ML19309D014
| ML19309D014 | |
| Person / Time | |
|---|---|
| Site: | Pilgrim |
| Issue date: | 03/27/1980 |
| From: | Brandon R, Gridley R, Kiss E GENERAL ELECTRIC CO. |
| To: | |
| Shared Package | |
| ML19309D008 | List: |
| References | |
| 80NED262, NEDO-24224-1-S1, NUDOCS 8004090476 | |
| Download: ML19309D014 (107) | |
Text
{{#Wiki_filter::= _. . - . . O NEDO-24224-1 Supplement 1 80NED262 Class I March 1980 SUPPLEMENT 1 TO SUPPLEMENTAL RELOAD LICENSING SUBMITTAL FOR PILGRIM NUCLEAR POWER STATION UNIT 1 RELOAD 4 Approved: 4 3 p/go b Approved: , yf7/,'
- O MD R. J. andon, Manager E. Kiss, Manager Nucle r Services Engineering Applied Mechanics Nuclear Fuel and Services Nuclear Technology Department Engineering Department Approved: y[27/fo R. L. Gr diey, M ager Fuel and Services Licensing Safety and Licensing Operation NUCLEAR POWER SYSTEMS olVISION e GENERAL ELECTRIC COMPANY SAN JOSE, CALIFORNI A 96125 GENER AL $ ELECTRIC 80040 90 g f(
- a NECO-24224-1 IMPORTANT NOTICE REGARDING CONTENTS OF THIS REPORT PLEASE READ CAREFULLY This report wa~ epared by General Electric solely for Boston Edison Company (BECo) ror RECo's use with the U.S. Nuclear Regulatory Commission (USNRC) for amending BECo's operating license of the Pilgrim Nuclear Power S** . .. The information contained in this /eport is believed by General Electric to be an accurate and true representation of the facts known, obtained or provided to General Electric at the time this report was prepared.
The only undertakings of the General Electric Company respecting infor-mation in this document are contained in the General Electric Company Proposal No. 416-TY293-HKI, (GE letter No. G-HK-0-33, dated February 19, 1980 and GE letter No. G-HK-0-52, dated March 19, 1980). The use of this information except as cafined by said contract or for any purpose other than that for which it is intended, is not authorized; and with respect to any such unauthorized use, neither General Electric Company
, nor any of the contributors to this document makes any representation or warranty (express or implied) as to the completeness, accuracy or usefulness of the information contained in this document or that such use of such information may not infringe provately owned rights; nor do they assume any responsibility for liability of damage of any kind which may result from such use of such information.
l ii
NEDO-24224-1 TABLE OF CONTENTS Page
- 1. INTRODUCTION AND
SUMMARY
l-1
- 2. CHANGES TO NED0-24224 2-1 APPENDIX G - CORE SPRAY SPARGER STRUCTURAL INTEGRITY G-1 G.1 Sparger Configuration G- 1 G.2 Fabrication Sequence G-2 G.3 Installation Sequence G-3 G.4 Performance History G-4 G.5 Potential Sources of Stress G-5 G.5.1 Fabrication Stresses G-5 G.5.2 Installation Stresses G-10 G.5.3 Stresses During Normal Operation G-ll G.5.4 Stresses During Core Spray Injection G-13 G . 5 . 4 .1 Water Hammer Loads G-13 G.6 MaterialsAspectsofCracking G-16 G.6.1 Potential Causes of Cracking G-16 G.6.2 Effects of Cold Work on IGSCC G-18 of Stainless Steels G.6.3 Conclusions on Causes of Sparger Cracking G-20 G.7 Crack Arrest Assessment G-21 G.7.1 Stresses Due to Bracket Restraint G-21 G.7.2 Fabrication Residual Stresses G-22 G.7.3 Conclusions on Crack Arrest G-24 G.8 Structural Integrity With Cracks G-24 G.9 References G-25 G.10 Stress Analysis of the Pilgrim Station G-46 Core Spray Sparger '
APPENDIX H - LOOSE PIECE ANALYSIS H-1 H.1 Introduction H-1 H.2 Loose Piece Description H-1
~
H.3 Safety Concerns . H-1 l H.4 Safety Evaluation H-1 l H.4.1 , General Description H-2 1 H.4.2 Postulated Loose Pieces H-2 iii
^'
NEDO-24224-1 TABLE OF CCNTENTS (Continued) v Page H.4.2.1 Sparger Pipe H-2 H.4.2.2 Spray No :le H-3 H.4.2.3 Small Pieces H4 H.5 Conclusions H-7 H.6 References H-7 H. 7 Flow '/elocity Calculations H-8 APPENDIX I - LOSS-0F-COOLANT ACCIDENT ANALYSIS WITH NO CORE SPRAY HEAT TRANSFER CREDIT I-l I I.1 Introduction I-l I.2 Input Changes to LOCA Analysis I-l I.3 Depressurization Rate Sensitivity I-2 I.4 Analysis Results 1-2 I.4.1 Large Break Analysis I-2 I.4.2 Small Break Analysis I-3 I.5 Conclusions I-3 I.6 Re ferences I-4 l l t 1 l iv
=
NEDO-24224-1 LIST OF TABLES Number Title Pace G.6.1-1 Possible Causes of Cracking G-26 I-l Pilgrim Large areak Results I-5 I-2 Pilgrim Small Break Uncovery Time I-6 LIST OF ILLUSTRATIONS Numoer Title a ge P_a G.1-1 Core Spray Sparger - Elevation View G-27 G.1-2 Core Spray Sparger - Plan View G.1-3 G-28 Sparger to Shroud Attachment Method g.29 G.1-4 Sparger Nozzles G.1-5 G-30 Sparger Support Method G.2-1 G-31 Pipe Bending Method G.5.1-1 G-32 Sequence of Events Leading to the Residual Stress Distribution G-33 G.5.1-2 Bilinear Stress-Strain curves for Type 304 Stainless Steel G.5:1-3 G-34 Stress and Strain Distribution in the Pipe Under Applied Moment G.5.1-4 G-35 Moment vs Outer Fiber Strain G-36 G.5.1-5 Resultant Residual Stress Distribution After Fabrication G- 37 G.5.2-1 Postulated Installation Stresses G.6.2-1 G-38 Effects of Cold Work on IGSCC of Type 304 Stainless Steel g_39 G.6.2-2 Effects of Cold Work on IGSCC G.6.2-3 G 40 Effects of Cold Work on IGSCC g.41 G.6.2-4 Effects of Cold Work on IGSCC G.6.2-5 G-42 Stress vs Percent Cold Work G-43 G.7.1-1 Compliance Change, Cracked Pipe g.44 G.7.1-2 Assumed Stress Distribution on the-Crack Face H-1 G-45 Reactor Vessel H-2 Steam Separator H-13 H-3 H-14 Largest Piece That Can Fit Thr6 ugh Turning Vane with Long Oimension in the Horizontal Plane H-15 H-4 Orificed Fuel Support H-16 H-5 - Fuel Assemblies and Control Rod Module H-17 H-6 Flow Paths H-18 V
. o ,
NEDO-24224-1
- l. INTRODUCTION AND
SUMMARY
One of the scheduled tasks during the Reload 4 refueling and maintenance outage in January, 1980 at the Pilgrim Nuclear Power Station was the performance of a visual inspection of the Core Spray Spargers using underwater tel-evision cameras. This inspection was conducted as part of the 10 year Inservice Inspection Program at Pilgrim Station, and also as recommended by General Electric in Service Information Letter (SIL) 289 (Reference 1-1). During the inspection of the core spray spargers, indications were observed on the upper and lower sparger headers in the area adjacent to the T-box. The discovery of these indications was reported to the Nuclear Regulatory Commission (NRC) on February 1, 1980 (Reference 1-2). On February 29, 1980 representatives of the Boston Edison Company (BEco) and General Electric (GE) met with members of the NRC Staff to describe the observed indications and describe the intended course of action (Reference 1-3). On March 13, 1980, BEco and GE met with members of the Staff to present the technical justification to support the startup plan proposed at the February 29, 1980 meeting. This presentation (Reference 1-4) included the technical basis to establish the continued structural integrity of the core spray spargers for all normal, transient, and accident conditions. A discussion of the methods and assumptions to be used in the Loss-of-Coolant Accident (LOCA) analysis to determine the MAPLHGRs, assuming no credit for core spray heat transfer, was also presented. In addition, a discussion of the possible consequences of a potential loose piece from a broken sparger was presented at the request of the Staff. This supplement to the Pilgrim Reload 4 Licensing Submittal (Reference 1-5) documents the information presented at the February 29 and March 13, 1980 meetings with the Staff to support operation of the plant with the reported core spray sparger indications. The resulting changes l to the original Pilgrim Reload 4 Licensing Submittal (Reference 1-5) l l 1-1
o > NED0-24224-1
- are given in Section 2 in this report, and are supported by the infor-mation presented in Appendices G, H, and I.
4 Appendix G describes the potential sources of stresses in the soargers resulting from fabrication, installation, normal operation, and operation during postulated Loss of-Coolant Accidents (LOCA). Potential causes of cracking are also discussed, and it is concluded that the structural integrity of the sparger will be maintained for all conditions of operation. 4 However, if breakage of the sparger is postulated, the loose piece evaluation presented in Appendix H concludes that the potentia.1 for unacceptable flow blockage of a fuel assembly, or for unacceptable c control rod interference, is essentially zero. It is also shown that . loose pieces are not expected to cause damage to the other reactor pressure vessel internals. Appendix I presents the results of LOCA analyses assuming no core spray i heat transfer credit in the calculations. This corresponds to a postulated worst case core spray sparger break in which the water flowing into the spargers does not spray onto the core, but rather floods into the shroud from the broken sparger. The resulting MAPLHGRs were calculated at the request of Boston Edison, but are considered by General Electric to be excessively conservative, based on the calculations which support the continued structural integrity of the spargers, and the many conservatisms in the current LOCA models. References 1-1 Inspection of Core Spray Spargers, General Electric Service Information Letter (SIL) No. 289, dated January 1979. 1-2 Licensing Event Report (LER) No. 80-004/012-0, dated February 1, 1980 1-2
NED0-24224-1 1-3 Boston Edison Company meeting witn the NRC, Bethesda, MD, February 29, 1980 1-4 Boston Edison Company meeting with the NRC, Bethesda, MD, March 13, 1980. 1-5 Supplemental Reload Licensing Sucmittal for Pilgrim Nuclear Power Station Unit 1 Reload 4, NED0-24224, dated November 1979. 6 a O J t i . I l-3
o - I NECO-24224-1 3
- 2. CHANGES TO NECO-24224 i
l ihe reload reanalysis results for Cycle 5 operation, assuming no credit ' for core spray heat transfer, are presented below using the same format as NE00-24224. These sections are intended to replace the corresponding sections in NEDO-24224. The results presented in the other sections of NEDO-24224 are unchanged. i t I 1 4 7 2-1
NEDO-24224-1
- 1. PLANT-UNIOUE ITEMS (1.0)*
Margin to Opening of Unpiped Spring Safety Valves: Appendix A GETAB Analysis Initial Conditions: Appendix 3 ATWS Recirculation Pump Trip: Appendix C New Bundle Leading Error Analyses Procedures: Appendix 0 Linear Heat Generation Rate for Bundle Loading Error: Appendix E Densification Power Spiking: Appendix F Core Spray Sparger Structural Integrity: Appendix G Loose Parts Analysis: Appendix H Loss-of-Coolant Accident Analysis with no Core Spray Heat Transfer Credit: Appendix I
- 14. LOSS OF COOLANT ACCIDENT RESULTS "*(5.5.2)
For the 5 fuel types in the Pilgrim Reload 4 core, MAPLHGR multipliers in the Table 14-1 below will be applied to the MAPLHGR values reported in Section 14 in NEDO-24224. The multipliers were developed assuming no core spray heat transfer credit in the LOCA analysis, as described in - Appendix I. Table 14-1 MAPLHGR Multioliers Assuming No Core Soray Heat Transfer Credit Fuel Type Core Flow t 90% Rated Core Flow < 90% Rated 808219L 0.93 0.85 808219H 0.93 0.85 , 808262 0.94 0.86 P80RB265L 0.91 0.84 P80RB282H 0.92 0.85 l *( ) refers to areas of discussion in " General Electric Boiling Water l Reactor - Generic Reload Fuel Application, "NEDE-24011-A-1, l July 1979. 2-2 .
NED0-24224-1 APPENDIX G CORE SPRAY SPARGER STRUCTURAL INTEGRITY G.1 Soarcer Conficuration The core spray sparger configuration is shown in Figures G.1-1 through G.1-5. The spargers are mounted in the upper shroud, as shown in Figure G.1-1. Vertical spacing is 10 inches between header pipe centerlines. The upper sparger has bottom-mounted nozzles and the lower sparger has top-mounted elbows. The plan view (Figure G.1-2) shows that the spargers are asymmetric. The shorter header pipe has an arc length of 75*, and the longer header pipe has an arc length of 105 . The T-boxes for the spargers are located 15 from the vessel 0* and 180 azimuths. Figu,e G.1-3 shows the attachment of the T-box to the shroud. ' The T-box is a 5" schedule 40 section of pipe with an end plate toward the vessel centerline. The 5" pipe extends through the shroud wall and is butt welded to external piping. The T-box pipe is attached to the shroud by the seal ring with the attachment welds to the 5" pipe and the exterior surface of the shroud wall. The sparger flow nozzles are shown in Figure G.1-4. The picture is not correct as shown. It shows two types of nozzles, one on each sparger. Actudlly there are two types, as shown, except that at any azimuthal location, the top and bottom configuration is the same and the nozzle types alternate along the header. As shown, the top nozzle represents an open one-inch elbow while the lower' configuration shows the 1HH12 90 G-1
NEDO-24224-1 nozzle used on these spargers. The 105* header pipe is supported at three locations and the 75' neacer pipe is supported at two locations. Figure G.1-5 shcws the support arrangement at locations other i ( than at T-box locations. The brackets are -inch thick and are welded to the shroud. The pipe-to-bracket mating surfaces are not welded to allow circumferential relative motion between l the header pipe and the shroud during a core spray injection of cold water into a system at reactor operating temperature. The header pipe is 3 " Schedule 40 Type 304 austenitic stainless steel. The street elbows, 90* elbows, half-couplings and the l close nipples (used to connect the elbows and orifice the elbows) are all Type 304 stainless steel. i G.2 Fabrication Secuence 1 Fabrication records show that the Pilgrim spargers were fabri-cated as follows:
- 1. The pipe was bent using a four roll bending process as '
shown in Figure G.2-1. The rollers have 2" radius grooves and rollers 3 and 4 are adjustable to accommodate the pipe size and to bend the pipe to the required radius. In this case the design radius is R = 94.25 inches. The maximum strain in the pipe is calculated to be 2.1%. L
- 2. After the pipe is bent to the proper radius, it is
! placed in the shroud to verify that the pipe fits th! j shroud as-built conditions. During this fit-up proce s, the T-box five-inch pipe is marked for drilling the header pipe holes. l l l l G-2 l l i i
NEDO-24224-1
- 3. After removing the pipe from the shroud, the headers are welded to the T-box.
- 4. The holes for each nozzle are drilled in the header pipes.
- 5. Stainless steel half-couplings are fillet welded at each nozzle opening. -
- 6. The elbows are screwed into the assembly and roughly aimed.
G.3 Installation Secuence The sparger is then installed in the shroud. This includes:
- 1. Welding the brackets to the shroud thereby positioning and holding the spargers. It also includes attaching the T-box to the shroud by welding the seal ring to the T-box and the shroud. It is assumed that because of interference between sparger ends, one'or more of the spargers would be cold sprung during installation. This operation was not addressed in the fabrication records. ,
- 2. The next operation was to aim the nozzles as required by the sparger drawing.
- 3. The elbows were then tack welded to assure that the threaded connections remain intact.
G-3
. - . . .. _ . . - - .= .. .-. .. . . -.-. - .
i ,
, NEDO-24224-1 G.4 Performance History -
t Pilgrim. Station first went critical in June 1972. There have been no inadvertent core spray injections. Pilgrim Station does flush the core spray spargers curing each refueling outage. Water is pumped from Condensate Storage at a temperature of approximately 70*F. The maximum AT that has occured is i 130*F. This aT is sufficiently low that fatigue is not a
,- concern.
i b i I I I L r { r G-4 1 i
,._ , - . . , - . . . , , . ': ' ~: ~ . _
~
NE00-24224-1 G.5 Potential Sources of Stress The potential sources of stress in the core spray sparger which could result from fabrication, installation, normal plant operation, and operation of the core soray system during postulated loss-of-coolant accidents are presented in this section. G.5.1 Fabrication Stresses Residual stresses are developed when an initially straight pipe is subjected to a moment sufficient to cause yielding and later unloaded as would occur during the fabrir.cion of the core spray spargers. The fabrication operation is idealized in Figure G.S.1-1. The steps involved in the calculation of the residual stresses are:
- 1. Determine the moment-curvature curve for the pipe assuming simple beam theory.
- 2. Calculate the applied moment, M , corresponding to the t
final unloaded radius of curvature. Determine the stress distribution associated with this moment.
- 3. To describe the unloading, calculate the elastic stress
~
distribution corresponding to the applied moment (-M )* t
- 4. The residual stress in the pipe is the algebraic sum of the elastic plastic stresses due to M and the elastic t
stresses due to (-Mt )' G-5
i
' ~
NECO-24224-1 r In calculating the moment-curvature curve for the pipe, thin shell theory was assumed and a representative bilinear stress-strain curve (Figure G.5.1-2) was used. As shown in Figure 3.5.1-3, the strain varies linearly through the section while the' stress follows the bilinear curve for angles greater than 9. The applied moment (M t
) is given by M
t
=2 fe (Ec o Sin $)(aSin$)(2adot)
#o
+2 n/2 {(cg Sin $yc ) Et + Ec y }(aSin$)(2ad$t) t e Eqn. (1) where ,
c,=f=outsidestrain a = radius of pipe R = radius of curvature cy, fy = yield strain and stress l E, Et = elastic and plastic modulus The first term in Eqn. (1) is the contribution from the elastic part of the stress distribution, and the second term corresponds to the plastic portion of the stress distribution. After integration and rearrangement, Eqn. (1) becomes , M =Mo (1-E/E)[20-Sin t Sine 20+ 4 Cos 0 - + - t l E Sine n I. s and Sin 0 = yc /c, = cy a/R Eqn. (2) ' G-6
NEDO-24224-1 where Mg = Moment corresponding to the first onset of yielding on the outside surface = fyn a2g Clearly .for fully elastic behavior, 0 = ',} and Mg = M, Figure G.5.1-4 shows the variation of the applied moment with the outside fiber strain and also the bend radius R. As shown in the figure, in order to get a final radius of 94.5 in., the outer fiber strain during bending is 2.33%. The corresponding moment is 1.43 fy a a2 t. The residual stress distribution can now be determined by combining the elastic stress corresponding to (-M ) and the ' t elastic plastic stress during bending. Figure G.5.2-5 she.a the resulting stress distribution. A correction for the thin shell theory assumption is included in the results. Figure G.5.1-5 shows that the pipe is subjected to high residual stresses (approaching 'the yield stress), and that the stress distribution varies around the circumference of the pipe. In particular, it shows tensile stresses on the surface facing the center line of the vessel. It should be noted that the actual stresses could be higher due to local yielding at locations where Hert::ian contact stresses (between the roller and the pipe) occur during bending. Since this would be most likely to occur on the surface of the sparger facing the center of curvature, higher stresses could be expected at this location. The residual stresses shown here were calculated for room temperature conditions. However, for reactor operating temperatures 2550*F, the residual stresses are expected to P G-7
NEDO-24224-1 relax to the yield value at that temperature (18.8 KSI at 550*F). Knowing the applied stress, one can calculate the minimum crack size that could prcpagate under sustained load stress corrpsion cracking. Using the following worst case assumotions: I 1. KISCC=6KSIfin
- 2. A long continuous crack
- 3. Sustained stress up to yield = 18.8 KSI the minimum crack depth for crack growth is given by a
KISCC = 1.12 [n min ! or
" min
- 18 8 x 1.12 30.025 in This shows that under worst case conditions, a 25 mil crack could propagate due to stress corrosion cracking.
The conclusions from the evaluation of fabrication stresses presented in this section are summarized below: l ( 1. Stresses due to fabrication could be significant and l l would exist throughout plant operation. . l
- 2. A possible synergistic combination of adverse metallurgical condition (e.g. sensitization, cold work) and high residual stresses may explain the observed cracking.
G-8
NED0-24224-1
- 3. Since the stresses change sign (beccme comprehensive) around the circumference, a crack that initiates in the tensile region can be expected to arrest in the compressive regions.
e e G-9
NEDO-24224-1
- G. 5. 2 Installation Stresses Stresses sufficient and necessary to cause initiation and propogation of cracks my intergranular stress corrosion crack-ing can be identified by postulating certain installation variables. Figure G.S.2-1 shows two cases which might be postulated.
In Case 1 it is postulated that differential weld shrinkage occurred during welding of the header pipes to the T-box. The outer bracket would provide a force to cause the header to contact the shroud wall. For simplicity, the arm is assumed to have an arc length of 90*. A 1/8-inch differential weld shrinkage is assumed. The deflection resulting at the header end would be approximately: l' 1/8 = a ; a = 2.95 inch 4 94.25 Then, from Reference G-3, Table 13.4 Case 1 a = WR 3 (2$ - sin 24) where Q = 90* IET ' Solving, W = 602 lb, assuming: R = 94.25 ! E = 28 x 10 6 I = 4.79 Since M = WR i o = WRC = 602 x 94.25 x 2 l I 4.79 o = 23679 PSI m 24000 PSI (elastic) l l G-10
NE00-24224-1 For Case 2 it is assumed that R1 is incorrectly facricated to a radius of 93.25 inches. It is further assumed that the vessel support brackets cause a uniform moment on the pipe increasing the radius to 94.25 inches. The initial inner length is n/2 x 91.25 = 143.34. After forming, the inner length is n/2 x 92.25 = 144.91. Strain = c = alinner = 144.91 - 143.34 = 0.011 2 inner 143.34
= 1.1%
Using a stress strain curve for Type 304 stainless steel, the
- resulting secondary stress is found to be 38,300 psi for 1.t%
strain. For the postulated conditions, these two examples show that high deflection limited tensile stresses can occur during installation. These stresses have not been confirmed. Installation stresses considered in conjunction with the material considerations discussed later (in Section G-6) may explain the cracks that have been observed. It should be emphasized that the installation stresses postulated above are all deflection limited secondary stresses that will relax to the elevated temperature yield strength of the material during normal plant operation. G.S.3 Stresses During Normal Ooeration All identified stresses during normal operation were found to be negligible. Loadings that were considered include impingement loads, (i.e., flow past the spargers), seismic loadings, pressure, thermal mismatch, stagnant line top-to-bottom temperature gradients, stagnant line through-wall temperature G-11
NE00-24224-1 gradients and weignt. The calculated stresses are shown in Section G.10. It should be noted that during normal plant operation there is no core spray flow. The sparger AP = 0 and AT = 0. Impingement loads are only 0.63 lbf/ inch of header arm, resulting in negligible stresses. Weight of the spargers and water is only 1.12 lbf/ inch, again resulting in negligible stresses. Stagnant line temper-ature gradient calculations are not provided since the maximum aT for top-to-bottom gradients and for through-wall gradients were found to be less than 8*F, which would result in insignif-icant stresses. It should be noted, however, that the AT for core spray injection is aJdressed in Section G.5.4. It is concluded that the normal operating loadings do not result in stresses that will explain the cracks observed in the Pilgrim Core Spray Spargers. I t G-12
a NEDO-24224-1 G. 5. 4 Stresses Durino Core Soray Infection Stresses during core spray injection are the design stresses for the spargers. Design loadings include all those discussed in Section G.5.3 plus those that occur because the system is no longer a passive system. The pressure differential in the sparger at rated flow is approximately 15 psid. The hoop stress in the pipe is about 160 psi. Impingement load stresses during spray injection are less than during normai operation. Thermal stresses due to the through-wall temperature gradient are high and are known to be
= Ea AT 2(1 p) .
These stresses are not a concern for one or a few cycles. The radius of the sparger shrinks when the sparger is cooled, resulting in secondary bending stresses of approximately 4500 psi. The axial stress in the pipe due to AP and bracket friction is low, only 300 psi. Flow through the nozzles results in a torsional stress which is low, less than 200 psi. Weight stresses are negligible. Water hammer is not expected because the pipe is essentially an open pipe, and the nozzle opening areas are approximately equal to the pipe internal area, even for the short leg. However, water hammer is addressed in the following section. G.5.4.1 Water Hammer Loads Water hammer loads as discussed herein are those loads associated with injection of core spray water into a core spray system where the system piping downstream of the check valve in primary containment is assumed empty (or filled with steam) G-13
s NEDO-24224-1 because of the draining of water from the spargers and/or the flashing of water to steam curing depressurization prior to core spray injection. For the purpose of maximizing injection loads, primarily on the core spray spargers, it is assumed that reactor pressure is essentially atmospheric (as for a large LOCA) enabling
. system flow to increase to runout controlled only by the injection valve opening characteristic. Upon valve opening, the head (H) is available to accelerate the flow, but as the velocity increases, the acceleration head is reduced by friction and local losses. If L, is the equivalent length of the pipe system, the final velocity V isf given by application of the energy equation.
l e Vf 2 H=f 5 2g The maximum velocity attainable is limited to that at system runout flow (4500 gpm) which produces a velocity of 40 ft/sec. in the sparger (at the entrance to the long sparger arm to be more concise; the velocity at the ends is zero). Actually, the velocity of the water first entering the sparger will be less than runout velocity because of the relatively slow opening characteristics of the injection valve. The injected water fills the pipe line between the injection valve and the sparger at a time prior to full valve opening and, therefore, before the final runout velocity is attained. Assuming the maximum velocity attainable, the resulting momentum load in the spargers is: 4
^
G-14
NE00-24224-1 2 P* = V = (40)2 = 21.6 psi 144gv 144(32.174)(0.0160) or F,= P,Ap = 21.6(9.89) = 213 lb. Where: P , = momentum pressure, (psi); Fm = momentum load, (lb); V = velocity, (ft.sec); g = c gravitational acceleration, (32.174 ft/sec2); v = specific volume, 0.0160 ft3/lb (s90 F water); Ap = pipe flow area, 9.89 in2 (3 " sch 40 pipe) . If the end plates at the ends of the spargers were removed, it is obvious there would be no impact load. Now cap the ends and also plug the sparger nozzles and elbows. Again, there would be no water impact load, because the trapped gas in the line acts as a surge tank. The actual end condition of the spargers is somewhere in between these two extremes. It is much closer to the open end condition, except that there are several " ends" instead of one end, and they are located along the length of the sparger arms. The open flow area of the sparger nozzles and elbows is computed as follows: Area Total Area 2 Number (in ) (in 2) Open Elbow 28 0.617 17.28 HH12 Nozzle 28 0.173 4.84 , 2 TOTAL Open Flow Area Per Sparger = 22.12 in The open flow area of the nozzles and elbows is actually 12% greater than the flow area of the two sparger arms (2 x 9.89 = 2 19.781n ). G-15
NEDO-24224-1 nn estimate of pressures induced in the sparger at the end of the filling time of the spargers and piping can be made by considering a sparger with only one open elbow located at the end of each arm. Steam would be pushed ahead of the oncoming front of water, exiting the sparger through the assumed single nozzle. The developed differential pressure to expel the steam would be approximately 2.0 PSID. Adding all sparger elbows and nozzles to this logic clearly demonstrates that the sparger indeed behaves like an open ended pipe, and con-ventional water hammer loads of any significant magnitude would not be present. Injection conditions at higher reactor pressure would clearly be bounded by the runout case presented here. G.6 MATERIALS ASPECTS OF CRACKING The potential causes of Pilgrim core spray sparger cracking are discussed in this section. A general discussion of the effects of cold work on the IGSCC susceptibility of Type 304 stainless steel is also presented.
, G.6.1 Potential Causes of Cracking The potential causes of core spray sparger cracking which are currently considered to be most probable are indicated in Table G.6.1-1. The evidence supporting each possible cause is also indicated.
Near the T-box, four possible causes of sparger cracking have been identified. First, sensitization by welding the sparger arms to the tee is supported by the patterns of cracking near the heat-affected zone of this weld. Intergranular Stress Corrosion Cracking (IGSCC) may result if stresses are suffi-ciently high in this area. Such crackin3 has been observed in l piping inciden.ts in the past. G-16
NED0-24224-1 Second, IGSCC resulting from cold work inherent in ' arm forming followed by weld sensitization may oe the cause of the sparger indications. As discussed below, sufficient cold work is present for enhancement of cracking tendencies. Third, fatigue induced by thermal variations in the environ-ment are being evaluated. However, the variations in tempera-ture during operation of the reactor (8 F, see Section G.S.3) are expected to be small. No evidence of a driving force for ' thermal fatigue has been identified. j Finally, fatigue resulting from flow-induced vibrations could be hypothesized. However, the natural frequencies of the - sparger are anticipated to be high relative to any flow-induced excitation sources, and the sparger brackets restrain the amplitudes of any vibrations. In the arms remote from the T-box by distances greater than 2 inches, welding cannot be considered a major influence on cracking. Sensitization may still be present if the original solution heat treatment were inadequate, either in temperature or quench rates. No direct evidence exists of this condition. Secondly, if cold work from arm bending were followed by local heating, a sensitized condition would more readily result. Again, no direct evidence exists. Thirdly, surface cold work
. resulting from arm bending or straightening could initiate cracks and subsequent growth could occur from residual or installation stresses. No documentation exists to support this possible cause. Finally, fatigue by either of the sources cited above for the T-box area could induce cracking, although there is no confirmed source of fatigue loading.
The most probable cause of cracking adjacent to the T-box area is currently considered to be cold work followed by weld sensitization leading to IGSCC. Approximately 5% cold work ,_
. could result from sparger arm fabrication and installation.
G-17
NE00-24224-1 4 Stresses in excess of the yield stress may be present, and weld sensitization could occur during arm to T-box joining. Sufficient conditions for cracking may therefore be present. In the arms away from the weld, determination of the cause of cracking awaits further metallurgical confirmation. However, the most likely causes hypothesized include sensitization from inadequate heat treatment, combined with residual or installa-tion stresses to provide IGSCC. Alternatively, local cold work from pipe bending or straightening and the above mentioned stresses could initiate cracking by transgranular means followed by intergranular crack propagation beneath the hardened outer layer of sparger material. G.6.2 Effects of Cold Work on IGSCC of Stainless Steel The mechanisms of cold work enhanced cracking are complex but can be visualized through the illustration in Figure G.6.2-1. In this illustration, factors influencing susceptibility to cracking are shown as increasing or decreasing susceptibility by lying to the left or right of the diagram, respectively. Cold work serves to increase the material yield strength. This enhances susceptibility if stresses in the material result from imposed strains; since the resulting stress state of the material would also be higher, consistent with the increased yield stress. If stresses are fixed as the result of imposed loads, susceptibility may decrease; since the stress state of the hardened material is a lower fraction of
.the yield stress.
l Cold work serves to promote chromium activity in the material matrix which reduces susceptibility through the more rapid , recovery of chromium-depleted regions. However, sufficier.t time at higher temperatures (>500*F) is necessary for the l recovery phenomenon, and such thermal treatment was not I practical for the spargers, nor deemed necessary. G-18
NEDO-24224-1 The most significant influence of cold work is in the trans-formation of austenite to martensita phases through defor-mation. Martensite, if present in sufficient cuantity, can assist in recrystallization of the material upon subsequent thermal treatment. The strain energy induced in the lattice promotes recrystallization. The result of recrystallization is migration of grain boundaries away from chromium depleted regions, with attendant benefits in reducing sensitization. However, the presence of martensite increases the tendency for carbide precipitation and local chromium depletion during subsequent weld sensitization. A wider heat affected zone can result from welding stainless steel with prior cold work-induced marter.<.i te. If sufficient cold work is present, transgranular cracking can occur in oxygenated water environ-ments with or without subsequent sensitization. Environmental tests conducted on tensile, bent beam and pressurized tube specimens are illustrated in Figures G.6.2-2 through G.6.2-5 (which are based on information from References G-1 and G-2). In Figure G.6.2-2 it can be seen that the time to failure in 0.2 ppm20 water for sensitized and cold worked and sensitized material varies with stress. Specimens tested at cold worked plus sensitized conditions (at higher stres::es) produced failure times (by (IGSCC) comparable to samples which con-tained no cold work prior to sensitization. Cold worked samples without subsequent sensitization, tested at compa'rable stresses did not fail. IGSCC failures could be induced at very high stresses in cold work /no sensitization samples, as illustrated in Figure G.6.2-3. ' If the data from Figures G.6.2-2 and 3 are plotted on a basis - normalized by the material yield strength, a more clear picture is formed of the results of deflection-induced stresses in stainless steel (Figure G.6.2-4). Material cold G-19 i
NE00-24224-1
- worked to various levels and subsequently sensitized can undergo stress corrosion at substantially lower percentages of the material yield strength, with cracking as low as 80% c ys in quarter-hard stainless steel furnace sensitized).
An equivalence must be established between plastic strain during sparger arm forming and the cold work condition of the test specimens. The yield strengths of specimens receiving 5, 8, and 15% cold work are illustrated in Figure G.6.2-5. The stress-strain curves for the same heat of material without prior cold work indicate the amount of plastic strain necessary to create a comparable yield stress to the uniformly cold worked specimens. Thus 2.1% plastic strain calculated for arm bending corresponds to approximately 1% cold work and stresses near yield may or may not result in cracking (data are insufficiently clear). The strain con-centration from localized bending, if a factor of 4 is considered, would be comparable to 5% cold work. A reduction of the cracking threshold to 0.80 0and cracking under residual and installation stresses could occur. G.6.3 Conclusions on Soarger Cracking Core spray sparger cracking at the Pilgrim Station can be hypothesized by the influence of weld sensitization or prior sensitization of the arm material and subsequent cold work of the arms during forming and installation. Sources of stress for IGSCC are dependent on residual stresses from arm bending and deflection during installation. The principal factors suspected of causing cracking are i considered to be highly variable from one plant to another. The absence of one or several key factors may explain the lack ! of reported indications in 19 of the 21 domestic BWR operating , plants inspected to date (March 1980). G-20 l l
)
__ ____O
NE00-24224-1 G.7 Crack Arrest Assessment In assessing the possibility of crack arrest, the following sources of stress are considered:
- 1. Stress due to pressure, mechanical loads and thermal gradients. These stresses have been shown to be negligible and are not considered in the crack growth assessment.
- 2. Stresses due to bracket restraint: these are displacement controlled (secondary) stresses and would be expected to relax as the crack propagates.
- 3. Residual stress due to fabrication: as the crack pro-pagates into a region of compression, the stress intensity factor can be expected to decrease, thereby resulting in arrest. -
- 4. Weld residual stress: weld residual stresses at the T-box - sparger welds would influence crack propagation.
These stresses are likely to vary circumferentially and also relax as the cracks become larger.
- 5. Stresses due to vibration are assumed to be negligible.
In considering crack arrest, the stresses due to bracket restraint and the fabrication residual stress are significant and are evaluated in detail. G.7.1 Stresses Oue to Bracket Restrain't Stresses due to bracket restraint are governed by the applied displacement and the compliance of the' pipe. Since the displacement is fixed, the compliance change with crack growth could lead to crack arrest. This is G-21
NEDO-24224-1 ccmparable to crack arrest in a bolt loaded 'a0L specimen in stress corrosion tests. Figure G.7.1-1 shows the variation of compliance with crack length for a pipe subjected to bending. The compliance was determined using the relationship between the strain energy release rate G and the compliance change per unit area of crack extensionf{(ReferenceG-4).For the cracks in the spargerL/disexpectedtobeintherange0<h<40. Figure G.7.1-1 shows that when more than 30% of the pipe is cracked the compliance of the pipe increases by a factor of ten. Therefore, for the given initial dis-placement, the stress in the sparger and the applied stress intensity factor would decrease by a factor of ten when more than 30% of the pipe circumference is cracked. Clearly, if the crack length exceeds this value, the restraint stresses become negligible and crack arrest is expected. G.7.2 Fabrication Residual Stress The residual stresses due to fabrication vary around the circumference and a precise calculation of the stress intensity is not possible. Nevertheless, a conservative representation of the stress is used to calculate the stress intensity factor. The assumptions made are as i follows:
- 1. The crack in the sparger is modeled as a thru crack
! in an infinite plate. l i l l 2. It is assumed that the tensile stress (a) is uniform and is applied on the crack face over a length (2b). (Later this will be conservatively taken as 25% of l the circumference). i l I G-22
NE00-24224-1
- 3. The remaining portion of the crack is assumed to be subjected to a compressive stress which is half the tensile stress. (See Figure G.7.1-2)
- 4. The crack length (?a) for which the combined stress intensity factor reduces to zero is calculated.
The stress intensity factor due to the tensile stress can i be shown to be
~
Ktension = 2ca sin W a)
. The stress intensity factor due to the compressive stress
{isgivenby
-1 '
Kcompression ,'-2 (c/2) a I yna' n 2 - sin (a @J Setting Kfens +Kcomp = 0, we get sin ~I j = sin-1
=f
~I or, sin
*{
or, b = 0.5a If we assume, b = 25% of the circumference is under tension oy and the rema'ining portion of the crack is under compressive stress (equal to half the tensile stress), the applied stress intensity factor becomes zero when the crack length is equal to 50% of the circum-ference. Thus, even under extremely conservative assumptions, crack arrest is expected. G-23
NE00-24224-1 G.7.3 Conclusions on Crack Arrest Based on the above material, the following conclusions may be drawn:
- 1. Since the applied loading is predcminantly dis-placement controlled, the stresses relax as the
- cracks grow. Crack arrest is therefore expected.
l
- 2. The residual stresses due to fabrication vary from tension to compression. As the cracks propagate into regions of compressive stress, the applied K value reduces to zero. Even for extremely conser-vative assumptions crack arrest can be shown for a 50% circumference crack.
l l 3. The above conclusions are valid as long as there is no stress cycling due to vibration (e.g. Flow Induced Vibration). G.8 Structural Integrity With Cracks From the discussions of the potential stresses in the core spray sparger, (Sections G.5 and G.6), it is concluded that only deflection limited secondary stresses approach 25 percent of the material yield strength (except for self equilibrating thermal stresses). If a 360* through wall crack is postulated at any location on any sparger arm, the remaining stresses will not produce a failure at any other location on the sparger. The AP stress and the stress resulting from an axial
- load in the pipe due to bracket friction are proportional to j the cross-sectional area of the pipe. The load from aP and friction was found to be <1000 lbs. Assuming a yield strength of 30,000 psi at core spray flow temperature, an area of less than 0.035 square inches is required to maintain continuity.
This area is much less than the original pipe metal area of G-24
NEDO-24224-1 2.68 square inches. The bending type stresses are all deflection limited secondary stresses. The discussion in Section G.7 shows that cracks are expected to arrest since the driving stress will be relieved. The bending loads may, however, in a worst case, cause an existing crack to open up by an additional .005 inch, assuming the existing crack has - progressed 360*. This is a geometry limited condition. It is concluded that no loadings have been identified which could result in stresses that would cause the spargers to break during normal plant operation, transients, or postula?.ed loss-of-coolant accidents. G.9 References G-1 A. E. Pickett and R. G. Sim, "The Effect of Stress and Cold Work on Intergranular Stess Corrosion," Materials Protection and Performance, Vol. 12, No. 6, June 1973 G-2 G. M. Gordon and R. E. Blood, " Reactor Structural Materials Environmental Exposure Program," Symcosium on Materials Performance i.n n Ooeratina Nuclear Systems, Ames Laboratory, Ames, Iowa, August 28-30, 1973. G-3 Hopkins, Design Analysis of Shafts and Beams, McGraw-Hill Book Company. G-4 E. Kiss, J.D. Heald, DA Hale, " Low Cycle Fatigue of Prototype Piping" GEAP-10135, January 1970. l G-25 l
. _ _ ~
NE00-24224-1 TABLE G.5.1-1 Possible Causes of Cracking Location Possible Cause Evidence
- 1. Sparger Arms Near T-Box Sensitization by Location of Cracks Welding Cold Work Followed by Estimated 5% Col /. Werk Weld Sensitization Near T-Box Fatigue (Thermally Under Evaluation Induced) (a T's Are Low)
Fatigue (Flow-Induced Under Evaluation Vibration) Amplitudes Are Limited ,
- 2. Sparger Arms Away From T-Box
] Sensitization From None* .l Fabrication Cold Work Followed Pipe Bend Forming *, By Sensitization No Evidence of Sensitization Local Heavy Cold None* Work 4 Fati gue Same As In 1. Above
- Sensitization and cold work state of spargers not yet known.
G-26
" - ~ ' - ~~
j
0 NE00-24224-1 f M
.b n
b L.LJ 3
-J
- N >
- N c3 c z .2 s
% b d
/ / > k
~ -
m W
/ '
/ u db .
m t.A J.ewd - f{ - x S . Fe-s~'
.C 2 s
~
gg/ et I a Z L-LLJ O t.3 V C:::
<C CL. -
M l ~ T
~
9
~__ c 4 e
\ 5 N o
% 'c gl0 b$
_u m-y G-27
~
NECO-24224-1 0* 00 a , 7
\ </
-l N. , , f g /
I
< / /
~ ,f' /
CORE SPRAY SPARGER E-US _J. m, _ k 1\\ )j% E
\\
e ll r( s s s
^~
,.cd)
/ / %\
l Y 00 Y. 180* Figure G.1-2. Core Spray Sparger - Plan View G-28
j!EAO-24224-1 SEROUD _* "' i,"#S__. i i 4 I l 1, l g p 1/32 in. G AP TYPtCAL Y l 4 i O LOWER SPARGER T l l TOP Gul0E
/
4 l Figure G.1-3._ ;Sparger to Shroud Attachment Method G-29 *
- m. -, .
NEDO-24224-1 sH,dQQQ l l l l Figure G.1-4. Sparger Nozzles
!G-30
'NE00-24226-1 f
/""- ;
ELDED TO SHROUD su [ ONE E ) vu LOWER BRACKET Figure G.1-5. Sparger Support Method
~.
i G-31
.,. . - . -- I~ _
jEDO-24je.1 1
^
j i
. _ _R = 94 25.~
~
+.- 3' b
L
.T.
L 1 4
! Figure G.2-1. Pipe Bending Method J 8 t
e i G-32 l
NEDO-24224_1_ I t INITIAL CONFIGUR ATION l M *M J . __t, LOAD APPLICATION OURING FABRICATION
/
\
/
/ /N s /
y s N N \
/ N ' N N ,
s N f N N s*%'- - _ _ _ . _ _ _
' ' ,/
,_,t,, *N / (-M i
/
's '
/
's
%*%~~ ~ __ .,_ __ .wa # s #
. FINAL DEFO_RMED SH APE AFTER LOAD REMOVAL Figure G.5.1-1. Sequence of Events Leading to the Residual
__ . Stress Distribution ,, _ __ _ . _ _ _ _ _ . . G- 33 9
t i NEDO-24224-1 RCCM TEVP (74'F)__
'E
__7_
'O__
1 I l a E_. 0 " 3#
- e0 = 0.0012 E = 28.3 x 103 ksi 3
{ET = 0.26 x 10 g,3 l
'#0:
.e
' (7 550*F l
l
.1 ' --'O -
r l joo = 18.0 kai e 0 = 0.0007 E = 25.8 x 103 kai 3
,E7 = 0.26 x 10 kai l
l'o i f_ Figure G.5.1-2. Bilinear Stress-Strain Curves for Type-304 l __ _ _ Stainless Steel G-34 l
^
~'
'NEC0-2422C1' i
e fC > j l eO" R {a=a y + E t I'0- 'y ] A --J l , ELASTIC-PLASTIC
.. ,_ # N - - ---____
4 p _ Jk 4 ', .
# EOSTid ' i+ 'y .
' I A <
1
- # STRE'SS
_---- 'i DISTRIBUTION
'o/E
/
STRAIN
! DISTRIBUTION ,
,l Figure G.5.1-3. ' Stress and Strain Distribution in the Pipe i Under Applied Moment 6
e h G-35 l l . ,_._. . . _ _ . ..
a i FINAL HADIUS (m ll l200i 1100 :94 5 M ti! I I I I I 6 CURVATURE K ! [000bt } 0014 ;0 0lb I I A i I I 11.6 - l __ l l 1.43 M i > jij _ e l 2.33% g l 3 ,. l 1 n l12 -
,; 12 iy s m ls e a i o O j 1.0 - I g e, , b m
l I i 8 I ro. g i to l 'd'.8 - I e l i I I e i I tiis - ! I I I I I _ _ . . I I I
'0.4 -
t I I i I ji.i - 8 i a 1 i 1
- l l i
!'Y l l l
- 0. ~ ~
[10; j 2.0 e l30j j OUTEf t FIBEll STHAIN (%)l
~-
I --. - _ . . . .
; Figure C.S.1-4. Momenit Versus Outer Fiber Strain i i
m n- + _- NE00-24224-1 y
.E 29
_ 3 u :s -
.s-
- ER .
w$ 'Y S edluu'
~
iG " l$C
,a 3 -
I3- . . :5.a
!:g
.w- 53 u
' La
-- a
*5 J3-
~~
! ,A
.u
\ '
U u C
\
o w O i
,.- 1
,w
. d 1-
,g_ 'Me
~ o m
M
.8- s .W g
,-- u
. r,n
.V. .
I w im iO
' f4 ha iC f0
' O!
'u-
.e=4 i]w 3 *h3 gg- im x @
l O' M
, m 4
lC t = 15 E ' t o. o i T lo Z l- E: lws
'55'
{*d
'" i i*
o
!M -
$7 l73' ' b
}T.$ - -
y I9a. S 5
'G-37
_ , ,,j_ me *=W* * ' ' " ' '
I _NE00 2 24224'-Y INSTALLATICN A ADIAL MISMATCH OlF F ERENTiAEWEL3 SHRINKAGE AT T 80x
. - . - . 'R1-
-- -R 1 = R 2, CASE 1 IN
'EQRCE FOR SHRINK = 1/8 in.
- o AT T 80x a 24 ksi(ELASTICJ 1
ASSUME R1 = 93.25
. R2 = 94.25
'g UNIFORM FORMING g e a 1.1 %
i a m '8.3 ksi (FROM o - e CURVE) i cISET } Cat < R2_ N MI Figure G.5.2-1. Postulated Installation Stresses - G-38
-c-.
_ - ..NECO-24224-1~ - - INCREASES CECREASES SUSCEPTIBILITY SUSCEPTIBILITY _. TO CR ACKING ' TO CR ACx!NG 1 1 I ^ I { - IMPOSEO- - - -STRAIN
' IN,CR E ASES
.t VIELD _
IMPOSE _D LOAD 5 ] (STR ENGT,H ,
.~A 4 PRCMOTES
_ j CHROMlUM 8 RECOVERY _ _ _ . _ _ _ y JNCREASES
- ' COLD WORK I a CH,ROMlUM - - - - +
ACTIVITY V j PRODUCES J1ARTENSITE ~ ~ "" - ~ ~ ] I 1 I IGSQQ IF_ SENSITIZED a y T AIDS RECRYSTALLIZATION IF
, _ AFTER MODER ATE C01,0,_ lCW > 15% AND HEAT TREATED WOR K (STR ESS REQUIREDF j AT HIGH TEMPERATURE ~~
~~---~~-~~
INCREASES CARBIDE - PRECtPtTATION: p - _. __ _ __ _ l l I t . __. TGSCC AFTER HE AVY,,, _ _
. COLD WORK ISTRESS REQUIRED)
Figure G.6.2-1. Effects of Cold Work en IGSCC of Type-304 Stainless Steel
. G-39
~. e,,-
w+ -an-.-
*10?
YlELD,1/4 HAftD qg g
" J T YPE-304 ST AINt LSS SIC ELF j 0 2 ppm 0 t
$ tJO COLD WuH K ' l 00 -
l 1/4 HAND, r M b% cot D WoHK dittNACL AND elijftADIATED l 18% COL D WOHg Stiga l l2L D i
$ 1/4 itAHD l C O 5% COtO Woe <K t
d 8% COLD WOHK . SE N' SIT IZ A IlOrd h 1/4 ilAllD , M kYlELD.Il% COI.D WOftK j
=
_I to . . _ _ _ .. ._ _.
$ t YlELD,5% COLO WOR K l ',F 8% COLO WOflK AND [
UltNACE SENSITIZED i y so 2 m O O Q aA5 e l m - 5% COL D , N { d . 2 m i WORK AND FtJHNACE i o $ . SENSITIZED j N 30 - F UHN ACE SENSlTIZED) G> g _;YlE_tt5Fi1HtJACE_ _ _ _'l2EdlSENSIT APPitOXIMATELY SAME TIME TO FAILUHE IN SENSITIZED, ' VEHSUS COLD WOHKED AND SENSITIZED MATEHIAL l' Y E AH ' l10 YE AHS l JL JL 1 I i
- 10. l 1,0ik),um) 1001 10001 10,000 l 100,000 i
' TIME TO F AILt1HE (Sol'
~ + ' ~ * * - . . . . . . - - - , _ _ _ _ .
l Figure G.6.2-2. Effects of Cold Work on ICSCC i a
> 1401 h -h 170 ksi ,
TYPE do4 STAINL ESS SIEEL : o2$4=n O 550'F (E LASTIC), l ; h 1[4 IIAHD UNyLNStil2LD e
\
\
100 -
\
k l COLD WsORK AND'l NO SENSITIZATION l
~ ~
hi YlELD 1/4 HAHDL h
=+
,8a
\g l ej -
E w K o o
+
f g ,60 - h> ?> d 4^ e 4 y w ,8% COL D WOH K t
+ IE'L'D'85 COLD'50f1K l e
~
+YlEl'D 5% CO'L'D WORK. '5% COLD WOftK - _ _
40 - t, t FUHNACE SENSITf2ED YlELD[FUitNACE SENSITIZED:
~~
,20
~
CHACKING IN COLD-WOHKED MATERIACON- 'AT VEHY HIGH STHESS WHEN NOT SENSITIZED !
,~o l l l 100; j 1.000 l . 10,000 1100,000, l3 to)3xx) ?
TIME TO F AILUHE ths). lFigureG.6.2-3. Effect.sofColdWorkonICSCCl
s 2.4 l
~ I -- FURNACEi TYPE-304 STAltdLESS 0 2 mo. 0,,
tcu* F
- STE ELl
^22! -
l SENSITIZED 1 '
' $ t40 COLD WORK
%> COI D WOHK l AfdD NO SENSITJZATION E ts COLD WOH A i PL US i 25; -
tss COLD WOH A AfdD '.FUHNACL l MMTWd 1/4 HARD g h I/4 HAHD ; ia - OssCOtDWoHA, g so, h Il4 88AHO SE rdsl TIZE D '
\
.1.6 -
.____.. _ ___.___ . ____ \
- o COLD WOHK AND
'$ FUHfdACE SEf4SITIZATIOtd;l
- E _
$ l 1.4 , -
A \ O s O h
'O 12 -
h \
\ ~
S m .r ,a e i er O o m g :o it.5, , a M i4 ie 08 - # V 05 - 0.'4 -
~
. O 2' - D E LUdEI E'L'D CH Addt bG'lh'Cb t DMIO R K E D~ ND SEF4 SIT 12ED COrdDa TBON ON LY
, I I i 100{ !1,000 1 :10.000 j1,000.000
- l100.000
, TIME TO F AILUHEiiEl'
~
~
, Figure G.6.2-4. Effects of Colt! Work on'IGSCC!
e
. - . . . ~ _
NEDO-24224-1
.i
=
- $ _ m 1 -
5 4
;: a a w 9 a-_
E. ! ' d *S w w w=o-E z
=.!
34 s
- w .a.32a.w -
CC m, m i
'uw-[ :,
D' $ ** P- -
=
p t o
- O
- g. m-w E_. : '$
\u z- e ,; v G.
w-E e
.!i t0 0-0 z
8:s !3 m w :- m
,g A
\
= w
'O U n >
\ 'c 8
a m
\ za U
t m w
\ -
- o ,
3 , t,.
..n
\ +1 CN
\ ,o.
o
!v s :b
.a a)\ --- - - -
', ~, el*a'~;; b ,xvW
~
L c. lh>
.g c-c I
.w C i
- s. w 1 t,mOv.-
rs W
\l
- __o
\
\ <sWuvi
\ 4
- - - NIY. _H.1S__G_ _N_.3 9. W_n_Wl.X_VW - 4 WN
.. . I. I I I I I .I- % 1 ,o
'S S.
S R S S_ . 'S- -8_ -P_._ .2 - U_81) S.S3 W1S G 43 r i l
IlECO-24224-1 iE000 - 1000- --. . P
~
wg
.o M
2 C/C 0
=1+8 (1 - n2 f O)dJ [
a n' L m
* ~
{1 - y ) Kf ~ 5 _ 01 l ~0d E
;' : x=
u Q
~ <,.,(wa3),,u, t .
di - { i [ t W?% 10, -_________ _
- l I
1
- i l
\ i Ud a 5' 25 l100
-- -~1000
~
l I I
~
3 -
'l ! ! ! ! l !I l l l l l!If
~ 0.01. 0.1 :10 g106 FRACTION OF CRACKED CIRCUMFERENCE U/d l
Figure G.7.1-1. Co=pliance Change, Cracked Pipe _ __ l G-44 l I 1 l
fiE00_-24224-1
= M- =
TENSILE iSTRESS n A A A a A A A A A A I I I I I I I I l l l l l 1 YYYYYYY Y Y YYYYY AAAAAAA A Ai1AAA I I I t i I I I I I I i i i VY Y Y YY Y Y Y YY . COMPR ESSIV E
;STR ESS 4
- 5 *
; Figure G.7.1-2. Assumed Stress Distribution on the Crack Face f ~
i G-45 1 I
^^: .-- ~"
NEDO-24224-1 > G.10 STRUCTURAL ANALYSIS OF THE PILGRIM STATION CORE SPRAY SPARGER STRUCTURAL ANALYSIS
. OF THE PILGRIM STATION CORE SPRAY SPARGER t
Prepared by: 123,/ A. N.M urrell I Verified by: ,d D. F. Holland Approved by: - < R. E. Legata() i ' 1 l r-46 e
. - - . -- . . - . - . . .- .~ . - . . -.
NE00-24224-1 INDEX TO SECTION G.10 Pace Introduction G-48 Sumary of Stresses G-49
- 1. Design Loads G-50 1.1 Impingement G-50 1.2 Differential Pressure G ,il 1.3 Torsion Due to Nozzle Thrust G-53 1.4 Weight G-54 1.5 Mismatch Due to Themal Expansion G-55 1.6 Natural Frequency - FIV G-59 '
- 2. Stresses During Core Spray Injection G-61 2.1 Sparger Pipe G-61 2.1.1 Impingement Load and Seismic G-61 2.1.2 Differential Pressure G-65 2.1.3 Mismatch Due to Themal AR G-65 2.2 Nozzles G-66 2.2.1 Torsion G-66 2.2.2 Differential Pressure G-67 2.3 Bracket Stresses G-68 2.3.1 Seismic and Impingement G-68 2.3.2 Thermal Mismatch G-69 2.4 Middle Bracket G-70 2.4.1 Differential Pressure G-70 2.4.2 Themal Mismatch G-72 G-47
Oi NEDO-24224- 1 INTRODUCTICN This structural analysis contains e loads and stress analysis of the pilgrim Station Core Spray Sparger and attat.5 ment brackets. The report has been prepared to show that during normal plant operation and for the core spray injection event, stresses are well below the material allowable stresses. Further, certain failure modes have been postulated to , demonstrate that even with a 3600 through wal1 break in a header arm (worst break location) the sparger segments and brackets will not be overstressed, t I k P b L I > ! G-48 se.- . y ,. , m - - , -- -
_NED0-24224-1 ENGINEERING CALCULATidN SHEET NUMBER b #S O ~ ~
# ' C ~i'd CATE SUBJECT ' 'O E 'M Ch BY U SHEET Eb CF Somets bre.GM , Ps's
). S C A R /:: G f2. 9i oG B EMo suc. - S e s m i c. .
2.7 3 N. tite *o
.! rAGt wac, E,MO.# -ICG
'E E uoiur., -
Sesmsc i?62.
** %pru Gsm Ew -
TCR Gsuo NG - m ertwas W3eurco 44 ti 2 Uch te Ct.d.c c s S ewDi wG 2.571 t
$ 8 CAn 3 53
- 3. t e_AC.tc.ET (i ct0Er?}
43 E W oiUG 4 7 'c 3 S *cAra 2%
- 4. 'Cet. Acx.ET (ms oot El G EM DiMG 43CO 5 As.n, e tei 66NOsuG (w sto') 3:IR SAEAR (td EL.0) If3 G-49 FCAM NO. NED 87 (4 75)
~ NE00-24224-1 ENGINEERING CALCULATION SHEET 0Y NUMBER b 'l O " 27 OATE " ' O ~ 90 SUBJECT
- fM ' **
- bO#A 96V b d 5"f 4 R ay _+*zA[ SHEET
' _CF
). TG5 t Gu L.o A C S e
1.1. noie<smeur loa.o Cro pecueer rec a n o-( = ,0 v'A
. # tr LOL.
gc. gc 7 a 5- s -v2 o gu f T~ P
= .p = 45.2,ib-/gs @ sco'r
, Cr o= 4. o > ~ (sh.s~so)
-x-
- :. 4.o G/3, E.-
(4<.77)( a.o f( %) n 2. f- : z, w M /g y : ,c3 % 4 A c.-rom u u nuo c = 1,7 ~1 Frf sec., l l l G-50 _ y _ .. y. . .. ... .-
NED0'24224-1 ENGINEERING CALCULATION SHEET NUMBER b '3 0 - I DATE " - ' '> D SU NECT ' ' G
- T ' Al bb SY -
SHEET L CF
-: e
).2 M CC6A.TMTt 6 kES$tJT?_4 fAAr. 9 W 4600 Lapw (, f&w60 Cud
- % C C af m )
Q - q fo O f# /\ mig x \w. y @ ucsce '7.c%ga\
= 10 0 0 /gu b uJ ArcEA 6(, u e E E- 4 5 etw 8 k3 *[DiA. N J 5.
An = (SQ($(p[
= 8. m ,e iOsJees c o.G orecica + 4 6 cuscas 'sc o.t7C oeicces Ac = io (d(oast' + 9t,(-Q(o.eif e 3 4. C'l i#
Q i
- Och
= ,A A. + 2G A,
= To (Am + Ae) 1; c C~ _. t o. o 3 ArvAe (n.uGo u.s,%s
- 3 2. Gf [Yjee :
l G-51 i i i FORM NO. Nfb47 f t.75)
'NE00-2422i-1 ENGINEERING CALCULATION SHEET NUMBER S O O 9 7 'I DATE #0~#0 sua; sci @ 'A c ' A C5 sy 4! sseer 3 CF Gs= v,4s = ($2..c f3(6. u Au d
. 2.is fe b fe< a u n les
(% =. L :-. m A%e.T( 90 s ec/mM( 7.4e $NStB 56 u n tes I 7. ff 6 P ^/nc ple-h3cetr 6.oe vs AP n.4 ys a 2o gi ff) , qo(\18')ALsc \.4 ] bP = t v. 3 7 ys\ F: APA G = a.3Q(s.rae7 9 -- Lo \ . 75 Ihe l l i l G-52 POmM No. NEO.47 (0.F5)
NE00-24224-1 ENGINEERING CALCULATION SHEET NUMBER WC BOO ~T21 DATE I~'O' O SUEUECT 4 4'*^ bb SY SHEET CF i .3. Te r<_s to y mer n ue w_s TAzusT d I 'n
@ 3.7s Et \ s C
g/ ' @ Ac7 = o 23- p, A , = (-m ) (- f e e..) = f e % 23 = p dA, + R A i 8 I fy = 0 R3 " 8. A t = x( p-A) R3= p.4 -A 2 t + fl. A t G-53 l .an.= n..w u r, in _. __ - . . _ .
.... .. .. . . - ._ JEC0_-24224-1 ENGINEERING CALCULATION SHEET NUMBER O O"Y1 ' cATE *-'o-93 SUBJECT ST'M
- CS gy M SHEET I CF f= 62. 4 lb/g s
~ ~
g- 32. G C h/se= Yg * (l .C YR ~ b . T& $$ IA E
\b.33 96 (79 - (Gz.d( 3z. csp (.fG43Ao) 6c.37)(e. gen) 32.2.
Q5 =. 10.G lkl Ars= $ 2.o f Wse c. A ,_ : o.t4 % i a Ps..o (GI.0(32.osD( ,ws A q)
%: 32.1 f2p :- 11.4 o i b.C l
l 4 l, 1. 4 LQ Eld Hf 1 Sh. ' S c.L_ 4 0 Rec
% g pe =- 9. \\ \V4-
.i W ms = 9. t? Ibd 'l w) = 17.39 lb/A o< 3. t i g Ibp;
' G-54
_ .n . . n n ,, ,.. . ._ . . _ _ . . _ . - . - -
~ ' ~ ' ' ~
..--..L.
-- ~
fiECO- 24224-l' ENGINEERING CALCULATION SHEET NUMBER b OO ~ i27 CATE 3*'O~TO SUBJECT ' ' /> v i M.
. R$ BY
/J.,2 SHEET f-CF r
)* 3 - 15 st AT C 4 DOE e Me.m A, t E')c c A u s te c
\.._ 30' /
g .- , p..
/ ..,
i
'A e, es .c o
'W g,s .
/ g --
s, . u, -, . 8'f.7 - , . 34 fo
\ .,\ '\
- l. , t ~.
i k )
\ l F
U - N . ,
- q. i.c.i 2s= 1. 1 L S +.
i f z-
- 9 7 ,m 1
kr., I 41 I .- 1. o : 'l 4.1$ 2. SAecoo = 550*F CS Ac)G -
'7 0
- f Ar = 'i;'c - F i
\
G-55 1 8'OftM NO. NED.87 88 793 . . _ _ _ _ .
NEDO-24224-1 ENGINEERING CALCULATION SHEET Nuueca 9Y1C- A- C O '< 'O DATE I ~ ' O ~ E0 SUBJECT 'm" ' ' ' ' C ' * \
- Cb sy M*f SHEET 7 CF Ge_ ss.
AR : m R 6T o< 6.h 8 s o ' ~ /i / f c m.sco*=
~
At ::- (614 810 9(97)( '4?o) sa = c. u s u . w ca\ v ud Po rt. 56LM.ENT ASS UM E AR : dtmq(\- u h
/ .
e l \ 5 m bk %y l l l l i , G-56 l PCAM NO. PeEO 47 (4 75)
NEC0-24224-1 ENGINEERING CALCULATION SHEET NUMBER - OO~IV CATE E ~ ' d "EO SUBJECT ' ' 0 2 ' 'A I5
- sy M SHEET 7 CF
$ cs M.em oc coe_F spra.AY Pi ps s W ese A W AC.NMGMTS l
\
\
i 4 i 4
\
5
\
\
Rc; . h 7 R a a:.7 cea.- s ,u p e <t.T u c.uc s . AR w s e.4.e_r cT5 cee_ 5stewew rs A4 = A%,( 1 - ces4) = . 4 N ( t - e e s +) s.,* Lo' ,if" A= E.f"- is.2.c e s= 31?? n 4-
- 2. - t AR, 0. o i ff im ARt = c.o2.f4 is -
a% = o. c 2.oT is G-57 ro== ~o. w n.ar m . _ _ _ . _m . . . _ _ _ _ _ .
..NE00324224-1 ENGINEERING CALCULATION SHEET NUMBER NC O O "I DATE 3 " # *
- I'3 SUBJECT '*O4'#^ bb BY Nd SHEET i CF
- .F c.
h P P
> J d l
Coe ss eerAm. . 6. - N ( 291J ' M Ef = 9t EI 9: 4TGTL , 14 E T s
- 2. 2 ' y E3 27 8to' ps; I = -[ (4 '- 3.9 tb : 4?H J P = (32.s,C y son $
f= 2(44.25)sim e 4,= s- 4, =. in. zr + = n.q-A,=- 47.7'1 or L= b 2. s C > 1. Ag- S0.37 m 8*6 t oc1L[ 8 : (? 2..n f y to*) _( .%wy , fl=(32.i7Gio)pz.iG P-(n.noio()(oo(n 3% 3 9, - 42.cl ib6 Pu= 3 2 lbf P3 =374 \bp G-58 I
m a m.
,,O W "
ENGINEERING CALCULATION SHEET NUMBER C~ '- CATE #0' SUBJECT
- Ob SY /I SHEET 3 CF,
- 1. 6 M Af 0 R A L. bG O EutT
&= "~ , .
GT 2w yju 6 :. W % $ 0 ' p$ T . 4 . 7 ? T( ie d y= 3. G 1 s o *3
/
/
bA$C 1 d 4 :. l $. d O i G13 I % , a "4" f 2 $ / I C
- C S Y g ) (g ,
- 1 ~~ 'I ( 3. s g ,C- ? ':( 17: '
5ld 4 s% ^x c~+' tisc 1 / u - 9. e, 3 :: c 2. i s . N
\t 1
- " S 67 fi t. 7 3. ( '.
. 1<7't?/ o .ION ,. s o ('l7??\ x a ., # I e os.Ju a, - 9.37 _f : n f. 5 2. 4 CASE 3 i h4 oc '2@" (2.Tvio@(c.11'@ g 21,(o N $ (it.yis9(1)%.ftf G-59 FCftM NC. NECH7 (4 75)
_ NE00-24224-1 ENGINEERING CALCULATION SHEET m A OO" TU CATE
#-'O~EC NUMBER SUBJECT
- 's4'^^ O BY SHEET 'l CF VCCTG% Pt6DDM6 hf CG GE M C.Y .
hV s *22'l V5 4 Tec : 4 ? L N'g g 3 o =. u
/ O *1 NI : 1. (p N q -
G-60 t FO AM NO. NEO 47 (8-75) ~ . - . . . . .
NECO-24226-1 ENGINEERING CALCULATION SHEET NUMBER C OO~$71 CATE I~ # # 40 SUS.lECT 'A 'E ' % 04 SY b SHEET IL OF 2 hmS5f5 su bMG ER - l Cov2.iuc C o rt C SFe AY iV4GencbN
- .i SoAesc<t. b c
%. i. i a m.9s u GE. fAEMT LcAo is. $ctSus c 4
- $ 615wd c., C ucY Ge.oevo A c.cscert Ancy (p 9 corti % S cT5mit . CSE DBE Non.i ?ouTAu c.c1 f c.if
\/E.rtTtc AL (F 3 doeit) C.cfC o.io u usE Vcst S t* h G' 4 E tt.
- A ssu m G $ g c,e. A e.s.z u : $[,i o) s .f g scis m c -
e s = az.ra e= i.oc'K (scei.Q
. L5 6 - uh = td +. C O : i . fd 4 i,0 .fw s . (td w.)3 1. C (i. i i G) = 1, v74 'N4 G-61 l
l l i
- roans No. sto ar ts 7s>
~
NEDO-24224-1 ENGINEERING CALCULATION SHEET NUMBER OO~$1 DATE E"#d*YC SUBJECT G fA O5 Sy M SHEET 'S CF k Stur4 C Me f,Rir.\ <. cownuvous t5cm w = 1.'A 4 ' L
! ! & ! l ! I L 4 LL k I k L $ 1 ! I t1
% gA _ ' ' <_a
_ qm s '
' ' 's d i' ~ _ _ _ _ , . _ _ . ' ' A l
.0, act. t in . h : c t.t. a j
- r. 11
'Tm(R.EM oC 'TWI.E7 Mo vnE L% '.
uJla. NMi+.2.%($,[s L Is
.; M s Os = #
I. Is 9 I. s Ys 0$b$ === 0 T : I I i o a 2.s ( Ag-) o = g(_o,b.07) g l 3r_( T R4.5 i + A. 4 ') I 3
.e- (.2 ~2.' -3' M 3 ~1~7
~ 1. tA 4 [ W.7 -
T \ 4t. t v G2.1 VTT f.h : Is~l 2. i v - v 6 l l ONib. ~~
~ ~ *= '
Ih $ . [ 4. y bmceost: q: 2.t 3 2 t c c, pd
'G-62 rom No. smer min
NE00-24224 1 ENGINEERING CALCULA f1ON SHEET NUMBER ,. O O - D7 CATE I~ # " "IO SUBJECT ' ' C' G ' W Ob sy M SHEET '* CF Assu rw G %c" SaCAs A.f T' sa,1.r.1 (sea.woest cA sc g /Awnto) d : 1. 0-4. .W
'I I. l '
, [ } .
l . [ [ . . . l A A % R, WiLn 1 q - et, L m
. I R. , = -2S S \hS e
- 4. di t. 3 lbf
\ %. \ '
0 y G 1
-2M ,
,y = o% 4t.t V - 2 5. f - 1, vNg
_ g. 2. M : -2f.C y - Ig(2-I u - 4 9.7 to ni. V: - 25. f - t .51 - 2. t l . 3
.. m . -1(I1 - i4218% :.11.3 (y-4al s
l 4 , i f
- 32.3% we M ar 4: 4 t. T
, i ~ . n, i
G-63
. . . ~ . ,......
~
NEDO-2d224-1 ENGINEERING CALCULATION SHEET NUMBER A OC ~ k71 DATE I~'O~EO SUILJECT @' M 5't' M O< SY I SHEET 'f CF s M c. (32at)(-J e = 2. o . e ?- rM37 y Sct t.f
- 4*79 Is 4 ."M , # ) Scr 7 : 13 f 2- pso due. % d e ss.w . t.
Lm ot o C C MEET ~
- O s (,o b bEC s 1e\
n r= = id9 PN 2:.k.3 1.G74 ( # 3 ct} i 1 rdz. - Q t m oicof - I
' b3 - l'i f I b i
~
Rt
- (-2113Y u.m i
%I L$e6 # 1 iMm eq P
Cd3 - 2 t i d - 76. 5 ; Qq: 1St.S (tA sco io c. A u e.v v Ari o a oc . c u ecsT srrtm) , i G-64 s l'ORM NO. PoEE>47 (0-79) .
NECO-24224-1 ENGINEERING CALCULATION SHEET NUMBER b 0 0 - I7- 7 CATE I~IU~O^ j SUBJECT ' L 6 4 i eA C$ SY g SPEET 'I' CF 1.1.1 @ ? 9 MEN 7i A v T.S t) 2.0 t W g . .
~
ef 543 2. ss o . mt
=M79u $ !
%,s=0.ATis .
rt.- ss ,~ A>i usm as ~- Pe 2- , Y~ Q'{~ (ivaC(t. m Y-(2.cr- O.,, cT-
- - (oO. 4 p:
2 .1. 3 W 5 f4WCN 'pV C 7p ro 5ea3a u R 4 . a n < P P (scc S e c. Ld l 5scumr pf: l NM M- h ' S C' (st4(st.% r :.oq 3 :
% > -. Str (3ebt.+ i-i ai u.- (mxn.30 2icu. t q'= M c-I h2.1 :
C. : 2.C ivt. j fS 9.*M i M !
- ? ! t 3 1 ( 2.o_) l
- 2. e.,m :
2 49 Il 9 51 l l 1 G-65 8'ORM NO. NED 47 (8 79)
o . _ .'lE00-26224-1 ENGINEERING CALCULATION SHEET NUMBER kC C 0 ~ k 2- CATE #~#0"IU suoEcr O' t 6 d W^ CS av M SHEET CF
- 2. 2 . do M:t C 1,2,t -To ras, s o u l
L i
' L
$ o' or I sca 4 0 pipe p, 1.3i<
/ .-. N. 3.75 Si s 1 0981
; s
~ i.i I - TA(i.3id-i.oW) = 0.cf 7,d b[ % -/-
E% - :- I:. . C.C77 -
,13 2 3 A 3
/
13i r/t
/
A2 y(i.sich-i.c#)=.eiy fr a.o ist (s ec i.e m=4 - (n.0( s.v $ = 22.s a-thc e = z E4 .1 2a = is v s;
.t323 Y: d'N s 2{ p5[
A: .ey l G-66 comu uo. mme , ni
e - aw - NEDO.-24224-1 ENGINEERING CALCULATION SHEET NuueER 4- 0 0 - S' M DATE 3~'O~ O SUELJECT 'b"# Ob BY SHEET 'T OF
- 2. 2.2 T i p f s e s u n A u 9 ca. Essure _ r d55c<v E woc
- G RE'M Mo t e t.4. 3
'choED E'( Ge.Ac.v ET
[ F
\
~
l u , e 1.1 J I
-/
J n. F: 161.76 ILG (scc iQ .
'C ::.A b- 16 \ .?(
o.4%
= 32X p 13cuo sec, q k* b ., (Ibl ib().9)
.ss z3
= 2 2. 0 2 p s t G-67 roa u mo. neo ores.rsi . _ _ . . . _ . . _ _ _ _ _ . .
NEDO-26224-1 ENGINEERING CALCULATION SHEET NUMBER I A00~@l7 OATE 3-#0*iG SUBJECT t L/ W '*A- CS gy Md SHEET i4 CF
%3 R A CC ET M E55ES TV e t u <. c., $ Ctow.
- 2. 3.l " b E \ $ M I C. 3 ~ Ymf5 0GE ME'MT SacA c-q: 2" A ag,o f o g,{ !
kq: \31.3 lh[ ($ t t.11. b s koeto *
- 2. (0.26)(2):.7C7*E e
/ b
^
I
~ 1. (, i u q =. .i s t .7
. ~7o7
- 1. 0 c ~
4 R: -- 19'(o.4 9v
- u. + - /
\
.tr y smess wy w wcomo emccr G 0 %% kg : ( l 31.N )( 3.5) : 9 (f l.3 s A lh a- bf- t(4ho.2mf o.nc sa L
K: N '1- - 1955'
~
psl
.uc G-68 l
con, m - em
~
_ ..NE00-24224-1 ENGINEERING CALCULATION SHEET NUMBER kOC- T DATE 3~d~IO SUBJECT ' 4 O d 5 ** CS BY SHEET M CF 2.1, 2 T+Ettrw --e Mi s m. A re.M S a css vos Tc sescius f4 : Ry d :- 4 2.9 (i.t,) : gig 4 i v -ib rr - M - c.9 6. 4 2 .2.14 TT,, 2.9 O V CcA6tuED 70 199 p>l er : GT+ %.= 4Tb3 p' r= 4 ?to 3 -/ (4 Tr t. 3) * + 4 ( 2.a,M2-2 F :. 4iTl(psb , G-69 vonmuo.n werte.nn .. ... .-. .-. .. .-
, .l. -
. NED0-24224-1 ~
ENGINEERING CALCULATION SHEET NUMBER OC~ $ OATE 5 ~ ' O' 'Y O SUEJECT i L /, r2.s % C$ SY k SHEET I' CF 2.4 . YV{coc6 Ort.A cK ET 1.4.i Ti sferager AL NC550RC l ~ ( O t l A kW m i
\ . _
)
j i
)
Js- . m
\
1 l 10.0 in g > ~ i
/
N .z r i l u-as.- -- l l l l ! Y Y.. I G-70 e=w no. =u>.7 es.ru
' ~
' NEDO-2452Aii-1 ENGINEERING CALCULATION SHEET NUMBER 9CC l- C C - T7 "7 CATE *~ ' O ~I O SUBJECT *.O'*
Cb BY k SHEET 1^-- CF Re= 16i.75 tb? (su.i.:.3 3xxe = 2* A vJeuo: Gn A cxcr
\ to \ .Z( _
y' \lo \ 5S 9' 1(q)(.:<Kf) (sb.C
\
= 153 pi = 10 7 psi (coederm3 iee\A \mR}
STE6% 906 To 15 m o C G Mc.m G ui 9teess = $ m= Re.cl =. (1st.td(2.J) = 3237 m-ib We,. o : e ne.cer
,- j I
- g- (3K.Sh _ , g gg I b L _ . .i kl
.'7 Q=1_ 3 .I8 ~~.SO 1 . , 9g y
- c. . .n.
@~ $ 13.3 {
- h 13.1 s LW s\ W
= il 64 es'- = 2590 ed G-71 i
i e FORM NO. NEl>47 (8-75)
NED0-24224-1 ENGINEERING CALCULATION SHEET NUMBER - CO~kM DATE I"" ' '3 ~ IU SUBJECT O "'VA #5 SY Nd SHEET U OF 7..k 1 4C12.W1 A L. k SM A(T~ $L N f4.6% cos o .%.8.i A u tcAn streggs =. D % = 4 2.#l lbG ($ e c. i.d A osuo: Gre A cx_sr . p 42.9 4 19 7' (HT(.ul(4 (3x.s)
~
l svests rue ro smom momac
$f4.E %
+ M: 2 d 7 : (4M)( 2.b e i O'l 1.5 O.Y i
WC L O .' 6EACEET ! g,b _ (2 7 yd(.26(3p g.. g, (,gy 3)2-u u u
.O,53 in 5 _ 0.2 f iA' i
l y . b . 1021.$ C =- 'O ' E .g3 .7 f e- = loli ps's 77 ) 4 3 o tra t G-72 r:au wo. uu>er se.7n
.NE00-24224 1 ENGINEERING CALCULATION SHEET NuusEs 9C L c '5 - ? '2. CATE I- W ~ ?'d SUBJECT ' t / d ' At
- Ob gy M SEET 'd CF Cousiweo Srtess es GLP 1 Cort A c.cEf TC 163 95L Cs 1oT pp i f
F : iI (o 4 + 9 0 9 4 2O 2.4 g : ._(qc+ ;% + iq3a
- 369'2- ps. : 43 c e, p 5[
! r- : 3 69 2_ - M (3 59 2.') 2 + 4(ist')'- 2._ t P- 3591 f5t ,1 i
- G-73 t
I F0mu No.Naber es.7si _
NEDO-24224-1 - APPENDIX H LOOSE PIECE ANALYSIS H.1 Introduction As stated in Appendix G, is is expected that the Pilgrim core spray spargers will not break, resulting in loose pieces in the reactor. However, an evaluation of the possible consequences of a potential loose piece is presented in this Appendix. H.2 Loose Piece Descriotion Since a piece has not been lost, it cannot be uniquely described. Three different types of loose prices are postulated in Section H.4.2. They are; a section of sparger pipe, an outlet nozzle and a small piece cf the sparger. The entire sparger is fabricated of Type 304 stainless steel. H. ' SAFETY CONCERN The following safety concerns are addressed in this safety analysis:
- 1. Potential for corrosion or other chemical action to reactor materials.
- 2. Potential for fuel bundle flow blockage and subsequent fuel damage.
- 3. Potential for interference with control' rod operation.
H.4 SAFETY EVALUATION The potential for postulated loose pieces to cause the safety concerns identified in Section H.3 is addressed in this section. H-1
NEDO-24224-1 The effect of these concerns on safe reactor operation is also addressed. H.4.1 General Descriotion The core spray spargers are attached to the inside of the core shroud (see figure H-1) between the top guide and the shroud head and steam separator assembly. For a piece of the sparger to reach the fuel assembly inlet orifices, it would have to be carried by the fluid flow up through the steam separators then outward to the downcomer annulus then through the jet pump nozzle into the lower plenum then make a 180*F turn and be carried upward to the fuel assembly inlet orifices. A part of the sparger cannot reach the fuel assembly inlet orifices by falling down inside the core shroud as the core plate and control rod guide tubes will prevent it. For a part of the core spray sparger to reach a control rod, it must migrate across the flow velocity gradient and then fall through the restrictive passage between two fuel channels. Since all parts of the core spray sparger are designed for in reactor service, there is no possibility that any loose part will cause any corrosion or other chemical action to any reactor material. H.4.2 POSTULATED LOOSE PIECES H.4.2.1 SPARGER PIPE - The sparger pipe 13 3-1/2 inch schedule 40 pipe and is attached to the core shroud at six location (tee box plus 5 brackets). The maximum span between supports is 38-1/2* which corresponds to approximately 62 inches. In order to generate a loose piece of pipe, 2 through wall cracks would have to propagate 360 around the sparger. The weight of the largest pipe segment would be approximately 65 pounds. H-2
NEDO-24224-1 A pipe segment could come to rest in any of three locations; the top surface of the top guide outboard of the fuel assemblies, the top surface of the feel assemoly handles, or in a very unlikely event the top surface of the core plate. In all three of these locations the flow velocity is low. The velocity at the core plate is less than 1 ft/sec. At the top guide elevation outboard of the fuel assemblies the velocity is 2.9 ft/sec. At the fuel assembly handles the average single phase velocity is 2.7 ft/sec. The two phase multiplier in the region above the fuel assemblies is 3.0. The maximum lifting force due to these velocities is 6.0 lb/ft which is significantly less than the weight of the sparger in
. water, which is 12 lb/ft. The supporting flow velocity calculations are presented in Section H.7. Thus a piece of pipe cannot be lifted by the flow, and therefore it will remain at one of the above mentioned locations.
A 65 lb piece of pipe which falls from the core spray sparger will not harm the core plate, top guide or fuel assembly handles as these components are designed for much larger loads. Since the pipe cannot be lifted by the flow and since the pipe cannot fit through either the steam separator or the jet pump, it will not cause any flow blockage at the fuel inlet orifice. Since the pipe is too large to fit between fuel channels, it will not cause any interference with control rod operation. H.4.2.2 SPRAY N0ZZLE Each spray nozzle consists of two 1" - 150 pound elbows fabricated of 304 stainless steel, which are welded to l the sparger. In order to generate a loose nozzle a l H-3
NED0-24224-1 through wall crack would have to propagate 360 around the nozzle. The weight of each nozzle assembly would be approximately 1-3/4 pounds. A loose nozzle would most likely come to rest on the top surface of the core plate or on the top surface of the top guide. The flow velocities in these regions are insufficient to lift tne nozzle thus it will remain at one of the above mentioned locations. Since the nozzle cannot be lifted by the flow and since the nozzle cannot fit through the steam separator, it will not cause any flow blockage at the fuel assembly inlet orifices. The nozzle is too large to fit between two fuel channels, thus it cannot cause any control rod interference. H.4.2.3 SMALL PIECES r A small piece of the sparger could become loose if both longitudinal and circumferential through wall cracking occurred. A small piece could be lifted by the flow if it maintained an orientation with its ma).imum projected area perpendicular to the flow. Due to flow turbulence and nonsysmetry of the loose part, the aart will tend to rotate so that the minimum projected area will be perpen-dicular to the flow. With this orientar.icn all parts with a length of greater than approxim'ately 0.7 inches will sink (Figure 3-7 of Reference H-1). Thus, most piecer, will not be carried by the flow toward the steam separetor. However, in the unlikely event that a piece reachea the steam separator, it would have to pass through the steam separator turning vane (Figure H-2). The turning vane has eight curved vanes. The outlet of each vane overlaps the inlet of the adjacent vane. The longest straight piece that can fit through the turning vane is approximately 6" long and it must be oriented with the long dimension in the vertical direction. The H-4 I
,w. .m -
.m++ -+ emn= ~ - *"
NEDO-24224-1 largest piece that can fit through the turning vane with its long dimension in a horizontal plane is shown in Figure H-3. It is very unlikely that the flow velocities would carry. either of these maximum sized pieces through the turning vane. After passing through the turning vane, the fluid momentum is reduced as the water is removed. At the separator exit the fluid is almost entirely steam. A typical water content is I weight percent. Thus it is very unlikely that any piece could be carried out of the separator by the steam. If any piece were carried through the separator by the steam, then it could be carried into the downcomer annulus, through the jet pump and enter the lower plenum. A piece that entered the lower plenum would most likely be thrown to the bottom of the reactor pressure vessel where it would most likely remain. However, per reference 1, a small piece could be carried by the flow up to the fuel inlet orifices. The orifice sizes are 1.388 and 2.211. It is extremely unlikely for a piece larger than the 1.388" orifice and essentially impossible for a piece larger than the 2.211" orifice to be carried through the steam separator. The outside diameter of the sparger is 4" while the fuel inlet orifices are located on the surface of the control rod guide tubes (Figure H-4) which
. have an outside diameter of 10-7/8". Due to the different radii of curvature, flow would be able to enter the fuel assemblies. Thus unacceptable flow blockage as defined by Reference H-1 would require that more than one loose piece be carried to the same inlet orifice. The probability of unacceptable flow blockage of any fuel orifice is judged to be insignificant.
H-5 1 i
NE00-24224-1 The flow velocities near the sparger are lower than those above the fuel assemblies. Thus it is unlikely that a small piece would be carried over the fuel assemolies. If the piece were carried over the fuel assemblies and then rotated so that the flow could no longer carry it, the piece could fall on top of a fuel assembly or between , fuel assemblies. Figure H-5 shows a typical unit cell of four fuel assemblies and one control rod. The control rod moves in the gap between the fuel channels. The gap between fuel channels has a thickness of 0.75". The length of the gap between the channel spacer and the channel fastener is 2.3". Thus any piece larger than 2.3" by 0.75" cannot cause control rod interference. The control rod thickness is 0.312" and the diameter of the control rod rollers is 0.520". Thus pieces r,maller than 0.337" will fall past the control rod without causing any interference. A piece of precisely the right size could be in contact with the control rod and 1 or 2 fuel channels. Such a piece might be detected during the normal control rod exercising. The rods are inserted 1 notch and withdrawn 1 notch each day. It is also possible, though unlikely, that a piece might wedge between 2 fuel channels above the control red and thus not be detected by normal control rod operation. If the rod were to be inserted the control rod mechanism has enough force to lift 1 fuel assembly with the reactor at normal operating pressure. If the fuel assembly were lifted 1 or 2 inches it would be able to move horizontally at both the bottom and the top, thus most likely relieving any interference. The rod would then insert and the fuel assembly would fall back into place. Thus it is very unlikely that any control rod will fail to insert. One of the licensing bases of the reactor is that the highest worth control H-6
NE00-24224-1 rod can be fully -stuck out and the reactor can be safely shut cown. Thus unacceptacle control rod interference will require multiple precisely sized pieces. The proba-bility of this is judged to be insignificant. H.S Conclusion The probability for unaccpetable corrosion or other chemical action due to a loose piece is zero. The potential for unacceptable flow blockage of a fuel assembly is essentially zero. The potential for unacceptabla control rod interference is essentially zero. H.6 References H-1. NED0-10174, Rev. 1, " Consequences of a Postulated Flow Blockage Incident in a Boiling Water Reactor", October 1977. H-7
NEDO-24224-1 H. 7 Flow Velocity Calculations This section describes the calculations for the flow vel.ocities given in Section H.4.2.1. H.7.1. Flow Velocity in Byoass Region Assumptions:
- 1. The plant is operating at rated power (1198 MW )
t 6 and flow (69.0 x 10 1b/hr).
- 2. The flow in the bypass regions is homogeneous.
- 3. The bypass flow fraction is 12% (8.28 x 106 1b/hr).
- 4. The water in the bypass regions is saturated.
S. There is no down flow in the bypass region. This assumption is discussed later. , There are two parallel flow paths in the bypass region. One is between the fuel channels and the other is between the Core Shroud and the outermost fuel assemblies. The flos areas for these paths are shown schematically in Figure H-6. Path 1 is between the Core Shroud and the outermost fuel channels. 2 Aj = 1700. in A5 is the area immediately above the top guide. A5 = 1952 in2 Path 2 is between fuel channels. A2 is the area between channels. H-8 l
NEDO-24224-1 2 A2 = 3735. in A3 is the area at the top guide elevation. 2 A3 = 2482. in f .I' A4 is the area immediately above the fuel channels. A 4 = 20,880.0 in.2 The total pressure drop in both paths are equal. The loss coefficient in Path I was estimated as 0.1. The loss coefficient in Path 2 was estimated as 0.944 (K)W) )/A)2 ,(g W2 2 )/A2
/
(0.1)W)2 (1700)2 = (0.944)W 2 /(2482)2 0.475 W) = W2 6 , 1.475 W) = 8.28 X 10 6 W) = 5.61 X 10 1b/hr i The velocity in the bypass region between the core spray sparger and the fuel assemblies is then: 6 V = W)/pA = 5.61.X 10 /(3600 X 45.8 X (1700/144))
= 2.88 ft/sec The fluid in this region is water (no steam); thus there is no two phase-multiplier. .
H-9
, O NEDO-24224-1 The fluid velocity in the periphery of the core bypass region was conservatively estimated at 2.9 ft/sec. In actuality, there probably is downflow in this region. The total pressure drop across the top guide is predominately due to the elevation head. In some portions of the core bypass region, boiling occurs, reducing the elevation pressure drop. Because there are no heat sources in the non-fueled peripheral regions of the core bypass, boiling would not be expected in the vicinity of the shroud. Thus downflow in the perpherial regions would be anticipated to balance the density differences. H.7.2 Flow Velocity at Too Surface of Core Plate Y*bTotal Bypass)/ pA Since WTotal Bypass = 8.28 X 101b/hr 6 A=$(D 2 - 2 Nd ) 0 = inside diameter of shroud = 181 in. N = number of control rod guide tubes = 145 d = outside diameter of control rod guide tube = 10.875 in. p = density = 45.8 lb/ft H-10
NE00-24224-1 Then 6 2 V = (8.28 x 10 )/(3600 x 45.8 x - (181 - 145(10.875)2)j(4 x )44)) V = 0.59 ft/sec H.7.3. Flow Velocity at the Too of the Fuel Assembly Handles 6 6 WTotal = 69. x 10 - 5.6 x 10 = 63.4 x 106 lb/h'r 4 A = na n = number of fuel assemblies = 580 a = area associated with each fuel assembly = (6)2 = 36 in2 The equivalent single phase velocity is: V = (WTotal)/PA Then 6 V = (63.4 x 10 )/(3600 x 45.8 x (580 x 36/144))
= 2.65 ft/sec At this location the fluid is a mixture of steam and water. Therefore, to calculate the lifting force due to the mixture a two phase multiplier must be used.
M = 1 + q (p /P ~ I) L g q = quality = mass flow rate of steam total mass flow rate 6 6
= /7.6 x 10 )/(69 x 10 ) = 0.11 3
pg = 45.8 lb/ft 3 p = 2.35 lb/ft g H- 11
^ - - ~ ^^
NEDO-24224-1 t M = 1 + 0.11((45.8/2.35)-1) = 3.03 4 The total lifting force on a section of core spray pipe per unit length is a 2 l F=C 3Ap Mg V /(2g) Where: ; Cd = drag coefficient = 1.2 A = area = (4in. x 1(ft/ft.))/12 (in/ft.) = 0.33 ft 2 jfg Then: F = 1.2 x 0.33 x 45.8 x 3.03 x (2.65)2/(2 x 32.2) l
= 6.0 lb
, ft , 6 6 't H-12 4
.- _ e - .
,m- .w,n- , - , . - - - , - - . m.c w -
. NEDO-24224-1 O
sftaasOavga vsfinec M
/
,' 's $
s v t wt amo at 40 F*a v /
,s** '%
{ gN g i h\] ' sit Aaa onv e m N a 9' M j '. : 4;. ;; ./ - " * * ' ' sitaas ouf t s ? , i n t
, (
,D_
. ~
9/
=.t:;a. f-I "
i 11 [. CQat yeaf tenL87 '%
, l{ 8 8 EDW ATI A FA Act R li L-~ .f F LOW Pet 3lRJRt COOLA#ef 4* , k g
N iCTceslatif . ,,ii co ng y eay p ,,g
~
Coat spear paACin p KT W A85th'eLY I l ccat seemouc l l l*
- w g!! ,A PV6L AaBeas4 List ,
,. % g i ! CCAtP(Aft 4T auw,ascincut.Af te saf tn issL87 i
% l '
ig , , atCapCULAf tose ba Waftm W7Lif
,,i , ,
f [, i n.i.i ytmat SJPromt Salaf .
,t] Smit LO WALL 0., ) '. .i l
. . i '.
#' l,
- l. f 'j w ' .!
y[.@{ ;f dk, CDerf olCl aco canytg '
, N
( s\ x, , CCaef mCL ACO O Alvt mvC AAut4 Limis iso < cat PLux hooself04 .) Figure H-1. Reactor Vessel H-13 L ,1 i
NE00-24224-1
.,/,. *
' . . 7.
\U.
WET STEAM
' M A, " ETURNING W ATE 9
! h Q ,
I R {\
, u.
T [l( i 7 WATER LEvEt
\ /
',' STEAM WATER vtxTURE j
I 1 g" ' AETUANING WATER - i A
/
~-
y\xi
\
i i k V '
/
Y Y Tu AniNG v ANES ciutE7 No::1. E i h ; l' t a k I i r STAN CPIPE
# \
TO CCWN CCMER AAEA A AND AEC1 ACULATION I ANO JET PUMP $UCTICNS #g d 1
- l .t i
I F1gure H-2. Steam Separator H-14
' ~ - - ~ _ ._. _. __ , ,
NED0-24224-1 4 t@ t P i ? 4 1 i 1 s 9 2.23 in, d,78 in. I 188 in i p,, 4 i i ,} + 4 i i I 4 4 4 3 _ . _ - . , , _ - - _ _ Figure H-3. Largest Piece That Can Fit .Through the Turning Vane ifth' thi Ldng^ ' Dimension. ,in a Horizontal Plane H-15
~ '~
.- . , , . . - - - - - . . . ,. , . - ~ . - 'A - , 4
D NED0-24224-1 i(
\h v) ;
0 M( l
, g CPENING 8CR CONTROL 200 S'. ACE FLOW PATH Ak Ak IN CORE GutOE TueE p2 ~
a p r 1 i
/ ; \
; g: ,
y , ; _ , s cc.e ,mTE
)
i 2";"b.?
- c. ncE
~ ~
v FigureH-45 Orificed Fuel Support H-16 l
N*%.ea w-. .g-e h-*,.e,.:c h NED0-24224-1 p' c- -- - . . , e'.
. x, .,i . . .a" ,, ,.
+ . . . ' .*:,w* .
K.Yji .n n . .,[f,p.l_, ta.
, e, .
n d3 p$y%. v ;n. kY - i' } q} % 61 5. "n48sly \
,s u n s .e :. % 1
\ /.
n;m plg:#1&g.i 9 M { fff? ym4 y .WMi W . mtir; t i ggGhptd:d.R.s.way + MM . 2 L.r.u. et - x:'$ w9.c . . i. . ;w- ~=w 1,y ,, r+ w;p4
&y,d, . t i .,L:
g .g.1,+ .;y _ nS, a u ' nm .1 ; W:, w; FUEL ;i'(,'j-.
+ 4 Jim *O i
t .,, *
.c. 4
~
M*ln,
- 5 ASSEMBLIES 1U'~ -
. %.... . ?,,f[ , :.!Y;! - ~
4 4 > f t
& CONTROL -
i
. , ..,. k il
.t
; B
'M .
t.n[
, 3 ROD MODULE jf.M . ,.. A 1 L1 -
'!S #
J.$ Y' Q. . , .-0 *k
; = ,
~ ,a h}is G
tr ' p,
- i. T O ,
..E I 6 h i% p,c '
i i f k ..*?- :ss.s >-
- 1. TOP FUEL GUIDE <, ;,%,a. - -
-{p . $ -Pf lk , # 7 2.C H ANNEL (1;) b, i %gp'-g _
- . .c - m
. -;s- ~
. -;,. rrr.
FASTENER f 9-;@ ~ h ( , r~.h.3 } . ;r :
- 3. UPPER TIE p%. s,',k. .. .ii .k '4 .N<d .'.', ,'
'a i
j ,ij %. h "<? ..,.1:c' //d; 34'4./ / a
.gi- , '//wsg o PLATE '
3 , h ' c .#,. 'fjN,.h A;, ,
%]
4.EXP AN SIO N G y.,2' q' 7p i
' '3, ' ~ J '.$: /f.' W , ';; lMd i .
j SPRING . w y/ ") W .: . ' .%8 L'. 1,F'". ..~1 i
+ l
; 1 PJ L'
n
.'.? :t
" 'b* . "jY();kd,j
. w 16. 'i
- 5. LOCKING TAB i .NSy: .'z b. 'T a I .j i
'97h'7 'B'
[ WMY.5 { ; 'jpl
- i. , .
- 6. CHANNEL '
skS.U.S *J ' J '
.hgy).
3#
'M
- 1.E'j,'f.[[![G.
4 ,- '
- 7. CONTROL RCD } 4 j "!' 2 .-) iA j ,
J .:q .I E f.w/1;
- 8. FUEL ROD '
'@ A g i gD f G ) lI;lj 'I
$ i.S
~'I' L i
l l 4$.If o'.W'/4M;/; l
.s
.I,!' . J 9.S P AC ER !* i g i h.),\$ '~w q 'S',$'..g% ,,,: , ,7 y<
{- .-
," f3 .,-
- 10. CORE PLATE ASSEMBLY
' S' h'.' 4
,d 1,j f ';
Y1 1 Ji' O - 3-J. ,
'l j
OiSl'd.p)RN,IM'f*.'gJ /$j -- lh
- 11. LOWER d . (k 4 i "' [
TIE PLATE , Mh"j'd [[ ' h ;\ '( ,[, } -
..~y' Fi g-
'J
,N2 fjp;f,[j)I [f ~
f
,,f, :
- 12. FUEL SUPPORT
$ $ M . Y , . 'I'i - J. ri . ~.;
g li.OM;('bd. } 4.
.,i, PIECE .i *<'3
$ : /,' s
- 13. FUEL PELLETS y l, . .
E
- / 'l~( ).M .1f y; *E'l'< 'iq l! ,'
Nj 14.END PLUG 15.C HANNEL if; p 9}(p_ ' ,
','q,. _] p'*(d
'/jj{1 'l y$5 %] ? , ' ..
c .o.s 4 ,;![ , 2 . SPACER Q;g i,i-Q' - i.. ,h. 33 :'g;; 4 g! . %r 2 ' :lf g.' , 16.PLEN U M I ". .73 . "4- e1 41 I.
,N , D M ,/ /jq N}. kf) b SPRING l $. j "A. I "'
# fr Q Q$. .% %$ -
Q Y r GENER At h ELECTRIC t ;
\') l [?.l* [a/[l! .
k, -u s,m
, .t ,
-yM,1 r ?,o
-eir .:r. : .- ,
4 l s Ek, , es
/x ;*
1} pf f ' ' i gw:4 hpm,u?.R'dtQL,do. + . .!d / / T. ' ; i 2:n. ,a n ; .4 Figure H-5. Fuel Assemblies and Control Rod Vodule H-17
smM-s NE00!24224Tf i 1 A 5 % A > _g_ _4 i
,'^3 - m
{ PATH 1 [ PATH 2 I
+Aj+ _
q ,A 2 i Figure H-6. Flow Paths ~ a . i 1 u .i 9 J l H- 18m wa h
NE00-24224-1 APPENDIX I LOSS-OF-COOLANT ACCIDENT ANALYSIS WITH NO CORE SPRAY HEAT TRANSFER CREDIT I.1 Introduction This Appendix describes the methods by which conservative MAPLHGR multipliers were determined for application to the MAPLHGR values previously reported in the Pilgrim Reload 4 licensing submittal (Reference I-1), assuming that no credit is taken for core spray heat transfer. The input changes to the approved 10CFR50 Appendix K computer codes are described in Section I.2, the depressurization rate sensitivity is discussed in Section I.3, the results of the . analysis are given in Section I.4, and the conclusions are presented in Section I.S. I.2 Inout Changes to the LOCA Analysis The approved versions of the SAFE, REFLOOD, and CHASTE codes were applied to the Pilgrim Plant as described in Section 4 of Reference I-2, with the input changes described below. No changes were made to the approved computer codes. The postulated effect of cracks in the core spray spargers is to deprive the hot assembly of adequate spray flow during a LOCA. This effect is represented by setting the spray heat transfer coefficients in the CHASTE heatup code to zero, from their Appendix K values of 3.0, 3.5, and 1.5 GTU/(hr-ft2 - F) for the corner, other [ outside, and inside rods, respectively. l l In the standard Appendix K analysis, the non-zero spray heat transfer coefficients are applied from the time rated core spray flow is achieved until the time that the hot node in the hot
- assembly is reflooded. After the reflooding time, a heat transfer coefficient of 25 BTU /hr ft2 *F) is applied to all fuel rods, which is sufficient to cause the peak cladding temperature to decrease.
I-l
NE00-24224-1 In the present analysis, credit was taken for a heat transfer coefficient of 25 BTU /(hr F ft2) applied to the outside of the channel starting at the time when the water level in the bypass (space between the channels) fills to the elevation of the hot node. Normally this outside channel cooling credit is not taken because it is not needed when core spray cooling is cresent. In the present analysis, credit for outside channel cocling is appropriate because the cool channel will act as a sink for the heat radiating from the uncooled rods to the channel The two input changes described above were made to the CHASTE heatup code, and no changes were made to the SAFE blowdown code, or the REFLOOD refill code. I.3 Deoressurization Rate Sensitivity The sensitivity of the depressurization rate (which is calculated by the SAFE code) to the global core spray heat transfer coeffi-cient was investigatad. If it is postulated that, in the worst case, all assemblies, are deprived of spray flow, then it is appro-priate to set the SAFE spray heat transfer coefficient to zero. This was done in a comparison case for the Design Base Accident (DBA), which is later shown to still be the limiting break. The change in the calculated uncovery time of the hot node, and in-the curve of pre'ssure versus time, was found to be negligible. It was therefore concluded that no input changes in the SAFE code were appropriate for the present analysis. I.4 Analysis Results I.4.1 Large Break Analysis The identification of the limiting break for this analysis follows the approach in Reference I-2. Table I-l'shows the total uncovered time for several breaks, taken from computer runs used to generate Figure 6 of Reference I-2. These uncovered times are expected to be unchanged for I-2
NED0-24224-1 this analysis, since the effect on the depressurization rate was found to be negligible in the previous section. The r'esults of CHASTE calculations show that the DBA remains the limiting break. I.4.2 Small Break Analysis Break sizes smaller than 1.0 2ft are not limiting as shown in the comparison of the uncovered times in Table I-2. All breaks in Table I-2 have an uncovered time which.is equal to or less than the 154 second uncovered time for the DBA, except for the 0.100 ft2break. For all these breaks (except the 0.100 ft2 , which is described later) the Peak Cladding Temperature (PCT) will be less than the 2200 *F calculated for the DBA, because the time of hot node uncovery decreases with decreasing break size, and the decay heat decreases with increasing time. Thus the 0.900 ft2break will have a lower PCT than the DBA, even though both breaks have equal uncovered times, because the 0.900 ft2 break has an uncovery time of 71 seconds, which is much later than the DBA uncovery time of 20.6 seconds. The decay heat at 7,1 seconds is less than at 20.6 seconds, so the heatup in the uncovered period for the 0.900 ft2 break is less than the heatup for the DBA. 2 For the 0.100 ft break, which has an uncovered time only l 3 seconds more than the DBA, the PCT is calculated by the small break model to be less than 1700*F. I.5 Conclusions Using the DBA as the limiting break, a bounding analysis was i performed with the CHASTE code. The results of the calculations show that for each of the five fuel types in the Pilgrim Reload 4 core, a MAPLHGR multiplier which is independent of exposure must be I-3
, o .
NEDO-24224-1 applied to the MAPLHGR values given in Section 14 in NE00-24224 (see Table 14-1 in Section 2 in this report for values of the multipliers). The use of these multipliers conservatively determines the MAPLFGR required to keep the PCT below 2200 F for the CBA, with no credit assumed for core spray heat transfer. The calculations described in this Appendix were performed at the request of the Boston Edison Company (BECO) in order to support BECO's proposal to return to service taking no credit for core spray heat transfer. The technical justification for such cal-culations has been presented in this Appendix. However, General Electric considers the assumption of no core spray heat transfer 8 credit to be excessively conservative, based on the calculations which support the continued structural integrity of the core spray spargers (presented in Appendix G in this report), and the many recognized conservatisms in the current LOCA models. I.6 References I-1 " Supplemental Reload Licensing Submittal for Pilgrim Nuclear Power Station Unit i Reload 4," NEDO-24224, dated November, 1979. I-2 " Loss-of-Coolant Accident Analysis Report for Pilgrim Nuclear Power Station, NEDO-21696," dated August,1977. I-4
~ ~ - . = . . . . - . _ .
o , NE00-24224-1 TABLE I-1 Pilgrim Large Break Results LPCI Injection Valve Failure; 2LPCS + HPCI + ADS Available No Core Spray Heat Transfer with Channel Cooling Break Size ' Uncovered Chase (ft.2) Time (sec) PCT (cf)* 4.343 (DBA) 154 2200 3.474 (80% DBA) 145 2142** 2.606 (60% OBA) 133 2062** 1.000 149 2068**
*MAPLHGR = 11.55 kw/ft. for 80262 fuel at 10,000 mwd /t
** Conservatively estimated by using temperature differences for a MAPLHGR of 11.08 kw/ft.
l
- I-5
=== .-
e 9 NE00-24224-1 Table I-2 Pilgrim Small Break Uncovery Time Break Uncovered Failure Size (ft2) Time (sec) Assumed 0.900 154 LPCI Injection Valve 0.800 133 0.700 131 0.600 102 0.500 95 0.400 90 0.300 75 0.200 107 II 0.150 109 HPCI Injection Valve 0.100 , 157 0.080 151 0.'060 155 0.040 132 II 6 I-6 [_ _ _.
--_ c- -}}