ML20209E789
| ML20209E789 | |
| Person / Time | |
|---|---|
| Site: | Zion File:ZionSolutions icon.png |
| Issue date: | 06/23/1999 |
| From: | Peterson R COMMONWEALTH EDISON CO. |
| To: | |
| Shared Package | |
| ML20209E743 | List: |
| References | |
| 22S-110M-0063, 22S-110M-0063-R02, 22S-110M-63, 22S-110M-63-R2, NUDOCS 9907150152 | |
| Download: ML20209E789 (60) | |
Text
1 Attachment 5 1
Calculation Number 22S-110M-0063 Maximum Cladding Temperature for Uncovered Spent Fuel Rod June 23,1999
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CALCULATION TITLE PAGE l
Calculation No.: 22S-0-110M-0063 DESCRIPTION CODE: N00 DISCIPLINE CODE: N i ZION NUCLEAR STATION l SYSTEM CODE: SF TITLE: Maximum Claddina Temocrature for Uncovered Spent Fuel Rod l
l l C Safety Related O Augmented Quality C Non-Safety Related REFERENCE NUMBERS Type Number Type Number PROJ 10436-002.10436-003 AEDV PSED COMPONENT EPN: DOCUMENT NUMBERS:
EPN Compt Type Doc Type /Sub Type Document Number l
1 l
REMARKS:
REV. REVISING APPROVED DATE NO. ORGAN 17ATION PRINT / SIGN 0 Sargent & Lundy Robert J. Peterson / Signature on file 8/31/98 i Sargent & Lundy Robert J. Peterson / Signature on file, , , 12/30/98 ,
2 Sargent & Lundy Robert J. Peterson / [h((hyg v ng --
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i COMMONWEALTil EDISON COMPANY CALCULATION REVISION PAGE CALCULATION NO. 22S-0-110M-0063 PAGE NO.: 2 REVISION SUMMARIES REV: 0 REVISION
SUMMARY
Initial Issue: Main Text: 31 pages; Attachment pages as shown in Table of Contents ELECTRONIC CALCULATION DATA FILES:
(Program Name, Version, File name ext / size /date/ hour /: min)
COBRA-SFS, Cycle 3 peakassm.inp 29591 8/28/98 4:07:36pm peakassm.out 1491703 8/28/98 4:10:22pm radgen.inp 207 8/27/98 1:00.04pm radgen.out 857407 8/27/98 1:24:26pm Prepared by: Mark Handrick/ Signature on file 8/31/98 Print / Sign Date The reviewer's signature indicates compliance with SOP-0402 and the venfication of the following minimum items: correctness of math for hand prepared calculations, appropdateness ofinput data, appropdateness of assumptions, and appropdateness of the calculation method.
Reviewed by: Kawei Chan/ Signature on file 8/31/89 Print / Sign Date Type of Review
@ Detailed O Alternate O Test DO ANY ASSUMIrrlONS IN THIS CALCULATION REQUIRE LATER VERIFICATION O YES @ NO Tracked by:
COMMONWEALTII EDISON COMPANY CALCULATION REVISION PAGE
! CALCULATION No. 22S-0-110M-0063 PROJECT No. 10436-003 PAGE NO.: 2.1 REVISION
SUMMARY
REV:1 l
REVISION
SUMMARY
- Revision 1 of this calculation is issued to incorporate a model of the Fuel Handling Building to determine the air space temperature following a postulated loss of all spent fuel pool cooling water (see Attachment D). This air space temperature is then used as a boundary condition to the COBRA-SFS model, used to calculate the l maximum cladding temperature of the spent fuel assemblies. The decay heat load of the spent fuel assemblics is taken as that corresponding to a date of 2/1/99 for the postulated loss of spent fuel cooling water, as opposed to a date of 7/1/98 which was used in Revision 0. The computer files for the revised COBRA-SFS model are included in Attachment E.
The hardcopy printout of Revision 0 computer files (Attachments B and C) are deleted and replaced by microfiche in Revision 1. Changes to the calculation text are denoted by revision bars.
Revised Pages 1,3-4,6,8-11,15,17-19,24-28,30, Al-A3, A5-A6, B1, Cl-C2 Added Pages 2.1,25.1,25.2, Dl-D16, El Deleted Pages B2-B405, C3-C172 Total Pages Issued as Revision 1: 60 pages Electronic Calculation Data Files:
(Program Name Version, I ile name ext sizeate hourl: min)
COBRA-SFS, Cycle 3 zirc-rvl.inp 29591 kb 12/22/98 9:51:24am zirc-rvl.out 1491352 kb 12/22/98 10.00:56am KITTYlS zion-fhb.out 20811 kb 12/21/98 1:35:24pm l Prepared by: Mark Handrick / Signature on file 12/30/98
! Print / Sign Date The reviewer's signature indicates compliance with S&L Procedure SOP-0402 and the verification of, as a minimum, the following items: correctness of mathematics fo, anual calculations, appropriateness of input data, appropriateness of assumpt:ons, and appropriateness of the calculation method.
Reviewed by: Helmut Kopke / Signature on file 12/30/98 Print / Sign Date Type of Review I x l Detailed [ } Altemate [ ] Test DO ANY ASSUMPTIONS IN THIS CALCULATION REQUIRE LATER VERIFICATION [ ] YES [x ] NO Tracked by:
I This box is the responsibility of Comed -
Supplemental Review Required YES (NEP-12-05 documentation) NO Supervisor / l l
F COMMONWEALTII EDISON COMPANY CALCULATION REVISION PAGE CALCULATION NO. 22S-0-110M-0063 PAGE NO.: 2.2 REVISION SUMMARIES l l REV: 2 l REVISION
SUMMARY
- Revision 2 of this calculation is a complete revision. Revision 2 evaluates the ,
l maximum spent fuel cladding temperature following a loss of spent fuel pool cooling water, postulated to ocxur on a reference date of 06/30/99. The maximum decay heat load of the spent fuel assemblies is computed I externally using the computer code ORIGEN2.1, corresponding to the reference date of 06/30/99. The Fuel Handling Building air temperature is computed in Attachment D for the reference date of 06/30/99, using an HVAC inlet flow rate of 21,000 cfm. The computer files for the revised COBRA-SFS models are included in Attachment B. Changes to the calculation text are denoted by revision bars. 1 Revised Pages 1,3-31, Al-A4, B1, Cl-C2, D1 - D16 Added Pages 2.2,32 - 34 Deleted Pages A5 - A6, El j Total Pages issued as Revision 2: 59 pages ELECTRONIC CALCULATION DATA FILES:
(Program Name, Version, File name ext / size /date/ hour /: min) I COBRA-SFS, Cycle 3 cob 63099.inp 29622 kb 06/15/99 4:33 pm 1
cob 63099.out 1491469 kb 06/15/99 4:40 pm !
KITTYlS thb63099.out 23781 kb 06/14/99 9:33 am Prepared by: Mark Handrick 2D4 b 88 ,9 Print / Sign Date Reviewed by: Kawei Chan (C(t, C-d. % 6/2z./99 Print / Sign Date Type of Review
@ Detailed O Alternate Test Su le tal Review R umenta )
kNO
, , ,,7, ,- , ,
DO ANY ASSUMPTIONS IN THIS CALCULATION REQUIRE LATER VERIFICATION O YES @ NO Tracked by:
I l
COMMONWEALTil EDISON COMPANY CALCULATION TABLE OF CONTENTS PROJECT # 10436-003 CALCULATION NO. 22S-0-110M-0063 REV. NO. 2 PAGE NO. 3 SECTION PAGE NO. SUB-PAGE NO.
TITLE PAGE 1 Rev.2 l REVISION
SUMMARY
2 2.1, 2.2 TABLE OF CONTENTS 3
- 1. PURPOSE / OBJECTIVE 4
- 2. METHODOLOGY / ACCEPTANCE CRITERIA 4-8 Rev.2 3. ASSUMPTIONS AND LIMITATIONS 9 - 10
- 4. DESIGN INPUT I1 - 14
- 5. REFERENCES 15 - 16 l
- 6. CALCULATIONS 17 - 29
- 7.
SUMMARY
AND CONCLUSIONS 30 Tables and Figures 31 - 34 ATTACHMENTS:
Rev. 2 A Spreadsheet Cell Formulas for Tables Al - A4 i
B COBRA-SFS Input / Output Files B1 Input File " cob 63099.inp"(Microfiche)
Output File " cob 63099.out"(Microfiche)
C RADGEN Input / Output Files l Input File "radgen.inp" Cl Output File "radgen.out"(Microfiche) C2 i
D Computation of Fuel Handling Building Homogeneous D1 - D16 Air Temperature 1
i l
i l
t.
Exhibit E NEP-12 COMMONWEALTH EDISON COMPANY
, CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 4 i
l
- 1. PURPOSE / OBJECTIVE -
The purpose of this calculation is to determine the maximum zircaloy cladding temperature in the spent fuel pool at Zion Nuclear Station under postulated conditions where all of the cooling water is lost and the fuel is exposed to an air environment. It is desired that the maximum claddmg temperature is kept below the critical temperature {
necessary for oxidation of the cladding. The critical temperature is defined in Reference 5.1 as 565 *C (1049 'F).
J Rev.2 This calculation will determine the maximiun cladding temperature at a reference date after the final reactor Ishutdown and subsequent fuel discharge to the spent fuel pool, assuming a sudden loss of cooling water. The time duration for spent fuel cooling corresponds to a final reactor shutdown date of 2/21/97 (Reference 5.6), and a postulated loss of spent fuel pool cooling water on a reference date of 06/30/99.
1 If the calculated maximum claddmg temperature does not exceed the critical value for the reference date, then a l postulated loss of spent fuel pool cooling water occurring after the reference date will not result in oxidation of the zircaloy cladding.
I
- 2. ' METHODOLOGY / ACCEPTANCE CRITERIA 2.1 - - Methodology l Under normal conditions, the spent fuel in the spent fuel pool (SFP) is cooled by an active cooling water system.
This calculation postulates a sudden loss of all pool water, exposing the spent fuel rods to an air environment. This greatly reduces the mechanism for heat removal from the spent fuel rods, and consequently results in higher fuel rod temperatures as the decay heat of the spent fuel is dissipated to the air heat sink at steady-state conditions.
Under these conditions, the primary mechanism for heat transfer away from the spent fuel rods is by convective
[ .
heat transfer to the ambient in the spent fuel pool room via air circulation currents established in the spent fuel pool Rev. 2 l cavity and Fuel Handling Building. The HVAC system serving the Fuel Handling Building provides a constant i source of relatively cold air to the exposed spent fuel rods. The air is heated as it passes upwards along tim fuel rods, then leaves the Fuel Handling Building via the return air ducts. A schematic representation of this flow pattern is'shown in Figure 1.
l l REVISION NO. 2 i
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 5 Figure 1 Schematic Representation of Convective Air Currents .
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Fuel Handling Building
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I REVISION NO. 2
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 6 Each fuel assembly stored in the spent fuel pool resides in a dedicated square storage cell, with each cell connected to a lower plenum via a 5-inch diameter hole in the cell base-plate. The storage cells in the spent fuel pool are one of two types, called Region I and Region Il type cells. There are no significant differences with respect to the interior of these cell types, as each type have an 8.94 inch inner dimension,0.075 inch cell wall thickness, and a 5-inch diameter flow hole in the base-plate (pages 3-2 through 3-5 of Reference 5.4). All cells are open to a common plenum above the fuel assemblies. Therefore, the spent fuel pool itselfis a system of parallel flow channels. The side walls of the storage cells provide a radial heat transfer path towaids the heat sink associated with the spent fuel i pool walls. A complete model of the entire spent fuel pool, with all fuel assemblies accounted for, would be I prohibitively large, in terms of available computer resources.
The computer code COBRA-SFS Cycle 3, " Thermal Hydraulic Analysis of Spent Fuel Casks", will be used to determine the maximum cladding temperature following a postulated loss of cooling water. The COBRA-SFS computer program was developed at Pacific Northwest Laboratory for thermal-hydraulic analysis of multi-assembly spent fuel storage and transportation systems. The code has undergone an independent technical review as part of a submittal to the Nuclear Regulatory Commission for a generic license to apply the code to spent fuel storage system analysis. The companion computer code RADGEN is used to calculate the gray body view factors employed in the radiation heat transfer module of the COBRA code. The RADGEN code validation is included as part of COBRA-SFS code package (Ref. 5.3).
To determine the maximum cladding temperature, the fuel assembly most recently discharged to the spent fuel pool, with the highest decay heat, will be modeled explicitly. To conservatively maximize the calculated fuel / cladding temperature, this assembly is assumed to have an adiabatic boundary condition along the sides of the storage cell.
This precludes any radial heat transfer towards the pool wall heat sink. This is representative of an assembly ;
located in the center of the spent fuel pool, surrounded by other assemblics of equal decay heat. Under these conditions, only heat transfer from the fuel rods to the air flowing axially upwards is applicable. Figure 2 illustrates the COBRA model of the storage cell, including the numbering scheme for the fuel rods and flow channels within the assembly.
Rev.2 he maximum decay heat generation rate associated with the spent fuel assemblies is calculated in an external analysis (Reference 5.25) using the computer code ORIGEN2.1. A volumetric heat generation rate is computed for the limiting spent fuel assembly based on the maximum assembly decay heat documented in Reference 5.25 and the active fuel length. The volumetric decay heat generation rate is used as input to the COBRA model. Details of these calculations are provided in Section 6.2 and Table 1.
To account for axial peaking with respect to decay heat, an axial power profile is applied to the average decay heat values for the most recent fuel assemblics discharged to the spent fuel pool. A sinusoidal distribution is employed, consistent with PWR axial power profiles, as cited in References 5.4 and 5.12. The axial profile is employed in the COBRA model to capture any localized peak fuel / cladding temperatures. Details of the axial power distribution are provided in Section 6.3.
REVISION NO. 2
Exhibit E i NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 7 Figure 2 !
COBRA Spent Fuel Assembly Model l l
Slab No.1 Slab No.2 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 j 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 17 18 19 20 21 22 23 2s 25 26 27 28 29 30 31 32 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 l 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 I 44 31 49 50 32 51 52 34 53 35 54 f 5537 56 38 57 39 58 59 41 60 42 61
-i 62 45 63 64
)
-g gg 46 47 48 49 50 51 52 n 54 55 56 57 58 59 60 No. 8 No. 3 65 66 67 68 69 70 71 72 73 74 75 76 77 78 79 80 1 61 62 63 64 66 67 68 69 70 72 73 74 75 81 82 83 64 85 86 87 88 89 90 91 92 93 94 95 96 76 77 79 80 81 82 83 84 85 86 87 89 90 97 96 99 100 101 102 103 104 105 106 107 108 109 110 111 112 91 92 93 94 95 96 97 98 99 00 101 102 03 104 105 113 tid 115 11 6 117 118 11 9 120 121 122 123 124 ?25 126 127 28
- 106 107 108 '~ 110 111 112 ' 114 115 116 118 119 120 -
129 130 131 132 133 134 135 136 ' 137 138 139 140 141 142 143 44 I 21 122 123 124 125 126 127 128 129 130 131 132 133 134 136 l 145 146 147 148 149 150 151 152 153 154 155 156 157 158 159 160 136 137 139 140 141 142 143 144 145 146 147 149 150 161 162 163 164 165 166 167 168 189 170 171 172 173 174 175 76 151 52 53 154 156 157 158 159 160 162 163 164 65 j g 177 178 179 180 181 182 183 164 185 186 187 188 189 190 191 92 g No.7 166 167 16e 169 170 171 172 174 175 176 177 178 179 180 No. 4 193 194 195' 196 197 198 199 200 201 202 203 204 205 206 207 81 82 -
184 185 187 188 189 191 192 194 96 209 10 211 212 213 214 215 216 217 218 219 220 221 4 222 4
^^ - 223 196 197 198 199 1 10 ,
225 226 227 228 229 230 231 232 233 234 235 236 237 238 239 40 211 12 13 14 15 16 1 218 21 221 4 241 242 243 244 245 246 247 248 249 250 251 252 253 254 255 256 s <
g Slab No. 6 blab No.5 l g controt Rod:
REVISION NO. 2
L Exhibit E NEP-12-02
- COMMONWEALTH EDISON COMPANY l L
CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 8 ,
1 2.2 Acceptance Criteria 1 In order to prevent the rapid oxidation of the zircaloy cladding following a loss of spent fuel pool cooling water, the j i
cladding temperature must be kept below the critical temperature in Reference 5.1 as 565 C (1049 'F). If the Rev 2 . maximum calculated cladding temperature at the reference date of 06/30/99 is below the critical temperature, then the spent fuel rods are not expected to oxidize as a result of the postulated loss of spent fuel pool cooling water accident, provided the postulated accident occurs on or after the reference date.
l 2.2.1 Computer Programs Used The following computer programs were used in the preparation of this calculation: ;
COBRA-SFS Cycle 3,"A Thermal-Hydraulic Analysis Code for Spent Fuel Storage And Transportation Casks",
Sargent & Lundy Program No. 03.7.672-1.0.
Microsoft Excel version 5.0c, Sargent & Lundy Program No. 03.1.138-5.0 KITIYlS, " Thermal-Hydraulic Steady States in Arbitrary Solid and/or Fluid Channel Configurations", S&L Program No. 03.7.171-5.11.
These computer programs are maintained by S&L's Software Center for use. The computer programs were run on S&L PC No. 5765, which is attached to S&L file server SNL2. COBRA-SFS Cycle 3 and KITTYlS have been validated per S&L's procedures. The Excel spreadsheet validation is implicit in the detailed review of the calculation and requires no additional documentation.
l REVISION NO. 2
Exhibit E NEP-12-02 f COMMONWEAloTH EDISON COMPANY l I
CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 9
- 3. ASSUMPTIONS AFD ! IMITATIONS 3.1 Fuel Assembiv/ Rod Assumptions 3.1.1 The thermal conductivity ofirradiated UO is 2 assumed to be reduced by 60 percent from that of
{
unirradiated UO 2per she statement in Table 5 1 of Reference 5.12. This is conservative with respect to calculated fuel / cladding temperatures and is representative of the fuel stored in the spent fuel pool.
3.1.2 The fuel rod grid spacers are assumed to be uniformly distributed along the fuel rod length. This assumption simplifies the COBRA model, with negligible impact on calculated cladding temperatures. ,
Rev.2 3.1.3 The grid spacer frictional loss coefficients are conservatively assumed to have a value of 4.4. This is twice the maximum value shown in Figure 9-26 of Reference 5.26, which illustrates the laminar loss coefficient for transverse grid spacers. The use of Reference 5.26 as a basis for the grid spacer loss coefficient is recommended by Reference 5.27. The factor of two is applied to account for possible unceitainties in the data of Reference 5.26. This assumption leads to conservatively high local loss coefficients, resulting in reduced mass flow rates and consequently higher air temperatures for a specified decay heat load. Higher air temperatures will lead to conservatively higher calculated cladding temperatures.
3.1.4 The exit loss coefficient at the top of the fuel assembly is assumed to be 1.0, consistent with an abrupt area expansion into an infinitely large area, per page A-26 of Reference 5.21.
3.1.5 The storage cell geometry is modeled after the Region 11 storage cells shown in Reference 5.4. This is representative of the vast majority of stored fuel assemblies, since Region 11 cells comprise 2670 of the 3012 total storage locations in the spent fuel pool (Design Input 4.1). The flor hole at the bottom of each storage cell is 5 inches in diameter for both types of storage cells, per pages 3-1 through 3-5 of Reference ]
5.4. The dimensions of the storage cell (inner diameter and wall thickness) are identical for the Region I i
- and Region 11 cells. Since the storage cell is modeleci as radially adiabatic, the minor differences in cell-to- i cell pitch are not a factor in this analysis, and the Region I and Region 11 storage cells can be considered ;
identical for the purposes of this calculation. I Rev.2 3.1.6 The spent fuel assembly model is applicable to the VANTAGE 5 fuel design, which is the type of fuel used in fuel cycles 14 and higher, per Reference 5.18. The M lmtch of spent fuel discharged to the pool from Unit I was Cycle 15, as shown in Reference 5.6. De i.u.c ofiload from Unit 2 contained a mixed batch of VANTAGE 5 and OFA fuel assemblies. He maximum decay heat for spent fuel is applicable to the Unit 1 VANTAGE 5 fuel, per Reference 5.25. Derefore, the modeling of VANTAGE 5 fuel is appropriate for determining the maximum cladding temperatures. l 3.2 Heat Transfer Modelina Assumptions ;
3.2.1 The fuel assembly storage cells are assumed to be adiabatic in the radial direction. This is conservative !
with respect to calculated cladding temperature since the neglected heat transfer mechanism would act in
. parallel with convection to the air in the axial direction.
REVISION NO. 2
)
Exhibit E NEP-12-02 COMMONWEALTII EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.10 Rev.3 3.2.2 The air temperature at the 11VAC inlet of the Fuel Handling Building is assumed to be 95 F, consistent ;
with the recommendation of Reference 5.27. Higher inlet air temperatures will result in higher calculated i
' fuel / cladding temperatures, and are therefore conservative.
3.2.3 The emissivity of the stainless steel storage cell walls is assumed to be 0.2, consistent with the values listed in Table A.8 of Reference 5.22 for stainless steel. The emissivity of the zircaloy cladding is assumed to be
' O.8, consistent with the value used in the COBRA validation case titled "TN24P", documented in Reference 5.3. This validation case models a Transnuclear, Inc. spent fuel storage cask loaded with standard Westinghouse 15 x 15 PWR spent fuel, with documented results showing that the COBRA model accurately predicts peak cladding temperatures with respect to the measured test temperatures.
3.2.4 It is assumed that the upper and lower plenum flows are one-dimensional, and that the flow mixes ,
instantaneously in the plenum regions. I 3.2.5 The environment in the Fuel Handling Building is assumed to be at atmospheric pressure, or 14.7 psia.
This pressure value is used to determine the inlet air density via the Auxiliary Building Ventilation system.
Variations of the ambient pressure are expected to be negligible, and have an insignificant effect on the results of this calculation.
3.2.6 The specific heat of air is assumed to be 0.253 Btu /lb.- F for the purpose of determining the temperature rise along the heated length of the fuel rod, as shown in Table 3, " Boundary Conditions for COBRA Model". This specific heat value is consistent with an air temperature of 700 F as shown in Design Input 4.5.1, and is representative of the average air temperature for flow along the heated length of the fuel rod.
l l
REVISION NO. 2
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PRJJECT NO. 10436-003 PAGE NO.11
- 4. DESIGN INPUT 4.1- Ecactor Core. Fuel Assembly. and Soent Fuel Storace Cell Parameters The following data is applicable to the reactor core, fuel assemblies, and spent fuel storage cells as Zion Station, with references as provided.
Parameter Value Reference Reactor Core and Spent Fuel Pool i Rated Core Thermal Power 3250 MW Refs. 5.7, 5.14, 5.15 Number of Assemblies per Core Loading 193 assemblics Refs. 5.6, 5.14 Date of Last Discharge to Spent Fuel Pool 2/21/97 Ref. 5.6 Total Pool Heat Load on 06/04/99 5.5E6 (Btu /hr) Ref. 5.27 Fuel Asssmblies Fuel Assembly type
- VANTAGE 5 Ref. 5.18 Fuel Rod Locations per Assembly 225; 15 x 15 array Refs. 5.4 (page 4-24),5.15 Active Fuel Rods per Assembly 204 Refs. 5.4 (page 4-24),5.15 Fuel Rod Pitch: 0.563 inches Refs. 5.4 (page 4-24),5.16 Fill Gas Helium Ref. 5.14 Fuel Rod Length 152.17 inches Ref. 5.16 Active Fuel Length 144 inches Ref. 5.4 (page 5-25)
Fuel Pellet Diameter 0.3659 inches Refs 5.4 (page 4-24),5.16 FuelTheoretical Density (TD) 10,97 g/cm' Ref. 5.12 (Table 4-2) '
Fuel Pellet Density 95% ofTD Ref. 5.4 (page 4-24) I Cladding Material = Zircaloy-4 Refs. 5.4 (page 4-24),5.16 Cladding OD 0.422 inches Refs. 5.4 (page 4-24),5.16 Cladding ID 0.3734 inches Ref. 5.4 (page 4-24)
Cladding Thickness 0.0243 inches Refs. 5.4 (page 4-24), 5.16 Number of Grid Spacers per Assembly
- 10 Refs. 5.15 (page 4.2-8 ),5.17 Spent Fuel Pool Storace Cells Total Storage Cells in Spent Fuel Pool 3012 Ref. 5.4 (page 1-1)
Total Storage Cells in Region 1 342 Ref. 5.4 (pnge 2-14)
Total Storage Cells in Region 11 2670 Ref. 5.4 (page 2-15)
Storage Cell Material Stainless Steel Type 304 Ref. 5.4 (page 2-3)
Storage Cell Square Box ID 8.94 inches Ref. 5.4 (page 4-29)
Storage Cell Wall Thickness 0.075 inches Ref. 5.4 (page 4-29)
Base-Plate Flow Hole Diameter 5.0 inches Ref. 5.4 (page 3-5)
Storage Cell Height above Baseplate 168 inches Ref. 5.4 (page 2-16)
Rev.2
- These are Unit 1 parameters applicable to the VANTAGE 5 fuel, per Assumption 3.1.6.
REVISION NO. 2
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY l
CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.12 1
4.2 Auxiliary Buildine Ventilation System Parameters Rn2 4.2.1 Deleted
- 4.3 LLQduel Physical Properties References 5.12 and 5.13 provides the following physical properties of UO fuel.
Temperature Thermal Conductivity Thermal Conductivity
- Specific Heat (Ref. 5.12) (Ref. 5.12) (Ref. 5.13)
(F) (Btu /hr-ft- F) (Bru/hr-ft- F) (Bru/lbm- F) 75 0.06 200 4.5 1.80 0.063 400 3.5 1.40 500 0 07 600 2.8 1.12 800 2.5 1.00 1000 2.2 0.88 1200 2.0 0.8 1
- Calculated UO 2thermal conductivity after 60% reduction following irradiation 1
Note: The density and specific heat values are not used in the steady-state solution for maximum cladding temperature, but are required input to the COBRA code.
4.4 Physical Properties of Metals !
4.4.1 Reference 5.13 provides the following physical properties of Zircaloy cladding.
l Temperature Density Specific Heat Thermal Conductivity (F) (ib./ft') (Btu /lbm- F) (Btu /hr-ft- F) 75 409 0.071 6.7 200 6.9 400 7.I 600 7.2 4.4.2 The thermal conductivity of Stainless Steel type 304 is given by Reference 5.10 as the following. ;
I 9,4 Bru/hr-ft 'F at 212 F 10.9 Btu /hr-ft- F at 572 F REVISION NO. 2
Exhibit E L
NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.13 i
r l 4.5 Physical Properties of Gases 4.5.1 Air Properties Reference 5.10 (Table .A-3) provides the following physical properties of air at atmospheric pressure. The l enthalpy data is provided by Reference 5.11 (Table A-13E).
l Temperature Enthalpy Thermal Specific Heat Specific Viscosity Conductivity Volume (F) (Btu /lbm) (Btu /hr-ft- F) (Btu /lb.- F) (ft'/lb.) (Ib./ft-hr) 0 109.9 0.0133 0.239 11.628 0.0400 l 100 133.9 0.0154 0.240 14.085 0.0463 200 157.9 0.0174 0.241 16.667 0.0518 300 182.1 0.01933 0.243 19.231 0.0580 400 206.5 0.0212 0.245 21.739 0.0630 l 500 231.1 0.0231 0.247 24.272 0.0680 600 256.0 0.0250 0.250 26.810 0.0720 700 281.1 0.0268 0.253 29.326 0.0770 800 306.7 0.0286 0.256 31.847 0.0810 900 332.5 0.0303 0.259 34.364 0.0850 1000 358.6 0.0319 0.262 36.900 0.0889 1500 493.6 0.040 0.276 49.505 0.1080 !
l !
1 The molecular weight of air is given as 28.97, per Table A-1E of Reference 5.11. )
l l 4.5.2 lielium Properties l Reference 5.10 (Table A-3) provides the following physical properties of helium.
Temperature Thermal Conductivity l (F) (Btu /hr-fi *F) l 200 0.097 400 0.115 600 0.129
- 800 0.138 1 4.6 Miscellaneous 4.6.1 The Universal gas constant is given as 10.73 psi-ft'/lb ele *R per Reference 5.22. The gas constant for air l is obtained using the molecular weight of 28.97, and h: a value of 53.34 fl-lb/lb. *R.
l l
REVISION NO. 2
7 2 Exhibit E l, NEP-12-02 COMMONWEALTH EDISON COMPANY l
CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.14 l
l Rev.3 4.7 Decav Heat for Spent Fuel Assemblics Per Reference 5.25, the spent fuel assembly with the highest decay heat thermal output is assembly F50D.
The calculated heat loads from fuel assembly F50D on 06/30/99 is 8561 Btu /hr, as documented in I Reference 5.25. !
l 4.8 Misgellaneous Zion Fuel Building and Spent Fuel Pool Assembly Information Reference 5.27 provides the following information:
4.8.1 The supply of HVAC air is from the Auxiliary Building Ventilation System, and is expected to be at ambient conditions of the Auxiliary Building. With no equipment running, this air temperature is expected to be below 95 F.
4.8.2 The appropriate assembly nonle drawings, for the limiting fuel assemblies, discharged into the spent fuel pool are as follows:
Unit 1 Region 15 - 15x15 VANTAGE 5 w/o IFMs Top Nozzle - 1125E34 (3 sheets), Revision 19 Bottom Nonle - 1125E22 (2 sheets), Revision 14 l
Unit 2 Region 14 - 15x15 0FA i Top Nonle .- 1123E36 (2 sheets), Revision 22 Bottom Nonle - 1541E27 (2 sheets), Revision 25 l
4 l
l l
l REVISION NO. 2
n Exhibit E NEP-12-02
' COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. IC 436-003 PAGE NO.15
- 5. REFERENCES 5I NUREG/CR-6451, "A Safety and Regulatory Assessment of Generic BWR and PWR Permanently Shutdown Power Plants", August 1997.
l 5.2 NUREG/CR-6441 BNL-NUREG-52494, " Analysis of Spent Fuel Heatup Following Loss of Water in a )
Spent Fuel Pool", Draft Report for Comment, May 1998.
3 5.3 COBRA-SFS, "A Thermal-Hydraulic Analysis Code for Spent Fuel Storage And Transportation Casks", i
. September 1995, Prepared for the U.S. Depanment of Energy by Pacific Northwest Laboratory, Sargent &
Lundy Computer Program No. 03.7.672-1.0. j 5.4 Commonwealth Edison Letter from S. F. Stimac, Nuclear Licensing Administrator. 9 Dr. Thomas E.
Murley, Office of Nuclear Regulatory Commission, dated January 15,1992; Attachment B, " Zion Station Spent Fuel Pool Modification Licensing Repon for Proposed Changes to Facility operating Licenses DPR-39 and DPR-48" 5.5 Safety Evaluation by the Office of the Nuclear Reactor Regulation Related to Amendment No.142 to Facility Operating license No. DPR-39 and Amendment No.131 to Facility Operating License No. DPR-48, Commonwealth Edison Company Zion Nuclear Power Station, Units I and 2, Docket Nos. 50-295 and - l 50-304, dated February 23,1993. /
5.6 Calculation No. 22N-0-110M-0058, Revision 0, " Zion Decommissioning Spent Fuel Pool Heat Load and Time to Saturation Calculation",4/20/98. 1 5.7 Calculation No. 22N-0-110X-0057, Revision 0, " Fuel Handling Accident Offsite Dose Calculation with Extended Radioactive Decay and no AB Filtration",3/24/98.
i Rev.2 5.8 Deleted 5.9 Deleted 5.10 Principles of Heat Transfer, Frank Kreith,3rd Edition,1973.
' 5.11 Fundamentals of Engineering Thermodynamics, Michael J. Moran, Howard N.
Shapiro,1988.
5.12 - Nuclear Heat Transport, M. M. El-Wakil,1978.
5.13 Nuclear Reactor Engineerinn. Samuel Glasstone & Alexander Sesonke,1963.
5.14 Zion Station UFSAR, Section 4.1, July 1993.
i l
REVISION NO. 2-l
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p , 1 Exhibit E NEP-12-02
' COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.16 -
I 5.15 Zion Station UFSAR, Section 4.2, July 1993.
i 5.16 Zion Station UFSAR, Table 4.2-2, July 1993.
5.17 Zion Station UFSAR, Figure 4.2-10, July 1993.
i 5.18 Zion Station UFSAR, Section 4.3.4.1, July 1993.
{
5.19 Zion Station UFSAR, Section 9.4.3, May 1996.
5.20 Comed Drawing No. M-78, Rev. Y, Zion Station Unit 1 & 2 Diagram of Auxiliary Building Vent System.
l 5.21 Flow of Fluids Through Valves. Fittinas. and Pioe, Crane Technical Paper No. 410, Twenty Fifth Printing,1991.
5.22- Introduction to Heat Transfer, Frank P. Incropera & David P. DeWitt, Second Edition,1990.
5.23 Deleted Rev. 3 5.24 Deleted 1
5.25 - Calculation 22S-B-110X-0068, Rev. O, " Determination of the ficat Load for Assemblics F50D and Z33A". l 5.26 Nuclear Systems I - Thermal liydraulic Fundamentals, Neil E. Todreas, Mujid S. Kazimi, Hemisphere Publishing Corporation.
5.27 NDIT NFM9900121, Seq. No. O,
Subject:
Zion Spent Fuel Pool Data,6/4/99.
1 i
l REVISION NO. 2' l
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p L
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.17 1
6.' CALCULATIONS 6.1 Homenclature 2
A : Cross-sectional flow area (ft )
c, Specific heat (Btu /lbm- F)
D Fuel' rod outer diameter (inches) 2 l g Gravitational acceleration (32.2 ft/sec )
' 2
- g. Conversion factor (32.2 lbm-ft/lbrsec )
Gr ' Grashof number
. h -- Fluid enthalpy (Btu /lb.), per context h Heat transfer coeHicient (Btu /hr-ft' *F), per context k Thermalconductivity(Btu /hr-ft *F)
.K Frictional loss coeHicient ,
L Length (ft) )
' #t Mass flow rate (Ib./sec)
MW Molecular weight N Number of assemblics Pr Prandt! number P Fuel rod pitch (inches) ;
q Decay heat for a single assembly (Btu /hr) l . G: Total decay heat from a group of assemblies (Btu /hr)
R Universal gas constant t Time (sec)
T Temperature ( F)
P Volumetric flow rate (cfm)
Vol Volume (ft')
I r Axial coordinate in the z-direction c Emissivity L o Stephan-Boltzmann Constant (Blu/hr-ft2 - R) p'. Dynamic viscosity of fluid (ib./hr-ft) 6.2 Spent Fuel Decav Heat Generation Rev.2 The maximum decay heat generation rate for the spent fuel assemblies is provided by Design input 4.7, as documented in Reference 5.25. The spent fuel assembly with the highest decay heat thermal output is assembly
. . F50D. The calculated heat load from fuel assembly F50D on 06/30/99 is 8561 Btu /hr. This decay heat loads is
! converted to a volumetric heat generation rate, required as input to the COBRA code, with the computations shown in Table L i
l REVISION NO. 2 l
n Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. I8 6.3 Decay Heat Generation Axial Profile
- The decay heat generation axial profile is assumed to follow a sinusoidal distribution along the active fuel length. l The peak-to-average weighting factor, x, is given as the following function of axial position : and active fuel length l
L: i y(:) = EsinE 2 L An appropriate distribution obeys the following normalization:
L L., y(r)d: = 1.0
. When this distribution is iutegrated over the active fuel length, the correct normalization is obtained.
W' xL 1 ~L' 'LY L
.1 x'-
(:)d:L= 12x
--- cos - = = - = 1.0 L <,
<Li,_ L_2 s2) _ L The decay heat generation at any location : is then the average decay heat per fuel rod times the peak-to-average weighting factor y(z). Values of the peak-to-average weighting factor are shown in Table 2 for a nodalization of 24 axial segments over the 144 inches of active fuel. The value of y(z) at the axial centerpoint is 1.571. j l
i l
l 6.4 Fuel Assembly Geometry 6.4.1 Parameters for Square Array Fuel Rod Assemblics l
The COBRA code requires information on the geometry of each channel modeled for a given assembly. The !
required subchannel data is derived from the following assembly p meters:
]
D= Rod diameter = 0.422 inches P= Rod pitch = 0.563 inches N= Number of rods in one row = 15 W= Rod diameter plus gap between the rod and assembly wall; for an assembly inner dimension of 8.94 inches (Design input 4.1), W = 0.422 + 0.5'[8.94-(14*P+D)] = 0.422 + 0.318 = 0.740 inches Per Table 4.2 of Reference 5.2, subchannel parameters are calculated as shown below.
REVISION NO. 2
c Exhibit E NEP-12-02 COMMONWEALTH EUISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO.19 2
Subchannel Number of Channel Flow Area (in ) Channel Wetted Perimeter (in)
Type Channels Central (N - 1)2 = 196 xD =1.326 P-x D = 0.1771 2 2 4
Wall r x 4*(N - 1) = 56 W - If P-x D = 0.2279 2
-D + P = 1.2259 s 2s 8 2 Corner 4 r # * '
W _ D'
- 2 D = 0.2449 -D + 2 W - D' = 1.3894 s 2s 16 4 s 2s
- The hydraulie, or equivalent, diameter is defined as four times the cross-sectional area divided by the wetted perimeter. The assembly-average equivalent diameter, weighted by each subchannel type, is computed as:
196(.1771) + 56(.2279) + 4(.2449) 48.4536
_4 = 0.580 inches = 0.04834 feet 196(1.326) + 56(1.2259)+ 4(1.3894) 334.104 The heated perimeter is the perimeter of all fuel rods seen by each channel, with values given below:
1 Subchannel Type Heated Perimeter (in)
Central xD = 1.326 Wall (x/2)D = 0.6629 Corner (x /4)D = 0.3314 6.4.2 Fuel AssemblyWallGeometry The fuel assembly walls are modeled using the slab solid structure component of the COBRA code, with slab identification numbers as shown in Figure 2. Dimensions of the slab nodes are given in the following table.
2 Slab Node Node Length (in) Node Width (in) Node Axial Area (in )
Number l i
1 4.545 0.075 0.3409 2 4.545 0.075 0.3409 3 4.47 0.075 0.3353 4 4.47 0.075 0.3353 ,
l 5 4.545 0.075 0.3409 6 4.545 0.075 0.3409 l 7 4.47 0.075 0.3353 l l 8 4.47 0 075 0.3353 I l
REVISION NO. 2 i t
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 20 l l
Geometry factors are specified in the COBRA code input file for node-to-node thermal connections pairs. The ,
geometry factors, defined per Equation 6.2 of Reference 5.3, are shown below for the corresponding node-to-node i thermal connections pairs.
w =
distance from the solid node center to the cdge facing the adjacent node
=
1 length of the solid node at the face in contact with the adjacent node, perpendicular to the axial direction
= geometry factor for Slab node A of the pair = (w/l)4 FG.A l
= geometry factor for slab node B of the pair = (w/l)a l Fa3 l
l Node A Node B l Node Pair w 1 FO,A w I Fa2 (A.B) (in) (in) (N/A) (in) (in) (N/A) i 1,2 2.2725 0.075 30.30 2.2725 0.075 30.30 2,3 0.0375 0.075' O.500 2.235 0.075 29.800 3,4 2.235 0.075 29.800 2.235 0.075 29.800 4,5 2.235 0.075 29.800 0.0375 0.075 0.500 5,6 2.2725 0.075 30.30 2.2725 0.075 30.30 6,7 0.0375 0.075 0.500 2.235 0.075 29.800 7,8 2.235 0.075 29.800 2.235 0.075 29.800 8,1 2.235 0.075 29.800 0.0375 0.075 0.500 l
l 6.4.3 Miscellaneous Fuel Assembly Parameters The thermal conductivity of the UO 2fuel is assumed to be 0.8 Btu /hr-ft 'F. This is a lower bound on the values given in Design Input 4.3 for irradiated fuel. Lower fuel thermal conductivities result in higher fuel temperatures, which is conservative with respect to calculated cladding temperatures. l The UO 2fuel density is computed at 95 percent of theoretical density, per Design Input 4.1.
, pw = 0.95* 10.97 = 10.422 g/cm' *(62.428 g/cm' per Ibm /ft') = 650.6 lb./ff i l The fuel / cladding gap conductance is computed by dividing the thermal conductivity of helium by the gap widdi. )
For helium at 800 F, the thermal conductivity is 0.138 Btu /hr-ft- F per Design Input 4.5.2. The gap width is calculated by subtracting the cladding thickness and fuel pellet diameter from the total cladding outer diameter as l shown:
d
[0.422 - 2*(0.0243) - 0.3659] / 2 = 0.00375 inches = 3.125 x 10 ft gap conductance = 0.138 / 3.125 x 10" = 441.6 Btu /hr-ft2 ,op REVISION NO. 2
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 21 l
l It is conservative to maximize the gap conductance for the purposes of calculating the maximum cladding 2
temperature. Therefore, a gap conductance of 500 Bru/hr-ft - F will be used in the COBRA model.
l
! 6.5 Frictional Pressure Drop 6.5.1 Friction Factor for Subchannel Flow The friction factor is given as a function of Reynolds Number according to the expression (see Equation 4.55 of Ref. 5.2):
C1 f = Re" where Cr is a friction constant as defined below, per Equation (4.54) of Reference 5.2 n = 1 for laminar flow n = 0.18 for turbulent flow r 8 r a2 Cf= a + b, p -I + b2 p -I
<D > <D >
For interior subchannels with P/D = 1.334, Table 4.1 of Reference 5.2 provides the following coefficients for computing both laminar and turbulent friction constant Cr.
a bi b2 Laminar 35.55 263.7 -190.2 Turbulent 0.1339 0.09059 -0.09926 Substituting these coefficients into the expression for the friction constant gives:
Laminar Flow: Cr = 102.4 Turbulent Flow: Cr = 0.153 6.5.2 Local Frictional Loss Coefficients 6.5.2.1 Base-plate Flow Hole Loss Coefficient l The loss coeflicient associated with the 5-inch diameter flow hole it the base-plate of the storage cell, K., is given l
by Equation (4.59) of Reference 5.2.
- 1 K* =C ^' hf.2 l
l REVISION NO. 2 l l
l
Exhibit E j
! NEP-12-02 i l COMMONWEALTil EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 22 where: A. is the cell cross sectional area An is the base-plate hole area -
Co is the discharge coefficient ,
l The areas are computed as: l 2 2 l-A. = 8.94 - 225*n/4*(.422)2 = 48.453 in 2
As = n/4*(5.0)2 = 19.635 in l
1 l
With the discharge coefficient taken as 0.6 per Reference 5.2, the value of K, is 14.14.
Rev.2 A twenty percent margin is conservatively added to the computed value, giving K, = 1.2*(14.14) = 16.97.
i 6.5.2.2 Fuel Assembly Grid Spacer Loss Coeflicient Rev.2 The loss coeflicients for the grid spacers are conservatively assumed to have a value of 4.4, consistent with Assumption 3.1.3. The ten grid spacers are assumed to be uniformly distributed along the heated length of the fuel, as shown below.
J Ra 2 -
6.5.2.3 Fuel Assembly Top / Bottom Nozzle Loss Coefficients The loss coefficients for the spent fuel assembly top / bottom nozzles are computed based on the methodology given in Reference 5.21 for a sharp-edged orifice. The loss coefficients are based on the cell cross-sectional flow area of 2
i 48.453 in . The ratio of open cross-sectional area to the total flow area is computed by examination of the drawings I
listed in Design input 4.8.2 for the VANTAGE 5 and OFA fuel assembly designs for Units I and 2 respectively.
The nozzle loss coefficient is computed for both the VANTAGE 5 and 0FA fuel assembly types, with the higher loss coefficient used in the COBRA model. The evaluation of both the VANTAGE 5 and OFA fuel assembly designs is conservative, as the limiting fuel assembly decay heat is applicable to the VANTAGE 5 fuel assembly l design, per Reference 5.25.
- Unit 1 Region 15: 15x15 VANTAGE 5 w/o IFMs l l
l Bottom Nozzle (Ref. Westinghouse Drawing No. I 125E22, Rev.14)
Open Area for Bottom Nozzle Flow lioles l
l l Iloie Type 11 ole Diameter Number ofIIoles Flow Iloie Area (inches')
l (inches) i liole D 0.190 656 18.60 1
Total Flow Area. A,, 18.60 l
REVISION NO. 2 l
L
7 l
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 23 Rev.3 e Unit 2 Region 14: 15x15 OFA Bottom Nozzle (Ref. Westinghouse Drawing No.1541E27. Rev. 25) l Total Cross-Sectional Area of Bottom Nozzle Open Area for Bottom Nozzle Flow Holes j k
Hole Type Hole Diameter Number of Holes 2 Flow Hole Area (inches ) 1 (inches) l Hole L -0.453 112 18.05 l Hole N 0.250 76 3.73 Hole G l
0.370 4 0.43 i Total Flow Area. A,, 22.21 1
The limiting ratio of open flow area to total area, applicable to the VANTAGE 5 fuel, is computed as:
2 = A" = 18.60 / 48.453 = 0.384 A,
= [ = 0.620 g = (p2)' = 0.147 Tlh flow coefficient, C, is taken as 0.6, consistent with the value for square-edged orifices shown on page A-20 of Reference 5.21. This flow coefficient value is a conservative lower bound for Reynolds numbers in the laminar range of 20 < Re < 2000 as shown on page A-20 of Reference 5.21. Justification for use of the laminar Reynolds number range is provided in Section 6.7 of this calculation. Lower flow coefiicient values result in higher loss coefficients, which is conservative for the purposes of this analysis. The loss coefficient for the bottom nozzle, Kwuo,,,, is computed according to the relation given on page A-20 of Reference 5.21:
I 2 K = l- = 2
( 1 - 0.384) / [ (0.6 )*(0.147) J = 11.64 c2y e Unit 1 Region 15: 15x15 VANTAGE 5 w/o IFMs Top Nozzle (Ref. Westinghouse Drawing No. I12SE34, Rev.19) ;
Open Area for Top Nozzle Flow Openings i
j' Opening Type Dimension (inches) Number of Holes 2 Flow Hole Area (inches )
A Oval . 0.439 x 0.765 16 4.71 B Oval 0.439 x 1.328 4 2.17 C Ova! 0.439 x 1.890 4 3.15 D Oval 0.439 x 0.967 12 4.60 I
E Oval 0.439 x 2.383 4 4.02 F Oval 0.439 x 1.529 6 3.78 H Oval 0.439 x 2.653 2 2.25 J Oval 0.485 x 1.869 8 6.85 Total Flow Area. A,, 31.53 REVISION NO. 2 L
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 24 Rev. 2 . -
e Unit 2 Region 14: 15x15 OFA Top Nonle (Ref. Westinghouse Drawing No. I123E36, Rev. 22)
Dimensions of the flow openings are not provided on the referenced drawing, however, the openings appear similar to those shown on the Unit 1 Region 15 top nozzle drawing. To account for any possible reduction in open flow area, the open area of the Unit I top nozzle will be reduced by ten percent for use in the computation of the limiting l top nozzl: loss coeflicient.
Limiting Open Area for Top Nozzle Flow Openings = (0.9)*(31.53 in2) = 28.38 in2 The ratio of open flow area to total area, applicable to the OFA fuel, is computed as:
2=A - = 28.38 / 48.453 = 0.586 A,
p = [ = 0.766; g =(B )2 2
= 0.343 The loss coeflicient for the top nozzle, K ,, is computed as:
K= 0 = 2
( 1 - 0.586 ) / [ (0.6 )*(0.343) ] = 3.35 C'$
The local loss coefficients used in the COBRA-SFS model are shown in the table below.
Description Axial Coordinate Non-dimensional Loss Coefficient Axial Coordinate (inches) (N/A) l Base Plate Flow Hole, Ko 0.0 0.0 16.97 Bottom Nozzle 0.0 0.0 11.64 Bottom of Active Fuel 0.0 0.0 Grid Spacer 7.2 0.05 4.4 Grid Spacer 21.6 0.15 4.4
, Grid Spacer 36.0 0.25 4.4 l
IMF Grid Spacer 50.4 0.35 4.4 Grid Spacer 64.8 0.45 4.4 IMF Grid Spacer 79.2 0.55 4.4 Grid Spacer 93.6 0.65 4.4 IMF Grid Spacer 108.0 0.75 4.4 Grid Spacer 122.4 0.85 4.4 Grid Spacer 136.8 0.95 4.4 Top of Active Fuel 144.0 1.0 Top Nozzle 144.0 1.0 3.35 Exit Loss 144.0 1.0 1.0 REVISION NO. 2
Exhibit E N EP-12-02 COMMONWEALTil EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 25 6.6 IJeat Transfer Correlations for Assembly Subchannels Convective heat transfer coeflicients for heat exchange between the fuel rods and the fluid are specified in the COBRA code using the following relation (see Equation 6.8 of Reference 5.3):
h = (a, Re"2 Pr+ a.)I D,
where: ai,a2,as,a4 == cmpirical coeflicients Re = Reynolds Number Pr = Prandtl Number k = thermal conductivity of the fluid D, = channel hydraulic diameter The heat transfer correlations are entered in pairs, with one set of coefficients for laminar flow and one set for turbulent flow. The COBRA code then employs the maximum value obtained with the laminar and turbulent equations, where the local Reynokis number and Prandtl number are calculated internally by the code.
Reference 5.3 (page 7.5) cites the following correlations for subchannel flow.
hu,,,,,,, = (0.83 Re " Pr ")I v.
h,,,,,,,, = (0.33 Re" o. Pr") k These correlations were used in the PNL validation case entitled Single-Assembly lleat Transfer Test (SAliTT) as documented in Ref. 5.3. The use of these correlations resulted in very close agreement between the calculated temperatures by COBRA and the measured test data. The temperatures predicted by the COBRA code conservatively bound the test data, per the validation documented in Reference 5.3.
6.7 {}ogndary Conditions fpr COBRA Modgl Boundary conditions required for the COBRA model include the following:
- 1) Inlet air temperature at the base of the spent fuel storage cell, T,,,
- 2) inlet average mass flux at the base of the spent fuel storage cell, G The air temperature at the inlet to the spent fuel storage cell is taken as the Fuel llandling Building air temperature.
Rev.2 This temperature is computed as 282.17 F for reference date of 06/30/99, as documented in Attachment D of this calculation. This temperature is conservatively rounded up to 283.0 F for the COBRA model.
Rev. 2 The average inlet mass flux is selected by an iterative procedure, as described below and shown in Table 3. At steady-state conditions, natural circulation pattems will be established in the Fuel fiandling Building as shown in REVISION NO. 2
r Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 26 Figure 1, with the inlet air being drawn towards the spent fuel pool cavity as the heated air rises past the fuel assemblies, heating the Fuel Handling Building environment. Under these conditions, the total pressure drop for l upward flow along the fuel rods will match the pressure gain for the return flow as the air returns to the bottom of
- the spent fuel pool. Equating these pressure differences provides a means of determining the mass flux for flow in l the modeled spent fuel storage cell.
At steady-state, the one-dimensional conservation of momentum equation for upwards flow along the heated length I-of the fuel assembly is given below, with positive taken as the upwards direction.
Rev.3 OpVV cP
= pg -local loss per unit length Substituting for the mass flux G, where G = pV results in:
Rev.2 e
= pg -local loss per unit length
&<ps &
Integrating the momentum equation over the control volume from the bottom of the fuel assembly, subscript o, to the top of the fuel assembly, subscript L, yields the following equation.
Ret 2 eggt d:-(g,+g,,,,,,,)g2 - [ g,g 2 (g,,+g,}G - p(:)gd:
- ;2 t,
= -(P,' - P,) - _
<pi, , 2p(:)D, 2p, , 2p, 2p, y Defining AP = (P, - P ),tthe above equation is rearranged as 1' + ! f;
+[g*g: + [g,,+g,)g2
' 2 g AP = G,' 1
-- d: + (K* + K,,"**)G + p(:)gd: l
<pt p,, ,2p(:)D, 2p, , 2p, 2pt ,
where the total pressure drop for upflow from o to L , AP, is comprised of the following components: l AP = AP ims. + APw rno,. + APs,,,piot, + AP,w .p.m. + AP,mo. + AP,.yu. l Since the precise nature of the temperature and density profile along the fuel rod heated length is not known a priori, the following approximaaons are used in place of the integrations.
For the friction component along the fuel rod surface:
JG' g; g fG*
- 2p( )D, 2p,D, For the gravitational term
4 p(:)gd: 2 p,gL An average air density is also used for the local form losses at the grid spaccrs, giving:
REVISION NO. 2
Exhibit E NEP-12-02 i COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 27
\
[K'G' s G' ZK, s 2Ps 2Pn o Substituting these approximations into the total pressure loss equation yields the following expression for upwards flow from point o to point L:
Rev.2 f j }S + flg2 + G' g" + g'""""' Z'e g"' + K' AP = G* --
+' + +p"gl
\Pt P) 2p ,D, 2p, 2p , 2pt The return flow path is from point L back to point o, through the plenum region above the spent fuel pool cavity, with recirculation back to the base of the spent fuel storage cell. Frictional losses are neglected, since the return flow is through the relatively large open areas of the Fuel Handling Building and open regions of the spent fuel pool. Under these conditions, the momentum equation for the return flow has the following form.
O 'G ' 5'
&\pi & #
integrating from point L to point o gives rg23 *
= -(P, - Pt )- f p(:)gd: '
epi t i which simplifies to:
AP = G'
'1 1' "
p(:)gd: i
\ Pt Po) 't 1 Upon exiting the top of the spent fuel assembly, the air mixes with the other parallel flows from the other spent fuel storage cells, reaching the homogeneous air temperature of the Fuel Handling Building documented in . Attachment j D. Therefore, the density of the return air flow can be considered as constant, consistent with the Fuel Handling Rev.2 Building air temperature computed in Attachment D. The gravitational term simplifies as a result of the constant l density, with the appropriate length scale for downwards flow taken as the hel Horizontal flow along the bottom of the spent fuel pool cavity does not contribute to the gravitational integral, as no elevation changes occur. Therefore, the pressure difference for return flow from point L to point o is expressed as:
e AP = G'
+ pnraRH
\ Pt P)
Equating the expressions of pressure difference for the upwards flow and the return flow gives:
j Rev.2 - -
2 JL K ,+ K,,,,, , + , K" K ,,+ K, G + + + p,gL = pniegH ,
2P, D, 2p, 2p , 2pt '
REVISION NO. 2
l i
Exhibit E NEP-12-02 COMMONWEALTH EDISON C6MPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 28 i
Solving for the mass flux G results in the following:
Rev.2 - -
U2 '
g, (PmaH - p,L)g ,
JL + K, + uK ,,,,,,, + [,+ K,,,+ K K,
_ 2p, D, 2p, 2p , 2pt _
All of the quantities in the above expression are known with the exception of the friction factorf, the average air )
density p,, and the outlet air density at the top of the heated fuel assembly, pt. For a given guess value of l assembly inlet air mass flux, the outlet and average air temperatures (and densities) can be determined. The friction factor can be evaluated from the correlation presented in Section 6.5.1. The iterative procedure used to determine the appropriate assembly inlet air mass flux is shown in Table 3, and follows the algoritlun described below.
- 1) Choose guess value of assembly inlet mass flux, G (Ibm /sec-ft2 )
Rev.2
- 2) Compute outlet air temperature, based on an inlet air temperature of Attachment D. The peak decay heat for a fuel assembly, q.,,,,, (Bru/hr), is obtained from Table 1. The outlet air temperature can then be computed according to the equation:
]
T", = 7l" + 4"""' !
3600sec/ hr
- G
- c,
- A ,,,,
>vhere the specific heat, c,(Btu /lb.- F) corresponds to a temperature of 700 F, which is representative of the average air temperature along the heated length, per Assumption 3.2.6.
- 3) The average air temperature for upwards flow in the assembly is then obtained.
7L = 0.5 * (7l,, + T.,)
- 4) The outlet and average air densities are obtained using the ideal Gas law at the respective air temperatures:
p(T) = pE-.
- 5) The friction factor is determined by the correlation given in Section 6.5.1, where the Reynolds number is
' based on the guess value of the assembly inlet air mass flux, calculated as shown below.
Re = GD' P
The air viscosity is evaluated from a curve fit to the data provided in Design input 4.5.1. The curve fit is shown in Figure 3 as a function of air temperature, and matches well with the viscosity data.
REVISION NO. 2.
Exhibit E NEP-12-02 '
l COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 29 i
l 6) Using the air densities and friction factor determined in steps 4 and 5 respectively, the expression for G derived earlier by equating the pressure differences along the upwards and return flow path, is evaluated i and denoted as G'.
l i v2 i
'g_ . (PnmH - p,,L)g JL + K,+K,,,,,,,+[,K'r+ K,, + K,
_ 2 p,,, D, 2p, 2p,, 2pt _
The guess value of assembly inlet air mass flux, G, is adjusted until convergence is obtained with the calculated value of G'
. Rev.2 The calculation details of the iterative procedure are chown in Table 3 for the reference date of 6/30/99, with the boxed values being used as input to the COBRA-SFS model. For a reference date of 6/30/99 with an inlet air temperature of 283 F, the inlet air mass flux for the modeled spent fuel assembly is 1.6582E-04 Mlbm/hr-f1,2 as shown in Table 3. Also shown in Table 3 is the total pressure drop for upwards air flow along the spent fuel assembly, d', as well as each of the pressure loss components. The total pressure drop computed in Table 3 of i 0.00514 psi will be compared to the COBRA-SFS output file as a check on the accuracy of the modeling described in this section of the calculation. Also shown in Table 3 is the Reynolds number for flow in the spent fuel storage cell. The Reynolds number is in the laminar range, with a value of 113.2.
REVISION NO. 2
c t
Exhibit E NEP-12-02 COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. 30
- 7.
SUMMARY
AND CONCLUSIONS l
, Rev.2 The COBRA-SFS computer code is used to determine the maximum fuel rod cladding temperatures following a i postulated loss of cooling water from the Zion Station spent fuel pool. The analysis is performed for a reference date of 6/30/99. The air temperature of the Fuel Handling Building is computed in Attachment D for the 6/30/99
{'
reference date.
The results of the COBRA-SFS analysis for determining the maximum spent fuel cladding temperature following a (
complete loss of spent fuel pool cooling water are presented in the table below.
)
1 COBRA-SFS Reference FHB Air Maximum Cladding Output Filename Date Temperature Temperature cob 63099.out 6/30/99 283.0 F 899.3 F This COBRA-SFS input file cob 63099. inp, as well as the COBRA output file listed in the table above, are i
' documented in Attachment B. Also documented in the COBRA-SFS output file are the pressure drop for upwards flow in the spent fuel assembly, and the assembly average outlet air temperature. For the reference date of 6/30/99, the COBRA code predicts a pressure drop 0.00501 psi and an assembly average outlet air temperature of 895.0 F.
These values compare favorably with the values shown in Table 3,0.00514 psi and 889.5 F respectively.
It is shown that the peak cladding temperature following loss of cooling water is below the critical oxidation temperature of 1049 F for the reference date of 6/30/99. Therefore, rapid oxidation of the zircaloy cladding will not occur as a result of a sudden loss of spent fuel pool cooling water, provided the postulated accident occurs after 6/30/99.
Jabks
- Table 1 Decay Heat Generation for Spent Fuel Rods
. Table 2 Axial Power Profile
. Table 3 Boundary Conditions for COBRA Model (Reference Date of 6/30/99) l l REVISION NO. 2 Ll
Comed Calc. No. 22S-0-110M-0063 Zion Str. tion Units 1 and 2 Revision 2 i Tcble 1 -
Project No. 10436-003 Page No. 31 Decay Heat Generation for Spent Fuel Rods Parameter Units Value Basis Reference Date 6/30/99 Peak Decay Heat per Assembly, q, (Btu /hr) 8561 Design Input 4.7 Total number of rods per Assembly 225 Design Input 4.1 Number of Unheated Rods 21 Design input 4.1 Number of Heated Rods,Nw 204 Design input 4.1 Peak Decay Heat per Heated Rod,qa (Blu/hr) 41.97 q_/N w Rod Outer Diameter, d, (inches) 0.422 Design input 4.1 Rod Heated Length, L (inches) 144 - Design input 4.1 Fuel Rod Volume based on Heated Length, Vol (ft') 0.011656 =n/4*do :
- L
. Peak Rod Volumetrie Heat Generation Rate (Blu/hr-ft') 3600.5 = qa / Vol 8
Peak Rod Volumetrie Heat Generation Rate (MBtu/hr-f1 ) 0.003600 Unit conversion t
S I
Ziozirc2.xis Decay Heat Generation
I
.. OImEd f.alc. No. 22S-0-110M-0063 Zion Station Units 1 and 2 Revision 2 Table 2 Project No. 10436-002 Page No. 32 Axial Power Profile Active Fuel Length, L (inches) 144 Number of axial segments, N . 24 Axial Segment Height in Active Fuel Region, Az (inches) 6 Nondimensional Relative Power
- Description' Elevation Elevation Factor Power Fraction z z/L x(z) x(z) * (Az/L)
(inches) (N/A)
I Bottom of Active Fuel 0.0 0.000 0.000 0.000 6.0 0.042 0.205 0.009 l 12.0 0.083 0.407' O.017 18.0 0.125 0.601 0.025 24.0 0.167 0.785 0.033 30.0 0.208 0.956 0.040
, 36.0 0.250 1.I11 0.046 42.0 0.292 1.246 0.052 48.0 0.333 1.360 0.057 54.0 0.375 1.451 0.060 j 60.0 0.417 1.517' O.063 66.0 0.458 1.557 0.065 72.0 'O.500 1.571 0.065 78.0 0.542 1.557 0.065 84.0 0.583 1.517 0.063 90.0 0.625 1.451 0.060 96.0 0.667 1.360 0.057 102.0 0.708 1.246 0.052 108.0 0.750 1.111 0.046 !
114.0 0.792 0.956 0.040 ;
i 120.0 0.833 0.785 0.033 126.0 0.875 0.601 0.025 132.0 0.917 0.407 0.017 138.0 0.958 0.205 0.009
' Top of Active Fuel 144.0 1.000 0.000 0.000 Total of Segment Power Fractions 1.00 l
1 ZIOZlRC2.XLS AxialPowerDistsibution ,
i
7' 1-Comed Calc. No. 22S-0-110M-0063 Zion Station Units I cnd 2 - TEble 3 Revision 2 Nect No. 10436 003 Page No. 33 Boundary Conditions for COBRA Model (Reference Date of 6/30/99)
Description Units Value Basis Peak Decay Heat Generation Rate per Assembly, q.,,, (Btu'br) 8561 Table 1 Assembly-AverageIlydraulic Diameter D. (ft) 0.04834 Section 6.4.1 Fuel Rod Length, L (ft) 12.68 Design input 4.1 Total Fuel Assembly Flow Area A (ft*) 0.33648 Section 6.4.1 i l~
l Storage Celf licight above Baseplate,11 (ft) 14.0 Design Input 4.1 1
Air inlet Conditions Fuel Handling Building Air Temperature, T. ('F) l 283 l Attachment D Pressure, P (psia) 14.7 - Assumption 3.2.5 !
Air Gas Constant, R (ft-lb,1b. 'R) 53.34 Design Input 4.6.1 )
l Density, p. (Ib,./ft') 0.0534 -P.*144/R/(T +459.67) )
Total Mass Flux for Assembly, G (ib./sec-ft') 0.04606 Selected by twration Total Mass Flux for Assembly, G (Mlb./hr-ft') l 1.6582E-04 l=G'3600/IE6 Air Outlet Conditions Outlet Air Temperature, T ('F) 889.5 -T.+q /(3600*G*A *c,)
Density, p., (Ib./ft') 0.0294 -P.*144/R/(T,,.+459.67)
Air Average Conditions j Ttmperature, T. ('F) 586.2 -(T. + T y2 I Density, p. (lbsft') 0.0379 -P.*144/R/(Tm+459.67)
Dynamic Viscosity, p . (lb./ft-sec) 1.97E-05 Curve fit (Figure 3)
Specific Ileat, e, (Bru'Ib.- F) 0.253 ' Assumption 3.2.6
)
Pressure Drop for Upwards Flow Along F:sel Rod Reynolds Number, Reo , !!3.2 -G'D/p y Friction Factor based on Reynolds Number, f 0.904 -102.4/ Rep.
Coc5cient for Base-Plate Hole Losses, K. 16.97 Section 6 5.2.1 Coemcient for Bottom Nozzle, Km I1.64 Section 6.5.2.3 l Coefficient for Spacer losses, K., 4.40 Assumption 3.1.3 Number of Spacers 10 Design Input 4.1 Chamel Loss Coeflicient, Ka = f*L/D. + IK 281 -f*L/D. + 10*K,,
Cocilicient for Top Nozzle, K., ' 3.35 Section 6.5.2.3 Coeflicient for Exit Losses, K. 1.0 Assumption 3.1.4 Solve for Total Assembly Mass Flux, G' (Ibssec-ft') 0.04606 Section 6.7 Check forConvernence Dif!'erence in Mass Fluxes (G G') 0.0000 -G G' Pressure Droo Components Acceleration Pressure Drop, AP w (psia) 0.000007 -G'*(1/p - 1/p.)'(144*32.2)
Fuel Rod Frictional Pressure Drop, APr (psia) 0.00143 -(f*llD.)*G'/(144'2*p *32.2)
-G'*[(K xw )'(2'p.)+ 10'K.,T2*p.)+(K,,,+ K.) l' local Form Loss Pressure Drop, APx (psia) 0.00036 /(2*p )}/(144'32.2)
Gr:vitational Pressure Drop, AP,,, (psia) 0.00334 pm*32.2*l/(144'32.2)
Total Pressure Drop for Upwards Flow along Fuel Rod, AP (psia) 0.00514 -AP., + APr + APx + AP,,,
ZlRC-RV2.XLS Boundary Conditions (6-30 99) iI
3 2l a 6
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Comed Calc. No. 22S-0110M4063 Zion statica Units 1 aul 2 Tr ble A l- RMsion 2 Iwject No.1006 003 Spreadsheet Cell Formulas for Table 1 * * #*8* ~^I 4 I a i c I D 2
Parameter Units Value Basis 1
g Reference Date 36341 6 Peak Decay 11 eat per Assembly , % (Btu /hr) 8561 Design input 4.7 1
s Total number ofrods per Assembly 225 Design Input 4.1 9 Number of Unheated Rods 21- Design Input 4.1 Number of11eated Rods, Nwa =C8-C9 Design input 4.1 3
2 12 Peak Decay 11 eat per lleated Rod, % (Btu /hr) =C6/C10 =%/ Nwa 2
14 Rod Outer Diameter, d, . (inches) 0.422 Design input 4.1 3 Rod lleated Length, L (inches) 144 Design input 4.1 Fuel Rod Volume based on IIcated Length, 3 Vol (it') =PIO/4*C14^2*C15/12^3 =x/4 *d,'
- L 2 .
Peak Rod Volumetric IIcat Generation 3 Rate (Btu /hr-ft') =C12/C16 =%/Vol Peak Rod Volumetric Heat Generation 20 Rate (MBtu/hr-ft') =C18/1000000 Unit conversion
- Ziozirc2. mis Decay Heat Generation (Eqns)
p I
Comed Cele. No. 22S-0-110M4)063 l Zion Station Units 1 and 2 Table A2 Revision 2 Project No.1043W2 Atta hment A Page No. A2 i
Spreadsheet Cell Formulas for Table 2 l
A l B l C l D l E
+ l Active Fuel Length, L I
144 l 1(inches) l Number of axial .
1 segments, N 24 Axial SegmentIIcight in Active Fuel Region, Az l 5 (inches) =D3/D4
_.6_
2 Nondimensional i
8 Description Elevation Elevation Relauve Power Factor Power Fraction j 9 z z/L x(z) x(z) * (Az/L) 10 (inches) (N/A)
I!
12 Bottom of Active Fuel 0 =B12/$D53 =PI(y2* SIN (PI()*Cl2) =D12*$D$5/$D$3 i
j =B12+$D$5 =B13/$D$3 =PI(y2* SIN (PI()*Cl3) =D13'$D$5/$D53 14 =B13+$D$5 =B14/$D$3 =PI(y2* SIN (PI()*C14) =D14'$D$5/$D$3 3 =B14+$D$5 =B15/$D$3 =PI(y2* SIN (PI()*C15) =D15*$D$5/$D$3 Y =B15+$D$5 =B16/$D$3 =PI(y2* SIN (PI()*C16) =D16'$D55/$D$3 T7~ =B16+$D$5 =B17/$D$3 =PI(y2' SIN (PI()*C17) =D17'$D$5/$D$3 3 =B17+$D$5 =B18/$D$3 =PI(y2* SIN (PI()*C18) =D18*$D$$/5D53 y =B18+5D55 =B19/$D$3 =PI(y2* SIN (PI()*C19) =D19'$D55/$D$3 20 =B19+$D$5 =B20/$D$3 =PI()/2* SIN (PI()*C20) =D20'$D$5/$D53
[22
=B20+$D$5
=B21+$D$5
=B21/$D$3
=B22/$D$3
=PI(y2* SIN (PI()*C21)
=PI(y2* SIN (PI()*C22)
=D21 *$D$5/$D$3
=D22*$D$5/$D$3 23 =B22+$D$5 =B23/$D$3 =PI(y2* SIN (PI()*C23) =D23'$D$5/$D$3 5 =B23+$D$5 =B24/$D$3 =PI(y2* SIN (PI()*C24) =D24'$D$5/$D$3 Y =B24+$D$5 =B25/$D$3 =PI(y2* SIN (PI()*C25) =D25*$D$5/$D$3 5 =B25+$D$5 =B26/$D$3 =PI(y2' SIN (PI()*C26) =D26'$D$5/$D$3 3 =B26+$D$5 =B27/$D$3 =PI(y2' SIN (PI()*C27) =D27'$D$5/$D$3 28 =B27+$D$5 =B28/$D$3 =P)(y2* SIN (PI()*C28) =D28'$D$5/$D$3 29' =B28+$D$5 =B29/$D$3 =PI(y2' SIN (PI()*C29) =D29'$D$5/$D$3 Y =B29+$D$5 =B30/$D$3 =PI(y2' SIN (PI()*C30) =D30*$D$5/$D$3 Y =B30+$D$5 =B31/$D$3 =PI(y2' SIN (PI()*C31) =D31'$D$5/$D$3 Y =B31+5D$5 =B32/$D$3 =PI(y2* SIN (PI()'C32) =D32*$D$5/$D$3 3 =B32+$D$5 =B33/$D$2 =PI(y2* SIN (PI()*C33) =D33*$D$5/$D$3 Y =B33+$D$$ =B34/$D$3 =PI(y2' SIN (PI()*C34) =D34'$D$5/$D$3 3 =B34+$D$5 =B35/$D$3 =PI(y2' SIN (PI()'C35) =D35'$D$5/$D53 36 Top of Active Fuel =B35+$D$5 =B36/$D$3 =PI(y2* SIN (PI()*C36) =D36*$D$$/$D$3 3.1 Total of Segment Power 38 Fractions = SUM (E13:E37)
ZIRCREV2.XLS Axial Power Distribution (Eqns)
r l l
l Comed l Colc. No 2254110M&53 Lo Staso Uruto I and 2 Table A3 Revision 2 Project No 10436 003 . Anadant A Page No A3 Spreadsheet Cell Formulas for Table 3 A I H I e i n Description Units Value Canis 2-peak Decay llent Generatson Rate per Assembly, 4 q (Btu /hr) 8561 Table 1 i
5 Assembly-Average Hydraubc Diameter, D. (R) 0 04534 Section 6 41 l t
1 Fuel Rod Length L (A) 12 68 Design input 41 7 3 TotalFuel AssemblyFlow Area, A (A ) 0 33648 Section 6 41 3 Storage Celllleight above Baseplate,11 (ft) 14 Design input 41 Air inlet Condet6 ens 1 Fuel Handhng Duilding Air Temperature T. (T) 223 Attachment D 3 pressure, P (psia) 14 7 Assumptusi 3 2 5 3 Air Oas Constant, R (A-lbylb *R) , 53 34 Design input 4 61 3 Density, gh (lb'ft') =Cl2+ I44tCl3*(Cl1+459 67)) =P.*144 R/(T.+459 67)
E 1 Total Mass ILx for Assembly, O (lbssec-ft') 0 0460623997308449 Selected by Iteration 3 Total Mass Flux for Assembly, O (Mib%r A') l=C16*360&l000000 l-O*360W1E6 5
3 Air Outlet Conditions y Outl* AirTemperature.T (T) =Cil+C4/(3600*C16*C7*C27) =T.+Q3600*G*A *c,)
3 Deck J (Ib/ft') =Cl2*144'(Cl3*(C20+459 67)) =P.*144%TTs459 67)
E g Air Average Conditions 3 Tunperatta 6. T (T) =(Cil+C20y2 =(T.* T y2 g Density, p (ibit') =Cl2*l44TCl3*(C24+459 6 T)) =P.*144RTT +459 67) 3 Dynamic Viscosity, p (lb'flac) =0 000012 t000000001308+C24 Curve fit (Figure 3)
E Specific iteal, c, (Btu'Ib T) 0 253 Assumption 3 2 6 5
29 Pressure Ihp for Upwards Flow Along Fuel Red n 1 3 Reynolds Number, Rep. =C16*CS/C26 =0*D.'p 1
l g Friction Factor based on Reynolds Number, f -102 4/C31 alm 4Re ue 33 Cocmcient for Base Plate liole Losses K. 16 97 - Section 6 5 21 M Coemcient for Bottom Nozzle, Km 11.64 Section 6 5 2 3 M Coemcient for Spacer Losses, K , 4d Assumption 313 .
M Number of Spacers 10 Design input 41 37 Channel las Cnemeient. K,s = f*UD, + EK *C32*C6/C5 +C36*r't5 af*1]D, + 10*K,,
Z.NC RV2 XL5 Doundary Conditions Eo.a (6 30)
i-CornFA Calc No 22 Sol 10M 0063 Zion Susan Umts 1 and 2 Table M Revision 2
' Nject Na W36-W3 Attachment A Page No A4 Spreadsheet Cell Formulas for Table 3 i' fWO A I H l C l D M Coefreient for Top Nozzle, K , . 3.35 Section 6 $ 2 3 M Coefficient for Exit Losses, K. 1 Answnpuan 314
=SQRT((Cl4*C8 .
. . C25*C6)*3217/(C37/(2*C25)+(C33+C34y(2*Cl g Solve for Total Assembly Masa Flux, O' (ib./sec-ft') 4)+(C38 +C39y(2*C21))) Section 6.7 ;
s a.ua,-- i i
M DifTerence m Mass Fluxes (0.G) C16 C40 =0. O I N
g bessee Drop Comoonents g Acceleration Pressure Drop, AP.a (psia) =l/(144*3217)*C16^2*(IC21-lCl4) =O4 (I/p I/psy(144*32.2) g Fuel Rod Fnchonal Pressure Drup, APr (psia) = 1 T 144* 32.17PC32 *C6*C16^2/(1*C5 +C25) -(l'UD.)*G'/(144*2*p ,*32.2) 8
=1T144*3217)*C16^2*(C31/(2*Cid)+C36*C35/ =O *[(Kramy(2*pw)+10*Kg(2*p P g local Form Loss Pressure Drop. APa (psia) (2*C25)+C38'(2*C21)) (K ,+K,y(2*p.)}T144*32.2) 1 49 Oravuauanal Pressac Drop, AP,,, (psia) =l/(144*3217PC25*32.17'C6 =p *32 2*L(144*32 2)
E Total Presswe Drop for Upwards Flow abng Fuel 51 Rod, AP (pain) =C46+C47+C48 *C49 =AP,,a + APr + APu + AP,,,
l l
t l
I l
Z1RC-RV2 XLS Boundary Condstions Eqns (&30)
Exhibit E NEP-12-02 COMMONWEALTII EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. B1 Final ATTACIIMENT B i Rev. 2 COBRA-SFS Input / Output Files Loss of Spent Fuel Pool Cooling Water 1
(Microfiche) 1 Reference Date: Files Included: l l
Refbrence Date of 6/30/99 " cob 63 099. inp" (1 sheet ofmicrofiche) l " cob 63 099. ou t - (2 sheets of microfiche)
I i
l i
$ AN ACOMP' REVISION NO. 2
F 1 1 '
Exhibit E !
NEP-12-02 I COMMONWEALTH EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. C1 i
i ATTACIIMENT C l
Rev.2 RADGEN Input / Output Files i RADGEN Input File File:
"radgen.inp" l
0 RADGEN input; Calculation No. 22S-0-110M-0063 Rev. 0 01.334 15 15 0 0.4220.3180.3180.3180.318 0.8000.200 8
0.2000.2000.2000.2000.2000.2000.2000.200
-1 l
l REVISION NO. 2
Exhibit E NEP-12-02 COMMONWEALTII EDISON COMPANY CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. C2 Final ATTACllMENT C Rev.2 RADGEN Input / Output Files RADGEN Output File (Microfiche) filt "radgen.out" (1 sheet ofmicrofiche) l l
i I
\
$ ANACOMP' l l
REVISION NO. 2
COMMONWEALTII EDISON COMPANY l
l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. Di l Attachment D Computation of Fuel Handling Building Homogeneous Air Temperature DI PURPOSE / OBJECTIVE The purpose of this attachment is to compute the air temperature of the Fuel Handling Building (FHB)
Rev.2 following a loss of all water in the spent fuel pool. The FHB air temperature is determined for a reference date of 6/30/99, where the FHB is supplied with 21,000 cfm of air by the HVAC system. The FHB air temperature determined in this attachment is used as the inlet air temperature boundary ;
condition for the COBRA-SFS model described in the main body of this calculation.
D2 METHODOLOGY / ACCEPTANCE CRITERIA Methodology The air temperature in the FHB following a loss of spent fuel pool cooling water is computed using the Sargent & Lundy computer program KITTYlS, " Thermal-Hydraulic Steady States in Arbitrary Solid l and/or Fluid Channel Configurations", S&L Program No. 03.7.171-5.11. This computer program l calculates node temperatures and path heat /enthalpy transfer rates for user-defined configurations under steady-state conditions. The spent fuel pool heat load is modeled as a heat source to the FHB air space.
Enthalpy transport to/from the FHB air space via the HVAC system is modeled by specifying the mass flow rate of the HVAC inlet flow and the inlet air temperature. Presersing this mass flow rate at the outlet, the KITTYlS program will iterate on the FHB air space temperature until the net energy balance Rev. 2 on the FHB is satisfied. Natural convection heat transfer paths from the FHB air space to the venical interior walls and ceiling of the FHB are included, as well as convection heat transfer paths from the exterior FHB walls and ceiling to the ambient. Conduction heat transfer paths through the concrete walls and ceiling are also included in the KITTYlS model. The FHB concrete walls are of varying thickness, with values of 1.5,2.83, and 3.0 feet as shown in the Zion Station FHB Foundation drawings (Reference Rev.2 l D5.3). Each vertical wall type (1.5, 2.83, or 3.0 feet total thickness), and ceiling (1.0 feet thickness) is )
partitioned into three nodes of equal surface area. The surface nodes facing the FHB air space and the ambient are modeled as 1 inch in thickness. Natural convection heat transfer to/from these surface nodes is modeled, and conduction heat transfer through the central node is modeled to establish the temperature gradient through the concrete wall.
A schematic diagram of the Fuel Handling Building is shown as Figure D1 on the following page.
l REVISION NO. 2 l
p COMMONWEALTH EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D2 l Figure D1
, Fuel Handling Building l
l l
Bottom of Roof; Elev. 663' Elev. 617'
,-.~---- y
/ Spent /I
/ Fuel /l
/ Pool / i p--- ---4 l '
l i l )
l l l /
ll L ---- J
/
\
Elev. 592' North Rev. 2 The KITTYlS model conservatively neglects any heat transfer to the concrete floor. Also, radiative heat transfer from the air space to the vertical wall or ceiling surfaces is neglected, as well as radiative heat transfer from the heated outside walls or ceiling to the environment.
A schematic diagram of the FHB KITTYlS modelis shown on the following page.
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l REVISION NO. 2 l
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l l COMMONWEALTil EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D3 l Figure D2 KITTY 1S Model of FIIB Air Temperature Conduction !
3 1.5 feet thick walls HVAC Convection Convection 4 5 6 Outlet Air --
n
^
h 2.83 feet thick walls 7 8 9 (y) j 1
FHB h h @ >
13 !
Ambient
_ 3.0 feet thick walls v
b 10 11 12 i
h h h) s 2 ^ 1.0 feet thick ceiling -
u (15) (18)
HVAC 14 15 16 Inlet Air --
(N h Nodes O Paths l
l Acceptance Criteria The are no specific acceptance criteria applicable to this attachment. The FHB air space temperature computed in this attachment is used as input to the COBRA-SFS model.
l REVISION NO. 2 l
1 COMMONWEALTH EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D4 l D3.0 ASSUMPTIONS Rev. 2 D3.1 - The Fuel Handling Building HVAC system is assumed to maintain a volumetric flow rate of 21,000 cfm into the FHB, consistent with the value cited in Design Input D4.5.
D3.2 The ambient air temperature is assumed to be 95 F. This conservatively bounds the summer i design dry bulb temperature of 92 F for Waukegan, IL, per Chapter 24 (Table 1) of the 1993 )
ASHRAE Handbook, Reference D5.1.
D3.3 The FHB is assumed to be at a pressure of 14.7 psia. This pressure, as well as the FHB air temperature, is used to compute air density via the Ideal Gas Law. Small deviations in pressure have a negligible efTect on the computed density, relative to the large changes in air temperature.
Also, the presence of the operational HVAC system is expected to limit pressure changes in the FHB.
Rev. 2 l D3.4 Deleted l
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4 l l REVISION NO. 2 1 l
I COMMONWEALTH EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. DS l D4.0 DESIGN INPUT l Rev. 2 D4.1 Dimension for the vertical walls and ceiling of the Fuel Handling Building are listed below, with appropriate Zion Station FHB Foundation drawings (Reference D5.3) as cited.
Thickness length Height Area Reference North Wall Elev. 592' to 617', 3.0 feet 71.17 feet 25 feet 1779.3 ft 2
Dwg. Nos. B-107, Columns X - W B-121, B-122, B-123 j Shield Door at N/A 18.0 feet 20.0 feet -360.0 ft2 Dwg. No. B-654 Elev 592' Elev. 617' to 663', 1.5 feet 71.17 feet 46 feet 3273.8 ft 2
Dwg. No. B-420 <
Columns X - W l
Elev. 617' to 663', 1.5 feet 35.0 feet 46 feet 1610.0 R 2
Dwg. No. B-421 Columns R - X South Wall Elev. 592' to 617', 3.0 feet 71.17 feet 25 feet 1779.3 ft 2
Dwg. Nos. B-107, l
1 Columns X - W B-121, B-122, B-123 1
Elev. 617' to 663', 1.5 feet 71.17 feet 46 feet 3273.8 ft 2
Dwg. No. B-420 Columns X - W
! Elev. 617' to 663', 1.5 feet 35.0 feet 46 feet 1610.0 ft Dwg. No. B-421 Columns R - X West Wall Elev. 592' to 617', 2.83 feet 84 feet 25 feet 2100.0 ft: Dwg. Nos. B-107, Columns 17 - 23 B-118, B-119, B-120 2
Elev. 617' to 663', 2.83 feet 87 feet 46 feet 4002.0 ft Dwg. No. B-420 Columns 17 - 23 l Total Wall Surface Areas 2
l 1.5 feet Wall Thickness 9767.6 ft 2.83 feet Wall Thickness 6102.0 n2 1 2 3.0 feet Wall Thickness 3198.6 fl Rev. 2 Ceiling Thickness I,ength Length Area Reference 2
Columns 17 - 23 1.0 feet 87 feet 13137.0 ft Drawing. Nos.
Columns R - W 151 feet B-404, B-405 l l REVISION NO. 2 l 1
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L
COMMONWEALTH EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D6 l D4.2 The physical properties of concrete are taken from Reference D5.4, and are shown below.
Density 145 lbm/ff I Specific Heat 0.156 Btu /lb.- F Thermal Conductivity 0.92 Btu /hr-ft- F D4.3 The following thermophysical properties of air at atmospheric pressure are taken from Reference D5.2. Additionally, the air properties cited as Design Input 4.5.1 (see main body of the calculation) are also used in the KITTY 1S model.
Temperature k gpp g2j 2 Pr (F) (Btu /hr-ft- F) (1/ F-ff) (N/A) 6 100 0.0154 1.76 x 10 0.72 6
200 0.0174 0.85 x 10 0.72 6
300 0.0193 0.444 x 10 0.71 6
l 400 0.0212 0.258 x 10 0.689 6
500 0.0231 ~0.159 x 10 0.683 Rev.2 6 D4.4 The total spent fuel pool heat load due to spent fuel is 5.5 x 10 BTU / hour, as computed in Reference D5.5.
D4.5 Per Reference 5.20 (see main body of the calculation), the Auxiliary Building HVAC System I currently supplies 21,000 cfm of air to the Fuel Handling Building Space.
Rev.2 D4.6 The maximum temperature of the HVAC inlet air to the Fuel Handling Building is 95 F per Reference D5.5.
l l REVISION NO. 2 l
t COMMONWEALTil EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D7 l D
5.0 REFERENCES
D5.1 1993 ASHRAE Handbook Fundamentals, I-P Edition.
D5.2 Principles of Heat Tramfq, Frank Kreith, Third Edition,1976 Impression D5.3 Zion Station Fuel Handling Building Foundation Drawings i
B-107, Rev. K, FHB Foundation Plan El. 617'-0" West Area B-108, Rev. P, FHB Foundation Plan El. 617'-0" East Area l B-109, Rev. J, FHB Foundation Plan El. 602'-0" West Area i B-118, Rev. L, FHB Foundation Section A-A B-119, Rev. R, FHB Foundation Section B-B B-120, Rev. N, FHB Foundation Section C-C B-121, Rev. H, FHB Foundation Section D-D B-122, Rev. E, FHB Foundation Section E-E -
)
B-123, Rev. M, FHB Foundation Section F-F B-420, Rev. A, FHB Wall Plan El. 617'-0" West Area B-421, Rev. A, FHB Wall Plan El. 617'-0" East Area B-654, Rev. G, FHB Ground Fl. Plan El. 592'-0" Rev.2 B-404, Rev. G, FHB Roof Framing Plan West Area B-405, Rev. L, FHB Roof Framing Plan East Area D5.4 NUREG 0800, Standard Review Plan, Rev.1, Section 6.2.1.5 Rev.2 D5.5 NDIT NFM9900121, Sen. No. O,
Subject:
Zion Spent Fuel Pool Data,6/4/99.
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I l REVISION NO. 2 l
COMMONWEALTII EDISON COMPANY ,
I CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D8 l D6.0 CALCULATIONS D6.1 Air Mass Flow Rate of Auxiliary Building HVAC System To compute the mass flow rate ofinlet air (required by the KITTYlS model), the inlet air density must be determined using the Ideal Gas Law as:
P P=g where P = 14.7 psia, per Assumption D3.3 an. 2 R = 1545(ft-lb/lbmole- R) / 28.97(lb /lbmole) = 53.33(ft-lb/lbm- R)
T= 95 + 460 = 555 R, per Design Input D4.6 Solving for the air density gives the following:
Rn2 2 p = 14.7(psi)
- 144(in#/ft ) / [ 53.33(ft-lb/lbm- R)
- 555( R) ] = 0.0715 lbm/R 5 The volumetric flow rate of FHB inlet air,21,000 cfm, is converted to a mass flow rate.
For 21,000 cfm:
di = 21,000 (fl'/ min)
- 0.0715 (Ibm /ft')
- 60(min / hour) = 90090 lbm/ hour This value is used as input to the KITTY 1 S model for heat transfer paths 1 and 2.
D6.2 Volume of Vertical Concrete Walls and Ceiling A required KITTY 1S input parameter is the volume of solid nodes. The concrete wall and ceiling node volumes are determined by multiplying the node thickness by the surface area, for each respective wall or ceiling node type. The surface nodes are 1 inch in thickness, with the central node thickness set to the total minus 2 inches as shown in the table below.
Wall Thickness Area Surface Nodes Central Node Thickness Volume Thickness Volume 2
1.5 feet 9767.6 f1 1 inch 814.0 n' l.33 feet 13020.2 fl' 2 3 2.83 feet 6102.0 ft 1 inch 508.5 ft' 2.67 feet 16274.0 fl 2
3.0 feet 3198.6 fl 1 inch 266.6 ft' 2.83 feet 9052.0 n' Ceiling Thickness 2
1.0 feet 13137.0 R 1 inch 1094.8 n' O.833 feet 10943.1fl' l REVISION NO. 2 l
1 l COMMONWEALTH EDISON COMPANY l CALCULATION NO. 22S-0-110M-0063 PROJECT NO. 10436-003 PAGE NO. D9 l l
l Rcy.2 D6.3 Natural Convection for Vertical and Horizontal Surfaces Reference D5.2 provides the following correlation for turbulent natural convection along vertical planes or cylinders, whare the local heat transfer coemeient is nearly constant over the surface:
Mit = = 0.13(Ge t Pr) >
for Gr > 10' For horizontal surfaces with a surface warmer than the surrounding media facing upward, or a cooler surface thcing downward (the orientation of the FHB ceiling), Reference D5.2 recommends the following relation for turbulent natural convection:
Mit =
= 0.14(Gr t Pr)
7 l for 2 x 10 < Grt Since these correlations are similar, a common expression will be used for evaluating the heat transfer coemcients at both the vertical walls and the horizontal ceiling surfaces. This is conservative with
- resnect to the ceiling heat transfer coemeients, and therefore results in conservative estimates of the Fuel Handling Building air temperature.
Justification for the use of the turbulent natural convection heat transfer correlation will be provided in l Section D7.0.
l
~
Substituting for Grt and Pr gives the following expression for the heat transfer coemeient for natural Rev.2 l convection to/from the vertical walls and ceiling of the FHB E, = 0.13 b Pr #
(7; - T,,,)
L < p >
with the air properties evaluated at the mean temperature, T,,,, between the air and the wall surface.
The heat transfer coemcient can be simply expressed as:
l E, = 0.13k Pr (T, - T,,,) = C(7; - T ,)
10'. The Grashof number is defined I as /gp//p'J
- L'
- AT. A lower bound on Gr is computed using the smallest value of/gp// ]shown in Section D6.3, the smallest height of the vertical walls listed in Design Input D4.1, and the minimum temperature difference between the wall surface and the air. The minimum value of ATis found at the l
exterior surface of the 3.0 feet thick concrete wall, per examination of the KITTY 1S output file. This value is 136.48 F - 95 F = 41.48 F. Computing the lower bound value of Gr gives: 6 Gr = [0.159 x 10 (1/ F-ft')] * [25.0 ft]' * [41.48 F] = 1.03 x 10" This lower bound value of Gr is above the required Grashof number necessary for turbulent natural convection along vertical surfaces. ! l For the horizontal ceiling, a lower bound on Gr is computed using the smallest value of /g //p'J shown I in Section D6.3, the smallest dimension of the ceiling listed in Design Input D4.1, and the minimum temperature difference between the ceiling surface and the air. The minimum value of ATis found at the i exterior surface of the ceiling, per examination of the KITTY 1S output file. This value is 157.19 F - 95 F = 62.19 F. Computing the lower bound value of Gr gives: 6 i2 Gr = [0.159 x 10 (1/ F-ft')] * [87.0 ft]' * [62.19 "F] = 6.51 x 10 This lower bound value of Gr is above the required Grashof number necessary for turbulent natural convection along horizontal surfaces. l REVISION NO. 2 I _ 32D2 - 6 1 0 nt iD - s v. 0io Miv no i O Ra 1 ed eN 1 t g APa t 0-S 3 2 0 2 0-o 6 N 3 4
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