ML20053E892

Core Spray Sparger Crack Analysis at Peach Bottom Atomic Power Station,Unit 2

ML20053E892
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Site:	Peach Bottom
Issue date:	05/31/1982
From:	Brandon R, Gridley R, Kiss E GENERAL ELECTRIC CO.
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ML20053E884	List: ML20053E883 ML20053E892
References
82NEDO60, NEDO-22139, NUDOCS 8206100183
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{{#Wiki_filter:, NED0-22139 DRF E21-0065-- [ 82NED060 Class I May 1982 CORE SPRAY SPARCER CRACK ANALYSIS AT PEACH BOTTOM ATOMIC POWER STATION UNIT 2 O \\ U Approved:[ [g7Mo-Approved: randon, Manager E. Kiss, Manager N ciear Services Engineering Plant Materials and Mechanics Section Operation Nuclear Power Systems . Engineering Department 4 Approved: R. . Gridley, Manager Fuel and Services Licensing Safety and Licensing Operation 8206100183 820604 PDR ADOCK 05000277 ) NUCLEAR POWER SYSTEMS DIVISION. GENERAL ELECTRIC COMPANY SAN JOSE. CALIFORNIA 95125 GENER AL h ELECTRIC

NEDO-22139 O IMPORTANT NOTICE REGARDING CONTENTS OF THIS REPORT PLEASE READ CAREFULLY This report was prepared by General Electric solely for Philadelphia Electric Company (PECo) for PECo's use with the U.S. Nuclear Regulatory Commission The information (USNRC) for the Peach Bottom Atomic Power Station Unit 2. contained in this report 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 information in this document are contained in the General Electric Company Proposal No. 424-TY727-HE2 (" Core Spray Sparger Crack Analysis for Philadelphia Electric Co., Peach Bottom Atomic Power Station Unit 2", dated April 1982), and PECo P.O. No. PB306903, Change Order No. 3. The use of this information except as O, defined 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 complete-ness, accuracy or usefulness of the information contained in this document or that such use of such information may not infringe privately owned rights; nor do they assume any responsiblity for liability of damage of any kind which may result from such use of such information. O 11 t

~ l NEDO-22139 O CONTENTS Page l-1 1. INTRODUCTION AND

SUMMARY

1-1 1.1 Structural 1-1 1.2 Lost Parts 1-1 1.3 Effect on LOCA Analysis 1.4 Reference; ~ 2-1 CORE SPRAY SPARGER STRUCTURAL INTEGRITY 2-1 ] 2. 2.1 Sparger Configuration 2-2 2.2 Fabrication Sequence 2-3 2.3 Installation Sequence 2-3 2.4 Performance History 2-3 2.5 Potential Sources of Stress 2-3 2.5.1 Fabrication Stresses 2-7 2.5.2 Installation Stresses 2-9 2.5.3 Stresses During Normal Operation 2-9 2.5.4 Stresses Dbring Core Spray Injection 2-12 2.6 Materials Aspects of Cracking 2-12 2.6.1 Potential Causes of Cracking 2.6.2 Effects of Cold Work on IGSCC of Stainless r 2-14 Steel 2-16 Conclusions of Sparger Cracking ' 2-16 2.6.3 (~)N A-2.7 Crack Arrest Assessment / 2-17 2.7.1 Stresses Due to Bracket,'Res'traint ~ j 2-17 ] 2.7.2 Fabrication Residual Stress 2-19 2.7.3 Conclusions on Crack Arrest 2-19 2.8 Structural Integrity with Cracks ,/ 2-20 2.9 References l 3-1 3. LOST PARTS ANALYSIS 'r / 3-1 3.1 Introduction 3-1 3.2 Loose Piece Description 3-l' 4 3.3 Safety concern 3-1 3.4 Safety Evaluation 3-1 3.4.1 General Description 3-2 j 3.4.2 Postulated Loose Pieces 3-5 3.5 Conclusion 3-6 3.6 Reference LOSS-OF-COOLANT ACCIDENT ANALYSIS WITH NONUNIFORM 4-1 4. SPRAY IN ONE SPRAY SPARGER 4-1 4.1 Introduction 4-1 j Input to the LOCA Analysis 4-3 4.2 Sensitivity of LOCA Analysis to Non Uniform Spray 4-4 4.3 4.4 Analysis Results 4-5 4.5 Conclusions 4-5 4.* 6 References 111 l

..._____m.._- 4 l 3 NEDO-22139 I l CONTENTS (Continued) P. ale l APPENDICES A-1 STRUCTURAL ANALYSIS OF THE PEACH BOTTOM 2 CORE SPRAY SPARGER A. B-1 B. SPARGER TEMPERATURE CALCULATIONS C-1 [- C. FLOW VELOCITY CALCULATIONS l i I l i i s l I i t P O iv r 1 t _,.. n, a- -.an ..e,n,.--

NEDO-22139 O ILLUSTRATIONS Figure Title Page 2-22 2-1 Core Spray Sparger - Elevation View 2-23 2-2 Core Spray Sparger - Plan View 2-24 2-3 Sparger to Shroud Attachment Method 2-25 2-4 Sparger Nozzles 2-26 2-5 Sparger Support Method 2-27 2-6 Pipe Bending Method 2-7 Sequence of Events Leading to the Residual Strebs 2-28 Distribution 2-8 Bilinear Stress-Strain Curves for Type-304 Stainless 2-29 Steel 2-9 Stress and Strain Distribution in the Pipe Under 2-30 Applied Moment 0 2-31 2-10 Moment Versus Outer Fiber Strain 2-11 Resultant Residual Stress Distribution After Fabrication 2-32 2-33 2-12 Postulated Installation Stresses 2-13 Effects of Cold Work on ICSCC of Type-304 Stainless 2-34 Steel 2-35 2-14 Effects of Cold Work on IGSCC 2-36 2-15 Effects of Cold Work on ICSCC 2-37 2-16 Effects of Cold Work on IGSCC 2-38 2-17 Stress Versus Cold Work 2-39 2-18 Compliance Change, Cracked Pipe 2-40 2-19 Assumed Stress Distribution on the Crack Face 3-7 3-1 Reactor Vessel 3-8 3-2 Steam Separator O V

NEDO-22139 O f ILLUSTRATIONS (Cor.tinued) i I Figure Title Pm 3-3 Largest Piece That Can Fit Through the Turning Vane with the Long Dimension in a Horizontal Plane 3-9 3-10 3-4 Orificed Fuel Support 3-5 Fuci Assemblies and Control Rod Module 3-11 A-6 .A-1 Flow Paths r I t l [ I l [ I.lO t l 6 O vi vr o rav sww- - w m -w-s~ m =- m m mmen ,->r---= m.e-w=----s-mwa -w m-mm -=amm-.-w-v*m-m -e+

NEDO-22139 1. INTRODUCTION AND

SUMMARY

One of the scheduled tasks during the Reload 5 refueling and mainte.aance outage in March 1982 at the Peach Bottom-2 Atomic Power Station was the performance of a visual inspection of the core spray spargers using underwater television This inspection was conducted as required by IE Bulletin No. 80-13 cameras. (Reference 1-1). During this inspection, a 180 degree, circumferential1y oriented crack in the header to T-box weld heat-af fected zone of the lower core spray sparger was found. General Electric reviewed this condition and is providing justification for continued operation without the installation of additional hardware by address-ing the following items. 1.1 STRUCTURAL A structural analysis is presented in Section 2, which describes the potential sources of stress in the spargers resulting from fabrication, installation, normal operation, and operation during postulated loss-of-coolant accidents Potential causes of cracking are also discussed, and it is concluded (LOCAs). that the structural integrity of the sparger will be maintained for all condi-tions of operation. 1.2 LOST PARTS If breakage of the sparger is postulated, the lost parts evaluation presented in Section 3 concludes that the potential for unacceptable flow blockage of a fuel assembly, or for unacceptable control red interference, is essentially It is also shown that loose pieces are not expected to cause damage to zero. the other reactor pressure vessel internals. 1.3 EFFECT ON LOCA ANALYSIS Section 4 presents the results of LOCA analyses assuming no core spray heat This corresponds transfer credit from the cracked sparger in the calculations. 1-1

r NEDO-22139 to a postulated worst-case core spray sparger break in which the water flowing The through the cracked sparger does not spray uniformly onto the core. resulting increase in peak cladding temperature (PCT) was calculated at the request of Philadelphia Electric Company assuming no clamping repair of the sparger is implemented. The analysis is considered to be conservative by General Electric based on the calculations which support the continued struc-tural integrity of the sparger and large conservatisms in the LOCA analysis as demonstrated by large-scale tests. The analysis nevertheless justifies con-tinued operation with no change in MAPLHGR limits. 1.4 REFERENCE USNRC Cracking in Core Spray Spargers, IE Bulletin 80-13. 1-1 O O 1-2

NEDO-22139 2. CORE SPRAY SPARGER STRUCTURAL INTEGRITY 2.1 SPARGER CONFIGURATION The core spray sparger configuration is shown in Figure 2-1 through 2-5. Vertical The spargers are mounted in the upper shroud, as shown in Figure 2-1. The upper sparger has spacing is 12 inches between header pipe centerlines. The plan bottom-mounted nozzles and the lower sparger has top-mounted elbows. The shorter header view (Figure 2-2) shows that the spargers are asymmetric. pipe has an arc length of 82.5*, and the longer header pipe has an arc length The T-boxes for the spargers are located 17.5* from the vessel 0* of 97.5*. and 180* azimuths. The T-box is a Figure 2-3 shows the attachment of the T-box to the shroud. 6-in. Schedule 40 section of pipe with an end plate toward the vessel center-The 6-in. pipe extends through the shroud wall and is butt-welded to line. The T-box pipe is attached to the shroud by the seal ring external piping. ) with the attachment welds to the 6-in. pipe and the exterior surface of the shroud wall. The Peach Bottom-2 upper The sparger flow nozzles are depicted in Figure 2-4. core spray sparger header uses 1-in. shielded VNC nozzles alternating with The lower header uses 1-in. shielded VNC nozzles alter-SPRACO 3101 nozzles. nating with 3/4-in. open elbows. header pipe The 97.5* header pipe is supported at three locations and the 82.5 Figure 2-5 shows the support arrangement at is supported at two locations. The brackets are 3/8-in. thick and ~ locations other than at T-box locations. The pipe-to-bracket mating surfaces are not welded are welded to the shroud. d to allow circumferential relative motion between the header pipe and the shrou during a core spray injection of cold water into a system at reactor operating Schedule 40 Type-304 stainless steel. The header pipe is 4-in. temperature. 2-1 i -. ~. y v

NEDO-22139 11 s aod tae c1ese ieries O The street e1bews. 90 18ews. a 1r-ce 9 (used to connect the e1 bows and orifice the elbows) are all Type-304 stainless steel. 2.2 FABRICATION SEQUENCE Fabrication records show that the Peach Bottom-2 spargers were fabricated as follows: The pipe was bent using a four-roll bending process as shown in 1. Figure 2-6. 'The rollers have 2-1/4 in. radius grooves, and rollers 3 i and 4 are adjustable to accommodate the pipe size and to bend the pipe In this case, the design radius is to the required radius. The maximum strain in the pipe is calculated R = 105.75 inches. to be 2.1%. j I Af ter the pipe is bent to the proper radius, it is placed in the 2. tse rive fits the esreed s-e 11t ceeditie s. O During this fit-up process, the T-box 6-in, pipe is marked for drilling sareed te verif7 tw t the header pipe holes. After removing the pipe from the shroud, the headers are welded to l 3. the T-box. The holes for each nozzle are drilled in the header pipes. 4. l Stainless steel half-couplings are bevel welded at each nozzle 5. opee l'ag. The elbows are screwed into the assembly and roughly aimed. i' 6. l l l i ' O ( i 2-2

r NEDO-22139 ( m. d 2.3 INSTALLATION SEQUENCE The sparger is installed in the shroud in the following manner: The brackets are welded to the shroud, thereby positioning and holding 1. It also includes attaching the T-box to the shroud by the spargers. It is assumed that, velding the seal ring to the T-box and the shroud. because of interference between sparger ends, one or more of ti.e This operation spargers would be cold sprung during installation. was not addressed in the fabrication records. The next operation was to aim the nozzles as required by the sparger 2. drawing. The elbows were then tack welded to assure that the threaded connec-3. tions remain intact. \\ 2.4 PERFORMANCE HISTORY There have been Peach Bottom-2 Station first went critical in September 1973. no inadvertent core spray injections. Peach Bottom-2 does flush the core Water is pumped from condensate spray spargers during refueling outages. The maximum AT that has storage at a temperature of approximately 70*F. l occurred is 120'F. This AT is sufficiently low that fatigue is not a concern. 2.5 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 spray system during postulated loss-of-coolant accidents are presented in this section. 2.5.1 Fabrication Stresser: () Residual stresses are developed when an initially straight pipe is subjected to a moment sufficient to cause yielding and later unloaded, as would occur 2-3

r NEDO-22139 during the fabrication cf the core spray spargers. The fabrication operation is idealized in Figure 2-7. 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 final unloaded g radius of curvature. Determine the stress distribution associated with this moment. 3. Calculate the elastic stress distribution corresponding to the applied moment (-M ) to describe the unloading. 4. Determine the residual stress in the pipe which is the algebraic sum of the elastic-plastic stresses due to M and~the elastic stresses g due to (-M ). O In calculating the moment-curvature curve for the pipe, thin shell theory was applied and a representative bilinear stress-strain curve (Figure 2-8) was used. As shown in Figure 2-9, the strain varies linearly through the section, while the stress follows the bilinear curve for angles' greater than 0. The applied moment (M ) is given by: g 0 r M =2 (Ec Sin 4) (a sin &) (2atd$) l t 9 r /2 n +2 { (c Sin $ - c ) E + Ec } (a Sin $) (2atd4) (2-1) Jg o y t y where O c = a/R = outside strain a = radius of pipe 2-4 )

r NEDO-22139 (m) R = radius of curvature c,o = yield strain and stress E, E = elastic and plastic modulus The first term in Equation 2-1 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, Equation 2-1 becomes: E "(1 - E /E) (20 - Sin 20) t g + c s0 + E sine Nt o n sinD c R and Sine = c /c =1-y o a n moment corresponding to the first onset of yielding on the outside s_s M = g surface = a r a't. y Clearly, for fully elastic behavior, 0 = r/2, and M = M,. Figure 2-10 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 105.75 inches, the outer fiber strain during bending is 2.33%. 2 The corresponding moment is 1.43 o na t. y The residual stress distribution can now be determined by combining the elastic stress corresponding to (-M ) and the elastic-plastic stress during bending. t ~ A correction for the thin Figure 2-11 shows the resulting stress distribution. shell theory assumption is included in the results. Figure 2-11 shows that the pipe is subjected to high residual stresses (approaching the yield stress), and that the stress distribution varies [) In particular, it shows tensile stresses around the circumference of the pipe. 2-5

e NED0-22139 on the surface facing the centerline of the vessel. It should be noted that the actual stresses could be higher due to local yielding at locations where Hertzian 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 tempi.ature condi-tions. However, for reactor operating temperatures =r550*F, the residual stresses are expected to 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 propagate intergranular stress corrosion cracking (IGSCC) under sustained load. Using the following worst-case assumptiens: si dn., 1. K = IGSCC O 2. A long continuous crack, 3. Sustained stress up to yield = 18.8 ksi, the minimum crack depth for crack growth is given by: .1 pa K = min IGSCC l l l or = 0.026 inches " min " 18.8 1.12 This shows that under worst-case conditions, a 26-mil crack could propagate due to stress corrosion cracking. C) l l f 2-6 I

r-NEDO-22139 The conclusions from the evaluation of fabrication stresses presented in this section are summarized below: Stresses due to fabrication could be significant and would exist 1. throughout plant operation. A possible synergistic combination of adverse metallurgical conditions 2. (e.g., sensitization, cold work) and high residual stresses may explain the observed cracking. Since the stresses change sign (become comprehensive) around the 3. circumference, a crack that initiates in the tensile region can be expected to arrest in the compressive regions. 2.5.2 Installation Stresses Stresses sufficient and necessary to cause initiation and propagation of cracks by intergranular stress corrosion cracking (IGSCC) can be identified by postulating certain installation variables. Figure 2-12 shows two cases which might be postulated. In Case 1, it is postulated that differential weld shrinkage occurred during The outer bracket would provide a welding of the header pipes to the T-box. 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-in. differential weld shrinkage l The deflection resulting at the header end would be approximately: l is assumed. 105.75 ; A = 2.94 inch. = Then, from Reference 2-1, Table 13.4, Case 1: A = WR4EI (2$ - sin 24), where $ = 90*. l l 2-7 i t

NEDO-22139 [ Solving W = 648 lb, assuming: R = 105.75 in. 6 E = 28.3 x 10 ksi I = 7.23 in.4 Since M = WR WRC 648 x 105.75 x 2.25 I 7.23 o = 21300 psi o = 21000 psi (elastic) For Case 2, it is assumed that R is incorrectly fabricated to a radius of It is further assumed that the vessel brackets cause a uniform 104.75 inches. moment on the pipe, thus increasing the radius to 105.75 inches. O The initial inner length is n/2 x 102.5 = 161.01. After forming, the inner length is w/2 x 103.5 = 162.58: 162. = 0.010 = Strain = c = 10 inner = 1.0% Using a stress strain curve for Type-304 stainless steel, the resulting second-ary stress is found to be 38,000 psi for 1.0% strain. For the postulated conditions, these two examples show that high deflection These stresses have limited tensile stresses can occur during installation. not been confirmed. In addition, the welding process produces residual stresses in the pipe near the weld. The magnitude and sign of the stresses vary with dis-These stresses tance from the weld and depend on pipe size and welding speed. are likely to vary circumferential1y. Maximum tensile residual stresses in the range of 18 ksi to 40 ksi have been measured in weld pipe tests (Reference 2-2). 2-8 ~' m

r-NEDO-22139 l Installation stresses considered in conjunction with the material considera-tion discussed later (Section 2.6) may explain the cracks that have been It should be emphasized that the installation stresses postulated observed. above are all deflection-limited secondary stresses that will relax to the elevated temperature yield strength of the material during normal plant operation. 2.5.3 Stresses During Normal Operation 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, pressure, thermal mismatch, stagnant line top-to-bottom temperature gradients, stagnant line throughwall temperature gradients and Stress calculations are given in Appendices A and B. weight. 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 4.45 lbf/in. c1 the header arm, resulting in negligible stresses. Weight of the spargers and Stagnant water is only 1.36 lbf/in., again resulting in negligible stresses. line temperature gradient calculations are not provided since the maximum AT for top-to-bottom gradients and for through wall gradients were found to be It should be less $an 10*F, which would result in insignificant stresses. noted, however, that the AT for core spray injection is addressed in Section 2.5.4. in stresses It is concluded that the normal operating loadings do not result that could explain the crack observed in the Peach Bottom-2 core spray sparger. 2.5.4 Stresses During Core Spray Injection Stresses during core spray injection are the design stresses for the spargers. Design loadings include all those discussed in Section 2.5.3 plus those that occur because the system is no longer a passive system. The pressure differen-tial in the sparger at rated flow is approximately 25 psid. The hoop stress in the pipe is about 210 psi. Impingement load stresses are less during spray J injection than during normal operation. Thermal stresses due to the throughwall temperature gradient are high and are known to be: 2-9

r NEDO-22139 E + a AT 2(1 - p) The radius of the These stresses are not a concern for one or a few cycles. sparger shrinks when the sparger is cooled, resulting in secondarp bending stresses of approximately 3000 psi. The axial stress in the pipe due to AP and bracket friction is low--only 420 psi. Flow through the nozzles results in a torsional stress which is low--less than 100 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. 2.5.4.1 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) because of the draining of water from the spargers and/or the flashing of water to steam during depressurization prior to core spray injection. For the purpose of maximizing injection loads, primarily on the core spray is assumed that reactor pressure is essentially atmospheric (as spargers, it 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 accel-eration head is reduced by friction and local losses. If L, is the equivalent is given by application of length of the pipe system, the final velocity Vf the energy equation: L V 2g The maximum velocity attainable is limited to that at system runout flow (8000 gpm), which produces a velocity of 55 ft/sec in the sparger (at the entrance to the long sparger arm to be more concise; the velocity at the ends is zero). 2-10

NEDO-22139 Actually, the velocity of the water first entering the sparger will be less than runout velocity because of the relatively slow opening characteristics The injected water fills the pipe line between the of the injection valve. 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 sparger is: V (55) = 40.8 psi m 144 gv 144(32.2) (0.0160) or F =P A = 40.8(12.73) = 519 lb. m m p where P = momentum pressure (psi); F = momentum load (1b); V = velocity (ft/sec); 2 = gravitational acceleration (32.2 ft/sec ); g 3 = specific volume (0.0160 ft /lb) (480*F water); and v = pipe flow area (12.73 in. ) (4-in. Schedule 40 pipe). A If the end plates at the ends of the spargers were removed, it is obvious there Now cap the ends and also plug the sparger nozzles. would be no impact load. Again, there would be no vater 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 there are is much closer to the open end condition, except that extremes. It O several " ends" instead of one end, and they are located along the length of the sparger arms. 2-11

NEDO-22139 O The exit flow area of the sparger nozzles is computed as follows: Area Total Area Number (in.2) ( t,,,2 ) 1-in. VNC Nozzle 32 1.018 32.6 3101 Nozzle 33 0.307 10.1 Total Open Flow Area Per Sparger = 42.7 The exit flow area of the nozzles and elbows is actually 68% greater than the flow a:ea of the two sparger arms (2 x 12.73 = 25.46 in.2), An 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 Steam would be pushed ahead of the open elbow located at the end of each arm. oncoming front of water, exiting the sparger through the assumed single nozzle. The developed differential pressure to expel the steam would be approximately 7 psid. Adding all sparger elbows and nozzles to this logic clearly demon-strates that the sparger indeed behaves like an open-ended pipe, and conven-tional 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. 2.6 MATERIALS ASPECTS OF CRACKING The potential causes of Peach Bottom-2 core spray sparger cracking are dis-cussed in this section. A general discussion of the effects of cold work on the IGSCC susceptibility of Type-304 stainless steel is also presented. 2.6.1 Potential Causes of Cracking The potential causes of core spray sparger cracking which are considered to be The table addresses cracking near most probable are indicated in Table 2-1. the heat-affected zone (HAZ) of the T-box to sparger arm weld, as well as The crack in the Peach cracking in the sparger arm remote from the weld. The evidence Bottom-2 core spray sparger is located within the HAZ of the weld. of each possible cause is also indicated. 2-12

NEDO-22139 Near the T-box, four possible causes of sparger cracking have been identified. sensitization by welding the sparger arms to the tee is supported by Fir st, the patterns of cracking near the HAZ of this weld. IGSCC may result if stresses are suf ficiently high in this area. Such cracking has been observed in piping incidents in tha past. Second, cold work inherent in arm forming followed by weld sensitization may As discussed below, increase the susceptibility of IGSCC in the spargers. sufficient cold work is present for enhancement of cracking tendencies. Third, fatigue induced by thermal variations in the environment may be the cause of the sparger indications. However, the variations in temperature during operation of the reactor (10*F, see Section 2.5.3) are expected to be No evidence of a driving force for thermal fatigue has been identified. small. Finally, fatigue resulting from flow-induced vibrations could be hypothesized. However, the natural frequencies of the sparger are high relative to any flow-induced excitation sources, and the sparger brackets restrain the amplitudes of any vibrations. In the arms remote f rom 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 was 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 susceptible condition would more readily exist. Again, no direct evidence Thirdly, surface cold work resulting from arm bending or straightening exists. could hasten crack initiation and subsequent growth could occur from residual No documentation exists to support this possible or installation stresses. Finally, fatigue by either of tbe sources cited above for the T-box area cause. 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 in a region of weld residual stress. Approximately 5% cold work could result from sparger arm fabrication and installation. 2-13

NEDO-22139 O 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. Effects of Cold Work on ICSCC af Stainless Steel 2.6.2 The mechanisms of cold work enhanced cracking are complex but can be visualized In this illustration, factors through the illustration in Figure 2-13. influencing susceptibility to cracking are shown as increasing or decreasing susceptibility by lying to the left or right of the diagram, respectively. This enhances Cold work serves to increase the material yield strength. susceptibility if stresses in the material result from imposed strains because the resulting stress state of the material would also be higher, consistent If stresses are fixed as the result of with the increased yield stress. imposed loads, susceptibility may decrease because the stress state of the hardened material is a lower fraction of the yield stress. Cold work serves to promote chromium activity in the material matrix, which reduces susceptibility through the more rapid recovery of chromium-depleted However, sufficient time at higher temperatures (>500*F) is necessary regions. for the recovery phenomenon, and such thermal treatment was not practical for the spargers, nor deemed necessary. The most significant influence of cold work is in the transformation of Martensite, if present austenite to martensite phases through deformation. in sufficient quantity, can assist in recrystallization of the material upon The strain energy induced in the lattice promotes subsequent thermal treatment. The result of recrystallization is migration of grain recry stallization. boundaries away from chromium depleted regions, with attendant benefits in However, the presence of martensite increases the reducing sensitization. tendency for carbide precipitation and local chromium depletion during A wider HAZ can result from welding stainless subsequent weld sensitization. If suf ficient cold work is steel with prior cold work-induced martensite. 2-14

NEDO-22139 O present, transgranular cracking can occur in oxygenated water environments with or without subsequent sensitization. Environmental tests conducted on tensile, bent beam and pressurized tube specimens are illustrated in Figures 2-14 through 2-17 (which are based on information from References 2-3 and 2-4). f# In Figure 2-14 it can be seen that the time to failure in 0.2 ppm 02 Speci-sensitized and cold-worked and sensitized material varies with stress. mens tested at cold-worked-plus-sensitized conditions (at higher stresses) produced failure times (by 1GSCC) comparable to samples which contained no work prior to sensitization. Cold-worked samples without subsequent sensitira-IGSCC failures could be tion, tested at comparable stresses, did not fail. induced at very high stresses in cold-work /no-sensitization samples, as illustrated in Figure 2-15. If the data from Figures 2-14 and 2-15 are plotted on a basis normalized by the material yield strength, a more clear picture is formed of the results of Material cold deflection-induced stresses in stainless steel (Figure 2-16). worked to various levels and subsequently sensitized can undergo stress corro-substantially lower percentages of the material yield strength, with sion at in quarter-hard stainless steel (furnace sensitized). cracking as low as 80% s An equivalence must be established betwcen plastic strain during sparger arm The yield strengths forming and the cold-work condition of the test specimens. of specimens receiving 5, 8, and 15% cold work are illustrated in Figure 2-17. 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 cal-culated 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 concentration from localized bending, if a f actor of 4 is considered, would be comparable to 5% cold work. A reduction of the cracking threshold and cracking under residual and installation stresses could occur. to 0.8 o 2-15

NEDO-22139 O Conclusions of Sparger Cracking 2.6.3 d by Core spray sparger cracking at the Peach Bottom-2 plant can be hypothesize f the arm material the influence of weld sensitization or prior sensitization o i Sources and subsequent cold work of the arms during forming and installat on. bending, weld of stress for IGSCC are dependent on residual stresses from arm residual stresses, and deflection during installation. hly The principal factors suspected of causing cracking are considered to be h The absence of one or several key factors variable from one plant to another. f the BWR may explain the lack of reported indications in the majority o operating plants inspected to date (May 1982). 2.7 CRACK ARREST ASSESSMENT the following sources of stress In assessing the possibility of crack arrest, are considered: Stress due to pressure, mechanical loads and thermal gradients. These stresses have been shown to be negligible and are not considere 1. in the crack growth assessment. these are displacement controlled Stresses due to bracket restraint: 2. (secondary) stresses and would be cxpected to relax as the crack propagates. as the crack propagates into a Residual stress due to fabrication: 3. region of compression, the stress intensity factor can be expected to decrease, thereby resulting in arrest. weld residual stresses at the T-box - sparger 4. Weld residual stress: I These stresses are likely to welds would influence crack propagation. vary circumferential1y and also relax as the cracks become larger. ) Stresses due to vibration are assumed to be negligible. 5. 2-16

NEDO-22139 O n considering crack arrest, the stresses due to bracket restraint and I the fabrication residual stress are significant and are evaluated in

detail, 2.7.1 Stresses Due to Bracket Restraint Stresses due to bracket restraint are governed by the applied displacement Since the displacement is fixed, the and the compliance of the pipe.

This is compliance change with crack growth could lead to crack arrest. in a bolt-loaded wedge-opening-loading (WOL) comparable to crack arrest Figure 2-18 shows the variation of specimen in stress corrosion tests. The compliance compliance with crack length for a pipe subjected to bending. was determined using the relationship between the strain energy release rate G and the compliance change per unit area of crack extension de/dA For the cracks in the sparger, L/d is expected to be (Reference 2-5). in the range of 0 < L/D < 40. Figure 2-18 shows that, when more than 30% ( of the pipe is cracked, the compliance of the pipe increases by a factor the stress in the Therefore, for the given initial displacement, of 10. sparger and the applied stress intensity factor would decrease by a factor

Clearly, of 10 when more than 30% of the pipe circumference is cracked.

the the crack length exceeds this value, the restraint stresses become negligible and crack arrest is expected. 2.7.2 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 The assumptions made are as calculate the stress intensity f actor. follows: The crack in the sparger is modeled as a through crack in an 1. () infinite plate. 2-17

NEDO-22139 It is assumed that the tensile stress (a) is unifois and is applied 2. on the crack face over a length (2b). (Later this will be conserva-tively taken as 25% of the circumference.) The remaining portion of the crack is assumed to be subjected to a 3. compressive stress, which is half the tunsile stress (Figure 2-19). The crack length (2a) for which the combined stress intensity factor 4. reduces to zero is calculated. The stress intensity factor due to the tensile stress can be shown to be: -1 @ tension, 2ca sin K \\a/ I g The stress intensity f actor due to the compressive stress a/2 is given by: c mpression,-2(o - sin ~ a K Setting K *"8 + K

• P = 0, we get C

7 7 =f' f-sin' ~1 sin 1 or, sin =f ~1 or, b = 0.5a and the remaining If we assume b = 25% of the circumference is under tension oy portion of the crack is under compressive stress (equal to half the tensile) stress), the applied stress intensity factor bec'omes zero when the crack length Thus, even under extremely conservative is equal to 50% of the circumference. assumptions, crack arrest is expected. 2-18

NEDO-22139 2.7.3 Conclusions on Crack Arrest Based on the above material, the following conclusions may be drawn: 1. Since the applied loading is predominantly displacement controlled, Crack the stresses can be expected to relax as the cracks grow. arrest is therefore expected. 2. The residual stresses due to fabrication vary from tension t3 compres-sion. As the cracks propagate into regions of compressive stress, the K value reduces to zero. Even for extremely conservative assump-tions, crack arrest can be shown for a 50% circumferential crack. The above conclusions are valid as long as there is no stress cycling 3. due to vibration (e.g., flow-induced vibration). 2.8 STRUCTURAL INTEGRITY WITH CRACKS From the discussion of the potential stresses in the core spray sparger (Sections 2.5 and 2.6), it is concluded that only deflection limited secondary stresses approach 25% of the material yield strength (except for self equili-brating thermal stresses). If a 360* throughwall crack is postulated at any location on any sparger arm, the remaining stresses will not produce a failure The AP stress and the stress resulting at any other location on the sparger. from an axial load in the pipe due to bracket friction are proportional to 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 0.033 in.2 is required to maintain continuity. The This area is much less than the original pipe metal area of 3.17 in.. The bending type stresses are all deflection limited secondary stresses. since the discussion in Section 2.7 shows that cracks are expected to arrest, driving stress will be relieved. The bending loads may, however, in a worst case, cause an existing crack to.open up by an additional 0.005 in., assuming the existing crack has progressed 360*. This is a geometry limited condition. 2-19

I NEDO-22139 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, l f transients, or postulated loss-of-coolant accidents.

2.9 REFERENCES

Hopkins, Design Analysis of Shaf ts and Beams, McGraw-Hill Book Company. 2-1. t H. H. Klepfer, et al, " Investigation of Cause of Cracking in Austenitic 2-2. Stainless Steel Piping," NEDO-21000-1, July 1978. l i A. E. Pickett and R. G. Sim "The Effect of Stress and Cold Work on 2-3. Intergranular Stress Corrosion", Materials Protection and Performance, Vol. 12, No. 6. June 1973. i G. M. Gordon and R. E. Blood. " Reactor Structural Materials Environmen 2-4. Exposure Program", Symposium on Materials Performance in Operating Nuclear 28-30, 1973. Systems, Ames Laboratory, Ames, Iowa, August /~T E. Kiss, J. D. Heald, D. A. Hale, " Low Cycle Fatigue of Prototype Piping", %/ 2-5. GEAP-10135, January 1970. i 1 i l O 2-20

NEDO-22139 Table 2-1 () POSSIBLE CAUSES OF CRACK 1!.- Evidence Possible Cause Location -ation of cracks 1. Sparger Arm Sensitization by Welding. Near T-Box

imate 5% Cold Work Cold Work Followed by

.. war T-Box Weld Sensitization iacation of cracks Weld Residual Stresses 1.T's are Low Fatigue (thermally Induced) Fatigue (Flow-Induced Vibration) Amplitudes are Limited None.* Sensitization fron Fabrication 2. Sparger Arms Pipe Bend Forming *, Cold Work Followed by No Evidence of Away from Sensitization Sensitization T-Box None* Local Heavy Cold Work Same as in 1. Above Fatigue ()

• Sensitization and cold work state of spargers not yet known.

O 2-21

geNCc O g b e ._UT I N rI SE L w Z e Z i 0 V N no Y i y A t R a s P ve \\ S l \\ '\\ E \\ r m e U g r ^ a [.I O p tg 4,. S l F [J y L a A rp H S R e E r G o R C AP S 1 2 f e r ug i / F . i T I 'C O I o~ !tf f

NEDO-22139 C' O K in

818 r

y i / -= \\

N

/ / / N [ CORE SPRAY SPARGER ses 1 .s 4. ) \\ N \\\\ ^ / \\ '\\ / 6-1h!?h"g O 180* Figure 2-2. Core Spray Sparger - Plan View 2-23

NEDO-22139 O s N{ _Ea~00D comE A j \\\\ l!'Oh ,-- Lc2.n. GIP' 6PidAI. 1 n . SEAL _LOWERSPARGER TOP GUIDE %=) O Figure 2-3. Sparger to Shroud Attachment Method 2-24

NEDO-22139 d y@ A ) O 6 O ure 2~0' sparger NO M 2-25

f NEDO-22139 O r i lSMRouc l 9 l \\ h [ LDE To ovo /(U o [ ONE E El h~n -_o_ER BR ACKE.T. LW O Figure 2-5. Sparger Support Method 2-26

i 1 NEDO-22139 1 0 l 105.75 in. 8 1 l

T Figure 2-6.

Pipe Bending Method O 2-27

=-. NED0-22139 O

• INITIAL CONFIGUR ATION 1

i l M

• M 2

_t ~ O LOAD APPLICATION DURING P ABRICATION / \\ / S \\ / / /s / \\ g s / N e' N s / s s' \\ 's

1. s' N

/ N %'-"~~__ ,/ p u,1 km,1 .s ,/ ~ s' % "* ein, - - " es* # .s** .PINAL DEPORMED SHAPE AFTER LOAD REuCVAL Figure 2-7. Sequence of Events Leading to the Residual Stress Distribution 2-28

{ NEDO-22139 O ACOM Tggp gy4.,, Q =- Y O 34.3 ksi E 'y

• 0.0012

'y

• 3 E=

28.3 x 10 k,; 3 E* 0.26 x t0 ksi T f 'y E (a) O ,550*F "Y - il 18.8 ksi 8y 00007 'y

• 3 E=

25.8 x 10 ksi 3 0.26x10 a,; ET* l E (b) O -Strain Curves for Type-304 Figure 2-8. Bili Stainless Steel 2-29

NED0-22139 O

l.. a r_

r eO" ~ "i {e=e,+EO~'y t k

• ELASTIC #LASTE

! / O g.e d)Tu h..e = - w5 y 7p 5TRE'SS t ' DISTRIBUTION e e/E E I = STRAIN l l DisTaisVTION O Figure 2-9. Stress and Strain Distribution in the Pipe Under Applied Moment 2-30

UH$ go O i t ee 3 s l I n u e s

ss l

a. ]M ' 2 l 34 ? 1 l s II ll, f 8IleI8 i Aa et eglIl**: s e I si in i l a o' s r o i c l t il c S I l re g k b ( i N F I l A l g R r .p x T e I S e S t U n R u u E O ID T ' I s ll A A F R V u L R R s - O A u h r T N c 'U e O V I F e e j tne tno M ) 0 s 1 o 0 1 l c I 1 o l 2 e rug i F l y s!. s !tl,i:lI, ( ~ 6 = 2 2 .Dl si 2 i 3 c e b i b 1 l l l ( 8 I ~ lt O , >l3 t 'l N s !I i t:,l! !l, I

NED0-22139 L 8 s- .iga -ue g g$ .a1 .w-la e IE h - e ii- !!E $n sa k \\ k\\ ~ \\' 8 3' - M i .c \\ t i.. a L W e i 1 .L f a M 3 'g e C ] .w I' cx = 15 8 2m !G E Ikb ~

e =

kk 0 a id E i q q' Io i 2-32

NEDO-22139 J INSTALLATIONJI ADI AL MISM ATCH OtP P ERENTI'AL Wf LO (SHRINKAGE AT T-40X ) l R1 R1 = R2_ AAM 5 f

• SR5 7

FOR SHRINK = 1/8 in. o AT T-BOX a 21 ksi (ELASTIC) O 1 ASSUME R1 = 104.75 R2 = 105.75 UNIFORM FORMING e a 1.0% o a 38 ksi (FROM o - e CURVE) i.R1 < R2 C.A.sO ~J, W M O Figure 2-12. Postulated Installation Stresses 2-33

NEDO-22139 O DECREASES ~ 7pi6'ITA'SES , SUSCEPTIBILITY SUSCEPTitlLITY TO CR ACKING TO CRACKING A 4 I I I l ______] IMPOSED LOADS '.888 CREASES _ Ylm _ IMPOSED STRAIN u sp pggoyg ~h PRCMOTES

CHROMlUM

._ INCESES I CHROMlUM - - - - + 'gLD WORK 'ACTl5/iTY ~ l O V PRO'6UCES e I MARTENSITE "" ~ ~ ~ ~ ~ ] I l l V. i AIDS RECRYSTALLIZATION IF 4GSCC IF SENSLTIZED,_ 8 CW > 15% A,ND HEAT TREATED (_AFTER MODER_ ATE COLD, 1 AT MiGH TEMPE'RATURE wor K ISTRESS REcutREDI -*__u==sr.=__.__ ,,C,,,,,, CAmalDE PRECIPITATION p _ __ __ __ __ _ l I l h TG5CC AFT _ER_ME AVY __ _ _ ,. COLD WCR K (STR ESS R EQUlm ED) O Effects of Cold Work on ICSCC of Type-304 Figure 2-13. Stainless Steel 2-34

r O O O l 1 I t70' YlELD. 814 HARDli l TYPE.304 ST AINLESS STEEle M 6s4 l 82m g i o ~ M. l e = =D I l F'URNACE ANDl E 6% COLD WORK .1/4 HARD, !RRAQ@TED. A Il% COLD WORK SENSITIZED jSO, " $ 1/4 HAHD 4 g O 6% COLD WORK ' 8 8% COLDWORK h 114 HARD ( ~ ~ l ise-mVidD.e%'COtD WORK l i 3% COLD WORK ANO[ .g 4 YlELD,5% COLD WORK) IFURNACE SENSITIZED l j j i 7 0 2 s 8 l 6% COL D, $j lWOHK AND FURNACE l

• u g

l $ENSITIZED 1 4 j

30', -

FURN ACE SENSlilZEU'l l 4 ~ N g _._,',,_ _'"SE M5IITIZED l YIELTIURNA'EE 1 APPROXIMATE LY SAME TIME TO F AlttWlE IN SENSITIZED! lYEAR8 liOYEARSl VERSUS COLDWORKED AND SENSITIZED MATERIAL i r l l l ,1 06 8000l ,10,000-l 100.000 l l1,000,000 i l 1001 , TIME TO F ALLURE ikall 1 Figure 2-14. Effects of Cold Work on ICSCC i +

Zbo4m O 'O L E f E T T SF IS 0 0 S* S0 N 0 E 8 E l 0 L5 S 0 N 0 N 70 U I A D I a Ts R s P a ep A t 2 H h 3- 0 4 P t 1 E 1 Y Q T I ~ 000 i o C c CS i l jl G ~O N I O NI D 'A T n A E o K Z Z I RI T k T OI N I r S iS o N W N OE \\ r l+ '$ d ~E W LS O O Q C l C E e o l C l lN A h N C t R E k U R f O g F. 0 U o \\{\\ 0 l s l 0 i A o t ) I F C K c l D C e a k R T T hmA l W S O f E f T M E L K i R O E I gI O L T [C W O - A it 5 YD O 6 1 6 i ( LE NZ l C -OT 2 I i I i S e 'S. A N r i E R u S E g I T T i D 5O F l MN E Z 0 D N 0 I i T E 0 E I KH S i N RW I E. OS l l s S _. WS O 'R ~F E E DR E A R R C 'H 6 O A ET i OS '4 i W N Ci V '6 R ? l L E U _ NG I d O O F. 'G H iI i I R ~Yl N (C 'E D NY L K S E D E l C 6 Y V 'l f A H E I I I C Y. V C ^ - to .l n Ot 3 e 04 m S S. 10 O 4 1 i. i'e N u K o5;# s N6e ii i! 'l j i i ,l I iIt} llllI jllllli

hoh[' 000 0 o 0 0 4 l i 'D D E E E Z Z I CI T A I L NIS TN g S E S E U RN OE T LU NE PF S 'E SS L A K K Kl N '2 I T R R R S O O O WW W 0 4 D 0 0 O D D 0 3 D D R R R l L 0 E L L A A A O 0 P O O S Y C C I H C 1 I I l 4 4 4 T = O % s/ / / N 5 t 1 a t $E $OQ' CCS G I no . y Q-l r k i \\ IN V i \\ . E d 0 R l 0 U o 6 0 L C l 9-l F f 0 IA 1 \\ o o l' O E t N T s O I D M TA \\ E I c K Z \\ Z T e RI IT l f T f OI WS IS E N N O E ES S I O O D 6 C N N 1 D A 2 N O A. 'E k e I HY r iD l N u ML l i EE Cl O N g I Ai O 0 T T O 0 F NI D A LN HS _N l 0 Z OO b N _AI 1 I _KI eB T C E T l F .R S WO l O N WS A 'G N l E 5 OC DE 1 .O A A .L C CA CN R R C U 'D F L E l Y N E b ~E S l , 00 t 0 l l _0} i . S. il i~' 6, s. 5 4 '2l ,1 T 4 2 .. s i i' o b. o 0 e s 2 2 2 l 't e l . $ S.E3E3!, i' !I n$ llllff l j, iil l

i NEDO-22139 5 '2 O F g ~ = I

• l

= C l 3 l l' wh l E.

• s E 3 h!k i

g E o = ~ t 5 E. i 8e g d u O

s j'

e '.s: o t'" c: E* u

u.

,.. -a e z -13 0 2 a u E 2 u u \\ T 4 'c s \\ m -o u O .u o \\ __ z -$s i \\ t' u \\ a \\ b e. e S in 1 _ _._ _1,= Nivwis cNas nnexvn g (xos33u + w\\ a y g-E. I g C. 3 I e 'ou I \\ vs = P- \\ \\ isnuvi \\ - NtVW15 CNEO WnntXVW 4 N 1 i i i N _. 2 _i i i ,= - O s.a a. .z. s s 's m .a t.A. l _SS3.w.a. _s 2-38

l NEDO-22139 i O " : M M ~ I 1000 m m M = T .r 5 M ut .j ac -> (') 6 - 2)/,o 2 <! uia o b a I~ 3 (1 - w ) Kf 2 %Q y 5'- G= E s - 1 uf [ \\ tu\\ w-K-@ f UI 2 g 2 f \\t%9 "~ ~ _ g-We+ i 10 __- I I I 3 l100 1000

L/d - 5

~ 1 l ~ l I I I I Ill ) i 1 I-1.0 0.1 031 ~ F R ACTION OF CR ACKED CIRCUMF ER ENCE (3/wl Figure 2-18. Compliance Change, Cracked Pipe 2-39

NEDO-22139 O s = = TENSILE .STR ESS ]L ]L 0 ]k ]k k 4k h Ak Ak ]k l I l l 't 't 'r'r'f'rY YY'T'T'T.'T 'T O h Ah h h A A A AAAAAA I I I l I l T 'T if 1rY IfY 17 17 t V ' COMPR ESSIV E .sTR ESS = 5 = Figure 2-19. Assumed Stress Distribution on the Crack Face O 2-40

i NED0-22139 3. LOST PARTS ANALYSIS ()

3.1 INTRODUCTION

Based on the structural analysis given in Section 2, it is expected that the Peach Bottom core spray sparger willnot break and result in loose pieces in However, an evaluation of the possible consequences of a poten-the reactor. tial loose piece is presented in this section. 3.2 LOOSE PIECE DESCRIPTION Since a piece has not been lost, it cannot be uniquely described. Three (1) a sec-different typets of loose pieces are postulated in Section 3.4.2: (2) an outlet nozzle; and (3) a small piece of the tion of sparger pipe; The entire sparger is fabricated of Type-304 stainless steel. sparger. 3.3 SAFETY CONCERN The following safety concerns are addressed in this safety analysis: Potential for corrosion or other chemicai reaction to reactor materials. 1. Potential for fuel bundle flow blockage and subsequent fuel damage. 2. Potential for interference with control rod operation. 3. 3.4 SAFETY EVALUATION The above safety concerns for the postulated loose pieces are addressed in The effect of these concerns on safe reactor operation is also this section. addressed. 3.4.1 General Description The core spray spargers are attached to the inside of the core shroud () For a piece of the sparger to reach and (Figure 3-1) in the upper plenum. 3-1

NED0-22139 /~T V potentially block the inlet of a fuel assembly, it would have to be carried out of the upper plenum and pass down into the lower plenum. To accomplish this, it would have to be carried by the f'uid flow in the upper plenum 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* turn and be carried upward to the fuel assembly inlet orifices. A part of the sparrer cannot reach the fuel assembly inlet orifices by falling down inside the core shroud this. For a part as the core support plate and the loaded core will prevent of the core spray sparger to reach a control rod, it must first traverse the upper plenum from the outer region of the shroud toward the center, which is unlikely, 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 reaction to any reactor material. l 3.4.2 Postulated Loose Pieces pJ 3.4.2.1 Sparger Pipe l The sparger pipe is 4-in. Schedule,40 pipe and is attached to the core shroud f l at six locations (T-box plus five brackets). The maximum span between supports In order to 38-1/2*, which corresponds to approximately 71 inches. is about generate a loose piece of pipe, two throughwall cracks would have to propagate f The weight of the largest pipe segment would be 360* around the sparger. l approximately 90 lb. (1) the top A pipe segment could come to rest in any of three locations:(2) the top surface surface of the top guide outboard of the fuel assemblies; the top surface of the fuel assembly handles; or (3) in an unlikely event, In all three of these locations, the flow velocity is of the core plate. low and insufficient to lif t a segment of the pipe. Therefore, it will remain at one of these locations (see Appendix C for flow velocity calculations). 3-2 i

NEDO-22139 { A 90-lb piece of pipe which falls from the core spray sparger will not harm the core plate, top guide or fuel assembly handles, since 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 operations. 3.4.2.2 Spray Nozzle Each spray nozzle consists of two 1-in elbows fabricated of Type-304 stain-In order to generate a loose less steel, which are welded to the sparger. nozzle,a throughwall crack would have to propagate 360* around the nozzle. A loose nozzle The weight of each nozzle assembly is approximately 1-3/4 lb. likely come to rest on the top surface of the core plate or on the would most top surface of the top guide. The flow velocities in these regions are [; insufficient to lif t the nozzle, thus, it will remain at one of the above mentioned locations. fit Since the nozzle cannot be lifted by the flow and since the nozzle cannot through the steam separator, it will not cause any flow blockage at the fuel The nozzle is too large to fit between two fuel assembly inlet orifices. channels; thus, it cannot cause any control rod interferences. 3.4.2.3 Small Pieces A small piece of the sparger could become loose if both longitudinal and cir-cumferential throughwall cracking occurred. A small piece could be lifted by the flow if it maintained an orientation with its maximum projected area Due to flow turbulence and nonsymmetry of the perpendicular to the flow. the part will tend to rotate so that the minimum projected area loose part, will be perpendicular to the flow. With this orientation, all parts with a 1ength of greater than approximately 0.4 in, will sink (Figure 3-7 of Refer-f'C} Thus, most pieces will not be carried by the flow toward the ence 3-1). 3-3

NEDO-22139 ( ) steam separator. However, in the unlikely event that a piece reaches the steam separator, it would have to pass through the steam separator turning vane (Figure 3-2). 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 inches long and it must be oriented with the long dimension in the vertical direction. The largest piece that can fit through the turning vane with its long dimen-sion in a horizontal plane is shown in Figure 3-3. It is very unlikely that the flow velocities would carry either of these maximum sized pieces through the turning vane. After passing through the At the turning vane, the fluid momentum is reduced as the water is removed. separator exit, the fluid is almost entirely steam. A typical water content is 1 weight percent. Thus, it is very unlikely that any piece could be car-ried 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 driven by jet pump flow to the bottom of the reactor pressure vessel where it would most likely remain. However, per Reference 3-1, a small piece could be carried by the flow up to the flow inlet orifices. The orifice sizes are 1.244, 1.469 and 2.211 inches. It is extremely unlikely for a piece larger than the 1.244-in. orifice and essentially impossible for a piece larger than the 2.211-in. orifice to be The outside diameter of the sparger is carried through the steam separator. 4.5 in., while the fuel inlet orifices are slightly recessed relative to the surf _~e of the control rod guide tubes (Figure 3-4), which have an outside diameter of 10-7/8 inches. Due to the different radii of curvature, flow l would be able to enter the fuel assemblies. Thus, unacceptable flow blockage as defined by Reference 3-1 wnuld require that more than one loose piece be carried to the same inlet orifice. This is based on the size of the piece (s) in a highly unlikely circumstance, have the potential of reaching the

that, vessel lower plenum. The probability of unacceptable flow blockage of any fuel orifice is judged to be insignificant.

Ix_ / 3-4

NEDO-22139 ()Theflowvelocitiesnearthespargerarelowerthanthoseabovethefuel assemblies. Thus, it is unlikely that a umall piece would be carried over the fuel assemblies. 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 3-5 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 is 0.75 inch. The length of the gap between the channel spacer and the channel fastener is 2.3 inches. Thus, any piece larger than The cor. trol rod 2.3 in, by 0.75 in, cannot cause control rod interference. thickness is 0.312 in. and the diameter of the control rod rollers is the control 0.520 inches. Thus, pieces smaller than 0.334 in, will fall past rod without causing any interference. A piece of precisely the right size Such could be in contact with the control rod and one or two fuel channals. The rods a piece might be detected during the normal control rod exercising. are inserted one notch and withdrawn one notch each day. It is also possible, ( } though unlikely, that a piece might wedge between two fuel channels above the If the control rod and thus not be detected by notual control rod operation. rod were to be inserted, the control rod mechanism has enough force to lift If the fue'. one fuel assembly with the reactor at normal operating pressure. 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 rods can be fully stuck out and the reactor can be safely shut down. Thus, unacceptable centrol rod interference will require multiple precisely-sized pieces interfering simultaneously with control rods that are in close prox-imity to each other. The probability of this is judged to be insignificant.

3.5 CONCLUSION

() The probability for unacceptable corrosion or other chemical reaction due to a loose piece is zero. The potential for unacceptable flow blockage of a 3-5

NEDO-22139 i l fuel assembly is essentially zero. The potential for unacceptable control rod l i interference is essentially zero, i I 3.6 REFERENCE 1 " Consequences of a Postulated Flow Blockage Incident in a Boiling Water i 3-1 ) Reactor", October 1977 (NEDO-10174, Rev.1), i e 1

O 4

i } ) i 4 4 i!O 3-6 }

NEDO-22139 0 st s.es g o. g a..e ' ac..r. / \\ / s. O., m .a ya.. / e ,/* ' m N 8 gT,am0e',A , h.'dby W 8'#b ' 7 Q [';eY ,-( +o!,,.L., {. ses eve 6st 2,. I 37 as ,,sa?Os k. i 8 i s. e -' - '" N, %,. $. gb, [,..........., i 3' 'f (=,tJg_g'geaegg...i -...=s., l CCat F4AYkset andeCTIOm t.T yg _m k _---Z M h(I,

b. v - - - M-

'f m aa rear asas n ^ l f 100 cuson l i i e 7 !k l. !,i, j. - i N co r m.u.. PUSL m ammaasig e W 'i

• • • • * '8

y,;

i I .,l / e w...eacuur. Sr.f 9 tab 4T j,e. .SC!aCVLAtuose A g A7 2OUTbit t .g j l 1 / 8 i i I P'" ':b N,,.....m _ _,,,, s t....I .w y a c Wirta06 age servts j 'N \\s l CDetect ROO Desv, mT0mauk4 414 8 awines,tva me to. Figure 3-1. Reactor Vessel 3-7

NEDO-22139 ., (. 0 ..>. v. At? 3 E Au E 9t LRNisc wa;gn r / / f' oh \\ Y h r n l

• ATEM LEVEL T

y I / " ' ' " * * " ' ~ " ' ) !3 O k j R ET.JamiNc testgm h, y l- '1 f TumNINC V ANE$ ItNLET NcZZ(g) p l e j l' t i h v STAN PtPE TO 00*N Coutm Amt A / A40 atcincygatieN . AMO atT PWWF SUCTIONS ), h O Figure 3-2. Steam Separator 3-8

I ~ NEDO-22139 O 0.78 n. 2.23 a. ' 18e.., ^ O 3.o32 1 l Figure 3-3, Largest Piece That Can Fit Through the Turning Vane with the Long Dimension in a Horizontal Plane 3-9

NEDO-22139 O ~ N_# b ] D \\ .bN( i (O ~ ^e C CatNisc sca esN?act ace BLaot Ten ~ ~ W ,-L f 'D( l )

• Coat GulDE Tust b

c,% x (c0*E plats A/f

;;n:,=c Oairac O

Figure 3-4. Orificed Fuel Support 3-10

NEDO-22139 -.:,::,+pt,.y,.._,~ s. i., , m mr=w -.w% y.*

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1.....? ?..

' =.. I >.1., t...,yi I ..-w~ ..a y . q:SMy, h.. '

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= :. w

u r.. <

% m).. :%,.. C-f,.a.p.. r.v1:}.. 5 i. ..1

.t 4p::..u,~W,g-y r;-T@w::.ff F f

1 1... W a u:: .g- .e-s. x.. ..'c u FUEL 1:- ~: s ..F, sss. '5 cf. ASSEMBLIES i L, 'j .. p%, s & CONTROL ROD MODULE l.' M '3 4 zg ,9 1 m.

• %{

g y;S ,r. +{ 8'Jb.q;.' s o % c.- e. a L 'k $~'5 N Ih b. '. h i '} k ll

1. TOP FUEL SutDE

~~I "E s N ((-4 'i 2.CH A NNEL ' [sh .si .,,,h;.. - . f.3 -?* i. k,.[$ ,I ;* *[.s'y ( .A

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'j jJ sn '.y. .s.r e.i

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. ; ' sk(.)i.' . s,,I, i.) II'. 4 EXPANSION /;p j j j ,f,,,',4 J SPRING l i r i. j,. g ..Q[, fi i., C *)l ,] i d r[e'I )J 1 a , '4. g. *. ' * * !.'. S. LOCKING TAS t i .'./(* y.7,, J

6. CHANNEL

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J = 9 'l,'-d.it, r.?. j '.,; - .;+ Q !.9-'i y g. 4,'. ;,.'.,$j f. '# Je (

7. CONTROL ROD gl

- ' ',fe,W.,,'-{a://p.?

• J<

4 j 5 s.PUEL RCD ' a d ' *7*+5 ,e 'J.f '*,.. b J. 45 ~.M 9 SPACER I,

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! '.75 f-I 3 W." ' 5';.i g-f +.6. 'I.* Q.g, jgN, 'n,., j i.Ys f! ASSEMBLY \\ ' ' 'Y 9.i 4 H 11.L O w E R k. iS.i.Y54}i : { 'I ,,i. ]. .\\ ";.ig T1E PLATE !s ,t i'T l\\lI;i R:.',.. I -?

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• y _.

Figure 3-5. Fuel Assemblies and Control Rod Module 3-11/3-12

NEDO-22139 1 4. LOSS-OF-COOLANT ACCIDENT ANALYSIS '41TH NONUNIFORM SPRAY IN ONE SPRAY SPARGER

4.1 INTRODUCTION

This section describes the methods used to evaluate the MAPLHGR requirements to meet 10CFR50 Appendix K for the Peach Bottom Reload 5, Cycle 6, assuming no The input, credit for core spray cooling from the cracked core spray sparger. to the apt-oved 10CFR50 Appendix K computer codes are discussed in Section 4.2; the general sensitivity of the loss-of-coolant accident analysis (LOCA) results to the spray cooling is discussed in Section 4.3; the results of the analysis are given in Section 4.4 and the conclusions are presented in Section 4.5. 4.2 INPUT TO THE LOCA ANALYSIS The approved versions of SAFE, REFLOOD, and CHASTE codes were used to evaluate the impact of a cracked core spray sparger in Peach, Bottom-2. The potential ef fects of cracks in one core spray sparger is to cause nonuniform If the second sparger is injecting flow spray diutribution from the sparger. (i.e., the other core spray system is operable), the postulated effect could only reduce the amount of spray flow to the hot fuel assembly by the contribu-This effect is conservatively modeled by setting the tion from one sparger. spray heat transfer coef ficients in the CHASTE heatup code to one-half of their This is the same assumption used in standard Appendix K Appendix K values. analysis to model a core spray system out of service (Reference 4-1). If one core spray system is rendered inoperable due to the assumed single failure per Appendix K, the remaining sparger may be assumed to be the one with The bounding ef fect (the assumed loss of all spray to the hot fuel cracks. assembly) can then be represented by setting the spray heat transfer coefficients in CHASTE to zero. in summary, the effect of cracks in one sparger is represented Therefore, conservatively in this calculation by setting the spray heat transfer coef ficients to zero or to one half their standard value, depending on the single failure analyzed. 4-1

.39 wllowing This representation is very conserv d :e as discussed i~ AU paragraphs. . ned with a steam Counter current flow limiting (CCFL) is the phenomenon 4.

ich, in this updraft limiting the downflow of water through a flow pa-case, is the fuel assembly. The steam updraft in the f*

- embly (due to rods subsequently) flasning during blowdown and to spray evaporation on the

er to an amount can, under certain conditions, limit the downflow of sp-

..;s CCFL is a smaller than the spray injection rate in the upper plenc

ecause subcoolfng function of the subcooling of the water in the upper plc can quench the steam updraft and cause the CCFL to "bren z en," eliminating the " holdup" of the coolant downflow.

Currently-approved Appendix K LOCA models assume saturated water CCFL condi-tions and conservatively ignore the inventory buildup of coolant in the upper Recent large-scale tests confirm that the CCFL " breakdown" can occur plenum. soon af ter spray initiation, causing downflow of the upper plenum inventory and rapid reflooding of the core. Following this, a residual pool of water remains in the upper plenum, ensuring uniform delivery of coolant to the individual fuel bundles. The present core reflood time fr om Appendix K models does not model CCFL breakdown or the residual pool in the upper plenum. The effects of saturated CCFL modeled in the REFLOOD model produce an overly conservative estimate of l If a crack, or cracks, forms in one sparger to the the core reflooding time. the flow rate through the spray nozzles is reduced, then more extent that the core periphery which will most likely cause localized l l injection will occur at j subcooling and CCFL breakdown. This would reduce the reflooding time for f Peach Bottom-2 up to 100 seconds from the value calculated with the standard Appendix K models resulting in PCTs up to 700*F lower. On the other hand, if no CCFL breakdown occurs, the upper plenum inventory builds up rapidly and ensures no reduction in coolant delivery from the core spray sparger system to the bundle and subsequently no degradation in cooling heat transfer. In addition to the above conservatisms, the 1973 ANS + 20% decay heat correla-l tion was used in the analysis per Appendix K. The technical community at this 4-2

NEDO-22139 1979 ANS decay heat correlation provides a time recognizes that the subsequent This decay heat e.orrela-more realistic basis for evaluating ECCS performance. D tion would further reduce calculated steaming rates and CCFL effects, as well as the core heatup rate, which would reduce the calculated PCT an additional 200* to 400*F. SENSITIVITY OF LOCA ANALYSIS TO NON UNIFORM SPRAY 4.3 For the Peach Bottom plant, there are no single failures for any break location (other than a core spray line break) that can render both core spray systems For core spray line break, there are always at least two low pres-inoperable. For medium and large sure ECCS pumps available, ensuring timely reflooding. break sizes (which depressurize relatively fast), the most limiting failures are those that result in the least number of emergency core cooling system (ECCS) pumps remaining operable. The two single-failure candidates that are potentially limiting for medium to large break sizes are: \\ Diesel Generator Failure - 1 core spray (LPCS) + 1 Low Pressure Coolant A. Injection (LPCI) + HPCI + the ADS operable LPCI Injection Valve Failure - 2 core spray + HPCI + the ADS operable B. Since the HPCI (High Pressure Coolant Injection) is steam turbine powered, is not a significant contributor to mitigating medium to large breaks. it Also, since the function of the ADS (Automatic Depressurization System) is l to depressurize the reactor as a backup to the HPCI, it contributes litt e toward mitigating medium and large break LOCAs. failure candidates A and B each results in a dependence on only Therefore, two ECCS. Per the Reload 5 analysis, failure candidate B (LPCI Injection valve failure) is limiting because of the conservative modeling of CCFL at the fuel assembly O) upper tie plates, which limits the downflow from the core spray systems and ( prolongs reactor reflooding. 4-3

NEDO-22139 These two single-failure candidates were re-examined for larger breaks assum-ing a cracked spray sparger as described in Section 4.2. The limiting single failure, break size and location does not change, since the calculated core uncovery and recovery times and the reactor depressurization rates do not change with the methods described in Section 4.2. For smaller break sizes, the limiting single failure is the high-pressure ECCS (HPCI), since the LOCA transient is a high pressure transient that is limited by the time required to either reflood the rear. tor with the high pressure system or the time to depressurize the reactor so that the low pressure systems become effective. Furthermore, the effects of CCFL in limiting coolant delivery to the core are not as large at higher reactor pressures. The small break LOCA transient is therefore insensitive to spray cooling and reflooding occurs very rapidly once any one or two of the six low pressure ECCS begin injecting coolant into the reactor vessel. Only medium and large break LOCAs are significantly af fected by core spray [} sparger cracking, and the effect is only significant with the conservative assurption of no CCFL breakdown in the peripheral bundles coupled with an assumed nonuniform spray distribution. 4.4 ANALYSIS RESULTS The most limiting fuel type and exposure combination for the limiting LOCA This is for per the Reload 5 analysis results is a calculated PCT of 1965*F. prepressurized 8x8R fuel at an exposure of 20,000 mwd /t and a MAPLHGR of 12.3 kW/ft. A reanalysis of this limiting case with the unrealistically conservative assump-tions discussed in Section 4.2 results in a calculated PCT of 2075*F. Therefore, a maximum increase in PCT of Il0*F bounds the ef fect of a cracked spray sparger for all fuel types and exposures. i A calculation of the maximum PCT for the limiting break with a single failure [} of a diesel generator using the cracked sparger assumptions of Section 4.2 results in a PCT of less than 1700*F. 4-4

NEDO-22139

4.5 CONCLUSION

S A conservative analysis of the effect of one cracked core spray sparger in the Peach Bottom-2 BWR results in a maximum increase in PCT of 110*F. Since the Reload 5 analysis shows a minimum margin af 235'F to the 10CFR50 Appendix K limit of 2200*F, the maximum increase in PCT of 110*F.can be accommodated with no change in MAPLHCR limit. Thus, with cracks in one core spray sparger and with the MAPLHGR limits unchanged, Peach Bottom-2 retains a minimum of 125'F margin to the 2200*F This PCT margin is still in excess of the PCT margin taken credit PCT limit. for in the generic study on the effect of increased fission gas at higher exposures (References 4-2 and 4-3).

4.6 REFERENCES

l SER, 0.D. Parr (NRC) to G. G. Sherwood (GE), " Review of General Electric 4-1 Topical Report NEDO-20566, Amendment 3." June 13,1978. R. E. Engel (GE) to T. A. Ippolito (NRC), " Extension of ECCS Performance 4-2 Limits," MFN-077-81, May 6,1981. j R. E. Engel (GE) to T. A. Ippolito (NRC), " Additional Informati,n 4-3 j Regarding Extension of ECCS Performance Limits, MFN-102-81, May 28,1981. !O 4-5/4-6 . ~.

J NEDO-22139 I () APPENDIX A i STRUCTURAL ANALYSIS 0F THE 1 PEACH BOT *)M 2 CORE SPRAY SPARGER Summary i Stress 9 (lb/in.2) 1. Sparger Pipe 853 Bending - Seismic 698 i (No Break) - Impingement l 901 Bending - Seismic 737 (Break) - Impingement l 2980 1 - Thermal Mismatch 1 Bending 2. Nozzle 5460 Normal (Weld) 5700 Shear (Weld) [) 1 3. Bracket (Lower) i 5140 Normal l 1502 Shear 3540 Normal (Weld) 633 Shear (weld) 4. Bracket (Middle) 9030 Normal 201 Shear 2233 Normal (Weld) 215 Shear (Weld) O A-1 l -. _ - _.., _ _... _. _ _ _ _ _ _ _ _ _ _ _., _.. _ _. _. _ _. -. _

NEDO-22139 A.1 DESIGN LOADS A.1.1 Impingement Loads (to deflect flow 90*) DL 4-m. sCMDULE 40 PPE y,pg,o Sc h F_,o V D g L Sc V p = 45.87 lb/ft @ 550"F D= h 12 h ]L V = 10.0 ft/sec* F_, 45.87 (10.0) (4.5/12) L 32.2 f=53.4lb/ft=4.45lb/in. A.1.2 Pressure / Flow Loads Maximum Flow = 8000 gpm** (Rated Flow = 6250 gpm) Q = 8000 gal / min x 60

• 7.4 gal 3

= 17.83 ft /sec

• Very conservative - more realistic value is %2 ft/sec.
  • See page B-5, Appendix B.

A-2

NEDO-22139 Maximum pressure in sparger arm e meas " 29 Psig @ 6068 gpm AP l 29(6 )? = 50.4 psig AP = Pressure load on spatz,er segment e A=fd =f(4.026)2 = 12.73 in.2 F = 6P A 50.4 (12.73) = 642_ lb l F = Maximum nozzle flow e The 1-in. VNC nozzle has the highest flow rate and will produce the maximum nozzle thrust. - 0.313 ) = 1.018 in.2 (min.) 2 + + W16 n 2 = 7 (1.181 Ay N = f (1.75 - 0.375 ) = 1.804 in.2 2 A 4 2 3 0 80*F = 62.2 lb/ft p g = 72 gpm @ 6068 gpm test flow 2 max =72[(6068/000j=95gpm l> a l-74 q max 4 1.t al ww = 13.2 lb/sec max " k8) W 14N 13.2(144) = 30 ft/sec @ nozzle exit max V = -= 62.2(1.018) A-3

F NEDO-22139 A Nozzle Thrust O .1.3 Y MMR Y l PsPE I l () -Z -x h o. Y ) [ i s n -g v = \\A-J- q Fy 1 W.VNC NOZZLE ^ O F' = AP A + # AP = 25 psi @ 6068 gpm test flow AP = 25 = 43.5 psi @ runout A=fd, where d = 1.181 in. (the minor dia. of 1-in. straight internal threads) A = f (1.181) 1.095 in. = = 28 ft/sec @ exit from header Y' = 6 095) = 43.5(1.095) + 62.2 8) 1 095), g Fy A-4

NED0-22139 A + gA V = 30 ft/sec (see Section A.1.2) p F = , 62.2(30)2(1.018) = 12.3 lb yz 32.2(144) A.l.4 Weight 4-in. Schedule 40 pipe W = 10.8 lb/ft pipe 5.5 lb/ft W water W = 10.8 + 5.5 = 16.3 lb/ft 6.3 = 1.36 lb/in. = l 12 A.l.5 Mismatch Due to Thermal Expansion o* 362.5* i s* \\ ,E. 38.4' 8 g4, 314.1' BRACKET T ( INLET TEE i I 5 DRACKET BRACKET ( #4ROUD CENTERLINE 38'" 38.4 CENTERLINE RC (C3 PIPE 27s.7* p SRACKET BRACKET l CENTERLINE l l CENTERLINE A-5

NEDO-22139 , 1. 5 = 108.75 in. R s 6_ _ 4.5 = 105.75 in. R = Shroud = 550*F CS Pipe = 198'F (See page B-2, Appendix B) AT = 352*F -6 AR = a R AT a = 9.6 x 10 in./in. *F for SST R = 216 = 108 in, at shroud-to-pipe interface 2 For 90* arc.

90. = 9.6 x 10-6 (108)(352) = 0.365 in.

AR For segmente assume.

90. (1 - c so) = 0.365 (1 - cose)

AR = AR c' is' ( INLET TEE (FIXED) -sa4* / uf \\ l I ( \\ l --M.s' B O i -- A A-6

NEDO-22139 (' - - ' - '2' O ^^15 53.4. = 0.365 (1 - cos 53.4') = 0.1474 in. AR 97,g. = 0.365 (1 - cos 91.8') = 0.3765 in. AR -38.4. = 0.365 (1 - cos 38.4') = 0.0790 in. AR -76.8. = 0.365 (1 - cos 76.8*) = 0.2817 in. AR / //// w l / l e M i } Assume the AR is resisted only by each bracket support in turn: O l AR = (20 - sin 20) - (cos 20 - 4 cose + 3) 4EIAR l g, R (20 - sin 20 - pcos 20 + 4 ucos 0 - 3p) E = 28 x 10 lb/in.2 R=R = 105.75 in. 0 I=h(4.5 - 4.026 ) = 7.23 in. p = 0.2 (coefficient of friction) 6 4(28 x 10 )(7.23)AR y, (105.75)3(20 - sin 20 - 0.2 cos 20 + 0.8 cos 0 - 0.6) O A-7 i ---v,,,n e----,-,- ---r,-w_--- ,, - - - -,, - - - +. -,. -. - - -. - -, - -, - - - - - - -.

=_ - - _, E i NEDO-22139 684.7 AR y, (20 - sin 20 - 0.2 cos 20 + 0.8 cose - 0.6) 4 684.7(0.0124) = 367 lb a -sin 30'-0.2cos30'+0.8cos15'-0.6). g15,,(2wx 15 18g i ) 684.7(0.1474) l -sin 106.8*-0.2cos106.8*+0.8cos53.4*-0.6) g 53*4*,(2wx 53 4 l 180 i I = 120 lb 1 i 684.7(0.3765) (2wx -sin 183.6*-0.2cos183.6*+0.8cos91.8*-0.6) y 1 9* 80 4 J ) = _91_ lb i i 684.7(0.0790) (2nx38' -sin 76.8*-0.2cos76.8*+0.8cos38.4*-0.6) g ( - 8.4 g0 = 155 lb 684.7(0.2817) (2wx76* -sin 153.6*-0.2cos153.6*+0.8cos76.8*-0.6) w = - 6.8 380 t = 9J5 lb 2 4 A.1.6 Flow-Induced Vibration - Natural Frequency _ 1 \\ 7 The vortex shedding frequency, fn, is given by: 4 i ) f D y = 0.21 V } 1 I A-8 1 l

4 NEDO-22139 V = velocity past the shroud wall = 10 ft/sec* 4.5 ft D = sparger pipe diameter = g 0.21(10)-= 5.6 Hz f v 4.5/12 General Electric design basis requires natural frequency: i ) f l n v ) Assume Calculate the natural frequency of the unsupported sparger segment. i the segment acts as a cantilever and has a uniform load, w (1bs per unit length): K / k V k fn" h A O 6= x - 3 52 n l I = 7.23 in.' g = 32.2 ft/see 6 E = 25.75 x 10 lb/in. x w x 105.75 = 28 in. (distance from crack to nearest support 15 L = 180 bracket) w = 1.36 lb/in. (Section A.1.4) 6 , 3.52 25.75 x 10 (7.23)(32.2)(12) = 167_ Hz f 1.36(28)' n 2n Ratio = 6 >3 = v O

• Very conservative - more realistic value is N2 ft/sec.

A-9 l---

NEDO-22139 Calculate the natural frequency of the sparger by examining the longest seg-(] ment between support brackets. Assume this section has a uniform load, e, and both ends are simply supported: Kn EIa g n 2n 4 Ifififif if If if1f if 1f if uL b K = 9.87 h n I = 7.23 in. 6 2 E = 25.75 x 10 lb/in.2 g = 32.2 ft/sec L= x n x 105.75 = 71 in, 180 w = 1.36 lb/in. OG D 9.87 25.75 x 10 (7.23)(32.2)(12) " fn" 2n 1.36(71)4 f Ratio =[=5.6 >3 v Calculate the natural frequency of the sparger by examining the longest seg-l This case is ment between support brackets ignoring an intermediete support. l the same as the above case except that L = 2 x 71 = 142 inches. l 72 f = 3 = 18 Hz 2

1. 3 Ratio =

= 5.6 V A ) v l A-10

NEDO-22139 Calculate the natural frequency of the sparger by considering the sparger arm j as a " free-free" beam (or floating ship). Assume the arm has a uniform load, w, and is free to rotate at the three support brackets as shown below: K f =2 S n 2x 4 wL 7 / / '= K = K,, = 61.7 n 4 p / / E = 25.75 x 10 lb/in. I 2 I = 7.23 in.4 g = 32.2 ft/sec I \\ t\\ w = 1.36 lb/in. 5 L= x w x 105.75 = 180 in. l 80 O 6 ,61.7 25.75 x 10 (7.23)(32.2)(12) = 70 Hz f 1.36(180) '1 >3 Ratio = = 5.6 1 v A.2 STRESSES DURING NORMAL OPERATION AND DURING CORE SPRAY INJECTION i A.2.1 Sparger Pipe l l f A.2.1.1 Impingement Load and Seismic i Impingement Only. = -4.45 lb/in. (upward) (Section A.l.1) w O A-ll

NEDO-22139 Seismic Only - Assume 3g (Very Conservative) w =w 3a w = 1.36 lb/in. (Section A.l.4) s = 1.36 - 3(1.36) = -2.72 lb/in. (upward) ws = 1.36 + 3(1.36) = 5.44 lb/in. (downward) ws Impingement + Seismic = -4.45 - 2.72 = -7.17 lb/in. (upward) w I = -4.45 + 5.44 = 0.99 lb/in. (downward) w l Assume No Break O For simplicity, assume continuous beam - three equal spans. j w i 1fIflfIflfIf lf II lfIfif IfIfifIfIf ifI I II AL AL Ak ll 4 t 7 C e C t = 0.4 R = 1.10 J R

• I*I D

R, = 0.40 we B C l 0.00 we 0.40 we 0.50 we SHEAROb) l M, N 7 l 0.80 we 0.50 we 0.40 wt 2 2 -4.10 s.d -4.10 wt /\\ k ) \\ J \\ O \\- / W+0.026 wt % / MOMENT Gn.4b) N / %~_s'/ %s.a/ 2 ,,2 +0.M we A-12

i NED0-22139 L= x n x 105.75 = 71 in. 18O M = 0.10(5.44)(71) = 2742 in.-Ib I = 7.23 in.' = 2.25 in. c= o= 2742(2 25) = 853 lb/in.2 (Seismic) o = o = 853I - = 698_lb/in.2 (Impingement) Assume Break and force P Assume two equals spans, uniformly loaded with end moment M3 3 at the third support. P3 w f/f1 f Ifif If1 17 If If if I .,\\ A,. 7-~~ 2+ 4 e e 3 R 3 "I 2 From the theorem of three moments. 2 3 1 2 "11 "22 ES N MA11+ T + Tj + 3 2 + 41 ~ I 2 I 41 1 1 2 2 1 2 M =0 1 =1 Il"I2 "1 * "2 y 1 2 A-13

..m NEDO-22139 u 4M L ME 3 2 3, ut I I 21 i 2 M M2" 8 - i is caused by the cantilevered section of pipe between the support M3 bracket and the break: } 2 wt3 3" 2 i Likewise, P is caused by the cantilevered section: 3 P3 " "L3 For Seismic. I, w = 5.44 lb/in. 1 = 28 in. t = 71 in. l 3 t P = 5.44(28) = 152 lb f 3 l l 5.44(28)2 g = 2132 in.-lb l 3 2 l l l , 5.44 71) _ 2132 = 2895 in.-lb (max) M2 4.5 o= I = 7.23 in. c= - 2.25 in. , 2895(2.25) = 901 lb/in. ,max 7.23 i!O 1 I l } A-14 i

. _. _ ~.. _.. _. NEDO-22139 I For Impingement

O l-4.45l=4.45lb/in.

w - T j P = 152 l= 125 lb 3 5 i M = 2132 = 1744 in.-lb 3

5. 4 i

= 2368 in.-lb M2 = 2895 ,44 o,,x = 2368 }25 = 737 lb/in.2 3 Determine reaction loads for seismic + impingement: l-w = 7.17 lb/in. l O s M + = 201 + 7.17( }* l - 1 = 441 lb 3 + "2 I R =P 2 3 7.17(71),38 5 = 201 lb Ry 2 - = 2 815 _ 2 1 = 577 lb (Mj = M ) = 7.17(71) + 2 R = wt +2 2 1 2 = 1219 lb Ry+R2+R3 checks w (2t + 1 ) = 1219 lb 3 A.2.1.2 Differential Pressure AP = 50.4 psi O R,= 4.5/2 = 2.25 in. A-15

NEDO-22139 O = 4.026/2 = 2.013 in. R g t = 0.237 notn. = 0.237 - 2[0.003 (corrosion allowance)] = 0.231 in. t,g e Hoop Stress 50.4(2 13) = 440 lb/in.2 = a= o Axial Stress 9 220 lb/in. = = o= 2t A.2.1.3 Mismatch Due to Thermal Expansion M = WR sine - WWR (1 - cos0) )u = WR [ sine - p(1 - coa 0)} / 4 t /. R w e mw Assume p = 0.2 (Coefficient of friction) R = 105.75 in. See Section A.1.5 for loads at each bracket: i A-16 ? . =

4 i, NED0-22139 r i i e 15' Bracket

15. = 360(105.75) [ sin 15* - 0.2(1 - cos 15*)]

M l = 9590 in.-lb (Maximum) i e 53.4* Bracket i M = 111(105.75) [ sin 53.4* - 0.2(1 - cos 53.4*)] 53.4 i = 8580 in.-lb j t e 91.8* Bracket 91.8. = 79(105.75) [ sin 91.8* - 0.2(1 - cos 91.8*)] i M I = 6630 in.-lb l 1 l e -38.4* Bracket 4 \\ -38.4. = 147(105.75) (sin 38.4 - 0.2(1 - cos 38.4*)] M = 8980 in.-lb e -76.8* Bracket -76.8. = 86(105.75) (sin 76.8* - 0.2(1 - cos 76.8*)] i M [ j = 7450 in.-lb i Mc 4.5 2.25 in. = = --- c= i o ex I i f I = 7.23 in.' , ax, 9590(2.25) = 2980 lb/in. i m 7.23 I A-17 i i _-.---c,----- .--i--.,.-..e-------~_,-,_ .. - ~,.,,, - v--,,-..

NEDG-22139 A.2.2 Nuzzles A.2.2.1 Nozzle Thrust Y Y Z' l y A - z x 1.N + + 1.bd r L l f \\ ( c* 0.11 VARIES - p i f ASSUME 15* , 3.29 4 15 O Weld properties: 1=y(1.76 - 1.52 ) = 0.209 in.' 0 4 K=h(1.76'-1.52)r 0.418 in.' 0 A={(1.76 - 1.52 ) = 0.618 in.2 2 l 1. 1* c= = 0.88 in. t = 0.12 in. r= = 0.88 in. 2 2 Loads: F = 60 lb f ' = 12.3 lb (See Section A.1.3) O y z P = 43.5 psi A-18

i NEDO-22139 1 The resulting loads at the weld are. F =F = 60 lb = F ' = 12.3 lb F, T = 3.29 F ' = 3,29(12.3) = 40.5 in.-lb torsion z M = 1.96 F ' = 1.96(12.3) = 24.1 in.-lb moment z The stresses are conservatively calculated as. "m 60 4.35(0.88) c Pr, 24.1(0.88), 0.618 4 o ,3 2t 0.209 2(0.12) y I A O - *1o1 + 97 + 160 1 erin 2 = 358 lb/in.2, 156 lb/in.2 oy Tc F, g xy K +"I a = 2 (thin wall cyl.) T = ,40.5(0.88) + 2 12.3 = 85 + 40 0.418 0.618 = 125 lb/in.2 Txy )O A-19

NEDO-22139 A.2.2.2 Differential Pressure Assume 360* break, nozzle loaded by bracket. 1 f' SECTION 1.2) } F = 842 La !r p t = L LJ 1.se JL - N j+ o.12 / l 2.88 l ! O l The resulting loads at the weld are. t F = F = 642 lb shen T = 2.68F = 2.68(642) = 1720 in.-lb orsion . 6F = 1.96(642) = 1260 in.- n M = moment The stresses are conservatively calculated as. + + + y"I 09 (0 2 o = -5,140 lb/in. , 5,460 lb/in. )O A-20 l

NEDO-22139 Tc F a = 2.0 T =- +a x 1720(0.88) + 2 642 = 3,620 + 2,080 0.418 0.618 T = 5,700_lb/in. A.2.3 Bracket (Lower) I O s (~ ~ s U l-p 7 7,! "2 1f l If RX / = l) + l H O gn _/ e II \\ h LOWER BRACKET L = (1 + cos 45*)(2.25) = 3.84 in. 1 = 3.26 in L'= + (1 - sin 45*)(2.25) = 2.29 in, 2 b = 0.38 in h = 0.25 in. ' O A-21

NEDO-22139 A.2.3.1 Seismic and Impingement R = 577 lb R =0 R, = 0 (Section A.2.1.1) (Conservatisms - Uses highest bracket load at weakest bracket and assumes seismic and impingement downward.) Maximum stresses in the fillet weld.. ' Avg " ~~ 0. 3 26) " i , 36 M, y, 3 6 (3.84)(577) = 3540 lb/in. , Bending 2 2 0.25(3.26)2 h1 h1 A.2.3.2 Mismatch Due to Thermal Expansion j i R = 367 lb R =pR = 0.2(367) = 73.4 lb z (Section A.1.5) i R =0 l Y Maximum shear stress in the fillet weld is. ( i 6 x+Rz+g g 2 ht 2 (b + h)(1 - h)h where M = t' R z 6 2.29(73.4) 6 (367 + 73.4) + T (0.38 + 0.25)(3.26 - 0.25)(0.25)

• " T 0.25(3.26)

O A-22 L

NEDO-22139 w = 382 + 251 = 633 lb/in. Maximum normal stress in weld is.. 2, (b + h), 36 M R , ax, _d x, ht(b + h) z y 2 2 m 2 ht h1 where M = t' Rx 0.25(3.26),3 6 (2.29)(367) , _d (367) 0.25(3.26)2 y max 2 N (3.84)2 + (0.38 + 0.25)2 ~ I 73.4 2 2 + 0.25(3.26)(0.38 + 0.25) = 318 + 1342 + 779 = 2439_lb/in.2 f a Maximum normal stress in the plate is. l r x xy z 2x x+ max A I I ~ zx xy A = 0.38(3.26) = 1.24 in.2 = 1.097 in.4 bt I = xy 12 2 = 1.63 in. C = xy 2 326g38)-=0.01491in.' 1,, = 48 = 1, O l A-23

_.m_ i, NEDO-22139 4 4 1 0.38 ! h -= 0.19 C = zx 2 i. q I 1.24 + 2.29(367)(1.63)- + 3.84(73.4)(0.19) = 300 + 1250 + 367 i 1.097 0.01491 o = max i s = 5140 lb/in.' o I max 1 1 Maximum shear stress in plate is. i z (31 + 1.8b) 367 + 73.4 L' R Rx+Rz bt 2 2 , 0.38(3.26) l g b , 2.29(73.4)(3 x 3.26 + 1.8 x 0.38) (3.26)2(0.38)2 = 356 + 1146 = g lb/in.2 I l T l l l b O A-24 _ _ _ _ _ _ _. _ _ _ _ _ _ _ _ _ _ -. _, _ _ _ _. _. _ _ _ _ _ - _ _. _. _ _ _ _ _ _ _ _ _ _ _ _. _. - _ _ _. _. _ -. _ ~ -

NED0-22139 A.2.4 Bracket (Middle) / sA E _n ~, T ~ v 's N J L / t' ,F 7 p II / S / y 1T y m ,S \\/ / / V N2 = T ( O x LJ J I w -4 L b = 0.38 in. L = 12.12 - 2(2.25) = 7.62 in. I h = 0.25 in. t'= + (1 - sin 30*)(2.25) = 4.94 in. 2 L = (1 + cos 30*)(2.25) = 4.20 in. (2.25 + 1.96) cos 15* + (1.50 + 1.18) sin 15* - 2.25 = 2.51 in. 1 = 7 A-25

NEDO-22139 i h = 2.25 - (2.25 + 1.96) sin 15' + [1.50 + 1.182 / cos 15' LF g l LF " 3'10 I"* i A.2.4.1 Pressure Load Only i F = 642 lb (Section A.1.2) Shear Stress (Neglect torsion - small) j .i = 238 lb/in. (Weld) "~ " Avg "

0. 5 62)

= 222 lb/in.2 (Bracket)

• Avg "

" 0.38 7.62) Stress Due to Bending (b + h)2 F 2 g max hf.(b + h) F 2 4 i t (3.18)2 + (0.38 + 0.25) 642 2 " 0.25(7.62)(0.38 + 0.25) = 2420 lb/in.2 (Weld) omax 0. o = c= = 0.19 in. 2 max 1 = p, = 762g38) = 0.03484 in.' M = (L - h) F p O A-26

NEDO-22139 (3.18 - 0.25)(642)(0.19) = 10.260 lb/in. (Bracket) O max 0.03484 A.2.4.2 Mismatch Due to Thermal Expansion Only Assume: R =R = 155 lb (Section A.1.5) x 1 2 R = 0.2(155) = 31 lb R e 1 2 Shear Stress 6(R

• 1 + R*2 6

62 " avg 2 h1 2 (0.25)(7.62) n,yg = 2_3_ lb/in. (Weld) 3 (R +R

• 1
• 2 62

" av g * - bt " 0.38(7.62) (Bracket) n,yg = 21 lb/in. Normal Stress R +R R +R , 5

• 1
• 2
• 1
• 2 2,(b + h)2 2L 2

ht ht(b + h) 2 g 2(4.20)2,(0.38 + 0.25)2 310 61 ,d_ 0.25(7.62) 0.25(7.62)(0.38 + 0.25) x 2 2 i 302 = 445 lb/in.2, -187 lb/in.2 (Weld) = 115 ' O I A-27 ,_..__...._.___._._,__,_.._..,m .. ~..

NEDO-22139 R +R (L - h) R -R c A b1 I V I = 0.0348 in.' c = 0.19 in. 310 . (4.20 - 0.25)(62)(0.19) = 110 + 1340 ,, 0.38(7.62) - 0.0348 o = 1450 lb/in. , -1230 lb/in. (Bracket) A.2.4.3 Combined Stresses During CS Injection Shear Stress u = 238 - 23 = 215 lb/in. (Weld) gyg = 222 - 21 = 201 lb/in.2 (Bracket) wAvg Normal Stress a = 2420 - 187 = 2233 lb/in.2 (Weld) o = 10,260 - 1230 = 9030 lb/in. (Bracket) A.3 REFERENCES i GE Drawing 731E779, " Core Spray Sparger." i GE Drawing 761E506, " Core Spray Sparger." j Roark, R. J., " Formulas for Stress and Strain, McGraw Hill, Fourth Edition, 1965. l t O A-28 i

NEDO-22139 + t Hopkins, R. B., "De-ign Analysis of Shafts and Beams," McGraw Hill. "Machinu, 1 Handbook," The Industrial Press, 16th Edftion, 1959. i Blevins, R. D., " Flow-Induced Vibration," Van Nostrand Reinhold Co., 1977. 1 j l t Shields, C. M., Wade, G. E., " Core Spray Distribution No. 17, 251 Standard i j Plant," NEDE-13006-4, December 1, 1970. h i t l ? i O I i i e !O + A-29/A-30 l:

NEr0-22139 O APPENDIX B SPARGER TEMPERATURE CALCULATIONS B.1 SPARGER TEMPERATURE Heat transfer coefficients for inadvertent spray injection are from pages B-3 and B-4. 4 D, y h = 5037 Btu /hr - ft2,.7 1 T. T, 2,.7 h, = 365 Btu /hr - ft "i K = 10 Btu /hr - ft

• F 3e (304 sst @ 200*F)

"o t = 0.237 in. D = 4.026 in. D,= 4.5 in. 1 O T = water in sparger = 80*F 1 T = water outside = 550*F o = n (4.026 1 = 1.054 ft2 (1 ft long section) A l i 2 4}5 1 = 1.178 ft (1 ft long section) A =n =n(4.026+4.5\\ 1 = 1.116 ft2 (1 ft long section) A 2(12) ) p I The thermal resistance, R, is: 1. 1 R 1 + 1 + t Ah AK oo p B-1

NEDO-22139 O Q= o-i 1 1 t Ah,Ah,AK ii oo p j ~ Ah o i film outside " O Ah 1 1 + 1 AT Ah Ah AK fg 9 p 1) ~ Ah o ATfilm inside " 1 1 t Ah +Ah +AK ii oo p

• ~Ah +Ah +AK p

1 + 1 0.237/12 O-

• " 1.054 (5037) 1.178 (365) + 1.116 (10) x = 0.000188 + 0.002326 + 0.001770 = 0.004284 0.002326 (550 - 80) = 255*F

= ATfilm outside 0.004284 Outside metal temperature = 550 - 255 = 295*F 0.000188 (550 - 80) = 21* AT = film inside 0.004284 Inside metal temperature = 80 + 21 = 101*F 285 + 101, p Average sparger (pipe) temperature = = 550 - 198 = 352*F 0 Bracket to pipe ( B-2 _.__--,._.,-__.______,,_.___._,.,,.--.~_,.-_,.,__,.,y,, ---,-,,,y_ ,,.m

NEDO-22139 In practice, the core spray pumping system cannot inject into the reactor until the pressure reaches 300 psia, where T,,g = 417'F. In this case, the

• "" # *~

AT ~ Bracket to Pipe tion bounds the inadvertent injection case. It also bounds the case of core spray operation during LOCA for the same reason. B.2 CONSERVATISMS

1. Bounding for reason described above.

2. Assumes steady-state conditions (Q

=Q =Q).

3. Neglects heat conduction from pipe.

4. Assumes runout flow.

B.3 REFERENCES

1. Kreith, " Principles of Heat Transfer", International,1969.

2. Welty, et.al., " Fundamentals of Momentum, Heat and Mass Transfer",

John Wiley, 1969. B.4 HEAT TRANSFER COEFFICIENTS B.4.1 Inside Sparger Arm (Near T-Box on Long Side) l Assume average film temperature = 90*F =f = 0.0884 ft D,= 4.026/12 ft fygg A p = 62.1 lb/ft3 @ 90*F i ' O film inside pfp, are ignorei

• AT and AT B-3

) NEDO-22139 2 v = 0.833 (10-5) gt /sec i l W = 7980 gpm = 1102 lb/sec Total W = 1102 $ = 298.5 lb/sec Am = 54.4 ft/sec S V= = 0.0 8 62.1) N = 0.023 R,

• P y

r R, = 54.4 6 (0.833(10-5 { = 2.19 x 10 f l R = l O l P = 5.20 0 90*F r j l l-(5.20)l/3 = 4,707 0 N = 0.023 2.19 x 10 D= ft =N K = 0.359 Btu /hr - ft

• F @ 90*F

'359) = 5037 Btu /hr - ft - "F h = y 2 2 B.4.2 Outside of Sparger Arm

1. Assume average water velocity is < f t/sec.

2. Assume average film temperature = 420*F.

O

3. Assume that heat transfer is like a cylinder in cross flow.

B-4 " ~ ~ ' " ~ e <e,v,n,_

NEDO-22139 y = 0.169 x 10 ft /sec @ 420*F ~ K = 0.375 Btu /ft - hr

• F P = 0.932 r

0.31 hD P C 0.35+0.56(R,) = ft D = g 2(4.5) 10 12(0.169x10~ ) = 4.438 :: R = 0.5 5 0.932 31 0 0.35 + 0.56 4.438 x 10 h, = 2 h,= 365 Btu /hr - ft - *F B.5 PUMP HEAD / RUNOUT Shutoff Head = 300 psia (Q = 0) = 6250 gpm @ 125 psia QRated P=P ~ SH where P = shutoff head SH 300 - C (6250)2 125 o O E-5

i '2 l NEDO-12239 1 1 / -61 2 l C = 300 - 125 = 4.48 \\10 / psi /gpm 2 6250 i I @ P = 14.7 psia (Runout) J r i P -P SH 300 - 147 m = 7980 gpm = 8000 gpm j< Q 4.48(10-6) Runout C 4 i 1 1 i t 7980 (62) = 1102 lb/sec W = Runout 60 (7.48) i 1 1 5 ) [ IO i o. f l s O B-6

NED0-22139 APPENDIX C FLOW VELOCITY CALCULATIONS This appendix describes the calculations for the flow velocities supporting statements in Section 3.4.2.1 of the text. C.1 FLOW VELOCITY IN BYPASS REGION Assumptions: The plant is operating at rated power (3293 MWt) and flow 102.5 x 1. 6 10 lb/hr. i 2. The flow in the bypass regions is homogeneous. 6 3. The bypass flow fraction is 12% (12.3 x 10 lb/hr). The water in the bypass regions is saturated. 4. This assumpcion is dis-5. There is no down flow in the bypass region. cussed later. There are two parallel flow paths in the bypass region--one is between the fuel fuel channels, and the other is between the core shroud and the outermost The flow areas for these paths are shown schematically in Fig-assemblies. The simple analysis that follows will give an estimate of the rela-ure C-1. tive flow velocity in the neighborhood of the spray sparger. fuel channels. The flow Path 1 is between the core shroud and the outermost area along path 1 changes from A, between the bottom and top of the active y immediately above the top guide: feel, to A at the top guide to A6 5 = 3720 in.2,A

• 9140 I"*

= 5261 in.2,A A 6 5 y O C-1

NEDO-22139 The flow area along path 2 changes Path 2 is between the fuel channels. above the fuel channels: from A t A at the top guide to A4 2 3 = 2028 in.2,A = 27504 in. A = 3918 in. ,A 4 2 3 From the geometry and the flow areas, K for path 1 is approximately 0.3 and K for path 2 is approximately 1.0: 2 (KWyy)/A A 5 22 3 (0.3)W (1.0)W y 2 3720 2028 2.5 W W 1 2 1 0.40 W =W y 2 6 1.40 W = 12.3 x 10 y 6 8.8 x 10 lb/hr W = y The velocity in the bypass region between the core spray sparger and the fuel assemblies is then: - 8.8 x 10 /[3600 x 45.8 x (3720/144)] = 2.1 ft/sec 6 V = W /0A y 5 The fluid in this region is primarily saturated liquid. The fluid velocity in the periphery of the core bypass region was conserva-In actuality, there probably is downflow in tively estimated at 2.1 ft/sec. The total pressure drop across the top guide is predominantly this region. In some portions near the top of the core bypass due to the elevation head. Because region, boiling may occur, reducing the elevation pressure drop. C-2

NEDO-22139 heat sources in the non-fueled peripheral regions of the core there are no

Thus, bypass, boiling would not be expected in the vicinity of the shroud.

some downflow or crossflow in the peripheral regions toward the central region would be anticipated to balance the density differences. C.2 FLOW VELOCITY AT TOP SURFACE OF CORE PLATE 0^

• ( Total Bypass 6

= x 0 lor nce W i Total Bypass i A = w/4 (D - Nd)2 D = inside diameter of shroud - 204 in. N = number of control rod guide tubes = 185 d = outside diameter of control rod guide tube = 10.875 in. p = density = 45.8 lb/ft Then 6 - 185(10.875)2)/(4 x 144)) V = (12.3 x 10 )/(3600 x 45.8 x n (204 V = 0.69 ft/sec l FLOW VELOCITY AT THE TOP OF THE FUEL ASSEMBLY HANDLES l C.3 6 6 6 W = 102.5 x 10 - 12.3 x 10 = 90.2 x 10 lb/hr To M A = na l n = number of fuel assemblies = 764 = 36 in.2 a = area associated with each fuel assembly = (6) O I C-3 i

= NED0-22139 1 The equivalent single phase velocity is: O V = (WTotal)/pA } Then f 6 V = (90.2 x 10 )/(3600 x 45.8 (764 x 36/144)) = 2.86 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 friction multiplier i 1 must be used: 2,,y,x(ff ) 8 f mass flow rate of steam x = quality = total mass flow rate 6 6 (13.4 x 10 )/(102.5 x 10 ) = 0.131 = l I 3 45.8 lb/ft o = g .i 3 2.35 lb/ft o = 8 l f n t' = 1 + 0.131(45. 8/ 2.35 - 1) = 3.42 f i lifting force on a section of core spray pipe per unit length is: The e, l 2 F'= C Ao 4 V /(2g) D f 1 l C-4 m

NEDO-22139 O where C = drag coefficient = 1.2 2 DA = area = (4.5 in. x 1(ft/ft))/12 (in./ft) = 0.375 ft fgg Then: F = 1.2 x 0.375 x 45.8 x 3.42 x (2.86)2 (2 x 32.2) = 9.0 lb/ft / O O c-5

NEDO-22139 i O A 7 6 O C 4 A + _ 3ml m O . P AT H 1 PATH 2 A +A+ c _2

• s Figure C-1.

Flow Paths O C-6 ._}}