ML112380574

Licensing Rept for Spent Fuel Storage Capacity Expansion Duane Arnold Energy Center.

ML112380574
Person / Time
Site:	Duane Arnold
Issue date:	03/26/1993
From:	IES Utilities, (Formerly Iowa Electric Light & Power Co)
To:
Shared Package
ML112380575	List: ML112380574 ML20035C448 NG-93-0566, Application for Amend to License DPR-49,revising TSs to Allow Additional Storage Capacity of 323 Fuel Assemblies in Proposed Fuel Rack Temporarily Located in Cask Loading Area of Cask Pit
References
HI-92889, NUDOCS 9304070241
Download: ML112380574 (286)

Text

LICENSING REPORT FOR SPENT FUEL STORAGE CAPACITY EXPANSION DUANE ARNOLD ENERGY CENTER IOWA ELECTRIC LIGHT AND POWER COMPANY Holtec Project 20170 Holtec Report HI-92889 Report Category: A 9304070241 930326 PDR ADOCK 05000331 P PDR

TABLE OF CONTENTS SECTION PAGE

1.0 INTRODUCTION

1-1 2.0 MODULE LAYOUT FOR INCREASED STORAGE 2.1 Module Layout 2-1 3.0 RACK FABRICATION AND APPLICABLE CODES 3.1 Design Objective 3-1 3.2 Anatomy of the Rack Module 3-2 3.3 Materials of Construction 3-4 3.4 Codes, Standards and Practices for the 3-6 Spent Fuel Pool Modification 4.0 CRITICALITY SAFETY ANALYSIS 4.1 Introduction 4-1 4.2 Summary and Conclusions 4-2 4.3 Abnormal and Accident Conditions 4-4 4.4 Input Parameters 4-4 4.4.1 Fuel Assembly Design 4-4 Specifications 4.4.2 Storage Rack Cell 4-5 Specifications 4.5 Analysis Methodology 4-5 4.6 Criticality Analyses and Tolerance Variations 4-6 4.6.1 Nominal Design Case 4-6 4.6.2 Uncertainties due to 4-7 Manufacturing Tolerances 4.6.2.1 Boron Loading Variation 4-7 4.6.2.2 Boral Width Tolerance Variation 4-8 4.6.2.3 Storage Cell Lattice Pitch 4-8 Variation 4.6.2.4 Stainless Steel Thickness 4-8 Tolerances 4.6.2.5 Fuel Enrichment and Density 4-8 Variation 4.6.2.6 Zirconium Flow Channel 4-9 4.6.3 Uncertainty in Depletion 4-9 Calculations 4.7 Higher Enrichments and GE-11 Fuel 4-10 4.8 Abnormal and Accident Conditions 4-10 4.8.1 Temperature and Water Density Effects 4-10 4.8.2 Abnormal Locationof a Fuel Assembly 4-11 4.8.3 Eccentric Fuel Assembly Positioning 4-11 4.8.4 Zirconium Fuel Channel Distortion 4-11 4.8.5 Dropped Fuel Assembly 4-12 i

TABLE OF CONTENTS (continued) 4.8.6 Fuel Rack Lateral Movement 4-12 4.9 Comparison with Other Recently 4-12 Licensed US Plants 4.10 References for Section 4 4-13 Appendix A to Section 4 5.0 THERMAL-HYDRAULIC CONSIDERATIONS 5.1 Introduction 5-1 5.2 Spent Fuel Cooling and Cleanup System 5-2 Description 5.3 Decay Heat Load Calculations 5-3 5.4 Discharge Scenarios 5-3 5.5 Bulk Pool Temperatures 5-5 5.6 Local Pool Water Temperature 5-9 5.6.1 Basis 5-9 5.6.2 Model Description 5-10 5.7 Cladding Temperature 5-11 5.8 Results 5-13 5.9 References for Section 5 5-14 6.0 STRUCTURAL/SEISMIC CONSIDERATIONS 6.1 Introduction 6-1 6.2 Background and Analysis Outline 6-1 6.3 Articifial Time-Histories 6-6 6.4 Rack Modeling for Dynamic Simulations 6-9 6.4.1 General Remarks 6-9 6.4.2 The 3-D 22 DOF Model for Single Rack Analysis 6-11 6.4.2.1 Assumptions 6-11 6.4.2.2 Model Details 6-13 6.4.2.3 Fluid Coupling Details 6-13 6.4.2.4 Stiffness Element Details 6-15 6.4.3 Whole Pool Multi-Rack (WPMR)

Model 6-16 6.4.3.1 General Remarks 6-16 6.4.3.2 Whole Pool Fluid Coupling 6-17 6.4.3.3 Coefficients of Friction 6-17 6.4.3.4 Modeling Details 6-18 6.5 Acceptance Criteria, Stress Limits 6-19 and Material Properties 6.5.1 Acceptance Criteria 6-19 6.5.2 Stress Limits for Various 6-21 Conditions 6.5.2.1 Normal and Upset Conditions 6-21 (Level A or Level B) 6.5.2.2 Level D Service Limits 6-23 6.5.2.3 Dimensionless Stress Factors 6-23 ii

TABLE OF CONTENTS (continued) 6.5.3 Material Properties 6-24 6.6 Governing Equations of Motion 6-24 6.7 Results of 3-D Nonlinear Analyses 6-26 of Single Racks 6.7.1 Racks in the Fuel Pool 6-27 6.7.1.1 Impact Analyses 6-28 6.7.1.2 Weld Stresses 6-29 6.7.2 Rack in the Cask Pit Area 6-30 6.8 Results from Whole Pool Multi-Rack 6-30 (WPMR) Analyses 6.9 Bearing Pad Analysis 6-32 6.10 References for Section 6 6-33 7.0 ACCIDENT ANALYSIS AND MISCELLANEOUS STRUCTURAL EVALUATION 7.1 Accident Analysis and Miscellaneous 7-1 Structural Evaluations 7.2 Refueling Accidents 7-1 7.2.1 Dropped Fuel Assembly 7-1 7.3 Local Buckling of Fuel Cell Walls 7-2 7.4 Analysis of Welded Joints in Rack 7-3 due to Isolated Hot Cell 7.5 References for Section 7 7-4 8.0 FUEL POOL STRUCTURE INTEGRITY CONSIDERATIONS 8.1 Introduction 8-1 8.2 General Features of the Model 8-3 8.3 Factored Loadings for the Fuel 8-5 Pool Structure 8.4 Finite Element Analyses 8-6 8.4.1 Load Cases 8-6 8.4.2 Applied Loads 8-8 8.4.3 Postprocessing of Finite 8-9 Element Load Cases 8.5 Results of Bending Evaluation 8-10 8.6 Results from Shear Evaluation on Floor 8-10 Slab Using Bounding Loads 8.7 References for Section 8 8-10 9.0 RADIOLOGICAL EVALUATION 9.1 Fuel Handling Accident 9-1 9.1.1 Assumptions and Source Term Calculations 9-1 9.1.2 Results 9-4 9.2 Solid Radwaste 9-4 9.3 Gaseous Releases 9-4 iii

TABLE OF CONTENTS (continued) 9.4 Personnel Exposures 9-5 9.5 Anticipated Exposure During Re-racking 9-6 10.0 BORAL SURVEILLANCE PROGRAM 10.1 Purpose 10-1 10.2 Coupon Surveillance Program 10-2 10.2.1 Coupon Description 10-2 10.2.2 Surveillance Coupon 10-3 Testing Schedule 10.2.3 Measurement Program 10-4 10.3 References for Section 10 10-6 11.0 ENVIRONMENTAL COST-BENEFIT ASSESSMENT 11.1 Introduction 11-1 11.2 Imperative for Reracking 11-1 11.3 Appraisal of Alternative Options 11-1 11.4 Cost Estimate 11-7 11.5 Resource Commitment 11-8 11.6 Environmental Considerations 11-8 11.7 References for Section 11 11-9 iv

LIST OF TABLES 1.1 Past and Projected Discharge Schedule 2.1.1 Design Data for New Racks 2.1.2 Existing Licensed Racks in the Duane Arnold Fuel Pool 2.1.3 Reracked Pool Configuration After 1992-94 Campaign (Campaign I) 2.1.4 Raracked Pool Configuration After the Intermediate Campaign (Campaign II) 2.1.5 Module Data After Final Rerack 2.2.1 Heavy Load Handling Compliance Matrix (NUREG-0612) 3.1 BoralTm Experience List (Domestic and Foreign) 3.2 Boron Carbide Chemical Composition, Weight %

Boron Carbide Physical Properties 4.2.1 Summary of Criticality Safety Analyses 4.3.1 Reactivity Effects of Abnormal and Accident Conditions 4.4.1 Fuel Assembly Design Specifications 4.6.1 Reactivity Uncertainties due to Manufacturing Tolerances 4.8.1 Effect of Temperature and Void on Calculated Reactivity of Storage Rack 5.4.1 DAEC Existing and Projected Fuel Discharge Schedule 5.4.2 Data for Discharge Scenarios 5.6.1 Data for Local Temperature Analysis 5.7.1 Peaking Factors 5.8.1 Major Design Input 5.8.2 SFP Bulk Pool Temperature 5.8.3 Results of Loss-of-Cooling 5.8.4 Maximum Local Pool Water and Fuel Cladding Temperature for the Limiting Case (Case 3) 6.1.1 Listing of Plants Where DYNARACK Was Applied V

6.3.1 Cross-Correlation Coefficients 6.4.1 Degrees-of-Freedom 6.4.2 Numbering System for Gap Elements and Friction Elements 6.4.3 Spent Fuel Pool Loading 6.5.1 Rack Material Data (200 0 F) and Support Material Data 6.7.1 Results of Single Rack Analyses 6.7.2 Summary of Worst Results From 18 Runs of Single Rack Analysis 6.7.3 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.4 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.5 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.6 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.7 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.8 Summary Results of 3-D Single Rack Analysis for Rack Module: G-11x21 6.7.9 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 6.7.10 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 6.7.11 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 6.7.12 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 6.7.13 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 6.7.14 Summary Results of 3-D Single Rack Analysis for Rack Module: J-14x21 vi

6.7.15 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.16 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.17 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.18 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.19 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.20 Summary Results of 3-D Single Rack Analysis for Rack Module: R-17x19 6.7.21 Comparison of Calculated and Allowable Loads/Stresses at Impact Locations and at Welds 6.8.1 Maximum Displacements from Whole Pool Multi-Rack Runs 6.8.2 Maximum Impact Force of Each Gap Spring 6.8.3 Maximum Pedestal Stress Factors of All Racks in Pool 6.8.4 Maximum Rack Displacements, Pedestal Loads and Pedestal Stress Factor in Single Rack Analyses and in WPMR Analyses (DBE Seismic; Reg. Fuel, Fully Loaded) 6.8.5 Results of Pool Wall Dynamic Pressures 6.8.6 Total Static Load and Dynamic Adder on Whole Slab 6.9.1 Average Bearing Pad Pressure - Comparison of Calculated and Allowable Stresses 8.1.1 Buoyant Weight of Existing Racks + Fuel Compared to New Racks + Fuel (680 lb. Fuel, All Cells Loaded) 8.1.2 Total Dead Load Over Pool Slab (Existing vs. New) (40' of Water in Pool) 8.5.1 Bounding Safety Margins-Bending Moments in Duane Arnold Pool - Maximum Bending Moment in Concrete 9.1 Results of Origen-2 Calculations for Radionuclides of Iodine, Krypton, and Xenon at 24-Hours Cooling Time 9.2 Radionuclide Properties Used in the Fuel Handling Accident Analysis Vii

9.3 Data and Assumptions for the Evaluation of the Fuel Handling Accident 9.4 Typical Concentrations of Raionuclides in Spent Fuel Pool Water 9.5 Preliminary Estimate of Person-Rem Exposures During Re-Racking 10.1 Coupon Measurement Schedule Viii

LIST OF FIGURES 2.1.1 Pool Layout - Existing Racks 2.1.2 Pool Layout - Campaign I 2.1.3 Pool Layout - Campaign II 2.1.4 Pool Layout - Campaign III 2.1.5 Pictorial View of DAEC Fuel Rack Module 3.1 Seam Welding Precision Formed Channels 3.2 Sheathing Shown Installed on the Box 3.3 A Cross Sectional View of an Array of Storage Locations 3.4 Three Cells in Elevation View 3.5 Adjustable Support 4.2.1 Change in Rack Reactivity with 4% Enriched Fuel 4.2.2 K-Infinite Limits in the Standard Core Geometry 4.4.1 Cross-Section of Typical Storage Cell (Calculational Model) 4.6.2 Correlation Between Reactivities in the Core and in the Rack 4.7.1 Correlation Between Reactivities in the Core and in the Rack 5.4.1 DAEC Discharge Scenario Case 1 5.4.2 Discharge Scenario Case 2 5.4.3 Discharge Scenario Case 3 5.4.4 Discharge Scenario Case 4 5.5.1 Evaporation Heat Loss 5.6.1 Idealization of Rack Assembly 5.6.2 Thermal Chimney Flow Model 5.6.3 Convection Currents in the Pool ix

5.8.1 Time After Reactor Shutdown, Hrs. - DAEC, SRP Normal Discharge, One Train - Case 1 5.8.2 Time After Reactor Shutdown, Hrs. - DAEC, SRP Abnormal Discharge, Two Trains - Case 2 5.8.3 Time After Reactor Shutdown, Hrs. - DAEC, SRP EOC Full Core Offload, Two Trains - Case 3 5.8.4 Time After Reactor Shutdown, Hrs. - DAEC, BOC Full Core Offload, Two Trains - Case 4 5.8.5 Time After Reactor Shutdown, Hrs. - DAEC, SRP Normal Discharge, One Train - Case 1 5.8.6 Time After Reactor Shutdown, Hrs. - DAEC, SRP Abnormal Discharge, Two Trains - Case 2 5.8.7 Time After Reactor Shutdown, Hrs. - DAEC, EOC Full Core Offload, Two Trains - Case 3 5.8.8 Time After Reactor Shutdown, Hrs. - DAEC, BOC Full Core Offload, Two Trains - Case 4 5.8.9 Time After Reactor Shutdown, Hrs. - DAEC, SFP Loss of Cooling - Case 1 (SRP Normal Discharge, One Train) 5.8.10 Time After Reactor Shutdown, Hrs. - DAEC, SFP Loss of Cooling - Case 2 (SRP Full Core Offload, Two Trains) 5.8.11 Time After Reactor Shutdown, Hrs. - DAEC, SFP Loss of Cooling - Case 3 (EOC Full Core Offload, Two Trains) 5.8.12 Time After Reactor Shutdown, Hrs. - DAEC, SFP Loss of Cooling - Case 4 (BOC Full Core Offload, Two Trains) 6.3.1 Horizontal Response Spectrum for DAEC 6.3.2 Vertical Response Spectrum for DAEC 6.3.3 Seismic Time History, OBE, N-S 6.3.4 Seismic Time History, OBE, E-W 6.3.5 Seismic Time History, OBE, VT 6.3.6 Seismic Spectrum: OBE, N-S, Damping 1%

6.3.7 Seismic Spectrum: OBE, E-W, Damping 1%

6.3.8 Seismic Spectrum: OBE, VT, Damping 1%

6.4.1 Pictorial View of Rack Structure x

6.4.2 Schematic Model for DYNARACK 6.4.3 Rack-to-Rack Impact Springs 6.4.4 Fuel-to-Rack Impact Springs 6.4.5 Degrees-of-Freedom Modeling Rack Motion 6.4.6 Rack Degree-of-Freedom for Y-Z Plane Bending 6.4.7 Rack Degree-of-Freedom for X-Z Plane Bending 6.4.8 2-D View of Rack Module 6.4.9 Pool Layout - Campaign III 6.8.1 Gap Time History, Gap Between Rack-N1 and Rack N2, West Corner, Top, DBE, Fully Loaded with 680# Reg. Fuel Assemblies 6.8.2 Gap Time History, Gap Between Rack N1 and North Wall, West Corner, Top, DBE, Fully Loaded with 680# Reg. Fuel Assemblies 6.8.3 Gap Time History, Gap Between Rack F and Rack E, North Corner, Top, DBE, Fully Loaded with 680# Reg. Fuel Assemblies 6.8.4 Gap Time History, Gap Between Rack E and Rack D, North Corner, Top, DBE, Fully Loaded with 680# Reg. Fuel Assemblies 6.8.5 Gap Time History, Gap Between Rack D and East Wall, North Corner, Top, DBE, Fully Loaded with 680# Reg. Fuel Assemblies 6.8.6 Total Slab Load Time-History, DBE Seismic, Fully Loaded with 680# Reg. Fuel Assemblies 7.4.1 Welded Joint in Rack 8.3.1 Duane Arnold Fuel Pool, Mode 1, N-S Direction 8.3.2 Duane Arnold Fuel Pool, Mode 2 8.3.3 Duane Arnold Fuel Pool, Mode 3 xi

1.0 INTRODUCTION

The Duane Arnold Energy Center (DAEC) is a Boiling Water Reactor installation (1658 MWt thermal output) supplied by the General Electric Company of San Jose, California. The DAEC reactor features a core consisting of 368 fuel assemblies. The plant is located on a site near Palo in Linn County, Iowa, and consists of approximately 500 acres adjacent to the Cedar River. The site is jointly owned by the Iowa Electric Light and Power Company, the Central Iowa Power Cooperative and the Cornbelt Power Cooperative.

The Iowa Electric Light and Power Company (IELP) is the principal owner and operator of the plant, henceforth also referred to as the Owner or the Licensee.

DAEC achieved initial criticality on March 23, 1974 and began commercial operation on February 1, 1975. The DAEC spent fuel pool is of 240" x 480" (nominal) rectangular planform section. Although the present licensed capacity of the DAEC pool is 2050 storage cells, the pool is currently equipped with 1898 accessible storage cells in 19 rack modules which were installed in the previous rerack campaign in 1979.

All existing racks in the DAEC pool are of bolted anodized aluminum construction with the Boral neutron absorber providing neutron attenuation. The neutron absorber, Boral, is sealed within two concentric square aluminum tubes forming the "poison can". Large levelling screws are located at the rack corners to adjust for variations in the pool liner surface. The racks are of the free standing design. Each rack is supported by four floor bearing pads which are attached to the rack corner levelling screws. The bottom of the racks are 7.25" to 8.25" above the floor in order to clear the interferences protruding from the liner and provide coolant flow under the racks.

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This licensing application is for reracking the DAEC pool with new maximum density racks. The reracking of the DAEC pool is planned in three campaigns over the next 15 years, with the first rack change-out scheduled to be carried out in early 1994. The first rerack campaign will increase the installed capacity in the DAEC pool to 2411 cells extending the projected end-of-full-core offload capacity date from ca. 1998 to the year 2005. The subsequent rerack campaigns will be undertaken only if the U.S. Department of Energy fails to assume custody of the spent nuclear fuel as envisaged by federal law. At the present time, two subsequent rerack campaigns, in approximately 2003 and 2005, are projected. However, depending on the changes in the fuel cycle durations and the permanent repository situation, the number of campaigns may be increased or diminished. Regardless of the number of rerack campaigns, ultimate storage capacity in the DAEC fuel pool will rise to 2829 cells.

Further, a 323 cell rack can be installed in the cask pit for temporary off loading of the reactor core. This rack will be needed after the inventory in the spent fuel pool begins to reach the ultimate capacity. The so-called cask pit rack will provide a temporary repository of the fuel in the core at that time. Should the cask pit rack be installed, several additional controls will be implemented. The drain from the cask pit will be sealed, the gate between the cask pit and SFP will be kept out to assure cooling water flow into the cask pit and no heavy loads will be allowed over the cask pit. These restrictions will be in place for as long as fuel is in the cask pit rack.

The pool configurations after the three projected campaigns are illustrated in Section 2 of this Licensing Report. Referring to Table 1.1, the end-of-full-core discharge capability dates for each pool configuration are as follows:

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Projected Date of End-of-Full Projected Date of Pool Core Reserve End-of-Normal Configuration Capacity Discharge Capacity (a) Existing pool (1898 1998 2001 accessible cells)

(b) After Campaign I 2005 2007 (2411 cells)

(c) After Campaign II 2007 2010 (2589 cells)

(d) After Campaign III 2014 End of Plant and Cask Pit Rack License Life (3152 cells)

It is concluded from the foregoing table that the proposed reracking scheme for the DAEC pool affords in-pool storage with full-core-offload-reserve almost to the end of the presently licensed life-of-the-reactor (ca. 2014). The Owner of DAEC has contracted with Holtec International of Cherry Hill, New Jersey to execute the present (Campaign I) rerack on a turnkey basis. IELP has also worked out a long-term turnkey arrangement with Holtec International to carry out the future rerack campaigns, if deemed necessary by the circumstances. The first reracking campaign is proposed to be carried out in early 1994.

The new spent fuel storage racks, like the existing ones, are free standing and self supporting. The principal construction materials for the new racks are ASME 240-Type 304 stainless steel sheet and plate stock, and SA564 (precipitation hardened stainless steel) for the adjustable support spindles. The only non-stainless material utilized in the rack is the neutron absorber material which is a boron carbide aluminum cermet manufactured under a U.S. patent and sold under the brand name BoralTm by AAR Advanced Structures, Livonia, Michigan.

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The new racks are designed and analyzed in accordance with Section III, Division 1, Subsection NF of the ASME Boiler and Pressure Vessel Code. The material procurement and fabrication of the rack modules conforms to 10CFR50 Appendix B requirements.

This Licensing Report documents the design and analyses performed to demonstrate that the new spent fuel racks satisfy all governing requirements of the applicable codes and standards, in particular, "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", USNRC (1978) and 1979 Addendum thereto.

Safety analyses were performed for the cumulative storage capacity of 3152 cells, encompassing the present as well as all future campaigns.

The safety assessment of the proposed rack modules involved demonstration of its hydrothermal, criticality and structural adequacy. Hydrothermal adequacy requires that fuel cladding will not fail due to excessive thermal stress, and that the steady state pool bulk temperature will remain within the limits prescribed for the spent fuel pool. Demonstration of structural adequacy primarily involves analyses showing that the free-standing modules will not impact in the cellular region under the postulated seismic events, and that the primary stresses in the module structure will remain below the ASME Code allowables. The structural qualification also includes analytical demonstration that the subcriticality of the stored fuel will be maintained under accident scenarios such as fuel assembly drop, etc.

The criticality safety analysis presented in Section 4 of this report shows that the neutron multiplication factor for the stored fuel array is bounded by the USNRC limit of 0.95 under assumptions of 95% probability and 95% confidence. Consequences of the inadvertent placement of a fuel assembly are also evaluated as part of the criticality analysis. The criticality analysis also sets 1-4

the requirements on the length of the B-10 panel and the areal density of the B-10 isotope.

This Licensing Report contains documentation of the analyses performed to demonstrate the large margins of safety with respect to all regulatory criteria pertaining to the reracking of spent fuel pools.

The analyses presented herein clearly demonstrate that the rack module arrays possess wide margins of safety from all four vantage points: thermal-hydraulic, criticality, structural and radiological. The No Significant Hazard Consideration evaluation presented in a companion document to this report is based on the descriptions and analyses synopsized in the subsequent sections of this report.

This document has been prepared for submission to the U.S. Nuclear Regulatory Commission for securing regulatory approval of the modification of the DAEC pool as proposed herein.

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Table 1.1 PAST AND PROJECTED DISCHARGE SCHEDULE Number of Cumulative Cycle Date of Assemblies Assemblies Number Discharge Discharged Discharged 1A 6/1975 4 4 lB 2/1976 84 88 2 3/1977 100 188 3 3/1978 88 276 4 2/1980 88 364 5 3/1981 84 448 6 2/1983 128 576 7 2/1985 120 696 8 3/1987 128 824 9 9/1988 120 944 10 6/1990 104 1048 11 2/1992 104 1152 12 7/1993 128 1280 13 2/1995 128 1408 14 9/1996 116 1524 15 3/1998 116 1640 16 9/1999 116 1756 17 3/2001 116 1872 18 9/2002 116 1988 19 3/2004 116 2014 20 9/2005 116 2220 21 3/2007 116 2336 22 9/2008 116 2452 23 3/2010 116 2568 24 9/2011 116 2684 25 3/2013 96 2780 26 2/2014 368 3148 (End of Plant License Life) 1-6

2.0 MODULE LAYOUT FOR INCREASED STORAGE 2.1 Module Layout This section provides general information on the new storage modules for the DAEC spent fuel pool. The information presented in this and the next section provide the basis for the detailed criticality, thermal-hydraulic and seismic analyses presented in the subsequent sections of this report.

The DAEC high density spent fuel storage racks consist of individual cells with 5.90" (nom.) inside square dimension, each of which accommodates a single Boiling Water Reactor (BWR) fuel assembly. The fuel assembly can be stored in the storage locations in channelled or unchannelled configuration. Table 2.1.1 gives the essential storage cell design data.

The existing licensed module layout in the DAEC pool is illustrated in Figure 2.1.1. The cell count for each of the 21 modules shown in Figure 2.1.1 is presented in Table 2.1.2. The pool configuration after the first rerack campaign (1992-94) is shown in Figure 2.1.2. As noted in Table 2.1.3, the cumulative storage capacity in the DAEC pool after the present rerack campaign will reach 2411 cells. Furthermore, 96 cells in the northeast corner of the pool (Module Al in Figure 2.1.2) are convertible into 24 square prismatic cells of approximately 12"x12" opening which can be used to store miscellaneous objects of larger cross-section such as the defective fuel container, control rods, etc. Modules A2, A3, N1 and N2 are also designed to support a specially engineered overhead platform which permits storage of miscellaneous objects up to five tons total weight without interfering with the normal function of the module as an assemblage of spent fuel storage cavities. The structural and thermal-hydraulic qualification of these racks includes the appropriate consideration of the overhead platform.

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Figures 2.1.3 and 2.1.4, along with their associated Tables 2.1.4 and 2.1.5, provide the storage information on the DAEC pool for the tentatively planned rerackings in the next century. The third campaign also includes the so-called "cask pit rack" which, as mentioned in Section 1, is merely a means to retain full-core offload capability after such a capacity is exhausted in the spent fuel pool itself. The cask pit rack is identical in its anatomical details to the racks in the spent fuel pool and as such, all analyses (seismic, thermal-hydraulic and criticality) have been carried out for the cask pit rack in an identical manner to the fuel pool racks.

The new modules for the DAEC fuel pool are qualified as quasi impacting free-standing racks, i.e., each module is free-standing and is shown to undergo minimal kinematic displacements during the postulated seismic events. Thus, rack-to-rack impacts are limited to the baseplate region. Impact between the racks in the cellular region is not permissible.

The rack module support legs are of remotely adjustable type.

Figure 2.1.5 shows a typical new rack module for the DAEC fuel pool.

IELP has developed a "defense-in-depth" approach to execute the DAEC fuel pool reracking which places a strong emphasis on equipment redundancy, personnel training and proceduralized execution.

A remotely engagable lift rig, meeting NUREG-0612 stress criteria, will be used to lift the empty modules. The rig designed for handling the DAEC racks is identical in its physical attributes to the rigs utilized to rerack Millstone Point Unit One (1988), Vogtle Unit Two (1989), Indian Point Unit Two (1990), Ulchin Unit Two (1990), Hope Creek (1990), Zion (1993), Laguna Verde Unit One (1990), La Salle Unit One (1993), and Kuosheng (1991). The rig 2-2

consists of independently loaded lift rods with a "cam type" lift configuration which ensures that failure of one traction rod will not result in uncontrolled lowering of the load being carried by the rig (which complies with the duality feature called for in Section 5.1.6(1) of NUREG 0612). The rig has the following additional attributes:

a. The stresses in the lift rods are self limiting inasmuch as an increase in the magnitude of the load reduces the eccentricity between the upward force and downward reaction (moment arm).

b. It is impossible for a traction rod to lose its engagement with the rig in locked position. Moreover, the locked configuration can be directly verified from above the pool water without the aid of an underwater camera.

c. The stress analysis of the rig is carried out using a finite element code, and the primary stress limits postulated in ANSI 14.6 (1978) are shown to be met.

d. The rig is load tested with 150% of the maximum weight to be lifted. The test weight is maintained in the air for one hour. All critical weld joints are liquid penetrant examined, after the load test, to establish the soundness of all critical joints.

The DAEC Reactor Building crane will be used for the reracking operation. The installation procedures call for all modules to be empty while being handled.

The Reactor Building crane meets the requirements of NUREG-0554 and NUREG-0612. The crane trolley is a single-failure-proof trolley with a 100-ton load capacity which has undergone a 125-ton in-place static load test.

Pursuant to the defense-in-depth approach of NUREG-0612, the following additional measures of safety will be undertaken for the reracking operation.

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(i) The crane and hoist will be given a preventive maintenance checkup and inspection prior to reracking and in accordance with station procedures.

(ii) Safe load paths have been developed for moving the old and new racks in the Reactor Building. The "old" or "new" racks will not be carried over any region of the pool containing fuel.

(iii) The rack upending or laying down will be carried out in an area which is not overlapping to any safety related component.

(iv) Crew members involved in the rigging of all heavy loads associated with the DAEC rerack project shall be trained in proper rigging techniques as well as safe travel path requirements for the loads.

Lifting and upending of the new racks will be done in accordance with Holtec International's design requirements to prevent potential damage during handling. All training of personnel shall be documented.

(v) All heavy loads will be lifted in such a manner that the center of the lift points is aligned with the center of gravity of the load being lifted.

(vi) Turnbuckles are utilized to "fine tune" the verticality of the rack being lifted.

In addition to the above design, testing and operation measures, IELP has also considered the consequences of a postulated rack drop on the integrity of the pool structure. The following analyses were performed:

a. The heaviest rack module (out of all existing and new racks) was postulated to free fall from the top of the water surface level to the pool floor.

b. The fall of a rack is assumed to occur in its normal vertical configuration which minimizes the retarding effect of water drag.

c. The falling rack is assumed to impact the pool slab undergoing an elastic/plastic impact.

d. The maximum impact load is compared with the gross seismic slab impact load during the SSE event (presented in Section 6 of this report).

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Analyses show that the maximum load due to the rack drop event is well below the cumulative impact load produced during the seismic event. Thus, the pool structure integrity analyses performed in support of this submittal and documented in Section 8 of this report bound the rack drop scenario.

The "old" racks will be "hydrolased" while underwater in the pool, and approved for shipping per the requirements of 10 CFR Part 71 and 49 CFR Part 171-178 before being brought to the Reactor Building Trackway. They will be housed in special shipping containers, and transported to a processing facility for volume reduction. Non-decontaminatable portions of the racks will be handled in accordance with 10CFR Part 61.

All phases of the reracking activity will be conducted in accordance with written procedures which will be reviewed and approved by IELP.

Our proposed compliance with the objectives of NUREG-0612 follows the guidelines contained in section 5 of that document. The guidelines of NUREG-0612 call for measures to "provide an adequate defense-in-depth for handling of heavy loads near spent fuel...".

The NUREG-0612 guidelines cite four major causes of load handling accidents, namely, (i) operator errors (ii) rigging failure (iii) lack of adequate inspection (iv) inadequate procedures The DAEC rerack program ensures maximum emphasis to mitigate the potential load drop accidents by implementing measures to eliminate shortcomings in all aspects of the operation including the four aforementioned areas. A summary of the measures specifically planned to deal with the major causes is provided below.

2-5

Operator errors: As mentioned above, IELP plans to provide comprehensive training to the installation crew for the rerack project.

Riging failure: The lifting device designed for handling and installation of the racks in the DAEC fuel pool has redundancies in the lift legs, and lift eyes such that there are four independent load paths. Failure of any one load bearing member would not lead to uncontrolled lowering of the load. The rig complies with all provisions of ANSI 14.6 - 1978, including compliance with the primary stress criteria, load testing at 150% of maximum lift load, and dye examination of critical welds.

The DAEC rig design is similar to the rigs used in the rerack of numerous other plants, such as Hope Creek, Millstone Unit 1, Indian Point Unit Two, Ulchin II, and Laguna Verde, among others.

Lack of adequate inspection: The designer of the racks will develop a set of inspection points which have thus far proven to have eliminated any incidence of rework or erroneous installation in numerous prior rerack projects.

Inadequate procedures: IELP plans procedures to cover the entire gamut of operations pertaining to the rerack effort, such as mobilization, rack handling, upending, lifting, installation, verticality, alignment, dummy gage testing, site safety, and ALARA compliance.

The series of procedures planned for the DAEC rerack are the successor of the procedures implemented successfully in other projects in the past.

In addition to the above, a complete inspection of the fuel handling crane and relubrication of its moving parts in accordance with station procedures prior to reracking operations is planned.

Safe load paths have been developed as required by NUREG-0612.

Table 2.2.1 provides a synopsis of the requirements delineated in NUREG-0612, and our intended compliance thereto.

All reracking operations will be carried out with foremost consideration of ALARA. Diving operations will be minimized unless the dose associated with a particular activity is judged to be minimized through a diving operation. All diving activities will comply with Reg. Guide (draft) DG-8006.

2-6

In summary, the measures implemented in DAEC reracking are similar to those utilized in the most recent successful rerack projects (such as Indian Point Unit 2, concluded in October, 1990; Hope Creek concluded in March, 1992; and TMI Unit One, concluded in September, 1992).

2-7

Table 2.1.1 DESIGN DATA FOR NEW RACKS I.D. 5.90 inch (nom.)

(inside dimension)

Cell Nominal Pitch 6.06 inch Boral Loading (nom.) .015 gm per sq.cm. (B-10)

Boral plate (nom.) width: 5 inch Boral picture frame (bounding) size 5.125" x 144-1/2" Boral length: 144 inches Cell height: 169 inch Baseplate thickness: 3/4 inch Bottom plenum height: 6.75" (nominal)

Number of supports per module: Four (minimum)

Support Type: Remotely adjustable 2-8

Table 2.1.2 EXISTING LICENSED RACKS IN THE DUANE ARNOLD FUEL POOL Cell Count Rack I.D. mxn = N Number of Cells El 8x11 88 E2 8xl 88 E3 8x8 64 E4 llxll 121 E5 1lx11 121 E6* 11x8 88 E7 1lxl0 110 E8 llxll 121 E9 11x8 88 E10 11x10 110 El llxll 121 E12 11x8 88 E13 8x10 80 E14 8x11 88 E15 8x11 88 E16 8x10 80 E17 10x11 110 E18 10x11 110 E19 lOx1l 110 E20 8x11 88 E21 8x11 88 Total Number of Cells: 2050 Modules E3 and E6 are currently not installed in the DAEC pool, although the pool is licensed to incorporate them.

2-9

Table 2.1.3 RERACKED POOL CONFIGURATION AFTER 1992-94 CAMPAIGN (CAMPAIGN I)

Total Cell Rack Existing (E) Cell Count Number of Count for I.D. or New (N) mxn = N Modules This Rack Type Al (N) 12x12=144 1 144 A2 (N) 12x12=144 1 144 A3 (N) 12x12=144 1 144 Bl (N) 10x12=120 1 120 B2 (N) 10x12=120 1 120 C (N) 12x12-10xl=134 1 134 D (N) 12x14=168 1 168 E (N) 10x14-7x4=112 1 112 F (N) 12x14-1x4=164 1 164 E10 (E) 11x10=l10 1 110 Ell (E) llxll=121 1 121 E12 (E) 11x8=88 1 88 E13 (E) 8x10=80 1 80 E14 (E) 8x11=88 1 88 E15 (E) 8x11=88 1 88 E16 (E) 8x10=80 1 80 E17 (E) 1Ox11=110 1 110 E18 (E) 1Ox11=110 1 110 E19 (E) 1Ox11=110 1 110 E20 (E) 8x11=88 1 88 E21 (E) 8x11=88 1 88 Total Cell Count for This Rack Type: 2411 2-10

Table 2.1.4 RERACKED POOL CONFIGURATION AFTER THE INTERMEDIATE CAMPAIGN (CAMPAIGN II)

Total Cell Rack Existing (E) Cell Count Number of Count for I.D. or New (N) mxn = N Modules This Rack Type Al (N) 12x12=144 1 144 A2 (N) 12x12=144 1 144 A3 (N) 12x12=144 1 144 Bl (N) 10x12=120 1 120 B2 (N) 10x12=120 1 120 C (N) 12x12-10xl=134 1 134 D (N) 12x14=168 1 168 E (N) 10x14-7x4=112 1 112 F (N) 12x14-1x4=164 1 164 G (N) 21x11=231 1 231 H (N) 21x12=252 1 252 J (N) 21x14-6x4=270 1 270 E16 (E) 8x10=80 1 80 E17 (E) 1Ox11=110 1 110 E18 (E) 1Ox11=110 1 110 E19 (E) 10x11=110 1 110 E20 (E) 8x11=88 1 88 E21 (E) 8x11=88 1 88 Total Cell Count for This Type Rack: 2589 2-11

Table 2.1.5 MODULE DATA AFTER FINAL RERACK Rack Number Size (Horizontal Shipping Weight I.D. of Cells Cross-Section) (in nearest KIP)

Al 144 72.96" x 72.96" 10300#

A2 144 72.96" x 72.96" 10300#

A3 144 72.96" x 72.96" 10300#

B1 120 72.96" x 60.8" 8600#

B2 120 72.96" x 60.8" 8600#

C 134 72.96" x 72.96" 9600#

D 168 85.12" x 72.96" 14000#

E 112 85.12" x 60.8" 8000#

F 164 85.12" x 72.96" 11700#

I: Total: 1250 Cells G 231 66.88" x 127.7" 16500#

H 252 72.96" x 127.7" 18000#

J 270 85.12" x 127.7" 19300#

II: Total: 753 Cells R

(cask 323 115.52" x 103.36" 23200#

pit)

K 110 66.88" x 60.8" 7900#

L 120 72.96" x 60.8" 8600#

M 140 85.12" x 60.8" 10000#

N1 144 72.96" x 72.96" 10300#

N2 144 72.96" x 72.96" 10300#

P 168 85.12" x 72.96" 12000#

III: Total: 1149 Cells Grand Total: 19 Racks - 3152 Cells 2-12

Table 2.2.1 HEAVY LOAD HANDLING COMPLIANCE MATRIX (NUREG-0612)

Criterion Compliance

1. Are safe load paths defined for Yes the movement of heavy loads to minimize the potential of impact, if dropped on irradiated fuel?

2. Will procedures be developed to Yes cover: identification of required equipment, inspection and acceptance criteria required before movement of load, steps and proper sequence for handling the load, defining the safe load paths, and special precautions?

3. Will crane operators be trained Yes and qualified?

4. Will special lifting devices meet Yes the guidelines of ANSI 14.6-1978?

5. Will non-custom lifting devices Yes be installed and used in accordance with ANSI B30.9-1971?

6. Will the cranes be inspected and Yes tested prior to use in rerack?

7. Does the crane meet the intent of Yes ANSI B30.2-1976 and CMMA-70?

2-13

N 480W NOM.

7 1 I '1 E19 E16 E13 El EXISTING EXIST ING EXISTING E10 E7 E4 EXISTING RACK RACK EXISTING EXISTING RACK EXISTING 8ACK 10 X 11 8 X 10 RACK RACK 8 X 10 RACK 6 X 11 11 X 10 11 X to 1I X 11 E20 E17 E14 Eli EXRSINGE EXISTING EXISTING E2 240' NOM.

4\) B X1l RACK 18 RACK X 11 RACK 11 X 11 E8 E5 EXISTING EXISTING RACK EXISTING RACK 8 X II RACKX x 11 X 11 E12 E6 E3 E21 E18 E15 EXISTING E9 EXISTING EXISTING EXISTING EXISTING EXISTING RACK EXISTING RACK RACK RACK RACK RACK RACK 11 X 9 11 X B 8 Xa 8 X 11 10 X 11 8 X 11 11 X 8

& _______________ A _______________ I___________

L I .1 (21) PACKS (2050) CELLS FIG 2.1.1 POOL LAYOUT - EXISTING RACKS

N

- ___ -- --- ---- 4Y80' NUM.

( 477,7' MIN. )

-- 207.72' -

FT mr14TTH4VE-q4- F 7 E19 E16 E13 4.46' EXISTING EXISTING NOM.

RACK EXISTING EIO RACK RACK EXISTING 10 X II 8 X 10 RACK a X 10 11 X 10 E20 E17 E14 El1 EXISTING EXISTING RCI IRACK RACK RACK 240' NOM.

aXI 1 tXI 8X i IIXII (238.34*MIN.)

nI E12 E21 E18 E15 EXISTING EXISTING EXISTING EXISTING RACK RACK RACK RACK 11 X 8 a X II 10 X II 8 X 11

& ________ J IllIllIllIll I _____

EIO-E21 - EXISTING RACKS (12) EXISTING RACKS - 1161 CELLS (9) NEW RACKS - 1250 CELLS 101AL 2411 CELLS FIW .. 2.. I .2 POOL LAYOUT - CAMPA IGN 1I

N 480' NUM.

( 477.7' MIN.)

127.7' E16 ING EXISTING RACK- - - - --

ae x 10 - - -- -- -------

E17 EXISING RACK 240"NOM.

(238.34' MIN.)

EIB EXISTING RACK 10 X II EliGE21 - EXISTING RACKS (6) EXISTING RACKS - 586 CELLS (12) NEW RACKS - 2003 CELLS 2589 TOTAL CELLS FIG. 2.1.3 POOL LAYOUT - CAMPAICN 11

N

'I

• MIN.

(19) RACKS - (3152) CELLS INCL. (1) RACK - 323 CELLS IN CASK PIT 14, FIG 2.1.4 POOL LAYOUT - CAMPAIGN III

TYPICAL TYPICAL DOUBLE PANEL CELL CAVITY CONTAINING POISON 304S/S CELL WALLS BEARING PAD ADJUSTABLE SUPPORT LEGS FIG. 2.1.5 PICTORIAL VIEW OF DAEC FUEL RACK MODULE 2-18

3.0 RACK FABRICATION AND APPLICABLE CODES The purpose of this section is to provide a comprehensive resume of the concepts and features underlying the design of the rack modules for the Duane Arnold Energy Center (DAEC) spent fuel pool.

3.1 Design Objective Some of the key objectives governing the design of the new high density storage racks for the DAEC spent fuel pool are defined in the following six criteria:

(1) The rack module is fabricated in such a manner that there is no weld splatter on the storage cell surfaces which would come in contact with the fuel assembly. Weld splatter on the lateral surface of the storage cell, which can come in contact with fuel assemblies, can be detrimental to its structural integrity.

(2) The storage locations are designed and constructed in such a way that redundant flow paths for the coolant are available in case the main designated flow path is blocked.

(3) The fabrication process based on the rack design involves operational sequences which permit immediate and convenient verification by the inspection staff to ensure that the "poison" panels are correctly installed.

(4) The storage cells are connected to each other by autogenously produced corner welds which leads to a honeycomb lattice construction. The extent of welding is selected to "detune" the racks from the ground motion such that the rack displacements are minimized.

(5) The baseplate provides a conformal contact surface for the "nose" of the fuel assembly.

(6) The module design affords built-in flexibility in the fabrication process so as to maintain the desired cell pitch even if certain "boxes" are slightly oversize.

The foregoing objectives are fully realized in the module design for the DAEC racks as described in the following.

3-1

3.2 Anatomy of the Rack Module The new high density rack module design employs storage cell locations with a single poison panel sandwiched between adjacent austenitic stainless steel surfaces.

A complete description of the rack geometry is best presented by first introducing its constituent parts. The principal parts are denoted as: (1) the storage box subassembly, (2) the baseplate, (3) the neutron absorber material, (4) picture frame sheathing, and (5) support legs.

Each part is briefly described below with the aid of sketches.

(1) Storage cell box subassembly: The so-called boxes are fabricated from two precision formed channels by seam welding them together in a seam welding machine equipped with copper chill bars, and pneumatic clamps to minimize distortion due to welding heat input. Figure 3.1 shows the box.

The minimum weld penetration is 80% of the box metal gage which is 16 gage. The boxes are manufactured to 5.9 inches nominal I.D. (inside dimension).

As shown in Figure 3.1, each box has two lateral holes punched near its bottom edge to provide auxiliary flow holes. In the next step, a picture frame sheathing is press formed in a precision die. The sheathing is shown attached to the box in Figure 3.2. The sheathing is made to an offset of 0.077 inch to ensure an unconstrained installation of BoralTM plates. The "picture frame sheathing" is attached to each side of the box with the poison material (BoralTM) installed in the sheathing cavity. The top of the sheathing is connected using a smooth continuous fillet weld near the top of the box.

The edges of the sheathing and the box are welded together to form a smooth lead-in edge. The box with integrally connected sheathing is referred to as the "composite box".

3-2

The composite boxes are arranged in a checkerboard array to form an assemblage of storage cell locations (Figure 3.3). The inter-box welding and pitch adjustment is accomplished by small longitudinal austenitic stainless connectors shown as small circles in Figure 3.3.

An assemblage of box assemblies thus prepared is welded edge to edge as shown in Figure 3.3 resulting in a honeycomb structure with axial, flexural and torsional rigidity depending on the extent of intercell welding provided. Figure 3.3 shows that two edges of each interior box are connected to the contiguous boxes resulting in a well-defined path for "shear flow".

(2) Baseplate: The baseplate, 3/4-inch thick, provides a continuous horizontal surface for supporting the fuel assemblies. The baseplate has a concentric hole with a 450 taper in each cell location to provide a seating surface conforming to the fuel assembly.

The baseplate is attached to the cell assemblage by fillet welds. The materials of construction of the baseplate and all other constituent elements of the rack module are presented in Section 3.3 of this document.

(3) The neutron absorber material: BoralTM is used as the neutron absorber material. BoralTM is manufactured by AAR Brooks and Perkins of Livonia, Michigan. More on this material follows in the next section.

(4) Picture Frame Sheathing: The sheathing is a part of the "composite box assembly" described earlier. The sheathing serves as the locater and retainer of the poison material. As such, it is made in repeatable precise dimensions. This is accomplished by press-forming stainless sheet stock in a specially high tolerance die.

The schematic of the sheathing is shown in Figure 3.2.

Figure 3.4 shows three storage cells in elevation with the fuel assembly shown in phantom in one cell. The poison screen extends over 144 inches vertical distance, straddling the active fuel length of all fuel assemblies to be used in the DAEC reactor.

(5) Support Legs: As stated earlier, all support legs are the adjustable type (Figure 3.5). The top (female) position is made of austenitic stainless steel material.

The bottom part is made of 17:4 Ph series stainless steel to avoid galling.

3-3

The support leg is equipped with a socket to enable remote leveling of the rack after its placement in the pool.

The baseplate projects beyond the cellular region of the rack modules by 0.125 inch (nominal). These baseplate projections serve as the designated impact locations for the racks in the event that the modules undergo kinematic movements during a seismic event.

3.3 Materials of Construction The principal material of construction utilized in the fabrication of the DAEC high density racks is austenitic stainless steel (ASME 240 and 479-304). One notable exception is the support spindle material which is made out of a special high strength (precipitation hardened) stainless steel (A564-630).

In addition to the structural and non-structural stainless material, the racks employ BoralTM, a patented product of AAR Brooks and Perkins, as the neutron absorber material. A brief description of BoralTm and a list of fuel pools in which it is used follows.

BoralTM is a thermal neutron poison material composed of boron carbide and 1100 alloy aluminum. Boron carbide is a compound having a high boron content in a physically stable and chemically inert form. The 1100 alloy aluminum is a light-weight metal with high tensile strength which is protected from corrosion by a highly resistant oxide film. The two materials, boron carbide and aluminum, are chemically compatible and ideally suited for long term use in the radiation, thermal and chemical environment of the spent fuel pool.

Boral's use in the spent fuel pools as a preferred neutron absorbing material can be attributed to the following reasons:

3-4

(1) The content and placement of boron carbide provides a very high absorption cross-section for thermal neutrons.

(2) Boron carbide, in the form of fine particles, is homogeneously dispersed throughout the central layer of the BoralTM panels.

(3) The boron carbide and aluminum materials in BoralTM are not detrimentally affected by long-term exposure to gamma radiation.

(4) The neutron absorbing central layer of BoralTM is clad with permanently bonded surfaces of aluminum.

(5) BoralTM is stable, strong, durable, and corrosion resistant.

BoralTM is manufactured under the control and surveillance of a Quality Assurance/Quality Control Program that conforms to the requirements of 10 CFR 50 Appendix B, "Quality Assurance Criteria for Nuclear Power Plants".

As indicated in Table 3.1, BoralTm has been licensed by the USNRC for use in numerous BWR and PWR spent fuel storage racks.

BoralTM Material Characteristics Aluminum: 1100 alloy aluminum is the metallic ingredient of BoralTM. The excellent corrosion resistance of the 1100 alloy aluminum is provided by the protective oxide film that develops on its surface from exposure to the atmosphere or water. This film prevents the loss of metal from general corrosion or pitting corrosion and the film remains stable between a pH range of 4.5 to 8.5.

Boron Carbide: The boron carbide contained in BoralTM is a fine granulated powder that conforms to ASTM C-750-80 nuclear grade Type III. The particles range in size between 60 and 200 mesh and the material conforms to the chemical composition and properties listed in Table 3.2.

3-5

A large body of test data and plant operating experience data is available in the publications in the public domain by Boral's manufacturer.

3.4 Codes, Standards, and Practices for the Spent Fuel Pool Modification The fabrication of the rack modules is performed under a strict quality assurance program which meets 10 CFR 50 Appendix B requirements.

The re-rack project will be designed, fabricated and installed in accordance with the latest approved revisions of the applicable industry codes and standards for ease of review. The use of these revisions does not constitute a change to previous IELP commitments for the DAEC. The revisions of codes and standards referenced in this document apply to the DAEC SFP re-rack modification only.

The following codes, standards and practices were used for all applicable aspects of the design, construction, and assembly of the spent fuel storage racks. Additional specific references related to detailed analyses are given in each section.

A. Design Codes (1) AISC Manual of Steel Construction, 8th Edition, 1980 (provides detailed structural criteria for linear type supports).

(2) ANSI N210-1976, "Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations" (contains guidelines for fuel rack design).

(3) American Society of Mechanical Engineers (ASME),

Boiler and Pressure Vessel Code,Section III, 1986 Edition.

(4) ANSI/AISC-N690-1984 - Nuclear Facilities - Steel Safety Related Structure for Design, Fabrication and Erection 3-6

(5) ASNT-TC-1A, 1984 American Society for Nondestructive Testing (Recommended Practice for Personnel Qualifications).

(6) ACI 349 Code Requirements for Nuclear Safety Related Concrete Structures B. Material Codes - Standards of ASME or ASTM, AS NOTED:

(1) E165 - Standard Methods for Liquid Penetrant Inspection (2) SA240 - Standard Specification for Heat-Resisting Chromium and Chromium-Nickel Stainless Steel Plate, Sheet and Strip for Fusion-Welded Unfired Pressure Vessels (3) A262 - Detecting Susceptibility to Intergranular Attack in Austenitic Stainless Steel (4) SA276 - Standard Specification for Stainless and Heat-Resisting Steel Bars and Shapes (5) SA479 - Steel Bars for Boilers & Pressure Vessels (6) C750 - Standard Specification for Nuclear-Grade Boron Carbide Powder (7) C992 - Standard Specification for Boron-Based Neutron Absorbing Material Systems for Use in Nuclear Spent Fuel Storage Racks (8) SA312 - Specification for Seamless and Welded Austenitic Stainless Steel Pipe (9) SA564 - Specification for Hot Rolled and Cold Finished Age-Hardening Stainless and Heat Resisting Steel Bars and Shapes (10) American Society of Mechanical Engineers (ASME),

Boiler and Pressure Vessel Code, Section II-Parts A and C, 1986 Edition (11) ASTM A262 Practices A and E - Standard Recommended Practices for Detecting Susceptibility to Intergrannular Attack in Stainless Steels (12) ASTM A380 - Recommended Practice for Descaling, Cleaning and Marking Stainless Steel Parts and Equipment 3-7

C. Welding Codes (1) ASME Boiler and Pressure Vessel Code,Section IX Welding and Brazing Qualifications, 1986 Edition (2) AWS D1.1 - Welding Standards D. Quality Assurance, Cleanliness, Packaging, Shipping, Receiving, Storage, and Handling Requirements (1) NQA-2-Part 2.2 1983 - Packaging, Shipping, Receiving, Storage, and Handling of Items for Nuclear Power Plants (During Construction Phase)

(2) NQA-1-1983 - Basic Requirements and Supplements (3) ASME Boiler and Pressure Vessel, Section V, Nondestructive Examination, 1986 Edition (4) ANSI - N45.2.6 - Qualifications of Inspection, Examination, and Testing Personnel for Nuclear Power Plants (Regulatory Guide 1.58)

E. Governing NRC Design Documents (1) "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," dated April 14, 1978, and the modifications to this document of January 18, 1979 (2) NRC Standard Review Plan NUREG 0800 - 9.1.2, Spent Fuel Storage, Rev. 3, July 1981 (3) NRC Standard Review Plan NUREG 0800 - 9.1.1, New Fuel Storage, Rev. 2, July 1981 (4) NRC Standard Review Plan NUREG 0800 - 3.8.4, Other Seismic Category I Structures, Rev. 1, July 1981 (5) NRC Standard Review Plan NUREG 0800 - 3.8.5, Foundations, Rev. 1, July 1981 (6) NRC Standard Review Plan NUREG 0800 -3.7.1 Seismic Design Parameters, Rev. 1, July 1981 (7) NRC Standard Review Plan NUREG 0800 - 3.7.2 Seismic System Analysis, Rev. 1, July 1981 (8) NRC Standard Review Plan NUREG 0800 - 3.7.3 Seismic Subsystem Analysis, Rev. 1, July 1981 3-8

(9) NRC Standard Review Plan , NUREG 0800 - 9.1.3 Spent Fuel Pool Cooling and Cleanup System, Rev. 1, July 1981 F. Other ANSI Standards (not listed in the preceding)

(1) ANSI/ASN 8.1 - 1983, Nuclear Criticality Safety in Operations with Fissionable Materials Outside Reactors (2) ANSI/ASN 8.7 - 1974, Guide for Nuclear Criticality Safety in the Storage of Fissile Materials (3) ANSI/ANS 8.11 - 1975, Validation of Calculation Methods for Nuclear Criticality Safety G. Code-of-Federal Regulations (1) 10 CFR 21 - Reporting of Defects and Non-compliance (2) 10 CFR 50 - Appendix A - General Design Criteria for Nuclear Power Plants (3) 10 CFR 50 - Appendix B - Quality Assurance Criteria for Nuclear Power Plants and Fuel Reprocessing Plants (4) 10CFR Part 20 - Radiation Protection Standards (5) 29CFR Section 1910.401 - OSHA Standards for Commercial Diving Operations H. Requlatory Guides (1) RG 1.13 - Spent Fuel Storage Facility Design Basis (2) RG 1.25 - Assumptions Used for Evaluating the Potential Radiological Consequences of a Fuel Handling Accident in the Fuel Handling and Storage Facility of Boiling and Pressurized Water Reactors (3) RG 1.28 - (endorses ANSI N45.2) - Quality Assurance Program Requirements, June, 1972 (4) RG 1.29 - Seismic Design Classification (5) RG 1.38 - (endorses ANSI N45.2.2) Quality Assurance Requirements for Packaging, Shipping, Receiving, Storage and Handling of Items for Water-Cooled Nuclear Power Plants, March, 1973 3-9

(6) RG 1.44 - Control of the Use of Sensitized Stainless Steel (7) RG 1.58 - (endorses ANSI N45.2.6) Qualification of Nuclear Power Plant Inspection, Examination, and Testing Personnel. Rev. 1, September, 1980 (8) RG 1.64 - (endorses ANSI N45.2.11) Quality Assurance Requirements for the Design of Nuclear Power Plants, October, 1973 (9) RG 1.74 - (endorses ANSI N45.2.10) Quality Assurance Terms and Definitions, February, 1974 (10) RG 1.88 - (endorses ANSI N45.2.9) Collection, Storage and Maintenance of Nuclear Power Plant Quality Assurance Records. Rev. 2, October, 1976 (11) RG 1.92 - Combining Modal Responses and Spatial Components in Seismic Response Analysis (12) RG 1.123 - (endorses ANSI N45.2.13) Quality Assurance Requirements for Control of Procurement of Items and Services for Nuclear Power Plants (13) NRC Regulatory Guide 3.41 Rev., May 1977 Validation of Calculation Methods for Nuclear Criticality Safety (14) NRC Regulatory Guide 1.26 Rev. 3, Feb. 1976, Quality Group Classifications and Standards for Water, Steam and Radioactive Containing Components of Nuclear Power Plants I. Branch Technical Position (1) CPB 9.1 Criticality in Fuel Storage Facilities (2) ASB 9-2 - Residual Decay Energy for Light-Water Reactors for Long-Term Cooling J. Other DAEC Updated Final Safety Analysis Report (UFSAR) 3-10

Table 3.1 BORALTM EXPERIENCE LIST (DOMESTIC AND FOREIGN)

Pressurized Water Reactors Vented Construc- Mfg.

Plant Utility tion Year Bellefont 1, 2 Tennessee Valley Authority No 1981 D.C. Cook 1,2 Indiana & Michigan Electric No 1979 Indian Point 3 NY Power Authority Yes 1987 Maine Yankee Maine Yankee Atomic Power Yes 1977 Salem 1, 2 Public Service Elec & Gas No 1980 Seabrook New Hampshire Yankee No Sequoyah 1,2 Tennessee Valley Authority No 1979 Yankee Rowe Yankee Atomic Power Yes 1964/1983 Zion 1,2 Commonwealth Edison Co. Yes 1980 Byron 1,2 Commonwealth Edison Co. Yes 1988 Braidwood 1,2 Commonwealth Edison Co. Yes 1988 Yankee Rowe Yankee Atomic Electric Yes 1988 Boiling Water Reactors Browns Ferry 1,2,3 Tennessee Valley Authority Yes 1980 Brunswick 1,2 Carolina Power & Light Yes 1981 Clinton Illinois Power Yes 1981 Cooper Nebraska Public Power Yes 1979 Dresden 2,3 Commonwealth Edison Co. Yes 1981 Duane Arnold Iowa Elec. Light & Power No 1979 J.A. FitzPatrick NY Power Authority No 1978 E.I. Hatch 1,2 Georgia Power Yes 1981 Hope Creek Public Service Elec & Gas Yes 1985 Humboldt Bay Pacific Gas & Electric Yes 1986 LaCrosse Dairyland Power Yes 1976 Limerick 1,2 Philadelphia Electric No 1980 Monticello Northern States Power Yes 1978 Peachbottom 2,3 Philadelphia Electric No 1980 Perry, 1,2 Cleveland Elec. Illuminating No 1979 Pilgrim Boston Edison No 1978 Shoreham Long Island Lighting Yes Susquehanna 1,2 Pennsylvania Power & Light No 1979 Vermont Yankee Vermont Yankee Atomic Power Yes 1978/1986 Hope Creek Public Service Elec & Gas Yes 1989 3-11

Table 3.1 (continued)

BORALTM EXPERIENCE LIST (DOMESTIC AND FOREIGN)

Foreign Installations Using Boralm France 12 PWR Plants Electricite de France South Africa Koeberg 1,2 ESCOM Switzerland Beznau 1,2 Nordostschweize rische Kraftwerke AG Gosgen Kernkraftwerk G osgen-Daniken AG Taiwan Chin-Shan 1,2 Taiwan Power Co mpany Kuosheng 1,2 Taiwan Power Co mpany Mexico Laguna Verde Comision Federa 1 de Electricidad 3-12

Table 3.2 Boron Carbide Chemical Composition, Weight %

Total boron 70.0 min.

Bw0 isotopic content in 18.0 natural boron Boric oxide 3.0 max.

Iron 2.0 max.

Total boron plus 94.0 min.

total carbon Boron Carbide Physical Properties Chemical formula B4C Boron content (weight) 78.28%

Carbon content (weight) 21.72%

Crystal Structure rombohedral Density 2.51 gm/cc-0.0907 lb/cu in Melting Point 2450 0 C-4442 0 F Boiling Point 3500 0 C-6332 0 F Microscopic capture 600 barn cross section

• Provided by AAR Brooks & Perkins.

3-13

K I; ~N NN N~N

>1 NN C.) /

- N WIIII SI-AM ~

N N NN - /7/

Ni NN FIGURE 3.1 SEAM WELDING PRECISION FORMED CHANNELS

,jI 2 I

I

- I

/

x x

-x x

CA) cI-LI F

FIGURE 3.2 SHEATHING SHOWN INSTALLED ON THE BOX

FORMED CELL BOX CELL

/

/

/

/

/

I p/

I :1 Iv

-74 4

7' A II II

~II I

.11 *Ij

___________________ ________________________ ___________________ ______________________ i ________________________

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• 1 I 'IJ_____________________________

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__________________ I ___________________

I .1 I I I I I FiG. 3.3 A CROSS SECTIONAL VIEW OF AN ARRAY OF STORAGE LOCATIONS 3-16

TTLH F -

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- - EN IF J

FIGURE 3.4 THREE CELLS IN ELEVATION VIEW 3-17

CELL (4) 3/4 0 FLHW H1LES

'" "1 IBASE PLATE 5" NIM 4 I/4 "

Io (4 ) GIUSSETS 4 1/4 41 UN, CLASS IA.

I. h"SIT FIC, 3.5" AUiTAlLE SUP PORT

4.0 CRITICALITY SAFETY ANALYSIS 4.1 Introduction The high density spent fuel storage racks for Duane Arnold Energy Center (DAEC) are designed to assure that the neutron multiplication factor (kff) is equal to or less than 0.95 with the racks fully loaded with fuel of the highest anticipated reactivity and the pool flooded with unborated 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 kff will be equal to or less than 0.95 with a 95%

probability at a 95% confidence level. Reactivity effects of abnormal and accident conditions have also been evaluated to assure that under credible abnormal conditions, the reactivity will be less than the limiting design basis value.

The design basis fuel is an 8x8 fuel rod assembly with a uniform initial enrichment of 4% U-235 containing 2% gadolinia (Gd2 0 3 ) in eight fuel rods. Criteria for the acceptable storage of higher enrichment fuel and for the GE-11 9x9 rod array were also developed based upon the k, in the standard core geometry (defined as an infinite array of fuel assemblies on a 6-inch lattice spacing at 20 0 C, without any control absorber or voids).

Applicable codes, standards, and regulations, or pertinent sections thereof, include the following:

• General Design Criterion 62, Prevention of Criticality in Fuel Storage and Handling.

• USNRC Standard Review Plan, NUREG-0800, Section 9.1.2, Spent Fuel Storage.

USNRC letter of April 14, 1978, to all Power Reactor Licensees - OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, including modification letter dated January 18, 1979.

4-1

• USNRC Regulatory Guide 1.13, Spent Fuel Storage Facility Design Basis, Rev. 2 (proposed), December, 1981.

• USNRC Regulatory Guide 3.41, Validation of Calculational Methods for Nuclear Criticality Safety (and related ANSI N16.9-1975).

• ANSI-8.17-1984, Criticality Safety Criteria for the Handling, Storage and Transportation of LWR Fuel Outside Reactors.

To assure the true reactivity will always be less than the calculated reactivity, the following conservative assumptions were made:

The racks were assumed to contain the most reactive fuel authorized to be stored in the facility without any control rods or any uncontained burnable poison, and with the fuel at the burnup corresponding to the highest reactivity during its burnup history.

Moderator is pure, unborated water at a temperature within the design basis range corresponding to the highest reactivity (4 0 C).

Criticality safety analyses are based upon the infinite multiplication .factor (k,,,), i.e., lattice of storage racks is assumed infinite in all directions. No credit is taken for axial or radial neutron leakage, except in the assessment of certain abnormal/accident conditions where neutron leakage is inherent.

• Neutron absorption in minor structural members is neglected, i.e., spacer grids are replaced by water.

4.2 Summary and Conclusions The design basis fuel assembly is a standard 8x8 array of BWR fuel rods containing U0 2 clad in Zircaloy, assumed to be uniform 4.0 wt%

U-235 enrichment initially, with 2% gadolinia burnable poison (Gd 20 3 ) in 8 fuel rods. The criticality safety was evaluated at the maximum reactivity over burnup where the gadolinium is nearly consumed. As shown in Figure 4.2.1 the maximum reactivity of the 4-2

reference fuel occurs at a burnup of approximately 8 MWD/KgU with a calculated k,, of 0.9085 (bias corrected). Adding the effect of calculational and manufacturing tolerances, the maximum k. in the storage rack is 0.935 (95% probability at the 95% confidence level) including all known uncertainties and plus a conservative allowance of 0.01 Sk for possible differences between fuel vendor calculations and those reported here.

The basic calculations supporting the criticality safety of the DAEC fuel storage racks for the design basis fuel are summarized in Table 4.2.1. For the design basis fuel, the fuel storage rack satisfies the USNRC criterion of a maximum kf less than or equal to 0.95, with a substantial margin. This margin is utilized in developing the acceptance criteria for storage of fuel of higher enrichment and different design in terms of the km, in the standard core geometry.

Calculations were made for fuel of 4.25%, 4.6% and the GE-11 9x9 design at 4.6% enrichment in the storage rack and in the standard core geometry (6.00-in. assembly pitch, 20 0 C). Since the higher enrichment fuel will contain a higher - but as yet unspecified

-- gadolinia content, these calculations were made to define the maximum acceptable km in the standard core geometry for fuel of various initial (average) enrichments and design concepts.

Calculations for the reference and higher-enriched fuel are summarized in Figure 4.2.2. The GE-11 fuel was lower in reactivity and the curves in Figure 4.2.2 bound the 9x9 fuel array at 4.6%

enrichment.

The limiting k. in the standard core geometry decreases with increasing initial average enrichment. The lower curve is for the maximum calculated reactivity (maximum rack k. of 0.935) and the upper curve is the acceptance limit for a kff of 0.95 in the storage rack, both curves with all uncertainties included.

Although the upper curve in Figure 4.2.2 could be used to define the acceptable reactivity, the most simple and conservative 4-3

criterion for acceptable storage of fuel in the DAEC storage racks is that (1) the fuel must have an average enrichment of 4.6% or less and (2) the k.. in the standard core geometry, calculated at the maximum over burnup, must be less than or equal to 1.31.

Parallel calculations also showed that any fuel of 3.1% average enrichment or less is acceptable for storage regardless of the gadolinium content or the k, in the standard core geometry. These criteria are expected to bound all present and future fuel designs anticipated to be used at DAEC.

4.3 Abnormal and Accident Conditions None of the abnormal or accident conditions that have been identified will result in exceeding the limiting reactivity (keff of 0.95). The effects on reactivity of credible abnormal and accident conditions are summarized in Table 4.3.1. The double contingency principle of ANSI N16.1-1975 (and the USNRC letter of April 1978) specifies that it shall require at least two unlikely independent and concurrent events to produce a criticality accident. This principle precludes the necessity of considering the occurrence of more than a single unlikely and independent accident condition concurrently.

Other hypothetical abnormal occurrences were considered and no credible occurrences or configurations have been identified that might have any adverse effect on the storage rack criticality safety.

4.4 Input Parameters 4.4.1 Fuel Assembly Design Specifications The design basis fuel assembly is a standard 8x8 array of BWR fuel rods containing U0 2 clad in Zircaloy (62 fuel rods with 2 water rods). For the nominal design case, fuel of uniform 4.0 wt% U-235 enrichment was assumed, with eight fuel rods containing 2%

4-4

gadolinia. The GE-11 fuel design, a 9x9 array of fuel rods with 7 rods replaced by two zircaloy water channels, was also evaluated.

Design parameters for both types of fuel are summarized in Table 4.4.1.

4.4.2 Storage Rack Cell Specifications The design basis storage rack cell consists of an egg-crate structure, illustrated in Figure 4.4.1, with fixed neutron absorber material (Boral) of 0.0162 g/cm 2 boron-10 areal density (0.015 gms B-10/cm 2 minimum) positioned between the fuel assembly storage cells in a 0.077 inch channel. This arrangement provides a nominal center-to-center lattice spacing of 6.060 inches. Manufacturing tolerances, used in evaluating uncertainties in reactivity, are indicated in Figure 4.4.1. The 0.060-in. stainless-steel box which defines the fuel assembly storage cell has a nominal inside dimension of 5.90 in. This allows adequate clearance for inserting/removing the fuel assemblies, with or without the Zircaloy fuel channel. Boral panels are not needed or used on the exterior walls of modules facing non-fueled regions, i.e., the pool walls.

4.5 Analysis Methodology In the fuel rack evaluation, criticality analyses of the high density spent fuel storage racks were performed with the CASMO-3 code [4.5.1], a two-dimensional multi-group transport theory code.

Independent verification calculations were made with the KENO-5a computer package [4.5.2], using the 27-group SCALE* cross-section library [4.5.3] with the NITAWL subroutine for U-238 resonance shielding effects (Nordheim integral treatment).

SCALE is an acronym for Standardized Computer Analysis for Licensing Evaluation, a standard cross-section set developed by the Oak Ridge National Laboratory for the USNRC.

4-5

Benchmark calculations are presented in Appendix A and indicate a bias of 0.0000 +/- 0.0024 for CASMO-3 and 0.0101 +/- 0.0018 (95%/95%)

for NITAWL-KENO5a. In the geometric model used in the calculations, each fuel rod and its cladding were explicitly described and reflecting boundary conditions (zero neutron current) were used in the axial direction and at the centerline of the Boral and steel plate between storage cells. These boundary conditions have the effect of creating an infinite array of storage cells in all directions.

The CASMO-3 computer code was used as the primary method of analysis as well as a means of evaluating small reactivity increments associated with manufacturing tolerances. Burnup calculations were also performed with CASMO-3, using the restart option to describe spent fuel in the storage cell. KENO-5a was used to assess the reactivity consequences of eccentric fuel positioning and abnormal locations of fuel assemblies.

4.6 Criticality Analyses and Tolerance Variations 4.6.1 Nominal Design Case With 2% Gd 2O 3 initially present in eight of the fuel rods, the reactivity of a fuel assembly of 4% enrichment increases with burnup to a maximum at 8 MWD/KgU (or slightly more) as shown in Figure 4.2.1. At 8 MWD/KgU burnup, , the nominal infinite multi plication factor, k.., in the storage racks is 0.909. With a Sk of 0.0263 for all known uncertainties statistically combined and all allowances (Table 4.2.1), the maximum km, is 0.935 which is less than the design basis limit for keffof 0.95. Reactivity effects of the natural uranium blanket normally located at the ends of the assemblies were [conservatively] neglected since an infinite fuel length was used.

4-6

Calculations were also made at low enrichments without any gadolinia present. These results, with all uncertainties and allowances included, are listed below, and show that the GE-11 9x9 array fuel exhibits a lower reactivity in the storage rack than the 8x8 fuel design for the same average enrichment.

3.0%E 3.1%E Ref. 8x8 Fuel 0.9427 0.9511 GE-11 9x9 Fuel 0.9340 0.9431 With axial leakage (-0.0020 6k) included, these data indicate that any fuel with an enrichment of 3.1% or less is acceptable for storage in the racks, regardless of the gadolinium content, fuel burnup or fuel rod array. The K-factor for 95% probability at a 95% confidence level was determined from NBS Handbook 91 [4.6.1].

Independent check calculations with NITAWL-KENO at an enrichment of 3.1% gave a k. of 0.9243 +/- 0.0025 (95%/95%) which confirms the CASMO-3 calculation at 3.1% enrichment (k. of 0.9248 without uncertainties). Similar calculations for the GE-11 assembly at 3.1% enrichment gave a k, of 0.9168 for CASMO-3 and 0.9230 +/- 0.0024 for KENO-5a.

4.6.2 Uncertainties Due to Manufacturing Tolerances The reactivity effects associated with manufacturing tolerances are listed in Table 4.6.1 and discussed below.

4.6.2.1 Boron Loading Variation The Boral absorber panels used in the storage cells are nominally 2

0.070-inch thick, with a B-10 areal density of 0.0162 g/cm . The manufacturing tolerance limit is +/- 0.0012 g/cm 2 in B-10 content, including both thickness and composition tolerances. This assures that the minimum boron-10 areal density will not be less than 4-7

0.015 gram/cm2 . Differential CASMO-3 calculations indicate that this tolerance limit results in an incremental reactivity uncertainty of +/- 0.0056 Sk.

4.6.2.2 Boral Width Tolerance Variation The reference storage cell design uses a Boral panel width of 5.00.

The tolerance on the Boral width is +/- 1/16 inch. Calculations (CASMO-3) showed that this tolerance corresponds to a 0.0019 Sk uncertainty.

4.6.2.3 Storage Cell Lattice Pitch Variation The design storage cell lattice spacing between fuel assemblies is 6.060 in. Decreasing the lattice pitch increases the reactivity.

For the manufacturing tolerance of +/- 0.03 inches, the corresponding uncertainty in reactivity is +/- 0.0039 Sk as determined by differential CASMO calculations.

4.6.2.4 Stainless Steel Thickness Tolerances The nominal thickness of the stainless steel box is 0.060 inches and 0.0235 inch for the steel backing plate. The maximum positive reactivity effect of the expected stainless steel thickness tolerances was calculated (CASMO-3) to be +/- 0.0009 Sk.

4.6.2.5 Fuel Enrichment and Density Variation CASMO-3 calculations of the sensitivity to small enrichment variations yielded an average coefficient of 0.0061 6k per 0.1 wt%

U-235, in the enrichment range from 4.0 to 4.25% enrichment. For an estimated tolerance on percent U-235 enrichment of +/-0.05, the maximum uncertainty is +/- 0.0031 Sk and becomes smaller for higher enriched fuel.

4-8

Calculations were also made to determine the sensitivity to the tolerance in U0 2 fuel density (+/- 0.14 g/cc). These calculations indicate that the storage rack k. is increased by 0.0019 6k for the expected maximum 10.59 gm/cc stack density. A lower fuel density results in correspondingly lower values of reactivity. Thus, the maximum uncertainty due to the tolerances on U02 density is +/- 0.0019 Sk.

4.6.2.6 Zirconium Flow Channel Elimination of the zirconium fuel channel results in a small (less than 0.01 6k) decrease in reactivity. More significant is a positive reactivity effect resulting from potential bulging of the zirconium channel, which moves the channel wall outward toward the Boral absorber. For the maximum expected bulging based on estimates provided by GE and conservatively assumed to be uniform throughout all assemblies, an incremental reactivity of + 0.0039 Sk could result (determined by differential CASMO calculations). Fuel assembly bowing results in a negative reactivity effect and is treated as an abnormal condition below).

4.6.3 Uncertainty in Depletion Calculations Since critical experiment data with spent fuel is not available for determining the uncertainty in depletion calculations, an allowance for uncertainty in reactivity was assigned based upon other considerations. The reactivity decrement in the absence of gadolinium is approximately 0.007 6k. Assuming the uncertainty in depletion calculations is less than 5% of the total reactivity decrement, an uncertainty in reactivity* equal to 0.0035 6k would result. At the burnup of maximum reactivity, the gadolinium is essentially depleted. However, for conservatism, the uncertainty in burnup was increased to +/- 0.0040 Sk as to compensate for

• Only that portion of the uncertainty due to burnup. Other uncertainties are accounted for elsewhere.

4-9

uncertainty in the residual gadolinium. Although the reactivity uncertainty due to depletion may be either positive or negative, for further conservatism it was added to the calculated kff rather than being combined statistically with other uncertainties. Over long periods of storage, the reactivity will continually decrease due to the decay of Pu-241 and growth of Am-241 which provides additional conservatism.

4.7 Higher Enrichments and GE-11 Fuel Calculations were made for fuel of 4.25%, 4.6% and the GE-11 (9x9 array) fuel at 4.6% enrichment. The reactivities in the rack, including calculational and manufacturing uncertainties, are shown in Figure 4.7.1 as a function of the fuel assembly k., in the standard core geometry at 20 0 C. By reviewing Figure 4.7.1, it can be concluded that a k. of 1.31 in the standard core geometry will conservatively bound both the 8x8 and 9x9 fuel designs for enrichments up to 4.6%. These curves use the uncertainties and allowances presented in Tables 4.2.1 and 4.5.

4.8 Abnormal and Accident Conditions 4.8.1 Temperature and Water Density Effects The moderator temperature coefficient of reactivity is negative.

Using the minimum temperature of 4 0 C therefore assures that the true reactivity will always be lower than the calculated value regardless of the temperature. Temperature effects on reactivity have been calculated and the results are shown in Table 4.8.1.

Introducing voids in the water in the storage cells (to simulate boiling) decreased reactivity, as shown in the table. Boiling at the submerged depth of the racks would occur at approximately 122 0 C.

4-10

4.8.2 Abnormal Location of a Fuel Assembly It is hypothetically possible to suspend a fuel assembly of the highest allowable reactivity outside and adjacent to the fuel rack, although such an accident condition is highly unlikely. The exterior walls of the rack modules facing the outside (where such an accident condition might be postulated to exist) is a region of high neutron leakage. With neutron leakage included, the reactivity with an extraneous fuel assembly of the maximum reac tivity, located outside and adjacent to the fuel rack, is less than the reference k,. Thus it is concluded that the abnormal location of a fuel assembly will have a negligible reactivity effect.

4.8.3 Eccentric Fuel Assembly Positioning The fuel assembly is normally located in the center of the storage rack cell with bottom fittings and spacers that mechanically restrict lateral movement of the fuel assemblies. Nevertheless, calculations with the fuel assembly moved into the corner of the storage rack cell (four-assembly cluster at closest approach),

resulted in a small negative reactivity effect. Thus, the nominal case, with the fuel assembly positioned in the center of the storage cell, yields the higher reactivity.

4.8.4 Zirconium Fuel Channel Distortion Consequences of bulging of the zirconium fuel channel are treated as a mechanical deviation in Section 4.6.2.6. Bowing of the zirconium channel (including fuel rods) results in a local negative reactivity effect analogous to that of eccentric positioning the fuel assembly toward one side of the storage cell. Thus, any bowing that might occur will result in a reduction in reactivity.

4-11

4.8.5 Dropped Fuel Assembly 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 from the active fuel region of more than the 12 inches which is sufficient to preclude neutron coupling (i.e., an effectively infinite separation). Maximum expected deformation under seismic or accident conditions will not reduce the minimum spacing to less than 12 inches. Consequently, fuel assembly drop accidents will not result in a significant increase in reactivity due to the separation distance.

4.8.6 Fuel Rack Lateral Movement Normally, the individual rack modules in the spent fuel pool are separated by a water gap which would normally eliminate the need for poison panels between rack modules. However, as an added precaution against possible mis-alignment or seismically-induced movement, Boral panels are installed in the rack wall along one side of the water gap. With this configuration, the maximum reactivity of the storage rack is not dependent upon the water-gap spacing between modules.

4.9 Comparison with Other Recently Re-racked NRC-Approved US Plants The list below gives the computed k. for recently executed high density storage rack projects. This listing indicates the range of km values which have been previously licensed for reracking:

Plant Year Utility Maximum km, Pilgrim (BWR) 1985 Boston Edison 0.935 Vogtle-2 (PWR) 1988 Georgia Power 0.936 Diablo Canyon (PWR) 1986 Pacific Gas and Electric 0.938 4-12

St. Lucie (PWR) 1988 Florida Power & Light 0.944 Byron 1 and 2 (PWR) 1988 Commonwealth Edison 0.947 Oyster Creek (BWR) 1984 GPU Nuclear 0.947 Millstone (BWR) 1989 Northeast Utilities 0.944 FitzPatrick (BWR) 1992 New York Power Authority 0.933 Nine Mile Point (BWR) Niagara Mohawk Power Corporation 0.935 Duane Arnold (BWR) Iowa Electric Light & Power 0.935 4.10 References for Section 4

[4.5.1] A. Ahlin, M. Edenius, H. Haggblom, "CASMO - A Fuel Assembly Burnup Program," AE-RF-76-4158, Studsvik report (proprietary).

A. Ahlin and M. Edenius, "CASMO - A Fast Transport Theory Depletion Code for LWR Analysis," ANS Transactions, Vol. 26, p. 604, 1977.

M. Edenius et al., "CASMO Benchmark Report,"

Studsvik/RF-78-6293, Aktiebolaget Atomenergi, March 1978.

[4.5.2] 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.

L.M. Petrie and N.F. Cross, "KENO-IV, An Improved Monte Carlo Criticality Program,"

ORNL-4938, Oak Ridge National Laboratory, November 1975.

[4.5.3] R.M. Westfall et al., "SCALE: A Modular Code System for Performing Standardized Computer Analyses for Licensing Evaluation," NUREG/CR 0200, 1979.

[4.6.1] M.G. Natrella, Experimental Statistics, National Bureau of Standards, Handbook 91, August 1963.

4-13

Table 4.2.1

SUMMARY

OF CRITICALITY SAFETY ANALYSES Temperature assumed for analysis 40 C Fuel Enrichment (average) 4.0%

Gadilionia Content, wt% Gd 2O3 2% in 8 rods Reference k. (CASMO) 0.9085 Calculational Bias 0.0000 Uncertainties Calculational +/-0.0024 Removal of flow channel negative Eccentric assembly location negative Tolerances (Table 4.6.1) +/-0.0080 Statistical combination) of uncertainties +/-0.0084 Effect of Channel Bulge 0.0039 Uncertainty in Depletion calculations 0.0040 Allowance for Vendor Calculations 0.0100 Total 0.9264 +/- 0.0084 Maximum reactivity 0.9348 Square root of sum of squares of all independent tolerance effects.

4-14

Table 4.3.1 REACTIVITY EFFECTS OF ABNORMAL AND ACCIDENT CONDITIONS Accident/Abnormal Condition Reactivity Effect Temperature increase Negative (Table 4.8.1)

Void (boiling) Negative (Table 4.8.1)

Assembly dropped on top of rack Negligible Misplacement of a fuel assembly Negligible Seismic Movement Negligible 4-15

Table 4.4.1 FUEL ASSEMBLY DESIGN SPECIFICATIONS FUEL ROD DATA 8 x 8 9 x 9 Cladding outside diameter, in. 0.483 0.440 Cladding inside diameter, in. 0.419 0.384 Cladding material Zr-2 Pellet diameter, inch 0.410 0.376 Enrichment (design basis) 4.0 +/-0.05 Maximum 4.6 4.6 U0 2 density (stack), g/cc U0 2 10.45 +/- 0.14 WATER ROD DATA Number of Water Rods 2 2 Inside diameter, inch 0.414 0.920 Outside diameter 0.484 0.980 Material Zr-2 FUEL ASSEMBLY DATA Fuel rod array 8 x 8 9 x 9 Number of fuel rods 62 74 Fuel rod pitch, inch 0.640 0.566 Fuel channel, material Zr-2 Inside dimension, inch 5.278 5.278 Outside dimension, inch 5.478 5.408 4-16

Table 4.6.1 Reactivity Uncertainties' due to Manufacturing Tolerances Nominal Quantity Value Tolerance Sk.m Boron Loading 0.0162 g/cm 2 +/-0.0012 g/cm 2 +0.0056 Boral width 5.00 inches +/-1/16 inches +/-0.0019 Lattice spacing 6.06 inches +/-0.03 inches +/-0.0039 SS thickness 0.06 and +/-0.004 inches +/-0.0009 0.0235 inches mean Fuel enrichment 4.0% U-235 +/-0.05% U-235 +/-0.0031 3

Fuel density 10.45 g/cm 3 +/-0.14 g/cm +/-0.0019 Statistical combination +/-0.0080 of uncertainties Square root of sum of squares of all independent tolerance effects.

4-17

Table 4.8.1 EFFECT OF TEMPERATURE AND VOID ON CALCULATED REACTIVITY OF STORAGE RACK Case Incremental Reactivity Change, Sk oC Reference 20 0C -0.003 40 0C -0.006 80 0C -0.015 122 0 C -0.026 122 0 C + 20% void -0.050 4-18

0.95

0. 94 '

0.93 0.92 0.91

( 0.90 0.89

0. 88

0. 87 0.86 0.85 5 6 8 9 10 11 12 13 Fuel Surnup, MWO/KgU FLg. 4.2.1 Change Ln Rack ReectLv Lty w Lth 4% Enr tched FuelI 4-19

1 .330 1.

0 E

0 1 L C oL 1

L 7a)

C>

o) 0 1.

U) E

( X

-~ 0 1

Cu C

1.

1 .295 1 . 290 ----r-i-- -r 4.0 4.1 4.2 4.3 4.4 4.5 4.6 In tt tel Average Enr ichmen t, wt% U-235 FLg. 4.2.2 k-InftnLte Limits in the Standard Core Geometry 4-20

BORAL 0.0162 g B-10/CM CENTERLINE THRU 5.00 +/- 1/16" WIDE

-'1 CENTER OF BORAL 0.070

• 0.004" Thk 4 SIDES /-SS CELL IN 0.077" CHANNEL 5.90 0.03" ID 0.060"/0.0235" Thk (MEAN 0.04525")

0.08025" WATER RODS 0.484" OD 0.414" I D NOT TO SCALE

• I I II 5.90" CELL ID .

I6 L S I6.060" LATTICE SPACING .l Fig. 4.4.1 CROSS-SECTION OF TYPICAL STORAGE CELL (CALCULATIONAL MODEL) 4-21

0.97 0.96-4.6 0

%4.0 E 0.91 F 4.6 E A)

'4 0.91 LL0.92

1. 28 1.29 1.30 1.31 1.32 1.33 1.34 1. 35 1.36 K-INFINITE IN CORE GEOMETRY FIG. 4.6.2 CORRELATION BETW-.EEN REACTIVITIES IN THE CORE AND IN THE RACK 4-22

~o CD,

-- 4I;

'-4 Ei xD K-INFINITE IN CORE GEOMETRY FIG. 4.7.1 CORRELATION BETWEEN REACTIVITIES IN THE CORE AND IN THE RACK 4-23

APPENDIX A BENCHMARK CALCULATIONS by Stanley E. Turner, PhD, PE HOLTEC INTERNATIONAL November, 1992

1.0 INTRODUCTION

AND

SUMMARY

The objective of this benchmarking study is to verify both the NITAWL-KENO5a(12 methodology with the 27-group SCALE cross-section library and the CASMO-3 codeO) for use in criticality safety calculations of high density spent fuel storage racks. Both calculational methods are based upon transport theory 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 are consistent with corresponding calculations reported in the literature.

Results of the benchmark calculations show that the 27-group (SCALE) NITAWL-KENO5a calculations consistently under predict the critical eigenvalue by 0.0101 +/- 0.0018 Sk (with a 95%

probability at a 95% confidence level) for critical experimentso 4) that are as representative as possible of realistic spent fuel storage rack configurations and poison worths.

Extensive benchmarking calculations of critical experi ments with CASM03 have also been reported(-, giving a mean kg of 1.0004 +/- 0.0011 for 37 cases. With a K-factor of 2.140 for 95%

probability at a 95% confidence level, and conservatively neglect ing the small overprediction, the CASMO3 bias then becomes 0.0000

+/- 0.0024. CASMO3 and NITAWL-KENO5a intercomparison calculations of infinite arrays of poisoned cell configurations (representative of typical spent fuel storage rack designs) show very good agreement, confirming that 0.0000 +/- 0.0024 is a reasonable bias and uncertain ty for CASMO3 calculations. Reference 5 also documents good agreement of heavy nuclide concentrations for the Yankee core isotopics, agreeing with the measured values within experimental error.

A - 1

The benchmark calculations reported here confirm that either the 27-group (SCALE) NITAWL-KENO or CASMO3 calculations are acceptable for criticality analysis of high-density spent fuel storage racks. Where possible, reference calculations for storage rack designs should be performed with both code packages to provide independent verification. CASMO3, however, is not reliable when large water gaps ( > 2 or 3 inches) are present.

2.0 NITAWL-KENO 5a BENCHMARK CALCULATIONS Analysis of a series of Babcock & Wilcox critical experiments(4 ), including some with absorber panels typical of a poisoned spent fuel rack, is summarized in Table 1, as calculated with NITAWL-KENO5a using the 27-group SCALE cross-section library and the Nordheim resonance integral treatment in NITAWL. Dancoff factors for input to NITAWL were calculated with the Oak Ridge SUPERDAN routine (from the SCALEO system of codes). The mean for these calculations is 0.9899 +/- 0.0028 (1 a standard deviation of the population). With a one-sided tolerance factor corresponding to 95% probability at a 95% confidence levelsO, the calculational bias is + 0.0101 with an uncertainty of the mean of +/- 0.0018 for the sixteen critical experiments analyzed.

Similar calculational deviations have been reported by ORNLm for some 54 critical experiments (mostly clean critical without strong absorbers), 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 NITAWL-KENO5a calculational model for use in criticality analysis of high density spent fuel storage racks. No trends in k, with intra-assembly water gap, with absorber panel reactivity worth, with enrichment or with poison concentration were identified.

Additional benchmarking calculations were also made for a series of French critical experiments(9 ) at 4.75% enrichment and for several of the BNWL criticals with 4.26% enriched fuel.

A - 2

Analysis of the French criticals (Table 2) showed a tendency to overpredict the reactivity, a result also obtained by ORNL 10 . The calculated km values showed a trend toward higher values with decreasing core size. In the absence of a significant enrichment effect (see Section 3 below), this trend and the overprediction is attributed to a small inadequacy in NITAWL-KENO5a in calculating neutron leakage from very small assemblies.

Similar overprediction was also observed for the BNWL series of critical experiments"), which also are small assemblies (although significantly larger than the French criticals).

In this case (Table 2), the overprediction appears to be small, giving a mean k,,, of 0.9959 +/- 0.0013 (1 a population standard deviation).

Because of the small size of the BNWL critical experiments and the absence of any significant enrichment effect, the overprediction is also attributed to the failure of NITAWL-KENO5a to adequately treat neutron leakage in very small assemblies.

Since the analysis of high-density spent fuel storage racks generally does not entail neutron leakage, the observed inadequacy of NITAWL-KENO5a is not significant.

Furthermore, omitting results of the French and BNWL critical experiment analyses from the determination of bias is conservative since any leakage that might enter into the analysis would tend to result in overprediction of the reactivity.

3.0 INTERPOLATION ROUTINE An interpolation routine was obtained from ORNL and is intended to interpolate the hydrogen scattering matrices for temperature in order to correct for the deficiency noted in NRC Information Notice 91-66 (October 18, 1991). Benchmark calcula tions were made against CASMO3, based on the assumption that two independent methods of analysis would not exhibit the same error.

Results of these calculations, shown in Table 3, confirm that the trend with temperature obtained by both codes are comparable.

This A - 3

agreement establishes the validity of the interpolation routine, in conjunction with NITAWL-KENO5a, in calculating reactivities at temperatures other than 200 C (the reference library temperature).

The deficiency in the hydrogen scattering matrix does not appear except in the presence of a large water gap where the scattering matrix is important. However, the absolute value of the km from CASMO3 is not reliable in the presence of a large water gap, although the relative values should be accurate. In the calculations shown in Table 3 and in Figure 1, the absolute reactivity values differ somewhat but the trends with temperature are sufficiently in agreement to lend credibility to the interpola tion routine.

4.0 CLOSE-PACKED ARRAYS The BAW close-packed series of critical experiments(2 ) intended to simulate consolidated fuel, were analyzed with NITAWL-KENO5a.

Results of these analyses, shown in Table 4, suggest a slightly higher bias than that for fuel with normal lattice spacings.

Because there are so few cases available for analysis, it is recommended that the maximum bias for close-packed lattices be taken as 0.0155, including uncertainty. This would conservatively encompass all but one of the cases measured.

Similar results were obtained by ORNLa, 5.0 CASMO3 BENCHMARK CALCULATIONS The CASMO3 code is a multigroup transport theory code utilizing transmission probabilities to accomplish two-dimensional calculations of reactivity and depletion for BWR and PWR fuel assemblies. As such, CASM03 is well-suited to the criticality analysis of spent fuel storage racks, since general practice is to treat the racks as an infinite medium of storage cells, neglecting leakage effects.

A - 4

CASMO3 is a modification of the CASMO-2E code and has been extensively benchmarked against both mixed oxide and hot and cold critical experiments by Studsvik Energiteknik 5 ). Reported ana lysesa5 of 37 critical experiments indicate a mean k. of 1.0004

+/-

0.0011 (la). To independently confirm the validity of CASMO3 (and to investigate any effect of enrichment), a series of calculations were made with CASMO3 and with NITAWL-KEN05a on identical poisoned storage cells representative of high-density spent fuel storage racks. Results of these intercomparison calculations* (shown in Table 5 and in Figure 2) are within the normal statistical variation of KENO calculations and confirm the bias of 0.0000

+/-

0.0024 (95%/95%) for CASMO3.

Since two independent methods of analysis would not be expected to have the same error function with enrichment, results of the intercomparison analyses (Table 5) indicate that there is no significant effect of fuel enrichment over the range of enrich ments involved in power reactor fuel. Furthermore, neglecting the French and BNWL critical benchmarking in the determination of bias is a conservative approach.

A second series of CASMO3-KENO5a intercomparison calculations consisting of five cases from the BAW critical experiments analyzed for the central cell only. The calculated results, also shown in Table 5, indicate a mean difference within the 95% confidence limit of the KENO5a calculations. This lends further credence to the recommended bias for CASMO3.

0Intercomparison between analytical methods is a technique endorsed by Reg. Guide 5.14, "Validation of Calculational Methods for Nuclear Criticality Safety".

A - 5

6.0 REFERENCES

TO APPENDIX A

1. Green, Lucious, Petrie, Ford, White, and Wright, "PSR-63

/NITAWL-1 (code package) NITAWL Modular Code System For Generating Coupled Multigroup Neutron-GAmma Libraries from ENDF/B", ORNL-TM-3706, Oak Ridge National Laboratory, November 1975.

2. R.M. Westfall et. al., "SCALE: A Modular System for Performing Standardized Computer Analysis for Licensing Evaluation",

NUREG/CR-0200, 1979.

3. A. Ahlin, M. Edenius, and H. Haggblom, "CASMO - A Fuel Assembly Burnup Program", AE-RF-76-4158, Studsvik report.

A. Ahlin and M. Edenius, "CASMO - A Fast Transport Theory Depletion Code for LWR Analysis", ANS Transactions, Vol. 26,

p. 604, 1977.

"CASMO3 A Fuel Assembly Burnup Program, Users Manual",

Studsvik/NFA-87/7, Studsvik Energitechnik AB, November 1986

4. M.N. Baldwin et al., "Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel", BAW-1484-7, The Babcock & Wilcox Co., July 1979.

5. M. Edenius and A. Ahlin, "CASMO3: New Features, Benchmarking, and Advanced Applications", Nuclear Science and Engineering, 100, 342-351, (1988)

6. M.G. Natrella, Experimental Statistics, National Bureau of Standards, Handbook 91, August 1963.

7. R.W. Westfall and J. H. Knight, "SCALE System Cross-section Validation with Shipping-cask Critical Experiments", ANS Transactions, Vol. 33, p. 368, November 1979

8. S.E. Turner and M.K. Gurley, "Evaluation of NITAWL-KENO Benchmark Calculations for High Density Spent Fuel Storage Racks", Nuclear Science and Engineering, 80(2):230-237, February 1982.

9. J.C. Manaranche, et. al., "Dissolution and Storage Experiment with 4.75% U-235 Enriched U02 Rods", Nuclear Technologv, Vol.

50, pp 148, September 1980

10. A.M. Hathout, et. al., "Validation of Three Cross-section Libraries Used with the SCALE System for Criticality Analy sis", Oak Ridge National Laboratory, NUREG/CR-1917, 1981.

A - 6

11. S.R. Bierman, et. al., "Critical Separation between Sub critical Clusters of 4.29 Wt. %2U Enriched U02 Rods in Water with Fixed Neutron Poisons", Battelle Pacific Northwest Laboratories, NUREG/CR/0073, May 1978 (with August 1979 errata).

12. G.S. Hoovler, et al., "Critical Experiments Supporting Underwater Storage of Lightly Packed Configurations of Spent Fuel Pins", BAW-1645-4, Babcock & Wilcox Company (1981).

13. R.M. Westfall, et al., "Assessment of Criticality Computation al Software for the U.S. Department of Energy Office of Civilian Radioactive Waste Management Applications", Section 6, Fuel Consolidation Applications, ORNL/CSD/TM-247 (undated).

A - 7

Table 1 RESULTS OF 27-GROUP (SCALE) NITAWL-KENO5a CALCULATIONS OF B&W CRITICAL EXPERIMENTS Experiment Calculated a Number ke IX 0.9922 +/- 0.0006 II 0.9917 +/- 0.0005 III 0.9931 +/- 0.0005 IX 0.9915 +/- 0.0006 XI 0.9903 +/- 0.0006 XI 0.9919 +/- 0.0005 XII 0.9915 +/- 0.0006 XIII 0.9945 +/- 0.0006 XIV 0.9902 +/- 0.0006 XV 0.9836 +/- 0.0006 XVI 0.9863 +/- 0.0006 XVII 0.9875 +/- 0.0006 XVIII 0.9880 + n.0006 XIX 0.9882 +/- 0.0005 XX 0.9885 +/- 0.0006 XXI 0.9890 +/- 0.0006 Mean 0.9899 +/- 0.0007')

Bias (95%/95%) 0.0101 +/- 0.0018 (1) Standard Deviation of the Mean, calculated from the kf values.

A - 8

Table 2 RESULTS OF 27-GROUP (SCALE) NITAWL-KENO5a CALCULATIONS OF FRENCH and BNWL CRITICAL EXPERIMENTS French Experiments Separation Critical Calculated Distance, cm Height, cm ke 0 23.8 1.0302 +/- 0.0008 2.5 24.48 1.0278 +/- 0.0007 5.0 31.47 1.0168 +/- 0.0007 10.0 64.34 0.9998 +/- 0.0007 BNWL Experiments Calculated Case Expt. No. kg No Absorber 004/032 0.9942 +/- 0.0007 SS Plates (1.05 B) 009 0.9946 +/- 0.0007 SS Plates (1.62 B) 011 0.9979 +/- 0.0007 SS Plates (1.62 B) 012 0.9968 +/- 0.0007 SS Plates 013 0.9956 +/- 0.0007 SS Plates 014 0.9967 +/- 0.0007 Zr Plates 030 0.9955 +/- 0.0007 Mean 0.9959 +/- 0.0013 A - 9

Table 3 Intercomparison of NITAWL-KENO5a (Interpolated) and CASMO3 Calculations at Various Temperatures Temperature CASMO 3 W-N-KENO5a )

4 0C 1.2276 1.2345 +/- 0.0014 17.5 0C 1.2322 1.2328 +/- 0.0015 25 0C 1.2347 1.2360 +/- 0.0013 50 0C 1.2432 1.2475 +/- 0.0014 75 0C 1.2519 1.2569 +/- 0.0015 120 0C 1.2701 1.2746 +/- 0.0014 Corrected for bias A - 10

Table 4 Reactivity Calculations for Close-Packed Critical Experiments Calc. BAW Pin Module Boron Calculated No. Expt. Pitch Spacing Conc. keff No. cm cm ppm KS01 2500 Square 1.792 1156 0.9891 +/- 0.0005 1.4097 KS02 2505 Square 1.792 1068 0.9910 +/- 0.0005 1.4097 KS1 2485 Square 1.778 886 0.9845 +/- 0.0005 Touching KS2 2491 Square 1.778 746 0.9849 +/- 0.0005 Touching KT1 2452 Triang. 1.86 435 0.9845 +/- 0.0006 Touching KTIA 2457 Triang. 1.86 335 0.9865 +/- 0.0006 Touching KT2 2464 Triang. 2.62 361 0.9827 +/- 0.0006 Touching KT3 2472 Triang. 3.39 121 1.0034 +/- 0.0006 Touching A - 11

Table 5 RESULTS OF CASMO3 AND NITAWL-KENO5a BENCHMARK (INTERCOMPARISON) CALCULATIONS Enrichment(') koo Wt. % U-235 NITAWL-KENO5ac) CASMO3 Skj 2.5 0.8376 + 0.0010 0.8386 0.0010 3.0 0.8773 + 0.0010 0.8783 0.0010 3.5 0.9106 + 0.0010 0.9097 0.0009 4.0 0.9367 + 0.0011 0.9352 0.0015 4.5 0.9563 + 0.0011 0.9565 0.0002 5.0 0.9744 +/- 0.0011 0.9746 0.0002 Mean 0.0008 Expt. No.()

XIII 1.1021 +/- 0.0009 1.1008 0.0013 XIV 1.0997 +/- 0.0008 1.1011 0.0014 XV 1.1086 +/- 0.0008 1.1087 0.0001 XVII 1.1158 +/- 0.0007 1.1168 0.0010 XIX 1.1215 +/- 0.0007 1.1237 0.0022 Mean 0.0012 (1) Infinite array of assemblies typical of high-density spent fuel storage racks.

km from NITAWL-KENO5a corrected for bias.

(3) Central Cell from BAW Critical Experiments A - 12

.28 1 . 27 1.24 11.224

.23 1 . 24 1 20 40 1

Terrperntura , eg se Fg. I COMPARISON OF CASMO-3 tnd KENCEe TEMPERATURE C77ENENCE A-1 3

1.00 z

z 0.

FUEL ENRICHMENT, WT% U-235 Ftg, 2 COMPAR:SON OF CASMO AND KENO-5e CALCULATIONS AT VAIULS ENREC-IENT$ IN REPRESCNTATIVE FUEL. S'ORAGE RAC' A-14

5.0 THERMAL-HYDRAULIC CONSIDERATIONS 5.1 Introduction This section provides a summary of the methods, models, analyses and numerical results to demonstrate the compliance of the reracked DAEC spent fuel pool with the provisions of Section III of the USNRC "OT Position Paper for Review and Acceptance of Spent Fuel Storage and Handling Applications", (April 14, 1978).

Similar methods of thermal-hydraulic analysis have been used in previous licensing efforts on high density spent fuel racks for Fermi 2 (Docket 50-341), Quad Cities 1 and 2 (Dockets 50-254 and 50-265), Rancho Seco (Docket 50-312), Grand Gulf Unit 1 (Docket 50-416), Oyster Creek (Docket 50-219), Virgil C. Summer (Docket 50-395), Diablo Canyon 1 and 2 (Dockets 50-275 and 50-323), Byron Units 1 and 2 (Dockets 50-454 and 50-455), St. Lucie Unit One (Docket 50-335), Millstone Point I (Docket 50-245), Vogtle Unit 2 (Docket 50-425), Kuosheng Units 1 & 2 (Taiwan Power Company),

Ulchin Unit 2 (Korea Electric Power Company), and J.A. FitzPatrick (Docket 50-333), Sequoyah Units 1 & 2 (Docket Nos. 50-327 and 50 328), D.C. Cook Units 1 & 2 (Docket Nos. 50-315 and 50-316), and Zion Units 1 & 2 (Docket Nos. 50-295 and 50-304).

The analyses to be carried out for the thermal-hydraulic qualification of the rack array may be broken down into the following categories:

(i) Pool decay heat evaluation and pool bulk temperature variation with time.

(ii) Determination of the maximum pool local temperature at the instant when the bulk temperature reaches its maximum value.

(iii) Evaluation of the maximum fuel cladding temperature to establish that bulk nucleate boiling at any location around the fuel is not possible with cooling available.

5-1

(iv) Evaluation of the time-to-boil if all heat rejection paths from the cooler are lost.

(v) Compute the effect of a blocked fuel cell opening on the local water and maximum cladding temperature.

The following sections present a synopsis of the methods employed to perform such analyses and final results.

5.2 Spent Fuel Cooling and Cleanup System Description The Spent Fuel Pool Cooling and Cleanup System contains two parallel loops. Each loop consists of one full capacity (450 gpm) 6 pump, one filter-demineralizer, and one 4.0x10 Btu/hr heat exchanger. The system takes suction from the pool skimmer surge tanks and circulates the pool water through the system and returns the cooled and purified water to the pool through diffusers.

Cooling water is supplied to the heat exchangers from the Reactor Building Closed Loop Cooling Water System, which is maintained below 95 0 F.

The heat exchangers in the RHR system are used in conjunction with the Fuel Pool Cooling and Cleanup System to supplement pool cooling when the RHR is not needed for cooling the reactor cavity. The arrangement of piping and valves also permits the use of the RHR system as an independent backup system.

Makeup water is normally provided to the skimmer surge tanks by the Condensate System. Makeup is also available from the Emergency Service Water system. A hose connection is provided on the Emergency Service Water system to ensure a Seismic Category I water supply. Makeup water can also be supplied directly to the Spent Fuel Pool through fire water hoses.

5-2

5.3 Decay Heat Load Calculations The decay heat load calculation is performed in accordance with the provisions of "USNRC Branch Technical Position ASB9-2, "Residual Decay Energy for Light Water Reactors for Long Term Cooling", Rev.

2, July, 1981.

5.4 Discharge Scenarios Four discharge scenarios are considered: the first two scenarios, presented in the following, are intended to demonstrate compliance of the DAEC spent fuel pool cooling system with the NUREG-0800, SRP 9.1.3 provisions. The latter two scenarios correspond to the actual plant refueling and abnormal discharge practices, respectively.

The background heat load from the previously stored fuel is computed using the data provided in Table 5.4.1. It is recognized that the actual discharge dates in the future will deviate from those assumed in Table 5.4.1. However, the effect of such deviations on the heat load and pool water temperature profile is negligible and therefore will not be required to be revisited in the future.

(i) Case 1 This scenario simulates the Standard Review Plan (SRP) discharge condition labelled as the normal scenario with appropriate modification for 18 month operating cycles. This is not intended to represent the actual refueling of the DAEC reactor.

The fuel pool is assumed to contain three normal batches (116 assemblies per batch) of discharged fuel with 4.5 years of exposure at full power. The previous two discharges occurred at scheduled 18 month intervals. The final discharge is assumed to occur one year after the second discharge (see Figure 5.4.1).

5-3

The discharge of fuel assemblies to the pool begins after 150 hours0.00174 days <br />0.0417 hours <br />2.480159e-4 weeks <br />5.7075e-5 months <br /> of decay in the reactor, and proceeds at the rate of 144 assemblies for each 24 hour2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> period. Only one fuel pool cooler is assumed to be operating. It is necessary to demonstrate that the peak pool water temperature is less than 140 0 F.

(ii) Case 2 This case simulates the full core offload scenario of the SRP. This case, like Case 1, is also intended to illustrate compliance of the spent fuel pool cooling system with the provisions of SRP 9.1.3. It does not represent an actual discharge scenario at DAEC.

The pool contains three batches of discharges. The heat load of two normal batches and one full core offload is considered. One normal batch (116 assemblies) is assumed to have 19 month decay and another normal batch (116 assemblies) has decayed for 36 days. One full core is offloaded to the pool after 150 hours0.00174 days <br />0.0417 hours <br />2.480159e-4 weeks <br />5.7075e-5 months <br /> decay (at the rate of 144 assemblies per day, see figure 5.4.2). Assume that both coolers are operating. The pool water temperature should be kept below boiling.

(iii) Case 3 This scenario corresponds to the actual discharge practice at DAEC. To bound the heat load in the pool, the calculations consider a total of 3152 locations for the fuel storage in the pool and are carried out at the point in time when the stored fuel inventory is such that the addition of a normal batch to the pool will leave it with insufficient capacity to accept another batch while maintaining the full core discharge reserve capability. The discharge of fuel is assumed to have occurred in accordance with the schedule presented in Table 5.4.1.

The discharge consists of first offloading the entire core to the pool followed by re-transfer to the reactor of all except the "burned" normal batch (116 assemblies).

The transfer to the pool begins after 120 hours0.00139 days <br />0.0333 hours <br />1.984127e-4 weeks <br />4.566e-5 months <br /> of in-core decay and is conducted at 144 assemblies per 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> (see Figure 5.4.3). Two cooling trains are assumed to be in operation. The peak pool 5-4

water temperature is sought to be limited to 180aF (which is well below the regulatory limit of 212 0 F).

(iv) Case 4 This scenario may be termed as the practical "abnormal" condition.

Assume that after the refueling described in the foregoing case, the full core is offloaded after 36 days of reactor operation. The duration of reactor shutdown during the previous refueling is assumed to be 45 days (see Figure 5.4.4). It is assumed that full core offload begins 120 hours0.00139 days <br />0.0333 hours <br />1.984127e-4 weeks <br />4.566e-5 months <br /> after reactor shutdown and is transferred at 144 assemblies per 24 hours. Both cooling trains are in operation. The acceptance criterion for this case is identical to that of Case 3.

Key discharge data for all cases may be found in Table 5.4.2.

5.5 Bulk Pool Temperatures In this section, we present the methodology for calculating the bulk pool temperature as a function of the time coordinate.

Further, the method to calculate the rate of temperature rise of the pool water and the time-to-boil, when all forced cooling paths are unavailable, is also presented.

In order to perform the analysis conservatively, the heat exchangers are assumed to be fouled to their design maximum. Thus, the temperature effectiveness, p, for the heat exchanger utilized in the analysis is the lowest postulated value calculated from heat exchanger thermal-hydraulic codes. The temperature effectiveness, p, is assumed to remain constant in the calculation.

The mathematical formulation can be explained with reference to the simplified heat exchanger alignment of Figure 5.5.1.

5-5

Referring to the spent fuel pool/cooler system, the governing differential equation can be written by utilizing conservation of energy:

C dT = QL - QHX (5-1) d7-QL = Pcons + Q (Q) - (T, ta) where:

C: Thermal capacitance of the pool (net water volume times water density and times heat capacity), Btu/oF.

QL: Heat load to the heat exchanger, Btu/hr.

Heat generation rate from recently discharged fuel, which is a specified function of time, T, Btu/hr.

Pcons = t Po: Heat generation rate from "old" fuel, Btu/hr. (8 = Dimensionless fraction of operating power and P0 = average assembly operating power, Btu/hr.)

Heat removal rate by the heat exchanger, Btu/hr.

QEV (T, a): Heat loss to the surroundings, which is a function of pool temperature T and ambient temperature ta, Btu/hr.

Qvg is a non-linear function of time if we assume the temperature effectiveness p is constant during the calculation. Qm< can, however, be written in terms of effectiveness p as follows:

Qux = W, Ct p (T - ti) (5-2) to - ti T - ti where:

Wt: Coolant flow rate, lb./hr.

Ct: Coolant specific heat, Btu/lb.OF.

5-6

p: Temperature effectiveness of heat exchanger.

T: Pool water temperature, OF ti: Coolant inlet temperature, OF to: Coolant outlet temperature, OF p is obtained by rating the heat exchanger on a Holtec proprietary thermal-hydraulic computer code. Q(T) is specified according to the provisions of "USNRC Branch Technical Position ASB9-2, "Residual Decay Energy for Light Water Reactors for Long Term Cooling", Rev.

2, July, 1981. Q(T) is a function of decay time, number of assemblies, and in-core exposure time. During the fuel transfer, the heat load in the pool will increase with respect to the rate of fuel transfer and equals Q(T) after the fuel transfer.

QEv is a non-linear function of pool temperature and ambient temperature. QEV contains the heat evaporation loss through the pool surface, natural convection from the pool surface and heat conduction through the pool walls and slab. Experiments show that the heat conduction takes only about 4% of the total heat loss

[5.5.1], therefore, can be neglected. The evaporation heat and nature convection heat loss can be expressed as:

QEv = m r A, + he A, e (5-3) where:

m: Mass evaporation rate, lb./hr. ft. 2 r: Latent heat of pool water, Btu/lb.

As: Pool surface area, ft. 2 hr: Convection heat transfer coefficient at pool surface, Btu/ft. 2 hr. OF E = T-ta: The temperature difference between pool water and ambient air, OF 5-7

The mass evaporation rate m can be obtained as a non-linear function of e. We, therefore, have m = hD () (WP 5 - Was) (5-4) where:

Wps: Humidity ratio of saturated moist air at pool water surface temperature T.

Was: Humidity ratio of saturated moist air at ambient temperature ta h D (8): Diffusion coefficient at pool water surface. hD is a non-linear function of 8, lb./hr. ft. 2 OF The non-linear single order differential equation (5-1) is solved using Holtec's Q.A. validated numerical integration code "ONEPOOL".

The next step in the analysis is to determine the time-to-boil if all forced cooling paths become unavailable.

Clearly, the most critical instant of loss-of-cooling is when pool water temperature has reached its maximum value. It is assumed that makeup water is added at the rate of G lb./hr. The makeup water is at temperature, tool. The governing enthalpy balance equation for this condition can be written as dT

[C + G(Ct)(T - TO)] - = + Q(T + Ti) + G (Ct) (tcoo - T) dT where water is assumed to have specific heat of unity, and the time coordinate 7 is measured from the instant maximum pool water temperature is reached. TO is the time coordinate when the makeup water application is begun. Tms is the time coordinate measured from the instant of reactor shutdown to when maximum pool water temperature is reached. T is the dependent variable (pool water temperature). For conservatism, QEV is assumed to remain constant after pool water temperature reaches and rises above 170 0 F.

5-8

A Q.A. validated numerical quadrature code is used to integrate the foregoing equation. The pool water heat up rate, time-to-boil, and subsequent water evaporation-time profile are generated and compiled for safety evaluation.

5.6 Local Pool Water Temperature In this section, a summary of the methodology for evaluating the local pool water temperature is presented.

5.6.1 Basis The local water temperature analysis uses the bulk pool temperature as the datum and establishes the maximum incremental water temperature which may exist adjacent to the most heat emissive fuel assembly in the pool.

In order to determine an upper bound on the maximum local water temperature, a series of conservative assumptions are made. The most important assumptions are listed below:

The fuel pool will contain spent fuel with varying time after-shutdown ('r). Since the heat emission falls off rapidly with increasing r,, it is conservative to assume that all fuel assemblies are from the latest batch discharged simultaneously in the shortest possible time and they all have had the maximum postulated years of operating time in the reactor. The heat emission rate of each fuel assembly is assumed to be equal and maximum.

As shown in the pool layout drawings, the modules occupy an irregular floor space in the pool. For the hydrothermal analysis, a circle circumscribing the actual rack floor space is drawn (Figure 5.6.1). It is further assumed that the cylinder with this circle as its base is packed with fuel assemblies at the nominal layout pitch.

The actual downcomer space around the rack module group varies. The nominal downcomer gap available in the pool is assumed to be the total gap available around the idealized cylindrical rack; thus, the maximum resistance to downward flow is incorporated into the analysis 5-9

(Figures 5.6.2 and 5.6.3) (i.e. minimum gap between the pool wall and rack module, including seismic kinematic effect).

No downcomer flow is assumed to exist between the rack modules.

5.6.2 Model Description In this manner, a conservative idealized model for the rack 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 5.6.2 shows a typical "flow chimney" rendering of the thermal-hydraulics model. The governing equation to characterize the flow field in the pool can now be written. The resulting integral equation can be solved for the lower plenum velocity field (in the radial direction) and axial velocity (in-cell velocity field), by using the method of collocation. The hydrodynamic loss coefficients which enter into the formulation of the integral equation are also taken from well-recognized sources (Ref. 5.6.1) and wherever discrepancies in reported values exist, the conservative values are consistently used. Reference 5.6.2 gives the details of mathematical analysis 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 of the storage location with the minimum axial flow (i.e, maximum water outlet temperatures). This is called the most "choked" location.

In order to find an upper bound on the temperature in a typical cell, it is assumed that it is located at the most choked location.

Knowing the global plenum velocity field, the revised axial flow through this choked cell can be calculated by solving the Bernoulli's equation for the flow circuit through this cell. Thus, an absolute upper bound on the water exit temperature and maximum fuel cladding temperature is obtained. In view of these aforementioned assumptions, the temperatures calculated in this 5-10

manner overestimate the temperature rise that will actually occur in the pool. Holtec's proprietary computer code THERPOOL, based on the theory of Ref. 5.6.2, automates this calculation. The analysis procedure embodied in THERPOOL has been accepted by the Nuclear Regulatory Commission on several dockets. The Code THERPOOL for local temperature analyses includes the calculation of void generations. The effect of void on the conservation equation, crud layer in the clad, flux trap temperature due to gamma heating, and the clad stress calculation when a void exists, are all incorporated in THERPOOL. The major input data are given in Table 5.6.1.

5.7 Cladding Temperature This section contains a description of the method to calculate the temperature of the fuel cladding.

The maximum specific power of a fuel array qA can be given by:

qA q (1) where:

F = radial peaking factor q = average fuel assembly specific power The peaking factors are given in Table 5.7.1. The maximum temperature rise of pool water in the most disadvantageously placed fuel assembly, defined as one which is subject to the highest local pool water temperature, is computed 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.

A fuel rod can produce Fz times the average heat emission rate over

• THERPOOL has been used in qualifying the spent fuel pools for Enrico Fermi Unit 2 (1980), Quad Cities 1 and 2 (1981), Oyster Creek (1984), Virgil C. Summer (1984), Rancho Seco (1983), Grand Gulf Unit 1 (1985), Diablo Canyon 1 and 2 (1986), St. Lucie Unit One (1988), J.A. FitzPatrick (1991), Three Mile Island Unit One (1992), among others.

5-11

a small length, where Fz is the axial rod peaking factor. The axial heat distribution in a rod is generally a maximum in the central region, and tapers off at its two extremities.

It can be shown that the power distribution corresponding to the chopped cosine power emission rate is given by 77 (a + x) q(x) = qA sin 1 + 2a where:

1: active fuel length a: chopped length at both extremities in the power curve x: axial coordinate with origin at the bottom of the active fuel region The value of a is given by 1 z a=

1 - 2z where:

1/2 1 1 1 2 z= +

S Fz 72 z2 72 where Fz is the axial peaking factor.

The cladding temperature T, is governed by a third order differential equation which has the form of d3 T d2 T dT

- + oi - - 02 - = f (X) d x' d x2 dx where a,, C2 and f(x) are functions of x, and fuel assembly geometric properties. The solution of this differential equation with appropriate boundary conditions provides the fuel cladding temperature and local water temperature profile.

5-12

In order to introduce some additional conservatism in the analysis, we assume that the fuel cladding has a crud deposit with .005 OF sq.ft.-hr/Btu thermal resistance, which covers the entire surface.

5.8 Results This section provides a synopsis of the results for the maximum values of the bulk pool temperature, local water temperature, and fuel cladding temperature. Calculated results for the time available for operator action in the aftermath of loss-of-forced cooling of the pool water are also provided.

(i) Bulk Temperature Results Table 5.8.1 presents the major design input for bulk pool temperature analysis. Calculations show that the maximum bulk temperature for Case 1 (normal discharge per the SRP scenario) is calculated to be 137.1 0 F with one of the two pumps and one of the two heat exchangers (one train) in operation, which is below the SRP limit of 140 0 F. The decay heat load coincident to the maximum temperature is 5.66 x 106 Btu/hr (excluding 0.37 x 106 evaporation heat losses). The maximum SRP abnormal 160.4 0 F with both cooling temperature is calculated to be trains in operation, which is also below the NRC criterion of no nucleate boiling. The coincident heat load (excluding 1.14 x 106 Btu/hr. evaporation heat losses) is found to be 17.59 x 106 Btu/hr. For the end-of-storage-life refueling full core offload, the maximum pool temperature is calculated to be 0

164.6 F with both cooling trains running. The full core offload is assumed to start 120 hours0.00139 days <br />0.0333 hours <br />1.984127e-4 weeks <br />4.566e-5 months <br /> after reactor shutdown and to transfer at the rate of 144 assemblies per day. The coincident heat load is 18.73 x 106 Btu/hr. (less 1.38 x 106 Btu/hr heat losses). The full core offload at the beginning of the cycle is also analyzed (Case 4 in Table 5.8.2) and is found to result in slightly lower temperature than the foregoing full core offload case (maximum pool water temperature = 163.2 0 F, maximum heat load = 18.37 x 106 Btu/hr).

The design of the existing fuel pool cooling system and the RHR system permit the operations of the systems in parallel to maintain the bulk pool temperature below 150 0 F. The results of bulk pool temperatures are summarized in Table 5.8.2.

Temperatures and heat loads are plotted vs. time-after reactor-shutdown in Figures 5.8.1 through 5.8.8.

5-13

(ii) Results for Loss-of-Cooling Scenarios The postulated loss-of-cooling events are also considered. A massive loss of water event due to the failures of the gates separating the fuel pool from the cask pool and the reactor cavity is also postulated, together with the event of loss-of cooling. That is, the loss-of-forced cooling occurs when the level of water in the fuel pool is only 16 feet above the pool liner. The assumptions provide a most critical accident scenario. The minimum time-to-action is calculated to be 5.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> and the maximum boil-off rate is 43.11 gpm. The makeup systems can provide up to 75 gpm makeup water directly to the pool. The results of loss-of-cooling for all cases are summarized in Table 5.8.3. Figures 5.8.9 through 5.8.12 show the variation of the pool water elevation during the postulated loss-of-cooling events. It is demonstrated that the water level can be maintained above the top of the active fuel if the makeup water can be initiated 21.8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> after the loss-of-cooling during a normal batch discharge or 5.5 hours5.787037e-5 days <br />0.00139 hours <br />8.267196e-6 weeks <br />1.9025e-6 months <br /> after the loss-of-cooling during a full core offload.

(iii)Results of Local Pool Water and Cladding Temperature The maximum local water temperature of the limiting case (Case

3) is calculated to be 216.6 0 F and the maximum local fuel cladding temperature is 265.1 0 F. If the limiting cells are 50% blocked on the top, the maximum local water temperature becomes 227.5 0 F and the maximum fuel cladding temperature is 274.0aF (see Table 5.8.4). No nucleate boiling is indicated at any location. It is therefore concluded that the reracked Duane Arnold spent fuel pool complies with all thermal hydraulic regulatory criteria.

5.9 References for Section 5

[5.5.1] Wang, Yu, "Heat Loss to the Ambient From Spent Fuel Pools: Correlation of Theory with Experiment", Holtec Report HI-90477, Rev. 0, April 3, 1990.

[5.6.1] General Electric Corporation, R&D Data Books, "Heat Transfer and Fluid Flow", 1974 and updates.

[5.6.2] Singh, K.P. et al., "Method for Computing the Maximum Water Temperature in a Fuel Pool Containing Spent Nuclear Fuel", Heat Transfer Engineering, Vol. 7, No. 1-2, pp.

72-82 (1986).

5-14

Table 5.4.1 DAEC EXISTING AND PROJECTED FUEL DISCHARGE SCHEDULE Cycle Date Number Batch Size Number Bundles Discharged 1A 4 4 6/75 1B 84 88 2/76 2 100 188 3/77 3 88 276 3/78 4 88 364 2/80 5 84 448 3/81 6 128 576 2/83 7 120 696 2/85 8 128 824 3/87 9 120 944 9/88 10 104 1048 6/90 11 104 1152 2/92 12 128 1280 7/93 13 128 1408 2/95 14 116 1524 9/96 15 116 1640 3/98 16 116 1756 9/99 17 116 1872 3/01 18 116 1988 9/02 19 116 2104 3/04 20 116 2220 9/05 21 116 2336 3/07 22 116 2452 9/08 23 116 2568 3/10 24 116 2684 9/11 25 116 2800 3/13 26 116 2916 9/14 Note: All fuel bundles in the pool are assumed to have 4.5 years full power operation in reactor. This table provides the basis for decay heat evaluations. The assumptions made at the end of the plant licensing life, which deviate from Table 1.1, are solely to provide a bounding analysis condition.

5-15

Table 5.4.2 DATA FOR DISCHARGE SCENARIOS Number of assemblies in refueling batch: 116 Number of assemblies in full core: 368 Number of fuel pool cooling trains in parallel: 2 (one cooling train contains one pump and one heat exchanger)

Fuel normal exposure time, hrs.: 39420 5-16

Table 5.6.1 DATA FOR LOCAL TEMPERATURE ANALYSIS*

Type of fuel assembly GE 8x8 Fuel cladding outer diameter, inches 0.484 Fuel cladding inside diameter, inches 0.414 Storage cell inside dimension, inches 5.90 Active fuel length, inches 150 Number of fuel rods/assembly 62 Operating power per fuel assembly 15.67 Po x 106, Btu/hr Cell pitch, inches 6.06 Cell height, inches 169 Bottom height, inches 5 Plenum radius, feet 25 Nominal gap between pool wall 2.0 and outer rack periphery, inches

. The fuel bundle data corresponds to an 8x8 GE assembly.

However, the resistance characteristics of other assembly types, such as GE10, used in the DAEC pool, are only marginally different.

Therefore, the 8x8 assembly data is used for all prototypical runs.

5-17

Table 5.7.1 PEAKING FACTORS FACTOR VALUE Radial 1.5 Total 3.0 5-18

Table 5.8.1 MAJOR DESIGN INPUT Net water volume of pool, ft.3 21600 Fuel pool thermal capacity, 106 Btu/oF 1.32 Average operating power of a fuel assembly, 106 Btu/hr 15.67 Coolant (SW) inlet temperature, OF 95 Coolant (SW) flow rate, 106 lb/hr 0.397 Fuel pool water surface area, ft. 2 800 Maximum fuel pool building ambient temperature, OF 100 Assumed humidity of building air, % 100 5-19

Table 5.8.2 SFP BULK POOL TEMPERATURE Coincident Coincident Coincident Maximum Time After Heat Load Evaporation Pool Reactor shut to SFP Hxs Heat Losses Temp., OF down, hrs. 106 Btu/hr 106 Btu/hr Discharge 137.1 197 5.66 0.37 in Case 1 Discharge 160.4 221 17.59 1.14 in Case 2 Discharge 164.6 190 18.73 1.38 in Case 3 Discharge 163.2 189 18.37 1.30 in Case 4 0-5-20

Table 5.8.3 RESULTS OF LOSS-OF-COOLING Case Time Required for Maximum Evaporation Number Operator Action (hours) Rate (GPM) 1 21.8 12.81 2 6.0 40.15 3 5.5 43.11 4 5.7 42.14 5-21

Table 5.8.4 MAXIMUM LOCAL POOL WATER AND FUEL CLADDING TEMPERATURE FOR THE LIMITING CASE (Case 3)

Maximum Local Maximum Local Pool Water Fuel Cladding Temp. OF Temp., OF No Blockage 216.6 265.1 50% Blockage 227.5 274.0 5-22

BATCH ONE BATCH TWO BATCH THREE

-J 18 MONTHS 12 MONTHS 150 HOURS p.

EFILO 0

ONE REFUELING LOAD id In w ONE REFUELING LOAD ONE REFUELING LOAD REACTOR SHUTDOWN REACTOR SHUTDOWN REACTOR SHUTDOWN FIGURE 5.4.1 DAEC DISCHARGE SCENARIO CASE 1

BATCH ONE BATCH TWO BATCH THREE

-J 0

0 O.

w U-U' I

REACTOR SHUTDOWN REACTOR SHUTDOWN REACTOR SHUTDOWN FIGURE 5.4.2 DISCHARGE SCENARIO CASE 2

EOC FULL CORE OFFLOAD

-j 0

0 POOL FILLED EXCEPT STORAGE CAPACITY FOR ONE FULL CORE REMAINING 0

z RELOA D252 FAS BACK TO REACTOR 120 HRS U. FULL CORE OFFLOAD us AT 6 FAS/HR I

45 DAYS REACTOR SHUTDOWN . END OF OUTAGE FIGURE 5.4.3 DISCHARGE SCENARIO CASE 3

DISCHARGE IN CASE 3 ONE FULL CORE OFFLOAD RELOAD 252 FAS BACK TO REACTOR z

O )REACTOR OPERATING w

FULL CORE (368 FAS) OFFLOAD FULL CORE (368 FAS) OFFLOAD AT6FAS/HR 45 DAYS 36 DAYS REACTOR SHUTDOWN END OF OUTAGE REACT OR SHUTDOWN FIGURE 5.4.4 DISCHARGE SCENARIO CASE 4

EVAPORATION HEAT LOSS tI W FIGURE 5.5.1 5-27

Idealized Outline Actual Ot Rack Assembly Outline of Rack Assembly Idealized Outline of Assumed Added Pool Boundary Fuel Assemblies FIGURE 5.6.1 IDEALIZATION OF RACK ASSEMBLY 5-28

TOUT 0O TIN H

u*(- Q HEAT ADDITION FIGURE 5.6.2 THERMAL CHIMNEY FLOW MODEL 5-29

RACK 1 I I

I U I I I I DOWN COMER PLENUM FIGURE 5.6.3 CONVECTION CURRENTS IN THE POOL 5-30

HOLTEC INTERNATIONAL DAEC, SRP NORMAL DISCHARGE, ONE TRAIN - CASE I REACTOR SHUTDOWN U'

I w

l.a i

200 400 600 TIME AFTER REACTOR SHUTDOWN, HRS PIGURE 5.8.1

HOLTEC INTERNATIDNAL DAEC, SRP ABNORMAL CONDITION, TWO TRAINS - CASE 2.

BATCH HEAT LOAD FILL CORE -EAT LOAD IO 140 U'

I 0120 100 SO . 1 1 I C I I I I I j l i l l l l l 0 400FE 120 TIME AFTER REACTOR SHUTDOWN. HRS FIGURE 5.8.2

HOLTEC INTERNATIONAL DAEC, EOC FULL CORE OFFLOAD, TWO TRAINS - CASE 3 REACTOR SHUTDOWN 110 160 140.

120 10600 M200 400 RS0 600 TIME AFTER REACTOR SHUTDOWN. HRS FIGURE 5.8.3

HOLTEC INTERNATIONAL DAEC, BOC FULL CORE OFFLOAD, TWO TRAINS - CASE 4 REACTOR SHUTDOWN REACTOR SHUTDOWN 190 LjL 160 Ut 140 w

ab 120 109 0 BO0 1000 1500 2000 2500 TIME AFTER INITIAL REACTOR SHUTDOWN, HRS FIGURE 5.8.4

HOLTEC INTERNATIONAL DAEC, 8RP NORMAL CONDITION. ONE TRAIN - CASE 1 REACTOR SHUTDOWN B. 0E+6 NET HEAT LOAD 600C 6 Ik K

I U' S4. OE+ 6 w

U'

~J.p~-E

-. ~-.-

EVAPORATION HEAT LOSSES 0.00C+0 I I II I II ij iI I I i i Ii i l ri i I i 0 200 400 600 B00 TIME AFTER REACTOR SHUTDON1 , RS FIGURE 5.8.5

HOLTEC INTERNATIONAL DAEC, SRP ABNORMAL CONDITION, TWO TRAINS - CASE 2 REACTOR SHUTDOWN REACTOR SHUTDOWN 2.OEs 7 BATCH HEAT LOAD FULL CORE HEAT LOAD 1.50E+7 m 1.00E07 at EVAPORATION HEAT LOSSES 0.00C.0 111111111 111111111 111111111 TM400 B 12 TIME AFTER INITIAL REACTOR SHUTDOWN, HRS FIGURE 5.8.6

HOLTEC INTERNATIONAL DAEC. EOC FULL CORE OFFLOAD, TWO TRAINS - CASE 3 REACTOR SHUTDOWN 2.00E+7 NET HEAT LOAD 1 .E6E+7 1 .0E+7-TIME AFTER REACTOR SHUTDOWN, HRS FIGURE 5.8.7

HOLTEC INTERNATIONAL DAEC. BOC FULL CORE OFFLOAD. TWO TRAINS - CASE 4 REACTOR SHUTDOWN REACTOR SHUTDOWN e.00C-7 1 . 50E+7 w 1.0E+7 w

0 NET HEAT LOAD 6.00OE+ 6 EVAPORATION IFEAT LOSSES Il1i111l 0 50 1000 1500 2000 2500 TIME AFTER INITIAL REACTOR SHUTDOWN, HRS FIGURE 5.8.8

HOLTEC INTERNATIONAL DAEC SFP LOS OF COOLING - CASE I (8P N(R1AL DISCHARGE, OE TRAIN)

LOSS OF COOLING 40.0-1 MAX. POOL WATER LEVEL V

w w

~0 3 TIff AFTER LOSS OF COOLING, HRS FIGURE 5.8.9

HOLTEC INTERNATIONAL DAEC SFP LOSS OF COOLING - CASE 2 ( 8P FLLL CORE OFFLOAD, TWO TRAINS)

LOSS OF COOLING 4.e.-1 MAX. POOL WATER LEVEL is On P

TIME AFTER LOSS OF COOLING. HRS FIGURE 5.8.10

HOLTEC INTERNATIONAL DAEC SFP LOSS OF COOLING - CASE 3 CEOC FL1.L CORE OFFL.OAD, 40 TRAINS)

LOSS OF COOLING Ii:

u' TIME AFTER LOSS OF COOLING, HRS FIGURE 5.8.11

HOLTEC INTERNATIONAL DAEC SFP LOSS OF COOLING - CASE 4 (BOC FLL CORE OFFLOAD, TWO TRAINS)

LOSS OF COOLING 40.9 MAX. POOL WATER LEVEL 3.

UI 28.0- ADD MAKE LP WATEF TOP OF RACKS TOP OF ACTIVE FLEL 2s 40 TIME AFTER LOSS OF COOLING. HRS FIGURE 5.8.12

6.0 STRUCTURAL/SEISMIC CONSIDERATIONS 6.1 Introduction This section contains analyses to demonstrate structural adequacy of the high density spent fuel rack design under seismic loadings postulated for the plant spent fuel pool. Analyses and subsequent evaluations are in compliance with the requirements of the OT Position Paper,Section IV [6.1.1], and follow the USNRC Standard Review Plan (SRP) [6.1.2]. The dynamic analyses employ a time-history simulation code used in previous licensing efforts listed in Table 6.1.1. This section provides details of the method of analysis, modeling assumptions, numerical convergence studies and parametric evaluations performed to establish the required margins of safety.

Results reported herein show that the high density spent fuel racks are structurally and kinematically adequate to meet requirements defined in references [6.1.1], [6.1.2], and [6.1.3] with large margins of safety.

6.2 Background and Analysis Outline A spent fuel rack is a seismic category I structure [6.2.1].

Furthermore, it is a free-standing structure consisting of discrete storage cells which are loaded with free-standing fuel assemblies. As a result, the response of a rack module to seismic inputs is highly nonlinear involving a complex combination of motions (sliding, rocking, twisting, and turning), resulting in impacts and friction effects. Linear methods such as modal analysis and response spectrum techniques cannot accurately simulate the structural response of such a highly nonlinear structure to seismic excitation. A correct simulation is obtained only by direct integration of the nonlinear equations of motion using actual pool slab acceleration time-histories to provide the loading. Therefore, the initial step in spent fuel rack 6-1

qualification is to develop synthetic time-histories for three orthogonal directions which comply with the guidelines of USNRC SRP

[6.1.2]. In particular, the synthetic time-histories must meet the criteria of statistical independence and enveloping of the design response spectra.

As stated above, a free-standing spent fuel rack, subject to a seismic loading, executes nonlinear motions - even when isolated. The motion of an array of closely spaced racks in the spent fuel pool involves additional interactions due to fluid coupling between adjacent racks and between racks and adjacent walls. Further mechanical interactions between racks occur if rack-to-rack impacts take place during the event. To demonstrate structural qualification, it is required to show that stresses are within allowable limits and that displacements remain within the constraints of the contemplated design layout for the pool.

This implies that impacts between rack modules, if they occur, must be confined to locations engineered for this purpose, such as the baseplate edge for these fuel racks. Similarly, rack-to-pool wall impacts, if engineered into the rack design (not contemplated for these racks), must be within stipulated limits. Accurate and reliable assessment of the stress field and kinematic behavior of the rack modules calls for a comprehensive and conservative dynamic model which incorporates all key attributes of the actual structure. This means that the model must feature the ability to execute concurrent sliding, rocking, bending, twisting and other motion forms available to the rack modules. Furthermore, it must possess the capability to effect the momentum transfers which occur due to rattling of the fuel assemblies inside the storage cells and impacts of support pedestals on the bearing pads. Finally, the contribution of the water mass in the interstitial spaces around the rack modules and within storage cells must be modeled in an accurate manner because erring in the 6-2

quantification of fluid coupling on either side of the actual value is no guarantee of conservatism. Similarly, the Coulomb friction coefficient at the pedestal-to-pool liner (or bearing pad) interface may lie in a rather wide range and a conservative value of friction cannot be prescribed a' priori. In fact, a perusal of results of rack dynamic analyses in numerous dockets (Table 6.1.1) indicate that an upper bound value of the coefficient of friction, g, often maximizes the computed rack displacements as well as the equivalent elastostatic stresses. Further, the analysis must consider that a rack module may be fully or partially loaded with fuel assemblies or entirely empty.

The pattern of loading in a partially loaded rack may also have innumerable combinations. In short, there are a large number of parameters with potential influence on the rack motion. A comprehensive structural evaluation should deal with all of these without sacrificing conservatism.

The 3-D single rack dynamic model introduced by Holtec International in the Enrico Fermi Unit Two rack project (ca. 1980) and used in some twenty rerack projects since that time (Table 6.1.1) tackles the above mentioned array of parameters in a most appropriate manner. The details of this classical methodology are published in the permanent literature [6.2.2] and have been widely replicated by other industry groups in recent years. Briefly speaking, the single rack 3-D model handles the array of variables as follows:

Interface Coefficient of Friction Parametric runs are made with upper bound and lower bound values of the coefficient of friction. The limiting values are based on experimental data.

6-3

Impact Phenomena Compression-only gap elements are used to provide for opening and closing of interfaces such as the pedestal-to-bearing pad interface.

Fuel Loading Scenarios The fuel assemblies are conservatively assumed to rattle in unison which obviously exaggerates the contribution of impact against the cell wall. The different patterns of possible fuel assembly loadings in the rack are simulated by orienting the center of gravity column of the assemblage of fuel assemblies with respect of the module geometric centerline in an appropriate manner.

Fluid Coupling The contribution of fluid coupling forces is ascertained by prescribing the motion of the racks (adjacent to the one being analyzed). The most commonly used assumption is that the adjacent racks vibrate out-of-phase with respect to the rack being analyzed.

Despite the above simplifying assumptions, targeted for accuracy and conservatism, a large menu of cases is run to foster confidence in the calculated safety margins. Most of the safety analyses reported in the previous dockets (Table 6.1.1) over the past decade have relied on single rack 3-D model. From a conceptual standpoint, all aspects of the 3-D single rack model are satisfactory except for the fluid coupling effect. One intuitively expects the relative motion of the free-standing racks in the pool to be poorly correlated, given the random harmonics in the impressed slab motion. Single rack analyses cannot model this interactive behavior between racks. However, as described later, analytical and experimental research in this field has permitted rack analyses to be extended to all racks in the pool 6-4

simultaneously. Holtec International had successfully extended Fritz's classical two body fluid coupling model to multiple bodies and utilized it to perform the first two dimensional multi-rack analysis (Diablo Canyon, ca. 1987). Subsequently, laboratory experiments were conducted to validate the multi-rack fluid coupling theory. This technology was incorporated in the computer code DYNARACK which now could handle simultaneous simulation of all racks in the pool. This development marked a pivotal expansion in the rack structural modeling capability and was first utilized in Chin Shan, Oyster Creek and Shearon Harris plants [6.2.3]. The Whole Pool Multi-Rack (WPMR) 3-D analyses have corroborated the uncanny accuracy of the single rack 3-D solutions in predicting the maximum structural stresses. The multi rack analyses also serve to improve predictions of rack kinematics.

In order to ensure confidence in the results of structural safety analyses, results are presented for both single rack 3-D and WPMR 3-D analyses. The intent of this parallel approach is to foster added confidence and to uncover any peculiarities in the dynamic response which are germane to the structural safety of the storage system.

In the following, we summarize the sequence of model development and analysis steps that are undertaken. Subsequent subsections provide model detail, limiting criteria for stress and displacement, and results of the analyses.

a. Prepare three-dimensional dynamic models of individual fuel racks which embody all elastostatic characteristics and structural nonlinearities of the plant specific free standing rack modules.

b. Perform 3-D dynamic analyses on limiting module geometry types (from all those present in the spent fuel pool) and include various physical conditions (such as coefficient of friction, extent of cells containing fuel assemblies, and proximity of other racks).

6-5

c. Perform detailed stress analysis for the limiting case of all the dynamic analysis runs made in the foregoing steps.

Demonstrate compliance with ASME Code Section III, subsection NF [6.1.3] limits on stress and displacement.

d. Perform a degree-of-freedom (DOF) reduction procedure on the single rack 3-D model such that kinematic responses calculated by the Reduced DOF model (RDOFM) are in agreement with responses obtained using the baseline single rack models of step (b). The RDOFM is also truly three dimensional.

e. Prepare a whole pool multi-rack dynamic model which includes the RDOFM's of all rack modules in the pool, and includes all fluid coupling interactions among them, as well as fluid coupling interactions between racks and pool walls. This 3 D simulation is referred to as a WPMR model.

f. Perform 3-D WPMR analyses to demonstrate that all kinematic criteria for the spent fuel rack modules are satisfied, and that resultant structure loads confirm the validity of the structural qualification. The principal kinematic criteria are (i) no rack-to-pool wall impact, and (ii) no rack-to rack impact in the cellular region of the racks.

6.3 Artificial Time-Histories Section 3.7.1 of the SRP [6.1.2] provides guidelines for establishing seismic time-histories. Subsection 3.7.1.II.l.b gives applicable criteria for generation of time-histories from design response spectra.

A generated artificial time-history is acceptable if the response spectrum in the free field at the specified level of the site, obtained from the generated time-history, envelops the design response spectrum at the same location for all damping values used in the analysis.

The acceptance criterion for spectrum enveloping is that no more than five points of the spectrum obtained from the time-history fall below, 6-6

and no more than 10% below, the design response spectrum. The SRP states that an acceptable method of comparison is to choose a set of frequencies such that each frequency is within 10% of the previous one. The nature of the spent fuel rack structure is such that primary response is to excitations above 5-8 HZ. Within the 5-33HZ range, discrete check points are established from the above 10% frequency separation criterion.

Generated artificial time-histories must also be statistically independent. Any two time-histories are considered to be statistically independent if their normalized correlation coefficient is less than 0.15.

A set of seismic response spectra is provided for the Duane Arnold Energy Center (DAEC) for use in qualification analyses [6.3.1]. In accordance with [6.3.1], the requirements for generating time histories suitable for use in qualification of the fuel racks and spent fuel pool is as follows:

(1) According to Section 2.4.2b of Appendix D of [6.3.1], the appropriate OBE response spectrum for horizontal motions of the pool slab, which is at elevation 812', will be Curve C on Figure 6.3.1 (Figure D1 on Ref. [6.3.1], the same for two perpendicular directions in horizontal plane). The zero period acceleration (ZPA) is taken as 0.1125g for OBE.

(2) According to Section 2.5.2b of Appendix D of [6.3.1], the appropriate OBE response spectrum for vertical motion of the pool slab will be Curve F on Figure 6.3.2 multiplied by 0.8 (Figure D2 on Ref. [6.3.1]). The multiplier is 0.8 instead of 2/3. This is shown in the office memo IE-77-2194, dated December 6, 1977, which is accompanied with Ref. [6.3.1].

The ZPA should be 0.057 x 0.8 = 0.0456g.

(3) According to the asterisk notes in Sections 2.3.2b and 2.5.2b of Appendix D of [6.3.1], the DBE response spectra are obtained by doubling the corresponding OBE spectra.

6-7

In addition to the conservatism built-in the spectra, the set of OBE seismic acceleration time-histories developed by Holtec incorporates additional conservatism, as outlined below:

(1) The appropriate OBE response spectrum for horizontal motions is curve B on Figure 6.3.1 (Ref. [6.3.1] (the same for two perpendicular directions in horizontal plane)). The zero period acceleration (ZPA) is taken as 0.141 for the OBE.

(2) The DBE earthquake horizontal spectrum is obtained by doubling curve B on Figure 2.1 and using a zero period acceleration 0.282 for DBE.

(3) The appropriate vertical response spectrum curves for OBE and DBE are obtained from the corresponding OBE and DBE curves by multiplying them by 0.8.

The foregoing approach produces a considerably more severe synthetic acceleration set than that mandated by the plant spectra.

GENEQ [6.3.2] was used to generate three synthetic, statistically independent time-histories for two horizontal and the vertical directions, respectively, from the given design response spectrum.

Figures 6.3.3 - 6.3.5 show OBE time-history plots. Response spectra are re-generated and overlaid on the design spectra in Figures 6.3.6 6.3.8. The normalized correlation coefficients pj between time histories i and j are provided in Table 6.3.1 and demonstrate compliance with the statistical independence requirement for the OBE event (all pj 5 .15).

Comparing the developed set of time-histories with the requirements of generating time-histories suitable for use in qualification of fuel racks and spent fuel pool, it is obvious that the developed set of time-histories is more conservative because much higher response spectra in the three orthogonal directions were used to develop that set. For conservatism, the developed set of time-histories is used in the rack and pool analysis reported in this section.

6-8

As noted above, amplification factors of 2.0 can be applied to the time-history results to obtain suitable DBE events. 2% structural damping is associated with a DBE event.

6.4 Rack Modeling for Dynamic Simulations 6.4.1 General Remarks Spent fuel storage racks are Seismic Class I equipment. They are required to remain functional during and after a DBE event. The racks are free-standing; they are neither anchored to the pool floor nor attached to the sidewalls. Individual rack modules are not interconnected. Figure 6.4.1 shows a pictorial view of a typical module. The baseplate extends beyond the cellular region envelope ensuring that inter-rack impacts, if any, occur first at the baseplate elevation; this area is structurally qualifiable to withstand any large in-plane impact loads.

A rack may be completely loaded with fuel assemblies (which corresponds to greatest total mass), or it may be completely empty.

The coefficient of friction, A, between pedestal supports and pool floor is indeterminate. According to Rabinowicz [6.4.1], results of 199 tests performed on austenitic stainless steel plates submerged in water show a mean value of A to be 0.503 with standard deviation of 0.125. Upper and lower bounds (based on twice standard deviation) are 0.753 and 0.253, respectively. Analyses are therefore performed for coefficient of friction values of 0.2 (lower limit) and for 0.8 (upper limit), and for random friction values clustered about a mean of 0.5.

The bounding values of p = 0.2 and 0.8 have been found to bracket the upper limit of module response in previous rerack projects.

6-9

Since free-standing racks are not anchored to the pool slab, not attached to the pool walls, and not interconnected, they can execute a wide variety of motions. Racks may slide on the pool floor, one or more rack support pedestals may momentarily tip and lose contact with the floor slab liner, or racks may exhibit a combination of sliding and tipping. The structural models developed permit simulation of these kinematic events with inherent built-in conservatisms. The rack models also include components for simulation of potential inter-rack and rack-to-wall impact phenomena. Lift-off of support pedestals and subsequent liner impacts are modeled using impact (gap) elements, and Coulomb friction between rack and pool liner is simulated by piecewise linear (friction) elements. Rack elasticity, relative to the rack base, is included in the model with linear springs representing a beam like action. These special attributes of rack dynamics require strong emphasis on modeling of linear and nonlinear springs, dampers, and compression only gap elements. The term "nonlinear spring" is a generic term to denote the mathematical element representing the case where restoring force is not linearly proportional to displacement. In the fuel rack simulations, the Coulomb friction interface between rack support pedestal and liner is typical of a nonlinear spring.

3-D dynamic analyses of single rack modules require a key modeling assumption. This relates to location and relative motion of neighboring racks. The gap between a peripheral rack and adjacent pool wall is known, with motion of the wall prescribed. However, another rack, adjacent to the rack being analyzed, is also free standing and subject to motion during a seismic event. To conduct the seismic analysis of a given rack, its physical interface with neighboring modules must be specified. The standard procedure in analysis of a single rack module is that neighboring racks move 1800 out-of-phase in relation to the subject rack. Thus, the available gap before inter-rack impact occurs is 50% of the physical gap. This 6-10

"opposed phase motion" assumption increases likelihood of intra-rack impacts and is thus conservative. However, it also increases the relative contribution of fluid coupling, which depends on fluid gaps and relative movements of bodies, making overall conservatism a less certain assertion. 3-D WPMR analyses carried out in support of recently submitted rerack applications on numerous dockets (such as Zion, LaSalle Unit 1, Sequoyah, D.C. Cook, among others) indicate that single rack simulations predict smaller rack displacement during seismic responses. Nevertheless, 3-D analyses of single rack modules permit detailed evaluation of stress fields, and serve as a benchmark check for the much more involved, WPMR analysis.

Particulars of modeling details and assumptions for 3-D Single Rack analysis and for WPMR analysis are given in the following subsections.

6.4.2 The 3-D 22 DOF Model for Single Rack Module 6.4.2.1 Assumptions

a. The fuel rack structure is very rigid; motion is captured by modeling the rack as a twelve degree of-freedom structure. Movement of the rack cross-section at any height is described by six degrees-of-freedom of the rack base and six degrees-of-freedom at the rack top. Rattling fuel assemblies within the rack are modeled by five lumped masses located at H, .75H, .5H, .25H, and at the rack base (H is the rack height measured above the baseplate). Each lumped fuel mass has two horizontal displacement degrees-of-freedom.

Vertical motion of the fuel assembly mass is assumed equal to rack vertical motion at the baseplate level. The centroid of each fuel assembly mass can be located off center, relative to the rack structure centroid at that level, to simulate a partially loaded rack.

b. Seismic motion of a fuel rack is characterized by random rattling of fuel assemblies in their individual storage locations. All fuel assemblies are assumed to move in-phase within a rack. This exaggerates computed dynamic loading on the rack 6-11

structure and therefore yields conservative results.

c. Fluid coupling between rack and fuel assemblies, and between rack and wall, is simulated by appropriate inertial coupling in the system kinetic energy. Inclusion of these effects uses the methods of [6.4.2] and [6.4.3] for rack/assembly coupling and for rack-to-rack coupling, respectively. Fluid coupling terms for rack-to-rack coupling are based on opposed phase motion of adjacent modules.

d. Fluid damping and form drag is conservatively neglected.

e. Sloshing is negligible at the top of the rack and is neglected in the analysis of the rack.

f. Potential impacts between rack and fuel assemblies are accounted for by appropriate "compression only" gap elements between masses involved. The possible incidence of rack-to-wall or rack-to-rack impact is simulated by gap elements at top and bottom of the rack in two horizontal directions. Bottom elements are located at the baseplate elevation.

g. Pedestals are modeled by gap elements in the vertical direction and as "rigid links" for transferring horizontal stress. Each pedestal support is linked to the pool liner by two friction springs. Local pedestal spring stiffness accounts for floor elasticity and for local rack elasticity just above the pedestal.

h. Rattling of fuel assemblies inside the storage locations causes the gap between fuel assemblies and cell wall to change from a maximum of twice the nominal gap to a theoretical zero gap. Fluid coupling coefficients are based on the nominal gap.

6-12

6.4.2.2 Model Details Figure 6.4.2 shows a schematic of the model. S i = 1,.. .,4) represent support locations, pi represent absolute degrees-of-freedom, and qi represent degrees-of-freedom relative to the slab.

H is the height of the rack above the baseplate. Not shown in Figure 6.4.2 are gap elements used to model pedestal/liner impact locations and impact locations with adjacent racks.

Table 6.4.1 lists the degrees-of-freedom for the single rack model.

Translational and rotational degrees-of-freedom 1-6 and 17-22 describe the rack motion; rattling fuel masses (nodes 1, 2, 3, 4, 5 in Figure 6.4.2) are described by translational degrees-of-freedom 7-16. Ui(t) represents pool floor slab displacement seismic time-history.

Figures 6.4.3 and 6.4.4, respectively, show inter-rack impact springs (to track potential for impact between racks or between rack and wall), and fuel assembly/storage cell impact springs at one location of rattling fuel assembly mass.

Figures 6.4.5, 6.4.6, and 6.4.7 show the modeling technique and degrees-of-freedom associated with rack elasticity. In each bending plane a shear and bending spring simulate elastic effects rA.4.4].

Linear elastic springs coupling rack vertical and torsional degrees of-freedom are also included in the model.

Additional details concerning fluid coupling and determination of stiffness elements are provided below.

6.4.2.3 Fluid Coupling Details The "fluid coupling effect" [6.4.2],[6.4.3] is described as follows:

6-13

If one body (mass mi) vibrates adjacent to a second body (mass M 2 ), and both bodies are submerged in frictionless fluid, then Newton's equations of motion for the two bodies are:

if nf (MI + MI) X1 + M 12 X2 = applied forces on mass mi + 0 (X 2 )

M21 X1 + (M2 + M22 ) X2 = applied forces on mass M2 + 0 (X22 )

if it X 1, X2 denote absolute accelerations of masses mi and M21 respectively, and the notation 0(X2 ) denotes nonlinear terms.

M 11, M12, M21 , and M22 are fluid coupling coefficients which depend on body shape, relative disposition, etc. Fritz [6.4.3] gives data for Mj for various body shapes and arrangements. The fluid adds mass to the body (M11 to mass mi), and an external force proportional to acceleration of the adjacent body (mass M 2). Thus, acceleration of one body affects the force field on another. This force field is a function of interbody gap, reaching large values for small gaps.

Lateral motion of a fuel assembly inside a storage location encounters this effect. For example, fluid coupling is between nodes 2 and 2* in Figure 6.4.2. The rack analysis also contains inertial fluid coupling terms which model the effect of fluid in the gaps between adjacent racks. Terms modeling effects of fluid flowing between adjacent racks are computed assuming that all racks adjacent to the rack being analyzed are vibrating 1800 out of phase from the rack being analyzed.

Thus, the modeled rack is enclosed by a hydrodynamic mass computed as if there were a plane of symmetry located in the middle of the gap region. Rack-to-rack gap elements (Figure 6.4.3) have initial gaps set to 50% of the physical gap to reflect this symmetry.

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6.4.2.4 Stiffness Element Details The cartesian coordinate system associated with the rack has the following nomenclature:

x = Horizontal coordinate along the short direction of rack rectangular planform y = Horizontal coordinate along the long direction of the rack rectangular planform z = Vertical coordinate upward from the rack base Table 6.4.2 lists all spring elements used in the 3-D 22 DOF single rack model.

If the simulation model is restricted to two dimensions (one 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 6.4.8. This simpler model is used to elaborate on the various stiffness modeling elements.

Gap elements modeling impacts between fuel assemblies and rack have local stiffness K, in Figure 6.4.8. In Table 6.4.2, for example, gap elements 5 through 8 act on the rattling fuel mass at the rack top.

Support pedestal spring rates Ks are modeled by elements 1 through 4 in Table 6.4.2. Local compliance of the concrete floor is included in Ks. Friction elements 2 plus 8 and 4 plus 6 in Table 6.4.2 are shown in Figure 6.4.8. Friction at support/liner interface is modeled by the piecewise linear friction springs with suitably large stiffness Kfup to the limiting lateral load, gN, where N is the current compression load at the interface between support and liner. At every time step during transient analysis, the current value of N (either zero if the 6-15

pedestal has lifted-off the liner, or a compressive finite value) is computed. Finally, support rotational friction springs KR reflect any rotational restraint that may be offered by the foundation. The rotational friction spring rate is calculated using a modified Bousinesq equation [6.4.4] and is included to simulate resistive moment by the slab to counteract rotation of the rack pedestal in a vertical plane. The nonlinearity of these springs (friction elements 9, 11, 13, and 15 in Table 6.4.2) reflects the edging limitation imposed on the base of the rack support pedestals and the shift in location of slab resistive load as the rack pedestal rotates.

The gap element Ks, modeling the effective compression stiffness of the structure in the vicinity of the support, includes stiffness of the pedestal, local stiffness of the underlying pool slab, and local stiffness of the rack cellular structure above the pedestal.

The previous discussion is limited to a 2-D model solely for simplicity. Actual analyses incorporate 3-D motions and include all stiffness elements listed in Table 6.4.2.

6.4.3 Whole Pool Multi-Rack (WPMR) Model 6.4.3.1 General Remarks The single rack 3-D (22 DOF) model outlined in the preceding subsection is used to evaluate structural integrity, physical stability, and to initially assess kinematic compliance (no rack-to rack impact in the cellular region) of the rack modules. Prescribing the motion of the racks adjacent to the module being analyzed is an assumption in the single rack simulations. For closely spaced racks, demonstration of kinematic compliance is further confirmed by modeling 6-16

all modules in one comprehensive simulation using a WPMR model. In WPMR analysis, all racks are modeled, and their correct fluid interaction is included in the model.

6.4.3.2 Whole Pool Fluid Coupling The presence of fluid moving in the narrow gaps between racks and between racks and pool walls causes both near and far field fluid coupling effects. A single rack simulation can effectively include only hydrodynamic effects due to contiguous racks when a certain set of assumptions is used for the motion of contiguous racks. In a WPMR analysis, far field fluid coupling effects of all racks are accounted for using the correct model of pool fluid mechanics. The external hydrodynamic mass due to the presence of walls or adjacent racks is computed in a manner consistent with fundamental fluid mechanics principles [6.4.5] using conservative nominal fluid gaps in the pool at the beginning of the seismic event. Verification of the computed hydrodynamic effect by comparison with experiments is also provided in

[6.4.5]. This formulation has been reviewed and approved by the Nuclear Regulatory Commission during post-licensing multi-rack analyses for the Diablo Canyon Unit I and II reracking project. The fluid flow model used to obtain the whole pool hydrodynamic effect reflects actual gaps and rack locations.

6.4.3.3 Coefficients of Friction To eliminate the last significant element of uncertainty in rack dynamic analyses, the friction coefficient is ascribed to the support pedestal/pool bearing pad interface consistent with Rabinowicz's data

[6.4.1]. Friction coefficients, developed by a random number generator with Gaussian normal distribution characteristics, are imposed on each pedestal of each rack in the pool. The assigned values are then held 6-17

constant during the entire simulation in order to obtain reproducible results.* Thus, the WPMR analysis can simulate the effect of different coefficients of friction at adjacent rack pedestals. The friction coefficients at the interface between rack support pedestals and pool liner is assumed distributed randomly with a mean of 0.5 and permitted to vary between the limits of 0.2 - 0.8.

6.4.3.4 Modeling Details Figure 6.4.9 shows a planform view of the spent fuel pool after Campaign III of the reracking project, which includes rack and pedestal numbering scheme and the global coordinate system used for the WPMR analysis. Table 6.4.3 gives details on number of cells per rack, and on rack and fuel weights. In WPMR analysis, a reduced degree-of-freedom (RDOF) set is used to model each rack plus contained fuel. The rack structure is modeled by six degrees-of-freedom. A portion of contained fuel assemblies is assumed to rattle at the top of the rack, while the remainder of the contained fuel is assumed as a distributed mass attached to the rack. The rattling portion of the contained fuel is modeled by two horizontal degrees-of-freedom.

Thus, the WPMR model involves all racks in the spent fuel pool with each individual rack modeled as an 8 degree of freedom structure. The rattling portion of fuel mass, within each rack, is chosen to insure reasonable agreement between displacement predictions from single rack analysis using a 22 DOF model and predictions from 8 DOF analysis under the same conditions.

• Note that DYNARACK has the capability to change the coefficient of friction at any pedestal at each instant of contact based on a random reading of the PC-clock cycle. However, exercising this option would yield results that could not be reproduced. Therefore, the random choice of coefficients is made only once per run.

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The WPMR model includes gap elements representing compression only pedestals, representing impact potential at fuel assembly-fuel rack interfaces, and at rack-to-rack or rack-to-wall locations at top and bottom corners of each rack module. Each pedestal has two friction elements associated with force in the vertical compression element.

Values used for spring constants for the various stiffness elements are equal to the values used in the 22 DOF model.

6.5 Acceptance Criteria, Stress Limits, and Material Properties 6.5.1 Acceptance Criteria There are two sets of criteria to be satisfied by the rack modules:

a. Kinematic Criteria The rack must be a physically stable structure and it must be demonstrated that there are no inter-rack impacts in the cellular region. The criteria for physical stability is that an isolated rack in water exhibit no overturning tendency when a seismic event of magnitude 1.1 x DBE is applied [6.1.2].

b. Stress Limit Criteria Stress limits must not be exceeded under certain load combinations. The following loading combina tions are applicable [6.1.3].

Loading Combination Service Level D + L Level A D + L + To D + L + To + E D + L + Ta + E Level B D + L + T +Pf D + L + Ta + E' Level D D + L + Fd The functional capability of the fuel racks should be demonstrated.

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Abbreviations are those used in Section 3.8.4 of the Standard Review Plan and the "Review and Acceptance of Spent Fuel Storage and Handling Applications" section:

D = Dead weight-induced internal moments (including fuel assembly weight)

L Live Load (not applicable for the fuel rack, since there are no moving objects in the rack load path)

Fd = Force caused by the accidental drop of the heaviest load from the maximum possible height (See Section 7 of this report.)

P = Upward force on the racks caused by postulated stuck fuel assembly (see Section 7)

E = Operating Basis Earthquake (OBE)

E = Design Basis Earthquake (DBE)

To = Differential temperature induced loads (normal operating or shutdown condition based on the most critical transient or steady state condition)

Ta = Differential temperature induced loads (the highest temperature associated with the postulated abnormal design conditions)

Ta and To cause local thermal stresses to be produced. For fuel rack analysis, only one scenario need be examined. The worst situation is obtained when an isolated storage location has a fuel assembly generating heat at maximum postulated rate and surrounding storage locations contain no fuel. Heated water makes unobstructed contact with the inside of the storage walls, thereby producing maximum possible temperature difference between adjacent cells. Secondary stresses produced are limited to the body of the rack; that is, support pedestals do not experience secondary (thermal) stresses. For rack qualification, T, Ta are the same.

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6.5.2 Stress Limits for Various Conditions Stress limits are derived from the ASME Code, Section III, Subsection NF [6.1.3]. Parameters and terminology are in accordance with the ASME Code.

6.5.2.1 Normal and Upset Conditions (Level A or Level B)

a. Allowable stress in tension on a net section is:

Ft= 0.6 Sy (SY = yield stress at temperature)

(F, is equivalent to primary membrane stress)

b. Allowable stress in shear on a net section is:

F, = .4 Sy

c. Allowable stress in compression on a net section (kl)2 2

[1 - - /2Co S, r 2 Fa 5 k1 k1 3 3 J(-) + [3 (- ) /8Cj - [(- ) /8C c ]I 3 r r where:

(272 E) 1/2 C" [ I SY 1 = unsupported length of component k = length coefficient which gives influence of boundary conditions; e.g.

k = 1 (simple support both ends)

= 2 (cantilever beam)

= 1/2 (clamped at both ends) 6-21

E = Young's Modulus r = radius of gyration of component kl/r for the main rack body is based on the full height and cross section of the honeycomb region.

d. Maximum allowable bending stress at the outermost fiber of a net section, due to flexure about one plane of symmetry is:

Fb = 0.60 Sy (equivalent to primary bending)

e. Combined flexure and compression on a net section satisfies:

fa Cmx fbx myfby

- + + <1 Fa DxFbx DyFby where:

fa = Direct compressive stress in the section fbx= Maximum flexural stress along x-axis fby= Maximum flexural stress along y-axis Cx= Cmy = 0.85 f&

Dx = 1 - - -

F'ex f

Dy = 1 - - -

F F1ey

'e 12 ff2 E F F'xt ey kl 23 ( - )

r xly and subscripts x,y reflect the particular bending plane.

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f. Combined flexure and compression (or tension) on a net section:

f. fbx fby

- + -+ -<1.0

0. 6SY Fbx Fby The above requirements are to be met for both direct tension or compression.

6.5.2.2 Level D Service Limits Section F-1370 (ASME Section III, Appendix F), states that limits for the Level D condition are the minimum of 1.2 (SY/F,) or (0.7Su/F,)

times the corresponding limits for the Level A condition. Su is ultimate tensile stress at the specified rack design temperature. For example, if the material is such that 1.2S, is less than 0.7S,, then the multiplier on the Level A limits, to obtain Level D limits, is 2.0.

6.5.2.3 Dimensionless Stress Factors Stress results are presented in dimensionless form. Dimensionless stress factors are defined as the ratio of the actual developed stress to the specified limiting value. Stress factors are only developed for the single rack analyses. The limiting value of each stress factor is 1.0 for OBE and 2.0 (or less) for the DBE condition. Stress factors reported are R = Ratio of direct tensile or compressive stress on a net section to its allowable value (note pedestals only resist compression)

R2= Ratio of gross shear on a net section in the x direction to its allowable value 6-23

R = Ratio of maximum bending stress due to bending about the x-axis to its allowable value for the section R4= Ratio of maximum bending stress due to bending about the y-axis to its allowable value for the section R5 = Combined flexure and compressive factor (as defined in 6.5.2.1e above)

R = Combined flexure and tension (or compression) factor (as defined in 6.5.2.1f)

R7 = Ratio of gross shear on a net section in the y direction to its allowable value.

6.5.3 Material Properties Physical properties of the rack and support materials, obtained from the ASME Boiler & Pressure Vessel Code,Section III, appendices, are listed in Table 6.5.1. Maximum pool bulk temperature is less than 200'F; this is used as the reference design temperature for evaluation of material properties. Stress limits for Levels A and D, corresponding to conditions in Section 6.5.2 above, are evaluated using given yield strength data.

6.6 Governing Equations of Motion Using the structural model for either 22 DOF single rack analysis, or the set of simplified 8 DOF models that comprise a WPMR model, equations of motion corresponding to each degree of freedom are obtained using Lagrange's Formulation [6.6.1]. The system kinetic energy includes contributions from solid structures and from trapped and surrounding fluid. The final system of equations obtained have the matrix form:

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[M] {q"} = {Q} + {G}

where:

[M] - total mass matrix (including structural and fluid mass contributions)

{q} - the nodal displacement vector relative to the pool slab displacement; (double prime stands for second derivatives with respect to time)

{G} - a vector dependent on the given ground acceleration;

{Q} - a vector dependent on the spring forces (linear and nonlinear) and the coupling between degrees-of-freedom The equations can be rewritten as:

{q"} = [M]' {Q} + [M]-' {G}

This equation set is mass uncoupled, displacement coupled at each instant in time; numerical solution uses a central difference scheme built into the proprietary computer program "DYNARACK" [6.6.2 6.6.5]. As indicated earlier, this program has been used in the licensing effort for a considerable number of reracking projects.

DYNARACK has been validated against exact solutions, experimental data, and solutions obtained using alternate numerical schemes

[6.6.5]. These solutions are chosen to exercise all features of DYNARACK. It is demonstrated there that well known classical nonlinear phenomena (subharmonic resonance, bifurcation, stick-slip) can be reproduced using DYNARACK.

The application of DYNARACK to the spent fuel rack analysis requires the establishment of a time step to ensure convergence and stability of the results. DYNARACK utilizes the classical central difference algorithm [6.4.4]. Stability of the results is assured as long as 6-25

the time step is significantly below the smallest period of the equivalent linear problem. Convergence is obtained by performing a series of rack analyses with different time steps to ascertain the upper limit on time step that will provide converged results. This is done by taking a typical rack module and subjecting it to the given time-histories using different integration time steps. Once an appropriate time step is determined, it is used in subsequent simulations.

Results of the dynamic simulations are time-history response of all degrees-of-freedom of the particular model, and of all forces and moments at important sections of the structure. From these results, maximum movements and stresses can be ascertained for the event, and appropriate structural qualifications can be carried out. Where required, DYNARACK automatically tracks maximum values of dimensionless factors R, to R7 defined above in Section 6.5, and reports results for the rack cross section just above the baseplate and for each pedestal cross section just below the baseplate. These are the critical sections which develop the highest stresses due to the geometry of a fuel rack structure. From the archived results, time-histories of all rack-to-rack fluid gaps, all rack-to-wall fluid gaps, and motion of any point on any rack can be generated.

Sections 6.7 and 6.8 present results obtained from single and multi rack analyses, respectively. The results demonstrate satisfaction of all requirements on structure and kinematic integrity.

6.7 Results of 3-D Nonlinear Analyses of Single Racks This section focuses on results from all 3-D single rack analyses.

In the following section, we present results from the whole pool 6-26

multi-rack analysis and discuss the similarities and differences between single and multi-rack analysis.

From the list of racks given in Table 6.4.3, those chosen to be analyzed are rack G (the rack with maximum aspect ratio), rack J (the largest rack in the pool), and rack R (the rack in the cask pit). Altogether, 18 runs are carried out for governing cases using Holtec proprietary computer program DYNARACK. Results are abstracted from output files and presented here for the governing cases. Analyses have been carried out for regular fuel (680 lb. dry weight) and for opposed-phase motion assumption.

6.7.1 Racks in the Fuel Pool A summary of results of all analyses performed for racks in the pool and in the cask pit as well, using a single rack model, is presented in summary Tables 6.7.1-6.7.20. Table 6.7.1 lists all runs carried out. Table 6.7.2 presents the bounding results from all runs, and Tables 6.7.3-6.7.20 give details for each run. Tables 6.7.3-6.7.14 give results for racks in the fuel pool; Tables 6.7.15-6.7.20 present results for Rack R in the cask pit. The tabular results for each run give maximax (maximum in time and in space) values of stress factors at important locations in the rack. Results are given for maximum rack displacements (see Section 6.4.2.2 for x,y orientation), maximum impact forces at pedestal-liner interface, and rack cell-to-fuel, rack-to-rack, and rack-to-wall impact forces. It is shown that no rack-to-rack or rack-to-wall impacts occur in the cellular region of the racks.

In the single rack analysis, kinematic criteria are checked by confirming that no inter-rack gap elements at the top of the rack close (see Figure 6.4.9). By virtue of the symmetry assumption 6-27

discussed in subsection 6.4.2.4, impact is assumed to occur if the local horizontal displacement exceeds 50% of the actual rack-to-rack gap.

Structural integrity at various rack sections is considered by computing the appropriate stress factors Ri. Results corresponding to the SSE event yield the highest stress factors. Limiting stress factors for pedestals are at the upper section of the support and are to be compared with the bounding value of 1.0 (OBE) or 2.0 (DBE). Stress factors for the lower portion of the support are not limiting and are not reported. From Table 6.7.2, all stress factors are below the allowable limits.

Additional investigation of important structural items is carried out and results are summarized in Table 6.7.21. A discussion of these items follows:

6.7.1.1 Impact Analyses

a. Impact Loading Between Fuel Assembly and Cell Wall Local cell wall integrity is conservatively estimated from peak impact loads. Plastic analysis is used to obtain the limiting impact load. Table 6.7.21 gives the limiting impact load and compares the limit with the highest value obtained from any of the single rack analyses. The limiting load is much greater than the load obtained from any of the simulations reported in Tables 6.7.3-6.7.20.

b. Impacts Between Adlacent Racks No non-zero impact loads are found for the rack-to rack gap elements (in the cellular region), or for the rack-to-wall elements; it is concluded that no impacts between racks or between racks and walls are likely to occur during a seismic event. This is confirmed by the WPMR results in Section 6.8.

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6.7.1.2 Weld Stresses Weld locations subjected to significant seismic loading are at the bottom of the rack at the baseplate-to-cell connection, at the top of the pedestal support at the baseplate connection, and at cell-to cell connections. Results from dynamic analyses of single racks are surveyed and maximum loading used to qualify the welds.

a. Baseplate-to-Rack Cell Welds and Baseplate-to-Pedestal Welds Reference (6.1.3] (ASME Code Section III, Subsection NF) permits, for the SSE event, an allowable weld stress T =

.42 S,. A comparison of this allowable value with the highest weld stress predicted is given in Table 6.7.21.

The highest predicted weld stress is less than the allowable weld stress value.

The weld between baseplate and support pedestal is checked using limit analysis techniques [6.7.1]. The structural weld at that location is considered safe if the interaction curve between net force and moment is such that:

G = Function(F/FYM/My) < 1.0 FY, My are the limit load and moment under direct load only and direct moment only. These values depend on the configuration and on material yield strengths. F, M are absolute values of actual force and moments applied to the weld section. The calculated value of G for the pedestal/baseplate weld is presented in Table 6.7.21 and is less than the limit value of 1.0. This calculated value is conservatively based on instantaneous peak loading.

This value also conservatively neglects the gussets that are in place to increase pedestal area and inertia.

b. Cell-to-Cell Welds Cell-to-cell connections are by a series of spot welds along the cell height. Stresses in storage cell to storage cell welds develop along the length due to fuel assembly impact with the cell wall. This occurs if fuel assemblies 6-29

in adjacent cells are moving out of phase with one another so that impact loads in two adjacent cells are in opposite directions; this tends to separate the two cells from each other at the weld. Table 6.7.21 gives results for the maximum allowable load that can be transferred by these welds based on the available weld area. An upper bound on the load required to be transferred is also given in Table 6.7.21 and is less than the allowable load. This upper bound value is obtained by using the highest rack-to-fuel impact load from Table 6.7.2 (for any simulation), and multiplying the result by 2 (assuming that two impact locations are supported by every weld connection).

6.7.2 Rack in the Cask Pit Area The cask area of the fuel pool is a separate pit area with a 108" x 120" horizontal envelope. Analyses have been carried out for a 17x19 free-standing rack (Rack R) installed in the cask pit area.

To evaluate the rack in the cask pit, DYNARACK analysis is performed using fluid gaps between rack and cask pit wall that reflects the actual dimensions of the cask pit area, and the rack envelope. Runs were carried out for coefficients of friction of 0.2 and 0.8 and for different rack fuel loading scenarios. From all analyses performed for a spent fuel rack in the cask pit area, the bounding structural and kinematic results are given in Table 6.7.2. The details for each run are summarized in Tables 6.7.15 to 6.7.20.

It is noted from Table 6.7.2 that the 17x19 rack in the cask pit is the governing case for all the racks examined in the pool and in the cask pit for both structural and kinematic results.

6.8 Results from Whole Pool Multi-Rack (WPMR) Analyses Figure 6.4.9 shows the DAEC spent fuel pool with 18 new Holtec spent fuel racks. In the WPMR analysis, a reduced degree-of-freedom (8 DOF) model for each rack and its contained fuel is employed. The 6-30

WPMR dynamic model for DAEC contains 144 degrees-of-freedom and requires a nonlinear analysis. All racks are assumed to be fully loaded with 680-pound fuel assemblies. Thirty-percent of the fuel load is assumed to be rattling and impacting the rack top.

Table 6.8.1 shows maximum corner absolute displacements at both the top and bottom of each rack in global x and y directions (refer to Figure 6.4.9) from the multi-rack runs. As noted previously, a random set of friction coefficients in the range of 0.2 - 0.8 with mean value being 0.5 is used. The seismic loadings are the DBE earthquake time-histories which are the corresponding OBE time histories multiplied by a factor of 2.0 . Table 6.8.2 summarizes the maximum impact force of each gap spring. For each rack in the table, the first 4 or 5 springs are the pedestal vertical springs; the last 4 springs are the fuel-to-cell impact springs. The springs from No. 146 to No. 235 are rack-to-rack/wall impact springs at rack top, and the springs from No. 236 to No. 325 are rack-to-rack/wall impact springs at baseplate level. No non-zero values found for impact springs from No. 146 to No. 325 indicate that there is no impact between racks and between rack and pool wall during a DBE seismic event. Table 6.8.3 shows the maximum pedestal stress factors of all racks in the pool from the WPMR analysis. In Table 6.8.4, the maximum displacement, pedestal vertical loads, and pedestal stress factor obtained from the DBE multi-rack simulation are compared with the limiting single rack analyses. The absolute displacement values are higher than those obtained from single rack analysis. Thus, it appears essential to perform WPMR analyses to verify that racks do not impact or hit the wall. Figures 6.8.1-6.8.5 show the time histories of rack-to-rack and rack-to-wall gaps at typical locations (see Figure 6.4.9 for locations). A survey of all of the rack-to rack and rack-to-wall impact elements confirms that there are no rack-to-rack or rack-to-wall impacts in the cellular region of any 6-31

rack in the spent fuel pool. The inter-rack gap elements in the whole pool analysis have an initial gap equal to the actual gap.

Table 6.8.5 gives the results of dynamic pressures on the pool walls. Table 6.8.6 presents the total load and dynamic load adder on the whole pool slab. Figure 6.8.6 shows the time-history of the total slab vertical load. These data are used as loadings in qualifying the spent fuel pool structure.

The WPMR analyses confirms that no new concerns are identified; overall structural integrity conclusions are confirmed by both single and multi-rack analyses.

Because the values of all the stress factors obtained for DBE are less than 1.0 and no rack-to-rack/wall impacts are found, it is not necessary to perform the WPMR analysis for OBE seismic.

6.9 Bearing Pad Analysis To protect the slab from high localized dynamic loadings, bearing pads are placed between the pedestal base and the slab. Fuel rack pedestals impact on these bearing pads during a seismic event and pedestal loading is transferred to the liner. Bearing pad dimensions are set to ensure that the average pressure on the slab surface due to a static load plus a dynamic impact load does not exceed the American Concrete Institute [6.9.1] limit on bearing pressures.

Pedestal locations are set to avoid overloading of leak chase regions under the slab. Time-history results from dynamic simulations for each pedestal are used to generate appropriate static and dynamic pedestal loads which are then used to develop the bearing pad size.

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Section 10 of [6.9.1] gives the design bearing strength as fb = 0 (.85 fe') e where 4 = .7 and ff is the specified concrete strength for the spent fuel pool. E = 1 except when the supporting surface is wider on all sides than the loaded area. In that case, e = (A 2 /Al)-, but not more than 2. A, is the actual loaded area, and A2 is an area greater than A, and is defined in [6.9.1]. Using a value of e > 1 includes credit for the confining effect of the surrounding concrete.

Bearing pads are sized so as to provide sufficient margin on average bearing pressure. Table 6.9.1 summarizes the limiting result.

6.10 References for Section 6

[6.1.1] "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", dated April 14, 1978, and January 18, 1979 amendment thereto.

[6.1.2] USNRC Standard Review Plan, NUREG-0800 (1981).

[6.1.3] ASME Boiler & Pressure Vessel Code,Section III, Subsection NF, appendices (1989).

[6.2.1] USNRC Regulatory Guide 1.29, "Seismic Design Classification," Rev. 3, 1978.

[6.2.2] Soler, A.I. and Singh, K.P., "Seismic Responses of Free Standing Fuel Rack Constructions to 3-D Motions", Nuclear Engineering and Design, Vol. 80, pp. 315-329 (1984).

[6.2.3] Singh, K.P. and Soler, A.I., "Seismic Qualification of Free Standing Nuclear Fuel Storage Racks - the Chin Shan Experience, Nuclear Engineering International, UK (March 1991).

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[6.3.1] "Duane Arnold Energy Center Reactor Building Earthquake Analysis", J.A. Blume and Associates (1973).

[6.3.2] Holtec Proprietary Report - Verification and User's Manual for Computer Code GENEQ, Report HI-89364, January, 1990.

[6.4.1] Rabinowicz, E., "Friction Coefficients of Water Lubricated Stainless Steels for a Spent Fuel Rack Facility," MIT, a report for Boston Edison Company, 1976.

[6.4.2] Singh, K.P. and Soler, A.I., "Dynamic Coupling in a Closely Spaced Two-Body System Vibrating in Liquid Medium:

The Case of Fuel Racks," 3rd International Conference on Nuclear Power Safety, Keswick, England, May 1982.

[6.4.3] Fritz, R.J., "The Effects of Liquids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, Trans. of the ASME, February 1972, pp 167 172.

[6.4.4] Levy, S. and Wilkinson, J.P.D., "The Component Element Method in Dynamics with Application to Earthquake and Vehicle Engineering," McGraw Hill, 1976.

[6.4.5] Paul, B., "Fluid Coupling in Fuel Racks: Correlation of Theory and Experiment", Holtec Proprietary Report HI 88243.

[6.6.1] "Dynamics of Structures," R.W. Clough and J. Penzien, McGraw Hill (1975).

[6.6.2] Soler, A.I., "User Guide for PREDYNAl and DYNAMO", Holtec Proprietary Report HI-89343, Rev. 2, March, 1990.

[6.6.3] Soler, A.I., "Theoretical Backqround for Single and Multiple Rack Analysis", Holtec Proprietary Report HI 90439, Rev. 0, February, 1990.

[6.6.4] Soler, A.I., DYNARACK Theoretical Manual", Holtec Proprietary Report HI-87162, Rev. 1, January, 1988.

[6.6.5] Soler, A.I., "DYNARACK Validation Manual, Holtec Proprietary Report HI-91700, Rev. 0, October, 1991.

[6.7.1] Singh, K.P., Soler, A.I., and Bhattacharya, S., "Design Strength of Primary Structural Welds in Free Standing Structures", ASME, Journ. of Pressure Vessel Technology, August, 1991.

6-34

[6.9.1] ACI 349-85, Code Requirements for Nuclear Safety Related Concrete Structures, American Concrete Institute, Detroit, Michigan, 1985.

6-35

Table 6.1.1 LISTING OF PLANTS WHERE DYNARACK WAS APPLIED PLANT DOCKET NO.

Enrico Fermi Unit 2 USNRC 50-341 Quad Cities 1 and 2 USNRC 50-254, 50-265 Rancho Seco USNRC 50-312 Grand Gulf Unit 1 USNRC 50-416 Oyster Creek USNRC 50-219 Pilgrim Unit I USNRC 50-293 V.C. Summer USNRC 50-395 Diablo Canyon Units 1 and 2 USNRC 50-275, 50-323 Byron Units 1 & 2 USNRC 50-454, 50-455 Braidwood Units 1 & 2 USNRC 50-456, 50-457 Vogtle Unit 2 USNRC 50-425 St. Lucie Unit 1 USNRC 50-335 Millstone Point Unit 1 USNRC 50-245 D.C. Cook Units 1 & 2 USNRC 50-315, 50-316 Indian Point Unit 2 USNRC 50-247 Three Mile Island Unit 1 USNRC 50-289 J.A. FitzPatrick USNRC 50-333 Shearon Harris Unit 2 USNRC 50-401 Kuosheng Units 1 & 2 Taiwan Power Company Chin Shan Units 1 & 2 Taiwan Power Company Ulchin Unit 2 Korea Electric Power Laguna Verde Units 1 & 2 Comision Federal de Electricidad Zion Station Units 1 & 2 USNRC 50-295, 50-304 Sequoyah USNRC 50-327, 50-328 6-36

Table 6.3.1 CROSS-CORRELATION COEFFICIENTS VALUES OF pij Time-History Group OBE N-S and E-W (1,2) -. 1213 N-S and Vertical (1,3) -.01238 E-W and Vertical (2,3) -. 004586 6-37

Table 6.4. 1 DEGREES-OF-FREEDOM Displacement Rotation Location Ux UY UZ ex ey e (Node) 1 Pl P2 P3 q4 q5 q6 2 P17 P18 P19 q20 q21 q22 Point 2 is assumed attached to rigid rack at the top most point.

2 P7 P8 3 P9 Pl0 4 P11 P12 5 P13 P14 1 P15 P16 where:

Pi = qi(t) + Ul(t) i = 1,7,9,11,13,15,17

- qi(t) + U 2 (t) i = 2,8,10,12,14,16,18

= qi(t) + U 3 (t) i = 3,19 Ui(t) are the 3 known earthquake displacements.

6-38

Table 6.4.2 NUMBERING SYSTEM FOR GAP ELEMENTS AND FRICTION ELEMENTS I. Nonlinear Springs (Gap Elements) (64 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 4 Support S4 Z compression only element 5 2,2* X rack/fuel assembly impact element 6 2,2* X rack/fuel assembly impact element 7 2,2* Y rack/fuel assembly impact element 8 2,2* Y rack/fuel assembly impact element 9-24 Other rattling masses for nodes 1*, 3*, 4* and 5*

25 Bottom cross Inter-rack impact elements section of rack (around edge)

Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements 44 Inter-rack impact elements 45 Top cross-section Inter-rack impact elements of rack Inter-rack impact elements (around edge) Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements Inter-rack impact elements 64 Inter-rack impact elements 6-39

Table 6.4.2 (continued)

NUMBERING SYSTEM FOR GAP ELEMENTS AND FRICTION ELEMENTS II. Friction Elements (16 total)

Number Node Location Description 1 Support S1 X direction friction 2 Support S1 Y direction friction 3 Support S2 X direction friction 4 Support S2 Y direction friction 5 Support S3 X direction friction 6 Support S3 Y direction friction 7 Support S4 X direction friction 8 Support S4 Y direction friction 9 S1 X Slab moment 10 S1 Y Slab moment 11 S2 X Slab moment 12 S2 Y Slab moment 13 S3 X Slab moment 14 S3 Y Slab moment 15 S4 X Slab moment 16 S4 Y Slab moment 6-40

Table 6.4.3 SPENT FUEL POOL LOADING Fuel Cell Configuration Rack Assembly (No. of Cells in Weight Weight Rack NS x EW Direction) (lbs) (lbs)

Campaign I:

Al 12x12 10300 680 A2 12x12 10300 680 A3 12x12 10300 680 BI 12x10 8600 680 B2 12x10 8600 680 C 12x12-1x10 9600 680 D 14x12 (dual 14000 680 purpose)

E 14x10-4x7 8000 680 F 14xl2-4x1. 11700 680 Campaign II:

G 11x21 16500 680 E 12x21 18000 680 J 14x21-4x6 19300 630 Campaign III:

K 11xI0 7900 680 L 12x10 8600 680 M 14x10 10000 680 N1 12x12 10300 680 N2 12x12 10300 680 P 14x12 12000 680 R (in cask pit) 19x17 23200 680 6-41

Table 6.5.1 RACK MATERIAL DATA (200aF)

Young's Yield Ultimate Modulus Strength Strength Material E (psi) Sy (psi) Su (psi) 304 S.S. 27.6 x 106 25000 71000 ASME Section III Table Table Table Reference 1-6.0 1-2.2 1-3.2 SUPPORT MATERIAL DATA Young's Yield Ultimate Modulus Strength Strength Material E (psi) Sy .(psi) Su (psi) 1 ASTM-240, 27.6x10 6 25,000 71,000 Type 304 psi psi psi (upper part of support feet) 2 ASTM 564-630, 27.6x10 6 106,300 140,000 (lower part of psi psi psi support feet; age hardened at 1100aF)

ASME Section III Table Table Reference 1-2.1 1-3.1 6-42

Table 6.7.1 RESULTS OF SINGLE RACK ANALYSES List of All Runs Holtec Rack Fuel Fuel Loading Seismic Coefficient Motion Run I.D. I.D. I.D. Condition Loading of Friction Mode drgsseo.rf8 G regular Fully Loaded SSE(=2*OBE) 0.8 opp-phase

.Region-2 680# 231 cells drgsseo.rf2 G regular Fully Loaded SSE(=2*OBE) 0.2 opp-phase Region-2 680# 231 cells drgsseo.rh8 G regular Half Loaded SSE(=2*OBE) 0.8 opp-phase Region-2 680# 110 cells drgsseo.rh2 G regular Half Loaded SSE(=2*OBE) 0.2 opp-phase Region-2 680# 110 cells drgsseo.re8 G regular Symmetrically SSE(=2*OBE) 0.8 opp-phase Region-2 680# 22 cells drgsseo.re2 G regular Symmetrically SSE(=2*OBE) 0.2 opp-phase Region-2 680# 22 cells 6-43

( to be continued )

Table 6.7.1 ( continued )

Holtec Rack Fuel Fuel Loading Seismic Coefficient Motion Run I.D. I.D. I.D. Condition Loading of Friction Mode drjsseo.rf8 J regular Fully Loaded SSE(=2*OBE) 0.8 opp-phase Region-2 680# 270 cells drjsseo.rf2 J regular Fully Loaded SSE(=2*OBE) 0.2 opp-phase Region-2 680# 270 cells drjsseo.rh8 J regular Half Loaded SSE(=2*OBE) 0.8 opp-phase Region-2 680# 140 cells drjsseo.rh2 J regular Half Loaded SSE(=2*OBE) 0.2 opp-phase Region-2 680# 140 cells drjsseo.re8 J regular Symmetrically SSE(=2*OBE) 0.8 opp-phase Region-2 680# 24 cells drjsseo.re2 J regular Symmetrically SSE(=2*OBE) 0.2 opp-phase Region-2 680# 24 cells

( to be continued )

6-44

Table 6.7.1 ( continued )

Holtec Rack Fuel Fuel Loading Seismic Coefficient Motion Run I.D. I.D. I.D. Condition Loading of Friction Mode drrsse.rf8 R regular Fully Loaded SSE(=2*OBE) 0.8 in cask Region-2 680# 323 cells pit drrsse.rf2 R regular Fully Loaded SSE(=2*OBE) 0.2 in cask Region-2 680# 323 cells pit drrsse.rh8 R -regular Half Loaded SSE(=2*OBE) 0.8 in cask Region-2 680# 153 cells pit drrsse.rh2 R regular Half Loaded SSE(=2*OBE) 0.2 in cask Region-2 680# 153 cells pit drrsse.re8 R regular Symmetrically SSE(=2*OBE) 0.8 in cask Region-2 680# 34 cells pit drrsse.re2 R regular Symmetrically SSE(=2*OBE) 0.2 in cask Region-2 680# 34 cells pit 6-45

Table 6.7.2

SUMMARY

OF WORST RESULTS FROM 18 RUNS OF SINGLE RACK ANALYSIS LOADED WITH 680# REGULAR FUEL ASSEMBLIES; SSE EARTHQUAKE TIME-HISTORIES )

Item Value Run I.D.

1. Maximum total vertical pedestal load: 337,949 lbs. drrsse.rf2

2. Maximum vertical load in any single pedestal: 115,152 lbs. drrsse.rf2

3. Maximum shear load in any single pedestal: 16,759 lbs. drrsse.rf2

4. Maximum fuel assembly-to-cell wall impact load at one local position: 240 lbs. drrsse.re2

5. Maximum rack-to-wall impact load at baseplat level: 0 lbs.

6. Maximum rack-to-wall impact load at the top of rack: 0 lbs.

7. Maximum rack-to-rack impact load at baseplat level: 0 lbs.

3. Maximum rack-to-rack impact load at the top of rack: 0 lbs.

9. Maximum corner displacements Top corner in x direction: 0.0928 in. drrsse.rf2 in y direction: 0.1477 in. drrsse.rh2 Baseplate corner in x direction: 0.0055 in. drrsse.rf2 in y direction: 0.0096 in. drrsse.rf2

10. Maximum stress factors Above baseplate: 0.111 (R6) drrsse.rf2 Support pedestals: 0.469 (R6) drrsse.rf 2 6-46

Table 6.7.3

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.rf8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 231 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynam0/dynamo.fav $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 22 8922.9 (2) Maximum vertical load in any single pedestal: 6 9890.0 (3) Maximum shear load in any single pedestal: 8584.1 (4) Maximum fuel-cell impact at one local position: 183.3 (5) Maximum rack-to-wall impact at baseplata: .0 (6) Maximum rack-to-wall imoact at rack ton: .0 (7) Maximum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0290 .0592 Baseplate corner: .0008 .0015 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .028 .016 .059 .020 .077 .086 .018 Support pedestal: .256 .037 .039 .034 .280 .287 .043

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-47

Table 6.7.4

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.rf2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 231 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynam0/dynamo.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynam0/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamo/dynas2. fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 22 8926.8 (2) Maximum vertical load in any single pedestal: 6 9873.6 (3) Maximum shnear load in any single pedestal: 2032.2 (4) Maximum fuel-cell impact at one local position: 183.3 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate: -0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0294 .0619 Baseplate corner: .0013 .0032 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .028 .015 .062 .020 .075 .084 .022 Support pedestal: .256 .046 .055 .042 .298 .311 .061

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-48

Table 6.7.5

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.rh8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 110 cells loaded; Fuel centroid X,Y: .0,-33.4 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo.fov $

$Revision: 2.5 $

$Loqfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 6341.5 (2) Maximum vertical load in any single pedestal: 4 2054.5 (3) Maximum shear load in any single pedestal: 5418.7 (4) Maximum fuel-cell impact at one local position: 167.4 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Too corner: .0218 .0571 Baseplate corner: .0004 .0006 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .016 .009 .024 .010 .032 .036 .009 Support pedestal: .154 .025 .022 .023 .179 .185 .025

• See Section 6.5.2.3 of the Licensing Report for definitions.

,6-49

Table 6.7.6

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.rh2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 110 cells loaded; Fuel centroid X,Y: .0,-33.4 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 1 6341.5 (2) Maximum vertical load in any single pedestal: 4 2315. 3 (3) Maximum shear load in any single pedestal: 7349.5 (4) Maximum fuel-call impact at one local position: 171.4 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maxi mum rack-to-rack imDact at basenlate:

(8) Maximum rack-tc-rack impact at rack top: .0 MAXIMUM'CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0212 .0571 Baselate corner: .0007 .0011 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .016 .007 .024 .010 .033 .036 .010 Support pedestal: .155 .026 .035 .023 .191 .198 .039

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-50

Table 6.7.7

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.re8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 22 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamo.fov $

$':-evision: 2.5 $

$Logfile: C:/racks/dynamO/dynasi.fov $

$Ravision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 3 7224.8 (2) Maximum vertical load in any single pedestal: 1 1872.2 (3) Maximum shear icad in any single pedestal: 1623.0 (4) Maximum fuel-cell impact at one local position: 175.7 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate:

(8) Maximum rack-to-rack immact at rack tot: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0083 .0122 Baseplate corner: .0001 .0003 MAXIMUM STRESS FACTORS

• Stress factor: RI R2 R3 R4 R5 R6 R7 Above baseplate: .008 .003 .012 .004 .017 .019 .003 Support pedestal: .043 .007 .008 .006 .048 .049 .009

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-51

Table 6.7.8

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: G-11x21 Holtec Run I.D.: drgsseo.re2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 22 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo.fov $

$Revision: 2.5 $

$Lcgfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 3 7234.9 (2) Maximum vertical load in any single pedestal: 1 1873.7 (3) Max mum shear load in any single pedestal: 1852.6 (4) Maximum fuel-cell impact at one local position: 175.7 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall imoact at rack too: .0 (7) Maximum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0084 .0122 Baseplate corner: .0003 .0004 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .008 .002 .012 .004 .017 .019 .004 Support pedestal: .043 .010 .009 .009 .051 .053 .010

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-52

Table 6.7.9

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.rf8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 270 cells loaded; Fuel centroid X,Y: -2.7, -4.1 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamO.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 S

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 267926.5 (2) Maximum vertical load in any single pedestal: 73061.0 (3) Maximum shear load in any single pedestal: 9305.2 (4) Maximum fuel-cell impact at one local position: 173.4 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximi -n -. - impact at baplate: .0 (3) Maximum rack-to-rack impact at rack tor: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Locaticn: X-direction Y-direction Top corner: .0505 .0573 Baseplate corner: .0010 .0013 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .026 .019 .040 .021 .054 .061 .016 Support pedestal: .268 .048 .037 .044 .302 .308 .040

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-53

Table 6.7.10

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.rf2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 270 cells loaded; Fuel centroid X,Y: -2.7, -4.1 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamO.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 26 7914.7 (2) Maximum vertical load in any single pedestal: 7 3034.0 (3) Maximum shear load in any single pedestal: 1579.7 (4) Maximum fuel-cell impact at one local position: 130.4 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall imract at rack toe: .0 (7) Maximum rack-to-rack impact at baseplat-e: .0 (8) Maximum rack-tc-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0485 .0635 Baseplate corner: .0012 .0016 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .026 .018 .045 .021 .058 .065 .017 Support pedestal: .268 .053 .055 .048 .329 .340 .061

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-54

Table 6.7.11

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.rh8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 140 cells loaded; Fuel centroid X,Y: .0,-33.4 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamO.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/rAcks/dvnam0/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total verzical pedestal load: 14 6142.0 (2) Maximum vertical load in any single pedestal 5 2344.9 (3) Maximum shear load in any single pedestal: 3805.8 (4) Maximum fuel-cell impact at one local position: 177.6 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate: .0 (8) Max imum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0182 .0774 Baseplate corner: .0005 .0008 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .015 .010 .026 .009 .036 .040 .010 Support pedestal: .191 .035 .031 .032 .214 .218 .035

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-55

Table 6.7.12 SUMMIARY RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.rh2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 630.0 (lbs.)

Fuel Loading: 140 cells loaded; Fuel centroid X,Y: .0,-33.4 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 S

$Logfile: C:/racks/dynamo/dynamo.fov S

$Revision: 2.5 $

SLogfile: C:/racks/dynamO/dvnasl.fov $

$Revision: 3.36 $

$Lcgfile: C:/racks/dynamo/dynas2.fov 5 DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal 'cad: 14 6324.7 (2) Maximum vertical load in any single pedestal: 5 2323.3 (3) Maximum shear load in any single pedestal: 3743.5 (4) Maximum fuel-cell impact at one lcca positin: 167.0 (5) Maximum rack-to-wall impact at baserlate: .0 (6) Maximum rack-to-wall impact at rack t: .0 (7) Maximum rack-to-rack 1mcact at baseclata: .0 (3) Maximum rack-to-rack imDact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top ccrner: .0163 .0773 Baseplate corner: .0011 .0011 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .015 .009 .026 .008 .035 .040 .010 Support pedestal: .191 .044 .041 .040 .227 .234 .046

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-56

Table 6.7.13

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.re8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 24 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo. fov $

$Revision: 2.5 $

$Lcgfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 4 2028.5 (2) Maximum vertical load in any single pedestal: 1 0707.2 (3) Maximum shear load in any single pedestal: 1822.6 (4) Maximum fuel-cell impact at one local position: 186.6 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wa1 1 impact at rack top: .0 (7) Max imum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0140 .0103 Baseplate corner: .0002 .0002 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .007 .003 .006 .004 .013 .014 .003 Support pedestal: .039 .009 .006 .008 .043 .044 .007

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-57

Table 6.7.14

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: J-14x21 Holtec Run I.D.: drjsseo.re2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 24 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo.fav $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 4 2028.3 (2) Maximum vertical load in any single pedestal: 1 0695.9 (3) Maximum shear load in any single pedestal: 163 5.1 (4) Maximum fuel-cell impact at one local position: 182.0 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall imnact at rack too: .0 (7) Maximum rack-to-rack im-act at baselate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0150 .0104 Baseplate corner: .0003 .0003 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .007 .003 .006 .003 .013 .014 .002 Support pedestal: .039 .008 .007 .007 .044 .045 .008

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-58

Table 6.7.15

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.rf8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 323 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamO.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C: /racks/dynamo/dynas2.fov $

DYNAMIC IMPACT LOADS (1bs.)

(1) Maximum total vertical pedestal load: 33 7948.9 (2) Maximum vertical load in any single pedestal: 10 7940.0 (3) Maximum snear load in any single pedestal: 1 1650. 6 (4) Maximum -fuel-cell impact at one local position: 168.8 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum racx-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate! .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0891 .1223 Baseplate corner: .0027 .0036 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .033 .015 .061 .038 .083 .093 .017 Support pedestal: .395 .049 .050 .044 .435 .443 .055

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-59

Table 6.7.16

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.rf2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 323 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynamo.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamo/dynasi.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 33 7948.9 (2) Maximum vertical load in any single pedestal: 11 5152.3 (3) Maximum shear load in any single pedestal: 1 6759.0 (4) Maximum fuel-cell impact at one local position: 177.6 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maxi mum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0928 .1215 Baseplate cc rner: .0055 .0096 MAXIMUM STRESS FACTORS

• Stress facto r: R1 R2 R3 R4 R5 R6 R7 Above basepl ate: .033 .015 .060 .038 .099 .111 .020 Support pede stal: .421 .061 .082 .055 .459 .469 .091

• See Secti on 6.5.2.3 of the Licensing Report for definiti ons.

6-60

Table 6.7.17

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.rh8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 153 cells loaded; Fuel centroid X,Y: .0,-30.4 (in.)

Ccefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Lcgfile: C:/racks/dynamo/dynamO.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dvnamO/dynas1.fov $

$Revision: 3.36 5

$Logfile: C:/racks/dvnamO/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 16 6620.1 (2) Maximum vertical load in any single pedestal: 7 7494.7 (2) Maximum shear lad in any single pedestal: 1 4663.5 (4) Maximum fuel-cell impact at one local position: 156.7 (5) Maximum rack-tc-wall impact at baseplate: .0 (6) Maximum rack-tc-wall impact at rack top: .0 (7 Maximum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0751 .1413 Baseplate corner: .0022 .0016 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .017 .008 .028 .030 .056 .063 .009 Support pedestal: .283 .041 .071 .037 .305 .309 .079

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-61

Table 6.7.18

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.rh2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 153 cells loaded; Fuel centroid X,Y: .0,-30.4 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamo.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamo/dvnas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 16 6622.2 (2) Maximum vertical load in any single pedestal: 7 7380.0 (3) Maximum shear lead in any sincle pedestal: 1 3084.5 (4) Maximum fuel-cell impact at one local position: 174.9 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall immact at rack ton: .0 (7) Max imu rack-to-rkitact at basemlate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0668 .1477 Baseplate corner: .0023 .0043 MAXIMUM STRESS FACTORS

• Stress factor: RI R2 R3 R4 R5 R6 R7 Above baseplate: .017 .008 .030 .027 .056 .063 .011 Support pedestal: .282 .049 .064 .044 .315 .325 .071

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-62

Table 6.7.19

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.re8 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 34 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.8

$Revision: 3.46 $

$Logfile: C:/racks/dynamO/dynamo.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamO/dynas1.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamo/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load: 57107.5 (2) Maximum vertical load in any single pedestal: 19607.9 (3) Maximum shear load in any single pedestal: 2186.8 (4) Maximum fuel-cell impact at one local position: 212.8 (5) Maximum rack-to-wall impact at baseplate: .0 (6) Maximum rack-to-wall impact at rack top: .0 (7) Maximum rack-to-rack impact at baseplate: .0 (8) Maximum rack-to-rack impact at rack top: .0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0257 .0183 Baseplate corner: .0008 .0005 MAXIMUM STRESS FACTORS

• Stress factor: R1 R2 R3 R4 R5 R6 R7 Above baseplate: .009 .003 .009 .011 .020 .023 .003 Support pedestal: .072 .011 .010 .010 .083 .085 .011

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-63

Table 6.7.20

SUMMARY

RESULTS OF 3-D SINGLE RACK ANALYSIS FOR RACK MODULE: R-17x19 Holtec Run I.D.: drrsse.re2 Seismic Loading: SSE(2xOBE)

Fuel Assembly I.D. and Weight: 680#reg.  ; 680.0 (lbs.)

Fuel Loading: 34 cells loaded; Fuel centroid X,Y: .0, .0 (in.)

Coefficient of friction at the bottom of support pedestal: 0.2

$Revision: 3.46 $

$Logfile: C:/racks/dynamo/dynam0.fov $

$Revision: 2.5 $

$Logfile: C:/racks/dynamo/dynasl.fov $

$Revision: 3.36 $

$Logfile: C:/racks/dynamo/dynas2.fov $

DYNAMIC IMPACT LOADS (lbs.)

(1) Maximum total vertical pedestal load:

57107.5 (2) Maximum vertical load in any single pedestal: 19596.8 (3) Maximum shear load in any single pedestal:

3024.5 (4) Maximum fuel-cell impact at one local position:

240.1 (5) Maximum rack-to-wall impact at baseplate:

.0 (6) Maximum rack-to-wall impact at rack top:

.0 (7) Maximum rack-to-rack impact at baseplate: .0 (3) Maximum rack-to-rack impact at rack top:

.0 MAXIMUM CORNER DISPLACEMENTS (in.)

Location: X-direction Y-direction Top corner: .0256 .0202 Baseplate corner: .0010 .0013 MAXIMUM STRESS FACTORS

• Stress factor: RI R2 R3 R4 R5 R6 R7 Above baseplate: .009 .003 .010 .010 .020 .023 .003 Support pedestal: .072 .015 .014 .014 .087 .089 .015

• See Section 6.5.2.3 of the Licensing Report for definitions.

6-64

Table 6.7.21 COMPARISON OF CALCULATED AND ALLOWABLE LOADS/STRESSES AT IMPACT LOCATIONS AND AT WELDS VALUE Item/Location Calculated Allowable Fuel assembly/ 610.9 2643 cell wall impact, lbs.

Rack/Baseplate weld 9277 29820 psi Pedestal/Baseplate .283 1.0 weld (dimensionless limit load ratio)

Cell/Cell welds 1222 5271 6-65

Table 6.8.1 MAXIMUM DISPLACEMENTS FROM WHOLE POOL MULTI RACK RUNS DUANE ARNOLD PLANT, IOWA ELECTRIC LIGHT AND POWER CO.

WPMR Analysis, 18 Holtec Racks in Pool, Fully Loaded with 680# Reg.Fuel; Random Friction; Seismic: DSE ( =OBEx2.0 ); dt=0.00005 sec.

Rack uxt uyt uxb uyb No. (in.) (in.) (in.) (in.)

1 .2420E+0 .2685E+00 .9074E-02 .1247E-01 2 .1353E+00 .2421E+00 .5904E-02 .8088E-02 3 .1774E+00 .1840E+00 .2153E-01 .2292E-01 4 .1824E+00 .1623E+00 .5806E-02 .5116E-02 5 .1904E+00 .1858E+00 .6006E-02 .5906E-02 6 .2770E00 .2420E+00 .9019E-02 .8036E-02 7 .23711+0.0 .2170E+00 .9101-02 .7007E-02 8 .2038E+00 .1975E+00 .68132-02 .6609E-02 9 .1650E+00 .1334E+00 .8933E-02 .1097E-01 10 .1872E+00 .1498E+00 .9188E-02 .7514E-02 11 .1539E00 .1685E+00 .4907E-02 .5301E-02 12 .1549E+00 .2382E+00 .8102E-02 .8195E-02 13 .1932E+00 .2335E+00 .1202E-01 .8841E-02 14 .2463E-00 .2273E+00 .8342E-02 .7257E-02 15 .1505E+00 .1976E+00 .4703-02 .6413E-02 16 .1807E+00 .1433E+00 .5708E-02 .4509E-02 17 .4086E-00 .1570E+00 .1462E-01 .5349E-02 18 .2819E+00 .2126E+00 .6593E-01 .6871E-01

$Revision: 1.8 $

$Logfile: C: /racks/multirac/maxdisp.fov 6-66

Table 6.8.2 MAXIMUM IMPACT FORCE OF EACH GAP SPRING DUANE ARNOLD PLANT, IOWA ELECTRIC LIGHT AND POWER CO.

WPMR Analysis, 18 Holtec Racks in Pool, Fully Loaded with 6807 Reg.Fuel; Random Friction; Seismic: DBE ( =OBEx2.0 ); dt=0.00005 sec.

GAP ELEMENT MAX.FORCE TIME RACK-1:

1 6. 428E+04 1.069E+01 2 6.672E+04 1.013E+01 3 5.156E+04 1.377E+01 4 6.626E+04 5.817E+00 5 2.974E+04 5.733E+00 6 3.092E+04 4.332E+00 7 2.729E+04 7. 365E+00 8 3.630E+04 7.635E+00 RACK-2:

9 4.660E+04 9.333E+00 10 5.748E+04 6.554E+00

1. 4.213E+04 9.917E+00 12 4. 550E+04 1. 154E+01 13 2.124E-04 4.852E+00 14 1.768E+04 4.529E+00 15 1.923E-04 9.553E+00 16 2.431E+04 6. 714E+00 RACK-3:

17 9.580E+04 5.554E+00 13 8.633E+04 1. 094E+01 19 7.940E+04 8.342E+00 20 8.726E+04 1.114E+01 21 2.910E+04 2.333E+00 22 3.333E+04 2.683E+00 23 4.870E+04 1.188E+01 24 4.533E1+04 7.303E+00 RACK-4:

25 6.237E+04 4.648E+00 26 5.722E+04 5.806E+00 27 5.202E+04 9.015E+00 28 5.416E+04 1.289E+01 29 22.189E+04 7.210E+00 30 1.908E+04 7. 552E+00 31 1.911E+04 1.088E+01 32 2.631E+04 1.052E+01

( to be continued )

6-67

continued RACK-5:

33 4. 942E+04 8. 529E+00 34 5. 120E+04 6. 739E+00 35 5. 730E+04 1. 173E+01 36 5. 844E+04 7. 202E+00 37 1. 705E+04 1. 025E+01 38 1.557E+04 1. 066E+01 39 1. 070E+04 9. 759E+00 40 1. 278E+04 1. 025E+01 RACK-5:

41 7. 097E+04 5. 477E+00 42 6. 154E+04 1. 196E+01 43 6. 048E+04 5. 253E+00 44 6. 247E+04 8. 038E+00 45 2. 680E+04 1. 188E+01 46 2. 155E+04 1. 138E+01 47 3.001E+04 9.836E+00 48 2.922E+04 1. 01.5E--01 RACK-7:

49 6. 845E+04 5.516E+00 50 5. 336E+04 5. 250E-00 51 8.081E+04 4.242E+00 52 6. 480E+04 9.877E+00 53 1. 935E+04 5. 754E+00 54 2.158E+04 1. 233E+01 55 2.950E+04 7.078E+00 56 3. 016E+04 9.338E+00 RACX-8:

57 5. 028E+04 5.505E+00 58 6. 155E+04 1.176E+01U 59 4. 497E+04 9.879E+00 60 4.263E+04 9.702E+00 61 1. 669E+04 1.266E+01.

62 1. 931E+04 7.833E+00 63 1. 321E+04 1.043E+01 64 1.511E+04 1.005E+01 RACK-9:

65 9. 915E+04 9.431E+00 66 8. 826E+04 6.741E+00 67 7. 134E+04 1.069E+01 68 8. 539E+04 5.569E+00 69 2. 737E+04 4.817E+00 70 3. 832E+04 4.360E+00 71 4. 142E+04 1.207E+01 72 3. 836E+04 1. 243E+01

( to be continued 6-68

continued RACK-10:

73 5. 055E+04 6. 543E+00 74 6. 174E+04 8.995E+00 75 5.4 19E+04 8.276E+00 76 5.436E+04 9.261E+00 77 1.388E+04 1.274E+01 78 2.002E+04 1.090E+01 79 1.793E+04 1. 101E+01 80 1. 664E+04 1. 059E+01 RACK-11:

81 4. 519E+04 1. 098E+01 82 4. 531E+04 9 .441E+00 83 5. 320E+04 9. 876E+00 84 5.821E+04 1. 119E+01 85 1. 409E+04 4.215E+00 86 1. 454E+04 1. 025E+01 87 2.132E+04 4. 766E+00 83 2.244E+04 5. 147E+00 RACK-12:

89 6.284E+04 7.215E-00 90 6.717E+04 4.223E+00 91 5.764E+04 3.941E+00 92 5.995E+04 5.483E+00 93 1.665E+04 1.056E+01 94 2.427E+04 1. 091E+01 95 2.228E+04 5. 563E+00 96 2.660E+04 5. 863E+00 RACK- 13 97 6.097E+04 5.481E+00 98 .901E+0+/-4 9.881E+00 99 6.666E+04 4.227E+00 100 7.835E+04 9.708E+00 101 2.122E+04 7. 006E+00 102 2.390E+04 6. 479E+00 103 3.439E+04 5. 576E-00 104 3.336E+04 5. 133E+00 RACK-14:

105 6.044E+04 9. 425E+00 106 5.281E+04 4.228E+00 107 6.060E+04 4. 229E+00 108 4.973E+04 1. 024E+01 109 2.090E+04 1. 255E+01 110 2.018E+04 1. 206E+01 111 3.123E+04 1.038E+01 112 4.002E+04 1. 067E+01

( to be continued 6-69

( continued )

RACK-15:

113 6. 348E+04 1. 114E+01 114 9.482E+04 1. 136E+01 115 1. 035E+05 5. 048E+00 116 7. 754E+04 8. 843E+00 117 5. 757E+04 5. 236E+00 118 2.457E+04 6. 035E+00 119 1.752E+04 1. 155E+01 120 5. 116E+04 6. 954E+00 121 4 .977E+04 7.288E+00 RACX-16:

122 5.376E+04 6. 754E+00 123 7.305E+04 1. 136E+01 124 6. 332E+04 7.392E+00 126 6. 215E+04 5.474E+00 126 3. 690E+04 5.561E+00 127 3. 736E+04 8. 513E+00 123 1. 190E+04 1. 345E+01 129 1. 450E+04 1. 266E+01 RACK-17:

130 5. 535E+04 8. 569E+00 13 5. 133E+04 1. 137E+01 132 4. 916E+04 8. 116E+00 123 5.351E+04 1. 120E+01 124 3. 579E+04 7.992E+00 13 2.790E+04 1. 155E+01 126 2. 419E+04 1. 109E+01

1. 497E+04 1.141E+01 RACK-18:

1383 7. 992E+04 5.498E+00 139 6. 048E+04 6.501E+00 140 7. 165E+04 5.028E+00 141 5. 909E+04 9.941E+00 142 3. 511E+04 8.896E+00 143 4.13 3E+04 8. 516E+00 144 3. 997E+04 5. 333E+00 145 4. 093E+04 5. 106E+00

( to be continued )

6-70

( continued )

RACK-TO-RACK/WALL IMPACT SPRINGS AT RACK TOP 146 0. OOOE+00 0. OOOE+00 147 0. OOOE+00 0. OOOE+00 148 0. OOOE+00 0. OOOE+00 149 0. OOOE+00 0. OOOE+00 150 0. OOOE+00 0. OOOE+00 151 0. OOOE+00 0. OOOE+00 152 0.000E+00 0. OOOE+00 153 0. OOOE+00 0. OOOE+00 154 0. 000E+00 0. OOOE+00 155 0. OOOE+00 0. OOOE+00 156 0. OOOE+00 0. OOOE+00 157 0.000E-00 0. OOOE+00 158 0.000E+00 0. OOOE+00 159 0.000E+00 0. OOOE+00 160 0.OOOE+00 0. 000E+00 161 0.000E+00 0.000E+00 162 0.000E-00 0.000E+00 163 0.OOOE+00 0. OOOE+00 164 0.000E+00 0. OOOE+00 165 0.OOOE-00 0. OOOE+00 166 0.000E-00 0. 000E+00 167 0.000E+00 0. 000E+00 168 0.000E+00 0. OOOE+00 169 0.000E+00 0. OOOE+00 170 0.000E+00 0. OOOE+00 17 . 0.000E+00 0. 000E+00 172 0.000E+00 0. 000E+00 173 0.000E+00 0. OOOE+00 17 A 0.000E-+00 0.000E-00 175 0.000E-00 0. 000E+00 176 0.000E+00 0. OOOE+00 177 0.000E+00 0. OOOE+00 178 0.000E-00 0. OOOE+00 179 0.000E+00 0. OOOE+00 180 0.000E+00 0. OOOE+00 181 0.000E+00 0. OOOE+00 182 0.000E-00 0. OOOE+00 183 0.000E+00 0. OOOE+00 184 0.OOOE+00 0.000E+00 185 0.OOOE+00 0.OOOE+00 186 0.OOOE+00 0.OOOE+00 187 0.OOOE+00 0.OOOE+00 188 0.000E+00 0.OOOE+00 189 0.OOOE+00 0.OOOE+00 190 0.000E+00 0.OOOE+00

( to be continued )

6-71

( continued )

191 0. OOOE+00 0. OOOE+00 192 0. OOOE+00 0. OOOE+00 193 0. OOOE+00 0. OOOE+00 194 0. OOOE+00 0. OOOE+00 195 0. OOOE+00 0. OOOE+00 196 0. OOOE+00 0. OOOE+00 197 0. OOOE+00 0. OOOE+00 198 0. OOOE+00 0. OOOE+00 199 0. OOOE+00 0. OOOE+00 200 0. OOOE+00 0. OOOE+00 201 0. OOOE+00 0. OOOE+00 202 0. OOOE+00 0. OOOE+00 203 0. OOOE+00 0. OOOE+00 204 0. OOOE+00 0. OOOE-+00 205 0. OOOE+00 0. OOOE+00 206 0. OOOE+00 0. OOOE+00 207 0. OOOE+00 0. OOOE+00 208 0. OOOE+00 0. OOOE+00 209 0. OOOE+00 0. OOOE+00 210 0. OOOE+00 0. OOOE+00 211 0. OOOE+00 0.OOOE+00 212 0. OOOE+00 0. OOOE+00 213 0.000E+00 0. OOOE+00 214 0. OOE+00 0. OOOE+00 215 0. OOOE+00 0. OOOE+00 216 0. OOOE+00 0. OOOE+00 217 0. OOOE+00 0.000E+00 213 0. OOOE+00 0.000E+00 219 0.000E+00 0. OOOE+00 220 0.00OE-00 0.000E+i00 0 .000E+00 221 0. OOOE+00 222 0. OOOE+00 0.OOOE+00 223 0. OOOE+00 0.000E+00 224 0. OOOE+00 0.OOOE+00 225 0. OOOE+00 0.OOOE+00 226 0. OOOE+00 0.000OE+00 227 0. OOOE+00 0.OOOE+00 228 0. OOOE+00 0.000E+00 229 0. OOOE+00 0.OOOE+00 230 0. OOOE+00 0.OOOE+00 231 0. OOOE+00 0.OOOE+00 232 0. OOOE+00 0.OOOE+00 233 0. OOOE+00 0.OOOE+00 234 0.000E+00 0.OOOE+00 235 0.00OE+00 0.OOOE+00

( to be continued )

6-72

continued )

RACK-TO-RACK/WALL IMPACT SPRINGS AT RACK BOTTOM 236 0. OOOE+00 0.OOOE+00 237 0. OOOE+00 0.OOOE+00 238 0. OOOE+00 0.OOOE+00 239 0. OOOE+00 0.OOOE+00 240 0. OOOE+00 0.OOOE+00 241 0. OOOE+00 0.OOOE+00 242 0. OOOE+00 0.000E+00 243 0. OOOE+00 0.OOOE+00 244 0. OOOE+00 0.000E+00 245 0. OOOE+00 0.000E+00 246 0. OOOE+00 0.OOOE+00 247 0. OOOE+00 0.000E+00 248 0. OOOE+00 0.OOOE+00 249 0. OOOE+00 0. 000E+00 250 0. OOOE+00 0.OOOE+00 251 0. 000E+00 0.OOOE+00 252 0.000E+00 0.OOOE+00 253 0 . OOOE+00 0.000E+00 254 0. OOOE+00 0.OOOE+00 255 0. OOOE+00 0.OOOE+00 256 0. OOOE+00 0.OOOE+00 257 0.000E+00 0.000E+00 258 0. OOOE+00 0.000E+00 259 0.000E-00 0.OOOE+00 260 0. OOOE-00 0.000E+00 261 0. OOOE+00 0.OOOE+00 262 0 . OOOE+00 0.000E+00 263 0. OOOE+00 0.OOOE+00 0.000E+~00 264 0 . OOOE+00 0.000E+00 265 0 . OOOE+00 0.OOOE+00 0.000E+00 267 0. OOOE+00 0.000E+00 268 0.000E+00 0.000E+00 269 0.000E+00 0.OOOE+00 270 0.000E+00 0.OOOE+00 271 0.000E+00 0.OOOE+00 272 0.000E+00 0.000E+00 273 0. OOOE+00 0.OOOE+00 274 0. OOOE+00 0.000E+00 275 0. OOOE+00 0.OOOE+00 276 0. OOOE+00 0.OOOE+00 277 0.000E+00 0.OOOE+00 278 0. OOOE+00 0.OOOE+00 279 0.OOOE+00O 0.OOOE+00 280 0. OOOE+00 0.OOOE+00 281 0. OOOE+00 0.OOOE+00 282 0. OOOE+00 o.OOOE+00

( to be continued )

6-73

continued )

283 0. OOOE+00 0. OOOE+00 284 0. OOOE+00 0. OOOE+00 285 0. OOOE+00 0. OOOE+00 286 0. 000E+00 0. OOOE+00 287 0. 00OE+00 0. OOOE+00 288 0.OOOE+00 0. OOOE+00 289 0. OOOE+00 0. OOOE+00 290 0. OOOE+00 0. OOOE+00 291 0. OOOE+00 0. OOOE+00 292 0. OOOE+00 0. OOOE+00 293 0. OOOE+00 0. OOOE+00 294 0. OOOE+00 0. OOOE+00 295 0. OOE+00 0. OOOE+00 296 0. OOOE+00 0. OOOE+00 297 0. 000E+ 00 0. OOOE+00 293 0. OOOE+00 0. OOOE+00 299 0. OOE+00 0. OOOE+00 300 0. OOOE+00 0. OOOE-+00 301 0. OOOE+00 0.000E+00 302 0. 000E+00 0. OOOE+00 303 0. 000E+00 0. OOOE+00 304 0.000E+00 0. OOOE+00 305 0. OOOE+00 0. OOOE+00 306 0. OOOE+00 0. OOOE+00 307 0.000 E+00 0. OOOE+00 308 0. oooE+00 0.000E+00 309 0. OOOE+00 0. OOOE+00 310 0. OOOE+00 0. OOOE+00 311 0. OOOE+00 0.OOOE+00 312

0. OOOE+00 0.OOOE+00

0. OOOE+00 0.OOOE+00 314 0. OOOE+00 0.000E+00 3 15 0. 000E+00 0.000E+00 316 0. OOE+00 0.000E+00 3 17 0.OOOE+00 0.OOOE+00 318 0. OOOE+00 0.OOOE+00 319 0.OOOE+00 0.OOOE+00 320 0.OOOE+00 0.OOOE+00 321 0.OOOE+00 0.OOOE+00 322 0.OOE+00 0.OOOE+00 323 0. OOOE+00 0.OOOE+00 324 0.OOOE+00 0.OOOE+00 325 0.OOOE+00 0.OOOE+00 6-74

Table 6.8.3 MAXIMUM PEDESTAL STRESS FACTORS OF ALL RACKS IN POOL DUANE ARNOLD PLANT, IOWA ELECTRIC LIGHT AND POWER CO.

WPMR Analysis, 18 Holtec Racks in Pool, Fully Loaded with 680# Reg.Fuel; Random Friction; Seismic: DBE ( =OBEx2.0 ); dt=0.00005 sec.

• • • • • • • • • • • • • • • • • • INPUT DATA ******************

File name of FX Time history :frictiox File name of FY Time history :fricticy File name of FV Time history :p tfy10

$Revision: 1.0 $

$Logfile: C:/racks/multirac/sfmr2.fov S

$Date: 28 May 1992 13:08:26 $

File name of result output :s sse. rfr Number of racks in the pool . 8 Height of the pedestal, in. 5.50 Offset of FV from center, in. 1.06 Area of female pedestal, in**2.  : 8.19 Inertia of female pedestal, in**4. :221.81 Distance of extrame fiber in X, in. 3.00 Distance of extrame fiber in Y, in. 3.00 Yield stress of female pedestal, psi. : 25000.

Number of pedestals of each rack  : 4 4 4 4 44 4 4 4 4 4 4 5 4 4 4 MAXIMUM VALUES OF STRESS FACTORS, R1 -- R7, FOR ALL RACK PEDESTALS Maximum Values of RI -- R7 R1,R3,R4 for Max.R6 RI R2 R3 R4 R6 R7 R1R6M R3R6M R4R6M

.378 .256 .093 .153 .439 .451 .183 .374 .026 .051 Rack No.:

15 9 13 9 15 15 13 Pedestal No.:

3 2 4 2 3 3 4 Time (sec.):

5.040 11.290 9.690 11.290 11.550 11.550 9.690 6-75

Table 6.8.4 MAXIMUM RACK DISPLACEMENTS, PEDESTAL LOADS, AND PEDESTAL STRESS FACTOR IN SINGLE RACK ANALYSES AND IN WPMR ANALYSES (DBE Seismic; Reg. Fuel, fully loaded)

Maximum Maximum Pedestal Maximum Rack Corner vertical Pedestal Displacement Load Stress Run I.D. Remarks (in.) (1bs) Factor drjsseo.rf8 single rack 0.0505 73061 0.308 analysis in x-dir. (R6)

Rack-I cof. = 0.8 drjsseo.rf2 Single rack 0.0635 73034 0.340 analysis in y-dir. (R6)

Rack-J cof. = 0.2 dwpmrsse.rfr WPMR 0.4086 103500 0.451 analysis; in x-dir. (Rack-J,' (R6) cof. = random Rack-E (17) Foot-3) Rack-J (mean cof.=.5) 0.2685 Foot-3 in y-dir.

Rack-N1 (1) cof. = coefficient of friction between pedestal and pool liner.

6-76

Table 6.8.5 RESULTS OF POOL WALL DYNAMIC PRESSURES DUANE ARNOLD PLANT, IOWA ELECTRIC LIGHT AND POWER CO.

VPM.R Analysis, 18 Holtec Racks in Pool, Fully Loaded with 680# Reg.Fuel; Random Friction; Seismic: DBE ( =OBEx2.O ); dt=0.00005 sec.

$Revision: 1.0 $

$Logfi le: C:/racks/multirac/wallpres.fov $

(1) AVERAGE DYNAMIC PRESSURES ON POOL WALLS (psi.):

Average dynamic pressure on (-x) wall: -2.297742E-04 Average dynamic pressure on (+x) wall: -2.297742E-04 Average dynamic pressure on (-y) wall: 1.761151E-03 Average dynamic pressure on (+y) wall: 1.761151E-03 (2) PEAK DYNAMIC PRESSURES ON POOL WALLS (psi.):

Positive peak pressure on (-x) wall: 6.310000 Negative peak pressure on (-x) wall: -6.160000 Positive peak pressure on (+x) wall: 6.310000 Negative peak pressure on (+x) wall: -6.160000 Positive peak pressure on (-y) wall: 10.400000 Negative peak pressure on (-y) wall: -8.060000 Positive peak pressure on (+y) wall: 10.400000 Negative peak pressure on (+y) wall: -8.060000 (3) DYNAMIC PRESSURE ADDERS ON POOL WALLS (psi.):

Positiv-e pressure adder on (-x) wall: 1.519534 Negative pressure adder on (-x) wall: -1.420336 Positive pressure adder on (+x) wall: 1.519534 Negative pressure adder on (-x) wall: -1.420336.

Positive pressure adder on (-y) wall: 2.471790 Negative pressure adder on (-y) wall: -2.490374 Positive pressure adder on (+y) wall: 2.471790 Negative pressure adder on (-y) wall: -2.490374 (4) NUMBER OF TIME POINTS:

• Total number of time points in file: 1501

• Number of time points where dynamic pressure on (-x) wall was positive: 725

• Number of time points where dynamic pressure on (-x) wall was negative: 776

• Number of time points where dynamic pressure on (+x) wall was positive: 725

• Number of time points where dynamic pressure on (+x) wall was negative: 776

• Number of time points where dynamic pressure on (-y) wall was positive: 754

• Number of time points where dynamic pressure on (-y) wall was negative: 747

• Number of time points where dynamic pressure on (+y) wall was positive: 754

• Number of time points where dynamic pressure on (+y) wall was negative: 747 6-77

Table 6.8.6 TOTAL STATIC LOAD AND DYNAMIC ADDER ON WHOLE SLAB DUANE ARNOLD PLANT, IOWA ELECTRIC LIGHT AND POWER CO.

WPMR Analysis, 18 Holtec Racks in Pool, Fully Loaded with 680# Reg.Fuel; Random Friction; Seismic: DBE ( =OBEx2.0 ); dt=0.00005 sec.

(1) THE TOTAL STATIC LOAD OF RACKS AND FUEL ON SLAB IS:

1350900.00 lbs.

(2) THE TOTAL DYNAMIC LOAD ADDER ON THE SLAB IS:

165284.70 lbs.

6-78

Table 6.9.1 AVERAGE BEARING PAD PRESSURE - COMPARISON OF CALCULATED AND ALLOWABLE STRESSES STRESS (psi Max. Load Pad Size (lb.) Calculated Allowable 12.0 x 12.0 103500 719 2380*

(Table 6.7.4)

(based on concrete strength fc'

= 4000 psi)

• factor E = 1 6-79

O (1)

DUANE ARNOLD ENEIRGY CENTER z4 REACTOR BUILDING N-S AND E-W EQS.

n E[PT. DESIGN SPECTRA DAMPING = 0.010 R

C -_

c0 0 oL caN I

T,00 0.25 0,50 0 1.00 1 .25 150 1.75 .

PERI10 IN SECONOS FIGURE 6.3.1 HORIZONTAL RESPONSE SPECTRUM FOR DAEC

JOHN A. BLUME AND ASSOCIATES. ENGINEERS DURNE ARNOLD ENERGY CENTER RERCTOR BUI LDING N-S AND E-W EOS.

EOPT. DESIGN SPE CTRA DAMPING = 0.010 U3 xjc) 0 E

F-00  :)

I-A F

Ck3 Lc) a I!

4.00 0.50 0.75 1-00 I -; 5 2.00 PERIO0 IN SECONDS FIGURE 6.3.2 VERTICAL RESPONSE SPECTRUM FOR DAEC

IEL&PDuane Arnold Fuel Pool. Slob.

0.15 Seismic Time History, OBE, N-S.

0.10 0.05

-0.00

-0.1O

-0.15

-0.20 1 0.00 5.00 10.00 15.OO Time, sec.

Figure 6.3.3

IEL&P Duane Arnold Fuel Pool Slob.

0.15 Seismic Time History, OBE, E-W.

0.00

.' -O.O I

w o

O

<c

-0.10

-0.15 -5

-0.20.:

O.1 Tirne, sec.

Figure 6.3.4

IEL&P,Duane Arnold Fuel Pool Slab.

0.15 Seismic Time History, OBEVT.

0.10 0.05

-" -0.00

-0.05 0

<0

-O 10

-0.15

-0.20 15.00 Time, sec.

Figure 6.3.5

IEL&P,Duane Arnold fuel pool,

'~ EZC~ Seismic Spectrum: OBE, N-S.

l/.

2.00 1.50 I'

0 1.00 0.50 U.U - I a i ar-rr-ri .

I I I 1511

'I I

-1I 10 1 10 102 Frequency, Hz.

FIGURE 6.3.6

IEL&P,Duane Arnold fuel pool, 2.50 Seismic Spectrum: OBE, 2.00 CI) 1.50 U

a' 0 D

0) 1.00 C-)

C-)

0.50 0.00 - I I I I lU ll I,-..----.-.--.---. U I I. I 11111 I I I sI s10 10-' 1 10 102 Frequency, Hz.

FIGURE 6.3.7

IEL&P,Duane Arnold fuel pool, 2.50 Seismic Spectrum: OBE, VT.

2.00

- 1.50 0 0 1.00 0.50 -

0.00 -

I I I I II I~II I I I I I1I11I1

-1 10 1 10 10 Frequency, Hz.

F1GU)RE 6.3.8

TYPICAL CELL WALLS BASEPLATE FIGURE 6.4.1 PICTORIAL VIEW OF RACK STRUCTURE 6-88

a2 I

/

pI qZ H/4 ear CailERLIIC H

H/4 41rp1 P14 Q

05 19sa FIGURE 6.4.2 SCHEMATIC MODEL FOR DYNARACK 6-89

I YI'll Al 11' INI'All flUMINI PALKS illE a

Ia 0o i~l'l(Al-111111

]WIT 4/~l 77

,f'7 7 FIGURE 6.4.3 RACK-TO-RACK IMPACT SPRINGS

Y IfL ASS1LY/[1ELL IMPAET SIPING I

to X - - - -x FIGURE 6.4.4 FUEL-TO-RACK IMPACT SPRINGS

A'I (]jq17 q20 zL

q. (I In FIGURE 6.4.5 DEGREES-OF-FREEDOM MODELING RACK MOTION

10 (2

L/2 9l 1-0 FIGURE 6.4.6 RACK DEGREE-OF-FREEDOM FOR Y-Z PLANE BENDING

q17 L/2 j

FIGURE 6.4.7 RACK DEGREE-OF-FREEDOM FOR X-Z PLANE BENDING

IV2 FRIIllN INIERFAEE SPRING, K

-51PPloR TLuli SI'INI, K

-FillN)JAIIPN PIlAIIONAL Uh IUNEE SPRING, Kp FIGURE 6.4.8 2-D VIEW OF RACK MODULE

I - 400' NI1M.

F (477.7" MIN.)

240'NOM.

(238.34' MIN.)

J (19) RACKS - (3152) CELLS INCL. (1) RAEK - 323 CELLS x IN EASK PIT POOL LAYOUT - CAMPAIGN III FIGURE 6.4.9

0 GAP TIME HISTORY, DUANE ARNOLD PLANT, I.E.L.& P.CO.

Gap between Rack-N1 and Rack-N2, West Corner,Top, 0 Fully Loaded with 680# Reg.Fuel Assemblies, Friction coefficient =: random ( mean = 0.5 ). File:G1-7YW.DAT.

0 00 0-O0 0O-O-

-S - I U I 11.l 1 I I CI). 0 2.0 4.0 6. 8.0 10.0 12.0 14.0 lime, sec.

FIGURE 6.8.1

0 0

GAP TIME HISTORY, DUANE ARNOLD PLANT, I.E.L.& P.CO.

Gap between Rack-N1 & North Wall, West CornerTop, DBE Fully Loaded with 680# Reg.Fuel Assemblies, Friction coefficient =: random ( mean = 0.5 ). File:G1 -WYW.DAT.

0 0-

~0 C 0

0-C

~

III TTT111 1i ' i IIIIIII III1III 1111111li ii 11III i ' 'r us

eI' ru''  !"

c).0 2.0 4.0 6.0 8.0 10.0 12.0 14.0 Time, sec.

FIGURE 6.8.2

0 0

GAP TIME HISTORY, DUANE ARNOLD PLANT, L.E.L.& P.CO.

Gap between Rack-F and Rack-E North Corner, Top, DBE, Fully Loaded with 6804 Reg.Fuel Assemblies, Friction coefficient == random ( mean = 0.5 ). File:G16-17XN.DAT.

0~v

, - Vi v

.0 614 C

c- -T--I--1-- 111111 11111 ~II1 1 111 111 Jill"11T1 11 11 11111 111 111 ji CO.0 2.0 4.0 6.0 8.0 10.0 12.0 14.0 Time, sec.

FIGURE 6.8.3

0 0

GAP TIME HISTORY, DUANE ARNOLD PLANT, I.E.L.& P.CO.

Gap between Rack-E and Rack-D North Corner, Top, DB Fully Loaded with 650.4 Reg.Fuel Assemblies, Friction coefficient =: random ( mean 0.5 ). File:G17-18XN.DAT.

0 6-0 O~

O-00-0 O-O -V

- II I CIO.0 2.0 4.0 6.0 8.0 10.0 12.0 14.0 Time, sec.

FIGURE 6.8.4

0 0-GAP TIME HISTORY, DUANE ARNOLD PLANT, I.E.L.& P.CO.

Gap between Rack-D and East Wall,North Cornerjop, DBE, Fully Loaded with 680# Reg.Fuel Assemblies.

Friction coefficient = random ( mean 0.5 ). File:G18-WXN.DAT.

0 0

ath I"

C')

0 0-0 -

- I T F i jI I I I I I I I I 1 1I 1II1 1 I1 1 1 1 l l l i s si5t i s l aI I aI II I I I I I c).0 2.0 4.0 6.0 8.0 10.0 12.0 14.0 Time, sec.

FIGURE 6.8.5

0 0

0 0

0 Lfl O

0 C

0*

0.

0 0

<n IJA 0 a N 0

-Q 0 0*

('0 0 O In o~

Total Slab Load Time-History DUANE ARNOLD PLANT,I.E.L.& P.CO.

DBE Seismic, Fully loaded with 680# reg.fuel assemblies.

Friction coefficient = random, File: slabld.dat.

O-I 0O O-T

.0o 6.0 8.0 Time, sec.

FTii lil: 6.8.6

7.0 ACCIDENT ANALYSIS AND MISCELLANEOUS STRUCTURAL EVALUATIONS 7.1 Introduction This section provides results of accident analyses and miscellaneous evaluations performed to demonstrate regulatory compliance of the new fuel racks.

The following limiting accident and miscellaneous structural evaluations are considered:

• Refueling accidents - drop of fuel assembly to top of rack or through a cell to the baseplate

• Local cell wall buckling

• Analysis of welded joints due to isolated hot cell 7.2 Refueling Accidents 7.2.1 Dropped Fuel Assembly The consequences of dropping a fuel assembly as it is being moved over stored fuel is discussed below. It is assumed that the lowest part of the fuel assembly being carried is 18" above the top of the new spent fuel racks [7.2.1]. The fuel assembly weighs 680 lbs. and associated handling equipment is assumed to weigh 120 lbs.

a. Dropped Fuel Assembly Accident (Deep Drop Scenario)

An 800 lb. fuel assembly plus handling equipment is dropped from 18" above the top of a storage location and impacts the base of the module. Local failure of the baseplate is acceptable; however, the rack design should ensure that gross structural failure does not occur and the subcriticality of the adjacent fuel assemblies is not violated. Calculated results show that there will be no change in the spacing between cells. Local deformation of the baseplate in the neighborhood of the impact will occur, but the dropped assembly will be contained and not impact the liner. Calculations also show that even if there is local cell-to-baseplate weld overstress in individual cells, the maximum movement of the baseplate toward the liner after the impact is at most between .94" 7-1

and 1.52". The load transmitted to the liner through the support by such an accident is well below the loads caused by seismic events (given in Section 6).

b. Dropped Fuel Assembly Accident (Shallow Drop Scenario)

One fuel assembly plus the channel is dropped from 18" above the top of the rack and impacts the top of the rack. Permanent deformation of the rack is acceptable, but is required to be limited to the top region such that the rack cross-sectional geometry at the level of the top of the active fuel (and below) is not altered. Assuming a minimal area of impact, it is shown that damage, if it occurs, will be restricted to a depth of less than or equal to 1.09" below the top of the rack. This is above the active fuel region.

7.3 Local Buckling of Fuel Cell Walls This subsection and the next one present details on the secondary stresses produced by buckling and by temperature effects.

The allowable local buckling stresses in the fuel cell walls are obtained by using classical plate buckling analysis. The following formula for the critical stress has been used based on a width of cell "b" [7.3.1]:

2 B 7TI E t acr 12 b 2 (1 _ p2) a,, is the limiting vertical compressive stress in the panel, E =

27.6 x 106 psi, p = 0.3, (Poisson's ratio) , t = .06", b = 5.96". The factor 8 is suggested in (Ref. 7.3.1) to be 4.0 for a long panel.

For the given data, gcr = 10112 psi 7-2

It should be noted that this stability calculation is based on the applied stress being uniform along the entire length of the cell wall. In the actual fuel rack, the compressive stress comes from consideration of overall bending of the rack structures during a seismic event and as such is negligible at the rack top and maximum at the rack bottom. It is conservative to apply the above equation to the rack cell wall if we compare acr with the maximum compressive stress anywhere in the cell wall. As shown in Section 6, the local buckling stress limit is not violated anywhere in the body of the rack modules. The maximum compressive stress in the outermost cell is obtained by multiplying the limiting value of the stress factor R6 (for the cell cross-section just above the baseplate) by the allowable stress (.6 x yield stress). Thus, from Table 6.7.2, a =

R6 x allowable stress = .111 x (.6 x 25000) = 1665 psi. This is less than 5% of the allowable secondary stress in the panel.

7.4 Analysis of Welded Joints in Rack due to Isolated Hot Cell In this subsection, in-rack welded joints are examined under the loading conditions arising from thermal effects due to an isolated hot cell.

A maximum thermal gradient between cells will develop when an isolated storage location contains a fuel assembly emitting maximum postulated heat, while the surrounding locations are empty. We can obtain a conservative estimate of weld stresses along the length of an isolated hot cell by considering a beam strip (a cell wall) uniformly heated and restrained from growth. The strip is subject to a uniform temperature rise AT = 62.9 0 F. The temperature rise has been calculated from the difference of the maximum local water temperature and bulk water temperature in the spent fuel pool. (see Tables 5.8.2 and 5.8.4). Then, using a shear beam theory, we can calculate an estimate of the maximum value of the average shear stress in the strip (see Figure 7.4.1).

7-3

The final result for wall maximum shear stress, under conservative restraint assumptions is given as [7.5.1]:

E a AT max

.931 where a = 9.5 x 106 in/in OF.

Therefore, we obtain an estimate of maximum weld shear stress in an isolated hot cell as Tmax = 17715 psi Since this is a secondary thermal stress, it is appropriate to compare this to the allowable weld shear stress for a faulted event r < .42S, = 29820 psi (S, defined in Table 6.5.1) . In the fuel rack, this maximum stress occurs near the top of the rack and does not interact with any other critical stress.

7.5 References for Section 7

[7.2.1] UFSAR/DAEC-1, p. 9.1-19.

[7.3.1] "Strength of Materials", S.P. Timoshenko, 3rd Edition, Part II, pp 194-197 (1956).

[7.5.1] "Seismic Analysis of High Density Fuel Racks, Part III -Structural Design Calculations Theory", Holtec Report No. HI-89330, Revision 1, 1989.

7-4

t Heated Cell Wall x H KL I WedLine T

y FIGURE 7.4.1 WELDED JOINT IN RACK 7-5

8.0 FUEL POOL STRUCTURAL INTEGRITY CONSIDERATIONS 8.1 Introduction The Duane Arnold Energy Center (DAEC) spent fuel pool is a safety related seismic category I, reinforced concrete structure. In this section, the analysis to demonstrate the structural adequacy of the pool structure, as required by Section IV of the USNRC OT Position Paper [8.1.1], is summarized.

Structural analyses have been performed to demonstrate that the reracked DAEC spent fuel pool structure satisfies the provisions of the USNRC Standard Review Plan Section 3.8.4 [8.1.2]. The analyses are carried out for the point in time when the DAEC pool is filled with fuel bundles (2829 storage cells) i.e, the inertia loading in the pool is at its maximum. A rack will be put in the cask pit.

The cask pit is already qualified for a cask that is heavier than any rack fully loaded with fuel. The cask pit is therefore not part of this analysis and the loads used are based only on the racks actually in the spent fuel pool. The compliance criteria used here is more conservative than that required by the plant UFSAR [8.1.3].

The original design basis for the spent fuel pool was the ACI Concrete Code (1963) [8.1.7].

A detailed 3-D finite element model of the pool slab and confining walls has been developed. The analysis approach and structural integrity evaluation is also in compliance with the relevant American Concrete Institute Standards (ACI-318.89, ACI-318R.89

[8.1.4], ACI-349.85 [8.1.5], and ACI-3491R.80 [8.1.6]).

The following procedures and assumptions have been used to ensure a state-of-the-art pool slab analysis with a large element of conservatism.

8-1

1. The slab structure and surrounding containment is modelled using a 3-D finite element code [8.1.8]

incorporating plate elements to form continuous walls and floor. Steel reinforcing I-beams are modeled using beam elements. The structure is supported from the reactor wall and also from I-beams which extend out to the exterior walls of the reactor building.

2. The increased vertical dynamic loading from the racks during a seismic event is calculated using the non-linear dynamic model encompassing all the racks in the pool (Chapter 6). The dynamic model of the racks moving in the excited pool incorporates the effect of fluid coupling and the interaction of all the racks with each other and with the pool structure. The results of the non-linear whole pool multi-rack analysis are the time histories of every rack support pedestal vertical load. These vertical pedestal load time histories are then used to develop the appropriate equivalent slab pressure which correctly models the dynamic effect of the racks plus their fuel load on the pool slab.

3. The finite element analysis is carried out using loads which bound the actual calculated loads so as to provide sufficient margins for potential future fuel loadings using a heavier fuel assembly.

4. The factored load combinations are performed using absolute value sums to provide the most conservative estimate of safety margins. This approach builds in a large margin of conservatism in the calculated results.

5. Mechanical and thermal stress resultants on the pool walls and slab are computed assuming that the concrete sections do not crack. This gives a conservative estimate of the effect of thru-thickness thermal gradients and mean temperature [8.1.6].

6. Certain boundary restraints are neglected in mechanical load analyses and included in the thermal load analyses; thus, the most conservative estimate of force and moment resultants are obtained for each class of loading.

7. The analysis is linear elastic; no nonlinear elements are used in the model.

In the finite element analysis, we consider:

1. Static loads from structure, water, and racks plus fuel.

8-2

2. Dynamic loads from the structure and contained fluid during a seismic event.

3. Dynamic loads from the racks plus stored fuel.

4. Thermal loads due to temperature gradients across the walls and slab.

These loadings are combined in accordance with the aforementioned design criteria and critical bending moments are evaluated for each load combination. The structural evaluation is then made by comparing calculated moments with the allowable moment at the same location. Both moment and shear capacities are checked.

An insight into the magnitudes of the proposed change may be obtained by examining Tables 8.1.1 and 8.1.2. In Table 8.1.1 it is shown that in terms of racks plus full fuel load there is a 26.7%

increase in load on the slab. However, Table 8.1.2 demonstrates that when the water plus the slab weight is included (to simulate an operating environment), the overall increase over the existing configuration is only 9%.

8.2 General Features of the Model The Duane Arnold Energy Center (DAEC) spent fuel pool is located in the reactor building immediately adjacent to the south wall of the reactor shield wall. The pool floor is located at elevation 812' and the pool walls extend to elevation 855'. The pool floor slab thickness is 6' with a 24' x 40' planform. The fuel pool north wall is the reactor shield wall. The fuel pool south wall is 6' thick. The fuel pool west wall is 6' thick from elevation 812' to 833.5' and 5' thick from elevation 833.5' to elevation 855'. The fuel pool east wall is 6' thick to elevation 833.5' and 4.5' from 8-3

elevation 833.5' to elevation 855'. The east and west walls extend beyond the south wall of the pool out to the south wall of the reactor building.

The finite element model of the fuel pool and surrounding supports includes the slab, the south wall, and the east and west walls out to the exterior wall of the reactor building. The reactor shield wall forms the north wall of the pool and is not modeled. The slab and the east and west walls are assumed fixed where they attach to the reactor shield wall. Vertical support to the fuel pool floor slab is provided by the east and west walls acting as deep beams supported by the reactor shield wall and the exterior wall of the building. Steel reinforcing I beams are located under the floor slab and along the long edges of the walls at levels 812', 833.5' and 855'. These major reinforcing members are included in the finite element model. In the interest of conservatism, no additional supports are included in the model when mechanical loads are considered, but additional restraint was modeled, where necessary, for the cases where thermal loadings are imposed so as to amplify the effect of temperature gradients. For example, the additional north-south directed wall between the south wall of the pool and the south wall of the building is not modeled; however, the south wall is restrained against movement laterally when thermal loads are applied. In other words, restraints are only included when they add to the conservatism of the modeling.

STIF63 shell elements (8.1.8] are used to model walls and slab.

STIF63 is a 3-D quadrilateral shell element and has both bending and membrane capabilities. This element has been independently validated under Holtec's QA program. STIF63 has options for modeling different thickness and elastic foundation supports.

Right angle connections between adjacent walls and between walls and slab are modeled by shell elements. Figures 8.3.1-8.3.3, developed in the course of the frequency analysis, show the extent of the model.

8-4

The concrete compressive strength f, used is 4000* psi and the reinforcement strength is 60,000 psi. This data was obtained from the plant UFSAR [8.1.3]. Reference [8.1.4] is used to compute the Young's Modulus for uncracked concrete as E, = 57000 V'fe' = 3.605 x 106 psi This value is assumed as the modulus for all elements for both the mechanical and thermal load cases. The Poisson's ratio of concrete used is 0.14. For the thermal calculations, a coefficient of linear thermal expansion of 5.0 x 10-6 in./in.OF is used for all components.

8.3 Factored Loadings for the Fuel Pool Structure SRP Section 3.8.4 [8.1.2] is used to set forth the factored load combinations for concrete structure. Since there are no live loads, of all load combinations listed in [8.1.2], the applicable ones are:

A. 1.4D + 1.9E B. 0.75 (1.4D + 1.9E + 1.7To)

C. D + Ta + 1.25E D. D + Ta + E' While the original load combination applicable to the plant was based on ACI 318-63 which contains a larger multiplier on the dead load (1.5D instead of 1.4D in the foregoing), the overall severity of the SRP load combinations is considerably stronger than those provided for in the ACI Code. Therefore, in the interest of conservatism, the factored load combinations have been upgraded to SRP 3.8.4 in this submittal.

• The actual core sample data, however, indicates the concrete compressive strength to be considerably higher.

8-5

In addition (not a requirement of [8.1.2]), we consider the additional accident condition of loss-of-cooling and develop the additional load combination:

E. D + Tb In the above combinations, the following notations are used:

D = dead load E' = Design Basis Earthquake (DBE) load E = Operating Basis Earthquake (OBE) load To = steady state thermal load during normal operating or shutdown conditions. To be conservative, use To

= T..

Ta = thermal load in the postulated abnormal design conditions Tb = thermal load in the loss of cooling accident 8.4 Finite Element Analyses 8.4.1 Load Cases To develop the foregoing factored load combinations, five separate load cases for the finite element model are considered:

Case 1 Dead loading from concrete, reinforcement and 40' of hydrostatic head. The loading is applied as a 1g vertical gravitational load for the structure and a surface pressure on the slab and walls for the hydrostatic head. The hydrostatic head on the walls is applied as a gradient from surface of the water down to the slab.

The dead weight of racks plus a full fuel load is also considered.

The equivalent pressure acting on the slab is used to represent these static loads.

Case 2 Seismic excitation of the pool structure and contained water mass in both DBE and OBE events. The structural lowest natural frequencies, giving rise to motions predominately in x, y and z directions, are determined using finite element modal analysis.

The response spectrum g loadings at these frequencies are used as multipliers on the dead weight to determine the seismic response from a static finite element analysis.

8-6

A frequency analysis of the pool model shows that the lowest natural frequency in x direction is 5.368 Hz, in y direction is 6.09 Hz, and in z direction is 6.31 Hz. Figures 8.3.1, 8.3.2 and 8.3.3 show the lowest three mode shapes of the pool structure (including the contained water mass). From the pool response spectra, we obtain the appropriate OBE g levels at these frequencies as:

N-S E-W Vertical OBE 0.6g 0.5g 0.32g For our analysis here, we use 0.6g for horizontal excitation and 0.48g (0.8 of horizontal) for vertical excitation. The DBE case is double the OBE per the specification.

To account for pool fluid sloshing, the method in Reference [8.3.1]

is used to establish an additional effective pressure component on the pool walls.

This "sloshing" pressure is included as a load component in Case 2 in addition to the seismic components.

Case 3 Dynamic load adders due to racks fully loaded with fuel assemblies in DBE (2xOBE) events. The vertical dynamic load adders are based on the totality of the vertical pedestal load adders obtained from the 3-D Whole Pool Multi-Rack Analysis. These vertical adders are then converted to an equivalent pressure acting on the slab by utilizing the principle of conservation of impulse.

Case 4 Thermal loading. The maximum pool water temperature is limited to 165 0 F (Table 5.8.2). The pool wall external air temperature is assumed to be 700 F. The thermal bending moments are obtained from the equivalent linear temperature distribution. The linear temperature distribution is separated into a pure gradient AT and a difference between the mean and the base (stress free) temperature (Tm - Tr). We use Tr as 70 0 F. Appropriate heat transfer coefficients are used to establish the three concrete gradients.

Because it is the temperature difference between the two faces of the pool wall, rather than the absolute values of the temperatures themselves which determines the thermal moment, the temperature difference of 95 0 F (165 minus 70) is the governing design basis for the DAEC pool for normal operating condition.

8-7

Case 5 Thermal loading due to accident conditions (loss of all cooling).

In this case, the pool water temperature is assumed as 212 0 F and pool outside air temperature is 70 0F.

Note that to ensure conservatism in results, material properties for the concrete are assumed to be based on uncracked sections.

This yields higher stress resultants for Cases 4 and 5.

8.4.2 Applied Loads The pool structural analysis is carried out using a bounding set of loads which are greater than the actual loads expected. This is done to ensure that the spent fuel pool contains additional margins sufficient to support the use of heavier fuels at some later date.

For the purpose of this analysis, we use these bounding loads as a means of introducing additional conservatism into the analyses.

For load cases described in 8.4.1, we show below the actual load expected and the load used in the finite element model to obtain results. This illustrates the additional conservatism built into the results in certain load cases.

Case 1 The effect of the new racks plus full fuel load is equivalent to a uniform slab pressure of 16.1 psi. The uniform vertical slab pressure component used to obtain numerical results is 30.3 psi, which introduces substantial conservatisms into the analysis. The higher value corresponds to storing consolidated fuel in the DAEC racks. While this rerack application is not for consolidated storage, the larger dead weight and seismic loading corresponding to consolidated storage are used in the pool structure evaluation.

However, from the perspective of the present amendment request, the increased loadings should be viewed as merely another element of conservatism.

Case 3 The calculated dynamic adders obtained from an analysis of the whole pool using a multi-rack model (see Figure 6.8.6) is equivalent to a pressure of 1.435 psi. Pursuant to the foregoing comment, the dynamic adder actually used to obtain pool resultant forces and moments for this load case is 2.18 psi.

8-8

Case 4 The calculated thermal gradient thru the concrete walls of the spent fuel is 70OF when appropriate film coefficients are considered. For conservatism, an 80aF thru-thickness gradient is used to obtain numerical results.

Case 5 The calculated thru-wall gradient based on actual loading and film coefficients is 108 0 F for the pool boiling case; additional conservatism is provided in the numerical computations by using a 0F.

bounding gradient of 142 The use of bounding mechanical and thermal loads to predict section forces and moments ensures additional conservatism in the predicted results.

8.4.3 Postprocessing of Finite Element Load Cases The load combinations described above are performed using the ANSYS postprocessor POST1. We note that the thermal load and seismic load may reduce other loads in certain areas. To introduce additional conservatism, absolute value summation is used for all load combinations. If we denote Mi as the section moment appropriate to load case i, i = 1,2.. .5, the five load combinations can be obtained through the following summations performed by the postprocessor.

A = 1.4 1MI + 1.9 1M21 + (1.9x.5) lM3 1 B = 1.051M1 + 1.4251M 21 + (1.425x.5) lM31 + 1.2751M41 C = 1.OM I + + 1.251M 21 + (1.25x.5) M31 + 1.01M 41 D = 1.0MI + 2.0lM2 1 + 1.01M 31 + 1.01M 41 E = 1.01M1 + 1.01M51 8-9

8.5 Results of Bending Evaluation To check the adequacy of the structure, the whole pool structure is divided into regions with different strength limits based on available reinforcement. For the structural qualification, we classify the structure into four regions: slab, east wall, west wall, and south wall. The maximum bending moments in each region in both directions are determined for every load combination. The values are then compared with the minimum value of strength in the corresponding direction for that area to get conservative safety margins. The results are shown in Table 8.5.1 and are bounding safety margins in that they are based on absolute value load combinations using conservative bounding loads. From the results (safety margins) presented in Table 8.5.1, it is clear that the pool structure is adequate to accommodate the new spent fuel racks loaded with regular fuel assemblies.

8.6 Results from Shear Evaluation on Floor Slab Using Bounding Loads The floor slab is subject to shearing forces. Using the bounding loads, the factored shear load around the perimeter is calculated as 121.11 KIPS/ft. Load combination A governs for shear. The allowable factored shear loading is 239 KIPS/ft. Thus, the available shear capacity is actually in excess of that required under the limiting load case using larger bounding loads.

8.7 References for Section 8

[8.1.1] OT Position for Review and Acceptance of Spent Fuel Handling Applications, B.K. Grimes, USNRC, April 1978.

[8.1.2] NUREG-0800, Standard Review Plan for Review and Safety Analysis Reports for Nuclear Power Plants, Section 3.8.4, July 1981.

[8.1.3] UFSAR/DAEC-1, 1990, Chapter 3.

8-10

[8.1.4] ACI 318-89, ACI 318R-89, Building Code Requirements for Reinforced Concrete, American Concrete Institute, Detroit, Michigan, 1989.

[8.1.5] ACI 349-85, ACI 349R-85, Code Requirements for Nuclear Safety Related Concrete Structures, American Concrete Institute, Detroit, Michigan, 1985.

[8.1.6] ACI 349.1R-80, Reinforced Concrete Design for Thermal Effects on Nuclear Power Plant Structures, 1981.

[8.1.7] ACI 318-63, Building Code Requirements for Reinforced Concrete, American Concrete Institute, Detroit, Michigan, 1963.

[8.1.8] ANSYS User's Manual Swanson Analysis, Rev. 4.1, 1987.

[8.3.1] TID-7024, Nuclear Reactors and Earthquakes, 1963.

8-11

Table 8.1.1 BUOYANT WEIGHT OF EXISTING RACKS + FUEL COMPARED TO NEW RACKS + FUEL (680 LB. FUEL, ALL CELLS LOADED)

EXISTING NEW CELLS 2050 3152-323 = 2829 FUEL WEIGHT 1,212,780 lb. 1,673,636 lb.

RACK WEIGHT 242,556 lb. 169,911 lb.

TOTAL 1,455,336 lb. 1,843,547 lb.

% Increase over existing = 26.7

• Only cells in spent fuel pool included.

8-12

Table 8.1.2 TOTAL DEAD LOAD OVER POOL SLAB (EXISTING VS. NEW)

(40' OF WATER IN POOL)

EXISTING NEW RACKS + FUEL 1,455,336 1,843,542 WATER 2,048,000 2,048,000 WEIGHT OF 6' SLAB (165 lb/cu.ft.) 792,000 792,000 TOTAL 4,295,336 lb. 4,683,542 lb.

% Increase over existing = 9.04%

8-13

Table 8.5.1 BOUNDING SAFETY MARGINS-BENDING MOMENTS IN DUANE ARNOLD POOL UNDER Maximum Bending Moment in Concrete" (in-lb./in)

Minimum****

Minimum* Load Load Load Load Load CaLculated Loca- Direc- Strength Comb. Safety Comb. Safety Comb. Safety Comb. Safety Comb. Safety Safety tion tion*** in.db./in. A Margin B Margin C Margin D Margin E Margin Margin Stab M, 1.404E6 0.299E6 4.70 0.309E6 4.54 0.273E6 5.14 0.319E6 4.40 0.245E6 5.73 4.40 M, 1.404E6 0.932E6 1.51 0.817E6 1.72 0.747E6 1.88 0.831E6 1.69 0.670E6 2.10 1.51 East M, 0.890E6 0.576E6 1.55 0.481E6 1.85 0.422E6 2.11 0.614E6 1.45 0.210E6 4.24 1.45 WalttM 0.688E6 0.306E6 2.25 0.305E6 2.26 0.273E6 2.52 0.309E6 2.23 0.251E6 2.74 2.23 00 Wa t M West M. 0.999E6 0.650E6 1.54 0.537E6 1.86 0.471E6 2.12 0.693E6 1.44 0.209E6 4.78 1.44 A WatL M, 0.772E6 0.286E6 2.70 0.290E6 2.66 0.259E6 2.98 0.289E6 2.67 0.251E6 3.08 2.66 South M, 1.151E6 0.285E6 4.04 0.281E6 4.10 0.244E6 4.72 0.329E6 3.50 0.148E6 7.78 3.50 Wal My 0.772E6 0.387E6 2.00 0.361E6 2.14 0.324E6 2.38 0.376E6 2.05 0.275E6 2.81 2.00

• The lowest anywhere in the region considered.

The highest in the region considered.

For the slab x = N/S direction; y = E/W direction; for the walls x = horizontal, y = vertical.

Safety margin defined as Maximum Allowable Moment/Calculated Moment

........... I...........

SUPPORT AT EXTERIOR WALL SUPPORT AT EXTERIOR WALL NORTH co IA 1111 11'.lII Illi: IF11I c1111111 11 I1T II.I1 II: .. 1 , 11'.lilT IIIII"1*11 I F II i , SI l - '.1 1 1 1 1 1i 1 .,0 11 1 1'.1 I~*

[CIJPI~ 8 3 1

II~~~ll~~~i i'~ II i ...

... i FICIJPE 8.3.2

I I III.~ ~....... ... .. I I II

-~. ..... ....

IF-GRE 8.3.3

9.0 RADIOLOGICAL EVALUATION 9.1 Fuel Handling Accident The potential radiological consequences at the Duane Arnold Energy Center (DAEC) exclusion area boundary (EAB) of a fuel handling accident in the secondary containment have been determined.

9.1.1 Assumptions and Source Term Calculations Evaluations of the accident were based on fuel of 4.6 wt% initial enrichment burned to 60,000 Mwd/mtU. The reactor was assumed to have been operating at 1692 Mw thermal power (102% of rated power) prior to shutdown; this yields a specific power of 25.82 kw/kgU. The fuel handling accident was assumed to result in the release of the gaseous fission products contained in the fuel/cladding gaps of 124 fuel rods in peak-power, 8x8R fuel assemblies. Gap inventories of fission products available for release were estimated using the release fractions identified in Regulatory Guide 1.251 except for Iodine-131, for which the release fraction is increased 20% in accordance with NUREG/CR 50092. Cooling time for the failed fuel prior to the accident was 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />.

The gaseous fission products that have the greatest impacts on the off-site doses following short fuel cooling 1 Regulatory Guide 1.25 (AEC Safety Guide 25), "Assumptions Used For Evaluating The Potential Radiological Consequences Of A Fuel Handling Accident In The Fuel Handling And Storage Facility For Boiling And Pressurized Water Reactors", March 23, 1972.

2 C. E. Beyer, et al., "Assessment of the Use of Extended Burnup Fuel in Light Water Power Reactors", NUREG/CR 5009, Pacific Northwest Laboratory, February, 1988.

9 - 1

reach saturation inventories during in-core operation. These inventories depend primarily on the fuel specific power over the few months immediately preceding reactor shutdown. In the highest power assemblies, the specific power and hence the inventory of iodine and xenon will be proportional to the peaking factor (taken as 1.50, per Reg Guide 1.25).

At the cooling time of 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br /> used in the DAEC calculations, three fourths of the thyroid dose comes from Iodine-131, while more than 90 percent of the whole-body dose comes from Xenon-133 and Xenon-135. Though these iodine and xenon isotopes are the major contributors to off-site doses, the contributions from other radionuclides are calculated and included in the overall dose values.

The present evaluation uses values for atmospheric diffusion factor (x/Q) and for filter efficiencies that have been specified previously for the DAEC. Core specific inventories (Curies per metric ton of uranium) of fission products were estimated with the ORIGEN-2 code 3 , based upon parameters stated earlier (specific power of 25.82 kw/kgU, initial enrichment of 4.6 wt% U, burnup of 60,000 Mwd/mtu, and a cooling time of 24 hours2.777778e-4 days <br />0.00667 hours <br />3.968254e-5 weeks <br />9.132e-6 months <br />). The results of the ORIGEN calculations for isotopes that contribute to the thyroid and whole-body doses are given in Table 9.1, while Table 9.2 lists pertinent data for the isotopes of interest. Data and assumptions used in the dose calculations are given in Table 9.3.

The following equation, from Reg Guide 1.25, was used to calculate the thyroid dose (D, in rads) from the inhalation of 3 n"ORNL Isotope GENeration and Depletion", ORNL/TM-7175, Oak Ridge National Laboratory, July, 1980.

9 - 2

radioiodine. Values for many of the terms in the equation are given in Table 9.2 and Table 9.3.

F9 Ii F P B Ri (X/Q)

Dose = -, where i DFP DF, Fg= fraction of fuel rod iodine Rj= dose conversion inventory in gap space factor (rads per curie)

I,= core iodine radionuclide inventory at time of the X/Q= atmospheric diffusion accident (curies) factor (sec per cubic meter)

F = fraction of core damaged so as to release iodine in the DF= effective iodine rod gap decon. factor for pool water P = core peaking factor DF,== effective iodine B = breathing rate (cubic meters decon. factor for per second) filters The equation given below was used to calculate the external whole-body dose from gamma radiation in the cloud of radionuclides released in the fuel-handling accident. The equation contains several of the terms defined above.

Dose, = E 0.25 (X/Q) F P Gi Ep.

In this expression, Gi is the gap inventory of the gaseous radionuclides of krypton and xenon, while the E term is the average energy per disintegration of each radionuclide (in Mev per disintegration, as given in Table 9.2). The equation assumes the noble gas decontamination factors in water and the charcoal filters are 1.0. The gap inventories of radioiodine make 9 - 3

negligible contributions to the whole body dose, D, , because of the large decontamination factors applicable to the iodines.

9.1.2 Results The doses at the DAEC EAB from the specified fuel handling accident are tabulated below. The doses are based on the release of all gaseous fission product activity in the gaps of 124 fuel rods in highest-power assemblies.

Thyroid dose, rad = 0.559 Whole-body dose, rem = 0.410 These potential doses are well within the exposure guideline values of 10 CFR Part 100, paragraph 11.

9.2 Solid Radwaste The necessity for resin replacement is determined primarily by the requirement for water clarity, and the resin is normally changed about once a year. No significant increase in the volume of solid radioactive wastes is expected with the expanded storage capacity. During re-racking operations, a small amount of additional resins may be generated by the pool cleanup system on a one-time basis, and the old racks themselves will be a form of solid radwaste.

9.3 Gaseous Releases Gaseous releases from the fuel storage area are combined with other plant exhausts. Normally, the contribution from the fuel storage area is negligible compared to the other releases and no significant increases are expected as a result of the expanded storage capacity.

9 - 4

9.4 Personnel Exposures During normal operations, personnel working in the fuel storage area are exposed to radiation from the spent fuel pool.

Operating experience has shown that the area radiation dose rates, which originate primarily from radionuclides in the pool water, are generally 1.0 to 2.0 mrem/hr, with a few areas such as the pool bridge showing dose rates of 4.0 mrem/hr.

Radiation levels in zones surrounding the pool are not expected to be affected significantly. Existing shielding around the pool (water and concrete walls) provide more than adequate protection, despite the slightly closer approach of the new racks to the walls of the pool and the increased amount of fuel to be stored. Preliminary calculations of dose rates in areas adjacent to the SFP show no long term increase. The calculation conservatively assumed a SFP full of freshly discharged fuel for a bounding analysis. The results of that analysis show that, due to the shielding of the concrete walls, dose rates in adjacent areas are not significantly increased. This analysis was very conservative in assumptions and, therefore, the conclusion is also very conservative.

Radionuclide concentrations typical of those found in pool water are shown in Table 9.4. During fuel reload operations, the concentrations might be expected to increase due to crud deposits spalling from spent fuel assemblies and to activities carried into the pool from the primary system. However, experi ence to date has not indicated a major increase as a consequence of refueling.

Operating experience has also shown that there have been negligible concentrations of airborne radioactivity and no increases are expected as a result of the expanded storage capacity.

9 - 5

No increase in radiation exposure to operating personnel is expected; therefore, neither the current health physics program nor the area monitoring system needs to be modified.

9.5 Anticipated Exposure During Re-racking All of the operations involved in re-racking will utilize detailed procedures prepared with full consideration of ALARA principles. Similar operations have been performed in a number of facilities in the past, and there is every reason to believe that re-racking can be safely and efficiently accom plished at the DAEC, with minimum radiation exposure to personnel.

Total occupational exposure for the re-racking operation is estimated to be between 6 and 12 person-rem, as indicated in Table 9.5. While individual task efforts and exposures may differ from those in Table 9.5, the total is believed to be a reasonable estimate for planning purposes.

Divers will be used only if necessary, but the estimated person rem burden includes a figure for their possible exposure.

The existing radiation protection program at the DAEC is adequate for the re-racking operations. Where there is a potential for significant airborne activity, continuous air samplers will be in operation. Personnel will wear protective clothing and, if necessary, respiratory protective equipment.

Activities will be governed by a Radiation Work Permit, and personnel monitoring equipment will be issued to each individual.

As a minimum, this will include thermoluminescent dosimeters and pocket dosimeters. Additional personnel monitoring equipment (i.e., extremity badges or alarming dosimeters) may be utilized as required.

9 - 6

Work, personnel traffic, and the movement of equipment will be monitored and controlled to minimize contamination and to assure that exposures are maintained ALARA.

In re-racking, the existing storage racks will be removed, then washed down in preparation for packaging and shipment. Estimates of the person-rem exposures associated with washdown and readying the old racks for shipment are included in Table 9.5. Shipping containers and procedures will conform to Federal DOT regulations and to the requirements of any state through which the shipment may pass, as set forth by the State DOT office.

9 - 7

11.5 Resource Commitment The expansion of the DAEC spent fuel pool capacity is expected to require the following primary resources:

Stainless steel: 75 tons Boral neutron absorber: 20 tons, of which 10 tons is Boron Carbide powder and 10 tons are aluminum.

The requirements for stainless steel and aluminum represent a small fraction of total world output of these metals (less than .001%).

Although the fraction of world production of Boron Carbide required for the fabrication is somewhat higher than that of stainless steel or aluminum, it is unlikely that the commitment of Boron Carbide to this project will affect other alternatives. Experience has shown that the production of Boron Carbide is highly variable and depends upon need and can easily be expanded to accommodate worldwide needs.

11.6 Environmental Considerations Due to the additional heat-load arising from increased spent fuel pool inventory, the anticipated maximum bulk pool temperature 0

increases from a previously-licensed 150Ot 164.6 F, as detailed in the calculations described in Section 5.0 of this report. This temperature is arrived at by assuming one fully-degraded spent fuel pool cooler (fully fouled, maximum allowable number of tubes plugged out of two available units). The assumption of only one maximum fouled cooler in-service is extremely conservative given the actual condition of the DAEC coolers. The resultant total heat load (worst case) is less than 19 million Btu/hr, which is less than .05% of the total plant heat loss to the environment and well within the capability of the plant cooling system.

11-8

The increased bulk pool temperature will result in an increase in the pool water evaporation rate. This has been calculated as sufficient to increase the relative humidity of the Fuel Handling Building atmosphere by less than 10%. This increase is within the capacity of Reactor Building Ventilation System. The net result of the increased heat loss and water vapor emission to the environment is negligible.

11.7 References for Section 11

[11.1] OT Position Paper for Review and Acceptance of Spent Fuel Storage and Handling Applications, USNRC (April, 1978).

(11.2] Electric Power Research Institute, Report No. NF-3580, May, 1984.

[11.3] "Spent Fuel Storage Options: A Critical Appraisal", Power Generation Technology, Sterling Publishers, pp. 137-140, U.K. (November, 1990).

11-9

Table 9.1 RESULTS OF ORIGEN-2 CALCULATIONS FOR RADIONUCLIDES OF IODINE, KRYPTON, AND XENON AT 24-HOURS COOLING TIME Radionuclide Curies per mtU 1-131 6.485 x 10' 1-132 8.143 x 105 1-133 6.261 x 105 1-134 3.180 x 102 1-135 1.038 x 10 Kr-85 1.464 104 Kr-85m 3.423 103 Kr-87 5.385 101 Kr-88 1.019 106 Xe-131m 7.803 101 Xe-133 1.320 106 Xe-133m 3.909 104 Xe-135 3.279 10s Xe-135m 1.663 104 9 - 8

Table 9.2 RADIONUCLIDE PROPERTIES USED IN THE FUEL HANDLING ACCIDENT ANALYSIS Dose Conversion, Radionuclide Rads/Curie E, (Mev/dis.)

Iodine-131 1.48 x 106 Iodine-132 5.35 x 104 Iodine-133 4.00 x 10' Iodine-134 2.50 x 104 Iodine-135 1.24 x 10' Krypton-85 0.002 Krypton-85m 0.158 Krypton-87 0.793 Krypton-88 1.980 Xenon-131m 0.020 Xenon-133 0.042 Xenon-13 3m 0.046 Xenon-135 0.248 Xenon-135m 0.431 9 - 9

Table 9.3 DATA AND ASSUMPTIONS FOR THE EVALUATION OF THE FUEL HANDLING ACCIDENT Core power level, Mw(t) 1692 Fuel enrichment, wt% U 4.6 Fuel burnup, Mwd/mtU 60,000 Specific power, kw/kgU 25.82 Fuel cooling time, hrs 24 Power peaking factor 1.50 No. of failed fuel rods 124 Core inventory released to gap, %

Iodine-131 12 Other iodines 10 Krypton-85 30 Xenon-133 10 Other xenons 10 Pool decontamination factors Iodine 100 Noble gases 1 Filter decontamination factors Iodine 100 Noble gases 1 Atmospheric diffusion factor (X/Q), sec/m' 2.07 x 10 Breathing rate, m'/sec 3.47 x 104 9 - 10

Table 9.4 TYPICAL CONCENTRATIONS OF RADIONUCLIDES IN SPENT FUEL POOL WATER Concentration, Nuclide A,/ml 4

Co-58 7 x 10 Co-60 4 x 10-'

Cs-134 2 x 10 4

Cs-137 4 x 10 9 - 11

Table 9.5 PRELIMINARY ESTIMATE OF PERSON-REM EXPOSURES DURING RE-RACKING Estimated Number of Person-Rem Step Personnel Hours, Exposureml Remove empty racks 5 40 0.5 to 1.0 Wash racks 3 10 0.08 to 0.2 Clean and Vacuum Pool 3 25 0.3 to 0.6 Remove underwater 4 5 0.4 to 0.8 appurtenances Partial installation 5 20 0.25 to 0.5 of new rack modules Move fuel to new racks 2 150 0.8 to 1.5 Remove remaining racks 5 120 1.5 to 3.0 Wash racks 3 30 0.2 to 0.4 Install remaining new 5 35 0.4 to 0.8 rack modules Decon and prepare old racks for shipment 4 80 1.0 to 2.0(2)

Total Exposure, person-rem 6 to 12 Assumes minimum dose rate of 2-1/2 mrem/hr (expected) to a maximum of 5 mrem/hr except for pool vacuuming operations, which assume 4 to 8 mrem/hr, and possible diving operations, which assume 20 to 40 mrem/hr.

(2) Maximum expected exposure, although details of preparation and packaging of old racks for shipment have not yet been deter mined.

9 - 12

10.0 BORALTM SURVEILLANCE PROGRAM 10.1 Purpose BoralTM, the neutron absorbing material incorporated in the spent fuel storage rack design to assist in controlling system reactivity, consists of finely divided particles of boron carbide (B4C) uniformly distributed in type 1100 aluminum powder, clad in type 1100 aluminum and pressed and sintered in a hot-rolling process. Tests simulating the radiation, thermal and chemical environment of the spent fuel pool have demonstrated the stability and chemical inertness of BoralTM (References [10.1.1]-[10.1.3]).

The accumulated dose to the BoralTM over the expected rack lifetime is estimated to be about 3 x 1010 to 1 x 101 rads depending upon how the racks are used and the number of full-core off-loads that may be necessary.

Based upon the accelerated test programs and a large body of in pool data, BoralTM is considered a satisfactory material for reac tivity control in spent fuel storage racks and is fully expected to fulfill its design function over the lifetime of the racks.

Nevertheless, the USNRC requires the Licensee to establish a surveillance program to monitor the integrity and performance of BoralTM on a continuing basis and to assure that slow, long-term synergistic effects, if any, do not become significant. The April 14, 1978 USNRC letter to all power reactor licensees (Reference

[10.1.4]), specifies that "Methods for verification of long-term material stability and mechanical integrity of special poison materials utilized for neutron absorption should include actual tests."

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The purpose of the surveillance program presented herein is to characterize certain properties of the BoralTM with the objective of providing data necessary to assess the capability of the BoralTM panels in the racks to continue to perform their intended function.

The surveillance program is also capable of detecting the onset of any significant degradation with ample time to take such corrective action as may be necessary.

In response to the need for a comprehensive BoralTM surveillance program to assure that the subcriticality requirements of the stored fuel array are safely maintained, a surveillance program has been developed incorporating certain basic tests and acceptance criteria. The BoralTM surveillance program depends primarily on representative coupon samples to monitor performance of the absorber material without disrupting the integrity of the storage system. The principal parameters to be measured are the thickness (to monitor for swelling) and boron content.

10.2 COUPON SURVEILLANCE PROGRAM 10.2.1 Coupon Description The coupon measurement program includes coupons suspended on a mounting (called a "tree"), placed in a designated cell, and surrounded by spent fuel. Coupons will be removed from the array on a prescribed schedule and certain physical and chemical properties measured from which the stability and integrity of the BoralTM in the storage cells may be inferred.

Each surveillance coupon will be approximately 4 inches wide and 6 inches long. The coupon surveillance program will use a total of 8 test coupons with each coupon mounted in a stainless steel jacket, simulating as nearly as possible, the actual in-service geometry, physical mounting, materials, and flow conditions of the BoralTM in the storage racks. The jacket (of the same alloy used in manufac-10-2

ture of the racks) will be closed by screws or clamps to allow easy opening with minimum possibility of mechanical damage to the Boralnm specimen inside. In mounting the coupons on the tree, the coupons will be positioned axially within the central 8 feet of the fuel zone where the gamma flux is expected to be reasonably uniform.

Each coupon will be carefully pre-characterized prior to insertion in the pool to provide reference initial values for comparison with measurements made after irradiation. The surveillance coupons will be pre-characterized for weight, length, width and thickness. In addition, two coupons (which need not be jacketed) will be preserved as archive samples for comparison with subsequent test coupon measurements. Wet chemical analyses of samples from the same lot of BoralTM will be available from the vendor as reference data for boron content.

10.2.2 Surveillance Coupon Testing Schedule To assure that the coupons will have experienced a slightly higher radiation dose than the BoralTM in the racks, the coupon tree is surrounded by freshly discharged fuel assemblies at each of the first five refuelings following installation of the racks.

Beginning with the fifth load of spent fuel, the fuel assemblies will remain in place for the remaining lifetime of the racks. A sample coupon management schedule is shown in Table 10.1.

At the time of the first fuel off-load following installation of the coupon tree, the (8) storage cells surrounding the tree shall be loaded with freshly-discharged fuel assemblies that had been among the higher specific power assemblies in the core. Shortly before the second reload, the coupon tree is removed and a coupon removed for evaluation. The coupon tree is then re-installed and, at reload, again surrounded by freshly discharged fuel assemblies.

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This procedure is continued for the third, fourth, and fifth off loading of spent fuel (except that a coupon is not pulled at the fourth refueling). From the fifth cycle on, the fuel assemblies in the (8) surrounding cells remain in place.

Evaluation of the coupons removed will provide information of the effects of the radiation, thermal and chemical environment of the pool and by inference, comparable information on the Boral TM panels in the racks. Over the duration of the coupon testing program, the coupons will have accumulated more radiation dose than the expected lifetime dose for normal storage cells.

Coupons which have not been destructively analyzed by wet-chemical processes, may optionally be returned to the storage pool and re mounted on the tree. They will then be available for subsequent investigation of defects, should any be found.

10.2.3 Measurement Program The coupon measurement program is intended to monitor changes in physical properties of the BoralTm absorber material by performing the following measurements on the pre-planned schedule:

Visual Observation and Photography,

• Neutron Attenuation (optional),

Dimensional Measurements (length, width and thickness),

Weight and Specific Gravity, and

• Wet-chemical analysis (Optional).

The most significant measurements are thickness (to monitor for swelling) and boron content. Boron content should be measured in 10-4

the event that loss of boron is suspected, in which case neutron attenuation* or wet-chemical analysis (a destructive gravimetric technique for total boron only) may be performed on the coupon to evaluate the boron content.

The acceptance criteria for these measurements are as follows:

• A decrease of no more than 5% in Boron-10 con tent, as determined by neutron attenuation, is acceptable. (This is tantamount to a requirement for no loss in boron within the accuracy of the measurement.)

• An increase in thickness at any point should not exceed 10% of the initial thickness at that point.

Changes in excess of either of these two criteria would require investigation and engineering evaluation which may include early retrieval and measurement of one or more of the remaining coupons to provide corroborative evidence that the indicated change(s) is real. If the deviation is determined to be real, an engineering evaluation shall be performed to identify further testing or any corrective action that may be necessary.

The remaining measurement parameters serve a supporting role and should be examined for early indications of the potential onset of Boraln degradation that would suggest a need for further attention and possibly a change in measurement schedule. These include (1) visual or photographic evidence of unusual surface pitting, corrosion or edge deterioration, or (2) unaccountable weight loss in excess of the measurement accuracy).

• Neutron attenuation measurements are a precise instrumental method of chemical analysis for Boron-10 content using a non destructive technique in which the percentage of thermal neutrons transmitted through the panel is measured and compared with pre determined calibration data. Boron-10 is the nuclide of principal interest since it is the isotope responsible for neutron absorption in the Boral panel.

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Procedures for coupon surveillance measurement have been provided to the Licensee by the rack supplier's vendor, Nusurtec, Inc.

10.3 References for Section 10

[10.1.1] "Spent Fuel Storage Module Corrosion Report",

Brooks & Perkins Report 554, June 1, 1977

[10.1.2] "Suitability of Brooks & Perkins Spent Fuel Storage Module for Use in PWR Storage Pools",

Brooks & Perkins Report 578, July 7, 1978

[10. 1.3] "BoralTm Neutron Absorbing/Shielding Material Product Performance Report", Brooks & Perkins Report 624, July 20, 1982

[10.1.4] USNRC Letter to All Power Reactor Licensees, transmitting the "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", April 14, 1978 10-6

Table 10.1 COUPON MEASUREMENT SCHEDULE Cou on Refuelinac~) After Rerack 2) 1 1st(

2 2nd 2 )

2 3 3rd )

4 5th2 )

5 8th 6 11th 7 14th 8 16th C') Remove coupons for evaluation within the 1 or 2 months preceding the next refueling.

(2) Place freshly discharged fuel in the 8 surrounding cells at the beginning of the 1st, 2nd, 3rd, 4th, and 5th refueling cycles after completion of reracking.

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11.0 ENVIRONMENTAL COST-BENEFIT ASSESSMENT 11.1 Introduction Article V of the USNRC OT Position paper [11.1] requires the submittal of a cost/benefit analysis for the chosen fuel storage capacity enhancement method. This section abstracts the analyses and evaluations made by IELP before selecting reracking as the most viable alternative.

11.2 Imperative for Reracking The specific need to increase the limited existing storage capacity of the DAEC spent fuel pool is based on the continually increasing inventory in the pool, the prudent requirement to maintain full core off-load capability, and a lack of viable economic alternatives.

Reference is made to Section 1 of this report wherein the current and projected fuel discharges in the DAEC spent fuel pool are tabulated. It is seen that the DAEC fuel pool will lose the capacity to discharge one full core (368 fuel assemblies) in 1998.

The projected loss of storage capacity in the DAEC pool would affect IELP's ability to operate the DAEC reactor. DAEC does not have an existing or planned contractual arrangement for a third party fuel storage or fuel reprocessing.

11.3 Appraisal of Alternative Options IELP has determined that reracking is by far the most viable option for the DAEC pool in comparison to other alternatives.

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The key considerations in evaluating the alternative options were:

• Safety: minimize the number of fuel handling steps

• Economy: minimize total installed and O&M cost

• Security: protection from potential saboteurs, natural phenomena

- Non-intrusiveness: minimize required modification to existing systems

• Maturity: extent of industry experience with the technology

• ALARA: minimize cumulative dose due to handling of fuel Reracking was found by IELP to be the most attractive option in respect to each of the foregoing criteria. An overview of the alternative technologies considered by IELP is provided in the following:

Rod Consolidation Rod consolidation has been shown to be a feasible technology. Rod consolidation involves disassembly of spent fuel, followed by the storage of the fuel rods from two assemblies into the volume of one and the disposal of the fuel assembly skeleton outside of the pool (this is considered a 2:1 compaction ratio). The rods are stored in a stainless steel can that has the outer dimensions of a fuel assembly. The can is stored in the spent fuel racks. The top of the can has an end fixture that matches up with the spent fuel handling tool. This permits moving the cans in an easy fashion.

Rod consolidation pilot project campaigns in the past have consisted of underwater tooling that is manipulated by an overhead crane and operated by a maintenance worker. This is a very slow and repetitive process.

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The industry experience with rod consolidation has been mixed thus far. The principal advantages of this technology are: the ability to modularize, compatibility with DOE waste management system, moderate cost, no need of additional land and no additional required surveillance. The disadvantages are: potential gap activity release due to rod breakage, potential for increased fuel cladding corrosion due to some of the protective oxide layer being scraped off, potential interference of the (prolonged) consolidation activity which might interfere with ongoing plant operation, and lack of sufficient industry experience. The drawbacks associated with consolidation are expected to diminish in time, and IELP views rod consolidation as a possibly viable sequel to reracking for the DAEC pool.

On-Site Cask Storage Dry cask storage is a method of storing spent nuclear fuel in a high capacity container. The cask provides radiation shielding and passive heat dissipation. Typical capacities for BWR fuel range from 40 to 60 assemblies that have been removed from the reactor for at least five years. The casks, once loaded, are then stored outdoors on a seismically qualified concrete pad. The pad will have to be located away from the secured boundary of the site because of site limitations. The storage location will be required to have a high level of security which includes frequent tours, reliable lighting, intruder detection (E-field), and continuous visual monitoring.

The casks, as presently licensed, are limited to 20 year storage service life and are for storage only. Once the 20 years has expired the cask manufacturer or the utility must recertify the cask or the utility must remove the spent fuel from the container.

Work is also continuing on providing a dual purpose cask that will 11-3

be capable of long storage and then transport. These casks tend to have a reduced capacity, thus requiring more containers in order to provide the same amount of storage space.

The plant must provide for a decontamination facility where the outgoing cask can be decontaminated for release.

There are several plant modifications required to support cask use.

Tap-ins must be made to the gaseous waste system and chilled water to support vacuum drying of the spent fuel and piping must be installed to return cask water back to the spent fuel pool/cask pit. A seismic concrete pad must be made to store the loaded casks. This pad must have a security fence, surveillance protection, a diesel generator for emergency power and video surveillance.

Presently no cask has dual certification; it is only good for either storage or transportation. Dual purpose casks, if and when certified, will have reduced capacity thus increasing the quantity required. As a result, dual purpose casks are considered to be inferior overall to present day casks.

Modular Vault Dry Storage Vault storage consists of storing spent fuel in shielded stainless steel cylinders in a horizontal configuration in a reinforced concrete vault. The concrete vault provides radiation shielding and missile protection. It must be designed to withstand the postulated seismic loadings for the site.

A transfer cask is needed to fetch the storage canisters from the fuel pool. The plant must provide for a decontamination bay to decontaminate the transfer cask, and connection to its gaseous waste system and chilled water systems. A collection and delivery system must be installed to return the pool water entrained in the 11-4

canisters back to the fuel pool. Provisions for canister drying, helium injection, handling and automatic welding are also necessary.

The storage area must be designed to have a high level of security similar to that of the nuclear plant itself. Due to the required space, the vault secured area must be located outside the secured perimeter. Consideration of safety and security requires it to have its own video surveillance system, intrusion detection, and an autonomous backup diesel generator power source.

Some other concerns relating to the vault storage system are:

inherent eventual "repackaging" for shipment to the DOE repository, the responsibility to eventually decommision the new facility, large "footprint" (land consumption) and high cost.

Horizontal Silo Storage A variation of the horizontal vault storage technology is more aptly referred to as "horizontal silo" storage. This technology suffers from the same drawbacks which other dry cask technologies do, namely, i) No fuel with cladding defects can be placed in the silo.

(ii) Potential for eventual repackaging at the site (iii) Relatively high cumulative dose to personnel in effecting fuel transfer (compared to reracking)

(iv) Compatibility of reactor/fuel building handling crane with fuel transfer hardware (v) Potential incompatibility with DOE shipment for eventual off-site shipment (vi) Potential for sabotage 11-5

New Fuel Pool Constructing and licensing a new fuel pool is not a practical alternative for DAEC since such an effort may take up to 10 years.

Moreover, the cost of this option is prohibitively high.

An estimate of relative costs in 1991 dollars for the aforementioned options is provided in the following:

Reracking: $6 million Horizontal Silo (NUHOMS): $8 to 12 million Rod consolidation (rental): $10 million Metal cask: $20 to 40 million Modular vault: $30 million New fuel pool: $100 million IELP's estimate of comparative costs of various options is consistent with other published industry data [11.2,11.3]. EPRI's assessment [11.2] is tabulated below.

Type of Storage Canital Costs (S/KaU)

Rerack $10-25 Fuel consolidation $25-55 Dry cask storage $88 Storage vault $79 New pool $115 To summarize, there are no acceptable alternatives to increasing the on-site spent fuel storage capacity of DAEC. First, there are no commercial independent spent fuel storage facilities operating in the U.S. Second, the adoption of the Nuclear Waste Policy Act (NWPA) created a de facto throw-away nuclear fuel cycle. Since the 11-6

cost of spent fuel reprocessing is not offset by the salvage value of the residual uranium, reprocessing represents an added cost for the nuclear fuel cycle which already includes the NWPA Nuclear Waste Fund fees. In any event, there are no domestic reprocessing facilities. Third, the alternative storage options are either much more expensive or relatively unproven.

11.4 Cost Estimate The proposed construction contemplates the reracking of the DAEC spent fuel pool using free-standing, high density, poisoned spent fuel racks. The engineering and design is completed for full reracking of the DAEC pool. This rerack will provide sufficient DAEC pool storage capacity to maintain a full core off-load capability to the end of the licensed reactor life.

The total capital cost is estimated to be approximately $6 million as detailed below.

Engineering, design, project management $2 million Rack fabrication $3 million Rack installation $1 million As described in the preceding section, many alternatives were considered prior to proceeding with reracking, which is not the only technical option available to increase on-site storage capacity. Reracking does, however, enjoy a definite cost and reliability advantage over other technologies.

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