ML20214P046
| ML20214P046 | |
| Person / Time | |
|---|---|
| Site: | Byron  |
| Issue date: | 08/31/1986 |
| From: | COMMONWEALTH EDISON CO. |
| To: | |
| Shared Package | |
| ML20214P037 | List: |
| References | |
| NUDOCS 8609170238 | |
| Download: ML20214P046 (173) | |
Text
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l LICENSING REPORT l .ON I HIGH DENSITY SPENT FUEL RACKS
$ FOR BYRON UNITS I AND 2 NRC DOCKET NO. 50 - 454 I 50 - 455 I
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O COMMONWEALTH EDISON COMPANY CHICAGO, ILLINOIS 60603 g
I AUGUST,I986 I
I pru888a 8889;r
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LICENSING REPORT
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.ON HIGH DENSITY SPENT FUEL RACKS FOR BYRON UNITS I AND 2 -
L NRC DOCKET NO. 50 - 454
[
50 - 455
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COMMONWEALTH EDISON COMPANY CHICAGO, ILLINOIS 60603 AUGUST,I986 e
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- TABLE OF CONTENTS Section .
Page
1.0 INTRODUCTION
1-1 2.0 GENERAL ARRANGEMENT 2-1 3.0 RACK CONSTRUCTION 3-1 3.1 Fabrication Details 3-1 3.1.1 Region 1 3-1 3.1.2 Region 2 3-3 3.2 Codes, 5tandards, and Practices for the Spent Fuel Pool Modification 3-4 4.0 NUCLEAR CRITICALITY ANALYSIS 4-1 4.1 Design Bases 4-1 4.2 Summary of Criticality Analyses 4-3 4.2.1 Normal Operating Conditions 4-3 I 4.2.2 Abnormal and Accident Conditions 4-7 4.2.3 New Fuel Storage 4-8 4.3 Reference Fuel Storage Cell 4-10 4.3.1 Reference Fuel Assembly 4-10 4.3.2 Region 1 Storage Cells 4-10 4.3.3 Region 2 Storage Cells 4-10 4.4 Analytical Methodology 4-15 4-15 lI .
4.4.1 Reference Analytical Methods and Bias 4.4.2 Fuel Bur..ap Calculations 4.4.3 Effect of Axial Burn-up Distribution 4-17 4-21 4-22
! 4.4.4 Long-term Decay 4.5 Region 1 Criticality Analysis and Tolerance Variations 4-24 4.5.1 Hominal Design Case 4-24 l
4.5.2 Boron Loading Variation 4-24 .
l i 4.5.3 Storage Cell Lattice Pitch Variation 4-25 .
4.5.4 Stainless Steel Thickness Tolerances 4-26 4.5.5 Fuel Enrichment and Density Variation 4-26 f 4-26 4.5.6 Boraflex Width Tolerance Variation
, 4.5.7 Axial Cutback of Boraflex 4-27
! 4.6 Region 2 Criticality Analysis and Tolerance '
Variations 4-28 4.6.1 Hominal Design Case 4-29 l$ 5 4.6.2 Boron Loading Variation 4.6.3 Storage Cell Lattice Pitch Variations 4-29 4-30 4.6.4 Stainless Steel Thickness Tolerance 4-30 l
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I TABLE OF CONTENTS (Continued) -
Section Page 4.6.5 Fuel Enrichment and Density Variation 4-30 l 4.6.6 Boraflex Width Tolerance 4-30 m 4.7 Abnormal and Accident Conditions 4-31 3 4.7.1 Eccentric Positioning of Fuel Assembly 4-31 g in Storage Rack 4.7.2 Temperature and Water Density Effects 4-31 4.7.3 Dropped Fuel Assembly Accident 4-32 4.7.4 Abnormal Location of a Fuel Assembly 4-34 l 4.7.5 Lateral Rack Hovement 4-35 4.8 New Fuel Storage 4-36 4.8.1 Storage in Region 1, Dry 4-36 4.8.2 Storage in Region 2, Flooded 4-36 4.8.3 Storage in Region 2, Dry 4-37 References to Section 4 4-38 5.0 THERHAL-HYDRAULIC CONSIDERATIONS 5-1 5 5.1 Decay Heat Calculations for the Spent Fuel 5-1 E 5.1.1 Basis 5-1 E 5.1.2 Model Description 5-3 5.1.3 Decay Heat Calculation Results 5-6 5.2 Thermal-Hydraulic Analyses for Spent Fuel 5-7 Cooling 5.2.1 Basis 5-7
. 5.2.2 Hodel Description 5-8 5.2.3 Results 5-9 References to Section 5 5-11 6.0 STRUCTURAL ANALYSIS 6-1 6.1 Analysis Outline ,
6-1 6.2 Fuel Rack-Fuel Assembly Model 6-3 6.2.1 Outline of Model 6-4 6.2.2 Model Description 6-6 6.2.3 Fluid Coupling 6-7 E 6.2.4 Damping 6-8 5 6.2.5 Impact 6-8 6.3 Assembly of the Dynamic Model 6-9 11 I
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F L TABLE OF CONTENTS (Continued) l Page Section 6.4 Time Integration of the Equations of Motion 6-11 l 6.4.1 Time-History Analysis Using 14 00F Rack Model 6-11 6.4.2 Evaluation of Potential for Inter-Rack 6-13 Impact t
6.5 Structural Acceptance Criteria 6-13 6.6 Material Properties 6-15 6.7 Stress Limits for Various Conditions 6-15
" 6.7.1 Normal and Upset Conditions (Level A or Level B) 6-15 -
6.7.2 Level D Service Limits 6-18 ;
6.8 Results 6-18 6.9 Impact Analyses 6-21 6.9.1 Impact Loading Between Fuel Assembly and Cell Wall 6-21 6.9.2 Impacts Between Adjacent Racks 6-21 r
L 6.10 Weld Stresses 6-22 6.11 Summary of Mechanical Analyses 6-22
{ 6-24 6.12 Definition of Terms Used in Section 6 References to Section 6 6-26 7.0 ENVIRONHEHTAL ANALYSIS 7-1 7.1 Summary 7-1 7.2 Characteristics of Stored Fuel 7-1
, 7.3 Related Industry Experience 7-2 7.4 Byron Nuclear Power Station Experience 7-4 l' Spent Fuel Pool Cooling and Cleanup System 7-4 l
7.5 7.6 Fuel Pool Radiation Shielding 7-5 7.6.1 Source Terms 7-5 7.6.2 Radiation Shielding 7-6 111
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TABLE OF CONTENTS (Continued)
Section Page 7.7 Radiological Consequences 7-7 7.8 Reracking Operation 7-8 7.9 Conclusions 7-9 References to Section 7 7-11 8.0 IN-SERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING MATERIAL 8-1 8.1 Program Intent 8-1 8.2 Description of Specimens 8-1 8.3 Specimen Evaluation 8-2 I
9.0 COST / BENEFIT ASSESSMENT 9-1 9.1 Specific Needs for Spent Fuel Storage 9-1 9.2 Cost of Spent Fuel Storage 9-2 9.3 Alternatives to Spent Fuel Storage 9-2 9.4 Resource Commitments 9-3 References to Section 9 9-5 10.0 QUALITY ASSURANCE PROGRAM 10-1 10.1 Introduction 10-1 10.2 Ceneral 10-1 10.3 System Highlights 10-1 10.4 Summary 10-3 Appendix A - Benchmark Calculations A-1 E:
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d u LIST OF TABLES F Page Table Byron Unit 1 & Unit 2 Fuel Assembly Discharge 1-3 l 1.1 (Tentative Schedule) 2-2 F 2.1 Design Data L
2-3 2.2 Module Data 3.1 Boraflex Experience for High Density Racks 3-7 Summary of Criticality Safety Analyses 4-5 4.1 l
' 4.2 Reactivity Effects of Abnormal and Accident 4-7 Conditions Fuel Assembly Design Specifications 4-12 4.3 Comparison of Cold, Clean Reactivities 4-19 4.4 Calculated at 25,000 MWD /HTU Burnup and 3.2%
Enrichment 4.5 Estimated Uncertainties in Reactivity due to 4-21
{ Fuel Depletion Effects Long-term Changes in Reactivity in Storage Rack 4-23 4.6 (Xenon-Free) 4.7 Fuel Burnup Values for Required Reactivities 4-28 (k.) with Fuel of Various Initial Enrichments 4.8 Effect of Temperature and Void on Calculated 4-32 Reactivity of Storage Rack 5.1 List of Cases Analyzed 5-12 5.2 Haximum Pool Bulk Temperature, t, Coincident 5-13 l
Total Power, Qi, and Coincident Specific Power, q, for the Hottest Assembly 5.3 Time (Hrs) to Boiling and Boiling Vaporization Rate From the Instant All Cooling is Lost 5-14 5.4 Maximum Local Pool Water Temperature and Local Fuel Cladding Temperature at Instant of Maximum Pool Bulk Temperature 5-15 y
s
I LIST OF TABLES (Continued)
Table Page 5.5 Pool and Maximum Cladding Temperature at the Instant Fuel Assembly Transfer Begins 5-16 6.1 Degrees of Freedom 6-27 6.2 Numbering System for Cap Elements and Friction 6-28 Elements 6.3 Rack Material Data 6-29 6.4 Support Material Data 6-29 6.5 Byron Racks - Bounding Values for Stress Factors 6-30 7.1 Photon Energy Production Rates of an Average 7-12 Spent Fuel Assembly 7.2 Photon Energy Production Rates of Peak 7-13 Spent Fuel Assembly 7.3 Calculated Dose Rates in Areas Adjacent to 7-14 the Spent Fuel Pool I
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s L LIST OF FIGURES Figure Page
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2.1 Pool Layout 2-4 2.2a Typical Rack Elevation - Region 1 2-5 H 2.2b Typical Rack Elevation - Region 2 2-6 3.1 3x3 Typ. Array Region 1 3-8 3.2 Channel Element - Region 1 & 2 (2 for Square Cell) 3-9
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3.3 Cap Elements Region 1 3-9 L 3.4 Composite Box Assembly . Region 1 '3-10 3.Sa Typical Cell Elevation - Region 1 3-11 7
3.5b Typical Cell Elevation - Region 2 3-12 3.6 Adjustable Support 3-13 3.7 3x3 Typical Array Region 2 3-14 4.1 Acceptable Burnup Domain in Region 2 of the 4-6 Byron Station Spent Fuel Storage Racks 4.2 Region 1 Storage Cell Geometry 4-13 6.3 Region 2 Storage Cell Geometry 4-14 4.4 Comparison of Depletion Calculations for 4-18 Fuel of 3.2% and 4.2% Initial Enrichments L 4.5 Reactivity Effect of water Spacing Between 4-33 Fuel Assemblies 5.1A Spent Fuel P.ool Bulk Temperature (0-700 hours) 5-17 (Hormal Refueling Discharge) f 5.18 Spent Fuel Pool Bulk Temperature (0-44 days) 5-18 L (Hormal Refueling Discharge) 1 5.1C Power Discharged in Spent Fuel Pool (0-700 hours) 5-19 (Normal Refueling Discharge) vil
I LIST OF FIGURES (Continued)
Figure Page 5.10 Power Discharged in Spent Fuel Pool (0-44 days) 5-20 (Normal Refueling Discharge) 5.2A Spent. Fuel Pool Bulk Temperature (0-1200 hours) 5-21 (Full Core Discharge) 5.2B Spent Fuel Pool Bulk Temperature (0-115 days) 5-22 (Full Core Discharge) 5.2C Power Discharged in Spent Fuel Pool (0-1200 hours) 5-23 (Full Core Discharge) 5.2D Power Discharged in Spent Fuel Pool (0-115 days) 5-24 5.3 (Full Core Discharge)
Idealization of Rack Assembly 5-25 I Thermal Chimney Flow Model 5-26 5.4 Byron Fuel Rack - SSE East / West 6-34 6.1 Byron Fuel Rack - SSE North / South 6-35 6.2 6-36 6.3 Byron Fuel Rack - SSE Vertical 6-37 E 6.4 Dynamic Model Cap Elements to Simulate Inter-Rack Impacts 6-38 6.5 Impact Springs and Fluid Dampers 6-39 6.6 Spring Mass Simulation for Two-Dimensional Motion 6-40 6.7 8-3 8.1 Test Coupon I
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1.0 INTRODUCTION
! This report describes the design, fabrication, and safety lg analysis of high density spent fuel storage racks manufactured by l3 3oseph Oat Corporation (Oat) for the Byron Station Unit 1 and Unit 2. The plant, which is located two miles east of the Rock River and approximately three miles southwest of Byron in Ogle l County, is owned and operated by Commonwealt'h Edison Company (CECO).
Byron is a two-unit pressurized water reactor (PWR) with a net design capacity of 1120 megawatts electric for each unit. Each of the two reactor cores contains 193 fuel assemblies and is rated to produce 3411 thermal megawatts (HWt). At present, there are no spent fuel assemblies stored in the spent fuel pool. Unit 1 went into commercial operation in September of 1985. Unit 2 is scheduled to go into commercial operation in June, 1987.
I The two units share one common spent fuel storage pool which is currently licensed for the storage of 1060 spent fuel assemblies. As shown in Table 1.1, the storage pool would lose full core discharge capability in 1994. The proposed reracking I will increase the number of pool storage locations to 2940 (includes six failed fuel locations). Table 1.1 indicates that the new racks will provide adequate storage with full core discharge capability well into the next century (circa 2011).
Table 1.1 is based on an estimated 18-month fuel cycle. Current trends toward longer cycles, extended burnup, and higher enrichment would further extend the time span 'of onsite storage.
The proposed racks are free-standing and self-supporting. The I principal construction materials are ASTH A-240, members and Type 304L shapes, and stainless steel for the structural "Boraflex," a patented product of BISCO (a division of Brand, Inc.), for neutron attenuation.
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I The specifications for design, construction, and quality assurance for the high density spent fuel storage racks were prepared by Sargent & Lundy Engineers (S&L) of Chicago, Illinois.
The mechanical design, seismic / structural analysis, thermal-hydraulic analysis, and other related calculations, and the fabrication of the hardware, were performed by Oat. S&L provided the seismic response spectra and performed the spent fuel pool structure evaluation. S&L performed the radiation shielding analysis. Southern Science, a division of Black & Veatch, served as a consultant to Oat in the area of criticality analysis. The analyses performed by Oat in conjunction with Black and Veatch and S&L demonstrate that acceptable margins of safety exist with respect to appropriate NRC and ASME acceptance criteria. A cost-benefit comparison of several potential spent fuel disposition alternatives indicates that (1) reracking of the Byron pool is the lowest risk and most cost-effective alternative, and (ii) that neither the reracking , operation nor ,
the increased onsite storage of irradiated material pose an undue hazard to the plant staff or the public.
The following sections provide a synopsis of the design, fabrication, nuclear criticality analysis, thermal / hydraulic analysis, structural analysis, accident analysis, environmental ,
analysis, and cost-benefit appraisal of the high density spent fuel racks. In particular, the integrity of the rack structure the specified combinations of inertial, seismic, and under mechanical loads and , thermal gradient per HUREG-0800 is demonstrated. -
Also included are descriptions of the rack In-Service Surveillance Program and the Oat Quality Assurance Program. This Quality Assurance Program has been reviewed and found acceptable for engineered fabrication of ASME Section III, Class 1, 2, 3 and HC Components by both ASME and the NRC.
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E Table 1.1 BYRON UNIT 1 & UNIT 2 FUEL ASSEMBLY DISCHARGE L (TENTATIVE SCHEDULE)
Remaining Remaining
- Total Discharged Storage Storage Assemblies in Capability Capability
( Number of Spent Fuel Pool Without With Assemblies Following Proposed Proposed Refueling Date Discharged Refueling Expansion Expansion w
February 1987 88 88 972 2852
[~ Unit #1 W 884 2764 l Unit #1 August 1988 88 176 88 264 796 2676 Unit #2 December 1988 Unit #1 February 1990 88 352 708 2588 440 620 2500
' Unit #2 3une 1990 88 88 528 532 2412 Unit #1 August 1991 88 616 444 2324 Unit #2 December 1991 February 1993 S4 700 360 2240 (Unit #1 88 788 272 2152 Unit #2 Dune 1993 Unit #1 August 1994 84 872 188** 2068 956 104 1984 Unit #2 December 1994 84 Unit #1 February 1996 84 1040 20
- 1900 84 1124 1816 Unit #2 June 1996 Unit #1 August 1997 84 1208 1732 1292 - 1648 Unit #2 December 1997 84 -
Partial core discharge capability lost - 84 assenblies
( *** Full core discharge capability lost - 193 assemblies 1-3
~ - - - - _ . _ _ _ _ - - - - ~ . _ - _ _ _ _ _ _ _ _ _____ __
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E Table 1.1 (Continued)
BYRON UNIT 1 & UNIT 2 FUEL ASSEMBLY DISCHARCE I: i (TENTATIVE SCHEDULE)
I' Remaining Remaining Total Discharged Storage Storage g Assemblies in Capability Capability B Number of Spent Fuel Pool Without With ,
Assemblies Following Proposed Proposed l Refueling Date Discharged Refueling Expansion Expansion Unit #1 February 1999 84 1376 1564 Unit #2 June 1999 84 1460 1480 Unit #1 August 2000 84 15 % 1396 Unit #2 December 2000 84 1628 1312 Unit #1 February 2002 84 1712 1228 Unit #2 June 2002 84 1796 11 %
Unit #1 August 2003 84 1880 1060 Unit #2 December 2003 84 1964 976 Unit #1 February 2005 84 2048 892 Unit #2 June 2005 84 2132 808 g Unit #1 August 2006 84 2216 724 5 Unit #2 December 2006 84 2300 640 g
Unit #1 February 2008 84 2384 556 3 Unit #2 June 2008 84 2468 472 Unit #1 August 2009 84 2552 388 Unit #2 December 2009 84 2636 ,
304 Unit #1 February 2011 84 2720 220 Unit #2 June 2011 84 2804 136**
Unit #1 August 2012 84 2888 52*
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2.0 CENERAL ARRANCEMENT The high density spent fuel racks consist of individual cells with 8.85-inch by 8.85-inch (nominal) square cross-section, each of which accommodates a single Westinghouse PWR fuel assembly or equivalent. A total of 2940 cells are arranged in 23 distinct 5 modules of varying sizes in two regions. Region 1 is designed for storage of new fuel assemblies with enrichments up to 4.2 L weight percent U-235. Region 1 is also designed to store fuel assemblies with enrichments up to 4.2 weight percent U-235 that have not achieved adequate burnup for Region 2. The Region 2 cells are capable of accommodating fuel assemblies with various initial enrichments which have accumulated minimum burnups within an acceptable bound as depicted in Figure 4.1. Figure 2.1 shows the arrangement of the rack modules in the spent fuel pool.
The high density racks are engineered to achieve the dual objective of maximum protection against structural loadings
{' (arising from ground motion, thermal stresses, etc.) and the maximization of available storage locations. In general, a greater width-to-height aspect ratio provides greater margin against rigid body tipping. Hence, the modules are made as large as pbssible within the constraints of transportation and site handling capabilities.
As shown in Figure 2.1, there are 23 discrete modules arranged in the fuel pool. Each rack module is equipped (see Figures 2.2a and 2.2b) with girdle bars, 5/8-inch-thick for Region 1 (1/4-inch-thick for Region 2) by 3-1/2 inches high. The nominal gap between adjacent modules is 1-1/4 inches for Region 1 and 3/4-inch for Region 2. The modules make surface contact between
[ their contiguous walls at the girdle bar locations and thus maintain a specified gap between them. Table 2.1 gives the relevant design data on each region. The modules in the two h
regions are of 11 different types. Table 2.2 summarizes the
( physical data for each module type.
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Table 2.1 DESIGN DATA (Cell Pitch) Flux Trap Nominal Min. B-10 Cap (Nominal)
Region in. Loading in.
1 10.32 H&S .020 gm/cm 2 1.16
& 10.42 E&W 1.26 2 9.011 010 gm/cm 2 0.0 I
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Table 2.2 MODULE DATA Approximate Module Number of Cells per Module Weight Rogion Type Modules Module Size (lb/ module) 1 A1 1 104 13x8 20,800.0 1 B1-3 3 96 12x8 19,200.0 I 26,900.0 2 C1-8 8 168 14x12 2 01-3 3 126 14x9 20,150.0 2 D4 1 116 14x9 18,550.0
-(2x2+3x2) 2 D5 1 114 14x9-(4x3) 18,250.0
[ 2 E1-2 2 112 14x8 17,900.0 2 F1 1 132 11x12 21,100.0 2 C1 '1 143 11x13 22,900.0 2 H1 1 56 7x8 8,950.0 2 31 1 35+6 7x5 10,150.0 failed fuel containers
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3.0 RACK CONSTRUCTION I 3.1 FABRICATION DETAILS I 3.1.1 Region 1 I The rack module is fabricated from ASTM A-240-304L austenitic l
stainless steel sheet and plate material, SA351CF3 and SA217CA15 I casting material. The weld filler material utilized in body welds .
is ASME SFA-5.9, Type 308L and 308LSI. Boraflex serves as the neutron absorber material. The detailed neutronic properties of Boraflex may be found in Section 4. The Boraflex experience list is given in Table 3.1.
A typical module contains storage cells which have an 8.85-inch nominal square cross-sectional opening. This dimension ensures fuel assemblies with maximum expected axial bow can be I that inserted and removed from the storage cells without any damage to the fuel assemolles or the rack modules.
Figure 3.1 shows a horizontal cross-section of a 3 x 3 array.
The cells provide a smooth and continuous surface for lateral contact with the fuel assembly. The anatomy of the rack modules is best explained by describing the components of the design, namely:
0 Internal square tube O Neutron absorber material (Boraflex)
I O Poison sheathing 0 Cap element 0 Baseplate O Support assembly 0 Top lead-in 3-1 I
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- a. Internal Square Tube This element provides the lateral bearing surface to the I
fuel assembly. It is fabricated by joining two formed channels (Figure 3.2) using a controlled seam welding operation. This element is an 8.85-inch square (nominal) cross-section by 168-7/8 inches long.
- b. . Neutron Absorber Material (Boraflex)
Boraflex is placed on all four sides of a square tube over a length of 139-1/2 inches, which covers the active fuel length except the top and bottom 3 inches, for Region 1. For Region 2 the Boraflex length is 144 inches and it covers the entire length of the active -
fuel length.
- c. Poison Sheathing .
The poison sheathing (cover plate)', shown in Figure 3.4, serves to position and retain the poison material in its designated space. This is accomplished by spot welding the ' cover sheet to the square tube along the former's edges at numerous (at least 20) locations. This manner of attachment ensures that the poison material will not l a
sag or laterally displace during fabrication processes
.and under any subsequent loading condition.
- d. CapLElement Gap elements, illustrated in Figure 3.3, position two inner boxes at a predetermined distance to maintain the minimum flux trap gap required between two boxes. The gap element is welded to the inner box by fillet welds.
An array of composite box assemblies welded as indicated l in Figure 3.1 form the honeycomb gridwork of cells which 5 harnesses the structural strength of all sheet and plate type members in an efficient manner. The array of g composite boxes has overall bending, torsional, and -
5 axial rigidities which are an or,er dof magnitude greater than configurations utilizing , grid bar type of construction. ,
- e. Baseplate The baseplate is a 5/8-inch thick plate type member which has 6-inch diameter holes concentrically located l
' with respect to the internal square tube, except at g support leg locations, where the hole size is 5 inches g in diameter. These holes provide the primary path for coolant flow. Secondary flow paths are available between adjacent cells via the lateral flow holes (1 inch in diameter) near the root of the honeycomb (Figures 3.5a and 3.5b) which preclude flow blockages.
The honeycomb is welded to the baseplate with 3/32-inch g fillet welds. E i
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- f. Support Assembly I Each module has four support legs.
adjustable in length to enable levellingconsists The supports are of the rack.
The variable height support assembly of a I flat-footed spindle which rides into an threaded cylindrical member.
internally-The cylindrical member is underside of the baseplate through attached to the I fillet and partial penetration welds. The base of the flat-footed spindle sits on the pool floor. Levelling of the rack. modules is accomplished by turning the square sprocket in the spindle using a long arm (approximately I 46 feet long) square head wrench. Figure 3.6 shows a support vertical cross-section of the adjustable assembly.
The supports elevate the module baseplate approximately 7-1/2 inches above the pool floor, thus creating t.he I water plenum for coolant flow. The lateral holes in path the cylindrical member provide .the coolant entry leading into the bottom of the storage locations.
I g. Top Lead-in Lead-ins are provided on each cell. .to facilitate fuel I assembly insertion. The lead-ins of contiguous walls of adjacent cells are structurally connected at the lead-in. These lead-in joints aid in reducing the
'I lateral deflection of the inner square tube due to the impact of fuel assemblies during the ground motion (postulated seismic motion specified in the FSAR). This I type of construction' leads to natural venting locations for the inter-cell material is located.
space where the neutron absorber I
3.1.2 Region 2 .
I The rack modules ' in Region 2 are fabricated from the same material as that used' for Region 1 modules, i.e., ASTM A-240-304L austenitic stainless steel.
I A typical Region 2 module storage cell also has an 8.85-inch Figure 3.7 shows a nominal square cross-sectional opening.
?
horizontal cross-section of a 3 x 3 array. The rack construction varies from that for Region 1 in as much as the stainless steel I
3-3 1
I I
and top lead-ins are eliminated. I cover plates, gap elements, Hence, the basic components of this design are as follows:
0 , Inner tube j 0 Neutron absorber material 0 Side strips O Baseplate O Support assembly In this construction, two channel elements form the cell of an 8.85-inch nominal square cross-sectional opening. The poison material is placed between two boxes as shown in Figure 3.7.
Stainless steel side strips are inserted on both sides of the poison material to firmly locate it in the lateral direction.
The bottom strip positions the poison material in the vertical direction to envelope the entire active fuel length of a fuel assembly (Figure 3.5b). Two adjacent boxes and the side strip between boxes are welded together as shown in Figure 3.7, to form the honey-comb rack module.
The baseplate and support assemblies are incorporated in for Region la the exactly the same manner as described 1 preceding section.
3.2 CODES, STANDARDS, AND PRACTICES FOR THE SPENT FUEL POOL MODIFICATION The fabrication of the rack modules is performed under a strict quality ass'urance system suitable for ASME Section III, Class 1, 2, and 3 manufacturing which has been.in place at Oat for over 10 years.
The following codes, standards, and practices were used as applicable for the design, construction, and assembly of the spent fuel storage racks and analysis of the pool structure.
Additional specific references related to detailed analyses are given in each section.
3-4 E
- a. Design Codes (1) AISC Manual of Steel Construction, 8th Edition, 1980.
(2) ANSI N210-1976, " Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations."
l (3) American Society of Mechanical Engineers (ASME),
I Boiler and Pressure Vessel Code,Section III, 1983 Edition up to and including Summer 1983 Addenda I, (Subsection NF).
I (4) ASNT-TC-1A June, 1980 American Society for I Hondestructive Testing (Recommended Practice for l
Personnel Qualifications).
- b. Material Codes (1) American Society for Testing and Materials (ASTM)
Standards - A-240.
(2) American Society of Mechanical Engineers (ASME),
l Boiler and Pressure Vessel Code,Section II - Parts I A and C, 1983 Edition, up to and including Summer 1983 Addenda.
l
- c. Welding Codes ASME Boiler and Pressure Vessel Code,Section IX -
Welding and Brazing Qualifications, 1983 Edition up to and including Summer, 1983 Addenda.
l Cleanliness, Packaging, Shipping,
- d. Quality Assurance, Receiving, Storage, and Handling Requirements I
(1) ANSI N45.2.2 - Packaging, Shipping, Receiving, Storage and Handling of Items for Nuclear Power Plants.
Cleaning of Fluid Systems and g (2) ANSI 45.2.1 -
Associated Components during Construction Phase of 5
Nuclear Power Plants.
l (3) ASME Boiler and Pressure Vessel, Section V, Nondestructive Examination, 1983 Edition, including Summer and Winter 1983.
(4) ANSI - N16.1-75 Nuclear Criticality Safety Operations with Fissionable Materials Outside Reactors.
l 3-5
I
- E (5) ANSI - N16.9-75 Validation of Calculation Methods for Nuclear Criticality Safety.
(6) ANSI - N45.2.11, 1974 Quality Assurance Require-ments for the Design of Nuclear Power Plants.
- e. Other References (1) NRC Regulatory Guides, Division 1, Regulatory Guides 1.13, Rev. 2 (proposed); 1.29, Rev. 3; 1.31, Rev. 3; 1.61, Rev. 0; 1.71, Rev. 0; 1.85, Rev. 22; 1.92, Rev. 1; 1.124, Rev. 1; and 3.41, Rev.1.
(2)- General Design Criteria for Nuclear PowerPart Plants, 50, Code of Federal Regulations, Title 10, g Appendix A (CDC Nos. 1, 2, 61, 62, and 63). g (3) HUREC-0800, Standard Review Plan (13781).
"0T Position for Review and Acceptance of Spent I,i (4) dated Fuel ' Storage and Handling Applications,"
modifications to this April 14, 1978, and the document of 3anuary 18, 1979.
I E
I 3.
I I
I. :
I!
3-6 Ii E
FL Table 3.1 L
BORAFLEX EXPERIENCE FOR HIGH DENSITY RACKS I Licensing Plant NRC Type Status
. Site Docket No.
L Point Beach 1 and 2 PWR 50-226 & 301 Licensed Nine Mile Point 1 BWR 50-220 Licensed l
Oconee 1 and 2 PWR 50-269 & 270 Licensed-Prairie Island 1 and 2 PWR 50-282 & 306 Licensed I Calvert Cliffs 2 PWR 50-318 Licensed Quad Cities
- 1 and'2 BWR 50-254 & 265 Licensed Watts Bar 1 and 2 PWR 50-390 & 391 Pending Waterford 3 PWR 50-382 Pending L BWR 50-341 Licensed Fermi
- 2 H. B. Robinson 2 PWR 50-261 Licensed
{
River Bend 1 BWR 50-458 Licensed Rancho Seco
- 1 PWR 50-312 Licensed Nine Mile Point 2 BWR 50-410 To be ap-plied for Shearon Harris 1 PWR 50-400 To be ap-plied for Millstone 3 PWR 50-423 To be ap-plied for Crand Gulf
- 1 BWR 50-416 Pending Oyster Creek
- BWR 50-219 Licensed
{
V. C. Summer
- PWR 50-395 Licensed Diablo Canyon
- 1 and 2 PWR 50-275 & 323 -Licensed
- 3oseph Oat Corporation fabricated racks
~
3-7
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. COVER GAP ELEMENT '
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FIGURE 3.7 3X3 TYPICAL ARRAY REGION 2 3-14
- E'
E 4.O NUCLEAR CRITICALITY ANALYSIS l 4.1
~
DESIGN BASES The high density spent fuel storage racks for the Byron Nuclear Power Station are designed to assure that the neutron multiplication factor (kegg) is equal to or less than 0.95 with the racks fully loaded with fuel of the highest anticipated reactivity in each of two regions, and flooded with unborated j water at a temperature corresponding to the highest reactivity.
The maximum calculated reactivity includes a margin for uncertainty in reactivity calculations and in mechanical tolerances, statistically combined, such that the true keff Will I, be equal to or less than 0.95 with a 95% probability at a 95%
l confidence level.
I l Applicable codes, standards, and regulations, or pertinent sections thereof, include the following:
e General Design Criterion 62, Prevention of Criticality in I Fuel Storage and Handling.
g e USNRC Standard Review Plan, NUREG-0800, Section 9.1.1, g New Fuel Storage, and Section 9.1.2, Spent Fuel Storage, e USNRC letter of April 14, _1978, to all Power Reactor Licensees -
OT Position for Review and Acceptance of I Spent Fuel Storage and Handling Applications, including l modification letter dated January 18, 1979.
e USNRC Regulatory Guide 1.13, Spent Fuel Storage Facility Design' Basis, Rev. 2 (proposed), December 1981.
l e USNRC Regulatory Guide 3.41, validation of Calculational Methods for Nuclear Criticality Safety (and related ANSI N16.9-1975).
I e ANSI /ANS-57.2-1983, Design Requirements for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Plants.
I 4-1 1
I o ANSI N210-1976, Design Objectives for Light Water Reactor I.
Spent Fuel Storage Facilities at Nuclear Power Plants.
e ANSI N18.2-1973, Nuclear Safety Criteria for the Design 5 of Stationary Pressurized Water Reactor Plants.
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 cor-responding to the highest reactivity.
e Lattice of storage racks is assumed infinite in all di-rections, i.e., no credit is taken for axial or radial neutron leakage (except in the assessment of certain abnormal / accident conditions).
e Neutron absorption in minor structural members is-neglected, i.e., spacer grids are replaced by water.
The design basis fuel assembly is a 17 x 17 Westinghouse g
optimized fuel assembly containing U02 at a maximum initial 5 enrichment of 4.2% U-235 by weight, corresponding to 48.6 grams U-235 per axial centimeter of fuel assembly. Two separate storage regions are provided in the spent fuel storage pool, with separate criteria defining the highest anticipated reactivity in each of the two regions as follows:
e Region 1 is designed to accommodate new fuel with a maximum enrichment of 4.2 wt% U-235, or spent fuel regardless of the discharge fuel burnup.
e Region 2 is designed to accommodate fuel of various initial enriclunents which have accumulated minimum g burnups within an acceptable bound as depicted in Fig. g 4.1.
I I
4-2 I
I
~ 4.2
SUMMARY
OF C'.ITICALITY P ANALYSES m
4.2.1 Normal Ocerating Conditions The criticality analyses of each of the two separate regions of the spent fuel storage pool previously described are sum-marized in Table 4.1 for the anticipated normal storage condi-tions. The calculated maximum reactivity in Region 2 includes a burnup-dependent allowance for uncertainty in depletion calcu-J lations and, furthermore, provides an additional margin of more than,2% t.k below the limiting effective multiplication factor (kefg) of 0.95. As cooling time increases in long-term storage, d decay of Pu-241 results in a significant decrease in reactivity, which will provide an increasing subcriticality margin and tends I to further compensate for any uncertainty in depletion calcula-tions. Spacing between the two different rack modules is suf-ficient to preclude adverse nuclear interaction between modules.
I Region 2 can accommodate fuel of various initial enrichments and discharge fuel burnups, provided the combination falls within the acceptable domain illustrated in Fig. 4.1. For convenient I reference, the minimum burnup values in Fig. 4.1 have been fitted by linear tangents at various values and the results are I- tabulated below, I
j Initial Minimum Initial Minimum Enrichment, % Burnup, MWD /MTU Enrichment, % Burnup, MWD /MTU I s 1.52 0 3.00 22,490 1.80 5,230 3.20 25,000 8 2.00 8,570 3.40 27,510
} 2.20 11,340 3.60 30,020 I 2.40 14,390 3.80 32,540 2.60 17,050 4.00 34,960 2.80 19,700 4.20 37,370 4-3 5
I Linear interpolation between the ' tabulated values will always yield values on or conservatively above the curve of limiting burnups. j These data will be . implemented in appropriate administrative procedures to assure verified burnup as specified in draf t Regu-latory Guide 1.13, Revision 2. Administrative procedures will also be employed to confirm and assure the presence of soluble poison in the pool water during fuel handling operations, as a further margin of safety and as a precaution in the event of fuel misplacement during fuel handling operations as discussed in Section 4.2.2.
I I
I I
I
. I I
I I
I 4-4 I
I
~
Table'4.1 I
SUMMARY
OF CRITICALITY SAFETY ANALYSES I Region 1 Region 2 lI Minimum acceptable burnup 0 37,370 MWD /MTU
@ 4.2% initial enrichment Temperature assumed O'C 0*C for analysis l Reference k, (nominal) 0.9374 0.8999 Calculational bias 0.0013 0.0013 Uncertainties I
i Bias iO.0018 i0.0018 l B-10 concentration *0.0021 i0.0028 Boraflex thickness iO.0047 iO.0078 Boraflex width i0.0007 i0.0009 Inner box dimension iO.0018 t0.00ll Water gap thickness k0.0038 NA SS thickness iO.0025 *0.0001 l
Fuel enrichment iO.0024 i0.0024 Fuel density iO.0026 i0.0026 Eccentric assembly negative negative l position '
Statistical i0.0082 0.0093 combinatign(l)
Allowance for NA +0.0187 burnup uncertainty Total 0.9387
- 0.0082 0.9199 i 0.0093 Maximum reactivity 0.9469 0.9292 I
(1) Square root of sum of squares.
I 4-5 I
l
.... .... .... . ... .... .... I i -
E 35 a -
AC OEPTAHLE l
r - BURNUP OC' MAIN -
% 30 o / -
l
" ~
~
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o .
- e. 20 o .
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- 5 z -
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- - UNAC CEPTAIBLE -
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' ' ' ' ' ' 3 0
2.0 2.5 3.0 3.5 4.0 4.5 1.5 INITIAL ENRICHMENT,WTX U-235 rw.1 accz,Tietz eusse, ocsx1s zu STATION S ENT ,UEL STORAGE RACKS.
ese10s 2 c, Tes exscs I 4-6
(
4.2.2 Abnormal and Accident Conditions
( .
f Although credit for the soluble poison normally present in the spent fuel pool water is permitted under abnormal or accident conditions,* most abnormal or accident conditions will not result L. in exceeding the limiting reactivity (k egg of 0.95) even in the absence of soluble poison. The effects on reactivity of credible r abnormal and accident conditions are summarized in Table 4.2 L
below. Of these abnormal / accident conditions, only one has the potential for a more than negligible positive reactivity effect.
(
' Table 4.2
{
REACTIVITY EFFECTS OF ABNORMAL AND ACCIDENT CONDITIONS l
l Accident / Abnormal Conditions Reactivity Effect l
l Temperature increase Negative in both regions l
Void (boiling) Negative in both regions Assembly dropped on top of rack Negligible Lateral rack module movement Negligible l
Misplacement of a fuel assembly Positive l
l l .
I
- Double contingency principle of ANSI N16.1-1975, as specified in the April 14, 1978 NRC letter (Section 1.2) and implied in the proposed revision (draft) to Reg. Guide 1.13 (Section 1.4, Appendix A).
4-7
The inadvertent misplaccment of a new fuel assembly (either li.to a Region 2 storage ' cell or outside and adjacent to a rack mauule) has the potential for exceeding the limiting reactivity siould there be a concurrent and independent accident condition tasulting in the loss of all soluble poison. Administrative procedures to assure the presence of solubla poison during fuel handling operations will preclude the possibility of the simul-taneous occurrence of these two independent accident condi-tions. The largest reactivity increase occurs .for accidentally placing a new fuel assembly into a Region 2 storage cell with all other cells fully loaded. Under this condition, the presence of g only 300 ppm soluble boron assures that the infinite' multipli- a cation factor would not exceed the design basis reactivity for Region 2. With the nominal concentration of soluble poison present (2000 ppm boron), the maximum reactivity, k, , is less than 0.95 even if Region 2 were to be fully loaded with fresh fuel of 4.2% enrichment.
4.2.3 New Fuel Storage Region 1 of the storage racko is designed to safely' accommo-date new unirradiated fuel of 4.2% enrichment, when fully flooded i with clean unborated water. Under certain circumstances, it may be desirable to store new fuel in the dry condition in Region 1 or to utilize Region 2 for the temporary storage of new fuel, either dry or fully flooded. These conditions were analyzed to assure the acceptability of Region 1 in the dry condition and to determine an arrangement in Region 2 that would assure criti-cality sa'f ety in conformance with the requirements of SRP 9.1.1, "New Fuel Storage."
Criticality analyses confirmed that Region 1 does not exhibit a peak in reactivity at low moderator densities (e.g.,
fog or foam moderation) and that the optimum moderation (highest kegg) occurs for the fully flooded condition. This condition is 4-8 I
I
[
r the design basis for Region 1 where the maximum k , including all I
- uncertainties, is less than 0.947. ,
I L In Region 2~, it was determined that a checkerboard pattern (fuel assemblies aligned diagonally) provided an acceptable k, in either the fully flooded or the dry (low density moderation) condition for new fuel assemblies of 4.2% enrichment. .These calculations indicated a nominal k, of 0.813
- 0.014 ( la ) when fully flooded with clean unborated water--a value substantially less than the limiting k,gg of 0.95, even with an additional allowance for uncertainties (maximum k, of ~0.86 at 95%/95% tol-erarice limits) .
Calculations, using Monte Carlo techniques, did not reveal a peak. in reactivity at low moderator densities, and the fully flooded condition corresponds to the highest reactivity (optimum moderation). Thus, the checkerboard pattern of new 4.2% enriched fuel ~in Region 2 represents a safe configuration in conformance with SRP 9.1.1 and 9.1.2.
{
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4.3 REFERENCE FUEL STORAGE CELL I
4.3.1 Reference Fuel Assembly The design basis fuel assembly, illustrated in Fig. 4.2, is a 17 x 17 array of fuel rods with 25 rods replaced by 24 control rod guide tubes and 1 instrument thimble. Table 4.3. summarizes the Westinghouse optimized fuel assembly (OFA) design specifi-cations'and the expected range of significant variations.
4.3.2 Region 1 Storage Cells' The nominal spent fuel storage cell used for the critic'ality analyses of Region 1 storage cells is shown in Fig. 4.2. The rack is composed of Boraflex absorber material sandwiched between a 0.060-inch inner stainless steel box and a 0.020-inch outer l stainless steel (SS) coverplate (0.125-inch coverplate for module periphery cell walls). The fuel assemblies ~are centrally located in each storage cell on a nominal lattice spacing of 10.320
- 0.050 inches in one direction and 10.420 t 0.050 inches in the other direction. Stainless steel gap channels connect one storage cell box to another in a rigid structure and define an outer water space between boxes. This outer water space con-stitutes a flux-trap between the two Boraflex absorber sheets g
i that are essentially opaque (black) to thermal neutrons. The E Boraflex absorber has a thickness of 0.075
- 0.007 inch and a nominal B-10 areal density of 0.0238 gram per cm 2, 4 . 3 . 3' Region 2 Storage Cells Region 2 storage cells were initially designed for fuel of 3.2 wt% U-235 initial enrichment burned to 25,000 MWD /MTU and extended to encompass fuel of 4.2% initial enrichment burned to g
37,370 MWD /MTU. In this region, the storage cells are composed W of a single Boraflex absorber sandwiched between the 0.060-inch 4-10 I
I
- - . - - - . , . . ~ - -m
lI
- stainless steel walls of adjacent storage cells. These cells, shown in Fig. 4.3, are located on a lattice spacing of 9.011 i l
- 0.040 inches.
!I lI 1
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I 4-11 I
Table 4.3 FUEL ASSEMBLY DESIGN SPECIFICATIONS I
Fuel Rod Data Outside diameter, in. 0.360' Cladding thickness, in. 0.0225 Cladding material Zircaloy-4 Pellet diameter, in. 0.3088 UO2 pellet density, % TD 95
- 2 UO2 stack density, g/cm 3 10.288
- 0.217 Enrichment, wt% U-235 4.2
- 0.05 I
Fuel Assembly Data Number of fuel rods 264 (17 x 17 array)
Fuel rod pitch, in. 0.496 I Control rod guide tube 3 Number 24 5 Outside diameter, in. 0.474 Thickness, in. 0.016 Material Zircaloy-4 l Instrument thimble Number 1 Outside diameter, in. 0.474 Thickness, in. -
0.016 Material Zircaloy-4 U-235 loading g/ axial cm of assembly 48.6
- 1.0 I
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r 8.85" 1 0.032"
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1.26" 2 0.040" Water Cap y Cap Channel
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Fig. 4.2 REGION 1 STORAGE CELL GEOMETRY.
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! ! S s.se0 2 0.032=. sax I.o. =j 5 j=
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Fig. 4.3 REGION 2 STORAGE CELL GEOMETRY.
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E L
7 4.4 ANALYTICAL METHODOLOGY L
4.4.1 Reference Analytical Methods and Bias The CASMO-2E computer code (Refs. 1, 2, and 3), a two-L dimensional multigroup transport theory code for fuel assemblies, has been benchmarked (see Appendix A) and is used both as a primary method of analysis and as a means of evaluating small reactivity increments associated with manufacturing tolerances.
CASMO-2E benchmarking resulted in a calculational bias of 0.0013
- 0.0018 (95%/95%).
L In fuel rack analyses, for independene vericication, criei-cality analyses of the high density spent fuel storage racks were also performed with the AMPX-KENO computer package (Refs. 4 and 5), using the 27-group SCALE
- cross-section library (Ref. 6) with the NITAWL subroutine for U-238 resonance shielding effects (Nordheim integral treatment). Details of the benchmark calcu-
{ 1ations with the 27-group SCALE cross-section library are also I presented in Appendix A.
in a bias of 0.0106 i 0.0048 (95%/95%).
These benchmark calculations resulted In the geometric model used in KENO, each fuel rod and its cladding were described explicitly. For two-dimensional X-Y analysis, a zero current (white albedo) boundary condition was l applied in the axial direction and, for Region 1, at the center-line through the outer water space (flux-trap) on all four sides of the cell, effectively creating an infinite array of storage I
cells. In Region 2, the zero current boundary condition was applied at the center of the Boraflex absorber sheets between SCALE is an acronym for Standardized Computer Analysis for Licensing _E_ valuation.
l 4-15 I
I storage cells. The AMPX-KENO Monte Carlo calculations inherently I
include a statistical uncertainty due to the random nature of neutron tracking. To minimize the statistical uncertainty of the KENO-calculated reactivity, a total of 50,000 neutron histories is normally accumulated for each calculation, in 100 generations l
l of 500 neutrons each.
I CASMO-2E is also used for burnup calculations, with inde-I pendent verificat, ion by EPRI-CELL and NULIF calculations. In tracking long-term (30-year) reactivity effects of spent fuel stored in Region 2 of the-fuel storage rack, EPRI-CELL calcula-tions indicate a continuous reduction in reactivity with time l
(after Xe decay) due primarily to Pu-241 decay and Am-241 growth.
l A third independent method of criticality analysis, util- g izing diffusion / blackness theory, was also used for additional a co,nfidence in results of the primary calculational methods, although no reliance for criticality safety is placed on the l
reactivity value from the diffusion / blackness theory technique.
l This technique, however, is used for auxiliary calculations of small incremental reactivity effects (e.g., axial cutback or mechanical tolerances) that would otherwise be lost in normal KENO statistical variations, or would be inconsistent with CASMO-2E geometry limitations.
Cross sections for the diffusion / blackness theory calcula-tions were derived from CASMO-2E or calculated by the NULIF com-puter, code (Ref. 7), , supplemented by,a blackness theory routine that effectively imposes a transport theory boundary condition at the surface of the Boraflex neutron absorber. Two different spatial diffusion theory codes, PD007 (Ref. 8) in two dimensions I
I 4-16 I
.- - ..-6
f L
and SNEID in one diniension, we're used to calculate reactivi-ties. The two-dimensional PD007 code was used to describe 'the actual storage cell geometry, with NULIF cell-homogenized con-stants representing each fuel rod and its associated water moderator. SNEID is a one-dimensional model, in cylindrical or L slab geometry, used for the calculation of axial cutback re-activity effects and in the assessment of abnormal occurrences.
I 4.4.2 Fuel Burnuo Calculations L
Fuel burnup calculations in the hot operating condition were performed primarily with the CASMO-2E code. However, to enhance the credibility of the burnup calculations (in lieu of critical r
experiments), the CASMO-2E results were independently checked by L calculations with the NULIF code (Ref. 7) and with EPRI-CELL (Ref. 9). Figure 4.4 compares results of these independent methods of burnup analysis under hot reactor operating condi-tions. The results agree within 0.008 ak in the hot operating condition.
In addition to' depletion calculations under hot operating conditions, reactivity comparisons under conditions more repre-sentative of fuel to be stored in the racks (cold, xenon-free) are also significant in storage rack critical'ity analyses. Table 4.4 compares the cold, xenon-free reactivities calculated by CASMO-2E, NULIF/PD007, and EPRI-CELL. In the cold condition, the CASMO-2E calculations gave a slightly higher reactivity value for
. the Region 2 fuel storage cell; and the good agreement generally observed lends credibility to the calculations, particularly in view of the known bias and uncertainty in CASMO-2E calculations (Appendix A).
SNEID is a one-dimensional dif fusion theory routine developed by Black & Veatch and verified by comparison with PDQ07 one-dimen-sional calculations.
4-17 e
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r-L Table 4.4 g COMPARISON OF COLD, CLEAN REACTIVITIES CALCULATED AT 25,000 MWD /MTU BURNUP AND 3.2% ENRICHMENT
(
[
k_ Xe-free 0*C Fuel Assembly In Region 2 Cell
( Calculational Method CASMO-2E 1.1206 0.9061
{ 1.1294 0.9017 NULIF/PD007 EPRI-CELL 1.1201-(l) -
( (1)EPRI-CELL k,at maximum value during long-term (30-year) storage.
[
No definitive method _ exists for determining the. uncertainty
[.- in . burnup-dependent reactivity calculations. All of the codes discussed above have been used to accurately follow reactivity h loss rates in operating reactors. CASMO-2E has been extensively benchmarked (Appendix A; Refs. 1, 2, 3, and 10) against cold,
(- clean, critical experiments (including plutonium-bearing fuel),
Monte Carlo calculations, reactor operations, and heavy-element concentrations in irradiated fuel. In particular, the analyses
{ (Ref. 10)'of 11 critical experiments with plutonium-bearing fuel gave an average k gg of 1.002
- 0.011 (95%/95%), showing adequate treatment of the plutonium nuclides. In addition, Johansson (Ref. 11) has obtained very good agreement in calculations of close-packed, high-plutonium-content, experimental configura-tions.
E Since critical-experiment data with spent fuel is not avail-it is necessary to assign an uncertainty in reactivity
{ able,
{ 4-19
~
I based on other' considerations, supported by the close agreement between different calculational methods and the general industry experience in predicting reactivity loss rates in operating plants. Over a considerable portion of the burnup, the reac-tivity loss rate in PWRs is approximately 0.01 Ak for each 1,000 MWD /MTU, becoming somewhat smaller at the higher burnups. By conservatively . assuming an uncertainty in reactivity of 0.5 x -
10-6 times the burnup in MWD /MTU, a burnup-dependent uncertainty is defined that increases with increasing fuel burnup, as would be reasonably expected. This assumption provides an estimate of the burnup uncertainty that is more conservative and bounds estimates frequently employed in other fuel rack licensing applications (i.e., 5% of the total reactivity decrement). Table 4.5 summarizes results of the burnup analyses and estimated uncertainties. These uncertainties are appreciably larger, in general, than would be suggested by the industry experience in predicting reactivity loss rates and boron let-down curves over many cycles in operating plants. The increasing level of conservatism at the higher fuel burnups provides an adequate margin in the uncertainty estimate to accommodate the possible existence of a small positive reactivity increment from the axial distribution in burnup (see Section 4.4.3). In addition' E
although the burnup uncertainty may be either positive or 5 negative, it is treated as an additive term rather than being combined statistically with other uncertainties. Thus, the allowance for uncertainty in burnup calculations is believed to be a conservative estimate, particularly in view of the j substantial reactivity decrease with aged fuel as discussed in Section 4.4.4.
Only that portion of the uncertainty due to burnup. Other un-certainties are accounted for elsewhere.
4-20 I
I
r -
L Table 4.5 f.
L- ,
ESTIMATED UNCERTAINTIES IN REACTIVITY DUE TO FUEL DEPLETION EFFECTS
(
r L Design 0.5 x 10-6 Initial Burnup Times Enrichment MWD /MTU Burnup, Ak Reac.tivit{1)
Loss, Ak 1.8 5,230 0.0026 0.0475.
2.5 15,720 0.0079 0.1575
{ 3.2 25,000 0.0125 0.2337 ,
[ 3.7 31,280 37,370 0.0156 0.0187
. 0.2757 0.3107 L 4.2
( (1) Total reactivity decrease, calculated for the cold, Xe-free condition in the fuel storage rack, from the beginning-of-life to the design burnup.
[ 4.4.3 Effect of Axial Burnup Distribution Initially, fuel loaded into the reactor will burn with a
{ slightly skewed cosine power distribution. As burnup progresses, the burnup distribution will tend to flatten, becoming more highly burned in the central regions than in the upper and lower ends. This effect may be clearly seen in the curves compiled in Ref. 12. At high burnup, the more reactive fuel near the ends of
.the fuel assembly (less than average burned) occurs in regions of Consequently, it
( lower reactivity worth due to neutron leakage.
is expected that distributed-burnup fuel assemblies would exhibit
~
a slightly lower reactivity than that calculated for the average
{ burnup. As burnup progresses, the distribution, to some extent, tends to be self-regulating as controlled by the axial power distribution, precluding the existence of large regions of signi-ficantly reduced burnup.
4-21
I A number of one-dimensional diffusion theory analyses have I
been made based upon calculated and measured axial burnup distri-butions. These analyses confirm the minor and generally negative reactivity effect of axially distributed burnup. The trends observed, however, suggest the possibility of a small positive reactivity effect at the high burnup values, and the uncertainty in k, due to burnup, assigned at the hig.her burnups (Section 4.4.2), is adequately conservative to encompass the potential for a small positive reactivity effect of postulated axial burnup distributions. Furthermore, reactivity decreases with time in storage (Section 4.4.4), and, in addition, there is a large margin in reactivity (>0.02 A k) below the limiting' ke gg value (0.95) which can accommodate any reasonabl'e reactivity effects that might be larger than expected.
4.4.4 Long-teim Decay Since the fuel racks in Region 2 are intended to contain spent fuel for long periods of time, calculations were made using EPRI-CELL (which incorporates the CINDER code) to follow the l long-term changes in reactivity of spent fuel over a 30-year l period. CINDER tracks the decay and burnup dependence of some 179 fission products. Early in the decay period, xenon grows in (reducing reactivity) and subsequently decays, with the reactiv-ity reaching a maximum at 100-200 hours. The decay of Pu-241 l
(13-year half-life) and growth of Am-241 substantially reduce l
reactivity during long term storage, as indicated in Table 4.6.
The reference design criticality calculations do not take credit l
for this long-term reduction in reactivity, other than to indi-cate an increasing subcriticality margin in Region 2 of the soent fuel storage pool.
I 4-22 I
f t
c Table 4.6 L
LONG-TERM CHANGES IN REACTIVITY IN STORAGE RACK
( (XENON-FREE) b Ak from Shutdown (Xenon-free)
Storage
{ Time, 3.2%E 4.2%E years 925,000 Mwo/MTU 037,000 MWD /MTU 0.5 -0.0046 -0.0057 L.0 -0.0080 -0.0103-
{ 10.0 -0.0406 -0.0529 20.0 -0.0588 -0.0756 30.0 -b.0692 -0.0886
[ .
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4-23
{
a mi um i i
I 4.5 REGI'ON 1 CRITICALITY ANALYSIS AND TOLERANCE VARIATIONS 4.5.1 Nominal Design Case Under normal conditions, with nominal dimensions, the k, values calculated by the three methods of analysis are as fol-lows.
Maximum k l
Analytical Method Bias-corrected k, (95%/95%)"
CASMO-2E 0.9387 i 0.0018 0.9405 AMPX-KENO 0.9301 i 0.0061 0.9362 Diffusion blackness 0.9393 0.9393 theory The AMPX-KENO calculations include a one-sided tolerance factor (Ref. 13) of 1.799 corresponding to 95% , probability at a 95% con-fidence limit. For the nominal ' design case, the CASMO-2E calcu-lation yields the highest (most conservative) reactivity, and, therefore, the independent verification calculations substantiate CASMO-2E as the. primary calculational method.
4.5.2 Boron Loading Variation The Boraflex ' absorber sheets used in Region 1 storage cells are nominally 0.075-inch thick, with a B-10 areal density of O.0238 g/cm 2. Independent manufacturing tolerance limits are i0.007 inch in thickness and t0.0017 g/cm 2 in B-10 content. This assures that at any point where the minimum boron concentration (0.0221 gram B-10/cm 2) and minimum Boraflex thickness (0.068 inch) may coincide, the boron-10 areal density will not be less than 0.020 gram /cm 2. Differential CASMO-2E calculations indicate that these tolerance limits result in an incremental reactivity uncertainty of *0.0021 ak for boron content and *0.0047 for Bora-flex thickness variations.
I 4-24
I ll 4.5.3 Storage cell Lattice Pitch variation The design storage cell lattice spacing between fuel assem-blies in Region 1 is 10.32 inches in one direction and 10.42 inches in the other direction. A decrease in storago cell lattice spacing may or may not ihcrease reactivity depending upon other dimensional changes that may be associated with the decrease in lattice spacing. Increasing the water thickness between the fuel and the inner stainless steel box results in a small increase in reactivity. The reactivity effect of the flux-thickness, however, is more significant, and de-I trap water creasing the flux-trap water thickness increases reactivity.
Both of these effects have been evaluated for independent design tolerances.
The inner stainless steel box dimension, 8.850 i 0.032 inches, defines the inner water thickness betwee5 the fuel and q
the inside box. For the tolerance limit, the uncertainty in reactivity is *0.0018 ak as determined by differential CASMO-2E:
calculations, with k, increasing as the inner stainless steel box dimension (and derivative lattice spacing) increases.
The design flux-trap water thicknesses are 1.160 i 0.040 inches and 1.260
- 0.040 inches, which result in an uncertainty of *0.0038 Ak due to the tolerance in flux-trap water thickness, assuming the water thickness is simultaneously reduced on all four sides. Since the manufacturing tolerances on each of the four sides are statistically independent, the actual reactivity I uncertainties would be less than *0.0038, although the more con-servative value has been used in the criticality evaluation.
I 4-25 I
I l
l 4.5.4 Stainless Steel Thickness Tolerances l The nominal stainless steel thickness in Region 1 is 0.060 inch for the inner stainless steel box and 0.020 inch for the Boraflex coverplate (0.125 inch on module bor.ndary). The maximum positive reactivity effect of the expected stainless steel thick-ness ' tolerance variations, statistically combined, was calculated (CASMO-2E) to be *0.0025 Ak.
4.5.5 Fuel Enrichment and Density Variation I
The design maximum enrichment is 4.20
- 0.05 wt% U-235.
Calculations of the sensitivity to small enrichment variations by CASMO-2E yielded a coefficient of 0.0047 Ak per 0.1 wt% U-235 at the design enrichment. For a tolerance on U-235 enrichment of i0.05 in wtt, the uncertainty on k, is *0.0024 Ak.
Calculations were made with the UO2 fuel density increased to a maximum value of 97% theoretical density (TD). For the mid-range value (95% TD) used for the reference design calculations, the uncertainty in reactivity- is *0.0026 ok over the range of UO2 densities expected.
4.5.6 Boraflex Width Tolerance Variation The reference storage cell design for Region 1 (Fig. 4.2) uses a Boraflex blade width of 7.75 t 0.0625 inches. A positive increment in reactivity occurs for a decrease in Boraflex absorber width. For a reduction in width of the maximum toler-ance, 0.0625 inch, the calculated positive reactivity increment is +0.0007 Ak.
I I
4-26 I
~
4.5.7 Axial Cutback of Boraflex w
~ The axial length of the Boraflex poison material is less than the active fuel length by three inches at the top and at the c bottom of the Region 1 storage rack modules. To account for the
+ reactivity effect of this axial cutback, one-dimensional (slab) diffusion theory calculations were made using flux-weighted L homogenized diffusion theory constants edited from CASMO-2E cal-culations of the array of storage cells, with and without Boraflex present. In the one-dimensional calculations, an
[
infinite (30-cm) water reflector was used above and below the p fuel assembly, with the lengths of the unpoisoned " cutback" regions, top and bottom, varied in a series of parametric cal-7 culations. Results of these calculations showed that the k,gg U remains less than the k, of the reference central storage cell region, until the axial cutback exceeds four inches top and i bottom. Thus, the actual axial neutron leakage more than compen-sates for the three-inch design cutback, and the reference I infinite multiplication factor, k, , remains a conservative over-estimate of the true reactivity.
I i
)
1 -
I I
I 4-27 I
4.6 REGION 2 CRITICALITY ANALYSIS AND TOLERANCE VARIATIONS 4.6.1 Nominal Design Case The principal method of analysis in Region 2 was the CASMO-2E code, using the restart option in CASMO to transfer fuel of a specified burnup into the storage rack configuration at a refer-ence temperature of 0*C. Calculations were made for fuel of several different initial enrichments and, for each enrichment, a limiting k, value established which included an additional factor for uncertainty in the burnup analysis and for the axial burnup g distribution. The restart CASMO-2E calculations (cold, clean, E rack geometry) were then interpolated to define the burnup value yielding the limiting k, value for each enrichment, as indicated in Table 4.7. These converged burnup values define the boundary of the acceptable domain shown in Fig. 4.1.
Table 4.7 FUEL BURNUP VALUES FOR REQUIRED REACTIVITIES (k,)
WITH FUEL OF VARIOUS INITIAL ENRICHMENTS Fuel Initial Reference Uncertainty (1) Design Burnup, Enrichment k, in Burnup, Ak Limit k, MWD /MTU l.58 0.9186 0 0.9186 0 1.8 0.9186 0.0026 0.9160 5,230 2.5 0.9186 0.0079 0.9107 15,720 3.2 0.9186 0.0125 0.9061 25,000 3.7 0.9186 0.0156 0.9030 31,280 4.2 0.9186 0.0187 0.8999 37,370 I
(1)See Section 4.4.2.
4-28 I
l t
r L
r At a burnup of 37,000 MWD /MTU, the sensitivity to burnup is calculated to be -0.0079 Ak per 1000 MWD /MTU. During long-term storage, the k, values of the Region 2 fuel rack will decrease continuously from decay of Pu-241 as indicated in Section 4.4.4.
( Two independent calculational methods were used to provide additional confidence in the reference Region 2 criticality
[ analyses. Fuel of 1.5% initial enrichment (approximately equiva-lent to the reference rack design for burned fuel) was analyzed by AMPX-KENO (27-group SCALE cross-secti6n library) and by the
{ CASMO-2E model used for the Region 2 rack analysis. For this case, the CASMO-2E k, .(0.9014) was within the statistical uncer-tainty of the bias-corrected value (0.9043 i 0.0030 (la))
obtained in the AMPX-KENO calculations. This agreement confirms the validity of the primary CASMO-2E calculations.
The second independent method of analysis used the NULIF code for burnup analysis, and for generating diffusion theory constants (cold, clean) for the NULIF-calculated composition at 25,000 MWD /MTU with fuel of 3.2% initial enrichment. These constants, together with blackness theory constants for the Boraflex absorber, were then used in a two-dimensional PDOO7 I calculation for the storage rack configuration. Results of this calculation (k of 0.9017) compared favorably with the CASMO-2E calculation for the same conditions (k, of 0.9061) and thus tend to confirm the validity of the primary calculational method.
I 4.6.2 Boron Loadino variation I The Boraflex absorber sheets used in the Region 2 storage cells are nominally 0.041 inch thick with a B-10 areal density of I 0.0130 g/cm 2. Independent manufacturing limits are i0.007 inch in thickness and i0.0009 g/cm 2 in B-10 content. This assures that at any point where the minimum boron concentration (0.01206 2
B-10/cm ) and the minimum Boraflex thickness (0.034 inch) may I 4-29 I 1 l
f coincide, the boron-10 areal density will not be less than 0.010 g/cm 2. Differential CASMO-2E calculations ' indicate that these tolerance limits result in an incremental reactivity uncertainty of *0.0028 Ak for boron content and i0.0078 Ak for Boraflex thickness.
4.6.3 Storage cell Lattice Pitch variations The design st,orage cell lattice spacing between fuel assem-blies in Region 2 is 9.011
- 0.040 inches, corresponding to an uncertainty in reactivity of 0.0011 Ak.
4.6.4 Stainless Steel Thickness Tolerance The nominal thickness of the stainless steel box wall is 0.060 inch with a tolerance limit of *0.005 inch, resulting in an uncertainty in reactivity of *0.0001 Ak.
4.6.5 Fuel Enrichment and Density variation I
Uncertainties in reactivity due to tolerances on fuel en-richment and U02 density in Region 2 are assumed to be the same as those determined for Region 1.
4.6.6 Boraflex Width Tolerance The reference storage cell design for Region 2 (Fig. 4.3) uses a Boraflex absorber width of 7.25 i 0.0625 inches. For a reduction in width of the maximum tolerance, the calculated posi-tive reactivity increment is 0.0009 Ak.
4-30
' ~ '
4.7 ABNORMAL AND ACCIDENT CONDITIONS 4.7.1 Eccentric Positioning of Fuel Assembly in Storace Rack The fuel assembly is normally located in the center of the storage rack cell with bottom fittings and spacers that mechan-ically limit lateral movement of the fuel assemblies. Neverthe-I less, calculations were made with the fuel assemblies moved into the corner of the storage rack cell (four-assembly cluster at l
closest approach). These calculations indicated that the reac-tivity decreases very slightly in both regions, as determined by PDOO7 calculations with diffusion coefficients generated by The highest reactivity I NULIF and a blackness theory routine.
therefore corresponds to the reference design with the fuel assemblies positioned in the center of the storage cells.
I 4.7.2 Temperature and Water Density Effects I The moderator temperature coef ficient of reactivity in both regions is negative; a moderator temperature of O'C, with a water density of 1.0 g/cm 3 , was assumed for the reference designs, I which assures that the true reactivity will always be lower, regardless of temperature.
Temperature effects on reactivity have been calculated and the results are shown in Table 4.8. Introducing voids in the water internal to the storage cell (to simulate boiling) de-creased reactivity, as shown in the table. Voids due to boiling will not occur in the outer (flux-trap) water region of Region 1.
"This calculational approach was necessary since the reactivity l effects are too small to be calculated by KENO, and CASMO-2E geometry is not readily amenable to eccentric positioning of a fuel assembly.
4-31
Table 4.8 EFFECT OF TEMPERATURE AND VOID ON CALCULATED i REACTIVITY OF STORAGE RACK Case Incremental Reactivity Change, Ak Region 1 Region 2 l
0*C Reference Reference j 20*C -0.0022 -0.0047 j 50*C -0.0084 -0.0081 l 80*C -0.0165 -0.0121 l 120*C -0.0298 -0.0178 120*C + 20% void -0.0953 -0.0520 I
I With soluble poison present, the temperature coefficients of reactivity would be expected to dif fer from those inferred from the data in Table 4.8. However, the reactivities would also be substantially lower at all temperatures with soluble boron present, and the data in Table 4.8 is pertinent to the higher-reactivity unborated case.
l 4.7.3 Dropped Fuel Assembly Accident l
l To investigate the possible reactivity ef fect of a postu-l lated fuel assembly drop accident, calculations were made for unpoisoned assemblies separated only by clean unborated water.
Figure 4.5 shows the results of these calculations. From these data, the reactivity (k ) will be less than 0.95 for any water
! 4-32 I
9 ,
1.50 i i ii i6 6 6 6 6 6 6 6 6 6 6 6 i i a i6 6 i
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=
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[ =
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E _ =
1.20 m.
2 -
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.m m
. W i.00 um 05 0 2 4 6 8 to 12
( WATER CAP BETWEEN ASSEMBLIES, In.
Fig. 4.5 REACTIVITY EFFECT OF WATER SPACING BETWEEN
[ fuel assemblies.
~
4-33
gap spacing greater than 6 to 7 inches in the absence of any absorber material, other than water, between assemblies. For a drop on top of the rack, the fuel assembly will come to rest horizontally on top of the rack with a minimum separation distance of more than 12 inches. Maximum expected deformation under seismic or accident conditions will not reduce the minimum spacing between fuel assemblies to less than ~12 inches. Con-sequently, fuel assembly drop accidents will not result in an increase in reactivity above that calculated for the infinite nominal design storage rack. Furthermore, soluble boron in the pool water would substantially reduce the reactivity and assure g that the true reactivity is'always less than the limiting value E for any conceivable fuel handling accident.
4.7.4 Abnormal Location of a Fuel Assembly I
The abnormal location of a fresh unirradiated fuel assembly I
of 4.2% enrichment ' could, in the absence of soluble poison, result in exceeding the design reactivity liriitation (k 'of 0.95). This could occur if the assembly were to be either posi-tioned outside and adjacent to a storage rack module or loaded I
into a Region 2 storage cell, with the latter condition producing the larger positive reactivity increment. Soluble poison, how-ever, is normally present in the spent fuel pool water (for which credit is permitted under these conditions) and would maintain the reactivity substantially less than the design limitation.
The largest reactivity increase occurs for accidentally
~
placing a new fuel assembly into a Region 2 storage cell with all other cells fully loaded. Under this condition, the presence of 300 ppm soluble boron assures that the infinite multiplication factor would not exceed the design basis reactivity. With the nominal concentration of soluble poison present (2000 ppm boron),
the maximum reactivity, k,, is less than 0.95 even if Region 2 l
were to be fully loaded with fresh fuel of 4.2% enrichment.
4-34
Administrative procedures will be used to confirm and assure the
(
continued presence of-soluble poison in the spent fuel pool water r during fuel handling operations.
4.7.5 Lateral Rack Movement Lateral motion of the rack modules under seismic conditions
( could potentially alter the spacing between rack modules. How-ever, girdle bars on the moduleo prevent closing the spacing to less than 1.25 inches, which is approximately the normal flux-trap water gap in the Region 1 reference design. Region 2 storage cells do not use flux-trap .and the reactivity is insensitive to the spacing between module's. Furthermore, soluble poison would assure ~ that a reactivity less than the design limi-tation is maintained under all conditions.
[
[
G -
t,
[ ,
4-35
I 4.8 NEW FUEL STORAGE I.
4.8.1 Storage in Region 1, Dry Region 1 is normally designed to accommodate new unirra-diated fuel assemblies in the fully flooded condition. For storage in the dry condition, the racks must also conform to the requirements of SRP 9.1.1 which specify a limiting keff value of 0.98 under optimum low density moderation. Calculations were made, using AMPX-KENO, for several hypothetical low-moderator densities down to 0.05 g/cc simulating fog or foam moderation.
These calculations showed a . continuously decreasing k, as the moderator density decreased, yielding a k, of 0.546 i 0.008 (la) at 10% moderator density. Axial leakage was neglected in these calculations, but would substantially reduce the already -
low k, values. These results are consistent with the general observation that a low-density optimum-modera' tion peak in reactivity does not exist in poisoned racks (Ref. 14).
4.8.2 Storage in Region 2, Flooded In a succession of trial-and-error calculations, it was found that a checkerboard storage pattern in Region 2 would allow new fuel assemblies of 4.2% enrichment to be safely accommodated without exceeding the limiting 0.95 k,gg value. In this checker-board loading pattern, the fuel assemblies are located on a diagonal array, as illustrated' on the next page, with alternate storage cells empty of any fuel.
I I
I I
4-36 I
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//// '///s //// //// '///d r/// '///, //// //// '///' (empty) r/// ///s //// f///. .///s r/// '///) //// f/// '///s
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//// '/// , //// //// '///.
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l l
Monte Carlo calculations (AMPX-KENO) resulted in a k, of l 0.8133 i 0.0138. With a one sided K-factor (Ref. 13) for 95%
probability at a 95% confidence level and a Ak of 0.009 for uncertainties (Table 4.1 for Region 2), the maximum k, is 0.863, l which is substantially less than the 0.95 limiting value. Thus, Region 2 may be safely used for the temporary storage of new fuel assemblies provided the storage configuration is restricted to l
the checkerboard pattern indicated above, f
4.8.3 Storace in Region 2, Dry As indicated in Section 4.8.1 above, a peak in reactivity (kegg) at low moderator densities is not expected for poisoned rack designs. AMPX-KENO calculations confirmed the absence of a -
low-moderator-density peak in Region 2 with 4.2% enriched fuel arranged in the checkerboard pattern. At 10% moderator density, the calculated k, was 0.552, which would be substantially reduced if axial leakage were to be included. Thus, Region 2 conforms to the requirements of SRP 9.1.1 (k. <0.98 at optimum moderation) for the safe storage of 4.2% enriched fuel, dry, in the checker-board loading pattern.
4-37 I
I REFERENCES I
- 1. A. Ahlin, M. Edenius, H. Haggblom, "CASMO - A Fuel Assembly Burnup Program," AE-RF-76-4158, Studsvik report (proprietary).
- 2. A. Ahlin and M. Edenius, "CASMO -A Fast Transport Theory Depletion Code for LWR Analysis," ANS Transactions, Vol. 26,
- p. 604, 1977.
- 3. M. Edenius et al., "CASMO Benchmark Report," Studsvik/RF 6293, Aktiebolaget Atomenergi, March 1978.
- 4. 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.
- 5. L. M. Petrie and N. F. Cross, " KENO-IV, An Improved Monte Carlo Criticality Program," ORNL-4938, Oak Ridge National Laboratory, November 1975.
- 6. R. M. Westfall et al., " SCALE: A Modular Code System for '
Performing Standardized Computer Analyses for Licensing Evaluation," NUREG/CR-0200, 1979.
- 7. W. A. Wittkopf, "NULIF -
Neutron Spectrum Generator, Few-Group Constant Generator and Fuel Depletion Code," BAW-426, The Babcock & Wilcox Company, August 1976.
- 8. W. R. Cadwell, PDOO7 Reference Manual, WAPD-TM-678, Bettis Atomic Power Laboratory, January 1967.
- 9. W. J. Eich, " Advanced Recycle Methodology Program, CEM-3,"
Electric Power Research Institute, 1976. -
- 10. E. E. Pilat, " Methods, for the Analysis of Boiling Water Reactors (Lattice Physics)," YAEC-1232, Yankee Atomic Electric Co., December 1980.
- 11. E. Johansson, " Reactor Physics calculations on Close-Packed Pressurized Water Reactor Lattices," Nuclear Technology, Vol. 68, pp. 263-268, February 1985.
- 12. H. Richings, Some Notes on PWR (W) Power Distribution Probabilities for LOCA Probabillstic Analyses, NRC Memorandum to P. S. Check, dated July 5, 1977.
I 4-38 I
[
L REFERENCES (Continued)
- 13. M. G. Natrella, Experimental Statistics National Bureau of Standards, Handbook 91, August 1963.
- 14. J. M. Cano et al., "Supercriticality Through Optimum Moderation in Nuclear Fuel Storage," Nuclear Technolocy, Vol. 48, pp. 251-260, May 1980.
F 9
9 e
~
4-39 J
[
5.0 THERHAL-HYDRAULIC CONSIDERATIONS A primary objective in the design of the high density spent fuel storage racks 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.
Simila'r analysis has been used in previous licensing reports on
" high density spent fuel racks for Fermi 2 (Docket 50-341), Quad c
Cities 1 and 2 (Dockets 50-254 and 50-265), Rancho Seco (Docket k 5'0-312), Grand Gulf Unit 1 (Docket 50-416), Oyster Creek (Docket 50-219), Virgil C. Summer (Docket 50-395), and Diablo Canyon 1 and 2 (Docket Hos. 50-275 and 50-323).
5.1 DECAY HEAT CALCULATIONS FOR THE SPENT FUEL This report section covers requirement III.1.5(2) of the NRC's "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications" issued on April 14, 1978. This r
requirement states that calculations for the amount of thermal energy removed by the spent fuel cooling system shall be made in l accordance with Branch Technical Position APCSB 9-2, " Residual Decay Energy for Light Water Reactors for Long Term Cooling" l (Ref. 1). The calculations contained herein have been made in accordance with this requirement.
l 5.1.1 Basis f ~
The Byron Nuclear Power Station Units 1 and 2 reactors are both I
f rated at 3411 megawatts thermal (HWt). Each core contains 193 Thus, the average operating power per fuel fuel assemblies.
assembly, Po, is 17.6736 MW. The fuel discharge can be made in one of the following two modes:
0 Normal refueling discharge 0 Full core discharge 5-1 j
I' I
An equilibrium reload consists of 84 assemblies (with 18-month cycles). The four transitional reloads for each unit consist of 88 assemblies. The fuel transfer begins after 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br /> of decay time in the reactor (time after shutdown). It is assumed that the time period of discharge of this batch is 28 hours3.240741e-4 days <br />0.00778 hours <br />4.62963e-5 weeks <br />1.0654e-5 months <br /> (three assemblies transferred to the pool per hour). The cooling system consists of a Seismic. Category I spent fuel cooling circuit. The bulk temperature analysis assumes a 105'F coolant inlet temperature to the spent fuel pool heat exchanger for these refueling cases.
For the full core discharge, it is assumed that the total time period for the discharge of the full core is 64 hours7.407407e-4 days <br />0.0178 hours <br />1.058201e-4 weeks <br />2.4352e-5 months <br /> (after 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br /> of shutdown time in the reactor). The discharge rate to the pool is assumed to be continuous and uniform.
The fuel assemblies are removed from the reactor after a maximum 3 postulated time'of 4.5 years of cumulative operating time. Since the decay heat load is a monotonically increasing function of the cumulative reactor operating time, to, it is conservatively assumed that every fuel assembly discharged has had the maximum l
postulated To of 4.5 years for the batch size of 84.
The water inventory in the reactor cavity cooled by the residual heat removal (RHR) heat exchanger exchanges heat with the fuel pool water mass through the refueling canal. This source of heat removal is neglected in the analysis. Thus, the results obtained for both . normal refueling discharge and full core discharge are conservative.
I E
5-2 l I
l I
l The fuel pool cooling system (FC) consists of two independent trains, each consisting of one pump and heat exchanger. Either -
train is capable of providing sufficient cooling for the pool.
I The following list identifies all relevant design data for the spent fuel pool heat exchangers:
0 Type ,
Tube and shell 0 Quantity 2 0 Performance data
- Heat transferred 15.833 x 106 Btu /hr Tube Side 6
- Fluid flow 2.23 x 10 lb/hr
- Pool water inlet temperature 120*F
- Outlet temperature 112.9'F
- Fouling Factor .0005 l Shell Side 6
- Fluid flow '2.72 x 10 lb/hr j
- Coolant inlet temperature 105*
- Outlet temperature 110.82*F
- Fouling factor 0.0005 The above data enables complete characterization . of the thermal performance of a fuel pool heat exchanger.
I 5.1.2 Model Description I Reference 1 is utilized to compute the heat dissipation requirements in the pool. The total decay heat consists of I fission product and heavy element decay heat. Total decay heat, P, for a fuel assembly is given as a linear function of Po and I 5-3 I
- s
as an exponential function of a and s:
i P = Po f( o, s) (5.1-1) l where:
l P = total decay heat per fuel assembly, linear I
function of P o l P o
= average operating power per fuel assembly _
= cumulative exposure time of the fuel assembly in the reactor
, = time elapsed since reactor shutdown E
The appropriate uncertainty factor, K, was applied in accordance with HUREC-0800 (Ref. 1). Furthermore, the operating power, P o, is taken equal to the rated power, even though the reactor may be operating at less than its rated power during much of the I exposure period for the batch of fuel assemblies. Finally, the computations and results reported here are based en the discharge l taking place when the inventory of fuel in the pool will' be at '
l its maximum resulting in an upper bound on the computed decay l heat rate.
Having determined the heat dissipation rate, the next task is to evaluate the time-dependent temperature of the pool water. Table 5.1 identifies the loading cases examined. The pool bulk temperature is determined using the first law of thermodynamics (conservation of energy).
A number of simplifying assumptions are made which render the analysis conservative, principally:
E 5-4 I
l l
O The heat exchangers are assumed to have maximum fouling.
Thus, the temperature effectiveness, S, for the heat exchanger utilized in the analysis is the lowest postulated value: S = .3875 for fuel pool cooler. S is calculated from heat exchanger technical data sheets. No heat loss is assumed to take place through the concrete floor.
O No credit is taken for the improvement in the film coefficients of the . heat exchanger as the . operating temperature rises. Thus, the film coefficient used in the computations are lower bounds.
l 0 No credit is taken for heat loss by evaporation of the pool water.
l 0 No credit is taken for heat loss to pool walls and pool I
floor slab.
I for the pool heat The basic energy conservation relationship l
exchanger system yields:
1 C
t
- 0 1 -0 2 (5.1-2) dT where:
1 C = Thermal capacity of stored water in the pool t
t = Temperature of pool water at time, T Qg = Heat generation rate due to stored fuel assemblies l In the pool; Q1 is a known function of time, T from the preceding section.
Q = Heat removed in the fuel pool cooler The pool has a total water inventory of 63444.0 cubic. feet when all racks are in place in the pool and every storage location is occupied.
I I
5-5
I 5.1.3 Decay Heat Calculation Results The calculations were performed for the pool, disregarding the additional thermal capacity and cooling system available in the transfer canal, and the reactor cavity.
For a specified coolant inlet temperature and flow rate, the quantity Q2 is shown to be a linear function of T in a recent (Ref. 3). As stated earlier, is an
, paper by Singh Q1 exponential function of T. Thus, Equation 5.1-2 can be integrated to determine t directly as ' a function of T. The results are plotted in Figures 5.1 through 5.2. The results show that the pool water never approaches the boiling point even with 3 the most adverse heat load, under normal operating conditions. E These figures also give Q1 as a function of T. Four plots are generated for each case. The first and third plots for each case power respectively, for a shows temperature and generation, period extending from T = 0 + T =2T n where T n is the total time of fuel transfer. The second and fourth plots show the same quantities (i.e., temperature and power generation, respectively) over a longer period. The long-term plots are produced to show the temperature drop with time. Summarized results are given in E
5 Table 5.2.
Finally, computations are made to determine the time interval to boiling after all heat dissipation paths are lost. Computations are made for each case under the following two assumptions:
0* All cooling systems lost at the instant pool bulk a temperature reaches the maximum value. 5 0 All cooling systems lost at the instant the heat dissipation power reaches its maximum value in the pool.
Results are summarized in Table 5.3. Table.5.3 gives the bulk vaporization rate for all cases at _the instant the boiling boiling commences. This rate will decrease with time due to reduced heat generation in the fuel. In all cases, adequate time exists to take corrective action.
5-6 I
1 5.2 THERHAL-HYORAULIC ANALYSES FOR SPENT FUEL COOLING F
This report section covers requirement III.1.5(3) of the NRC's "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," issued on April 14, 1978. Conservative 4 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 p fuel rods occurs only at the locations where freshly discharged fuel assemblies are stored.
l 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:
0 As stated above, the fuel: pool will contain spent-fuel h with varying time-after-shutdown (Ts). SinceT theitheat is emission falls off rapidly with increasing s, I ~
obviously assemblies conservative are fresh (T s to assume
= 100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br />) that and they all all fuel l have had 4.5 years of operating time in the reactor for I cases 1 and 2. The heat emission rate of each fuel assembly is assumed to be equal (Ref. 2).
0 As shown in Figure 2.1 in Section 2, the modules occupy For the 5 an irregular floor space in the pool.
circumscribing the f
hydrothermal analysis, a circle actual rack floor space is drawn. It is further assumed I
I l
that the cylinder with this circle as its base isofpacked with fuel assemblies at inches (see Figure 5.3).
the nominal pitch 9.011 O The downcomer space around the rack module group varies, as shown in Figure 2.1. The nominal downcomer gap available in the pool is assumed to be the total gap I available around the idealized cylindrical rack; thus, the maximum resistance to downward flow is incorporated into the analysis.
5-7 I
l -s l
0 No downcomer flow is assumed to exist between the rack I
l modules.
5.2.2 Model Description I
In this manner, a conservative idealized model for the rack I
assemblage is obtained. The water flow is axisymmetric about the vertical axis of the circular rack assemblage, and thus, the flow is two-dimensional (axisymmetric three-dimensional). Figure 5.4 5
shows a typical " flow chimney" rendering of the thermal B hydraulics model. The governing equation to characterize the flow field in the pool can now be written. The resulting integral l
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 that the hydrodynamic loss coefficients which enter into the formulation of the integral equation are also taken from well-recognized sources (Ref. 4) and wherever discrepancies in reported values exist, the conservative values are consistently
, used. Reference 5 gives the details of mathematical analysis 4
used in this solution process.
After the axial velocity field is evaluated, it is a straight-forward matter to compute the fuel assembly cladding temperature. The knowledge of the overall flow field enables pinpointing the storage location with the minimum axial flow (i.e., maximum water outlet temperature). This is called the most " choked" location. In order to find an upper bound on the
,emperature in a typical cell, it is assumed that it is located
~E at the most choked location. Knowing the global plenum velocity B field, the revised axial flow through this choked cell can be calculated by solving the Bernoulli's equation for the~ flow circuit through this cell. Thus, an absolute upper bound on the water exit temperature and maximum fuel cladding temperature is obtained. It is believed that, in view of the aforementioned assumptions, the temperatures calculated in this manner overestimate the temperature rise that will actually occur in the pool.
5-8
i I
i I The maximum pool bulk temperature, t, is computed in Section
. 5.1.3 and reported in Table 5.2. The corresponding average power output from the hottest fuel assembly, q, is also reported in i
that table. The maximum radial peaking factor, F xy, is 1.55 I for the Byron Nuclear Power Station. Thus, it is conservative to assume that the maximum specific power of a fuel assembly, qA, is given by:
qA = q Fxy (5.2-1) where:
Fxy = 1.55 1 The maximum temperature rise of pool water in the most disadvantageously placed fuel assembly is given in Table 5.4 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 conservative _1y assumed that the total peaking factor F is 2.32. Thus, a fuel rod can produce 2.32 times the average heat emission rate over a small length. The axial heat dissipation in a rod is known to reach a maximum the central region, and taper off at its two I
in 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 equations which directly give the maximum local cladding temperature, tc.
5.2.3 Results I
Table 5.4 gives the maximum local cladding temperature, te, at the instant the pool bulk temperature has attained its maximum value. It is quite possible, however, that the peak cladding temperature occurs at the instant of maximum value of qA, 5-9
I I: !
the instant when the fuel assembly is first-placed in a storage I
location. Table 5.5 gives the maximum local cladding temperature at T = 0. The local boiling temperature near the top of the fuel cladding is 240*F. However, the cladding temperature must be somewhat higher than the boiling temperature to initiate and sustain nucleate boiling. The above considerations indicate that E a comfortable margin against the initiation of localized boiling 5I exists in case 1. For full core discharge (case 2) un' der the described assumptions, the maximum cladding temperature will give rise to localized nucleate boiling, but not to bulk pool boiling (5.4)..
E I
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5-10
k r
L
( 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.
L 2. FSAR, Byron Nuclear Power Station.
- 3. 3ournal of Heat Transfer, Transactions of the ASME, August L 1981, Vol. 103, "Some Fundmental Relationships for Tubular Heat Exchanger Thermal Performance," K.P. Singh.
[ 4. General Electric Corporation, R&D Data Books, " Heat' Transfer and Fluid Flow," 1974 and updates.
- 5. 4th National Congress of the ASME, "A Method for Computing
[
L the Maximum Water Temperature in a Fuel Pool Containing Spent Nuclear Fuel," by K.P. Singh and A.I. Soler, paper 83-NE-7, Portland, Oregon (3une 1983).
O e
[. .
E
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[
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J 5-11
Table 5.1 LIST OF CASES ANALYZED No. of Total Time Fuel to Transfer Assemblies Fuel Into Decay Time Discharged, the Pool Refore Transfer Case No. Condition N th, hrs Begins, hrs 1 Normal refueling discharge
- 84 28 100 v' 2 Full core
- 193 64 100 I. discharge m
- Discharge is assumed to be into a pool containing fuel from 17 previous discharges of 168 assemblies.
M M M M m M- ~M W M M M m m m m-m m Table 5.2 HAXIMUM POOL BULK. TEMPERATURE, t, COINCIDENT TOTAL POWER, 01, AND COINCIDENT SPECIFIC POWER, q, FOR THE HOTTEST ASSEMBLY
. Coincident Coincident Time to Maximum Time (After Coincident Total Transfer Pool Bulk Initiation Specific Power Case No. of Fuel Into Temp., t, of Fuel Power, q, Qt (10 6)
No. Assemblies Pool, hrs *F Transfer), hrs Btu /sec Btu / hour Hotes 1 84 2 8, ' 138 37.0 -
55.15 .1985 Normal w
discharge" C 2 193 64 155 71.0. 50.30 .1811 Full Core Discharge"
- Discharge is assumed to be into a pool containing fuel from 17 previous discharges of 168 assemblies.
1
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l Table 5.3 l
TIME (HRS) TO BOILING AND BOILING VAPORIZATION RATE FROM THE INSTANT ALL COOLING IS LOST CONDITION 2 I
CONDITION 1 Loss of Cooling at Loss of Cooling at Maximum Maximum Power Discharge Pool Bulk Temperature Rate g 5
( Case Time (Hrs) Vap. Rate Time (Hrs) Vap. Rate-No. lb/hr ib/hr 35788.0 I
1 9 35329.0 9 i 2 4 54233.0 4 54827.0 l I I
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I I Table 5.4 I MAXIMUM LOCAL POOL WATER TEMPERATURE AND LOCAL FUEL CLADDING TEMPERATURE AT INSTANT OF MAXIMUM POOL BULK TEMPERATURE Maximum Local Maximum Local Fuel Cladding I Case No.
Water Temperature, 'F Temperature, *F Case Identified 194 239 84 assemblies 1
208 250 193 assemblies 2
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Table 5.5 POOL AND MAXIMUM CLADDING TEMPERATURE AT THE INSTANT FUEL ASSEMBLY TRANSFER BEGINS Coincident Pool Cladding Temperature, 'F
' Case No. Temperature, *F Bulk Local 1 236.6 122.6 186.0 2 236.6 122.8 186.2 I
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L STRUCTURAL ANALYSIS 6.0 .
The purpose of this section is to demonstrate the structural r
adequacy of the spent fuel rack design under normal and accident loading conditions. The method of analysis presented herein uses a time-history integration method similar to that previously used in the Licensing Reports on High Density Fuel Racks for Fermi 2 (Docket No. 50-341), Quap Cities 1 and 2 (Docket Hos.'50-254 and 50-265), Rancho Seco (Docket No. 50-312), Grand Gulf. Unit 1 (Docket H No. 50-416), Oyster Creek (Docket No. 50-219), V.C. Summer (Docket No. 50-395), and Diablo Canyon 1 and 2 (Docket Nos. 50-275 and c 30-323). The results show that the high density spent fuel racks sre structurally adequate to resist the postulated stress n ,w nh i n a t i o n s associated with level A, B, C, and D conditions as dafined in References 1 and 2.
6.1 ANALYSIS OUTLINE The spent fuel storage racks are Seismic Cate' gory I equ.ipment.
L Thus, they are required to remain functional during and after a Safe ~ Shutdown Earthquake (R e f . 3). As noted previously, these racks are neither anchored to the pool floor nor 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 completely empty. The coefficient of friction, p , between the supports and pool floor is another indeterminate factor. According to Rabinowicz (Ref. 4) the results of 199 tests performed on austenitic stainless steel plates submerged in water show a ~ mean value of u to be 0.503 with a standard deviation of 0.125. The upper and lower- bounds (based on twice the standard deviation) are thus 0.753 and 0.253, respectively. Two separate analyses are performed for the rack assemblies with values of the coefficient of friction equal to 0.2 (lower limit) and 0.8 (upper limit),
[ respectively. Analyses performed for the geometrically limiting
[ 6-1
E on limiting values of the coefficient of rack modules focus friction, and the number of fuel assemblies stored. Typical cases studied are 0 Fully loaded rack (all storage locations occupied),
I q = 0.8; 0.2 ( = coefficient of friction) 0 Nearly empty rack = 0.8, 0.2
= 0.8 :
0 Rack half full,
. The method of analysis employed is the time-history method. The
=
data were developed from the response pool slab, acceleration by the Sargent and Lundy Company, Chicago, spectra provided Illinois.
]
E the seismic analysis is to determine the The objective of structura) response (stresses, deformation, rigid body motion, etc.) due to simultaneous application of the three independent, orthogonal excitations. Thus, recourse to approximate statistical summation techniques such as the " Square-Root-of-the-Sum-of-the-Squares" method (Ref. 5) is avoided. For nonlinear analysis, the only practical method is simultaneous application.
Pool slab ' acceleration data are provided for two, earthquakes:
Earthquake Operating Basis Earthquake (OBE) and Safe Shutdown (SSE). Figures 6.1 - 6.3 show the time-histories corresponding to the SSE condition.
I The seismic analysis is performed in three steps, namely:
- 1. Development of a nonlinear dynamic model consisting of inertial mass elements and gap and friction elements.
I e-z I
[
- 2. Generation of the equations of motion and inertiaE coupling and solution of the equations using the "c~omponent element
[ time integration scheme" (References 6 and 7) to determine nodal forces and displacements
- 3. Computation of the detailed stress field in the rack (at the critical location) and in the support legs using the nodal forces calculated in the previous step. These stresses are g
checked against the design limits given in Section 6.5.
l.
s A brief description of the dynamic acdel follows.
6.2 FUEL RACK - FUEL ASSEMBLY MODEL Since the racks are not anchored to the pool slab or attached to the pool walls or to each other, they can execute a wide variety of
{ rigid body motions. For example, the rack may slide on the pool floor (so-called " sliding condition"); one or more legs may
[ momentarily lose contact with the liner (" tipping condition"); or the rack may experience a combination of sliding and tipping conditions. The structural model should permit simulation of these
- kinematic events with inherent built-in conservatisms. Since the Byron racks are equipped with girdle bars to dissipate energy due to inter-rack impact (if it occurs), it is also necessary to model ,
the inter-rack impact phenomena in. a conservative manner.
Similarly, lift off of the support legs and subsequent ~1mpacts must be modelled using appropriate impact ele . nants , and Coulomb friction between the- rack and the pool liner .must be simulated by appropriate piecewise linear springs. These special attribut?s of' the rack dynamics require a strong emphasis on the modeling of the linear and nonlinear springs, dampers, and stop elements. The model
( outline in the remainder of this section, and the model description in the following section describe the detailed modeling technique to simulate these effects, with emphasis placed on the nonlinearity
{ of the rack seismic response.
[
6-3
I E
6.2.1 Outline of Model
- a. The fuel _ rack structure is a folded metal plate assemblage welded to a baseplate and supported on four legs. The E
rack structure itself is a very rigid structure. Dynamic E analysis of typical multicell racks has shown that the motion of the structure is captured almost completely b.y the behavior of a six degrees-of-freedom structure; there-fore, the movement of the rack cross-section at any height is described in terms of the six degrees-of-freedom of the rack base.
- b. The seismic motion of a fuel rack is characterfred by random rattling of fuel assemblies in their indis dual storage locations. Assuming that all assemblies vibrate in phase obviously exaggerates the computed dynamic loading on the rack structure. This assumption, however, greatly reduces the required degrees-of-freedom needed to model the fuel assemblies which are reprecented by two lumped masses located at different levels of . the rack.
The centroid of each fuel assembly mass can be located, relative to the rack structure centroid at that level, so as to simulate a partially loaded rack.
l l
5
- c. The local flexibility of the rack-support interface is modeled conservatively in the analysis.
the pool I
- d. The rack -base support may slide or lift off floor.
- e. The pool floor has a specified time-history of seismic accelerations along the three orthogonal directions,
- f. Fluid coupling between rack and assemblies, and between and adjacent racks, is simulated by introducing rack E
6 I+
[
appropriate inertial coupling into the system kinetic energy. Inclusion of these effects uses the methods of References 4 and .6 for rack / ass'embly coupling and for rack / rack coupling (see Section 6.2.3 of this report). J
- g. Potential impacts between rack and assemblies are accounted for by appropriate " compression only" gap elements between masses involved.
- h. Fluid damping between rack and assemblies, and between rack and adjacent rack, is conservatively neglected.
1 1
( i. The supports a r.e modeled as " compression only" elements for the vertical direction and as " rigid links" for dynamic analysis. The bottom of a support leg is attached
{ The to a frictional spring as described in Section 6.2.2.
cross-section inertial properties of the support legs are computed and used in the final computations to determine support leg stresses.
- f. The effect of sloshing can be shown to be negligible at b the bottom of a pool and is hence neglected.
Inter-rack impact, if it occurs, is simulated oy a series
( k.
of gap elements at the top and bottom of the rack in the two horizontal directions. The most conservative case of adjacent rack movement is ast.med; each adjacent rack is assumed to move completely out of' phase with the rack being analyzed.
- 1. The form drag opposing the motion of the fuel assemblies in the storage locations is conservatively neglected in the results reported herein,
- m. The form drag opposing the motion of the fuel rack in the water is also conservatively neglected in the results reported herein.
[
6-5
I
- n. The rattling of the fuel assemblies inside the storage locations causes the " gap"'between the fuel assemblies and cell wall to change from a maximum of twice the the nominal gap to a theoretical zero gap. However, the fluid g coupling coefficients (Ref. 8) utilized are based on 5 linear vibration theory (Ref. 9). Studies in the literature show that inclusion of the nonlinear effect (viz. vibration amplitude of the same order of magnitude as the gap) drastically lowers the equipment response .
(Ref. 10).
Figure 6.4 shows a schematic of the model. Six degrees-of- freedom are used to track the motion ' of the rack structure. Figures 6.5 and 6.6, respectively, show the inter-rack impact springs and fuel assembly / storage cell impact springs.
motion I
-The production run model for simulating fuel assembly incorporates five lumped masses. The lower mass is assumed to be attached to the baseplate and to move with the baseplate. The four rattling masses are located at quarter height, half height, three quarter height, and top of the rack. Two degrees-of-freedom are used to track the motion of each rattling mass.
6.2.2 Model Description I
The ab' solute degrees-of-freedom associated with each of the mass locations are identified in Figure 6.4 and Table 6.1. Note that for clarity, only the top rattling mass (node 2*). is shown in the figure. The remaining rattling masses (nodus 3*,4*, 5*) are located at z = 3/4H, 1/2H, 1/4H, respectively, and are described by translational degrees-of-freedom q9-q n.
U t(t) is the pool floor slab displacement seismic time-history.
Thus, there are fourteen degrees-of-freedom in the system. Not 'W shown in Fig. 6.4 are the gap elements used to model the support legs and the impacts with adjacent racks.
- 6-6 I
c i
r 6.2.3 Fluid Coupling L *
[
L An ef f ect . of some significance requiring careful modeling is the so-call'ed " fluid coupling effect". If one body of mass (m1)
[ vibrates adjacent to another body (mass m 2) , and both bodies are submerged in a frictionless fluid medium, then Newton's equations
- of motion for the two bodies have the forms (m i + M 11) X1- M 12 X2= applied forces on mass mi L
-M 21 X i+ (m 2 + H 22) X 2= applied forces on mass m2 L
X 1, X2 denote absolute accelerations of mass m1 and m 2, F respectively.
L M i t, H 12, M 21, and H 22 are fluid coupling coef ficients which depend on the shape of the two bodies, their relative, disposition, etc.
Fritz (Ref. 6) gives data for Mj i for various body shapes and
[ arrangements. It is to be noted that the above equation indicates that the effect of the fluid is to add a certain amount of mass to body body 1), and an external force which is
( the (M11 to proportional to the acceleration of the adjacent body (mass m2).
Thus, the acceleration of one body -af fects the force field on
{ another. This force is a strong function of the interbody gap, reaching large values for very small gaps. This inertial coupling is called fluid coupling. It has an important effect in rack dynamics. The lateral motion of a fuel assembly inside the storage location wil.1 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 2 and 2+
in Figure 6.4. Furthermore, the rack equations contain coupling terms which model the effect of fluid in the gaps between adjacent
{ racks. The coupling terms modeling the effects of fluid flowing between adjacent racks are computed assuming that all adjacent racks are vibrating 180* out of phase from the rack being 6-7
I analyzed. Therefore, only one rack is considered surrounded by -a hydrodynamic mass computed as if there were a plane of symmetry located in the middle of the gap region.
Finally, fluid virtual mass is l'ncluded in the vertical direction vibration equations of the rack; virtual inertia is also added to the governing equation corresponding to the rotational degree-of-freedom, qs(t).
6.2.4 Damping I
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). In the analysis, a maximum of 4% structural damping is imposed on elements of the rack structure during SSE seismic simulations. This is in accordan,ce with the FSAR and NRC guidelines (Ref. 11). Material and fluid damping'are conservatively neglected. The dynamic model has the provision to incorporate fluid damping effects; however, no fluid damping has been used for this analysis.
6.2.5 Impact Any fuel assembly node (e.g. 2*) may impact the corresponding structural mass node 2. To simulate this impact, four comp re s siort-o nly gap elements around each rattling fuel assembly node are provided (see Figure 6.6). As noted previously, fluid dampers may also be provided in parallel with the springs. The compressive loads developed in these springs provide the necessary data to evaluate the integrity of the cell wall structure and stored array during the seismic event. Figure 6.5 shows the location of the impact springs used to simulate any potential for inter-rack impacts. Section 6.4.2 gives more details on these additional impact springs.
I .
6-8
r L
( 6.3 ASSEMBLY OF THE DYNAMIC HODEL The cartesian coordinate system associated with the rack has the following nomenclature L
i
, 0 x = Hor'izontal coordinate along the short direction of rack I rectangular platform L
0 y = Horizontal coordinate along the long-direction of the rack rectangular platform 0 z = Vertically upward
[
As described in the preceding section, the rack, along with the base, supports, and stored fuel assemblies, is modeled for the
{ general three-dimensional (3-0) motion simulation by a fourteen degree-of-freedom model. To simulate the impact and sliding phenomena expected,. 60 nonlinear. gap elements and 16 nonlinear friction elements are used. Cap and friction elements, with their
.connectivity and purpose, are presented'in Table 6.2.
restricted to two dimensions (one
( If the simulation model is horizontal motion plus vertical motion, for example) for the purposes of model clarification only, then a descriptive model of the simulated structure which includes gap and friction elements is shown in Figure 6.7. (Note that only the top rattling mass is shown for clarity.)
The impacts between fuel assemblies and rack show up in the gap elements, having local stiffness KI, in Figure 6.7. In Table
( 6.2, gap elements 5 through 8 are for the vibrating mass at the top of the rack. The ' support leg spring rates K 6
are modeled by elements 1 through 4 in Table 6.2. Note that the local compliance
{ To simulate sliding of the concrete floor is included in K 6 potential, friction elements 2 plus 8 and I+ plus 6 (Table 6.2) are shown in Figure 6.7. The friction of the support / liner interface is modeled by a piecewise linear spring with a suitably large 6-9
I.
~ E stiffness Kf up to the limiting lateral load, tN , where N is the and current compression load at the interface between support At every time step during the transient analysis, the liner. E current value of N (either zero for liftoff condition, or a 5 l
value) computed. Finally, the support compressive finite is rotational friction springs KR reflect any rotational restraint by the foundation. This spring rate is thu .nay be offered calculated using a modified Bousinesq equation (Ref. t+ ) and is included to sianula te the resistive moment of the support to counteract rotation of the rack leg in a vertical plane. This rotation spring is also nonlinear, with a zero spring constant value assigned after a certain limiting condition of slab moment loading is reached.
The nonlinearity of these springs (friction elements 9, 11, 13, and l
15 in Table 6.2) reflects the edging limitation imposed on the base of the rack support legs. In this analysis, this effect is bending, induced by liner / baseplate neglected; any support leg forces, is resisted by the leg acting as a beam j friction _
cantilevered from the rack baseplate.
t The spring rate K6 modeling the effective compression stiffness of the structure in the vicinity of the support, is computed from S'
the equation:
l 1 1
=_
1
._ . _1 K1 K2 K3 I
K l
6 l
! where:
spring rate of the support leg treated as a I
K1 =
tension-compression member = ESUPPORT
- ASUPPORT/h 3 (h = length of support leg) g I
. 10 g
[
2 pool slab K2= 1.05Ec8/(1 v ) = local spring rate of (Ee = Young's modulus of concrete, and B =
length of bearing surface)
K3= spring rate of folded plate cell structure above support leg (same form as K 2 with E chosen to reflect the local
[
stiffness of the honeycomb structure above the leg)
L (listed in Table For the 3-D simulation, all support elements e the two L 6.2) are included in the model. Coupling between horizontal seismic motions is provided both by the off set of the fuel assembly group centroid which causes the rotation of the
(
entire rack and by the possibility of liftoff of one or more support legs. The potential exists for the rack to be supported
{ on one or more support legs or to liftoff completely 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.
6.4 TIME INTEGRATION OF THE EQUATIONS OF MOTION
[
6.4.1 Time-History Analysis Using 14 DOF Rack Model 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 by using Lagrange's equation. The system of equations can be represented in matrix notation as:
(M ] {q } = {Q } + {G }
displacements and where the vector {Q } is a fun'etion of nodal velocities, and {G } depends on the coupling inertia and the ground acceleration. Premultiplying the above equations by (M ]-1 renders the resulting equation uncoupled in mass.
6-11
I We haves {q } = [M ]-1 {Q } + [M }-1 {G}
I
~
As noted earlier, 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 wil.1 provide maximum impact force level, and induce additional conservatism in
~
the time-history analysis.
This equation set is mass uncoupled, displacement coupled, and is ideally suited for numerical solution using a central difference scheme. The computer program "0YNAHIS"* is utilized for this purpose.
Stresses in various portions of the structure are computed from known element forces at each instant of time.
Dynamic analysis of typical multicell racks has shown that the motion of the structure is captured almost completely by the behavior of a six-degree-of-freedom structure; therefore, in this analysis model, the movement of the rack cross-section at any height is described in terms of the rack base degrees-of-freedom (qi(t),...q6(t)). The remaining degrees-of-freedom are associated with horizontal movements of the fuel assembly masses. In this dynamic model, four rattling masses are used to represent fuel g assembly movement. Therefore, the final dynamic model consists of 5 six degrees-of-freedom for the rack plus eight additional mass degrees-of-freedom for the four rattling masses. The remaining portion of the fuel assembly mass is assumed to move with the rack Thus, the totality of fuel mass is included in the ,
base.
simulation.
I;1
- This code has been previously utilized in licensing of similar a racks for Fermi 2 (Docket No. 50-341), Quad Cities 1 and 2 g (Docket Nos. 50-254 and 265), Rancho Seco (Docket No. 50-312),
Oyster Creek (Docket No. 50-219), V.C. Summer (Docket No.
50-395), and Diablo Canyon 1 and 2 (Docket Nos. 50-275 and 50-323).
6-12
?
L e 6.4.2 Evaluation of Potential for Inter-Rack Impact L
Since the racks are closely spaced, the simulation includes impact
~
I
. springs to model the potential for inter-rack impact, especially for low values of the friction coefficient between the support and the pool liner. To account for this potential, yet still retain the simplicity of simulating only a single rack, gap elements were and the L located at the corners of the rack at the top at baseplate. Figure 6.5 shows the location of these gap elements.
e Loads in these elements, computed during the dynamic analysis, are L used to assess rack integrity if inter-rack impact occurs.
r .
L 6.5 STRUCTURAL ACCEPTANCE CRITERIA
( There are two sets of criteria to be satisfied by the rack modules:
[ a. Kinematic Criterion
~
I This criterion seeks to ensure that the rack is a physically stable structure. Byron racks are designed i
to sustain certain inter-rack impact at designated l locations in the rack modules. Therefore, physical stability of the rack is considered along with the I localized inter-rack impacts. Localized permanent deformation of the module is permissible, so long as the subcriticality of the stored fuel array is not l
violated.
- b. Stress Limits The stress limits of the ASHE Code,Section III, Subsection NF, 1983 Edition are used since this code provides the most appropriate and consistent set of l limits for various stress types and various loading conditions. The following loading combinations are applicable (Ref. 1).
6-13
r I
l I
Loading Combination Stress Limit i
D+L Level A service limits l
'D + L + To l 0+L+To+E D + L + Ta + E Level B service limits l D+L+To+Pf 5 D + L + Ta + E' Level D service limits 3
. D+L+Fd The functional capability g of the fuel racks should be demonstrated I
where:
D = Dead weight-induced stresses (including fuel g i assembly weight) g L = Live Load (0 for the structure, since there are no moving objects in the rack load path).
Fd = Force caused by the accidental drop of the i heaviest load from the maximum possible height Pf = Upward force on the racks caused by postulated stuck fuel assembly E = Operating Basis Earthquake E' = Safe Shutdown Earthquake T o = Dif ferential temperature induced loads (normal or upset condition)
T = Differential temperature induced loads a (abnormal design conditions) be I
l The conditions Ta and To cause local thermal stresses to l produced. The worst situation will be obtained when an isolated l storage location has a fuel assembly which is generatlag heat at I
I 6-14 I
Q k
the maximum postulated rate. The surrounding storage locations are p
L essumed to contain no fuel. The heated water makes unobstructed contact with the-inside of the storage walls, thereby producing the temperature difference between the adjacent caxiinum possible cells., The secondary stresses thus produced are limited to the
( body of the rack; that is, the support legs do not experience the secondary (thermal) stresses.
F 6.6 MATERIAL PROPERTIES The data on the physical properties of the rack and support materials, obtained from the ASME Boiler & Pressure Vessel Code,Section III, appendices, and supplier's catalog, are listed in Tables 6.3.and 6.4. Since the maximum pool bulk temperature (except for the full ~ core discharge case) is 150*, this is used as the
(
reference design temperature for evaluation of material properties.
[
6.7 STRESS LIMITS FOR VARIOUS CONDITIONS Tht following stress limits are derived from the guidelines of the ASME Code,Section III, Subsection NF, in conjunction with the y
- k. material properties data of the preceding section.
( 6.7.1 Normal and Upset Conditions (Level A or Level B)
- a. Allowable stress in tension on a net section
( .
=F t = 0.6 S y or F t= (0.6) (23,150) = 13,890 psi (rack material)
(
( 6-15
(
i
I F = is equivalent to primary membrane stresses t
F (.6) (27,500) = 16,500 psi (upper part of t= support feet)
= (.6) (62,400) = 37,440 psi (lower part of support feet)
- b. On the gross section, allowable stress in shear is:
F y
= .4 S
(.4) (23,150)
= 9,260 psi (main rack body) -
F t= (. (27,500) = 11,000 psi (upper part of support feet)
= (.4) (62,400) = 24,960 psi (lower part of support feet)
- c. Allowable stress in compression, F a:
I (1 - (U) 2 2 C, * ]Sy F =
( b )+ (3 ($) 8C, ) - [(b ) 8Cc 3]
3 r r where:
1/ 2 Cc = [ }
S y
k t/ r for the main rack body is based on the full height and cross section of the honeycomb regio.n. Substituting numbers, we obtain, for both support leg and honeycomb region:
Fa = 13,890 psi (main rack body)
Fa = 16,500 psi (upper part of support feet) g
= 37,440 psi (lower part of support feet) E I
6-16
[
- d. Maximum allowable bending stress at the outermost fiber
[. due to flexure about one plane of symmetry:
( Fb = 0.60 Sy = 13,890 psi (rack body).
Fb = 16,500 psi (upper part of support feet)
= 37,440 psi (lower part of support feet)
[ e. . Combined flexure and compression:
C,x fbx C,yfby
[ x bx y by where:
(
f, = Direct compressive stress in the section fbx = Maximum flexural stress along y-axis f by = Maximum flexur$1 stress along y-axis C
mx = C,y = 0.85 f
D x
=1 -
[ 7,ex t
a
[ D =1 -
y 7, ey where
[
=
12w 2E F'ex,8Y z
{
g*'YD 23 ( )
bx,y and the subscripts x,y reflect the particular bending f plane of interest.
6-17 l
- f. Combined Flexure and compression (or tension):
- +
D* + < 1.0 0.6 S y F F bx by The above requirement should be met for both the direct tension or compression case.
I Level D Service Limits 6.7.2 F-1370 (Section III, Appendix F), states that the limits for the I
Level D condition are the minim.um. of 1.2 (Sy /Ft) or limits for Level A times the corresponding (0.7Su/Ft )
condition. Since 1.2 Sy is less .than 0.7 Su for the rack upper part of the support feet, the material, and for the multiplying factor for the limits is 2.0 for the SSE condition for g 5
the upper section. The factor is 1.68 for the lower section under SSE conditions.
I Instead of tabulating the results of these six different stresses as dimensioned values, they are presented in a dimensionless form.
These so-called stress factors are defined as the ratio of the actual developed stress to its specified limiting value. With this 7
definition, the limiting value of each stress factor is 1.0 for OBE and 2.0 or 1.68 for the SSE condition.
l 6.8 RESULTS F.igu re s 6.1, 6.2, and 6.3 show the pool slab motion in horizontal x, horizontal y, and vertical directions.
This motion is for the i SSE earthquake.
! Results are abstracted here for a 12x14 module (the largest I
module) and for an 8 x 14 module (largest aspect ratio).
' I 6-18 I
[
r A complete synopsis of the analysis of the 12x14 and the Sx14 L
module subject to the'SSE earthquake motions is presen'ted in a I summary table 6.5 which gives the bounding values of stress factors R1 (i = 1,2,3,4,5,6). The stress factors are defined ast R t = Ratio of direct tensile or comprese,1ve stress on a net section to its allowable value (nots support feet only ]
{ support compression) l R 2 = Ratio of gross shear on a net scotion to its allowable
' ^
r value L
R 3 = Ratio of maximum bending stress due to bendlng about the
( x-axis to its. allowable value for the 'section Rg = Ratio of maximum bending stress due to bending abo'ut the y-axis to its allowable value
{
R 5 = Combined flexure and compressive 'f actor (as defined in 6.7.1e above)
R 6 = Combined flexure and tension (or compression) factor (as defined in 6.7.1f above)
As stated before, the allowable value of R i '.1 =1,2,3,4,5,6) is
[
1 for the OBE condition, (except for the lower section of the 5 support where the factor is 1.68). and 2 for the SSE.
The . dynamic analysis gives the maximax (maximum in time and in space) values of the stress factors at critical locations in the rack module. SLnce these maximax values are subject to minor (under 5%) variation if the input data (viz rack baseplate height, cell inside dimension) is perturbed within the range of manufacturing tolerances, the bounding values, instead of .the actual values, are presented in Table 6.5. The terms in Table 6.5
( have the following meanings r a implies Ri < 1.0 L b. implies R i < 1.5 e implies Ri < 1.75 d implies R i < 2.0
( 6-19 f - - - -
It is found that'the results corresponding to SSE are~most critical vis-a-vis the 'orresponding allowable limits. The results given herein are for the SSE. The maximum stress factors (R i) are below the limiting value for the SSE condition for all sections. It is noted that the critical load factors reported for the support _ feet are all for the upper segment of the foot and are to be compared with the limiting value of 2.0.
Analyses (not included here) have been carried out to show that significant margins of safety exist against local deformation of the fuel storage cell due to rattling impact of fuel assemblies and against local overstress of imp.act bars'due to inter-rack impact.
Analyses (not presented here) have also been carried out for the OBE condition to demonstrate that the stress factors are below 1.0. Results obtained for all rack sizes and shapes are enveloped by the data presented herein. Overturning has also been considered for the cases where racks are cdjacent to open areas.
I I
6-20
I 6.9 IMPACT ANALYSES 6.9.1 Impact Loading Between Fuel Assembly and Cell Wall The local stress in a cell wall is estimated from peak impact loads obtained from the dynamic simulations. Plastic analysis is used to obtain the limiting impact load that can be tolerated. Including a I safety margin of 2.0, number of cells (NC) is:
we find that the total limit load for the I
QL = 9030 NC 6.9.2 Impacts Between Adjacent Racks All of the dynamic analyses assume, conservatively, that adjacent racks nove completely out of phase. Thus, the highest potential for inter-rack impact is achieved. Based on the dynamic loads I obtained in the gap elements simulating adjacent racks, we- can study rack integrity in the vicinity of'the impact point. The use of framing mate-ial around the top of the rack allows us to withstand impact loads of 57000 lbs. at a corner of the rack prior to reaching the fully yielded state above the active fuel region.
It is shown that tack-to-rack corner impact loads' can be accommodated. Thus, impacts between racks can be accommodated without violating rack integrity.
I 6-21 e
Y I
6.10 WELD STRESSES ,
weld locations under seismic loading are at the The critical connection of the rack to the baseplate and in the support leg welds. For the rack welds, the allowable weld stress is the ASME Code value of 24000 psi (T able NF-3324.5(a)-1, Subsection NF). For the support legs, the allowable weld stress is governed by the levels outlined in Section 6.7 (see NF-3324.5 for partial penetration welds).
Weld stresses due to heating of an isolated hot cell are also computed. The assumption used is that a 91ngle cell is heated, over its en':fre length, to a temperaturu above the value associated with all surrounding cells. No thermal gradient in the vertical direction is assumed so that the results are conservative. Using the temperatures associated with this unit, we show that the skip welds along the entire c' ell length do tio t exceed the allowable value for a thermal loading condition.
hs 6.11
SUMMARY
OF MECHANICAL ANALYSES The mathematical model constructed to determine the impact velocity of falling objects is based on several conservative assumptions, such,as:
- 1. The virtual mass (see ref. 8-10 for further ma tie r ial on this subject) of the body is conservatively assumed to be equal to its displaced fluid mass. Evidence in the i
literature (Ref. 12), indicates that the virtual mass can be many times higher.
I
- 2. The minimum frontal area is used for evaluating the drag coefficient.
I 6-22
~
b
- 3. The drag coef ficients - utilized in the analysis are the lower bound values reported in the literature (Ref. 13).
r- at the beginning of the f all when the L In particular, velocity of the body is small, the corresponding Reynolds number is low, resulting in a large drag coefficient.
(
bodies are assumed to be rigid for the
- 4. The falling
{ purposes of impact stress calculation on the-rack. The solution of the immersed body motion problem is found analytically. .The impact velocity thus computed is used to determine the maximum stress generated due to stress wave propagation.
( With this model, the following analyses are performed:
I a. Dropped Fuel Accident I h .
A fuel assembly (weight = 1545 pounds with control rod
, assembly) is dropped from 36 inches above the module and impacts the base. The final velocity of the dropped fuel assembly (just prior to impact) is calculated and, thus, the total energy at impact is known. To study baseplate integrity, we assume that this energy is all directed punching the baseplate in shear and thus toward of work done 'by the supporting shear transformed into
, stresses. It is determined that shearing deformation of the baseplate is less than the thickness of the base-plate so that we conclude that local piercing of- the baseplate will not occur. Direct impact with the pool i liner does not occur. The subcriticality of the adjacent fuel assemblies is not violated.
l i
[ !
,6 23
l I
- b. Dropped Fuel Accident II One fuel assembly drops from 36 inches above the rack and hits the top of the rack. Permanent deformation of the rack is found 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 below) is not altered. The region of local permanent deformation does not extend below 6 inches from the rack top. An energy balance approach is used here to obtain the results.
I
- c. 3ammed Fuel-handling Equipment A 4000-pound uplift force is applied at the top of the I
rack at the " weakest" storage location; the force is assumed to be applied on one wall of the storage cell boundary as an upward shear force. -T h e plastic deformation is found to be limited to the region well above the top of the active fuel.
These analyses prove that the rack modules are engineered to provide maximum safety against all postulated abnormal and accident conditions.
6.12 DEFINITION OF TERHS 'JSED IN SECTION 6 S1, S2, S3, S4 Support designations Pi Absolute degree-of-freedom number i qi Relative degree-of-freedom number i u Coefficient of friction Ui Pool floor slab displacement time history in the 1-th direction x,y coordinates horizontal direction z coordinate vertical direction I
6-24
F L
r Impact spring between fuel L K1 ,
assemblies and cell Kf Linear component of friction spring i K6 Axial spring of support leg locations N Compression load in a support foot Rotational spring provided by the f KR pool slab E
Subscript i When used with lJ or X indicates direction (i = 1 x-direction, 1=2 l y-direction, i = 3 z-direction)
I L
b i ~
E L
F I
l I
l .
6-25
I REFERENCES TO SECTION 6 I
- 1. USHRC Standard Review Plan, HUREG-0800 (1981).
ASME Boiler & Pressure Vessel Code,Section III, Subsection 3
- 2. W NF (1983).
- 3. USNRC Regulatory Guide 1.29, " Seismic Design Classification,"
Rev. 3, 1978.
- 4. " Friction Coefficients of Water Lubricated Stainless Steels for a Spent Fuel Rack Facility," Prof. Ernest 1976.
Rabinowicz, MIT, a report for Boston Edison Company, USNRC Regulatory Guide 1.92, " Combining Modal Responses and g
- 5. 1, W Spatial Components in Seismic Response Analysis," Rev.
February 1976.
- 6. "The Component Element Method in Dynamics with Application to Levy and 3.P.D.
Earthquake and Vehicle Engineering," S.
Wilkinson, McGraw Hill, 1976.
- 7. " Dynamics of Structures," R.W. Clough and 3. Penzien, McGraw -
Hill-(1975). .
- 8. " Mechanical Design of Heat Exchangers and Pressure Vessel Soler, Arcturus Components," Chapter 16, K.P. Singh and A.I.
Publishers, Inc., 1984.
- 9. R.3. Fritz, "The Effects of Liquids on the Dynamic for Motions Industry, of Immersed Solids," Journal of Engineering Trans. of the ASME, February-1972, pp 167-172.
Spaced Two-Body System
- 10. " Dynamic Coupling in a Closely The Case of Fuel Racks ,"- K.P . 3 Vibrating in Liquid Medium: 3; Singh and A.I. Soler, 3rd International Conference on Nuclear Power Safety, Keswick, England, May 1982.
USNRC Regulatory Guide 1.61, " Damping Values for Seismic
- 11. l Design of Huclear Power Plants," 1973. ,
- 12. " Flow Induced Vibration," R.D. Blevins, VonNostrant (1977). .
" Fluid Mechanics," M.C.. Potter and 3.F. Foss, Ronald Press, 13.
p 459 (1975).
I I
6-26
--"-~ - ___
E F-Table 6.1 DECREES OF FREEDOM
{
Displacement Rotation u 0 Location u* Y u*' 6* Y 0*
[- (Hode) 1 pt p2 p3 94 45 'q s 1* Point 1* is assumed fixed to base at XB,YB, Z=0 2 Point 2 is assumed attached to rigid rack at the top most point.
2* p7 pa Pi = qi(t) + U i(t)
[ Otter p 3, p to Rattling { p it, p 12 Hode points 3*, 4*, 5*
Masses p is, p is l
4 e
[
6-27 l
I Table 6.2 HUMBERING SY3 TEM FOR GAP ELEMENTS AND FRICTION ELEMENTS I. Nonlinear Springs (Cap Elements) (60 total)
Number Node Location Description 1 Support S1 Z compression only element 2 Support S2 Z compression only element 3 Support S3 Z compression only element g 4 Support S4 Z compression only element 3 5 2,2* X rack / fuel assembly impact element 6 2,2* X rack / fuel assembly impact element E 7 2,2* Y rack / fuel assembly impact element g 8 2,2* Y rack / fuel assembly impact element 9-20 Other rattling masses 21 Bottom cross-section Inter-rack impact elements of rack (around edge) g Inter-rack impact elements 5
. Inter-rack impact elements
. Inter-rack impact elements g
. Inter-rack impact elements g
- Inter-rack impact elements
- . Inter-rack impact elements 40 Inter-rack impact elements 41 Top cross-section Inter-rack impact elements l .
of rack (around edge) Inter-rack Inter-rack impact impact elements elements 5
W
= Inter-rack impact elements
. Inter-rack impact elements 5 Inter-rack Inter-rack impact impact elements elements 5
l 60 Inter-rack impact elements II. 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 E 4 Support S2 Y direction friction g 5 Support S3 X direction friction 6 Support S3 Y direction friction 7 Support S4 X direction friction i 8 Support S4 Y direction friction I
9 S1 X Slab moment 10 S1 Y Slab moment g 11 S2 X Slab moment 3 12 S2 Y Slab moment 13 S3 X Slab moment E 14 S3 Y Slab moment g 15 S4 X Slab moment 16 So Y Slab moment 6-28
I .
I .
l Table 6.3 RACK MATERIAL DATA Young's Yield Ultimate Modulus Strength Strength (psi)
Haterial E (psi) S y (psi) S 8 68100 304L S.S. 27.9 x 10 23150 Section III Table Table Table Reference I-6.0 I-2.2 I-3.2 I
3 Table 6.4 SUPPORT HATERIAL DATA Young's Yield Ultimate Material Modulus Strength Strength I
6 68,100 psi, 1 SA-351-CF3 27.9 x 10 psi 27,500 psi (upper part of support feet) 6 90,000 psi 2 SA-217-CA15 27.9 x 10 62,400 psi (lower part of support feet) 1 I 6-29 l .
--- - , .--- .,,,,,.- _.,..--,--, .--,,.n, , . - - - , - - , - . - - ---
i l Table 6.5
! BYROH RACKS - BOUNDIHC VALUES FOR STRESS FACTORS i
l Stress Factorst R1 R2 R3 Rg R5 R6 Run No. (Upper values for rack base - lower i values fo'r support feet)
C012, SSE a a a a a a y= .8, full a a b b b b 12 x 14 i
I C013, SSE a a a a a a
- i p= .2, full
\$ 12 x 14 i
C014, SSE a a a a a a p= .8, 16 a a a a b b cells filled j 12 x 14 1
i t The terms a, b, c and d Imply the stress factors R1 (i = 1,2.. 6) are bounded by the following limiting values l a 1.0 .
} b: 1.5 c: 1.75 d 2
! M M W W W W W W W M M M M W M M M M
- mm r- r, vm r- r - _r rw Table 6.5 (continued) .
BYRON RACKS - BOUNDING VALUES FOR STRESS FACTORS Stress Factors Rt R2 R3 Rg R3 R6 Run No. (Upper values for rack base - lower values for support feet)
C015, SSE a a a a a a y= .2, 16 a a a a a a l
i cells filled 12 x 14 C016, SSE a a a a a a
!m '
p= .8, 1/2 Full a a a a b b l Pos. X Quadrant 12 x 14 C017, SSE a a a a a a p= .2, 1/2 Full a a a a a a Pos. X Quadrant 12 x 14 C018, SSE a a a a a a p= .8' a a b b c c C1 Edge Rack
w, s, Table 6.5 (continued) ,
BYROH RACKS - BOUNDING VALUES FOR STRESS FACTORS Stress Factors Rt R2 R3 Rg Rs R6 Run No. (Upper values for rack base - lower values for support feet)
C019, p= .2 a a a a a a C1 Edge Rack a a a a a a 12 x 14 m CO20, SSE a a a a a a O p= .8, Full a a b b c c n
8x 14
! CO21, SSE a a a a a a I p= .2, Full
- a a a a a a
! 8x 14 CO22, SSE a a a a a a p= .2, a a a a b b I
11 Cells Filled
- 8 x 14 i
l M W W W W W m M M M M M M
TR V F
Table 6.5 (continued)
BYROH RACKS - BOUNDING VALUES FOR STRESS FACTORS ,
Stress Factors Rt R2 R3 Rg R3 R6 Run No. (Upper values for rack base - lower values for support feet)
CO23, SSE a a a a a a
"* *0 a a a a b b 11 Cells filled 8 x 14 T
U CO24, SSE a a a a a a p= .8, 1/2 Full a a a a b b in Heg. X Half 8 x 14 CO25, SSE a a a a a a .
p= .2, 1/2 Full a a' a a a a I
in Heg.X Half 8 x 14
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- XB, YB - Location of fuel rod l ,
g q, group mass centroid - relative to centerline of fuel rack I '
FIGURE e.4 Dynamic Model 6-37
TYP.TOPIMPACT l ..
ELEnMNT ll
\ -
x~ } j' ) l w mg A *_ s~,v - mH l fff - ff I n f [ _BAR,M STRUCTURE l 1 i TYP. BOTTOM IMPACT pr ur Il I w r .
./ ./ ./
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i:) " g
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fr- [ { { b, g I I I FIG. 6.5 G AP. ELEMENTS TO SIMUL ATE INTER-R ACK IMP ACTS 6-38 I I
9 L r [ b y impact Springs I [ , 6 J b r. n y P Mass [ w ~ 2* LJ 5 b 1
. \ Fluid
[ Dampers 4 ' (not used in [ this analysis) ~ Rigid [ Frame [ X FIGURE 6.6 IMPACT SPRINGS AND FLUID D AMPERS { (
~
6-39 l
l I f "' : I Kw K-w / _g
;- 2
-4 2
;rm~g a E
UGap' Elements To Simulate Inter-Rack Impacts 5 Rack Centroid
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E I
- H I
[ Rigid Rack I Baseplate ' Kw K, l W) ,' 3
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Kh Kg h K9 _ wj ,, th Y rr-K V RfK f I FIGURE e.7 Sprina Mass Simulation For I Two-Dime n sio nal Motion g 6 40 I
J b 7.0 ENVIRONMENTAL EVALUATION 7.1
SUMMARY
Installation of high density spent fuel storage racks at the F Byron Nuclear Power Station will increase the licensed storage capacity of the spent fuel pool from 1060 to a maximum of 29 t+ 0 e assemblies. Radiological consequences of expanding the capacity have been evaluated with the objective of determining if there is a significant additional onsite or offsite radiological impact relative to that previously reviewed and evaluated (Ref. 1). In addition, radiological impact to operating personnel has been evaluated to ensure that exposures remain as low as is reasonably achievable (ALARA). F L The decay heat loading and the radiological burden to the spent fuel pool water are determined almost entirely by refueling g L operations. The frequency of refueling operations N and the conduct of refueling are independent of the increased capacity of the storage pool, except that the increased capacity should reduce fuel movement and allow continued normal operation. Since the fuel assemblies which will utilize the bulk of the storage capacity (and will ultimately fill all incremental capacity above r that of the existing design) are aged, their contribution to L either the peak decay-heat load or the increased radiological impact, in terms of increased doses, is negligible. A study performed by the NRC (Ref. 2) supports this conclusion. Consequently, the increase in the storage capacity of the spent fuel pool will neither significantly alter the operating characteristics of the current pool nor result in a measurable change in impact on the environment. 7.2 CHARACTERISTICS OF STORED FUEL Because of radioactive decay, the heat generation rate and the intensity of gamma radiation from the spent fuel assemblles 7-1
I I decreases substantially with decay time. After a cooling time of about 4 years (Ref. 3), the decay heat generation rate is less than 2% of the rate at 7 days, the nominal time at which depleted fuel assemblies are transferred to the spent fuel pool. The intensity of gamma radiation is very nearly proportional to the decay heat and decreases with cooling time in a similar manner. The bulk of the heat load is due to freshly discharged fuel; g aged fuel contributes relatively little to the total heat load. 5 Therefore, this expansion will not 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 will be no significant increase in gamma radiation beyond the pool environment due to the expanded storage. It is important to note .that the aged fuel in the expanded storage capacity will not contain significant amounts of radioactive lodine or short-lived gaseous fission products, since would have decayed during the storage period. The these Krypton-85 which might escape from defective fuel assemblies has been shown to do so quickly (Ref. 2) (i.e., within a short time after discharge from the. core). Further, the residual Krypton-85 l will be contained within the fuel pellet matrix and hence any leakage would occur at very low rates (Ref. 2). Cesium 134/137 (Ref. 2) is strongly bound within the fuel pellet matrix and its dissolution rate in water is extremely small. Any Cesium l l dissolved in the pool water is easily controllable in ,the cleanup . system (demineralizer-lon exchanger resin bed) (Ref. 2). Thus, the planned storage expansion will not significantly increase the release of gaseous radionuclides. l l 7.3 RELATED IHOUSTRY EXPERIENCE I Experience with storing spent fuel underwater has been substantial (Refs. 2, 3, and 4). These references show that the 7-2 l - t .
[ pool water activity, normally low, experiences small increase dur'ing refueling periods, which then decays rapidly with time. I Typical concentrations (Ref. 5) of radionuclides in spent fuel pool water range from 10 pC1/ml to 10 3 pC1/ml, with the higher value associated with refueling operations. References 2 and 5 also state that the increase in pool water activity during refueling can be attributed tot 7 O Dislodging (sloughing off) of corrosion products on the fuel assembly transfer and during handling operations. O The possible short-term exposure of fuel pellets to pool water via a cladding defect. O Mixing of the spent fuel pool water with the higher L activity reactor coolant. Upon cessation of the refueling operations, the fuel pool water and the reactor coolant system would be isolated from each r other, thereby terminating transports of corrosion L products from the reactor coolant system. Thus, deposition of crud is a function of refueling operations and is not impacted by the expanded storage. O Furthermore, it has been shown (Ref. 6) that release of fission products from failed fuel decreases rapidly after shutdown to essentially negligible levels. The dissolution of exposed fuel pellets (made of UO 2) is very slow in water at fuel pool temperatures and the corrosion of the cladding (Zircaloy 4) at spent fuel [ pool water temperatures is virtually nil (Refs. 2 and Another mechanism available for the release of the 5). gaseous fission products is diffusion through the UO 2 has been shown that at low water ( pellet. temperatures It (<150*F), the diffusion coefficient is extremely small (Ref.7). 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 refueling operations. The expansion of fuel pool storage capacity will not cause a significant increase in doses either onsite or offsite. The corrosion properties of irradiated Zircaloy cladding ( have been reviewed in References 2 and 4 and the conclusion is drawn that the corrosion of the cladding in spent fuel pool water is negligible. The minor incremental heating of pool water, due to the expansion of storage capacity, is for too small to materially affect the corrosion properties of Zircaloy cladding. 7-3
I I 7.4 BYRON NUCLEAR POWER STATION EXPERIENCE I At present there are'no spent fuel assemblies in the spent fuel pool. 7.5 SPENT FUEL POOL COOLING AND CLEANUP SYSTEM It has been shown previously in Section 5 of this licensing report that the cooling system at Byron is adequate to handle the expected heat loads and maintain the pool temperature peaks within acceptable limits. It has been shown in Section 5 that a the small increase in heat load due to the storage capacity expansion will neither significantly increase the thermal dissipation to the environment nor increase the propensity for corrosion of the cladding. It has also been shown that the crud deposition in the spent fuel pool water occurs during refueling outages and that the planned expansion will not increase long-term crud deposition'.' The fuel pool cleanup system (filter and demineralizer) is designed to maintain fuel pool water clarity and is operated and maintained in accordance with the Byron operating procedures. The cleanup l system takes a surface skim from the fuel pool and cleans it through a process of filtration and demineralization to prevent crud buildup on the fuel pool walls at the water-to-air interface. l The sp'ent fuel pool water is sampled'and analyzed periodically to confirm proper operation of the pool cleanup system. The spent l fuel pool modification will not result in a significantly higher quantity of solid radwaste. ' E I 7-4 I
r L Y r ( 7.6 FUEL POOL RADIATION SHIELDING 7.6.1 Source Terms ( E The spent fuel gamma source terms used for the fuel pool shielding evaluation were generated using the point reactor fission product inventory code RIBD. The f ollowing ' assumptions were used in the analysis: l 0 Initial fuel enrichment = 4.2% O Het Reactor core power level = 3411 HWt Het k 0 Average assembly discharge burnup = 38,000 MWD /HTU p 0 Power level for average assembly = 17.67 MWt L 0 Power level for peak power assembly = 29.16 HWt (peaking factor of 1.65) O Burnup for peak power assembly = 33,000 MWD /HTU (one~18-month cycle at maximum power) O Start of Refueling. Process = 100 hours after shutdown [ The peaking factor of 1.65 and burnup for peak power assembly of 33,000 HWD/HTU for one cycle were chosen to produce the highest r possible gamma source term attainable under operational b conditions. The average assembly burnup and power level were y chosen to represent a conservative gamma source term for the L spent fuel. The 100-hour decay time is the minimum permitted period before refueling can begin per the Technical Specifications. - The photon energy production rates of an average assembly and of a peak power spent fuel assembly are given in Tables 7.1 and 7.2, respectively. [ , 7-5
I' 7.6.2 Radiat an Shielding For an equilibrium refueling cycle 84 fuel assemblies will be I discharged into the pool starting no earlier than 100 hours after reactor shutdown at a rate of 1 fuel assembly per hour. For a full core discharge, 193 assemblies are discharged. To evaluate the adequacy of the shielding capability of the spent fuel pool walls, the radiation dose from 104 freshly discharged fuel assemblies arranged in the A1 Region 1 fuel storage rack (see Figure 2.1) is calculated. It is not necessary to consider the effects of the remaining 89 fuel assemblies in the 81 Region 1 rack on the north pool wall since the A1 rack effectively shields the adjacent area to the north from the B1 storage rack assemblies. The radiation effects on the east / west walls are based on 386 assemblies. The photon production rates used in the calculations are tabulated in Table 7.2. The north and south perimeter walls of the spent fuel pool are 5 feet thick. The east and west perimeter walls of the pool are 6 feet thick. The wall separating the spent fuel pool from the transfer canal is 5 feet 6 inches thick. These walls, together with the water and fuel storage racks are incorporated in the dose rate calculations for the adjacent areas. The results of the calculations are provided in Table 7.3. Except for the area immediately adjacent to the A1 rack, the maximum calculated dose rates through the pool walls are less than the currently designated radiation level limit of 20-mrem /hr for these areas. The area adjacent to the A1 rack is normally occupied only briefly during the local and integrated leak rate tests which are conducted every 5 years. Access to the area is controlled. I 7-6
l lI~ il The dose rates in Table 7.3 are considered an upper limit since they are calculated for freshly discharged fuel. The dose rates will reduce by a factor of six, 60 days after the fuel is discharged into the pool. above discussion indicates that the shielding available I The around the spent fuel pool is adequate for installation of the high density storage racks. I . The radiation dose level at the side of the pool and on the spent fuel pit crane bridge due to the transfer of a peak power fuel assembly are 2.5 mrem /hr and 2.0 mrem /hr, respectively. These , calculations assume a minimum of 10 feet of water cover over the active fuel. If the transfer of fuel assemblies into,the spent I fuel cask pit becomes necessary, the water cover will be less than 10 feet over the active fuel, however, the dose rate during this fuel movement will be much lower than for the transfer of I the peak power assembly noted above since the radiation level' of the fuel assembly will have had time to decay to a level which would more than compensate for the loss of water cover shielding. Since the fuel transfer operation normally lasts less than 4 days (88. assemblies at 1 assembly per hour), the above radiation field does not create excessive operator exposure. 7.7 RADIOLOGICAL CONSEQUENCES I The design, basis fuel handling accident (dropped assembly) in the Fuel Handling Building in Section 15.7 of the FSAR was reviewed for possible effects on radiological dose consequences. The review determined that the conclusions in the FSAR were still valid and that offsite radiological dose consequences were well within 10CFR100 limits. I E '-'
I I 7.8 RERACKING OPERATION Installation of the fuel racks will
~
include removal of the I existing racks, making minor pool modifications, and cleaning and installing the new racks. The existing racks are bolted to the pool floor. The new racks will- be cleaned prior to installation. The fue 5 handling building overhead crane will be used to place the racks W in the pool. This effort is scheduled to be performed prior to the first refueling for Unit 1, which will allow a " dry" l l installation with no water or spent fuel in the pool. In this ( case, the existing fuel racks will not have been exposed to spent fuel and will only be nominally contaminated, if at all. Therefore, doses to individuals involved in the reracking will be negligible. If there is a. delay'in ins 0alling the high density racks until I after the first refueling, then a " wet" installation will be required. All pool modificat,'.ons tha! can be comoteted prior to filling the pool with water will be done to minimize underwater work. Although divers may be needed for some tasks, all of the
, work associated with the inatallation will be sequenced to minimize potential radiation exposure c, f personnel due to the
' spent fuel located in the pool. ALARA considerations will be fully incorporateo in the installation prucedures for this condition. If the fuel handling building overhead crane is used over the pool, electrical interlocks will be adjusted on the crane to preclude carrying racks over any stored fuel assemblies. Exact disposition of the existing racks har not been determigid. They will be decontaminated and/or packaged a r, d disposed of in accordance with the applicable Federal and State regulatio3s. ; I 7-8
L
7.9 CONCLUSION
S Based upon the industry experience and evaluations discussed in previous sections, the following conclusions are made: 'S : { Minor increases in radiological burden to the pool
- 'f 7 water, if any, can be adequately handled by the fuel I
pool cleanup system (filter and demineralizer), thereby maintaining the radionuclide concentration in the water at an acceptably low level. No appreciable increase in solid radioactive wastes (i.e., filter media and demineralizer resin) is { anticipated. 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 o,f residence in ' the pool. [ Further, Vol. 1, ' Reference 3 (pp. 4-16) has shown airborne in 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 ga'ses. I The existing spent fuel pool coolihg system will keep the pool water temperature at an acceptable level (Se.e Section 5, Thermal-Hydraulic Considerations). The existing radiation protection monitoring systems and program are adequate to detect and to 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. I 7-9 b--- .____ _ _ _ _
I For a dry reracking operation, radiation exposures will be extremely low. If the reracking occurs after the first refueling, procedural controls and necessary precautions will be taken to reduce radiation exposure to as low as is reasonably achievable, and hence, radiological impact will be minimized. Expanding the storage capacity of the spent fuel 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 nonradiological impact anticipated. E
~
I I I I I I I I I 7-10
. - - . - - - - - . , _ _ _ ~ . , - - - _ _ _ ~ _ , _ . - _ . . .- __ . - - . _ - _ . _ , . . - , , _ . . - . , _ _ _ . _ _ . . . . _ _ _ _ - _ . . - . - _ _ _ . - .
F REFERENCES TO SECTION 7 I
- 1. FSAR, Byron Nuclear Power Station.
, 2. NUREC-0575, " Handling and Storage of Spent Light Water L Power Reactor Fuel," Vol. 1, Executive Summary and Text, USNRC, August 1979.
- 3. HUREC-0800, USHRC Standard Review Plan, Branch Techni-cal Position ASB9-2, Rev. 2, 3uly 1981.
- 4. 3. R. Weeks, " Corrosion of Materials in Spent Fuel Storage Pools," BHL-NUREG-2021, July 1977.
- 5. A. B. Johnson, 3r., " Behavior of Spent Nuclear Fuel in j
y Water Pool Storage," BNWL-2256, September 1977.
- 6. 3. M. Wright, " Expected Air and Water Activities in the Fuel Storage Canal," WAPD-PWR-CP 1723 (with Addendum),
( undated. I 7. ANS 5.4 Proposed Standard, " Method for Calculating the L Fractional Release of Volatile Fission Products from Oxide Fuel," American Nuclear Society, issued for r review, 1981. 1
- 8. " 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.
[
- 9. " Licensing Report for High Density Spent Fuel Storage Racks," Rancho Seco Nucle =r Generating Station, Sacra-mento Municipal Utilities District, Docket No. 50-312, June 1982.
- 10. Final Safety Analysis Report, Limerick Generating Station Units 1 and 2, Section 9.1
- 11. Safety Evaluation Report Related to the Operation of Limerick Generating Station Units 1 and 2, NUREG-0991, August 1983.
- 12. Source Term Data for Westinghouse Pressurized Water Reactors, WCAP-8253, July 1975.
[ 7-11
-- ---- - - _ _ ~ _ _ _ _ _ _ . _ _ _ _ _ _ _ _ _
I I Table 7.1 PHOTON ENERGY PRODUCTION RATES OF AN AVERACE SPENT FUEL ASSEMBLY Photon Energy Photon Energy Production Rate I (MeV) (MeV/sec) 1.50E-002 2.38E+016 2.50E-002 9.64E+015 3.50E-002 1.52E+016 3 5.50E-002 9.94E+015 . 3 8.50E-002 1.08E+016 1.50E-001 1.75E+016 2.50E-001 4.82E+015 3.50E-001 9.70E+015 5.60E-001 4.72E+016 9.10E-001 4.10E+016 1.35E+000 1.50E+016 1.80E+000 2.07E+014 2.20E+000 4.09E+014 3 2.60E+000 4.96E+014 5 3.00E+000 2.05E+013
. TOTAL Z.05E+017 I
~
I 3 g L l E I I l 7-12 I
. _ _ _ _ _ _ . _ _ . . . _ _ _ _ _ . - _ _ _ _ _ _ . _ _ . . . _ _ _ . _ . _ ___._.'__._______1.__
F L Table 7.2 PHOTON ENERGY PRODUCTION RATES OF PEAK SPENT FUEL ASSEMBLY ( F Photon Energy Photon Energy Production Rate L (MeV) (MeV/sec) 1.50E-002 3.92E+016 2.50E-002 1.59E+016 3.50E-002 2.50E+016 , 5.50E-002 1.64E+016 l 8.50E-002 1.79E+016 1.50E-001 2.88E+016 2.50E-001 7.96E+015 3.50E-001 1.60E+016 5.60E-001 7.78E+016 9.10E-001 6.77E+016 1.35E+000 2.47E+016 [ 1.60E+000 3.41E+014 2.20E+000 6.75E+014 2.60E+000 8.18E+014 3.00E+000 2.60E+013 TOTAL 3.39E+017 [ E 7-13
I Table 7.3 I CALCULATED DOSE RATES IN AREAS A03ACENT TO THE SPENT FUEL POOL High Density Rack Location Dose Rate (mrem /hr) E Floor el. 401 ft, O in., areas 54.0 adjacent to the north walls
- Floor ei. 426 ft, O in., areas 2.3 adjacent to the edge of the pool.
Floor el. 401 ft, O in., spent fuel 4.0 pool heat exchanger area ** Fuel transfer canal ** - 20.0 I
- A design water gap of 4-7/8 inches between the high density rack and the wall is used.
- A design water gap of 4 inches between the high density rack and the wall is used.
I I I I I E 7-14 I
k e l L 8.0 p -SERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON' ABSORBING MATERIAL J l
.8.1 PROGRAM INTENT b
A sampling program to verify the integrity of the neutron c absorber material employed in the high density fuel racks in the L long-term environment is described in enis section. I ? L The program is conducted in a manner which allows access to the representative absorber. material samples without disrupting the integrity of the e n t i'r e ' f u e l storage system. The program is tailored to evaluate the material in normal use mode and- to forecast future changes using the data base developed.
8.2 DESCRIPTION
OF SPECIMENS The absorber material used in the surveillance program,
~
henceforth. referred to as poison, is representative of the material used within the storage system. It is of the same composition, produced by the' same method, and certified to the ( The sample coupon is same criteria as the production lot poison. of similar thickness as the poison used within the storage system { and not less than 4 by 2 inches on a side. Figure 8.1 shows a typical coupon. Each poison specimen is encas'ed 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 similac to the design used for the storage system. The jacket has to be closed by tack welding in such a manner as to retain its form throughout the test period and still allow rapid and easy opening without causing mechanical The jacket {' damage to the poison specimen contained within. should permit wetting and venting of the specimen similar to the actual rack environment.
. 8-1
8.3 SPECIMEN EVALUATION After the removal of the jacketed poison specimen from the cell at a designated time, a careful evaluation of that specimen should be made to determine its actual condition as well as its apparen't 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 be visually examined for any effects of environmental exposure. Specific attention should be directed to the examina-tion of the st inless steel jacket for any evidence of physical degradation. Functional evaluation of the poison material can be accomplished by the following measurements: 0 A neutron radiograph of the poison specimen aids in the determination of the maintenance of . unif ormity of the boron distribution. . O 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. O A measurement of the hardness of the poison material and I will establish the continuance of physical structural durability. The hardness acceptability criterion requires that the s'pecimen hardness will not reduce the hardness listed in the qualifying tegtg document for laboratory test specimen irradiated to 10 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 pre-irradiated sample. O Measurement of the length, the width, and the average l thickness and comparison with the preexposure data will 5 indicate dimensional stability within the variation range reported in the Boraflex laboratory test reports. A procedure will be prepared for execution of the test procedure and interpretation of the test data. 8-2
] q L e ,, W .07 5 THK. FILLER - ' E F R REGION 1. / o 040"THK FILLER . r { FO R--REGION 2 i L o. ( - 3 04 L S.S T. ).
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>- TACK WELD /
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i . c ./ /- [ NEUTRON - ! s%3 7 .
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48
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FIG. 8'.1 - TEST C0ljPON
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8-3
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9.0 COST / BENEFIT ASSESSMENT L [ A cost / benefit assessment has been prepared in accordance with the requirements of Reference 1, Section V - Part 1. The assessment demonstrates that the installation of high density spent fuel storage racks is the most advantageous means of handling spent fuel. The material is presented merely for informational purposes. It is CECO's position that an environmental impact statement need not be prepared because the installation of high density fuel racks provides no significant impact on the environment. NRC precedent establishes that alternatives and economic costs need not be discussed when there is no significant environmental impact. However, for completeness, alternatives to reracking for additional spent fuel storage capacity are discussed in Section 9.3. 9.1 SPECIFIC NEED'S FOR SPENT FUEL STORACE Disposal of Byron nuclear fuel is scheduled to be carried out by the Department of Energy in or after 1998 in accordance with Public Law 97-425, Huclear Waste Policy Act of 1982. As Byron spent . fuel may not be accorded a high priority under the DOE Program, CECO is seeking to provide a spent fuel storage capacity to support approximately 25 years of nominal operation. No other contractual arrangements exist for the interim storage, or c'eprocessing of spent fuel from Byron Nuclear Power Station. Therefore, increased storage capacity in the spent 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 2011, 17 years beyond the ' current capability and 13 years beyond the scheduled repository fuel receipt date, per the. . DOE Hission Plan.
\
9 9-1
I 9.2 COST OF SPENT FUEL STORAGE 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 $2.6 and
$2.9 million.
I 9.3 ALTERNATIVES TO SPENT FUEL STORAGE I CECO has considered the various alternatives to the proposed onsite spent fuel storage. These alternatives are discussed belows
- a. Shipment of Fuel to a Reprocessing or Independent Spent Fuel Storage / Disposal Facility No commercial spent fuel repr_ocessing facilities are presently operating in the United States. CECO has made arrangements whereby spent nuclear fuel.
' contractual and/or high level nuclear waste will be accepted and disposed of by the U.S. Department of Energy. However, such services are not expected to be available before 1998. The existing Byron spent fuel storage capacity g will not provide full core discharge capability beyond 3 1994. Spent fuel acceptance and ' disposal by the Department of Energy is not, therefore, an alternative g to increased onsite pool storage capacity. g
- b. Shipment of Fuel to Another Reactor Site Shipment of . Byron fuel to another reactor storage site could capacity provide short-term relief to the problem. However, . transshipment of spent fuel merely g serves to transfer the problem to another site and.does 3 not result in any additional net long-term storage capacity. Accordingly, CECO does not consider the transshipment of spent fuel to be an appropriate alternative to high density spent fuel storage at the site.
- c. Not Operating the Plant after the Current Spent Fuel Storage Capacity is Exhausted 3
As indicated in NUREG-0575, " Final Environmental Impact Statement on Handling and Storage of Spent Light Water Power Reactor Fuel," (Ref.2) the replacement of E 9-2
( .. l I ,I nuclear power by coal generating capacity would cause excess mortality to rise from 0.59 - 1.70 to 15 - 120 per year for 0.8 GWY(e). Based on these facts, not ! operating the plant or shutting down the plant after exhaustion of spent fuel discharge capacity is not a I viable alternative to high density storage in the spent. fuel pool. The prospective 1986 expenditure of l I approximately $ 2.6 million for the high density racks is small compared to the estimated value of replacement power equivalent to the plant's energy output: i I approximately $ 21 million per month in 1994 and $ 32 million per month in 2011. The subject of the comparative economics associated with various spent fuel options is the subject of Chapter 6 of NUREG-0575 (Ref. 2). Although the material presented is generic, it is of value in comparing the costs of the various options. Of the options presented in that chapter, high density spent fuel storage at the site is the most economic option at $ 3.50 per KgU. The price of Independent Fuel Storage Facilities (IFSF), if available, would be $54.35 per Kgu. 9.4 RESOURCE COMMITMENTS The expansion of the Byron spent fuel storage capacity will require the following primary rescuces: . O Stainless steel - 284,815 lb/ unit 0 Boraflex neutron absorber - 22,000 lb/ unit of which 10,925 lb is boron carbide (BgC) powder I The requirement for stainless steel represents a small fraction of the total domestic production for 1986 (Ref. 3). Although the fraction of domestic production of B gC required for the fabrication.is somewhat higher than that for stainless steel, it I I I 9-3 I .
I' I' is unlikely that the commitment of B gC to this project will affect other alternatives. Experience has shown that the pro-duction of B gC is highly variable and depends on need but could : easily be expanded to accommodate additional domestic needs. I : I: I! i I l' I I I I I I. I E I 9 t+
L h REFERENCES TO SECTION 9
- 1. B. K. Grimes, "0T Position for Review and Acceptance of
- Spent Fuel Storages and Handling Applications," April 14, 1978.
e L 2. NUREG-0575, " Final . Environmental Impact Statement on Handling and Storage of Spent Light Water Power Reactor Fuel," Vols. 1-3, USNRC, August 1979. 7
- 3. " Mineral Facts and Problems," Bureau of Mines Bulletin 671, 1980.
l L r L . [ [ [ [ 9-5
F . l l 10.0 QUALITY ASSURANCE PROGRAM l
10.1 INTRODUCTION
This chapter provides a general description of the quality r assurance program that is implemented to assure that the quality objectives of the contract specification are met. L 10.2 GENERAL r b The quality assurance program used on this project is based upon [ the system described in Oat's Nuclear Quality Assurance Manual. This system is designed to provide a flexible but highly controlled system for the design, manufacture, and testing of customized components in accordance with various codes, specifications, and regulatory requirements. The Oat Nuclear Quality Assurance Program has been accepted by ASME and has been approved by the CECO Quality Assurance Department and placed on CECO's Qualified Suppliers List. [ 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 day-to-day basis. The system readily adapts to different r designs and component configurations, making possible the construction of many varied forms of equipment. The following paragraphs provide an overview of the system and how it has been applied to Commonwealth Edison's specifications. 10.3 SYSTEM HIGHLIGHTS The design control is organized to provide for careful review of { all contract requirements to extract each individual design and [ 10-1
I quality criterion. These criteria'are transl~ated into design and quality control documents customized to the contract requirement and completely reviewed and approved by responsible and qualified personnel. The system for control of purchased material includes 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 over-check specimens, special treatments or conditioning of material, source inspection, and provision of performance documentation on 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 traveller is prepared for each subassembly and assembly in each contract. The traveller is generated specifically to provide step-by-step inspection, testing, cleaning, instructions for fabrication, packaging, etc., which address all standard and special
- requirements of the contract specifications. Special attention g l
1s given to deployment of fabrication sequence and inspection 5 steps to preclude the possibility of missing poison sheets or incorrect sheets (incorrect B 10 loading). All nondestructive examination procedures and test procedures are custom written to apply to CECO's requirements. The system provides for qualification and written certification of personnel performing quality-related activities including l l nondestructive examination and fabrication inspection, welding, 5 engineering, production supervision, and auditing. 5 I E 10-2
r' .' L Other CECO requirements are fully covered in the Quality Assurance Program, including document control, control of measuring and test equipment, control of nonconforming material and parts, corrective action auditing, and other areas as specified by CECO. 10.4
SUMMARY
r Oat'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 s inspection of material and witnessing of material testing by the f engineer, furnishing of material certifications and test reports within 5 days of shipment, and obtaining verification of qualification testing of poison materials. F , [ [ [ [ 10-3
c'
- O b
L r L ( APPf:NDIX A BENCHMARK CALCULATIONS E E [ [ l [ [ l [ A-1 ii
I
- 1. INTRODUCTION AND
SUMMARY
The objective of this benchmarking study is to verify both the AMPX (NITAWL)-KENO (Refs. 1 and 2) methodology with the 27-group SCALE cross-section library (Refs. 3 and 4) and the CASMO-2E code (Refs. 5, 6, 7, and 8) for use in criticality calcula-tions of high density spent fuel storage racks. Both calcu-lational methods are based on transport theo'ry and have been benchmarked against critical experiments that simulate typical spent fuel storage rack designs as realistically as possible. Results of these benchmark calculations with both methodologies E are consistent with corresponding calculations reported in the 5 literature and with the requirements of Regulatory Guide 3.41,* Rev. 1, May 1977. Results of these benchmark calculations show that the 27-group (SCALE) AMPX-KENO calculations consistently underpredict the critical eigenvalue by 0.0106
- 0.0048 Ak (with a 95% proba-bility at a 95%. confidence level) for critical experiments l
selected to be representative of realistic spent fuel storage rack configurations and poison worths. Similar calculations by Westinghouse suggest a bias of 0.012
- 0.0023, and the results of ORNL analyses of 54 relatively " clean" critical experiments show a bias of 0.0100
- 0.0013.
l Similar calculations with CASMO-2E for clean critical I experiments resulted in a bias of 0.0013
- 0.0018 (95%/95%).
CASMO-2E and AMPX-KENO intercomparison calculations of infinite arrays of poisoned cell configurations show very good agreement and suggest that a bias .o f 0.0013 i 0.0018 is the reasonably expected bias and uncertainty for CASMO-2E calculations. I Validation of Calculational Methods for Nuclear Criticality Safety. (See also ANSI N16.9-1975.) A-2 I I
L The benchmark calculations reported here indicate that either the 27-group (SCALE) AMPX-KENO or CASMO-2E calculations are acceptable for criticality analysis of high density spent [ fuel storage racks. The preferred methodology, however, is to perform independent calculations with both code packages and to utilize the higher, more conservative value for the reference { design infinite multiplication factor. E L
- 2. AMPX (NITAWL)-KENO BENCHMARK CALCULATIONS Analysis of a series of Babcock & Wilcox (B&W) critical experiments (Ref. 9), which include some with absorber sheets typical of a poisoned spent fuel rack, is summarized in "'able 1, as calculated with AMPE-KENO using the 27-group SCALE cross-section library and the Nordheim resonance integral treatment in NITAWL. The mean for these calculations is 0.9894 i 0.0019, conservatively assuming the larger standard deviation calculated from the k,fg values. With .a one-sided tolerance factor (K = 2.502), corresponding to 95% probability at a 95% confidence level (Ref. 10), the calculational bias is +0.0106 with an uncer-tainty of *0.0048.
Similar calculational deviations reported by Westinghouse ( (Ref. 11) are also shown in Table 1 and suggest a bias of 0.012 i 0.0023 (95%/95%). In addition, ORNL (Ref. 12) has analyzed some 54 critical experiments using the same methodology, obtaining a { mean bias of 0.0100
- 0.0013 (95%/95%). These published results are in good agreement with the results obtained in the present analysis and lend further credence to the validity of the 27-group AMPX-KENO calculational model for use in criticality analy-sis of high density spent fuel storage racks. Variance analysis of the data in Table 1 suggests the possibility that an unknown
( factor may be causing a slightly larger variance than might be expected from the Monte Carlo statistics alone. However, such a m A-3
I Table 1 RESULTS OF 27-GROUP (SCALE) AMPX-KENO CALCULATIONS OF B&W CRITICAL EXPERIMENTS Westinghouse Experiment Calculated Calculated-meas. Number k,gg a kegg I 0.9889 *0.0049 -0.008 II 1.0040 *0.0037 -0.012 III 0.9985 *0.0046 -0.008 IXI1) 0.9924 iO.0046 -0.016 X 0.9907 *0.0039 -0.008 XI 0.9989 0.0044 +0.002 XII 0.9932 *0.0046 -0.013 XIII 0.9890 *0.0054 -0.007 XIV 0.9830 i0.0038 -0.013 XV 0.9852 i0.0044 -0.016 XVI 0.9875 i0.0042 -0.015 XVII 0.9811 i0,0041 -0.015 l XVIII 0.9784 *0.0050 -0.015 XIX 0.9888 i0.0033 -0.016 XX 0.9922 *0.0048 -0.011 XXI 0.9783 *0.0039 -0.017 Mean 0.9894 *0.0011(2) -0.0120 i 0.0010 Bias 0.0106 *0.0019(3) 0.0120
- 0.0010 Bias (95%/95%) 0.0106 *0.0048 0.0120 i 0.0023 Maximum Bias 0.0154 0.0143 (1) Experiments IV through VIII used B 4 C pin absorbers anc were (2)not considered representative of poisoned storage racks.
individual standard deviations. (3) Calculated Calculated from from k,gg values and used as reference. I. l A-4 I
L factor, if one truly exists, is too small to be resolved on the 7 L basis of critical-experiment data presently available. No trends in k,gg with intra-assembly water gap, with absorber sheet reactivity worth, or with soluble poison concentration were identified. E u .
- 3. CASMO-2E BENCHMARK CALCULATIONS 3.1 GENERAL
+ The CASMO-2E code is a multigroup transport theory code utilizing transmission probabilities to accomplish two-dimen-sional calculations of reactivity and depletion for BWR and PWR fuel assemblies. As such, CASMO-2E is well-suited to the criti- [ cality analysis of spent fuel storage racks, since practice is to treat the racks as an infinite medium of storage general cells, neglecting leakage effects. { CASMO-2E is closely analogous to the EPRI-CPM code (Ref. 13) and has been extensively benchmarked against hot and cold crit-ical experiments by Studsvik Energiteknik (Refs. 5, 6, 7, and l 8). Reported analyses of 26 critical experiments indicate a mean keff of 1.000 i 0.0037 (la). Yankee Atomic (Ref. 14) has also [ reported results of extensive benchmark calculations with CASMO-2E. Their analysis of 54 Strawbridge and Barry critical experi- { ments (Ref. 15) 0.9987 t 0.0009 using the reported buckling indicates a mean of (la), or a bias of 0.0013 i 0.0018 (with 95% probability at a 95% confidence level). Calculations were " repeated for seven of Strawbridge and Barry experiments the s ~ a Significantly large trends in keff with water gap and with ab-sorber sheet reactivity worth have been reported (Ref. 16) for AMPX-KENO calculations with the 123-group GAM-THERMOS library. A-5 i
1 l I . selected at random, yielding a mean ke gg of 0.9987
- 0.0021 (le),
. thereby confirming that the cross-section library and analytical f methodology being used for the present calculations are the same as those used in the Yankee analyses. Thus, the expected bias ;
for CASMO-2E in the analysis of " clean" critical experiments is 1 0.0013
- 0.0018 (95%/95%).
l 3.2 BENCHMARK CALCULATIONS CASMO-2E benchmark calculations have also been made for the B&W series of critical experiments with absorber sheets, simu-lating high density spent fuel storage racks. However, CASMO-2E, l I as an assembly code, cannot directly tepresent an entire core configuration
- without introducing uncertainty due to reflector constants and the appropriateness of their spectral weighting.
For this reason, the poisoned cell configurations of the central assembly, as calculated by CASMO-2E, were benchmarked against corresponding calculations with the 27-group (SCALE) AMPX-KEN E l code package. Results of this comparison are shown in Table 2. E Since the differences are well within the normal KENO statistical variation, these calculations confirm the validity of CASMO-2E calculations for the typical high density poisoned spent fuel rack configurations. The differences shown in Table 2 are also consistent with a bias of 0.0013 i 0.0018, determined in Section 3.1 as the expected bias and uncertainty of CASMO-2E calcula-tions. l
~
I l Yankee has attempted such calculations (Ref. 14) using CASMO-2E-I generated constants in a two-dimensional, four-group PDQ model, l obtaining a mean kegg of 1.005 for 11 poisoned cases and 1.009 l for 5 unpoisoned cases. Thus, Yankee benchmark calculations suggest that CASMO-2E tends to slightly overpredict reactivity. l l A-6 l I L
(; Table 2 RESULTS OF CASMO .IE BENCHMARK (INTERCOMPARISON) CALCULATIONS k,III B&W Experiment No.II) AMPX-KENO (2) CASMO-2E Ak { XIX- 1.1203
- 0.0032 1.1193 0.0010 XVII 1.1149
- 0.0039 1.1129 0.0020 XV 1.1059
- 0.0038. 1.1052 0.0007 Interpolated (3) 1.1024
- 0.0042 1.1011 0.0013 XIV 1.0983
- 0.0041 1.0979 0.0004
[ XIII l.0992
- 0.0034 1.0979 0.0013 Mean
- 0.0038 0.0011 Uncertainty *0.0006 BWR fuel rack
- 0.9212
- 0.0027 0.9218 -0.006
[ (1) Infinite array of central assemblies of 9-assembly B&W criti-(2) cal from configuration (Ref. 9). AMPX-KENO corrected ~ for bias of 0.0106 Ak. (3)kInterpolated from Fig. 28 of Ref. 9 for soluble boron concen-tration at critical condition. [ [ [ { A-7 e
I h REFERENCES TO APPENDIX A
- 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," g ORNL-TM-3706, Oak Ridge National Laboratory, March 1976. 3
- 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. R. M. Westfall et al., " SCALE: A Modular Code System for l Performing Standardized Computer Analyses f ,., r Licensing W Evaluation," NUREG/CR-0200, 1979.
- 4. W. E. Ford, III et al., "A 218-Neutron Group Master Cross-section Library for Criticality Safety Studies," ORNL/TM-4, 1976.
S. A. Ahlin, M. Edenius, H. Haggblom, "CASMO - A Fuel AE-RF-76-4158, Assembly report I Burnup Program," Studsvik (proprietary)..
- 6. A. Ahlin and M. Edenius, "CASMO - A Fast Transport Theory Depletion Code for LWR Analysis," ANS Transactions, Vol. 26, g
- p. 604, 1977. g
- 7. M. Edenius et al., "CASMO Benchmark Report," Studsvik/RF-78/6293, Aktiebolaget Atomenergi, March 1978.
- 8. "CASMO-2E Nuclear Fuel . Assembly Analysis, Application Users Manual," Rev. A, Control Data Corporation, 1982.
- 9. M. N. Baldwin et al., " Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel," BAW-1484-7, 3 The Babcock & Wilcox Company, July 1979. I
- 10. M. G. Natrella, Experimental Statistics,. National Bureau of Standards, Handbook 91, August 1963.
- 11. B. F. Cooney et al., " Comparisons of Experiments and Calculations for LWR Storage Geometries," Westinghouse NES, 3 ANS Transactions, Vol. 39, p. 531, November 1981. 5
- 12. R. M. Westfall and J. R. Knight, " Scale System Cross-section Vclidation with Shipping-cask Critical Experiments," ANS Transactions, Vol. 33, p. 368, November 1979.
- 13. "The EPRI-CPM Data Library," ARMP Comouter Code Manuals, l
=
Part II, Chapter 4, CCM3, Electric Power Research Institute, November 1975. I A-8
b ( l REFERENCES TO APPENDIX A (Continued) [ .
- 14. E. E. Pilat, " Methods for the Analysis of Boiling Water YAEC-1232, Yankee Atomic
{ Reactors (Lattice Electric Co., December 1980. Physics)," p 15. L. E. Strawbridge and R. F. Barry, " Criticality. Calculations L for Uniform, Water-moderated Lattices," Nuclear Science and ! Engineering, Vol. 23, p. 58, September 1965. [ 16. S. E. Turner and M. K. Gurley, " Evaluation of AMPX-KENO Benchmark Calculations for High Density Spent Fuel Storage
. Racks," Nuclear Science and Engineerina, 80(2): 230-237,
{ February 1982. [ [ r l 1 1. 1 l 1 l l 1
-}}