress Analysis for Induction Heating of Welded BWR Pipes", December 1982.
4.
EPRI Report, NP-81-4-LD, " Residual Stress Improvement by Means of Induction Heating", March 1981. -
5.
BWROG IGSCC Research Program Status Report presented by T. Umemoto and A. Tanaka, " Application of Induction heating Stress Improvement to Pipe Branches", December 9, 1980.
- 6. Buchalet, C.B., and Bamford, W. H.,
" ASTM 8th National Symposium on Fracture Mechanics,1974", ASTM STP-590, pp. 385-402,1975.
7.
NUREG 1061, " Investigation and Evaluation of Stress Corrosion Cracking in Piping of Boiling Water Reactor Plants", U.S. Nuclear Regulatory Commission, March, 1984.
s<
- 8. " Guidelines for Flaw Evaluation and R' emedial Actions for Stainless Steel Piping Susceptible to IGSCC", Final Report for EPRI Project T303-1, Report No. SIR-84-005, April 13, 1984.
- 9. Bickford, R. L., et al, " Nondestructive Evaluation Instrument Sur-veillance Test on 26-Inch Pipe", EPRI NP-3393, January,1984.
- 10. Ranganath, S., and Norris, D.M.,
" Evaluation Procedure and Acceptance Criteria for Flaws in Austenitic Steel Piping", Draf t No. 10, Sub-committee 1983. on Piping, Pumps, and Values of the PVRC of the WRC, July f 11. Ranganath, S., Mehta, H. S., and Norris, D.M.,
" Structural Evaluation
', of Flaws in Power Plant Piping", ASME PVP-Vol. 94, Circumferential Cracks in Pressure Vessels and Piping - Vol. I, pp. 91-116, 1984.
12. ASME Boiler and Pressure Vessel Code, Section XI, 1983.
- 13. ASME Section XI Meeting Minutes, May 25, 1984.
- 14. Kumar, V., et al.,
"An Engineering Approach for Elastic-Plastic Fracture Analysis", EPRI NP-1931, July, 1981.
- 15. Kumar, V.,
EPRI NP-3607, et Aug., al., " 1984. Advances in Elastic-Plastic Fracture Analysis",
- 16. Hutchinson, J. W., and Paris, C.,
P. " Stability Analysis of J-Controlled Crack Growth", in Elastic-Plastic Fracture. ASTM 668, American Society for Testing and Materials,1979, pp. 37-64. STRUCTUIULL 6-1 INTEGRITY ASSOCIATEStINC v- ~
~
17. Westinghouse Test Data, presented by J. Landes at the meeting of ASME Boiler & Pressure Vessel Code Section XI, Task Group on Piping Flaw Evaluation, San Antonio, Texas, April 23, 1984. , r
- 18. Gudas,-J.P., and Anderson, D. R., "J 1-R Curve Characteristics of Piping Material and Welds", NSRDC, presented at U,5. NRC 9th Water Reactor Safety Research Infr,rmation Meeting, Washington, D.C., Oct. 29, 1981.
19. NSRDC Test Data, presented by M. Vassileros at the meetir.g of ASME Boiler & Pressure Vessel Code Section XI, Task Group on Piping Flaw Evaluation, San Antonio, Texas, April 23,1984. I
- 20. Paris, P.C., Brunetti, J. V., and Cotter, K. H., "The Effect of Large Crack Extension on the Tearing Resistance of Stainless Stael Piping Materials", Presented at the CSNI Specialist Meeting on " Leak-Before-Break in Nuclear Reactor Piping Systems", Sept. 1-2, 1983, Monterey, CA. ,
. 21. McCabe, D. E.,
Weldment Tests,Westinghouse Aug. 29, 1984. letter to J. F. Copeland, Stainless Pipe
- 22. Metals Handbook - Ninth Edition, Volume 6 - Welding, Brazing, and Soldering, American Society for Metals, Metals Park, Ohio, c. 1983.
- 23. Tetelman, A.S., and McEvily, Jr., A. J.,
,; Fracture of Structural Materials, John Wiley & Sons, Inc., New York, c.19o1, pp. 212-222.
s 24 Landes, J. D., 1 et al, " Elastic-Plastic Methodology to Establish R-Curves and Instability Criteria", Sixty Semi-annual Report, Jan. 1, 1982 to June 30, 1982, EPRI Contrac t No. RP 1238-2, Aug. 4,1932. I I
- 25. Ernst, H. A.,
" Material Resistance and Instability Beyond J Controlled Crack Growth", presented at the Second International Synposium on Elastic-Plastic Fracture Mechanics, Philadelphia, PA, Oct. 1981.
\
I. l i . ! L ! 6-2 STRUCTURAL . INTEGRITY ASSOCIATES INC
ATTACHMENT 3 ( Browns Ferry Nuclear Plant Unit 2, Cycle 5 Induction Heating Stress Improvement (IHSI) of IGSCC Susceptible 304 Stainless Steel (SS) Welds 1.0 Introduction
~
The results'df the ultrasonic (UT) examinations performed on the recirculation, residual heat removal (RHR), core spray, and reactor water cleanup (RWCU) piping systems indicated that only five welds contained IGSCC. It was decided to perform induction heating stress - improvement (IHSI) on all accessible, susceptible 304 SS Class 1 welds in those systems to prevent the initiation of IGSCC . IHSI was also performed on four of. the uelds with IGSCC indications to prevent the propagation of cracking. General Electric Company was contracted to perform IHSI under a two-phase workplan. Phase I consisted of a site survey to evaluate the implementation of IHSI on candidate welds. Phase II included coil development, scheduling, equipment setup, and all other work necessary to complete the IHSI treatments on welds identified as treatable in Phase I.
~
2.0 Phase I - Site Survey ( The site survey was conducted from December 10, 1984 through December 20, 1984. The following work was performed during the survey:
- evaluation of candidate welds designated by TVA for treatment
- collection of weld contour data
- verification of weld accessibility and identification of obstructions
- measurement of piping systems - study of potential IHSI equipment locations The survey information was then evaluated and a workplan for Phase II was laid out. *
*s .g e
e D H S@ G
~
2.1 IHST Workscone
' It was determined that IHSI could be implemented on 156 welds.
The treatable welds are listed in Tables 1 through 5. During the course of IHSI implementation, welds DSRWC-2-7 and DRWC-2-4 were deleted from the workscope. The elbow containing these welds was cut out and replaced to effect repair of the crack in weld DRWC-2-4. The total number of welds in the IHSI workscope was therefore reduced to 154. Twelve recirculation, 22 core spray, and 6 RHR welds were excluded from the workscope; these are listed in table 6. The recirculation nozzle-to-safe end welds and welds DCS-2-12, DCS-2-3, DRHR-2-12, and DRHR-2-3 were excluded because they were untreatable by the IHSI methods ^ generally available when the requisition was prepared. As IHSI techniques to treat these configurations become available, these welds will be treated. The other core spray and RHR welds which were excluded are carbon steel or low-carbon SS and are not considered susceptible to IGSCC. They will require no further disposition. f 2.2 Induction Coils The survey results indicated that 62 induction coils would be needed to perform the 154-weld IHSI work scope. This required 3 new coils in addition to the 59 coils already available to GE. 23 Interferences
- Sixty-seven interferences were identified during the site survey.
The list below gives the type and number of each obstruction identified. Type No. of Obstructions Structural Steel 2 Hanger Lug 5 Hanger Pad 6 Hanger Rod 8 Hanger Clamp 6 Electrical Conduit 8 Chain Falls and Wire Rope 8 Snubber Lug 5 Lead Blanket 2
~ Pipe Bracket 8 Penetration Insulation 1 Chain 1
Pump Housing 2 Thermocouples _
-
- 2 Instrument Lines 2 '
Painted Pipe 1 All interferences were removed prior to the treatment of each weld. Plant equipment, such as hanger components and conduit, was restored following treatment of the associated weld.
- i l
2.4 Equipment The equipment locations were also determined during the survey. Equipment needed for IHSI consisted of a 4160/480V three-phase
~
transformer, a frequency converter (power supply), work stations, a coolidg water system, and a data acquisition system. Each work station consisted of a voltage-reducing transformer, a capacitor bank, and a variable transformer that matches the converter output power to the impedance of the induction coil. The cooling water system was a self-contained closed loop supplying cooling water to the frequency converter, work station, coils and electrical cables. The data acquisition system monitored and documented the pipe temperature during each IHSI treatment. Thermocouples were attached to the pipe's outer surface and connected to the data acquisition system. Two work stations were located outside of the drywell, one at each equipment hatchway. .The IHSI control room, which housed the data aquisition hardware as well as the process control panel, was located at the personnel air lock. The power supply and cooling supply system pump skid were placed on elevation 593. In addition, a direct line communication system was established-( between the power supply, pump skid, heat station, and IHSI control room. A communication line between the IHSI control station and the reactor control room was also established. 3.0 Phase II - IHSI Treatments The IHSI treatments were performed from January 14 through , March 31, 1985. each system. The following table shows the time taken to complete No. of Date First Date Last Successful Thermocouple . Thermocouple l System Treatments Installed Removed l RWCU 12 1/10 3/24 CS 9- 1/15 1/24 Recire 99 1/21 RHR 3/31 29 3/6 1 3/29 , ,, 4 l An overall average of 2.9 treatments were performed each day. GE was ~ unsuccessful in treating recirculation welds KR-2-4, KR-2-1, KR-2-26, KR-2-23, and RWCU weld DRWC-2-5A. A total of 149 welds were treated successfully. In general, the treatment sequence for each weld included thermocouple (TC) installation, coil installation, ' low-power idle run, coil (_ adjustment, welds. treatment, coil removal, TC removal, and PT of TC tack the IHSISelected treatment. welds were also ultrasonically examined following
~
) . 49
. . - , , . , ~
~ , , , . , . . , . , - - - , , , - - . ~ , - , , , , . -,.-,-n.-- ..,n., , ,,,v,-, an-. . - - , - - - . . . - ,
'i 31 Thermocouoles Eleven TCs were attached to each weld to record temperature data i during IHSI treatment. Five TCs were positioned on one azimuth, parallel to the center axis of the pipe, with one centered on the weld crown and two on either side placed in the heat-affected zone (HAZ) and at the edge of the IHSI heat zone. Two TCs were also attached on the HAZ on the three remaining azimuths spaced 900 apart. On some welds, a twelfth TC was used to monitor the temperature of permanent obstructions positioned close to the IHSI heat zone. The data acquisition system had a 12-channel input, allowing all data to be recorded on tapes, and provided individual TC temperature printouts every 4 seconds.- A temperature profile plot was also provided during each IHSI treatment. The TCs were resistance welded to the pipe in accordance with ASME Section III, NB4311-3 Following the IHSI treatment, the TCs were removed and the affected areas were blended smooth and liquid penetrant examined in accordance with ASME Section III NB5000. 32 Low-Power Idle Run s ( A low-power pre-treatment at 2500C+500C (4820F+900F) was performed on each weld just prior to the full IHSI treatment to verify that the TCs were operative, the coil was positioned correctly, the water 5as cooling effectively, and load controls were operative. On some welds several low-power tests were
. required to precisely align the coil.
33 IHSI Treatment To obtain a successful IHSI treatment, the minimum throughwall temperature difference of 2750C (527oF) was effected within the treatment zone for the minimum heating time (see Table 7 for
- process control parameters). This was achieved by heating the pipe outer surface within the treatment zone to between 4000C (7520F) and 5750C (10670F) while simultaneously cooling the inner surface with system water flowing at the specified rates.
Several welds required more than one attempt to obtain a successful treatment. In- the treatment of 14 welds, there were - s' deviations from the process control parameters; these were all
, analyzed by GE engineering and documented on~ NCRs and FDDRs. The analyses showed 'that all fourteen welds obtained sufficient
- comprehensive stress to qualify for full treatment. TVA disagreed with the GE disposition of welds KR-2-36 and KR-2-37.
~ These welds were retreated within the specified process control parameter. limits.
L S 4 r- "
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g ww-, ->w y w v-- y r a=+ y . - - - g -1 +c- -w - y v --+ y e .m,
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~
3.4 _ Post IHSI Ultrasonic Examination A 25-percent sample of IGSCC suscep.tible welds were ultrasonically examined following the IHSI treatments. The welds were selected for examination based on the following factors: 1. Welds which had recordable indications and/or underwent evaluation and were found to have geometric reflectors during initial examination for IGSCC. 2. Welds in the same location wherc defects were found during the unit 1, cycle 5 IGSCC examinations. The welds in the sample are listed in Attachment 1. 4.0 _ Conclusions Despite schedule delays caused by labor chortages, weather, and loss of cooling water, the IHSI program undertaken on Browns Ferry unit 2 was successfully completed. Most of the IGSCC susceptible 304 ss welds in board of the penetrations on the subject systems received successful IHSI treatments. The _ _ . . - - susceptible welds which were excluded from the scope and those that were unsuccessfully treated have complicated or unconventional configurations. These welds which are listed in table 8 will be treated as the technology becomes available. IHSI has been shown to offer a level of mitigation against IGSCC. Treatment of these recirculation, RHR, core spray, and
~
RWCU will be cost effective by providing one or more cycles of operation repair with relative freedom from cracking and associated activities. Current speculati with other mitigation measures, e.g. ,on is that IHSI combined is required to provide lifs-of-plant immunity. alternate water chemistry O e e 8 y - e
BFN-2 TABLE 1 RECIRCULATION LOOP A SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 28 STP/SE GR-2-53 28 STP/LREL KR-2-45 28 STP/LREL - GR-2-54 28 STP/LREL - KR-2-47 28 STP/LREL KR-2-2 28 STP/ TEE GR-2-55 28 STP/ TEE KR-2 46 28 STP/ TEE KR-2-3 28 VLV/LREL GR-2-56 28 VLV/LREL GR-2-3 28 VLV/STP GR-2-57 28 VLV/STP GR-2-2
. 28' STP/SREL , KR-2-48 28 PMP/SREL GR-2-58 28 STP/PMP GR.2-1 28 CRS/ RED KR-2-11 28 CRS/ TEE GR-2-8
- 1 22 HDR/ECP KR-2-15 !
22 HDR/CRS
. KR-2-12 22 HDR/CRS GR-2-18 o
B220gE.PD
BFN-2 TABLE 1 RECIRCULATION LOOP A (Continued) SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 22 HDR/VLV GR-2-25 22 HDR/VLV . GR-2-26 22 HDR/SQL . KR-2-14' 22 HDR/ SOL KR-2-13 22 HDR/ SOL KR-2-19 22 HDR/SQL KR-2-20 12 STP/SQL GR-2-9 12 STP/SQL _ GR-2-12 12 STP/ SOL GR-2-19 12 STP/SQL GR-2-22 12
, 'STP/SE GR-2-11 12 STP/SE GR-2-14 12 STP/SE GR-2-17
^
12 STP/SE GR-2-21 12 STP/SE GR-2-24 12 STP/ RED GR-2-15 12 STP/LREL GR-2-10
' 12 STP/LREL GR-2-13
)
12 i STP/LREL GR-2-16' i i 12 STP/LREL GR-2-20 12 STP/LREL GR-2-23 12 STP/LREL KR-2-16 12 STP/LREL KR-2-17 12 STP/LREL
., - - < - - , , - ,,-,r. e. . , . ,
KR-2-18 - ,e
*"4
~
BFN-2 TABLE 1
~
RECIRCULATION LOOP A (Continued) SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 12 STP/LREL KR-2-21 12 STP/LREL KR-2-22 4 ECP/WLT _ GR-2-7 4 ECP/WLT GR-2-4 4 WLT/STP KR-2-4 4 WLT/STP KR-2-1 6 FLN/STP KR-2-49 Weld with indication of crack
~ ~
Titroughwall crack discovered after IHSI 94
~
h i O O O e I
l
-- l BFN-2 TABLE 1
_ RECIRCULATION LOOP B SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 4 WLT/ECP GR-2-33 4 ' WLT/ECP GR-2-30 4 WLT/STP KR-2-26 4 WLT/STP _ KR-2-23
~
6 . FLN/STP KR-2-53 22 HDR/SQL KR-2-41* 22 EDR/ SOL KR-2-42 12 STP/SQL GR-2-35 12' STP/ SOL GR-2-38 12 STP/SQL GR-2-45
- 12 STP/SQL GR-2-48 12 STP/SE GR-2-37 12 STP/SE GR-2-40 12 i
'STP/SE GR-2-43 12 STP/SE GR-2-47 12 STP/SE GR-2-50
_ 12-STP/ RED 1 GR-2-41 12 STP/LREL GR-2-49 12 STP/LREL
. GR-2-46 12 STP/LREL GR-2-42
- 12
. STP/LREL GR-2-39 12 .
STP/LREL GR-2-36 12 . STP/LREL KR-2-44 [ _ , . , . _ , - .- - ' ~ ~ ' - - ~ ~ * " ' " " ~ ^ ~ ,
BFN-2 TABLE 1
~
RECIRCULATION LOOP B (Continued) SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 28 STP/SE GR-2-59 28 STP/LREL KR-2-50 28 STP/LREL - GR-2-60 28 STP/LREL KR-2-51 28 STP/LREL KR-2-24 28 STP/ TEE KR-2-25 . 28 STP/STP GR-2-61 , 28 VLV/LREL GR-2-62 28 VLV/LREL ) ._ . GR-2-29 28 VLV/STP GR-2-63 28 VLV/STP GR-2-28 28 STP/SREL KR-2-52
~
28 PMP/SREL GR-2-64 28 STP/PMP GR-2-27 28 CRS/ TEE GR-2-34 28 CRS/ RED KR-2-33 22 HDR/ECP KR-2'-37 ' 22 HDR/CRS KR-2-34 l 22 HDR/CRS GR-2-44 I 22 HDR/VLV ' GR-2-51 l 1 22 HDR/VLV GR-2-52 ,
)
2 22 HDR/SQL KR-2-35 22 HDR/ SOL KR-2-368 O
-,rr. - - , + ,, ,
- n. n ,-- -- e n---- - - , . . . , , - , r. -
. BFN-2
, TABLE 1 RECIRCULATION LOOP B (Continued)
' SIZE (IN ) .
CONFIGURATION TVA-WELD IDENTIFICATION 12 STP/LREL KR-2-43
- 12 STP/LREL KR-2-40
. 12 STP/LREL KR-2-39 12 STP/LREL KR-2-38 5 WLT/ECP GR-2-63A
- 5 WLT/STP GR-2-63B Weld with indication of crack
- O e
i e e e 4 1 O 8 l
s
' BFN-2 TABLE 2 RHR LOOP A (SUCTION)
' SIZE (IN )
CONFIGURATION TVA-WELD IDENTIFICATION 20 STP/ TEE DRHR-2-19 20 STP/LREL DSRHR-2-9 20 STP/LREL - DSRHR-2-10
. 20 STP/LREL
~
DSRHR-2-11 20 LREL/VLV DRHR-2-21 ' ~ 20 STP/VLV DRHR-2-22
- 20 STP/VLV DRHR-2-23 20 STP/SCL
. DSRHR-2-8 J
' e D
4 4 e 1 o i s eN
- BFN-2
- TABLE 3
_ RHR LOOP B (DIS fARGE SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 24 TEE /STP DRHR-2-18 24 STP/VLV DRHR-2-17 24 VLV/SREL - DRHR-2-16
. 24 ,
SREL/STP DSRHR-2-7 24 STP/STP DSRHR-2-6 24 STP/VLV DRHR-2-15 24 SREL/VLV DRHR-2-14
. 24 SREL/STP/SREL DSRHR-2-SA 24 SREL/STP/SREL,, DSRHR-2-5 24 STP/SREL DRHR-2-13
, RHR LOOP A (DISCHARGE 24 STP/ TEE . DRHR-2-9 24 STP/VLV DRHR-2-8 24 SREL/VLV DRHR-2-7 24 SREL/LREL DSRHR-2-4A 24 STP/LREL DSRHR-2-4 24 , STP/STP DSRHR-2-3 24 STP/VLV DRIIR-2-6 24 VLV/LREL DRHR-2-5 24 STP/LREL DSRHR-2-2 24 STP/SREL DSRHR-2-1 ,
24 , STP/SREL DRHR-2-4 e . .d et
_. ~_ - 1 3 BFN-2 TABLE 4 CORE SPRAY - J . . ~
. ' SIZE (IN.) C0!! FIGURATION i TVA-WELD IDE!JTIFICATION
.l . 12 STP/STP i ' DCS-2-13 12 STP/LREL DCS-2-13A 12 LREL/LREL -
; DCS-2-7 12
, STP/LREL -
DSCS-2-9 . 12 STP/VLV DCS-2-14
. 12 STP/STP DCS-2-4 12 STP/LREL DSCS-2-1 i
d
. 12 STP/LREL DSCS-2-2 12 STP/VLV DCS-2-5 i .-
i 1 .4 1 i i ) 1 l i , . 1 1 i e 1- 8- ) . . h) O
...e;.
BFN-2 TABLE 5 REACTOR WATER CLFAN-UP SIZE (IN.) CONFIGURATION TVA-WELD IDENTIFICATION 6 SOL /VLV #DRWC-2-1A/DSRWC-2-1B 6 VLV/STP DRWC-2-1
. 6 STP/LREL- DSRWC-2-1
. 6 LREL/VLV DRWC-2-2 6 VLV/STP DRWC-2-3
~
6 STP/LREL DSRWC-2-1A 6 LREL/STP DSRWC-2-2
; 6 STP/LREL
. DSRWC-2-3
. . 6 STP/LREL DSRWC-2-4 6 STP/LREL -
DSRWC-2-5 6 LREL/STP DSRWC-2-6
~
6 STP/LREL DSRWC-2-7 6 STP/LREL , DRWC-2-4 6 FLUED HEAD /STP DRWC-2-5A 6 STP/ VALVE DRWC-2-53
*/0NE WELD ONLY 9
8 9
- m. . _ _ _ _ ~ _ . , _ , _ _ _ . . . _ _ _ . - . _ , _ . . _ _ _ - . _ _ _ . , _ , , . , , , - _ , , _ _ , . _ . . , _ . . _ _ . _ _ . . , . _ , . . _ _ , _ . _ . . . _ , , . , ,_ ._
j
. . TABLE 6 j ..
WELDS EXCLUDED FROM IH5I WORKSCOPE i t a r Recirculation N-2 Nozzle-to-Safe End Welds (10) - l
~
Recirculation N-1 Nozzle-to-Safe End Welds (2) t . I _ Core Sorav System i DCS-2-12 TCS-2-422 TCS-2-402 DCS-2-3 TCS-2-423 TSCS-2-404 r TCS-2-417 TSCS-2-424 .. TCS-2 405 TSCS-2-418 TSCS-2-425 . TCS-2-406
- TCS-2-419 TCS-2-426 i TCS-2-407 3
TSCS-2-420 TCS-2-401 I TSCS-2-408 i TCS-2-421 TCS-2-403 TSCS-2-409 1 t TCS-2-410 i . l RHR System 4 i TRHR-2-191 , TRHR-2-192 , , DRHR-2-12 DRHR-2-3 ..1 TRHR-2-194 o TRHR-2-193 o
..-..,.-m. -, . - ,.. . . _ . _ . . - . - _ . .
-c.. ,,..._.,...#-.-___,-__-,._,....m,.-.,,,,,,--1
, ,. -.,.., ,m .~., - - . - . . ,
TABLE 7 4 IHSI PROCESS CONTROL PARAMETERS - STAINLESS TO STAINLESS STEEL JOINTS
- 1. - Pipe Outer Surface Temperature 5000(+750,-1000) within Treatment Zone (Notes 1, 2) ,
1A. Maximum Weld Crown Temperature 6000C
- 2. Minimum Throughwall Temperature 2750C Difference 0$T) -
3 Minimum Width of Zone Heated to 1.5g/Rtorcoillength/2, T Minimum (Note 3) whichever is less (R = Radius to mid-wall, t = wall thickness)
- 4. Minimum Distance from Weld Center 15 mm (0.6 inch) or t/2 to Boundary or21T Minimum (whichever is larger, but not less than edge of weld crown)
- 5. Minimum Heating Time to 2 seconds 0.7 t /a Temperature (a = Thermal diffusivity, t = wall thickness)
. 6. Maximum total time for outer 20 minutes
- surface above temperature of 4250C -
1
- 7. Nominal Frequency 3 to 4 kHz
- 8. Minimum Induction Coil Length 3\/Rt (R = Radius to mid-wall,
, t = wall thickness) l 4
I t l l
- l D
-_y,,y ,,m.m, , . , . , -. _ . _
TABLE 8 WELDS STILL REQUIRING IMSI TREATMENT - Recirculation "N-1 nozzle to safe end (2) N-2 nozzle to safe end (10) KR-2-23
, KR-2-26 .
KR-2-4 KR-2-1 Core Soray DCS-2-12
, DCS-2-3 RHR DRBR-2-12 DRHR-2-3 RWCU DRWC-2-5A e
e 6 O
* '44 5
Attach =ent 4
- R*?..
STRUCTURAL JUSTIFICATION FOR THE OVERLAY REPAIR ON WELD GR-2-15 ( Overlay Sizing Calculations Weld overlay sizing calculations were performed based on a 360 0 ! through-wall circumferential crack in the 12-inch end of the 28 X 12-inch reducer. The thickness at this joint is 0 579 inch. The resultant overlay is 0 35-inches thick and is depicted in Figure 1. The 0 35-inch thickness
. .~ .
is in addition to the seal weld which is applied over the crack and the first weld layer that clears dye-penetrant testing (PT) inspection. 4 Axial stresses at this joint are given as: Pressure = 6,321 psi Dead' Weight = 1,990 psi Seismic = 6,000 psi Thermal Expansion = 14,000 psi . The primary stresses include pressure, dead weight, and seismic stresses; thus the resultant stress is 14,311 psi. the design temperature of 5750F is 16,675 pai.The allowable stress, Sm, at i The primary stress ratio, (Pm + P b }/S m , is about 0.858, which results y in an allowable flaw depth to thickness ratio, a/t, of 0.495 for a 3600 crack, from ASME Section XI, Table IWB-3641-1. Therefore, the unrepaired
~
joint is unacceptable; however, an overlay repair of 0 35-inch thickness results in several effects which render the repaired joint acceptable. The ( a/t ratio is reduced from 1 to 0.6232, and the primary stress ratio is reduced from 0.858 to 0.5348 because of the increased pipe wall thickness. For this stress ratio of 0.5348, an allowable a/t ratio of 0.6626 is obtained from IWB-3641-1 for a 3600 crack, and the allowed crack depth, a, is determined to be 0.6155-inch deep. , Fatigue Crack Growth . Consideration of fatigue crack growth during service is required to show that the original 3600 crack of 0.579-inch depth will not extend past the allowed 0.6155-inch. depth. O 0365 inch. Thus, the allowance for fatigue crack growth is , 4 Axial stresses at this joint for heatup/cooldown cycles include pressure, and thermal stresses, thus the resultant stress is'20,321 psi. This stress can be reduced by the unoverlaid-to-overlaid-thickness ratio, 0.6232, and this reduced stress is 12,665 psi. The EPRI DRIVE Computer 4 Program was used to compute the stress intensity factor, K, for a 3600 , - 0.579-inch deep flaw having a stress of 12,665 psi. The resulting stress intensity factor is approximately 38 kai yinen. A weld metal fatigue crack growth curve is assumed equal to the upper bound 3 of solution anncaled Type 304 for DWR environments, as shown in the attached figure from EPRI Report NP-2423-LD. Th , correspondingtoK=38kaiVinchisabout4X10gcrackgrowthrate in/ cycle. Because k. e p* ..
* . g* *
^ changes in K are negligible for small amounts of crack growth, it is ' estimated that it would take 91 heatup Nooldown cycles to use up the 0.0365 inch allowance for fatigue. _ Conclusion - Based on a conservative estimate of 10 heatup/cooldown cycles pe:f year, it would take about 9 years for the crack to extend from 0.579 inch to the limit of 0.6155 inch. Thus, the joint is suitable for service with the weld overlay for at least 2 fuel cycles which is the maximum that is currently accepted by NRC.
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' GENERAL NOTESPERTAINING TO WELOING ANO NDE
! FOR OVERL A Y OF HELO GR-2-IS l1 '
- 1. sELDING ANONONDESTRUCTIVE EXAkf/ NATION SHAL L i' , . . . BEPERFONAfEOIN ACCOROANCE K1THASAtE SECHON II,1974 WITHSUMMER1975 ADDENDA, AND THE ADDIT-IONAL REOUIREMENTS OF THIS DRAWING. N 2 WELDINGPROCEOURES MELDERSANOKELDINGDPERA- .
TORS SHAL L BE OUAL dtEO TO THEREOUIREI,*ENTS OF a. [k~b O.S ' ASA?ESECT LE. H[LO!NGPROCEECRESHALL GE AFFIC-s k. j'
'Q ,, VED BY TVA PRIOR TOUSC 2: <So- 3 Al L WEL DING EXCEPTASPROV/DEDINNOTES SHAL L n 2 .
BECGvE SY THE GAS TUNGSTEN ARC WEl. D!NG PRO- O H
- s. s CESSUSINGER30GL FILLERMETAL CONfORA!/NG TO
\ CENTER PUNCH BEFORE WELO ASAtE SFA S9 DELTA FEHRITECCNIENTOFDEPOSITED
'%. ~8 PLACES ATEACHLOCAHON ssELO METAL SHALL BEBFNMIN ASDETERMINEO/sY
%. (4 LOCAHONS1 THE MAGNEHC INSTRUA!ENT MED100 OFASA1ESECT 2. N n
~.s NB-2400. ,l
? {:,'Q.
- t. N':3 r ,*a -
- 4 DURI*/G WELDING, A MAXIMUMINTERPASS TEMPERATu%.
- . STER C - x%. a J' =
N OF3SO*FAND A Af4YlMUMHEATthRUTOFSOK/lOJWLES PERINCH SHALL BE OBSERVED N . 2 _Q ;\ j ,
, c h '. STEP B s.= -
r ',
. S PR:OR TOCEPOS! HON OF DrE STRUCTURAL OVERL Af TC sin PESIGNDI.1 TENSIONS, THE AREA CONTAIN!NG hiiru nAll ~
\ l,l
~
CRACKS SHALL BESEALEDBYMELCING. l $.% ': STEP A
\ '
^
$^[f-0 *
- Q, y
*l THICXNESS CF THE SEAL WEL O NEED fl
~
'h
, WELD CRACK
\ '
a.
//
['-
+
. ^^'
V--
'/
NOT EXCEED ONE L A YER PROvlDED THE RECuffrEMENTSn ' OFNOTE 6 AREME T PIPE INTERIOR SHALL BE DRY OvRD 1
, ;- EXISHNG KELOBuffOUP
//\_f. _d*
-/
=
'h ING SEAL WELDING. SEAL WELD MAYBE MAOCBY THE '
PROCESS OF NOTE 3 ORBY THESHIELDEOMEIAL A TrC 5 r
/
S //
y, ,f' /
PROCESS USil./G EJOBL-IS CN-16 ELECTisODES Cf ASt i s SFA 5 4. THE FERRITE REQUIREMENTS OF NOTE 3 APPCT l C THESEAL 1tELD LAYER AND AUJACENT SUPfACE TOBL*
~ GR-2-15 l
- OVERL A YED SHALL BELIQUID PENETRANTEYAMirEO PRIOR TO START OF THE STEPB OVERL AY L AYER, THE STEPS A [EPOSIT ABOUT 3 TO 4 BEADS TODRYSEAL (PT REO'O) STEPB L AYER SHALL Al SO BELIOulO PENETRANT B AUTO (GTAW), ONE L AYER, DRY (P T REO'DJ EXAMINEO PRIOR TO BEGINNING THESTRUCTORAL CVERL AY (STL P C).
C AufO(GTAW), HETPIPE,O3SINCHMIN. THICKNESS (E2CLUDING L AYERS A a DJ yJ . -
~
!L. . s . y .:.q T x.y j' g 7 It-E flRSTL AYER OF THESTRUCTURAL OL ERL AY SHALL BE MACE WITH PIPE INTERIOR :; FA-Th TRf :vRtNG DEPOSIHON OFSUBSEOuENTL A YERS REOulRED 10 CEL L LCPE I.[CL 55 arf CVERL AY THICKNESS, PIPE!NT[RIORSHALL CONTAIN 5 TANDING CRfLOWING WATETs.
.:3GtitE f~f.l ~~{Ja ,
' .g
- B CCMPLETEU OVERLAY SHALL BE EXAMINEO BYLIOUID PENETRANT AND BYUL TRASONIC RO ISSUE FOR ECNPS2/S 90wERK3uSL ne ncson aunto\uc.umi t
.L ~~ TEST 5 MPROPRIATEFOR DETERMINA TION OF WELD SOUNDNESS AND BOND TO THE ~~. ORIGINAL PjP[/MELO SURFACE. NOT TO SCALE VECHANICAL
- ALL nELD LAYERSINCLUOpVGSEALWELO TOEXTEND 360*AROUND11PECIRCUUFER[NCE ,' RECiRCULATDON S1 STEM
- 9. . WELD GR 2-15 OVERLAY r c"mw;y.yp."g j
. - _ _ . ip
=. -
F" r~ ' g w m m .,
STRUCTUREL INTEGRITY ASSOCIATES,INC. C.6 socrates
) frecerica Copeaand.Ph D.
Thomas L Gert cr. Ph D. Anthony J Giannur.s.Ph D. Anthony N Mucciards. Ph D. Pats r C. Riccardella. Ph D. PCR-85-032 March 27, 1985 Mr. James E. Wilson Tennessee Valley Authority 1420 Chestnut Street Tower Il Chattanooga, Til 37401
Subject:
Independent Review of the Overlay Repair for Weld GR-2-15, Browns Ferry, Unit 2
Dear Ed:
Our independent review of the overlay repair on weld GR-2-15 shows that the ( weld overlay design on the subject weld is adequate. Highlights of the review for the subject weld overlay are summarized as follows:
. Axial stresses at this joint were calculated and tabulated in Table 1. Resulting stresses are very close to those used in the TVA analysis. We concur with your approach of not using Code stress indices, as this is the standard approach used on all Browns Ferry, Unit 1 overlays, as well as those at most other plants.
. Based on the stresses in Table 1, a minimum thickness of 0.31 inch is required for the overlay (Table 2). The 0.35 inch thick designed overlay provides an extra 0.04 inch allowance for fatigue crack growth.
. Stress intensity factor for a 0.929 inch thick cylinder with a 0.579 inch deep, 3600 circumferential crack was calculated to be 35.1 ,,
KsivTn (Figure 1) which is compatible with 38 Ksifin i~- given in the design analysis.
. The f atigue curve used in the design analysis was judged to be adequate and the 4x10-4 in/ cycle crack growth rate was reconfirmed.
b 3150 ALM ADEN EXPRESSWAY SUITE 220 = SAN JOSE.CALIFORiflA 95118. (408)978 8200 TELEX 171G18 STRUCT
,. ,._ _ _ ,, ,, , s mm. r--rr~ -^ -
Page 2 PCR-85-032 '
. Allowable flaw size after the 0.35 inch overlay repair was evaluated and tabulated in Table 3. It was also reconfirmed that fatigue crack growth from more than 90 heatup/cooldown cycles can be tolerated within the extra 0.04 inch thickness allowance.
Additional mar. gin on cycles could also be obtained by taking credit for part of the first weld overlay layer if needed. Should you have any further questions, please call me. Very M 1 yours, ji '. YhV/ P. C. Riccardella
/sl enc.
cc: Frank flovak Welding Services, Inc. ( W e STRUCTURAL INTEGRITY ASSOCIATES INC
b TABLE 1 Calculation of Applied Stresses , TVA-06 WELD GR-2-15 Pressure =1150 psi OD=12.75 inches
~- -
Z =64.5 in**3 LOAD M:: liy fin l1b A:: i a l 5ig CASE (ft-lbf) (ft-lbf) (ft-lbf) (f t-l b f ) (psi) PRESSURE 6~.30.96 DW 381.00 139.00 10661.00 10661.91 1933.61 TE1 ~.1 6 2 . 0 0 701.00 12590.00 12617.49 2347.44 TE2 6621.00 73762.00 3232.00 73832.77 13736.~3 ODE-xy 1129.00 5317.00 31487.00 31932.77 5940.90 - ODE-y: 549.00 2486.00 125AO.00 12003.66 2:32.00 SSE-xy 1610.00 7922.00 46071.00 47535.76 8843.86 SSE y: 786.00 3509.00 17068.00 10209.30 3337.78
, e a 6 4
L 4
,.e 9 .
TABLE 2 - Weld Overlay Sizing - pcCRACK STRUCTURAL INTEGRITY ASSOCIATES, INC. VERSION 1.0, AFRIL 1985 SAN JOSE, CA (408)978-8200 WELD OVERLAY SIZING OVERLAY SIZING FOr: CIRCUMFERENTIAL CRACK:- TVA-06, WELD GR-2-15 (JALL THICKNESS = 0.5790 STRESS RATIO = 0.8550 L/ CIRCUMFERENCE
- 0. 0 0.1 0.2 0. 3 0. 4' O.5-->1.O FINAL A/T 0.7500 0.7500 0.7500 0.7500 0.7500 0.6514 OVERLAY THICKNESS 0.1930 0.1930 C.1930 0.1930 0.1930 0.3099 W h l
l b Y e
N: TABLE 3 Allowable Flaw Size for Pipes with 0.35" Weld Overlay pcCRACK STRUCTURAL INTEGRITY ASSOCIATES, INC. VERSION 1.0, APRIL 1995 SAN JOSE, CA (405)978-8200 CRITICAL FLAW SIZE EVALUATION CRITICAL FLAW SIZE FOR CIRCUMFERENTIAL CRACK:- TVA-06, WELD GR-2-15 ALL THICI'fJESS= a.9290 STRESS RATIO = 0.5330 L/ CIRCUM
.O .1 .2 .3 ALLOWABLE A/T
.4 . 5-- :- 1. 0 O.7500 0.7500 0.7500 0.7500 0.7500 0.6635 .
z**
- 1;..RE+TE2 P .
J . o. 50 ........................
......_....................................................................................>1.......
. .p 4 ............................... ................................. . . . . . . . . . . . . . . . . . . . . . . . . . . . ....._....
I I n0 _ . . . . . . . .. . . . . . . . . . . . . . . . . . . . . . . _ 1
- u -
- l x . .
g, 0 ..............................
....................... p ..... ............................. .
10 .....................-..._.,.._f' . G I t 0 0,2 0,4 0,6 O,8 IUCHES CRACl! DEPTH PRESSURE +IHERHAL: 12,5071(SI FIGURE 1. Stress Intensity factor Versus Crack Depth for a 0.929" Thick Cylinder (R/t=10) r. __ _.-._--_.-a}}