ML17301A117

Forwards Addl Info Re Spent Fuel Rack Submittal & Summary of 840816 Meeting Re Fuel Rack Structural Analysis,In Response to 840713 Request

ML17301A117
Person / Time
Site:	Saint Lucie
Issue date:	08/29/1984
From:	Williams J FLORIDA POWER & LIGHT CO.
To:	John Miller Office of Nuclear Reactor Regulation
References
L-84-219, NUDOCS 8408310145
Download: ML17301A117 (120)

Text

II REGULATOR NFORMATION DISTRIBUTION S EM (RIDS)

ACCESSION NBR 84083101/5 DOC ~ DATE! 84/08/29 NOTARIZED'O DOCKET FACIL,'50, 389 St-, Lucia Planti, Unit '2i Florida Power 8 Light Co. 05000389 AUTH's NAME'UTHOR AFFILIATION HILLIAMSiJeH. Flor ida>> Power 8 Light Co, RECIPE NAME'ECIPIENT AFFILIATION MILLER' ~ Re Operating Reactors Branch 3 r v,

• info r e'pent fuel'rock submi t tal L summary,of V P ~0 >>

SUBJECT:

For wards, addi 840816'..meeting re'uel~ rack structural anal ysisiin response to 840713'request ~

DI'STRIBUTION CODE: A001D COPIES RECEI/ED:LTR 'NCL .

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. BOX 14000, JUNO BEACH, FL 33408 Qll//y FLORIDA POWER & LIGHT COMPANY August 29, 1984'-34-2I9 Office of Nuclear Reactor Regulation Attention: Mr. James R. Miller, Chief Operating Reactors Branch 83 Division of Licensing U. S. Nuclear Regulatory Commission Washington, D. C. 20555

Dear Mr. Miller:

Re: St. Lucie Unit No. 2 Docket No. 50-389 Request for Additional Information S ent Fuel Rack Submittal In response to your letter of July l3, l984, the additional information you requested is attached.

Also attached is a summary of the meeting held on August l6, l984, with Franklin Research Center, NRC, Combustion Engineering and FPL, regarding the fuel rack structural analysis.

Very truly yours, J. W. Williams, Jr.

Group Vice President Nuclear Energy JWW/R JS/law Attachments 84083k 0145 05000389 840829'DR ADOCK P PDR PEOPLE... SERVING PEOPLE P

ATTACHNENT 1 CORE PERFORHANCE BRANCH QUESTIONS/RESPONSES

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NRC Item No. I:

The statement regarding dry storage of new fuel has been deleted from Technical Specification 5.6.l. Where will fresh (unirradiated) fuel be stored for each reload core and in what condition (wet or dry)?

FPL Res onse:

The statement regarding dry storage of new fuel was applicable for the first core loading stored dry in the spent fuel pool, and therefore, it is no longer applicable.

However, Technical Specification 5.6. lb. (see attached page 5-4) should be added for dry storage of unirradiated fuel in the new fuel storage racks.

i I

Fresh (unirradiated) fuel for reload cores will be stored dry in the new fuel storage racks or wet in Region I of the new spent fuel storage racks.

The new fuel storage area is shown in Figures l.2-I8 and l.2-I9 of the St. Lucie Unit 2 FSAR. The new fuel storage racks are shown in Figure 9. I-I of the FSAR.

The new fuel storage racks are designed to store eighty-l6XI6 fuel assemblies containing fuel of up to 4.5 w/o enrichment while maintaining kef f < 0.98 under the most reactive condition. Refer to the attached analysis.

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FPL Res onse to NRC Item No. 1:

DESIGN FEATURES VOLUME 5.4.2 The total water and steam volume, of the reactor coolant system is l0,93 I + 275 cubic feet at a nominal Tavg of 572 F.

5 5 METEOROLOG(CAL TOWER LOCATION 5.5.I The meteorological tower shall be located as shown on Figure S.l-l ..

5 6 FUEL STORAGE CRITICALITY 5.6.I a. The spent fuel storage racks are designed and shall be maintained witht I. A keff equivalent to less than or equal to 0.95 when flooded with unborated water, which includes a conservative allowance of 0.024 hkeff for Total Uncertainty.

2. A nominal 8.96 inch center-tormenter distance between fuel assemblies placed in the storage racks.

3. A b'oron concentration greater than or equal to 1720 ppm.

Region I can be used to stare fuel which has a U-235 enrichment less than or equal to 4.5 weight percent. Region II can be used to store fuel which has achieved sufficient burnup such that storage in Region I is not required. The initial enrichment vs. burnup requirements of Figure 5.6-1 shall be met prior to storage of fuel assemblies in Region II.

b. 'Ihe new fuel storage racks are designed for dry storage of unirradlated fuel assemblies having a U-235 enrichment less 'than or equal to 4.5 weight p rcentE while maintaining a keff of less than or.

equal to 0.98 under the most reactive condition.

DRAINAGE 5.6.2 The spent fuel storage pool is designed and shall be maintained to ~

prevent inadvertent draining of the pool below elevation 56" feet.

CAPACITY 5.6.3 The spent fuel storage pool is designed and shall be maintained with a storage capacity limited to no more than l 076 fuel assemblies. I 5.7 COMPONENT CYCLIC OR TRANSIENT LIMITS 5.7.l The components identified in Table 5.7-I are designed and shall be maintained within the cyclic or transient limits of Table 5.7-I.

ST. LUCIE - UNIT 2 Amendment No.

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FPL,Res onse, to t/RC I m No; 1:

1 ~ 0 Introduction The existing new fuel. storage rack in St. Lucie 2 was reanalyzed to determine ifThefuel assemblies enriched to 4.5 wt$ U-235 could be stored in the racks. previous licensed limit was 3.7 wt$ U-235 (Reference II-1).

2.0 Desi n Bases The new fuel storage racks are designed to:

(a) store 80 16x16 fuel assemblies containing fuel of up to 4.5 w/o enrichment, (b) provide suf ficient spacing between the fuel assemblies to maintain a subcritical (Keff < 0.98) array assuming the most reactive condition.

3 ' S stem Descri tion The location of the new fuel storage racks is shown on the Fuel Handling Building general arrangement drawings, Figures 1.2-18 and 1.2-19 of Reference II-1. The new fuel storage racks are shown in Figures 9.1-1 of Reference II-1.

The new fuel storage racks consist of 80 square cavities fixed together in two 4 by 10 arrays. The design data for the racks is listed in Table 3-1.

Provisions are'ade for the future storage of 16 additional fuel assemblies. Each storage cavity is fabricated from four stainless steel angles connected by horizontal ties and each can contain one new fuel assembly. Each cavity is provided with a hinged checkered plate cover.

4 ' Method of .Anal sis The method of analysis used is a four group infinite array DOT (Reference II-2) transport analysis with the thermal group cross sections obtained by considering the spatial arrangements of the components. 'In order to calculate the effect of leakage on the eigenvalue, cases have been reanalyzed using the 3-D KENO code (Reference II-3). The macroscopic cross sections using in the KENO input were obtained from DOT.

New fuel in the storage racks has been analyzed for uranium enrichments up to 4.5 weight percent and with consideration of 'a full range of moderator (non-borated water) densities from mist to full immersion. Assumptions for the cr iticality analysis is shown in Table 4-1. Credit was taken for the corner angle irons which are shown in Reference II-1 on Figure 9.1-1.

5.0 Results The resulting eigenvalue for the new fuel rack for 4.5 wt% U-235 is 0.8707

+ 0.0058 for the full flooded condition and 0.7952 + 0.0083 for the optimum mist condition. These results show sufficient margin to cover all other uncertainties such as rack dimension and material tolerances.

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6.0 References lI-1. St. Lucie Nuclear Power Plant Unit Two, Final Safety Analysis Report in Support of Docket No. 50-380, License No. NPF-16.

II-2. "DOT-II M Discrete Orindates Transport Computer Code," MANL-THE-1982.

II-3. "KENO IV Hultigroup Honte Carlo Criticality Code Syst: em," ORNL-4938.

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TABLE 3-1 ASSUMPTIONS FOR CRITICALITY ANALYSIS FOR NB( FUEL RACKS

1. Enrichment 4.5 w/o U-235

2. Fuel Fresh and Non-depleted

3. Burnable Poison Rods No Credit Assumed

4. Control Rods No Credit Assumed

5. Minor Structural Hembers No Credit Assumed

6. Moderator Effects of full range of water densities (mist conditions) considered 1-5

TABLE 4-1 NEW FUEL STORAGE RACKS Number of Square Cavities 80 Size of the Cavity 8-11/16 inches square Center to Center Spacing Between Cavities 23 inches Separation Between Arrays 69 inches 1-6

4; NRC Item No. 2:

Since storage in Region II is dependent on initial enrichment and burnup, Figure3-4 showing allowable burnup for storage in Region II should be included as a Technical Specification figure.

FPL Res onse:

Figure 3-4, showing allowable burnup for storage in Region II, is included with the administrative procedure for evaluating burnup and determining placement of fuel in either Region I or Region II. This figure should be included in the Technical Specifications as Figure 5.6-I. Also Technical Specification 5.6.I should be changed to include reference to Figure 5.6-l (see attached pages 5-4 and 5-4A).,

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FPL Res onse to NRC Item No. 2:

DESIGN, FEATURES VOLUME 5.4.2 The total water and, steam volume of. the reactor. coolant system. Is:

I0,93l + 275 cubic feet at a nominal~ Tavg of 572. F".

5.5 METEOROLOGICAL TOWER LOCATION.

S.S.I The meteorological tower shall be located as shown on Figure S.l-l. ~

5.6 FUEL STORAGE CRITICALITY 5.6.I a. 'Ihe spent fuel storage racks are designed and shall be maintained with:

I. A keff equivalent to less than or equal to 0.95 when flooded with unborated water, which includes a conservative allowance of 0.024 hkeff for Total Uncertainty.

2. A nominal 8.96 inch center-to-center distance between fuel assemblies placed in the storage racks.

3. A"l boron concentration greater than or equal to 1720 ppm.

Region I can be used to store fuel which has a U-235 enrichment less than or equal to 48 weight percent., Region II can be used to store fuel which has achieved sufficient burnup such that storage in Region I is not required. The initial enrichment vs. burnup requirements of Figure 5.6-I shall be met prior to storage. of fuel assemblies in Region II.

b. The new fuel storage racks are designed for dry storage of

'unirradlated fuel assemblies having a U-235 enrichment less- than or equal to 48 weight percent, while maintaining a kef f of less than or equal to 0.98 under the most reactive condition.

DRAINAGE t

5.6.2 The spent fuel storage pool is designed and shall be maintained to prevent inadvertent draining of the pool below elevation 56 feet.

CAPACITY 5.6.3 The spent fuel storage pool is designed and shall be maintained with a storage capacity limited to no more than I076 fuel assemblies.

5.7 COMPONENT CYCLIC OR TRANSIENT LIMITS 5.7.1 The components identified in Table 5.7-I are designed and shall be maintained within the cyclic or transient limits of Table 5.7-I.

ST. LUCIE - UNIT 2 Amendment No.

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FPL Res onse to >1RC Iten No. 2:

40,000 I

30,000 CI BURNUP REQUIRED FOR STORAGE IN REGION II tel ED

+

lC 20,000 W

BURNUP (lli200) X (ENRICHHENT M/0)

-15i900 w 10,000 1.5 2.0 28 3.0 38 4.0 486,0 INITIAL U 235 ENR ICHHENTg N/0 FIGURE 5o6-1 INITIAL ENRICHHENT YS BURNUP REQUIREHENTS FOR STORAGE'OF FUEL ASSEHBLIES IN REGION II STe LUCIE - UNIT 2'-4A AHENDHENT NOo 1-9

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NRC ITEM NO. 3:

Describe the calculational cells used for criticality analysis of Regions I and II.

~FPL P For both Regions I and II, an infinite array of storage cells was assumed to exist in the directions transverse to the axis of the fuel assembly. In addition, an infinite length active fuel region was assumed. Thus, no credit for neutron leakage from the spent fuel rack was taken. The nominal center-to-center spacing of adjacent fuel cells was taken to be 8.965 inches." The nominal thickness of the stainless steel monolith wall structure was taken to be O.I88 inches. The attached figure shows the geometry employed in the DOT calculations for Regions I and II. The Region I calculations employed periodic boundary conditions because of the asymmetric L insert in each fuel storage part. For Region II, two types of calculational cells were employed. Type a was used for the majority of the analyses because of the savings in computer resources and.it is a conservative approximation to the actual geometry illustrated as type b. Selected analyses were run at different burnups for fuel assemblies of different initial enrichments to verify that the type a calculational cell is a conservative approximation to that of type b. The difference in infinite multiplication factor is between 0.00I5 and 0.0020.

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K FPL Res onse to NRC Item No. 3:

REGION I PERIODIC B. C. - 4 SIDES 1/2 MONOLITH WALLTHICKNESS FULL MONOLITH WALLTHICKNESS FUEL ASSEMBLY L-INSERT REGION II (SEE TEXT)

TYPE a

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TYPE I3 FUEL ASSEMBLY I

I MONOLITH WALLS I I I I

I I I L

REFLECTING B,C, ~ I

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NRC Item No. 4 What temperature was used for the design basis condition to yield the highest reactivity? How does keff vary with pool temperature over the range of normal and abnormal temperature conditions'?

FPL Res onse:

Design analyses were run at a nominal pool temperature of 98.6 F. In addition, analyses were run 'at l49, l97.6 and 248oF to determine the temperature dependence of the multiplication factor of the stored spent fuel. The difference between the multiplication factor at 98.6 F and the maximum value over the temperature range was included in the evaluation of the calculational uncertainty. For Region I, the maximum multiplication factor occurred at 248 F and it was approximately 0.00l higher than the nominal condition value. For Region II, the maximum multiplication factor. again occurred at 248 F and it was approximately 0.005 higher than the nominal condition value.

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NRC Item No. 5 Describe and quantify all the uncertainties applied to the calculated k ff for both regions. Verify that these uncertainties are at least 95/95 probabilityfconfidence level values.

~FPL R Uncertainties applied to the calculated ke ff for both regions were evaluated in the following manner. First, an uncertainty and calculational bias were evaluated for the basic calculational model (see response to question 6); the results of this evaluation indicated an uncertainty'at the 95/95 confidence level of 0.007l and a calculational bias of +0.00I4; that is, the calculational model overestimated the mean multiplication factor by the latter amount.

Next, calculations were carried out to evaluate the contributions of mechanical tolerances, off-center placement of fuel assembly in adjacent storage cells, and maximum temperature of spent fuel and coolant. The results of these calculations are summarized in the following table.

Chanche ~Re ion I ~Re ion Il Minimum center-tormenter +0.0034 +0.0046 distance of storage cells Minimum L-insert thickness +0.0015 Minimum monolith wall thickness -0.0005 +0.00 I 8 Off center placement of fuel assemblies in adjacent cells +0.0 I 80 -0.0094 Temperature Change . +0.00 I 0 +0.0044 R.M.S. Value 0.0 I 84 O.OI l5 The individual contributions were combined in a root mean squared manner to yield the values shown as the last entry in the table. When the latter values were then combined in a direct additive manner with the bias .and calculational uncertainty, noted above, an overall uncertainty of 0.024 and 0.0I7 was obtained for Regions I and II, respectively. All components in the overall uncertainty'are at least 95/95 confidence level.

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0 NRC Item No. 6:

A calculation bias is mentioned which has been determined from comparisons between calculations and experiments. Specify what experiments were used and justify their adequacy in representing the characteristics of the St. Lucie spent fuel pool. What organization performed these benchmark calculations?

FPL Res onse:

The following, "Qualification of Analytical Methods Used ln Spent Fuel Storage Rack Analyses", summarizes the analysis of a broad range of experiments to qualify the calculational model. Table II includes experiments which are very relevant to the St. Lucie spent fuel storage racks.

1-14

FPL Res onse to NRC Item No. 6 QUALIFICATION OF ANALYTICAL METHODS USED IN SPENT FUEL STORAGE RACK ANALYSES I. ~Pur oie The purpose of this memo is to provide qualification of the calculational model and evaluation of calculational uncertainties and/or bias factors

'used in analyzinq spent'uel storage racks, especially the HI-CAPT" racks employing steel boxes and super HI-CAPs containing boron carbide poison.

This is based on the analysis of a variety .of reactor and laboratory experiments. The methods of cross..section generation are essentially those of C-E's physics design procedures modified appropriately for use in four group transport, discrete ordinate method criticality calcula-tions, and Honte Carlo codes.

II. Calculational Uncertaint and Bias The results of the analysis of.a series of U02 critical experiments are summarized in Table I. These are calculated using the methods described by Gavin (Reference 1} for CEPAK 2.3, which is used fn present storage rack calculations. Table I includes the mean and standard deviation for this CEPPK model.

Although the spatial. solution. for the flux distribution was obtained by use of a diffusion theory code such as PDg-7, transport corrections. for the reflector and heterogeneous lattice effects were employed. Thus, for example, in Reference 8 the 4.3 w/o infinite lattice of close packed assemblies in room temperature water had a Keff of l.4547 in PDg and 1.4568 in DOT, the conservative bias in DOT of 0.0021 will be ignored.

These calculations support use of the differential cross-section data base and broad group cross section generation codes.

Since fuel storage arrays do involve the spacing of the fuel assemblies at larger separation distances than in'ypical PMR reactor lattices, the pre-dictive capability of the calculational model was tested on the following experiments. In these analyses done for this memo, the spatial flux solution was obtained directly with the transport code, ANISN. To assess the accuracy of the calculational model in predicting the multiplication factor of fuel assemblies having a separation distance sufficiently large so as to be isolated, analyses were carried out for a group of subcritical, 1-15

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exponential experiments on clusters of 3.0 w/o U02 fuel pins clad with type 304 S.S. and moderated by H20 (page 165 of Reference 9). The cluster sizes analyzed vary from 181 to 301 fuel rods so.as to encompass the range typical of current PhR fuel; assemblies. The multiplication factorsof'izes for the lattices analyzed using axial bucklings deduced from the reported re 1 axa ti on 1 eng ths a re tabu1 a ted be 1 ow.

No. of Fuel Rods Keff 181 0.9966 211 1.0011 235 0.9966 265 0.9988 301 0.9984 These results indicate that the calculational model predicts the multiplica-tion factor for small clusters of fuel rods in a water environment to a high degree of accuracy, i.e. a bias of -.0017.

To ascertain whether the calculational mode can predict the reactivity characteristics of thick stainless'teel plates and boron poisoned plates an analysis (Reference 10) was made of PNW experimental (Reference ll) critical separations of 2.35 w/o U-235 U02 subcritical clusters. The results using the Honte Carlo code KENO IV are shown in Table II.

Method of Calculation "The calculation methods for these experimental covparisons which are als'o used to determine reactivity for fuel rack storage, fuel shipping containers plus other fuel configurations found in fuel manufacturing areas are based on CEPAK 2.3 (Reference 1) cross sections. Using an appropriate buckling .

value and taking proper account of resonance absorption, three fast groups

.are collapsed from 55 fine energy mesh groups in FORM and the one'hermal group is collapsed, rom 29 thermal energy groups in THERMOS. In addition, each component such as water gap, or poison plate has its thermal

. cross section determined by a slab THERMOS calculation employing the proper fuel environment. FORt1 ar d THERMOS are sub-prcgr ms of CEPAK.

For one dimensional analyses such as the BNL exponential experiments the discrete ordinates code ANISN (Reference 12) is used. For two dimensional analyses OOT-2M (Reference 13) is used. For three dimensional analyses (such as the critical separation experiments] KENO IV (Reference 14) is used.

Results The above analyses indicate a bias between. predicted mean and measured multiplication factors of +.00t38 and a calculational uncertainty of

.00714 at the 95/95 confidence level for the complete series of UO<

experiments. 4 Pa 1-16

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Thus, using CEPAK 2.3 cross sections we conclude the following-Total Number of Results 41 Mean Value (u) 1.00138 Standar'd Oeviation = 6 0.00337 0 Multiplier or 95/95 confidence 2.118 95/95 Confidence Level Uncertainty 0. 00714 Bias (u -1.0) +.00138 Uncertainty Minus Bias .00575

~

It wi11 be noted that the seven no boron steel cases have a bias of 0.00207 .

(i.e. the calculated value is .00207 greater. than the critical keff value of unity) which is greater than the mean bias. The three boral cases have a bias of -0.00435 with unity particle'self-shielding factor for the B4C.

Because of the size and distribution of the boron carbide particles the boron allows more transmission than an equivalent homogeneous boron carbide mixture.

Neutron transmission exoeriments conducted by the University of Michigan for Brooks & Perkins Inc. (Reference 15) are, consistent with using a 0.9 sel-shielding factor in the third of four CEPAK neutron group and a 0.75 self-shielding factor in the thermal group. These self-shielding factors which are used in designing boron containing fuel racks make the bias for these.

boral cases +0.00008.

References:

P. H. Gavin,"CEPAK 2.3 Mod 0," C-E Internal Report, 12/14/76.

2. T. C. Engelder, et al, "Spectral Shift Control Reactor, Basic Physics Program," B&W-1273, flovember 1963.

3.

R. H. Clark, et al, "Physics Yerifjcation Program Final Report,"

B&W-3647-3, triarch 1967.

E

4. P. W. Oavison, et al, "Yankee Critical Experiments," YAEC-94, April, 1959.

5. W. J. Eich and W. P. Rocacik, "Reactivity and Neutron Flux Studies in Multi-Region Loaded Cores," WCAP-1443, 1961.

6. F. J. Fayers, et al, "An Evaluation of Some Uncertainties 'in the Comparison Between Theory and Experiments for Regular Light Water Lattices, Brit. Nuc. En. Soc. J., 6, April 1967.

7. J. R. Brown, et al, "Kinetic and Buckling Measurements on Lattices of Slightly Enriched Uranium and U02 Rods in Light Water," WAP0-176, 1958.

8. J. Handschuh, L. C. Noderer, R.C. for "Compact Spent Fuel Storage Criticality Analysis for Arkansas Power and Light, Unit 2 at 68',"

C-E Internal Report, April 8, 1975.

0

9. G. A. Price, "Uranium - Water Lattice Compilation Part I, BNL Exponential Assemblies," BNL-50035 (T-4"9), Oecember 1966.

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J 1

lp. L. C. Noderer, "Analysis of Critical Separation of Low Enriched Subcritical Clusters," C-E Internal Report, May 11, 1979.

11. S. R. Bierman, E. D. Clayton and B. M. Durst, "Critical Separation Between Subcritical Clusters of 2.35 w/o U-235 Enriched U02 Rods in Mater with Fixed Neutron Poisons," PNL-2438, October 1977.

12. Mard.M. Engle, Jr., "A Users Manual for ANISN, A One Dimensional Discrete Ordinates Transport Code Mith Anisotropic Scattering K-1693; March 30, 1967.

13. R. G. Sottesy, R. K. Disney, A Collier, "User's Manual for the DOT-IIM Discrete Ordinates Transport Computer Code," MANL-TME-1982, December 1969.

14. L. H. Petrie and N. F. Cross, "KENO IV, An Improved Honte Carlo Criticality Program," ORNL-4938, November 1975.

15. James M. Bryson, John C. Lee and R. Robert Burn, "Neutron Transmission Through Boral Shielding ttaterial: Theoretical Model and Experimental Comparison," University of tlichigan, Dept. of nuclear Engineering, tlichigan Memorial-Phoenix Project, prepared for Brooks tt Perkins, Inc. April 1978.

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TABLE I Results of Analysis of Critical. U02 Systems tlo. Q2 Lattice tot Blitt (2) I .08-2 1.00121 2 II .172-2 1.00534 3 'X

.79-2 .99838

,XI I I .701-2 1.00419 5 XX .. 202-2 1.00550 6 BaW (3) .861-2 1.00269 7 2 .420-2  : 1.00443 8 Yankee (4) 1 .408-2 '.00088 9 2' .531-2 1.00115 10 ~

.633-2 1.00136 11 Yankee (5) 4 .688-2 1.00244.

Minfrith (6) 12 ~

Rl -.20 .660-2 1.00214 13 'Rl-80 .626-2 .99942 14 R3 .510-2 1.00422 15 Bettis(7) 1 .326-2 1.00053 16 2 .355-2 l.ooo46 17 3. .342-2 1.00106 Average 1.00208

+.00206

• Using calculat ated radial bucklings and measured axial bucklings.

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TABLE II Calculated keff Values For Separation Experiments t"onte Carlo Expt f Type Poison Plate Keff 6'(STD Deviation) 15 None 1.00227 .00534 o4 None 0.99912 :0054n 49 'one 1.00221 .00473 18 None 1. 00813 .00489 21 None 0.99589 .00461 28 304 S Steel 0.0 w/o Boron 1.00393 .00308 05* 304 S Steel 0.0 w/o Boron 1.00329 .00303 29 304 S Steel 0.0 w/o Boron 1.00271 .00302 27 304 S Steel 0.0 w/o Boron l.on418 .00273 26 ~

304 S Steel 0.0 w/o Boron 0.99811 .00279 34 304 S Steel 0.0 w/o Boron 0.99793 .00297 35 304 S Steel 0.0 w/o Boron 1.00436 .00290 32 304 S Steel 1.05 w/o Boron . 0.99970 33 304 S Steel- 1.05 w/o Boron 1.01173

.00524'00491 38 304 S Steel 1.62 w/o Boron 1.00289 .00512 39 304 S Steel 1.62 vi/o Boron 1.00208 .00506 20 Boral 0.99585 . 00301 16 Boral 1'.00020 .00288 17 Boral 0.99519 .00286 Mean Keff Value 1.00157 Std. deviation .00419 1-20

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NRC Item No. 7 Describe in detail how the allowable burnups as a function of initial enrichment were determined for Region II.

~FL tl The discussion in'he response to question 8 describes how fuel nuclide cross sections were determined for the DOT analyses of Region II; the response to question 3 describes the calculational geometry employed in these DOT calculations. The resulting multiplication factors, after upwards adjustment to include calculational uncertainties and biases (net value of +0.0 I 72 in ke ff) yield a family of curves of keff versus burnuP for a range of enrichments as shown in Figure 3-3 of the St. I ucie 2 Spent Fuel Pool Rerack Safety Analysis Report. The curves can be used to define the minimum burnup, for fuel of a given initial enrichment, which will result in a an'effective multiplication factor of 0.95 when Region I I is ful ly loaded with fuel assemb lies of this type. These burnup/enrichment data points can then be used to plot a curve of burnup versus initial fuel enrichment similar to Figure 3-4 of the subject report.

'The latter data is applicable to fuel of a given average burnup since it is derived by planar DOT analyses. Since discharged fuel assembilies will exhibit a characteristic trend in the planar average burnup distribution along the longitudinal axis of the fuel assembly, KENO calculations were run for an infinite array of individual storage ports containing fuel with one of two typical axial burnup distributions. One distribution was typical of fuel discharged after one burnup cycle and the other was typical of fuel discharged after three burnup cycles. The former distribution resulted in a smaller multiplication factor than for the case of a flat axial burnup distribution equivalent to the average of the shaped burnup distribution whereas the axial shape typical of fuel discharged after three cycles resulted in a higher multiplication factor than the equivalent flat axial distribution. Although the difference in multiplication factors was only 0.0I I, a burnup decrement of l900 MWD/MTU was applied to the final version of Figure 3-4 of the subject report to assure conservatism in the'riterion for discriminating between fuel assigned to Regions I and II.

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NRC Item No. 8:

Discuss and verify the accuracy of the burnup dependent isotopics used in the Region II analyses. Include a discussion on the ability to predict actinide depletion, fission product accumulation, and the applicability of the calculations for long-term storage.

FPL Res onse:

The CEPAK lattice program is described in the Analytical Methods section (4.3.3) of the St. Lucie 2 FSAR. Figure 4.3-52, of the FSAR shows good agreement between the predicted and measured plutonium isotopic concentrations for a Yankee-Rowe asymptotic fuel pin. Fission products are calculated by the CINDER module in CEPAK. Fourteen fission product chains representing the more important fission products are used instead of the 69 chains employed in WAPD-TM-334+; the difference in fission product poisoning between the condensed and detailed fission product chain is treated as a lumped fission product for each fissioning nuclide.

The possible reactivity changes in the fuel subsequent to removal of the fuel from the reactor for storage in the spent fuel racks are calculated in the following manner. At various time points during the CEPAK full power depletion calculation, nuclide concentrations are computed after a 3000 hour0.0347 days <br />0.833 hours <br />0.00496 weeks <br />0.00114 months <br /> cooling period. The cooling period is simulated by a nearly zero power operating Jevel in

'CEPAK. A restart calculation is then initiated at the temperature conditions appropriate to thespent fuel pool .and with the fuel nuclide concentrations appropriate to the end of the 3000 hour0.0347 days <br />0.833 hours <br />0.00496 weeks <br />0.00114 months <br /> cooling period to calculate the microscopic cross sections for use in the DOT analyses of the effective multiplication of Region II. Thus, these analyses include appropriate changes in fission products and key actinides for long term storage.

+T. R. England, "CINDER - A Point Depletion and Fission Product Program,"

WAPD-TM-334. Revised June I 964.

1-22

FPL Res onse to NR tern No. 8:

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NRC Item No- 9:

With reference to the DOT-2W code, what order of angular quadrature was used, how many energy groups were assumed, and what order of Legendre expansion was for anisotropic scattering? 'sed

~FPL ~

The order of angular quadrature employed in the DOT-2W code was S-6 and the order of Legendre expansion was P>. These are the same approximations 'as employed in the anlaysis of those critical experiments (see attachment referred to in response to question 6) which employed DOT for the spatial solution. The number of neutron energy groups was four.

1-24

NRC Item No. IO:

Since administrative controls are used to evaluate the burnup of each spent fuel assembly and determine its placement in either Region I or II, we will require the Technical Specifications to include requirement of an independent check of burnup and storage. allocation as well as complete records of fuel assembly burnup analysis during the entire onsite storage period.

FPL Res onse:

FPL is developing procedures which will require an independent verification of burnup and storage allocation of fuel stored in the spent fuel racks. These procedures will also specify records requirements. The procedures will be available for NRC review prior to placement of any spent fuel in the spent fuel racks.

I -25

ATTACHNENT 2 STRUCTURAL AND GEOTECHNICAL ENGINEERING BRANCH QUESTIONS/RESPONSES

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NRC QUESTION 1:

Please provide documents on CESHOCK computer code for review.

FPL RESPONSE:

CESHOCK is Combustion Engineering's (C-E's) extensively modified proprietary version of the SHOCK computer code developed by V.K.

Gabrielson and R.T. Reese of Sandia Laboratories (Reference 1).

CESHOCK has been used "by C-E for nonlinear time-history analyses of reactor internals and fuel since 1972. These analyses are documented in CENPD-42, CENPD-178, and CENPD178 Revision 1 '(References 2,3, and 4)..

CESHOCK has been used in the seismic analysis of all C-E spent fuel racks. The same basic methodology was used in the seismic analyses for all these plants as well, as for Florida Power F Light's St. Lucie Unit 2.

C-E spent fuel racks have licensed at the following sites; Duke Power Oconee 1 (1979), 2 (1979), and,. 3 (1976), Arkansas Power and Light 2 (1977), Northeast Utilities Millstone Unit 2 ( 1977), Florida Power Light St. Lucie Unit 1 (1977), and Turkey Point Units 3 (1978) and 4 (1978). C-E fuel racks have also been supplied for Arizona Public Service's Palo Verde Nuclear Generating Station which has been reviewed by the NRC and is expected to be licensed in the near future.

Documentation for the submittal and NRC review of the above licensing efforts provides ample indication that the NRC has reviewed and accepted C-E fuel rack seismic analysis methodology (see References 7-10).

2-1

NRC QUESTION 2:

Please provide detailed analysis report showing

a. 'athematical models based on both CESHOCK and SAP IV computer codes, and FPL RESPONSE:

A typical SAP IV finite element model of a spent fuel rack is shown in Figure 1. The model consists primarily of plate elements with beam elements used to represent the fuel support bars.

The CESHOCK model is shown in Figure 2. Node 1 represents the spent fuel pool. Mass nodes 2 through 11 represent the fuel rack and nodes 12 through 21 represent the fuel. The hydrodynamic couplings between the rack and the fuel and the pool and the rack are designated by an H.

Nonlinear gap-spring elements represent the possibility of impacting between the fuel and the rack. A friction element couples the base of the fuel to the rack. The coupling element shown between the rack and the pool floor represents a friction element in a sliding analysis and a nonlinear torsion spring in a'ocking ana'lysis.

Separate CESHOCK models were developed for fuel assembly storage and consolidated fuel storage. Appropriate values for fuel element weight, beam stiffness, hydrodynamic coupling terms, gap and impact spring stiffness were used in each case.

2-2

Figure 1 Typical SAp EV Spent Fuel Rack Finite Element Model I

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NRC QUESTION 2:

Please provide detailed analysis report showing

b. Stress and displacement results FPL RESPONSE:

Reference 5 p'rovides detailed seismic loads and fuel rack stresses. The stresses reported in Reference 5 are'ummarized and compared below to the allowable stress limits as defined -in the ASME Code;Section III, Subsection NF, paragraphs 3220 and 3230:

Maximum Stress Intensities Found in the Racks Design Allowable (psi) (psi)

Normal Operation and OBE:

Primary Membrane (Pm) 19,713 20,000 Primary Membrane and Bending (Pm+Pb) 29,670 30,000 Primary and Secondary (Pm+Pb+Pe) 42,020 60,000 Faulted Conditions:

Primary Membrane (Pm) 28,056 30,000 Primary Membrane and Bending (Pm+Pb) 33,262 45,000 Maximum Stresses Found in the Fuel Support Bars Faulted Condition:

Bending Stress 4,930 33,000 Shear Stress 414 22,000 The displacements of the fuel racks during a seismic event due to the structural deflection, sliding, and tipping were examined in a series of analyses. Sliding analyses were performed for an empty 'rack and full rack. Tipping was examined for fully loaded, partially loaded, and empty racks. Structural deflections were obtained from analyses of fully loaded racks. The maximum relative displacement of any two fuel rack modules obtained from the sliding/tipping analysis was 1.88 inches. The nominal rack-to-rack spacing is 2 inches; therefore, adjacent rack modules do not contact each other.

2-5

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NRC QUESTION 3:

Please explain how the linear dynamic characteristics of the SAP IV model in air could be incorporated into the nonlinear CESHOCK model in water.

FPL RESPONSE:

The first step in the analytical procedure is to determine the dynamic characteristics of the rack structure. This was done by developing a three-dimensional finite element model of the rack and solving for its natural frequencies and mode shapes in air. The SAP IV -finite element code (Reference 6) was used for this purpose.,

The next step was the development of a dynamically equivalent model for use with the CESHOCK code. The detailed finite element'odel was reduced to a simple veritcal array of springs and masses which duplicated the weights, natural frequencies and mode shapes of the original three-dimensional modei. The mass nodes are allowed both translational and rotational degrees of freedom and because of the importance of fuel impacting, were located at the same elevations as the fuel assembly

, spacer grids.'eparate dynamically equivalent models were developed for each horlzonta1 direction.

The third step was to add a model representing the fuel assemblies. Gap-spring elements at every spacer grid elevation are used to model impacting of the fuel assemblies against the rack cell.

Finally, hydrodynamic coupling terms are included which couple the rack to the pool and the fuel assemblies to the rack.

2-6

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NRC QUESTION 4:.

Please discuss how the time step of integration was selected in the CESHOCK model for time history analysis in regard to the solution stability and convergence.

FPL RESPONSE:

The CESHOCK code numerically integrates the equations of motion using a Runge-Kutta-Gill technique. The initial integration time step, calculated by CESHOCK, is one-twentieth of the period of th'e highest individual mass-spring frequency in the model. The time step is continually checked and adjusted by the code as a function oF the rate of change of the linear and angular accelerations. The time step is held within the bounds of one-fifth times the initial time step to two times the initial time step. With this procedure for selecting the time step, the CESHOCK numerical solution has been shown to be stable and convergent.

2-7

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NRC QUESTION 5:

Please indicate the coefficient of friction of the pool liner used in the analysis and Justify its application.

FPL RESPONSE:

Friction between the pool liner and the rack is addressed in two ways.

In the first approach, the rack is not permitted to slide relative to the pool. This approach models the possibility of adhesion between the rack and the pool which could occur over the life of the racks due to one of several mechanisms. The fixed-base model provides conservative base shear loads for both the rack and the pool liner.

The second approach uses a sliding-base model in which a friction element connects the rack to the pool. Realistic values for the coefficient of friction (0.55 static and 0.28 dynamic), obtained from experimental data, are used. The sliding-base models are used to determine how far the rack will slide relative to the pool.

2-8

rt4 NRC QUESTION 6:

Please indicate what type of damping ( I.e., structural material, and fluid) was used in the analysis.

FPL RESPONSE:

Structural damping, in the form of Rayleigh'ass and stiffness proportional damping, was the sole type of damping used in the analysis.

2-9

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NRC QUESTION 7:

Please elaborate on the nonlinear elements used in the CESHOCK model.

FPL RESPONSE:

Two types of nonlinear elements were used in the CESHOCK models.

The first type of nonlinear element that was used is a slip-stick friction element. This type of element coupled the base of the fuel model with the rack in all cases and coupled the base of the rack with the pool in the fuel rack sliding response analysis. The friction element utilizes a velocity dependent coefficient of fr ict ion, equal to a static value when the relative velocity between model nodes is zero and which is a function of the nodal relative velocity for non-zero velocities.

A second type of nonlinear element, a piecewise -linear elastic spring, was used to model the nonlinear behavior of impacting between the fuel and the rack, and to model the nonlinear rocking characteristics of the rack module.

A piecewise-linear gap-spring element was used in the CESHOCK model to couple the rack and fuel nodes at each discrete spacer grid location.

The gap-spring element also allows a coefficient of restitution to be specified. Where appropriate, a coefficient of restitution, based on test results, was specified for the gap-spring elements in the St. Lucie 2 CESHOCK models.

A nonlinear torsion spring, represented by a piecewise-linear moment versus rotation curve, was used in the CESHOCK model. This element coupled the base of the rack with the pool in all cases. Further detail concerning this particular nonlinear element is included in the response to Question 9.

2-10

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NRC QUESTION 8:

Please elaborate on the procedure to estimate the hydrodynamic coupling effects between adjacent racks, and between fuel cell and fuel assembly.

FPL RESPONSE:

In the CESHOCK models, hydrodynamic coupling is specified between the rack and pool, and between the fuel and the rack. Potential theory

( Incompressible invicid theory) is employed, using simple two-dimensional models of the structures coupled by the fluid, to estimate the hydrodynamic virtual mass terms based on the model configuration. Three-dimensional end effects were then accounted for by modifying the calculated hydrodynamic mass terms.

For the rack-to-pool hydrodynamic element, the potential theory model consisted of two rigid, bodies: the fuel rack modules and the pool walls.

For the rack-to-pool hydrodynamic element, the potential theory model also consisted of two rigid bodies: the fuel cell and either a consolidated fuel canister assembly or a standard duel assembly, depending on the case under consideration.

For both types of elements, a finite element analysis was used to calculate the hydrodynamic masses of two-dimensional bodies with arbitrary eros's-sectional shapes with fluid finite elements between the bodies.

2-11

4,+

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NRC QUESTION 9-Please explain how the tipping-off analysis was conducted.

FPL RESPONSE Nonlinear time-history analysis using the CESHOCK computer code were used to conduct the tipping-off, or rocking analysis. The basic CESHOCK model, as described in Responses 3 and 7, was used to qualify the seismic rocking response of the racks. This was accomplished by using a nonlinear torsion spring at the base of the rack to model the moment versus rotation relationship of a rocking rigid body. The first portion of this moment-rotation curve is linear and respresents the linear rotational stiffness of the module up to the point of impending lift-off.

At greater rotations, when the module could tip up,and lift off of the pool floor, the moment-rotation curve is nonlinear and represents the restoring moment of an assumed rigid module having tipped up on one edge.

Several analyses were made to quantify the rocking response for fully loaded, partially loaded, and empty modules. Results are summarized in the response to question 2b.

2-12

NRC QUESTION 10:

Please indicate the mode of vibration in assessing the hydrodynamic coupling effects between adjacent racks (I.e., symmetric or antlsynmetrlc).

FPL RESPONSE:

An in-phase mode of vibration was conservatively considered in assessing the hydrodynamic coupling effects between adjacent racks. Because of the character of the site specific St. Lucie 2 seismic excitation, the higher rack frequencies resulting from the in-phase mode analysis were conservative.

An out-of-phase mode of vibration would have resulted in significantly larger hydrodynamic mass terms, thereby lowering the coupled fuel rack natural frequencies.

2-13

NRC QUESTION 11:

Please explain how seismic loads were applied and how the results were combined in the analysis.'PL

RESPONSE

Seismic loads were computed for all shock directions. The lateral analysis (North-South) and (East-West) was done by a nonlinear time history technique. A separate response spectrum analysis'as performed in the vertical direction.

The resultant load was applied to a finite element model of the fuel rack module at each cell, for each shock direction. The resultant'omponent stress on each element from the application ot each shock direction load was combined by the square root sum of the squares method. The results were compared to stress allowable in accordance with the rules of ASME Boiler 8 Pressure Vessel Code Section III Subsection NF-3220 and are shown in the response to question 2b and the spent fuel rack design analysis report.

2-14

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Please provide analytical model used for the structural (static and dynamic) analysis of the fuel pool floor system, and give the loading systems and summary of results.

FPL RESPONSE:

Conventional lumps mass mathematical models are used in the'ynamic analysis of the Fuel Handling Building (FHB). Attached Figures 1 to 3 are the detailed models used.

Figure 4 is our .finite element model of the fuel pool utilizing plate and rigid bar elements.

The loading combinations and structural acceptance criteria for the FHB are specified in FSAR Section 3.8.4. In this rerack, each of the loading conditions were reviewed with the increased fuel rack loads as specified in paragraph 4.3 of the Spent Fuel Rerack Safety Analysis Report. The critical load condition was the following load combination:

U = 1.0D + 1.0 (FC+L) + 1.0E + 1.0TN Where D = Dead Load L = Live Load E'= SSE TN Thermal Load FC = Cask Drop The required strength (~ ) of the maximum stressed element is less than the furnished strength capacity.

2-15

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2-17

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NRC QUESTION 13:

Detailed information is required on the dynamic analysis of floor system under OBE and SSE events. Please provide response amplif ications and rack impact loads.'PL

RESPONSE

Detail information on the dynamic analysis of the Fuel, Handling Building is contained in FSAR Section 3.7.2. FSAR Tables 3.7-31. through -3.7-33 provides the structural responses of the building under a SSE event.

Attached Tables 1 and 2 provides the structurai responses of the building under an OBE event.

The fuel rack impact loads on the fuel pool floor from the tipping of the fuel rack module are not significant compared to the total vertical seismic plus deadweight load used to evaluate local concrete stresses under the rack legs. The analysis has shown that the racks tip enough to transfer loads from four pads to two pads, but do not significantly lift from the f loor.

2-20

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NRC UESTION 14:

The rack impact loads cause high localized stresses on the concrete underneath the rack legs. Please indicate the stress levels and justifications if necessary.

FPL RESPONSE:

The maximum ultimate vertical load (based on the loading in Section 4.3 of the Spent Fuel Rerack Safety Analysis Report) on one rack leg bearing pad is 294.3K. The maximum bearing stress of 3.26ksi is less than the allowable bearing stress of 4.76ksi as specified in ACI 318-77 paragraph 10.16.12.

The fuel rack impact loads on the fuel pool floor from the tipping of the fuel rack module are not significant compared to the total vertical seismic plus deadweight load used to evaluate local concrete stresses under the rack legs. The analysis has shown that the racks tip enough to transfer loads from four pads to two pads, but do not significantly lift off from the floor.

2-23

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NRC QUESTION 15:

Please outline the procedure for thermal analysis indicating whether the analysis was based on cracked or uncracked section and provide a sample calculation.

FPL RESPONSE:

In the thermal analysis of the FHB, a 3-dimensional finite element model of the building was constructed based on uncracked sections. Figure 4 (see response 1) is a portion of our 3-dimensional model. The design temperatures inside and outside the building were input into the finite element model. STARDYNE was used to perform the thermal analyses. The resulting forces/moments are reduced by the ratio I<< where Icr = cracked Iunc section and Iunc uncracked section. These design forces/momen'ts are then used in various load combinations (see FSAR Section 3.8.4) in the design of the building. The reduction of the thermally induced forces/moments by the ratio Icr was reviewed and found acceptable by the Iunc NRC during the July 1981 St. Lucie Unit 2 Structural Audit in connection with the. Reactor Building Exterior Shield Wal I design.

Attachment A is a sample calculation.

2-24

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REFERENCES

1. Gabrielson, V.K., "SHOCK" - A Computer Code for Solving Lumped-Mass Dynamic Systems," Sandia Corporation, Livermore Laboratory, SCL-DR-65-34,

,January 1966.

2. CENPD-42, Topical Report on Dynamic Analysis of Reactor Vessel Internals Under Loss-Of-Coolant Accident Conditions With Application To C-E 800 Mwe Class Reactors, August 1972 (Proprietary).

3. CENPD-178, Structural Analysis of Fuel Assemblies For Combined Seismic and Loss of Coolant Accident Loading, August 1976 (Proprietary).

4. CENPD-178, Revision 1, Structural Analysis of Fuel Assemblieq for Seismic and Loss of Coolant Accident Loading, August 1981 (Proprietary).

5. Design Analysis Report For The Florida Power and Light St. Lucie Generating Station Unit 2 Spent Fuel Storage Racks, June 29, 1984.

6. Bathe, K.J., Wilson, E.L., and Peterson, F.E., "SAP IV - A Structural Analysis Program for Static and Dynamic Response of Linear Systems,"

Earthquake Engineering Research Center Report No. EERC-73-11, University of California, Berkeley, June 1973.

7. Letter form W.O. Parker (Duke Power Company) to H.R. Oenton (NRC), Oconee Nuclear Station Dockets 50-269 and 50-270, February 2, 1979.

8. Letter from W.O. Parker (Duke Power Company) to H.R. Denton (NRC), Oconee Nuclear Station Dockets 50-269 and 50-270, April 20, 1979.

9. NRC Safety Evaluation: "Safety Evaluation by the Office of Nuclear Relating to the Modification of Oconee Units I/2 Common Spent Reactor'egulation Fuel Storage Pool", Facility Operating License Numbers DPR-38 and DPR-47,

'ated June 19, 1979.

10. Palo Verde SQRT Visit Report (Final), Prepared for the U.S. NRC, DOE Contract DE-AC07-76-1001570, Idaho National Engineering Laboratory, Report Number EGG-EA-6427, October 1983.

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ATTACHMENT 3 AUXILIARY SYSTEMS BRANCH QUESTIONS/RESPONSES

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NRC QUESTION I:

Provide revised heat generation rate calculations using NURGE-0800, Standard Review Plan, Branch Technical Position ASB 9-2 and Section 9.I.3 for the following cases:

(a) Normal refueling until the spent fuel pool is full.

(b) . Same as case (a) above except the last 2I7 locations are filled by a core off

'load.

FPL RESPONSE:

The decay heat loads were calculated consistent with the documents'identified above, and using the following assumptions:

(a) Normal Refueling I. Twelve refueling batches are in the pool (984 assemblies, which prevents a ful I core off-load).

2. The latest batch has decayed 5 days.

3. The other eleven batches have decayed I8 months, 36 months, etc.

4. All assemblies received 48 months of full power burn.

5. Table 5-I of the Spent Fuel. Pool Rerack Safety Analysis Report was used for the number of assemblies per batch.

The total decay heat load for6the normal refueling case is I 6.9xl0 BTU/hr. This heat load is less than 17.2xl0 BTU/hr which results in a fuel pool temperature of l37 F with only one spent fuel pool cooling pump operating. Therefore, for the normal refueling case, the fuel pool temperature will be less than l37 F.

(b) Full Core Off-load I. Eleven refueling batches and the full core off-load are in the pool (I I I3 assemblies, which exceeds the l076 storage capacity).

2. The full core off-load has decayed for 7 days.

3. The other eleven batches have decayed 18 months, 36 months, etc.

4. All assemblies received 48 months of full power burn.

5. Table 5-I of the Spent Fuel Pool Rerack Safety Analysis Report was used for the number of assemblies per batch.

The decay heat load for t)e full core off-load case is 31.7xl0 BTU/hr. This heat load is less than 32.0xl0 BTU/hr which results in a fuel pool temperature of I50 F with both spent fuel pool cooling pumps operating. Therefore, for the full core off-load case, the fuel pool temperature will be less than l50 F.

Assuming loss of all fuel pool cooling, the following conditions bound the two cases above:

Time to Reach Minimum Acceptable Water Level Heat oad Time to Boil Boil-off Rate Above the Fuel 32.0x I 06 U/hr 2.9 hours1.041667e-4 days <br />0.0025 hours <br />1.488095e-5 weeks <br />3.4245e-6 months <br /> ~ 3 gpm I 7.2x I 0 BTU/hr 9.0 hours0 days <br />0 hours <br />0 weeks <br />0 months <br /> 35.6 gpm 2.49 days

pp NRC QUESTION 2:

Provide the following information for the spent fuel pool cooling system:

(a) Heat exchanger tube surface area (square feet).

(b) Heat exchanger conductance (BTU/ft2-Hr-oF).

(c) Spent fuel pool water flow rate, per pump (Lb/Hr) ~

(d) Component cooling water flowrate, per heat exchanger (Lb/Hr).

(e) Design component cooling water inlet temperature to the heat exchanger (oF).

FPL RESPONSE:

(a) The heat exchanger tube surface area is 4380 ft (b) The heat exchanger conductance is 60 Btu/ft2-Hr-oF.

.(c) The spent fuel pool water flow rate is 0.75x10 Lb/Hr per pump.

(d) The component cooling water flow rate ,is 1.78x106 Lb/Hr per heat exchanger.

(e) The design component cooling water inlet temperature to the heat exchanger is 100oF.

As indicated in an April 16, 1984 phone call to the NRC, the design inlet temperature on the tube side (hot side) of the heat exchanger is 150oF and the logarithmic mean temperature difference from the tube side to the shell side is 28. 1oF. The design capacity of the heat exchanger is (conductance) x (heat transfer area) x ( logarithmic mean temperature difference) = 32x10 Btu/hr. IA k

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Attachment 4 Heetin Summar St. Luci e 2 Fuel Rack Seismic Anal sis August 16, 1984 Windsor, CT

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8 16 84 SL-2 Fuel"Rack Heetin Attendees Stan Ritterbusch C-E - Project Licensing (203) 285-5206 Ronald J. Stevens FPL - Nuclear Licensing (305) 863-3620 Dave Baisley C-E - Structural Analysis (203) 688-1911 (ext. 4113)

Art Johnson C-E - Structural Analysis

( 203) 688-1911 (ext. 2726)

Donald E. Sells NRC/OL (301) 492-9735 Dan Ferris C-E - Mechanical Design (203) 688-1911 (ext. 4222)

Bob Longo C-E - Structural Analysis Mike Falzarano C-E Fuel Rack Coordinator Sang Bo Kim NRC/DE/SGEB Vincent K. Luk Franklin Research Center R. Clyde Herri ck Franklin Research Center 4-1

CONFERENCE ON SPENT FUEL RACK ANALYSIS, ST. LUCIE 2, COMBUSTION ENGINEERING, USNRC,. FRC .

AUGUST 16, 1984 DISCUSSION TOPICS

1. Discussion of the modeling and analysis procedure

a. Generation of Mass Elastic Properties of a Rack Module

b. Development of the Non-Linear Dynamic Rack Model o Rack mass-elastic modeling o Fuel assembly modeling o Hydrodynamic coupling of fuel assemblies and rack modules o Friction modeling at rest and sliding o Mounting pad lift-off capability (gap elements)

'c. Solution of Non-Linear Dynamic Model o Loadings employed (seismic) o Solution method (Runge-Kutta-Gill?)

o Time step increment size and how chosen?

o What assures that the solution is stable and uslj conve ges?

d. Stress Model o Model type, 2-D or 3-D, elements employed, linear?

o How were the results of the 2-D non-linear impact dynamic model incorporated into the stress model. Give detailed description.

o How are seismic loads applied what sequence?

o What assures that this procedure provides conservative estimate of stress? Does it provide an upper hand?

o Procedures to determine peak stresses at high stressed points.

Stress Levels and Allowable Values a.'dentify the allowable stresses and their authoritive source.

b. Justify all low margin stresses with adequate conservation of analysis procedures.

c. Identify which low margin stresses may result from the use of consolidated fuel.

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3. Fuel Drop Accident Provide detail drawing of rack cell base and fuel mounting interfaces.

b. Provide assurance that a full drop accident will not damage'the spent fuel pool liner.

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Res onse to Discussion To ics A more detailed plot of the SAP-IV model is provided in the attached figure 1.

Fuel rack structural damping was 4/ for the SSE and Q for the OBE. The value of structural damping for the fuel assemblies is conservative when compared to the results of full scale assembly forced vibration tests.

The maximum relative displacement between rack modules from the seismic sliding and tipping analyses is 1.88 inches. This is an upper bound value determined from a subset of nonlinear time history sliding and tipping analyses. Cases were run for a fully loaded rack without sliding, a partially loaded rack without sliding, an empty rack without sliding, a fully loaded rack with sliding, and an empty rack with sliding. The sliding cases yielded relative displacements between modules that were significantly less than the inter-module gap of 2.0 inches. For the tipping cases, the fully loaded rack did not lift off the pool floor. The empty rack and the partially loaded rack cases provided the highest relative deflections due to tipping and elastic displacement. To obtain the upper bound value of 1.88 inches, the displacements from empty and partially loaded racks were added assuming that they were moving totally out of phase. The maximum tipping displacement'f the empty racks was combined with the partially loaded rack displacement at the same time during the earthquake.

A copy of the ADMASS documentation was provided to FRC at the meeting, however, the report given, ANL-CT-78-49 dated 9/78, is not applicable. A microfiche copy of the applicable report, ANL-CT-78-42 dated 9/78, is attached.

The peak base shear, load from the seismic analysis is obtained by conservatively assuming the fuel rack cannot slide relative to the pool (infinite coefficient of friction). The peak tipping occurs for a loaded rack and is also determined from an analysis .in which the

'artially fuel rack does not slide. The peak relative displacement of a rack module due to sliding is obtained from an analysis in which realistic friction values are assumed. The friction element is a slip-stick friction element with a velocity dependent coefficient of friction. The friction values are based upon the textbook "Friction and Wear of Materials" by Ernest Rabinowi cz, data from Combustion Engineering laboratory tests, and data obtained through a technical exchange agreement with Kraftwerk Union (KWU) of West Germany.. '

The time steps used in the seismic analysis provided a stable and converged solution. Confidence in the solution is based upon approximately 15 years of experience at Combustion Engineering using CESHOCK for seismic analyses of spent fuel racks, reactor internals, fuel and other complex nonlinear dynamic problems. Comparisons of typical fuel rack analysis have been made with.other computer codes and good agreement was shown. Nonlinear time history analysis parameter runs with different time steps showing convergent solutions have been performed.

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7. The maximum stresses occur at plate elements of the fuel cell wall at an elevation near the fuel lower end fitting and support bar interface.

Maximum stress points are clustered near the rack/floor support points (see attached figure 2).

8. Small margins calculated through realistic analyses are acceptable based upon conservatisms inherent in code allowables. In addition, the following conservative assumptions were made in this analysis:

a) Stress is computed assuming that the rack base is totally fixed to the fuel pool floor and not allowed to slide.

b) Stress analysis results assume 1005 fuel loading in each rack.

Actual loading is between 50-75/ for normal storage.

c) The value of fuel assembly damping used in the analysis is significantly less than that measured by test.

d) Peak base shear and tipping analyses were conducted assuming racks cannot slide (infinite coefficient of friction).

e) -

Tipping representation is based on rigid body rotation and does not account for bending flexibility of rack.

f) Peak broadening is done in accordance with Reg. Guide 1.122.

9. The fuel drop accident was evaluated to determine the effect of a dropped assembly on the functional and structural integrity of the racks. The analysis indicated that the impact of a fuel assembly .on the support bars caused plastic deformation of the support bars and the fuel cell wall supporting the bars. For conservatism it was assumed that further..

displacement of the bars occurs, resulting in the fuel and support bars potentially resting on the pool floor. Neither functional nor structural integrity of the racks was impaired.

Impact on the fuel pool liner was not analyzed; however, a dummy fuel assembly was dropped during gaging of the St. Lucie 2 racks. This drop,

'which. occurred in air as opposed to water, resulted in some deformation of the support bars, but did not .impact the fuel pool liner. This supports the assumption that a dropped fuel assembly will deform the support structure but not result in impact to the fuel pool'iner.

10. The seismic analysis indicates that a fully loaded rack does not lift off the spent fuel pool floor for the conditions postulated. Analysis has shown that lift off can occur in the case of a partially loaded rack.

Further analysis was performed which demonstrates that loads resulting from the tipping and subsequent impact of a partially loaded rack are bounded by t'e maximum loads of the fully loaded rack.

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11. 'The method for transferring the loads from the CESHOCK "stick" model to SAP IV stress model is as follows:

A one-G response spectrum load is applied in each orthogonal

'direction to the three-dimensional SAP-IV stress model.

Component stresses derived from this procedure are multiplied by a load factor determined from the results of the CESHOCK stick model. Typical load factors are as follows:

Normal Stora e OBE SSE Maximum Horizontal 7.7 10.0 Maximum Vertical 12.0 13.2 Consolidated Stora e Maximum Morizontal 5.7 10.5 Maximum Vertical 17. 3 19.0 The component stress on each element resulting from the applic'ation of each directional load is combined by the square root sum of the squares method. The results are compared to stress allowables in accordance with the rules of ASME Boiler 8 Pressure Vessel Code Section III Subsection NF-3220.

Both horizontal and vertical load factors are derived from the ratio of the peak load from the CESHOCK stick model to the empty rack load.

The two horizontal peak loads are determined from the stick model non-linear time history analysis.

The vertical peak load is determined from the stick model response spectrum analysis. . The loaded rack vertical

'atural frequency is so high that it behaves like a rigid, body. The peak vertical seismic load is determined by multiplying the mass of the loaded rack times the zero perio'd accelleration of the response spectrum.

All the horizontal and vertical loads used in the stress analysis were maximums.

The typical ratio of a loaded cell weight to the empty cell weight is about a factor of 10.

4-6

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Figure I Plot of the St. Lucie II Spent Fuel Rack SAP-IV Finite Element Model

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Rack Str'esses Maximum stresses found in the fuel supoort bars.

Maximum stress intensities found in the canisters: Bending Stress: Faulted ~ 4,930 Normal Operation 5 OBE Shear Stress: Faulted 414 psi Pm '9.713 psi P + Pb K 29,670 psi P + Pb + P

• 45,020 psi m

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P ~ 28,056 psi m

~ m Pb 33,262 psi Skewer I

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.~ fuel Assem "ly Suoport Plate Slot Flovr Passaoes Maximum Stresses and Stress Intensities in Fuel Storage to Earthquake Loading Conditions. Packs Gue Figure 2 4-8

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