ML20005A992
| ML20005A992 | |
| Person / Time | |
|---|---|
| Site: | Quad Cities  |
| Issue date: | 06/30/1981 |
| From: | JOSEPH OAT CORP. |
| To: | |
| Shared Package | |
| ML20005A990 | List: |
| References | |
| NUDOCS 8107060177 | |
| Download: ML20005A992 (101) | |
Text
{{#Wiki_filter:-- .. -._ . . . . - . . ~ -. . . _ . . . _ . - - . _ . . = - - .. . _ , i Revision 1 O i 4 4 1 Licensing Report on High-Density Spent Fuel Racks t For i i Quad Cities Units l and 2 I i Docket Nos. 50-254, 50-265 1 O i I 'l i June 1981 i l l I I l O 8107060177 810 4 PDR ADOCK 0500 54 P g,DR
. TABLE OF CONTENTS O Page
- 1. INTRODUCTION ............................................. 1-1
- 2. GENERAL ARRANGEMENT ....................................... 2-1
- 3. RACK CONSTRUCTION AND ASSEMBLY ........................... 3-1 3.1 Construction ........ ....*. - - - .. - 3-1 3.2 Codes, Standards, and Practices for the Spent Fuel Fool Modification .....................-... 3-3 3.3 Installation and Leveling ........................... 3-4
- 4. NUCLEAR CRITICALITY ANALYSIS ...............-............. 4-1 4.1 Design Basis ....................... .... .... . .. 4-1 4.2 Geometric and Calculational Models .................. 4-2 4.2.1 Ref_erence Fuel Assembly ...................... 4-2 4.2.2 Alt:ernative Fuel Assembly Designs ............ 4-3 4.2.3 Calculational Models ......................... 4-5 4.2.3.1 Analytical Methods .................. 4-5 s
4.2.3.2 Calculational Bias and Uncertainty .. 4-5 4.2.3.3 Trend Analysis ...................... 4-6 4.2.4 Reference Fuel Storage Cell ................... 4-6 4.3 Reference Subcriticality and Mechanical Tolerance Variations .. .................................... 4-7 4.3.1 Nominal Case .................................. 4-7 4.3.2 Maximum Enrichment Capability . . . . . . . . . . . . . . . . 4-7 4.3.3 Alternative Fuel Assemblies .................. 4-7 ! 4.3.3.1 Alternative Geometry and Enrichments ...................... 4-7 4.3.3.2 Plutonium-Bearing Experimental Fuel Asseinb11es ..................... 4-8 4.3.4 Boron Loading Variation ....................... 4-9 4.3.5 Storage Cell Lattice Pitch Variation ... ... 4-9 l 4.3.6 Stainless Steel Thickness Variations .......... 4-10 4.3.7 Fuel Enrichment and Density Variation ........ 4-10 4.3.8 Boraflex Width Tolerance Variation ............ 4-10 4.3.9 Effect of Zirconium Fuel Channel ............. 4-10 4.3.10 Summary of Statistical variations ............ 4-11 4.4 Abnormal and Accident Conditions ..................... 4-11 4.4.1 Fuel Assembly Positioning in Storage Rack .. . 4-11 4.4.2 Effect of Zirconium Flow Channel Distortion ... 4-13 () 4.4.3 Temperature and Water Density Effects ........ 4-13 i
Page n/ s_ 4.4.4 Abnormal Positioning of Fuel Assembly Outside Storage Rack ......................... 4-14 4.4.5 Missing Absorber Plate ....................... 4-14 4.4.6 Dropped Fuel Assembly Accident ............... 4-14 4.4.7 Fuel Rack Lateral Movement ................... 4-15 4.5 Summary ............................................. 4-15 References ................................a.......... 4-18
- 5. THERMAL HYDRAULIC CONSIDERATIONS .......................... 5-1 5.1 Heat Generation Calculations ......................... 5-1 5.2 Analysis of Pool Thermal Hydraulics .................. 5-1 5.3 Results ............................................. 5-3 i
References * * . *= = ** * * *' '* ' 5-4
- 6. STRUCTURAL ANALYSIS ....................................... 6-1 6.1 Analysis outline ...... - - - - .... - - - . 6-1 6.2 Fuel Rack - Fuel Assembly Model .................... 6-3 l
() 6.2.1 6.2.2 Assumptions .................................. Model Description ............................ 6-3 6-4 1 6.2.3 Fluid Coupling ............................... 6-5 6.2.4 Damping ...................................... 6-6 6.2.5 Impact ..................................... 6-6 6.2.6 Assembly of Dynamic Model .................... 6-6 6.3 Stress Analysis ........................... ... . .. 6-10 6.3.1 Stiffness Characteristics .................... 6-10 6.3.2 Combined Stresses and Corner Displacements ... 6-11 6.4 Time Integration of the Equations of Motion . ...... 6-12 6.5 Structural Acceptance Criteria ....................... 6-15 6.6 Accidents Associated with Rack Integrity ............ 6-19 . References ............ ..... ....... . . . 6-20 t i
- 7. ACCIDENT ANALYSES (Results Later) ....... ......... ...... 7-1 1
- 8. RADIOCHEMICAL CONSEQUENCES - . - .- .- 8-1 8.1 Objectives and Assumptions .................... .. .. 8-1
() 8.2 Operating Experience and Nature of Stored Fuel ...... 8-2 11
Page () 8.3 Consequences of Failed fuel ............... .. . 8-3 8.3.1 Methods of Analysis .......................... 8-5 8.3.2 Fission-Product Radionuclide Concentrations .. 8-6 8.3.3 Gaseous Releases from Failed Fuel . . . . . . . . . . . . . 8-7 8.4 Exposure for the Installation of New Racks .......... 8-8 8.5 Conclusions ......................................... 8-9 References .......................................... 8-10
- 9. POOL STRUCTURAL CALCULATIONS (Later) ...................... 9-1
- 10. INSERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING MATERIAL ........................................ 10-1 10.1 Program Intent ...................................... 10-1 10.2 Description of Specimens .................<>........ 10-1 10.3 Test ................................................ 10-1 10.4 Specimen Evaluation .................................. 10-2
- 11. COST / BENEFIT ASSESSMENT ................................... 11-1 11.1 Specific Needs for Spent Fuel Storage . . . . . . . . . . . . . . . . 11-1 11.2 Cost of Modification ............................... 11-2
() 11.3 Alternatives to Spent Fuel Storage Expansion ........ 11-2 11.4 Resource Commitments ................................ 11-4 11.5 Environmental Effects ...........................- ... 11-4 References .......................... . . . . . . . . . . 11-6 1 l 1 O 111
l I
- 1. INTRODUCTION O The purpose of this report is to provide descriptive information and performance and safety analyses on the installation and use of y ,
high-density spent fuel storage racks at Quad Cities station Units 1 and 2. The corresponding request to change the Quad Cities Terbrical Specifications to allow the use of high-density spent fuel storage racks was submitted to the NRC via letter dated March 26, 1981. Quad Cities nuclear power station consists of two generating units (Unit 1 und Unit 2), each with a General Electric BWR-3 reactor. The station is owned by Commonwealth Edison Company (75%) and Iowa- l1 Illinois Gas and Electric Company (25%), and is operated by Common-wealth Edison Company. The two utilities share the electrical output in proportion to the ownership. At the present time, the Quad Cities units have the following capacity in their spent fuel pools: e Storage racks for 2280 fuel assemblies. e Storage racks for 354 control rods. e Storage allocation for 40 channels, e Special containers for 6 defective fuel assemblies. Table 1.1 shows the previous and projected fuel discharge schedule for Quad Cities Units 1 and 2. After each operating cycle, approximately 150 to 200 fuel assemblies are transferred from the reactor to the spent fuel storage pool. Considering the present com-bined, spent fuel storage capacity of 2280 fuel assemblies, Table 1.1 indicates that following the fall 1981 refueling outage, insufficient fuel storage capacity will exist to receive a full core discharge of 724 fuel assemblies. Furthermore, following the 1933 refueling out-e;e, insufficient fuel storage capacity exists for a subsequent refueling discharge in excess of 158 hel assemblies. A limited number of spent fuel racks of the presently approved design are avail-able for installation in the Quad Cities Units 1 and 2 spent fuel storage pools, which could expand the storage capacity to 2920 fuel assemblies. As shown in Table 1.1, with additional spent fuel storage
- o 1-1
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s.) V b' Table 1.1 Ouad cities station, Units 1 ani.' Fuel Assembly Discharges Discharged Assemblies Total Discharged Remaining Storage Capacity
- Year Unit 1 Unit 2 Assemblies In Pool With Additional High-Density Existing Licensed Racks Racks 1974 64 144 208 2072 - -
1975 0 4 212 2068 - - 1976 156 164 532 1748 - - 1977 184 0 716 1564 - - 1978 0 180 896 1384 - - 1979 19? 180 1258 1012 - - 1980 224 0 1492 788 - - 1981 0 224 1716 564 1204 - 1982 210 0 1926 354 994 5758 ta 1983 0 196 2122 158 798 5562 1 1984 184 184 2490 0 430 5194
" 1985 192 2682 0 -
238 5002 1986 0 204 2886 - 34 4798 1987 200 200 3286 - 0 4398 1988 200 0 3486 - - 4198 1989 0 200 3686 - - 3998 1990 200 200 4086 - - 3598 1991 200 0 4286 - - 3398 1992 0 200 4486 - - 3198 1993 200 200 4886 - - 2798 l 1994 200 0 5086 - - 2598 1995 0 200 5286 2398 ~ 1996 200 200 5686 - - 1998 1997 200 0 5886 - - 1798 1998 0 200 6086 - - 1598 , 1999 200 200 6486 - - 1198 i 2000 200 0 6606 - - 998 2001 0 200 6886 - - 798 2002 200 200 7286 - - 398 2 2003 200 0 7486 - - 198 y 2004 0 200 7686 - - 0 j 2005 200 200 8086
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TThe number of locations available atter the completion of that year's scheduled refueling outage.
racks of the present design, full core discharge and refueling dis-charge capabilities would be lost after the 1984 and 1986 refueling outages, respectively. No further expansion of the Quad Cities spent fuel storage capacity is possible using the presently approved spent 1 fuel storage rack design. In contrast, high-density spent fuel storage racks have a capacity of 7684 fuel assemblies. Therefore, full core discharge is possible until the refueling outage of 2002 is completed. Refuel discharge capability would be lost after the refueling outage of the year 2003, l l Commonwealth Edison Company, in its function as operator, pro-poses to increase the spent fuel storage capacity by replacing the present spent fuel storage racks with now, high-density storage racks. This modification will include the use of a neutron absorber material in the racks, at an increase of k fr m 0.90 to 0.95. The March 26, eff 1981, letter to the NRC requests a modification Quad Cities 1 Technical Specification 5.5B, " Fuel Storage," to .:plement this change in k eff' ! The specification for design, construction, and quality assurance of the high-density racks was prepared by Quadrex, a San Jose-based company. The mechanical design, seismic analysis, thermohydraulic l1 analysir,4 &nd other related calculations as well as the fabrication of the hardware will be performed by Joseph Oat Corporation. Joseph Oat j Corporanion, based in Camden, N.J., possesses ASME Code stamps for Section III, Classes 1, 2, and 3, and MC pressure vessels and co.nponent s . Southern Science Applications, Inc., of Dunedin, Florida, is serving as a consultant to Joseph Oat Corporation in the areas of criticality analysis and other radionuclide evaluations. Consulting support on the overall e' fort is provided by NUS l Corporation of Rockville, Maryland. A V l 1-3
- 2. GENERAL ARRANGEMENT O The high-density spent fuel racks consist of individual cells with a 6-inch-square cross section, each of which accommodates a single BWR fuel assembly. The cell walls consist of a neutron absorber sandwiched between sheets of stainless steel. The cells are arranged in modules of varying numbers of cells with a 6.22-inch center-to-center spacing.
The high-density racks are engineered to achieve the dual objec-tive of maximum protection against structural loadings (such as ground motion) and the maximization of evailable storage locations. In general, a greater width to height aspect ratio provides greater margin against rigid body tipping. Hence, the modules are made as wide as possible within the constraints of transportation and site-handling capabilities. The high-density spent fuel racks will be installed in the Unit 1 and Unit 2 spent fuel pools, each of which is 33 feet t'ide by 44 feet long. The Quad Cities Unit 1 pool will contain 19 high-density fuel racks in 7 different module sizes. The module types are labelled A through G in Figure 2.1, which also shows their relative placement. There will be a total of 3714 storage locations in the Quad Cities Unit 1 pool. The Quad Cities Unit 2 pool will contain 20 high-density ruel racks in 6 different module sizes. The module types are labelled A through F in Figure 2.2, which also s'iows their relative placement. There will be a total of 3970 storage locations in the Quad Cities Unit 2 pool. Table 2.1 gives the detailed module data (e.g. , weight, quantity, and number of storage locations). The spent fuel tack modules are not anchored to the pool floor or connected to the pool walls. The minimum gap between any two spent O fue1 recx mod =1ee wi11 ee 2.0 1 caee et 11 1ocetio#e. rae mi=1 mum eee between the fuel pool wall and spent fuel rack modules will be 1 I 2-1 l l
-l j o Table 2.1 Module Data Approximate Number of Weight T_ype Quantity Cells / Module Array Size Lbs/ Module , e A 12 210 14 x 15 27,100 B 8 195 14 x 14 25,300 1 C 8 182 14 x 13 23,500 D 4 135 9 x 15 17,500 E 4 224 14 x 16 29,000 F 2 256 16 x 16 33,100 G 1 192 12 x 16 24,800 O l t O l 2-2
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9 inches. The minimum gap between the spent fuel rack modules and pool wall fixtures will be 3 inches, except near the contaminated 1 equipment storage pit where the gap is 2-3/8 inches. Adequate clearance from other existing pool hardware will also be provided. Due to the gaps provided, the possibility of interrack impact, or rack collision with pool walls or other pool hardware during the postulated ground motion events will be precluded. i O
; O 2-3
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- Below Water JI (14, g3) 100%" Sq.
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=, Quad Cities Unit 2 (3970 Cells)
- 3. RACK CONSTRUCTION AND ASSEMBLY O 3.1 Construction The racks will be constructed frco ASTM 240-304, austenitic 1 stainless steel sheet material, ASTM 204-304 austenitic stainless steel plate material, and ASTM 182-F304 austenitic steel forging material. Boraflex,* a patented brand name product of BISCO,** will serve as the neutron absorber material.
A typical module will contain storage cells which have a 6-inch minimum internal cross-sectional opening. This dimension ensures that y fuel assemblies with maximum expected channel deformations can be inserted into, and removed from, the storage cells without any damage to the fuel assemblies or storage racks. Figure 3.1 shows a horizontal cross-section of an array of 3 x 3 cells. The cells provide a smooth and continuous surface for lateral contact with the fuel assembly. The construction of the rack modules ( may best be described by exposing the basic building blocks of this design, namely the " cruciform," " ell," and " tee" elements, shown in Figure 3.2. The cruciform element is made of 4 angular sub-elements, "A" (Figure 3.3) with the neutron absorber material tightly sandwiched between the stainless sheets. The cruciform assembly has 4-inch-high l1 stainless strips at the bottom, which ensure against slippage of the
" poison" material downwards due to gravitation loads or operating con-ditions. The fabrication procedure leads to 100% surface contact (in macroscopic sense) between the poison and the stainless sheets. The top of the cruciform is also end welded using a spacer strip as shown in Figure 3.3. Skip welding at the top ensures proper venting of the 11 sandwiched space in the cruciform spokes.
- Boraflex was approved by the NRC for use in spent fuel storage racks at Point Beach Units 1 and 2 by NRC letter dated April 4, 1979 (Docket Nos. 50-266 and 50-301) and for use at Oconee Units 1, 2, and 3 on 1
(~; January 24, 1981 (Docket Nos. 50-269, 50-270, and 50-287) . G ** Brand Industrial Services, Inc., is a subsidiary of Brand Insulations, Inc., 1420 Renaissance Drive, Park Ridge, Illinois 3-1
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The " ell" and " tee" elements are constructed similarly using angular sub-element "B," and flat sub-element "C". Having fabricated the required quantities of the " cruciform," " tees," and " ells," the assembly is performed in a specially designed fixture which serves the function of maintaining dimensional accuracy while welding all the contiguous spokes of all elements using fillet welds. Figure 3.4 l1 shows the fillet welds in a 4 x 4 array. In this manner, the cells are produced which are bonded to each other along their long edges, thus, in effect, forming an " egg-crate". The bottom ends of the cell walls are welded to the base plate. Machined sleeve elements are positioned in the holes drilled in the base plate concentric with the cell center lines, and attached to the base plate through circular fillet welds (Figure 3. 5) . The conical machined surface on the sleeve provides a contoured seating surface for the " nose" of the fuel assembly. Thus, the contact stresses at the fuel assembly nose bearing surface are minimized. The central hole in the sleeve provides the coolant flow path for heat transport from the fuel assembly cladding. Lateral holes in the cell walls (Figure 3.5) provide the additional flow path in the 11 unlikely event that the main coolant flow path is clogged. The rack assembly is typically supported on four plate-type sup-ports. The supports elevate the module base plate 6.5 inches above the pool floor level, thus creating the water plenum for coolant flow. Figure 3.6 shows vertical and horizontal cross sections of a typical support leg. The box-shaped support structure is treated as a " linear , type" support for stress analysis purposes. The welds are . sized to y produce large margins of safety consistent with the rest of the sup-port region. Four lateral holes in the support foot plates provide flow paths for coolant flow to the storage locations (see Figure 3.6). -) 3-2
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3.2 Codes, Standards, and Practices for the Spent Fuel Pool () Modification The following are the public domain codes, standards, and practices to which the fuel storage racks are designed, constructed, and. assembled, and/or pool structure analyzed. Additional problem-specific references related to detailed analyses are given at the end of each section. I. Design Codes (a) AISC Manual of Steel Construction, 7th Edition, 1970, including supplements 1, 2, and 3 to the AISC Specification. (b) ANSI N210-1976 Design Objectives for Light Water Reactor Spent Fuel Storale Facilities at Nuclear Power Stations. 1 (c) ACI 318-77 Building Code Requirements for Reinfo'ced r Concrete. II. Material Codes (a) American Society for Testing and Materials (ASTM) Standards (b) American Society of Mechanical Engineers (ASME) , Section III, Div. 1, Subsection NF (1980). III. Welding Codes (a) ASME Boiler and Press re Vessel Code, Section IX-1980 Welding and Brazing Qualifications. IV. Quality Assurance, Cleanliness, Packaging, Shippino, Receiving, Storage, and Handling Recuirements The quality assurance program for the design and () installation of the new spent fuel storage racks will be the 3-3 l t J
._ = . _. . ._, _ _ ..
applicable portions of Revision 15 of. topical report CE-1-A, commonwealth Edison Company Quality Assurance Program for Nuclear Generating Stations. Revision 15 of this report, dated January 2, 1981, was approved by the NRC in Fabruary 1981. V. Other References (a) NRC Regulatory Guides, Division 1, Regulatory Guides 1.13, 1.29, 1.71, 1.85, 1.92, and 1.124 (Revisions effective as of April 1980). (b) General Design Criteria for Nuclear Power Plants, Code of Federal Regulations, Title 10, Part 50, Appendix A (GDC Nos. 1 1, 2, 61, 62, and 63). . (c) NRC Standard Review Plan, Sections 3.8.3 and 3.8.4. , (d) NRC Standard Review Plan, Section 9.1.2 (as applicable to spent fuel racks). (e) "NRC Position for Review and Acceptance of Spent l'uel l Storage and Handling Applications," dated April 14, 1978, and the modifications to this document of January 18, 1979. !O 3.3 Installation and Leveling i To be supplied. i O 3-4
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- 4. NUCLEAR CRITICALITY ANALYSIS O This section provides information concerning the spent fuel rack design features that ensure that ample margin to criticality will be maintained. This material concludes that under the most adverse 1 effects of manufacturing tolerances, considering normal and accident conditions, the most reactive fuel will not achieve a keff greater than 0.95 with a 95% probability at a 95% confidence level.
4.1 Design Basis The spent fuel storage racks are designed to ensure that a k gg e equal to or less than 0.95 is maintaincd with the racks fully loaded with fuel of the highest anticipated reactivity, flooded with unborated water, at a temperature corresponding to the highest reactivity, under normal or abnormal conditions. The maximum calcu- l1 lated reactivity includes a margin for uncertainty in reactivity calculations and in mechanical tolerances, statistically combined, l such that the true keff will be equal to or less than 0.95 with a 95% N- probability at a 95% confidence level. Applicable codes, standards, and regulations or pertinent sections thereof include the following: l e General Design Criterion 62 - Prevention of Criticality in l Fuel Storage and Handling. o NRC letter of April 14, 1978, to all Power Reactor Licensees - OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, including modification letter dated January 18, 1979. e NRC Standard Review Plan, Sections 3.8.4 and 9.1.2, as they apply to spent fuel racks. e Regulatory Guide 3.41, Validation of Calculational Method for Nuclear Criticality Safety (and related ANSI N16.9-1975.) e ANSI N210-1976, Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Plants. () e ANSI N18.2-1973, Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants. 4-1
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The design basis fuel assembly is an 8 x 8 array of fuel rods (BWR type) containing UO2 at a maximum nodal enrichment of 3.2% U-235 by
\- weight. In the analysis, all axial nodes were assumed to be uniformly 1 enriched to this value. Fuel assemblies containing gadolinium burn-able poison or assemblies of other configurations or enrichments, e.g., 7 x 7 array, may also be safely accommodated in the spent fuel storage racks provided the maximum reactivity is less than or equal to the reactivity of the design basis fuel assembly.
To ensure the true reactivity will always be less than the calcu-lated reactivity, the following conservative assumptions were made: e Moderator is considered to be pure, demineralized, unborated y water at a temperature corresponding to the highest reactivity. e Lattice of storage racks is considered to be infinite in all l 1 directions; i.e., no credit is taken for axial or radial neutron leakage. f") (s j e Neutron absorption in minor structural members is neglected; l 1.e., spacers and Inconel springs are replaced by water. e Pure zirconium is considered to be used for cladding and 1 flow channel; i.e., higher neutron absorption of alloying materials in Zircaloy is neglected. e Each fuel assembly is at maximum enrichment and reactivity. l1 4.2 Geometric and Calculational Models 4.2.1 Reference Fuel Assembly i i The ref erence design basis fuel assembly, illustrated in Figure 4.1, is an 8 x 8 array of fuel rods with two of the central rods replaced by Zircaloy " water-rods." The square Zircaloy channel sur-1 rounding the fuel has walls 0.080 inches thick and has a nominal out-
^
side dimension of 5.438 inches. A maximum nodal enrichment of 3.2% 4-2 1
/ I
U-235 by weight was assurced to exist uniformly over the entire fuel O V assembly length for the design basis fuel. In actuality, this fuel' assembly has its top and bottom 6-inch nodes composed of natural uranium while its center zone nodes are enriched to 3.2% U-235 by weight. Thus, the design basis fuel assembly conservatively approxi-mates the reactivity of the actual fuel assembly. 1 Because the curved surface, corresponding to the maximum expected distortion of the Zr channel cannot be adequately described in the two-dimensional computer codes used for analysis, an approximation, preserving the Zr thickness and weight, was necessary. This should represent a reasonable approximation, since the reactivity effect due to distortion of the Zr channel is small (see Section 4.3.9). 4.2.2 Alternative Fuel Assembly Designs The spent fuel storag<e racks are also intended to accommodate fuel assemblies consisting of a 7 x 7 array and an 8 x 8 array (1 water
- rod), both containing fuel with a maximum nodal enrichment less than 3.2% U-235. Specifications for these fuel assemblies and for the reference fuel assembly are listed in Table 4.1, which permits compar-ir ons of the designs. The reactivity of the alternative fuel assemblies is lower than the reactivity of the design basis fuel assembly, primarily because of the lower enrichments in the alterna-tive assemblies. Consequently, the design basis fuel assembly is the assembly of highest anticipated reactivity and is the limiting case.
In the present analysis, gadolinium burnable poison is not included in the fuel. However, the spent fuel storage racks can safely accommodate gadolinium-bearing fuel of higher U-235 enrichment than that specified for the design basis, provided the reactivity of the fuel assembly is less than or equal to that of the reference design basis fuel assembly. For comparison purposes, the calculated reactiv-ity (by AMPX-KENO, see Section 4.2.3 below) of the design basis fuel assembly on a 6.00-inch lattice spacing is 1.362 1 0.004 (lo) with unborated water in the standard reactor core geometry. O 4-3
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Table 4.1 Fuel Assembly Design Specifications
- O Fuel Assembly Designation 8 x 8R (Reference) 7 x 7/7 x 7R 8x8 Fuel Rod Data Outside diameter, in. 0.483 0.563 0.493 C1sdding thickness, in. 0.032 0.032/0.037 0.034 Cladding material Zr-2 Zr-2 Zr-2 Pellet density, gm UO 2 /cc 10.41 10.19/10.41 10.41 l1 Pellet diameter, in. 0.410 0.488/0.477 0.416 Muximum nodal enrichment, wt% U-235 3.2* 2.12/2.30 2.62 l1 Water Rod Data Outside diameter, in. 0.591 -
0.493 Wall thickness 0.030 - 0.034 Material Zr-2 - Zr-2 () Number per assembly 2 none 1 Fuel Assembly Data Number of fuel rods 62 49 63 Fuel rod pitch, in. 0.640 0.738 0.640 Fuel channel outside dimension, in. 5.438 5.438 5.438 ( Fuel channel wall thickness, in. 0.080 0.080 0.080 Fuel channel material Zr-4 Zr-4 Zr-4
- Actual fuel assemblies have 6 inches of natural uranium at both ends 1 of fuel rod.
4-4
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4.2.3 Calculational Models O)
~#
~4.2.3.1 Analytical Methods Nuclear criticality analyses of the high-density spent fuel storage rack were performed with the AMPX1-KENO 2 computer package, using the 123-group XSDRN cross-section set and the NITAWL sub-routine for U-238 resonance shielding effects (Nordheim integral treatment) .
l AMPX-KENO has been extensively benchmarked against a number of critical experiments.3,4,5 For the investigation of small reactivity effects (e.g., mechanical tolerances), a four-group diffusion /blackaess theory method of analysis (NULIF-CNROD-PDQ7)$ was used to calculate small incremental reactivity changes. This model has been used previously with good results and is normally used only to evaluate trends and small incremental reactivity effects that would otherwise be lost in the KENO statistical variation. Where possible, trends calculated by AMPX- KENO and by dif fusion / blackness theory were compared and found to be in good agreement, well within the statistical uncertainty of KENO calculations. 4.2.3.2 Calculational Bias and Uncertainty Results of benchmark calculations 5 on a series of critical exper-iments indicate a calculational bias of 0, with an uncertainty of 10.0028 (95% probabili?.; at a 95% confidence level). In addition, a il small correction in the calculational bias is necessary to account for j the sltghtly larger gap thickness (1.1 inches) between fuel assemblies in the Quad Cities spent fuel rack compared to the corresponding thickness (0.644 inch) in the benchmark critical experiments. Based upon the correlation developed in Reference 5, the correction for water-gap thickness in the Quad Cities spent fuel storage rack is an underprediction of -0.0036 ak. Thus, the net calculational bias is 0.0036 1 0.0028, including the effect of the water-gap thickness. l1 O 4-5
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( 4.2.3.3 Trend Analysis 5 Trend analysis of benchmark calculations on critical experiments with varying boron content in the absorber plate between fuel assem-blies indicates a tendency to overpredict keff with higher reactivity worth of the boron absorber. In the Quad Cities spent fuel rack, the boron worth is about 40% ak, or. approximately 2.5 times the highest boron worth (15.9% ak) in the critical experiments analyzed in Reference 5. Based upon the extrapolation of the trend analysis, AMPX-KENO calculations of the Quad Cities rack would be expected to overpredict k, by an estimated 3.1% ak, including allowance for water-gap thickness. Statistically combining the standard deviation of the regression analysis 5 ( 0.0027, 1a) and a typical standard deviation of the KENO variation of the mean ( 0.005, le ) , the maximum uncertainty would be 10.0116, including a one-sided tolerance factor 6 of 2.03 (95% probability at a 95% confidence level) for an assumed 60 generations in a KEl;0 calculation. Thus, to the extent extrapolation of the linear regression analysis is valid, the AMPX-KENO calculation Quad Cities rack will be high (overprediction) by 0.031 i lc of the 0.012 Ak, or a minimum overprediction of 0.019 ak, including calculational uncertainty. Although extrapolation of the regression trend much beyond the range of the measurements may be questionable, ! the analysis does indicate that AMPX-KENO calculetions would be expected to overpredict kegg when strong boron absorbers are present. No credit is taken for the expected overprediction other than to indicate an additional level of conservatism in the criticality analysis of the Quad Cities spent fuel storage rack. 4.2.4 Reference Fuel Storage Cell The nominal spent fuel storage cell model used in the criticality analyses is shown in Figure 4.1. The rack is composed of Boraflex absorber material sandwiched between two 0.075-inch stainless steel plates. The fuel assemblies are centrally located in each storage cell on a nominal lattice spacing of 6.22 inches. For two-dimensional X-Y analysis, a zero current (reflecting) boundary condition was O 99 tied to the exie1 directio- e=d et eae oe=ter-11=e tarecea tae Boraflex absorber on all four sides of the cell, effectively creating 4-6
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an infinite array of storage cells. The Boraflex absorber has a nomi-nal thickness of 0.070 inch and a nominal B-10 areal density of 0.01728 gram B-10 per cm . 4.3 Reference Subcriticality and Mechanical Tolerance Variations 4.3.1 Nominal Case (8 x 8R Fuel Assembly of 3.2 wt% U-235, Maximum y Nodal Enrichment) Under normal conditions, with nominal dimensions, the calculated k, is 0.9155 i 0.0036 (la with 140 generations) . For a one-sided tol-erance factor of 1.879, corresponding to 95% probability at a 95% con-fidence limit with 140 generations, the maximum deviation of k,is 10.0067. 4.3.2 Maximum Enrichment Capability For the nominal nodal enrichment of 3.2 wt% U-235 and the g selected Boraflex loading (0.31728 g B-10/cm2 nominal), there is a U margin between the limiting reactivity (k. of 0.95) and the calculated reactivity including all uncertainties. Additional calculations were performed to estimate the maximum uniform enrichment which the racks could safely accommodate without exceeding the limiting reactivity. Results of these calculations indicate a limiting nodal enrichment of 3.40 wt% U-235. For this enrichment, the estimated k,is 0.9305 +, 1 0.0036 (la). With a uniform enrichment of 3.40 wt% U-235, the maximum k ,would not exceed 0.95, including all uncertainties, with 95% probability at a 95% confidence level (see Section 4.5.2 below) . i 4.3.3 Alternative Fuel Assemblies i 4.3.3.1 Alternative Geometry and Enrichments The alternative 8 x 8 fuel assembly of 2.62 wt% U-235 maximum nodal enrichment will have an appreciably lower reactivity than the l reference 3.2% enriched assembly, because of the lower enrichment.
- O rer eae 7 x 7 eeeem817 et e co==ervetive1r eeeumed eede1 eerioemeat oe 11 2.8 wt% U-235, AMPX-KENO calculations with nominal dimensions yielded a Pegg of 0.890 1 0.005, which is substantially less than that of the 4-7
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reference 8 x 8 fuel assembly. For the enrichments indicated in Table 4.1, the reactivity would be even lower. Thus, the reference 8 x8 assembly, with 3.2% U-235 maximum nodal enrichment, is the limiting case. 4.3.3.2 Plutonium-Bearing Experimental Fuel Assemblies Five special experimental fuel assemblies (7 x 7 array) have been fabricated and irradiated. In four of these experimental fuel assemb-lies, 10 fuel rods contain plutonium as th: oxide, mixed with natural uranium oxide; the fifth experimental fuel assembly includes only 8 PuO 2-bearing fuel rods. In both cases, the remaining fuel rods in the assembly are UO2 with enrichments ranging from 1.33 wt% to 3.33 wt% U-235. Two-dimensional calculations (PDQ) of the k, were made with diffusion / blackness theory, explicitly describing each fuel rod cell i and associated enrichment for the assembly containing 10 Pu-bearing rods. For conservatism, the Gd 023 that is contained in five of the UO2 rods was not included. Similarly, for the Pu-bearing rods, the Pu-240 g and Pu-242 content was neglected and the assumption made that all V fissile Pu was Pu-239. 1 The calculated k, for the Pu-bearing assembly (10 Pu rods) in the storage rack is 0.885 (including a transport theory correction) compared to the nominal design k, of 0.9155 1 0.0036 (Section 4.3.1 above). In the typical BWR core arrangement, the calculated k,is 1.331 compared to the nominal case k, of 1.362. Thus, the four 10-Pu-rod experimental assemblies have a k, substantially lower than the nominal design case and, therefore, may safely be stored in the spent l fuel rack. The fifth experimental fuel assembly contains only eight Pu-bearing rods and the average U-235 enrichment is lower. Conse-quently, it will have an even lower reactivity and can also be stored safely in the spent fuel rack. Omitting the neutron absorption in Gd, Pu-240, and Pu-242 in the calculation should yield conservative l results. In addition, these fuel assemblies have already accumulated significant burnup histories and the actual reactivities should be appreciably lower than that calculated for the initial loading without O ca eer=e81e voiso=- 4-8
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4.3.4 Boron Loading Variation 11 The Boraflex absorber plate is nominally 0.070 inch thick with a B-10 areal density of 0.01728 g/cm2 Manufacturing tolerance limits are 10% in both thickness and boron content. This ensures that, at any point where the minimum boron loading (0.01555 g B-10/cm2) and minimum Boraflex thickncss (0.063 inch) may coincide, the boron areal density will not be less than 0.014 g B-10/cm2 i Calculations were made of k, with variations in Boraflex absorber loading and thickness. Results of these cilculations, given in Figure 4.2, indicate that the k, can be described by the following regression l1 fit (least squares) to the date over the range of B-10 loading from 0.010 to approximately 0.020 g/cm2,
-In k, = 0.06494 in (B-10, g/cm2) + 0.3519 Within the precision of the calculations, this relationship indi-cates that the 10% tolerance limit on either boron content or Boraflex O eatex=e re u1e= i= eue e me i= creme e 1 reeceivier cae=ee of o oo83 ak. The trend calculated both by AMPX-KENO and by dif fusion / blackness theory is the same within the analytical uncertainty.
4.3.5 Storage Cell Lattice Pitch Variation i 1 1
- The design storage cell lattice center-to-center spacing between fuel assemblies is 6.220 inches. For manufacturing tolerances of
+0.125 or -0.000 inch, increasing the lattice pitch from the minimum 6.220 inches to 6.345 inches (maximum tolerance) reduces reactivity by 0.0113 1 0.006 k , as calculated by AMPX-KENO or by 0.0094 ok calcu-lated by diffusion / blackness theory. Thus, the nominal case exhibits the largest keff and the effect on reactivity of a lattice pitch increase is negative. A larger increase in lattice pitch produces an even larger negative effect. Results of calculations at several lat-tice spacings and boron loadings are shown in Figure 4.3 in terms of l1 l
the overall fuel region volume fraction in the spent fuel storage cell (0.6775 for the nominal design). l 4-9
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4.3.6 Stainless Steel Thickness variations I1 O b The nominal stainless steel thickness is 0.075 inch. The reac-tivity effect of the expected stainless steel thickness tolerance 1 variation (10.004 inch) was calculated to be 10.0009 ok by the diffusion / blackness theory, since the reactivity increment is too small to be calculated by AMPX-KENO. 4.3.7 Fuel Enrichment and Density Variation 1 The design basis maximum nodal enrichment is 3.2% U-235 by weight. Calculations of the sensitivity to small enrichment varia-tions by diffusion / blackness theory yielded an average coefficient of 0.0075 ak per 0.1 wt% U-235. For an estimated tolerance on U-235 i enrichment of 10.015%, the uncertainty is 10.0011 Ak. Calculations raade with the UO2 fuel density reduced from the nominal design value of 10.41 g/cm 3 to 10.25 g/cm 3 indicate that the l1 . storage rack k ,is reduced by 0.0022 ak (dif fusion / blackness theory) . For an assumed tolerance of 10.05 g/cm3 in UO2 density (10.46 g/cm3 maximum), the uncertainty in kegg becomes 10.0007. A lower fuel density would result in lower values of reactivity. 1 1 4.3.8 Boraflex Width Tolerance Variation i l The calculational model (Figure 4.1) uses a Boraflex blade width of 5.86 inches for each cell wall. This is coTservative since, in the i1 final design of the storage cell, the minimum Boraflex absorber width is nominally 5.91 inches, including tolerances. The calculational model thus results in the highest reactivity (0.9155 1 0.0036), and the greater width of the actual absorber would further decrease reac-tivity. 4.3.9 Effect of Zirconium Fuel Channel l1 Elimination of the zirconium fuel channel results in a small de-O cre theory. e 1" re ctivier (-o oo35 2x) ce1c=1 tea er aire" i =/81 cx"e== More significant is a small positive reactivity effect 4-10
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resulting from bulging of the zirconium channel, which moves the chan-nel wall outward toward the Boraflex absorber. For the maximum {"} analyzed bulging 5.925 inches outside dimension) uniformly (to throughout the assembly, an incremental reactivity of +0.0039 ok would result. This was calculated by diffusion / blackness theory using the approximate geometric model for the flow channel indicated by the dotted lines in Figure 4.1. Actual channel bulging will be much less than the magnitude analyzed. Because actual bulging of the flow channel would not be the maximum everywhere in all assemblies, the reactivity effect can be statistically combined with the reactivity effect of other mechanical ~ deviations. Fuel assembly bowing yields a negative reactivity effect and is treated under abnormal conditions (Section 4.4 below) . 4.3.10 Summary of Statistical variations l1 Calculated reactivity increments from mechanical and fabrication tolerances are summarized in Table 4.2. 4.4 Abnormal and Accident Conditions l1 Although credit is permitted for absorption by other absorbers under abnormal conditions, the following evaluaticns were made without any additional absorber material in the spent fuel storage pool. To the extent any additional absorbers may be present in the realistic case, the following analyses are even more conservative. 4.4.1 Fuel Assembly Positioning in Storage Rack The fuel assembly is normally located in the center of the stor-age rack call with bottom fittings that mechanically prevent lateral movement of the fuel assemblies. Nevertheless, calculations were made with adjacent fuel assemblies (each assumed to be located on one side of its cell with the zirconium fuel channel touching the SS-Boraflex plate) creating an infinite series of two-assembly clusters separated (} only by the SS-Boraflex plate. For this case, the calculated reactiv-ity was slightly less than the nominal design case (by 0.0011 Ak) . 4-11
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Table 4.2 Calculated Statistical variations in Reactivity (Mechanical) fl J Case Telerance Incremental Reactivity, ak_ Boron concentration 110% 10.0063 Boraflex thickness 110% 10.0063 5 Lattice pitch [0.,0 inch Zero to negative SS tolerance 10.004 inch 10.0009 Channel bulge 0.49 inch max. +0.0039 Fuel enrichment +0.015% U-235 +0.0011 1 Fuel density 0.05 g/cm3 f0.0007 Boraflex width
- Zero to negative Statistical average 10.0098 (Root-mean-square of positive increment)
,() *Boraflex width used is conservatively less than the minimum width expected, including tolerances. O l 4-12
/
Calculations were also made with the fuel assembly mov,ed into the p corner of the storage rack cell (four-assembly cluster at closest approach), resulting in an even lar uer negative reactivity effect (calculated decrease in k, of approximately 0.01. With the zirconium fuel channel removed, the reactivity effect cv off-set fuel assemblies is even more negative. Thus, the nominal cave, with the fuel assembly positioned in the center of the storage rack cell, yields the maximum reactivity. 4.4.2 Effect of Zirconium Fuel Channel Distortion Consequences of bulging of the zirconium fuel channel have been treated as a statistical deviation in Section 4.3.9 above. Bowing of y the zirconium channel results in a negative reactivity effect analagous to that of positioning the fuel assembly toward one side of the storage cell as described in Section 4.4.1 above. Thus, bowing will result in a reduction in reactivity. 4.4.3 Temperature and Water Density Effects Decreasing temperature from the nominal 680F to 390F (maximum water density) is calculated to increase reactivity by 0.0007 Ak, as indicated in Table 4.3 (reactivity ef fects calculated by diffusion / blackness theory). Increasing the water temperature or introducing voids (to simulate boiling) decreases reactivity, as s,hown in the table. Table 4.3 Ef f ect of Temperature and Void on Calculated Reactivity of Storage Rack l Case ak. Comment l 390F +0.0007 Maximum water density 680F 0 Reference 1040F -0.004 p (H2 O) = 0.992 1760F -0.013 p (H2 O) = 0.972 2120F -0.020 p(H 2 O) = 0.958 2120F with 50% void -0.175 Simulates boiling 4-13
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4.c.4 Abnormal Positioning of Fuel Assembly Outside Storage Rack O
~
Since the storage rack criticality calculations were made assuming an infinite array of storage cells, positioning a fuel assembly outside and adjacent to the actual rack has been effectively analyzed; such positioning would result in a reactivity that is lower than the k, calculated for the infinite array. This has been l1 confirmed by two-dimensional PDQ analysic of finite racks with a new fuel element positioned outside and adjacent to the rack. 4.4.5 Missing Absorber Plate Should a Boraflex absorber plate be missing from between fuel assemblies, the reactivity will be slightly higher than the reference case. Calculations performed in two dimensions (PDQ7) indicate the largest reactivity increment is less than +0.0031 ak due to the loss of a single plate. Because of mesh size limitations in PDQ7, symmetry considerations (with reflective boundary conditions) effectively resulted in the loss of an absorber plate from one side of every 15 , storage cells. Thus, the calculated incremental reactivity addition due to the loss of an absorber plate should be conservative. l 4.4.6 Dropped Fuel Assembly Accident A postulated fuel assembly drop into and to the bottom of a stor-age rack cell results in a. configuration that is that of the nominal cell configuration; therefore, a drop will not result in a reactivity greater than that of the nominal design case. Three additional dropped fuel assembly configurations were considered: e The dropped fuel assembly enters a storage cell already occupied by a fuel assembly (straight drop). 1 e The dropped fuel assembly leans against the pool wall or ! other structure (inclined drop). () e The dropped assembly does not enter horizontally on top of the spent fuel racks. a cell but lies 4-14
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The above configurations result in the separation of fuel 1 assembly only by water. Figure 4.4 shows the results of calculations for unpoisoned fuel assec lies separated only by water. From these data, the reactivity (k.) will be less than 0.95 for any spacing greater than 8 inches. For a straight drcp on top of the l1 rack, an inclined drop, or a fuel assembly li.ng horizontally on the top of the rack, the minimum separation distance is approximately 9 inches. Maximum expected deformation under seismic or accident conditions (see Sections 6 and 7) will not reduce the minimum spacing to less than 8 inches. In addition, the upper 6 inches of fuel is natural uranium oxide, which affords a further effective separation from the higher-enriched active fuel in the storage racks. Finally, a three-dimensional PDQ analysis, with a new fuel element immediately above the active fuel in the storaga rack (neglecting structural material) confirms that the reactivity is less than that of the design basis infinite array. 1 Fuel assembly drop accidents will not result in an Increase in reactivity above that calculated for the infinite nominal design stor-age rack. 4.4.7 Fuel Rack Lateral Movement f Normally, the individual racks in the spent fuel pool are i separated by a water gap of 3 inches. For finite fuel racks, this l1 l separation would reduce the actual maximum reactivity of the racks. Should lateral motion of a fuel rack occur, for whatever reason, I closing the gap between racks, the reactivity would, in the limit, only approach the limiting reactivity of the reference infinite array. 4.5 Summarv l l The criticality analyses of the spent fuel storage rack under normal and abnormal conditions are summari=ed in Table 4.4. I O l 4-15
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_,, ,<-m, . . _ _ _ _ _ _ . . - - , . - _ _ _ . . . . _ . . . . - , . , , . , , , , . . _ _ . _ _ - . _ ._ _ _ . . - - . -
Table 4.4 Summary of Criticality Calculations O Case k. or Ak, Comment Netmal Conditicns k , reference 0.9155 Section 4.3.1, includes gap correction Calculations 1 bias +0.0036 Section 4.2.3.2 Uncertainties Bias 10.0028 Section 4.2.3.2 Calculational 10.0067 Section 4.3.1 1 Mechanical. 10.0098 Section 4.3.10, Table 4.2 10.0122 Statistical combination Total 0.9191 + 0.0122 O Meximum x. 0.931 l Abnormal and Accident Conditions Decreased temperature +0.0007 Maximum water density Increased temperature or void negative Fuel element positioning negative Fuel channel bowing negative Lost / missing absorber plate +0.0031 Conservative Fuel handling accident negligible Lateral rack movement negligible 'O 4-16 l
/
,,_ --_.~ _ _. ,~ , ,__.._. ._ _ , , _ , m. _ , , _ _ , , _ . , , , , . . , , _ _ . . _ . , , _ _ .
----e , . - - _ , - . . _ . _ , , . . . . .__m_ -
Thus, a k, of 0.9313 is conservatively estimated to be the maximum k, under the worst combination of calculational and mechanical uncertainties (normal conditions), with a 95% probability at a 95% confidence level. Under the worst combination of abnormal and accident conditions, the maximum k ,could be no more than 0.935. l1 Removal of the zirconium fuel channel from all assemblies would reduce the maximum k,to 0.927 (norinal conditions). If the trend 1 toward overprediction with baron worth (Section 4.2.3.3) is valid, the maximum expected k, under normal conditions would be 0.905. O 'O 4-17
/
REFERENCES TO SECTION 4 O 1. Green, Lucious, Petrie, Ford, White, Wright, PSR-63/AMPX-1 (code package) , "AMPX Modular Code System for Generating Coupled Multi-group Neutron-Gamma Libraries from ENDF/B," ORNL-TM-3706, Oak Ridge National Laboratory, March 1976.
- 2. L. M. Petrie and N. F. Cross, KENO-IV, "An Improved Monte Carlo Criticality Program," ORNL-4938, Oak Ridge National Laboratory, November 1975.
- 3. S. R. Bierman et al., " Critical Separation Between Subcritical Clusters of 4.29 wt% U-235 Enriched UO2 Rods in Water with Fixed Neutron Poisons," NUREG/CR-0073, Battelle Pacific Northwest Laboratories, May 1978, with errata sheet issued by the NRC August 14, 1979.
- 4. M. N. Baldwin et al., " Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel," BAW-1484-7, Babcock & Wilcox Company, July 1979.
- 5. S. E. Turner and M. K. Gurley, " Benchmark Calculations for Spent Fuel Storage Racks," Report SSA-127, Revision 4, Southern Science Applications, Inc., April 1981.
I 6. M.-G. Natrella, " Experimental Statistics," National Bureau of Standards, Handbook 91, August 1963. l l r O 4-18
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- 5. THERMAL-HYDRAULIC CONSIDERATIONS l1 (S)
A central objective in the desian of the high-density fuel rack is to ensure adequate cooling of the fuel assembly cladding. In the following, a brief synopsis of the design basis: the method of analysis, and computed results is given. 5.1 Heat Generation Calculations To be supplied. 5.2 Analysis of Pool Thermal Hydraulics In order to determine an upper bound on the maximum fuel cladding temperature, a series of conservative assumptions are made. The most important assumptions are listed below:
- a. As stated above, the fuel pool will contain spent fuel with varying " time-after-shutdov " (ts). Since the heat emission
( falls off rapidly with increasing ts, it is obviously con-servative to assume that all fuel assemblies are fresh (ts " 100 hours), and they all have had 4 years of operating time in the reactor.1 The beat emission rate of each fuel assembly is assumed to be equal.2
- b. As shown in Figures 2.1 and 2.2, the modules occupy an irregular floor space in the pool. For purposes of the hydrothermal analysis, a circle circumscribing the actual rack floor space is drawn. It is further assumed that the cylinder with this circle as its base is packed with fuel assemblies at the nominal pitch of 6.22 inches (see Figur6 5.1) .
- c. The downcomer space around the rack module group varies, as shown in Figure 5.1. The minimum dcwncomer gap (9 inches) l1 available in the pool is assumed to be the total gap avail-able around idealized cylindrical rack; thus,
(]) the the 5-1
- - - . .. . ~. - -. - , - . . . -- _ _ . . . - - . __
m maximum resistance to downward flow is incorporated into the ( analysis,
- d. No downcomer flow is assumed to exist between the rack modules.
In this manner, a conservative idealized model for the rack assemblage is devised. The water flow is axisymmetric about the vertical axis of the circular rack assemblage, and thus, the flow is two-dimensional (axisymmetric three-dimensional). The governing equation to characterize the flow field in the pool can now be written. The resulting integral equation can be solved for the lower plenum velocity field (in the radial direction) and axial velocity (in-cell velocity field) , by using the method of collocation. It should be added here that the hydrodynamic loss coefficients which enter into the formulation of the integral equation are also taken from well-recognized sources and wherever discrepancies in reported values exist, the conservative values are consistently used.
\
After the axial velocity field is evaluated, it is a straight-forward matter to compute the fuel assembly cladding temperature. The knowledge of the overall flow field enables pinpointing the storage location with the minimum axial flow (i.e. , maximum water outlet temp-erature). This is called the most " choked" location. It is recog-nized that these storage locations, where rack module supports are located, have some additional hydraulic resistance not encountered in other cells. In order to find an upper bound on the temperature in such a cell, it is assumed that it is located at the most " choked" location. Knowing the global plenum velocity field, the revised axial flow through this choked cell can be calculated by solving the Bernoulli's equation for the flow circuit through this cell. Thus, an absolute upper bound on the water exit temperature and maximum fuel cladding temperature is obtained. It is believed that in view of the precading assumption, the temperatures calculated in this manner over-estimate the temperature rise that will actually be obtained in the pool. O 5- 2
.. _ . _ _ _ _____ _ . _ _ _ . _ . . _ _ _ . _ _ _ _ . _ _ _ _ . . _ . _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ __ _ _ _ . . ___~. _. __ __ _ _ __
I i 5.3 RESULTS i lO To be supplied. 1 I I ! l. i l i f l $ l i l i 4 I lO i i l 1 l l I i I I i l l l [ ; r I i l0 1 t 5-3 l w w w --ure=~ 'eww _____---nu---wene. ..,e. e-, er% ---- cw r w n-- --Tet-- w w- w----- --= -g---e- =
1 6 REFERENCES TO SECTION 5 0 1. . - FSAR, Quad Cities, Section 10, Auxiliary and Emergency Systems.
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- 2. U.S. Nuclear Regulatory Commission, Standard Review Plan, Branch .
Technical Position, APCSB 9-2, Rev. 1, November 1975.
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h O O O j IDE ALIZ EU OU T LIN E l OF POOL 8OUNDARY ACTUAL OU T LlH E OF 3 IDE ALIZE D OUTLIN E OF R ACK ASSEMBLY f [ RACK AS S E MBLY 4 AC TU AL OUTLINE /,
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ASS UMED ADDED FUEL ASSEMBLIES I.[.. ,_. :. N. Figure 5.1 Rack Space Enveloping Cylinder (Quad Cities Station Units 1 and 2) i
- 6. STRUCTURAL ANALYSIS
\_/
The purpose of this section is to demonstrate the structural adequacy of the spent fuel rack design under high-density normal and The results show that the high-density 1 accident loading conditions. spent fuel racks are structurally adequate te resist the postulated stress combinations associated with normal and accident conditions. 6.1 Analysis outline The spent fuel storage racks are seismic category I equipment. Thus, they are required to remain functional during and after an SSE (Safe Shutdown Earthquake) .1 As noted previously, these racks are
.neither anchored to the pool floor, nor are they attached to the side walls. The individual rack modules are not interconnected. Further-more, a particular rack may be completely loaded with fuel assemblies (which corresponds to greatest rack inertia), or it may be partially loaded so as to produce maximum geometric eccentricity in the struc-r~m ture. The coefficient of f riction, p , between the supports and pool floor is another indeterminate factor. According to Rabinowicz, the results of 199 tests performed on austenitic stainless plates sub-merged in water show a mean value of g to be 0.503 with a standard deviation of 0.125. The upper and lower bounds ( 2a) are thus 0.753 and 0.253, respectively. Two separate analyses are performed for this rack assembly with values of p equal '.o 0.2 (lower limit), and 0.8, respectively. In summary, the following six separate analyses are performed:
- 1. Fully loaded rack (all storage locations occupied),
I 4 = 0.8 (# = coefficient of friction).
- 2. Fully loaded rack, p = 0.2.
- 3. Half-loaded rack to produce maximum geometric asymmetry about the major dimension of the rectangular rack,g = 0.8.
- 4. Half-loaded rack to produce maximum geometric asymmetry about the major dimension of the cectangular rack, g 0.2
- 5. Half-loaded rack to produce maximum loading asymmetry about a p)
(_ diagonal, g= 0.8. 6-1
- 6. Half-loaded rack to prc3uce maximum loading asymmetry about a O -di se# 1 - o.2.
The method of analysis employed is the time history method. The ground acceleration.coincidently in three directions is specified by CECO. The object of the seismic analysis is to determine the structural response (stresses, deformation, rigid body motion, etc.) due to simultaneous application of the three orthogonal excitations. Thus, recourse to approximate statistical summation techniques such as SRSS method 3 is avoided and the dependability of computed results is ensured. i The seismic analysis is performed in four steps, namely 1
- 1. Development of nonlinear dynamic model consisting of beam, gaps, spring, damper, and inertia coupling ele-ments.
- 2. Derivation and computation of element stiffnesses using a sophist icated elastostatic model.
1
- 3. Layout of the equations of motion, and inertial '
- decoupling and solution of the equations using the 4 1
" component element time integration" procedure to
! determine node and element forces and displacement of
- nodes.
i
- 4. Computation of the detailed stress field in the rack structure using the detailed elastostatic model from the nodal forces calculated in Step III above. Deter-i mine if the stress and displacement limits (given in Section 6.5) are satisfied.
j A brief description of the dynamic model now follows. .O 6-2
~ _ _
l 6.2 Fuel Rack - Fuel Assembly Model 6.2.1 Assumptions
- a. The fuel rack metal structure is represented by five lumped masses connected by appropriate elastic springs. (Refer to Figure 6.1) . The spring rates simulate the elastic behavior 1 of the fuel racks as a beamlike structure,
- b. The fuel assemblies are represented by five lumped masses located, relative to the rack, in a manner which simulates either full or partially filled conditions.
- c. Tne fuel rack base is considered as a rigid body supported at four points.
- d. The rack base support may slide or lift off the pool floor.
- e. The pool floor is assumed to have known time history ground
(~
~/ accelerations in three orthogonal directions,
- f. Fluid coupling between rack and assemblion, and between rack and adjacent racks is simulated by introducing appropriate inertial coupling into the system kinetic energy.
- g. Potential impacts between rack and assemblies are accounted 4 for by appropriate spring gap connectors between masses involved.
- h. Fluid damping between rack and assemblies, and between rack and adjacent rack is simulated by the inclusion of appro-priate equivalent linear damping.
j i. The supports are modeled as rigid beams for dynamic analy-sis. The bottom of a support leg is attached to a fric-tional spring as described in Section 6.2.2. The elastic () properties of the support beams are derived and used in the final computations to determine support leg stresses. 6-3
- j. The effect of sloshing is shown to be negligible and is hence neglected. It is to be noted that the top of the rack is over 20 feet below the free water surface.
6.2.2 Mode: Description The absolute degrees of freedom associated with each of the mass locations i, i* is as follows (see Figure 6.1): Location _ _ Displacement Rotation (Note) u uy ug ex ey az x 1 Py P P 95 96 2 3 94 1* Point is assumed fixed to base at XB,YB, Z=0 2 911 912 7 9 2* Pg P 10 3 P # 918 13 15 917 3* P 14 P 16 4 P yg P 21 923 924 4* .P 20 P 22 5 P E 25 27 32 929 930 931 l 5* P 26 28 Thus, there are 32 degrees of freedom in the system. Note that elastic motion of the rack in extension is represented by generalized coordinates P3 and P32 This is due to the relatively high axial rigidity of the rack. Torsional motion of the rack rela-tiv4 to its base is governed by q31' The members joining nodes 1 to 2, 2 to 3, etc., are beam elements l1 l with deflection due to be, ding and shear capability (see Reference 4, pp. 156-161). The elements of the stiffness matrix of these beam ele-ments are readily computed if the effective flexure modulus, torsion modulus, etc., for the rack structure are known. These coefficients follow from the elastostatic model as described later. The node points 1 (i = 1,2. . 5) denote the fuel rack mass at the 5 elevations. The node points i* (i* = 1,2.. 5) denote the cumulative mass for all 1 6-4 i
the fuel assemblies distributed at 5 elevations. Referring to the () General Electric specification,5 the bending and torsional stiffnesses of the fuel assembly (channeled or unchanneled) are several orders of magnitude smaller than the stiffnesses of the rack beam elements. Hence, it is reasonable to neglect the spring elements joining these lumped masses. In order to demonstrate that fuel assembly structural springs can be disregarded to produce conservative resul's, the case (refer to Section 6.1) which yields maximum rack primary stress is also run with beam springs connecting fuel assembly lumped masses. The nodes i* are located at X =XB, Y =YB in the global coordinate l1 system shown in Figure 6.1. The coordinates (XB, Y) B are determined by the center-of-mass of the set of fuel assemblies. For a completely loaded rack XB"YB = 0. 6.2.3 Fluid Coupling An effect of some significance requiring careful modeling is the so-called " fluid coupling effect." If one body of mass mi vibrates adjacent to another body (mass m2) , and both bodies are submerged in a frictionless fluid medium, then the Newton's equation of motion for the two bodies have the form (mi + M11) M3 - M12 N2 = applied forces on mass mi l1
-M21 N1+ (m2 + M2 2) N2 = applied forces on mass M2 M11, M12, M21, and M22 are fluid coupling coef ficients which depend on the shapes of the two bodies, their relative disposition; etc,. Fritz gives data for Mij for various body shape and arrangements. It is to be noted that the above equation indicates that effect of the fluid is to add a certain amount of mass to the body (Mil to body 1), and an external force which is proportional to the acceleration of the adjacent body (mass m2) . Thus, the acceleration of one body affects the force field on another. This force is a strong function of the interbody gap, reaching large values for very small gaps. This inertial coupling is called fluid coupling. It has an important effect in rack dynamics. The lateral motion of a fuel assembly inside the storage location will encounter this effect. So will the motion of a rack adjacent to another rack. These effects are included in the
/~% (,/ equations of motion. The fluid coupling is between nodes i and i* (i = - 2,3. . 5) in Figure 6.1. Furthermore, nodal masses i contain 1 6-5
coupling terms which model the effect of fluid in the gaps between [] v adjacent racks. Finally, fluid virtual mass is ir:cludea in vertical direction vibration equations of the rack; virtual inertia is added to the governing equations corresponding to rotational degrees of freedom, such as q4, q5' 96 ' 911' "
- 6.2.4 Damping In reality, damping of the rack motion arises from material hysteresis (material damping), relative intercomponent motion in structures (structural damping), and fluid drag effects (fluid damping). The fluid damping acts on the i and i* nodal masses. In the analysic, a maximum of 4% structural damping is imposed on elements of 1
the rack structure during SSE seismic simulations. This is in accord-ance with NRC Specifications.7 6.2.5 Impact 3 ~J The fuel assembly nodes i* will impact the corresponding struc-tural mass node i. To simulate this impact, 4 impact springs around each fuel assembly node are provided (see Figure 6. 2) . The fluid l 1 dampers are also provided in parallel with the springs. The spring constant of the springs is equal to the local stiffness of the vertical panel computed by evaluating the deflection of a 6-inch-diameter circular plate (0. 07 5-inch) uniformly loaded and built in around the edge. The spring constant calculated in this manner should provide an upper bound on the local stiffness of the structure. 6.2.6 Assembly of the Dynamic Model The dynamic model of the rack, rack base plus supports, and internal fuel assemblies is aodelled for the general three-dimensional 1 (3-D) motion simulation, by five lumped masses and inertia nodes for the rack, base, and supports, and by five lumped masses for the p y! assemblage of fuel assemblies. To simulate the connectivity and the elasticity of the configuration, a total of 37 linear spring dampers, 6-6
20 nonlinear gap elerants, and 18 nonlinear friction elements are ; O eeed. ^ emmarv of grine, eae, end friction eremene ita their connectivity and purpose is presented in Table 6.2. If we restrict the simulation model to two dimensions (one hori-zontal motion plus vertical motion, for example) for the purposes of model clarification only, then a descriptive model of the simulated structure which includes all necessary spring, gap, and friction ele-ments is shown in Figure 6.3. The beam springs KS, KB at each level, which represent a rack segment treated as a structural beam,4 are located in Table 6.2 as linear springs 2, 3, 6, 7, 10, 11, 14, and 15. The extensional spring KE, which simulates the lowest elastic motion of the rack in extension relative to the rack base, is given by linear spring 37 in Table 6.2. The remaining springs either have zero coef ficients (if fluid damping is neglected) , or do not enter into the two-dimensional (2-D) motion shown in Figure 6.3. The rack mass and inertia, active in rack cending, is apportioned to the five levels of 1 rack mass; the rack mass active for vertical motions is apportioned to locations 1 and 5 in the ratio 2 to 1. The mass and inertia of the rack base and the support legs is concentrated at node 1. The impacts between fuel assemblies and rack show up in the gap elements, having local stiffness KI, in Figure 6.3. In Table 6.2, these elements are gap elements 3, 4, 7, 8, 15, 16, 19, and 20. The l support leg spring rates Ks are modelled by elements 9,10 in Table 6.2 for the 2-D case. Note that the local elasticity of the concrete floor l is included in Kg. To simulate sliding potential, f riction elements 2 plus 8 and 4 plus 6 (Table 6.2) are shown in Figure 6.3. The local spring rates Kf reflect the lateral elasticity of the support legs. Finally, the support rotational f riction springs KR, reflect the rota-tional elasticity of the foundation. The nonlinearity of these
- springs (friction elements 9 plus 15 and 11 plus 13 in Table 6.2) reflects the edging limitation imposed on the base of the rack support legs.
O 6-7
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Table 6.2 Numbering System for Sprinas, Gao Elements, Friction Elements I. Spring Dampers (37 total) N, umber Node Location Description 1 1-2 X-Z rack shear spring 2 1-2 Y-Z rack shear 3 1-2 Y-Z rack bending spring 4 1-2 X-Z rack bending 5 2-3 X-Z rack shear 6 2-3 Y-Z 7 2-3 Y-Z rack bending 8 2-3 X-Z 9 3-4 X-Z d rack shear 10 3-4 Y-Z h 13 3-4 Y-Z J. rack bending g 12 3-4 X-ZI U 1 13 4-5 X-Z d rack shear 14 4-5 Y-Zh 15 4-5 Y-Z rack bending 16 4-5 X-Z 17 l-5 Rack torsion spring' 18* 1 Fluid damping of rack in torsion 19* 1 Fluid damping of rack in X direction 20* 1 Rack fluid damper in Y direction 21* 2 X direction rack fluid damper 22* 2 Y direction rack fluid damper 23* 3 X direction rack fluid damper , 24* 3 Y direction rack fluid damper 25* 4 X direction rack fluid damper 26* 4 Y direction rack fluid damper 27* 5 X direction rack fluid damper 28* 5 Y direction rack fluid damper 29* 2,2* X rack / fuel assembly damper 30* 2,2* Y rack / fuel assembly damper 31* 3,3* X rack / fuel assembly damper 32* 3,3* Y rack / fuel assembly damper 33* 4,4* X rack / fuel assembly damper s 34* 4,4* Y rack / fuel assembly damper
- Note: Dampers 18-36 assumed inactive.
6-8
Table 6.2 (continued) s Number Node Location Description
'~] 35* 5,5* X rack / fuel assembly damper 36* 5,5* Y rack / fuel assembly damper 37 1-3 Z rack extensional spring
- Note: Dampers 18-36 assumed inactive for zero fluid damping runs.
II. Nonlinear Springs (Gap Elements) (20 total) Number Node Location Description 1 2,2* X rack / fuel assembly impact spring 2 2,2* X rack / fuel assembly impact 3 2,2* Y rack / fuel assembly impact 4 2,2* Y rack / fuel assembly impact 5 3,3* X rack / fuel sGsembly impact 6 3,3* X rack /fac1 assembly impact 7 3,3* Y rack / fuel assembly impact 8 3,3* Y rack / fuel assembly impact 9 Support S1 Z compression spring 10 Support S2 Z compression spring 11 Support S3 Z compression spring 12 Support S4 Z compression spring l 13 4.4* X rack / fuel assembly impact spring i 14 4,4* X rack / fuel assembly impact spring 15 4,4* Y rack / fuel assembly impact spring 1 !() 36 17 4,4* 5,5* 5,5* Y X X rack rack rack
/
/
/
fuel fuel fuel assembly assembly assembly impact impact impact spring spring spring 18 19 5,5* Y rack / fuel assembly impact spring l 20 5,5* Y rack / fuel assembly impact spring l III. Friction Elements (16 total) - Number Node Location Description 1 Support S1 X direction support friction 2 Support S1 Y direction friction 3 Support S2 X direction friction l 4 Support S2 Y direction friction 5 Support S3 X direction friction 6 Support S3 Y direction friction 7 Support S4 X direction friction 8 Support S4 Y direction friction 9 S1 X Floor Moment ! 10 S1 Y Floor Moment I 11 S2 X Floor Moment l 12 S2 Y Floor Moment l 13 S3 X Floor Moment 14 S3 Y Floor Moment 15 S4 X Floor Moment 16 S4 Y Floor Moment , (h ( \) 6-9
For the 3-D simulation, carried out in detail for this analysis, additional springs and support elements (listed in Table 6.2), are included in the model. Coupling between the two horizontal seismic motions is provided by the offset of the fuel assembly group centroid which causes the rotation of the entire rack. The potential exists for the assemblage to be supported on 1 to 4 rack supports during any instant of a complex 3-D seismic event. All of these potential events may be simulated during a 3-D motion and have been observed in the i results. A brief description of the elastostatic model follows. This detailed model is used to obtain overall beam stiffness formulae for the rack dynamic model, and to determine detailed stress distributions in the rack from a knowledge of the results of the time history analysis. 6.3 Stress Analvsis
/^ 6.3.1 Stiffness Characteristics O
The fuel rack is a multicell, folded-plate structure which has what is colloquially called an " egg-crate" configuration. This type of construction is very similar to the so-called " stressed-skin" con-struction of ribs, spars, and cover plates which are widely used in aircraft construction. Techniques developed in the field of aircraft i structural analysis are utilized herein to find the stresses and deformations in such structures. These methods have been thoroughly [ l tested and their reliability has been documented in a number of well-known publications.8-12 l Figure 6.4 shows two cross sections of the fuel rack which is modelad as a rectangular network of plates interconnected along nodal lines shown as points in Figure 6.1. Anarbitraryloadwithcompo-l1 nents Fxi, Fyi , Fzi acts as an arbitrary elevation on one of the nodal lines. We find the displacements and stresses due to such a typical load according to the stressed-skin model as follows. I U l l 6-10 l
The torsional deformations are solved for by using the classical (S kl theory of torsion for multicelled, thin-wa'. led, cross sections.13 The bending deformation is found by using the theory of shear ficw l2 wherein all axial stresses are carried by the ef fective flanges (or stringers) formed by the intersections of the plates and all transverse shears are carried by the plates modeled as shear panels. From a knowledge of the shear flows, the bending and torsional deformations, it is possible to provide a set ,of influence functions or the following section properties for the fuel rack as a whole: (EI)eq = Bending rigidity (in two places) (GJ)eq = Torsional rigidity (AE)eq = Extensional rigidity ks = Shear deformation coefficient ()) Such properties are used for the dynamic analysis of seismic loads. l1 6.3.2 Combined Stresses and Corner Displacements The cross-sectional properties and the Timoshenko shear correc-tion factor calculated in the previous section are fed into a dynamic analysis of the system shown in Figure 6.5 with a specified ground motion simulating earthquake loading. From the dynamic analysis, the l stress resultants (Fx, Fy, F2, Mx , My, Mz) action as shown in Figure 6.6 are computed for a large number of times t = At , 2at, etc., at a The displacements (Ux, Uy, Jz) at selected number of cross sections. selected nodal points on the z axis are also provided by the dynamic analysis as well as rotations (ex, oy, az) of the cross sections at the nodes. O kJ Figure 6.7 shows a typical subdivision of the structure into ele-ments, nodes, and sections. The stresses are calculated at all 6-11 L
sections and the displacements at all four corners of the rack are () calculated at these elevations. Since a varies linearly over the crosa section and achieves its extreme values at one of the four corners of the rack, the shear stres-ses due to torsional loads (Mz) achieve their extreme values near the middle of each side. The shear stresses due to lateral forces (Fx, F y) will achieve their extreme values at the center of the cross section or at the middle of each side. Thus, candidates for the most critical point on any section will be the points labelled 1 through 9 in Figure 6.8. The expression for the combined stress and kinematic displace-ment for each of these points is written out. Simi.'.arly, the stresses in the so, port legs are evaluated. A proprietary computer program "EGELAST" computes the stresses at l1 the candidate points in each level. It sorts out the most stressed location in space as well as time. The highest stress, and maximum , kinematic displacement are thus readily found. 6.4 Time Integration of the Equations of Motion Hav..;g assembled the structural model, the dynamic equations of motion corresponding to each degree of freedom can be written by using Newton's second law of motion; or using Lagrange's equation. For example, the motion of node 2 in y-direction (governed by the gener-alized coordinate Pa is written as follows: l The inertial mass is l
*22 + A222 + B211 where m 22 is the mass of node 2 for y-directional motion.
A is the fluid coupling mass due to interaction with node 2*, and 222 B 211 is the fluid coupling mass due to interaction of node 2 with the () reference frame (interaction between adjacent racks). 6-12
Hence, Newton's law gives
.O
(*22 + A222 + B211) E a +A222 Elo + L212 U = 09 where 09 represents all the beam spring and damper forces on node 2, and A212 is the cross term fluid coupling effect of node 2*; B2 12 is the cross term fluid coupling effect of the adjacent racks. u (+) represents the ground acceleration. Let 99=pg - u, q10 " 10
-u That is, q9 is the relative displacement of node 2 in x-direction with respect to the ground. Substituting in the above equation, and rear-ranging, we have I"22 + A222 + B211 IN9+A212 910 = 09 (m22 " ^222 + 8211 + ^212 +
B212)N O A similar equation for each one of the 32 degrees of freedom can be written out. The. system of equations can be represented in matrix notatior. as: [M] @ } = [0] + l G } nodal displacement and where the vector [Q]is a function of velocities, and {G} depends on the coupling inertias and the ground acceleration. Premultiplying above equation by [M]-1 renders the resulting equations uncoupled in mass. We have:. - { 5 l = [M]-1 [Q] + [M]-1 {G } The generalized force 0, 9 which contains the effects of all spring elements acting on node 2 in the " direction" of coordinate qg l (the relative displacement of node 2 in the y direction) , can easily
, /~T be obtained from a free body analysic of node 2. For example, in the
(_/ 6-13
2-D model shown in Figure 6.3, contributions to 09 are obtained from
/' ~' the two shear springs of the rack structure, and the two impact b) springs which couple node 2* and node 2. Since each of these four spring elements contain couplings with other component deformations through the spting force-deformation relations, considerable static coupling of the complete set of equations results. The lev e.1 of static coupling of the equations further increases when 3-D motions are considered due to the inclusion of rack torsion and general fuel assembly group centroid offset.
For example, referring to Figure 6.3, a 2-D simulation introduces static coupling between coordinates 2, 9, and 15 in the expression for 1 0; 9 this coupling comes from the shear springs simulating the rack elasticity which have constitutive relations of the form /F/= Ks/9 9-9 2 /, Ks/915'9 9/. Further, t.he impact springs introduce two additional forces having constitutive equations o' the form /F/= K I/9 8-910/. Of course, at any instant, these forces may be zero if the local gap is open. The local gap depends on the current value 9 9-910, b) N' It should be noted that in the numerical simulations run to verify structural integrity during a seismic event, all elements of the fuel assemblies are assumed to move in phase. This will provide maximum impact force level, and hence induce additional conservatism in the time history analysis. This equation set is mass uncoupled, displacement coupled; and is ideally suited for numerical solution using the central difference scheme. The computer program, natued "DYNAHIS" developed by General 1 Electric Company, performs this task in an efficient manner.4 Having determined the internal forces as a function of time, the computer program "EGELAST" computes the detailed stress and displace-ment field for the rack structure as described in the preceding sec- ! tion. s l l 6-14
. ,~,. - -. _.-_- . . . - . - . - . . . _ , _ .
6.5 Structural Acceotance Criteria m b There are two sets of criteria to be satisfied by the rack modules (a) Kinematic Criterion: This criterion seeks to ensure that adjacent racks will not impact during SSE (condition E'14), assuming the lower bound value of the pool floor surface friction coefficient. It is further required that the factors of safety against tiltingl5 are met (1.5 for OBE, 1.1 for SSE). (b) Stress Limits (1) The stress limits of the ASME Code, Section III, Subsection 1 NF, 1960 Edition were chosen to be met, since this Code provides the most consistent set of limits for various stress types, and various loading conditions. The tc11owing loading casesl4 have 1 been analyzed. (~'/)
'- SRP Designation ASME Designation (i) D+L Level A (normal condition)
(ii) D+L+E Level B (upset condition) (iii) D + L + To No ASME designation. Primary membrane plus bending stress required to be limited to lesser of 2 S y* and Su* (iv) D + L + To + E No ASME designation. Stress limit same as (iii) above , (v) D + L + Ta + E No ASME designation. Stress limit same as above (vi) D + L + Ta + E' Level D (f aulted condition) where l D = Dead weight induced stresses i L = Live load induced stresses; in this case stresses developed during lifting I
*Sy: Yield stress of the material; Su: ultimate stress.
6-15 l l [
E = OBE (time history loading) A E' = SSE '] To = Stresses due to asymmetric heat emission from the fuel assemblies Ta = Thermal stresses due to accidents The conditions Ta and To cause local thermal stresses to be produced. The worst situation will be obtained when an isolated stor-age location has a fuel assembly which is generating heat at the maxi-mum postulated rate. The surrounding storage locations are ascumed to contain no fuel. Furthermore, the loaded storage location is assumed to have unchanneled fuel. Thus, the heated water makes unobstructed contact with the inside of the storage walls, thereby producing maximum possible temperature difference between the adjacent cells, b The secondary stresses thus produced are limited to the body of the rack; that is, the support legs do not experience the secondary (thermal) stresses. (2) Basic Data: The following data on the physical properties of O ehe reck meterie1 are obtained from the ^S E code. Seceien 111, appe'idices. Table 6. 3 Physical "roperty Data
- Property Young's Yield Ultimate Allowable Modules Strength Strength Stress E Sy Su S Value 28.3 x 106 25 KSI 71 KSI 17.8 KSI psi Reference Table Table Table Table I-6.0 I-2.2 I-3.2 I-7.2
- Evaluated at 200eF. This temperature is higher than the pool water bulk temperature under any of the loading conditions under
() g consideration. 6-16
(3) Stress limits for normal and upset, and faulted conditions: The following limits are obtained from NF-3230 in conjunction with Appendix XVII as modified by the NRC Regulatory Guide 1.124.16 (3.1) Normal and upset conditions (level A or level B): (i) Allowable stress in tension on a net section = Ft" 0.6 Sy or Ft= (0.6) (25000) = 15000 psi j Ft is equivalent to primary membrane stresses (ii) On the gross section, allowable stress in shear is Fy = 0. 4 Sy
= (0.4) (25000) = 10000 pai (iii) Allowable stress in compression, Fa 7 ,
1~ C 'y _
+ 8C -
8C c c - where Cc " s = 147.81 Y l Substituting numbers, we obtain, for both support leg and 4
" egg-crate" region:
Fa = 15000 psi (iv) Maximum bending stress at the outermost fiber due to flexure about one plane of symmetry: Fb = 0.60 Sy = 15000 psi l i (v) Corabined flexure and compression: f f j _fa_ , C,x bx , C,y by < y F DF a x bx y by a 6-17 l
where fa: Direct compressive stress in the section fbx: Maximum flexural stress x-axis Fby: Maximum flexural stress y-axis Cmx = Cmy = 0.85 Dx = 1 EA-- Fe'x f a Dy =1 - F'ye where 12,2E p, , 23g/k1\ b I
\ *b )
(vi) Combined flexure and compression (or tension) ,p I f I a bv 0.6 S y +Fbxbx
+byI
- The above requirement should be met for both direct tension or compression case.
(3.2) Faulted Condition: F-1370 (Section III, Appendix F) , states that the limits for the faulted condition are 1.2 times the corresponding limits for normal condition. Thus, the multiplication factor is Factor = (1.2) = 2.0 ! (3.3) Thermal Stresses: There are no stress limits for thermal (self-limiting) stresses in Class 3-NF Structures for linear-type supports. However, the range of primary and secondary stress intensity {) is required to be limited to 3 Sm in the manner of Class 1 6-18
components; Sm is the allowable stress intensity of the rack (3 material at the maximum operating temperature.
%)
6.6 Accidents Associated with Rack Integrity In addition to the ground motion analyses, the following mechani-cal loads have been analyzed:
- a. Dropped Fuel Accident I A fuel assembly (weight -
600 pounds) dropping from 36 inches above a storage location and impacting the base. Local failure of the base plate is acceptable; however, a substantial impact with the pool liner is not acceptable. The suberiticality of the adjacent fuel assemblies is not to be violated,
- b. Dropped Fuel Accident II One fuel assembly dropping from 36 inches above the rack
, and hitting the top of the rack. Permanent deformation of the rack is acceptable, but is required to be limited to the top region such that the rack cross-sectional geometry at the level of the too of the active fuel (and below) is not altered.
- c. Jammed Fuel-Handline Equipment and Horizontal Force A 2000-pound uplift ,cce and a 1000-pound horizontal force are applied at the top of the rack at the " weakest" storage location; the force is assumed to be applied on one wall of the storage cell boundary as an upward shear force. *he damage, if any, is required to be limited to the region
! above the top of the active fuel. , The above loading conditions were analyzed to determine an upper bound on the plastic deformation zones. For the above conditions, it has been shown that the plastic deformation is limited to the rack l structure well removed from the active fuel regions. Thus, the sub-fh J criticality of the fuel arrays is not modified or violated. 6-19 l - .-- . . - -. -
REFERENCES TO SECTION 6 g
- 1. Regulatory Guide 1.29, " Seismic Design Classification," Rev. 2, February 1976.
- 2. " Friction Coef ficients of Water Lubricated Stainless Steels for a Spent Fuel Rack Facility," by Prof. Ernest Rabinowicz, MIT, a report for Boston Edison Company.
- 3. U.S. Nuclear Regulatory Commission, Regulatory Guide 1.92,
" Combining Modal Responses and Spatial Components in Seismic Response Analysis," Rev. 1, February 1976.
- 4. "The Component Element Method in Dynamics with Application to Earthquake and Vehicle Engineering" by S. Levy and J. P. D.
Wilkinson, McGraw Hill, 1976.
- 5. General Electric specification 2 2A 586 6, Rev. 1, Appendix II,
" Fuel Assembly Structural Characteristics."
- 6. R. J. Fritz, "The Effect of Liguids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, Trans. of the ASME, February 1972, pp. 167-172.
U.S. Nuclear Regulatory Commission, Regulatory Guide 1.61. l1 {} 7.
- 8. J. T. Oden, " Mechanics of Elastic Structures," McGraw-Hill, N.Y.,
1967.
- 9. R. M. Rivello, " Theory and Analysis of Flight Structures,"
McGraw-Hill, N.Y., 1969.
- 10. M. F. Rubinstein, " Matrix Computer Analysis of Structures,"
Prentice-Hall, Englewood Cliffs, N.J., 1960.
- 11. J. S. Przemienicki, " Theory of Matrix Structural Analysis,"
McGraw-Hill, N.Y., 1966.
- 12. P. Kuhn, " Stresses in Aircraft and Shell Structures," McGraw-Hill, N.Y., 1956.
- 13. S. P. Timochenko and J. N. Goodier, " Theory of Elasticity,"
McGraw-Hill, N.Y., 1970, Chapter 10.
- 14. U. S. Nuclear Regulatory Commission, 'tandard Review Plan, NUREG-75/087, Section 3.8.4.
- 15. U.S. Nuclear Regulatory Commission, Standard Review Plan, Section 3.8.5.
- 16. U.S. Nuclear Regulatory Commission, Regulatory Guide 1.124, k, " Design Limits and Loading Combinations for Class 1 Linear-Type Component Supports, November 1976.
6-20
O 3 1 7 a3 3
. o /
/
CO U PLIN G ELEMENTS 4 TYPICAL FU EL A S S EM B LY 3 GROUP MAS S H TYPIC AL FUEL RACK MAS S 2 FUEL R ACK B A SE 2 Ay , =
/ 1
\ /
p Ax "' ye l =Y l y 7-l
/- +- %* 8I 32 I / !
a a s e I I h l ys sY FUEL R ACK SUPPORT
~
1 X X8, YB - L O C ATION O F C E N'T R O I D O F FU E L ROD G ROUP M ASSES - REL ATIVE TO CENTER O F FU EL R A C K g ni = UNIT VECTORS Figure 6.1 Dynamic Model
O Y IMPACT k SPRIN G S
- i b .}
[d. MASS P
- v= ; 4* -
du } F LU ID
}
DAMPERS RIGID FR AM E
- =X I
Figure 6.2 Impact Springs and Fluid Dampers Revision 1
5 ' l
~~ '
O 5 K , K,(Typ ) 4 v 4 Seismic E l H W>W Motions 4 Z Fuel Assembly Group twTped Mass y (Typ.) v If kW WW 3 Rack Lumped Mass & Inertia For Horizontal C Motions (Typ.) y s .] l V4N N 2 K, (Typ.) JL K3 g@ - B --> K s h
/ i E ' f-
/t fmNV Mp x, A da x, A,
W 4 Figure 6.3 Spring Mass Simulation For A Two Dimer.:,ional Motion fievis.on 1
m O i j'I Y Y4 sk . ; e L =FA
. "X (a) TOP VIEW zh I
y yz F .
# I O "
=F x (b) AXIAL CROSS u-S ECTlON ( B-B )
l Figure 6.4 (a) Horizontal Cross Section of Rack (b) Vertical Cross Section of Rack i l 0
i i O ICW) CELL 0 WALLS ffc t -,nllllAfs' , n , 12
/ ownnons n h
' n^ ^
n" ,'/ ~ b= NyC 1 d a:NxC7 C . A' , y(V) A RIGID PLATE B / i ,7
*/I> = X (u) 1 i f *7 f "'
_ a:. +}
- , A j l_.,
4- /
/
SUPPORTS Figure 6.5 hMZ d
/ MY C' .
./ /
A./ _; M X
'- Fx B ,C
! A 8 l l l l Figure 6.6 l O l l l
l O oZ NODE I-
, - E L. I S E C. I - - - - - -
NODE 2 - / , EL.2 SEC.2 - NOD E 3 - E(3 S E C. 3 - -- , N O D E 4 -% - 3g l - - EL 4 ,F-p,, E L 5 EC.4 ---- Tz
> / -
f - S EC.5
' J,, *- ,
, -- a .
SEC 6 x- 4 . NODES k ' ,c-wp gR O OT O F R'A CK' - - E L.8 E L.7 - ~ y
-- S E C. 8 SEC 9 N O. OF ELEM EN TS = 8 N O. OF S EC TION S' =9 i
N O. 'O F N O'D E S =5 Figure 6.7 l l L__..._..__.__._.______.___.__._.__..__.___._ _ __
P' m O i l l Y m i @ .,
@x W
O "
- 0 =
l Figure 6.8 l l l 'O 2 l I
- 7. ACCIDENT ANALYSIS 1
O Information regarding heavy object drop to be supplied.
\
J l f O t i i O 7-1
- 8. RADIOLOGICAL CONSEQUENCES O Obiectives and Assumptions 8.1 The radiological consequences of expanding the storage capacity of the spent fuel storage pool have been evaluated wita the objective of determining if there is any significant additional radiological impact, onsite or offsite, relative to that of the currently author-ized spent fuel storage pool. The principal factors considered in evaluating the additional radiological consequences were the fol-lowing:
e Opurating experience and measurements, o Reduction in decay heat generating rate and fuel tempera-tures with time following removal from the reactor. o Age and nature of the additional fuel to be stored, y e Analyses of radionuclide releases to the pool water from failed fuel. In addition, the radiological impact to operating personnel has been evaluated to ensure that such exposure remains as low as is reasonably achievable. J Each spent fuel pool is currently authorized to store approxi-mately two full cores. By comparison, the expanded spent fuel storage capability can accommodate more than five full core loads in each pool. The additional storage will be aged fuel which has been out of the reactor 5 years or more. It is important to note that the dif-ference between the radiological impact for the currently authorized storage pool capacity and the expanded storage pool capacity is attributable entirely to the presence of additional aged fuel in the expanded spent fuel storage pool. The radiological consequences of storing the additional quantity of aged fuel has been evaluated. To ensure a conservative evaluation 8-1
- - - _ _ _ _ _ _ _ _ _ _ - _ . ._. )
of the storage of failed fuel, it was assumed that the spent fuel storage pool is entirely filled with high-burnup spent fuel (28,500 Mwd /MtU burnup) , ranging from newly removed fuel (1 core load of 724 fuel assemblies) to aged fuel with a cooling time of approximately 18 years. The maximum fission-product inventory in the stored fuel would result from an idealized fuel cycle in which approximately 181 spent fuel elements were removed from the core and placed in the pool annually. With this fuel cycle, the expanded storage pool capacity, when completely filled, would contain the following: (1) For currently authorized 724 newly removed assemblies storage capacity (full core load) and 4 refueling discharges with storage periods of 1, 2, 3, and 4 years, respectively. (2) Aged fuel in expanded 13 refueling discharges, with storage capacity storage periods of 5 to 17 1 {} years and capacity, any remaining containing fuel stored for 18 years. Reduced fuel burnup or increased cycle length would result in a lower fission-product invantory or longer storage (decay) periods. Thus, the assumed storage pool composition should result in a con-servative estimate of any additional radiological impact due to the expanded storage capacity. l 8.2 Operating Experience and Nature of Stored Fuel t l In a survey 1 of spent fuel storage pool experience, Johnson, at Battelle Pacific Northwest Laboratories, has shown that typical con-
-4 centrations of radionuclides in spent fuel pool water range from 10
-2 gCi/ml, with the higher value associated pCi/ml, or less, to 10 with refueling operations. Isotopic measurements of the nuclides con-firm that a major fraction of the coolant activity results from OV activated corrosion products dislodged from fuel element surfaces during refueling operations or carried into the spent fuel pool water 8-2
(with some fission-product radionuclides) by mixing the pool water with primary system water during refueling. These sources of storage (]}) pool radionuclides depend upon the frequency of refueling operations : and are basically independent of the total number of fuel assemblies in storage. Once fuel-handling operations are completed, the mixing of pool water with primary system water ceases and these sources of radio-nuclides decrease significantly; only dissolution of fission-products absorbed on the surface of fuel assemblies and low levels of erosion of corrosion-product (crud) deposits remain. With aged fuel (5 or more years storage), neither of these latter sources would be expected to contribute significantly to the concentrations of radionuclides in the storage pool. In view of the above, it is concluded that the additional ctorage capacity of the expanded spent fuel pool will not measurably alter the 1 currently approved radiological impact or impose any significant additional burden on the cleanup system as a result of corrosion-product radionuclides or fission-product carry-over from the primary system during refueling operations. During storage, the level of gamma radiation f rora fission products in the fuel decreases naturally due to radioactive decay. Because of this decay, the contribution of the aged fuel to the dose I rate at the pool surface by direct radiation will be very small ( < 5%) compared to that from the more-recently-discharged fuel. Thus, it is concluded that the occupational dose rate above the surface of the pool from direct radiation will be essentially the same as that for l the currently authorized storage pool. 8.3 Consecuences of Failed Fuel Escape of fission-products from failed fuel stored in the spent fuel pool will contribute to the radionuclide concentrations in the pool water. However, calculations described below indicate that the () radionuclide concentrations from failed fuel are considerably less l 8-3
than the concentrations of corrosion-product radionuclides and, there-fore, the aged fuel in the expanded storage pool will not contribute significantly to the onsite or offsite radiological impact. ' The decay heat generated in spent fuel rapidly decreases (by radioactive decay) following removal from the reactor and, in the aged fuel, will be very small ( < 5% of that in f reshly-remcVed fuel) . Fuel temperatures and internal gas pressures will correspondingly decrease with time. Johnson 1 also cites evidence to confirm that UO2 is inert to the relatively-cool water of spent fuel storage pools. Therefore, the release rate of fission-products from any defective rods among the aged fuel is expected to be negligibly small. Release of fission-products from failed fuel probably results from water leaching or diffusion of material plated out or absorbed in the fuel-clad gap of the fuel element during operation. Once the material in the gap is depleted, further release will be very small. Most of the fission-products are absorbed (retained) in the fuel matrix and can escape only by diffusion through the UO2 At the
> temperatures of the fuel in the spent fuel pool, the diffusion 1 coefficient will be extremely small.2 In his survey, Johnson indicates that numerous fuel assemblies with one or more defects have been stored in several spent fuel pools without requiring special handling. Detailed analysis of the spent j
fuel pool water confirmed that fuel elements with defects do not I continue to release significant quantities of radionuclides for long periods of time following removal from the reactor. Nevertheless, the calculations described here in Sections 8.3.1 and 8.3.2 were based on the very conservative assumption that the rate of fission-product l release remains the same as the rate for newly discharged fuel. l l , Both Johnson, at Battelle, and Weeks, at Brookhaven National Laboratory, have reviewed the corrosion properties of Circaloy cladding and the integrity of spent fuel elements stored for long periods of time. They conclude that the corrosion of Zircaloy O c1eddine in event eee1 voo1 weter te nee 11eib17 sme11 end ehee there 8-4 1
, , ,,- -v. .,- . , - . - - , . - - . - - - - - - - - -
is suf ficient evidence of satisf actory fuel integrity to justify long-V term storage. Consequently, there is not expected to be any signifi-cant deterioration of stored fuel that might lead to additional fuel failures in the expanded-capacity spent fuel storage pool. 8.3.1 Methods of Analysis To assess the oaximum potential radiological contribution from failed fuel, the inventory of fission-products in the spent fuel was calculated with the ORIGEN code, conservatively assuming that all 4 Experi-fuel was discharged from the core at 28,500 Mwd /MtU burnup. mental values of escape rate coefficients in cool water shortly after discharge, as derived by Westinghouse,5 were used to calculate the fractional release of fission-products from failed fuel, and it was These escape rate as3umed that there were 1% fuel element failures. in Table 8 .1) were assumed to be constant coefficients (listed throughout the storage period, although it is known that fission-product release from failed fuel is strongly dependent upon the 1 D) (, temperatures within the fuel pin. As natural radioactive decay occurs, decay heat generation in the fuel becomes less and, as a consequence, the fuel temperatures and internal gas pressures are reduced. Furthermore, the inventory of leachable fission-products becomes depleted and release from the bulk UO2 by diffusion becomes extremely low.3 Table 8.1 Escape Rate Coef ficients into Cool Water for Spent Fuel in Storage Pool Element Escape Rate Coefficient (sec~ )
-12 I
1.7 x 10
-12 Rb, Cs 3.0 x 10
-12 Mo 1.8 x 10 Te (0.9 x 10-12)*
-1 Sr 8.5 x 10
-16 Ba 5.8 x 10
** -16 Zr 1.2 x 10 D
(G
- Escape rate coefficient for Te assumed to be in same ratio to Mo, as
~
given in NUREG-0017.
** Assumed applicable to all other nuclides.
8-5
Thus, within a few months after discharge, the fuel temperatures () and effective leak rate coefficients decrease, and further leakage is reduced to relatively insignificant levels.1 The percentage of failed fuel that exists in the stored fuel, averaged over a large number of reactor cycles, is uncertain. Johnson1 estimates that an average of 0.013 should be achievable. The NRC, in NUREG-0017, cites 0.12% as a representative value.6 Nevertheless, to establish a conservative upper limit, the calcula-tions reported here were based on the assumptions that 1% of all stored fuel is failed and that constant leak rate coefficients, cor-responding to those measured shortly after shutdown, apply over the storage periods. Concentrations of released fission-products were calculated from the dynamic balance between the source term (leakage from the assumed failed fuel) and the rate of removal by (1) radioac-tive decay and (2) the spent fuel pool cleanup system (using demineralizer cleanup ef ficiencies cited in NUREG-0017) .6 This method of analysis is similar to that used in NUREG-0017. 1 8.3.2 Fission-Product Radionuclide Concentrations Based upon the method of analysis described above, the concentra-tions of fission-product radionuclides in the opent fuel pool were calculated at several times following unloading of a full core into the spent fuel pool, with the remainder of the pool assumed to be filled with older fuel. Results of these calculations are summarized in Table 8.2. These calculated concentrations of fission-product radionuclides are directly proportional to the assumed 1% failed fuel and would be a factor of 8 lower for the 0.12% failures estimated in NUREG-0017 6 as a typical weighted average value based on operating experience in a number of reactors. Of the fission-product radio-l nuclides released, Cs-137 is the dominant acti"ity from the aged fuel
-6 (calculated to be a mar'. mum of 2. 2 x 10 gCi/ml with 1%-6failed fuel).
-6 pCi/ml) and Mo-99 (3 x 10 gCi/ml) are Low levels of I-131 (8 x 10 calculated to be present as a result of leakage from a full core load of newly removed fuel. (with 1% f ailures) . However, in the aged fuel,
() these nuclides have decayed and there are no significant quantities of I-131 or Mo-99 remaining. 8-6
- Table 8. 2 Fission-Product Radionuclide Concentrations j in Fully-Loaded Spent Fuel Storage Pool with 1% Failed Fuel Concentration (gCi/ml)
For Currently Incremental Additions Approved Storage due to Expanded Time (days) Capacity
- Capacity
- Total
-5 -5 5 2.08 x 10 -5 2.62 x 10-6 -6 2.34 x 10 -5 10 1,43 x 10 2.62 x 10 1.69 x 10
-6 -6 -5 20 9.35 x 10 2.61 x 10 1.20 x 10 30 7.54 x 10-6 2.61 x 10-6 1.02 x 10-5 50 6.23 x 10
-6 2.60 x 10-66 8.83 x 10-66 5.67 x 10-6 8.26 7.94 xx 10 10-6 75 2.59 100 5.37 x 10
-6 2.57 xx 10 10-6
*See Section 8.1 for a description of composition.
Even with 1% failed fuel, the radionuclide concentrations in the spent fuel pool water are dominated by those from corrosion products and carry-over from the primary coolant system during refueling. Furthermore, since the release rate from the aged fuel will be consid-erably smaller than that indicated in Table 8.2 (due to the lower 1 temperatures and pressures in the fuel elements) , the actual contribu-tion from the aged fuel will be negligibly small in practice. It is also expected that the percentage of failed fuel, averaged over the reactor lifetime, will be considerably less than 1%. Thus, it is con-cluded that the expanded-capacity spent fuel storage pool will not increase the radionuclide concentrations in the pool water signifi-cantly above those for the currently approved spent fuel storage pool. Consequently, expanding the storage capacity of the spent fuel pool ( will neither alter the onsite or offsite radiological impact nor significantly increase the burden on the spent fuel pool cleanup system, as a result of failed fuel. 8.3.3 Gaseous Releases from Failed Fuel Because of the half-lives of the noble-gas radionuclides, only the release of Kr-85 (T of 10.76 years) has the potential of g ,U increasing the radiological impact to the spent fuel building atmosphere as a result of expanding the capacity of the spent fuel 8-7
l i
)
storage pool. (Short-lived noble-gas radionuclides and other volatile fission-products, such as iodine, are not present in the aged fuel.) Johnson 1 concludes that the radioactive fission gases will have been largely expelled from defective fuel rods during reactor operation and, therefore, are not available for release during fuel storage. This is expected, since the noble gases are chemically inert and there are no plate-out or hold-up mechanisms in the fuel-clad gap of the fuel element. The small amount of chemically inert Kr-85 that might be absorbed on the surface of a fuel assembly and released slowly during storage, is believed to be insignificant, particularly in the aged fuel. Since UO2 is chemically inert to cool water, diffusion of Kr-85 entrapped within the UO2 fuel matrix would be the remaining source for Kr-85 release. Based on the method outlined in the proposed ANS 5.4 2 standard on fission gas release, the diffusion coefficient in the aged fuel at spent fuel pool temperatures will be negligibly small (of the order of 10-40). Consequently, diffusion release of Kr-85 from aged fuel will be negligible in accord with Johnson's findings.1 1 it is concluded that the incremental radiological impact f rom the release of Kr-85 with the expanded-capacity spent fuel storage pool will be negligibly small. 8.4 Exposure for the Installation of New' Racks The existing spent fuel racks will be removed, and the new racks I will be installed in a manner which will maintain occupational expo-t sure to levels as low as reasonably achievable ( ALARA) . The following , methods for the disposal of the existing racks are currently under l review by CECO: i e Crating and shipment of the racks in "as-is" condition. e Decontamination and shipment. e Dismantle and volume reduction, with or without prior decon-tamination, and shipment of waste. 8-8 t L
. . _ , _ _ _ _ _ _ _ . _ ~ _ . _ . _ _ _ . _ _ - . _ _ _ _ - _ _ _ - . _
The final decision concerning the disposal of the existing spent fuel racks will be based on project needs and experience gained from rack disposal at Dresden. 8.5 Conclusions Based on operating experience - and the analysis of potential releases, it is concluded that expanding the capacity of the spent fuel storage pool will not significantly increase the onsite or off-site radiological impact above that of the currently authorized storage capacity. Similarly, the expanded storage capacity will not impose any significant additional burden on the spent fuel pool cleanup system. O O 8-9
REFERENCES TO SECTION 8'
- 1. A. B. Johnson, Jr. , " Behavior of Spant Nuclear Fuel in Water Pool Storage," BNWL-2256, September 19774
- 2. ANS 5.4 Proposed Standard, " Method for Calculating the Fractional Release of Volatile Fission Products from Oxide Fuel."
- 3. J. R. Weeks, " Corrosion of Materials in Spent Fuel Storage Pools," ENL-NUREG-2021 (Informal Report) , July 1977. 1
- 4. M. J. Bell, "ORIGEN-The ORNL Isotope Generation and Depletion Code," ORNL-4628, May 1973.
- 5. J. M. Wright, " Expected Air and Water Activities in the Fuel Storage Canal," WAPD-PWR-CP-1723 (with addendum), undated.
- 6. Of fice of Standards Development, NRC, " Calculation of Releases of Radioactive Materials in Gaseous and Liquid Effluents from Pres-surized Water Reactors (PWR-Gale) ," NUREG-0017, April 19 76.
O U a 8-10
- m. --
- 9. POOL STRUCTURAL CALCULATIONS To be supplied.
O O 9-1 l l
- 10. INSERVICE SURVETLLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING l1 MATERIAL h./
s_ 10.1 Program Intent l1 A sampling plan to verify the integrity of the neutron absorber material employed in the high-density fuel racks in the long-term environment is described in this section. The program is intended to be conducted in a manner which allows access to representative absorber material samples without disrupting the integrity of the fuel storage system. The program is tailored to evaluate the material in normal use mode, and to forecast future changes using the data base developed. 10.2 Description of Specimens l1 The absorber material, henceforth referred to as " poison," used in the surveillance program must be representative of the material
> used within the storage system. It must be of the same composition, produced by the same method, and certified to the same criteria as the production lot poison. The sample coupon must be of similar thickness as the poison used within the storage system and not less than 4 x 4 inches on a side. Figure 10.1 shows a typical coupon. Each poison specimen must be encased in a stainless steel jacket of an identical alloy to that used in the storage system, formed so as to encase the poison material and fix it in a position and with toler-ances similar to that designed into the storage system. The jacket would be closed by tack welding in such a manner as to retain its form throughout the use period yet allow rapid and easy opening without contributing mechanical damage to the poison specimen contained within. The jacket will permit wetting and venting of the specimen y similar to the actual rack environment.
i I 10.3 Test l1 t h. C The test conditions represent the vented conditions of the cruci-l form elements. The samples will be located adjacent to t.he fuel racks 10-1 l l
.- . ~ _ _ . . _ .
l and suspended frcm the spent fuel pool wall. Eighteen test samples () are to be fabricated'in accordance with Figure 10.1 and installed in 11 the pool when the racks are installed. The procedure for fabrication and testing of samples shall be as follows:
- a. Samples shall be cut to size and carefully weighed in milli-t l grams.
l b. Length, width, and average thickness of each specimen to be measured and recorded.
- c. Samples shall be fabricated in accordance with Figure 10.1 l1 and installed in the pool.
- d. Two samples shall be removed at each time interval per the schedule shown in Table 10.1. Il 10.4 Specimen Evaluation l1 l rg After removal of the jacketed poison specimen from the fuel pool at the designated time, a careful evaluation of that specimen will be made to determine its actual condition as well as its apparent durability for continued function. Separation of the poison from the stainless steel snecimen jacket must be performed carefully to avoid mechanically damaging the poison specimen. Immediately upon removal, the specimen and jacket section should be visually examined for any effects of environmental exposure. Specific at*.ention should be directed to the examination of the stainless steel jacket for evidence of physical degradation. Functional evaluation of the poison material is accomplished by the following measurements:
- a. A neutron radiograph of the poison specimen will allow for a determination of the maintenance of uniformity of the boron distribution.
l
- b. Neutron attenuation measurements will allow evaluation of 1 the continuing nuclear ef fectiveness of the poison. Con-() sideration must be given in the analysis of the attenuation measurements for the level of accuracy of such measurements 10-2
as indicated by the degree of repeatability normally () observed by the testing agency.
- c. A measurement of the hardness of the poison material will establish the continuance of physical and structural dura-bility. The hardness acceptability criterion' requires that the specimen hardness will not exceed the hardness listed in the qualifying test document for lab test specimen irradi-ated to 1011 rads. The actual hardness measurement should be made after the specimen has been withdrawn from the pool and allowed to air dry for not less than 48 hours to allow for a meaningful correlation with the preirradiated sample.
- d. Measurement of the length, width, and average thickness and comparison with the pre-,xposure data will indicate dimen-sional stability within the variation range reported in the Boraflex laboratory test reports.
A detailed procedure paraphrasing the intent of this program will be prepared for step-by-step execution of the test procedure and interpretation of the test data. I I t i l 10-3 l l [
3 TABLE 10.1 (G Date Installed INITIAL FINAL WEIGHT PIT WEIGHT WEIGl}T CHANGE PENETRATION (mg/Cm 2 -Yr) (mg/cm'-Yr) (mg/Cm'-Yr) mil /Yr _' SCHEDULE i 1 2 90 day y 3 1 4 180 day 1F 5 6 1 year wr
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- 11. COST / BENEFIT ASSESSMENT n
() A cost / benefit assessment has been prepared in accordance with the requirements of Section V, Part 1.1 The purpose of the assessment is to demonstrate that the installation of high-density spent fuel storage racks is the most advantageous means of handling spent fuel, considering the needs of our customers for a dependable source of electric power. The material is presented to satisfy the NRC's need for informa-tion; it is the position of CECO that no environmental impact state-ment need be prepared in support of the request, because there will be no significant impact on the human environment. Similarly, NRC precedent establishes that alternatives and economic costs need not be discussed when there is no significant environmental impact. 1 11.1 Specific Needs for Spent Fuel Storage l I l No centractual arrangements exist for the storage or reprocessing of spent fuel from Quad Cities. Accordingly, the storage of spent fuel from Quad Cities, in the Quad Cities spent fuel pool, is the only viable option being considered. Table 11.1 shows the schedule for loss of full core discharge capability (FCDC) and reload dier arge i capability (RDC) based on the following options: e Existing spent fuel rack configuration. e Addition of spent fuel rack using the current approved design. e High-density ' pent fuel racks. In addition to spent fuel, storage is available in each Quad l Cities spent fuel storage pool for other types of materials (see Section 1). i l r i 11-1
A recent survey of the Quad Cities spent fuel storage pool indi-
.r t cated the presence of the following non-fuel materials:
o Control rods. o Control rod guide tubes. e Incore instruments. e various pieces of highly radioactive hardware. Based on the present lack of an alternative to onsite spent fuel storage, it is not possible to predict how long the additional spent fuel storage capability will be required. It is unlikely that an alternative to onsite spent fuel storage will be available before 1990. Based on Table 11.1, the proposed increase in storage capacity would accommodate the refueling of Unit 1 in the year 2003, but would not accommodate the refueling of Unit 2 in the year 2004. 11.2 Cost of Modification The design and manuf acture of the spent fuel storage racks will {} be undertaken by the organizations described in Section 1. It is 1
'xpected that the total project cocc will be between 8 and 9 million j
dollars. 11.3 Alternatives to Sp?nt Fuel Storage Expansion Commonwealth Edison has considered the various alternatives to the proposed increase in spent fuel storage capacity. These alterna-l tives are as follows: e Shipment of fuel to a reprocessina or independent spent fuel storage facility l No commercial spent fuel reprocessing facilities are In addition, CECO presently operating in the United States. has not obtained commercial spent fuel storage commitments for fuel from Quad Cities. The Department of Energy Away-l ) From-Reactor Storage Program has been terminated, and no commercial independent spent fuel storage facilities exist. l 11-2
e Shipment te another reactor site O) CECO considers the storage of spent fuel at reactor sites to be a long-term option due to the lack of any reasonable alternatives. Shipment of fuel .to another site would provide short-term relief; however, transshipment of spent fuel does not contribute to the long-term goal of providing adequate storage capacity but merely serves to transfer the problem to another site. Accordingly, CECO does not consider the transshipment of spent fuel to be an appropriate alternative to high-density spent fuel storage at the site, o Shutting down the reactor Shutting down Quad Cities would result in an economic hard-1 ship that would be imposed ca CECO shareholders and customers. Moreover, as indicated in NUREG-0575, " Final 1 Environmental Impact Statement on Handling and Storage of V Spent Light Water Power Reactor Fuel," the replacement of nuclear power by coal-generating capacity would cause excess mortality to rise from 0.59-1.7 to 15-120 per year for
- 0. 8 GWY ( e ) . Based on the above, shutting down Quad Cities does not represent a viable alternative.
The subject of the comparative economics associated with various spent fuel options is the subject of Chapter 6 of NUREG-0575. Although the material presented is generic, it is of value in comparing the costs of the various options. Of the options presented in Chapter 6 of NUREG-0575, high-density spent fuel storage at the site is the most economic option at $18 per KgU. The price of "Away
- From Reactor (AFR) " fuel storage, if available, would be $115 per KgU.
l This corresponds to 0.5 mill /Kwh from a 1000 MWe power reactor for AFR stortge. Should the lack of spent fuel storage cause a shutdown, NUREG-0575 estimates that replacement power would be 12 mill /Kwh for coal and 23 mill /Kwh for oil in 1979. !O 1 l 11-3 _ __ . - ~ _ _ _ _ . _ _ _ _ _.._._. - . _ _ _ _
11.4 Resource Commitments The expansion of the Quad Cities spent fuel storage capacity will require the following primary resources: , o Stainless steel - 926,940 pounds e Boron Carbide (B 4 C) powder - 26,264 pounds , The requirement for stainless steel represents a small fraction of the total domestic production of 175 million tons for 1980.2 Although the fraction of domestic production of B4C, required for the modification, is somewhat higher than that for stainless steel, it is unlikely that the commitment of B4C to this project will affect other citernatives.* The total boron production estimated for 1985 is 275 to 350 thousand tons. 11.5 Environmental Effects l To be supplied.
- Experience has shown that the production of B 4 C is highly variable and depends on need, but could easily be expanded to accommodate
() additional domestic demands. 11-4 r
. . . _ . _ _ _ . _ . , . . , _ _ . , . , , _ , - _ , _ - , _ . _ . _ . _ . . , _ , . .. ~ _ _ _ _ . . _ . _ _ , _ _ . _ . . - _ . , . _ . . _ . ,._ . , .
Table 11.1 Quad Cities Station: Projection for Loss of Full Core and Reload Discharge Capability Currently Available Spent Fuel Racks Capacity = 2290 Capacity with FCDC = 1556 Capacity with RDC* = 2080 Loso FCDC - 9/81 Lose RDC - 3/83 1 Currently Licensed Spent Fuel Racks capacity = 2920 Capacity with FCDC = 2196 Capacity with RDC = 2720 Lose FCDC - 3/84 Lose RDC - 3/85 High-Density Spent Fuel Racks Capacity = 7684 Capacity with FCDC = 6960 O cer ciev tea noc Lose FCDC 7484 3/02 Loae RDC - 9/03
- Full core capacity = 724 l
'O 11-5 l l E. _ . . . _ . . . . , . _ _ . . _ . . . _. . _ _ _ _ _ . _ . _ . . _ _ . _ . , _ . . . _ . _ . _ _ _ . . . _ . . _ _ . _ _ _ _ _ _ .
REFERENCES TO SECTION 11 1 D d "OT Position for Review and Acceptance of Spent
- 1. B. K. Grimes, 1 Fuel Storage and Handling Applications," April 14, 1978.
- 2. " Mineral Facts and Problems," Bureau of Mines Bulletin 667, 1975.
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