ML20249B245
| ML20249B245 | |
| Person / Time | |
|---|---|
| Site: | Wolf Creek, Callaway  |
| Issue date: | 05/18/1998 |
| From: | Pellet S HOLTEC INTERNATIONAL |
| To: | |
| Shared Package | |
| ML20036E447 | List: |
| References | |
| HI-971769-ERR, HI-971769-R04-ERR, HI-971769-R4-ERR, NUDOCS 9806220223 | |
| Download: ML20249B245 (66) | |
Text
_ __ _ _ - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - -
Holtec Center,555 Lincoln Drive West, Marlton, NJ 08053 HOLTEC INTERNATIONAL Telephone (609) 797-0900 Fax (609) 797-0909 l
LICENSING REPORT FOR RERACKING OF THE CALLAWAY AND WOLF CREEK NUCLEAR PLANTS for UNION ELECTRIC AND WCNOC by HOLTEC PROJECT No: 70384 & 70814 HOLTEC REPORT No: 111-971769 REPORT CA. . ORY: A Report Class: SAFETY RELATED COMPANY PRIVATE This document version has all proprietary information removed and has replaced those sections, figures, and tables with highlighting and/or notes to designate the removal of such information. This document is to be used only in connection with the performance of work by Holtec International or its designated subcontractors.
Reproduction, publication or presentation, in whole or in part, for any other purpose by any party other than the Client is expressly forbidden.
9806220223 980527 PDR ADOCK 05000483 P PDR ,
Holtec Center,555 Lincc in Drive West, Marlton, NJ 08053 HOLTEC INTERNATIONAL Telephone (609) 797-0900 Fax (609) 797-0909 REVIEW AND CERTIFICATION im DOCUMENT NAME: I Licensing Report for Reracking of the Callaway and Wolf Creek Nuclear Plants I HOLTEC DOCUMENT 1.D. NUMBER: HI-971769 HOLTEC PROJECT NUMBER: 70384 and 70814 CUSTOMER / CLIENT: Union Electric and WCNOC REVISION BIACK REVISION' AUTilOR & REVIEWER & " QA APPROVED 8 NUMBER VTE DATE MANAGER & DATE
_ . , n O A & DATE . _
ORIGINAL '
5th" J z - 3)-9 7 ( @ I ? 7 /I S n.O.IA* 8 4 I* 7 REVISION 1
- W t-S-98 h3 'LlIl95 , h 3/3 int REVISION 2 ' '
sw 1 - us 98 ^*"^*h/e/re /NlM U9 *$- s 3-9r REVISION 3 9" g.es -n ^**'> 3/2/m $.2.;
- Su# 3 :'.15 i REVISION 4 '
1
< p,0 f . if.9 C 6 #D dr/ft s a i
" 'r'.M9K SM' S It-T h REVISION 5 REVISION 6 l
l l
l This document conforms to the requirements of the design specification and ne appbcable sections of the l
governing codes.
Note: Signatures and printed names are required in the review block. I 1
' 1
/-. t avision of this document will be ordered by the Project Manager and carried out if any of its contents i l is 1. wrially affected during evolution of this project. The determination as to the need for revision will l be nwle by the Project Manager with input from others, as deemed necessary by him.
l :
Must be Project Manager or his designee.
THE REVISION CONTROL OF THIS DOCUMENT IS BY A "
SUMMARY
OF REVISIONS LOG" PLACED BEFORE THE TEXT OF THE REPORT.
l l
, i i J
SUMMARY
OF REVISIONS Revision 4 contains the following pages, including figures:
COVER PAGE 1page REVIEW AND CERTIFICATION LOG 1page
SUMMARY
OF REVISIONS 1page TABLE OF CONTENTS 4 pages
1.0 INTRODUCTION
6 pages 2.0 OVERVIEW AND PROPOSED CAPACITY EXPANSION 23 pages 3.0 MATERIAL, HEAVY LOAD, AND CONSTRUCTION CONSIDERATIONS 15 pages 4.0 CRITICALITY SAFETY ANALYSES 35 pages APPENDIX 4A - BENCHMARK CALCULATIONS 25 pages 5.0 THERMAL-HYDRAULIC CONSIDERATIONS 27 pages 6.0 STRUCTURAL / SEISMIC CONSIDERATIONS 88 page 7.0 FUEL HANDLING AND CONSTRUCTION ACCIDENTS 22 pages 8.0 FUEL POOL STRUCTURE INTEGRITY CONSIDERATIONS 15 pages 9.0 RADIOLOGICAL EVALUATION 7 pages 10.0 INSTALLATION 11 pages 11.0 ENVIRONMENTAL COST / BENEFIT ASSESSMENT 9 pages -
TOTAL 290 pages Revision 1 contains changes from Wolf Creek and Callaway comment letters NE 98-0008, NED 98407, NED 98-010, NED 98-018. The changes are discussed in Holtec response letter 70384.SP6.
Revision 2 contains changes to typographical changes on Figure 2.1-1, Tables 5.4.1 and 9.4.1, pages 3-1,4-27,5-10,5-16, and 7-6. An additional sentence has also been added to the top of page 5-3 about emergency water makeup.
Revision 3 corrects the references of Section 3. Revision 3 also prepares two versions of the report; the version intended for NRC review contains all information with the proprietary information denoted by highlighting or notes, the other version intended for public viewing contains only highlighting with the proprietary information extracted.
Revision 4 revises the quantity of proprietary information designated in Sections 4 and 5. A typographical error was also corrected in Tabel 4.2.3 to change one symbol from " > " to "< ".
Holtec International Proprietary Information Report Hl.971769
4.0 CRITICALITY SAFETY ANALYSES 4.1 DESIGN BASES -
l The high density spent fuel storage racks for the Callaway and Wolf Creek Nuclear Power Plants are designed to assure that the effective neutron multiplication factor (k,) in the spent nuclear i fuel pool and cask loading pit is equal to or less than 0.95 with the racks fully loaded with fuel of the highest anticipated reactivity, and flooded with unborated water at the temperature within the
! operating range corresponding to the highest reactivity. The spent fuel storage racks are l
designed to accommodate any and all of the following Westinghouse fuel assembly types:
t
! 17x17 OFA ,17x17 Standard, and 17x17 Vantage SH (V5H), with a maximum nominal initial
! enrichment of 5.0 wt% 2"U and a minimum of 16 Integral Fuel Burnable Absorber (IFBA)
L ,
.'(1.5x) rods. Additional restrictions are specified to allow the storage of the aforementioned l fuel assembly types without IFBA rods. The maximum calculated reactivity includes a margin for uncertainty in reactivity calculations including mechanical tolerances. All uncertainties are ll
. statistically combined, such that the final k, will be equal to or less than 0.95 with a 95%
l probability at a 95% confidence level. Enrichments less than 5.0 wt% 2"U are also evaluated, I
and soluble boron concentrations necessary to protect against postulated accidents are I determined.
l- Applicable codes, standards, and regulations or pertinent sections thereof, include the following:
i i- e General Design Criteria 62, Prevention of Criticality in Fuel Storage and Handling.
i-
!: - USNRC Standard Review Plan,' NUREG4)800, Section 9.1.2, Spent Fuel Storage, Rev. 3 - July 1981 e- USNRC letter of April 14,1978, to all Power Reactor Licensees - OT Position for l Review and Acceptance of Spent Fuel Storage and Handling Applications, including modifk:ation letter dated January 18,1979.
(
I
' i I The OFA designation is used generically throughout this section and includes V-5 and V+ fuel.
Ilottec International Proprietary information : 41 Report Hi-971769 1
, T e
USNRC Regulatory Guide 1.13, Spent Fuel Storage Facility Design Basis, Rev. 2 (proposed), December 1981.
ANSI ANS-8.17-1984, Criticality Safety Criteria for the Handling, Storage and Transportation of LWR Fuel Outside Reactors.
USNRC guidelines and the applicable ANSI standards specify that the maximum effective rnultiplication factor, km , including bias, uncertainties, and calculational statistics, shall be less than or equal to 0.95, with 95 % probability at the 95 % confidence level.
To assure that the true reactivity will always be less than the calculated maximum reactivity, the following conservative assumptions were made:
Moderator is unborated water at a temperature that results in the highest reactivity (4 C, corresponding to the maximum possible moderator density).
No soluble poison or control rods are assumed to be present for normal operations, although the additional margin due to the presence of soluble boron is identified, e
The effective multiplication factor of an infinite radial array of fuel assemblies was ,
used except for the assessment of peripheral effects and certain abnormal / accident conditions where neutron leakage is inherent.
Neutron absorption in minor structural members is conservatively neglected, i.e.,
spacer grids are replaced by water, e
Depletion calculations assume conservative operating conditions; highest fuel and moderator temperature and an allowance for the soluble boron concentrations during in-core operation.
The assemblies with IFBA rods are assumed to contain the minimum possible number of IFBA rods (i.e.,16) in a conservative loading pattern with a conservative length of 120 inches. Further, the IFBA loading is assumed to be 1.5xIFBA (2.25 mg B-10/ inch) with an uncertainty of 5%. The IFBA loading used in the analyses is reduced by the 5% uncertainty.
Holtec international Proprietary information 4-2 Repon 111-971769
The spent fuel starage racks are designed to accommodate any and all of the following Westinghouse fuel assembly types: 17x17 OFA,17x17 Standard, and 17x17 Vantage 5H (V511), with a maximum initial enrichment of 5.0 wt% 235U. To assure the acceptability of the racks for storage of any and all of the above assembly types, the most reactive assembly type l was identified via independent criticality calculations. The results of these calculations show that at zero burnup the 17x17 OFA assembly has the greatest reactivity in the storage racks, and thus, is the design basis fuel assembly. The Mixed-Zene Three-Region (MZTR) configuration uses fuel assemblies with high d <, charge burnup as barrier fuel to isolate fresh fuel assemblies in order to achieve an acceptable k, in the spent fuel pool and cask loading pit.
Three separate storage regions are provided, with independent criteria defining the highest potential reactivity in each of the three regions as follows:
Region 1 is designed to accommodate new un-irradiated (fresh) fuel with a maximum nominal enrichment of 5.0 wt% 235U and a minimum of 16 IFBA (1.5x) rods, or fuel of equivalent reactivity (e.g.,4.6 wt% 235U maximum enrichment without IFBA rods). Further, Region 1 cells on the periphery of the pool, that are adjacent to a concrete wall, may accommodate fresh fuel assemblies with maximum nominal enrichment of 5.0 wt% 235U and no IFBA rods.
Region 2 is designed to accommodate fuel with a maximum nominal initial enrichment of 5.0 wt% 235U and high (2 50 mwd /kgU) discharge fuel burnup, or fuel of initial enrichment and burnup combinations yielding an equivalent reactivity.
Region 2 locations are used to isolate Region 1 fuel assemblies from other Region 1 and Region 3 fuel assemblies.
Region 3 is designed to accommodate fuel with a maximsm nominal initial enrichment of 5.0 wt% '"? and typical (40.75 s burnup s 50 mwd /kgU) discharge fuel burnup, but can accommodate any spent fuel with discharge fuel burnup greater than or equal to 40.75 mwd /kgU. Additionally, fuel of initial enrichment and burnup combinations yielding an equivalent reactivity are acceptable for storage in Region 3.
The water in the spent fuel storage pool normally contains soluble boron which would result in a large sub-criticality margin under actual operating conditions. However, the NRC guidelines, l
Ilottec International Proprietary information 4-3 Report H1-971769
__... i
i based upon the accident condition in which all soluble poison is assumed to have been lost, specify that the limiting 4 of 0.95 for normal storage be evaluated for the accident condition that assumes the loss of soluble boron. The double contingency principle of ANSI N-16.1-1975 and of the April 1978 NRC letter allows credit for soluble boron under other abnormal or accident conditions, since only a single independent accident need be considered at one time.
Consequences of abnormal and accident conditions have also been evaluated, where " abnormal" refers to conditions whid may reasonably be expected to occur during the lifetime of the plant and " accident" refers to conditions which are not expected to occur but nevertheless must be protected against.
l 1 I
l l
l I
Holtec International Proprietary Information 4-4 Repon HI-971769
?4.2L .
SUMMARY
OF CRITICALITY ANALYSES 4.2.1 Normal Oneratina Conditions In the MZTR configuration, the fresh fuel cells (Region 1) are located alternately along the psyhery of the spent fuel pool (where neutron leakage reduces reactivity) or along the boundary between two storage modules (where the water gap provides a flux-trap which reduces reactivi-ty). High burnup fuel in Region 2 affords a low-reactivity barrier between fresh fuel assemblies and Region 3 fuel ofintermediate bumup.
' Numerous configurations of the various assemblies within the spent fuel pool and cask loading pit are possible. The criteria for determining an acceptable loading arrangement in the MZTR
< configuration for fuel of different burnups are as follows:
Region 1 cells are only located along the outside periphery of the storage modules and must be separated by one or more Region 2 (burnup > or = 50 mwd /kgU for 5.0 wt% 2"U, or equivalent burnup/ enrichment combinations) cells.
Region 1 cells may be located directly across from one another when separated by e water gap. Along the interface between storage modules the water gap is 1.5" +/-
1/8"(excluding sheathing).
The outer rows of alternating Region 1 and Region 2 cells must be further separated (isolated) from the internal Region 3 cells by one or more Region 2 cells.
. - Fresh fuel assemblies without IFBA rods and a maximum enrichment of 5.0 wt% 2nU ,
may be stored in any periphery Region 1 cell location that is next to a concrete wall.
Figure 4.2.1 shows a representative sample of the multitude ofpossible MZTR pool cconfigurations that meet the previously stated criteria.
Prior to approaching the reactor end-of-life, not all storage cells are needed for spent fuel.
'Iherefore, an alternative'(interim) configuration may be used in which the cells of selected modules may be loaded in a checkerboard panem of fresh fuel (or spent fuel of any burnup) with Holtec international F;oprietary information _ 4-5 Repon Hi-971769
empty cells. Figure 4.2.2 illustrates the concept of using a checkerboard loading pattern. A checkerboard configuration is intended primarily to develop a simple configuration of Region 1 cells and facilitate storage of fresh (unburned) and low burnup fuel.
The principles involved in the design and specification of an acceptable loading arrangement in the interim checkerboard configuration are as follows:
Fuel with maximum nominal enrichment of 5 wt% 235U and a minimum of 16 IFBA (1.5x) rods, or fuel of equivalent reactivity (e.g.,4.6 wt% 235U maximum
]
enrichment without IFBA rods), is placed in an alternating checkerboard style pattern with empty cells (i.e., fuel assemblies are surrounded on all four sides by empty cells).
- Fuel assemblies may not be located directly across from one another, even when separated by a water gap.
e- So long as the checkerboard pattern is maintained in a linear array greater than or equal to 2x2, the arrangement may be used anywhere in the pool. More than one ,
checkerboard pattern may be used, as long as the limitations discussed herein are adhered to. .
- A checkerboard region may be bounded by either a water gap, empty rack cells, Region 2 fuel assemblies, or Region 3 fuel assemblies.
- MZTR and checkerboard storage shall not be developed within the same rack.
Non-fueled items such as trash baskets and dummy fuel assemblies may be stored anywhere in the spent fuel pool or cask loading pit, while damaged fuel storage baskets shall be stored in j
any cell which allows fuel assembly storage.
Figum 4.2.3 defines the acceptable burnup domains for spent fuel and illustrates the limiting burnup for fuel of various initial enrichments for both Region 2 (upper curve) and Region 3
- (lower curve), both of which assume that the fresh fuel (Region 1) has a maximum nominal enrichment of 5.0 wt% 235U. Criticality _ analyses demonstrate that the most reactive configuration Holtec International Proprietary Information 4-6 Report Hi-971769 l
l
__ . _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . . _ . . _ _ _ . _ _ _ _. . _ _ _ _ _ . _ _ _ _ _ _ . _ _ _ _______._a
t occurs along the boundary between modules where the water gap affords a neutron flux trap.
Along the periphery of the modules facing the concrete wall of the pool, the reactivity is substan-tially lower due to neutron leakage. The bounding criticality analyses are summartzed in Table i
4.2.1 for the design basis MZTR storage configuration and in Table 4.2.2 for the interim checkerboard storage configuration (in both cases, the single accident condition of the loss of all-soluble boron is assumed). The calculated maximum reactivity of 0.943 (corresponding to the design basis MZTR storane configuration) is within the regulatory limit of 0.95. This maximum l ' reactivity includes calculational uncertainties and uncertainties in reactivity due to manufacturing tolerances (95 % probability at the 95% confidence level), an allowance for uncertainty in depletion calculations, and the evaluated effect of the axial distribution in burnup.
L l The value of1% ni Table 4.2.1 assumes no soluble boron to be present. For normal operations, a minimum soluble boron concentration of 2000 ppm is maintained in both the Callaway and L Wolf Creek spent fuel pools. This concentration of soluble boron reduces the maximum reactivity by 0.210 Ak, thus providing a large safety margin for sub-criticality.
As cooling time increases in long-term storage, decay of *Pu (and growth of"Am) results in a continuous decrease in reactivity, which provides an increasing sub-criticality margin with time.
l No credit is taken for this decrease in reactivity other tlan to indicate conservatism in the l calculations. -
l The burnup criteria identified above (Figure 4.2.3) for acceptable storage in Region 2 and Region 3 can be implemented in appropriate administrative proceduits to assure verified burnup as specified in the proposed Regulatory Guide 1.13, Revision 2. Soluble poison is present in the pool water during fuel handlitig operations, and this serves as a further margin of safety and as a
. precaution in the event of fuel misplacement during fuel handling operations.
l.
T 'Ihe thick base plates on the rack modules extend beyond the storage cells and provide assurance that the necessary water gap between roodules is maintained.
Holtec international Proprietary information 4-7 Report HI-971769
For convenience, the minimum (limiting) burnup data in Figure 4.2.3 for unrestricted storage may be described as a function of the initial enrichment, E, in wt% 2"U by fitted polynomial expressions as follows:
Region 2 storage: B = -0.063xE' + 0.97xE'- 6.09xE' + 29.97xE - 28.95, and Reelon 3 storaue: B = -0.049xE' + 0.74xE' - 4.65xE' + 25.10xE - 30.19, 2
where B is the minimum burnup in mwd /kgU and E is the enrichment in wt% "U (for initial enrichments from 2.0 to 5.0 wt% 2"U). Fuel assemblies with enrichments less than 2.0 wt%
zuU will conservatively be required to meet the burnup requirements of 2.0 wt% 2"U assemblies as shown in Fig 4.2.3.
4.2.2 Abnormal nM Accident CoMitions Although credit for the soluble poison normally present in the spent fuel pool water is pennitted.
under abnormal or accident conditions, most abnormal or accident conditions will not result in exceeding the limiting reactivity (k, of 0.95) even in the absence of soluble poison. The effects on reactivity of credible abnormal and accident conditions are discussed in Section 4.6 and '
summanzed in Table 4.2.3. Of these abnormal or accident conditions, only two have the potential for a more than negligible positive reactivity effect. These include: (1) the inadvertent misplacement of a fresh fuel assembly and (2) the mis-location of a fresh fuel assembly into a position external and adjacent to a storage rack.
The inadvertent misplacement of a fresh fuel assembly has the potential for exceeding the limiting reactivity, should there be a concunent and independent accident condition resulting in the loss of all soluble poison. Assuring the presence of soiuble poison during fuel handling operations will preclude the possibility of the simultaneous occurrence of the two independent accident conditions. The largest reactivity increase would occur if a fresh fuel assembly of 5.0 wt%2 "U enrichment were to be inadvertently loaded into an empty cell in the checkerboard configuration Holtes international Proprietary Information 4-8 Report Hi-971769
l with the remainder of the rack fully loaded with fuel of the highest permissible reactivity. For the MZTR configuration, the situation in which a fresh fuel assembly of 5.0 wt% 2"U enrichment is inadvertently loaded into a Region 2 location with the remainder of the rack fully loaded with fuel of the highest permissible reactivity is slightly less reactive, but also exceeds the l
l limiting reactivity in the absence of soluble boron. Under these accident conditions, credit for the presence of soluble poison is permitted by the NRC guidelines,t and calculations indicate that 500 ppm soluble boron would be adequate to reduce the k, of either case to below the reference k, value (Table 4.2.1). This soluble boron concentration bounds all other accidents and is well below the 2000 ppm soluble boron concentration that is maintained in both the Callaway and Wolf Creek spent fuel pools.
The pool , layouts shown in Figs. 4.2.1 and 4.2.2 reveal that it is possible for a fuel assembly to be dropped or mis-located such that it may be situated outside and adjacent to a storage rack.
The calculated k, value for the worst case situation exceeds the limit on reactivity in the absence of soluble boron. However, this case is less severe than the misplaced fresh fuel assembly accident, and thus, requires less than 500 ppm soluble boron to reduce the k, to the reference value (Table 4.2.1).
I l
I l
' Double contingency principle of ANSI N16.1 1975, as specified in the April 14,1978 NRC letter (Section 1.2) and implied in the proposed revision to Reg. Guide 1.13 (Section 1.4, Appendix A).
Holtec International Proprietary Information 49 Report 111-971769
( _ _ _ __ __ _ _ . _ _ - - _ _ _ _ _ _ _ _ - - - _ - _ . _ _ _ --
4.3 REFERENCE FUEL STORAGE CELLS
'4.3.1 Reference Fuel Assembly 1 The spent fuel storage racks are designed to accommodate any and all of the following ,
Westinghouse fuel assembly types: 17x17 OFA,17x17 Standard, and 17x17 Vantage 5H !
(V5H), with a maximum nominal initial enrichment of 5.0 wt% zuU. Additional restrictions
~J
- are specified to allow the storage of any of the aforementioned fuel assembly types without IFBA rods. To assure the acceptability of the racks for storage of any and all of the above assembly types, the most reactive assembly type was identified by independent criticality calculations, The results of these calculations show that at zero burnup the 17x17 OFA assembly has the greatest reactivity in the storage racks, and thus, is the design bash fuel assembly. The Westinghouse OFA is a 17 x 17 array of fuel rods with 25 rods replaced by 24 control rod guide tubes and 1 instrument thimble, and is depicted in Figure 4.3.1. Table 4.3.1 summarizes the fuel assembly design specifications.
At burnups beyond approximately 25 mwd /kgU, the 17x17 Standard and 17x17 Vantage SH become the most reactive assembly types. These two assembly types are essentially identical. l Therefore, for the determination of the equivalent enrichments associated with Regions 2 and 3,-
the reactivity of the V5H assembly was related to an initial enrichment for the 17x17 OFA assembly.
i i
The fresh fuel assemblies were assumed to contain the minimum possible number ofIFBA rods l
(i.e.,16) in a conservative loading pattern with a conservative length of 120 inches. The IFBA l
rods are characterized by a thin ZrB2 coating on the outside of the fuel pellets. Because B-10 l l
. in ZrB 2is a strong neutron absorber, it reduces the assembly reactivity, and thus, enables the ;
storage of fuel with high initial enrichment. The IFBA loading was assumed to be 2.25 mg B- !
10/ inch with an uncertainty of 5%. The IFBA loading was assumed to be reduced by the 5%
uncertainty in this analysis. With 16 IFBA rods present, the reactivity of the assembly does 1 not exhibit a peak with burnup, and thus the calculated reactivity of the fresh assembly is I bounding. . The IFBA rods are modeled in the fresh fuel assemblies only; no credit is taken for
- residual IFBA in the Region 2 and Region 3 fuel assemblies.
i i
- Holtec International Proprietary information 4-10 Report Hi-971769 !
I' i
j 4.3.2 High Density Fuel Storane Cells De nominal spent fuel storage cell used for the criticality analyses of the Callaway and Wolf Creek l spent fuel storage cells is shown in Figure 4.3.1. Stainless steel boxes are arranged in an attemating pattern such that the connection of the box corners form.sto. rag
. _e cells between those of g:
the stainless steel boxes. The walls of the stainless steel boxes are gighuj inches thick with a
~-
boral panel (attached by al$mmgy stainless steel sheathing) centered on each side. The box I.D.
pen:Mugg t?mwEwM i is dysdy inches. Peripheral cells use a Ld inch stainless steel sheathing on the outside I wall to attach the Bomi panel. The fuel assemblies a,re normally located in the center of each r .
i storage cell on a nominallattice spacing of 8.99 !_ , 3 inches. The Boral absorber has a nummmeno per mme m thickness of jbbgj inches and a nominal B-10 areal' density of g6 1, ij g B-2 10/cm minimum).
l l
1 l I i
n i !
l
- Holtec International Proprietary Information 4-11 Report Hi-971769 I'
L
L 4.4 ANALYTICAL METHODOLOGY
~ 4.4.i Reference Desian Calcuhtions The principal method for criticality analysis of the high density storage racks is the three-dimensional Monte Carlo KENO 5a [4.4.1] code, as developed by the Oak Ridge National 12boratory as part of the SCALE 4.3 package. Independent verification calculations were
~ performed with the MCNP (version 4a) code [4.4.2], a continuous energy three-dimensional Monte Carlo code developed at the Los Alamos National Laboratory. The KEN 05a calculations
'used the 238-group SCALE cross-section library and NITAWL [4.4.3] for 2"U resonance 1 shielding e#ects - (Nordheim integral treatment). Benchmark calculations, presented in Appendix
- A, indicate a bias of 0.0030 with an uncertainty ofi 0.0012 for KENO 5a and 0.0009 i 0.0011 for MCNP4a, both evaluated at the 95% probability,95% confidence level [4.4.4].
Fuel depletion analyses during core operation were performed with CASMO-3, a two-
- dimensional multigroup transport theory code based on capture probabilities [4.4.5 - 4.4.7].
Restarting the CASMO-3 calculations in the storage rack geometry at 4*C yields the two-
~ dimensional infinite multiplication factor (k.) for the storage rack. Parallel calculations with
' CASMO-3 for the storage rack at various enrichments enable a reactivity equivalent enrichment-(fresh fuel) to be determined that provides the same reactivity in the rack as the depleted fuel. >
- CASMO-3 was also used to determine the small reactivity uncertainties (differential calculations) of manufacturing tolerances.
1 In the geometric models used for the calculations, each fuel rod and its cladding were described explicitly and reflecting boundary conditions were used in the radial direction which has the eNect of creating an infinite radial array of storage cells. KENO 5a and MCNP4a Monte Carlo calculations inherently include a statistical uncertainty due to the random nature of neutron l tracking. To minimize the statistical uncertainty of the KENO 5a<alculated reactivity and to
' assure convergence, a minimum of 5 million neutron histories in 1,000 generations of 5,000
. neutrons per generation were accumulated in each single assembly infinite array calculation and a L minimum of 20 million neutron histories in 2,000 generations of 10,000 neutrons per generation
'were een=nh ed t in each multiple assembly (MZTR and checkerboard) configuration.
Holtec Intemational Proprietary information 4-12 Repon HI-971769 l
L________________________________ . _ _ _ _
Figure 4.4.1 represents the reference MZTR geometric model used in the KENO 5a calculations.
This figure is intended to show the arrangement of fuel assemblies modeled, and not the specific details of the model. With reflecting boundary conditions, this model effectively describes the entire pool in the MZTR configuration, including the water gap between storage modules. In the axial direction, the full length 144-inch fuel assembly was described assuming 30-cm water reflector, top and bottom. In addition, the axial variation in burnup was explicitly modeled and resulted in a slightly lower reactivity than the reference design calculation (which assumes uniform axial burnup). Figure 4.4.2 represents the reference checkerboard geometric model used in the KENO 5a calculations. With reflecting boundary conditions, this model effectively describes the entire pool in the checkerboard configuration, including the water gap between storage modules. These large models were also used to investigate uncertainties in the configurations and the consequences of potential accident conditions, including a misplaced fresh fuel assembly.
Because NITAWL-KENO 5a does not have burnup capability, burned fuel was represented by fuel of equivalent enrichment as determined by CASMO-3 calculations in the storage cell (i.e. an l enrichment which yields the same reactivity in the storage cell as the burned fuel). In tracking long-term (30-year) reactivity effects of spent fuel, previous CASMO-3 calculations have
{
demonstrated a continuous reduction in reactivity with time (after Xe decay)[4.4.8] due primarily I to "Pu decay and "Am growth. -
4.4.2 Fuel Burnuo Calculations and Uncertainties I
CASMO-3 was used for burnup calculations in the hot operating condition. CASMO-3 has been extensively benchmarked [4.4.7, 4.4.9] against cold, clean, critical experiments (including plutonium-bearing fuel), Monte Carlo calculations, reactor operations, and heavy element concentrations in irradiated fuel. In addition to burnup calculations, CASMO-3 was used for evaluating the small reactivity increments (by differential calculations) associated with manufacturing tolerances and for determining temperature effects.
In the CASMO-3 geometric model, each fuel rod and its cladding were described explicitly and ;
reflective boundary conditions were used at the centerline of the Boral and steel plates between storage cells. These boundary conditions have the effect of creating an infinite array of storage cells in the X-Y plane and provide a conservative estimate of the uncertainties in reactivity attributed to manufacturing tolerances.
Holtec International Proprietary information 4-13 Report H1-971769
Conservative assumptions of moderator and fuel temperatures and the average operating soluble boron concentrations were used to assure the highest plutonium production and hence l conservatively high values of reactivity during burnup. Since critical experiment data with spent fuel is not avcdable for determining the uncertainty in depletion calculations, an allowance for t
uncertainty in reactivity was assigned based upon the assumption of 5% uncertainty in burnup.
At the design basis burnups of 40.75 and 50 mwd /kgU, the uncertainties in burnup are i 2.04 and 2.5 mwd /kgU respectively. These uncertainties correspond to approximately 0.013 Ak and 0.016 Ak in the fuel infinite multiplication factor.
To evaluate the reactivity consequences of the uncertainties in burnup, independent MZTR calculations were made with fuel of 38.5 and 47.5 mwd /kgU burnup in Regions 2 and 3, and the incremental change from the reference burnups assumed to represent the net uncertainties in reactivity attributable to uncertainty in depletion calculations. These calculations resulted in an incremental reactivity uncertainty in k, of i 0.0056 Ak for Region 2 and i 0.0001 Ak for Region 3. These effects would be lower for lower initial enrichments and burnups. The fresh unburned fuel in Region I strongly dominates the reactivity which tends to minimize the reactivity consequences of uncertainties in depletion calculations. The allowance for uncertainty in the burnup calculations is believed to be conservative, particularly in view of the substantial reactivity decrease with time as the spent fuel ages. -
4.4.3 Effect of Axial Burnun Distribution Initially, fuel loaded into the reactor will burn with a slightly skewed cosine power distribution.
As burnup progresses, the burnup distribution will tend to flatten, becoming more highly burned in the central regions than in the upper and lower regions. At high burnup, the more reactive fuel near the ends of the fuel assembly (less than average burnup) occurs in regions of high neutron leakage. Consequently, it is expected that over most of the burnup history, fuel assemblies with distributed burnups will exhibit a slightly lower reactivity than that calculated for the uniform average burnup. As burnup progresses, the distribution, to some extent, tends to be t
The majority of the uncertainty in depletion calculations derives from uncenainties in fuel and moderator temperatures and the effect of reactivity control methods (e.g., soluble boron). For depletion calculations, bounding values of these operating parameters were assumed to assure conservative results in the analyses.
lioltec international Proprietary Information 4-14 Report 111-971769
l self-regulating as controlled by the axial power distribution, precluding the existence of large regions of significantly reduced burnup.' Among others, Turner [4.4.10] has provided generic analytic results of the axial burnup effect based upon calculated and measured axial burnap distributions. These analyses confirm the minor and generally negative reactivity effect of the f
axially distributed burnup. j Based on axial burnup distributions of spent fuel (axial burnup data for assemblies from the Callaway plant with average burnups of 50.10 and 36.84 mwd /kgU were normalized to the Region 2 and 3 burnups,50 and 40.75 mwd /kgU, respectively), three-dimensional KENO 5a calculations were performed. In these calculations, the axial height of the Region 2 and 3 fuel was divided into 5 axial zones, each with an average enrichment equivalent to the burnup of that zone. The selection of the five axial zones was based on the shapes of the axial burnup distributions. The resulting k,,, was 0.007 Ak less than the reference k,n (which assumes
. uniform axial burnup). Fuel of lower initial enrichments (and lower burnup) would have a more negative reactivity effect as a result of the axial variation in burnup. These estimates are believed to be conservative since smaller axial increments in the calculations have been shown to result in lower incremental reactivities [4.4.10].
I l
l l l
Holtec international Proprietary information 4-15 Report Hi-971769 l l
1 L 4.5 CRITICALITY ANALYSES AND TOLERANCES L
o 1
1 l
4.5.1 Nomind Design l For the nominal MZTR storage configuration, the bounding criticality analyses are summarized in Table 4.2.1. The NITAWL-KEN 05a calculated k,, value is combined with all the known L uncertainties and corrected for bias and temperature (see Section 4.6.1 for temperature correction),
to determine the maximum k,, value with a 95% probability at the 95 % confiderice level [4.4.4]..
For the interim loading pattern of fresh fuel checkerboarded with empty cells, the bounding criticality analyses are summarized in Table 4.2.2. An alternate calculation with a 2x2 checkerboard pattern bordered on all sides with Region 3 fuel assemblies resulted in a maximum k,,
of 0.903 with a 95% probability at the 95 % confidence level. Therefore, the checkerboard loading L _ pattern may be used anywhere in any module provided that the checkerboard pattem is a linear array greater than or equal to 2x2 and is bordered by any of the following: the water gap between
- j. rack modules, the water gap between a rack module and the pool wall, empty rack cells, Region 2 fuel assemblies, and/or Region 3 fuel assemblies.
1 L 4.5.2 Uncertainties Due to Manufacturing Tolerances l
l The uncertainties due to manufacturing tolerances are summarized in Table 4.5.1 and discussed below.
4.5.2.1 Boron loading Tolerances
. n The Boral absorber panels used in the storage cells are nominally (mIss d inches thick,ingd inches l
l wide and C inches long, with a nominal B-10 areal density of hd g/cm 2. The vendors 2
manufacturing tolerance limit is %vm magj g/cm in B-10 content which assures that at any point, emm 2 the minimum B-10 areal density will not be less than k j g/cm . Differential CASMO-3 calculations for an infinite array of fresh assemblies with the minimum tolerance B-10 loading g . , -
results in an incremental reactivity uncertainty of bd,iLd A k. This value was conservatively assumed to be the B-10 loading uncertaint, l
l Holtcc International Proprietary Infonnation 4 16 Report HI-971769
l l
f
- 4.5.2.2 Boral Width Tolerance The reference storage cell design uses a Boral panel with a width of[i;[3. ' hj$ inches. Forthe
~ tolerance of @yanswzbgyj g inches, the differential CASMO-3 calculated reactivity uncertainty is avvas hkda A k-4.5.2.3 Telerances in Cell 12ttice Spacing The manufacturing tolerance on the inner box dimension, which directly affects the storage cell
- lattice spacing between fuel assemblies, is ! [ ;,2 inches. This corresponds to an uncertainty in dbad reactivity of pyy2MM@TQ$j$
r k A , determined by differential CASMO-3 calculations.
4.5.2.4 Stainless Steel Thickness Tolerances
<. ~ , , ~ ,
The nominal stainless steel thickness is [ _ 1j inches for the stainless steel box. The reactivity uncertainty of the expected stainless steel thickness tolerances was calculated with CASMO-3 to be b~ d Ak .
4.5.2.5 Fuel Enrichment and Density Tolerances The design maximum enrichment is 5.0 + 0.05 wt% 2"U. Separate CASMO-3 burnup calculations j 2
were made for fuel of the maximum enrichment (5.05 wt% nU) and for the maximum UO2 density (10.61 g/cm'). Reactivities in the storage cell were then calculated using the restart capability in CASMO-3. For fresh fuel, the incremental reactivity uncertainties were01 0023Ak for the enrichment tolerance and + 0.0026Ak for the tolerance in fuel density.
l i
- l. i
~
l 1
l i
)
Holtec International Proprietary Infonnation 4-17 Report Hi-971769
l 4.5.3 Water-Gao Soacine Between Modules i
The water-gap between modules, which is 1.5 inches (excluding sheathing), constitutes a neutron {
flux-trap for the storage cells of facing racks. KENO 5a calculations were made with the l reference MZTR model to determine the uncenainty associated with a water-gap tolerance. Due to the asynunetries in the MZTR pool configuration, the effect of the horizontal and vertical water gaps (see Fig. 4.4.1) were calculated separately. From these calculations, it was determined that the incremental reactivity consequence (uncertainty) for a water-gap tolerance of i 1/8 inches is i 0.0014 Ak (horizontal gap) and i 0.0003 Ak (venical gap). The racks are constructed with the base plate extending beyond the edge of the cells which assures that the minimum spacing between storage modules (1.5 inches, excluding sheathing) is maintained under all credible conditions. <
j 4.5.4 Eccentric Fuel Positioning The fuel assembly is assumed to be centered in the storage rack cell. Calculations were made i
- using KENO 5a assuming the fuel assemblies were located in the corners of the storage rack cells (four-assembly clusters at the closest possible approach). These calculations indicated that the reactivity effect is small and negative. Therefore, the reference case in which the fuel assemblies are centered is controlling and no uncertainty for eccentricity is necessary.
Holtec International Proprietary information 4-18 Report Hi-971769
1-
.4.6 ABNORMAL AND ACCIDENT CONDITIONS !
i
- 4.6.1 Temperature and Water Density Effects The moderator temperature coefficient of reactivity is negative; a moderator temperature of 4 C (39 F) was assumed for the reference calculations, which assures that the true reactivity will !
' always be lower over the expected re3e of water temperatures. Temperature effects on
.{
reactivity have been calculated (CASMO-3) and the results are shown in Table 4.6.1. In !
addition, the introduction of voids in the water internal to the storage cell (to simulate boiling)
{
decreased reactivity, as shown in Table 4.6.1. !
With soluble boron present, the temperature coefficients of reactivity would differ from those l
listed in Table 4.6.1. However, the reactivities would also be substamially lower at all l
I
' temperatures with soluble boron present. The data in Table 4.6.1 is pertinent to the higher-reactivity unborated case.
For the dominant Region 1 tuel, the value of Ak between calculations at 20*C and 4*C is O.0020 Ak. Since the KENO 5a code cannot properly handle temperature dependence, all I KENO 5a calculations were performed at 20*C and a temperature correction factor (+0.0020. i Ak) was applied to the results. l
- l. .4.6.2 Lateral Rack' Movement i
- ~ The possibility of reductions in the rack-to-rack gaps and the resulting criticality consequences !
i have also been reviewed. Criticality evaluations are sensitive to these gap dimensions, since l L ' the inter-rack gaps provide a flux trap which reduces the reactivity. Rack to rack gap reductions are a concern subsequent to dynamic events which are severe enough to displace the racks laterally or produce fuel to rack cell wall impacts of sufficient magnitude to exceed cell l- ~ wall material yield strength (i.e., produce plastic deformation).
u F The criticality analyses are based on the minimum nominal rack to rack gap of 1.5 inches (excluding sheathing). Thus, the outer sheathing wall-to-outer sheathing wall gap is 1.35
. Holtec International Proprietary Information 4 19 Report HI-971769 i
L__ = __:___ :___--- _ __--. - _ _ _ - -
~ inches. This gap dimension is maintained during initial installation and subsequent to dynamic loadings, and is ensured by fabrication of the 3/4 inch base plate extensions on each rack.
Momentary reductions in these gaps may be caused by the swaying of the tops of the racks during seismic events, during which the tops of the cells may actually come into contact. Even
- under these circumstances, the bottoms of the cells in adjacent racks are still maintained at the 1.5 inch dimension due to the base-plate extensions. Transient reduction in the inter-rack gap l dimension below 1.5 inches is acceptable because of the presence of soluble boron which may be credited during seismic events. Additionally, a time-history plot of the inter-rack gaps (see Figs. 6.8.1 through 6.8.3) indicates that the gaps are reduced for a very short period of time before being restored to the minimum of 1.5 inches.
4.6.3 Rack-Gao Channes Another consideration which could potentially reduce the inter-rack gap is the impact of the l fuel assembly on the inside of the cell wall during seismic events. If these impacts are of sufficient magnitude to allow plastic deformation of the cell wall membrane, then permanent i displacement of the cell would take place, thus reducing the inter-rack gap. The largest fuel
{
assembly to cell wall impact load is determined to be 840 pounds (see subsection 6.8.4.3). -
j Evaluations on the local cell wall integrity (see subsection 6.9.4.a) have determined that the load required to produce permanent deformation (i.e., exceed the cell membrane material yield strength) exceeds the calculated load of 840 pounds by a factor of approximately 4. Therefore, i there are no criticality concerns related to the reductions in inter-rack gaps from plastic
{
deformation of the cell wall.
l l
4.6.4 Abnormallocation of a Fuel Auembly L !
In the MZTR configuration, the abnormal location of a fresh unirradiated fuel assembly of 5.0 l wt% 2"U enrichment could, in the absence of soluble poison, result in exceeding the design reactivity limitation (k, of 0.95). This would occur if a fresh fuel assembly of the highest i permissible enrichment were to be inadvertently loaded into either a Region 2 or Region 3 storage cell.' Calculations (KENO 5a) confirmed that the highest reactivity, including uncenainties, for the worst case postulated accident condition (fresh fuel assembly in Region 2) l , Holtec International Proprietary information 4-20 Report HI-971769 L _-_-_-____ ___-_-__
l l
would exceed the limit on reactivity in the absence of soluble boron. Soluble boron in the spent !
l fuel pool water, for which credit is permir x1 under these accident conditions, would assure that l the reactivity is maintained substantially less than the design limitation. Calculations indicate that j a soluble poison concentration of 440 ppm boron would be required to limit the maximum reactivity to the reference k, value (Table 4.2.1), including all uncertainties and biases, under this maximum postulated accident condition.
In the checkerboard configuration, the worst case postulated accident condition (fre:h fuel l
assembly inadvertently loaded into an empt) cell) would also exceed the limit on reactivity in the absence of soluble boron. Soluble boron in the spent fuel pool water, for which credit is permitted under these accident conditions, would assure that the reactivity is maintained substantially less than the design limitation. Calculations indicate that a soluble poison concentra-tion of 500 ppm boron would be required to limit the maximum reactivity to the reference k, value (Table 4.2.1), including all uncertainties and biases, under this maximum postulated accident condition.
i l 4.6.5 Drooned Fuel Assembly For the case in which a fuel assembly is assumed to be dropped on top of a rack, the fuel -
assembly will come to rest horizontally on top of the rack with a minimum separation distance from the active fuel in the rack of more than 12 inches, including the potential deformation under seismic or accident conditions. At this separation distance, the effect on reactivity is insignificant. Furthermore, the soluble boron in the pool water assures that the tme reactivity is always less than the limiting value for this dropped fuel accident.
It is possible for a fuel assembly to be mis-located adjacent to a storage rack in the northwest l (area near the opening to the fuel transfer canal) and southeast (area near the opening to the cask loading pit) corners of the spent fuel storage pool. The worst case postulated accidents are: (1) in the southeast corner of the MZTR configuration, a fresh fuel assembly could be dropped anxi l come to rest in the corner made up by a fresh assembly to the north and a Region 2 assembly to the west and (2) in the northwest corner of the checkerboard configuration (shown in Fig. 4.2.2),
a fresh fuel assembly could be mis-located in a corner with fresh assemblies on two sides.
The k, values for these two cases are very similar, and exceed the limit on reactivity in tie absence of soluble boron. Soluble boron in the spent fuel pool water, for which credit is Holtec International Proprietary information 4-21 Report 111-971769 I
l
l l
i f
, permitted under these accioent conditions, would assure that the reactivity is maintained substantially less than the design limitation. These cases are less severe than the misplaced fresh I fuel assembly accidents, and thus, are bounded by them.
t 1
i l
l l l l
l l t
I 1
?
l l
1 l I l
l i
Holtec International Proprietary information 4-22 Repon 111-971769
4.7 - ' REFERENCES
~
4.4.1 L.M. Petrie and N.F. Landers, " KENO Va - An Improved Monte Carlo Criticality Program with Supergrouping," Volume 2, Section Fil from " SCALE: A Modular System for Performing Standardized Computer Analysis for Licensing Evaluation"
' NUREG/CR-0200, Rev. 4, January 1990.
4.4.2 J.F. Briesmeister, Editor, "MCNP - A General Monte Carlo N-Particle Transport Code, Version 4A," LA-12625, Los Alamos National Laboratory (1993).
4.4.3 ' ' N.M. Greene, L.M. Petrie and R.M. Westfall, "NITAWL-H: Scale System Module for Performing Shielding and Working Library Production," Volume 1, Section F1 from
" SCALE: A Modular System for Performing Standardized Computer Analysis for Licensing Evaluation" NUREG/CR-0200, Rev. 4, January 1990.
4.4.4 M.G. Natrella, Experimental Statistics. National Bureau of Standards Handbook 91,-
August 1%3.
4.4.5 . A. Ahlin, M. Edenius, and H. Haggblom, "CASMO - A Fuel Assembly Burnup Program", AERF-76-4158, Studsvik report.
4.4.6 A. Ahlin and M. Edenius, "CASMO - A Fast Transport Theory Depletion Code for LWR Analysis", ANS Transactions, Vol. 26, p. 604,1977.
4.4.7 M. Edenius, A. Ahlin and B.H. Forssen, "CASMO-3 A Fuel Assembly Burnup Program User's Manual," Studsvik/NFA-86/7, Studsvik Energitechnik AB, November 1986.
l . 4.4.8 ' S.E. Turner, "Waterford Criticality Analsysis," HI-%1562,1996.
- 4.4.9 M. Edenius and A. Ahlin, "CASMO-3
- New Features, Benchmarking, and Advanced
' Applications," Nucl. Sci. Eng., 100 (1988)
. 4.4.10 S.E. Turner, " Uncertainty Analysis - Burnup Distributions", presented at the l~ DOE /SANDIA Technical Meeting on Fuel Burnup Credit, Special Session, ANS/ ENS
- Conference, Washington, D.C., November 2,1988 Holtec International Proprietary information .4-23 Report Hi-971769 i
Table 4.2.1.
Summary of the Criticality Safety Analyses for the MZTR Storage Configuration Design Basis Burnups at 5.0 i 0.05 wt% "U 0 in Region 1 initial enrichment 50 in Region 2 40.75 in Region 3 Temperature for Analysis 20 C Uncertainties Manufacturing tolerances (Table 4.5.1) i 0.0059 Wer-gap (horizontal) i 0.0014 '
Water-gap (vertical) i 0.0003 ,
Burnup (Region 2) i0.0056 i Burnup (Region 3) i 0.0001 Eccentricity in position negative KENO 5a statistics (95%/95%) i 0.0003 Bias statistics (95%/95%). i 0.0012 Statistical combination of uncertainties i 0.0084 l 5.0 wt%"U 4.6 wt%"U Region 1 Fuel Description '
with 16 IFBA with no IFBA ds rMs
{
Reference km (KEN 05a) 0.9266 0.9294 Total Uncertainty (above) 0.0084 0.0084 i
Calculational Bias (see . Appendix A) 0.0030 0.0030 l Axial Burnup Effect negative negative Temperature Correction to 4*C (39 F) 0.0020 0.0020 Maximum k, 0.9400 0.9428 l Limiting k, 0.9500 0.9500 l
l'
' Square root of the sum of the squares.
Holtec International Proprietary Information 4-24 Report HI-971769 l
1 I
I Table 4.2.2 Summary of the Criticality Safety Arzalyses for the Interim Checkerboard Storage Configuration Temperature for Analysis 20 C l
l Uncertainties -
Manufacturing tolerances (Table 4.5.1) i 0.0059 Water-gap (horizontal) i 0.0014 l Water-gap (vertical) i 0.0003 L Burnup (Region 2)' N/A Burnup (Region 3) N/A Eccentricity in position ~ negative KENO 5a statistics (95%/95%) i 0.0004 Bias statistics (95%/95%) i0.0012 Statistical combination of uncertainties i 0.0062' Fuel Description 5.0 wt%usU ' 4.6 wt%u5U f
with 16 IFBA with no IFBA rods rods Reference k,,,(KENO 5a) 0.8439 0.8490 Total Uncertainty (above) 0.0062 0.0062 Calculational Bias (see Appendix A) 0.0030 0.0030 l- l Axial Burnup Effect negative negative l )
I Temperature Correction to 4*C (39 F) 0.0020 0.0020 1 Maxhnum k,, 0.8551 0.8602 i Limiting k., 0.9500 0.9500 l 1
Square root of the sum of the squares.
Holtec International Proprietary inforrnation 4-25 Report Hi-971769 l
l i
l Table 4.2.3 Reactivity Effects of Abnormal and Accident Conditions Abnormal / Accident Conditions Reactivity Effect Temperature Increase (above 4*C) Negative (Table 4.6.1)
Void (boiling) Negative (Table 4.6.1)
Assembly Drop (on top of rack) Negligible Positive - controlled by < 500 ppm Assembly Drop (adjacent to rack) soluble boron Lateral Rack Movement Included in Tolerances Positive - controlled by 500 ppm soluble Misplacement of a fresh fuel assembly boron i
l l
Holtec International Proprietary information 4-26 Report Hi-971769
_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - _ _ ~
l l
l Table 4.3.1 Design Basis Fuel Assembly Specifications Fuel Rod Data Assembly type OFA Standard Vantage-5H
. ~ ,
Fuel pellet outside diameter, in. *, <
l Cladding thickness, in.
Cladding outside diameter, in.
Cladding material Pellet density, % T.D.
- Maximum nominal enrichment, .
wt% "U 2 Fuel Assembly Data Fuel rod array 17 x 17 17 x 17 17 x 17 Number of fuel rods Fuel rod pitch, in.
Number of control rod guide and instrument thimbles -
Thimble outside diameter, in.
Thimble thickness, in.
Number ofIFBA rods Active fuel length, in.
Iloltec International Proprietary Information 4-27 Report HI-971769 i
l _ _ . _ _ _ _ _ _ . _ . . _ _ _ . _ _ _ _ _ _ _ _ _ _ _ _ . . . _ _ _ _
h l
Table 4.5.1
! Reactivity Effects of Manufacturing Tolerances Tolerance Reactivity Effect, Ak s.. . .
\ l' ~ , b . ." :
> r 4" t t3
+'
i ' v . . . /'
,..{ '
%% ,, 8 , ,
, zs - > s. s <
Q !:; j
,3 , s 3 ,Qg g'jygjik:" 4:n
' ^ ^
Mi! J Ri ' < B.9 '
EW f$s JM
" * @M " N
' ^}<' si+'s [ii3 E ....d.;k"jd ' j t@$i !MS, ., < < . %g ;O' s s
>. TW ^Es.
.Mpf QfjMV '
y
, N%y < jpppF ""' ' S' '7 yjy. s >
"f"'
ku.s. .re c i s~ ,
-' : q. _, ,,.;, _ . , . . , I,
, / ' {}bll{f .7""'i . $,"31.i;gg by .Mjf -!Ktd > , 1 - ,,s3 , , ,
v c .<
- sh.> >
s e < x 3 <
M \ ^ ^%
I. - ( v '., u < ', ' .s t .4
+- 4 <
x -
g < c .: a s s , ,> > y s',,
i v . .;,u < .:...
s #4 )> b y 4 4 ! ^ s .::
w 7,. 4 o <
. v 'c , a.g.. '
9gn; . .. ~ me a..,.. w .
/i < k,s ~ ' e g
.33.... y
' [fyif ' 4 y
- <,.
- < ;; (.
s' > s s
^
,,..,.,, 4 OM i' v 3
' " ' " 5 jij.}R:j:'v < .,.'. Oh, g. ,.;.,.,. ., fJ'"., ,
'< - j> s ,
v jjQ) o 3..
< Mty:t ;gy
'i?/ihL
' [p*;Ei: x giyjj-Total (statistical sum)t +0.0059 I
/
i l
l I
i
\
\
i Square root of the sum of the squares.
Holtcc International Proprietary Information 4-28 Report 111971769
]
Table 4.6.1 Reactivity Effects of Temperature and Void Reactivity Effect, Ak Region 1 Region 2 Region 3 Case (Fresh fuel) (50 mwd /kgU) (40.75 mwd /kgU)
, 4*C (39 F) reference reference reference 20*C (68*F) -0.002 -0.0036 -0.0034 60*C (140*F) -0.0095 -0.0137 -0.0134 120 C (248*F) -0.0252 -0.0314 -0.0313 120 C w/10%
void -0.04% -0.0484 -0.0501 l \
l l
l Holtec International Proprietary information 4 29 Report 111-971769 I
I l
l l
l-1 l
Heltec Proprietary
)
i l
l- ,
Figure 4.2.1; Pool. Layout for Callaway and Wolf Creek Mixed Zone Three Region Storage j Hollec International Proprietary Information 4-30 Repon HI-971769
Holtec Proprietary f
1 l
1, Figure 4.2.2; Pool Layout for Callaway and Wolf Creek Checkerboard Storage
' Hollec International Proprietary Information 4-31 Report HI-971769
l l
55 -
l 50 g : ACCEPTABLE BURNUP DOMAIN c(o a : FOR
( 45 REGIOrl 2 AND 3 S"0 RAGE g ,
, y h
d 40
^
q+
[ '
w : 5' '
s' e 5
m 35
- .4,
+
18 # i .
2 x
- V F 6 g.$
td .
30 -
25 m oc g ,-
8
- i j e' I
>; : f' v ,a 'n# '
j m
m
, Pf
/9 f>
3 10
- @9/g4 w
2
/ /p- UNAC CEPTABLE BUFtNUP DOMAIN
- h, FOR ltEGION 2 OR 3 ST0ltAGE 0 .... .... .... .. . .. .i.i i ..
1.5 2.0 2.5 3.0 3.5 4.0 4.5 5.0 NOMINAL INITIAL ENRICHMENT (w/o U-235) i Figure 4.2.3 Minimum Required Fuel Assembly Burnup as a Function of Nominal Initial Enrichment to Permit Storage in Regions 2 and 3 (Fuel assemblies
!I with enrichments less than 2.0 wt% "'U will conservatively be required to meet the burnup requirements of 2.0 wt% "'U assemblies).
f I'
Holtec International Proprietary Infonnation 4-32 F.eport 111-971769 l
s f-l l
l in t
(
t l
?
I I. ..
f I.
1 i
?
l i
i
" .Holtec Proprietary t
1 1..
h t
/
Figure 4.' 3.1 Fuel Storage Cell Cross Section
' Holtec International Proprietary Infonnation 4 33 Report HI-971769 t
.: i f
______.____m__.___ L__.._._ _ _ _ _ _ _ . _ _ _ . _ _
Region 2 Region 3 - Region 3:, Reg.
4 4
63'
" S:; t, ,.
, 4.,
Region 2 Region 2 Reden 3 Reglen"3[ _ Reg..
.3F 6<
s:;n ,
. , . , . . ;, . l ,
I'~ i A ..
.Y Region 2 " Redes 3( 5 3! Reg.
s T3;
.' !.. . g::t+e 3:g-; ,
~;
Region 2 Region 2 Region 2 Region 2 Reg.
2 Region 2 Region 2 Region 2 Region 2 Region 2
. I l
Region 2 Region 2 Region 2 Reg.
2
- r .?:
i ,. '
..v
Region 2 Region 2 ; $" xk',l@f' ' * 'l3] Rei.
ru,
'> qz ,$ { ?. c l'M i i i 9,3:sN
. ; '21,5- y0 .QOj
$[_ _
, e r. . , ,
'.., ,g . , ,
R$' :r9 4 " ddy j ,, [ dhS'g[2{tyv r',^
Region 2 rRegge ${
]@Si Reg. .
q;. L q p .3 g3j
- g[h.;%$h3,, b]3,h,jh
~ ..
- hh
~
~:p , x w geg, Region 2 i Region 2 Region $ ' Reglen 3 4 13 [
Figure 4.4.1 Representation of the KEN 05a Reference MZTR Calculational Model Holtec International Proprietary information 4-34 Report 111-971769 l
5.0 THERMAL-HYDRAULIC CONSIDERATIONS 5,1 Introduction f
T*nis section provides a summary of the methods, models, analyses, and numerical results to demonstrate the compliance of the reracked twin site Spent Fuel Pool and Cask Loading Pit and the Spent Fuel Pool Cooling and Cleanup System (SFPCCS) with the provisions of Section III of the USNRC "OT Position Paper for Review and Acceptance of Spent Fuel Storage and Handling
~ Applications",(April 14, 1978). Similar methods of thermal-hydraulic analysis have been used in I 1
other rerack licensing projects.
I The thermal-hydraulic qualification analyses for the rack arrays may be broken down into the l following categories:
- i. Evaluation of the maximum decay heat load limit as a function of the bulk temperature limit for the postulated discharge scenario.
ii. Evaluation of the postulated loss-of-forced cooling scenarios to establish that pool boiling will not occur.
iii. Determination of the maximum temperature difference between the pool local temperature and the bulk pool temperature at the instant when the bulk temperature reaches its maximum value, iv. Evaluation of the maximum temperature difference between the fuel rod cladding temperature and the local pool water temperature to establish that nucleate boiling at any location around the fuel is not possible with forced cooling available.
As stated previously in this licensing report, the Callaway and Wolf Creek spent fuel pools and cask loading pits are nominally identical. A survey of the pools by Holtec personnel showed that the actual pool cavity dimensions of the two plants deviates slightly (less than four inches).
Accordingly, to provide bounding thermal-hydraulic calculations, the pool water volume is conservatively based on the minimum east-west and north-south dimensions of the two pools.
Hokee International Proprietary Information 5-1 Report HI-971769 u-_________________________________-.______-._
- - __ _ _ _J
Thus, a lower bound thermal inertia and outer periphery downcomer dimension is used in the thermal-hydraulic calculations.
The following sections present the plant system description, analysis assumptions, a synopsis of the analysis methods employed, and final rew'is. Hereinafter, the term " plant" used in this section refers to both Callaway and Wolf Creek.
5.2 Svatem Description l The Spent Fuel Pool Cooling and Cleanup System (SFPCCS) at Callaway and Wolf Creek consist of two cooling trains, a cleanup loop, and 'a surface skimmer loop.
I The fuel pool cooling system consists of two 100% capacity cooling trains for the removal of l
decay heat generated by irradiated fuel stored in the spent fuel pool and cask loading pit. Each 1 train consists of a horizontal centrifugal pump, a shell, and U-tube heat exchanger, a strainer, ,
manual valves, and the instrumentation required for system operation. The decay heat generated by the stored fuel in the pool is transferred from the fuel pool cooling system through the fuel
~
pool cooling heat exchangers. The fuel pool cooling heat exchangers are serviced by the component cooling water system on the shell side with remote manual-operated isolation valves provided.
During normal system operation, one fuel pool cooling pump takes suction from the spent fuel pool and transfers the pool watcr through a fuel pool cooling heat exchanger back to the spent fuel pool. The fuel pool cooling pump suction is protected by a permanent strainer located at the terminal end of the suction piping within the spent fuel pool. Tne pump suction line penetrates the spent fuel pool wall, near the normal spent fuel pool water level. The return line terminates at the bottom of the spent fuel pool. In order to prevent the draining of the spent fuel pool by siphoning ;
i action, an antisiphon hole is located in each return line, near the surface of the pool water. i l
l Hokee International Propnetary Information 5-2 Report HI-971769 l
= _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ - . _ _ _ _ _ _ _ _ _ _ _ _ _ __ ____ _ _ _ _ _ _ _ _ _ _ _
Normal makeup water to the spent fuel pool is supplied by the reactor makeup water system. An I alternate source of makeup water is the RWST via the fuel pool cleanup pumps. Emergency .
makeup water is supplied from the Essential Service Water system. Boron addition to the spent
{
fuel pool is normally accomplished by supplying borated water from the boric acid tanks via the boric acid blending tee. Boron may also be added by using the RWST as the source ofmakeup water to the spent fuel pool. Isolation of non-safety related portions of the SFPCCS is a manual action.
The fuel pool cleanup system contains two inline centrifugal pumps and two filters in parallel, a mixed bed demineralized, and a wye-type strainer. The pumps and filters are designed for fifty- '
percent of the system capacity, and the demineralized and strainer are designed for one-hundred percent system capacity. The demineralized removes ionic corrosion impurities and fission products. The filters a e provided to remove particulate matter which would have otherwise entered the demineralized, and the wye strainer downstream of the demineralized removes resin fines which may be released from the resin bed.
The fuel pool cleanup system provides the capability for purification of the water in the spent fuel pool, the cask loading pit, the transfer canal, the refueling pool, and the RWST. The cleanup system is an essential adjunct to the SFPCCS system to maintain clarity and water chemistry controlin the spent fuel pool.
5.3 Discharge / Cooling Alignment Scenario Consistent with the current plant practice, two discharge scenarios are postulated. They are:
- i. partial core ofiload ii. ' full-core offload Hokee beernational Propnetary Infonnation 3 Report HI-971769 w_________-__-_____-__ _ _ _ _ . _ _ _ _ _ - _ _
l L.
In lieu of prescribing a batch size and cooling period for partial core ofiload, the twin plants seek to determine the maximum pool heat load resulting in a steady state bulk pool temperature limit of 140*F under this scenario with only one cooling train operating.
Similarly, the full core omoad scenario is required to b: executed such that the maximum pool heat load will not allow for bulk pool boiling at the end of a postulated 2 hour2.314815e-5 days <br />5.555556e-4 hours <br />3.306878e-6 weeks <br />7.61e-7 months <br /> loss of forced
. cooling transient which occurs immediately after the full core omoad. More specifically, the bulk l water temperature is sought to be limited to 207*F (which includes 5*F of margin) after two hours of pool heat-up in the absence of all forced cooling paths.
Evaluation of these two scenarios will allow maximum flexibility in batch sizes and cooling periods prior to omoad into the pool. In both scenarios, the component cooling water used to remove heat from the spent fuel cooler is assumed to be at its maximum design temperature.
During the partial core omoad scenario CCW flow is assumed to be at its nominal rate. During Full Core Omoad conditions CCW flow is assumed to be at its design basis fk,w rate. With the thermal effectiveness of the spent fuel pool cooler thus fixed, the requirement of the ceiling on the bulk pool temperature essentially translates into a limit on the total heat generation rate in the pool.
An additional evaluation is performed for a loss of cooling accident occurring some time after restart. This evaluation considers a four hour long loss of forced cooling in the SFPCCS followed by a twenty hour long period with cooling provided at one-half the normal coolant flow rate. It must be demonstrated that the spent fuel pool does not reach the bulk boiling temperature during the 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period. For this evaluation, the component cooling water to the heat exchanger is assumed to be at an elevated temperature and reduced flow rate.
Hokec Intemational Propnetary Information 5-4 Rep >rt HI-971769 1
I.
5.4 Decav Heat Load Limit In this section, we present the methodology for calculating the decay heat load limit for the scenario presented in the preceding section.
The heat load imposed on the pool is from the decay heat generated by fuel assemblies discharged
- into the pool. The primary safety function of the SFPCCS is to adequately transport this hea!
L load to the CCW system and thereby maintain the bulk pool temperature within specified limits.
Compliance with the limiting heat load will be ensured through adjustments to the cooling system performance and/or adjustments to the fuel offload rate. Commonly used decay heat calculation methods based upon ASB 9-2, ANS 5.1, or ORIGEN2 will be used to provide conservative estimates of decay heat values for specific fuel pool inventories.
The following conservatism are applied in the decay heat load limit calculations.
! e n , ,s ,,- n , ,, , , , , . . , . , .
rerse-- n nein>,z,,,,,~,- ~;.. .. . , , , . . , . . , , ,,
f' y
} < >
r ,
,/? <
i ?,<
.s
, s
,~ q' g
>' "~ < <
, / '
' , i >
3, g < ,~
,- . ,, a 0 - , ~
1
, i ,
,e ,
,o ,, < '
l
> < r j i v, / /< i < , , s
( <, , , . . . .. ,,,
.p ~:, ~. - , ' -
,s, . ,' ,: ,
,, ' , ' ' s ,, .; <,
- s < ,
s 5 4 x +' a s 3
v , + ,
s , , ',
<,',, n
- , s -
- :,'- '
l .. , ;
t t s
- ' > ,s
't
,.s ;
t-g ", 's's s j; p' , , s v>
> 2,
. 'i' ' * '
s /s
- ~ '
,' 1 s+ , ^
a v , ,
'f<
l yu" + <-
c x , -
l s ,4 ,s - ,
,, f g V < 7 , d #
/% 4 7 ()
4 < , ', , s r' , > >
$/>, s', f<
gc,
~
+
+
. . ~- -
+
v,
, ,i' t ,
r
~,
W
~
, , b ,
- ,2
~
3- , ,
t $', , ' s ^ * '
s t
L
- -r 4 - A
~ ,
ji ,
s v / > , r is ' ' ,
, i ;
['ce ,,
, ' <',s s e <
{ .
> s ,
Iwl/ w.<n,,v u nxw ,+s ms // v. *y va ' < was , e4d 4s v m v4
, ,x v > 4x / e n +m s , e < sI- w vs % * <%v ,- v, s < ww , , * <
Holtec International Proprietary information 5-5 Report HI-971769 d______..__________._.____
The mathematical formulation can be explained with reference to the simplified heat exchanger alignment of Figure 5.4.1. Referring to the spent fuel pool cooling system, the governing differential equation can be written by utilizing conservation of energy as:
CS = G(t) - G,y(7) - Gg,(7) dr where:
C = Pool thermal capacity, Btu /*F T = Pool bulk temperature, *F t = Time after reactor shutdown, hr Q(t) = Time varying decay heat generation rate, Btu /hr Qux(T) = Temperature dependent SFPCCS heat rejection rate, Btu /hr Qcv (T) = Temperature dependent evaporative heat loss, Btu /hr Subject to the second of the conservatism listed above, this differential relationship can be reduced to the following algebraic relationship:
0= G3,,,,, - G,y(T ,,,,,)
3 - Gy.(T ,,,,,)
3 where:
Tu is the maximum bulk pool temperature limit,*F Qw is the decay heat load limit, Btu /hr Qux(T) is a function of the bulk pool temperature and the coolant waier flow rate and temperature, and can be written in terms of the temperature effectiveness (p) as follows:
Q,g(7) = W, C, p (T - t,)
where:
W, = Coolant water flow rate, Ib/hr C, = Coolant water specific heat capacity, Btu /(Ibx *F) p = SFPCCS heat exchanger temperature effectiveness T = Bulk pool water temperature,*F t, = Coolant water inlet temperature, F Holtec International Proprietary Information 5-6 Report HI-971769
The temperature effectiveness, a measure of the heat transfer efficiency of the SiPCCS heat exchangers,is defined as:
10 - ti P=T-t l .
where tois the coolant outlet temperature (*F) and all other terms are as dermed above.
Qty (T)is a nonlinear function of the pool temperature and ambient temperature. Qsv contains the heat evaporation losses from the pool surface, natural convection and thermal radiation from the y pool surface, and heat conduction through the pool walls and slab. Experiments show that the 1
l l . heat conduction takes only about 4% of the total heat loss [5.4.1].
g3>;m ~, 7;( ~ vp > w <v ~ e- - r ,x7 ; g ,,+fy >s - a ~ s . ~ , , > m ~ ~ . m , s ,. > , ,,, ,, , t,z n < - , + , - < - - v ~ s ~
~
<, ,', -'s
+ n < 'st',
~ f, , , ,x > ' ^^ v
, y , , , ,, , .
r,, .
- ' ~
. y l g s , <
[
f
+, < s,$ / # ,, <
s' "'
15 , <
v <se > .,'2 z
i c< (>3,e , ,< e ,v >e > <v ,) $-s , .c f
u {
[e ,,, , ,," f l , - '
, ' ~ ,
~
>q
' ' + ,. t
,. < t s ' < , -
,3 jf ,$ ,, , , t ,
, ,, s 3 >
(
- f. / ','r~ ~
r ? ' '
' ,,, ', >, ,'< ", ,, e% ' ','
, ', 5,
',, s
?
f u, ,
> s + ,' ' ~ '
<' e, ,'
'"*' '> ss 5
, s a 'v. ' < > s e # > f 4 s< rj ^>s+, < s
) 's. - ^ , t s ; , <> ; -
y g,;',
, ,4
. - s
, ~
s s.
?, ,
- :~ " -
,,"s ,<
c, , ,, x <
- , ,,4 , sr 6 o 2
l1 g >> - v m r 1.g , , , s s < , ' >
r ; >< < ,
,/
[ , * > , #
,/,,',
> ,5 s ,,# A / # %, y
.3 % ,
,y yf j g ,*', s , <
7,
,. ,.s
- > , v ,9 p, e , < < / n s l', v >^ ' , y #
,.' * q s ' A? ' , , ', ,','s >
', ,'s j, '
z'$' s
^",
/ i
.e
~
, ^
i e > w, ,
s ,,4<
>' 's
, y: , '
- ~; g 6,; /
s -
, , < t ,, ; ,
,3 <
m . .
~ '# ~
< r 7, , ,, s , s !
,y
. ' ,>' , : ,, :(, ,
- ~
, s ., - , ; '
r >
f ',
g -
v g
, , , s, -
,<^- ' < < , > - +
- ' ve, eq
{ 'q v ;f$p ,f
- tf , , ,
V. ',a u.,s 'Ai a d,,',,- < . . ' v , : LL.',.
- - t r u', A ,,. ,1 ... >x , ,s , ,a n,a,+ a
'.',b The algebraic heat balance equation is solved for the decay heat load limit by rearranging the
' equation given above and substituting the maximum temperature limit for pool water temperature (T). The major input values for this analysis are summarized in Table 5.4.1.
l I
i Hohec International Propnetary Information 57 Repost HI-971769 w _ - - - - _ - __-- _ __- - __ -__ .>
5.5 Margin Agahist Boiling As stated previously, it is necessary to ensure that the pool bulk temperature will remain less than 207'F (i.e., adequate margin against boiling) if: (1) all forced cooling paths are lost following a full core ofiload and cooling is not restored for two hours, and (2) a loss of coolant accident occurs after restart and partial cooling is restored after 4 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br />. The SFPCCS system has two independent trains, both of which are seismically qualified and safety-related, so a complete loss
- of forced cooling is not possible under single failure criteria. Regardless of this fact, these '
evaluations are performed for postulated non-mechanistic loss of forced cooling accidents.
The following conservatism are applied in the heat-up calculations.
3;yyne w ;
e . .
ym ;m w e w mt- m-- :m x w w y> ,,
j'[ :$,.j: ) ,.c,.:.... s );"I o - 4'.f; ,.
x:.w p3 x A . .s . > < > ,s .
&$'5
' ^
5
'IN n'
< < >;?.- 5
' p:m~-
n -
>.p W:n h.if i:
9;'
g o
^
l[.p.;. > :.<g .v
> z . ...,.:.
^
h@" .
et
[ , , , ..
- wp
- ~.42 u- <
?:3
- ss gy , g gm^
< s ,
y - T ,
M ; , +.,,m>.
+ >4 -
' 4 pn +
3N s - n -
s , s .
>^ ' @g' '
,4T:Q '
- 4^,
^
^ "
y ...' f[dA:) ~ 'Ph5.
- T * * > 4 k' bf
- hiik.w.
4 g-
>..<y ew.
),4 ' '
?hi k JOMY g s
b NLinic
' ^
~ ' ',k'
. ng .
d!! ', :p r.r .!$,.2.s.
w is:n# ^
,j;.g n- <
=7n ' .- -::99_ y q:y s ,, 'Q:g- s
~~~ _
?l% +, , , , , , ,
^ '
$4S ' " Ei E N$ '
?Wi>
'. s. x < , > 3 ,
W, > ' My < .. > s .
(QM r &g / ,UP' > .i .,c..M; ($$ '
E:y K g+ v w
n~ frWW g.,.;
i i < , ;g j
.y.- -b ,
w,+:_-:1 < >
p Au.. .
+ . .
w.>
i R"*j. .jr i ..c nW g@ < 4> ,5 s <. M < J5 i
p.. timt: x
>+ - .;>
?hY * ' hh?b' bbY > '
~ ys?
- k
, w
^'
+
[
G(m; s ne e
<.m - ; :',. 3 5
. :a hh $$dk$02ihd N'4 .$$$.d5S.dsi#Mi d5I$fideXe ,4w ~ 5sMAM WGJiN55.SSNNs JM:ddOSh/.*N55$433A+4 & $$5 N:$>M::I $$5 bNN .@Er- NONIN$$
The temperature rise of the water in the pool over any period of time is a direct function of the average net decay heat load during that period. Therefore, maximizing the decay heat load will Hohec International Proprietary Information 5-8 Report Hi-971769 l l
maximize the pool temperature increase rate and minimize the corresponding time-to-boil. As a transient decay heat load would necessitate a reduced average net heat load, the steady-state assumptions are indeed conservative.
The governing enthalpy balance equation for this condition, subject to these conservative assumptions, can be written as:
C 7dT= G ,,,,,-Gy(7) 3 where t is the time after cooling is lost (hr) and all other terms are the same as defined in Section 5.4.
This differential equaiion is solved using a numerical solution technique to obtain the bulk pool temperature as a function of time. The inajor input values for these analysis are summarized in L
Table 15.1.
5.6 ' Local Pool Water Temperature In this section, a summary of the methodology for evaluating the local pool water temperature is presented. A single conservative evaluation for a bounding amalgam of conditions is performed.
The result'of this single evaluation is a bounding temperature difference between the maximum local water temperature and the bulk pool temperature.
In order to determine an upper bound on the maximum local water temperature, a series of conservative assumptions are made. The most important of these assumptions are:
l i
~ep 'wl gg
??cwq y egrgyvpgyAm ylg lmgglj%em ?** :
mpaw & m < ,;vye^Nm
^
QLl^lh'M y ef W,f M( , ,
g gy y"^ ,
,, . : - , y
{gj; :
gp" #
<. ^ ,.
x ." j 'jh,7 Ed o IT,!;sfjP ',., ;
d!db!$NS Id&MW X# esAa $$d.ds&S bN 31w 3dh/7\ MJ3: @M4!dMi+ Ad?2dl i: ,
Holtec International Proprietary information 5-9 Report HI-971769 l
r._ .. _ .
i i
- fh 'N'DNs N' , , .
[:}![f # # 4 F # " W WQ f'j'@ ,V j;. j, o.,
Y? <. <
9 li!Ej
~
i i
- (4:j };y
- h :n , i -;sig l
f fffh Nh j.:g.
'Og ' pg Mt ,
M" i.' - N < ,
W'<+
M*~'
gin W ,. ., . ,.
gg
>Jp e> ,,5g 0:FO , < ? f):(
ks: man m a s mi me ani , m , s - mar as , . .
aM. . j All rack cells are conservatively assumed to be 50% blocked at the cell outlet to account for drop accidents resulting in damage to the upper end of the cells. This blocked cell portion is conservative, since structural evaluations have shown that only about 20% of the cell is blocked subsequent to the impact of dropped objects.
enemmemmmmwwn m;; < : ., m m e,a,.n; m ee m
- m. m-- ~ ~me ~ ~ . = . . .
w w a n: m=<:um >
' ' V '
.fk[ j: 0M +> , gg.;
'~
j Q%N@/
b!bs:m meMoenwea:ssitsw wn . . m: .< we mi:: , aw a
i 5.6.1 Local Temocrature Evaluation Methodoloav i
The inlet piping which returns cooled pool water from the SFPCCS terminates above the level of the fuel racks. It is not apparent from heuristic reasoning alone that the cooled water delivered to the pool would not bypass the hot fuel racks and exit through the outlet piping. To demonstrate
- adequate cooling of hot fuel in the pool, it is therefore necessary to rigorously quantify the l l l velocity field in the pool created by the interaction of buoyancy driven flows and water l
' injection / egress. A Computational Fluid Dynamics (CFD) analysis for this demonstration is
- required. The objective of this study is to demonstrate that the principal thermal-hydraulic criteria of ensuring local subcooled conditions in the pool is met for all postulated fuel discharge / cooling l' l
l alignment scenarios. The local thermal-hydraulic analysis is performed such that partial cell blockage and slight fuel assembly variations are bounded. An outline of the CFD approach is described in the following.
l There are several significant geometric and thermal-hydraulic features of the twin site Spent Fuel .
Pools which need to be considered for a rigorous CFD analysis. From a fluid flow modeling Hohec International Proprietary Information 5-10 Report HI-971769
standpoint, there are two regions to be considered. One region is the bulk pool / cask loading pit region where the classical Navier-Stokes equations are solved with turbulence effects included.
The other region is the heat generating fuel assemblies located in the spent fuel racks located near the bottom of the Spent Fuel Pool. In this region, water flow is directed vertically upwards due to buoyancy forces through relatively small flow channels formed by the Westinghouse 17x 17 fuel assembly rod arrays in each rack cell. This situation shall be modeled as a porous solid region in which fluid flow is governed by the classical Darcy's Law:
dP,_ p 9 X; K(i) V'. - C p lVl 2 where ap/dXi is the pressure gradient, K(i), V i and C are the corresponding permeability, velocity and inertial resistance parameters and p is the fluid viscosity. The permeability and inertial resistance parameters for the rack cells loaded with Westinghouse 17x17 fuel were determined based on the friction factor correlations for the laminar flow conditions typically encountered due to the low buoyancy induced velocities and the small size of the flow channels.
The twin site pool geometry required an adequate portrayal oflarge scale and small scale features, spatially distributed heat sources in the spent fuel racks and water inlet / outlet configuration.
Relatively cooler bulk pool water normally flows down between the fuel rack outline and pool wall liner clearance known as the downcomer. Near the bottom of the racks, the flow turns from a vertical to horizontal direction into the bottom plenum supplying cooling water to the rack cells.
Heated water issuing out of the top of the racks mixes with the bulk pool water. An adequate modeling of these features on the CFD program involves meshing the large scale bulk pool region and small scale downcomer and bottom plenum regions with sufficient number of computational cells to capture the bulk and local features of the flow field.
Mainnbisd 6t"sourcesLinlihijp6dfusi hosilahkA;are rnodeled bd Holtec International Proprietary Inferrnation 5-11 Report HI-971769
- n ,, ~ :>~ ,,,, ~m , - - - ; ,, ~- ~~n,,,, . , > ~ , ~ , , , ~ . . ,
f'e,',s,- c ,r e n ~,,:v .,1~
' ' , ,' , v , < ' , 4 2 tg 9 .,,',,1,,
, < ~
/+ ,
' '~ ' ' ;, (;fff;ffy
{' i , ' -
} > y ,, ,'. ', Eg/jgg "f.,>
~<<
- s. o , , >v, ,
i a , ,
s 1
,8~~~,
~- '
t> ~ ,, ,,- , . ,
e,, , > ,
., ,> ,n 4 ,m& & ,4 40, ,v,<~w 1, m a 'anvis rn s. nu n u>+,<r,,> ,,t,~,,, v t < ', u ,ia ,. , en a, % + +~.; ,' , < < < - , , ,s>,, AA The CFD analysis was performed on the industry standard FLUENT [5.6.4] fluid flow and heat transfer modeling program. The FLUENT code enabled buoyancy flow and turbulence effects to be included in the CFD analysis. Turbulence effects are modeled by relating time-varying "Reynolds's Stresses" to the mean bulk flow quantities with the following turbulence modeling options:
(i) k-e Model (ii) RNG k-e Model (iii) Reynolds Stress Model The k-c Model is considered most appropriate for the twin site CFD analysis. The k-e turbulence model is a time-tested, general purpose turbulence model. This model has been demonstrated to give good results for the majority of turbulent fluid flow phenomena. The Renormalization Group
. (RNG) and Reynolds Stress models are more advanced models that were developed for situations j where the k-e Model does not provide acceptable results, such as high viscosity flow and supersonic shock. The flow regime in the bulk fluid region is such that the k-e Model will provide ;
acceptable results.
Rigorous modeling of fluid flow problems requires a solution to the classical Navier-Stokes equations of fluid motion [5.6.1]. The governing equations (in modified form for turbulent flows with buoyancy effects included) are written as:
Holtec International Proprietary Information 5-12 Report HI-971769 1
1
t f
Bp, u; + Sp,(ul up = B p a u, + a u, '
Bt. . d x, O x, ,
axj 8x, ,
B Spo(u',u', )
p - p, p (7 - T,) g, +
l dx, O x, l
where ui are the three time-averaged velocity components. p (u;' uj ') are time-averaged Reynolds stresses derived from the turbulence induced fluctuating velocity components u;#, p, is the fluid
! density at temperature T., p is the coefficient of thermal expansion, is the fluid viscosity, gi are the components of gravitational acceleration and x3 are the Cartesian coordinate directions. The Reynolds stress tensor is expressed in terms of the mean flow quantities by defining a turbulent
- j. viscosity , and a turbulent velocity scale k" as shown below [5.6.2]:
p ( u;uj' ) = 2/3p k 6; 3 p, +b 6x; Bx, The procedure to obtain the turbulent viscosity and velocity length scales involves a solution of two additional transport equations for kinetic energy (k) and rate ofenergy dissipation (c). This
- methodology is known as the k-c model for turbulent flows as described by Launder and Spalding t
[5.6.3]..
4 1
Some of the major input values for this analysis are summarized in Table 5.6.1. An isometric view of the assembled CFD model is presented in Figure 5.6.1. Figures 5.6.2 through 5.6.4 present temperature contours at the exit of the rack cells, velocity vectors near the inlet and outlet ;
piping, and velocity vectors in the cask loading pit region.
5.7 Fuel Rod Claddmg Temperature ;
In this section, the method to calculate the temperature of the fuel rod cladding is presented.
Similar to the local water temperature calculation methodology presented in the preceding Hokee International Proprietary Information 5-13 Report HI-971769 i
_. .____________________a
section, this evaluation is performed for a single, bounding scenario. The maximum fuel cladding superheat above the local water temperature is calculated.
p -
I. <
Q ,
y n
^ #
.4 p 4
> h ri-
+
k <
.it:
9 p
f:!.
E q.
b-. .
r h:
4
/
?
4 9
I ^
X
, bI h
c.::u , .s. .4 d;:-
Holtec International Proprietary Infonnation 5-14 Report HI-971769
l l
l g m . , ~,, . y - , ~ ,, , , , _ , :,_ . , ,,y, ,
,,,,< ~ s m . ,, g s~ . , , , %
s --
s v < >
y < < 4,q s s . <
l bl + , <
4 , ,
, < ,f
/ , +
I y o s ;
i e , > , - s < s, l
} ' ,,a i > /< ,c v i ,,
'< s '> < '
(<
>S
' k* '
i g; <,s , , , - + - < ,< <
, e ,: - < '
l f' a 's s' ', '
> r 5, f 4 . + 5 -
^
, s + < r k, 5 ##
', " 3 ' /
' ^ ' '
[,<
)
'r l, <
s r
')
< s > < .> + < .. s 4 ,
s s s
{
- a i
r$ ^;< ^
is s
- s s> .s s
jpl sv
> > 1
^
( ('2' ,
s'
, - ,r< ,
( <
'> ,' , < 0,
<4 s .,
- ' 't Lu.nw m s <A m ,k .2- ~ x>m ~ s o u, s., wm .- , - vu, ~
% + J uhn ,s, s ~ + sw > < . - d, ~w a + > - . ->
In order to introduce some additional conservatism in the analysis, we assume that the fuel cladding has a crud derosit resistance]
Y ~Lwhich covers the l
1 entire surface. Thus, including the temperature drop across the crud resistance, the cladding to l 1
water local temperature difference (AT,) is given by:
wwwm ~e n1 mmw - ~ r mv~ ~
i- m
- s~
~ - -
nem i
- m. m m
..l y
- O. h x
x l
,h.S ' 'r.' v. .,. " m,< , ~
> - t s.-
'I Cd #
s .,.s.,
g ega > m.m.z. ,
, m
.v p.)y.:- Nhh[ '
- p. 4 g m.y s
=:. ~
h adbNbk.hhhNk>< ar e> < hh A h .,. N N' NSd :< .v b < abfs v e $b$b f<
5.8 Results This section contains results from the analyses performed for the postulated discharge scenario.
5.8.1 Decav Heat Load Limits For the discharge / cooling scenarios postulated in Section 5.3, the calculated decay heat load limit i
is summarized in Table 5.8.1. Remembering that all transient effects were excluded from the evaluations, this decay heat load corresponds to the invariant heat load which results in a steady-l state bulk pool temperature which will not exceed the temperature limit for either the partial core or full core ofiload scenario.
Hohec International Proprietary Information 5-15 Report HI-971769
This calculated decay heat load limit is not based on any specific discharge conditions, but is a mathematically derived quantity. Any conservative decay heat calculation used to determine the operational limits (i.e. in-core hold time requirement) necessary to avoid exceeding this decay heat load will provide conservative operational limits. The operational limits will be determined based on the decay heat load limit in Table 5.8.1. Based on this limit, the twin plant cooling systems will remain in compliance with their existing FSAR and SER.
5.8.2 Time-to-Boil If all SFPCCS forced pool cooling becomes unavailable, then the pool water will begin to rise in temperature and eventually will reach the normal bulk boiling temperature of 212"F. In order to maintain some margin to this boiling condition, analyses are performed with the acceptance criterion of a bulk pool temperature that is s 207*F. The time to reach the boiling point will be the shortest when the loss of forced cooling occurs at the point in time when the pool bulk temperature is at its maximum calculated value. Although the probability of the loss-of-cooling event coinciding with the instant when the pool water has reached its peak value is extremely remote, the calculations are performed under this extremely unlikely scenario.
Analysis shows that, for postulated full-core discharge, and a maximum bulk temperature of 207'F after two hours without cooling, the maximum allowable decay heat load is 63.41 MBtu/hr. The steady-state SFP temperature at this heat load would be 169.68'F.
For the loss of coolant accident scenario, the bulk temperature after four hours without cooling would be 172.l *F. Once partial cooling is reestablished, the steady-state temperature would be less than 175 *F, thereby precluding the possibility of boiling even with continued reduced cooling capacity.
Holtec International Proprietary Infonnation 5-16 Report HI-971769
5.8.3 Local Water and Fuel Cladding Temperatures Consistent with our approach to make conservative assessments of temperature, the local water temperature calculations are performed for a pool with decay heat generation equal to the maximum calculated decay heat load limit. Thus, the local water temperature evaluation is a calculation of the temperature increment over the theoretical spatially uniform value due to local hot spots (due to the presence of a highly heat emissive fuel bundle).
The CFD study has analyzed a single bounding local thermal-hydraulic scenario. In this scenario, a bounding full-core discharge is considered in which the 193 assemblies are located in the pool, farthest from the cooled water inlet, while the balance of the rack cells are postulated to be occupied by fuel from old discharges.
In this analysis, the difference between the peak local temperature and the coincident bulk pool temprature was conservatively calculated to be 64.6*F.
The peak fuel cladding superheat is determined for the hottest cell location in the pool as obtained from the CFD model for the twin site pools. The maximum temperature difference between the fuel cladding and the local water (AT,) is calculated to be less than 67.4 *F. Applying this calculated cladding AT,, along with the maximum temperature difference between the local water temperature and the bulk pool temperature, to the bulk maximum normal operating pool temperature of 170*F yields a conservatively bounding 234.6*F maximum local water temperature and a conservatively bounding 302*F peak cladding temperature. The maximum local water temperature is lower than the 239'F local boiling temperature on top of the racks, thereby precluding nucleate boiling in the subchannel. The heat fluxes are too low to support a departure from nucleate boiling (DNB) condition. Thus, nucleate and departure from nucleate boiling do not occur anywhere within the Callaway and Wolf Creek pools.
Holtec International Proprietary Infonnation 5-17 Report HI-971769
\
5.9 References
[5.4.1] " Heat Loss to the Ambient from Spent Fuel Pools: Correlation of Theory with l Experiment". Holtec Report HI-90477, Rev. O, April 3,1990.
. [5.4.2] "An Improved Correlation for Evaporation from Spent Fuel Pools", Holtec Report HI-971664, Rev. O.
l
[5.6.1] Batchelor, G.K., "An Introduction to Fluid Dynamics", Cambridge University Press,1967. .
l
[5.6.2] Hinze, J.O., " Turbulence", McGraw Hill Publishing Co., New York,' NY,1975.
. [5.6.3] Launder, B.E., and Spalding, D.B., " Lectures in Mathematical Models of Turbulence",
Academic Press, London,1972. i
[5.6.4] "QA Documentation and Validation of the FLUENT Version 4.32 CFD Analysis Program", Holtec Report HI-961444.
[5.7.1] Rohsenow, N.M., and Hartnett, J.P., " Handbook ofHeat Transfer", McGraw Hill Book Company, New York,1973.
?
i Hohec International Propnetary Information 5-18 Report HI-971769 !
=_=____:__ __
TABLE 5.4.1
. DATA FOR DECAY HEAT LOAD LIMIT EVALUATION Length of Spent Fuel Pool (min.) 597.56 inch Width of Spent Fuel Pool (min.) 339 inch Pool Building Ambient 110*F Temperature Emissivity of Water 0.96 Specific Heat ofWater 0.998 Btu /(ibx*F)
HX Temperature Effectiveness 0.4981 (partial core) 0.3116 (full core)
Coolant Water Inlet Temperature 105'F 130"F (post LOCA) ,
Coolant Water Flow Rate 1.50x10'lb/hr (partial core) 0.75x10'lb/hr (post LOCA) 3.00x10'lb/hr (full core) l 1
l l
Holtec Intemational Proprietary Information 5 19 Report HI-971769 s
TABLE 5.5.1 DATA FOR TIME-TO-BOIL EVALUATION Length of Spent Fuel Pool 597.56 inch l
l Width of Spent Fuel Pool 339 inch Depth of Spent Fuel Pool 37.25 ft Total Fuel Rack Weight 411,320 lb Number ofFuel Assemblies 2,642 assys Bounding Assembly Weight 1,467 lb Pool Building Ambient i10*F Temperature Emissivity of Water 0.96 Pool Thermal Capacity 3.144x 10' Btu /*F Specific Heat of Water 0.998 Btu /(Ib x*F)
Holtec International Proprietary Information 5-20 Report HI-971769
I l
TABLE 5.6.1 l
DATA FOR LOCAL TEMPERATURE EVALUATION Bounding Assembly Weight 1467lb Maximum Fuel Assembly Heat Flux 1870 Btu /hr-ft2 Radial Peaking Factor 1.65 Total Peaking Factor 2.5 Number of Fuel Assemblies 2642 SFPCCS Water Flow Rate 3.0x 10' lb/hr I Type of fuel assembly Westinghouse 17x17 Std. j t
Fuel Rod Outer Diameter 0.374 in . l Rack CellInner Dimension 8.77 in Active Fuel Length 144 Number of Fuel Rods per 289 rods :
Assembly *
\
Rack Cell Length 169 m l
1 Minimum Bottom Plenum Height 5m l
- Note: Fuel assembly is modeled as a square array with all locations containing fuel rods.
l l
Holtec International Proprietary Information 5-21 Report HI-971769
p u
e e
_ k m a u l M Vp o )m 0 7 8 5
. de r g
_ 1 5 r e(
i t ua qW
. e R
N O
I T
A yt U ai cei0m ')
L A DL1 5
_ V ma d x 1 1
4 E r uoh
_ 7 3 T
I i mL/ t t u 2 6 x
M a aB e I
L MH(
- 1 D 8 A 5 O )
E L k d e
L T u e )
_ l t
B r it a A B ut m l
u A E ma i l c T H ue r ( l a
Y mp F
- (
c x em i
A 0 F 4
MaT C
1 0
E 7 D 1 F
O S f T oS s L rCn e
. U bCi a 1 1 S
E mP u FT r
R S N
e o
r o d e i
c a rd r
a - lo oa c o n lai l e l f Fu o l f c it r o f
S a P
c i 1
1 i
EVAPORATIONHEATLOSS SPENTFUEL POOL _
HEATEXCHANGER i
II P g' :
i COOLANT
\
i, tg %
Figure 5.4.1; SPENT FUEL POOL COOLING MODEL HI-971769
l e
d o
M D
- F C
l o
- y P o
r
- t a l e
u e F t
. i n r e p
- p S o
r f
o P i w
e c V t
e i c
- r l t e
o m H o s
I I.
6 5
e r
u ig F
y r
t a
. i e
r p
o r
P c
t e
l o
H
y r
t a
. i e
r
_ p o
r P
c t
e l
o H
y r
t a
i e
r p
o r
P c
t e
l o
H
ULNRC.03837 Enclosure 2 Proprietary Chapters 4 and 5 Licensing Report for Reracking of Callaway Spent Fuel Pool i