BVY-99-115, Forwards non-proprietary & Proprietary Responses to 990714 RAI Re Civil & Mechanical Engineering Considerations for Proposed Change to TS to Increase Spent Fuel Storage Capacity from 2,870 to 3,355.Proprietary Encls Withheld

Forwards non-proprietary & Proprietary Responses to 990714 RAI Re Civil & Mechanical Engineering Considerations for Proposed Change to TS to Increase Spent Fuel Storage Capacity from 2,870 to 3,355.Proprietary Encls Withheld

ML20212D497
Person / Time
Site:	Vermont Yankee File:NorthStar Vermont Yankee icon.png
Issue date:	09/16/1999
From:	Leach D VERMONT YANKEE NUCLEAR POWER CORP.
To:	NRC OFFICE OF INFORMATION RESOURCES MANAGEMENT (IRM)
Shared Package
ML20138E156	List: BVY-99-115, Forwards non-proprietary & Proprietary Responses to 990714 RAI Re Civil & Mechanical Engineering Considerations for Proposed Change to TS to Increase Spent Fuel Storage Capacity from 2,870 to 3,355.Proprietary Encls Withheld
References
BVY-99-115, NUDOCS 9909230139
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Text

,

1 l

VERMONT YANKEE NUCLEAR POWER; CORPORATION 185 Old Ferry Road, Brattleboro, VT 05301 7002 (802) 257-5271 September 16,1999 BVY 99-115 U.S. Nuclear Regulatory Commission ATTN: Document Control Desk Washington,DC 20555

References:

(a) Letter, VYNPC to USNRC, " Technical Specification Proposed Change No. 207, Spent Fuel Pool Storage Capacity Expansion," BVY 98-130, dated September 4, 1998.

(b) Letter, USNRC to VYNPC," Vermont Yankee Nuclear Power Station, Request for Additional Informatien Regarding Spent Fuel Pool Storage Capacity Expansion (TAC W. MA 3490)," NVY 99-68, dated July 14,1999.

Subject:

Verraont Yankee Nuclear Power Station Lieunse No. DPR-28 (Docket No. 50-271)

Hvsponse to Request for AdditionalInformatio: Regarding .

Spent Fuel Pool Storare Capacity Expansion In Reference (a), Vermont Yankee proposed a change to the Technical Specifications to increase the spent fuel storage capacity from 2,870 to 3,355 fuel assemblies. During the review of the proposed change, the Staff requested via Reference (b) the submittal of additional information with regard to Civil and Mechanical Engineering considerations. Attachment A, which includes related Attachments I A through IK provided our response to Staff Question number 1. It is noted that Attachments l A through IE contain proprietary information and it is requested that these attachments be withheld from public disclosure per 10CFR2.790(a)(4). Attachment 2 provides the response to Staff question number 2. Attachment 3 contains Holtec International's proprietary information affidavit for the proprietary information contained within Attachments l A through IE.

We trust that the enclosed information will enable you to complete your review of Reference (a).

If you have any questions on this transmittal, please contact Mr. Thomas B. Silko at (802) 258-4146.

Sincerely,

/

VERMONT YANKEE NU LEAR POWER CORPORATION c:9

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NOT.. r"~'Ac.rrae [ 3c,7j},,ggSEV MY L..;h. ..s.: C.g, r;J Af fJ 25,2@ 4 a vrumost nsur Necuan rowen couronu nis Docket No. 50-271 BVY 99-115 Attachment 1 Vermont Yankee Nuclear Power Station Response to Request for Additional Information Regarding Spent Fuel Pool Storage Capacity Expansion , (Civil and Mechanical Engineering Considerations) Response to Question 1 I . . . . i REQUEST FOR ADDITIONAL INFORMATION ON TiiE SPENT FUEL POOL STORAGE RACK MODIFICATION AT VEIB10NT YANKEE NUCLEAR STATION Item 1 l You indicated that the calculated seismic loading stresses in a fully-loaded rack will not exceed that of SRP Section 3.8.4 which was used as a guide. With respect to your stress calculations using the DYNARACK computer code presented in Chapter 6 of the submittal, provide the following: a) Explain how the simple stick model used in the dynamic analyses can represent accurately and realistically the actual highly-complicated nonlinear hydrodynamic fluid- . rack structure interactions and behavior of the fuel assemblies and the box-type rack l structures. Discuss whether or not a finite element (FE) model with 3-D plate, beam and fluid elements together with appropriate constitutive relationships would be a more realistic, accurate approach to analyze the fluid-structure interactions in contrast to the l stick model.  ! b) Provide the results of any prototype or experimental study that verifies the correct or adequate simulation of the fluid coupling utilized in the numerical analyses for the fuel , assemblies, racks and walls. If there is no such experimental study available, provide justification that the current level of the DYNARACK code verification is adequate for l engineering application without further experimental verification work. c) Provide the stiffness calculations for the rack used in single rack analysis for the three (North-South, East-West and Vertical) directions. d) Provide the physical dimensions of the racks, gaps between the racks, and the gaps between the racks and the walls. e) Demonstrate that the artificial seismic time histories used in the analyses satisfy the power spectral density (PSD) requi.rement of Standard Review Plan (SRP) Section 3.7.1. Response 1a As explained in Sections 6.2 and 6.5 of Holtec International Report HI-981932, the Whole Pool Multi-Rack (WPMR) model used to predict the dynamic behavior of the storage racks contains elements specifically designed to represent the attributes necessary to simulate rack motions during earthquakes. These elements include non-linar springs to develop the interaction between racks, between racks and walls, and between fuel assemblies and rack internal cell walls. Linear springs having the necessary characteristics to capture the lowest natural frequencies of the ensemble cf fuel cells acting as an elastic beam-like structure in extension / compression, two-plane bending, and twisting are used to simulate rack structural elastic action. Hydrodynamic effects within these interstitial spaces are accounted for using Fritz's classical method which relates the fluid kinetic energy in the annulus due to relative 1 \ projects \980414\ license \RAI motion to an equivalent hydrodynunic mass. In the next section, we present a historical overview of the fluid coupling effect as applied to the modeling of spent fuel racks in a seismic environment. The phenomenon of fluid coupling between rectangular planform structures was sparsely investigated until the 1980s. Fritz's classical paper (ca.1972) was used in the earliest version of ) DYNARACK to model rack-to-surrounding fluid effects in the so-called single rack 3-D simulation. Enrico Fermi Unit 2 (ca. 'i980) and Quad Cities Units 1 and 2 (ca.1982) were licensed using the Fritz fluid coupling terms embedded in DYNARACK. The Fermi 2 and Quad Cities 1 and 2 submittals were thefirst rerack applications wherein a rack module was analyzed using the 3-D time-history technique. The adeption of a nonlinear time-history approach helped quantify the motion of a rack under a 3-D earthquake event and as a byproduct, also served to demonstrate that solutions using the Response Soectrum Method (which, by definition, presumes a linear structure) can be woefully non-conservative. Practically all rerack licensing submittals from 1980 on utilized the 3-D time-history method. While the nonlinear 3-D time-history method was a huge improvement over the Response Spectrum (by definition, linear) approach, it nevertheless was limited inasmuch as only one rack could be modeled in any simulation. The analyst had to assume the behavior of the adjacent racks. Models, which postulated a priori the behavior of the contiguous racks in the vicinity of the subject rack (rack being analyzed), were developed and deployed in safety analyses. Two most commonly used models w;re the so-called " opposed phase" model and the "in-phase" model, the former used almost exclusiely to predict inter-rack impacts until 1985. Holtec Position Paper WS-115 (proprietary), included herein as A, provides a summary description of these early single rack 3-D models. The inadequacy of the single rack models (albeit nonlinear) to predict the response of a grouping of submerged racks arrayed in close proximity became an object of prolonged intervenors' contention in the reracking of PG&E's Diablo Canyon units in 1986-87. Holtec, with active participation from the USNRC, developed a 2-D multi-rack model for the Diablo Canyon racks; this model helped answer intervention issues, permitting PG&E to rerack. USNRC experts testified in support of the veracity of the 2-D multi-rack dynamic models at the ASLB hearings in Pismo Beach, California in June 1987. The Diablo Canyon intervention spurred Holtec to develop what later came to be known as the 3-D Whole Pool Multi-Rack (WPMR) analysis. A key ingedient in the WPMR analysis is quantification of the hydrodynamic coupling effect that couples the motion of every rack with every other rack in the pool. In 1987, Dr. Burton P.ml (Professor Emeritus, University of Pennsylvania) developed a fluid mechanics fonnulation using Kelvin's recirculation theorem that provided the fluid coupling matrix (2n x 2n for a pool containing n racks). As an example, refer to Figure RAI 1.1, where an array of N (N = 16) two-dimensional bodies (each with two degrees of freedom) is illustrated. The dynamic equilibrium equation for the i th mass in the x-direction can be written as: 2 \ projects \980414\ license \RAI p i [m; + Ma b, + f My2, + Ny y, = Q (t)

-i in the above equation, mii s the mass of body i (i = 1,2...N), and L is the x-direction acceleration vector of body i. My and Ny denote the " virtual" mass effects of body j on body i in the two directions of motion. The second d

rivative of y with respect to time represents the accelection in the y-direction.  !

)- The terms My are functions of the shape and size of the bodies (and the container boundary) and,  ; most important, the size of the inter-body gaps. My are analytically derived coefficients. Q, represents the so-called Generalized force that may be an amalgam of all externally applied loads on the mass i in the x-direction. The above equilibrium equation for mass i in x-direction translational motions can be written for all degrees of freedom and for all masses. The resulting second order matrix differential equation contains a' fully populated mass matrix (in contrast, dynamic equations without multi-body fluid coupling will have only diagonal non-zero terms). The above exposition explains the inclusion of fluid coupling in a multi-body fluid coupled problem using a simplified planar motion case. This explanation provides the building blocks to explain the more complicated formulation needed to simulate freestanding racks. Dr. Paul's - formulation is documented in a series of four reports written for PG&E in 1987. References [4] through [7], included herein as Attachments IB, IC, ID, and IE, contains the formulation l I information. The Paul multi-body fluid coupling theory conservatively assumes the flow of water to be irrotational (inviscid) and assumes that no energy losses (due to form drag, turbulence, etc.) occur. The USNRC personnel reviewed this formulation in the course of their audit of the Diablo Can3 >n rerack (ca.1987) and subsequently testified in the ASLB hearings on this matter, as  ! stated above. While the ASLB, USNRC, and Commission consultants (Brookhaven National Laboratory and Franklin Research Center) all endorsed the Paul multi-body coupling model as an appropriate I and conservative construct, the theory was still just a theory. Recognizing this perceptual weakness,-Holtec and' Northeast Utilities undertook an experimental program in 1988 to  ! benchmark the theory. The experiment consisted of subjecting a scale model of racks (from one to four at one time in the tank) to a two-dimensional excitation on a shake table at a QA qualified laboratory in Waltham, Massachusetts. . The Paul multi-body coupling formulation, coded in QA validated preprocessors to DYNARACK, was compared against the test data (over 100 separate tests were run). The results, documented in Holtec Report HI-88243, were previously provided to the Commission. The experimental benchmark work validated Paul's fluid mechanics model and showed that the ' theoretical model (which neglects viscosity effects) is consistently bounded by the test data. This experimentally verified multi-body ~ fluid . coupling is the central' underpinning of the L DYNARACK WPMR solution that has been employed in every license application since ! Chinshan (1989). The DYNARACK 3-D WPMR solution has been found to predict much greater rack' displacements and rotations than the previously used 3-D single rack results [8] (see - Attachment IF, included herein). 3 \ projects \980414\ license \RAI p L l l Ir, general, the advance from linearized analyses (response spectrum) in the late 1970s to the single rack 3-D analyses until the mid-1980s and, finally, to the 3-D WPMR analysis in the past eleven years has, at each technology evolution stage, led to some increase in the computed rack response. The stresses and displacements computed by the DYNARACK 3-D WPMR analysis for the Vermont Yankee racks, in other words, may be larger (and therefore more conservative) than the dock'eted work on similar instances from 15 years ago. The conservatism's built into the . WPMR, solution arises from several simplifying assumptions explicitly intended to establish an upper bound on the results, namely:

i. In contrast to the single rack 3-D models, the fluid forces on every rack in the pool consist of the aggregate of fluid coupling effects from all other racks located in the pool. No assumptions on the motion of racks need be made a priori; the motion of each rack in the pool is a result of the analysis, ii. The fluid coupling' terms are premised on classical fluid mechanics; they are not derived = from empirical reasoning. Further, fluid drag and viscosity effects, collectively referred to' as " fluid damping", are neglected. In short, while the transfer of fluid kinetic energ, wthe racks helps accentuate their motion, there is no subtraction of energy through damping or other means.

iii.' In the Vermont Yankee rack simulations, the dynamic model for the fuel-assemblies in a rack assumes that all fuel assemblies within a rack move in unison. Work in quantifying the effect of discordant rattling of fuel assemblies within a rack in other licensing applications by Holtec has shown that the " unified motion" assumption exaggerates the rack response by 25% to 60%, depending on the rack geometry details and earthquake harmonics. iv. The rack-to-rack and rack-to-wall gaps are taken as the initial nominal values. During the earthquake, these gaps will in fact change through the time-history duration. Strictly speaking, the fluid coupling matrix should be recomputed at each time-step with the concomitant gap distribution. The inversion of the mass matrix at each time-step (there are over four million time-steps in a typical WPMR run) would, even today, mandate use of a supercomputer. Fortunately, neglect of this so-called nonlinear fluid coupling effect is a conservative assumption. This fact is rigorously proven in a peer reviewed paper by Drs. Soler and Singh entitled "Dynenic Coupling in a Closely Spaced Two-Body System Vibrating in a Liquid Medium: The Case of Fuel Racks", published in 1982. The only docket where recourse to the nonlinear fluid coupling was deemed essential was Vogtle Unit 2 (in 1988) where the margin inherent in the nonlinear fluid effect, published in the above mentioned paper, was reaffirmed. Nonlinear fluid coupling effects due to.the use of current gaps at each time instant are not employed in this present application which imputes over 15% margin (in Holtec's analysts' estimate) in the computed rack response. In summary, the WPMR analysis utilizes a fluid coupling formulation-that is theoretically derived (without empiricism) and experimentally i 4 \ projects \980414\ license \RAI g validated. The assumptions built in the DYNARACK formulation are aimed to demonstrably exaggerate the response of all racks in the pool simulated in one comprehensive model. The DYNARACK fluid coupling model is neither approximate or empirical. A further elaboration of the details of the structural model used for the spent fuel racks and a mathematical explanation of . the manner in which fluid coupling is considered in the solution is provided below.

l. 'DYNARACK, developed in the late 1970s and continuously updated since that time to l incorporate technology advances such as multi-body fluid coupling, is a Code based on the i Component Element Method (CEM). The chief merit of the CEM is its ability to simulate friction, impact, and other nonlinear dynamic events with. accuracy. The high-density racks designed by Holtec Intemational are ideally tailored for the CEM-based Code because of their honeycomb construction (HCC). Through the interconnection of the boxes, the HCC rack l- essentially simulates a multi-flange beam. The beam characteristics of the rack (including shear, flexure, and torsion effects) are appropriately modeled in DYNARACK using the classical CEM

" beam spring". However, the rack is not rendered into a " stick" model, as implied by the staffs RAI. Rather, each rack is modeled as a prismatic 3-D structure with support pedestal locations l and the fuel assembly aggregate locations set to coincide with their respective C.G. axes. The rattling between the fuel and storage cells is simulated in exactly the same manner as it would be i experienced in nature: namely, impact at any of the ,four facing walls followed by rebound and impact at the opposite wall. Similarly, the rack pedestals can lift off or slide as the instantaneous l dynamic equilibrium would dictate throughout the seismic event. The rack structure can undergo l overturning, bending, twist, and other dynamic motion modes as determined by the interaction L between the seismic (inertia) impact, friction, and fluid coupling forces. Hydrodynamic loads, which can be quite significant, are included in a comprehensive manner, as we explain in more detail below, i [ As explained in the foregoing, the fluid coupling effect renders the mass matrix into a fully l . populated matrix. Modeling the fuel rack as a multi-degree of freedom structure, the following ~ j key considerations are significant:

i. Over 70% of the mass of the loaded rack consist of fuel assemblies, which are unattached to the rack, and resemble a loose bundle of slender thin-walled tubes I (high mass, low frequency).

I ii. In honeycomb construction (HCC) racks, as shown in a 1984 ASME paper [9],  ! the rack behaves like a stiff elongated box beam (End Connected Construction l l l racks, built 20 years ago and now obsolescent behave as a beam and bar l' assemblage). . Since the racks under inertial loading have overall structural characteristics of a multi-flange  ! l beam, it is computationally. wasteful (and, as we explain later, numerically hazardous) to model , such a structure as a plate assemblage. The DYNARACK dynamic model preserves the numerical stability. of the physical problem by representing the rack structure by an equivalent flexural and shear resisting " component element" (in the terminology of the Component Element j Method [10]).  ! 5 \ projects \980414\ license \RAI l p \ A detailed discussion of the formation of the fluid mass matrix is presented below. The problem to be investigated is shown in Figure RAI 1.1, which shows an orthogonal array of sixteen rectangles which represent a unit depth of the sixteen spent fuel racks in the Vermont Yankee Spent Fuel Pool. The rectangles are surrounded by narrow fluid filled channels whose width is much smaller than the characteristic length or width of any of the racks. The spent fuel

pool walls are shown enclosing the entire array of racks.

L ' The dimensions of the channels are such that an assumption of uni-directional fluid flow in a channelis an engineering assumption consistent with classical fluid mechanics principles. We consider that each rectangular body (fuel rack) has horizontal velocity components U and V parallel to the x and y axes, and that the channels are parallel to either the x or y axes. The pool walls are also assumed to move. , l \ i We conservatively assume that the channels are filled with an inviscid, incompressible fluid. Due i to a seismic event, the pool walls, and the spent fuel racks are subject to inertia forces that induce - ! motion to the rectangular racks and to the wall. This motion causes the channel widths to depart l from their initial nominal values and causes flow to occur in each of the channels. Because all of 1 the channels are connected, the equations of classical fluid mechanics can be used to establish L the fluid velocity (and hence, the fluid kinetic energy) in terms of the motion of the spent fuel i racks. For the case in question, there are 40 channels of fluid identified. Figure RAI 1.2 shows a typ; cal L rack (box) with four adjacent boxes and fluid and box velocities identified. The condition of I _ vanishing circulation around the box may be expressed as V = (v,ds = 0 or l st2 b/2 ("a ~"rN5+ kVa ~ vs. )dr7 = 0 -a/2 -6/2 l l where the subscripts (L, R, B, T) refer to the left, right, bottom, and top channels, respectively; 4,y are local axes parallel to x and y, and u, y are velocities parallel to 4, y. 6 \ projects \980414\ license \RAI 9 l . Continuity within each channel gives an equation for the fluid velocity as w = w,-( )s where w represents the velocity along the axis of a channel, wmrepresents the mean velocity in the channel, s is either ? or ti, and h is the rate ofincrease of channel width. For example, isi=U,-U ^ . From Figure RAI 1.2, four equations for us, ur, va, vt ni terms of the respective mean channel - velocities, can be developed so that the circulation equation becomes a (Un. - Ur ) + b (% - %) = 0 One such circulation ~ equation exists for each spent fuel rack rectangle. We see that the velocity -in any channel is determined in terms of the adjacent rack velocities if we can determine the mean fluid velocity in each of the 40 channels. Circulation gives 16' equations. The remaining equations are obtained by enforcing continuity at each junction as shown in Figure RAI 1.3. Enforcing continuity at each of the 25 junctions gives 25 equations of the general form s Ihc7w1IL6 2- -where w is the mid-length mean velocity in a connecting channel of length L and 5 is the - relative normal velocity at which the wails open. The summation covers all channels that meet at the node in question. The sign indicator o = 1 1 is associated with flow from a channel either into or out of ajunction. Therefore, there are a total of 25 + 16 = 41 equations which can be formally written; one circulation equation, however, is not independent of the others and reflecting the fact that the

sum total of the 25 circulation equations must also equal zero, representing circulation around a

' path enclosing all racks. Thus, there are exactly 40 independent algebraic equations to detennine

the 40 unknown mean velocities in this configuration.

Once'the velocities are determined in terms of the rack motion, the kinetic energy can be written and the fluid mass matrix identified ~ using the Holtec QA-validated pre-processor program CHANBP6.' The fluid mass matrix is subsequently apportioned between the upper and lower portions of the actual rack in a manner consistent with the assumed rack deformation shape as a function of height in each of the two horizontal directions. This operation is performed by the L L 7- \ projects \980414\ license \RAI \ Holtec QA-validated pre-processor code VMCIIANGE. Finally, structural mass effects and the hydrodynamic effect from fluid within the narrow annulus in each cell containing a fuel  ; assembly between fuel and cell wall is incorporated using the Holtec QA-validated pre-processor i code MULTIl55. The initial inter-rack and rack-to-wall gaps are provided in Figure RAI 1.2. These gaps, which directly figure in the computation of fluid mass effects in fluid coupling matrix, are assumed to apply for the entire duration of the earthquake. In reality, the gaps change throughout the seismic event and a rigorous analysis would require that the mass matrix be recomputed at every time- i step. Besides being numerically impractical, such refinement in the solution would reduce the conservatism in the computed results, as discussed earlier. The time variations in the inter-rack and rack-to-wall gaps are, however, tracked for the duration of the earthquake. Closure of any gap at any location results in activation of the compression gap spring at that location. The loads registered in the gap spring quantify the collis. an force at that location. The fuel-to-storage cell rattling forces and rack pedestal-to-pool liner impact forces (in the event of pedestal lift-off) are typical examples of collision forces that are ubiquitous in rack seismic simulations. The nonlinear contact springs in DYNARACK simulate these " varying gap" events during seismic events using an unconditionally convergent algorithm. In summary, the Whole Pool Multi-Rack (WPMR) analysis is a geometrically nonlinear formulation in all respects (lif1-off, sliding, friction, impact, etc.), except in the computation of the fluid coupling matrix, which is based on the nominal (initial) inter-body gaps. The modeling technique used (i.e. representation of the fuel rack and contained fuel by clastic beams and appropriate lumped masses) was chosen based on the application Codes, Standards and Specifications given in Section IV (2) of the NRC guidance on spent fuel pool modifications entitled, " Review and Acceptance of Spent Fuel Storage and Handling Applications," dated April 14,1978. This reference states that " Design...may be performed based upon the AISC specification or Subsection NF requirements of Section III of the ASME B&PV Code for Class 3 component supports." The rack modeling technique is consistent with the linear support beam-element type members covered by these codes. It is agreed that finite element models could be developed using plate and fluid elements, which may also provide satisfactory simulated behavior for a single rack. However, there is no known commercial finite element code which can treat multi-body fluid interaction correctly and sufficiently account for near and far field fluid effects involving many bodies (racks) in a closed pool. It is for this reason that the global dynamic analysis uses the fonnulation specifically developed and contained within DYNARACK. Response lh Holtec Report HI-88243 by Dr. Burton Paul reports comparison of DYNARACK fluid coupling fonnulation with over 100 experiments carried out in an independent laboratory under a 8 \ projects \980414\ license \RAI I '_10Cf R50 Appendix B program. These tests were performed with the sole purpose of validating the multi-body fluid coupling formulation based on Kelvin's recirculation theorem in classical fluid mechanics. These experiments, to our knowledge, are the only multi-body fluid coupling tests conducted and recorded under a rigorous QA program. The participating bodies used in the tests were carefully scaled to simulate rectangular planform fuel racks. The tests were run with a - wide range of seismic frequencies to sort out effects of spurious effects such as sloshing in the tank, and to establish that the fluid coupling matrix is (ought to be) independent of the frequency content of the impressed loading. The' University of ' Akron tests [3] performed some testing under the sponsorship of the predecessor company of U.S. Tool & Die, Inc. However, these tests were performed in the time when racks were still being analyzed using the Response Spectrum Method. We note that a theoretical model developed by Scavuzzo [2] is exactly that used in the Holtec WPMR analysis when the Holtec mass matrix is reduced to a single rectangular solid block surrounded by four rigid (pool) walls. That is, the work by Scavuzzo is a special case of a Holtec WPMR analysis for a spent fuel pool containing a single spent fuel rack. The Holtec WPMR fluid mass matrix for many racks in the pool is obtained by applying the same classical principles of fluid continuity, momentum balance, and circulation, to a case of many rectangular bodies in the pool with multi-connected narrow fluid channels. The experimental work performed by Scavuzzo, et al., does not attempt to model a free standing rack since many rack structures of that vintage were not free-standing. The experimental test is ~ equivalent to a single spring-mass-damper subject to a forced harmonic oscillation while submerged. If one accepts the fact that the fluid model used by Scavuzzo is a limiting case of the 1 more general Holtec formulation, then the good agreement of theory with experiment for the single " rack" modeled experimentally serves as additional confimtation that the Holtec theoretical hydrodynamic mass model, which is identical to the Scavuzzo model (for a simple rack) is reproducible by experiment. We have utilized the data supplied by Scavuzzo to simulate the experiment using the pre-processor CHANBP6 and the solver DYNARACK. The results of this comparison have been H incorporated into the Holtec validation manual for DYNARACK (HI-91700) as an additional conlirmation of the fluid coupling methodology. This validation manual, along with additional supporting ' documentation and discussions, was presented to the NRC in April,1992 under . dockets 50-315 and 50-316 for the D.C. Cook station and also was submitted in the licensing for l reracking of the Waterford 3. spent fuel pool. The submittal for Waterford contained the i evaluation of the Scavuzzo theory and experiment and demonstrated that the WPMR general formulation was in agreement with the experimental work presented in [2]. 1 Response Ic Rack "CA" geometry is used to compute the requested rack stiffnesses. The DYNARACK computer code requires stiffness inputs in the x, y and vertical direction in order to perform the ' dynamic analysis.  ; i i 9 \ projects \980414\ license \RAI l J F ' Attachment IG provides the calculations to develop the bending spring constants and the i extensional stiffness appropriate to the directions requested. Attachment 1H provides similar calculations for the shear springs and the torsional spring based on the methods presented in the Holtec Position Paper WS-126 which is contained in Attachment 1J. The following table summarizes the results calculated in Attaciunents 1G and lH. SPRING CONSTANTS FOR RACK "CA" Item Value Bending spring associated 4.614 x 10* with y deformation (rotation about x-x axis) (in.-lb) l Bending spring associated 2.511 x 10" i with x deformation (rotation about y-y axis)(in.-lb) 7 Extension spring vertical 3.958 x 10 direction (in.-lb)  ; 6 Shear spring associated 3.426 x 10 - ) with y deformation (iblin.) - ' 6 Shear spring associated 3.0 x 10 i with x deformation (Ib/in.) Torsion spring associated 5.079 x 10" with twisting (in.-lb) Response 1d The requested data, which has been excerpted from Holtec International Report HI-981932, is provided in attachment 1K. Response le Holtec Report HI-981969 provides the details of the development of the time histories used for l the Vermont Yankee pool from the design basis . Response Spectra. Figures RAI 1.4-1.9, reproduced from that report, demonstrate the required enveloping of the target PSD, over the frequency range important to spent fuel racks (3-7Hz) by the PSD regenerated from the ' developed time histories. 10 \ projects \980414\ license \RAI p:

References:

1. Deleted.

2. Scavuzzo, R.J., et al, " Dynamics Fluid Structure Coupling of Rectangular Modules in Rectangular Pools", ASME Publication PVP-39,1979, pp. 77-87.

3. Radke, Edward F., " Experimental Study of Immersed Rectangular Solids in Rectangular Cavities," Project for Master Science Degree, The University of Akron, Ohio,1978.

4.' " Evaluation of Fluid Flow for In-Phase and Out-of-Phase Rack Motions", by B. Paul, Holtec Report HI-98113, April 1987.

5. " Study of Non-Linear Coupling Effects", by A.I. Soler, Holtec Report HI-87102.

6. " Estimated Effects of Vertical Flow Between Racks and Between Fuel Cell Assemblies",

by B. Paul, Holtec Report HI-87114, April 1987.

7. " Fluid Flow in Narrow Channels Surrounding Moving Rigid Bodies", by B. Paul, Holtec Report HI-87112.

8. " Chin Shan Analyses Show Advantages of Whole Pool Multi-Rack Approach", Nuclear Engineering International, March 1991.

9. " Seismic Response of a Free Standing Fuel Rack Construction to 3-D Floor Motion",

Nuclear Engineering and Design,1984.

10. United States Patent No. 4,382,060, " Radioactive Fuel Cell Storage Rack", May 3,1983.

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RAI RESPONSES 1 for VERMONT YANKEE NUCLEAR STATION l RERACKING PROJECT iO i ATTACHMENT 1F

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R':sponse to Qu:stions on Technical Specification Chings 1 Rsqu::st NPF-38-193 ATTACHMENT 8 . SPENT FUEL STORAGE A outside of the gamma shield. The around 120r. Four trunnions are at- rate is estimated to be about 60 mrem /h.

neutron shield is enclosed by a thin tached to the cask body with the dose rate at any location outer shell. In addition to contiining the tiedown and rotation. Twofor of thlifting'e, accessible to the public well'within the resin, the aluminum also provides a trunnions are located near the top of the allowable limits.

conduction path for hear transfer from body and two near the lower end. The the cask body to the outer shell. lower trunnions may be used for rotat- MNSING The cask'is sealed usins redundant ing the unloaded cask benveen vertical Work on the project began last aurumn.

metallic seals. The cask cavity is pressur- and horizontal positions. The completion date depends largely on ized above atmospheric pressure with The long term cask surface tempera- the licensing process. A saferv-analysis ,

helium to preclude air inleakage in the cure has been calculated as being 191*F. report hu been submitted by Northern event of seal failure. The short term temperature on hot, States Power to the US NRC. whose ]

The casks will be about 16.5ft tall and sunny days has been calculated as 233*F. review is expected to be completed 8.5ft wide. A fully loaded cuk will weigh The maximum extemal contact dose around mid-April.

1 Chin Shan analyses show advantages of whole pool multi-rack approach By KP Singh and A ISoler l

Results from whole pool multi-rack (WPMR) analyses at Chin Shan and Oyster Creek point up '

the potential inadaquacies of single rack 3D analyses, and show just how important it is to carry out WPMR simulations, despite their abstruseness and high cost.

m

] Fuei storace racks are essentially thin-walled ceilular structures of prismatic proaches (viz the response spectrum method) are predicated on the assump-other with velocity u. The moving planes. for simpiicity of this illustration, cross-section. Althouch the details of tion that the structure is linear. A fuel are assumed to be infinitely lon5. such desi5 n varv from on~e supplier to an- rack, however, is the epitome of a that the motion of water exiting the other. certain ke non-linear structure (defined as one in inter-plane space is in the plane of the common to 21! c,v For esigns. physical attributes example, all are which the applied force does not have a page in the diagram below. For this l racks feature square cells of sufficient linear relacionship to the resulting dis- geometry, the velocity of water v ts opening size and height to enable placement). The stored fuel assemblies. computed by direct volume balance msemon and withdrawal of the fuel which constitute over 60 per cent of the (continuitv): l usembiv. weight of a fully loaded rack module. are 1 The cells (or " boxes") are arranged in free to rattle 'inside the storage cell In a time interval dt. {

a square (or rectangular) pattern and are durins a seismic event. The rack module fastened to each other using suitable itselfis not attached to the fuel pool slab. w (2u)(dt) = 2vd (dt) connectors and welds. The strav of cells Furthermore. the Coulomb friction re-is positioned in a vertical orientation sisting the sliding of the rack module on or v/u = w/d 4 and is supported off the pool slab surface the pool surface is, by defmition, a by four or more support legs. The spent non-linear force. This leads to the conclusion that the fuel pool is filled with the individual fuel el city of water exiting the fluid gap is TIME INTEGRATION TECHNIQUES racks. The plenum created by the w/d nmes the velocity of the approach support let:s is essential for proper in recognition of these highly non linear plane. In a typical spent fuel pool, the p cooline of the fuel assemblies stored in attributes of the dynamic behaviour of racks are about 250cm (100m) wide.

the rac'k. which relies on natural convec- fuel storage racks. 'their seismic simula- and are spaced at 4.5-7.5cm (2-3inJ rive cooling to extract the heat emitted tion hu been carried out using time by the spe'nt fuel. However, it hu the integration techniques. The state 'of-the-insalutary etTect of making it kinemati- art analvsis technique involves mod-cally less stable. Regulatory authorities require careful and comprehensive anal-elling s' single rack module as a in structure with features to capture the { .

$ v

/

ysis of the response of the racks under fuel assembly rattling, module sliding. A the seismic monons postulated for the rocking and twistmg monons. f1L D pool slab. Despite the versatilirt of the m 4 1

(~) Non-linear arructure. Such an analysis seismic model, the accuracy of the single rack simulations has been suspect due to

-) cannot be conducted in the manner of one key element: namelv. hvdrodvnamic anventional structural analyses for particip' ation of water irou'nd the racks.

,,ower plants, because the classical ap- This etTect is understood bv considering the motion of water between laree tiat j T!.e .suve .are with N,isu hirentanomd .* Pl anes of width w at a (sm2!D dist2r.ce d A Two submerged paraitel flat planes F.nre.n .ir e. rue Grrn Nd4.V O.W3 Ws C14 apart. which sre moving towards each approaching eacn otner.

_ J

1 _

i SPENT FUEL STORAGE (

i rack to-rack (or rack-to-pool walb colli- Shan analysis, the min analyses yielded

[)tervals. For the rack modules arranged a maximum kinematic displacement of a i 1 3 a typical spent fuel pool w - 100in sions dunng a severe seismic event.

Taiwan Power set out to determine the rack in the pool- 8.5 times the single l

1. and d 2in, so v/u 50. Since kineric  :

response to racks by a comprehensive rack analysts prediction. The impact I energv velocin. t is ke water exiting the inter-rackreportional whole pool analysis. to the square luads benveen of rack support pedestals space will have 2500 times the specific Under a consulting contract with and the pool slab decreased slighth from I Taiwan Power. Holtec International the values obtained from the single rack l kine.ic energy of the moving rack. analysis. In the Chin Shan analysis, the This hydraulic ener;;y is either drawn (USA) undertook to prepare a dynamic l i

from or'added to tne moving rack. model of the entire assemblage of racks coefficient of friction between pedestal and slab was about 0.2. Even though the I modifving its submerced motion in a (a total of 14 modules) in the pool, with sienificant manner. The dynamics of due consideration of fluid coupling rack displacements relative to th'e slab (

one rack. therefore, affects the motion of effects. Holtec's code imnack, which showed a large increase over the single l uses the component element Method for rack results, no rack-to-rack or rack-to-all others in the pool. A dynamic non-linear dynamic analysis. and has wall impacts were predicted.

sime i stion which treats onlv one rack. or a small grouping of racks.' therefore, is been used in over a dozen fuel rack l OYSTER CREEK ANALYSES intrinsically inadeuuate to predict the licensing projects. was used for this motion of rack modules with any quan- purpose. Subsequent to the Chin Shan analysis.

Holtec International completed some I tifiable level of accuraev. The results of this first ever so-called whole pool multi rack (wt'uR) anaitsis similar work for CPU Nuclear's Ovster l Creek piant located near Toms R'iver.

EXPERIENCE IN TAlWAN pmvided t'urther insicht into the in-p'ool rack dvnamic behas iour. Tracking oithe  !

Taiwan - no strancer to seismie trem- New Jerser. The Ovster Creek ansivses ors - has three n'uclear installationt inter rack cap showed that the presence were pert'ormed usinc coctiicients' of iricdon oi0.2 and 0.S. In this case. the l Kuoshene. Maanshan and Chin Shan. of ware has the effect of miecans a Taiwan I fower Company procured racks certain symmettv mto the monon' of maximum displacement of any rack in l for the Chin Shan site from General adiacent racks. although a cert 2in the pool predicted by the nun analyses l amount of out-of-phase monon occurs. was 1.4 times the sincie rack analysis Electne Company in 1986. These racks I are of the so-called honeycomb con- Companson with s ngic racn m anah> predienon. In this anaiysis. the pedestal struction. and were initially analysed by ses. however. pointed to the rather to slab impact loads predicted by the (

a single rack m seismic model. Recon- unsettiint conclusion that the sincie nun anaivses are slichriv creater than mzin'c the inadequaev oisuch a model En rack mcJets do not bound the results'oi the values obtained from 'the sincie' rack the whole pool simulations. !n the Chm analysis Orognosticate the potennal h2ard of l kJ New storage technology a Greifswald  !

f By W Fischer. 5 Standke, M Liin and K Hoclurrate

• t The Greifswald site in the forrner GDR boasts a large interim fuel store (as well as four now shut down WER-440s and four rnore in various stages of construction). In recent years the i east Germans have been working towards expanding the capacity of the store by re-racking the por'ds using locally-developed transport / storage baskets.

I in 1W. ths arrangements for shipping itally designed ha kets to accommod.ae centreco-centre distance of 226mm hc- l pent fuel from Rheinshcre I X '.m awemblics with faikd clemenh. A fourth tween ene tuel assemhhe>. j m.i and Grcibuaid u X h : .w in pond acts as a resen c. NEW BASKET l cavern Germam hack to the 'NR were At the end of IWO. Z Ah comamed amund 2a00 spent fuel awemblies from Work has been in pmgress a expand the changed. l'revidush. the spent fuel had been returned af:er conhnu for three the n i n... .m and the m um it n yeary at the power st.tNon Io mcctthe pl.InncJ that a further 2.;00 fuct anem- ,,3 , ,

hoviet rVqturemynt for a itve-ycar Cool- bltes stored in (.areifswald units I W 4 l :rse cw me time, the mterim uorace faciliev will be transferred to ZAll after a cooline 2'

• i

/..kll i/.u IsheniJuer fur a pebrannte time o[lbfee ye.fr% Ibkh Ir. ins [cr ks to hC g,,vv q f Orennstahes was hddt at (vrythu did. performed using the ik4 e TJ6l ifano , ,gl l J -

[A8 u.h sic tpncd hr the Nniet% Jnd porter the (krman Fuci y , ; yw +i!  !

Inunate,descloped l'reibert. at ,l'ht the nru stats na heen had in hold a.m meet s y i 8

4% undaniaged sp6m fuci aswmblies n.uional and intcEational tran pon eg- W1j.lll l l CM l l l

n i Mnt hcJs s metal M O, l'hs ancmblics ulations for lipe h packapet .. m.... . ,

are uorgd in thrse postd% vash .tunni- Nos is t dcs clopt d [utt haE.

f wne imth as inwrh for the e 6.id and p #

' mod.tfing 52 hadet % u ith mome sps -

I

.h storJye rask m :he /.AB ponds. I his  %.  %,  !

a t i eu s ,,, s s sa . s. *... .n md. t hg need to handk singk nic! mta C',0 e nrT l' e

'. .J ffe n s. <. fi . . 's. . . * '.* / to assmbhes in the [Ah la6ihn. I'.M  :

ha i som .U n s M fus! .sws +'I'nc' A Cross-section of the ZAB mterim

.. . . o . A ,n .. . . / /., , , e, i billl M s. Ths .rinsahn .afs n W SDent fuel store at Greifswald m eastern o

o"n, . .

Germany.

hadct a ashicsed bs hasmn a 4 -t.

NUCLEAR ENGINEERING INTERNATIONAL 20

i 1

RAI RESPONSES for VERMONT YANKEE NUCLEAR STATION RERACKING PROJECT ATTACHMENT 1G I

o ATTACHMENT 1G Stiffness Calculations The purpose of this document is to calculate the rack bending and extension spring constants, which characterize the Holtec fuel racks for Vermont Yankee Nuclear Station. The formulas are obtained from an article published in Nuclear Engineering and Design, which is titled " Seismic Response of a Free Standing Fuel Rack Construction to 3-D Floor Motion." The calculated values are used as inputs to the 22-DOF single rack analysis.

The following rack data, which is specific to rack "CA", are used in the calculations:

b := 6.00 in Cellinner dimension t := 0.075 in Cell wall thickness (minimum) p :=6.218 in Cell pitch Nx := 14 Number of storage cells in the x direction Ny := 19 Number of storage cells in the y direction H := 168 in Height of the storage cells E := 27.610'-psi Young's modulus of cell material First, the area and moment ofinertia properties of the gross cross section aie calculated.

For a BWR fuel rack, any two adjacent storage cells share one wall. Therefore, the determination of the storage cell metal area and moment ofinertia (per cell) are based on an effective cell wall thickness equal to one half of the actual wall thickness.

t e :=.'.

2 t , = 0.037 in Effective Cell Wall Thickness The cross sectional area and moments ofinertia of a single storage cell are 2

/ eeg; :=4-(b + te) te A,,;; = 0.906ain Area of Single Cell d

I .

Icell = 5.502 in Moment ofInertia of Single Cell cell := 12-(b + 2 t )'- b' e

Page1of3 \porojects\980414\ license \Rai

' ATTACHMENT 1G The cross sectional area and moments ofinertia of the gross cell structure are A rack:cNx NyAceti 2

Arack = 240.8%g Area of Gross Cross Section

/ INx - 1) Nx-1 ) I INx- 1) I no  :=if floor = ,1,0.5 row :=0.. : ceil '

-1 i ( 2 / 2 / (. t 2 j j

- I ,:= Ny- Iceii + Acelk'[("0 + row) p]2

~

IN x- 11 x Iyy rack :=il 11 r. =N - 1 (N Iycetit2D),2D i 2 j 2 5 4 lyy rack = 1.528 10 6 Y-Moment ofInertia of Gross Cross Section l

l l IN - 11 N - 1 INY-11 I n 0 :=if fl r = ,1,0.5 row :=0 cet! -1

\ \ 2 > 2 \ 2 1 I l

I

.1, := N x ! cell + ^ cell'[("0 + row) p]2

~

lurack :=if 11 r IN - 1)= N - 1 (N Ix celi + 2 D),2 D 2

i 2 /

5 lxxrack = 2.80910 =in' X-Moment ofInertia of Gross Cross Section Page 2 of 3 \Porojects\980414\ license \Rai

ATTACHMENT 1G The following stiffness values are calculated based on the material and geometric properties of the rack EArack 7 lbf KEXTENSION " KEXTENSION = 3.95810 9 11 m EIntack 1 K BENDINGX to KBENDINGX " g

.g = 4.61410 +Ibfin E lyyrack I KBENDINGY " K BENDINGY = 2.511 10'o.Ibf in g 'g l

~

page 3 of 3 Wojects\980414\ license \Rai L-

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I RAI RESPONSES for VERMONT YANKEE NUCLEAR STATION RERACKING PROJECT 1 ATTACHMFsNT 1H 1 l

l i

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ATTACHMENT 1H DYNARACK SHEAR DEFORMATION CALCULATION IN ACCORDANCE WITH HOLTEC WS 126 - CALCULATION OF COMPOSITE SHEAR DEFORMATION CONSTANT FOR RACK "CA", VYNPS Input Height of Storage Cell Ltotal := 168 in Length ofintermittant weld L 3 := 8 in Total Length Between Intermittant Wolds L2 (Ltotal- 6 L j) 5 h = 2%

I' 1 x;:= x ; = 0.048 I' total L 3+L2 L2 X

2 = 0.19 x 3 = 0.143 x2:* L total x 3 := total L

2 A2 :=x 3 + (x 2 + x j)2- x 2 + (2 x2 + x j)2- (2 x 2) + (3 x 2 + X 1) - (3 x 2) -

+ (4 x 2 + X 1)-(4'X2) + (5 x 2 + X 1) - (5 x2)

A2 = 0.286 2

B2 :=x 2 - x 1 + (2.x2) - (X 1+X2) + (3 x 2) - (2 x 2 + x j)2 + (4'X2)

+ (-1)-(3 x 2 + X I) + (5 x2) - (4 x2+ X I)

B2 = 0.714 We can now compute the effective shear deformation for each of two directions, using the information calculated in Attachment A.

Shear Deformation Coefficient for the "y" direction deformation (Attachment A)

Moment of inertia of rack cross section about the x-x axis Ixx .:= 280900 in' Metal area of rack cross section 2 A t:=240.896 in Moment ofinertia of a single cell I := 5.502 in' Page 1 of 4 Yrojects\980414\licenseiRai

r ATTACHMENT 1H Number of cells in rack n := 1419 n = 266 a :=2.276 (from calculation file "sheardef.mcd")

( g g := 31.2 a- -

$ ;g = 2.934 AtLtotal This is the shear deformation coefficient if the entire rack were of honeycomb construction For an individual fuel cell in an end connected region

$ ,:=31.2d b ' $ , = 0.395 5At L;2 Then the effective shear deformation coefficient for the assemblage of individual fuel cells is

$2 := (1n1+ $xx e) -1 I $ 2= 266.727 2

I stary 0 ll A2 + $ 2 B2.x 3 The effective shear deformation coefficient, for the full length composite section,is I stay = 4.726 SF. ear Deformation Coefficient for the "x" direction deformation Moment of inertia of rack cross section about the y-y axis I yy := 152800 in' Metal area of rack cross section A t:=240.896 in2 Moment ofinertia of a single cell d

I := 5.502 in Number of cells in rack n := 1419 n = 266 a :=2.245 ' (from calculation file "sheardef.med")

Page 2 of 4 $rojects\980414\ license \Rai

ATTACHMENT 1H I'

( ig := 31.2 a -

$ 33=1.574 AtLtotal This is the shear deformation coefficient if the entire rack were of honeycomb construction For an individual fuel cell in an end connected region

$ e :=31.2db I $ e = 0.395 5AtL2 Then the effective shear deformation coefficient for the assemblage of individual fuel cells is 42 '~ (1n1+ $g e)I l & 2= 144.634

$ m :=$ gi A2 + $ 2B2 x 3 The effective shear deformation coefficient, for the full length composite section, is I starx = 2.558 Page 3 of 4 \ projects \980414\ license \Rai

r ..

ATTACHMENT 1H DYNARACK TORSIONAL INERTIA PROPERTY CALCULATION IN ACCORDANCE WITH HOLTEC WS-126 -

For the effective torsional rigidity, we use the formula in WS-126. For the VYNPS CA Rack d

J g := 170800 in (from calculation file "sheardef.med")

The torsional inertia of the rack considered as a series of end connected tubes is d

J2 := 5820 in (from calculation file "sheardef.med")

Therefore, following WS-126, we have the composite moment of inertia as J

I 3 J := J = 8.03810 ain' f J ) -

A2+ j B2 l

( 2 / {

l SHEAR AND TORSION SPRING CONSTANTS I I

I Young's Modulus E :=27600000 psi H := 168 in  !

Poisson's Ratio y := 0.3

)

Using results from Attachment A Shear Spring Rate for y-direction deformation 6

K K shear _y = 3.42610 g shear _,y := (4.614 H- 10" in lbf) , (1 + 4 stary)

• i Shear Spring Rate for x-direction deformation I

K K shear x = 310'h shear _x := (2.511 10" in lbf)- ~

H -(1 + $ sta x)

Torsional Spring Rate K Ktorsion = 5.07910'albf in torsion '~ 2-(1 + v) H Page 4 of 4 \ projects \980414\licenseHal i

i HOLTEC POSITION PAPER WS-126 I

ATTACHMENT 1J 4

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THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK y HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 Scope The Holtec Spent Fuel Racks are modeled as a linearly clastic structure in the dynamic

~

simulation of a seismic event. The methodology for calculation of spring constants for the dynamic model has been developed for more than a decade. While the calculation of spring constants for bending and extension is well understood and simple to explain, the calculation of the spring constants for shear effects and for torsion merits additional comments. While l the development is set forth in appropriate Holtec theory reports, the pmpose of this position paper is to set down the theoretical basis in a form that is easily communicated to the general public on a "need-to-know" basis.

Evaluation of shear deformation in a composite beam representing a fuel rack i

The purpose of this development is to demonstrate how the effective shear deformation coefficient for a composite beam representing the Holtec spent fuel storage rack is 1 l

developed. Modeling the rack as a composite short beam section oflength L, area moment of inertia I, and metal cross section area A, the complementary energy U is written as f

yz gy2%

.U=Q + dx g 2EI 2GA, where M(x) = bending moment V(x) = shear force and x represents a length coordinate defined along the neutral axis of the beam.

Holtec International Proprietary Information Page1of11

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK O HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: Cetober 23,1998 i E, A are the Young's Modulus and Shear Modulus, respectively. Let us consider a beam that has uniform A and I but unknown shear deformation constant a along the beam. That is, in general, write '

U=

2El lfM' +aS

< GA V*

dx Define , = 12 EI 12 L,5-a,GA = T, L 9<

where i = 1,2 represent distinct regions along the length of the beam where at has two distinct values ai, a2. In particular, in region 1 oflength Li, we have fully welded honeycomb construction where ai = 1.6-12/5 which is appropriate for a rectangular shaped multi-cell grid beam. In region 2 oflength L , 2we have an effective shear deformation coefficient a2 hatt represents the assemblage ofindividual boxes. The complementary energy .

U can now be written as:

I U= j ,M (,), p(,)y2(x) 2 j,

2El *  ;

We consider a uniformly loaded cantilever beam with an additional concentrated load P at I

the tip and determine the tip deflection S using Castigliano's Theorem.

01 H a u nte m atie e ai e m g s et o fe m - ie n Page 2 of11

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 l

/ 1 v v v v v v v v v v v v v v v v

/ l

/ J P;S x  !

/

v i

Noting that M(x) = Px +qox2/2  ; V(x) = P + gox

~ ~

6=BU 1 BM BV BP p o

=El-l$ M OP

+ pV BP ,,o _

dx

_ r .n BM BV

=x -=1  ;

BP SP l

so that,in general

]

1 6 = gff 'M(x)x + p(r)V(x)(y,,,dr

( -

Holtec International Proprietary Information Page 3 of 11

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK

[]

V HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 We first compute this for a beam where there is a uniform connectivity so that 9 (x) = 0. The solution is:

S = q,1) ~1 + 49' Note: thatincreasein p leads tolower naturalfrequency.

8EI .

L' We see that in general, the term dealing with shear deformation is the second term. Thus, -

the uniform beam oflength L, for this special case, 98 9"' xdx O 2 EI 9 = j' p(x) El or BL'

= l$ p(x)xdx 2

Of course, for 9(x) = 0, the above relationship is an identity. However, we now postulate that for the composite beam consisting of 6 sections oflength Li and 5 sections of tength L2 ,

the umn relation can be used to define an effective 0 for a uniform beam and thus define an 2

effective 4 = (12/L ) 6.

Then in general, the effective $ is dermed as P. .

V- Holtec International Pronrietary Information )

l Page 4 of11

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998

$L'=lfp(x)xdx 24 orin termsof a coordinatez=1 L

2 L where L = totallength of beam

$ 24 = l' p(z)zdz Now, in each of the sections (i.e. honeycomb or end-connected) we have L'L or p2 = L#

p, = 12 , 12 ,

so that

< 32 r _s2

$=25,,L, t zdz + 259 $2 -L zdz

< <L, where

~ Jt ,( )dz means the totality of the integrals evaluated over all of the sections of length Li .

We note that the paper " Seismic Response of a Free Standing Fuel Rack Construction to 3-D

- Floor Motion", by Soler and Singh (Nuclear Engineering Design, Vol. 80 (1984) pp. 315-329) provides the appropriate equations for $i and $2. The work in that paper assumed that shear deformation effects could be neglected in the so-called "end-connected" construction because that construction was present over the entire length of the rack. In the Holtec rack 1

constmetion, the individual regions between the honeycomb welded sections can have a much more important shear deformation effect since the ratio of characteristic cross section length to length of end-connected construction is much larger. Modifying the formulas in the paper so as to include shear deformation in the individual segments of end-connected construction leads to the following results for (i and $2:

O neitee i ter tie i rrePriet 171rerm tie-l Page 5 of 1I 1 i

I

e ,

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK

-O

1. HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 12EI' a a = 6/5

,= GA*L*,

(1 + , )I*

,= _1 nl where l' = totalmomentofinertiaabout thecentroidof n fuelcells.

4 I = momentofinertiaabout thecenterof a singlefuelcell(or opening) n = numberof assembliesin therack A* = nA; A =areaof a singlecell If wedefine 12EI' n= .,a = sheardeformatim coefficiert if theentirerack were honeycombweldedconstructbn, then O $ = 2l t, $gzdz + 25 t, $2 v-zdz

<L,

. Now consider the spent fuel rack where we have six honeycomb (welded) sections and five end-connected sections. For the purpose of this analysis, we assume an average L2 (the bottom section has a smaller unconnected length) and define l x, = L,1 L x, = L' + L' where L

\

Holtec International Proprietary Information I

Page 6 of11 l

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK

("') HOLTEC POSITION PAPER WS-126 Author Alan I. Soler, Ph.D.

Revision 0: October 23,1998 the following sketch shows a representation of the fuel rack as a single beam with regions of honeycomb (welded) construction and end-connected individual fuel cells identified:

< L y li

(- -> < >

L2 (Typ.)

Li (Typ.)

Define geometric quantities A and B so that

(-

LJ A / 2 = l zdz + l,"' zdz + $ "' zdz + ll"' zdz,

+ ll"' zdz + .5^s,"'"' zdz B I 2 = l zdz + $2,, zdz + ll",',.,, zdz + lj,",',,, zdz

+SU',n,zdz Integrating yields the following results for A and B A=x 2i + (x, + x2 )' - X 22 + (2x2 + x,)2 -(2x )2 2

+ (3x 2+ X i)2 -(3X 2)* +(4X2 + Xi )* -(4X2 )2 + (5x2 + xi )2 -(5x2)*

B=X 2 -X i + (2X 2)* -(X2 + x,)2 + (3X2 )2 -(2X2 + X

. i )' + (42 2

)2 -(3X2 + iX )2 + (5x 2 )' -(4X2+X)'

i V, Holtec International Proprietary Information Page 7 of 11

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK 7 x, HOLTEC POSITION PAPER WS-126 t/ Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 Note that in the limit as Li =0, and L2 = US -

A=0 and B=1.

In the limit Li = U6 and L2 = 0. -

A = 1 and B = 0.

Therefore, in terms of the known values for hn s h2 the effective shear deformation coefficient for the composite beam representing the spent fuel rack is expressible in the final form p = pu A + N2 (

2

)2 g The determination of the effective shear deformation value for the rack in terms of the rack geometry and proportion of honeycomb and end-connected construction enables the appropriate spring constant to be calculated per the formulas in the reference text

" Component Element Method in Dynamics", by S. Levy and J. Wilkerson, McGraw-Hill, 1976. This is implemented in the Holtec computer code for simulating the dynamics of spent fuel racks.

Evaluation of effective Polar Moment ofInertia in a composite beam representing a fuel rack I 1

i O. Holtec Intemational Proprietary Information

%J Page 8 of 11 l

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 To compute the appropriate torsion spring constant to represent the resistance of the spent fuel rack to a twist about its longitudinal axis, the effective Polar Moment ofInertia needs to be computed. Consider the complementary energy U for the torsional behavior of the composite beam-like structure representing the spent fuel rack:

y (T'(x)g a 2GJ where T(x) is the twisting moment acting on the beam. The varicb!c x is the coordinate defined along the length of the beam in the same manner as was done for the shear deformation calculation above. Let ma be a distributed unifonn twisting moment (per unit length of beam) and To be a concentrated twisting moment at the tip of the beam (x=0).

Using Cutigliano's Theorem, the angle of twistst the tip of the beam is obtained as g, =L T(x) mo L x

-) G GJ' (x) > J(x)

For the beam with constant J, the following relation is an identity  !

L' L x 5

2J o J(x)

For a beam with variable J(x), the above relation can be used to define an effective uniform J

= J., to give the same angle of twist at the top of the beam ,

)

.O neitee i ter atienai rre914 tarv i fermatien Page 9 of 11

(;

THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK (O HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 L* _

L x 2J, 0

/(x)

As was done previously in the development of the effective shear deformation value, we introduce dimensionless variable z=x/L. Then the effective Polar Moment ofInertia J, can be written in the form 1 -

d: -

zd

= + .

2J, .L, J, .Ls J 2 where the integral notation represents integration over the appropriate sections of the Q composite structure representing the spent fuel rack. We recognize the integrals as the functions A and B defined previously. Therefore, J, is given in terms of the individual properties Ji and J2 (for the different construction along the length of the beam representing (

the spent fuel rack) as: j 1

1/J.= A/J i+B/J2 1

Therefore, the effective Polar Moment ofInertia is written in the form I J,= J' (A+ J B) 2 Holtec International Proprietary Information l

Page 10 of11 1

r THEORETICAL BASIS FOR SHEAR AND TORSIONAL SPRING CONSTANTS IN A HOLTEC SPENT FUEL RACK

( HOLTEC POSITION PAPER WS-126 Author: Alan I. Soler, Ph.D.

Revision 0: October 23,1998 With the effective Polar Moment ofInertia developed, the appropriate twisting spring rate is formed using the methodology in the text " Component Element Method in Dynamics" and implemented in the Holtec computer code for dynamics of spent fuel racks.

I O

Holtec Intemational Proprietary Information Page 11 of11

RAI RESPONSES for i VERMONT YANKEE NUCLEAR STATION RERACKING PROJECT ATTACHMENT 1K I

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Table 2.2 NUMBER OF STORAGE CELLS MODULE QTY. NUMBER OF CELLS Total Per I.D. Rack North-South East-West Direction Direction CA 1 19 14 266 SE 1 17 12 + 2 cells 238

• NE 1 12 14 168 9 - - 2,683 Racks m.8 Pool TOTAL: 12 - - 3,355 Includes use of eight dual-purpose storage cells for fuel.

SHADED REGIONS ARE HOLTEC PROPRIETARY INFORMATION Report HI-981932 2-10 i

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Table 2.3 MODULE DIMENSIONS AND WEIGHTS FOR VY RERACK Dimmsion (inches)t Module I.D. Shipping Weight in Pounds North-South East-West NE 74.991 87.427 16,810 D. 87.427 25,860 CA 118.517

()

23,800 SE 108.2 87.916 J

' Nominal rectangular planforrn dimensions.

SHADED REGIONS ARE HOLIEC PROPRIETARY INFORMATION 2-11 Report HI-981932

vrummr Ymut Nt cu:.ut Powrn Coluaunmos Docket No. 50-271 BVY 99-115 Attachment 2 Vermont Yankee Nuclear Power Station Response to Request for Additional Information Regarding Spent Fuel Pool Storage Capacity Expansion (Civil and Mechanical Engineering Considerations)

Response to Question 2

'"""N Y^h0E Ntut^n Powtu couronruns BVY 99-115 / Attachment 2 / Page I Questlos 2a:

You indicated that the design conditions described in SRP Section 3.8.4 and ACI 349-80 were used as a guidance in the calculations of the spent fuel pool (SFP) capacity. With respect to the SFP capacity calculations using the ANSYS computer code presented in Chapter 8 of the submittal:

a) Provide physical dimensions of the reinforced concrete slab and walls, .

liner plate and liner anchorage.

Response 2a:

De structural analysis that was performed to determine the capacity of the spent fuel pool (SFP) structure did not employ use of the ANSYS computer program. The ANSYS computer code was used in the fuel rack pedestal bearing pad analysis, as described in Section 8.5 of the submittal.

With the exception of the structural reanalysis performed by EBASCO in 1984, all follow-on analyses addressing the SFP capacity, including the structural analysis of record, utilize hand analysis techniques. The computer analysis performed by EBASCO used the finite element program, NASTRAN, and provided the basis for the scope of the structural evaluation in the follow-on hand analyses.

The SFP slab is 26'-0" wide by 40'-0" long and is approximately 4'-1" thick. It is supported on the north side by the 4'-5"(minimum) thick reactor shield wall and on the south side by a deep, S'-6" thick east-west concrete tie beam, which also serves as support for the south wall of the SFP. The east and west sides of the SFP slab, and the east west tie beam, are supponed by a pair of deep nonh-south concrete beams, 6'-0" thick, which, in turn, are supported by the reactor shield wall and the 3'-2" thick south exterior concrete wall of the building. The SFP is 39'-0 %"

deep.

Concrete floor slabs at elevations 303.00', 318.67', and 345.67' are contiguous with the SFP structure.

The interior of the pool is constructed with a stainless steel liner anchored to the concrete walls 1 and floor by stiffeners. The liner is designed as a leaktight membrane and is not relied upon as a structural member. The plate thickness of the floor liner is 1/4 inch, and the plate thickness of the wall 3/16 inch. The spacing for horizontal and venical stiffeners on wall liners is 6'-6" and l'-8",

{

. respectfully. On the floor liner, the stiffeners are only in the north-south direction having a I spacing of 4'-5" of 5'-7." Beneath the floor liner there is an 11" thick grout pad on top of the I rough concrete surface at El. 305.50'.

I

- Question 2b:

Provide the mesh used in the analysis.

Response 2b:

1 As noted in the response to part a) above, computer analysis techniques were not employed in the SFP capacity analysis of record. The SFP reanalysis performed by EBASCO in 1984 using the finite element program NASTRAN provided the basis for the scope of the follow-on structural analysis. The methodology employed in the structural analysis to determine the SFP capacity is

Wmm1 hen Nrcaan Powen Com on.u m BVY 99-115 / Attachment 2 / Page 2 contained in Section'3.3.3.6.2 of Vermont Yankee's licensing submittal Spent Fuel Storage Rack F2 placement Report for installation of the NES racks (dated April 1986) and is reprinted in pan

%Iow. (References noted below will be provided if requested.)

"The basic approach in determining fuel rack load capacity was to i;olate the controlling element of the load path and then determine the load capacity (in terms of ksf fuel rack floor load) with respect to this element. The strength design method for reinforced concrete .vas used in conjunction with conventional structural analysis procedures to determine capacities. Section strengths were determined using the methods and procedures contained in Reference 31 The allowable limit loads were converted to actual rack loads using the load equations contained in

. Reference 9.

The impact loading associated with predicted rocking motion of the racks was included as part of the seismic loading considerations. Determination ofimpact energy of the fuel racks was based on. maximum predicted gaps between the fuel rack support legs and the supporting surface of the fuel pool floor. Gap data were obtained from Table 3-8, which considered an eouivalent fuel rack load of 2.87 ksf. This data was a direct product of the NES Report identified as Reference 34.

The impact energy applied to the slab was detennined using vectorial partitioning of the total racking energy with respect to normal impact on the slab (impact energy based on velocity component normal to the slab).

He kinetic energy transmitted into the slab during impact and the resulting structural response was determined using the methods and procedures contained in References 31,32, and 33.

To account for the nonconcurrent aspects of the impacts from different racks (Reference 34), the slab kinetic energies from each impact were combined using the square-root-sum-of squares (SRSS) method to determine the maximum concurrent kinetic energy of the slab.

Kinetic energy was then equated to strain energy to determine the maximum structured response. )

This enabled determination of the differential uniform load capacity required for impact effects.

He fuel rack load capacity was then determined from the controlling load equation consiJering the fuel rack dead and seismic (with impact) loads to be proportional to the mass of the fuel racks."

The methodology described above was taken directly from Vermont Yankee's licensing submittal for installation of the NES racks in 1986. Although the methodology remained essentially the same for the structural analysis in support of the proposed installation of the Holtec racks, it is important to note two differences from the above methodology: 1) the equivalent fuel rack load was taken as 2.93 ksf(due to the weight of fuel being conservatively taken as 700 lbs, compared i to a fuel weight of 670 lbs used in the 1986 submittal), and 2) impact energies from individual racks were conservatively taken as acting concurrently and, consequently, cumulative impact effects were determined by means of absolute summation.

Question 2c:

Identify and describe the boundary conditions used in the mesh.

""" ^*E bCC L AH P(8WLH C(plus(phag H3N BVY 99115 / Attachment 2 / Page 3 Response 2c:

In the hand evaluation performed, the slab boundaries were taken as fixed due to the high relative stiffness of the supporting concrete beams and walls.

Question 2d:-

Provide the material properties used in the analysis.

Response 2d:

The properties of the reinforced concrete SFP used in the analysis are the same as the properties identified in Section 2.3.3.2 of Vermont Yankee's licensing submittal Spent Fuel Storage Rack Replacement Report for installation of the NES racks (dated April 1986) and is reprinted in part below, "The following symbols are used to describe the material properties used in the structural analysis:

f', = Compressive strength ofconcrete y = Specific weight p = Poisson's ratio a = Coefficient oflinear expansion E = Modulus of elasticity fy = Yield stress A. Concrete f' = 4000.0 psi * (all members except SFP slab)'

(f' = 6400.0 psi - SFP slab)*

y = 150.0 pcf ,

I p = (Q = 0.181 1 350 E = 3.834 x 10' ksi

• a = 5.5 x 104 per'F l- B. F_einforcind Sg.pJ y = 490 pcf fy = 40.0 ksi I

L

m-I . Vamost YAwamecu:An Powr.n Coimon mos BVY 99-115 / Attachment 2 / Pcge 4 i'

p = 0.3 E = 29 x 10' ksi ot =_ 6.5 x 104 per 'F For the fuel pool floor the in place concrete test data shown in Table 2-2 may be used."

Table 2-2 documents the testing program results that were the basis to upgrade compressive strength properties of the SFP slab. Table 2-2 is attached to this response.

Question 2e:

Describe the applied loading conditions including the magnitudes, and indicate their locations in the mesh.'

Response 2e:

The SFP structure is subjected to dead load, live load from the contiguous floors, hydrostatic load from a water depth of approximately 35'-9", the weight of the high density spent fuel racks plus fuel, and the effects of OBE and SSE seismic loads, as well as thermal loads. The loads are essentially unchanged from those described in Vermont Yankee's licensing submittal Spent Fuel Storage Rack Replacement Report, dated April 1986, which was prepared in support of the proposed installation of the NES racks. The loads and load combinations are defined in Sections 2.3.3.3 and 2.3.3.4, respectively, of that report and are reprinted in part below. (References noted below will be provided if requested.)

"2.3.3.3 ' Desian Loads A. Dead Load (D)

Dead loads consist of the dead weight of concrete, grout steel liner, fuel racks, fuel, and any equipment permanently attached to the pool.

Hydrostatic pressure loads acting on walls and' floor shall be included in this category. It should be noted that the hydrostatic load acting on the north wall of the pool is dependent on the operation condition.

Dead loads distributed from the contiguous floors and beams to the pool walls shall be considered in the analysis. An 80 psfintensity on all floors connected to the pool shall be used to account for piping weights as documented in Reference 13.

B. Live Load (L)

, Live loads are random, temporary load conditions during maintenance and operation. They shall include the following loads:

e . Weight of fuel cask

y VEHuuNT YANKF.E NUCl.EAH POWEH COHl' OHM mN BVY 99115 / Attachment 2 / Page 5

'e _ Weight of control rod storage racks and stored control rods e ~ Weight of refueling bridge aind service platform Live loads distributed from the contiguous floors and beams shall be calculated based on the following load densities:

• El. 345.67' 500.0 psf e El. 318.67' - 200.0 psf

e~ ' El. 303.00' 200.0 psf Values ofload densities are obtained froin Reference 13.

C. Normal Operatina Thermal Load (Ty These thermal loads are generated under normal operating or shutdown conditions. The following temperature data shall be used in the calculation of thermal gradients through walls and floor:

e Inside drywell temperature 135 F

e. Pool water temperature 150 F e Room temperature 60'F e- Outside ambient temperature:

Summer 100*F Winter 0F

- D. Accident Thermal Loads (Ta)

These thermal loads are due to the thermal conditions generated by the postulated accidents. The thermal accident temperature for the spent fuel pool is 212*F throughout the whole pool.

E. Seismic Loads (E. E')

E - = Seismic loads generated by Operating Basis Earthquake (OBE)

E' = Seismic loads generated by Safe Shutdown Earthquake (SSE) .

The SSE seismic loads are assumed to be twice that of the OBE seismic loads. The maximum accelerations for OBE are given in Figures 2-6 through 2-8. The vertical seismic effects shall be considered either upward or downward to result in the worst loading in the load combinations.

' " " " ^*'*""^" l'"* ' " Com onriins BVY 99ol15 / Attachment 2 / Page 6 The hydrodynamic loads of pool water acting on pool. walls shall be calculated in accordance with Chapter 6 of Reference (14).

The vertical seismic loads due to the refueling bridge and service platform and the vertical seismic loads distributed from the contiguous floors and beams shall be included in the analysis.

The seismic loads due to equipment mounted on pool walls shall be calculated using 150 percent of the peak floor response spectra.

The effects of.three components of earthquake shall be combined by the SRSS method.

Amplified response spectra provided by YAEC shall be used to determine the appropriate acceleration values for computing seismic loads.

The effect of rack impact on the pool floor during a seismic event should also be considered in the analysis. These loads should be combined with the loads due to venical seismic floor acceleration.

2.3.3.4 Load Combinations The following load combinations shall be considered in the spent fuel pool analysis (Paragraph 3.0, Section 3.8.4 of NUREG-0800M:

A. Concrete Pool Structure

1. Service Load Conditions

' a. 1.4D + 1.7L

b. 1.4D + 1.7L + 1.9E

c. (0.75) (1.4D + 1.7L + 1.7T.)

d. (0.75) (1.4D + 1.7L + 1.9E' + 1.7T.)

c. 1.2D + 1.9E

2. Factored Lead Conditions

a. D + L + T + E'

b. D+L+T.

c. D + L + T. + 1.2SE'

Vr.iutoNT YANK .i; Naur.An Pown;n CultPOHATION BVY 99115 / Attachment 2 / Page 7 Where any load reduces the effects of other loads, the corresponding coefficients for that load should be taken as 0.9 if it can be demonstrated that the load is always present or occurs r.imultaneously with other loads. Otherwise, the coefficient for that load should be taken as zero."

Question 2f:

Expl tin how the interface between the liner and the concrete slab is modeled, and also, how the liner' anchors are modeled. Provide the basis for using such modeling with respect to how they.

accurately represent the real structure behavior.

Response 2f:

In the structural analysis to determine the SFP capacity, no credit was taken for either the pool liner, and its anchorage to the concrete structure, or the 11" thick grout pad.

Question 2g:

Provide the calculated governing factors of safety in tabular form for the axial, shear, bending, and combined stress conditions.

Response 2g:

The controlling element in the structural analysis of the SFP structure is shear in the slab for the load combination involving OBE loading. Based upon the maximum loading from the proposed Holtec rack configuration, the minimum factor of safety is conseivatively calculated to be 1.07.

With respect to flexure, the factor of safety in the slab is conservatively calculated as 1.03. Wall stresses are governed by the south wall of the SFP. The factors of safety based upon shear and )

flexure for the south wall are 1.46 and 1.30, respectively.

These factors of safety are summarized below, f 1

SFP Structural Element Factor of Safety Shetr Flexure Slab 1.07 1.08 Walls 1,46 1.30 (governed by south wall)

- Question 2h: l l

What is the maximum bulk pool temperature at a full-core offload during a refueling outage? If I the temperature exceeds 150* F, provide the following:

i) Describe the details of the SFP structural analysis including the material properties (i.e.,

modulus of elasticity, shear modulus, poisson's ratio, yield stress and strain, ultimate m___

r, VEn\1oNT YANht.E Nrci.EAn l' owr.n CouronAitoN BVY 99-115 / Attachment 2 / Page 8 stress and strain, compressive strength) used in the analysis for the reinforced concrete slab and walls, liner plate, welds and anchorage in the analysis.

ii) ACI Code 349 limits a concrete temperature up to 150 F for normal operation or any other long-term period. Provide technical justifications for exceeding the required temperatureof150 F.

Response 2h:

The maximum bulk pool temperature at a full core offload during a refueling outage is 150* F.

An evaluation of the SFP slab was also made for an accident temperature of 212 F. Based upon data extracted from published test results, the strength in the top one third of the slab was reduced by 20% to account for the increased temperature in excess of 150* F. The slab was still found to be acceptable. However, it is noted that scenarios that could result in spent fuel pool bulk water temperatures greater than 150 F are beyond our licensing Basis.

l l

l l

= 1 Vt;ltMoNT YANhl.I'. Nt'ce.i:An l'ower.u Coni onxitoN Docket No. 50-271 BVY 99-115 l

i i

Attachment I A j Vermont Yankee Nuclear Power Station i

Response to Request for Additional Information Regarding Spent Fuel Pool Storage Capt. city Expansion (Civil and Mechanical Engineering Considerations)

. Holtec Position Paper WS-115, Revision 1,3D Single Rack Analysis of Fuel Racks Proprietary Information Enclosed f

1 l

l

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