ML20195J824

Proposed Tech Specs Re Rev to Increase Spent Fuel Pool Storage Capacity

ML20195J824
Person / Time
Site:	Wolf Creek
Issue date:	05/15/1997
From:	WOLF CREEK NUCLEAR OPERATING CORP.
To:
Shared Package
ML20138L753	List: ML20195J795 ML20195J824
References
NUDOCS 9811250076
Download: ML20195J824 (8)

Text

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. Attachment 111 to ET 98-0092 Page1of1 Des 1gn Features 4.0 4.0 DESIGN FEATURES 4.3 Fuel Storage (continued)  ;

c. A nominal 4-036 inch center to center distance between fuel assemblies placed in the fuel storage racks:

(Burnap Domain fee Qn 7.a,nd 3 inu7p.

d. "= =[rtially spent fuel assemblies with a J discharge burnup in the '/cceptable siangF of ,

Figure 3.7.171 may be allowed unrestricted storage i 1nt M r.c.rs.g c::' : and i l

ceptabis fush stennes leemti nS yh-e@. . New or partially spent fuel asseeblies with a

_ discharge burnup in the */nacceptablejenge" of

- - C Fiaure 3.7.171 will be stored in Region 1.

'brnuP boa f*r Region '2. .e 3 se.cagg f.(location) 4.3.1.2 The new fuel storage racks are designed and shall be maintained with:

a. Fuel assemblies having a maximum U 235 enrichment of ight percent:

b.  % s 0.95 if fully flooded with unborated water, which includes an allowance for uncertainties as described in Section 9.1 of~the USAR:

c. 4 s 0.98 if moderated by aqueous foam. which includes an allowance for uncertainties as described in Section 9.1 of the USAR: and

d. A nominal 21 inch center to center distance between fuel assemblies placed in the storage racks.

4.3.2 Drainane res<.

The spent fuel poof is designed and shall be maintained to prevent inadvertent draining of the pool below elevation 2040 ft.

FThe. ca.aw. Ioadvn3 pit is dut ar+ained uth a. .sto. rage. ca.pu q =a tai to

+Snea a,nA no -ere. shatt ihan be m m nat 4.3.3 Canacity !

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The spent fuel pool is designed and shall be maintained with a storage y capacity limited to no no than +34+ fuel assemblies. -

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Ihmitn 6r Region 3 Storau3c." ,F Trigure. 3.'1.rt-i n% be. auoweA u.nre.tricted storage. in acceptable. be.t storage. (oca.tions , ea cm.px. A Region 7. t oca.ttom l

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9811250076 981120 4*0 2 5/15/97 PDR ADOCK 05000482 P PDR

l Enclosure to ET 98-0092 Page 1 of 35

) ADDITIONAL INFORMATION REQUESTED FOR THE TOPICS DISCUSSED DURING THE CCTOBER 14 AND 15,1998 MEETING l item 9 Figures 7.2.1,7.2.2 and 7.2.4 in the applications show an additional bar around l the top perimeter of a fuel rack. This is not in agreement with Figure 2.1.1 of the applications. What is the current configuration?

Resoonse 9 Figures 7.2.1,7.2.2,7.2.3, and 7.2.4 were provided primarily to indicate the orientation of impactor and target during drop scenarios. Figure 2.1.1 was l

included as a pictorial view of a " typical" rack structure, as indicated by the title of the figure. The schematic of the fuel racks in all of these figures was not intended I to be a stmeturally accurate depiction of any specific rack. For example, all of the figures show a 7x7 cell rack, which is not consistent with any of the rack sizes ,

to be used for this project. TITE figures are provided for illustrative purposes only as discussed with the staff during the October 14 meeting. The bars shown at the top perimeter of the racks in Figures 7.2.1,7.2.3, and 7.2.4 were leftover from generic material previously developed for other projects. No such bars are included in the design of any racks for this project. FSAR updates will not contain these actual figures. All FSAR and USAR updates will include technically accurate and factual information pertinent to the specific projects at Callaway and l Wolf Creek, respectively, item 13 i

Part 1: Submit a frequency calculation demonstrating the contribution of load from the first mode frequency is adequate without considering higher mode contribution. A beam solution with shear included is acceptable.

Provide the equations and results from the calculations. ,

Part 2: The equations for torsional stiffness, and the calculation results, should also be supplied.

l B.esconse 13 i

l The method for determination of the shear and torsional stiffnesses was

! developed in the early stages of the development of the DYNARACK software

. program. A detailed description of the formulation of these stiffnesses is included here as Attachment A. Computations of the shear deformation and torsional i

rigidity terms discussed in Attachment A are provided in Attachment B. i Determination of rack frequency values are performed in Attachment C.

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, Enclosure to ET 98-0092 Page 2 of 35 t

l ltem 26 f

Ensure weld calculations adequately address the evaluation of shearloads.  !

Expand the information provided in the table of Section 6.9.1 to include the weld i stress factors of all the welds identified in Table 6.9.1.  !

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Response 26 I I

As discussed during the October 15 meeting, there are three welds required to be evaluated for the storage rack modules. I The female pedestal weld to the underside of the baseplate is evaluated by a combination of ANSYS model results and hand calculations. These calculations consider the weld stresses developed from loads along all three orthogonal axes.

The information discussed during the October 15 meeting presented the actual computation technique. Therefore, shear welds are already included in the results reported in Table 6.9.1.

The rack cell wall to baseplate weld stresses are determined through conversion of the cell wall stressee using appropriate factors. The computations which were performed in Sechon 6.9.5.a and reported in Table 6.9.1 of the Holtec Licensing Report (HI-971769) did not include the shear stress components. These components may be conservatively included by direct summation of stress I factors R2, R7, and R6 before the conversion is performed to develop weld

. stress values. This provides a conservative estimate of the cell-to-baseplate weld stress for the following reasons:

1) To calculate the actual stress below, we will conservatively use (R2)(0.6)(cy), l whereas the correct shear stress is (R2)(0.4)(cy). The same comment ,

applies to the conversion of stress factor R7 into actual stress.  !

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2) The directional stresses associated with the normal stress or and the two shear stresses Tx, and Ty should be combined using SRSS instead of di ect summation.

The shear stress factors are small in comparison to the maximum OBE R6 stress factor of 0.442 and SSE R6 stress factor of 0.389. The maximum R2 and R7 values corresponding to the simulation with the maximum OBE R6 value are 0.069 and 0.064, respectively. The maximum R2 and R7 values corresponding to the simulation with the maximum SSE R6 value are 0.095 and 0.078, respectively. Reproducing the formulas given in Section 6.9.5 of the licensing report to include the two shear stress terms gives:

- (0.442+0.069+0.064)[(0.6) * (21300)]

• 2.144 = 15,755 psi for OBE and 4

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Enclosure to ET 98-0092 j Page 3 of 35 l

l (0.389+0.095+0.078)[(0.6) * (2.0) * (21300)]

• 2.144 = 30,798 psi for SSE These computed weld stresses are lower than the corresponding allowables of 19,860 and 35,748 psi for OBE and SSE, respectively. Therefore, the we6:is are shown to be adequate.

The cell-to cell welds are evaluated using hand calculations and documented using MATHCAD. The computations and results reported in Table 6.9.1 did not include the stress components from shear stresses in the rack. The shear stress may be mnservatively included by considering the maximum shear stress at the base of the rack, which can be computed from the maximum R2 and R7 stress l factors. The maximum R2 or R7 stress factors computed under any simulation )

are 0.069 and 0.095 for OBE and SSE, respectively. These stress factor values l can conservatively be converted to weld stresses similar to the method used i above by assuming that the values occur for both stress factors simulataneously. l (0.069+0.069)[(0.6) * (21300))

• 2.144 = 3,781 psi for OBE and (0.095+0.095)[(0.6) * (2.0) * (21300)]

• 2.144 = 10,412 psi for SSE These stresses must be combined with the stress resulting from the OBE and SSE maximum fuel impact loads of 403 and 840 pounds. The stresses from these loads may be computed by assuming that the load is equally distributed to each side of the cell wall along 1* of weld.

For OBE the stress from the impact load, conservatively applied to 1" of weld, would be (403/2)/(1/16*0.707*1) = 4,572 psi. Adding this stress to that obtained above from the shear loads gives a total stress of 8,352 psi, which is below the OBE allowable weld stress of 19,860 psi, as given in Table 6.9.1 of the licensing report. [WCNOC editorial Note: The 4,572 and 8352 figures were obtained after conservatively rounding-up values (eg.,403.03 pounds to 404 pounds.).

For SSE the stress from the impact load would be (840/2)/(1/16*0.707)=9,505 psi. Adding this stress to that obtained above from the shear loads gives a total stress of 19,917 psi, which is below the SSE allowable weld stress of 35,748 psi, as given in Table 6.9.1 of the licensing report. Therefore, the welds are shown to be adequate.

Item 27 Provide a calculation to address fluid coupling affects with regard to cell buckling. !

Ensure that a bounding load from either fuel impact or hydrodynamic pressure is used, include an explanation of which load is bounding for this calculation.

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, Enclosure to ET 98-0092 Page 4 of 35 l

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( Response 27 The requested calculation is provided in Attachment D and is titled " Elastic l Stability of a Storage Module Cell Wall." The calculation shows that the l maximum compressive outside wall loading, from the computed hydrodynamic l pressure of 11.1 psi, does not produce buckling on any intemally supporting walls. Fuel impacts obviously do not occur on the exterior walls of the storage i module and will be shared by a number of rack interior cell walls with at least two l walls experiencing tension. Therefore, cell buckling from fuel impacts loads is not a concem.

Item 29 l Perform and submit a hot cell calculation to account for the thermal loading on I

the racks or provide a reference to a previously submitted hot cell calculation that would be applicable.

Response 29 l The thermal stresses developed in the cell due to the restriction of thermal expsnsion are discussed in Attachment E. The description provides the computation specific to the Callaway and Wolf Creek project, which shows that l

the thermally induced stresses are acceptable.

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Item 30 What is the maximum vertical force developed in t% support pedestal resulting l from the deep drop of a fuel assembly into a comer cell?

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Response 30 l

l The intention of the deep drop evaluation is primarily to determine the damage to the pool liner and underlying concrete not to evaluate pedestal loads. To conservatively determine loads to these pool structure components the LSDYNA3D model is prepared with very conservative assumptions. For example, the fuel assembly and female pedestal are considered to be rigid.

i Additionally, the fuel assembly velocity at impact is calculated by neglecting the j increased fluid pressures experienced as the assembly reaches the bottom of a j cell located over a pedestal location. The limited flow hole cross sectional area  !

for the cells over pedestals represents significant friction for the outflow of water.

These conservatisms combine to produce a pedestal load, which is much larger

! than would actually be experienced due to an assembly impact. It is recognized

in recent dry fuel storage project work that fuel assemblies can only withstand 1 approximately 63g of deceleration due to impact. Dedileration values above this

level result in collapse of the fuel rods and destruction of the fuel assembly 4

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, Enclosure to ET 98-0092 1 Page 5 of 35 I j

l integrity. The worst case pedestal load from the rack seismic evaluations is j reported as 291,000 lbs. This load would represent a deceleration of about 1779

for an assembly weighing 1647 lbs, if it were applied directly to the fuel. This j deceleration far exceeds the values, which a fuel assembly is capable of  !

i withstanding. Therefore, it is known that the real fuel assembly drop scenario l where impact loads are imparted to the fuel could not impart a load to the

! pedestal that exceed that calculated from seismic action.  ;

d i With all of the conservatisms included, the maximum vertical force developed in

the pedestal during the deep drop over pedestal scenario is computed to be 1.05E+06 lbf, which occurs at about 1.8E-03 seconds after initial contact of the assembly with the baseplate. This conservatively computed force is shown to be

adequately resisted by the rack structure, bearing pad, liner, and underlying concrete.

The numerical analysis of the bounding impact postulated during the " deep" drop over pedestal event shows that the duration of contact between the 2250 lb impactor and the bearing pad is only approximate 0.0028 seconds. The 1/4 inch pool liner is not pierced during the collision, since the maximum Von Mises stress of about 25 ksiis less than the failure stress of 71 ksi for the liner material. A very high Von Mises stress (54 ksi) is observed in the pedestal cylinder at the contact surface with the pad. However, this value is below the failure stress of 140 ksi for this component. The bearing pad registers a maximum Von Mises stress of 22 l ksi, located in the area where the contact with the pedestal takes place. The concrete stratum directly beneath the pedestal sustains a very localized (peak normal) compressive stress with a maximum amplitude of 25.6 ksi resulting in localized damage. This calculated localized maximum stress exceeds the concrete failure stress. Thus localized concrete crushing would be experienced which is self-limiting, since the failure strain (5.5E-02) of the region is not exceeded by the maximum calculated element strain value of 6.624E-03. LS-DYNA 3D has the capability of removing components from the model which experience stresses and/or strains above the failure values. All of the modeled concrete elements remain in the finite element model throughout the solution.

Therefore, the concrete experiences only localized crushing. The rest of the modeled area is in tension but the stress is only 108 psi, which can be supported by the concrete without cracking.

We apologize for any confusion that may have occurred from the statements I' made in section 7.4.2 of the license amendment report. In retrospect it is apparent that these statements, although true, may be misleading because of the '

terse manner in which the topic is covered. Hopefully, the more complete discussion provided above clarifies the intent of the statements made in the i licensing report. In future documents, our discussion will be expanded to i J

preempt this question.

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- Enclosure to ET 98-0092 Page 16 of 35 ATTACHMENT B Then the effective shear deformation coefficient for the assemblage of individual fuel cells is *

(1 + 4,; It

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$2 ,.3 $ 2 = 181.73 l t star "+ H A2 - $'2 B2 x 3 2 j The effective shear deformation coefficient, for the full length composite section, is 4 star = 3.528 i

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Enclosure to ET 98-0092 Page 17 of 35 ATTACHMENT B DYNARACK TORSIONAL INERTIA PROPERTY CALCULATION For the effective torsional rigidity of a 13 x 13 rack d

J g = ll3915 in Taken from the rack seismic structual calculation package For the torsional rigidity of an individual box around its own center, we use the formula for a hollow tube from Timoshenko and Goodier, Theory of Elasticity,3rd Edition, McGraw-Hill 1970, p.333.

t ~ = 0.075 in c = 8.77 in + t Ja = 169-(c tl 3

J a= 8.771* 103 *in' The contribution from the individual boxes acting like clamped beams to provide an overall resistance is

,4 Calculated from values internally generated 3

Jb -5.06210 *in by DYNARACK pre-processor.

Jb := 8925 1.763 ,

where we have added in the effect of shear deformation on an individual tube. The total torsional l moment of inertia for the assemblage of end-connected boxes is 4 d l j J 23 a+J b J 2 = 1.38310 *in c Therefore, we have the composite moment of inertia as '

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J' 4 J = 1.84310 *in' J.=

31 A2+J-B2} 2 l

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