ML20214R412

Forwards Response to NRC 870407 Request for Addl Info Re SER Open Item 16, Spent Fuel Pool Rack Design. W/12 Proprietary Oversize Drawings.Drawings Withheld

ML20214R412
Person / Time
Site:	Vogtle
Issue date:	05/22/1987
From:	Bailey J GEORGIA POWER CO., SOUTHERN COMPANY SERVICES, INC.
To:	NRC OFFICE OF ADMINISTRATION & RESOURCES MANAGEMENT (ARM)
Shared Package
ML20214R415	List: ML20214R412
References
GN-1369, NUDOCS 8706080124
Download: ML20214R412 (36)

Text

{{#Wiki_filter:_ _ - _ - _ _ _ _ _ - - _ 4 Georgiz Power Company Fbst Offica Box 282 g Wayncsboro, Georgia 30830 Telephone 404 554-9961 404 724 8114 Southern Company Services. Inc. Post Office Box 2625 Birmingham, Alabama 35202 reieonone 205 8" " Vogtle Project May 22, 1987 U.S. Nuclear Regulatory. Commission File: X7BC35 Atta: Document Control Desk Log: GN-1369 Washington, D.C. 20555 REF: Miller (NRC) to O'Reilly (GPC), dated April 7, 1987 NRC DOCKET NUMBER 50-424 OPERATING LICENSE NPF-68 V0GTLE ELECTRIC GENERATING PLANT-UNIT 1 SER OPEN ITEM 16: SPENT FUEL POOL RACK DESIGN Gentlemen: The referenced letter contained the NRC's request for additional information on Vogtle spent fuel pool rack seismic design. Attached are the responses to your information request. We anticipate that these responses will be sufficient to resolve this issue. Should your staff require additional information or clarification concerning these responses we request that a meeting be organized within two weeks of the receipt of this letter for the purpose of providing the necessary information in the most timely manner. Since rely. J. A. Bailey JAB /caa Enclosure xc: J. P. O'Reilly R. E. Conway P. D. Rice L. T. Gucwa l R. A. Thomas (w/o drawings) J. E. Joiner, Esquire B. W. Churchill. Esquire M. A. Miller (2) g G. Bockhold, Jr. 60 NRC Regional Administrator NRC Resident Inspector R. Goddard, Esquire j D. Feig I g R. W. McManus (w/o drawings) j Vogtle Project File 8706080124 870522 ~ ADO JAWMS To C..F@3 PDR ADOCK 05000424 E ppg

+ -e ENCLOSURE SER OPEN ITEM 16: SPENT FUEL POOL RACK DESIGN QUESTION 1: Provide the following regarding the spent fuel structure for review: 4 (a) Sketches and drawings that show the general layout and design of the spent fuel pool, details of liner plates and pads, and installation and arrangements of racks including the leak chase channels. (b) A description of how the dynamic interaction between the pool structure and the rack modules was considered, including the key assumptions used in assessing the interaction effects and the value of any dynamic amplification factors. Also include all assumptions made regarding the summation and phasing consideration of all rack module dynamic. loads. (c) Analysis results demonstrating the adequacy of the pool floor and liner under rack sliding and impact loads. (d) Identification of the critical regions of the pool structure from the standpoint of controlling loads or stresses resulting from the pool / rack interaction. List the loads or stresses as appropriate. Compare the loads and/or stresses to allowable values.

RESPONSE

(a) Enclosed is a sketch (figure 2a-1, page 17) that overlays the leak chase channela and the rack structure showing the support locations on the liner plate plan drawing. In addition, enclosed are the construction drawings showing the details of the leak chases and the liner plate (floor and wall), and the vendor drawing showing the installation of the racks in the pool. Drawing No. Description AX2D09J001 (annotated) Composite markup AX2D090J01, revision 13 Liner plate plan with embed details AX2D09CO20, revision 4 Leak chase details AX2D09C022, revision 5 Leak chase details AX2D09J013, revision 7 Liner plate details AX2D09J014, revision 4 Liner plate details AX2D09J015, revision 9 Liner plate details AX2009J003, revision 7 Liner plate elevations AX2D09J004, revision 6 Liner plate elevations X6AN10-45-2, 46-3, Spent fuel pool rack drawings and 47-2 with MFCRB-15,849 _i_

r_ ENCLOSURE (CONTINUED) (b) -The response of the spent fuel racks was determined by a nonlinear time-history analysis. The seismic time histories applied to the racks were synthesized from pool floor response spectra which account for the pool floor amplification. Applying the time histories to the nonlinear dynamic fuel rack model, the loads on the pool floor were determined. Since the dynamic model accounts for possible rack lift and then pad impact on the pool floor, the results of the nonlinear dynamic ' analysis provide the proper rack-to-pool interaction. Therefore, the use of additional dynamic amplification factors is not necessary. The maximum spent fuel rack floor loads derived for one rack module are for the condition of all racks fully loaded with fuel, responding in phase with the maximum coefficient of friction. The dynamic loads from two spent fuel rack modules are assumed to be acting in phase to provide the most severe combination of rack loads for spent fuel pool floor analysis. See the response to question 1 (c) and (d) for a description of how the loads were analyzed to evaluate their effects upon the pool floor. (c) Relative to the spent fuel pool floor, under seismic conditions, the and pool wall maximum displacement corresponding to the top of the rack (d) elevation is less than 0.3 in. The maximum displacement of the rack is described in Table 2c-2 (page 13). Since the minimum clearance of the rack from the pool wall is greater than the maximum rack displacement, plus wall displacement, there is no impact and, thus, no load imparted to the pool wall by the racks. The fuel pool liner and floor were evaluated for the rack sliding and impact loads calculated from the rack analysis. The spent fuel pool floor is part of the 6 ft. thick basemat of the fuel handling building. The controlling parameters are the local stresses in the pool liner and the concrete floor in the immediate vicinity of the rack support pads. The racks are placed in the pool such that the support pads are not over the leak chase channels and do not cause over stressing of the liner plate over the leak chase channels and the concrete in the vicinity of the leak chase channels. In addition, the distance of the support pad from the leak chase channel is maintained greater than the leak chase channel depth, i.e., the distance required to eliminate the transfer of stresses from the rack loads to the concrete floor through the leak chase channels. The maximum stresses in the liner plate and concrete floor under operating basis earthquake (OBE) and safe shutdown earthquake (SSE) conditions are as follows: - -

ENCLOSURE (CONTINUED) LINER PLATE Tensile Stress (in-plane stress) Loading Actual

• Allowable Dead load + OBE 13.42 ksi 15.0 ksi Dead load + SSE 15.58 kai 22.5 kai
• No credit is taken for friction between the liner plate and the concrete floor.

CONCRETE FLOOR Bearing Pressure Loading Actual Allowable Dead load + OBE 3.22 ksi* 4.76 kai Dead load + SSE 2.05 ksi 4.76 kai

• Stress is under factored OBE loading, i.e. 1.9 E.

ENCLOSURE (CONTINUED) QUESTION 2: The licensee's response of January 21, 1987, to the staff's November 28, 1986, request for additional information was incomplete. Please provide the following: (a) A set of fuel rack and fuel assembly drawings to verify the mathematical model. (b) Provide a detailed mathematical model of spent fuel rack modules for seismic analysis. Figures 1-1 and 1-2 of your January 21, 1987 response do not clearly show the coordinate axes, lines of symmetry, types of supports and restraints, application of forces and reactions, boundary conditions and other pertinent -information. In addition to this information, also provide the numerical values and bases of all analytical modeling/ parameters including - dimensions - weights or masses - spring constants - damping values of dashpots and dampers - gap element properties fluid coupling coefficients - coefficient for restriction for impacts - multi-rack to rack and multi-rack to pool wall impacts. (c) Calculations of the seismic responses and stress analysis of the spent fuel rack modules have not been provided as requested in question 5 of our November 28 letter. Specifically, we need (1) all supporting calculations for model parameters, (2) calculations supporting the assumption that the fuel racks are rigid, (3) detailed seismic analysis results, (4) thermal stress evaluations, and (5) various limiting single rack and multi-rack displacements due to seismic loads. (d) Regarding question 6 of our November 28 request, provide discussion of the mathematical model used and numerical values for physical properties of the fictitious springs and dampers in the gaps between the fuel and the cell. Also provide a comparison of the mathematical model with experimental results or other justification for the use of the model. (e) The allowable stresses in response to question 8 of the staff's November 28 request should be rechecked to comply with SRP Section 3.8.4, Appendix D (Rev. 1). The statement that the SSE seismic allowables are twice the OBE seismic allowables may not be true. Provide the basis of this statement and verify with the load combinations of the FSAR and the acceptance criteria of SRP Section 3.8.4, Appendix D. (f) In response to question 9 of the staff's November 28 request the licensee indicated that acceptance limits in FSAR Table 9.1.2-1 were based on the paper "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications." However, acceptance limits should be based on SRP Section 3.8.4, Appendix D (Revision 1) dated July 15, 1982. 7-ENCLOSURE (CONTINUED)

RESPONSE

(a) Enclosed figures 2a-1 and 2a-2 (pages 17 and 18) show the basic details of the VEGP fuel storage rack. Pertinent dimensions as used in the mathematical model are shown in figure 2a-2 which shows all the rack components and their interfaces with each other. Enclosed figure 2a-3 (page 19) shows the outline of the Westinghouse 17 x 17 fuel assembly. (b) Two mathematical models were developed for the fuel rack analysis: a structural model of the rack module and a non linear seismic model of an average single cell within a module. Enclosed figure 2b-1 (page 20) is a plot of the geometry of the structural model used in the analysis. The model is used to obtain effective structural properties of the single cell model and to determine stress gradients throughout the rack module. Due to the symmetry of the racks, only one quarter of each rack is modeled with the appropriate boundary conditions along the lines of symmetry to simulate the other quadrants of the rack. Enclosed figure 2b-3 (page 22) shows the one quarter of the rack actually modeled, as well as the lines of symmetry. The model shown in figure 2b-1 is evaluated for several different loading conditions. The loading conditions evaluated include an X-direction shock (response spectrum analysis), a Y-direction shock (response spectrum analysis), and a Z-direction load representing the deadweight condition. o For the X-direction shock loading condition, the following boundary conditions are used in the model: - Along the line of symmetry parallel to the X-axis UY = 0 RX = 0 RZ = 0 (UX, U and Y, and UZ are displacement degrees of freedom. RX, RY, and RZ are rotational degrees of freedom.) - Along the line of symmetry parallel to the Y-axis UY = 0 UZ = 0 RX = 0 9 ENCLOSURE (CONTINUED) - At the interface of the six support pads (1/4 model) to the pool floor UX = 0 RY = 0 UZ = 0 o For the Y-direction shock loading condition, the following boundary conditions are used in the model: - Along the line of symmetry parallel to the X-axis UX = 0 UZ = 0 RY = 0 - Along the line of symmetry parallel to the Y-axis UX = 0 RY = 0 RZ = 0 - At the interface of the support pads to pool floor UY = 0 UZ = 0 RX = 0 Enclosed figure 2b-2 is a plot of the geometry of the single cell non linear seismic model. This model is used to determine the seismic displacements and loads. The boundary conditions used in the model consist of constraining all degrees of freedom for those nodes at the pool floor and on the pool wall. The constrained nodes are at the end of the dynamic friction elements representing the support pads, and also at the end of the mass matrix elements representing the hydrodynamic mass between the cell and the pool wall. The sliding and lift off of the support pads (and impact with the floor following lif t off) is accounted for by 'the behavior of the dynamic friction elements and the relative displacemeuts of the nodes of each element. (See our Response to Question 4 for further detailed discussion.) !

ENCLOSURE (CONTINUED) The loading on the model is produced by imposing the acceleration time histories for the pool at the nodes on the pool floor and on the pool wall. The cells of the two models (figures 2b-1 and 2b-2, pages 20 and 21) are 168.25 in. long. For the quarter rack structural model shown in Figure 2b-1, the cell centerline-to-centerline spacing is 10.6 in. The pad span in the model shown in figure 2b-2 is 116.6 in., which is the actual span between corner support pads on a 12 x 12 rack. These dimensions are consistent with the information supplied as a part of the fuel rack drawings provided with our Response to Question 2(a). The weight of the fuel assembly used in the analysis is 1647 lbs. The weight of the rack structure used in the analysis is 235 lbs. per storage location. The pertinent spring constants used in the analysis are as follows: 5 lb/in. - Support pad vertical spring stiffness = 4.0 x 10 - Fuel grid impact spring stiffness = 7.2 x 103 lb/in. 4 lb/in. - Fuel top nozzle impact spring stiffness = 2.0 x 10 The pad stiffness used in the model shown in Figure 2b-2 is derived from results obtained from the rack structure model shown in figure 2b-1. Using load and displacement results, the combined stiffness of the pads and the base structure is extracted. The fuel impact stiffnesses are determined by combining in series the local cell wall stiffness and the fuel grid or nozzle impact stiffness. The fuel assembly impact stif fnesses have been determined through testing by the fuel supplier. The sampling values used in the analysis for the impact elements (fuel impact elements or support pad elements) are 4 percent of critical damping for the support pad impact elements and the fuel top nozzle impact element and 25 percent of critical damping for the fuel grid impact elements. The fuel damping values have been determined through testing by the fuel supplier. The fuel-to-cell gap elements are composed of impact stiffness, damping, and concentric gap. The impact stiffness and the damping values used in the analysis are those values discussed in the preceding paragraphs. The gap sizes used in the analysis 0.191 in. (gap per side) for the fuel grids and 0.198 in. (gap per side) for the fuel top nozzle. These values are determined from the inside dimension of the rack cell and the outside dimension of the fuel grid or top nozzle. The hydrodynamic coupling between the fuel and the cell, and between the rack and the pool wall is represented in the models by mass matrix elements. For the fuel-to-cell coupling, the value of the hydrodynamic mass term is 795 lbm. For the rack-to-pool wall coupling, the hydrodynamic mass terms are 2978 lbm (per cell, north-south direction) and 2182 lbm (per cell, east-west direction). ENCLOSURE (CONTINUED) The hydrodynamic mass between the fuel assembly and the cell walls is based upon the fuel rod array size and the cell inside dimensions using the technique of potential flow and kinetic energy. The hydrodynamic mass is calculated by equating the kinetic energy of the hydrodynamic mass with the kinetic energy of the fluid flowing around the fuel rods. The concept of kinetic energy of the hydrodynamic mass is discussed in a paper by D. F. DeSanto ("Added Mass and Hydrodynamic Damping of Perforated Plates Vibrating in Water," ASME Journal of Pressure Vessel Technology, May 1981). The hydrodynamic mass between the rack cells and the pool wall was calculated by evaluating the effects of the gap between the rack modules and the pool wall using the method outlined in the paper by R. J. Fritz ("The Effect of Liquids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, February 1972). Coefficients of restitution for impacts are not explicitly a part of the input values for the analytical model. The appropriate physical response is accounted for in the impact stiffness and the damping values used in the analysis. The spacing of the racks is such that impact between racks or between rack and pool wall does not occur. The analysis confirms that the distance between racks and between the rack and the pool wall is greater than the combination of rack deflection, sliding, and thermal growth. (See our Response to Question 2(c) for additional details.) (c) In the response to Questions 2(a) and 2(b), specific information has been provided describing the rack structure and the fuel assembly, as well as a discussion of pertinent input values used in the analysis and the basis for those input values. The detailed calculations performed to obtain the input values are on file at the Westinghouse site in Pensacola, and if desired, may be reviewed at your convenience. The fuel racks are modeled, as shown in figures 2b-1 and 2b-2, primarily by the use of beam elements. Each team element has the stiffness appropriate for the rack component it represents (cell, base plate, support pad, etc.). For example, each rack cell is modeled as a beam with area and inertia values of a single cell and with many degrees of freedom along its length. This method of modeling represents the rack structure not as rigid, but rather as a flexible structure which responds to the loading conditions, as appropriate. As mentioned in our Response to Question 1 of the NRC letter dated November 28, 1986, one instance where an assumption of rigid behavior is applied is in regard to the cell-to-base plate connection. Physically, the cell is welded to a stiff grid, which in turn is welded to the base plate. Rotationally, the connection has some flexibility, and this is included in the model. However, laterally and vertically the connection is very stiff, and it is appropriate to assume a rigid connection (stiffness value of 1.0 x 108 lb/in.). However, this local rigid connection does not cause the rack structure itself to respond as a rigid structure. ENCLOSURE (CONTINUED) Results of the seismic analysis include loads and rack displacements. Table 2c-1 (page 12) details some of the load information resulting from the seismic analysis. Rack displacement results from the analysis are listed in Table 2c-2 (page 13). In addition, figures 4b-1 through 4b-4 (pages 28 through 31) show time-history plots of displacements, as discussed in our Response to Question 4(b). Also provided is a sample-calculation showing how the margin of safety of a structural member is typically determined. The sample calculation is shown in Table 2c-3 (page 14), with the geometry of the component evaluated being shown in figure 2c-1 (page 23). Thermal stress calculations are performed for two loading scenarios. The first way that thermal stresses are produced in the rack structure is by the loading of one storage cell with hot fuel, while the surrounding storage locations remain empty. The temperature gradient of adjacent cells produces axial stresses in the cells and shear loads in the welds attaching the cells to the rest of the rack structure (the grids). The maximum thermal stress produced in the cell or welds is 4000 psi which, when combined with the seismic and deadweight stress, is below the allowable. Thermal stresses may also be produced in the rack by assuming that: the racks are installed in the pool at one temperature; the temperature of the racks then changes to a higher temperature; but the temperature of the pool floor does not change as much as the temperature of the racks changes. If the support pads are restrained by the maximum coefficient of friction ( p= 0.8), the relative thermal growth of the rack produces stresses, especially in the support pad structure. The calculated thermal stresses in the support pad structure are small in comparison to the seismic stresses, and the allowables are readily satisfied. Rack displacement results from the analysis are listed in Table 2c-2. The reported displacements are appropriate for each rack module. Since the rack displacements listed in Table 2c-2 (page 13) are considerably less than 1.0 in., it is apparent that rack-to-rack or rack-to-pool wall impact does not occur, because the distance between racks is approximately 4.0 in. and the distances between the racks and the pool wall is even greater. (d) The properties of the springs and dampers in the gaps between the fuel and the cell are detailed in our Response to Question 2(b). Please refer to that response for the requested data. The mathematical models used in the analysis are similar and employ the same methodology as that used on other spent fuel rack evaluations performed in recent years. These models and methodology have been reviewed by the NRC and found to be acceptable on fuel rack applications for the McGuire, Turkey Point, Peach Bottom, and Palisades nuclear plant sites. -9_

ENCLOSURE (CONTINUED) As previously discussed in our Response to Question 2 of the NRC's November 28, 1986, letter, the models used in the analysis have been formulated using the WECAN computer code which has also been reviewed and approved by the NRC. Additional discussion of the WECAN code and the application of its capabilities to fuel rack nonlinear dynamic analysis is available in the following documents: Shah, V. N., Gilmore, C. B., " Dynamic Analysis of a Structure with Coulomb Friction," ASME Paper No. 82-PVP-18, presented at the 1982 ASME Pressure Vessel Piping Conference, Orlando, Florida, June 1982. Gilmore, C. B., " Seismic Analysis of Freestanding Fuel Racks," ASME Paper Number 82-PVP-17, presented at the 1982 ASME Pressure Vessel Piping Conference, Orlando, Florida, June 1982. (e) The applicable edition of the ASME Code for the Vogtle site is the 1980 and Edition with Addenda through Summer of 1982. Per this edition of (f) the Code, Paragraph NF-3231.1(c) specifies Paragraph F-1370 of Appendix F for the Level D limits. Paragraph F-1370 specifies the Level D limits to be a factor times the Level A and B limits. The specified factor is defined as follows: Factor = lesser of 1.2 ( Sb[ ) or 0.7 ( Sb! ) Ft Ft Ft = allowable tensile stress =.6 Sy S = yield strength = 27,500 psi at 150*F for SA-240, Type 304 y S = ultimate tensile strength = 73,000 psi at 150*F for u SA-240, Type 304 Factor = lesser of [ 1.2 ( SZ )] or [ 0.7 (73000.6(27500) ) ) .6Sy Factor = 2.0 Therefore, it is concluded that the Level D (SSE) allowables are twice the Level A and B (OBE) allowables. Although the load combinations and acceptance criteria presented in Table 9.1.2-1 of the VEGP FSAR were extracted from the paper "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," the information contained in Table 9.1.2-1 is also consistent with the requirements of the SRP, Section 3.8.4, Appendix D. As shown in Table 2e-1 (page 16), the only difference between Table 9.1.2-1 of the FSAR and Table 1 of Appendix D is that the Table 9.1.2-1 load combinations list separately those load combinations with and without thermal loading, while Appendix D load combinations do not. Since the ASME Code distinguishes between primary and secondary stresses l J

ENCLOSURE (CONTINUED) (with thermal stresses being secondary stresses), the application of the Appendix D acceptance limits implies separate load combinations with and without thermal loading. Therefore, af ter application of the acceptance limits, Table 1 and Table 9.1.2-1 are in agreement. The end result is that Table 9.1.2-1 satisfies the requirements of Appendix D, as well as the requirements of the OT position paper. Therefore, the allowable stresses presented in our Response to Question 8 of the NRC's November 28 request are consistent with the Appendix D requirements, and the reported results remain valid. -t1-

ENCLOSURE (CONTINUED) TABLE 2c-1 GEORGIA POWER SPENT FUEL RACKS SUPPORT PAD LOADS

• Maximum Single Total Pad Load Rack Load (ib)

(1b) Vertical load: DW + SSE 48,600 530,000 Vertical load: DW + OBE 43,000 510,000 North-south horizontal load: SSE 37,400 301,000 East-west horizontal load: SSE 11,000 271,000 North-south horizontal load: OBE 32,100 258,000 East-west horizontal load: OBE 10,000 247,100 i

• All loads are for the maximum loading condition of 0.8 coefficient of friction.

ENCLOSURE (CONTINUED) TABLE 2c-2 GEORGIA POWER SPENT FUEL RACKS DISPIACEMENT RESULTS DISPIACEMENT (in.) Maximum support pad lift-off (p = 0.8) 0.053 Maximum support pad sliding (p = 0.2) 0.036 Maximum top of rack lateral deflection (p = 0.8) 0.100 Rack thermal growth (normal condition) 0.113 Rack thermal growth (accident condition) 0.184 4 <e. 2 ENCLOSURE (CONTINUED) l TABLE 2C-3 (Sheet 1 of 2) Sample Margin of Safety Calculation: Interior Bottom Grid Member (see Figure 2C-1 for section geometry) The following equations and parameters apply for interior bottom grid members. Box Beam Height = 3,500 in. H = Box Beam Width = 1,270 in. W1 = Box Beam Wall Thickness = 0.109 in. j T1 = Base Plate Width = 6,600 in. W2 = Base Plate Thickness = 0.500 in. T2 = I ABOX = Area of Box Beam = H*W1 - (H-2*T1)*(W1-2*T1) = 0.992 IN**2 } APLT = Area of Base Plate = W2*T2 = 3.300 IN**2 Total Area of Interior Bottom Grid Member = ABOX + APLT = 4.292 IN**2 i A = ZBAR = Centroidal Coordinate of Bottom Grid Member in Local Element Coordinate System (2BAR = 0.0 at Bottom Surface Bottom Plate) = (APLT*T2/2 + ABOX*(T2 + H/2) )/A = 0.712'in. Moment of Inertia About Local Element Y-Axis IYY = (U2*T2**3)/12 + APLT*(ZBAR-T2/2)**2 + = l ( (Wi*H**3) - (Wi-2*T1)*(H-2*T1)**3 )/12 I + AB0X*(T2+fi/2-ZBAR)**2 = 4.559 IN**4 Distance from Local Element V-Axis to Extreme Fiber l CZ = T2 + H - ZBAR = 3.288 In. = Moment of Inertia About Local Element Z-Arts IZZ = (T2*W2**3)/12 + ((H*Ul**3) - (H-2*T1)*(W1-2*T1)**3)/12 = 123.258 IN**4 = Distance from Local Element jZ-AXISTO Extreme Fiber = W1/2 = 0.635 In. CY = Area Within Mean Perimeter of Box Beam = (W1-TI)*(H-TI) = 3.937 IN**2 AT = SIGMAX1 = Axial Stress Due to FX = FX/A TAUY1 = Shear Stress Due to FY = FY/A I TAUZ1 = Shear Stress Due to FZ = FZ/A TAUZ2 = Torsional Stress Due to MX = MX/(2*T*AT) SIGMAX2 = Axial Stress Due to MY = MY*CZ/IYY SIGMAX3 = Axial Stress Due to MZ = MZ*CY/IZZ SIGMA = Combined Axial Stress = SIGMAX1 + SIGMAX2 + SIGMAX j TAU = Combined Shear Stress = (TAUYl**2 + (IAUZ1 + TAUZ2)**2)**.5 SIGMABR = Combined Seismic + Deadweight Axial Stress = (SIGMA E-W**2 + SIGHA N-S**2)**.5 + SIGMA DW i TAUBAR = Combined Seismic + Deadweight Shear Stress = (TAU EW**2 + TAU NS**2)**.5 + TAU DW j Allowable Stresses: SIGMALL = Allowable Axial + Bending Stress = 1.000*0.6*SY TAUALL = Allowable Shear Stress = 1.000*0.4*SY Margins of Safety: SFTY1 = Axial + Bending Stress Margin = (SIGMALL/SIGMABR) - 1 SFIY2 = Shear Stress Margin = (IAUALL/TAUBAR) - 1,

O ENCLOSURE (CONTINUED) TABLE 2C-3 (Sheet 2 of 2) Maximum Grid Member Loads North-South East-West Deadweight Fx (ib) 149" 2163 225 Fy (1b) 8/ 1325 11 Fz (1b) 2338 131 603 Mr (in.-lb) 2129 76 542 My (in.-lb) 9334 95 2380 Mz (in.-lb) 3818 589 148 Maximum Grid Member Stresses North-South East-West Deadweight SIGMAX 1 348 504 53 TAUY 1 21 309 3 TAUZ 1 545 31 140 TAUZ 2 2481 90 631 SIGMAX 2 6731 69 1717 SIGMAX 3 198 31 7 SIGMA 7277 604 1777 TAU 3026 331 771 Total axial stress = SIGMABR = [(7277)2 + (604)2]1/2 + 1777 = 9079 Axial stress allowable =.6 Sy - 16,500 Axial stress margin of safety = 16500 - 1 = 0.82 9079 Total shear stress = TAUBAR = [(3026)2 + (331)2]1/2 + 771 = 3815 Shear stress allowable =.4 Sy = 11,000 Shear stress margin of safety = 11000 - 1 = 1.88 3815 r, ENCLOSURE (CONTINUED) TABLE 2e-1 COMPARISON OF FSAR AND APPENDIX D LOADS AND LOAD COMBINATIONS FROM FSAR TABLE 9.1.2-1 Load Combination Acceptance Limit D+L Normal limits of NE 3231.la D+L+Pt Normal limits of NE 3231.la D+L+E Normal limits of NE 3231.la 0 D + L + To Lesser of 2S or S stress y u range (see note 3) D + L + To + E Lesser of 2S or S stress range (see note 3)u y D + L + Ta + E Lesser of 2Sy or S stress range (see note 3)u D + L + To + Pf Lesser of 2Sy or Su stress range (see note 3) D + L + Ta + E' Faulted condition limits of NF 3231.1c (see note 4) D + L + Fd The functional capability of the fuel racks shall be demonstrated TABLE FROM APPENDIX D Load Combination Acceptance Limit D+L Level A service limits D + L + To D + L + To + E D + L + Ta + E Level B service limits D + L + To + Pf D + L + Ta + El Level D service limits D+L+Fd The functional capability of the fuel racks should be demonstrated - -

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2, ~g w.- - - n. :. - x - m. _ >- x = <%g_ r ~ q _~ ~ C CELL TO CELL -~ Z " ~ CONNECTIONS ?= e = - e s ~ m x s x HYDRODYNAMIC m e MASS N_ m r ELEMENTS h-N - M _V 168.25 CELL s ,_s. g N - -n ~ 2 s s-w~ ~ = ~ ] = i ~ = m s s = = s. m l \\ y= - m ~ =,. ~ 2 h O b.',- RACK GRID ~, C - ~ - %~%,% ~ o l / DM'd h._ %4 $.lk," ~ Y RACK SUPPORT PADS i X FIGURE 2b-1 STRUCTURAL MODEL OF RACK MODULE '

l FUEL ASSEMBLY. TOP N0ZZLE j IMPACT ELEMENT i ) CELL TO POOL WALL HYDRODYNAMIC MASS u CELL n FUEL ASSEMBLY k {Af 9-0 [ FUEL ASSEMBLY GRID ROTARY SPRINGS I 4 l ,f ,4 4 FUEL ASSEMBLY GRID + T IMPACT ELEMENT "= \\ FUEL RACK GRID / N* (CELL TO CELL CONNECTI )N) ROTARY SPRING

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/w T si x Y 7 G Nf l x b i o tutt MLK PUtL MLK SUPPORT PAD SUPPORT PAD ~ --,,,,,,, ----,. 9.,.,,:-,,,,,,,,,,,,,,,,,,,,,,,,,,,1 1 6. 6 l ..,,,,,,,,,,,,,,,,_,,,,,, 2, r FIGURE tb-2 NONLINEAR SINGLE CELL MODEL

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y p 10.6 e e o a r e = s 6 = JC/ DC/ DC/ JCn JC JC' s s s /s m 10.s ' JC" OC" QC JC" JC "QC _L -s /s /s n -s m /s n -s /s /s -s N/ \\/ \\/ \\/ \\/ \\/ \\/ N/ N/ N/ N/ N/ /s k /N /s /s /s /s /s k /s T* Xv o' ov r v v v v v Jc/ JC s /s m m n m m v v v v \\/ v v y v v v JCm JC m /s n /s m /s m n /s y\\ g / v v v v \\/ v v v v ^ ^ - ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ LINE OF e \\/ N/ \\/ N/ N/ \\/ \\/ \\/ \\/ N/ N/ \\/ SYMMETRY, 'T 'N /N 'N 'N 'N 'N 'N 'N 'N 'N \\/ \\/ \\/ \\/ \\/ \\/ \\/ NJ \\/ \\/ \\/ %/ /s /s /s Ms /s /s /s /s /s /s g SUPPORT PAD ' m][" ][I '\\' V ^ )Cm )C-m n /s /s m n s m /s (TYPI C AL) S v v v v v v v gj v v DC O o JC m -s m n x n /s s /s v ig g g v v v v v v i v m /s m m /s m m n /s m m n \\/ \\/ \\/ \\/ \\/ \\/ \\/ \\/ N/ \\/ \\/ \\/ /s /s /s /s /s /s

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3r ac 3c. 3c v v v v v v / QUARTER OF RACK MODELED Y h -+- X LINE OF SYMMETRY FIGURE 2b-3 V0GTLE RACK: QUARTER MODELED AND SYMMETRY LINES l i -

Point A g1+ rur / / / / - T1 / Z j h H / / / f ',g y= y q .\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\\ n 7 I l W2 = = 3 50 in. H = 1.27 In. W1 = 0.1091n. T1 = 6.60 in. W2 = 0 50 in. T2 = FIGURE 2c-1 l INTERIOR BOTTOM GRID MEMBER..

ENCLOSURE (CONTINUED) QUESTION 3: It is not clear how the support legs of the racks are constructed. Please provide the following information with respect to the design, construction and installation of the support legs: 1. Sketches showing the upper and lower parts of the support legs. 2. Discussion of rack installation procedures when an adequate clearance to adjust the supports may not be available.

RESPONSE

(a) Figure 3a-1 (page 25) shows the two separate sections of the rack leveling pad assembly (" support legs") and how they fit together. Each leveling pad assembly is threaded into the bottom of a rack support block (figure 2a-2, page 18). The threaded engagement of the leveling pad assembly and the support block provides for up and down leveling adjustment of the rack. The spherical bottom of the leveling screw, interfaced with the spherical dish of the leveling pad, provides the screw with a swivel capability should the pool be out of level at the support pad location. (b) Clearance on the outside of the rack is not required to adjust the support feet. After the rack has been positioned in the pool, a long leveling adjustment tool can be inserted from above the rack into the cell opening at the support location to be adjusted. The lower end of the tool engages a machined slot in the top of the threaded leveling screw. The leveling tool functions exactly as a screwdriver, turning the leveling screw down or up to achieve rack levelness. i l )

L ~ N / INTO SUPPORT BLOCK THREADED FOR INSERTION I SCREW i SAME SPHERICAL RAD. a PAD I l.50 5.00 0 = 3* l { l FORM TO CAPTURE ITEM-01 5 i l l l l i i l l \\ FIGURE 3a-1 RACK LEVELING PAD ASSEMBLY

ENCLOSURE (CONTINUED) QUESTION 4: The following requests pertain to rack to rack interactions based on the bounding parameters that (i) use a coefficient of friction of 0.2 and 0.8, (ii) use a center to center spacing of 10.5 in, for rack arrays, and (iii) simulate the rocking behavior of racks, if multi-racks are used. (a) Provide a comparison of the impact forces resulting from an analysis of the single rack and multi-rack models. In the event that the results indicate that the multi-rack model impact force exceeds the single rack impact force, demonstrate that the integrity of the rack is maintained during the design earthquake. (b) In order to better understand the dynamic behavior of the simulated multi-rack model during the design earthquake, provide time-history data in tabular or graphical form. The data should include, as a minimum, the translation of each rack as a function of time and the impact forces on the wall as a function of time. (c) Provide a discussion of the fluid coupling analysis including the governing principles, general assumptions, controlling equations and derivations. Address the influence on the impact forces of the modeled racks when the adjacent rows of racks (1) move in phase, (2) move out of phase, and (3) are fixed. (d) For the simplified single and multi-rack models, compare the impact forces on the fuel assemblies. In the event that the results indicate that the impact forces on the fuel assemblies using the multi-rack model exceed those of the single rack model, demonstrate the integrity of the fuel assemblies for the design earthquake.

RESPONSE

(a) As previously discussed in our Responses to Questions 2(b) and 2(c), the racks are placed in the pool with sufficient spacing between racks and between the racks and the pool wall so that rack-to-rack or rack-to-pool wall impact does not occur. Therefore, the rack-to-rack and rack-to-pool wall impact forces are zero. The impact forces that occur within an individual rack have been determined in the analysis. These impact loads are the forces due to fuel-to-cell impact and the forces due to the support pad-to-pool floor impact following any pad lift-off. The maximum impact loads have been included in the structural evaluation of the rack, which has demonstrated j that the structural integrity of the rack is maintained through ) satisfaction of the ASME Code stress limits. As a point of information, it should be noted that the cell center-to-center spacing within a rack module is 10.6 in., rather than 10.5 in. j l

ENCLOSURE (CONTINUED) (b) As discussed in our Response to Question 2(c), the racks do not impact the pool wall. Therefore, the impact forces on the wall are zero. Also 3 included in our Response to Question 2(c) is a summary of the maximum rack displacements. (See Table 2c-2, page 13.) As an example of the rack response data reviewed to determine the maximum rack displacements, Figures 4b-1 through 4b-4 (pages 28 through 31) are presented for your info rmation. These plots depict the rack response during selected 5-second intervals of the time-history loading. Similar data exists for the remainder of the 20-second interval of the time-history loading. The first two plots are for the maximum loading and rocking with a 0.8 coefficient of friction, while the third and fourth plots are for maximum sliding with a 0.2 coefficient of friction. (c) The hydrodynamic mass between the rack cells and the pool wall was calculated by evaluating the effects of the gap between the rack modules and the pool wall using the method outlined in the paper by R. J. Fritz ("The Effect of Liquids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, February 1972). The close proximity of adjacent racks, as well as the size of the racks relative to the gap between racks, is such that extremely large hydrodynamic masses are produced if the racks attempt to respond out of phase. It is this large hydrodynamic mass which causes the racks to respond in phase. For purposes of calculating hydrodynamic mass, the seismic analysis treats the racks as if they are hydrodynamically coupled (move in phase). Once the individual rack displacements are obtained on a basis of in phase movement, the behavior of adjacent racks is assumed to be out of phase so that the possibility of rack to rack impact can be evaluated in a conservative manner. Using the individual rack displacements obtained from analysis (sliding, rocking, structural, and thermal displacement) and assuming out-of-phase movement by adjacent racks, it has been demonstrated that rack-to-rack and rack-to pool wall impact does not occur for the VEGP racks. (See our Response to Question 2(c) for more detailed results.) (d) Since there is no rack-to-rack impact, the loads on the fuel are obtained by evaluating the forces due to the interaction between fuel and cell within a single rack. It is the results of this analysis that are used to demonstrate that the integrity of the fuel is maintained for the seismic conditions. The analysis provides the impact loads at the fuel assembly grid and top nozzle locations. The maximum load on the fuel assembly at any time during the time-history loading was found to be 3.0 g (where: g = the acceleration due to gravity). Since the fuel supplier has qualified the fuel for loads much greater then 3.0 g, the loads obtained in the analysis are acceptable, and the fuel integrity is maintained. 1

• • IIII

~d. REG 10l ~ ~ ~ G i 1 RACK Mu+.8 UK DISP A1 ;ENENT Or TOP Or enCK l L 0.0750 l l 9.0500 t I l 9.8250 i 1 1 i l!Li !l i lla. h y;UI f.4h 1L a L A' l' Rj0i f= l f

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'Q in 1 ) I t.0250 1 i s. esse S 3 j e.s7se 8.9065 .g g = = e a = = = 9 9 9 9 8 9 9 9 e s n n p m e itME I FIGURE kb-1 TOP OF RACK LATERAL DEFLECTION: y = 0.8 (0-5 sec.) -

\\ r I' 1 e.R5'S - r 4L REGloi <G l i Vact ru :.e i 8.eSec VERTICAL l ilSPLACLMENT OF LEFT 'A0 PLOT NUMBER 4 n e.edee l l 3 e.esee e e.s2ee e.elea E 1 I I 4 l 3,g e YNMhi hh1/\\i khd/ / . s won __._ _ . e. aas., M E R M M. 2 E R o e e s ~ i H 9 = 9 9 = 9 a 9 9 m o n, e m e m .._ _.. _T.l4 l FIGURE kb-2 LEFT SUPPORT PAD VERTlCAL DISPLACEMENT: p = 0.8 (0-5 sec.) 31 -

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L d 9 i ~ u I i 1 I( l ~ P 9 s 8.0158 A a 5 g i W 8.8181 cmm i i i i i !Iiii! 5 5 d 4 / / 5 TIT FIGURE kb-4 RIGHT SUPPORT PAD SLIDING DISPLACEMENT: y = 0.2 (5-10 sec.) l l l l,,

J. MCLOSURE (CONTINUED) QUESTION 5: Provide the parameters and constants used in the analysis for impact loading due to the drop of a fuel assembly. Also, provide a summary of ductility ratios utilized to absorb kinetic energy in the tensile, flexural, compressive and shearing modes. Provide typical calculations indicating the input constants, equations used, and the results of the impact analysis.

RESPONSE

For fuel drop accident conditions, the double contingency principle stated in ANSI N16.1-1975 is applied. This states that one is not required to assume two unlikely, independent, concurrent events to ensure protection against a criticality accident. Thus, for accident conditions, the presence of soluble boron in the storage pool water can be assumed as a realistic initial condition since not assuming its presence would be a second unlikely event. The presence of approximately 2000 ppm boron in the pool water will decrease reactivity by about 30 percent a K. In perspective, this is more negative reactivity than is present in the poison plates (25-percent A K), therefore, Keff for the rack would be less than 0.95, even if the poison plates were removed by a drop accident. In fact, with 2000 ppm boron in the pool water, there is no deformation that could be achieved by the drop of a fuel assembly, thus causing the criticality acceptance criteria to be exceeded. Therefore, the requirement for the drop accident conditions that the fuel be maintained in a subcritical condition has been satisfied, independent of the racks deformation (or aven if the racks were removed altogether) due to the presence of the boron in the water. Consequently, a detailed structural evaluation of the rack for the drop accident conditions need not be performed. In addition to performing the evaluation to ensure that the fuel remains suberitical for the drop accident conditions, an evaluation has been performed to demonstrate that the pool liner will not be perforated for the condition of a dropped fuel assembly passing through an empty cell. For this evaluation, it is assumed that when the fuel hits on the base plate at the bottom of the cell, the base plate welds or a local region of the base plate fails and allows the fuel assembly to travel to the pool floor. The strain energy absorbed by the failed base plate or the welds is neglected. The energy due to the falling fuel assembly is assumed to be absorbed by crushing the base plate and the fuel assembly lower-end fitting. The fuel drop energy is the product of the total dropped weight (2300 lb) times the drop neight (210.7 in.). By equating the strain energy to the fuel drop energy, the value of strain and, in turn, stress and impact force are determined. Finally, the stress on the pool liner is calculated and is found to be less than the ASME Code allowable, thus ensuring that the liner is not perforated. l

ENCLOSURE (CONTINUED) QUESTION 6: Provide considerations regarding the potential impact on the functionality of fuel rack modules due to bowing and localized deformations of fuel assemblies and fuel rack cells. Provide discussions of all possible effects on racks due to earthquakes such as twisting, bouncing, rocking and their corresponding stresses versus allowable stress values.

RESPONSE

The rack cell and the fuel assembly are modeled as flexible beams in the analysis. As such, both structures are free to deflect, as dictated by the seismic loading and the beam stiffness properties. As the fuel moves from side to side inside the cell and frequently impacts the cell wall, the fuel assemblies are in an everchanging state of deflection and bowing. Therefore, whether the fuel is assumed to be straight or bowed at the start of the time-history loading is not a significant parameter, since the flexible fuel takes on a deflected shape early in the time history. In the analysis, the response of the rack structure to the appropriate time-history loading has been evaluated. The rack is free to respond by rocking, bounding, or twisting, as the loading dictates. The results obtained from the analysis (displacements, loads, and stresses) have been used to qualify the structural integrity of the racks by demonstrating that the ASME Code allowables are satisfied. A summary of the stresses and the appropriate Code allowables has previously been presented as a part of our Response to Question 8 of the NRC's November 28, 1986, letter. 1160V

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