ML20023B731

Forwards Response to 830311 Ltr Requesting Addl Info Re Proposed Spent Fuel Storage Rack Mod.New Analysis Indicates Spent Fuel Can Be Adequately Cooled Even w/4 H Loss of Spent Fuel Pool Cooling

ML20023B731
Person / Time
Site:	Fort Calhoun
Issue date:	04/29/1983
From:	William Jones OMAHA PUBLIC POWER DISTRICT
To:	Clark R Office of Nuclear Reactor Regulation
References
LIC-83-095, LIC-83-95, MR-FC-80-25, NUDOCS 8305060186
Download: ML20023B731 (40)

Text

,

e Omaha Public Power District 1623 Hamey Omaha Nebraska 68102 402/536-4000 April 29, 1983 LIC-83-095 Mr. Robert A. Clark, Chief U. S. Nuclear Regulatory Commission Office of Nuclear Reactor Regulation Division of Licensing Operating Reactors Branch No. 3 Washington, D.C. 20555 l

Reference:

Docke t No . 50-285 l

l

Dear Mr. Clark:

i Spent Fuel Storage Rack Modification Attached is Omaha Public Power District's response to the Commission's letter dated March 11, 1983 requesting addi-tional information concerning the Fort Calhoun Station's proposed spent fuel storage rack modification.

Sincerely, fIh$k h W.Division v

. Jones Manager Production Operations

WCJ/TLP

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l Attachment cc: Le Boe u f , Lamb, Leiby & MacRae 1333 New Hampshire Avenue, N.W. 1 Washington, D.C. 20036 ()(> f 1 i Mr. E. G. Tourigny, NRC Project Manager Mr. L. A. Yandell, NRC Senior Resident Inspector 8305060186 830429 p DR ADOCK 05000285 PDR 455124 Employment with Equal Opportunity Maleifeiriale

OPPD Response to NRC Questions letter dated March 11, 1983 March 30, 1983 MR-FC-80-25 l

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ATTACHMENT PART A NRC COMMENT

1. Provide justification for the conclusion that the degree of agreement with diffusion theory calculations of the Pacific Northwest Laboratories (PNWL) high leakage critical experiments can be properly extrapolated to the Fort Calhoun Spent Fuel Rack infinite array.

OPPD RESPONSE:

A spent fuel rack infinite array that is effective for storing high enrichment unirradiated fuel is similar in many ways to a high leakage critical array such as that represented by the PNWL experiments. The infinite multiplication fac-tor of the fuel' assembly used as a reference for the Fort Calhoun racks is about 1.45, and criticality safety is achieved in this design primarily by effectively decoupling each fuel assembly from its neighbors. The decoupling is achieved by the flux trap effect produced by fixing neutron poison sheets on both sides of a water filled gap located between each fuel assembly.

Criticality control is primarily achieved by leakage of fast neutrons that are born in the fuel assembly, slowed to thermal energies in the water gap, and subsequently absorbed in the neutron poison material. In addition, epithermal and thermal neutrons from the fuel assembly leak directly into the neutron poison material. Thus, the accuracy of the analysis model in calculating the PNWL critical experiments is rel evant to the accuracy of the model for criticality safety calculations of the Fort Calhoun spent fuel storage racks.

In addition, as described in the submittal, other experiments have been ana-lyzed and reported therein, and the extensive use of similar analysis models and methods in the Naval Reactor Program has been described. Additional quantitative justification of the analysis model is provided in the answers to questions 2 and 4 below.

I NRC COMMENT l

! 2. Since the absorber plate reactivity worths in the fixed neutron poison l critical experiments analyzed were much lower than the Boraflex worths in l the spent fuel racks, provide justification for the conclusion that the benchmarking signifies that absorption effects were properly treated.

OPPD RESPONSE:

l Of particular significance in the analysis of the experiments used for bench-marking the analysis model, is the apparent lack of any trends in the model bias for both the Westinghouse and PNWL high leakage critical experiments.

This lack of a trend is evidenced by a mean calculated ke rf of .9928 with a stedard deviation of .0012 for the 14 Westinghouse experiments and a mean calculated ke rf of .9931 with a standard deviation of .0011 for the 12 PNWL experiments.

In addition. as described in Reference 10 of the submittal, Section 4, the use of blackness theory for calculating effective cross sections for highly absorb-ing thin plates has been found to allow calculation of critical experiments -

I 9 containing strong absorbers with about the same degree of accuracy as that

> achieved in the analysis of critical experiments which do not contain strong absorbers. The attached Table A-1 provides further verification of the ability of the analysis model to accurately cal culate the multiplication factor of critical experiments containing strong neutron absorbers. Even though the fuel utilized in these experiments is not similar to commercial reactor fuel, it will be demonstrated that the adjustment to the nominal calculated multipli-cation factor for the rack as derived from analysis of these experiments (i.e.,

Akbias + 20-) is essentially the same as that derived from the experiments reported in the submittal. As shown in the last column of Table A-1, the worth of the absorbing plates are comparable to the Boraflex worth in the Fort Calhoun spent fuel racks. Column four shows the calculated results reported by the experimenters using blackness theory, and column five shows results obtained using the same LEOPARD /PDQ-7/ Blackness Theory model used for the analysis of the Fort Calhoun racks.

Although the standard deviation of the difference between calculated and experi-mental values is greater than that based on the PNWL experiments, the bias in the average calculated multiplication factor is significantly less; so that the derived adjustment to be made based upon the sum of the bias and 2a is nearly identical in both cases.

Table A-2 shows a revision to Table 4.4-1 of the submittal based on the use of the bias and standard deviation shown in Table A-1. In this case, the final maximum adjusted k is .9384 which is less than the value of .9415 reported in Taule 4.4-1 of the submittal .

NRC COMMENT

3. The experiments performed by Babcock and Wilcox (M. N. Bal dwin, et al, Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel, BAW-1484-7, July 1979) include much higher absorber worths.

t Have they been analyzed as part of your verification? If so please pro-vide results of the analyses.

OPPD RESPONSE:

i As described in the responses to questions 1 and 2 above, the analysis model

[ has been extensively benchmarked against a large number of critical experi-ments. However, the experiments performed by Babcock and Wilcox are not among those included in the benchmarking.

i NRC COMMENT l

4. Have comparisons been made to higher order calculations (e.g., KENO-IV with the AMPX-NITAWL 123 group cross section set)? If so provide results of such comparisons.

OPPD RESPONSE:

A comparison of k.'s calculated for fuel pin cells with the PLG version of LE0-PARD with k.'s resulting from a 123 group XSDRN library shows the XSDRN based k.'s to be consistently higher than the LEOPARD based k.'s by about 0.6% Ak.

The calculated results of KEN 0 IV calculations reported for the PNWL critical experiments (1) are rather consistently high by the order of 1-2% Ak. Since .

- .-. - __ L . - - _ . - - _ _ - _

i these results did not utilize a 123 group XSDRN based library, the experiment designated 002 in Table 4.1-3 of the submittal was analyzed with KEN 0 IV using a 123 group XSDRN l ibrary. The LEOPARD /PDQ-7 analysis model resul ted in a calculated Keff of .993 for this experiment while the Keff calculated with KEN 0 IV was 1.011 + .006. This result is consistent with the results of the KEN 0 IV - calculatioiis referenced above; i.e., KEN 0 IV appears to result in keff's that are too high by about 12% Ak. Although we continue to believe that bench-marking an analysis model against critical experiments is the most reliable method of determining model bias and uncertainties, a comparison was nade of

! results from the LEOPARD /PDQ-7 model with results from a KEN 0 IV calculation for a Point Beach type spent fuel rack which has neutronic characteristics similar to the Fort Calhoun racks. As expected, the results shown in Table A-3 indicate KEN 0 IV predicts a k, that is about 1% higher than that calculated wi th the LEOPARD /PDQ-7 model. However, no bias has been included in these results to represent the effects of cross section and model uncertainties. If, as indicated above, the bias for the KEN 0 IV calculation is of the order of 1%

Ak, then the agreement between the two models is excellent as shown in the final column of Table A-3. If the bias for the KEN 0 IV calculation is greater than 1%, then the final k. derived with KEN 0 IV could be even lower than the i corresponding value derived from the LEOPARD /PDQ-7 model.

NRC COMMENT

5. The statement is made that the use of the simple assembly average expo-sure can result in an over-estimate of the fuel assembly kerf by about +

l .015 Ak/k. Is this based on an actual calculation? If not, how does the k erf of the pancake region consisting of the lower exposure end of

! the fuel assemblies compare to that calculated based on an assembly l average exposure ?

! OPPD RESPONSE:

l l A one-dimensional PDQ-7 axial model of a fuel assembly was depleted to an aver-age burnup of 15,000 MWD /MTU. The relative burnup at the end of that depletion is shown in Figure A-1. The keff of the assembly at that time with zero radial leakage was 1.1334. The axial calculation was rerun using an axially uniform set of cross sections corresponding to an average burnup of 15,000 MWD /MT!! and the resulting kgff was 1.1509 or about .015 Ak/k higher than the more realistic non-uniform axial burnup case.

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TABLE A-1 SUFDtARY OF ABSORBING PLATE EXPERIMENTS

• Four Group Calculated keff Calculated Reactivity Critical Experimental Reported by Current LEOPARD / Worth of Absorbing Plate Dimensions (in.)3 Rod Free k ggg Dettis Lab PDQ-7 Model Absorbing Place 0.020 in. Cd 9 x 26 x 40I/4 1.002 1.005 1.0097 0.144 0.010 in. Au 7 x 32 x 395/8 1.000 1.001 1.0060 0.065 0.030 in. Au 8 x 32 5/8 x 39 5/8c 1.001 1.004 1.0005 0.121 0.050 in. Att 9 x 26 7/9 x 39 5/0 1.000 1.011 1.0118 0.142 0.059 in. lit 10 x 21 x 395/8 1.005 1.005 1.0095 0.151 0.204 in. lit 11 x 215/11 x 395/8 1.000 1.004 1.0056 0.200 0.204 in. lif at core reflector 8 x 22 x 395/8 1.003 0.992 0.995 0.086 interface Fully inserted, 0.059 in. thick, Figure 7 0.999 0.996 0.995 0.142 cruciform Ilf ~

control rod Average and Standard Deviation 1.0013 1.0023 + .0060 1.0041 +~ .0068

~

for Eight Experiments Dias in Average k 0.0010 0.0028 eff "If the smallest dimension is even, the plate is at center of slab; if odd, the plate is 1 in off center unless otherwise indicated.

Obtained from residual rod worth.

c325/8 means that the 33rd row contained only five bundles.

M. Goldsmith, et al., " Experimental and Theoretical Study of Critical Slabs - Ef fect of Absorbing Membrancs of Cadmium, Gold, and Boron," Bettis Atomic Power Division Report WAPD-170, (April 1957).

TABLE A-2 REVISED

SUMMARY

OF REACTIVITY BIASES AND UNCERTAINTIES FOR FORT CALHOUN REGION 1 BASED Of4 BIAS AND UNCERTAINTY OF TABLE A-1 REACTIVITY DESCRIPTION EFFECT Ak k.

Basic cell at 68'F 4.0 w/o 0.9258 U-235, 9.935 inch pitch, (see Figure 4.1-1)

Calculation Biases 4

Leopared/PDQ Model Bias .0028*

Hesh Spacing Effect + .0004 Net Axial Adjustment .0015

.0039* 0.9219*

Basic cell including Biases Tolerances and Uncertainties Minimum pitch (i.e., box to .0091 box spacing)

. Wrapper .0002

, Tolerance in SS box thickness .0009 Pellet density ( 0.15) .0017 Pellet diameter (t .0005) .0003 Calculational uncertainty (20-) .0136*

Total Uncertainty (statistical) .0165*

Maxinum, including all biases and 0.9384*

uncertainties Basic cell at 68'F with 1700 ppm boron 0.7411

• Revision from Table 4.4-1 of the submittal.

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TABLE A-3 KENO-IV AND LEOPARD /PDQ-7 CRITICALITY CALCULATIONS FOR A POINT BEACil TYPE SPENT FUEL STORAGE RACK (Design based on 68'F, 3.7 w/o U-235, and aB 10 loading corresponding to 46 w/o D4 C in Doraflex)

RANGE INCLUDING 23 BIAS AND MODEL k. UNCERTAINTY MODEL DIAS UNCERTAINTY LEOPARD /PDQ-7 0.898 0.0022 III + .0069(1I .903 .907 KENO-IV 0.908 0.012(2) - .011 I3I .885 - .909 (1) Based on twelve critical experiments dodeled with LEOPARD /PDQ; includes effects of cross sections and inaccuracies of diffusion theory.

(2) Includes only the statistical variation associated with the monte carlo cpproach; does not include bias effects of cross sections and model uncertainties.

( I Based on analyisis of PNWL experiments, bias in KENO IV calcu-lations appears to be about + .01 to + .02 ak.

FIGURE A-1 RELATIVE AXIAL BURNUP AND POWER DISTRIBUTIONS AFTER DEPLETION OF 15,000 MWD /MTU 4 - -

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PART B NRC COMMENT

1. Provide a sketch showing the thickness of material in the rack cans.

OPPD RESPONSE:

The rack can material is stainless steel sheet, .090 in. thick as marked on attached drawings: AD-345-13-1 Sheet 2, AD-34515-0 Sheet 2 and AD-34525-C.

Complete material specifications are provided by parts lists No.'s L-34513 and L-34515.

2. NRC COMMENT Describe or provide a more detailed sketch of the bounding bar or angle at the top of the rack around the perimeters. What are the stresses in this bar?

OPPD RESPONSE The bounding bar is a stainless steel strip 4" wide x 3/8" thick. Complete material specifications are provided by the applicable design drawings noted.

The highest stress in the perimeter bar is 2818 psi. This stress consists of two orthogonal bending components and an axial compressive stress of 837 psi.

3. NRC COMMENT Provide a tabulation to actual as well as allowable stresses for welds at key points in the racks for pertinent bending gonditions.

OPPD RESPONSE The work on structural analysis that deals with calculations of stresses on various wel ds is in progress. Stresses for welds at key points have been calculated. From the results of the preliminary analytical work completed thus far, the highest weld stress occurs at the joint between a fuel can and the grid. Consistent with par's past experience, the maximum stress index is computed as 14985 psi, to be compared with an allowable fillet weld stress in accordance with appendix XVII-2450 of 21 KSI. This stress occurs during the seismic loading. The complete weld stresses analysis will be documented in the final design report.

NRC COMMENTS

4. Was local buckling of the cans considered? What acceptance criteria was buckling compared against? Where is the potential for local buckling greatest in the cans?

OPPD RESPONSE:

Local buckling was considered both for seismic loading and the " fuel drop accident". In general, acceptance criteria used were those of ASME Section III, Div. 1, Appendix XVII-2235. In specific instances where support or loading conditions were not covered in that Appendix, independent buckling analyses were performed using closed form and finite element techniques as appropriate foll owing the procedures set forth in Section 1.9 of "AISC Commentary on the specification for the Design Fabrication and Erection of Structural Steel for Building".

The potential for local buckling is greatest just aLove the lower flare for seismic loading. Although the highest stresses occur within the flare itself, this area is made up of short, stiff panel segments which are not prone to load-ing. Just above the flare, the can wall is much softer because of its length and is most subject to buckling even though the stresses are numerically lower.

For local impact, buckling is confined to the area near the loaded portion of the upper flare. Even though the flare is quite stiff, compressive stress is too highly localized to buckle a large portion of the can. Euler buckling loads for the can structure are too high to be of concern.

It should be noted that the rules of Appendix XVII apply strictly to ' Linear Supports' and should be applied to panel structures like the fuel cans with care, since the stress fields are highly bi-axial. In particular the Qs fac-tors governing buckling design of compression elements assume constant compres-sive stress over a length three or more times the width, and will be over-con-servative for shorter lengths. At the same time bi-axial coompression may render the Qs factors nonconservative.

NRC COMMENT

5. Provide a tabulation of actual buckling stresses compared with allowable stresses in the cans for local buckling.

OPPD RESPONSE:

The highest compressive stress in the region subject to buckling is 5368 PSI.

In accordance with Appendix XVII the allowable compressive stress is 10600 PSI.

As noted above, the slenderness ratio of the can is too large for Euler buckl-ing to be of significance. Moreover, shearing stresses are too low for shear buckling to be a factor.

9

NRC C0tWENT

6. Provide a tabulation of actual and allowable stresses for the key structural components of the rack.

OPPD RESPONSE:

Actual and allowable stresses in accordance with the requirements of appendix XVII are as follows:

COMPONENTS ACTUAL STRESS ALLOWABLE STRESS Grid 16403 PSI 28800 PSI Perimeter Bar 2818 PSI 28800 PSI

. Fuel Can 5368 PSI 10600 PSI Note that the design stresses of welds are covered in HRC comment 3 above.

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NRC COMMENT

7. Describe the seismic input load for these racks. Were the three components of earthquake input to the ANSYS model ? Describe the method in detail.

OPPD RESPONSE:

A general description of the seismic input appears in paragraph 7.5 of the Licensing Submittal. The response spectra used to generate the time history were taken from the USAR, and are shown in attached figures 1, 2 and 3. The SSE spectrum is the more severe condition even when the increased allowable stresses are taken into account, so it was used for the dynamic analysis. The SIM0KE computer program (devel oped at MIT under National Science Foundation sponsorship) was used to generate the acceleration time histories shown. These time functions were integrated twice to provide displacement boundary condition input to the ANSYS time history run. Six thousand load steps were used in the ANSYS run to insure faithful response. Details of the SIM0KE synthesis process appear in " Simulated Earthquake Motions Compatible with prescribed Response Spectra," MIT Dept. of Civil Engineering Report R76-4, Order No. 527. The work was performed under NSF grant ATA 74-06935.

Both horizontal and the vertical motion components were applied simultaneously in the ANSYS time history run.

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C Q l u 4 m I q b

T PART C NRC COMMENT

1. Regarding the use of the shutdown cooling system to cool the spent fuel pool, provide the following:

The licensee states that the shutdown cooling system can be aligned to provide cooling for the spent fuel pool four (4) hours after receipt of a high pool temperature alarm. The licensee did not specify the condition of the reactor (operating mode) during the time the shutdown cooling system is aligned for spent fuel cooling. Verify that the reactor will be in cold shutdown prior to alignment of the shutdown cooling system for spent fuel pool cooling.

OPPD RESPONSE:

The District has re-evaluated its position on the time required to align the shutdown cooling system for spent fuel pool cooling. This evaluation indicates that the system can be aligned within 2 to 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> after the receipt of the high pool temperature alarm. The system alignment may be accomplished within two hours only if the appropriate personnel are available onsite. However, the loss of cooling analysis is bounded by the worst case, which would now be a four hour loss of cooling. The section of the licensing submittal (Rev. 1) which covers the fuel pool heat-up following a loss of spent fuel pool coolant was re-evaluated assuming a four hour loss of cooling instead of a two hour loss. The new analysis indicates that the spent fuel can be kept adequately cooled even with a four hour loss of spent fuel pool coolf eg as shown in the revised Table 5.4.1 of Revision 2 of the licensing submittal (see Attachment A).

In regard to the condition of the reactor when utilizing the shutdown cooling system for spent fuel pool cooling refer to p.86 Section 5.2 paragraph 2. The shutdown cooling system can only be used as an alternate cooling for the spent fuel pool when the reactor is in a cold shutdown condition.

NRC COMMENT

2. The licensee stated in his submittal that the spent fuel pool temperature would be maintained below 120*F. The licensee did not use NUREG-0800, Standard Review Plan, Section 9.1.3 and Branch Technical Position ASB 9-2 for calculating the decay heat loads. Consequently, we are not sure how much conservatism is in the licensee's analysis. Therefore, provide the following information with the heat exchangers expected fouling factor and pluggage factor for the life of the plant:

A discussion of the capability and procedure to remove the spent fuel pool cooling system heat exchanger from service for tube cleaning, tube plugging or retubing. The spent fuel pool cooling system consists of two pumps and a single heat exchanger. Include in the discussion of the time available to perform these tasks without exceeding any pool temperature alarm setpoints.

OPPD RESPONSE:

The decay heat loads were calculated by using the latest revision of the NRC Branch Technical Position ASB.9-2 (reference p.82 Section 5.1 paragraph three) in the revised licensing sul.nittal; thus, the District meets the NRC criteria for calculating heat loads and sufficient conservatism was used in the analysis.

The subject heat exchanger has been designed for a 40 year service life and no problems have been identified to date that would require cleaning, plugging or retubing. Therefore, the District does not presently have a specific procedure for correcting any of these postulated problems. If such problems were to arise, procedures would be developed at that time, which woul d provide reasonable assurance that SFP temperature would be maintained below 120*F.

NRC COMMENT

3. On April 14, 1978 a generic letter was sent to all licensees which provided guidance concerning the information to be provided by the utility when requesting spent fuel pool modifications for the purpose of increasing the number of fuel bundles to be stored in the pool . The licensee submittal did not contain all of the information requested by the generic letter. Therefore, provide the following information.

a. With respect to Section 1.2 , verify that no combination of events, and/or failures will result in a keff of the spent fuel storage arrangement of greater than .95.

OPPD RESPONSE:

As stated in the revised submittal p.40-41 Section 4.6 Accident Analysis Region 1 and 2.

"The lattice of the fuel assemblies results in an undermoderated configuration so that any crushing or compaction of the fuel assemblies would tend to reduce the multiplication factor of the spent fuel pool. Therefore, the dropping of heavy objects into the fuel pool or deformations from the effects of earth-quakes or tornadoes will not produce a criticality accident. The relative

! positions of the fuel racks are maintained by perimeter bars on the rack I

therefore a significant reactivity perturbation due to rack movement is pre-cluded. The criticality safety analysis is based on the mininum spacing be-

, tween adjacent racks, and any rack movement which would increase this spacing l would tend to reduce the k. of the racks.

l Loss of all cooling systems would result in a pool water temperature increase l and under the worst possible conditions, " boiling", of the pool water. Both l higher temperature and lower density water results in negative reactivity pertu-bations for the most limiting conditions for the criticality analysis and, l therefore, loss of all cooling systems would not produce a criticality acci-l dent.

l l Because of the well founded, conservative technique used for determination of the infinite multiplication factor, there is more than reasonable assurance that this spent fuel rack design will not cause undue risk to the public health and safety resulting froa criticality considerations. .

As reported in Tables 4.4-1 and 4.4-2 the largest k of the basic cell with 1700 ppm boron in the water is 0.7411 and therefore an extremely large reac-tivity perturbation (i.e., greater than .2589 Ak) would be required to produce a criticality accident.

The fuel racks are designed to prevent a dropped fuel bundle from penetrating and occupying a position other than a normal fuel storage location. The only potential significant positive reactivity effect of a dropped fuel bundle is the reduction in axial neutron leakage that could occur if the dropped bundle came to rest in a horizontal position on top of the racks. However, the top of the active fuel stored in the racks is about 27 inches below the top of the storage boxes. This thickness of water is greater than that required to neutronically decouple the dropped fuel assembly from the assemblies stored in the rack, and therefore the maximum possible keff for Region 1 remains unchanged for this assumed accident (i.e., a value of .9415 in unborated water and .7411 with 1700 ppm boron in the pool). Similarily, for Region 2 the maximum possible keff remains unchanged for this assumed accident (i.e., a value of 0.9459 in unborated water and 0.6380 with 1700 ppm boron in the pool).

The reactivity effect of a fresh fuel assembly located adjacent to a fully loaded spent fuel storage rack has been evaluated for all postulated locations other than normal fuel storage locations. In areas where there is sufficient space to physically allow a fresh fuel assembly to be placed immediately adja-cent to a rack storage box, stand offs have been provided; so that, in all cases, the spent fuel storage rack design assures that the multiplication factor will be less than 0.95."

Note that while the second paragraph above is strictly applicable only to Region 1, a uni form temperature increase or a uniform reduction in water density in Region 2 can result in some small increase in the calculated k. of the racks. However, this is not a concern because: (1) realistically, the temperature of the coolant will always be greater in the fuel storage locations than the temperature in the water boxes and for these more realistic nonuniform conditions the reactivity pertubation will be negative, and (2) loss of all cooling systems requires more than a single failure in which case credit is allowed for the presence of soluble boron in the coolant and under these conditions the k. of the rac'4 < *e.

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ATTACHf4EllT A Revision to Pages 86, 87, 90, 91, 93 and 98 of the District's submittal dated January 21, 1983

Under these same conditions, the variation of pool temperatures with mini-mum cooling time was determined. As shown in Figure 5.2-1, the normal re-fueling pool temperature limit is less than 120*F for any realistic cool-ing time (ts > 3 days). For full core discharge, the pool temperature drops below 120*F when cooled about two weeks (ts > 16 day s) . For a higher limit of 150*F, a minimum cooling time before discharge of only about four days is required.

It is also possible to make use of the shutdown cooling system to provide alternate or supplemental cooling of the spent fuel pool during a full core discharge. This system would use low pressure safety injection pumps and one shutdown heat exchanger. With approximately triple the heat removal capacity of the normal spent fuel cooling system, this system will be available to insure that the pool temperature limits stay well below 150*F for full core discharge. This system can only be used when the reactor is in a cold shutdown condition.

5.3 FUEL POOL HEAT-UP FOLLOWING LOSS OF COOLING For fuel pool heat-up transients, it has been assumed that the pool tem-perature limits are below 120*F and 140*F for normal refueling and full 2 core discharge corditions, respectively. Section 5.2 demonstrates that the SFP cooling system can be used to maintain the SFP temperature below 120*F. For a Full Core Discharge, shutdown cooling can be used or can supplement the SFP cooling system to keep the pool temperature below 140*F. 2 Initiating cooling using the shutdown cooling system will be possible within 2-4 hours after alarms notify the operators that the pool tempera-ture limits are exceeded. The SFP alarms at 150*F. Accordingly, a four- 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> pool heatup transient is analyzed in this section. The case where cooling can be restored (for either cooling node) within four hours is also bounded by this analysis.

86 Rev. 2 05/02/83

It was censervatively assumed that the pool water heats up adiabatically without accounting for evaporation losses at the pool surface or heat losses through the concrete walls and piping exterior to the pool. The previously established limits of 11 x 106 BTU /hr,114.5"F for normal 2

refueling and 22 x 106 BTU /hr,140*F for the full core discharge were used. Reductions in heat loads and temperatures during the transient were conservatively ignored.

The thermal capacity, pCp, was estimated to be 61 BTU /ft3 *F for the pool water, and the pool volume is approximately 215,000 gallons. The ,

total thermal capacity is thus,1.75 x 106 BTU /*F.

l Starting at 114.5'F, the pool will reach 140*F in 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> for normal refueling conditions. For full core discharge conditions starting at 2 I

139'F, the pool temperatures will reach 189'F in 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />.

l If the pool is assumed to reach boiling before cooling can be restored, it will take: 1) 15.5 hrs (normal refueling) and 2) 5.8 hrs (full core discharge) for boiling to occur. Water would boil away at approximate rates of 11,340 lbm/hr and 22,680 lbm/hr for normal refueling and full core discharge conditions, respectively.

Depending on the degree of subcooling, make-up water can be added to l

compensate for these low rates of boil off to maintain the pool water level and keep the pool surface at or below boiling. Normally, make-up water is provided by the fuel transfer canal drain pumps from the safety l

make-up water at these nodest rates should the normal source malfunction.

l The radiological consequences of a boiling pool are discussed in Section 8.5, with regard to compliance to 10 CFR 100.

l l

87 Rev. 2 05/02/83

The resulting calculations for the coolant temperature rise indicated a weak dependence on the combined pressure loss coefficient, with a stronger dependence on the decay heat load and flow areas. This was documented in our earlier submittal with the Wachter design, it is also true for the par design, especially since the par design provides a lar-ger flow area for the inlet to the fuel channels. The inlet flow area is about 6 times as great as the Wachter design, however the under rack clearance is only 4.25" compared to 5" in the Wachter design. The under rack clearance was the factor which determined how much the tenperature

change would vary. The temperatures calculated in the previous submittal based on the Wachter design were found to be lower than those calculated in this submittal due to the fact that the under rack f1qw areas were different.

Fluid properties were used at 115'F for normal refueling and 140*F for FCD, the coolant increases were found to be 2

ATh = 39.1*F Normal Refueling = 27.85 ATh = 50.3*F Full Core Discharge = 20.4 ATh is the coolant change in temperature in the hot channel.

l l The AT's calculated above were based only on a natural circulation analy-sis and did not take into account the forced flow from the spargers and flow down the four walls. The only mixing that occurs in this analysis is due to the imbalance in the fluid densities in the channels and above the racks. (The density difference between the hot and cold channels balanced the resistive losses due to friction, expansion contraction, and turns.)

90 Rev.2 05/02/83

AT's calculated for a hot channel using a natural circulation technique are not much higher than the AT's calculated across the heat exchanger.

It should be noted that the heat exchanger has a large forced flow heat transfer capacity, whereas, the pool was not analyzed for any forced fl ow. Since the AT's are not abnormally high, cooling by natural circu-lation will provide adequate local cooling, and keep the pool well mixed to a nearly uniform temperature above the racks.

Peak clad temperatures were calculated using equations on calculating 2 temperature profiles along fuel rods.(5) With the heat exchangers in operation, the coolant inlet temperatures to the hottest assemblies is taken to be the average heat exchanger temperature as determined in Section 5.2.

The results, together with the maximum inlet and outlet coolant tempera-ture are summarized in Table 5.4-1. Results are also given for the postu-lated 4 hour4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> loss of cooling transient. Above the fuel racks, the local 2 saturation temperature for all cases is below saturation. For the most extreme case (FCD with 4 hr. cooling loss), the clad could reach local 2 saturation. Considering the severity of this case and the extreme conser-vatisms used in the model, the fuel racks will provide an acceptable cool-ing geometry for heat removal by natural convection.

I 5.5 GAMMA HEATING AND COOLING 0F THE POISON CHANNEL AND INTER-CELL WATER i

Some of the gamma decay energy will be absorbed in the poison channel and inter-cell water; therefore, considerations must be given to heat removal requirements for these locations.

l 91 Rev. 2 05/02/83 l .. - -. - -

REFERENCES FOR SECTION 5.0

1. NRC Branch Technical Position ASB 9-2, " Residual Decay Energy for Light Water Reactors for Long-Term Cooling," Standard Review Plan, Section 9.2.5-8a, Rev.1,1978, (This was previously designated as APCSB 9-2 prior to a branch change. There are not technical differences between the two standards).

2. American National Standards Institute ANSI N210-1976, " Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations. "

3. " Evaluation of Spent Fuel Storage Racks for Fort Calhoun Including Disassembly and Compact Storage of Individual Fuel Rods", PLG and WAI, October 1980.

4. Updated Safety Analysis Report, Fort Calhoun Unit 1.

5. " Introduction to Nuclear Engineering", Lamarsh,1975, Chapter 8.4. 2 l

93 Rev. 2 05/02/83

i TABLC 5.4.1 Results of Natural Circulation Analysis Conditions Ti n( F) Tout ( F) Tclad(*F)

Normal With HX operation 108 122.7 197 Refueling 4 hr. loss of cooling 139.6 165.0 22.6 2 Full With HX operation 126 137.8 188 Core Discharge 4 hr. loss of cooling 189 197.1 248 Notes:

1. Tn i = Coolant inlet temperature

2. Tout = Maximum coolant temperature in the hottest assemblies Tout =

Tint constant [1+sinwZ/H](5)

Constant based on heat load l Z = height of channel 2 H = Active fuel height

3. Tclad = Maximum clad temperature Tclad = Tin + C 2[1+V1+s2] m = based on mass rate of coolant 2 2 and convective resistance.

4. The fuel assembly thermal powers are taken to be 71.6 KW (2.44x106 BTU /hr) for normal refueling and 48.4 KW (.17x106 BTU /hr) for the full core discharge.

5. Above the racks, Tsat >240*F. The coolant will be subcooled for the extreme case (*). The clad could reach local saturation, but will not exceed the 240*F cstimate.

! 6. 4.25" Clearance between the floor and the racks.

l l

, 98 Rev. 2 05/02/83 l

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