ML20117C283
| ML20117C283 | |
| Person / Time | |
|---|---|
| Site: | Grand Gulf  |
| Issue date: | 05/06/1985 |
| From: | MISSISSIPPI POWER & LIGHT CO. |
| To: | |
| Shared Package | |
| ML20117C244 | List: |
| References | |
| TAC-57619, NUDOCS 8505090377 | |
| Download: ML20117C283 (232) | |
Text
{{#Wiki_filter:- Attachment 2 Licensing Report High Density Spent Fuel Racks Grand Gulf Nuclear Station Unit 1 8505090377 850506 PDR ADOCK 05000416 P pon J0P14PMI85031504 - 2
l TABLE OF CONTENTS l l 1 REEE SECTION 1 - INTRODUCTION 1.0 Introduction 1-1 SECTION 2 - GENERAL ARRANGEMENT 2.0 General Arrangement 2-1 SECTION 3 - RACK CONSTRUCTION 3.1 Construction 3-1 3.2 -Codes, Standards and Practices 3-5 for the Spent Fuel Pool High Density Fuel Racks SECTION 4 - NEUTRONIC CONSIDERATIONS 4.1 Introduction 4-1 4.1.1 . Neutron Multiplication Factor 4-2 4.1.2 Analytical Methods 4-3 4.1.3 Calculational Bias and Uncertainty 4-5 4.1.4 Trend Analysis 4-6 4.2 Input Parameters 4-7 4.2.1 Fuel Assembly Design Specifications 4-7 4.2.2 Reference Design Spent Fuel Storage 4-7 Cell 4.3 Postulated Accidents and Abnormal Conditions 4-10 4.3.1 Temperature and Water Density Effects 4-10 4.3.2 Abnormal Positioning of Fuel 4-13 Assembly Outside Storage Rack 4.3.3 Fuel Assembly Positioning in Storage 4-13 Rack 5 - O i
t .- TABLE OF CONTENTS (Continued) 4.3.4 Effect of Zirconium Fuel Channel Distortion 4-14 4;3.5 Dropped Fuel Assembly Accident 4-14 4.3.6 Fuel Rack Lateral Movement 4-14 4.4 Criticality Analysis 4-16 4.4.1 Nominal Case 4-16 4.4.2 Maximum Reactivity Storage 4-16 Capability 4.4.3 Boron Loading Variation 4-18 4.4.4 Boraflex Width Tolerance Variation 4-20 4.4.5 Axial Cutback in Boraflex Length 4-20 4.4.6 Storage Cell Lattice Pitch Variation 4-20 4.4.7 Stainless Steel Thickness Variations 4-21 4.4.8 Fuel Enrichment and Density Variation 4-21 4.4.9 Effect of Zirconium Fuel Channel 4-21 , 4.5 Acceptance Criteria for Criticality 4-23 REFERENCES TO SECTION 4 4-24 SECTION 5 - THERMAL-HYDRAULIC CONSIDERATIONS 5-1 5.1 Forced Circulation Thermal Hydraulic Analysis 5-1 5.1.1 Basis 5-1 5.1.2 Model Description 5-3 5.1.3 Results and Discussion 5-7 5.2 Natural Circulation Thermal-Hydraulic Analysis 5-14 o 5.2.1 Basis 5-14 5.2.2 Model Description 5-15 ii , 6
TABLE OF CONTENTS (Continued) 5.2'.3 Results and Discussion 5-17 REFERENCES TO SECTION 5 5-20 SECTION 6 - STRUCTURAL ANALYSIS 6 6.1 Analysis outline 6-1
. 6.2 Fuel Rack - Fuel Assembly Model 6-3 6.2.1 Assumptions 6-3 6.2.2 Mcdel Description 6-5 6.2.3 Fluid Coupling 6-6 6.2.4 Damping 6-7 6.2.5 Impact ,
6-8 6.2.6 Assembly of the Dynamic Model 6-8 6.3 Stress Analysis 6-12 6.3.1 Stiffness Characteristics 6-12 6.3.2 Combined Stresses and Corner Displacements 6-13 6.4 Time Integration of the Equations of Motion 6-14 6.5 Structural Acceptance Criteria 6-17
' 6.6 Results 6-23 REFERENCES TO SECTION 6 6-32 SECTION 7 - ACCIDENTS ASSOCIATED WITH RACK INTEGRITY 7-1 AND LINER PLATE INTEGRITY 7.1 Dropped Fuel Accident I 7-1 7.2 Dropped Fuel Accident II 7-1 7.3 Jammed Fuel-Handling Equirment and 7-2 Horizontal Force 7.4 Liner Integrity Analysit, 7-2 I
7.5 Dropped Gate 7-6 i 111
TABLE OF CONTENTS (continued) Page SECTION 8 - SPENT FUEL POOL FLOOR STRUCTURAL ANALYSIS 8-1 8-1 8.1 Introduction Assumptions 8-1 8.2 8.3 Dynamic Analysis of Pool Floor Slab 8-3 to Obtain Maximum Floor Displacement 8.4 Results and Discussion 8-4 8.5 Conclusion 8-6 REFERENCES TO SECTION 8 8-12 SECTION 9 - ENVIRONMENTAL EVALUATION 9-1 9.1 Summary 9-1 19 . 2 Characteristics of Stored Fuel 9-2 9.3 Related Industry Experience 9-3 9.4 Operating Experience 9-5
- Spent Fuel Pool Cooling and. Cleanup Systems 9-5 9.5 (FPCC)
Radiological Consequences 9-7 9.6 Reracking Operation 9-7 9.7 conclusions 9-8
9.8 REFERENCES
TO SECTION 9 9-10 SECTION 10 INSERVICE SURVEILLANCE PROGRAM FOR 10-1 BORAFLEX NEUTRON ABSORBING MATERIAL 10.1 Program Intent 10-1 10.2 Description of Specimens 10-1 10.3 Test 10-2 10.4 Specimen Evaluation , ,10-2 - SECTION 11 COST / BENEFIT ASSESSMENT 11-1 11.1 Specific Needs for Spent Fuel Storage 11-1 iv O
4 TABLE OF CONTENTS (Continued) 11.2 Cost of Spent Fuel Storage 11-2 11.3 Alternatives to Spent Fuel Storage 11-2 11.4 Resource Commitments 11-4 REFERENCES TO SECTION 11 11-5 SECTION 12 OUALITY ASSURANCE PROGRAM 12-1 12.1 Introduction 12-1 12.2 General 12-1 12.3 System Highlights 12-1 12.4 Summary 12-3 Appendix I Report on Seismic Analysis of Spent Fuel Pools for High Density Spent Fuel-Racks prepared by Bechtel Power Corporation, Gaithersburg, Maryland. 9 e o 9 V
LIST OF FIGURES Page SECTION 2 Figure 2.1 Racks Arrangement in Spent Fuel Pool 2-4 2.2 Racks Arrangement in Containment Pool 2-5 SECTION 3 Figure 3.1 Array of Cells (4x4) 3-7 3.2 Elements Cross-Section 3-8 3.3 Angular Sub-Element "A" 3-9 3.4 Cruciform Element (Isometric View) 3-10 3.5 Sub-elements 3-11 (a) Angular Sub-element "B" (b) Flat Sub-element "C" - 3.6 Typical Cell Elevation 3-12 3.7 Support 3-13 3.8 Top View - Equipment Storage Rack 3-14 3.9 Elevation 3-15 , 3.10 Bottom Detail 10" Size Container 3-16 3.11 Bottom Detail 12" Size Container 3-17 3.12 Typical Module 3-18 SECTION 4 Figure 4.1 - Geometric Model of Grand Gulf Spent Fuel 4-9 Storage Rack Cell 4.2 Ak versus H 2 O Density 4-12 4.3 k. of Unpoisoned Fuel Assemblies as a 4-15 Function of Assembly Spacing 4.4 Reactivity of Spent Fuel Storage Rack as 4-17 a Function of Fuel Reactivity in Standard Reactor Core Geometry , vi
LIST OF FIGURES (continued) Pace 4.5 Log log plot of Calculated k. versus 4-19 B-10 Loading SECTION 5 Figure 5.1.1 Model for Grand Gulf Pools 5-21 5.1.2 Spent Fuel Pool Bulk Temperatures and 5-22 Power Discharged for Case A (Normal Discharge) 5.1.3 Spent Fuel Pool Bulk Temperatures and 5-23 Power Discharged for Case A (Abnormal Discharge) ' 5.1.4 . Fuel Pool Bulk Temperature Profile Using 5-24 Criteria from Case A 5.1.5 Fuel Pool Heat Load Profile Using 5-25
. Criteria from Case A 5.2.1 Rack Space Enveloping Cylinder (Grand 5-26 Gulf Unit One)
SECTION 6 Figure 6.1 Dynamic Model (lump mass) 6-33 . 6.2 Impact Springs and Fluid Dampers 6-34
' 6.3 Spring Mass Simulation for Two 6-35 Dimension Motion 6.4 (a) Horizontal Cross Section of Rack 6-36 (b) Vertical Cross Section of Rack 6.5 Dynamic Model (Rack) 6-37
'6.6 Stress Resultants Orientation
, 6-37 6.7 Subdivision of a Typical Rack 6-38 6.8 Finite Element Model Cross Section 6-39 6.9 Time History for Auxiliary Pool (N-S),SSE 6-40 6.10 Time History for Auxiliary Pool (E-W),SSE 6-41 vii
l l LIST OF FIGURES (continued) Page 6.11 Time History for Auxiliary Pool (Vertical, 6-42 SSE) 6.12 Time History Upper Containment Pool (N-S, 6-43 SSE) 6.13 Time History Upper Containment Pool (E-W, 6-44 SSE) 6.14 Time History Upper Containment Pool 6-45 , (Vertical, SSE) SECTION 8 Figure 8.1 Pool Slab Pictorial View 8-8 8.2 Node Numbers 8-9 . A 8.3 Plate El'ements 8-10 8.4 Beam Elements, Springs 8-11 SECTION 10 Figure 10.1 A Typical Coupon 10-5 Appendix I Figure 1 Auxiliary Building Lumped Mass Model 11 (East-West Direction) Figure 2 Auxiliary Building Lumped Mass Model 12 (North-South Direction) Figure 3 Ideal vs. Simplified Lunped Mass Model 13 Auxiliary Building (Partial) Figure 4 Containment Building Lumped Mass Model 14 Figure 5 Spectra of Horizontal Motion "H1" 15 (Damping 21) Figure 6 Spectra of Horizontal Motion "Hl" 16
,j (Damping 4%)
l I i ' i viii 8 I
,,c- 4 . ~ , . e - - . _ . -, - _ . , . -_ 7 ,_ . _ m,,. ..,, . _ , - - _ . - _ ,
t LIST OF FIGURES (continued) Page Figure 7 Spectra of Horizontal Motion "H1" 17 (Damping 5%)- Figure 8 Spectra of Horizontal Motio'n "H1" 18 (Damping 7%) i Figure 9 Spectra of Horizontal Motion "H2" 19 (Damping 2%) Figure 10 Spectra of Horizontal Motion "H2" 20 (Damping 4%) Figure 11 Spectra of Horizontal Motion "H2" 21 (Damping 5%) Figure 12 Spectra of Horizontal Motion "H2" 22 -
-(Damping 7% )
Figure 13 Spectra of Vertical Motion "Vertquake" 23 (Damping 2%) Figure 14 Spectra of Vertical Motion "Vertquake" 24 (Damping 4%) Figure 15 Spectra of Vertical Motion "Vertquake" 25 (Damping 5%) Figure 17 Horizontal Ground Motion Horquake 1 27 Figure 18 Horizontal Ground Motion Horquake 2 28 Figure 19 Vertical Ground Motion Vertquake 29 Figure 20 Seismic Response Time History at ell 66'-0" 30 of Auxiliary Building Spent Fuel Pool (Horizontal N-S OBE 2% Damping) Figure 21 Seismic Response Time History at EL166'-0" 31 . of Auxiliary Building Spent Fuel Pool (Horizontal N-S OBE 4% Damping) Figure 22 Seismic Response Time History at ell 66'-0" 32 of Auxiliary Building Spent Fuel Pool (Horizontal E-W OBE 2% Damping) ix 4
,e. , - - , - - - , , ~ -, , --, ... -, ,--v- , , - - , , - -
LIST OF FIGURES (continued) Page Figure 23 Seismic Response Time History at ell 66'-0" 33 of Auxiliary Building Spent Fuel Pool (Horizontal E-W OBE 4% Damping) Figure 24 Seismic Response Time History at EL166'-0" 34 of Auxiliary Building Spent Fuel Pool (Vertical OBE 2% Damping) Figure 25 Seismic Response Time History at ell 66'-0" 35 of Auxiliary Building Spent Fuel Pool (Vertical OBE 4% Damping)
- Figure 26 Seismic Response Time History at ell 66'-0" 36 of Auxiliary Building Spent Fuel Pool (Horizontal N-S SSE 5% Damping)
Figure 27 Seismic Response Time History at ell 66'-0" 37 of Auxiliary Building Spent Fuel Pool (Horizontal N-S SSE 7% Damping) Figure 28 Seismic Response Time History at ell 66'-0" 38 of Auxiliary Building Spent Fuel Pool (Horizontal E-W SSE 5% Damping) Figure 29 Seismic Response Time History at ell 66'-0" 39 of Auxiliary Building Spent Fuel Pool (Horizontal E-W SSE 7% Damping) Figure 30 Seismic Response Time History at EL166'-0" 40 of Auxiliary Building Spent Fuel Pool (Vertical SSE 5% Damping) Figure 31 Seismic Response Time History at EL166'-0" 41 of Auxiliary Building Spent Fuel Pool (Vertical SSE 7% Damping) l l
-Figure 32 Seismic Response Time History at ell 61'-10" 42 l
of Upper Containment Pool (Horizontal N-S OBE 2% Damping)' Figure 33 Seismic Response Time History at EL161'-10" 43 l of Upper Containment Pool l (Horizontal N-S OBE 4% Damping) X
LIST OF FIGURES (continued) Page Figure 34 Seismic Response Time. History at ell 61'-10" 44 of Upper Containment Pool (Horizontal E-W OBE 2% Damping) 4 Figure 35 Seismic Response Time History at ell 61'-10" 45 of Upper Containment Pool (Horizontal E-W OBE 4% Damping) Figure 36 Seismic Response Time History at ell 61'-10" 46 of Upper Containment Pool (Vertical OBE 2% Damping) Figure 37 Seismic Response Time History at EL161'-10" 47 of Upper Containment Pool (Vertical OBE 4% Damping) Figure 38 Seismic Response Time History at ell 61'-10" 48 of Upper Corstainment Pool . (Horizontal N-S 5% Damping) Figure 39 Seismic Response Time History at ell 61'-10" 49 of Upper Containment Pool (Horizontal N-S SSE 7% Damping) Figure 40 Seismic Response Time History at ell 61'-10" 50 of Upper Containment Pool (Horizontal E-W 5% Damping) - Figure 41 Seismic Response Time History at ell 61'-10" 51 of Upper Containment Pool (Horizontal E-W SSE 7% Damping) Figure 42 Seismic Response Time History at EL161'-10" 52 of Upper. Containment Pool (Vertical SSE 5% Damping) Figure 43 Seismic Response Time History at' ell 61'-10" 53 of Upper Containment Pool (Vertical SSE 7% Damping) xi
LIST OF TABLES Page SECTION 1 TABLE 1.1 - Grand Gulf Unit 1 Fuel Assembly Discharges 1-3 SECTION 2 Table 2.1 Module Data 2-2 SECTION 3 Table 3.1 Boraflex Experience for High Density Racks 3-2 SECTION 4 Table 4.1 Summary of Criticality Analysis 4-4 4.2 Fuel Assembly Design Specifications 4-8 4.3 Effect of Temperature and. Void on 4-11 Reactivity of Storage Rack SECTION 5 Table 5.1.1A List of Cases Analyzed 5-8 5.1.lB Cases Analyzed for Long Term Coolability 5-10 5.1.2A Maximum Spent Fuel Pool Bulk Temperature, 5-11 Coincident Total Power, Coincident Time and Time to Boiling (cases 1-16) 5.1.2B Maximum Spent Fuel Pool Bulk Temperature 5-12 Coincident Total Power, Coincident i Time, and Time to Boiling (cases A&B) 5.2.1 Maximum Local (Spent Fuel) Pool Water 5-18 Temperature and Local Fuel Cladding Temperature 5.2.2 Pool and Maximum Cladding Temperature 5-19 at the Instance Fuel Assembly Transfer From Upper Containment Pool to Spent Fuel Pool Begins SECTION 6 Table 6.1 Degrees of Freedom 6-5 xii I 4
m - LIST OF TABLES (Continued) 6.2 Numbering System for Springs, Gap Elements, Friction Elements 6-10 6.3 Physical Property Data 6-19 6.4 Maximum Rack Module Displacements (Damping = 2% (except where noted) 6-27 6.5 Maximum Values of Stress Factors R 1-R6 6-30 SECTION 7 Table 7.1 Support Reactions at the Instant when 7-4 x-shear is maximax (t=ll.29 seconds, Rack A, u = 0.8, fully loaded, loading Case 1. 7.2 Support Reactions at the Instant when 7-4 y-shear is Maximax. (t=10.52 seconds, Rack A, y= 0.8, fully loaded) 7.3 Support Reactions at the Time Instant 7-5 when x-shear is Maximax (t=13.61 seconds, Case 5 loading) 7.4 Support Reactions at the Time Instant when 7-5 y-shear is Maximax (t=14.74 seconds, Case 5 loading) 7.5 Liner Direct Stress 7-6 SECTION 8 Loading Data 8-5 Table =8.1 8.2 Synopsis of Acceptance Checks for Spent 8-7 Fuel Pool Floor - SECTION 10 Table'10.1 Time Schedule for Removing Coupons 10-4 xiii h . . _
1.0 INTRODUCTION
The purpose of this report is to describe the design, fabrication, and safety analysis of High Density Spent Fuel Racks produced by Joseph Oat Corporation for Grand Gulf Nuclear Station, Unit 1 (GGNS-1). GGNS-1 is a 1250 MWe generating unit containing a General Electric Nuclear Steam Supply System of the BWR-6 Design. It is owned jointly by Middle South Energy, Inc. and Southwest Mississippi Electric Power Association and operated by Mississippi Power and Light Company (MP&L). The spent fuel racks described in this report are intended to replace the existing GGNS-1 spent fuel racks in both the Upper Containment Pool and the Spent Fuel Pcol. When installed, these racks will increase the total GGNS-1 spent fuel storage capacity from the current 1440 fuel assemblies to 5148 fuel assemblies (800 in the Upper Containment Pool and 4348 in the Spent Fuel Pool). Based upon the projected GGNS-1 fuel discharge schedule presented in Table 1.1, full core discharge capability will be lost after twenty years of commercial operation with these racks installed. ,
- Joseph Oat Corporation, in conjunction with its nuclear physics consultant Southern Science (a division of Black &
Veatch), has performed detailed dynamic, seismic, thermal / hydraulic and criticality analyscs of the proposed rack design under both normal and postulated accident conditions. In all cases, the results of these analyses demonstrate that acceptable margins of safety exist with respect to appropriate NRC and ASME acceptance criteria. In addition, Middle South Energy, Inc. has performed an environmental impact evaluation and a cost-benefit comparison of several potential spent fuel disposition alternatives. The results of these analyses demonstrate that reracking of the GGNS-1 Upper Containment and Spent Fuel Pools would not present a significant impact. In addition, it is shown that re-racking is the most cost-ef fective , ! alternative and that neither the reracking operation nor the i 1-1
increased storage of irradiated material pose an increased hazard to the Plant Staff or the public. The following sections provide a synopsis of the design, fabrication, neutronics analysis, thermal / hydraulic analysis, structural analysis, accident analysis, environmental analysis and cost-benefit analysis of the High Density Spent Fuel Racks. In particular, the integrity of the spent fuel pool slab under the specified combinations of inertial, seismic, and mechanical loads and thermal gradient per NUREG-0800 is demonstrated. Also included are concise descriptions of the rack Inservice Surveillance Program and the Joseph Oat Corporation Quality Assurance program. The Joseph Oat Corporation Quality Assurance Program has been reviewed and found acceptable for engineered fabrication of ASME section III Class 1, 2, 3 ar.d MC components by both ASME and NRC. s 1-2
l Table 3.1 Grand Gulf Unit 1 Fuel Assembly Discharges Total Discharged Assemblies in Remaining Discharge Spent Fuel Pool Storage , Year Assemblies Following Refueling Capability 1986 280 280 4068 1987 240 520 3328 1988 208 728 3620 1989 228 956 3392 1990 228 1184 3164 1991 228 1412 2936 1992 228 1640 2708 1993 228 1868 2480 1994 228 2096 2252 1995 228 2324 2024 1996 228 2552 1796 1997 228 2780 1568 1998 228 3008 1340 1999 228 3236 1112 2000 228 3464 884 2001 228 3692 656 2002 228 3920 428 2003 228 4148 200 Full core disc'harge in upper containment pool. 1-3
I 2.0 GENERAL ARRANGEMENT 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 6.26" (nominal) center-to-center spacing. The high-density racks are engineered to achieve the dual objectives of maximum protection against structural loadings (such as ground motion) and the maximum utilization of available storage volume. 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 upper containment and spent fuel pools. The upper containment pool is only for interim cooling of the fuel assemblies discharged from the reactor and does not contain any fuel while the plant is operating. The fuel assemblies are transferred from this pool to the spent fuel pool for long term storage. The Grand Gulf Unit 1 spent fuel pool will contain 16 high-density fuel racks in 5 different module sizes. The module types are labelled A,B,C,D and H in Figure 2.1, which also shows their relative placement. There will be a total of 4393 storage locations in the spent fuel pool. The Grand Gulf Unit 1 upper containment pool will contain 8 high-density fuel racks in 3 different module sizes. The module types are labelled E through G in Figure 2.2, which also shows their relative placement. There will be a total of 800 storage locations in the upper containment pool. l Table 2.1 gives the detailed module data (e.g., weight, quantity, and number of storage locations). i 2-1 T
4 Table 2.1 Module Data Approximate Number of Weight Type Ouantity Pool Cells / Module Array Size lbs/ module A 12 SFP 304 19x16 32,400 B 1 SFP 256 , 16x16 27,360 C 1 SFP 216 12x18 23,160 D 1 SFP 228 12x19 24,420 E 4 UCP" 90 9x10 9,930 . F 2 UCP 99 9xil 10,880 G 2 UCP 121 llxil 13,190 H 1 SFP 45 9x5 11,300 i l { i Spent Fuel Pool (SFP) it Upper Containment Pool (UCP) 2-2
. _ . , _ _ . . _ . _ _ _ _ " T_ _
The spent fuel rack modules are free standing, i.e. they are not anchored to the pool floor or connected to the pool walls through snubbers or lateral restraints. The nominal gap between any two modules is 3 3/4 inches along the top edge for modules in the spent fuel pool and 3 15/16" for modules in the upper containment pool. The minimum gap between the spent fuel pool wall and the modules is 6 1/4 inches and 2 inches between the upper containment pool wall and the modules. Adequate clearance from other pool resident hardware is also provided. In this manner, the possibility of inter-rack impact, or rack collision with other pool hardware during the postulated ground motion events is precluded. Details on rack kinematics under seismic conditions may be found in Section 6.6. Of the 4393 storage locations in the spent fuel pool, 4348 locations are intended for spent fuel storage. Rack Type H is designated for storage of 27 control rods, 9 control rod blade guides and 9 defective fuel storage canisters. l i 2-3
1
- N 4 O'- I I/2" ;
1 61/4" - e- 119 I/8"~ ell 9 I/8" = * -1191/8" r *-100 3/8"-+ *-- 6 1/4 "
-+ +-3 3/4 " -* +- 3 3/4" + + 3 3/4" xxxxxxxxxxxxxxxxxxxxxxxxxxxxxxx at n JL s
d s Al A2 A3 BI '
-9 3/4" l00 3/8" 19 X 16 = 16 X 16 = [
s 304 CELLS 256 CELLS s s s 16 '-6"
x s d s HI s "
sxs= 45 cats s 65"REE s A4 A5 A6 " ' 100 3/8" s FUEL .R DS ' o {- 7REE 13/16" s i s n a "
<r 27 s Cl d h s 12 X 18 =
36'-l I
- 4 5/32" (TYP.) a s
A7 A8 A9 216 CELLS
\ 112 7/8 a
100 3/8 l sh - 4 9/16" i s N s s
\
N u o u d ' Di N s s 12X19 = \
\
n s 228 CELLS \ 119l/8" l00 3/8 A IO All Al2 N s s \ N u s xx x x xx x x s xx xx x x x sx xxxxxxxxxxxx\ N ,, t 4 5/32"---* - -
-93/4
---* 25" 6I/32'4
- FIA 9 I - A A Cl(R A RR o N GFFA EtJT I ?.! GPENT FilFf. POOL
= l 2 *~ 2"
-2" ,24"__ 561/2" _
-- 7 u
561/2" _ .--. 2,, I'N' 1r x x x x s Nx . x x x El E2 s 6213/id 9 X 10 = 90 CELLS s
- 1r s 3 15/16" s s 13'-O" s
n E3 E4 s 6213/I6 s s
\
jf
~
3 15/16 " n N jr FI F2 6 9 I/16" 9 X 11 =
+ +-- 21/2,' R EF.
l \ n 0 I'
- 9/16 v
u s N
\
99 CELLS 5 7/16" s a s s s GI G2 2" 691/16" s 11 X ll = 8 '- O" s 121 CELLS 2"-+- s*- o s l 2" \'
, e Nx x xN N ,,
il 3/8"---> c
691/16"
- 7" = -
i l l 15'-6"
=
FIG. 2.2- RACKS ARRANGEMENT IN CONTAINMENT POOL 2-5
- 1
i l i I l 3.0 RACK CONSTRUCTION 3.1 construction The racks are constructed from SA 240, Type 304, austenitic steel sheet material, SA 240, Type 304 austenitic steel plate material, and SA 182, Type F 304 austenitic steel forgings material. Boraflex, a patented brand name product of Bisco *
, is the neutron absorber material. The detailed radiological properties of Boraflex may be found in Section 4. The experience list of Boraflex is given in Table 3.1.
Figure 3.1 shows a horizontal cross-section of an array of 4x4 cells. As stated in the preceding section, the modules vary in array size from the smallest 9x10 array to the largest array size of 19x16. A typical module contains storage cells which have 6" nominal internal cross-sectional openings. The cells provide a smooth and continuous surface for lateral contact with the fuel assembly. The construction of the rack modules may be best 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
- Lisco, a Division of Brand, Inc., 1420 Renaissance Drive, Park Ridge, Illinois 3-1
Table 3.1 BORAFLEX EXPERIENCE FOR HIGH DENSITY RACKS Plant NRC Licensing Site Type Docket # Status Point Beach - 1 & 2 PWR 50-226 & 301 Issued Nine Mile Point --1 BWR 50-220 Issued Oconee - 1 & 2 PWR 50-269 & 270 Issued Prairie Island - 1 &2 PWR 50-262-& 306 Issued Calvert Cliffs _ 2 PWR 50-318 Issued
- Quad Cities - 1 & 2 BWR 50-254 & 265 Issued Watts Bar - 1 & 2 PWR 50-390 & 391 Pending Waterford - 3 PWR 50-382 Issued
- Fermi - 2 BWR 50-341 Issued H.-B. Robinson - 2 PWR 50-261~ Issued River Bend - 1 BWR 50-458 Pending
- Rancho Seco - 1 PWR 50-312 Issued Nine Mile Point - 2 BWR 50-410 Pending Shearon Harris --1 PWR 50-400 Pending
. Millstone - 3 PWR 50-423 Pending
- Joseph Oat Corporation fabricated racks.
3-2
I stainless sheets. The long edges of the cruciform are welded using a 3/8" thick stainless steel backing strip as shown in Figure 3.4. The bottom of the cruciform assembly has 7 7/8" high stainless strips, which ensure against slippage of the
" poison" material downwards due to gravitational loads or operating conditions. The fabrication procedure leads to one hundred percent surface contact (in a 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 3.4.
Continuous welding of the straight segments of the top edges produces a smooth lead-in surface. Ample venting is available through the roof openings of cell corners. This venting will preclude bulging of the cell walls due to gas entrapments. The " ell" and " tee" elements are constructed similarly using angular sub-elements "B", and flat sub-elements "C" (Figure 3.5). Other more complicated shapes have also been used to produce the same final " cruciform" & " tee" elements. Having fabricated the required quantities of the " cruciform", " tees", and " ells", the assembly is performed in a specially designed fixture which serves the vital function of maintaining dimensional accuracy while welding all the contiguous spokes of all elements using fillet welds. Figure 3.1 shows the fillet welds. The cells 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 baseplate. Machined sleeve elements are positioned concentric with the cell center lines above the holes drilled in the base plate, and attached to the base plate through circular fillet welds (Fig. 3.6). 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 baseplate provides the coolant flow path for heat transport from the fuel assembly cladding. Lateral holes (Nom, dia. = 3/4") in the cell walls (Figure 3.6) 3-3
provide the- redundant flow path in the unlikely event that the main coolant flow path is clogged. Each module is supported on four " plate-type" supports; a sketch of a typical support is shown in Figure 3.7. The defective cell container module (module H1, Figure 2.1) employs the same structural assemblage concept, although the physical details are different. Figure 3.8 shows a horizontal cross section of this module. The individual storage locations are composed of 10.5" O.D. x 0.125" wall tubing, except for 9 containers which are 12" schedule 20S pipe (12.75" O.D. x 0.25"
. wall). The latter type of containers is intended for storing.
control rod blade guides. These are located at the short sides of the defective cell rack. The cylindrical containers are - joined to each other near the top using specially shaped stainless steel stampings and at the mid-span by welding. A suitably machined baseplate provides the bottom support for the module. The construction of the cell assembly is an integral structure. which possesses extremely high flexural rigidity.
.Thus, the response of the module to seismic excitation simulates that of the Type A modules.
The defective fuel storage rack module is 150" high, thus, a portion of the defective fuel container will be above the module. Redundant flow paths (2 holes 1" diameter in each cell) are provided to alleviate concerns regarding main flow path clogging. Further construction details may be found in Figures 3.8 to 3.12. 3-4
s 3.2 CODES, STANDARDS, AND PRACTICES FOR THE SPENT FUEL POOL l MODIFICATION i 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, 8th edition (1980) (b) ANSI N210-1976, Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations. (c) American Society of Mechanical Engineers (ASME), Boiler & Pressure Vessel Code, Section III, 1980 Edition up to and including Winter 1981 addenda. (Subsection NF) Note: Where the weld stresses are less than 20% of yield, the vendor visual inspection criteria will apply in lieu of NF visual inspection criteria. (d) ASNT-TC-1A June, 1980, American Society for Nondestructive Testing (Recommended Practice for Personnel Qualifications) (e) ANSI N412-1975, The Determination of neutron reaction rate distributions and reactivity of nuclear reactors. II. Material Codes (a) American Society for Testing and Materials (ASTM) Standards - A240 & A262, Practice E.
(b) American Society of Mechanical Engineers (ASME), Boiler & Pressure Vessel Code, Section II, Subsection NA, 1980 Edition up to and including Winter 1981 addenda. III. Welding Codes (a) ASME Boiler and Pressure Vessel Code, Section IX- Welding and Brazing Qualifications, 1980 Edition up to and including Winter is31 addenda. IV. Quality Assurance, Cleanliness, Packaging, Shipping, Receiving, Storage, and Handling Requirements (a) ANSI 45.2.2, Packaging, Shipping, Receiving, - Storage and Handling of, Items for Nuclear Power Plants. (b) ANSI 45.2.1, Cleaning of Fluid Systems and Associated Components During Construction. (c) ASME Boiler and Pressure Vessel, Section V, Non-destructive Examination,' 1980 Edition, including Winter 1981. V. Other References (a) NRC Regulatory Guides, Division 1, Regulatory Guides 1.13, 1.29, 1.60, 1.61, 1.71, 1.85, 1.92, and 1.124 (revisions as applicable). (b) General Design Criteria for Nuclear Power Plants, Code of Federal Regulations, Title 10, Part 50, Appendix A (GDC Nos. 1, 2, 61, 62, and 63). (c) NUREG-0800, Standard Review Plan (1981). (d) "NRC Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," dated April 14, 1978, and the modifications to this document of January 18, 1979. 3-6
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- 4. NEUTRONIC CONSIDERATIONS 4.1 Introduction The spent fuel storage racks are designed to assure that a ke gg equal to or less than 0.95 is maintained with the racks fully loaded with fuel of the highest anticipated reactivity and flooded with unborated water at a temperature corresponding to the highest reactivity. The maximum calculated reactivity in-cludes a margin for uncertainty in reactivity calculations and in mechanical tolerances, statistically combined, such that the true kegg will be equal to or less than 0.95 with a 95% probability at a 95% confidence level.
Applicable codes, standards and regulations or pertinent sections thereof include the following: e General Design Criterion 62 - Prevention of Criticality in Fuel Storage and Handling. e NRC letter of April 14, 1978, to all Power Reactor Licensees - OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, includ-ing modification letter dated January 18, 1979. e USNRC Standard Review Plan, NUREG-0800, Section 9.1.2, Spent Fuel Storage. e Regulatory Guide 1.13, Spent Fuel Storage Facility Design Basis (Draft, Revision 2), December 1981. 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
= = ,--
l 1 The design basis fuel assembly is a standard 8 x 8 array of fuel rods (BWR type) containing 00 2 clad in Zircaloy. Fixed neutron absorbing material is used in the spent fuel storage racks ' to maintain the required margin in suberiticality. The . spent fuel racks are designed to safely. accommodate fuel assem-blies with a maximum infinite multiplication factor (k,) of 1.395, as calculated for the standard core configuration in the cold, unrodded condition. To assure the true reactivity will always be_-less than the calculated reactivity, the following conservative assumptions were made. e Moderator is pure, unborated water at a temperature corresponding to the highest reactivity, e Lattice of storage racks is infinite in all directions; i.e., no credit is taken for axial or radial neutron leakage, except in the evaluation of axial cutback and for certain abnormal conditions where neutron leakage is inherent. e No credit is taken for the presence of gadolinium burnable poison or for the reduction in reactivity that accompanies fuel burnup. e Neutron absorption in minor structural members is neglected; i.e., spacers and Inconel springs are re-placed by water. o Pure zirconium is used for cladding and flow channel; i.e. , higher neutron absorption of alloying materials in Zircaloy is neglected. 4.1.1 Neutron Multiplication Factor The nominal design case assumes fuel of uniformly distributed 3.5 wtt U-235 enrichment, corresponding to 16.49 grams U-235 per axial centimeter of fuel assembly. Fixed neutron absorber material (Boraflex) of 0.02041 g/cm 2 boron-10 areal density is positioned between fuel assemblies in an egg-crate 4-2
structure that provides a nominal center-to-center lattice spacing of 6.2585 inches for the storage cell locations. The maximum infinite multiplication factor ( k,, ) calculated -for the nominal design case is 0.936 including all uncertainties (95% probability at a 95% confidence level) for fuel of 3.5% uniform enrichment ( k, of fuel ' assembly in standard core geomet."
- of 1.3798 i 0.0037).
With the configuration described above, the spent fuel storage rack can safely accommodate fuel assemblies whose maximum ' infinite multiplication factor in the standard reactor core geom-etry (cold conditions) is 1.395, without exceeding the limiting design criterion for spent fuel storage racks (maximum k,, of 0.95 including all uncertainties). The calculations and uncerte'nties, supporting the criticality safety of the spent fuel storage racks for Grand Gulf Nuclear Station Unit 1, are summarized in Table 4-1, and de-scribed in Section 4.4, Criticality Analysis. 4.1.2 Analytical Methods The reference method for nuclear criticality analyses of the high density spent fuel storage rack is the AMPX -KENO 1 2 computer package, using the 123-group GAM-THERMOS cross-section set and the NITAWL subroutine for U-238 resonance shielding effects (Nordheim integral treatment). AMPX-KENO . has been extensively benchmarked against a number of critical experiments (e.g., Refs. 3, 4, 5 and 6). I In the standard reactor core geometry, fuel assemblies at 39.6*F are located on a 6.00 inch center-to-center spacing, surrounded by full density (p=1.0) water, with all control blades removed and with no = credit for the presence of gadolinium burnable poison. 4-3
----___.m__
Table 4-1
SUMMARY
OF CRITICALITY' ANALYSIS Nominal Design Maximum Case Storage Capabilit: k, in standard core geometry 1.3798 0.0037 (la) 1.395 k, in spent fuel storage rack 0.9216 0.935 Calculational bias, Ak 0.0036 0.0036 Uncertainties and tolerances Calculational bias 10. 0028 A k Calculation (statistical) 0.0048 A k Boraflex thickness 70.0059 Ak B-10 concentration 70.0027 Ak Boraflex width 7 0. 0016 A k Fuel enrichment i0.0032 A k Fuel density i0.0024 A k Pellet diameter negligible Lattice pitch 7 0.0040 Ak SS thickness i0.00ll A k Flow channel bulge 0.0047 Ak
*0.0114 Statistical combination 10.0114 i0.0114 Maximum k, 0.937 0.950 Abnormal / accident condition Ak Temperature increase ' negative A k Boiling negative Ak Reduced moderator density negative Ak Fuel assembly positioning negative Ak Assembly outside rack negligible A k Dropped fuel assembly negligible A k 4-4 I &
For two-dimensional X-Y _ analysis, a zero current (re- l flecting) boundary condition was applied in the axial direction and at the centerline through the Boraflex absorber on all four sides of the cell, effectively creating an infinite array of
~
storage cells-for analytical purposes. For investigation of small reactivity effects (e.g., mechanical tolerances), a four-group diffusion / blackness theory method of analysis was used (Ref. 5) to calculate small incre-mental- reactivity changes. This model has been used previously with good results and is normally used only to evaluate trends and small incremental reactivity ef fects that would otherwise be lost in the KENO statistical variation. Where possible, trends calculated by AMPX-KENO and by diffusion / blackness theory were compared and found to be in good agreement, within the statis-tical uncertainty of KENO calculations. 4.1.3 Calculational Bias and Uncertainty Results of benchmark calculations 6 on a series of critical experiments indicate a calculational bias of 0, with an uncer-tainty of 0.0028 (95% probability at a 95% confidence level). In addition, a small correction in the calculational bias is neces-sary to account for the slightly greater gap thickness (1.1 inches) between fuel assemblies in the Grand Gulf spent fuel rack compared to the corresponding thickness (0.644 inch) in the benchmark critical experiments. Based upon the correlation de-veloped in Ref. 6, the correction for water-gap thickness in the Grand Gulf spent fuel storage rack indicates a small underpredic-tion of 0.0036 ak. Thus, the net calculational bias is taken as 0.0036 i 0.0028, including the effect of the water-gap thickness. 4-5
l s l 4.1.4 Trend Analysis Trend analysis 6 of benchmark calculations on critical experiments with varying boron content in the absorber plate between fuel assemblies indicates a tendency to overpredict k egg with higher reactivity worth of the boron absorber. In the Grand Gulf spent fuel rack, the boron worth is about 40% Ak, or ~ 2. 7 times the highest boron worth (15. 9% A k) in the critical experi-ments analyzed in Ref. 6. Based upon extrapolation of the trend analysis, AMPX-KENO calculations of the Grand Gulf rack would be expected to overpredict k, by an estimated 3.1% A k, including allowance for water-gap thickness. Statistically combining the standard deviation of the regression analysis 6( 0.003,la) and a typical standard deviation of the KENO variation of the mean (i0.003,lo), the maximum uncertainty would be 0.008, including a one-sided tolerance factor 7 of 1.92 (95% probability at a 95% confidence level) for 100 generations in a KENO calculation. Thus, to the extent extrapolation of the linear regression anal-ysis is valid, the AMPX-KENO calculation of the Grand Gulf rack will be high (overprediction) by 0. 031 i 0. 00 8 A k , or a minimum overprediction of 0.023 A k 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 calculations would be expected to over-predict keff when strong boron absorbers are present. No credit is taken for the expected overprediction other than to indicate l I an additional level of conservatism in the criticality analysis of 'the Grand Gulf spent fuel storage rack. l e 4-6
4.2 Input Paramaters 4.2.1 Fuel Assumbly Design Specifications Design specifications for the fuel assembly as used in the criticality analysis are given in Table 4-2. Assemblies contain-ing fuel rods of slightly different dimensions or configuration may be safely accommodated in the high density spent fuel storage racks, provided~the assembly k, in the standard core-geometry is within the limit established in the criticality analysis. 4.2.2 Reference Design Spent Fuel Storage Cell The nominal spent fuel storage cell model used in the criti-cality analyses is shown in Fig. 4.1. The rack is composed of Boraflex absorber material sandwiched between 0.063-inch stain-less-steel plates. The fuel assemblies are centrally located in each storage cell on a nominal lattice spacing of 6.2585 inches. The Boraflex absorber has a nominal thickness of 0.070 inch and a-nominal B-10 areal density of 0.02041 grams B-10 per square centimeter. l f f 4-7
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' Cladding thickness, in. 0.035 Cladding material Zr Pellet density, g UO 2 /cc 10.367 1 0.165 Pellet diameter, in. 0.4055 Enrichment, wt% U-235 3.50* 0.05 Grams U-235 per axial centimeter 16.52 Water Rod Data Outside diameter, in. 0.591 Wall thickness 0.030 Material Zr-2 Number per assembly 2 Fuel Assembly Data Number of fuel rods 62 Fuel rod pitch, in. 0.636
! Fuel channel outside dimension, in. 5.458 0.0065 Fuel channel wall thickness, in. 0.120 Fuel channel material Zr-4 o
- Nominal design case 4-8
B O R AFLE X H A L F-THICK NESS. 5.705* WIDE 0.591* O.D., 0.531' l. D. WATER ROD 0.484* O.D., 0.414
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4.3 Postulated' Accidents and Abnormal Conditions 4.3.1 Temperature and Water Density Effects The nominal criticality analyses were performed for the maximum water density (p =1.0) corresponding to a temperature of ~39'F. As shown in Table 4-3, increasing temperature or introducing void (to simulate boiling) decreased reactivity of the spent fuel storage rack. Since the storage rack may also accommodate the initial core loading in the dry condition, calculations were made for modera-tion by water of reduced density to investigate the reactivity effect of hypothetical moderation (foam or spray) of the dry storage cells in conformance with the requirements of SRP 9.1.1. As indicated in Table 4-3, the maximum reactivity occurs for the fully flooded case (reference design configura-tion) and the reactivity is always substantially lower than the reference k,, for all other moderating conditions. As shown in Fig. 4.2, the calculated k, decreases continuously down to 10% moderator density, with no suggestion of a reversal in slope or the appearance of a second maximum in reactivity. These results are consistent with those reported by Cano et al.,8 who showed, in a parametric study, that the phenomenon of a second maximum in reactivity does not occur in a closely-spaced lattice with a strong neutron absorber present. Thus the spent fuel storage rack will accommodate the initial core loading in the dry condi-tion well within the requirements of SRP 9.1.1. , ' 1 4-10
-r
i Table 4-3 EFFECT OF TEMPERATURE AND VOID ON REACTIVITY OF STORAGE RACK Case Ak= Comment 39'F (~ 4'C) Reference Maximum water density 68'F (20'C) -0.0006 p (H O) = 0.998 2 104'F (40*C) -0.005 p(H 2O) = 0.992 176*F (80*C) -0.015 p (H2 O) = 0.972 212*F (100*C) -0.020 p (H2 O) = 0.958 212'F with 50% void -0.184 Simulates boiling 39'F (4*C), 50% density -0.166 Simulates hypothetical 39'F (4*C), 40% density -0.219 - reduced density moderation 39'F (4*C), 30% density -0.276 .of cells containing 39'F (4*C), 20% density -0.340 an infinite array of 39'F (4*C), 10% density -0.405 fuel assemblies of the nominal enrichment. 4-11 wa.... . , . . . . . . . . . . . . . . . . . _... . .N
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g..., 1.0 0.9 0.8 0.7 0.6 0.5 0.4 0.3 0.2 0. WATER MODERATOR DENSITY Fig. 4.2 Ak versus H O density. 2 4-12
4.3.2 Abnormal Positioning of Fuel Assembly ) Outside Storage Rack Since the storage rack criticality calculations were made assuming an infinite array of storage cells with no neutron leakage, positioning a fuel assembly outside and adjacent to the actual finite rack cannot add reactivity, but would, because of neutron leakage, result in a lower k,gg than the k, calculated for the infinite array. This has been confirmed by two-dimensional PDQ analysis of finite racks with a fresh fuel ele-ment positioned outside the rack. 4.3.3 Fuel Assembly Positioning in Storage Rack The fuel assembly is normally located in the center of the storage rack cell with bottom fittings and spacers that mechan-ically prevent lateral movement of the fuel assemblies. Never-theless, 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 reactivity was slightly less than the nominal design case (by 0.010 ak). Calcu-lations were also made with the fuel assembly moved into the corner of the storage rack cell (four-assembly cluster at closest approach), resulting in an even larger negative reactivity effect (calculated decrease in k, of 0.015 ak). With the zirconium fuel channel removed, the reactivity effect of off-set fuel assemblies is even more negative. Thus, the nominal case, with the fuel assembly positioned in the center of the storage rack cell, yields the maximum reactivity. e f 4-13 _ _ - . _ _ _ _ ______ . , m
c , 4,3.4 Effect_of Zirconium Fuel Channel Distortion Consequences of bulging of the zirconium fuel channel are treated as a mechanical deviation in Section 4.4.9 below. Bowing i of the zirconium channel (including fuel rods) results in a local' negative reactivity effect analogous to that of positioning the fuel asembly toward one side of the storage cell, as described in Section~4.3.3 above. Thus, bowing will result in a reduction in reactivity. 4.3.5 Dropped' Fuel Assembly Accident To investigate the possible reactivity effect of postulated fuel drop accidents, calculations were made for unpoisoned assem-blies separated only by water. Figure 4.3 shows the results of these calculations. From these data, the reactivity (k,) will be less than 0.95 for any spacing between unpoisoned assemblies greater than ~8 inches. For a dropped fuel assembly lying horizontally on top of the rack, the minimum separation distance is ~14 inches. Maximum expected deformation under seismic or accident conditions (see Section 6) will not reduce the minimum spacing to less than 8 inches. Thus, a dropped fuel assembly will not constitute a criticality hazard, and the storage rack infinite multiplication factor will not be materially altered. 4.3.6 Fuel Rack Lateral Movement Normally, the individual rack modules in the spent fuel pool are separated by a water-gap of several inches. For finite fuel racks, this separation would reduce the actual maximum reactivity of the racks. Should lateral motion of a fuel rack occur, closing the gap between racks (for whatever reason), the reactivity would, in the limit, only approach the limiting reactivity of the reference infinite array. 4-14
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8 7 8 9 10 11 12 L ATTIC E SP A CING. INCHES i Fig. 4.3 k co of unpoisoned fuel assemblies as a function of assembly spacing. 4-15
4.4 Criticality Analysis 4.4.1 Nominal Case Under normal conditions, with nominal dimensions, the calculated k,, is 0.9216 i 0.0026 (lo with 150 generations) with fuel of 3.5% uniform enrichment. For a one-sided tolerance factor of 1.868, corresponding to 95% probability at a 95% confidence limit for 150 generations, the maximum deviation of k,, is 10.0048. With the calculational bias and all uncertainties added, the reactivity (k,,) of the storage racks will always be less than 0.937 with 95% probability at a 95% confidence level. Calculations with fuel of distributed enrichments (average 3.5% enrichment), representative of BWR-type fuel assemblies, showed a lower reactivity than the corresponding calculation with uniform enrichment. 4.4.2 Maximum Reactivity Storage Capability For uniform enrichment of 3.5 wt% U-235 ( k,, of 1.3798 in standard core geometry) and the selected Boraflex loading (0.02041 g B-10/cm 2 nominal, 0.0175 g B-10/cm 2 minimum), there is a margin of 0.013 Ak between the limiting reactivity (k,, of 0.95) and the calculated reactivity including all uncertainties. Addi-tional calculations were performed to determine the maximum re-activity fuel which the racks could safely accommodate without exceeding the limiting reactivity. Figure 4.4 shows the calcu- ! lated relationship between the fuel k,, (standard core geometry) and the infinite multiplication factor of the spent fuel storage
- racks. These calculations indicate that fuel with a k,, of 1.395 (reactor core geometry) can be safely accommodated and that the maximum k,, of the storage rack will not exceed 0.950, including all uncertainties, with 95% probability at a 95% confidence level. A fuel k ,, o f 1.395 in the standard core geometry corresponds to a uniform fuel enrichment of ~3.7 wtt U-235.
4-16 e 2--
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0*00 ~ 5E [:: - 7:= - :4 = = i rrie _ __.2=r jn= ==4nu :iE:I f ___ . .ar - _ _ _ ._-~- '~ ~_' ~ _ _ _ . . ..--n.... i. . . . . . . _ _:t : --n :- _ . _. _ . . _ . . .. --. . J- --:--n_.,.._.._...___=---
...._._......_..,___.,gn_...-y.=....--__.-_.g._..3.-l=_=-7 -n =u _ - - - - - eu _ =
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--------~"'_-------"--'---"--'---'.__._rn.p~' -= r
-('=-~:="- =~= - ' - -::n - " = =:
0.89 1.34 1.36 1.36 1.37 1.38 1.39 1.40 1.41 k, OF FUEL IN STANDARD CORE GEOMETRY Fig. 4.4 Reactivity of spent fuel storage rack as a function of fuel reactivity in standard reactor core geometry, t 4-17 L
In practice, neutron leakage and a higher moderator temperature will reduce the actual storage rack reactivity below the values indicated above. In addition, the presence of gadolinium poison will further reduce the actual rack reactivity by an estimated 0.17 ok for fresh fuel and by 0.04 Ak for fuel at the point of maximum reactivity in fuel burnup (when the gadolinium has been essentially consumed). l 4.4.3 Boron Loading Variation The Boraflex absorber plate is nominally 0.070 inch thick with a B-10 areal density of 0.02041 g/cm 2. Manufacturing tolerance limits are +0.007 inch in thickness and 0.001 g/cm 2 in boron content. This assures that, at any point where the minimum boron loading (0.01944 g B-10/cm 2 ) and minimum Boraflex thickness (0.063 inch) may coincide, the boron areal density will not be less than 0.0175 g B-10/cm2, Calculations were made of k. with variations in Boraflex absorber loading and thickness. Results of these calculations, shown in Fig. 4.5, indicate that the k can be described by the following regression fit (least squares) to the data over the range of B-10 loading from 0.010 to 0.030 g/cm 2 , 2 k,,= 0.7280 EXP (-0.0606 in B-10, g/cm ) This relationship indicates that the tolerance limit on boron concentration and Boraflex thickness results in incremental reactivity changes of 10.0027 ok and 0.0059 ok, respectively. The trend calculated both by AMPX-KENO and by diffusion / blackness theory is the same within analytical uncertainty. s 4-18
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4.4.4 Boraflex Width Tolerance Variation A decrease in Boraflex plate width increases reactivity. For the manufacturing tolerance limit of i0.0623 inch on the width of the Boraflex absorber plate, the correspanding uncer-tainty in reactivity is 70.0016 Ak, calculated by diffu-sion/ blackness theory since the reactivity increment is too small to be calculated by AMPX-KENO. 4.4.5 Axial Cutback in Boraflex Length The axial length of the Boraflex absorber is less than the active fuel length by 3 inches at both the top and bottom of the fuel racks. This axial cutback occurs in the region of high neutron leakage. Axial calculations (1-dimensional) indicate that the incremental reduction in reactivity due to axial leakage (-0.0023 ak) is greater than the increase in reactivity due to the axial cutback (+0 .0013 A k) . Thus, the infinite multiplica-tion factor (k,,) used as the reference reactivity is the more conservative condition. The calculations were performed with diffusion / blackness theory since the incremental reactivity effects are too small to be determined by AMPX-KENO calculations. 4.4.6 Storage Cell Lattice Pitch Variation The design storage cell lattice spacing between fuel assemblies is 6.2585 i 0.0625 inches. Increasing the lattice pitch reduces reactivity. For the manufacturing tolerance of 10.0625 inch, the corresponding maximum uncertainty in reac-tivity is 70.0040 Ak, calculated by diffusion / blackness theory, since the reactivity increment is too small to be calculated by
~
AMPX-KENO. 1 4-20 mr
4.4.7 Stainless. Steel Thickness variations The nominal. stainless-steel thickness is 0.063
- 0.005 inch. The maximum . reactivity effect of the expected stainless-steel thickness tolerance variation (0.005 inch) was calculated to be '*0.0011 A k by diffusion / blackness theory, since the - reac-tivity increment is too small-to be calculated by AMPX-KENO.
4.4.8 Fuel Enrichment and Density Variation The nominal design enrichment is 3.5 wtt. U-235. Calculations of the sensitivity to small enrichment variations-by diffusion / blackness theory yielded an average coefficient of 0.0065 Ak per 0.1 wt% U-235. For an estimated tolerance on U-235 enrichment of i0.05%, the maximum uncertainty is 10.0032 A k due '
-to-the tolerance on fuel enrichment.
Calculations were also made to determine the sensitivity to tolerances in UO2 fuel density. These calculations indicate . that the storage rack k, is increased by 0.0024 Ak (diffusion / black-ness theory) for the expected tolerance ( 1.5% in ,T. D. ) in fuel density. A lower fuel density results in correspondingly lower
. values of reactivity. Thus, the maximum uncertainty due to the tolerances on 00 2 density is 10. 0024 A k.
4.4.9 Effect of Zirconium Fuel Channel i-Elimination of the zirconium fuel channel results in a small j decrease in reactivity (-0.0038 A k) as calculated by diffu- ! sion/ blackness theory. More significant is a small positive reactivity ef fect resulting from bulging of the zirconium chan- [' nel, which moves the channel wall outward toward the Boraflex
, Labsorber. For the maximum expected bulging (to 5.93 inches out-side dimension) uniformly throughout the assembly, an incremental reac'tivity of +0.0047 Ak would result, as calculated by 4-21 l
h L.
- diffusion / blackness theory. Since actual bulging of the flow channel would not be the maximum everywhere in all assemblies, the reactivity effect has been 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 4.3.4 above). 4 4-22
-~
4.5 Acceptance Criteria for Criticality The USNRC letter of April 14, 19 7 8 , . to all Power Reactor Licensees, and the draft revision to Regulatory Guide 1.13 specify that the neutron multiplication factor in spent fuel pools shall be less than or equal to 0.95, including all uncer-tainties, when fully loaded with fuel of the highest anticipated reactivity. For BWR type fuel with an infinite multiplication factor of 1.395 or less in the standard reactor core geometry, the spent fuel storage rack described herein satisfies this criterion. i 4-23
a I REFERENCES
- 1. Green, Lucious, Petrie, Ford, White, Wright, PSR-63/AMPX-1 (code package), AMPX Modular Code System for Generating Coupled Multigroup 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 UO Rods in Water with Fixed Neutron Poisons, NUREG/CR-0073 , 2Battelle Pacific Northwest Laboratories, May 1978, with errata sheet issued by the USNRC August 14, 1979.
- 4. M. N. Baldwin et al., Critica. Experiments Supporting Clone Proximity Water Storage of Power Reactor Fuel, BAW-1484-7, The Babcock & Wilcox Company, July 1979.
- 5. S. E. Turner and M. K. Gurley, Benchmark Calculations for Spent Fuel Storage Racks, Report SSA-127 (Rev. 4), Southern Science Applications, Inc., May 1981.
- 6. S. E. Turner and M. K. Gurley, Evaluation of AMPX-KENO Benchmark Calculations for High Density Spent Fuel Storage Racks, Nuclear Science and Engineering, Vol. 80, No. 2, pp.230-237, February 1982.
- 7. M. G. Natrella, Experimental Statistics, National Bureau of Standards, Handbook 91, August 1963.
- 8. J. N. Cano et al., Supercriticality Through Optimum Moderation in Nuclear Fuel Storage, Nuclear Technology, Vol. 48, pp.252-260, May 1980.
4-24
p ?
- 5. THERMAL-HYDRAULIC CONSIDERATIONS A central objective in the design 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 are given. Similar
- analysis has been used in previous licensing reports on high density fuel racks for Fermi II (Docket 50-341) and Quad Cities (Docket 50-254/265).
5.1 Forced Circulation Thermal-Hydraulic Analysis This report section covers requirement III.l.5(2) of the y NRC "OT Position for Review and Acceptance of Spent -Fuel Storage g[- and Handling Applications" issued on April 14, 1978. This requirement states that calculations for the amount of thermal energy removed by the spent fuel cooling system shall be made in accordance with Branch Technical Position ASB 9-2 " Residual Decay . Energy for Light Water Reactors for~ Long Term Cooling"1 The 1 calculations contained herein have been made in accordance with this requirement.
5.1.1 Basis
The Grand Gulf Unit 1 reactor is rated at 3833 Megawatt-Thermal (MWT). The core contains 800 fuel assemblies. . Thus, the average operating power per fuel assembly, Po , is 4.791 - MW. The fuel assemblies are removed from the reactor after the maximum burn-up of 30,000 Megawatt-days per short ton of uranium (MWD /STU). - - The Grand Gulf installation has two pools 2 equipped for fuel storage. The pool adjacent to the reactor cavity, hereinafter referred to as the " upper containment pool" has the capacity to accept the entire reactor core (800- fuel assemblies). This pool directly communicates Vith the reactor cavity and is flooded together with the reactor cavity prior to 5-1
initiation o' fuel transfer. The second pool, hereinafter referred to as the " spent fuel pool" is located in the Auxiliary Building. Fuel assemblies are transferred to this pool from the containment pool via the Fuel Transfer Canal for long term storage. The upper containment pool does not contain any fuel while the plant is operating. The cooling systems for the two pools are somewhat intertwined. Figure 5.1.1 shows the flow schematic for the two pools. The analytical characterization of this system for thermal analysis purposes is developed later in this section. The fuel discharge can be made in one of the following two modes: (i) Normal discharge - Mode (i) (ii) Full Core discharge - Mode (ii) As shown in Table 1.1 of Section 1, the equilibrium fuel assembly removal batch size for Mode (i) is 228 fuel assemblies. Mode (ii) corresponds to a full core discharge (800 assemblies). Figure 5.1.1 shows the model for the pool cooling system. The heat dissipation from the spent fuel pool is
- accomplished by two independent fuel pool cooler loops, each equipped with a pump rated at 1100 gpm. In addition, two Residual Heat Removal (RHR) heat exchangers, available for supplemental cooling, may be used in conjunction with the fuel pool coolers to boost the heat removal rate. Only one RHR heat exchanger is assumed to be available for supplemental cooling with a pump rated at 7450 gpm for this task. The RHR flow rate in spent fuel pool cooling mode was calculated to be 2550 gpm the other RHR acts as a stand-by unit.
In the following, all relevant performance data for the spent fuel pool and RHR heat exchangers is giv.en. 5-2
.s ,
. y
- a. Spent Fuel Pool Heat Exchanger:3 TEMA type 26-252 CEU 2087 sq. ft. effective surface on 253 U-tubes, 3/4" diameter x 18 BWG arranged on 0.9375" triangular pitch. Postulated fouling for tube and shellside surfaces is .0005 and 0.0005 sq.ft *F-Hr/ BTU reqpectively. Shellside pool water and tubeside cooling water flow rates are 550,000 and 532,500 lbs/hr. respectively. The temperature efficiency is 0.5.
- b. RHR Heat Exchanger:2 The effective surface is 21250 sq. ft. Postulated fouling for the shellside (po,ol water) of the tube
~
surface is 0.0005 sq.ft *F-hr/ BTU and that for the tubeside (cooling water) is .002 sq.ft. *F-Hr/ BTU. The shellside and tubeside design flow rates are 3.725 x 10 6 lbs/hr. and 3.95 x 10 6 lbs/hr. respectively. The overall heat transfer coefficient is 210 BTU /hr-ft2 _ F. The temperature efficiency is 0.455. The above data enables complete characterization of the thermal performance of the heat exchangers. 5.1.2 Model Description Reference (1) is utilized to compute the heat dissipation requirements in the pool. The total decay power consists of
" fission products decay" and " heavy elements decay". Total decay power P for a fuel assembly is given as a linear function of Po and an exponentional function of to and ts' i.e.: P = Po f (to,t,) (5.1.1) 5-3
^ W
where i P = linear function of Po Po = average operating power per fuel assembly to = cumulative exposure time of the fuel assembly in the reactor t, = time elapsed since reactor shutdown The uncertainty factor K, which occurs in the functional 7 relationship f (to,ts) is' set equal to 0.1 for ts > 10 sec in the interest of conservatism. Furthermore, the operating power Po is taken equal to the rated power, even though the reactor may be operating at a fraction of its rated power during most of the period of exposure of the batch-of fuel assemblies. The computations and results reported here are based on the discharges at the 13th loading cycle. This approach assures that an asymptotic peak temperature is attained, i.e. highest peak temperature, and the fuel thermal inventory in the spent fuel pool is at its maximum. The reactor operating time corresponding to 30,000 MWD /STU is less than 3.5 years at the nominal power level. However, since the decay power is a weakly increasing monotonic function 8 of the operating time to, we select to = 4 years (to = 1.26 x 10 secs) in evaluating the decay power. 5-4
- gr.
1 I Having determined the heat dissipation rate, the next task is to evaluate the time temperature history of the pool water. Table 5.1.1A identifies the loading cases examined. The pool bulk . 1 temperature time history is determined using the first law of thermodynamics (conservation of heat). The system to be analyzed is shown in Figure 5.1.1. A number of simplifying assumptions are made which render the analysis conservative. The principal ones are:
- 1. The cooling water temperature in the fuel pool heat exchanger (95'F) and the RHR heat exchangers (90*F) are based on the maximum postulated values given in the FSAR.
- 2. The heat exchangers are assumed to have maximum fouling. Thus, the temperature effectiveness, P, for the heat exchangers utilized in the analysis are the lowest postulated values: Pi =
0.50 for fuel pool coolers (i=1), 0.455 for RHR heat exchangers (i=2). Pi is calculated from FSAR and heat exchanger technical data sheets.
- 3. No heat loss is assumed to take place through the concrete floor.
- 4. No credit is taken for the improvement in the film coefficients of the heat exchangers as the operating temperature rises. Thus, the film coefficient used in the computations are lower bounds.
- 5. No credit is taken for heat loss due to evaporation of the pool water.
6 In the model return flow for the RHR system (in the supplemental cooling mode) from the UCP is through the reactor vessel and the re-circulation system. 5-5 7
This is depicted by W22 in Fig. 5.1.1. However, should this path not be available , flow arrangement of the RHR system would require mixing of the stream W22 (Fig. 5.1.1) with the Spent Fuel Pool water inventory. This represents communication between the two pools through the fuel transfer tube. Since the peak bulk temperature of the water in the UCP is lower than that in the SFP, this assumption has the net effect of over predicting the peak value of the SFP bulk temperature. Thus, the temperature plots presented in Section 5.2.3 are upper bounds. The basic energy conservation relationship for each pool and heat exchanger system has the form: C =O IN -000T (5.1.2) where C: Thermal capacity of stored water .in the pool. t: Temperature of pool water at time, t OIN: Heat generation rate due to stored fuel . assemblies in the pool. O IN is a known function of time T, from the preceding section. 000T: Heat removed via the heat exchangers. The up'per containment pool and the spent fuel pool in the Grand Gulf Unit I installation have total water inventory of 11844 and 51043 cubic feet respectively when all racks are in place in the pools and every storage location is occupied. The calculations were performed accounting for the interaction between the upper containment pool and the spent fuel pool (see Figure 5.1.1). . Let 02 and 02 denote the rate of heat generation in Pool 1 (spent fuel pool) and Pool 2 (upper containment pool) respectively. 04 and 02 are specified functions of time, T. 5-6
l l Furthermore, let Pt and P2 represent the " temperature ! efficiency"" of spent fuel pool heat exchangers and RHR heat exchangers, respectively. Referring to Figure 5.1.1: p3 = t 3 -tg = 0.50 t 3 -01 p , ts-ts = 0.455 t 5 -02 Therefore eq. (5.1.1) for each pool is: C2 1
=Oi+W ii (tu-t i) +W21 (ts-t i) (5.1.4) dT C2 = O2 + W12 (tg-t2) +W22(ts-t 2) (5.1.5) dT where ci and C2 are heat capacities of pools 1 and 2, respectively, and Wij (see figure 5.1.1) represents the thermal flow rate from heat exchanger i into pool j (RHR heat exchanger i=2, Fuel Pool Cooler, i=1).
Equations (5.1.4) - (5.1.5) can be cast into the form dt 3
=
[ by ) t) ; i = 1,2 (5.1.6)
- dT 3=1 where t3 in thd RHR has the implied value of unity. The matrix , coefficients bi j are functions of Wij, Pi, Ci, etc.
Runge-Kutta first order forward integration scheme is used to solve the two simultaneous first order differential equations 4 (-5.1. 6 ) . 5.1.3 Results and Discussion Sixteen typical scenarios (cases) are chosen for the first phase of .the analysis. These cases and their underlying l assumptions are presented in Table 5.1.lA. The results of the i 5-7
1 1 TABLE 5.1.lA LIST OF CASES ANALYZED Case No. Assumptions and Criteria I II III 1 (a) (b) (a) 2 (b) (b) (a) 3 (c) (b) (a) 4 (a) (c) (a) 5 (b) (c) (a) . 6 (c) (c) (a) 7 (c) (a) (a) 8 (a) (a) (a) 9 (c) .(a) (b) 10 (a) (a) (b) 111 (c) (c) (c) 12t (a) (c) (c) 13t (c) (a) (c) 14t (a) (a) (c) 15 (a) (b) (a) 16 (b) (b) (a)
- Please see next page for explanation of symbols (a,b, & c) used.
i t These cases are analyzed to fulfill the requirements of single failure criterion (i.e. one FPHX and its associated loop is out of service). 5-8
TABLE 5.1.lA (Continued) List of Assumptions and Criteria I. Time-(hrs) for Fuel Movement from Core after Reactor Shutdown (Cool-off time) (a) 90 (b) 110 (c) 130 II. Rate of fuel discharge (a) Instantaneous discharge to UCP - hold for 24 hours in UCP and then instantaneous discharge to SFP. (b) Discharge to UCP @ 4 bundles / hour, 24 hours after start of discharge into UCP, begin transfer from UCP into SFP @ 4 bundles / hour until all required bundles are discharged. (c) Similar process as in (b) above except that the transfer rate is 2 bundles / hour. III. Availability of fuel pool heat exchangers (FPHX) (a) 1 FPHX in continuous service and 1 FPHX put in service at start of refuelling and operating for 60 days thereafter. (b) 2 FPHX in continuous service (c) 1 FPHX in continuous service NOTE: Cases 1-14: Normal operation - Fuel discharge of 228 bundles for the 12th discharge (or 13th loading. cycle). Last refueling outage completed 12 months prior.
- Cases 15 & 16
- Abnormal operations - Full core discharge of 800
. bundles 90 days after the 12th refueling outage (or the lith discharge) l I
1 5-9 l _ _ _ ______ _ _ _ n % a _
O TABLE 5.1.lB CASES ANALYZED FOR LONG TERM COOLABILITY Case Condition No. of No. of RHR Total Time Cool-Off Fuel Spont Fuel Availability to Transfer Time Assemblies Pool HX's Fuel into before SFP (hrs) Transfer begins (hrs) A Normal 228 Note 1 . Note 2(a) 57 110 Discharge B Full Core 800 Note 1 Note 2(b) 200 110 Y Notes: 1 1 SFPHX in continuous service and 1 SFPHX put in service at start of refueling and operates for 60 days thereafter. 2 (a) 1 RHR loop available immediately 3fter shutdown and is in service for 30 days after start of discharge into upper containment pool. (b) Same as (a) but RHR remains in service.
I l TABLE 5.1.2A MAXIMUM SPENT FU3L POOL BULK TEMPERATURE t, COINCIDENT TOTAL POWER Qi, COINCIDENT TITE AND TIME TO BOILING (Cases 1 - 16)
-6 Case Peak Coincident Qix10 Time to Boil-Pool Bulk time (since BTU /hr ing from the No.
temp. *F initiation of instant all fuel transfer, cooling is into UCP), hrs. lost, hrst 23 i 105.5 84 15.0530 86 14.6332 23 2 105.2 104.9 86 14.3058 24 3 4 104.7 140 14.1007 24 104.4 142 13.8097 24 5 142 13.5721 25 6 104.2 36 15.2252 23 7 105.7 36 16.3065 21 i 8 106.7 36 15.2252 23 9 105.7 36 16.3065 22 10 106.7 142 13.5721 24 11* 106.1 140 14.1007 24 12* 106.6 38 15.1810 22 13* 107.8 36 16.3065 21 14* 109.1 15 126.1 226 37.4520 7 16 125.5 226 36.8038 7 i All cooling is assumed ..s at the time instant when pool bulk temperature peaks.
- Single failure critert - cases [see Table 5.1.1A].
9 5-11 m
TABLE 5.1.2B MAXIMUM SPENT FUEL POOL BULK TEMPERATURE t, COINCIDENT TOTAL POWER Qt, COINCIDENT-TIME AND TIME TO BOILING (Cases A&B) 1
-6 Case No. of- Time Point in Peak Coincident Time to Boil - Qix10 I
Assemblies to Transfer Fig. 5.1.2 Pool Bulk time (since ing from the BTU /hr Fuel into or 5.1.3 temp. *F initiation of instant all 4 Pool, hrs. fuel transfer, all cooling is j en into UCP, hrs. lost, hrs. I b A 228 57 Al 105.2 86 23 14.6632 , A2 121.9 830 29 9.8473 A3 125.9 1500 -30 8.5029
, B 800 200 B1 125.5 226 7 36.8038 B2 113.0 1455 14 21.4858 e
e
cnolysos cro prosanted in Table 5.1.2A. Thoso analysos are performed for an elapsed time (i.e. time since commencement of fuel transfer) of 400 hours. Results obtained from these analyses show that neither cool-off time (i.e. time elapsed between reactor shutdown and before commencement of fuel discharge from the core) nor mode of discharge (ramps or instantaneous) have any significant impact on bulk pool temperatures after 200 hours since commencement of discharge. Therefore, two cases, case 2 and case 16, re-labelled as cases A and B respectively, are chosen for analysis in Phase II for an extended time period so as to study the effect of operational parameters on bulk pool temperatures. Tr.e two cases are described in Table 5.1.1B, their results presented in Table 5.1.2B and graphically shown in Figures 5.1.2 and 5.1.3. These plots show that the spent fuel pool water never approaches the boiling point under the most adverse conditions. These figures also give 01 as a function of T. In addition, an analysis for each loading cycle is performed based on the refueling schedule (ref Table 1.1). These results are plotted in Figure (5.1.4) and (5.1.5). As stated earlier in this section the upper containment pool serves as a temporary storage pool and therefore the bulk pool temperatures in this pool are generally lower than those in the spent fuel pool. The exception, as can be expected, being for an instantaneous discharge for a normal batch (upper limit fuel discharge rate) where peak temperatures in the upper ! containment pool are higher. However, even for the most
- conservative condition (case 14) the calculated peak temperature in the upper containment pool is about 123'F. For the case of an abnormal discharge " Full core discharge (Cases 15 & 16)",
discharge rate of 4 bundles /hr is considered. An instantaneous discharge rate, for these cases, is not representative of refueling operations. Therefore, for purposes of analysis, the discharge rate considered coupled with scenario presented above provides an upper bound to bulk pool temperature. 5-13
It should be recognized' that the analysis described and l the results presented herein are conservative and do not represent realistic operational conditions. The purpose of those analyses is to show that under the most conservative scenario acceptable temperature limits will not be exceeded utilizing the available spent fuel pool cooling system with an appropriate backup from the RHR system. Moreover, the licensee intends to monitor the bulk pool temperature subsequent to refueling and take appropriate measures. 5.2 Natural Circulation Thermal-Hydraulics Analysis This report section covers requirement III.l.5(3) of the NRC "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications" issued on April 14, 1978. Conservative methods have been used to calculate the maximum fuel cladding temperature as required therein. Also, it has been determined that nucleate boiling or voiding of coolant on the surface of the fuel rods does not occur.
5.2.1 Basis
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-shutdown" (ts)+ Since the heat emission falls off rapidly with increasing t s, it is obviously conservative to assume that all 110 hours), and they all have had 4 years fuel assemblies are fresh (ts " of operating time in the reactor. The heat emission rate of each fuel assembly is assumed to be equal.2
- b. As shown in Figures 2.1 and 2.2 in Section 2, the modules occupy an irregular flpor space in the pool. For purposes of the hydrothermal analysis, a 5-14 -
m
4 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 6.2585 inches (see Figure 5.'2.1).
- c. The . downcomer space around the rack module group varies, as shown in Figure 5.2.1. The minimum downcomer gap (6.25 inches) available in the pool is assumed to be the total gap available around the idealized cylindrical rack; thus, the maximum resistance to downward flow is incorporated into'the analysis.
- d. No downcomer flow is assumed to exist between the rack modules.
5.2.2 Model Description In this manner, a conservative idealized model for the rack assemblage is devised. The water f' low 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 j coefficients which enter into the formulation of the integral s
l equation are also taken from well-recognized sources and wherever discrepancies in reported values exists, the
- j. conservative values are consistently used.
i After the axial velocity field is eva'luated, it is a straight-forward matter to compute the fuel assembly cladding l l l l 5-15
- ~
temperature. The knowledge of the overall flow field enables pinpointing the storage location with the minimum axial flow (i.e. maximum water outlet temperature). This is called the most
" choked" location. It is recognized that some 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 asumed 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. In view of the preceding assumption, the temperatures calculated in this manner over-estimate the temperature rise that will actually be obtained in the pool.
In this analysis results for cases A and B as presented in Section 5.1.3 were utilized. The radial peaking factor used in the analysis is 1.356. The maximum temperature rise of pool water in the most disadvantageously placed fuel assembly is given in . Table 5.2.1 for all loading cases. Having determined the maximum
" local" water temperature in the pool, it is now possible to determine the maximum fuel cladding temperature. It is conservatively assumed that the total peaking factor aT is 1.63. Thus, a fuel rod can produce 1.63 times the average heat emission rate over a small length. The axial heat dissipation in a rod is known to reach a maximum in the central region; and taper off at its two extremities. For the sake of added conservatism it is assumed that the peak heat emission occurs at the top where the local water temperature also reaches its maximum. Furthermore, no credit is taken for axial conduction of heat along the rod. The highly conservative model thus constructed leads to simple algebraic equatio,ns which directly give the maximum local cladding temperature, tc*
5-16 n
e 1 l l 5.2.3 Results and Discussion Table 5.2.1 gives the corresponding maximum local cladding temperature, tc, at the instances the spent fuel pool bulk temperature peaks. There are three such ' " peaks" for case A as l shown in Figure 5.1.2 and two for case B (only peak B1 is analyzed since peak B2 is significantly lower than peak B1) as indicated in Figure 5.1.3. It is quite possible, however, that the peak cladding temperature occurs at the instant of maximum value of . gr, i.e., at the instant when the fuel assembly is first placed in a storage location. Table 5.2.2 gives the maximum local cladding temperature at t=24 hours (i.e. the instance when fuel transfer from the upper containment pool to the spent fuel pool begins). It is to ,be noted that there are wide margins to local boiling in all cases. The local boiling temperature near the top of the fuel cladding is 243*F. Furthermore, the cladding temperature must be somewhat higher than the boiling temperature to initiate and sustain nucleate boiling. The above considerations indicate that a comfortable margin against the initiation of localized boiling exists in all Cases. 5-17 m
1 TABLE 5.2.1 MAXIMUM LOCAL (SPENT FUEL) POOL WATER TEMPERATURE AND LOCAL FUEL CLADDING TEMPERATURE Case Point in Fig. 5.1.2 Max. Local Pool Coincident Specific Maximum Coincident Local or Fig. 5.1.3 Water Temperature *F__ Power q, BTU /see Cladding Temperature 'F A A1 131.8 12.488 153.62 A2 138.4 6.741 150.77 Y A3 139.7 5.103 149.32 H B B1 147.6 10.509 165.88 T I _
TABLE 5.2.2 POOL AND MAXIMUM CLADDING TEMPERATURE AT THE INSTANCE FUEL ASSEMBLY TRANSFER FROM UPPER CONTAINMENT POOL TO SPENT FUEL POOL BEGINS, T=24 Case Cladding Coincident Specific Coincident Pool Temperature, 'F Temperature 'F Power, q, BTU /sec Bulk Local A 154.6 15.31 98.0 128.5 B 155.6 15.31 99.0 129.5 Y e
r n j REFERENCES TO SECTION 5 1.- NUREG 0800 U.S. Nuclear Regulatory Commission, Standard Review Plan, Branch Technical Position, ASB 9-2, Rev. 2, July 1981.
- 2. -FSAR, Grand Gulf Unit I, Section 9, Auxiliary and Emergency Systems.
3 " Technical Specifications for Design, Fabrication and' Purchase of'High Density Spent Fuel Storage Racks", NPE-M-181.1, Rev. 2, Dec., 1982. Mississippi Power & Light Company.
~4. "Some Fundamental Relationships for Tubular Heat Exchanger Thermal Performance", K.P. Singh, Journal of Heat Transfer, Transactions of the ASME, Vol. 103, No. 3,
. August, 1981.
- 5. General Electric Corporation, R&D Data Books, " Heat
' Transfer and Fluid Flow", 1974 and updates.
- 6. T.J. Helbling (ENC) to L.F. Dale (MP&L) " Grand Gulf Nuclear Station FSDD_ Response", Letter MPEX-82/57, July 3, 1982. (EXXON PROPRIETARY) 4 7
5-20 m
= ' '
ts A W22
.. ..f4 14
.. .. is W i2 Wll W21 U lI :_ --
If U - UCP SFP POOL-2 POOL-l O !l
.12 -
Il t2 _ _ __t ' W 22 W21 T H w . w 02 __ te W2 f i r 1r Os M l :
'T 1r is j
13 e[w(j I4 m W: WI R. H.R. HX 7 S. F. P. H X FIG. 5.1.1 - MODEL FOR GRAND GULF POOLS
J 140 - PEAK VALUE = lI4,633 x 10' Btu /hr. - 20 AT 86 HOURS (SPENT FUEL POOL)
~
h PEAK VALUE=125.9 Ji 130 - AT 1500 HOURS (SPENT FUEL - 15 POOL) s 32
~
SPENT FUEL t3ULK b
- POOL YEMP.
W g 120 -
'O 4
B m @ w y - 10 to 2 $ i
$ I o
y 11 0 - (n POWER DISCHARGED t' c i _J w 8
- A -
52 100 - 720 hrs.= TIME AT WHICH 1440 hrs. = TIME AT WHICH RHR LOOP IS (1)- FPHX IS " SHUT-OFF"
" SHUT-OFF" 90 ' ' ' " ' ' ' ' " ' O
. O 200 400 600 800 1000 1200 1400 1600 1800 TIME, HOURS =
l FIG. 5.l.2 - SPENT FUEL POOL BULK TEMPERATURES AND POWER DISCHARGED FOR CASE A - ( NORMAL DISCHARGE) 5-22 e
I 4 . 1 i s 37o _ PEAK VALUE = 36.804 x10 B t u / hr. _ 4o 1 AT 226 HOURS l 38 0 - h PE AK VALUE =125.5 g AT ~ 226 HOURS n j 150 - 30
'. I 14 0 - i
, i g
g POWER . R DISCHARGED 8 o g 130 - t 20 ; n' S b E i F 12 0 - B2 E i 5 E ! E E l 11 0 10 3 - m g SPENT FUEL POOL y BULK TEMPERATURE E 100 1440 HOURS = TIME AT WHICH (1)-FPHX iS " SHUT-OFF " 90 ' ' ' ' ' ' ' " ' O O 200 400 600 800 1000 1200 1400 1600 1800 l TIME, HOURS : FIG. 5.l.3 - SPENT FUEL POOL BULK TEMPERATURES AND POWER DISCHARGED FOR CASEB ( ABNORMAL DISCHARGE) i 5-23
1 I I
- ASYMTOTIC PEAK TEMPER ATURE = 125.9
'30 ~ w AT LOADING CYCLE 13 4
125- y , y l E LOADING i I20- $ CYCLE y 2 3 4 5 6 7 8 9 t 11 5 - 110- \ l !O t* 105-
\
100- \ \ \ j C5-I ' ' ' ' ' ' ' ' ' ' ' ' ' ' 90 O 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 95 10 0 ! MONTHS AFTER FIRST BATCH REMOVAL TO SPENT FUEL POOL i FIG. 5.l.4- FUEL POOL BULK TEMPERATURE PROFILE USING CRITERIA FROM CASE A i 1 1
Pc/Po, POWER RELEASED BY FUEL ASSEMBLIES / OPERATING POWER PER ASSE M BLY Po = 16.348 x 10' B tu/hr. g ASYMTOTIC PEAK LOAD = 0.90 Po
) AT LOADING CYCLE 13
, n. i j 1.0 - LOADING i CYCLE 2 3 4 5 6 7 8 9
, 0.8 -
E 0.6 - ; 4 0.4 -
! O.2 -
O ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' ' O 5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 95 100 MONTHS AFTER FIRST BATCH REMOVAL TO SPENT FUEL POOL FIG. 5.l.5- FUEL POOL HEAT LOAD PROFILE USING CRITERI A FROM CASE A i
i Idealized Outline _
-" ' " '- ~Q .
of Rack Assembly f' -
/
/ x Actual Outline
/ ,
of Rack Assembly A g ui
! RACK ASSEMBLY
/ -
\
E - Actual Outline m i of Pool
\\ ;/
\ s ,
. N ss
/ /
'Q s f - Assumed Added Fuel Assemblies Idealized Outline N-C_
~
of Pool Boundary Figure 5.2.1 Rack Space Enveloping cylinder - (Grand Gulf Unit One)
- 6. STRUCTURAL ANALYSIS The purpose of this section -is to demonstrate the structural
{ adequacy.of the spent fuel rack design under normal and accident loading conditions. The method of analysis presented herein is similar to that previously used in the Licensing Reports on High Density - Fuel Racks for Fermi II (Docket 50-341) and Quad Cities (Docket 50-254-265). The results show that the high density spent fuel racks are structurally adequate to 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 Ac 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. Furthermore, 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 structure. The coefficient of friction, y, between the supports and pool floor is another-indeterminate factor. According to Rabinowicz,2 the results of , 199 tests performed on austenitic stainless steel plates submerged i' in water show a mean value of y to be 0.503 with a standard deviation of 0.125. The upper and lower bounds (*20) are thus 0.753 and 0.253, respectively. Two separate analyses are performed
- for this rack assembly with values of y equal to 0.2 (lower limit) and 0.8 (upper limit) respectively. Initially, the following six separate analyses are performed on the largest rack module (Module A).
- 1. Fully loaded rack (all storage locations occupied),
y = 0.8 ( p = coefficient of friction).
- 2. Fully loaded rack, p= 0.2.
6-1
-,--,n---..,., - - - . - - - -
, s., --..-------,-.nn-n.. -.,----------,-n- ,, ..--.c - - -
- 3. Half-loaded rack to produce maximum geometric asymmetry about the major dimension of the rectangular rack, y = 0.8.
- 4. Half-loaded rack to produce maximum geometric asymmetry about the major dimension of the rectangular rack, y = 0.2.
- 5. Empty rack, u = 0.8.
- 6. Empty rack, p = 0.2.
Based on the results of these runs, additional analyses are performed. The actual studies performed for the different rack modules are summarized in Section 6.6. The method of analysis employed is the time history method. The pool slab acceleration data are developed from the original plant design as developed by Bechtel Power Corporation, Gaithersburg, Maryland. Bechtel's report is provided as Appendix 1 (proprietary) to this report. 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 " Square-Root-of-the-Sum-of-the-Squares" method. 3 is avoided and the dependability of computed results is ensured. The seismic analysis is performed in four steps, namely
- 1. Development of nonlinear dynamic model consisting of beam, ' gap, spring, damper and inertial coupling elements.
~
- 2. Derivation and computation of element stiffnesses using a sophisticated elastostatic model.
6-2 e
- 3. Layout of the equations of motion, and inertial ,
decoupling and solution of the equations using the
" component element time integration" procedure 4,5 to
. determine nodal and element forces and displacements of nodes.
- 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 .
Determine if the stress and displacement limits,given in Section 6.5, are satisfied. A brief description of the dynamic model follows. _ 6.2 Fuel Rack - Fuel Assembly Model s 6.2.1 Assumptions y'
- a. The fuel rack metal structure is represented by five e lumped masses connected by appropriate elastic springs as shown in Figure 6.1. The spring rates -
( _ simulate the elastic behavior of the fuel rack as a beamlike structure. p<
- b. The fuel assemblies are represented by five ,1 umped i
masses located, relative to the rack, in a manner which simulates either fully or , partially. loaded s conditions.
,. , , ~
- c. The local flexibility of the rack-support interface
~
is modeled conservatively in 'the analysis.
~
- d. The rack base support may slide or lif t off the pool floor.
~
.g.
6-3
._ -__ __ _ m i
s
- e. The pool floor is assumed to have a known time history of accelerations along the three orthogonal directions,
- f. Fluid coupling between rack and assemblies, 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 for by appropriate spring gap connectors between masses involved.
- h. Fluid damping between rack and assemblies, and between rack and adjacent rack is conservatively s
neglected.
- i. The supports are modeled as extensional elements for dynamic analysis. The bottom of a support leg is attached to a frictional spring as described in Section 6.2.2. The cross section properties of the support beams are derived and used in the final computations to determine suppor* 1eg stresses.
- j. The effect of sloshing can be shown to be negligible at the bottom of a pool and is hence neglected.
- k. Deformation due to shear is included in the rack springs, but neglected in the fuel assembly springs.
6-4
,7
f 6.2.2 Model Description The absolute degrees of freedom associated with each of the mass locations i, i* are as follows (see Figure 6.1): Table 6.1 Degrees of Freedom Location Displacement Rotation (Node) ux uy uz 8x O y 6z 1 p1 p2 P3 94 95 96 1* Point is assumed fixed to base at XBsYB,2=0 2 p7 pg 911 912 2* P8 P10 3 pl3 P15 917 918 # 3* Pl4 P16 4 pig p21 923 924 4* P20 P22 5 p25 P27 P32 929 930 931 5* P26 P28 i 6-5
f i 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 1 relatively high axial rigidity of the rack. Torsional motion of l the rack relative to its-base is governed by q31-The members joining nodes 1 to 2, 2 to 3, etc., are the beam elements with deflection due to bending and shear capability (see Reference 4,pp. 156-161.). The elements of the stiffness matrix of these beam elements are readily computed if the ef fective flexure modulus, torsion modulus, etc., for the rack structure are known. These coefficients follow from the elastostatic model as described later. The nodal points i (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 the fuel assemblies distributed at 5 elevations. The element stiffnesses of the fuel assembly are obtained from its geometry. The nodes i* are located at x = XB, Y = YB in the global coordinate system shown in Figure
'6.1.
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 j . vibrates adjacent to another body (mass m2), and both bodies are submerged in a frictionless fluid medium, then the Newton's l O i 6-6 i
Y I
- equations of motion for the two bodies have the form (m1 +M11) 1-M 12 2= ' applied forces on mass mi
-M21 X 1+ (m2 + M22) X2 = applied forces on mass m2 M11, M12e M21, and M22 are fluid coupling coefficients which depend on the shape of the two bodies, their relatives disposition, etc. Fritz6 gives data for Mij for various body shapes 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)-
i' 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 forms 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
- coupling terms which model the effect of fluid in the gaps between adjacent racks.
Finally, fluid virtual mass is included 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, etc. 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 scsses. In A - 6-7 o 4
the analysis, a maximum of 4% structural dcmping is imposed on j elements of the rack structure during SSE seismic simulations. I Actual structural damping values used in the analysis are provided l in Table 6.4. This .is in accordance with the Grand Gulf Nuclear Station Units 1 & 2 FSAR and NRC guidelines 7 Material and fluid damping ere conservatively neglected. J' 6.2.5 Impact The fuel assembly nodes i* will impact the corresponding structural mass node i. To simulate this impact, 4 impact springs around each fuel assembly node are provided (see Figure 6.2). The fluid dampers are also provided in parallel with the springs. The spring constant 'of the impact springs is assumed equal to the local , stiffness of the vertical panel computed by evaluating the peak deflection of a six inch diameter circular plate subject to a i' specified uniform pressure, and built in around the edge. The spring constant calculated in this manner should provide an upper bound on the local stiffnesses 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 modeled for. the general three dimensional (3-D) motion simulation, by five lumped masses and inertial nodes for the rack, base, and supports, and by five lumped masses for the assemblage of fuel assemblies. To simulate the i connectivity and the elasticity of the configuration, a total of 18 linear spring dampers, 20 nonlinear gap elements, and 18 nonlinear friction elements are used. A summary of spring-damper, gap, and friction elements with their connectivity and purpose is presented in Table 6.2. If we restrict the simulation model to two dimensions (one horizontal motion plus vertical motion, for example) for the purposes of model clarification only, then a descriptive model of 4 the simulated structure which includes all necessary spring, gap, 6-8 m e , ~
1 l
- and- f riction elements is shown in Figure 6.3. The beam springs, KA, KB at_each level, T 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, 14, ~and 15. The extensional spring, K, 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 spring-dampers either have zero coefficients (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 bending, is apportioned to the five levels of 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 support leg spring rates K3 are modelled by elements 9 and 10 in Table 6.2 for the 2-D case. Note that the local l
elasticity of the concrete floor is included in K A. To simulate sliding potential, friction elements 2 plus 8 and 4 plus 6-(Table 6.2) are shown in Figure 6.3. The local spring rates Kg reflect , the lateral elasticity of the support legs. Finally, the ' support i rotational friction springs K, R reflect the rotational elasticity of the foundation. The nonlinearity of these springs (friction elements 9 plos 15 and 11 plus 13 in Table 6.2) reflects the edging [ [. limitation imposed on the base of the rack support legs. For the 3-D simulation, carried out in detail for this i 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 l' 6-9 <
l
---w--w "-w+ '----w' =-+w e'
Table 6.2 Numbering System for Springs, Gap Elements, Friction Elements I. Spring Dampers (18 total) Number Node Location Description 1 1-2 X-2 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-2 rack shear 10 3-4 Y-Z 11 3-4 Y-Z rack bending 12 3-4 X-Z 13 4-5 X-Z rack shear 14 4-5 Y-Z 15 4-5 Y-2 rack bending 16 4-5 X-Z 17 1-5 Rack torsion spring 18 1-5 z rack extensional spring i 6-10 9
m Table 6.2 (continued) 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. assembly impact 6 3,3* X rack / fuel assembly impact 7 3,3* Y rack / fuel assembly impact 8 3,3* Y rack / fuel assemlby impact 9 Support S1 Z compression spring 10 Support S2 Z compression spring 11 Support S3 2 compression spring 12 Support S4 Z compression spring 13 4,4* X rack / fuel assembly impact spring 14 4,4* X rack / fuel assembly impact spring 15 4,4* Y rack / fuel assembly impact spring 16 4,4* Y rack / fuel assembly impact spring 17 5,5* X rack / fuel assembly impact spring 18 5,5* X rack / fuel assembly impact spring 19 5,5* Y rack / fuel assembly impact spring 20 5,5* Y rack / fuel assembly impact spring 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 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 11 S2 X Floor Moment 12 S2 Y Floor Moment 13 S3 X Floor Moment 14 S3 Y Floor Moment 15 S4 X Floor Moment 16 S4 Y Floor Moment 6-11
'T
I 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 results. A brief description of the elastostatic model now 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 Analysis 6.3.1 Stiffness Characteristics: The fuel rack is a multicell, folded-plate structure
. which has what is colloquially called a " honey-comb" configuration. This type of construction is very similar to the so-called " stressed-skin" construction of ribs, spars, and cover plates which are widely used in aircraf t construction. Techniques developed in the field of aircraft structural analysis are utilized herein to find the stresses and deformations in such structures.
These methods have been thoroughly tested and their reliability has been documented in a number of publications.8-12 Figure 6.4 shows two cross-sections of the fuel rack which is modeled as a rectangular network of plates interconnected along nodal lines shown as points in Figure 6.1. An arbitrary load with components Fx i, Fyi, Fi g acts at 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. The torsional deformations are solved for by using the classical theory of torsion for multicelled, thin-walled, cross sections.13 6-12 m
The bending deformation is found by using the theory of shear flow 12 wherein all axial stresses are carried by the effective
~
flanges (or . stringers) formed by the intersections of the plates and all transverse shears are carried by the plates modeled as shear panels. 4 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),q a Bending rigidity (in two places)
,[ (GJ),q = Torsional rigidity
- j. '
(AE)egs = Extensional rigidity I ks = Shear deformation coefficient Such properties are used for the dynamic analysis of seismic loads and serve to establish values for the spring rates of the elastic beam elements representing each rack section. 6.3.2. Combined Stresses and Corner Displacements i The cross-sectional properties and the Timoshenko shear correction 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 stress resultants (Fx, Fy, F 2, Mx, l M, y M) z act as shown in -Figure 6.6 are computed for a large number of times t= At, 2At, etc., at a selected number of cross sections. The displacements (U xe Ur y U) z at selected nodal points on the z axis are also provided by the dynamic analysis as
^
well as the rotations (e x, e, y
- 6z) of the cross sections at I the nodes. i
+ (
- 6-13
Figure 6.7'shows a typical subdivision of the structure into I elements, nodes, and' sections. The stresses are calculated at all sections and-the displacements at all four corners of the racks are calculated at these elevations. Since the axial stress varies linearly over the cross section and achieves its extreme values at one of the four corners of the rack, the shear stresses 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 displacement for each of these _ points is written out. Similarly, the stresses in the support legs are evaluated. A. validated Joseph Oat Corporation proprietary computer
. program "EGELAST"i computes the stresses at the candidate points at each level. It sorts out the most stressed location in space as
-well' as time. The highest stress and maximum kinematic displacements are thus readily found.
6.4 - Time Integration of the Equations of Motion Having 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
. generalized coordinate pg) is written as follows: . t This code has been previously utilized in licensing of similar 1
racks for Fermi II (Docket No. 50-341), and Quad Cities I and II i (Docket Nos. 50-254 and 265). 1-6-14 j-t
The inertial mass is: m22 + A211 + B211 where m22 is the mass of node 2 for y-directional motion. A211 is the fluid coupling mass due to interaction with node 2*, and B211 is the fluid coupling mass due to interaction of node 2 with the reference frame (interaction between adjacent racks). Hence, Newton's law gives (m22 + A211 + B211) P9 + A212 P10 + B212 ,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*; B212 is the cross term fluid coupling effect of the adjacent racks, and u represents ground acceleration. Let 99 " P9 - u 910 = P10 - 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 rearranging, we have: (m22 + A211 + B211) 99 + A212 910 = 09 - (m22 + A211 + B211 + A212 + B212) 'u' 6-15 T
t A similar equation for each one of the 32 degrees of freedom can be written. The system of equations can be represented in matrix notation as: (M) {q} = (0] + {G} l where the vector (0) is a function of nodal displacements and velocities, and {G} depends on the coupling inertia and the ground acceleration. Premultiplying above equation by (M )-1 renders the resulting equations uncoupled in mass. We have: {'q } = [g ]-1 [o] + [g]-1 {c} The generalized force 0,9 which contains the effects of all spring elements acting on node 2 in the " direction" of coordinate gg (the relative displacement of node 2 in the y direction), can easily be obtained from a free body analysis of node 2. For example, in the 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 springs which couple node 2* and node 2. Since each of these four spring elements contain couplings with other component deformations through the spring force-deformation relations, considerable static coupling of the complete set of equations results. The level 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 effect. For example, referring to Figure 6.3, and Table 6.1, a 2-D simulation introduces static coupling between coordinates 2,9 and 15 in the expression for 0,9 this coupling comes from the shear springs simulating the rack elasticity which have constitutive relations of the form F = Ks (99 - 92) , Ks (415 - 99) . Further, the impact springs introduce two additional forces having constitutive equations of the form F =KI (q9 - ql0) . Of course, at any instant, these forces may be zero if the local gap is open. The local gap depends on the current value of gg - glo . + 6-16
~
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 l 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 named "DYNAHIS"i, developed by General Electric Company, performs this task in an efficient-manner. Having determined the internal forces as a function of time, the computer program "EGELAST" computes the detailed stress and displacement fields for the rack structure as described in the preceding section. 6.5 Structural Acceptance Criteria 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 tilting 15 are met (1.5 for OBE, 1.1 for SSE). (b) Stress Limits (1) The stress limits of the ASME Code, Section III, Subsection NF, 1980 Edition up to and including Winter 1981 addenda were chosen to be met, since t This code has been previously utilized in licensing of similar racks for Fermi II (Docket No. 50-341), and Quad Cities I and II (Docket Nos. 50-254 and 265). ' 6-17 7
this Code provides the most consistent set of limits for various stress types, and various loading conditions. The following loading casesl4 have been analyzed. SRP Designation ASME Designation (1) .. D+L Level A (ii) D+L+T o Level A (iii) D+L+To + E. Level B (iv) D+L+Ta+E Level B (v) D + L + To + Pg Level D (vi) D + L + Ta + E' Level D D+L+Fd Level D These loads have the following connotation with respect to the fuel rack analysis: D = Dead weight of rack and contained fuel assemblies.
-L~ = Live load = 0 T = Stresses due to differential thermal expansion in the body of the rack. It is further explained below.
E = Operating basis earthquake T, = Identical to T o E' = Safe shutdown earthquake Pg = Upward force on the racks caused by postulated stuck fuel assembly F d
= Force caused by accidental drop of the heaviest load for the maximum possible height.
The conditions Ta and To cause local thermal stresses to be produced. The worst situation will be obtained when an isolated storage location has a fuel assembly which is generating heat at the maximum postulated rate. The surrounding storage locations are assumed 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. The secondary stresses thus produced are limited to the body of the rack, that is, the support legs do e not experience the secondary (thermal) stresses. , 6-18
(2) Basic Data: The following data on the physical properties of the rack material are obtained from the ASME Codes, Section III, appendices. Table 6.3 Physical Property Data
- Property . Young's Yield Ultimate Allowable Modulus Strength Strength Stress 4 2000F 92000F 92000F 6 2000F E Sy Su S Value 28.3 x 106 25 KSI 71 KSI 17.8 KSI psi Section III Table Table Table Table Reference I-6.0 I-2.2 I-3.2 I-7.2
- Evaluated at 2000F. This temperature is higher _ than the pool water bulk temperature under any of the loading conditionc under consideration.
(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 Ft is equivalent to primary membrane stresses (ii) On the gross section, allowable stress in shear is Fy = 0.4 S y
= (0.4) (25000) = 10000 psi 6-19 l 1
l l (iii) Allowable stress in compression, F a (1 - ( ) 2C c )S y. F = (( .) +(3 ( 8C ) 8C c ) 3 r c] - [(r where 1/2 2w2E C" = [ S
]
y substituting numbers, we obtain, for both support leg and " honey-comb" 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 (v) Combined flexure and compression: h,Cmx bx f
, C,y fby 4 y F, DFx bx DF y by where fan Direct compressive stress in the section.
fx b Maximum flexural stress along x-axis Fby: Maximum flexural stress along y-axis Cmx = Cmy = 0.85 6-20
~
f a x =1-D Ex D =1- f* Y F OY where 12: 2 E Fx= e 2 kl 23 ( b)
*b (vi) Combined flexure and compression (or tension)
"
- Y
+ + < l.0 0.6 S y F bx P by The above req.)irement should be met for both diret 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 (Sy /Ft) times the corresponding limits for normal condition. Thus, the multiplication factor is Factor 25000 = 2.0
= (1.2) 15000 i
6-21 m
l 1 (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 S m in the manner of Class 1 components; S m is the allowable stress intensity of the rack material at the maximum operating temperature. O 6-22 7
i 6.6 RESULTS The input time history' accelerations for seismic motion were developed by Bechtel Corporation (see Appendix 1). Plots of the time history motions, utilized in the present analysis are shown in in Figure 6.9 through Figure 6.14. These plots correspond to the Safe Shutdown Earthquake (SSE) with 5% damping. Since there are several rack module configurations (Figs. 2.1 and 2.2) it was decided to make an exhaustive analysis of one rack type. We note that rack A is an above-average size module, and hence will produce above-average floor reaction and support stress levels. Rack type A is also most numerous. Hence rack A is chosen for performing extensive analyses. Appropriate simulations are also carried out for other limiti,ng rack geometries (e.g. tipping study for rack with low cross section to height aspect ratio, stress evaluations for the heaviest module, etc.). To determine the magnitude of structural dampers, free lateral vibration plots of the top of rack A (in X and Y directions) for fully loaded and empty conditions were developed. The dominant natural frequency of vibration thus evaluated enables computations of the linear structural dampers. The maximum percentage structural damping for SSE condition is assumed to be 4% and modifications to the stiffness matrix to incorporate damping is based on the dominant frequency of 12.75 cps. Having determined the damper characteristic data, the dynamic analysis of the rack module is performed using the computer program DYNAHIS. Two components of the SSE horizontal acceleration are applied in two orthogonal directions concurrently with the. vertical seismic acceleration. Abstracted results for all simulations carried out are reported in Table 6.4. Table 6.5 gives the maximum values of stress factors Ri (i = 1,2,3,4,5,6). The values given in the tables are the maximum values in time and space (all sections of the rack). The various stress factors are listed below for convenience of reference. 6-23 7
R:1 Ratio of tensile stress on a net section to its allowable OBE value R:2 Ratio of maximum gross shear on a net section to its allowable OBE value R:3 Ratio of maximum bending stress in one plane' (x-y) to its allowable OBE value for the section R:4 Ratio of maximum bending stress in one plane (y-z) to its allowable value in OBE R:5 Combined flexure and compressive factor R:6 Combined flexure and tension (or compression) factor The allowable value of Ri (i = 1,2,3,4,5,6) is 1 for OBE condition, and is 2 for SSE condition (see Section 6.5). The displacement and stress tables given herein are for the SSE conditon. If necessary, OBE studies are run to qualify the - rack when the SSE simulation with damping of 21 does not . yield stress factors that would clearly indicate a safe margin under OBE conditions. Seismic simulations for the tipping conditions are carried out by increasing the horizontal SSE accelerations by 50g15 These calculations indicate ' that the rack remains stable, and the gross movement remains within the limit of small motion theory. Por rack type A (304 cells), the gap between racks is 4 5/32" in the x direction and 3 3/4" in the y direction. The maximum displacement in either direction for any case considered is 0.62" which is less than 50% of the spacing. Note that the direction along the smallest side is the local X direction of the rack. For rack Dl, the critical spacing is 4 5/32" in the local x direction and 4 9/16" in the y direction. The critical deflection for all cases considered is 1.64". 1 6-24
. ~
l l For rack type G in the containment pool, the critical dimensions are 5 7/16" and 7" respectively. Referring to Table 6.4, the maximum displacements for rack G under all cases is less than 1.6", although the artificial tipping analysis (1.5 x SSE on orizontal earthquake) gives 3.02". However, the purpose of the tipping analysis is simply to show the safety factor against overturning. For rack type F, the critical spacings are 7" and 3 15/16" respectively with the maximum displacement of 1.17 obtained being less than 50% of the inter-rack spacing. Note that all of the rack displacements presented in Table 6.4 are quoted for the top of the rack. Table 6.5 shows the maximum values of the stress factors obtained. For rack types A, Dl, all stress factors are less than 1.0 even using the SSE earthquake if we exclude the cases simulated for tipping. Even under 1.5 SSE horizontal, 1.0 vertical earthquakes, the stress factors remain under 2.0 for an SSE event. For racks G and F, the stress factors are below 2.0 for all SSE events simulated, even when only 2% damping is considered. The one OBE simulation shows that stress factors drop considerably when the less severe earthquake is used. It is noted that all of the stress factors shown are maximum in the rack supports. Stress coefficients in the rack proper are usually about 10% or less than the values quoted for the rack supports. As noted above, seismic simulations for the tipping conditions are carried out by increasing the horizontal SSE accelerations by 50% 15 The calculations indicate that the rack remains stable, and that the gross movement remains within the limit of small motion theory. Thus, the rack module is seen to satisfy both kinematic and stress criteria with large margins of safety. 6-25 M
I l Analysis of the rack for damaged fuel modules shows comparable results, and an analysis of the welded joints in the rack also indicates acceptable safety margins. Cases Considered (All SSE Events except Case 11) Case Number Description 1 Full rack, Damping > 2% p= .8 2 Tipping Analysis (1.5 SSE horizontal quake) 2% damping, p = .8 3 Full rack, p = .,8 , 2% damping 4 Full Rack,.y = .2, 2% Damping 5 Half load, Diag. Fill, p= .8 2% Damping 6 Half Load, Diag. Fill, p = .2 24 Damping 7 Half Load, Positive X Quadrant, y= .8, 2% Damping 8 Empty Rack, p = .8 2% or 4% Damping 9 Empty Rack, p= .2 2% or 4% Damping 10 Half load; Positive X Quadrant u = .2, 2% Damping 11 Full Rack, p= .8 24 Damping, OBE Quake 6-26
. ~
n . .. Table ~6.4 , Maximum Rack Module Displacements (Damping = 2% except where noted). Max. X - Dispt. & Time Max Y - Dispt. & Time Module Case Ux (max) Time Instant Uy (max) Time Instant Type No. u (inch) (sec) (inch) (sec) Comments A 1 .8 .034 6.54 .027 10.51 Full rack 4% Damping SSE A 2 .8 .063 8.07 .121 10.88 Full Rack (1.5 SSE) A 3 .8 .038 8.9 .033 10.98 Full Rack (SSE) A 4 .2 .167 15.21 .184 10.42 Full Rack (SSE) A 5 .8 .050 7.66 .016 6.76 Half full-diag. loading , A 6 .2 .174 10.53 .114 5.0 Half full- dia. loading A 7 .8 .056 10.6 .020 10.44 Half full Pos. x loading A 8 .8 .59 13.37 .62 9.04 Empty rack 4% Damping SSE D1 3 .8 .045 10.3 .021 8.38 Full Rack SSE D1 *2 .8 .122 10.5 .042 10.5 Full Rack 1.5 SSE D1 4 .2 .206 10.4 .200 10.5 Full rack SSE D1 8 .8 1.64 9.38 .611 7.31 Empty Rack 4% damping SSE
. D1 9 .2 .552 5.25 .650 6.75 Empty rack 4% Damping SSE
Table 6.4 (Continued) Maximum Rack Module Displacements (Damping = 2% except where noted) Max. X - Dispt. & Time Max Y - Dispt.& Time Module Case Ux (max) Time Instant Uy (max) Time Instant Type' No. p (inch) (sec) (inch) (sec) Comments D1 5 .8 .215 10.53 .032 10.57 Half full ding. loading - D1 6 .2 .244 10.43 .361 10.51 Half full diag. loading D1 7 .8 .053 9.31 .019 7.23 Half full pos. x loading
.2 D1 10 .179 10.6 .230 10.52 Half full pos. x loading G 1 .8 .257 7.44 1.27 11.31 Full rack 4% damping, SSE G 4 .2 .581 10.55 .612 10.48 Full rack SSE ,
G 2 .8 2.305 10.56 3.02 10.73 Full rack 1.5 SSE G 5 .8 1.52 11.86 1.55 11.22 Half full SSE diag. loading C 6 .2 1.042 8.20 .812 10.94 Half full, SSE, ding. loading G 10 .2 1.10 8.59 .791 10.48 Half full pos. x loading C 9 .2 1.452 8.63 0.774 8.86 Empty rack, SSE (2% damping) M
a Table 6.4 (Continued) Maximus Rack Module Displacements (Damping = 2% except where noted) Max. X - Dispt. & Time Hex Y - Dispt.& Time Module Case Ux (max) Time Instant Uy (max) Time Instant Type No. p (inch) (sec) (inch) (sec) Comuments G 11 .8 .027 5.03 .024 4.79 Full rack OBE Quake 2% Dainping F 1- .8 .147 11.36 1.52 11.35 Full rack 4% damping SSES F 8 .8 .951 15.57 1.17 10.49 Empty rack 4% damping SSE
.F 9 .2 .62 8.63 .74 10.72 Empty rack 4% damping SSE i H 1 .8 .494 11.63 .012 5.98 Full rack, 4% damping 1.5 K SSE E
H 2 .8 1.42 10.74 .38 10.52 Full rack 4% damping 1.5 x SSE H 8 .8 1.402 14.20 1.147 11.62 Empty Rack
Tcbis 6.5 Maximum Values of Stress Factors R1-R6 Module. Case y Rt R2 R3 Rg R5 R6 Type No. Comments A 1 .8 .285 .195 .293 .306 .649 .713 Full rack 4% damping (SSE) A. 2 .8 .439 .374 .706 .597 1.44 1.62 Full Rack (1.5 SSE) A. 3 .8 .321 .209 .334 .363 .647 .849 Full Rack (SSES) A 4 .2 .246 .064 .124 .149 .4353 .469 Full Rack SSE A 5 .8 .223 .179 .215 .386 .635 .712 Half full diag. loading A 6 .2 .172 .049 .081 .097 .280 .299 Half full ding. loading A 7 .8 .197 .151 .262 .304 .635 .713 Half full pos. x loading i e A 8 .8 .154 .176 .292 .231 .458 .516 Empty rack (4%) damping SSE D1 3 .8 .239 .184 .263 .317 .599 .666 Full rack SSE D1 2 .8 .348 .267 .462 .723 1.28 1.45 Full rack 1.5 SSE D1 4 .2 .206 .056 .123 .217 .422 .460 Full rack SSE D1 8 .8 .222 .192 .319 .213 .671 .750 Empty rack (4% Damping) SSE D1 9 .2 .025 .007 .014 .032 .050 .061 Empty rack (4% Damping) SSE D1 5 .8 .209 .148 .254 .409 .624 .698 Half Full diag. loading D1 7 .8 .185 .137 .227 .227 .539 .601 Half full pos. x loading
. , _ _ _ - - M
Table 6.5 (Continued) Maximue Values of Stress Factors R1 -R6 Module Case .y Rg R2 R3 Rg R5 R6 Type No. Comments D1 10 .2 .132 .039 .062 _,172 .299 .328 Half full pos. x loading G 1 .8 .250 .208 .565 .447 .768 .866 Full rack 4% Damping SSE G 4 .2 .165 .045 .224 .282 .454 .508 Full Rack SSE G S .8 .453 .216 .697 .342- 1.081 1.192 Half full, SSE, diag. loading G 7 .8 .172 .094 .403 .251 .549 .622 Half full, pos. x loading G 10 .2 .099 .029 .130 .143 .297 .308 Half full pos. x loading G 9 .2 .022 .006 .025 .026 .060 .067 Empty rack sSE G 11 .8 .135 .039 .123 .151 .273 .303 Full rack OBE Quake 2% damping i 0 F 1 .8 .385 .220 .743 .480 1.069 1.189 Full rack 4% damping SSE F 8 .8 .134 .067 .305 .105 .462 .520 Empty 4% damping F 9 .2 .025 .007 .028 .047 .071 .077 Empty 4% Damping H 1 .8 .184 .173 .207 .487 .588 .618 Full rack 4% damping H 2 .8 1.5 SSE - 2% Damping H 8 .8 .376 .496 .529 .577 .853 .854 Empty rack 2% damping
1 REFERENCES TO SECTION 6
- 1. USNRC Regulatory Guide 1.29, " Seismic Design-Classification,"
Rev. 3, 1978. 2.- " Friction Coefficients of Water Lubricated Stainless Steels for a Spent Fuel Rack Facility," by Prof. Ernest Rabinowicz, MIT, a report for Boston Edison Company, 1976.
- 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. " Dynamics of Structures" R.W. Clough & J. Penzien, McGraw Hill (1975).
- 6. R.J. Fritz, "The Effects of Liquids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, Trans. of the ASME, February 1972, pp. 167-172.
- 7. USNRC Regulatory Guide 1.61, Damping Values for Seismic Design of Nuclear Power Plants, 1973.
- 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., 1966.
- 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. Timoshenko and J.N. Goodier, " Theory of Elasticity,"
McGraw-Hill, N.Y., 1970, Chapter 10. 14 . - U.S. Nuclear Regulatory Commission, Standard Review Plan, NUREG-0800 (1981).
- 15. U.S. Nuclear Regulatory Commission, Standard Review Plan, Section'3.8.5, Rev. 1, 1981.
- 16. U.S. Nuclear Regulatory Commission, Regulatory Guide 1.124,
" Design Limits and Loading Combinations for Class 1 Linear-Type Component Supports, November 1976.
6-32
~
Z nh 3 5
, 5%
/
s' C O U PLIN G ELEMENTS 4# 4 . 3 TYPICAL FU EL AS S EM B LY GROUP M AS S H TYPICAL FUEL R A C K Id A S S . FUEL R ACK B A S E 2 i.' < Ay = 7
' ' / \ /
Ax _L y Iy B 7-
\ "' l .Y l
t /
/-+-y's EB N7 Rg t
e I i r a i 4 h ' 1
, ///
.1.
///
9 FUEL RACK SUPPORT
- I X -
XB, YB - LOCATION OF CENTROID d F FUEL ROD GROUP M ASSES - RELATIVE TO CENTER OF FUEL R ACK Di = UNIT VECTORS FI G. 6.1 D YN A M I C MODEL G-33 l
g . . Y
, ., IMPACT d SPRIN G S. .
O il on
~
[.H MASS .
,% i*
l
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? F LUID DAMPERS RIGID FR AM E t
= X 1
i FI G.. 6.2 l M PA C T S P R I N G S A N D FLUID D AMPERS 6-34
'E
5 , . C %/h*/ g s
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g 4 - Seismic . Motions - g E h WM
~
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. Z Fuct Assembfr y Group Lumped f. tass (ITP-)
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K (TYP-) [ o a
' u ,, _
a x TB + K. s h
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. FIDure 6.3 Spring Mass Simulation For A Two Dimensional Motion i
h 6-35 _ _ _ _ _ _ _ _ . _ _ _ TE
t - ()
' Fy(
Yd B i J B Fx;
~
=x '
(a) TO? V'EW Zi. . I Fz h
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. l
=F x ~
{ (b) AX'AL CROSS
//Niii////,/, -
S E C T ' Ol\ ( B-B .' FIG. f> .4 (a) - O R Z Ol\ TA _ C R O S S .
$.ECT ON O r RAC <
(b) VERT.CA _ C ROS S S ECT O \' 0.7 RAC<. . g e 6-36
CELL z(W) .
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9"79
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o
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) SUPPORTS FIG 6,5 DYNAMIC MODEL (RACK)
;Mz 1
jMy n n g Fz , . 4 / '{' B
=~M*
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. 6-37 I
y-t
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- E L,1 SEC.I ~ - - - -
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- a
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gROOT OF R ACK , -EL,8
?! e S E C,8 ' S E C,9 N O, O F E L E M E N T S = 8
. 'N O, O F S EC TI ON S' =9 N O. 0 F N O'D E S =5 FIG 6.7 SUBDIVISION OF A'TYR RACK 6-38 Y
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FIG .6,8 FINITE ELEMENT MODEL CROSS SECTidN e e 9 g h 6-39 , eng
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. . . . . . TIME. .(SEC. X .100 1. . . 6-40
......................................3
RUXILLIRRY POOL FIG. 6-10 E-W EXCITRTION (SSE) 9,25 , 0,20 - . e : l i l l i , l : 0,15 - T-- : - -- t - - - - - y 1 I i
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. . w n a i w a .. a i vvu FIG. 6-11 VERTICRL EXCITRTION (SSE) 0.11 7 , ; ,
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- 7. ACCIDENTS ASSOCIATED WITH RACK INTEGRITY AND LINER PLATE INTEGRITY In addition to the seismic analysis presented ir Section 6, the following computations for mechanical loads are also performed.
7.1 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 baseplate is acceptable; however, the rack design should preclude impact with the pool liner. The subcriticality of the adjacent fuel assemblies is not to be violated. Calculated results show that the baseplate is not pierced and the rack feet l
- loading on the liner is well below those caused by j seismic loads. The maximum depth of baseplate l penetration is conservatively estimated to be 0.446" (vs. 0.625" baseplate nominal thickness).
7.2 Dropped Fuel Accident II Onw fuel assembly dropping from 36 inches above the rack and hitting the top of the rack. Permanent deformation l 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 top of the active fuel (and l below) is not altered. Analysis indicates that the maximum local stress at the top of the rack is limited to 21000 psi which is less than material yield point. Thus, the functionality of the rack is,not affected. t . I 7-1 t m
7.3 Jammed Fuel-Handling Equipment and Horizontal Force A 4000-pound uplift force and a 1000-pound horizontal force are applied at the top of the rack at the
, " weakest" storage locations. The force is assumed to be applied on one wall of the storage cell boundary as an upward edge force over length 1 It is shown that if
, the length 1 is over 2.46" then no yielding will occur. If A is smaller than 2.46", the damage is limited to the
- region above the top of the active fuel. Horizontal force of 1000-pounds applied at the top edge of a cell wall produces plastic deformation over 2" depth - well removed from the zone of active fuel.
I 4 7.4 Liner Integrity Analysis
. The floor liner in both pools is 1/4" thick, ASTM 240-Type 304 plate material. Since the auxiliary pool has heavier rack modules, the liner loads developed in the auxiliary pool bound the problem. The time history loading of the liner due to shear between rack module and liner plate (pool floor) is obtained as a by-product of the seismic analysis (described in Chapter l 6). To investigate liner integrity, two load cases -
cases 1 6 5 from Chapter 6 - for rack A (heaviest rack module) are examined in detail. Case 1 refers to the condition when all storage locations are occupied, and the coefficient of friction y between liner and rack support feet is 0.8. Case 5 refers to the condition when all storage locations on one side of the diagonal (in plan form) are loaded leading to a highly asymmetric rack , loading. These two cases were identified in 7-2 _ . . _._ _ _____ L
Chapter 6 to be the most limiting in a kinematic and a structural sense.) Tables 7.1 and 7.2 give the instantaneous loadings (two shear and vertical) at each
-of the four support feet when x-shear and y-shear,
- respectively, at a support foot reach maximax value.
Referring to Table 7.1, maximum x-shear develops at foot 4 at 11.29 seconds. Similarly, maximum y-shear develops at foot 2 at 10.52 seconds into the earthquake. Tables 7.3 and 7.4 give sirailar data for loading cases 5 (diagonally loaded rack). In actuality, the shear loading of the liner due to ground excitation is a self-limiting loading since the liner displacement is kinematically limited. However, treating it as a primary loading, the maximum tensile stress in the liner can be calculated in a very conservative manner by considering the equilibrium of a 14" wide (width of support foot) strip. To add to the conservatism, friction between the liner and concrete grout is neglected. Table 7.5 gives the computed liner direct stress for data of Tables 7.1 thru 7.4 which are labeled Cases a thru d in Table 7.5. t 7-3
. e
l Table 7.1 Support reactions at the instant when x-shear is maximax (t=11.29 seconds, Rack A, p = 0.8, fully loaded, loading Case 1) Support X-Shear Y-Shear Vertical Notes Foot No. Gx(lb) G y(lb) Gg(lb) 1 0 0 0 Foot il has lifted off 2 -3437 -4007 -24590 - 3 26650 4648 -63280 - 4 51430 26520 -137600 Maximax reaction this foot Table 7.2 , Support Reactions at the instant when y-shear is maximax. (t=10.52 seconds, Rack A, u = 0.8, fully loaded) Support X-Shear Y-Shear Vertical Notes Foot No. Gx(lb) G (lb) G,(lb) 1 0 0 0 Support foot I has lifted off 2 -15410 -65660 -144200 Maximax reaction in this foot 3 -15000 -27090 -53970 - 4 -9393 5126 -14610 - 7-4 3_
Table 7.3 Support reactions at the time instant when x-shear is maximax (t=13.61 seconds, Case 5 loading) Support X-Shear Y-Shear Vertical Notes Foot No. Gx(lb) G y(lb) Gz(lb) 1 2546 6113 8277 - 2 0 0 0 Support foot 2 has lifted off. 3 17250 2543 -51390 - 4 41320 19910 -67180 This foot has maximax reaction. Table 7.4 Support reactions at the time instant when y-shear is maximax (t=14.74 seconds, Case 5 loading) Support X-Shear Y-Shear Vertical Notes Foot No. G,(lb) G y(lb) Gz(lb) 1 0 0 0 Support foot one has lifted off. 2 10700 14340 -55100 - - 3 12540 -46640 -70520 This foot has maximax y-shear 4 7823 -3640 -10790 - i e 7-5 7
c-i Table 7.5 Liner Direct Stress Case Foot G Vectorial Max. Liner Direct Sum of direct stress-(psi shear (1b) a 4 57741 16497 b 2 67444 19270 c 4 45866 13105 d 3 48296 13798
- Table 7.5 shows that the maximum liner stress is less than 19300 psi. The minimum tensile strength of 304 S.S. is 75000 psi 0 100"F. Therefore, the factor of safety against tearing is over 3.8.
These new results based on conservative analyses make consideration of liner strain analysis and cyclic loading unnecessary. 7.5 Dropped Gate The transfer canal gate is approximately four feet wide and weighs 7,000 lbs. The consequences of accidentally dropping this gate from a height of 15" (h=15") are considered. The gate l 1s assumed to fall with its fluid drag working on its smallest ! cross-sectional dimension. Furthermore, the impact on the top ! of the rack is assumed to take place along a lineal edge. l ! 7-6 l
Localized plastic deformation is not of concern here; however, gross buckling of cruciform panels is not acceptable. The following additional assumptions are made to render the analysis conservative. (i) The vertical walls of the cells are modeled as long ribbed plate columns 169" high x .063" wall. In reality, it is a composite ribbed plate column consisting of two plates .063" thick sandwiching an elastomeric material. The ribs are assumed to be 3" long and .063" thick at 6.26" pitch. (ii) The virtual mass of the gate in water is assumed to be equal to its displaced mass. (iii) Form and viscous drag of water is neglected. (iv) Top 1.25" of the cruciform walls under impact is considered to be crushed by the impact of the gate. With these assumptions, the equation of motion of the gate is given by
' (m + my ) dy ,g,,,) g dt where v downward velocity at time t t: time coordinate ma mass of gate m:y virtual mass . g: Acceleration due to gravity.
7-7 f I e
f Integrating these equations and utilizing the theory of impact loading of linear springs, the equivalent static load due to gate drop assuming no crushing of the top of the cruciform walls is given by FST = (2kEo) where Eo = (m-my )gh and, , k = spring constant of the " column" under consideration. However, the deformation of . 1/2 6=(E) = 0.63 k and a strain of c = O.5 results in part of the energy E being absorbed. Replacing Eo by the remaining energy Eo* ( = 3281 lb-in) and 8 lb/in, the pseudo static force is given by using k = .926 x 10
,' 1/2 r sr = (2kEo)
= 55120 lbs.
We now determine the critical buckling load using Eq. 3.75 p. 108 of " Buckling of Bars, Plates and Shells" 'by D.O. Brush, 8.0. Almroth, McGraw Hill, 1975. i l 7-8 m
b
=(h) [ (*-)
a Cgg +2 ("-) ab (C45 + C66) 4
+(1b ) C551 with Cgg =D+
EI* d x C55 = D C66 = (1-v) D + b ( GJ* ) 2 d X C45= vD Eh 3 DE 12(1-v2) G= 2(1+v) where D = bending stiffness parameter G = shear modulus E = Young's modulus I = stiffener moment of inertia relative to the plate x middle plane J x = torsional constant V= Poisson ratio i 7-9 ( 7
t 6 For this composite ribbed plate column, we have E=27.6x10 pgi, v=0.3, the plate length a=169", the width b=48", the thickness h=0.063", the stiffener or rib spacing d x =6.26" and the stiffener height and width equals 6" and 0.063", respectively. Using the above set of equations and data, we obtain the lowest critical load (with m=1) of P er
= 85008 lb.
The analysis presented in this section thus precludes the buckling of the rack under the impact of a gate drop from a height of 15" above the rack. Moreover, the licensee intends to develop and put in place administrative control for gate movement across the rack areas. In addition, redundant measures will be taken to preclude a gate drop. 7-10 . _O
8.0 SPENT FUEL POOL FLOOR STRUCTURAL ANALYSIS 8.1 Introduction The high density rack modules for long term fuel storage described in Section 2 are located in the spent fuel pool (auxiliary building). An overall schematic of the pool structure is given in Appendix I. In essence, the Grand-Gulf Unit I fuel pool slab is a reinforced concrete plate-type structure buttressed by a centrally located girder (I-beam). There are two column supports underneath the girder over the pool floor plan. The slab is further reinforced by imbedded I-beams in the lateral direction. Figure 8.1 shows a pictorial view of the slab cross section. Upper containment pool slab section has a considerably higher bending and shear strength than the spent fuel pool section. A comparison of Figures 2.1 and- 2.2 shows that the loading intensity on.the U.C.P. slab is significantly lower than that on the S.F.P. slab. Therefore, the analysis of the S.F.P. slab bounds the problem. In this section, pertinent results of pool slab analysis are presented to demonstrate its structural integrity for all postulated loading conditions. In particular, compliance with the strengh limits and load combination of ACI-349 Ill and NUREG-0800 I2I ,respectively, is shown. 8.2 Assumptions Seismic qualification of Spent Fuel Pool floor is carried out using the following conservative assumptions:
- 1. The pool floor is analyzed as a. simply supported, composite rectangular plate; no credit is taken for 8-1 7_
structural resistance offered by the adjacent pool walls. Cross beams and girders are modelled by plate elements.
- 2. Calculation of the stiffness and strength properties for- the concrete is based on the assumption of complete cracking _of the concrete in tension over the entire floor plan area.
- 3. The loading used to qualify the pool floor assumes
, that all racks are fully loaded with channelled fuel assemblies. 1
- 4. The ANSYS finite element code [3] is used to I determine the static stress state under a uniform pressure load on the floor.
i The input loading for dynamic analysis of the pool floor is l. obtained from the results of detailed ~ dynamic analysis of a , single fuel rack. The dynamic mass used on the floor slab analysis includes the concrete mass, the reinforcement mass, and the virtual miass of the water set in motion by the pool floor. j The pool floor stiffness properties assume that concrete is fully active on the compression side of the neutral axis, and is
- fully cracked on the tension side. The time history analysis of the poo1 floor is carried out using the Joseph Oat proprietary j computer code DYNAHIS. Output floor loads, obtained from a time history analysis of an individual fuel rack, are converted to a e floor pressure load time history acting on the entire floor slab, and used as the input dynamic load for the time history
! analysis of the pool floor. The results of the pool floor i analysis are scanned during computations, and the maximum floor
- deformation obtained during the complete seismic event is l considered as the primary output for further analysis. An -
equivalent s'tatic load, that yields the same value of maximum 1 8-2
.______R-
deformation, is then computed and used to perform the structural integrity checks by using a detailed finite element model of the f3oor under a static pressure. The finite element representation of the slab may be found in Figures 8.2-8.4. This effective static pressure is combined with the dead load pressure due to weight of the racks, the fuel assemblies, and the 40 foot head of water. 8.3 Dynamic Analysis of Pool Floor Slab to Obtain Maximum Floor Displacements With the dynamic model of the pool floor slab, considered as a simply supported rectangular orthotropic plate, a dynamic load history can be applied to the floor and be used to obtain the maximum displacements of the pool floor from the horizontal position. By equating the maximum displacements, obtained from an analysis over the total time of the seismic event, with the finite element solution for the static deformation of the floor configuration, we obtain a conservative estimate of the effective static pressure load on the pool floor slab. This effective pressure load can then be used in a standard strength qualification of the pool floor as outlined in SRP 3.8.4. The frequencies for the transient analysis are obtained from a detailed modal analysis of the pool floor using ANSYS. The dynamic load histories applied to the pool floor have been obtained from the results of dynamic analyses of a fully loaded Type A rack. The resulting load represents the algebraic sum, at each time point, of the loads in the four supports, and includes the dead load of the rack and fuel assemblies. This dead load is subtracted out prior to the floor transient analysis. For the purpose of a floor dynamic analysis, it is assumed that these load histories are representative of the averaged 8-3 7'
i pool floor loads from all of the different rack types, acting concurrently. This load is converted into a time history of floor pressure by dividing by the base area of a single rack A after removing the dead load component. Using the obtained l pressures as input to the floor slab time history analysis, the program DYNAHIS determines the pool floor displacement as a function of time, and as part of the output, gives the maximum displacement of the floor slab, 6 max. Structural damping, based on the lowest calculated pool floor natural frequency f i = 13 HZ, is incorporated into the model by modification of the structural stiffness matrix according to standard practices. To derive an effective static uniform pressure load, for subsequent strength analysis, 6 max is compared with the finite element solution for the statically loaded floor. The effective , pressure associated with the maximum dynamic deflection 6max is then obtained from the equation 9 9 s = 40 psi q* , g s) 6""* ; w, w, = max. deflection under 40 psi The following effective static pressures are obtained from the floor slab dynamic results: q,(SSE) = 2.92 psi 8.4 Results and Discussions Table 8.1 summarizes the loadings used in the qualification of the pool floor. The pressures are calculated based on a ra'ck footprint 122.5" x 104.25" = 12770.625 sq. in. ,
'g 8-4 a
i Tablo 8.1 Londing Dnto Loading Type, Computed Value (Pounds)
- 1. Dead weight of racks 22270
- 2. Weight of fuel assemblies (racks fully loaded) 202544
- 3. Weight of 40' head of water 227062
- 4. Dead weight of floor slab 60187.96 Total dead load pressure (Area = 12770.625 sq. in.) 40.093 psi
- 5. SSE Seismic load due to racks 1.57 psi
- 6. SSE Seismic load due to dead weight of floor slab 1.35 psi The time histories developed by Bechtel Power Corporation used a slightly smaller value of the dead loads on the slab (2 Ksf (Appendix I) vs. 2.53 Ksf given by above data). This difference, however, has an imperceptible effect on the pool floor time histories.
The fuel assembly weight utilized in the present analysis
> includes the weight of the fuel channel. Furthermore, all available fuel storage locations are considered to be occupied.
i In this manner, the total inertial mass of the fuel assembly is maximized and is treated as dead load (D) conjunctively with the pool slab mass and pool water mass. Using the notation of SRP 3.8.4, the following load combinations are deemed critical for the qualification of the pool floor. i B-5
}
I Load Combinations Concrete: The section moments and shear due to the following load combinations should not exceed the design strength, U. 1.4 D 1.05 (D + To) 1.05 D + To(D ++ 1.25 To) +E'j.75)(1.9)E Steel: The section moment and shear due to the following load combinations should not exceed the plastic moment, M p for the beam section. 1.7 (D + E) 1.3 (D + To + E) D + To + E' (See section 6 for definitions of above terms) Note that the dynamic impact loads on the pool floor, due to motion of the fuel racks, has been accurately included in the loadings.E'. We conservatively assume E = .667 E'. Table 8.2 shows a sampling of the structural integrity results that have been obtained for the pool floor. 8.5 Conclusion The pool floor has been shown to meet all structural acceptance requirements when conservatively analyzed as a simply supported rectangular plate with no credit taken for the supporting effects of the adjacent walls, or for the effect of hydrostatic loading on the walls causing a load reducing uplift on the floor. It is widely recognized that the term E' in this equation should be E. However, E' is used herein following the current text of SRP-3.8.4. This renders this load combination quite conservative. 8-6
Table 8.2 Synposis of Structural Acceptance Checks for Spent Fuel Pool Floor Governing Calculated Load Section Allowable Item Condition Loadings Strength Location Steel beams or 1.7 (D + E) 67.58x10 6 in.# 100.125x1g61nt Girder "Y" girders 1.7 (D + E) 5.86x10 6 in.# 20.88x10 in.# W36x150 Concrete slab 1.05(D+To)+1.42SE 196 KIP in/in 306 KIP in/in Element #4 (Fig. 8.3) Column 1.05D + 1.425E 5911 psi 16841 psi Column at Compression G&X function Girder shear 1.4D 679,610# 899100# Girder X Average edge shear 1.4D 9381#/in 57519/ int Slab (concrete edge) wall junction Edge shear on beam 1.4D 63245 307128# Beam W36 x 135 Shear Transfer 172 psi from concrete to beams (allowable panel pressure) t Capacity based on concrete alone.
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10 0 404 108 10 4 91 395 396 99 95 , Y ; 82 386 387 90 86 110 210 21E )( 73 377 378 81 77 64 368 369 72 68 55 359 360 63 59 - 46 350 3 51 54 50 37 341 342 45 41 Ell log 281 209 , 28 332 333 36 32 19 323 324 27 23 10 314 315 18 14 l 1 305 9 FIG. 8.2- NODE NUMBERS l 8-9 v- , . - . ,_wme-m----- - - - - , - - - - , - - - - . - - - - ,a--,,- n - - - - ,- - - - , - - - - - - "' --
i l 2 181 188 Y 17 3 18 0 X 14 8 15 5 126 133 109 11 6 92 99 75 82 52- t 59 35 42 18 25 I 8 FI G. 8.3 PLATE ELEMENTS 8-10
t ] 18 9 i 16 4 167 168 171 17 2 15 6 15 9 16 0 16 3 y l42 145,146 143,144 134 137 13 8 141 X 4 121
- 11 7 12 0 122 125 10 4 10 0 10 3 105 10 8 i
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References:
[1] American Concrete Institute, Code Requirements for Nuclear Safety Related Concrete Structures (ACI-349-76). [2] U.S. Nuclear Regulatory Commission, NUREG-0800, SRP 3.8.4., July, 1981. [3] Swanson Analysis Systems, Inc., ANSYS-User Manual, February 1982. f a e 8-12 1
9.0 ENVIRONMENTAL EVALUATION An environmental evaluation for the existing storage racks has been presented to the NRCl and has been approved.2 MP&L concludes that the expansion of storage capacity of the spent fuel pools will not have a significant radiological or non-radiological impact on the environment. 9.1
SUMMARY
Installation of High Density Spent Fuel Storage Racks at Grand Gulf Nuclear Station Unit-1, will increase the storage capacity of the spent fuel from 1270 to 4348 assemblies. In addition 45 defective fuel assemblies or control rod blades storage locations are provided. The upper containment pool temporary storage capacity is increased from 170 to 800 assemblies. Radiological consequences of expanding the capacity have been evaluated with the objective of determining if there is significant additional on site or off site radiological impact relative to that previously reviewed and evaluatedl -2 In addition, radiological impact to operating personnel has been evaluated to ensure that exposures remain As Low As is Reasonably Achievable (ALARA). The decay heat loading and the radiological burden to the spent fuel pool water are determined almost entirely by refueling operations. The frequency of refueling operations and the conduct of refueling are independent of the increased capacity of the storage pool, except that the increased capacity will reduce fuel movements and allow continued normal operation. Since the fuel assemblies which will utilize the bulk of the storage capacity (and 9-1
will ultimately fill all incremental. capacity above that of the existing design) are aged, their contribution to either the peak decay-heat load or the increased radiological impact, in terms of increased dose, would be negligibly small. A study performed by the NRC3 supports this conclusion. Consequently, -the increase in the storage capacity of both the spent fuel pool and the upper containment pool will neither significantly alter the operating characteristics of the current pool nor result in a measurable change in impact on the environment. 9.2 CHARACTERISTICS OF STORED FUEL The currently authorized storage capacity of the spent. fuel storage pool is 1270 assemblies and that of the upper containment pool is 170 assemblies. This capacity allows for two normal refueling discharges plus a full core discharge of 800 assemblies. The planned expansion of both the upper containment pool and the spent fuel pool would permit storage of 800 and 4348 assemblies in the respective pools. This expansion would allow storage of 18 annual discharges (see Table 1.1, Section 1) in the spent fuel pool with the capability to store a full core discharge in the upper containment pool. The decay heat generation rate af ter about 4 years of cooling 6 is only a small fraction (less than 2%) of the rate at 110 hours (the time at which fuel transfer from the core is assumed to commence) after shutdown. As shown in Figures 5.1.4 and 5.1.5 (Section 5), the highest decay heat loading and the corresponding peak temperatures in the spent fuel pool occur during the refueling period. At this time the Residual Heat Removal System provides additional i heat removal such that the bulk temperatures remain within l acceptable limits. Following discharge, the stored fuel i 9-2 L -
. ~
t I decay heat rate falls off substantially, hence the spent l fuel pool cooling system adequately handles the heat load. Because the majority of heat loading is due to freshly discharged fuel and since aged fuel contributes very little to the total heat load, the effect of expanding the spent fuel storage capacity is insignificant. Therefore, it is not expected that this expansion will significantly increase the thermal dissipation to the environment. Since the intensity of gamma radiation follows the decline in decay heat generation rate, it is similarly concluded, that there would be no significant increase in gamma radiation due to the expanded storage. It is important to note that the aged fuel in the expanded storage capacity pool will not contain significant amounts of radioactive iodine or short-lived gaseous fission products, since these would have decayed during the refueling period. The Krypton-85 which might escape from defective fuel assemblies has been shown to do no quickly3 (i.e. within a short time after discharge from . the core). Further, the residual Krypton-85 will be contained within the fuel pellet matrix and hence any leakage would occur at very low rates 3 Cesium 134/1373 is strongly bound within the fuel pellet matrix and its dissolution rate in water is extremely small. Any cesium dissolved in the pool water is easily controllable in the clean up system (demineralizer-ion exchange resin bed)3 Thus the planned storage expansion will not significantly increase the release of gaseous radionuclides. 9.3 RELATED INDUSTRY EXPERIENCE Experience with storing spent fuel underwater has been 9-3 n
substantial 3,7,8 These references show that the pool water activity, normally low, during refueling periods experiences a small increase which decays rapidly with ! time. ~ References 3 and 7 also' state that the increase in l pool water activity during refueling can be attributed to:
- a. dislodging (sloughing off) of corrosion products on the fuel assembly during transfer and handling operations.
- b. the possible short-term exposure of fuel pellets to pool water via a cladding defect, and
*c. mixing of the spent fuel pool water with the higher activity reactor coolant. Upon cessation of the refueling operations the fuel pool water and the reactor coolant system would be isolated from each other, . thereby terminating transport of corrosion products from the Reactor Coolant System. Thus, deposition of crud is a function of refuelling operations and is not impacted by the expanded storage.
Fur'hermore, t it has been shown9 that release of fission products from failed fuel' decreases rapidly after shutdown to essentially negligible levels. The fuel pellets are made of inert UO2 that have very low solubility in water and the propensity for corrosion of the cladding (Zircaloy 2) at spent fuel pool water temperatures is virtually nil.3r7 Thus the only mechanism available for the release of the gaseous fission matrix is diffusion through the UO2 pellet. It has been shown that at low water temperatures (<l50*F) the. diffusion coefficient is 9-4
extremely small10 Therefore, the small increase in activity of the spent fuel pool water is due to either crud transport, fission products release, or cross flow from the reactor coolant system and is only a function of refuelling operations. It is reasonable to assume that the increased capacity of the spent fuel pool will reduce fuel handling operations, since fuel assemblies will not have to be consolidated or shipped for an extended period, thereby reducing the probability of increased pool water activity due to crud dislodging. Thus, the expansion of fuel pool storage capacity will not cause a significant increase in dose either on site or off site. The corrosion properties of irradiated Zircalloy-2 cladding have been reviewed 3,8 and the conclusion is drawn that the corrosion of the cladding in spent fuel pool water is negligibly small. The minor incremental heating of pool water, due to the expansion of storage capacity, is far too small to materially affect the corrosion properties of Zircalloy-2 cladding. 9.4 OPERATING EXPERIENCE Grand Gulf Nuclear Station Unit 1 is not currently a commerical operating plant and, since no spent fuel has been discharged, an experience data base is not yet available. 9.5 SPENT FUEL POOL COOLING AND CLEANUP (FPCC) SYSTEM The fuel pool cleanup system at Grand Gulf Nuclear Station 9-5 7
Unit 1 is described in Section 9.1.3 of Reference 1. Radiological considerations are described in Chapters 11 and 12 and Section 15.7 of Reference 1. It has been shown previously (section 5) that the cooling systems at GGNS-1 are adequate to handle the expected heat loads with the additional heat removal capacity from the RHR system that would be required during refueling periods to maintain the temperature peaks within acceptable limits. The use of the RHR system to provide additional heat removal capability has been discussed for the GGNS-2 high density spent fuel storage pooll and was accepted by the NRC2 It has been shown earlier in this section that the small increase in heat load due to the storage capacity expansion of Unit-1 will neither significantly increase the thermal dissipation to the environment nor increase the propensity for corrosion of the cladding. It has been shown that the crud deposition in the spent fuel pool water occurs during refueling outages and that the planned expansion will not increase crud deposition. The fuel pool clean up system (filter and demineralizer) is designed to maintain fuel pool water clarity and is operated and maintained per Grand Gulf Nuclear Station Unit-1 operation procedures. The spent fuel pool filter demineralizer is backwashed and precoated when one of the two limits, either differential pressure or conductivity, is reached. The differential pressure set point is 25 psid and conductivity set point is 0.19 mhos. When these set points are exceeded an alarm sounds in the control room and the system operations procedure directs the operator to appropriate procedures for backwash and precoat. Furthermore, the Plant Chemistry
/
e 9-6
program at Grand Gulf Nuclear Station Unit 1, has a reading / recording schedule for differential pressure and conductivity. Neither inpact on the existing procedures nor a significant increase in activity of the cleanup system filters or resins is anticipated. i 9.6 RADIOLOGICAL CONSEQUENCES As stated earlier and confirmed by other studies 3,4,5,7,8,9 it can be shown that there will be no significant increase in activity due to Krypton-85, Cesium 134/137 or crud buildup on pool walls. It is concluded that the incremental impact from the release of either volatile fission products or crud with the expanded capacity of the spent fuel pool will be negligibly small. 9.7 RERACKING OPERATION The existing spent fuel racks, in the upper containment and spent fuel pool, are to be removed prior to the discharge of any spent fuel from the core. These racks have not been exposed to spent fuel and are not contaminated. Therefore, it is concluded that significant radiation dose to individuals involved in the raracking is not anticipated. 9-7
9.8 CONCLUSION
S Based upon the industry experience and evaluations discussed in previous sections, the following conclusions can be supported, o Minor increases in radiological burden to the pool water, if any, can be adequately handled by the fuel pool clean up system (filter and demineralizer), thereby maintaining the radionuclide concentration in the water at an acceptably low level. o No appreciable increase in solid radioactive wastes (i.e., filter media and demineralizer resin) is anticipated. o No increase in release of radioactive gases is expected, since any long-lived inert radioactive gas potentially available for release (i.e., Kr-85) will have leaked from the fuel either in the reactor core during operation or during ' the first few months of residence in the pool. Further, Vol. 1, Ref. 3 (pp. 4-16) has shown that airborne activity to be considerably lower than that allowable by Table 1 of 10CFR Part 20, Appendix B. Therefore, the planned expansion will not significantly increase the release of radioactive gases. o Based on industry experience to date, [ References 4 & 5 and the findings cited in Reference 3, pp. 6-16], no increase in radiological impact due to crud buildup either on the pool walls or by suspension in pool water is expected. 9-8
o The existing spent fuel pool cooling system, along with the added cooling capability achieved by the operation of the Residual Heat Removal System (RHR) . l during refuelling outages, will keep the pool water I temperature at an acceptable level [see Section 5-Thermal Hydraulic Considerations) o The existing radiation protection monitoring systems and program are adequate to detect and warn of any unexpected abnormal increases in radiation level. This provides sufficient assurance that personnel exposures can be maintained As Low As is Reasonably Achievable. o Since the re-racking operations will be performed prior to any spent fuel being placed in either pools, the existing racks are not expected to be contaminated. Hence removal and disposal of the existing racks will have no radiological impact. o Expanding the storage capacity of the spent fuel pool and the upper containment pool will not significantly increase the onsite or offsite radiological impact above that of the currently authorized storage capacity, nor is any significant increase in environmental radiological or non-radiological impact anticipated. 9-9 J
, , _ _ , v
REFERENCES
- 1. "FSAR", Grand Gulf Nuclear Station Units 1 & 2," Chapters 9, 11, 12 &-15, Docket No. 50-416 and 417.
- 2. NUREG 0831 " Safety Evaluation Report Related to the Operation of Grand Gulf Nuclear Station, Units 1 and 2".
Docket Nos. 50-416 and 50-417, Issued by USNRC, Office of Nuclear Regulation, Sept. 1981.
- 3. NUREG 0575, " Handling and Storage of Spent Light Water Power Reactor Fuel, Vol. 1, Executive Summary and Text, USNRC August 1979.
- 4. " Licensing Report on High-Density Spent Fuel Racks for Quad-Cities Units 1 and 2," Docket Nos. 50-254 and 50-265, Commonwealth Edison Company, June 1981.
- 5. " Licensing Report for High Density Spent Fuel Storage Racks", Rancho Seco Nuclear Generating Station, Sacramento Municipal Utilities District, Docket No. 50-312, June 1982.
- 6. NUREG 0800, USNRC Standard Review Plan - Branch Technical Position ASB9-2, Rev. 2, July 1981.
- 7. A.B. Johnson, Jr., " Behavior of Spent Nuclear Fuel in Water Pool Storage,:" BNWL-2256, September 1977.
- 8. . J.R. Weeks, " Corrosion of Materials in Spent Fuel Storage Pools", BNL-NUREG-2021, July 1977.
- 9. J.M. Wright, " Expected Air and Water Activities in the Fuel Storage Canal", WAPD-PWR-CP 1723, (with addendum) undated.
- 10. ANS 5.4 Proposed Standard, " Method for Calculating the Fractional Release of Volatile Fission Products from Oxide Fuel," American Nuclear Society, issued for review 1981.
9-10
.a.
- 10. INSERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING MATERIAL 10.1 Program Intent:
A sampling program to verify the integrity of the neutron absorber material employed in the hig h-density fuel racks in the l long-term environment is described in this section. The surveillance program is designed for the spent fuel pool since the Boraflex used in these racks will experience long term radiation. No surveillance program is planned for the Boraflex used in the upper containment (UCP) pool racks. These UCP racks are planned to be used for interim (short term) storage of spent
-fuels; therefore, they are not subject to long term radiation.
The program is intended to be conducted in a manner which allows access to the representative absorber material samples without disrupting the integrity of the entire 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. The surveillance program is to be utilized in the spent fuel racks auxiliary building pool. Since the spent fuel racks in the upper containment pool are utilized for interim (transitory) storage, a surveillance program for these racks is not needed. 10.2 Description of Specimens: . 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 5 3/4 x 3 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 tolerances similar to that design used for the storage system. The jacket has to be closed by tack welding.
10-1
10.3 Test: The test conditions represent the vented conditions of the box elements. The samples are to be located adjacent to the fuel racks and suspended from the spent fuel pool wall. Eighteen test samples are to be fabricated in accordance with Figure 10.1 and installed in the pool when the racks are installed. The procedure for fabrication and testing of samples is as given below:
- a. The samples should be cut to size and weighed carefully in milligrams.
- b. The length, width, and the average thickness of each specimen is to be measured and recorded.
- c. The samples should be fabricated in accordance with Figure 10.1 and installed in the pool.
- d. Two samples should be removed at each time interval according to the schedule shown in Table 10.1.
10.4 Specimen Evaluation: After the removal of the jacketed poison specimen from the fuel pool at a designated time, a careful evaluation of that specimen should be made to determine its actual condition as well as its apparent durability for continued function. Separation of the poison from the stainless steel specimen jacket must be performed carefully to avoid mechanical damage to the poison specimen. Immediately after the removal, the specimen and jacket section should visually be examined for any effects of environmental exposure. Specific attention should be directed to the examination of the stainless steel jacket for any evidence of physical degradation. Functional evaluation of the poison material can be accomplished by the following measurements: i 10-2
- a. A neutron radiograph of the poison specimen aids in the determination of the maintenance of uniformity of the boron distribution.
- b. Neutron ~ attenuation measurements will allow evaluation of the continued nuclear effectiveness of the poison. Consideration must be given, in the analysis of the attenuation measurements, for the level of accuracy of such measurements 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 durability. The hardness acceptability criterion requires that the specimen hardness will not exceed the hardness listed in the qualifying test document for laboratory test specimen irradiated 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, the width, and the average thickness and comparison with the pre-exposure data will indicate dimensional 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. 10-3
TABLE 10.1 Date Installed l INITIAL FINAL WEIGHT PIT WEIGHT WEIGHT . PENETRATION (mg/Cm2 -Yr) (mg/Cm2 -Yr) CHANGE-Yr) (mg/Cm SCHEDULE mil /Yr 1 2 90 day l f 3 4 180 day l f 5 6 1 Year U - 7 8 5 Year U 9 , 10 10 Year V 11 12 15 Year I I 13 I I 14 20 Year i 15 16 30 Year II 17 18 40 Year U s TIME SCH8DULE FOR REMOVING COUPONS 10-4 .
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11 0 COST / BENEFIT ASSESSMENT , A cost / benefit assessment has been prepared in accordance with the requirements of reference 1 Section V, Part 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 information; it 'is the position of MP&L that no environmental impact statement need be prepared in support of the request, because there will be no significant impact on the human environment. NRC precedent establishes that alternatives and economic costs need not be discussed when there is no significant environmental impact. However, for the sake of completeness, alternatives to re-racking, for additional spent fuel storage capacity, are' discussed in Section 11.3. 11.1 Specific Needs for Spent Fuel Storage Disposal of Grand Gulf spent nuclear fuel is scheduled to be carried out by the Department of Energy in or after 1998 in accordance with Public Las 97-425; Nuclear Waste Policy Act of 1982. As Grand Gulf spent fuel may not be accorded a high priority under the DOE program, MP&L is seeking to provide a spent fuel storage capacity to support approximately twenty years of nominal operation. No other contractual arrangements exist for the interim storage or reprocessing of spent fuel from Grand Gulf; therefore, increased storage capacity in the Grand Gulf fuel pool is the only viable option under consideration. Table 1.1, the fuel discharge schedule, indicates that with the high density spent fuel racks, loss of full core discharge capability
. (FCDC) will occur in 2003.
In addition to spent fuel, storage is available in the . Grand Gulf Unit 1 spent fuel storage pool for other types of ,/ materials (see Section 1), namely, 11-1
o Control rods
. -o Control rod-guide tubes.
o Defective fuel l 1 11.2 Cost of Spent Fuel Storage 4 The design ~and' manufacture of the spent fuel storage racks
.- will be undertaken by the organizations described in Section 1.
It is' expected that the total project cost will be between
. 3.3 and 3.5 million dollars.
11.3 Alternatives to Spent Fuel Storage d Mississippi Power & Light has considered the various alternatives to the proposed onsite spent fuel storage. These alternatives are as follows: o Shipment of fuel to a reprocessing or independent spent fuel storage / disposal facility No commercial spent fuel reprocessing. f acilities are presently operating in the United States. MSEI and h
- - MP&L have a -contractual arrangement whereby, spent nuclear fuel and/or high level nuclear waste will be .
i' accepted -and disposed of by the U.S. Department of Energy. However, such acceptance and disposal is not expected to begin before- 1998; the Grand Gulf Unit 1 i existing . fuel storage capacity will not provide full core discharge capability beyond 1987. Spent fuel acceptance and disposal by the Department of Energy is [ not, therefore, an alternative to increased on-site pool storage capacity. o Shipment of* fuel to another reactor site Shipment of Grand Gulf-l fuel to another reactor site 11-2 S
'- ~ - -r- e--n -
-w--ne -,y,g,.~g ,. _ _ _ , , _ , , , _ _ , , , , , , _ , p_, . _ , , _
could provide short term relief to the storage capacity problem. However, transshipment of spent fuel merely serves to transfer the problem to another site and does not result in ' any additional net long-term storage capacity. Accordingly, MP&L does not consider the
. transshipment of spent fuel to be an appropriate alternative to high-density spent fuel storage at the ; site.
o Not completing the reactor plant /not operating the plant after-the current spent fuel storage capacity is exhausted As indicated in NUREG-0575, " Final Environmental Impact Statement on Handling and Storage of Spent Light Water Power Reactor Fuel," the replacement of-nuclear power
"" cause by coal-generating capacity would excess
- ~ mortality to rise from 0.59-1.70 to 15-120 per year for 0.8 GWY(e). Based on these f acts, not operating the
' plant'or shutting down the plant after exhaustion of
- spent fuel discharge . capacity are not viable alternatives to high density storage in the spent fuel x
, pool. They rospective 1983 expenditure of approximately
, $3.5 15f ilion _for the high density racks is small-
-- compared to the estimated value of replacement power l .
equivalent to the plant's energy output: $590,000 per l day in 1983 and'S2,000,000 per day in 1990-1991.
,; e The sut ect of the comparative economics associated with svarioudspent fu'el options is the subject of Chapter 6- of' .
HUREG-0575. Although the material spresented is generic, it- v. . _is of L value in comparing the costs ots the various options. Of the options presented in Chapter 6 of NUREG-05752,high-density / spent , l fuel storage at the site is the most economic' option at $13 p$r
~- .
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I ~ , ef , w.: [ [ O . . , . . . _ _ , . _ _ , _ _ , _ _ _ _ _ _ _ . . _ _ _ ,_,_,N_.__ . _ , _ _ _ . _ _ . _ . _ _ _ _ _ _ _ _ _ , , _ . ,
l 1 KgU. The ' price of "Away From Reactor (AFR)" fuel storage, if available, would be $115 per KgU. This corresponds to 0.5 mill /Kwh from a 1000 MWe power reactor for AFR storage. The marginal cost per KgU of high density spent fuel racks for Grand Gulf-l is SS.18. 11.4 Resource Commitments The expansion of the Grand Gulf-l spent fuel storage capacity will require the following primary resources: o Stainless steel - 522,000 pounds o Boraflex neutron absorber - 41,000 pounds - of which 2800 pounds is Boron Carbide ( B i, C ) powder. The requirement for stainless steel represents a small fraction of the total domestic production for 1983.3 Although the fraction of domestic production of B i, C , required for the fabrication, is somewhat higher than that for stainless steel, it is unlikely that the commitment of B i, C to this project will affect other alternatives. Experience has shown that the production of B4C is highly variable and depends on need, but could easily be expanded to accommodate additional domestic needs. 11-4
- _ n
& ?
s REFERENCES TO SECTION 11, I s
- 1. . B.K. Grimes, "OT Position for Review and Ac'ceptance of Spent Fuel Storages and Handling Applications," April 14, 1978. t
- 2. NUREG-0575, " Handling and Storage of Spent Light-Water Power Reactor Fuel", Vol. 1-3, USNRC, August, 1979.
- 3. " Mineral Facts and Problems," Bureau of Mines Bulletin 671, 1980.
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- 12. OUALITY ASSURANCE PROGRAM l
12.1 Introduction This chapter provides a general description of the Quality Assurance Program that is implemented to assure that the quality objectives of the contract specification are met. 12.2 General The Quality" Assurance Program used on this project is based
.upon the system described in Joseph Oat's Nuclear Quality Assurance Manual; both of these meet the intent of 10CFR, Part 50, Appendix B. This system is designed to provide a controlled system for the design, manufacture and testing of customized components in accordance with various Codes, specifications, and regulatory requirements. The Joseph Oat Nuclear Quality Assurance
-Program has been accepted by ASME and found to be adequate by NRC audit team.
The philosophy behind Oat's Quality Assurance System is that it shall provide for all controls necessary to fulfill the contract requirements with sufficient simplicity to make it functional on a . As this system is applied to most of the day to day basis. contracts which Joseph Oat obtains, implementation of it is almost second nature to Oat's personnel. The system readily adapts to different designs and component configurations, making possible the construction of many varied forms of equipment. The highlights of this system, as addressed in the following paragraphs, provide an overview of the system and how it has been applied to the customer specifications and regulations. 12.3 System Hiahlichts: Design control is organized to provide for careful review of all contract requirements to extract each individual design and quality criteria. These criteria are translated into design and quality control documents customized to the contract requirements and completely reviewed and approved by responsible personnel. 12-1
~ The system for control of' purchased material entails generating detailed descriptions of each individual item of material along with specifications for any special requirements such as impact testing, corrosion testing, monitoring, or witnessing of chemical analysis, provision of overcheck specimens, special treatments or conditioning of matierial, source inspection,
-and provision of documentation of performance of any of the above.
Material receipt inspection includes a complete check of all material and its ' documentation. Upon acceptance, each item of material is individually listed on a control sheet issued once a week to assure that only accepted material goes into fabrication. The fabrication control system provides that a shop traveler is prepared for each subassembly and assembly in each contract. The. traveler is generated specifically to provide step by step instructions for fabrication, inspection, testing, cleaning,
-packaging, etc. which address all standard and special requirements of the contract specifications. Special attention is given' to deployment of fabrication sequence and inspection steps to preclude the possibility of missing poison cheets or incorrect sheets (incorrect B10 loading).
Due to the tendency of contract specifications to require special examination techniques or test procedures, all nondestructive examination procedures and test procedures are
. custom written to apply to each given component within a contract.
The system provides for- qualification and written certification of personnel performing quality related activities including- nondestructive examination and fabrication inspection, welding, engineering, production supervision and auditing. Other requirements of a solid quality conbrol system are fully covered as specified in the Quality Assurance Manual including 12-2 e
document control, cont.rol of measuring and test equipment, control of nonconforming material and parts, corrective action auditing and other areas as specified. 12.4 Summary: Joseph Oat Corporation's Quality Assurance System provides the full measure of quality assurance required by the contract. All special requirements of the specifications are covered including source inspection of material and witnessing of material testing by the Engineer, furnishing of material certifications and test reports within five days of shipment, and obtaining verification of qualification testing of poison materials. Oat has a long history of providing excellent quality control over a wide range of equipment types such as the high density fuel racks. 9 12-3 m
e G
- - . ..e -
b h APPENDIX I G O M
'I d
M " g S A
O REPORT ON SEISMIC ANALYSIS OF SPENT FUEL POOLS FOR HIGH DENSITY SPENT FUEL RACKS December 1982 GRAND GULF NUCLEAR POWER STATION, UNIT 1 MISSISSIPPI POWER AND LIGHT COMPANY 4 8 ' e e 0 e 4 e 1C0006
.e
.SEISMfC ANALYSIS OF SPENT FUEL POOLS FOR HfGH DENSITY SPENT FUEL RACKS ; 1. Introduction -
MP&L has decided to install High Density Spent Fuel Racks (HDSFR) in the t auxiliary building spent fuel pool at floor elevation 167'-6" and in the ~ l containment building upper pool at floor elevation 167'-6" in Grand Gulf l NuclearStation(GGNS) Unit 1(seeReference1). Seismic response time histories at the spent fuel pool floor level are required in three orthogonal directions (i.e., two horizontal and one vertical) for the design of HDSFR to withstand OBE and SSE loadings. To enable algebraic addition of the rack responses in the three
. directions, the earthquake motions specified in these directions will have to be statistically independent (Reference 2. Section 3.7.2). Also, the time histories will have to be in compliance with the Regulatory i Guide 1.60 (Reference 3) for Regulatory Guide 1.61 (Reference 4) dampings to be applicable.
Becht'el computer programs CE786, and CE917, CE931 CE920, and CE921 were used to modify input earthquake motion and to generate and verify the response time histories. ' ~
- 2. Mathematical Model
- 2.1 Auxiliary Buildino The lumped mass models shown in Figures 1 and 2 are obtained from the GGNS FSAR (Reference 5), Figures 3.7-19, and 3.7-20. All properties of these models are essentially the same as in the original models
. (Reference 6). Sloshing effects of water in the spent fuel pool subjected to horizontal excitations and new loads of the high density
, racks and fuel bundles (2000 psf per Reference 7) have been included in - the models.
. Figure 3A shows the actual lumped masses with W and W3 which representtheequivalentweightoffluidtopro8uceimpulsiveforceand
. equivalent oscillating weight to produce the convective force on the tank respectively. '
This model is simplified to the model on Figure 3B by lumping W to
. 'Mg and W tom,andignoringthespringstiffnessthatrepresentthe 1
effect of sloshing water. This simplified model will give structural response slightly more conservative than that of Figure 3A since the centroid of W and W in Figure 3A is slightly lower than the one on Figure 38. T$ewei (elevation 166'-0")gRtoftheractstevebeenappitedatthemasspoint4 1 I For vertical excitation, since there is no sloshing effect of water, the enly change in inass 4 from the sv in.11sedel is the new weight of racks y and fuel bundles. 100006
~ 2.2 Containment Building
- Figure 4 shews the lumped mass model of the containment building as obtained from the GGNS FSAR Figure 3.7-18. All properties of this model are essentially the same as in the original model (Reference 8). The only change in the model is the additional weight of the racks and fuel bundles to mass point 15 at elevation 161'-10" (2200 psf. per Reference 7). This elevation is four inches higher than the bottom of pool floor slab which has top of floor at elevation 167'-6" and bottom of floor at elevation 161'-6". Since the slab is rigid and has thickness of six feet, the displacements of top of floor and bottom of floor are almost identical. Thus, the response time histories at mass 15 represent the response at top of floor slab at elevation 167'-6" as required for design of the racks.
- 3. Ground Motion The basic ground motions have been derived from the Bechtel standard ground motions for two horizontal and one vertical directions, namely, H1, H2 and V as mentioned in the Bechtel Design Guide C-2.44 (Reference
- 9) " Seismic Analysis of Structures and Equipment for Nuclear Power Plants." These input ground motions are spectrum consistent in that they
' envelope the respective Regulatory Guide 1.60 Spectra. The ground
. motions were further evaluated to determine if they would envelope the Grand Gulf Design Spectra to comply with GGNS Licensing requirements.
Also, it was detemined that these Ground motions are statistically independent. The spectra of the horizontal ground motions H1 and H2 nomalized to 1 g for 2%, 4%, 5%, and 7% damping values are examined against the Grand Gulf Design Spectra as defined in the GGNS FSAR. The results showed that these horizontal ground motions will envelope the Grand Gulf Design Spectra designated as GGNS and the Regulatory Guide 1.60 Spectra, as shown in Figures 5 through 12. For vertical excitation, however, the ground motion V had to be modified to envelope the Grand Gulf Design Spectra. This is done by using a r spectrum raising procedure in accordance with Bechtel Computer Program CE
, 786. By using this program, the new vertical ground motion is obtained. -
e The corresponding spectra are shown in Figures 13 through 16. The'three ground motions, i.e., H1, H2, and modified V, are named
' . HORQUAKE 1. HORQUAKE 2, and VERTQUAKE for the two horizontals and one vertical directions respectively. Each time history has 4800 data points with an earthquake duration of 24 seconds (Time step = D.005 sec) and scaled to 1 g (see Figures 17,18, and 19).
The statistical independence between the three components of the earthquake ground acceleration time history has been verified by using the absolute values of correlation coefficients among the three . zamponents. Namely, in order to accept the three generated % m..h as
; I being statistically independent, the absolute values of all of the torrelation coefficients inust be less than or equal to 0.16 per Reference 6 of Regulatory Guide 1.92 (Reference 10).
1C0006 _ . - _, ._. _ _ - -i___ _ _ _ - - _ _ -. - N
o ! Absolute values computed as shownofbelow: correlation coefficients of these Ground Motions were f CORRELATION COEFFICIENT FOR Hy-H2 = .02413 CORRELATION COEFFICIENT FOR H -V = .03309 2 CORRELATION COEFFICIENT FOR V-H g = .12468 These values show that the input ground motions are statistically independent.
- 4. Analysis The time history analyses were performed by using the following Bechtel computer programs as shown in the flow chart below. The details of the analyses are given in Reference 11.
1 CE917 CE931 . CE917 Modal Properties Soil-Structure Interaction Modal Properties of of fixed base -e (composite damping & -9 flexible base structures participationfactors) structure (frequency
, and mode shapes)-
l Input: Time History ---D 9 P CE920 Response Time History
.:- V CE921 Response Spectra The andal ,..,-si.ies, i.e., natural frequencies, participation factors and mode shapes, of the lumped mass model with the bases considered fixed were obtained from CE 917. These modal properties together with appropriate structural damping are utilized in CE 931 to obtain the composite modal
- dampingandtheparticipationfactorsdueforsoilstructureinteraction(see 3ection 3.7 of Reference 5 for details). The romposite andal damping have been evaluated for two sets vf structural damping, i.e.,25 and 45 for DBE and 1
1C0006
5% and 7% for SSE events. The lower value of damping for each event (i.e., 2%
-and 5%) has been specified in the Section 3.7 of the GGNS FSAR while the .
d higher values (4% and 7%) have been obtained from the Regulatory Guide 1.61. The results of the analyses presented in the following sections indicate that the composite modal damping is rather insensitive to the indicated variations in the structural damping. The composite modal damping and the participation factors obtained from CE 931 are used in combination with the modal properties (frequencies and mode shapes) obtained from a flexible base analysis of the lumped mass model to determine the response time history for the input ground notion. However, Section 3.7.2.4 of the GGNS FSAR stipulates that, for the sake of conservatism, the computed composite modal dampings need not exceed 10% of. critical except for those modes that are clearly associated with rigid body translation or rotation of the structure. Since the input ground motions are consistent with Regulatory Guide 1.60 and the spectra used in the present analysis are clearly more conservative than that stipulated in the GGNS FSAR, it was decided that an exception to the GGNS FSAR may be considered if it were to be shown by analysis that the effect of the arbitrary cut off of damping at 10% is critical. The input ground motions are scaled to peak acceleration of 0.075g and 0.15g for DBE and SSE events. 4.1 Auxiliary Building 4.1.1 Free Vibration Analysis E-W Excitation: Mode Frea(cps) Participation Factor 1 4.01 69.42 2 - 8.69 -26.11 3 12.24 -1.06 t N-S Excitation:
. Mode Frea(cps) Participation Factor i
i 1 3.65 67.70 l .
. 2 8.91 29.78
- 3 14.46 -7.95 Vertite1 Excitatiom
! Mode Freq(cps) Participation Factor 3 6.08 74.45 2 25.64 -5.93 3 38.84 1.76 f . t -
' 1C0006
V , 4.1.2 Composite Madal Damping for Auxiliary Building E-W Excitation Structural Damping Mode 2% Damp. 4% Damp. 5% Damp. 7% Damp. 1 0.256 0.260 0.262 0.266 2 0.115* 0.120* 0.123* 0.128* 3 0.004 0.005 0.006 0.007 N-S Excitation Structural Damoing Mode 2% Damp. 4% Damp. 5% Damp. 7% Damp. 1 0.197 0.201 0.204 0.209 2 0.203* 0.209* 0.211* 0.217* 3 0.023 0.025 0.027 0.031 Vertical Excitation Structural Damping Mode 25 Damp. 4% Daup. 5% Damp. 75 Damp. 1 0.585 0.586 0.587 0.589 2 0.044 0.046 0.047 0.050 3 0.012 0.015 0.016 0.020
*The damping is suppressed to 10% for nonrigid body translations.
i _- IC0006
4.2
- Containment Building 4.2.1 Free Vibration Analysis E-W Excitation Mode Freq(cps) Participation Factor
- 1 0.21 8.98 2 2.49 50.15 3 4.80 4.86 4 5.34 21.78 5 6.54 6.97 6 7.57 -25.39 N-S Excitation Mode Frea(cps) Participation Factor
- 1 0.21 7.33 2 2.49 50.22 3 4.80 4.80 4 5.34 -
21.39 5 6.54 7.06 6 7.60 -25.20 Vertical Excitation Mode Freq (cps) Participation Factor 1 4.74 6.11 2 5.05 61.80 3 14.46 0.69 4 18.19 2.53
- Water Sloshing Mode e
'e e
g O O h
, 1C0006 y
4.2.2 Composite Modal Damping for Containment l l 0 Structural Damping i E-W Excitation Mode 2% Damp. 4% Damp. 5% Damp. 71 Damp. 1 .0128 .0123 .0121 .0118 2 .070 .0728 .0739 .0761 3 .0848 .0860 .0855 .0868 4 .0859 .9110 .1066 .1203 5 .0346 .0438 .'0474 .0548 6 .2232 .2277 .2300 .2354 Structural Damoing N-S Excitation
. Mode 2% Damp. 4% Damp. 5% Damp. 7% Damp.
1 .0273 .0223 .0209 .0194 2 .0704 .0726 .0737 .0759 3 .0844 .0855 .0849 .0863 4 .0823 .0889 .1048 .1174 5 .0344 .0436 .0472 .0548 6 .2222 .2269 .2292 Q348 Vertical Excitation
. Structural Damoing Mode 2% Damp. 4% Damp. 5% Damp. 7% Damp.
1 .0229 .0452 .0562 .0780 2 .6548 .6560 .6572 .6590 3 .0058 .0093 .0104 .0120
.;. 4 .0141 . .0190 .0215 .0270
- 5. ,
Res'ults The seismic response time histories for the auxiliary building and the containment at the specified locations for SSE with 5% and 7% structural dampings and DBE with 2% and 4% structural dampitigs are shown in Figures 20 to 43. e IC0006 m _
t The effect of cut off of composite model damping to 10% for nonri modes was rather small in increasing the response time histories.gid body Although the acceleration response time histories were computed both with and without cut off of composite modal damping, only the ones with cut off damping are shown here. It should be noted that for the containment building the composite modal dampings, which exceed 10%, are all associated with rigid body modes. M e o 8 , 3C0006 , m
References:
- 1. Letter from J. F. Pinto to R. S. Trickovic, BMP-82/224, dated June 23, 1982.
- 2. USNRCStandardReviewPlan(NUREG0800)
- 3. USNRC Regulatory Guide 1.60 Rev.1
- 4. USNRC Regulatory Guide 1.61.
- 5. Grand Gulf Nuclear Station Final Safety Analysis Report.
- 6. Bechtel Civil / Structural Calculation No. C-H002.5, Rev. O for Grand Gulf Nuclear Station.
- 7. Letter from J. F. Pinto to R. S.'Trickovic, BMP-82/263, date,d July 28, 1982.
- 8. Bechtel Civil / Structural Calculation No. C-G711.0, Rev.1, for Grand Gulf Nuclear Station.
S. Bechtel Power Corporation Design Guide C-2.44, Rev. O, dated August 1980.
- 10. USNRC Regulatory Guide 1.92, Rev. 1.
- 11. Bechtel Civil / Structural Calculation No. C-K600.0, Rev. O for Grand Gulf Nuclear Station. '
?
' 1C0006
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M ARK III CONTAINMENT ! MISSIS $3PPI POWER & LIGHT COMPANY GRAND GULF NUCLEAMSTATION
, UNITS 1 & 2 CONTAIPNENT 57RUCTutE l ! .50ll INTERACTED SEISMIC
, MATHEMATICAL tODEL Figure 4 CONTAINMENT BUILDING LUMPED MASS MODEL
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Attachnent 3 Neutron Absorber Traceability l A general description of the Quality Assurance (QA) Program utilized by the Joseph Oat Corporation during the fabrication of the GGNS high density l spent fuel racks (HDSFR) is described in Chapter 12 of the HDSFR Licensing Report. As described in Chapter 12, the Joseph Oat QA program was established to meet the intent of 10CFR50, Appendix B. In order to ensure proper inplanentation of the program requirements, MP&L's QA and Nuclear Plant Engineering organizations monitored various phases of the HDSFR construction and fabrication process. ) The acceptance criteria for criticality calculations is that k,ff be less than or equal to 0.95, including all uncertainties. To meet this criteria, the minim e B10 areal density of 0.0175 g/cm2 was established to determine the minima acceptable specific gravity of the neutron absorber (Boraflex). This ~ determination is explained in nore detail in Section 4.4.3 of the HDSFR Licensing Report. During the manufacture of the Boraflex, testing was conducted on sanples from each batch to establish that the mininun areal density was met. The Boraflex test docunentation for each batch, fabrication drawings, and shop travelers provide a basis for traceability of boron content throughout the rack manufacturing process. Thus, the Boraflex material used in each cruciform, 'T' sub-assecbly, and corner angle sub-assably of the racks can be traced to this batch docunentation. Based on the above discussion, MP&L is confident that the HDSFR manu-factured for GGNS meets the design criteria for boron content and criticality control.}}