ML20024E165
| ML20024E165 | |
| Person / Time | |
|---|---|
| Site: | Pilgrim |
| Issue date: | 12/31/1982 |
| From: | Cipolla R, Hayward J, Mcnaughton W APTECH ENGINEERING SERVICES |
| To: | |
| Shared Package | |
| ML20024E158 | List: |
| References | |
| AES-8102243, NUDOCS 8308090346 | |
| Download: ML20024E165 (64) | |
Text
{{#Wiki_filter:I AES 8102243 Final Report APTECH engineering servicer,Inc .~o,~eeRi~e cOusuL1A~1S 795 SAN ANTONIO ROAD . PALO ALTO . CALIFORNIA 94303 (415)858 2863 A FRACTURE MECHANICS EVALUATION OF CORE SPRAY SPARGER INDICATIONS Prepared by Russell C. Cipolla Warren P. McNaughton John A. Hayward
- Aptech Engineering Services, Inc.
795 San Antonio Road Palo Alto, California 94303 Prepared for Boston Edison Company Pilgrim Station Rocky Hill Road Plymouth, Massachusetts 02360 Attention: Mr. Joseph A. Nicholson 8308090346 830804 PDR ADOCK 05000293 December 1982 G PDR CMr. J. Hayward is now with Design Reliability Assoc., Mountain View, CA 94040 Services in Mechanical and Metallurgical Engineering, Welding, Corrosion, Fracture Mechanics, Stress Analysis
I:? i
! QUALITY ASSURANCE VERIFICATION RECORD SHEET i.
TITLE: A Fracture Mechanics Evaluation of Core Spray Sparger Indications. Originated by st/[- ht ./../s Russell C. Cipolla i Warren P. McNaughted M
'% lY ' Vlp' Approved and Verified by R-Geoffrey R. Egan 7
Quality Assurance Approval is - Jeffrey D. Byron 1.
TABLE OF CONTENTS Contents Page Section SYN 0PSIS INTRODUCTION 1 1 ANALYSIS METHOD 5 2 Introduction 5 Failure Behavior of Type 304 Stainless Steel 7 Fracture Mechanics Approach to Stress Corrosion 8 Cracking Crack Growth Rate Representation 9 Development of the Fracture Mechanics Model 10 3 ANALYSIS OF STRESS STATE 12 Sources of Stress 12 Rolling Fabrication Stresses 13 Analytical Determination 13 Experimental Measurements 13 Welding Fabrication Stresses 13 Service Stresses 17 MATERIAL PROPERTIES 22 4 Crack Growth Rates 22 Mid-Sparger Growth Rate 24 Near Weld Material Condition 26 Yield Strength 26
SUMMARY
OF INSPECTION RESULTS 28 S LIMIT LOAD ANALYSIS 29 6 Limit Load Model 29 Numerical Results 31 RESULTS OF STRUCTURAL INTEGRITY AND FRACTURE MECHANICS 34 7 ANALYSES Introduction 34 Region Away From Welds 35 Region Near the Weld 38 41 Back-Side Defect Evaluation 44 8
SUMMARY
AND CONCLUSIONS 46 REFERENCES APPENDIX - Proposed Pilgrim Core Spray Sparger 48 Evaluation Plan
ILLUSTRATIONS Figure Title Page 1-1 Boston Edison Core Spray Sparger Fracture Mechanics 3 Evaluation. 2-1 Schematic Showing the Relationship Between Failure Stress 6 and Flaw Size For Two Limiting Failure Modes. 2-2 Model of Center Cracked Plate. 11 3-1 Resultant Residual Stress Distribution After Fabrication 14 and Assumed Relaxation to 5500 Yield (23.7 ksi max). 3-2 Multi-Roll Pipe Bending Machine. 15 3-3 Residual Stresses Due to Rolling. 16 3-4 Maximum Surface Residual Stresses For a Four-Inch Pipe 18 Inside Surface at 0.1 Inch From Edge of Fillet of Pieces 1 and 2 Versus Azimuth. 4-1 Crack Velocity as a Function of Kmax 23 4-2 Assumed Crack Velocity Behavior. 25 6-1 Limit Load Model and Geometry 30 6-2 Critical Stress Results For Through-Wall Crack. 32 7-1 Effect of Stress Level and Material Crack Growth Rate on 36 Flaw Length vs. Time for Front-Side Indication Away From Weld Region. 7-2 Predicted Crack Growth For Indications on Front-Side of 37 Sparger Pipe Away From Welds For 18 Months of Operation. 7-3 Effect of Stress and Material Crack Growth Rate on Flaw 39 Length vs. Time For Indication Near a Weld. 7-4 Predicted Crack Growth For Indications Near Welds Over 18 40 Months of Operat.f on.
i
' TABLES Ta'ble Titie- Page 3-1 Core Spray Sparger Stress Summary 19 4-1 Comparison of Sparger Pipe Heat With Other Typical Heats . 27 of Type 304 7-1 Summary of Crack Growth Predictions For Back-Side Flaw 42 7-2 Computed Remaining Time For a Visible Back-Side Flaw 43 Indication
SYNOPSIS Surf ace Indications have been located in the core spray sparger of the Pilgrim Nuclear Power Station operated by the Boston Edison Company (BECO). The Indications were found by visual Inspection during a scheduled outage in January 1980. The results of an evaluation of the significance of those Indications on the sparger Integrity were presented to BECO in September and in November 1981, in Aptech Engineering Services Report AES-81-10-83, entitled, " Preliminary Report Structural Evaluations of the Pilgrim Station Core Spray Sparger Based Upon Results From the October 1981 Remote Visual inspection." This report has been reissued with changes as a final report in December 1982. The present report provides detailed analytical and computationsi background to that report. It also quantifies the potential degradattor. of the sparger to perform its design f unction and estimates the expected remaining service life of the sparger as a function of the size and location of Indications observed. The Indications have been divided into two general groups for purposes of l this evaluation: (1) those indications near the weld regions and heat af facted zones of various attachment weldments, and (2) those remote from welds. The following assumptions have been utilized in the analysis that was performed: e The principles of linear elastic fracture mechanics (LEFM) are appli-cable for predicting subcritical crack growth. e The cracking mechanism is Intergranular stress corrosion cracking (IGSCC) and, hence, the life prediction analysis was based upon an IGSCC model. l
- l. e The flaw model is based upon the solution for a center cracked flat plate geometry under general varying stress conditions.
l
e The representation of the crack growth rate or velocity is primarily based on upperbound da/dt behavior (i.e., Type 304 stainless steel in the f urnace sensitized condition and tested in water with 8 ppm 0 2 at 550*F). o Crack arrest is assumed when crack driving force (i.e., maximum stress Intensity factor K . vanishes (K = 0). max max e The maximum stress !evel is limited to yield strength at 550*F (o = 23.7 ksi). Y The analysis perf ormed shows that for the proposed stress state existing at these locations, a through-wall crack located away from the welds will arrest given " service" stresses less than 10 ksi tension. The arrest length is shorter than the critical size to cause failure by a limit load mechanism. A through-wall crack located in the region of the weld details, however, may not arrest prior to reaching critical length (~ 80% of pipe circumference) to f ail ure, even f or low primary stress loads. However, established maintenance criteria requires the placement of an external pipe clamp when a 180' crack Indication is detected or predicted. Consequently, f ailure by primary loading exceeding the Ilmit-load capacity will be mitigated. i (
1 Section 1 INTRODUCTION During a scheduled maintenance outage in January 1980, a remote video inspection founa indications in the core spray sparger of Boston Edison's Pilgrim Nuclear Power Station. Aptech Engineering Services was asked to participate in determining the significance of these Indications and did so in a two part study. First, a digital enhancement technique was used to enhance videotape inspection records to determine the validity of the suspected Indications. Second, once it was determined that probable cracks did in fact exist in the sparger, a fracture mechanics evaluation was performed. A preliminary report, " Structural Evaluation of the Pilgrim Station Core Spray Sparger Based on Results From the October 1981 Remote Visual Inspection," AES-81-10-83, was issued in November 1981 and included the inspection results and a summary of the fracture mechanics evaluations. This report has been subsequently reissued as a final report in December 1982 with some minor revisions. The detalis of that latter work are the subject of this report. The work detailed herein had four objectives: e To assess the significance of core spray sparger indications on the structural Intsgrity of the sparger e To estimate the expected remaining service life of the sparger given its present condition e To quantify the potential degradation of the sparger to perform its design function v g-P+ e-%y--v - - *- a, -
2 e To establish inspection acceptance criteria and develop a maintenance action plan (see Appendix) The digital enhancements showed that probable cracks did exist in the sparger (.1) . For the purposes of analysis, the Indications were classified into two distinct groups: (1) those indications that were associated with weld and cold heat af facted zones, and (2) those Indications that were away from the told regions. Details of typical indication sizes that were detected by the visual examination are given in Section 5. The of fact that the indications may have on the structural integrity of the core spray sparger has been determined using the Iimiting condition that the , Indications behave like cracks. Several questions were examined as outlined in the flow diagran of Figure 1-1. Given the presence of a flaw (in the torst case a through-thickness crack), could a growing flaw be expected to crrest under the combined actions of residual stress plus service induced stress? If so, how large would the flaws be when this occurred and would this length flaw affect integrity? If not, how long would it take a crack to grow before it did af fact integrity? To answer these questions, each region was examined in detail. The stresses cssociated with each region were critiqued and are summarized in Section 3. This discussion includes residual stresses due to sparger fabrication (i.e., ' pipe roliIng and welding), and postuiated Ievels of service-induced stress. A summary of the results of an experimental program performed by Teledyne Engineering Services (TES) to measure the rolling f abrication stresses is also included in Section 3. The probable crack growth rates are discussed in Section 4. These rates are a function of the material condition (history of f abrication) and stress levels and have been the subject of much experimental tork in the past five years. Section 6 presents the background for limit load analysis, as well as the numerical results in order to address the ! questions of sparger Integrity. The results of the analysis are presented in Section 7. l l l
3 BOSTON EDISON CORE SPRAY SPARGER FRACTURE MECHANICS EVALUATION Are video indications ck - Bound analysis by treating cjefeets ; as cracks or accept Yes Are indications Perform surface crack analysis, through-wall? No m Determine lifetime to throughwall flaw. Add this to analysis below. Yes u What region is indication in? Near , Input to Analysis Weld (1) Probable weld residual stress j distribution Away_ (2) Bounding growth rate From (3)-Through wall flaw (4) Parametric applied stresses-Input to Analysis (1) Fabrication (rolling) stresses v (2) Furnance sensitized growth rate 'lill arrest occur? (3) Through-wall flaw model At what length? (4) Parametric applied stresses u ' Will arrest occur? Find limit load critical flaw. At what length? m, Compare to arrest length. Determine failure mode. Find lifetime to failure. Find remaining life from NDE data. l Figure 1-1 M
. . . . - , , . . .-,.-4. ,--.-e -- - - - - - - , - - - - . - - . - - - , - - - - - - + - - - ,- -w. . .
4 Before we proceed to the analytical assessment and numerical results, the next section (Section 2) presents the background and t' asis to the analytical methods and fracture mechanics concepts utilized throughout the remainder of the report. l 9 a ( i l
- - - - - - - . - - - ~ - . - - - . - . _
5 Section 2 ! ANALYSIS ET}iOD l [ f 2.1 Introduction in assessing the significance of sparger defects in stainless steel piping, two f ailure modes were investigated: (1) ultimate capacity or residual strength under increasingly long cracks, and (2) subcritical crack growth by l an Intergranular stress corrosion cracking (IGSCC) mechanism. The failure l behavior of pipes under monotonic loading can be classified into three regimes in which a specific type of failure mode is appropriate. The discipiines required to assess these regimes ares e Linear Elastic Fracture Mechanics (LEFM) - The structure f alls in a brittle manner and, on a macroscale, the load to failure occurs within nminally elastic loading. l e Elastic-Plastic Fracture Mechanics (EPFM) - The structure falls in a ductile manner, and significant stable crack extension by tearing may procede ultimate failure, o Fully Plastic Instability or Limit Load - The failure event is charac-terized by large deflections and plastic strains associated with ultimate strength collapse. A diagram showing the relationship between critical and f ailure stress and flaw size for the three f ailure modes is shown in Figure 2-1. The shape and position of the failure locus will depend on the fracture toughness (K ) and strength properties (o and o ) of the material, as welI as the structural gemetry (t) and type of loading. LEFM is used most appropriately to describe the behavior of low toughness /high strength materials in which
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! Limiting Failure Modes. l i e
7 the plastic zone is small relative to the structural geometry and little ductility precedes fracture. With this method, no account is taken of increased material resistance to brittle failure when significant plasticity occurs. Under LEFN conditions, the most usef ul parameter for characterizing the behavior of cracks is the stress intensity factor, K, which characterizes the singular stresses near the crack tip. In contrast, plastic Instability, when it occurs without prior crack extension, is dominated by the flow properties of the material. In these circumstances, the failure condition is independent of fracture toughness and crack tip characteristics, an1 a limit load analysis is used to define the f ailure conditions. EPFM analysis can be used to predict failure behavior in the transitional regime between LEFN and Iimit load, and under EPFM conditions, the crack tip singularity, the material toughness, and net section strength are all important parameters for failure assessment. 2.2 Failure Behavior of Type 304 Stainless Steel in deformation studies of Type 304 stainless steel, under monotonic loading, significant ductIIe behavior has been observed. Specifically, a recent experimental program (2) on the Integrity of pipes containing stress corrosion cracks in Type 304 stainless steel pipes has found that: o "Because of the substantial crack tip blunting that precedes crack growth, the applied stress at f ailure is virtually Independent of the sharpness of the Initial flaw introduced into the material." e "The presence of a weld and any sensitization of the material sur-rounding the flaw does not significantly af fect the applied stress at
- f ail ure."
e "The exact shape of the flaw is of considerably less importance than the area of the flaw (or of the not flaw area when multiple flaws are present) in determining the applied stress at f ailure." l
8 The foregoing conclusions are highly supportive of a net-section f ailure criterion where a limit load analysis will provide a reasonably accurate cppraisal of the f ailure load and the necessary flaw size which could become critical during the service operation of the sparger system. A limit load f ! model is developed in Sec+1on 6 to calculate the conditions for final pipe l l failure. i l ( 2.3 Fracture Mechanics Approach to Stress Corrosion Cracking i Structural components containing defects and under combinations of static cnd alternating loads can experience time-dependent degradation in strength due to subcritical flaw growth. Since crack growth itself may be regarded as ! damage by determining the extent of crack progression under load, damage can be accumulated and the residual or remining service life can be predicted. ! the defect size, I The principles of LEFM of fectively link three parameters: crack growth rate of the material, and the applied stress, so that if any two of them are known, the third can be quantified. The most useful parameter in describing the character of the near-crack-tip stress distribution is the stress Intensity f actor. The stress intensity f actor, K, def ines the magnitude of stress distribution and is calculated in terms of the applied, nominally uniform stress, o, the crack length, a, and a factor that depends on the flaw geanetry, stress distribution, and structural displacement constraints, F(a), from the relation K = oF nra , (o < o ) (2-1) Y Stress corrosion cracking evaluation based on fracture mechanics assumes that flaws are present of size a and that the lifetime of a part is that required for a crack to grow from the initial size, a , to the critical size, a . It has been proposed (3) that crack growth rate data may be correlated to the crack tip stress Intensity factor for the given load cycle in the form of relations such as:
- ,_. .. _ .__ _ ~- - ... - .. -- . - . _ . , - - - - - .
9 da/dt = f(K), (2-2) where t is a time parameter. By integrating Eq. (2-2) with the appropriate component stress field to calculate K, the time (residual life) for a crack to grow from a to a is computed from:
*f da T= (2-3) da/dt "I
where T is the remaining life in units of time. The final flaw size expected at the end of the design life, a , can be determined by Eq. (2-3) with the appropriate stress distribution in the analysis for K by Eq. (2-1) and the ! desired life T from: i I
"f da T - =0 (2_4) o da/dt J
a; Equation (2-4) is an expression involving a that usually must be solved by an iterative process. 2.4 Crack Growth Rate Representation Many empirical relations to express crack growth behavior have been proposed; the earliest and most well known follows the Paris rule (4) as proposed by Egan and Cipolla (3) which takes a linear form on log-log paper and is represented as da/dt = CK (2-5) max where C and n are constants determined f rom the data, and K is the max maximum applled stress intensity factor computed from the maximum stress level in the loading cycle. In the remaining life analysis, a more general
10 rule was used for describing da/dt; specifically, the crack velocity was represented by a piecewise linear curve that best describes the data in a bounding way. This piecewise curve as a function of K was used in the max numerical Integration of Eq. (2-4). Section 4 presents the crack growth data used in constructing these curves. 2.5 Development of the Fracture Mechanics Model A through-thickness crack provides the largest stress intensity factor of the potential models which could have been chosen for this analysis. To provide an upper bound (that is worst case) analysis, this model was used to cxamine arrest and limit load conditions. This gemetry is shown in Figure'2-2. This model was used in both the weld region and away frm the teld.
11 IFI = 201 - CENTER. CRACKED PLATE IN MODE I o r 0 (m) y \# MODEL CE04tRY 2a
._a_ _a_. V N# n p - - q m, n '
rx :
- t(x)
I" : NOTE: Local crack (x'} and global model (a) coordinate systems are shown o n ,, ,' as being coincident (a g= 0) 4H j MODEL DESCRIPTION H0 DEL FEATURES PARAMETER OPTION FEATURED Model Inden Number IFI 201 Number of Degrees of Freedom IDOF 1 Crack Front Shape -- Straight Crack Cpening Mode -- Itode 1 Finite Width Effects w yes Variable Thickness Effects NTH yes l l Figure 2 Model of Center Cracked Plate (5). W
12 Section 3 ANALYSIS OF STRESS STATE 3.1 Sources of Stress The sources of stress in the core spray sparger f all into two categories: (1) the residual stress ramalr.ing af ter fabrication, and (2) the actively applied stresses due to service. The residual stresses result from the forming of the pipe bends by rolling, and the subsequent welding together of subassembl ies. The service stresses are due in part to heatup/cooldown loads; however, these stresses are small as will be discussed later. The total stress state can af fect the analysis in several ways as outlined below: e The failure mode and cracking mechanism will be af fected by stress level e The magnitude and way in which the stresses are distributed will af fect crack arrest considerations e The magnitude and distribution of stressos will af fect crack growth rates i e The magnitude and distribution of stresses will af fect final condi-tions for failure The dif ferent components of stress are discussed next in the remaining subsections. These are the stresses due to rolling f abrication, those due to welding, and finally service-induced stresses.
- _ - _ - _ ~ . _ _ _ __ . _ _ _ . . . _ _ _ _ __ _ _ _ _ _
13
'3.2 1 Ro! ling Fabricatico Stresses
- (
-3.2.'1 -Analytical Determina lon '
l The residual stress state due to rolling is the remaining stress condition resulting from the extent of plastic deformation during the rolling process and the subsequent elastic rebound. An analysis of the stress state in the
.sparger pipe was performed by General Electric (6) and the result is given here in Figure 3-1. This result was used for preliminary analyses performed
^
.except that an account was taken of the actual sparger material properties at operating temperature.' As discussed in Section 4, a 550'F yield stress of 23.7 ksi was used to bound the stress distributions. Hence, the stress distributions used in the crack growth analyses assume that the maximum stress level _.in' the stress gradlent is limited to 23.7 ksi (the yield strength of the riaterial at 550*F). It should be noted that in the analysis, the origin of cracking was initially taken at the ID location which is in a
! tensile stress region. Hcwever, an additional location on the backside of the pipe was also evaluated si.nce cracks starting from that region, under certain conditions, may become scmewhat longer than those located at the ID. 3.2.2 Experimental Measurements i i An experimental program was performed by Teledyne Engineering Services, Waltham, Massachusetts, to determine fabrication stresses. The program was performed using curved pipe segments formed on the same equipment used to fabricate the actual sparger. A schematic diagram showing the multi-roll bending procedure is given in Figure 3-2, and the results are shown in l l Figure 3-3. In Figure 3-3, the measured stresses are plotted along with the
' theoretical residual stresses calculated by simple beam theory for a curved j; -beam with a circular cross-section.
l: ' 3.3 Welding Fabrication Stresses
'1 Several major prograns (2, 8, .9) have been. concerned with the magnitude and distribution of residual stresses in welded stainless steel piping. In
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i i .. i Figure 3 Resultant Residual Stress Distribution After Fabrication (6) and
. Assum'ed Relaxation to 5500F Yield (23.7 ksi max). .,
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LEGEND (Teledyne Data) e Data points o Points based on symmetry
+ Single meridional element Figure 3-3 _ Residual Stresses Due to Rolling.
17 particular, extensive data'on butt welds in four-inch diameter lines were obtained as a part of EPRI program RP449-2 (A). The findings of this study, as well as results reported in (.1A), Indicate that yleid level tensile residual stresses could exist in four-inch line butt welds. The distribution in stress was non-axisymmetric with a periodic behavior around the pipe circimference. This has several profound influences on crack growth. First, the high tensile stresses can increase the crack growth rate. Second, since residual stresses are self-equilibrating, an axisymmetric distribution will assure that compressive stresses exist _at each circumferential location at a through-walI position to balance the tensile surface stresses and thus tend to slow and possibly arrest cracks growing through the pipe wal1. For an asymmetric distribution, this assurance is not present. In particular, for the highly asymmetric four-inch lines, through-wall distributions which are completely tensile are observed. Thus, it is not generally possible to arrest surface cracks before they become through-wall. On the other hand, to balance the tensile regions, there are extensive compressive fields which can act to arrest a through-wall crack as it grows around the pipe circumference. The competing mechanisms of through-walI tensile fields and circumferential compressive locations represent the crack driving force that will determine arrest of a growing crack. L The distribution chosen for this region was that developed for butt welds in I four Inch pipes as shown in Figure 3-4. The maximum stress of 23.7 ksi is allowed. 3.4 Service Stresses The service stresses were found to be small. From (1) "all identified stresses were found to be negligible." Loadings that were considered include I impingement-loads (i.e., flow past.the spargers), seismic loadings, pressure, thermal mismatch, stagnant iIne top-to-bottom temperature gradients, stagnant l line through-wall temperature gradients and weight." These stresses are presented in Table 3-1 taken from (6).
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-so Figure 3 Maximum Surface Residual Stresses For a Four-Inch t
Pipe Inside Surface at 0.1 Inch From Edge of Fillet of Pieces 1 and 2 Versus Azimuth (10). w * " TN--- ' Tw&PP---P+ "TM M- " ' '--"4---'^-'---m-- -,e'rT--""-'m-"-4----?w- 9N w" ,ww*T*-"'wvv e -ee-* *r r r Mww+-wmw'-e---- '" e "Nwa""'--""m-v--
19 Table 3-1 Core Spray Sparger Stress Summary Assumed Load Condition Stress Type Location Magnitude (psi) Design: Impingement P pipe at bracket 106 to 509(I) b P pipe at-bracket 283 to 1352 II) Seismic (DBE) b Pressure (Ap=16 psi) - 300(3) P, l Thermal Q pipe at bracket 4411
- Negligible (2)
Water Hammer P, Normal Plant Operation Impingement P pipe at bracket 106 to 509(1) b P Pipe at bracket 236 Seismic (0BE) b Cold Spring Q bracket or junction None box Pressure (Ap=0) - 0 P, Thermal Q
- Negligible (1)The higher magnitude was determined under the assumption that a break (3600 crack) exists in the sparger pipe.
(2) Pressure increase due to water, hammer is negligible as estimated by GE. (3) Axial stress due to pressure conservatively includes bracket friction effects l as determined by GE (6).
20 The service stresses are predominately bending in nature, and are derived . mostly be secondary type loads. For design conditions, the summary of stress by type of load is given below: Primary Secondary Total Tension 300 psi 0 300 pst Bending 389-1861 psi 4411 psi 4800-6272 psi Total 689-2161 psi 4411 psi 5100-6572 psi The design conditions include postulated Emergency Core Cooling Systems (ECCS) loadings that are considered in this report as the short-term loads l under which the sparger must safely function for one occurrence. Be.ad on the above summary, the maximum total service stress is approximately 5 ksi but could be 6.6 ksi If a sparger pipe is completely severed at one junction box. Secondary stresses wilI tend to relax upon crack propasstion, therefore, the maximum design stress for limit load considerations of a partially cracked sparger pipe would be approximately 700 psi. The normal operating loads are long-term service conditions under which a potential for SCC wilI exist. These stresses are very low and are alI primary bending. The absolute stress magnitude is 342 to 745 psi. Since the l sparger pipe was f abricated outside the vessel, then lowered into the vessel as a complete unit, no installation stresses caused by cold spring are believed to exist. Hence, only the residual stresses due to pipe rolling and welding provide the driving forces for SCC. Since it was felt that some questions may arise on the magnitude of the l- long-term service stresses, a parametric study that includes dif ferent levels of applled stress in addition to the residual stress was performed. Specifically, applied axial stress levels of 5 ksi,10 ksi, and 15 ksi tension and 5 ksi bending were studied. In the analysis, the service stresses were conservatively added to the residual stress state by simple elastic superposition while maintaining the condition of maximum applied l I
) l 21 - \ stress limit of 23.7 ksi (approximete yield strength at 550*F). These stress distributions were then used as input into the fracture mechanics analysis for predicting the remaining life of the sparger pipe and the conditions under which crack arrest would occur. h
22 I I Section 4 MATERIAL PROPERTIES 4.1 Crack Growth Rates A major factor in assessing the potential for fracture of the core spray sparger during its design lifetime is the rate at which SCC growth can occur. This mechanism is of most interest since preliminary work (fi) has shown an absence of fatigue loading, and past work on the SCC of Type 304 stainless steel (10,11,12) has shown a susceptibility to this f ailure mechanism. The use of crack growth rate information is two-fold. First, if crack arrest cannot be demonstrated, it is of ten possible to demonstrate analytical lifetimes far in excess of design life, depending on the assumptions for residual stress and the state of applied stress. Second, where such Information is available from field observations, it can serve as a check on asstaption about applled stresses. Where known crack growth patterns have occurred, it is possible from final flaw dimensions to verify the proposed stress state. Several major programs have been recently completed or are currently underway (ll, .L4,15) to assess the potential growth rate for SCC in Type 304 stainless. Figure 4-1 shows typical data associated with a range of conditions which have been tested in assessing SCC in Type 304. These include various metallurgical structures and environments: furnace sensitized, weld sensitized, annealed and low temperature sensitized, and oxygen content ranging f rom 0.2 ppm to 8 ppm. Yarlous specimen y configurations were also tested to include constant load and constant K loading geometries. Given the scatter in data, it is necessary to pick those data that are most appropriate for a.glven pipe region, geanetry, stress level and environment.
F 23
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Figure 4 Crack Velocity as a Function of bax*
24 As has been discussed, the fracture mechanics assessment has examined two - distinct regions: (1) those subject to fabrication stresses due to bending, and (2) those subject to subsequent welding residual stresses. Since the data in Figure 4-1 do not represent exactly the conditions that exist at these two locations, two da/dt curves were proposed: one to represent an upper bound (worst case) behavior, and another to provide a more reasonable "best estimate" for crack growth. These two curves are shown in Figure 4-2. Both curves were used in the analysis for remaining Ilfe in order to bound the expected service life. A discussion of the use of these curves with relation to the two sparger pipe locations follows next. f 4.2 Mid-Sparger Growth Rate The pipe away from welds has seen the following history. The mill-annealed pipe was bent to shape (see Section 3.1 for details). It was then installed 4 and has been subjected to approximately 5.5 x 10 hours of 550*F exposure in a BWR environment. -The fabrication process introduced residual stresses as well as certain metallurgical changes due to cold work. The effect of cold work on the susceptibility of Type 304 stainless has bosn the subject of a 1imited amount of work (16, .12). The recent paper by Pednekar and Salatowska (.11) provides insight as to its ef fects. They estabi tsh that cold work increases the dislocation density along slip planes and leads to ' formation of martensitic platelets. Further, with increasingly larger amounts of prior cold work (maximum at 10%), subsequent sensitization treahent results in progressively greater precipitation of carbides, first along the grain boundaries, then within the grains proper. They conclude that: "Small amounts of cold work, as little as 5%, followed by sensitization, can result in extreme SCC susceptibility in Type 304 steel." With these comments in mind, it is possible to establish the following scenario for material condition in the sparger. The mIII. annealing process performed on the pipe product may not have been suf ficient to prevent carbide
- formation. The subsequent cold rolling increased dislocation density along l
slip planes and lead to the formation of martensitic platelets, thus, a condition susceptible to crack growth by SCC was achieved. Additional j
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26 4 sensitization caused by 5.5 ' x 10 hours of exposure at 550*F'could only aggrevate the existing condition. To provide an upper bound for this analysis, the growth rate associated with furnace sensitized material (.15) has been chosen for the analysis. These data were isolated in a curve that represents the behavior as shown in Figure 4-2. 4.3 Near Weld Material Condition The condition of the material near welds is better characterized, due to the bulk of research which has been directed toward weld sensitization in the last few years. The furnace sensitized growth rates are again used as an upper bound since furnace sensitized material is clearly more susceptible
- than materials subjected to weld sensitization. It is expected though that l
l the "best estimate" curve in Figure 4-2 is still a reasonable approximation l of the material behavior near the f abrication welds. , 4.4 Yield Strength The rocen temperature yield was reported to be 37,615 psi and the corresponding ultimate tensile strength is 83,660 psi (18). Given a factor of 0.63 for-Type 304 on 550*F yield strength to 75'F yield (12), a yield strength of 23.7 ksi has been taken as the 550*F yield strength to bound the stress distributions employed. A comparison with other nuclear grade heats of Type 304 (20, ID) In Table 4-1 shows the chemistry and yield strength of the core spray sparger material is consistent with other piping products. It should be noted that the carbon content of 0.048% is on the low side, making the material less susceptible to weld sensitization and, hence, less susceptible to IGSCC. l
Table 4-1
- COMPARISON OF SPARGER PIPE HEAT WITH OTHER TYPICAL HEATS OF TYPE 304 l
i HEAT "Y , NUMBER SUPPLIER .C. Mn . Ni Cr & (ksi) l . M0063 --- 0.050 1.72 10.46 18.64 38.9'
- M7616 ---
0.060 1.72 10.92 18.83 45.8 M7772 --- 0.050 1.80 10.15 18.81 0.11 41.1 454659 --- 0.045 1.25 9.76 18.40 0.23 35.4 TH6656 --- 0.060 1.72 9.30 18.31 0.24 37.0 C3 4 78500 --- 2P6396 --- 0.040 1.65 10.30 18.66 0.20 38.4 i 834264 --- 0.060 1.58 9.12 18.30 0.30 38.5 2P6424 --- 0.040 1.65 9.61 18.37 0.25 39.0
- .454970 ---
0.042 1.09 10.10 18.10 .37.6 C-15056 Curtiss-Wright 0.040 1.20 9.86 19.82 46.7 i C-1513 Curtiss-Wright 0.065 1.25 10.25 19.85 41.1 i J-2500-3* SWEPC0 0.048 1.24 9.36 18.48 0.00 37.6 l *All sparger pipes were supplied in Heat J-2500-3. . i 4 4
28 ( Section 5 i
SUMMARY
OF INSPECTION RESULTS The results of the 1980 video tape enhancement have been provided in AES Report 81-10-83, Revision 1 (1) and subsequent letter reports (21, 22). This section will summarize those data to provide a basis for the subsequent j- Interpretation of the fracture mechanics analysis. l Two general regions of Interest were established af ter review and enhancement l of the 1980 video tapes. There were Indications associated with welds and Indications welI removed from weld influences. Those regions that had l Indications associated with weld details, included several that were found not to be cracks, based on work performed during and subsequent to the 1981 inspection. The weld regions which did have Indications that could be interpreted as cracks were the B sparger junction box (Figures A-10 through A-19 of (1)) and possibly the A sparger junction box area (Figures A-27 and A.28 of (1)). For these Indications, an upper bound length of 3.5 inches was assumed in the analysis. The crack was conservatively assumed to be through l the wall of the sparger. This will provide a ' worst-case" ef fect of the crack on structural Integrity. In the other general region, away from welds (also called mid-sparger), the Indication of Interest was originally considered to be a significant Indication but during the 1981 inspection was judged not to be a crack. The , area of Interest was on the 270* side of the A sparger junction box l \ (Figures A-29 through A-32 of (1)). Despite the finding that this Indication was not a crack, the full potential impact on structural integrity was assessed by assuming a crack-like defect equal to the Indication length. This vilI obviously result in a very conservative estimate of the potential [ l lor failure in this region. l
29 Section 6 LIMIT LOAD ANALYSIS 6.1 Limit Load Model The structural integrity of the sparger pipe was assessed with the principles of limit load theory. The criteria for failure is that condition l that causes the not or _ remaining section of the pipe to become fully plastic. ! The load required to form a plastic hinge in a thin-walled pipe with a through-walI crack was determined analytically for the geometry shown in Figure 6-1. The stress strain behavior of the pipe material is assmed to behave as a rigid perfectly-plastic material. To accommodate for strain hardening, the iimit stress (o ) for the material is assumed as the average flow stress: o = (a + ou)/2 (6-1) 4 y For the values tabulated as specified minimms at 550*F for o and o in y u
- the ASE Code (21), o is approximately 45.0 ksi.
! 4 The development follows that for a surf ace flaw where the shif t in the neutral bending axis, caused by both the presence of the crack and combined tension-bending, is calculated from a relationship which satisfies force equilibrita in the longitudinal direction given below as: Um
*_~ [n - a(a/t)]
w 2
~Y$ (6-2) where o is the applied axial stress for the uncracked section and a is the a
crack hal f angle. The requirement of moment equilibrium is satisfied through
30
- NOuiN A L STRESS IN THE UNCRACKED SECTION OF PIPE
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$f RESS Ol5TRiguTION IN THE CRACKED SECTION AT THE POINT OF COLLAPSE Figure 6 Limit Load Model and Geometry.
l
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31 Integration of the stress distribution across the section and equating with the applied bending moment to give: 2 i
=
(2 sin $-fsina) (6-3) I The simultaneous solution of Eq. (6-2) with Eq. (6-3) defines the failure locus for plastic collapse of the section. Since in the derivation of these expressions an assumption was made which idealizes the sparger as a thin-wall cylinder, the location of the flaw relative to the wall thickness (i.e., j surface versus subsurface), is not a variable in the solution and the above ! expressions could be applied to subsurf ace flaws when 2a/t is substituted for l .a/t in Eq. (6-2) and Eq. (6-3) . 6.2 Numerical Results In ali'Iimit load calculations, oni.y parametric service stresses are utilized. The reason is that the secondary weld residoal and service stresses are displacement controlled and will tend to relax during crack growth. 'The service-Induced stresses to be considered in the limit load assessment should therefore be those associated with primary loading. Any l Inclusion of secondary stress into this analysis will provide for a l conservative estimate of critical flaw size. For a crack completely through the thickness of the pipe, the critical crack angle for f ailure is shown in l Figure 6-2. By simple inspection, an applied stress of only 2 ksi will clearly give critical angles much greater than 180*. For a primary service stress of 300 ps! membrane and 389 psi bending determined.from design loads calculated by General Electric (1), the critical crack angle for limit load f allure is computed as 285* or a crack that is ( 10.0 inches long. If the sparger pipe is assumed to be severed at one r junction box, then limit load f ailure would be computed for a crack about 258' (or about nine inches long) around the pipe circumference. Hence, very L
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33 fong cracks are required to f all by limit load under primary design loads that include seismic, pressure, impingement and water hammer stress events all assumed to occur at the same time. It should be noted that the maintenance action requires an external pipe clamp to be Installed when a crack-like Indication extends 180' or when the Indication extends beyond the Inspector's fIold of view. l e d l L 7v-u- - - g- -y w e - ,3w -- -- we - m a w w ew -*v- - , - y , -wrew--w - w w w-
I 34 Section 7 RESULTS OF STRUCTURAL INTEGRITY AND FRACTURE EOiANICS ANALYSES 7.1 Introduction This section discusses the calculated results for remaining service Ife and structural integrity of the core spray sparger pipe. The residual and cpplled stresses from Section 3 were combined to determine the total stress ctate. A through-wall flaw was introducad with some small Initial length in
-Sach of the relevant regions as defined in Section 5. The crack growth behavior was modeled using LEFN techniques as described in Section 2 and using the appropriate crack growth rates of Section 4. Although the long-term service stresses during normal operation are expected to be low as reported in Section 3, the of fact of service stress on crack growth was studied parmetrically. It is anticipated that the results presented herein will be used to predict expected flaw sizes given the current state of the sparger as determined by periodic inspections.
Flaws were modeled as growing by a stress corrosion mechanism to one of three final states. These are: e A length such that the flaw exceeded the limit length calculated as described in Section 6 e The creck arrested owing to the distribution and characteristics of the stress state e The expected lifetime to State 1 or 2 exceeded the plant design life-time l
-35 The two regions of Interest, one away from any sparger pipe weldnents and the
, other near the pipe welds are discussed separately. A third region was also addressed. This was away from the weld but on the "back" side (not i inspectable) of the sparger. This analysis was performed for the sparger pipe region away from any weldnents. At each region, a through-wall circumferential crack was postulated at the intrados or core-side surf ace of the sparger pipe. The crack length as a function of time was computed by Integrating Eq. (2-3). Crack arrest is predicted when the stress Intensity factor becomes zero, and the flaw lengths when this occurs are Indicated in j the results. l l I 7.2 Region Away From Welds The long Indication observed near Junction Box A but away from any welds (mid-sparger) was determined not to be a crack. However, as a bound, this region was evaluated assuning a postulated through-wall crack, and the crack growth as a function of time was calculated for a range of applied stress levels. These results showing the corresponding ef fects of stress level and material crack growth rate on calculated flaw lengths, are plotted in Figure 7-1. This figure gives flaw length as a function of time for five 1 states of stress: the theoretical residual stress state only as welI as several cases of the residual stress plus parametric service load cases. The service loads were taken at levels of 5,10, and 15 ksi membrane and 5 ksi bending stresses. The calculated arrest lengths are shown for all cases, except residual stress plus 15 ksi membrane stress which will not arrest. For all other cases, arrest will occur before critical size for that stress level is reached as calculated from not section failure criteria. l The expected crack extension for an 18-month period was ~ determined f rom the ( results in Figure 7-1. For a given observed flaw size, the flaw size that would be observed af ter 18 months of additional service is shown in Figure 7-2. These results are based on the upper bound crack growth behavior f and can,be used to predict conservatively the amount of crack growth between l ref ueling outages if the time span between Inspections is 18 months or less.
10 M 10 TO THE CFNTIMETEf1 te u .' " 46 1510 14 E UF F E L (e f ME R LO mootoguga 8
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.L_ 2 Figure 7 Effect of Stress Level and Material Crack Growth Rate on Flaw Length vs. Time For Front-Side Indication Away From Weld Region.
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Figure 7 Predicted Crack Growth For Indications on Front-Side l of Sparger Pipe Away From Welds For 18 Months of Operation. 1
38 For example, an observed flow length of 3.2 inches would grow to approximately 3.3 inches if the applied stresses in that region are 5 ksi tension. The 10 ksi and 15 ksi applied stress curves would predict a final flew size of about 3.8 and 5.5 inches, respectively. The point of tangency between the curve and the 45-degree line represents the point where the crack is predicted to arrest. 7.3 Region Near The Weld A similar analysis was performed for a flew postulated to exist near a sparger pipe weld connection. The weld residual stress distribution was assumed to have a sinusoidal shape along the pipe circumference having a peak stress of yield strength level and a period equal to the pipe circumference as discussed in Section 3. The results for this analysis are given in Figure 7-3. A comparison of Figure 7-3 with Figure 7-1 Indicate that much longer cracks are possible in .the regions influenced by the weld. Calculations for the residual stress only case show crack arrest at a final
~ flaw length of about 9.8 inches. Crack lengths greater than 10 inches are possible for the applied stress level as shown in Figure 7-3. Crack arrest for the cases of 5 ksi,10 ksi, and 15 ksi is still possible if the applied service stresses are secondary in nature. The potential for limit load f ailure exists if primary stresses are present; however, for the primary stresses specified by General Electric (see Section 3), the critical flew size for limit load conditions could reach 10 Inches or a pipe with a 285*
circunferential crack. The predicted crack growth for an 18-month period based of operation as a f unction of the length of the observed Indications is shown in Figure 7-4. In this figure, both upper bound and mean crack growth rate results are plotted for comparison. The final flaw length is very sensitive to the material behavior and a wide range in crack length can be predicted depending on the crack growth rate assumed. For example, as a best estimate, a flaw of length 3.5 (can grow to approximately 4.4 inches over 18 months, or it can grow to as long as 8.4 inches under worst case material behavior, it is in - - _ _ _ . _ _
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41 Important that these predictions be used in conjunction with multiple inspection data gathered over a long period of time. Such comparisons will validate the evaluation method and provide a means of calibrating the model to the field observations. Comparison between the inspections of January 1 1980 and October 1981 Indicates insignificant changes in the observed Indications. Clearly, this suggests that the fracture mechanics model and upper bound input data are providing a conservative estimate for the amount of crack extension that could occur. 7.4 Back-Side Defect Evaluation l The potentialifor flaws to initiate and grow on the back-side of the sparger pipe was evaluated. The results obtained are given in Table 7-1. assuming upper bound crack growth rate behavior. If the measured tensile stresses on back-side (see Figure 3-3) era an Indication of a local surf ace stress
, ef fect, then' the theoretical stress distribution would be more representative i for long flaws. This model predicts crack arrest when the crack is visible
(~210'). However' , a conservative assessment of the tensile stresses "as measured" on the back-side does not predict crack arrest. It should be noted
~
that the limited test data shown in Figure 3-3 show considerable scatter, particularly on the back-side of the pipe. Because " hand Jacking" was used for final pipe curvature adjustments, it is believed that non-uniform and highly localized stresses may exist in the sparger pipe. Table 7-2 gives the l remaining tNae for cracking beyond 180', at which point the back-side crack just becomes visible from the front side. l l i
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A 14 { i; \; .: 42 [ N, , y e s; y t TABLE 7-1 s
SUMMARY
OF CRACK GROWTH PREDICTIONS
. FOR.BACK-SIDE FLAW TIME (MONTHS)
CRACK GROWTH . AS-f1EASURED II) THEORETICAL 2a o rONLY +5Ksi o ONLY o'p+5Ksi r r y 70 to 1800 :1484 160 _(2) _(2) x, 1800 to 360 1928 37.1 Arresgs Arresgs 6 204 0 226 1 e 1 See Figure 3-3.for assumed stress distribution. 2 Theoretical distribution gives negative stress on back-side so cracking will not initiate. If measured tensile surface stresses are local, as expected, the distribution shown in Figure 3-3 will also result in limited crack growth. FLAW N0 DEL
'$*y si , M ( !Ne
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43 i TABLE 7-2 COMPUTED REMAINING TIME FOR A VISIBLE BACK-SIDE FLAW INDICATION REMAININGTIME(MONTHS) AS-MEASURED (1) THEORETICAL CRACK ANGLE RESIDUAL RESIDUAL RESIDUALS RESIDUAL 2a OflLY +5Ksi ONLY +5Ksi 180 1928 37.1 112 38.8 0 36.8 183 1927 - 193 - - 3.0 36.8 204 - - 0' - 0 32.2 - 218 1913 226 - - 0
- 259 1770 22.0 308 14.5 6.6 360 0 0 (NoArrest) (No Arrest) 1 See Figure 3-3 for stress distribution.
FLAW t10 DEL Back nt ide g2 .
44 Section 8
SUMMARY
AND CONCLUSIONS e Fracture mechanics analyses have been performed for defects postu-lated to extend through wall in two general regions of the Pilgrim Station core spray sparger. These regions are away from welds and near welds. e The following conservatisms were present in this analysis: e Upper bound crack growth rates are used e Bounding stress fields are used e Assumption of through-wall cracks even with regions where Indica-tions were determined not to be cracks e A through-wall crack located away from a weld will arrest given
" service" stresses less than 10 ksi membrane tension. Further, this arrest length is shorter than the critical length to cause failure by a net section mechanism under the conservative assumption that the service-induced stress is due to a primary load.
e A througn-wall crack located in the region of a weld detail will arrest prior to reaching critical length to failure, given the low applied service stresses as specified by General Electric. Crack growth beyond approximately 10 inches will not assure protection against limit load failure; however, constraint in displacement and rotation may limit the concern for plastic collapse. The length of time required for substantial crack growth in the weld region may also allow a reasonable period from detection to arrest or potential
45 failure. As a precautionary measure, the maintenance action plan requires the Installation of an external pipe clamp when the length of the crack Indication is found or predicted to be greater than 180* l (6.3 Inches). l e The Identified Indications do not jeopardize structural integrity of the core spray sparger over the next 18 months. l e Crack growth between inspection periods (i.e.,18-month Interval) is l predicted to be small. For the Indication lengths observed away from any welds, crack extension is less than 0.1 Inch of growth. Near the sparger pipe weldments, it is predicted that the maximum crack growth will be less than 1 inch. l l l i I
1 46 REFERENES
- 1. Hayward, J. A., R.C. ClpolIa, and A.J. Sanders, " Computer Enhancement of Yideo Tapes of ISI of Pilgrim Core Spray Sparger", Phase 1, Preliminary Report, Revision 1, Aptoch Engineering Services, AES-81-04-60 (April 1981).
- 2. " Mechanical Fracture Predictions for Sensitized Stainless Steel Piping With Circumferential Cracks," Final Report, EPRI Research Project RP 585-1, Report NP-192 (June 30,1976).
- 3. Egan, G.R. and R.C. Cipolla, " Stress Corrosion Crack Growth and Fracture Predictions for BWR Piping," American Society of Mechanical Engineers, Paper 78-Mat-23 (1979).
- 4. Paris, P.C., M.P. Gomez and W.D. Anderson, "A Rational Analytical Theory of Fatt gue," .TAa Trand h Fnnt naarIng, Volune 13, No.1 (January 1961).
- 5. Besuner, P.M. , D.C. Peters, and R.C. Cipoi la, "BIGIF: Fracture Mechanics Code for Structures," EPRI Report NP-838 (July 1978).
- 6. "Supplamont I to Supplemental Reioad LIconsing Submittal for PIIgrim Nuclear Power Station Unit 1 Reload 4," General Electric Nuclear Power Systems Division, NED0-24224-1, Supplement 1 (March 1980) (CD-2).
- 7. Shack, W.J., W. A. Ellington and L.E. Pahls, "The Measurement of Residual Stress in Type 304 Stainless Steel Piping Butt Weldnents," Electric Power Research Institute, Finst Report No. RP449-1 (December 1978).
- 8. Giannuzzi, A.J., " Studies on AISI Type 304 Stainless Steel Piping Weldnents for Use in BWR Applications," EPRI NP-944, Research Project 449-2, Final Report.
- 9. Rybicki, E.F., P.M. McGuire and R.B. Stonest for, "Ef fect of Weld Parameters on Residual Stresses in Bolling Water Reactor Piping Systems," EPRI, Semlannual Report RP 1174, April 1,1978 to October 1,1978.
- 10. Klopfer, H.H., et al., " investigation of Cause of Cracking in Austenttic Stainless Steel Piping," General Electric Report NEDO-21000, 75-NED35, Class 1 (July 1975).
- 11. Clarke, W.L. and G.M. Gordon, " investigation of Stress Corrosion Cracking Susceptibility of Fe-NI-Cr Alloys in Nuclear Reactor Water Envirorunents," Corrosion, Volume 29, No.1, pp.1-12 (January 1973).
- 12. Park, J.Y., and W.J. Shack, " Corrosion Studies of Nuclear Piping in BWR Environments," Semi-Annual Report for Period ending 31 July 1979, EPRI
, Research Project 449-1.
1 47 l 13. Caligluri, R.D., et al., " Low Temperature Sensitization of Weld Heat Af fected Zones in Type 304 Stainless Steel," Stanford Research Institute, EPRI Contract T-110-1.
- 14. Park, J.Y., and W.J. Shack, " Crack Propagation Rate Studies," Argonne National Laboratories, EPRI Contract T117-1.
- 15. Horn, R.M., et al., " Crack Growth / Arrest Studies," General Electric Company, San Jose, EPRI Contract T118-1.
- 16. " Basic Studies on the Variables of Fabrication Related Sensitization Phenomena in Stainless Steels," General Electric Company, Schenectady, New York, EPRI Research Project RP 1072-1.
- 17. Pednekar, S. and S. Salalowska, "The Ef feet of Prior Cold work on the Degree of Sensitization in Type 304 Stainless Steel," Corrosion, 0010-9312/80/000189, p. 565-577.
- 18. O'Connor, H.W., Trip Report entitled " Core Spray Sparger Visit to Sun Ship Works and Philadelphia Pipe Bending," (April 14, 1981).
- 19. Peckner, D. and I.M. Bernstein, Handbook of Stainless steals, McGraw-Hili Book Company, New York (1977).
- 20. Egan, G.R. , R.C. Cipoi la and W.P. McNaughton, " Stress Corrosion Cracking Testing Methods Appilcability to Plant Performance," AES 81-07-77, Final Report EPRI Project 1562-1.
- 21. Letter report from Aptoch Engineering Services to F.N. Famulari, Manager, Pilgrim Nuclear Power Station, dated October 29, 1981.
- 22. Letter report from Aptoch Engineering Servics to F.N. Famulari, Manager, Pilgrim Nuclear Power Station, dated November 10, 1981.
- 23. American Society of Mechanical Engineers, Boller and Pressure Vessel Code, Section Ill,1980 Edition.
i
w - 48 APPENDIX PROPOSED PILGRIM 00RE SPRAY SPARGER EVALUATION PLAN I
\ 49 APTECH cnginacting farvici,Inc .N21NeenlN2 CON:oL1AN1S 795 SAN ANTONIO ROAD . PALO ALTO . CALIFORNIA 94303 (415)858 2863 Pag 2 1 PROPOSED PILGRIM CORE SPRAY SPARGER EVALUATION PLAN Areas of Interest
- 1. Sparger pipe away from weld
- 2. Sparger pipe to junction box weld HAZ (all welds)
- 3. Nozzle weld HAZ (all welds)
- 4. Core spray jumper pipe weld HAZ (all welds)
Strategy for identification of crack indication, crack growth rate, and crack arrest
- 1. The sparger crack indication identified during the 1980 ISI shall form the basis of reference cracks. Image enhanced photographs taken from video tape records will be used for this purpose.
- 2. It is essential that reference cracks be established for determining the rate of crack growth. In order to achieve this goal with some degree of confidence and accuracy, approximately 10 reference indications will be identified for thorough examination during the 1981 ISI. Inspection techniques will include the capability to measure crack length and width.
- 3. Since the quality of the 1981 ISI will be greatly improved and computer image enhancement information will be available, it is exgected that the extent of cracking can be quantified within approximately 180 of the exposed surface.
Accordingly, previously undetected or recently fomed cracking may be found.
- 4. Comparison of 1981 ISI results to 1980 reference indications will provide improved fracture mechanics evaluation basis for crack growth. The recent work by APTECH will be refined to include these data for crack growth and arrest predictions for fuel cycle 6 (March 1983). Structural integrity evaluations will be modified as necessary.
Maintenance Action Criteria Maintenance action decisions will be made on a case by case basis and may be dependent upon the ability to reliably predict crack growth rate for fuel cycle 6. In addition, there exists some uncertainty in the nature of residual stresses induced by sparger pipe fabrication and rolling. Therefore, cracking patterns In particular, the may not be identical in different regions along)the pipe. condition of the backside of the pipe experimentally measured tensile stresses of varying magnitudes where crack propagation could continue. That area is on the blind side (extrados) and cannot be inspected visually. Based upon the brief discussion above and analytical and experimental work performed or reviewed by APTECH, the following maintenance criteria are recomended Services in Mechanical and Metallurgical Engineering, Welding, Corrosion, Fracture Mechanics, Stress Analysis
. - - - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ \
APMHh SERVICES, INC, 50 m SAN ANTONIO ROAD Page 2 PALO ALTO. CALIFORNIA 94303
.to assure sparger operability throughout fuel cycle 6.
External Pipe Clamp Fix Ext:rnal sparger pipe clamps shall be designed so as to satisfy " full operability" requirements upon installation. The circumstances under which clamps will be installed over a crack in a local region on the sparger or jumper pipe are as follows:
- 1. In the event that crack arrest is not predicted for a given crack.
- 2. In the event that a crack indication is found or predicted to extend for 180 It cannot be demonstrated in the circumferential direction in the field of view.
by visual inspection methods that cracking does not continue around the backside of the pipe.
- 3. In the event that circumferential cracksIt are found at or near the top or bottom cannot be demonstrated by visual of the pipe within the field of view.
inspection methods that cracking does not continue around the backside of the pipe.
- 4. In the event that sparger design function and operability criteria are exceeded .
(flow rate, momentum and spray pattern operability criteria are being developed).*
- 5. If a significant number of areas are determined to require clamping based upon the above criteria, an optional air test of the sparger assembly may be performed to identify areas of through-wall cracking.
Replace Sparger Assembly Based upon present engineering knowledge, including empirical data and analytical model results, we do not predict or forsee any situation where the application of clamps will not provide a sound engineering fix. Clamping where warranted in local areas is considered to be the primary fix and the need for sparger replacement is not anticipated. However, the following maintenance criteria is proposed as a contingency. Replacement of the sparger, or a sparger pipe segment, will be implemented only if clamping is not practical or feasible. NOTE: APTECH has not performed a structural integrity evaluation However, cracknor operability / propagation safety evaluations of the 304ss jumper pipe system. studies performed for the sparger juncture box weld region are representative and applicable to jumper pipe welds which were not heat treated after welding.
- See pages 5-10.
APTECH GDDIHEALO ' SERVICES, INC, Page 3 51 ns SAN ANTONIO ROAD PALO ALTT, CAUFORNIA 94303 JAN. 1980 ISI INDICATIONS it CORE SPRAY SPARGER SAFETY EVALUATION
* ,r MATERIAL AND ISI VIDEO TAPE SERVICE WELD / FABRICATION IMAGE ENHANCEMENT STRESS RESIDUAL STRESS EVALUATION EVALUATION EVALUATION I
1r p IDENTIFY FRACTURE RELEVANT MECHANICS INDICATIONS EVALUATION 1r CRACK GROWTH / ARREST PREDICTIONS I 1lr 4 STRUCTURAL SPARGER INTEGRITY DESIGN FUNCTION EVALUATION EVALUATION 1r SPARGER SAFETY AND OPERABILITY EVALUATION ir ACTION PLAN FOR SEPT. 1981 ISI
?
Figure 1 - Pilgrim Core Spray Sparger Evaluation Flow Chart
APTECH ENGINEERING Page 4 SERVICES, IN3. 52 796 SAN ANTONIO ROAD PAL"> ALTG CAUFORNIA 94303 SEPT. 1981 CORE SPRAY SPARGER ISI AND EVALUATION ACTION PLAN u STATE OF ART VIDE 0 NDE AND COMPUTER IttAGE ENHANCEMENT
<r IDENTIFY RELEVANT INDICATIONS I
1 ,r C0tiPARIS0N WITH NEW 1980 REFERENCE INDICATIONS INDICATIONS 1P 1P I CRACK GROWTH REVISED FRACTURE PREDICTIONS BASED UPON MECHANICS CRACK OBSERVED da/dt GROWTH EVALUATION 1 ,r STRUCTURAL SPARGER INTEGRITY FUNCTION EVALUATION EVALUATION I I 4 SPARGER SAFETY AND OPERABILITY EVALUATION Unacceptable Acceptable 4F MAINTENANCE ACTION NO REQUIRED ACTION Figure 2 - Pilgrim Core Spray Evaluation Flow Chart
~
APTECH ENGINEERINQ SERVICES, INC. 53 ns SAN ANTcNio r.oAo PALO ALTO, CALIFORNIA 94303 Page 5 PROPOSED PILGRIM CORE SPRAY SPARGER EVALUATION PLAN Sparger Function Evaluation Sparger functional requirements include core reflood, heat transfer, sparger nozzle flow rates and sparger spray pattern. This evaluation study includes only the latter two items. The first two are being addressed by General Electric. Sparger Nozzle Flow Rate A fluid mechanics study was performed which considered the flow rate design requirements of 3600 gpm per sparger circle and the actual pump capacity reserve margin. Once those values were established, the pump system performance curves provided a besir for establishing the change in flow rates as a function of crack area. Pump performance curves are shown in Figure 1. Figure 2 shows expanded pump curves in the area of interest. This figure illustrates that the reserve pumping capacity is approximately 20% greater (4300 gpm compared to 3600 gpm) than the design flow requirement for the case of no leakage, (i.e. when no cracks are present). However, as crack leakage increases, the calculations show that a maximum leakage flow of 1400 gpa . (5000 gpm minus 3600 gpm) could be present and the pump would still provide flow through the nozzles at the design rated flow of 3600 gpm. For conservatism and to account for other system uncertainties, it is recommended that the leak criteria should be based upon half of the reserve pump capacity. The operating poirt would then be at point 3 in Figure 2. The nozzle flow would be about 3960 gpm, and the crack leakage would be 740 gpm. Table 1 summarizes the evaluation and presents the results in a parametric form of acceptable crack length versus crack width, based upon 50% of the reserve pump capacity. Table 2 shows the resulting clamp fix criteria. The system can tolerate a crack .010 in, wide and 57 feet long, which is equivalent to adding about 40 small nozzles to the sparger system. These calculations indicate that it is highly unlikely that crack clamping will be required because of insufficient flow capacity due to crack leakage.
] Nozzle Spray Distribution Fluid Mechanics calculations were performed to determine the relative q
effects of flow through a typical sparger nozzle as compared to a crack of varying length and width. Parametric calculations' are summarized in Table 2 as a function of ratios of velocity, flow rate and momentum.
/ APTECH ENGmurunu CERVICES, thC. g 795 cAN ANTCNI3 ROAD PALo ALTO, CAUFORNIA 94303 Page 6
=
Sparger inspection data from Oyster Creek indicated that the largest A measured circumferential crack was approximately .010 in, wide. i' boundingevaluationwaspgrformedassumingthatcrackwidthandaat the intrados, l locat circumferential crack 180The Rm value from Table 1, divided by the crack smallest nozzles. length, gives a momentum flow of approximately 35% of the magnitude for one Fulljet IM12 nozzle. Since there are 56 of these nozzles and 56 i larger nozzles in each sparger circle, the effect of bounding crack flow is seen to be negligible and would be expected to affect spray distribution only in a local region within a radial distance of a few feet from the nozzle sparger. Interference with nozzle spray pattern is only possible for through-wall Since structural criteria dictates that cragks 180 at the sparger intrados. cracks (5.5 in) shall be clamped, only smaller cracks at the sparger intrados need be considered. Rather than postulating a bounding crack, consider a more realistic circumferential crack length of 4 inches as predicted by fracture mechanics analysis based upon measured residual stress distribution due to pipe fabrication. Since IGSCC cracks are inherrently tight, assume a crack width of .004 in. Calculated momentum flow from Table 1 is approximately 8% of the momentum flow of nne Fulljet IM12 nozzle. This study indicates that an individual crack will have a negligible effect on the core spray pattern. However, the spray pattern could be affected if a number of through wall cracks were present ir. a local region between two nozzles. proposed Clamp Fix Extension for Spray Pattern An external pipe clamp will be installed over cracks when the cumulative crack lengths and widths result in total flow momentum equal to or greater than the flow momentum of the two adjacent nozzles. 1
Page 7 SERVICES, INC. 55 ns SAN ANTONIO ROAD PALO ALTO, CAUFORNIA M303 o CRACK WIDTH CRACK FLOW 10% RESEPVE CAPACITY (INCH) VELOCITY CRACK LENGTH * (FT/SEC) (INCH) 0.001 2.92 81,000 0.002 11.0 10,800 0.004 29.2 2,000 0.010 33.4 690 ' 1 0.015 37.8 420
- Length needed for 740 GPM crack flow.
TABLE 1 Core Spray Sparger Pump Leak Rate Criteria F 1 I
l l Page 8 . y 56
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d ' / 9gg CONg I I i e ! 9 J Design Flow 3960 ,'4300 i 4700 l s 400 , % N N ; 2500 3000 3500 4000 4500 5000 FLOW RATE GALLONS PER MINUTE 1 FIGURE 2 Pilgrim Core Spray System Pump S Performance Curves
58
;e 10 TABt.E 2 PILGRIM CORE SPRAY SPARGER e Crack Leak Rate Evaluation ASSUMPTIONS _: m Pressure Drop = 15 psi .,
Nozzle: Flow * = 17.1 GPM = 0.0381 ft.#/sec. Velocity = 31.7 Ft/sec. Homentum Rate = 1.21 p LBM Ft./sec.p Where p = Density in LBM/Ft.3 Crack: AP = 15 psi Length : 1 inch Depth : 0.226 inch R Flow R M Velocity Ry g Type Crack Width Crack Flow inch Ft./sec. 20,400 Laninar 10.9 1877 t 0.001 2.92 Laminar 249 712 11.0 2.9 0.002 51.3 Laminar 1.1 27 , 0.004 29.2 Turbulent 0.95 16 15.6 0.010 33.4 Turbulent 0.84 9.7 8.1 0.015 37.8 Turbulent 0.76 5.4 4.0 0.025 41.8 Nozzle Velocity _ R V
= Crack Velocity Nozzle Flow Rate _
R = Q Crack Flow Rate Nozzle Momentum Flow b~ Crack Momentum Flow APTECH ENGINEERING SERVICES, INC.
*: Fulljet Nozzle Catalog, page 14.
IM12 has 17.1 gal./ min. 0 15 psi res SAN ANTONIO ROAD PALO ALTO CAUFORMLA 94303 Nozzle Orifice 15/32 in dia. (112 nozzles in two full circles) _ _ - _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _}}