ML20246M471
| ML20246M471 | |
| Person / Time | |
|---|---|
| Site: | Saint Lucie  |
| Issue date: | 01/25/1989 |
| From: | Woody C FLORIDA POWER & LIGHT CO. |
| To: | NRC OFFICE OF ADMINISTRATION & RESOURCES MANAGEMENT (ARM) |
| References | |
| L-88-38, OLA-A-006, OLA-A-6, NUDOCS 8903270159 | |
| Download: ML20246M471 (155) | |
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NUCLEAR REGULATORY COMMISSION
- p. o sox 14000 a.,No BE A;m rL 3M;E X O Docket No. 5d ~ 3 6~6 A Official Exh. No.
In the rnatter of Ouidm /b V 4d - 5/, M ECO /_
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'89 ya 22 P7 2 Contractor -
DATE /VEb Other - Viitness EE 2 9. N Reporter _ l~'lllic- U ey w L-88-38 10 CFR 50.90 U. S. Nuclear Regulatory Commission Attn: Document Control Desk T P@ff $
Washington, D. C. 20555 - y qg l;1044 Gentlemen: ggy [([3 Re: St. Lucie Unit 1 Docket No. 50-335 i Scent Fuel Rerack b(
O By letter L-87-245, dated June 12, 1987, Florida Power &
Light Company (FPL) submitted a proposed license amendment to permit replacement of the spent fuel pool racks at St. Lucie Unit 1 to ensure that sufficient future capacity exists for ,
l storage of spent fuel. I l
In various correspondence subsequent to the above submittal I date, the NRC Staff requested additional information it needed to continue its review of the proposed license amendment. FPL has responded to all of these information requests. Additionally, FPL has met with the NRC on several occasions discussing various aspects of the St. Lucie Unit 1 spent fuel pool rorack.
Enclosed is Revision 1 to the Gpent Fuel Storage Facility Modification Safety Analysis Report submitted in FPL's June
.12 , 1987 application and replaces Attachment 3 to that letter in its entirety. This revision incorporates changes to certain sections resulting from the FPL to NRC correspondence l and the meetings with the Staff.
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8903270159 890125 l PDR ADDCK 05000335 C PDR \,
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EJW/020.SFR b an FPL Group compam L
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U. S. Nuclear Regulatory Commission L-88-38 Page two If additional information is required, please contact us.
Very truly yours,
(:. ce' C. O. y Execut ve Vice President COW /EJW/gp Enclosure cc: Dr. J. Nelson Grace, Regional Administrator, Region II, USNRC
( Senior Resident Inspector, USNRC, St. Lucie Plant
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OL Because of the size of attachments to NRC letter dated January 29, 1988
" Spent Fuel Rcrack", only th e cover letter is being distributed.
Copies are located in the Nuclear 1.icensing File Room and the General Office Executive Doc Files.
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FLORIDA POWER & LIGHT COMPANY ST LUCIE PLANT - UNIT NO. 1 SPENT FUEL STORAGE FACILITY MODIFICATION SAFETY ANALYSIS REPORT DOCKET NO. 50-335
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. TABLE OF CONTENTS PAGE SECTION
1.0 INTRODUCTION
1-1 1.1 License Amendment Requested 1-1 1.2 Current Status 1 1.3 Interfaces with Other Organizations 1-1
.)
1.4 Summary of Report 1-1 j 1.5 ' Conclusions 1-2' 1.6 References 1-2 2.0
SUMMARY
OF RACK DESIGN 2-1 2.1 Existing Racks 2-1 2.2 New High Density Racks 2-1' 3.0 NUCLEAR AND THERMAL-HYDRAULIC CONSIDERATIONS 3-1 3.1 Neutron Multiplication Factor 3-1 3
3.1.1 Normal Storage 3-1 3.1.2 Postulated Accidents 3-2 3.1.3 Calculation Methods 3-2 1 3.1.4 Rack Modification 3-9 3.1.5 Acceptance Criteria for criticality 3-10 3.2 Decay Heat Calculations for the Spent Fuel Pool ~
3-10 (Bulk) i 3.2.1 Spent Fuel Pool Cooling System Design 3-10 i 3.2.2 Decay Heat Analyses 3-11 3.2.3 Spent Fuel Pool Makeup 3-14 3.3 Thermal-Hydraulic Analyses for the Spent Fuel 3-15 Pool (Localized) 3.3.1 Basis. 3-15 3.3.2 Model Description 3-15 3.3.3 Cladding Temperature 3-16 O- 1
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TABLE OF CONTENTS (Cont'd)
PAGE SECTION
]N/ 3.4 Potential Fuel and Rack Handling Accidents 3-16 Rack Module Mishandling 3-17 3.4.1 3.4.2 Temporary Construction Crane Drop 3-17 3.4.3 Loss of fool Cooling (Storage Rack Drop) 3-17 3.5 Technical Specification Changes 3-17 3.6 References 3-18 4.0 MECHANICAL, MA7ERIAL,' AND STRUCTURAL CONSIDERATIONS 4-1 4
4.1 Description of Structure 4-1 4.1.1 Description of Fuel Handling Building 4-1 l l 4.1.2 Description of Spent Fuel Racks 4-1 4.2 Applicable Codes, Standards, and Specifications 4-5 4.2.1 NRC Documents 4-5 4.2.2 Industry Codes and Standards 4-7 4.3 Seismic and Impact Loads 4-8 )
l 4.4 Loads and Load Combinations 4-9 4.4.1 Spent Fuel Pool 4-9 4.4.2 Spent Fuel Racks 4-11 4.5 Design and Analysis Procedures 4-12 4.5.1 Design and Analysis Procedures for Spent 4-12 Fuel Pool 4.5.2 Design and Analysis Procedures for Spent 4-13 Fuel Storage Racks l
l 4.6 Structural Acceptance Criteria 4-20 ;
4.6.1 Structural Acceptance Criteria for Spent 4-20 Fuel Pool Structure 4.6.2 Structural Acceptance Criteria for Spent 4-23 Fuel Storage Racks 4.6.3 Fuel Handling Crane Uplif t Analysis 4-27 4.6.4 Impact Analysis 4-27 ,
4.6.5 Weld Stresses 4-27 !
4.6.6 Summary of Mechanical Analysis 4-28 4.6.7 Definition of Terms Used In Section 4 4-29 ;
4.6.8 Lateral Rack Movement 4-30 ii
TABLE OF CONTENTS (Cont'd)
PAGE SECTION Materials, Quality Control, and Special Con- 4-30 4.7 struction Techniques Construction Materials 4-30 4.7.1 4-30 4.7.2 Neutron Absorbing Material Quality Assurance 4-30 4.7.3 4-30 4.7.4 Construction Techniques 4-32 4.8 Testics and In-Service Surveillance 4-32 4.8.1 Program Intent Description of Specimens. 4 4.8.2 4-32 4.8.3 Specimen Evaluation 4-33 4.9 References 5-1 5.0 COST / BENEFIT AND ENVIRONMENTAL ASSESSMENT Cost / Benefit and Thermal Assessment 5-1 5.1 Need for Increased Storage capacity 5-1 5.1.1 5.1.2 Estimated Costs 5-1 Consideration of Alternatives 5-2 5.1.3 5-2 5.1.4 Resources Committed Thermal Impact on the Environment 5-2 5.1.5 5.2 Radiological' Evaluation 5-3 5.2.1 Solid Radwaste 5-3 5.2.2 Gaseous Releases 5-3 Personnel Exposure 5-3 5.2.3 .
5-4 5.2.4 Radiation Protection During Re-Rack Activities 5.2.5 Rack Disposal 5-5 5-6 5.3 Accident Evaluation 5,3.1 Spent Fuel Handling Accidents 5-6 1 5.3.2 Fuel Decay 5-8 5.3.3 Loads Over Spent Fuel 5-9 5.3.4 Temperature and Water Density Effects 5-9 5.3.5 Conclusions 5-9 5.4 References 5-10 iii O
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TABLE OF CONTENTS
-i LIST OF TABLES PAGE
-* TABLE TITLE Design Data 2-3 2-1 2-2 Table of Module Data 2-4 2 Module Dimensions and Weight 2-5 2-3 I 3-1 Summary of Critical Safety Analyses 3-20 Minimus Burnup Values 3-21 3-2 3-3 Reactivity Effects of Abnormal and 3-22 Accident Conditions I
3-4 Fuel Burnup Values for Required Reactivities (k. ) 3-23 with Fuel of Various Initial Enrichments 3-5 Comparison of Cold, Clean Reactivities 3-24 l Calculated at 36.5 Mwd /kgU Burnup and 4.5% 1 Enrichment l 3-6 Estimated Uncertainties in Reactivity 3-25 Due To Fuel Depletion Effects 3-7 Long Term Changes in Reactivity in Storage Rack 3-26 3-8 Design Basis (Limiting) Fuel Assembly 3-27 Specifications (CE 14 x 14) l 3-9 Thermal / Hydraulic Cases Treated 3-28 i 3-10 Peaking Factor Data _ 3-29 3-11 Essential Heat Transfer Data for the Fuel 3-30 Pool Heat Exchanger 3-12 Power Generation Ratio Previously 3-31 !
Discharged Batches 3-13 Bulk Pool Temperature vs. Time During 3-32 Normal Discharge 3-14 Pool Bulk Temperature vs. Time Subsequent to 3-33 Completion of Discharge iv A
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TABLE OF CONTENTS
. LIST OF TABLES (Cont'd)
TABLE TITLE PAGE 3-15 Loss of Cooling after Completion of 3-34 Normal Refueling Discharge 3-16 Bulk Pool Temperature vs. Time During 3-35 Full Core Discharge 3-17 Pool Bulk Temperature vs. Time Subsequent 3-36 to Completion of Full Core Discharge 3-18 ' Loss of Cooling.After Completion of 3-37 Full Core Discharge 3-19 Local and Cladding Temperature Data 3-38 4-1 Boraflex Experience for High Density Racks 4-34 4-2 Maximum Stress Summary 4-35 4-3 Streas/ Strain Summary for Liners and Anchors 4-36 4-4 Soil Bearing Stresses' 4-37 4-5 Stability Safety Factors 4-38 OO 4-6 Degrees of Freedos 4 4-7 Numbering System for Gap Elements and 4-40 Friction Elements 4-8 Rack Material Data 4-41 4-9 Adjustable Height Support Material Data 4-42 4-10 Bounding Values for Stress Factors 4-43 5-1 Nuclear Fuel Discharge Information 5-11 St Lucie Unit 1 5-2 Annual Fuel Sav.ings Attributed to 5-12 l St Lucie Unit No.1 !
5-3 Gaseous Releases From Fuel Handling Building 5-13 5-4 Gamma Isotopic Analysis Spent Fuel Pool Water 5-14 i
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TABLE OF CONTENTS LIST OF TABLES (Cont'd) g PAGE
- TABLE TITLE Anticipated Domes During Reracking 5-15 5
Effect of Temperature and Void on Calculated 5 5-6
. Reactivity of Storage Rack Spent Fuel Pool Purification ~ System 5-17 5 .
l Radionuclides Analysis Report Resin Activity .
Spent Fuel Pool Airborne Activity Radionuclides 18 5-8 Analysis Report O
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TABLE OF CONTENTS LIST OF FIGURES .
K L FIGURE TITLE 2-1 Pool Layout 2-2 Typical Rack Elevation - Region 1 2-3 Typical Rack Elevation - Region 2 3-1 Acceptable Burnup Domain in Region 2 of the l l
St Lucie Plant Spent Fuel Storage Racks 3-2 Region 1 Storage Cell Geometry 3-3 Region 2 Storage Cell Geometry 3 i
3-4 Comparison of Depletion Calculations for Fuel of 4.5% Initial Enrichment 3-5 Bulk Pool Temperature Model for Code BUIKTEM
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3-6 Idealization of Rack Assembly l 3-7 Thermal Chimney Flow Model Channel Element - Regions 1 and 2 f h 4-1 4-2 Composite Box Assembly - Region 1 4-3 Gap Element - Region 1 4-4 Typical Cell Elevation - Region 1 1
4-5 Typical Cell Elevation - Region 2 )
4-6 Adjustable Support -
l 4-7 3 x 3 Typical Array - Region 1 4-8 3 x 3 Typical Array - Region 2 4-9 Fuel Handling Building Spectra Envelope Curves 4-10 Mat Plan and Section j 4-11 Model Overall View 4-12 North South SSE .
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J TABLE OF CONTENTS LIST OF FIGURES ,
FIGURE TITLE 4-13 East West SSE 4-14 Vertical SSE 4-15 Schematic Model for DYNARACK 4-16 Rack to Rack Impact Springs 4-17 Impact Springs Arrangement at Node'i 4-18 Spring Mass ' Simulation for Two-Dimensional Motion 4-19 Test Coupon
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' REVISION PAGE REVISION PAGE I
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1.0 inn 0 DUCTION 1.1 LICENSE AMENDMENT REQUESTED t
( Florida Power & Light Company (FPL) has contracted for the design and manufacture of new spent fuel storage racks to be placed into the spent fuel pool of St Lucie Unit No. 1. The purpose of the new racks is to increase the amount of spent fuel that can be stored in the existing spent fuel pool. The racks are designed so that they can store spent fuel assemblies in a high density array. Therefore, FPL hereby requests that a License Amendment be issued to the St Lucie Unit No. 1 Facility Operating License DPR-67(1) to include installation and use of new storage racks that meet the criteria contained herein. This Safety Analysis Report (SAR) has been prepared to support this request for license amendment.
1.2 CURRENT STATUS The existing racks in the spent fuel pool at St Lucie Unit No.1 have 728 total storage cells. With the presently available storage cells, St Lucie Unit No.1 lost the full-core reserve storage capability after the seventh refueling, which was completed in the spring of 1987. To correct this situation and provide sufficient capacity at St Lucie Unit No.1 to store discharged fuel assemblies, FPL plans to replace the existing storage racks with new high density spent fuel storage racks. The design of the new racks will allow for more dense storage of spent fuel, thus enabling the existing pool to store more fuel in the spent fuel pool. The new high density racks j have a usable storage capacity of 1706 cells, extending the full-core-reserve storage capability until the year 2009.
l h If a full core offload is required in the interim, prior to the installation of the new racks, FPL intends to transfer enough of the oldest spent fuel from St. Lucie Unit 1 to St. Lucie Unit 2 to allow full core offload. A proposed license amendment to allow spent fuel transfer was submitted in July 1986(2) and is being reviewed by the NRC.
1.3 INTERFACES WITH OTHER ORGANIZATIONS FPL has overall responsibility for'this modification. Holtec International has designed the new spent fuel storage racks. Joseph Oat (J0) is responsible for the fabrication of the new spent fuel storage racks and the evaluation of those racks under accident conditions. Ebasco Services, Inc. is responsible for the building structural analysis, the evaluation of the spent fuel cooling 1 l
system and the related sceident evaluations. The installer, who will be chosen later, is responsible for the installation of the new spent fuel pool racks. -
1.4
SUMMARY
OF REPORT This Safety Analysis Report follows the guidance of the NRC position paper entitled, "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," ated April 14,.1978, as amended by the NRC letter dated January 18, 1979(3 Sections 3.0 through 5.0 of this report are consistent with the section/ subsection format and content of the NRC position paper, Sections III through V.
1-1 0076L/0011L
- q The nuclear and-therral-hydraulic espacts of the report (Section 3.0) address the neutron multiplication factor, considering normal storage and handling of spent fuel as well as postulated accidents with respect to criticality and the t ability of the spent fuel pool cooling system to maintain sufficient cooling.
Movement of spent fuel stored in the spent fuel pool during removal of the j present racks and installation of the new racks is also addressed. i l
Section 4.0, which describes the mechanical, material and structural aspects of the new racks, contains information concerning the capability of the fuel assemblies, storage racks, and spent fuel pool system to withstand the effects of natural phenomena and other service loading conditions.
The environmental aspects of the report (Section 5.0) concern the thermal and radiological release from the facility under normal and accident conditions. i This section also addresses the occupational radiation exposures, generation of radioactive waste, need for expansion, commitment of material and {
non-naterial resources, and a cost-benefit assessment.
1.5 CONCLUSION
S On the basis of the evaluations and information presented in this report, plus operating experience with high density fuel storage at St Lucie Unit 2 and Turkey Point Unit 3, FPL concludes that the proposed modification of St Lucie Unit No.1 spent fuel storage facilities provides safe spent fuel storage, andthatthemodificationiscongtentwiththefacilitydesignandoperating and operating license.
criteria as provided in the FSAR 1
1.6 REFERENCES
- 1. St Lucie Unit No.1 Facility Operating Licenses DPR 67, Docket j No. 50-335.
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- 3. Nuclear Regulatory Commission, Letter to All Power Reactor Licensees, from B. K. Grimes, April 14, 1978, "0T Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," as amended by the NRC letter dated January 18, l 1979.
I 4 St Lucie Plant Unit No.1 Updated Final Safety Analysis Report, Docket No. 50-335.
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0-1-2 0076L/0011L
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2.0
SUMMARY
OF RACK DESIGN 2.1 EXISTING RACKS The spent fuel pool at St. Lucie Unit 1 presently contains spent fuel assembly j storage racks which are designed to provide storage locations for up to 728 i fuel assemblies. The racks are designed to maintain the stored fuel in a safe, coolable, and suberitical configuration during normal and abnormal conditions.
The present storage racks are a rectangular array composed of 14 modules.
Each storage rack module is self supporting and rests on stainless steel pads. The present racks are free standing in that they are neither bolted nor welded to the floor, nor are they attached to the pool walls. The interface with the pool boundaries is designed to transfer normal and shear loads via the rack supports into the pool bottom slab.
Each fuel assembly storage module is composed of rectangular storage cavities fabricated from one quarter inch thick stainless steel plate, with each cavity capable of accepting one fuel assembly. The fuel assembly storage cavities have lead-in surfaces at the top to provide guidance for insertion of fuel assemblies. The cavities are open at the top and bottom to provide a flow path for convective cooling of spent fuel assemblies through natural circulation. The fuel assembly storage cavities are connected by a chevron i grid structure to form modules which limit structural deformations and maintain a nominal center-to-center spacing of 12.53 inches between adjacent storage cavities during design conditions including seismic. l For further information on the existing spent fuel storage racks see Section h 9.1.2 in the St Lucie Unit No.1 updated FSAR.
2.2 NEW HIGH DENSITY RACKS l The new high density spent fuel storage racks consist of individual cells with 8.65 inch by 8.65 inch (nominal) square cross-section, each of which accommodates a single Combustion Engineering or Exxon PWR fuel assembly or equivalent, f rom either St. Lucie Unit 1 or Unit 2. A total of 1706 cells are arranged in 17 distinct modules of varying sizes in two regions. Region 1 is !
designed for storage of new fuel assemblies with enrichments up to 4.5 weight i percent U-235. Region 1 is also designed to store fuel assemblies with enrichments up to 4.5 weight percent U-235 that have not achieved adequate burnup for Region 2. The Region 2 cells are capable of accommodating fuel assemblies with various initial enrichments which have accumulated minimum burnups within an acceptable bound as discussed in this report. For example, corresponding to 4.5 and 4.0 percent initial enrichments, the minimum required burn-ups for safe storage in Region 2 are 36.5 and 30.9 MWD /KgU, respectively. Figure 2-1 shows the arrangement of the rack modules in the spent fuel pool.
The high density racks are engineered to achieve the dual objective of maximum protection against structural loadings (arising from ground motion, thermal stresses, etc.) and the maximization of available storage locations. In general, a greater width-to-height aspect ratio provides greater margin against rigid body tipping. Hence, the modules are made as large as possible within the constraints of transportation and site handling capabilities.
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2-1 0076L/0011L
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As shown in Figure 2-1, there are 17 discrete modules arranged in the fuel aool. Each rack module is equipped (see Figures 2-2 and 2-3) with girdle
(% .rs, 3/4-inch thick by 3-1/2 inches high. The nominal gap between adjacent l j
kI module walls is 1-1/2 inches. The modules make surface contact between their contiguous walls at the girdle bar locations and thus maintain a specified gap I
)i between the cell walls. Table 2-1 gives the relevant design data on each region. The modules in the two regions are of eight different types. Tables 2-2 and 2-3 summarize the physical data for each module type. j l
i The poison in Regions 1 and 2 is Boraflex. The use of this absorber material is to preclude inadvertent criticality.
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0076L/0011L 2-2 Revision 1 i
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i TABLE 2-1 DESIGN DATA Min. B-10 Flux Trap Region Cell Pitch Loading Gap (nominal inch) (areal density) (nominal inch) l i
i 10.12 .020 ga/cm 2 1.12 1
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2 8.86 .007 ga/cm 2 0. 0 OO I
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'V TABLE 2-2 TABLE OF MODULE DATA NO. OF NO. OF CELLS CELLS 70TAL NO..
NO. OF IN N-S IN E-M OF CELLS
. MODULE'I.D. MODULES DIRECTION DIRECTION PER MODULE Region 1 2 9 9 81 Al and A2 Region 1 2 9 10 90 l-B1 and B2 Region 2 4 13 9 117 C1, C2, C3, C4 Region 2 3 13 8 104 l
h D1, D2, D3 Region 2 2 11 8 88 El and E2 l Region 2 1 12 8 96 F1
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Region 2 2 12 9 108 '
G1 and C2 Region 2* 1 13 8 96 H1
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- Cells missing in this module due to sparger. l Refer to Figure 2-1.
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0076L/0011L 2 -4 Revision 1 1
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TABLE 2-3 MODULE DIMENSIONS AND WEIGHT NOMINAL CROSS-6ECTION* ESTIMATED DRY l DIMENSIONS WEIGHT (1bs)
N -S E-47 PER MODULE MODULE I .D.
l Region 1 90-1/4" 90-1/4" 26,700 Al and A2 l
Region 1 90-1/4" 100-7/16" 29,800 l
B1 and B2 Region 2 115-11/16" 80-1/6" 24,100 C1, CZ, C3, C4 l Region 2 115-11/16" 71 -3/16" 21,500
{ D1, D2, D3 l Region 2 97-7/8" 71 -3/16" 18,200 El and E2 l Region 2 106 -3/4" 71 -3/16" 19,800 F1 Region 2 106 -3/4" 80-1/16" 22,300 l G1 and G2 l Region 2 115-11/16" 71 -3/16" 19,800 H1
- Excluding girdle bars l.
1 0076L/0011L 2 -5 Revision 1
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EVISION 1 FLORIDA POWER & LIGHT OOMPANY ST. LUCtE PLANT UNIT 1 l Q POOL LAYOUT FIGURE 21 i
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i FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 O TYP1 CAL R ACK ELEVATION REGION 2 FIGURE 2 3
3.0 NUCLEAR AND THERMAL-HYDRAULIC CONSIDERATIONS 3.1 NEUTRON MULTIPLICATION FACTOR q
()
The following subsections describe the conditions in the spent fuel pool which are assumed in calculating the effective neutron multiplication factor (keff), the analysis methodology, and the analysis restilts.
3.1.1 Normal Storage The criticality analyses of each of the two separate regions of the spent fuel storage pool are summarized in Table 3-1 for the anticipated normal storage conditions. The calculated maximum reactivity in Region 2 includes a burnup-dependent allowance for uncertainty in depletion calculations and, furthermore, provides an additional margin of 0.0065 dk below the limiting effective multiplication factor (k ff) of 0.95. As cooling time increases in long-term storage, decay of Pu-I4I results in a significant decrease in reactivity, which will provide an increasing suberiticality margin and tends to further compensate for any uncertainty in depletion calculations. Spacing between two different rack modules is sufficient to preclude adverse nuclear interaction, since the minimum spacing between racks is greater than the design water gap spacing.
Region 2 can accommodate fuel of various initial enrichments and discharge fuel burnups, provided the combination falls within the acceptable domain illustrated in Figure 3-1. For convenience of reference, the minimum burnup values in Figure 3-1 have been fitted by linear tangents at various values and the results are tabulated in Table 3-2. Linear interpolation between the
[()D ( tabulated values will always yield values on or conservatively above the curve of limiting burnups.
These data will be 1.nplemented in appropriate administrative procedures to ]
l assure verified burnup as specified in draf t Regulatory Guide 1.13, Revision l
- 2. Administrative procedures will also be employed to confirm and assure the
! presence of soluble poison in the pool water at all times, providing a further I margin of safety and assuring suberiti:Ality in the event of fuel misplacement during fuel handling operations, as discussed in Section 3.1.2. l 3.1.1.1 New Fuel Storage in Region 2 Criticality analyses confirm that a checkerboard pattern (fuel assemblies l aligned diagonally) provides an acceptable k., for the storage of fresh fuel assemblies of 4.5% enrichment in Region 2. These calculations indicate a nominal k., of 0.819 + 0.025 (95%/95%) when fully flooded with clean
~
unborated water. This value is substantially less than the limiting kefg of 0.95, even with the addition of a reasonable allowance for uncertainties.
With Boraflex absorber between assemblies, conditions do not exist for the appearance of a peak in reactivity at low moderator densities, and the fully flooded condition corresponds to the highest reactivity (optimum moderation).
Thus, the checkerboard pattern of new 4.5% enriched fuel in Region 2 represents a safe configuration in conformance with both Standard Review Plan l (SRP) 9.1.1 and 9.1.2.
O
\ /
3-1 0076L/0011L 1
l
3.1.2 PostulatQd Accidonts Although credit for the soluble poison normally present in the spent fuel pool 9 water is permitted under abnormal or accident conditions *, most abnormal or accident conditions will not result in exceeding the limiting reactivity (keff of 0.95) even in the absence of soluble poison. The effects on reactivity of credible abnormal and accident conditions are summarized in Table 3-3. Of these abnormal /
accident conditions, only one has the potential for a more than negligible positive reactivity effect.
The inadvertent misplacement of a fresh fuel assembly (either into a Region 2 storage cell or outside and adjacent to a rack module) has the potential for exceeding the limiting reactivity should there be a concurrent and independent accident condition resulting in the loss of all soluble poison. Administrative procedures assure the presence of soluble poison at all times and will preclude the possibility of the simultaneous occurrence of these two independent accident conditions. The largest reactivity increase occurs for accidentally placing a new fuel assembly into a Region 2 storage cell with all other cells fully loaded with fuel of the highest permissible reactivity. Under this condition, the presence of approximately 500 ppm soluble boron assures that the infinite multiplication factor would not exceed the design basis reactivity for Region 2. With the normal concentration of soluble poison present (1720 ppm boron), km is less than 0.80 and the storage racks would not be critical even if Region 2 were to be fully loaded with fresh fuel of 4.5% enrichment. This concentration of soluble boron also precludes the possibility of exceeding the criticality limit in the event of a dropped cask accident.
See Section 5.3 for discussions on Accident Evaluations.
3.1.3 Calculation Methods 3.1.3.1 Criticality Analysis for Region 1 3.1.3.1.1 Nominal Design Case Under normal conditions, with nominal dimensions, the k. values calculated by three different methods of analysis are as follows:
Maximum k.
Analytical Method Bias-corrected k. (95%/95%)
CASMO-2E 0. 9313 + 0.0018 0.9331 AMPX-KENO (27 gp SCALE) 0.9210 + 0.0084 0.9294 Diffusion / blackness 0.9313 0.9313 theory The AMPX-KENO calculations include a one-sided tolerance factor (13) corresponding to 95% probability at a 95% confidence limit. For the nominal design case, the CASMO-2E calculation yields the highest reactivity and, therefore, the independent verification calculations substantiate CASMO-2E as the primary calculational method.
- Double contingency principle of ANSI N16.1-1975, as specified in the April 4 1978 NRC letter (Section 1.2) and implied in the proposed revision (draft) to O Reg. Guide 1.13 (Section 1.4, Appendix A).
3-2 0076L/0011L
3.1.3.1.2 Boron Loading Variation The Boraflex absorber sheets used in Region 1 storage cells2 are nominally O 0.075 inch thick, with a B-10 areal density of 0.0238 g/cm . Independent manufacturingtolerancelimitsare10.007inchinthicknessand
+ 0.009 g/cm in B-10 content. Thisassurestha3atanypointwherethe liiinimum boron concentration (0.1158 gram B-10/cm ) and minimus Borafier thickness (0.068 inch) may coincide, the boron-10 areal density will not be less than 0.020 g/cm2 . Dif ferential CASMO-2E calculations indicate that these tolerance limits result in reactivity uncertainty of + 0.0021 Ak for boron content and 1 0.0044 Ak for Borafier thickness variations.
3.1.3.1.3 Storage Cell Lattice Pitch Variation The design storage cell lattice spacing between fuel assemblies in Region 1 is 10.12 inches. A decrease in storage cell lattice spacing may or may not increase reactivity depending upon other dimensional changes that may be associated with the decrease in lattice spacing. Increasing the water thickness between the fuel and the inner stainless steel bor results in a small increase in reactivity. The reactivity effect of the flux-trap water thickness, however, is more significant, and decreasing the flux-trap water thickness increases reactivity. Both of these effects have been evaluated for independent design tolerances.
The inner stainless steel box dimension, 8.650 1 0.032 inches, defines the inner water thickness between the fuel and the inside box. For the tolerance limit, the uncertainty in reactivii.; is 1 0.0011 A k as determined by differential CASMO-2E calculations, with k. increasing as the inner stainless steel box dimension-(and derivative lattice spacing) increases.
The design flux-trap water thickness is 1.120 1 0.040 inches, which results in an uncertainty of 1 0.0043 4 k due to the tolerance in flux-trap water thickness, assuming the water thickness is simultaneously reduced on all four cides. Since the manufacturing tolerances on each of the four sides are statistically independent, then actual reactivity uncertainties would be less than + 0.0043, although the more conservative value has been used in the criticality evaluation. ,
3.1.3.1.4 Borafier Width Tolerance Variation The reference storage cell design for Region 1 (Figure 3-2) uses a Borafier blade width of 7.50 1 0.0625 inches. A positive increment in reactivity occurs for a decrease in Boraflex absorber width. For a reduction in width of the maximum tolerance, 0.0625 inch, the calculated positive reactivity increment is +0.0017 Ak.
3.1.3.1.5 Stainless Steel Thickness Tolerances The nominal stainless steel thickness in Region 1 is 0.080 + 0.005 inch for the inner stainless steel box and 0.020 1 0.003 inch for the Borafier coverplate. The maximum positive reactivity effect of the expected stainless steel thickness tolerance variations, statistically combined, was calculated (CASMO-2E) to be + 0.0010 Ak.
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l 3.1.3.1.2 Boron Lo: ding variction The Boraflex absorber sheets used in Region 1 storage cells are nominally 0.075 inch thick, with a B-10 areal density of 0.0238 g/cm2 . Independent O
V manufacturing tolerance limits are 10.007 inch in thickness and
+ 0.009 g/cm in B-10 content. Thisassurestha3atanypointwherethe liiinimum boron concentration (0.1158 gram B-10/cm ) and minimum Boraflex thickness (0.068 inch) may coincide, the boron-10 areal density will not be )
I l
less than 0.020 g/cm 2
. Differential CASMO-2E calculations indicate that l
these tolerance limits result in reactivity uncertainty of + 0.0021 Ak for l boron content and 1 0.0044 dk for Boraflex thickness variations.
3.1.3.1.3 Storage Cell Lattice Pitch Variation The design storage cell lattice spacing between fuel assemblies in Region 1 is 10.12 inches. A decrease in storage cell lattice spacing any or may not increase reactivity depending upon other dimensional changes that may be associated with the decrease in lattice spacing. Increasing the water thickness between the fuel and the inner stainless steel box results in a small increase in reactivity. The reactivity effect of the flux-trap water thickness, however, is more significant, and decreasing the flux-trap water thickness increases reactivity. Both of these effects have been evaluated for independent design tolerances.
The inner stainless steel box dimension, 8.650 1 0.032 inches, defines the inner water thickness between the fuel and the inside box. For the tolerance limit, the uncertainty in reactivity is 1 0.0011 Ak as determined by differential CASMO-2E calculations, with k . increasing as the inner n stainless steel box dimension-(and derivative lattice spacing) increases.
(' The design flux-trap water thickness is 1.120 1 0.040 inches, which results in I
an uncertainty of 1 0.0043 A k due to the tolerance in flux-trap water thickness, assuming the water thickness is simultaneously reduced on all four sides. Since the manufacturing tolerances on each of the four sides are statistically independent, then actual reactivity uncertainties would be less than + 0.0043, although the more conservative value has been used-in the criticality evaluation. ,
3.1.3.1.4 Boraflex Width Tolerance Variation The reference storage cell design for Region 1 (Figure 3-2) uses a Boraflex blade width of 7.50 1 0.0625 inches. A positive increment in reactivity occurs for a decrease in Boraflex absorber width. For a reduction in width of the maximum tolerance, 0.0625 inch, the calculated positive reactivity increment is +0.0017 A k.
3.1.3.1.5 Stainless Steel Thickness Tolerances The nominal stainless steel thickness in Region 1 is 0.080 1 0.005 inch for the inner stainless steel box and 0.020 1 0.003 inch for the Boraflex coverplate. The maximum positive reactivity effect of the expected stainless steel thickness tolerance variations, statistically combined, was calculated (CASMO-2E) to be + 0.0010 A k.
1 3-3 0076L/00111.
3.1.3.1.2 Boron Lo: ding Variation The Boraflex absorber sheets used in Region 1 storage cells are nominally 0.075 inch thick, with a B-10 areal density of 0.0238 g/cm2 . Independent C(3 manufacturingtolerancelimitsare10.007inchinthicknessand
+ 0.009 g/cm in B-10 content. Thisassurestha5atanypointwherethe liiinimum boron concentration (0.1158 gram B-10/cm ) and minimum Boraflex thicknees (0.068 inch) may coincide, the boron-10 areal density will not be 1
less than 0.020 g/cm 2 . Differential CASMO-2E calculations indicate that these tolerance limits result in reactivity uncertainty of + 0.0021 Ak for i boron content and 1 0.0044 Ak for Boraflex thickness variations.
3.1.3.1.3 Storage Cell Lattice Pitch Variation The design storage cell lattice spcing between fuel assemblies in Region 1 is 10.12 inches. A decrease in storage cell lattice spacing may or may not increase reactivity depending upon other dimensional changes that may be I associated with the decrease in lattice spacing. Increasing the water thickness between the fuel and the inner stainless steel box results in a small increase in reactivity. The reactivity effect of the flux-trap water thickness, however, is more significant, and decreasing the flux-trap water thickness increases reactivity. Both of these effects have been evaluated for independent design tolerances.
The inner stainless steel box dimension, 8.650 1 0.032 inches, defines the inner water thickness between the fuel and the inside box. For the tolerance limit, the uncertainty in reactivity is 1 0.0011 Ak as determined by differential CASMO-2E calculations, with k. increasing as the inner l stainless steel bor dimension-(and derivative lattice spacing) increases.
V The design flux-trap water thickness is 1.120 1 0.040 inches, which results in an uncertainty of 1 0.0043 A k due to the tolerance in flux-trap water thickness, assuming the water thickness is simultaneously reduced on all four sides. Since the manufacturing tolerances on each of the four sides are statistically independent, then actual reactivity uncertainties would be less than 1 0.0043, although the more conservative value has been used in the criticality evaluation.
3.1.3.1.4 Borafier Width Tolerance Variation The reference storage cell design for Region 1 (Figure 3-2) uses a Boraflex blade width of 7.50 1 0.0625 inches. A positive increment in reactivity occurs for a decrease in Boraflex absorber width. For a reduction in width of the maximum tolerance, 0.0625 inch, the calculated positive reactivity increment is +0.0017 o k.
3.1.3.1.5 Stainless Steel Thickness Tolerances ,
l The nominal stainless steel thickness in Region 1 is 0.080 1 0.005 inch for j the inner stainless steel box and 0.020 1 0.003 inch for the Borafier l coverplate. The maximum positive reactivity effect of the expected stainless steel thickness tolerance variations, statistically combined, was calculated ,
(CASMO-2E) to be + 0.0010 A k.
(3 3-3 0076L/0011L
3.1.3.1.6 Fuel Enrichsent cod Density Variation The design maximum enrichment is 4.50 1 0.05 vt% U-235. Calculations of the sensitivity to small enrichment variations by CASMO-2E yielded a coefficient G of 0.0054 A k per 0.1 wt% U-235 at the design enrichment. For a tolerance on U-235 enrichment of 1 0.05 in wt%, the uncertainty on k. is 1 0.0027 A k.
Calculations were also made with the UO2 fuel density increased to the maximum expected value of 10.811 g/cm3 (smeared density). For the reference design calculations, the uncertainty in reactivity is 1 0.0005 d k.over the
} maximum expected range of UO2 densities.
3.1.3.1.7 Fuel Pin Pitch Normally, the fuel pins in the lattice are arranged on a 0.577 inch lattice spacing. For the maximum expected tolerance of 1 0.0023 inch, the calculated uncertainty is 1 0.0324 A k.
3.1.3.1.8 Eccentric Positioning of Fuel Assembly in Storage Rack The FuG Assembly is assumed to be normally located in the center of the storage rack cell. Calculations were also made with the fuel assemblies assumed to be the corner of the storage rack cell (four-assembly cluster at closest approach). These calculations indicated that the reactivity increases very slightly, as determined by differential PDQ07 calculations with diffusion coefficients
- generated by NULIF and a blackness theory routine. This uncertainty is included in the evaluation of the highest possible reactivity of the Region 1 storage cells.
3.1.3.1.9 Summary of Region 1 Criticality Results Table 3-1 demonstrates that the CASMO-2E calculated results for Region i storing fresh fuel at 4.50 w/o U-235 enrichment plus calculational bias and uncertainties exhibit a maximum k . of 0.9409 which allows a margin of 0.0091 A k below the limiting effective multiplication factor of 0.95.
3.1.3.2 Criticality Analysis.for Region 2 3.1.3.2.1 Nominal Design Case The principal method of analysis in Region 2 was the CASMO-2E code, using the restart option in CASMO to transfer fuel of a specified burnup into the storage rack configuration at a reference temperature of 400 (maximum moderator density). Calculations were made for fuel of several different initial enrichments and, at each enrichment, a limiting k. value was established which included an additional factor for uncertainty in the burnup analysis and for the axial burnup distribution. The restart CASMO-2E calculations (cold, clean, rack geometry) were then interpolated to define the burnup value yielding the limiting k. value for each enrichment, as indicated in Table 3-4. These converged burnup values define the boundary of the acceptable domain shown in Figure 3-1.
- This calculational approach was necessary since the reactivity effects are too small to be calculated by KENO, and CASMO-2E geometry is not readily l 9 amenable to eccentric positioning of a fuel assembly.
I 3-4 0076L/0011L
At a burnup of 36.5 Hwd/kgU, tha sensitivity to burnup is ccicul-5sd to ba
-0.0074 A k per Mwd /kgU. During long-term storage, the k. values of the Region 2 fuel rack will decrease continuously from decay of Pu-241, as indicated in Section 3.1.3.3.4.
Two independent calculational methods were used to provide additional confidence in the reference Region 2 criticality analyses. Fuel of 1.69%
initial enrichment (approximately equivalent to the reference rack design for burned fuel) was analyzed by AMPX-KENO (27 group SCALE cross-section library) and by the CASMO-2E model used for the Region 2 rack analysis. For this case, the CASMO-2E k . (0.9304) was within the statistical uncertainty of the bias-corrected value (0.9347 + 0.0064) (95%/95%) obtained in the AMPX-KEN 0
~
calculations. This agreement confirms the validity of the primary CASMO-2E calculations.
The second independent method of analysis used was the NULIF code for burnup analysis, and for generating diffusion theory constants (cold, clean) for the composition at 36.5 Hwd/kgU with fuel of 4.5% initial enrichment. These constants, together with blackness theory constants for the Boraflex absorber, were then used in a two-dimensional PDQ07 calculation for the storage rack configuration. The result of this calculation (k. of 0.8959) was somewhat lower than the corresponding CASMO-2E calculation for the same conditions (k.
of 0.9114) and thus also tends to confirm the validity of the primary calculational method.
3.1.3.2.2 Boron Loading Variation The Boraflex absorber sheets used in the Region 2 storage cells are nominally 0
2 Independent O h ma.031 inch thick with a B-10 areal density of 0.0097 g/cm .nufacturing lim
() content. This assures that at any point where the miniliium boron concentration (0.1158 g B-10/cm3 ) and the minimum Boraflex thickness (0.024 inch) may coincide, the boron-10 areal density will not be less than 0.007 g/cmd.
Differential CASMO-2E calculations indicate that these tolerance limits result in an incremental reactivity uncertainty of 1 0.0036 A k for boron content and 1 0.0111 A k for Boraflex thickness.
3.1.3.2.3 Boraflex Width Tolerance The reference storage cell design for Region 2 (Figure 3-3) uses a Boraflex absorber width of 7.25 + 0.0625 inches. For a reduction in width of the maximum tolerance, the calculated positive reactivity increment is 0.0011 Ak.
3.1.3.2.4 Storage cell Lattice Pitch Variations The design storage cell lattice spacing between fuel assemblies in Region 2 is 8.86 + 0.04 inches, corresponding to an uncertainty in reactivity of
- 0. 0013 A k. l 3.1.3.2.5 Stainless Steel Thickness Tolerance The nominal thickness of the stainless steel box wall is 0.080 inch with a tolerance limit of + 0.005 inch, resulting in an unceri:ainty in reactivity of
+ 0.0002 Ak.
3-5 0076L/0011L
)
3.1.3.2.6 Fuel Enrichnent, Density and Pin Pitch Variation Uncertainties in reactivity due to tolerances oc fuel enrichment, UO2 density, and pin pitch in Region 2 are assumed to be the same as those 9 determined for Region 1.
3.1.3.2.7 Eccentric Positioning of Fuel Assembly in Storage Rack The fuel assembly is assumed to be normally located in the center of the storage rack cell. Calculations were also made with the fuel assemblies assumed to be in the corner of the storage rack cell (four-assembly cluster at closest approach). These calculations indicated that the reactivity decreases very slightly, as determined by PDQ07 calculations with diffusion coefficients generated by NULIF and a blackness theory routine. The highest reactivity therefore corresponds to the reference design with the fuel assemblies positioned in the center of the storage cells.
3.1.3.3 Analytical Methodology 3.1.3.3.1 Reference Analytical Methods and Bias The CASMO-2E computer code (1 8 2' 3) , a two-dimensional multigroup transport theory code for fuel assemblies, has been benchmarked and is used both as a primary method of analysis, and as a means of evaluating small reactivity increments associated with manufacturing tolerance. CASMO-2E benchmarking resulted in a calculational bias of 0.0013 1 0.0018 (95%/95%).
h In fuel rack analyses, for independent verification, criticality cualyses of G C. the high density AMPX-KE!gcomputer spent fuegstgage package racks were
, using the 27-group SCALEalso cross-section with the NITAWL subroutine for U-238 resonance shielding effects perforged with the library (Nordheim integral treatment). Benchmark calculations resulted in a bias of 0.0106 1 0.0048 (95%/95%).
In the geometric model used in KENO, each fuel rod and its cladding were described explicitly. In Region 1 calculations, a reflecting boundary j condition (zero neutron current) was used in the axial direction and at the centerline of the water gap between storage cells. These boundary conditions have the effect of creating an infinite array of storage cells in all directions. In Region 2, the zero current boundary condition was applied at the center of the Borafier absorber sheets between storage cells. The AMPX-KENO Monte Carlo calculations inherently include a statistical uncertainty due to the random nature of neutron tracking. To minimize the statistical uncertainty of the KENO-calculated reactivity, a total of 50,000 neutron histories is normally accumulated for each calculation, in 100 generations of 500 neutrons each.
- SCALE is an acronym for Standardized Computer Analysis for Licensing Evaluation, a standard cross-section set developed by ORNL for the USNRC.
O 3-6 0076L/0011L
- - - _ _ _ _ l
CASMO-2E is clso uced for burnup calculations, with indep2ndsnt verification by EPRI-CELL and NULIF calculations. In tracking long-tern (30 year) reactivity effects of spent Suel stored in Region 2 of the fuel storage rack, EPRI-CELL calculations indicaca a continuous reduction in reactivity with time O (after Xe decay) due primarily to Pu-241 decay and Am-241 growth.
A third independent method of criticality analysis, utilizing diffusion / blackness theory, was also used for additional confidence in results of the primary calculational methods, although no reliance for criticality safety is placed on the reactivity value from the diffusion /bischness theory technique. This technique, however, is used for auxiliary calculations of the small incremental reactivity effect of eccentric fuel positioning that would otherwise be lost in normal KENO statistical variations, or would be inconsistent with CASMO-2E geometry limitations.
Cross sections for the diffusion / blackness theory calculations were derived from the NULIF computer code (7} , supplemented by a blackness theory routine that effectively imposes a transport theory boundary condition at the surface of the Boraf neutron absorber. Two dif fgreat spatial dif fusion theory codes, PDQ07 in two dimensions and SNEID in one dimension, were used to calculate reactivities.
3.1.3.3.2 Fuel Burnup Calculations Fuel burnup calculations in the hot operating condition were performed primarily with the CASMO-2E code. However, to enhance the credibility of the burnup calculations, the CASMO-25 results were independently checked by calculations with the NULIF codel7) and with EPRI-CELL (9). Figure 3-4 compares results of these independent methods of burnup analysis under hot The results agree with the CASMO calculation O h reactor b' withinoberating
- 0. 054 A conditions.
k in the hot operating condition. An archive calculation with the CHEETAH-P code is also presented in Figure 3-4 for additional confidence.
Similar comparisons were obtained in burnup calculations for other initial enrichments, as indicated in Figure 3-4.
In addition to depletion calculations under hot operating conditions,.
reactivity comparisons under conditions more representative of fuel to be stored in the racks (cold, zenon-free) are also significant in storage rack criticality analyses. Table 3-5 compares the cold, renon-free reactivities calculated by CASM0-2E, EPRI-CELL, and diffusion / blackness theory. In the rack under cold conditions, the CASMO-2E calculations gave a slightly higher ,
reactivity value for the Region 2 fuel storage cell, and the good agreement generally observed lends credibility to the calculations.
- SNEID is a one-dimensional diffusion theory routine developed by Black &
Veatch and verified by comparison with PDQ07 one-dimensional calculations.
U' 3-7 0076L/0011L
No definitive cathod exists for datsrmining tha uncertainty in i burnup-dependent reactivity calculations. All of the codes discussed above have been used to accurately follow reactivity loss rates in operating reactors.
CASHO-2E has been extensively benchmarked (1, 2, 3, 101 against v/ cold, clean, critical experiments (including plutonium-bearing fuel), Monte Carlo calculations, reactor operations, and heavy-element concentration in irradiated fuel. In particular, the anclysesllW of 11 critical experiments with plutonium-bearing fuel gave an average keff of 1.002 + 0.011 (95%/95%), ,
showing adequate treatment of the plutonium nuclides. In addition, Johansson(ll) has obtained very good agreement in calculations of close packed, high plutonium-content, experimental configurations.
Since critical-experiment data with spent fuel is not available, it is '
necessary to assign an uncertainty in reactivity based on other considerations, supported by the close agreement between different '
calculational methods and the general industry experience in predicting reactivity loss rates in operating plants. Over a considerable portion of the ;
burnup, the reactivity loss rate in PWRs is approximately 0.01 Ak for each Mwd /kgU burnup, becoming somewhat smaller at the higher burnups. By assuming an uncertainty in reactivity of 0.0005 times the conservatively burnup in Mwd /kgU, a burnup-dependent uncertainty is defined that increases with increasing fuel burnup, as would be reasonably expected. This assumption provides an estimate of the burnup uncertainty that is more conservative and bounds estimates frequently employed in other fuel rack licensing applications (i.e., 5% of the total reactivity decrement). At the design basis burnup of 36.5 Mwd /kgU, the estimate of burnup uncertainty is 0.0183 Ak; Table 3-6 summarizes results of the burnup analyses and estimated uncertainties at other burnups. These uncertainties are appreciably larger, in general, than would g3 - be suggested by the industry experience in predicting reactivity loss rates and boron let-down curven over many cycles in operating plants. The increasing level of conservatism at the higher fuel burnups provides an adequate margin in the uncertainty estimate to accommodate the possible existence of a small positive reactivity increment from the axial distribution in burnup (see Section 3.1.3.3.3). In addition, although the burnup uncertainty may be either positive or negative, it is treated as an additive term rather than being combined statistically with other uncertainties. Thus, the allowance for uncertainty in burnup calculations is considered to be a conservative estimate, particularly in view of the substantial reactivity decrease with aged fuel, as discussed in Section 3.1.3.3.4.
- Only that portion of the uncertainty due to burnup. Other uncertainties are accounted for elsewhere.
i 3-8 0076L/0011L l
3.1.3.3.3 Effcet of Arial Burnup Distributico Initially, fuel loaded into the reactor will burn with a slightly skewed cosine power distribution. As burnup progresses, the burnup distribution will Q
V tend to flatten, becoming more highly burned in the central regions than in the upper and lower ends. This effect may be clearly seen in the curves compiled in Reference 12. At high burnup, the more reactive fuel near the ends of the fuel assembly (less than average burned) occurs in regions of lower reactivity worth due to neutron leakage. Consequently, it is expected (
that distributed-burnup fuel assemblies would exhibit a slightly lower reactivity than that calculated for the average burnup. As burnup progresses, the distribution, to some extent, tends to be self-regulating as controlled by the axisi power distribution, precluding the existence of large regions of l
significantly reduced burnup.
A number of one-dimensional diffusion theory analyses have been made based !
upon calculated and measured axial burnup distributions. These analyses )
I confirm the minor, and generally negative, reactivity effect of the axially distributed burnup. The trends observed, however, suggest the possibility of a small positive reactivity effect at the high burnup values (estimated to be as much as 0.006 4 k at 36.5 Mwd /kgU); but the uncertainty in k., due to burnup, assigned at the higher burnups (Section 3.1.3.3.2), is adequately conservative to encompass the potential for a small positive reactivity effect of axial burnup distributions. Furthermore, reactivity significantly decreases with time in storage (Section 3.1.3.3.4), and, in addition, there is a further margin in reactivity (>0.006 Ak) since the maximum calculated value (0.9435) is reasonable below the reactivity ff value limiting k,ffects (0.95).
that mightThese factors be larger thanwould accommodate any e expected.
V 3.1.3.3.4 Long-term Decay Since the fuel racks in Region 2 are intended to contain spent fuel for long periods of time, calculations were made using EPRI-CELL (which incorporates the CINDER code) to follow the long-term changes in reactivity of spent fuel over a 305 year period. CINDER tracks the decay and burnup dependence of some 179 fission products. Early in th.e decay period, renon grows from iodine decay (reducing reactivity) and subsequently decays, with the reactivity reaching a maximum at 100-200 hours. The decay of Pu-241 (13 year half-life) and growth of Am-241 substantially reduce reactivity during long term storage, as indicated in Table 3-7.
The reference design criticality calculations do not take credit for this long-term reduction in reactivity, other than to indicate an increasing {
suberiticality margin in Region 2 of the spent fuel storage pool. l 3.1.4 Rack Modification The design basis fuel assembly, illustrated in Figure 3-2, is a 14 x 14 array of fuel rods with 20 rods replaced by 5 control rod guide tubes. Table 3-8 summarizes the design specifications and the expected range of significant variations. Independent calculations, with other potential fuel assembly specifications, confirmed that the 14 x 14 CE design exhibited the highest reactivity and was therefore used as the design basis.
O) v 3-9 0076L/0011L I
['_ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ _ I
l 3.1.4.1 Region 1 Storego Calls The nominal spent fuel storage cell used for the criticality analyses of
/Q < Region 1 storage cells is shown in Figure 3-2. The rack is composed of 0.080-inch Boraflex absorber material sandwiched between an 8.65-inch I.D.,
thick inner stainless steel box, and a 0.020-inch outer stainless steel coverplate. The fuel assemblies are centrally located in each storage cell on a nominal lattice spacing of 10.120 + 0.05 inches. Stainless steel gap channels connect one storage cell box to another in a rigid structure and define an outer water space between boxes. This outer water space constitutes a flux-trap between the two Boraflex absorber sheets that are essentially opaque (black) to thermal neutrons. The Boraflex absorber has a thickness 2
of 0.075 + 0.007 inch and a nominal B-10 areal density of 0.0238 g/cm .
3.1.4.2 Region 2 Storage Cells Region 2 storage cells were designed for fuel of 4.5 wt% U-235 initial enrichment burned to 36.5 Mwd /kgU. In this region, the storage cells are composed of a single Boraflex absorber sandwiched between the 0.080-inch stainless steel walls of adjacent storage cells. These cells, shown in Figure 3-3, are located on a lattice spacing of 8.86 + 0.040 inches. The Boraflex
~
absorber has a thickness of 0.031 + 0.007 inch and a nominal B-10 areal
~
density of 0.0097 g/cmd .
3.1.5 Acceptance Criteria for Criticality Criticality is precluded by spacing of the fuel assemblies, which ensures that a suberitical array of k egf less than or equal to 0.95 is maintained, The pool, however, will always contain borie C., assuming unborated pool water. acid at the refueling concentration of 1720 ppa whenever t v fuel in the pool.
The neutron multiplication factor in spent fuel pools shall be less than or equal to 0.95, including all uncertainties, under all conditions. Calculated maximum reactivity uncertainties for fuel stored in the racks are presented in Table 3-1.
Methods of initial and long-term verification of poison material stability and i mechanical integrity are discussed in Section 4.8.
3.2 DECAY HEAT CALCULATIONS FOR THE SPENT FUEL POOL (BULK) 3.2.1 Spent Fuel Pool Cooling System Design For normal refueling discharge conditions, one fuel pool pump and the fuel pool heat exchanger are in service. During abnormal refueling conditions, such as full core discharge, two fuel pool pumps and the heat exchanger are in service. The system is manually controlled and the operation monitored locally, except as follows. A pressure switch on the fuel pool pump discharge header annunciated low header pressure in the control room. The fuel pool high temperature alarm and low level alarms are annunciated in the control room. In the event the fuel pool pump breakers are opened, an alarm is annunciated in the control room. The component cooling water flow to the fuel pool heat exchanger is initially adjusted to the required flow. Further p adjustments of the component cooling water are not required. The component cooling water discharge line has a flow indicator. High and low component V cooling water flow alarms are annunciated in the control room.
3-10 0076L/0011L
The clarity and purity of the water fn the fual pool to esintained by the purification portion of the fuel pool system. The purification. loop consists of the fuel pool purification pump, . ion exchanger, filter, strainers and surface skimmers. Most of the purificat(n flow is drawn through the surface g'
skimmers to remove surface debris. A b4/d et strainer is provided in the purification line to the pump suction to casove any relatively. large particulate matter. The fuel pool wate" Ic circulated by the pump through a' filter, ~which removes particulate larset t.han 5 micron size, and through an -
ion exchanger to remove ionic material. Connections are provided for purification of the refueling water tank and refueling water cavity. . Fuel pool water chemistry is given in FSAR Table 9.1-2.
The fuel pool piping is arranged so that the pool cannot be inadvertently drained to uncover the fuel in the event of a supply or discharge pipe rupture. All fuel pool piping is arranged to prevent gravity draining the fuel pool. To prevent siphoning of the fuel pool, the fuel pool discharge and purification suction lines have 1/2" and 1/4" holes respectively 1 foot below the normal water level.
The only means of draining the pool below these' siphon breaker holes is through an open line in the cooling loop while operating the pool cooling..
pumps. In such an event the fuel pool water level can be reduced by only 6 feet since the pump suction connection enters near the top of the pool. The remaining water in the Spent Fuel Pool will provide adequate shielding and heat renoval capabilities at this point. The' temperature and level alaras would warn the operator of such an event.
l ll 3.2.2 Decay Heat Analyses 3.2.2.1 Basis The St. Lucie Plant Unit i reactor is rated at 2700 megawatts thermal (MWt).
The core contains 217 fuel assemblies. Thus, the average operating power per fuel assembly, Po, is 12.44 MW. The fuel discharge can be made in one of the following two modes:
Normal refueling discharge
- Full core discharge
~
Tables 3-9 through 3-11 give the parameters for bulk and local pool temperature analyses.
3.2.2.2 Model Description ,
NUREG-0800 Branch Technical Position ASB 9-2, " Residual Decay Energy For Light Water Reactors For Long Term Cooling =(15) is utilized to compute the heat dissipation requirements in the pool. l O; ;
3-11 0076L/0011L
Uith tha long term uncertainty factor, K, es spccified in SRP 9.1.3 (15) ,
j the operating power, Po , is taken equal to the rated power, even though the reactor may be operating at less than its rated power during auch of the exposure period for the batch of fuel assemblies. The computations and (Q_/ results reported here are based on the discharge taking place when the I inventory of fuel in the pool will be at its maximum resulting in an upper I bound on the decay heat rate.
Having determined the heat dissipation rate, the next task is to evaluate the tiae-dependent temperature of the pool water. Table 3-9 identifies the loading cases examined. This is a conservative representation of at.tual and BULKTEM future expected discharges such as those presented in Table 5-1.
treats the generalized pool cooling problem shown in Figure 3-5.
A number of simplifying assumptions are made which render the analysis conservative, including:
- The heat exchanger is assumed to have maximum fouling. Thus, the temperature effectiveness, P, for the heat exchanger utilized in the analysis is the lowest postulated value calculated from heat e2 changer technical data sheets. 1
- No credit is taken for the improvement in the film coefficients of the l heat exchanger as the operating temperature rises due to monotonic reduction in the water kinematic viscosity with temperature rise.
Thus, the film coef ficient used ic the computations are lower bounds.
- No credit is taken for heat loss by evaporation of the pool water. )
- No credit is taken for heat loss to pool walls and pool floor s12b. l 1
The basic energy conservation relationship for the pool heat exchanger system j yields:
l Ct U =
Q1 - Q2 d7 where:
C e
= Thermal capacity of stored water in the pool t = Temperature of pool water at time,7 Q1
- Heat generation rate due to stored fuel assemblies in the pool Q2
= Heat removed in the fuel pool heat exchanger This equation is solved as an initial value problem by noting that the cooler heat removal rate must equal the heat generation rate f rom previously discharged assemblies. Hence, Wcool P (Ti,- teool) = PCONS O
3-12 0076L/0011L
l-whara: -
1 PCONS: Heat generation rate from previously stored assemblies Weo,1:
Coolant thermal flow ra;:e P: Temperature effectiveness of the fuel pool cooler Coincident pool water temperature (initial value before Tin: beginning of discharge) te3,1: Coolant inia.t temperature The above equetion yields l
PCONS Tn" i +teool Weool P j
The value of T in computed from the above formula is the initial value of the j pool water temperature (at the start of fuel discharge),
BULKTEM automates the solution of the above equation using the theory presented in Reference 16. Tabulated results are presented in the next sub-section. J 3.2.2.3 Bulk Pool Temperature Results Table 3-12 gives the total dimensionless power generation ratio of all fuel f) / assembly batches previously stored in the pool consisting of a total of 18
's_/ batches. The first column in Table 3-12 gives the batch number, and the last column gives the dimensionless power, defined as the heat generation rate of the batch divided by the nominal operating power of one fuel assembly. It is noted from Table 3-12 that the cumulative power is 0.14 times the operating power of one fuel assembly. Tables 3-13/3-14 and 3-16/3-17 give the bulk i temperature vs. time data.
The following key output data is gleaned from these tables: j Maximum pool bulk temperature:
Normal discharge: 133.30F Table 3-14 Full core discharges 150.80P Table 3-17 Tables 3-13 and 3-18 give time-to-boil data.
Time-to-boil (if coolant flow is lost upon completion of discharge and when the bulk pool temperature is mar.imum):
Normal discharge condition: 13.43 hours4.976852e-4 days <br />0.0119 hours <br />7.109788e-5 weeks <br />1.63615e-5 months <br /> Table 3-15 Full core discharge condition: 5.04 hours4.62963e-5 days <br />0.00111 hours <br />6.613757e-6 weeks <br />1.522e-6 months <br /> Table 3-18 L
3-13 0076L/0011L 2
l i
3.2.2.4 Sp:nt Fu21 Pool Cooling Systcu Summary The spent fuel decay heat calculations were performed in accordance with the !
method provided in NRC Branch Technical Position ASB 9-2, Residual Decay O Energy for Light-Water Reactors f or Long-Term Cooling (15),
l
() )
1 j
The existing spent fuel pool cooling system is considered to be adequate. The j
spent fuel pool is designed to withstand stresses associated with a steady-state water temperature of 2170 F. As shown in Table 3-17 the pool I
l peak transient water temperature after full core discharge is less than 1510 F.
In the event of a complete loss of cooling capability, there is sufficient 1 time to provide an alternate means for cooling.
The total increase in heat load rejected to the environment through the cooling systems due to the increased spent fuel storage over the current heat load rejected is 1.7 x 10 6 Btu / hour. Thie represents an increase of The approximately 0.03 percent of the total heat rejected to the environment.
increase in heat rejected will have negligible impact on the environment.
The increase in heat load does not alter in any way the existing facility design bases. Thus, the heat load increase is acceptable. This decay heat analysis is also bounding for the temporary fuel storage configuration (see Section 4.7.4) that will be utilized during rack installation. ]
l 3.2.2.4.1 Safety Evaluation l l
The calculations for the amount of thermal energy that may have to be removed
) k. by the spent fuel pool cooling system are made in accordance with Branch v Technical Position ASB 9-2 (Reference 15). The resulting bulk spent fuel pool temperatures are acceptable.
3.7.3 Spent Fuel Pool Makeup There are several sources of fresh water on the site that are available to the fuel handling building; namely, refueling water storage tank, city water storage tank via the fire main, city water storage tanks via the portable fire pump, and primary water tank. The concurrent loss of these sarces and the fuel pool cooling system is remote. Due to the fuel pool's boil-off period, there is sufficient time to obtain makeup. It should be noted that a seismic Category I backup salt water supply is available f rom the intake cooling water intertie. A standpipe on the fuel handling building is provided from grade to the operating deck elevation and hose connections are provided at both ends of the standpipe. Thus, iria fire hose, the fuel pool makeup can be readily supplied by the intake cooling water pumps. The head provided by these pueps is sufficient to provide the required fuel pool make up. The structural and leaktight integrity of the fuel pool vill not be compromised by continuous fuel pool temperatures of up to 2170 F. The results of the bulk decay heat analys9s indicate that these temperatures are not exceeded. The intake cooling water system connection via the hose connections can provide 150 gpm of ma..eup. See FSAR Bubsection 9.1.3.4.
A v
3-14 0076L/0011L
3.3 THERMAL-HYDRAULIC ANALYSES FOR THE SPENT FUEL POOL (LOCALIZED)
The purpose of the thermal-hydraulic analyses is to determine the maximum fuel clad temperatures which may occur as a result of using the new high density
(]
V spent fuel racks in the St Lucie Unit 1 spent fuel pool. l l
3.3.1 Bases In order to determine an upper bound on the maximum fuel cladding temperature, a series of conservative assumptions are made. The most important assumptions are listed below:
- As stated above, the fuel pool vill contain spent fuel with varying time-af ter-shutdown ( itT is Since the heat emission falls off rapidly s).obviously with increasing 7,, conservative to assume that all fuel assemblies are fresh and they all have had the maximum postulated j
years of operating time in the reactor. The heat emission rate of each fuel assembly is assumed to be equal and maximum. j i
- As sh wn in Figure 2-1, 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 3-6). It is further assumed I I
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, as i shown in Figure 2-1. The nominal downcomer gap available in the pool l is assumed to be the total gap available around the idealized cylindrical rack; thus, the maximum resistance to downward flow is
[m i incorporated into the analysis (Figure 3-7).
%)
- No downcomer flow is assumed to exist between the rack modules.
3.3.2 Model Description Using the bases described above, 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 3-7 shows a typical " flow chimney" rendering of the thermal hydraulics model. The governing equation to characterize the flow field in the pool is an integral equation that can be solved for the lower plenum velocity field (in the radial direction ) and axial velocity (in-cell velocity field), by using the method of collocation.
It should be added that the hydrodynamic loss coefficients which enter into the forpu \1}etionoftheintegralequationarealsotakenfromwell-recognized sources > and wherever discrepancies in reported values exist, the conservative values are consistently used. Reference 18 gives the details of mathematical analysis used in this solution process.
O v
3-15 0076L/0011L
After th2 crial velocity field is avslutted, tha fusi ass:nbly citdding temperature can be calculated. 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" f 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 equation for the flow circuit through this cell. Thus, an absolute upper bound on the water exit temperature and maximum fuel cladding temperature is obtained. In view of the aforementioned assumptions, the temperatures calculated in this manner overestimate the temperature rise that will actually occur in the pool.
THERPOOL, based on the theory of Reference 18, automates this calculation.
Finally, the maximum specific power of a fuel array qA can be given by:
oA
= qF gy where:
q = average fuel assembly specific power l F xy = radial peaking factor l
The data on radial and axial peaking factors may be found in Table 3-10.
The maximum temperature rise of pool water in the most disadvantageous 1y placed fuel assembly is computed for all loading cases. Table 3-19, third column, gives the outputs f rom THERPOOL in tabular form.
3.3.3. Cladding Temperature 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 FTot times the average heat emission rate over a small length, where F
Tot is the total peaking factor. The axial heat dissipation in a rod is )
known to reach a maximum in the central region, and taper off at its two extremities. For added conservatism, it is assumed that the peak heat emission occurs at the top where the local water temperature also reaches its maximum. Futhermore, no credit is taken for axial conduction of heat along l the rod. The highly conservative model thus constructed leads to simple j algebraic equations which directly give the maximum local cladding l temperature, t e.
Table 3-19, fourth column, summarizes the key output data. It is found that the maximum value of the local water temperature is well below the nucleate boiling condition value. The incremental cladding temperature is too small to produce significant thermal stresses.
! 3.4 POTENTIAL FUEL AND RACK HANDLING ACCIDENTS The method for moving the racks into and out of the spent fuel pool is briefly discussed in Sectiota 4.7.4.2. The methods utilized ensure that postulated accidents do not result in a loss of cooling to either the spent fuel pool or the reactor, or result in a keff in the spent fuel pool exceeding 0.95.
O 3-16 0076L/0011L
3.4.1 Rack Module Mish.ndling The potential for mishandling of rack modules during the rerack operation has
/O ' been evaluated. At no time will the cask handling crane or the temporary construction crane carry a rack module directly over a rack containing spent V
fuel. The procedures and administrative controls governing the rerack operation will ensure the safe handling of rack modules. Both the temporary construction crane and the cask handling crane meet the design and operational requirements of Sectipn 5 1.1 of NUREG-0612, " Control of Heavy Loads at Nuclear Power Plants"t191 In the unlikely event that a rack should strike the side of another rack module containing fuel assemblies, the consequences of this postulated accident would be bounded by the cask drop evaluations described in Section 5.3.1.2.
3.4.2 Temporary Construction Crane Drop l
j During the rerack operation, a temporary construction crane will be installed in the Fuel Handling Building. This installation will be performed using lift i rigs which meet the design and operational requirements of NUREG-0612, j
" Control of Heavy Loads at Nuclear Power Plants." The consequences of a <
postulated accident during this installation are bounded by the cask drop evaluations described in Section 5.3.1.2.
3.4.3 Loss of Pool Cooling (Storage Rack Drop)
During the re-racking operation, it will be necessary to raise and maneuver the old racks out of the spent fuel pool in order to install the new spent g) f~ 4 fuel racks (See Section 4.7.4). The handling of these heavy loads will be s
I V accomplished by the use of'a temporary construction crane and the cask handling crane. Both of these cranes meet the design and operational requirements of Section 5.1.1 of NUREG-0612, " Control of Heavy Loads at Nuclear Power Plants."
The consequences of dropping a rack in the Spent Fuel Pool were determined by reviewing the analysis in FSAR Subsection 9.1.4 for dropping of the spent fuel ;
cask. The results of this cask drop analysis demonstrated that the pool floor l would remain elastic during impact and that cracks would not develop. This l cask weighs substantially more than a single rack assembly and has a smaller cross sectional area for load distribution. Therefore, the rack drop scenario is bounded by the previous analysis for a cask drop scenario, and loss of l spent fuel cooling from less of pool water inventory will not occur as a l result of a rack drop.
3.5 TECHNICAL SPECIFICATION CHANGES This proposed amendment permits replacement of the spent fuel pool racks to ensure that sufficient capacity exists for storage of spent fuel at St. Lucie Unit 1. The new racks increase the available storage to 1706 spent fuel assemblies and is expected to provide adequate storage space until the year 2009.
n 3-17 0076L/0011L
j I
The proposed Technical Specification changes are described below:
- 1. Specification 3/4.9.14 Bases is revised to reflect the assumptions used in calculations of doses based on the Decay Times. l C
- 2. Specification 5.6.1.a.1 is revised to correspond to the Standard i Technical Specifications for Combustion Engineering Pressurized Water !
Reactors (NUREG-0212 Rev 2).
- 3. Specification 5.6.1.a.2 is revised to show the nominal center-to-center distance for the high capacity spent fuel storage racks.
- 4. Specification 5.6.1.a.3 is edited to discuss the boron concentration 1 only.
- 5. Specification 5.6.1.a.4 is created to indicate the presence of Boraflex in the cells. 1
- 6. Specification 5.6.1.b and accompanying Figure 5.6-1 are created to define the fuel enrichment /burnup limits for storage in each region of the high capacity spent fuel storage racks.
- 7. Specification 5.6.lc is editorially changed from "b" to "c".
- 8. Specification 5.6.3 is changed to show the capacity of the high-capacity spent fuel storage racks.
3.6 REFERENCES
FOR SECTION 3
- 1. A. Ahlin, M. Edenius, H. Hagsblom, "CASMO - A Fuel Assembly Burnup Program," AE-RF-76-4158, Studsvik report (proprietary).
- 2. A. Ahlin and M. Edenius, "CASMO - A Fast Transport Theory Depletion Code for LWR Analysis," ANS Transactions, Vol. 26, p. 604, 1977.
- 3. M. Edenius et al., "CASMO Benc5unark Report," Studsvik/RF-78-6293, Aktiebolaget Atomenergi, March 1978.
- 4. Green, Lucious, Petrie, Ford, White, Wright, "PSR-63/AMPK-1 (code package), AMPX Modular Code System for Generating Coupled Multigroup Neutron - Gamma Libraries from ENDF/B," ORNL-TM-3706, Oak Ridge National Laboratory, March 1976.
- 5. L. M. Petrie and N. F. Cross, " KENO-1V, An Improved Monte Carlo Criticality Program," ORNL-4938, Oak Ridge National Laboratory, November j
1975.
- 6. R. M. Westfall et al., " SCALE: A Modular Code System for Performing gdardizedComputerAnalysesforLicensingEvaluation,"NUREG/CR-0200, n '
V 3-18 0076L/0011L
p
- 7. W. A. Vittkopf ."NULIF - Neutron Spectrua Ganarctor, Faw-Croup Constant Generator and Fuel Depletion Code," BAW-426,' The Babcock & Wilcor Company, August 1976.
- 8. W. R. Cadwell, PDQ07 Reference Manual, WAPD-TM-678, Bettis Atomic Power Laboratory, January 1967.
- 9. W. J. Eich, "i.dvanced Recycle Methodology Progras, CEM-3," Electric Power Research Institute, 1976.
- 10. E. E. Pilat, " Methods for the Analysis of Boiling Water Reactors (Lattice Physics)," YAEC-1232, Yankee Atomic Electric . Co. , December 1980,
- 11. E. Johansson, " Reactor Physics Calculations on Close-Packed Pressurized Water Reactor Lattices," Nuclear Technology, Vol. 68, pp. 263-268, February 1985.
- 12. H. Richings, Some Notes on PWR (W) Power Distribution Probabilities for LOCA Probabilistic Analyses, NRC Memorandum to P. S. Chack, dated. July 5, 1977.
- 13. M. G. Natrella, Experimental Statistics, National Bureau of Standards, Handbook 91, August 1963.
14 J. M. Cano et al., "Supercriticality Through Optimus Moderation in Nuclear Fuel Storage," Nuclear Technology, Vol. 48, pp. 251-260, May 1980.
(' 15. NUREG-0800, U.S. Nuclear Regulatory Commission, Standard Review Plan, Branch Technical Position ASB 9-2, Rev. 2, July 1981.
- 16. Singh, K. P., Journal of Heat Transfer, Transactions of the ASME, August 1981, Vol.1-3, "Some Fundamental Relationships for Tubular Heat Exchanger Thermal Performance."
- 17. General Electric Corporation, R&D Data Books,' " Heat Transfer and Fluid Flow," 1974 and updates.
- 18. 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).
- 19. Nucitar Regulatory Cosaission, " Control of Heavy Loads at Nuclear Power Plants, NUREG-0612, July 1980.
3-19 0076L/0011L
_-y- - . _ _ - -
l TABLE 3-1 e
SUMMARY
OF CRITICALITY SAFETY ANALYSES Region 1 Region 2 0 36.5 Mwd /kgU Minimum acceptable burnup j
0 4.5% initial enrichment 0
Temperature assumed 4C 40C for analysis Reference k (nominal) 0.9313 0.9114 .
l 0.0013 0.0013 Calculational bias 1 1 Uncertainties Bias +0.0018 +0.0018 70.0036 L B-10 concentration 70.0021 Boraflex thickness 70.0044 70.0111 !
Botaflex width I0.0017 70.0011 i
~
i Inner box dimension +0.0011 70.0016 Water gap thickness 70.0043 N/A OC- SS thicknesa Fuel enrichment
+0.0010 I0.0027
+0.0002 70.0027 Fuel density 70.0005 70.0005 Fuel element pitch I,0.0024 ~0.0024
+
Statistical combination (l) +0.0080 +0.0125 Eccentric assembly position T0.0003 negative
. i l
Allowance for N/A +0.0183 burnup uncertainty Total 0.9329 1 0.0080 0.9310 1 0.0125 Maximum reactivity 0.9409 0.9435 (with 1720 ppa soluble boron) (0.767) (0.760)
(1) Square root of sum of squares.
- l. N/A - Not Applicable l
l O
3-20 0076L/0011L
~
I TABLE 3-2 MINIMUM BURNUP VALUES Initial Minimum Enrichment, % Burnup, Mwd /kgU 1.63 0 1.75 2.30 e 2.00 6.00 2.25 9.70 2.50 12.90 1 2.75 16.10 '
i 3.00 19.15 i 3.25 22.20 3.50 25.15 3.75 28.10 4.00 30.90 4.25 33.70 4.50 36.50
~
OC O .
3-21 0076L/0011L
o
' TABLE 3-3 REACTIVITY EFFECTS OF ABNORMAL AND ACCIDENT CONDITIONS-
'I 4
Accident / Abnormal Conditions Reactivity Effect ;
l.
Temperature increase Negative in both regions ;
Void (boiling) Negative in both regions Assembly dropped on top of rack Negligible l l
Lateral rack module movement Negligible Misplacement of a fuel assembly Positive j O C' 1
i j
l O'
3-22 0076L/0011L f i
i l
TABLE 3-4 FUEL BURNUP VALUES FOR REQUIRED REACTIVITIES (k. )
WITH FUEL OF VARIOUS INITIAL ENRICHMENTS (Reference k . - 0.9297)
Calculated Initial Uncertainty (1) Design Burnu limit in Burnup, A k Limit k. Hwd/kgU Enrichment 1.6 0 0.9297 0 2.0 0.0030 0.9267 5.99 2.5 0.0064 0.9233 12.88 3.0 0.0096 0.9201 19.13 3.5 0.0126 0.9171 25.15 4.0 0.0154 0.9143 30.86 4.5 0.0183 0.9115 36.50 ec (1) See Subsection 3.1.3.3.2 e
s 3-23 0076L/0011L i
v e
TABLE 3-5 COMPARISON OF COLD, CLEAN REACTIVITIES CALCULATED AT 36.5 Mwd /kgU BUR.NUP AND 4.5% ENRICEMENT ,
k., Xe-free, 40C Assemblies in Calculational Method Infinite Array 9tlf ) Region 2 Cell 1 Fuel Assemblies in Reactor Spacing 0.9114 l CASMO-2E 1.1212 l
1.1306 0.8972 DIFFUSION / BLACKNESS THEORY EPRI-CELL 1.1281(2) f I
l (1) Cold, clean condition in contrast to hot operating conditions of Figure 3-4.
(2) EPRI-CELL k,,, at maximum value during long-term (30 year) storage.
I I
l 3-24 0076L/0011L 1
l l
-___-_-__-___________w
q l
TABLE 3-6 ESTIMATED UNCERTAINTIES IN REACTIVITY DUE TO FUEL DEPLETION EFFECTS Design 0.0005-Initial Burnup ~ Times Design Reactivit Loss, dk 1)
Enrichment Mwd /kgU Burnup, A k k.
0 0.9297 0 1.6 0 0.0579 2.0 5,99 0.0030 0.9267-2.5 12.88 0.0064 0.9233- 0.1284 0.0096 0.9201 0.1828 3.0 19.13 3.5 25.15 0.0126 0.9171 0.2262-30.86 0.0154 0.9143 0.2620 4.0 0.2924 4.5 36.50 0.0183 0.9115 (1) Total reactivity decrease, calculated for the cold, Xe-free
! condition in the fuel storage rack, from the beginning-of-life-to the design burnup.
1 i
O'
' 3-25 0076L/0011L l
1 l
TABLE 3-7
.)
LONG-TERM CHANGES IN REACTIVITY IN STORAGE RACK Storage d k from Shutdown
. Time, (Xenon-free) at 4.5% E years and 36.5 Mwd /kgU 0.5 -0.0047
- 1. 0 -0.0088 10.0 -0.0470 20.0 -0.0673 30.0 -0.0788 I
l OC l
O' 3-26 0076L/0011L
TABLE 3-8 N DESIGN BASIS (LIMITING)
FUEL ASSEMBLY SPECIFICATIONS (CE 14 x 14)
Fuel Rod Data Cladding outside diameter, in. 0.440 Cladding thickness, in. 0.028 Cladding material' Zircaloy-4 Pellet diameter, in. 0.377 UO2 stack density, g/cm 3 '10.281 + 0.031' Enrichment, wt% U-235 4.5,f 0.05 Fuel Assembly Data Maximum number of fual rods 176 (14 x 14 array) s Fuel rod pitch, in.
0.577 + 0.0023
.7 Control rod guide tube' Number 5 Outside diameter, in. 1.115 Inside diameter, in. 1.035 Material Zircaloy-4 U-235 Loading ,
grams / axial em of assembly 51.7 1 0.7 3-27 0076L/0011L i .
L-_-_____---_ _ _ _
TABLE 3-9 THERMAL / HYDRAULIC CASES TREATED *
- 1. Normal Batch Discharge Irradiation time: 54 months (1.42 x 108 ,,e,)
- Addition of the most recent batch : 150 hours0.00174 days <br />0.0417 hours <br />2.480159e-4 weeks <br />5.7075e-5 months <br /> after shutdown Batch size: 80 assemblies
- 2. Full Core Discharge Irradiation time: 73 assemblies 90 days 72 assemblies 21 months 72 assemblies 39 months
- Fuel transfer begins ys after shutdown.
o.c l j
i l
- The pool has total storage capacity of 1706 storage cells. It is conservatively assumed that 18 batches of 80 assemblies have been previously discharged at 18 month intervals. Each assembly in these ous discharges has had 54 months of exposure at full power (12.44 O
3-28 0076L/0011L
O. j- TABLE 3-10 PEAKING FACTOR DATA 1 Maximum Radial Maximum Axial Peaking Factor Peaking Factor Fuel St. Lucie Unit 1 1.67 1.32 4
CE 14 x 14 and Exxon 14 x 14 St Lucie Unit 2, 1.75 1.35 i
CE 16 x 16 1 l
1 I
l
[
j Ot 5
0 3-29 0076L/0011L
> TABLE 3-11 ESSENTIAL HEAT TRANSFER DATA FOR THE FUEL POOL HEAT EXCHANGER Number of heat exchangers: one Coolant'. flow rate:. 3560 sps Temperature effectiveness: 0.36 (two pumps)*-
0.263 (one pump) 4380 sq. ft.
Heat transfer surface area:
)
Overall heat transfer coefficient (fouled) )
(two pumps): 260 Btu /sq.ft.-hr OF O.
- Temperature efficiency of the heat exchanger is calculated in the following manner, using the information provided in the FSAR:
Cooling water outlet - inlet
=
P Pool water inlet - cooling water inlet 118-100~
150-100
= .36 i
l t
O 3-30 0076L/00111 ,
i
TABLE 3-12 POWER GENERATION RATIO PREVIOUSLY DISCHARGED BATCHES Batch Batch Time After Shut Reactor Exposure Non Dimensional Down in Days Time in Days Power Gen.. Ratio No. Size 80 9719.9 1643.5 .00487 1
9179.9 1643.5 .00505 2 80 3 80 8639.9 1643.5 .00523 80 8099.9 1643.5 .00542 4
5 80 7559.9 1643.5 .00562 7019.9 1643.5 .00582 6 80 7 80 6479.9 1643.5 .00603 80 5939.9 1643.5 .00624 8
9 80 5399.9 1643.5 .00647 80 4859.9 1643.5 .00670 10 11 80 4319.9 1643.5 .00694 80 3779.9 1643.5 .00720 12 13 80 3239.9 1643.5 .00746 80 2699.9 1643.5 .00776 14 C,.1 15 80 2159.9 1643.5 .00815 80 1619.9 1643.5 .00888 16 17 80 1079.9 1643.5 .01097 80 540.0 1643.5 .01893 18 CUMULATIVE DIMENSIONLESS POWER = 1.3374E - 01 O
3-31 0076L/0011L
~
TABLE 3-13 BULK POOL TEMPERATURE VS. TIME DURING NOPJfAL REFUELING DISCHARGE Time Bulk Fool Heat Generation (Hrs.) Temp. (DF)- Rate (Btu /hr) 150.00* 106.0 .5689E + 07 151.00 108.8 .1643E + 08-
- This table contains only two lines of output data. . This is due to the fact that the discharge is assumed to take place instantaneously, simulated by one hour in this computer run.
1
-5 O
3-32 0076L/0011L E - -_-_ _ _ _ _ _ _ _ _ _ - _ _ _ _
v- -
< j
.q TABLt 3-14 POOL BULK TEMPERATURE VS. TIME SUBSEQUENT TO COMPLETION OF NORMAL REFUELING DISCHARGE l
l l
Time Bulk Pool Heat Generation l (Hrs.) Temp. (OF) Rate (Btu /hr) I
[
151.00 108.8 .1642E + 08 ;
161.00 130.0 .1613E + 08 l 171.00 133.2 .1588E + 08 181.00 133.3 .1565E + 08 191.00 133.0 .1544E + 08 201.00 132.6 .1525E + 08 211.00 132.2 .1507E + 08 221.00 131.8 '
.1490E + 08 231.00 131.5 .1475E + 08 241.00 131.1 .1461E + 08 251.00 130.8 .1447E + 08 261.00 130.6 .1435E + 08 271.00 -
130.3 .1423E + 08 c C' 281.00 291.00 130.1 129.8
.1411E + 08
.1401E + 08 301.00 129.6 .1390E + 08 311.00 129.4 .1380E + 08 321.00 129.2 .1371E + 08 331.00 129.0 .1362E + 08 341.00 128.8 .1353E + 08 351.00 128.6 .1344E + 08 361.00 128.4 .1336E + 08 371.00 128.3 .1328E + 08 381.00 128.1 .1320E + 08 391.00 127.9 .1313E + 08 O
3-33 0076L/0011L
i f
O' !
TABLE 3-15 i LOSS OF COOLING AFTER COMPLETION OF NORMAL REFUELING DISCHARGE J Rate of Rate of !
Evaporation Level Change l Time to Boil Case (hrs) (1bm/hr) (inch /hr) ]
I 16.79 16933.0 2.67 When heat generation is maximum 2.57 j When the bulk 13.43 16294.0 pool temperature j La maximum 1
I l
l ~
l l
3-34 0076L/00111 E----_---.---_-----.--_-__
TABLE 3-16 BULK POOL TEMPERATURE VS TIME DURING FULL CORE DISCHARGE Time Bulk Pool Heat Generation (Hrs.) Temp. (OF) Rate (Btu /hr) 168.00* 113.6 .8690E + 07 169.00 117.8 .3371E + 08
- This table contains only two, lines of output data. This-is due to the fact that the discharge is assumed to take place instantaneously, simulated by one
( ,-.
' hour in this computer run.
4 O
3-35 0076L/0011L h_-- - __ _ _ _ _ _ _ _ _ _ _ ______.____m_. . _ . _ _ _ _ _ . _ _ - . _ _ _ _ _ _ _ _ _ - - _ . _ _ _ . _ _ _ .
() TABLE 3-17 POOL BULK TEMPERATURE VS TIME SUBSEQUENT TO COMPLETION OF FULL CORE DISCHARGE Time Bulk Pool Heat Generation (Hrs.) Temp. (OF) Rate (Btu /hr) 169.00 117.8 .3370E + 08 179.00 148.8 .3307E + 08 189,00 150.8 .3249E + 08 l
199.00 150.2 .3197E + 08 209.00 149.4 .3149E + 08 219.00 148.7 .3104E + 08 229.00 148.1 .3062E + 08 239.00 147.4 .3024E + 08 249.00 146.9 .2987E + 08 259.00 146.3 .2953E + 08 269.00 145.8 .2921E + 08 279.00 145.3 .2991E + 08 289.00 144.8 .2862E + 08
- 299.00 144.4 .2834E + 08
/',')' 309.00 144.0 .2807E + 08
\s -
143.6 .2782E + 08 l 319.00 329.00 143.2 .2758E + 08 339.00 142.8 .2734E + 08 349.00 142.5 .2712E + 08 359.00 142.1 .2690E + 08 369.00 141.8 .2668E + 08 379.00 -141.5 .2648E + 08 389.00 141.1 .2628E + 08 107 00 140.8 .2608e + 08
'+ 3. 00 140.5 .2589E + 08 3-36 0076L/0011L l
~
(
O' TABLE 3-18 LOSS OF COOLING AFTER COMPLETION OF FULL CORE DISCHARGE Rate of Rate of Time to Boil Evaporation Level Change Case (hrs) (1bs/hr) (inch /hr)
When heat 7.47 34742.2 5.47 generation is maximum When the bulk 5.04 33660.0 5.3 pool temperature is nazimum OC O 3-37 0076L/0011L
TABLE 3-19 LOCAL AND CLADDING TEMPERATURE DATA Maximum Local Maximum Water Cladding Temp. OF Temp. OF-Case Instant l
Normal When the pool heat 155.9 198 8 discharge generation rate is at its peak value When the pool bulk 179.2 219.4 Normal discharge- temperature is at-its peak value Full core When the heat 162.8 209.4 discharge generation rate in the pool is at the peak value Full core When the pool bulk 188.0 222.8 discharge temperature is at its peak value
~
l O 3-38 0076L/0011L
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' FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 ACCEPTABLE SURNUP DOMAIN IN O REGION 2 OF THE ST. LUCfE PLANT FENT FUEL STORAGE RACKS FIGURE 31
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l FLORIDA POWER & LIGHT COMPANY ST. LUCIE PLANT UNIT 1 l
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{ IDEALIZATION OF RACK ASSEMBLY FIGURE 34 L_________-____-___________.
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4.0 MECRANICAL, MATDIAL, AND STRUCTURAL CONSIDERATIONS
4.1 DESCRIPTION
OF STRUCTURE 4.1.1 Description of the Fuel Handling Building The Fuel Handling Building (F B) consists of cast-in-place reinforced concrete interior and exterior walls. It is completely isolated from all other structures. The floors and roof.are of beam and girder construction supported by coluans. A complete description of the F D is provided in Section 3.8.1.1.2 of the St Lucie Unit No. 1 updated FSAR. The F B general arrangement is shown on FSAR Figures 1.2-18 and 1.2-19.
The FB has been designed as a seismic Class I structure in accordance with the criteria outlined in Sections 3.8.1.1.2 and 3.8.1.4 through 3.8.1.7 of the updated FSAR. The building exterior walls, floors and interior partitions are designed to provide plant personnel with the necessary biological radiation shielding and protect the equipment inside from the effects of adverse environmental conditions including tornado and hurricane winds, temperature, external missiles and flooding.
The spent fuel pool is a cast-in place steel lined reinforced concrete tank structure that provides space for storage of spent fuel assemblies, control element assemblies, new fuel during initial core loading and a spent fuel shipping cask. The fuel pool portion of the FB including the walls and roof directly above the pool is designed to withstand, without penetration, the impact of high velocity external missiles that might occur during the passage of a tornado. The design missiles are further discussed in Section 3.5 of the St Lucie Unit No.1 updated FSAR.
} The spent fuel handling system includes interlocks, travel limits and other protective devices to minimize the probability of either mishandling or of
- equipment malfunction that could result in inadvertent damage to a fuel assembly and potential fission product release. The interlocks prevent movement into the walls while limit switches prevent the spent fuel handling machine from raising the fuel above a height where less than nine feet separates the surface of the water from the top of the active fuel length.
A leak detection system is provided in the spent fuel pool to. monitor 100 percent of the pool liner plate veld seams. This system consists of a network of stainless steel angles attached to the outside of the pool liner walls and the underside of the pool liner floor by means of welds and sealed with epoxy material. In the event that one of the weld seams develops a leak, the liquid enters the monitor channel system and flows to one of 19 collection points at the base of the pool, from which the leak can be traced back to a specific pool area.
4.1.2 Description of Spent Fuel Racks The function of the spent fuel storage racks is to provide for storage of spent fuel assemblies in a flooded pool, while maintaining a coolable geometry, preventing criticality, and protecting the fuel assemblies from excessive mechanical or thermal loadings. ,
1 0 4-1 0077L/0011L l
~
I A list of design criteria is given balows
- 1. The racks are designed'in accordance'with the NRC, "0T Position for Review and Acceptance of Spent Fuel Storage and Handling O Applications," dated April 14, 1978 (as amended by the NRC letter dated January 18, 1979) and SRP Section 3.8.4 [1].
- 2. The racks are designed to a2et the nuclear requirements of ANSI N210-1976. The effective multiplication factor, k ff, in the spent fuel pool is less than or equal to 0.95, including all ,
uncertainties and under all credible conditions. l i
- 3. The racks are designed to allow coolant flow such that boiling l in the water channels between the fuel assemblies in the rack l' does not occur. Maximum fuel cladding temperatures are calculated for various pool cooling conditions as described in Section 3.3.
- 4. The racks are designed to seismic Category I requirements, and' are classified as ANS Safety Class 3 and ASME Code Class 3 Component Support Structures. The structural evaluation and seismic analyses are performed using the specified loads and load combinations in Section 4.4.
- 5. The racks are designed to withstand loads without violating the criticality acceptance criteria which may result.from fuel handling accidents and from the nazimus uplift force of the spent fuel handling machine.
T) 6. Each storage position in the racks is designed to support and guide the fuel assembly in a manner that will minimize the '
possibility of application of excessive lateral, axial and bending loads to fuel assemblies during fuel assembly handling and storage. l
)
- 7. The racks are designed to preclude the insertion of a fuel assembly in other than design locations within the rack array.
- 8. The materials used in construction of the racks are compatible with the storage pool environment and will not contaminate the -
fuel assemblies. l 4.1.2.1 Design of Spent Fuel Racks 4.1.2.1.1 Region 1 The rack module is fabricated from ASME SA-240-304L austenitic stainless steel sheet and plate asterial, and SA-351-CF3 casting material and SA-564-630 precipitation hardened stainless steel (to 11000F) for supports only. The weld filler material utilized in body welds is ASME SFA-5.9, Classification ER 308L. Borafier serves as the neutron absorber material. Additional information on Boraflex say be found in Section 3.1.3. The Boraflaz experience list is given in Table 4-1.
O 4-2 0077L/0011L
A typical codule contains storage cells which hava en 8.65-inch noninsi square cross-sectional opening. This dimension ensures that fuel assemblies with maximum expected axial bow can be inserted and removed from the storage cells O without any damage to the fuel assemblies or the rack modules.
O Figure 4-7 shows a horizontal cross-section of a 3 x 3 array. The cells provide a smooth and continuous surface for lateral contact with the fuel assembly. The anatomy of the rack modules is best explained by describing the components of the design, namely:
- Internal Square Tube
- Neutron Absorber saterial (Boraflex)
- Poison sheathing
- Cap element Baseplate
- Support assembly Top Lead-In 4.1.2.1.1.1 Internal Square Tube This element provides the lateral bearing surface to the fuel assembly. It is fabricated by joining two formed channels (Figure 4-1) using a controlled seam welding operation. This element is an 8.65-inch square (nominal) cross-section by 169 inches long.
4.1.2.1.1.2 Neutron Absorber Material (Boraflex)
Boraflex is placed on all four sides of a square tube over a length of 143" (minimum), which provides the requisite B-10 screen for all stored assemblies including a four-inch shrinkage allowance.
4.1.2.1.1.3 Absorber Sheathing l The absorber sheathing (cover plate), shown in Figure 4-2, serves to position and retain the absorber saterial in its designated space. This is accomplished by spot welding the cover sheet to the square tube along the former's edges at numerous (at least 20) locations. This manner of attachment ensures that the absorber material will not sag or laterally displace during fabrication processes and under any subsequent loading condition.
4.1.2.1.1.4 Gap Element Cap elements, illustrated in Figure 4-3, position two inner boxes at a predetermined distance to maintain the minimum flux trap gap required between two boxes. The gap element is welded to the inner box by fillet welds. An array of composite box assemblies welded as indicated in Figure 4-7 forms the honeycomb gridwork of cells which harnesses the structural strength of all sheet and plate type sembers in an efficient manner. The array of composite boxes has overall bending, torsional, and axial rigidities which are an order of magnitude greater than configurations utilizing grid bar type of construction.
4-3 0077L/0011L I
4.1. 2 .1.1. 5 Base plate The baseplate is a 3/4-inch thick plate type member which has 6-inch diameter holes concentrically located with respect to the internal square tube, except f
at support leg locations, where the hole size is 5 inches in diameter. Rese holes provide the primary path for coolant flow. Secondary flow paths are available between adjacent cells via the lateral flow holes (1 inch in diameter) near the root of the honeycomb (Figure 4-4) which preclude flow blockages. De honeycomb is welded to the baseplate with 3/32-inch fillet welds.
4 .1. 2 .1.1. 6 Support Assembly Each module has at least four support legs. All supports are adjustable in l
length to enable leveling of the rack. De variable height support assembly consists of a flat-footed spindle which rides into an internally-threaded cylindrical member. The cylindrical member is attached to the underside of I the baseplate through fillet and partial penetration welds. The base of the flat-footed spindle sits. on the pool floor. Leveling of the rack modules is accomplished by turning the square sprocket in the spindle using a long arm (approximately 46 feet long) square head wrench. Figure 4-6 shows a vertical cross-section of the adjustable support assembly.
l De supports elevate the module bastplate approminately 5-5/8 inches above the l pool floor, thus creating the water pienus for coolant flow. The lateral
' holes in the cylindrical seaber provide the coolant entry path leading into the bottom of the storage locations.
i 4 .1. 2 .1.1. 7 Top Lead-In Lead-ins are provided on each cell to facilitate fuel assembly insertion.
}- Contiguous walls of adjacent cells are structurally connected at the lead-ins with a suitable vent opening. These lead-in joints aid in reducing the lateral deflection of the inner square tube due to the tapact of fuel assemblies during the ground action (postulated seismic motion specified in the FSAR). Bis type of construction leads to natural venting locations for the inter-cell space where the neutron absorber material is located. l 4 .1. 2 .1.2 Region 2 Design 2he rack modules in Region 2 are fabricated from the same material as that used for Region 1 modules, i.e., ASME SA-240-304L austenitic stainless steel. 1 As shown in Figure 4-5 a typical Region 2 module storage cell also has an 8.65-inch nominal square cross-sectional opening . Figure 4-6 shows a horizontal cross section of a 3 x 3 array. De rack construction varies from that for Region 1 inasmuch as the stainless steel cover plates, gap elements and top lead-ins are eliminated. Hence, the basic components of this design are as follows:
- Inner tube
- Neutron absorber material ,
- Side strips I
- Base plate
- Support assembly l 0077L/0011L 4 -4 Revision 1 s
I In this construction, two chennal elem nts form the cell of cn 8.65-inch f nominal square cross-sectional opening. The poison ccterial is pieced between t Stainless steel side strips are inserted on l two boxes as shown in Figure 4-8.
both sides of the poison material to firmly locate it in the lateral O direction. The botton strip positions the poison material in the vertical direction to envelope the entire active fuel length of a fuel assembly (Figure 4-5). Two adjacent boxes and the side strip between boxes are welded together 3
as shown in Figure 4-8, to form the honeycomb rack module.
The baseplate and support assemblies are incorporated in exactly the same manner as described for Region 1 in the preceding section.
4.1.2.2 Fuel Handling The design of the spent fuel racks.will not af fect the conclusions of the fuel handling accidents presented in the FSAR (Section 15.4.3) and summarized by the NRC in the Safety Evaluation Report. That is, the radiological doses for the postulated fuel cask and fuel assembly drop accidents are well within the 10 CFR 100 criteria.
4.2 APPLICABLE CODES, STANDARDS, AND SPECIFICATIONS The design and fabrication of the spent fuel racks and the analysis of the spent fuel pool have been performed in accordance with the applicable portions of the following NRC Regulatory Guides, Standard Review Plan Sections, and published standards:
4.2.1 NRC Documents April 14, 1978 NRC OT Position for Review and Acceptance of a.
(}
V Spent Fuel Storage and Handling Applications, as amended by the NRC letter dated January 18, 1979. l
- b. St Lucie Plant Unit 1 Updated Final Safety Analysis Report, Docket No. 50-335.
- c. NRC Regulatory Guides
~
1.13, Rev 2 Spent Fuel Storage Facility Design Basis Dec. 1981 (Draft) 1.25 Assumptions Used for Evaluating the Potential March 1972 Radiological Consequences of a Fuel Handling Accident in the Fuel Handling and Storage Facility for Boiling and Pressurized Water i i
Reactors 4-5 0077L/0011L l
1.26, Rev 3 Quality Grcup Cicssifications and Stendards Feb. 1976 for Water, Steam and Radioactive Waste Containing Components of Nuclear Power Plants 1.29, Rev 3 Seismic Design Classification Sept. 1978 1.31, Rev 3 Proposed Control of Ferrite Component in Stainless Steel Weld Material 1.71, Rev 0 Welder Qualification for Areas of Limited Accessibility 1.85, Rev 22 Material Code Case Acceptability ASME Section III Division I 1.92, Rev 1 Combining Modal Responses and Spatial Components in Seismic Response Analysis 1.124, Rev 1 Service Limits and Load Combinations for. Class Jan. 1978 1 Linear-Type Component Supports 3.41, Rev 1 Validation of Calculational Methods for Nuclear Criticality Safety.
- d. NRC Standard Review Plan - NUREG-0800 Rev 1, July 1981 Section 3.7, Seismic Design
/7* Rev 1, July 1981 Section 3.8.4, other Seismic Category I QU Structures, Appendix D Rev 3, July 1981 Section 9.1.2, Spent Fuel Storage Rev 1 July 1981 Section 9.1.3, Spent Fuel Pool Cooling System Rev 2, July 1981 NRC Branch Technical Position.
ASS 9-2, Residual Decay Energy for Light Water Reactors for Long Tara Cooling
- e. General Design Criteria for Nuclear Power Plants, Code of Federal Regulations, Title 10, Part 50, Appendis A (GDC Nos. 1, 2, 61, 62 and 63)
- f. NUREG-0612 control of Neavy loads at Nuclear Power Plants.
O 4-6 0077L/0011L
4.2.2 Industry Codes cod Secadards ANSI N14.6-1978 American National Standard for Special Lifting Devices for Shipping Containers Weighing 10,000 iO Pounds or More for Nuclear Materials ANSI N16.1-75 Nuclear Criticality Safety in Operations with Fissionable Matericis Outside Reactors ANSI N16.9-75 Validation of Calculation Methods for Nuclear Criticality Safety ANSI N18.2-1973 Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants ANSI N45.2.2 Packaging, Shipping, Receiving, Storage and Handling of Items for Nuclear Power Plants ANSI N45.2.1 Cleaning of Fluid Systems and Associated Components during Construction Phase of Nuclear Power Plants. j ANSI N45.2.11 1974 Quality Assurance Requirements for the Design of Nuclear Power Plants ANSI ANS-57.2-1983 Design Requirements for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Plants ANSI N210-76 Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations
(
'd{ ASKE Section III Nuclear Power Plant Components, Subsection NF (1983 Edition up to and in-cluding Summer 1984 Addenda ACI-ASME Code for Concrete Reactoi Yessels and Section III, Containment:
Division 2 (1977 Edition)
ACI 316-63 Building Code Requirements for Reinforced Concrete AISC 1980 Specification for the Design, Fabrication and Erection of Structural Steel for Buildings, Eighth Edition AWS D1.1 Structural Welding Code ASNT-TC-1A American Society for Nondestructive Testing June 1980 (Recommended Practice for Personnel Qualification)
O 4-7 0077L/co11L
ASME II Part A Material Sp2cifiestions Part A Forrous, Part C
&C Walding Rods, Elsctrodes cod Filler Metals .
(1983 Edition
-. up to and including k Summer 1984 j
Addenda)
ASHE IX_ Welding & Brazing Qualifications (1983 Edition up to'and in-cluding Summer f
1984 Addenda) l ASME Boiler and Non-destructive Examination Pressure Vessel,Section V, (1983 Edition up to and including Summer 1984 Addenda) 4.3 SEISHIC AV.D IMPACT LOADS The objective of the seismic analysis of the spent fuel racks is to determine the structural responses resulting from the simultaneous application of three orthogonal seismic excitations, n e method of analysis employed is the time !
history method.
4 Seismic floor response spectra for the spent fuel pool floor have been
- developed using the methods described in Subsections 3.7.1 and 3.7.2 of the St )
Lucie Unit No 1 Updated ESAR. The parameters of the original lumped mass l model of the Fuel Handling Building were adjusted to reflect the increased mass corresponding to the new high density spent fuel storage racks. The resulting floor response spectra are shown in Figure 4-9. These spectra were then used to generate statistically independent time history excitations, one for each of the three orthogonal directions. Since the spent fuel racks have no connection with the pool walls-or with each other, the pool floor time histories are used as input to the dynamic analysis of the racks, as described in Subsection 4.5.2.2.1. Fluid coupling is also considered as described therein.
Deflection or movements of racks under earthquake loading is limited by design such that the nuclear parameters outlined in Section 3.1 are not exceeded.
Impact loads have been considered as discussed in Subsection 4.6.4.
The interaction between the fuel assemblies and the rack has been considered, particularly gap effects. The resulting impact loads are of small magnitudes so there is no structural damage to the fuel assemblies.
The spent fuel pool structure has been reanalyzed for the increased dead, thermal and seismic loading resulting from the storage of additional fuel assemblies in the pool, as described in Subsection 4.5.1.
O 4-8 0077L/0011L
v - - - - - ____ _ _________ _ ___ _ _ _ _ _
4.4 10 ADS AND IDAD COMBINATIONS j
4.4.1 Spent Fuel Pool 4.4.1.1 Loads The following design loads were considered in the spent fuel pool analysis:
a) Structural Desd Losd (D)
Dead load consists of the dead weight of the spent fuel racks, the pool water and the concrete structure, superstructure, walls and miscellaneous building items within the Fuel Handling Building.
b) Live Load (L)
Live loads are random temporary load conditions for maintenance which include the spent fuel cask dead weight.
c) Seismic Loads (SSE and OBE)
Seismic loads include the loads induced by Safe Shutdown Earthquake (SSE) and Operating Basis Earthquake (OBE). The hydrodynamic load during the earthquake events was also considered.
d) Normal Operating Thermal Loads (T)
The load induced by normal thermal gradients existing between the building interior and the ambient external environment was e considered. The conditions are:
Gummer
- Interior water temperature 1500F
- Exterior air temperature 930F
- Soil temperature - 700F Winter
- Interior water temperature 1500F
- Erterior air temperature 320F
- Soil temperature 700F f
For all cases, the "as constructed" concrete temperature was assumed to be 700F. A linear gradient through the wall and sat was assumed.
O 4-9 0077L/0011L
~
e) Accid 2nt (Loss of Fuel Pool Ccoling) Thermal Ltd (TA )
(
The thermal accident temperature for the spent fuel pool water is 2170F throughout the pool. At this temperature, the exterior air temperature at 400 F was assumed for the critical thermal gradient through the wall. 70 F0 soil temperature was used. The thermal .
gradient was assumed to be linear. 1 i
f) Fuel Cask Drop Load (M)
A 25 ton cask drop from the maximum height of 58 feet above the pool floor (Elevation 79.50') was considered. (The cask bottom must )
attain Elevation 77.00' for entry into the building.)
l 4.4.1.2 Load Combinations In the spent fuel pool analysis, the following load combinations, from the St Lucie No.1 Updated FSAR, Section 3.8.1.5, were consideref, l
a) Normal Operation 1.5 (D + T) + 1.8 L l
b) OBE Condition !
1.25 (D + T + OBE + 0.2 L) i c) SSE Condition i
f-~r 1.05 (D + T + 0.2 L) + 1.0 SSE J d) Accident and Cask Drop 1.05 (D + TA + 0.2 L) 1.05 (D + 1 + 0.2 L) + 1.0 M For the evaluation of the liner and liner anchors, the above load combinations are applicable except that load factors for all cases may be taken equal to 1.0 (in accordance with Table CC-3230-1 of ACI-ASME Section III, Division 2) in conjunction with the structural acceptance criteria of this SAR subsection 4.6.1.1.b.
Linear analyses without iterations were performed initially to determine the critical load combinations. As a result, the following loading cases were selected for the non-linear concrete cracking analysis:
l i) 1.5 D + 1.8 L ,
ii) 1.05 (D + T winter + 0.2 L) + 1.0 SSE iii) 1.05 (D + T suaser + 0.2 L) + 1.0 SSE iv) 1.05 (D + 0.2 L) + 1.0 SSE v) 1.05 (D + Tg + 0.2 L) vi) 1.05 (D + T winter + 0.2 L) + 1.0 M vii) 1.05 (D + 0.2 L) + 1.0 M 4-10 0077L/0011L
4.4.2 Sp nt Furl Racks 4.4.2.1 Loads O The following loads were considered in the rack design:
Q Dead Load (D) = Dead weight-induced stresses (including fuel assembly weight).
(D') = Dead weight of empty rack.
Live Load (L) = 0 for the structure, since there are no moving objects in the rack load path.
Fuel Drop (Fd ) = Force caused by the accidental drop of the Accident heaviest load from the nazimum possible height.
Load (See Section 4.6.6. )
Crane (Pf ) = Upward force on the racks caused by postulated Uplift stuck fuel assembly (4000 lbs).
Load Seismic (E) = Operating Basis Earthquake.
Loads (E') = Safe Shutdown Earthquake.
Thermal (To) = Differential temperature induced loads (normal Loads condition).
Differential temperature induced loads !
(T,) =
(abnormal design condition). For upset and emergency conditions. T,is the differential temperature for the full core offload condition. For faulted conditions, Ta is the differential temperature for the loss of cooling condition.
and To cause local thermal stresses to be produced. The The conditions worst situation wT,ill be obtained when an isolated storage location has a fuel assembly which is generating heat at the maximum postulated rate. The !
surrounding storage locations are assumed to contain no fuel. The heated water makes unobstructed contact with the inside of the storage walls, thereby producing the maximum possible temperature difference between the adjacer.t celle. The secondary stresses thus produced are limited to the body of the rack; that is, the support legs do not experience the secondary (thermal) stresses.
l I
4.4.2.2 Load Combinations Each component operating condition has been evaluated for the applicable leading combinations listed below:
l' O 4-11 0077L/0011L l
j1
l l
c) Norani Condition D+L D+L+T o D+L+To+E l D' + T o b) Upset Condition D + L + T, + E D + L + T, + Pf j i
D + T, + Fp c) Emergency Condition D + T ,+ Pg + E D + T, + FD+E d) Faulted Condition D + L + T,+ E' {
\
D + L + Fp.
D+L+Pf 4.5 DESIGN AND ANALYSIS PROCEDURES 4.5.1 Design and Analysis Procedures for the Spent Fuel Pool I
4.5.1.1 Spent Fuel Pool Structure Finite Element Analysis V In this analysis, the EBS/NASTRAN program, developed by Ebasco and linked to l the commercially available NASTRAN program, was used. Various layers of I concrete and reinforcing bars were used to determine the effects of concrete cracking. The nonlinear analysis scheme based on the combination of stiffness iteration and load iteration methods, which were available in EBS/NASTRAN program, was used to automatically determine the stresses in the concrete and reinforcing bars after the concrete cracks. The finite element model used in this analysis can be summarized at follows:
a) Since the effect of the additional fuel rack load on the pool floor is limited to the sat in the pool area, the upper portion of the pool walls is not required for the re-evaluation. Therefore, the finite element model included the lower portion of walls, the pool floor (nat) and the underlying soil. The structural components included in the model are shown on Figure 4-10. The cut-off boundary of the walls is at EL. 45.25 ft. i b) The following boundary conditions were used at the model cut-off boundariest
- 1) South end of the sat - Rotational springs representing the bending resistance of the cut-off mat were provided.
O 4-12 co77t/co11L j l
1
a
- 11) Top of the walls - The rotation cbout th2 cris parcliel to the edge of the wall w:s restrain 2d to censidst the offect of the cut-off wall. This assumed boundary condition has little effect on the response of the pool mat, since the boundary is far above the mat. This was demonstrated in the linear analysis results.
iii) South end of east and west walls - Since the rig'.dity of the l cut-off walls is very small, a free boundary coradition was assumed.
A computer plot of the finite element model is presented in Figure 4-11 which I shows the overall view of the model indicating the composite of the four exterior and one interior walls.
4.5.1.2 Liner and Anchorage Analysis The liner and its anchors were evaluated for the temperature load, the strain induced load due to the deformation of the floor, and the horizontal seismic load. The program iPOSBUKF developed by Ebasco was used for the liner buckling analysis due to the temperature and strain induced loads. This program is capable of determining the post-buckling stress / strain if the liner plate buckles. The effect of the hydrostatic pressure was considered in this analysis. In calculating the in-plane shear due to the horizontal seismic loads transmitted from the fuel rack to the liner, the marinum assumed 4
friction coefficient of 0.8 was used.
i The liner anchors were evaluated for the unbalanced liner in-plane force due j to the temperature and strain induced loads, as well as the horizontal seismic j l
in-pla u shear force. i
- 4. 5.1. 3 Foundation Stability and Soil Bearing A detailed soil bearing evaluation was performed for the increased fuel rack ;
loading. The soil stresses were obtained at each sat corner and compared to the allowable value. Stability calculations were performed for overturning and sliding.
4.5.2 Design and Analysis Procedures for Spent Fuel Storage Racks i
The purpose of this subsection is to demonstrate the structural adequacy of the spent fuel rack design under normal and accident loading conditions. The method of analysis presented herein uses a time-history integration method similar to that previously used in the Licensing Reports on High Density Fuel Racks for Fermi 2 (Docket No 50-341), Quad Cities 1 and 2 (Docket Nos 50-254 and 50-265), Rancho Seco (Docket No 50-312), Grand Gulf Unit 1 (Docket No 50-416), Oyster Creek (Docket No 50-219), V C Summer (Docket No 50-395),
Diablo Canyon 1 and 2 (Docket Nos 50-275 and 50-323) and Byron Units 1 and 2 (Docket Nos 50-454 and 50-455). The results show that the high density spen
- fuel racks are structurally adequate to resist the postulated stress combinations associated with level A, B, C and D conditions as defined in References 1 and 2.
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4.5.2.1 Anslysis Outlina The spent fuel storage racks are seismic Category I equipment. Thus, they are required t emain functional during and after a Safe Shutdown Earthquake 3 As noted previously, these racks are neither anchored to the pool floor nor are they attached Furthermore,to the side walls. The individual rack modules are not interconnected. a particular rack may be completely loaded with fuel assemblies (which corresponds to greatest rack inertia), or it say be completely empty. The coefficient of friction, p, According to 1 betweenthgupportsandpoolfloorisdeterminedasfollows.
the results of 199 tests performed on austenitic stainless Rabinowicz steel plates submerged in water show a mean value of p to be 0.503 with a standard deviation of 0.125. The upper and lower bounds (based on twice the 1
standard deviation) are thus 0.753 and 0.253, respectively. Two separate analyses are performed for the rack assemblies with values of the coefficient of friction equal to 0.2 (lower limit) and 0.8 (upper limit), respectively.
Analyses performed for the geometrically limiting rack modules focus on limiting values of the coefficient of friction, and the number of fuel assemblies stored. Typical cases studied are:
i
- Fully loaded rack (all storage locations occupied),
- = 0.8, 0.2 ( g = coefficient of friction) )
- Nearly empty rack # = 0.8, 0.2
- Rack half full # = 0.2, 0.8 f Pool floor slab acceleration data developed for the Safe Shutdown Earthquake (SSE) are shown in Figures 4-12 through 4-14. The rethod of analysis employed q is the time-history method. The pool slab acceleration data were developed j from the building response spectra.
The objective of the seismic analysis is to determine the structural response (stresses, deformation, rigid body motion, etc) due to sisuistaneous application of the three independent, orthogonal excitations.
The seismic analysis is performed in three steps, namely: )
- 1. Development of a nonlinear dynamic model consisting of inertial sass elements and gap and friction elements. ,
- 2. Generation of the equations of motion and inertial coupling and solution of the equ tions using the " component element time '
integration schese 6, 7) to determine nodal forces and displacement.
- 3. Computation of the detailed stress field in the rack (at the =
critical location) and in the support legs using the nodal forces ,
calculated in the previous step. These stresses are checked against the design limits given in Section 4.6.2.2.
A brief descrfetion of the dynamic sodel follows.
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y i
4.5.2,2 Fual Rock - Fuel Assembly Model Since the racks are not anchored to the pool slab or attached to the pool walls or to each other, they can execute a wide variety of rigid body notions. For example, the rack may slide on the pool floor (so-called
" sliding condition"); one or more legs say momentarily lose contact with the liner (" tipping condition"); or the rack may experience a combination of sliding and tipping conditions. The structural model should permit simulation Since these f
j of these kinematic events with inherent built-in conservatisas. j racks are equipped with girdle bars to dissipate energy due to inter-rack impact (if it occurs), it is also necessary to model the inter-rack impact phenomena in a conservative manner. Similarly, lift off of the support less and subsequent impacts must be modelled using appropriate impact elements, and Coulomb friction between the rack and the pool liner must be sisuisted by appropriate piecewise linear springs. These special attributes of the rack dynamics require a strong emphasis on the modeling of the linear and nonlinear springs, despers, and stop elements. The model outline in the remainder of this section, and the model description in the following section describe the detailed modeling technique to simulate these effects, with emphasis placed an the nonlinearity of the rack seismic response.
4.5.2.2.1 Outline of Model
- a. The fuel rack structure is a folded metal plate assemblage welded to a baseplate and supported on four legs. The rack structure itself is a very rigid structure. Dynamic analysis of typical multicell racks has shown that the motion of the structure is captured almost completely by the behavior of a six degrees-of-freedos structure; therefore, the movement of the rack cross-section at any height is described in terms of the six degrees-of-freedom of the rack base.
I
- b. The seismic motion of a fuel rack is characterized by randos rattling of fuel assemblies in their individual storage locations.
Assuming that all assemblies vibrate in phase obviously exaggerates the computed dynamic loading on the rack structure. This assumption, however, greatly reduces the required degrees-of-freedos needed to model tre fuel assemblies which are represented by five lumped masses located at'different levels of the rack. The centroid i of each fuel assembly mass can be located, relative to the rack structure centroid at that level, so as to simulate a partially ,
)
loaded rack.
- c. The local flexibility of the rack-support interface is modeled conservatively in the analysis.
- d. The rack base support say slide or lift off the pool floor.
l
- e. The pool floor and walls have a specified time-history of seismic accelerations along the three orthogonal directions.
- f. Fluid coupling between rack and assemblies, and between rack and adjacent racks, is simulated by introducing appropriate inertial coupling into the systen kinetic energy, Inclusion of these effects uses the methods of References 4 and 6 for rack / assembly coupling and for rack / rack coupling (see Section 4.5.2.2.3 of this report).
O 0077L/0011L 4-15
m v-4
- 3. Potential impacts between rack and assemblies are accounted for by !
appropriate " compression only" gap elements between masses -involved.
l
- h. Fluid damping between rack and assemblies, and between rack and adjacent rack, is conservatively neglected.
- 1. The supports'are modeled as " compression cnly" elements for the vertical direction and as " rigid links" for dynamic analysis. The ;
bottom of a support leg is attached to a frictional element as described in Section 4.5 2.2.2. The cross-section inertial i properties of the support legs are computed and used in the final I computations to determine support leg stresses.
- j. The effect of sloshing has been shown to be negligible at the bottom ,
of a pool and hence is neglected. { l
- k. Inter-rack impact, if it occurs, is simulated by a series of gap elements at the top and bottom of of the rack in the two horizontal directions. The most conservative case of adjacent rack movement is assumed ; each adjacent rack is assumed to move, completely out of phase with the rack being analyzed.
- 1. The fora drag opposing the motion of the fuel assemblies in the storage locations is conservatively neglected in the results reported herein.
- a. The fora drag opposing the motion'of the fuel rack in the water is also conservatively neglected in the results reported herein.
I
- n. The rattling of the fuel assemblies inside the storage locations causes the " gap" between the fuel assemblies and the cell wall to changeHowever, from a maximum the fluidofcoupling twice thecoefficients nominal gap 8)
(toutilized a theoretical are zero ga p.
based on linear vibration theoryC9). Studies in the literature j show that inclusion of the nonlinear effect (viz., vibration amplitude of the same order of magnitude as the gap) drastically lowers the equipment response (10),
Figure 4-15 shows a schematic of the model. . Six degrees-of-freedom are used to track the motion of the rack structu)e. Figures 4-16 and 4-17 I respectively, show the inter-rack impset springs and fuel assembly / storage i
cell impact springs.
The model for simulating fuel assembly notion incorporates five lumped ;
~
saases. The lower mass is assumed to be attached to the baseplate and to move with the baseplate. The four rattling masses are located at quarter height, half height, three quarter height and top of the rack. Two degrees-of-freedom are used to track the motion of each rattling mass.
The solution procedure described in the following is implemented in computer code DYNARACK, which is a validated computer code under Holtec's Q A program.
O 0077L/0011L 4-16 Revision 1
4.5.2.2.2 Model Description ne absolute degrees-of-freedom associated with each of the mass locations are shown in Figure 4-15. As shown, the discrete mass fractions are located at O heights z=0, 0.25H, 0.5H, 0.75H and H respectively. Table 4-6 gives the degrees-of-freedom and the associated generalized coordinates.
Ug (t) is the pool floor slab displacement seismic time-history. Sus, as tabulated in Table 4-6 and shown in Figure 4-15, there are sixteen degrees-of-freedom in the system. Not shown in Figure 4-15 are the gap elements used to model the support legs and the impacts with adjacent racks.
4.5.2.2.3 Fluid Coupling An ef fect of some significance requiring careful modeling is the so-called
" fluid coupling e f fect" . If one body of mass (m) vibrates adjacent to another body (mass m), and both bodies are submerged in a frictionless fluid medium, then Newton's equations of motion for the two bodies have the form:
(my + M11) 1-M11 kg=appliedforcesonmassmi
-H 21 X1 + (m2 + M22) X2 = applied forces on mass m2 k[, k denote 2
absolute accelerations of mass mi and m2, respectively.
M11, M12e M21 and M22 are fluid coupling coefficients which depend on the shape of the two bodies, their relative disposition, etc. Fritz(9) gives data for Mi j for various body shapes and arrangements. It is to be rY noted that the above equation indicates that the effect of the fluid is to add Q) a certain amount of mass to the body (M11'to body 1), and an external force ,
which is proportional to the acceleration of the adjacent body (mass m2). I mus, the acceleration of the one body affects the force field on another. (
Ibis force is a strong function of the interbody gap, reaching large values l for very small gaps. Bis inertial coupling is called fluid coupling. It has 1
an important ef fect in rack dynamics. Be lateral motion of a fuel assembly inside the storage location will encounter this effect. So will the motion of a rack adjacent to another rack. R ese effects are included in the equations of motion. Se fluid coupling is between fuel array node i and cell wall in Figure 4-17. Furthermore, the rack equations contain coupling terms which model the effect of fluid in the gaps between adjacent racks. De coupling terms modeling the effects of fluid flowing between adjacent racks are computed assuming that all adjacent racks are vibrating 180 degrees out of Phase from the rack being an lyzed. Berefore, only one rack is considered surrounded by a hydrodynamic mass computed as if there were a plane of symmetry located in the middle of the gap region.
Finally, fluid virtual mass is included in the vertical direction vibration equations of the rack; virtual inertia is also added to the governing equation corresponding to the rotational degree-of-freedom, q(t).
f G
0077L/0011L 4 -17 Revision 1
4.5.2.2.4 Damping In reality, damping of the rack notion arises from material hysteresis (material damping), relative intercomponent motion in structures (structural
/]
C damping), and fluid drag effects (fluid damping). In the analysis, a maximu:n of 2% structural damping is imposed on elements of the rack structure during ,
SSE seismic simulations. This is in accordance with the St Lucie Unit 1 FSAR(13). Material and fluid damping are conservatively neglected. The dynamic model has the provision to incorporate fluid damping effects ; however, no fluid damping has been used for this analysis. i 1
4.5.2.2.5 Impact Referring to Figure 4-18, any fuel assembly node may impact the corresponding structural mass node. To simulate this impset, four compression-only gap -
elements around each rattling fuel assembly node are provided (see Figure 1 4 -17 ) . As noted previously, fluid dampers may also be provided in parallel ;
with the springs. The compressive loads developed in these springs provide l the necessary data to evaluate the integrity of the cell wall structure and l stored array during the seismic event. Figure 4-16 shows the location of the impact springs used to simulate any potential for inter-rack impacts. Section 4.5.2.4.2 gives more details on these additional impact springs.
- 4. 5.2.3 Assembly of the Dynamic Model Ihe cartesian coordinate system associated with the rack has the following nomenclature:
o x = Horizontal coordinate along the short direction of rack i rectangular platform o y = Horizonta1' coordinate along the long direction of the rack rectangular platform o z = Vertically upward As described in the preceding sec, tion, the rack, along with the base, supports, and stored fuel assemblies, is modeled for the general three-dimensional (3-D) motion simulation by a fourteen degree-of- freedom model. To simulate the impact and sliding phenomena expected, 60 nonlinear gap elements and 16 nonlinear friction elements are used. Gap and friction elements, with their connectivity and purpose, are presented in Table 4-7.
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 4-18. (Note that only the top rattling mass is shown for clarity.)
The impacts between fuel assemblies and rack show up in the gap element, having local stiffness K I, in Figure 4-18. In Table 4-7, gap elements 5 through 8 are for the vibrating mass at the top of the rack. The support leg spring rates Tg are modeled by elements 1 through 4 in Table 4-7. Nnte that l O
0077L/0011L 4-18 Revision 1
m -- .. -
the local compliance of tha concrete floor is included in K d . To siculate i I
sliding potential, friction elements 1 through 8 in Table 4-7 are employed.
Friction elements 2 and 8, and 4 and 6 (Table 4-7) are represented as Kf in Figure 4-18. he friction of the support / liner interface is modeled by a piecewise linear spring with a suitably large stiffness Kf up to the limiting lateral load, N, where N is the current compression load at the interface between support and liner. At every time step during the transient analysis, the current value of N (either zero for liftoff condition, or a compressive finite value) is computed. Finally, the support rotational ,
l friction springs KR reflect any rotational restraint that may be offered by the foundation. This spring rate is calculated using a modified Boussinesq equation (4) and is included to simulate the resistive soment of the support to counteract rotation of the rack leg in a vertical plane. This rotation spring is also nonlinear, with a zero spring constant value assigned after a certain limiting condition of slab soment loading is reached.
he nonlinearity of these springs (friction elements 9,11,13 and 15 in Table 4-7) reflects the edging limitation imposed on the base of the rack support legs. In this analysis, this effect is neglected; any support leg bending, induced by liner / baseplate friction forces, is resisted by the leg acting as a beam cantilevered from the rack baseplate.
For the 3-D simulation, all support elements (listed in Table 4-7) are included in the model. Coupling between the two horizontal seismic motions is provided both by the offset of the fuel assembly group centroid which causes (
the rotation of the entire rack and by the possibility of liftoff of one or more support lege. he potential exists for the rack to be supported or one or more support legs or to liftoff completely during any instant of a complex
~
i 3-D seismic event. All of these potentia 1 events may be simulated during a 3-D action and have been observed in the results.
- 4. 5.2.4 Time Integration of the Equations of Motion 4.5.2.4.1 Time-History Analysis Using 16 DOF Rack Model l l Having assembled the structural model, the dynamic equations of motion corresponding to each degree-of-f.reedom can be written by using Newton's second law of action; or by using Lagrange's equation. he system of equations can be represented in matriz notation as:
(M) (q) = (Q) + (G)
) where the vector (Q) is a function of nodal displacements and velocities, and
) (G) depends on the coupling inertia and the ground acceleration.
Premultiplying the above equations by [M]-1 renders the resulting equation uncoupled in mass.
We have: (q) = [M]-1 (Q) + [M]-1 (G)
As noted earlier, in the numerical simulations run to verify structural integrity during a seismic event, all elements of the fuel assemblies are assumed to move in phase. his will provide maximum impact force lev:1, and induce additional conservatism in the time-history analysis.
l O
0077L/0011L 4 -19 Revision 1 l
Ihis equation set is mass uncoupisd, displacco2nc coupled, cnd is idos11y The computer suited for numerical solution using a central difference scheme.
,rogram "DYNARACK"* is utilized for this purpose.
Stresses in various portions of the structure are computed from known element ;
forces at each instant of time.
l Dynamic analysis of typical multicell racks has shown that the motion of the structure is captured almost completely by the behavior of a six degree-of-freedom structure; therefore, in this analysis model, the movement ,
of the rack cross-section at any height is described in terms of the rack base I degrees-of-freedom (qi(t), ...q6(t)). We remaining degrees-of-freedomIn this f are associated with horizontal movements of the fuel assembly masses. l dynamic model, five rattling masses are used to represent fuel assembly movement. Therefore, the final dynamic model consists of six degrees-of-freedom for the rack plus ten additional mass degrees-of-freedom for the five rattling masses. The remaining portion of the fuel assembly is assumed to move with the rack base. Thus, the totality of fuel mass is included in the simulation.
l
- 4. 5.2 .4.2 Evaluation of Potential for Inter-Rack Impact Since the racks are closely spaced, the simulation includes impact springs to model the potential for inter-rack impact, especially for low values of the friction coefficient between the support and the pool liner. To account for I
this potential, five inter-rack gap elements were located at each side of the rack at the top and at the baseplate. Figure 4-16 shows the location of these gap elements. Loads in these elements, computed during the dynamic analysis, tre used to assess rack integrity if inter-rack impact occurs.
4.6 STRUCIURAL EVALUATION CRITERIA l 4.6.1 Structural Acceptance Criteria for Spent Fuel Poo. Structure 4.6.1.1 Criteria The stresses / strains resulting from the loading combinations described in Section 4.4.1 satisfy the following acceptance criteria:
a) Spent Fuel Pool Concrete Structure The design stress limits described in Section 3.8.1.6 of St l Lucie Unit No.1 Updated FSAR were used for the evaluation of the spent fuel pool reinforced concrete structural components.
The capacity of all sections was computed in accordance with ACI 318-63 Part IV-B, Ultimate Strength Design.
- The numerical procedure underlying DYNARACK has been previously utilized in licensing of similar racks for Fermi 2 (Docket No 50-341), Quad Cities 1 and 2 (Docket Nos 50-254 and 265), Rancho Seco (Docket No 50-312), Oyster Creek (Docket No 50-219), V C Summer (Docket No 50-395), and Diablo Canyon 1 and 2 (Docket Nos 50-275 and 50-323).
0077L/0011L 4-20 Revision 1
b) Liner cod Liner Anchors The acceptance criteria for the liner and liner anchors is in
/~'N accordance with the requirements specified in Paragraph CC-3720 U and CC-3730 of ACI-ASME Section III, Division 2, Subsection CC and can be summarized as follows:
i i) Liner The strain in the liner induced by thermal loads and the deformation of the pool structures is limited to the allowables presented in Table CC-3720-1 of ACI-ASME Section III Code. Load Combinations (a) and (b) presented in Section 4.4.1.2 of this report are considered as Service Load category and (c) and (d) as Factored Load category as these terms are used in that table.
ii) Liner Anchors The displacement of the liner anchors induced by thermal loads and deformation of the pool structures is limited.
to the allowable presented in Table CC-3730-1 of ACI-ASME Section III Code. Load Combinations (a) and (b) presented in Section 4.4.1.2 of this report are considered as Normal Load category, (c) as Extreme Environmental Load category and (d) as Abnormal Load category as these terms are used in that table.
. 4.6.1.2 Material Properties l The following material properties were used in the analysis of the spent fuel pool structure:
a) Concrete - (f'c = 5,200 psi)
Young's modulus Ec = 3.85 x 10 6 poi Poisson's ratio Fe = 0.17 Thermal Expansion coeff g e = 5.5 x 10-6 1/or b) Rebar Steel -
Young's modulus Es = 29 x 10 6 poi Poisson's ratio F s = 0.30 Thermal Expansion coeff a s = 6.5 x 10-6 1/or Yield Strength = 40,000 psi c) Liner Plate -
6 poi Young'smodulusE[p==28.8x10 Poisson's ratio 0.3 Thermal Expansion coeff a p = 6.5 x 10-6 1/or Yield Strength = 27,500 psi
- 4. 6.1. 3 Results a) Spent Fuel Pool Floor For the nonlinear analysis of the selected loading cases O (Section 4.4.1.2), the maximum stress results in the concrete and rebars are summarized in Table 4-2.
4-21 0077L/0011L
-v It.is observsd that the stresses in wsil reinforcement cre significantly affected by'those loading combinations which-include temperature effects. It is further observed that loading case y, which_has the. largest temperature gradients, has the worst effect on concrete compressive stresses fur both sat O' and wall locations. Also, the average _ reinforcement stress for locations is greatest for this loading case while the sat rebar stress at the cask storage area is greatest for loading case vi.
The safety factor (SF) is defined as ultimate stress divided by maximum actual stress including load factors. The safety factors for the maximum stress are also presented in Table 4-2.
The smallest safety factors for the reinforcement tension and concrete compression are 1.10 and 3.65, respectively, which resulted from loading case y, while the smallest safety factor for concrete shear is 1.05 resulting from loading case vi. This clearly indicates that the shear stress in the concrete is the governing component. The critical location of this shear stress is at the cask storage area of the sat, since the thickness of the sat is the smallest (5 feet) here.
b) Liner and Anchorage The critical loading case for the liner evaluation was loading case v which produced nazimum compressive stress in the liner plate. This compressive stress was due to temperature and the deformation of mat. The buckling analysis result indicated that the liner plate would not buckle, due to the stability effect of the hydrostatic pressure.
Two loading conditions were considered necessary in the liner anchor evaluation; one was the strain-induced load which produced the unbalanced in plane force at the edge of the pool area, and the other was the horizontal seismic load transmitted through the friction between the rack support and the liner.
This horizontal seismic load was assumed to the uniformly distributed at the liner anchors. A maximum friction coefficient of 0.8 was used in calculating this horizontal force. In the liner anchor analysis, the load-deflection relationship for the liner anchor subjected to the liner in-plane force, which is usually obtained from actual test data, is required. Since there were no test data available for the actual anchor size of W8 x 24 the load-deflection test data for a lesser strength anchor angle 3 x 2 x 1/4 were used in this analysis. This is considered to be a conservative approach. :
The results of the liner and liner anchor evaluation are summarized in Table 4-3. The minimum safety factors for liner 1 and liner anchor are 5.20 and 1.33 respectively. It should be noted that the actual safety factor for the liner anchor would be greater than 1.33 if the load-deflection data for the actual anchor size of W8 x 24 were used.
O 4-22 0077L/0011L' r
i
c) Foundation Stability and Soil Baaring A detailed soil bearing evaluation was performed. D e soil pd stresses were obtained at each sat corner and compared to the allowable value.
The results for the critical loading case are summarized in Table 4-4. The minimum safety factor for soil bearing for this loading condition is 1.0.
Stability calculations were performed for overturning and sliding. The results for the critical loading case are summarized in Table 4-5. The minimum safety factors for overturning and sliding for this loading condition are 4.59 and 3.10 respectively.
4.6.2 Structural Acceptance Criteria for Spent Fuel Storage Racks )
4.6.2.1 Criteria There are two sets of criteria to be satisfied by the rack modules:
l j
- a. Kinematic Criterion !
l This criterion seeks to ensure that the rack is a physically )
stable structure. St Lucie racks are designed to sustain )
certain inter-rack impact at designated locations in the rack j modules. Therefore, physical stability of the rack is ]
considered along with the localized inter-rack impacts. j Localized permanent deformation of the module is permissible, so j long as the suberiticality of the stored fuel array is not I violated. f 1
- b. Stress Limits ]
The stress limits of the ASE Code,Section III, Subsection NF, 1983 Edition up to and including Summer 1984 Addenda are used since this code provides the most appropriate and consistent set of limits for various stress types and various loading conditions.
4.6.2.2 Stress Limits for Specified Conditions The following stress limits are derived from the guidelines of the ASE Code,Section III, Subsection NF, in conjunction with the material properties data of Subsection 4.6.2.3.
O 4-23 co77L/0011L
t Normal and Upset Conditions (Level A or Livel B Service Limits) f
- 4. 6.2.2.1 - 1
- a. Allowable stress in tension on a net section = Fe = 0.6 Sy'or
'l Ft = (0.6) (23,150) = 13,890 psi (rack material)
Fg is equivalent to primary membrane stresses l
Fe = (.6) (23,150) = 13,890 psi (upper part of support feet)
= (.6) (101,040) = 60,600 psi (Iower part of support feet) l
- b. On the gross section, allowable stress in ahest is:
Fy = .4 S y
! 1
= (,4) (23,150) = 9,260 pai (main rack body) l l Fv = ( .4 ) (2 3,150) = 9,2 60 psi (upper part of support feet) l
= (.4) (101,040) = 40,400 psi (lower part of support feet) l l
- c. Allowable stress in compression, F,:
l F, = (1-1/2a 2 C{} 1 7 5 3a . a3 3 +7 e 8Cd 2x2 E 1/2 j
where a = k1/r 'and Ce= >
S y
k1/r for the main rack body is based on the full height and cross section of the honeycomb region. Substituting numbers, we obtain, for both support les and honeycomb region:
l F,= 13,890 psi (main rack body)
F, = 13,890 psi (upper part of support feet)
I
= 60,600 psi (lower part of support feet)
- d. Maxinua allowable bending stress at the outermost fiber due to flexure about one plane of symmetry:
l Fb = 0.60 Sy = 13,890 psi (rack only)
I Fs = 13,890 psi (upper part of support feet)
= 60,600 psi (lower part of support feet)
- e. C9abined flexure and compression:
fa Cmx fbx Cay fby O F,
+
Dr Fx b
+
Dy Fby
<1 4-24 Revision 1 0077L/0011L
I where:-
f a = Direct compressive stress in the section fbr = Maximum flexural stress about.x *xis fby = Maximum flexural stress about y-axis Cox = Cay = 0.85 fa D, =1-F',,
fa Dy =1-F',y where:
12 x 2 E F'ex, ey =
bx y 23
\ rbx,y /
lb and rb indicate unbraced length and radius of gyration about the concurrent plane (x or y) and the subscripts x,y e reflect the particular axis of bending.
- f. Combined flexure and compression (or tension):
fa fx b fby
+ + < 1.0 0.6sy Fbz Fby The above requirement abould be met for both the direct tension and compression case.
4.6.2.2.2 Faulted Condition (Level D Service limits)
Paragraph F-1370 (Section III, Appendix F)(2), states that the limits for the Level D condition are the minimum of 1.2 (Sy /Fe) or (0.75u/Ft) times the corresponding limits for Level A condition. Since 1.2 S y is less than 0.7 Su for the rack material, and for the upper part of the support feet, the multiplying factor for the limits is 2.0 for the SSE condition for the upper section. 1he factor is 1.62 for the lower section under SSE !
conditions.
Instead of tabulating the results of these six different stresses as dimensioned values, they are presented in a dimensionless fora.- 2hese so-called stress' factors are defined as the rat's of the . actual developed stress to its specified limiting value. With th.s definition, the limiting value of each stress factor is 1.0 for OBE and 2.0 or 1.62 for the-SSE l condition.
0077L/0011L 4-25 Revision 1
i l
1 l
j 4.6.2.3 Material Properties J The data on the physical properties of the rack and support materials, O obtained from the ASME Boiler & Pressure Vessel Code,Section III, appendices, and supplier's catalog, are listed in Tables 4-8 and 4-9. The reference f l
design temperature for evaluation of material properties is 2000F. )
j
- 4. 6.2 . 4 Results for Rack Analysis l I
Figures 4-12 through 4-14 show the pool slab notion in horizontal x, horizontal y and vertical directions. 2his motion is for the SSE.
Results are abstracted in Table 4-10 for modules B2, H1 and C1 (Figure 2-1). l A complete synopsis of the analysis of these modules subject to the SSE s earthquake motions is prertented in a sussary Table 4-10 which gives the I
bounding values of stress factors Rg (i = 1,2 ,3,4, 5,6 ) . 2he stress factors are defined as:
Ri = Ratio' of direct tensile or compressive stress on a net section to its allowable value (note support feet only support compression)
R2= Ratio of gross shear on a net section to its allowable value R3= Ratio of maximum bending stress due to bending about the x-axis l to its allowable value for the section R4= Ratio of maximum bending stress due to bending about the y-exis i to its allowable value i
R5= combined flexure and compressive factor (as defined in 4.6.2.2.le)
R6= Combined flexure and tension (or compression) factor (as defined in 4.6.2.2.lf)
As stated before, the allowable value of R1 (i = 1,2,3,4,5,6) is 1 for the OBE condition, (except for the lower section of the support where the factor is 1.62), and 2 for the SSE. l The dynamic analysis gives the maximax (saximus in time and in space) values of the stress factors at critical locations in the rack module. Since these i
maximar values are subject to sinor (under 51) variation if the input data (viz., rack baseplate height, call inside dimension) is perturbed within the range of manufacturing tolerances, the bounding values, instead of the actual ,I values, are presented in Table 4-10. The terms in Table 4-10 have the following meaning:
l a implies Ri < 1. 0 I b implies Ri < 1. 5 l c implies Ri < 1.75 .
- i. d impliats Ri < 2.0 1 1 O '
0077L/0011L 4-26 Revision 1 l
j It is found that the results corresponding to SSE are sost' critical when compared with the corresponding allowable limits. . We results given herein are for the SSE. We maximus stress factors (Ri ) are below the limiting O
i value for the SSE condition for all sections. It is noted that the critical load factors reported for the support feet are all for the upper segment of the foot and are to be compared with the limiting value of 2.0.
Analyses have been carried out to show that significant margins of safety exist against local deformation of the fuel storage cell due to rattling j impact of fuel assemblies and against local overstress of impact bars due to inter-rack impact.
Analyses have also been carried out for the OBE condition to demonstrate that the stress factors are below 1.0. Results obtained for all rack sizes and shapes are enveloped by the data presented herein. Overturning has also been !
considered for the cases where racks are adjacent to open areas. l 1
4.6.3 Fuel Handling Crane Uplift Analysis An analysis was performed to demonstrate that the rack can withstand an uplift load of 4,000 pounds produced by a jused fuel assembly. This load, which exceeds the capacity of the fuel handling crane, can be applied to any point of the fuel rack without violating the criticality or structural acceptance 1 criteria. Resulting stresses are within acceptable stress limits, and there I is no change in rack geometry of a magnitude which causes the criticality acceptance criterion to be violated.
4.6.4 Impact Analyses 4.6.4.1 Impact Loading Between Fuel Assembly and Cell Wall he local stress in a cell wall is estimated from peak. impact loads obtained from the dynamic simulations. Plastic analysis is used to obtain the limiting impact load that can be tolerated.' Including a safety margin of 2.0, the j total limiting load for the number of cells is over 5.5 times the actual .
3 saximax (saximus in time and space) value.
4.6.4.2 Impacts between Adjacent Racks All of the dynamic analyses assume, conservatively, that adjacent racks sove completely out of phase. Rus, the highest potential for inter-rack impact is achieved. Based on the dynamic loads obtained in the gap elements simulating adjacent racks, we can study rack integrity in the vicinity of the impact point. The use of framing material around the top of the rack allows the rack to withstand impact loads totaling 1x106 lbs applied along any edge of the rack at any instant of time reaching the fully yielded state above the active fuel region. The actual total rack-to-rack impact loads along any rack edge do not exceed 500,000 lbs, and thus, impacts between racks can be accommodated without violating rack integrity. It is found that pool walls are suf ficiently far away from the racks such that pool wall-to-rack impact does not occur.
4.6.5 Weld Stresses he critical weld locations under seismic loading are at the connection of the l
rack to the baseplate and in the support leg welds. For the rack welds, the allowable weld stress is the ASME Code value of 27,300 psi for fillet welds and 42,000 pai for groove welds (Table NF-3324.5(a)-1, Subsection NF, with reference to Section III, Appendix F).
0077L/0011L 4-27 Revision 1
For the support legs, the allowsble wald strcss is governed by th9 levels outlined in Section 4.6.2 (see NF-3324.5 for partial penetration welds).
Weld stresses due to heating of an isolated hot cell are also computed. Se I
\ assumption used is that a single cell is heated over its entire length to a temperature above the value associated with all surrounding cells. No thermal gradient in the vertical direction is assumed so that the results are conservative. Using the temperatures associated with this unit, the skip welds along the entire cell length do not exceed the allowable value for a thermal loading condition.
4.6.6 Summary of Mechanical Analyses he mathematical model constructed to determine the impact velocity of falling objects is based on several conservative assumptions, such as:
- 1. De virtual mass (see Refs 8-10 for further material on the subject) of the body is conservatively assumed to be equal to its displaced fluid mass. Evidence in the literature (11),
indicates that the virtual mass can be many times higher.
- 2. De minimum frontal area is used for evaluating the drag l
coefficient.
1
- 3. ne drag coefficients utilized in the analysis are the lower bound values reported in the liters::ure(12 J. In particular, at the beginning of the fall when the velocity of the body is small, the corresponding Reynolds number is low, resulting in a large dras coefficient.
' l
- 4. Se falling bodies are assumed to be rigid for the purposes of )
impact stress calculation on the rack. The solution of the l immersed body motion problem is found analytically. Be impact velocity thus computed is used to determine the maximum stress generated due to stress wave propagation.
With this model, the following analyses are performed:
- a. Dropped Fuel Accident I A fuel assembly (weight is conservatively analyzed as 1500 pounds with control rod assembly) is dropped from 36 inches above the module and impacts the base (actual height is 361/2 inches; 1/2 inch dif ference between analysis and actual is considered insignificant). Se final velocity of the dropped fuel assembly (just prior to impact) is calculated and, thus, the total energy at impact is known. To study baseplate integrity, it is assumed that this energy is all directed toward punching of the baseplate in shear and thus transformed into work done by the supporting shear stresses. It is determined that shearing deformation of the baseplate is less than the thickness of the baseplate so that we conclude that local piercing of the baseplate will not occur'. Direct ispect with the pool liner does not occur. We suberiticality of the adjacent fuel assemblies is not violated.
0077L/0011L 4-28 Revision 1
- b. Dropped Fuel Accident II
'\
One fuel assembly drops from 36 (actual maximum distance is 36 1/2 inches; 1/2 half inch difference between actual and analysis O
is deemed insignificant.) inches above the rack and hits the top of the rack. Permanent deformation of the rack is found to be limited to the top region such that the rack cross-sectional geometry at the level of the top of the active fuel (and below) f is not altered. he region of local permanent deformation does 'j' not extend below 6 inches from the rack top. An energy balance approach is used here to obtain the results.
- c. Jammed Fuel Handling Equipment A 4000-pound uplift force is applied at the top of the rack at ]
the " weakest" storage location ; the force is assumed to be applied on one wall of the storage cell boundary as an upward shear force. he plastic deformation is found to be limited to the region well above the top of the active fuel.
1 I
Rese analyses prove that the rack modules are engineered to provide maximum safety against all postulated abnormal and accident conditions.
4.6.7 Definition of Terms Used in Section 4 S1, S2, S3, S4 Support designations Pi Absolute degree-of-freedos number i 91 Relative degree-of-freedos number i p Coefficient of friction ui Pool floor slab displacement time history in the i-th direction z,y coordinates Horizontal direction z coordinate Vertical direction KI Impact spring between fuel assemblies and cell Kg Linear component of friction spring N Compression load in a support foot KR Rotational spring provided by the pool slab Subscript i When used with U or X indicated direction (i =
1 indicates x-direction, i = 2 indicates y-direction, i= 3 indicates z-direction)
[H] Mass Matrix (generic notation)
_ (Q) Generalized coordinate vector k Axial Spring of Support leg locations 0077L/0011L 4-29 Revision 1 l _ _ -
.a 1 J
4.6.8' Lateral Rack Move 2ent Lateral motion of the rack modules under seismic conditions could potentially alter the spacing between rack modules.- However, girdle bars on the modules prevent closing the spacing to less than 1.50 inches, which is greater than the normal flux-trap water-gap in the Region 1 reference design. Region 2 storage celle do not use a flux-trap and the reactivity is insensitive to the spacing between modules. Furthermore, soluble poison would assure that a reactivity less than the design limitation is maintained under all conditions.
4.7 MATERIALS, QUALITY CONTROL, AND SPECIAL CONSTRUCTION TECHNIQUES 4.7.1 Construction Materials Construction materials will conform to the requirements of the ASKE Boiler and Pressure Vessel Code,Section III, Subsection NF. All the materials used in the construction are compatible with the storage pool environment and will not contaminate the fuel assemblies or the pool water. The plates, sheets, I l
strips, bars and structural shapes used for rack construction are Type 304L stainless steel.
4.7.2 Neutron Absorbing Material The neutron absorbing material, Boraflex, used in the St Lucie spent fuel rack construction is manufactured by Bisco and fabricated to the safety-related nuclear criteria of 10 CFR 50, Appendix B. Boraflex is a silicone-based polymer containing fine particles of boron carbide in a homogeneous, stable matrix. The specification for the handling and installation of the poison asterial requires that it will not be installed in a stretched condition. The specification precludes the use of adhesives in the attachment of the borafier l ,
to the rack cell walls. FPL will require that the manufacturing process avoid I
techniques which could pinch the boraflex. The design of the racks requires that additional lengths of boraflex, i.e. greater than the active length of a fuel assembly, be installed to account for anticipated shrinkage of the boraflex.
4.7.3 Quality Assurance The design, procurement, and fabrication of the new high density spent fuel storage racks comply with the pertinent Quality Assurance requirements of Appendix B to 10 CFR 50 as implemented through: FPL's Topical Quality Assurance Report PPL-NQA-100A[9); the Joseph Ost Quality Assurance Plan as described in their QA Manual; the Holtec's Nuclear Quality Assurance plan as
{ described in their QA Manual; and the Ebasco Quality Assurance Program for Nuclear Plants, ETR-1001. All have been approved by the NRC.
4.7.4 Construction Techniques 4.7.4.1 Administrative Controls During Manufacturing and Installation The St. Lucie Unit I new spent fuel storage racks will be manufactured at the Joseph Oat Corp., Camden, New Jersey. This facility is a modern high quality shop with extensive experience in forming, nachining, welding, and assembling nuclear-grade equipment. Forming and welding equipment are specifically designed for fuel rack fabrication and all welders are qualified in accordance with ASME Code Section IX.
4-30 0077L/0011L
To avoid damage to the stored spent fuel during rack raplace.eent, all work on the racks in the spent fuel pool area will be performed using written and i l
approved procedures. These procedures will preclude the movement of the fuel A racks over the stored spent fuel assemblies. ]
Radiation exposures during the removal of the old racks from the pool will be controlled by procedure. For anticipated radiation doses see Table 5-5.
Water levels will be maintained to afford adequate shielding from the direct radiation of the spent fuel. Prior to rack replacement, the cleanup system will be operated to reduce the activity of the pool water to as low a level as can be practically achieved.
l 1
4.7.4.2 Procedure 4.7.4.2.1 Preinsta11ation The following sequence of reinstallation events is anticipated for the spent fuel storage rack replacement for Unit I.
- a. Design and fabricate new spent fuel storage racks.
- b. Prepare modification procedure.
- c. Fabricate and test all special tooling.
- d. Receive and inspect new spent fuel storage racks.
4.7.4.2.2 Installation
/~T The final configuration of the 17 new rack modules in the spent fuel pool is LL shown in Figure 2-1. The installation of these racks will be accomplished in accordance with the following considerations and guidelines:
o A temporary construction crane will be installed and operated in the spent fuel pool area to move new and existing rack modules within the spent fuel pool.
o At no time vill this temporary crane carry a rack module '
directly over another module which contains stored spent fuel.
o The temporary construction crane and all rack modules will be installed and/or removed from the spent fuel pool area with the fuel cask crane.
o All load handling operations in the spent fuel pool area will be conducted in accordance with the criteria of Section 5.1.1 of NUREG-0612, " Control of Heavy Loads at Nuclear Power Plants".
o Spent fuel relocations within the pool will be performed as required to maintain separation between the stored fuel and the rerack operatioes.
O 4-31 0077L/0011L
(.
4.8 TESTING AND IN-SERVICE SURVEILLANCE I i
l 4.8.1 Program Intent A sampling program to verify the integrity of the neutron absorber saterial employed in the high density fuel racks in the St Lucie fuel pool environment is described in the following paragraphs.
I j
ne program is conducted in a manner which allows access to the representative j absorber material samples without disrupting the integrity of the entire fuel i
storage systes. Se program is tailored to evaluate the saterial in normal use !
mode and to forecast future changes using the data base developed.
4.8.2 Description of Specimens The absorber material used in the surveillance program, henceforth referred to as 1
" poison", is representative of the Boraflex saterial'used within the storage systes. . It is of the same composition, produced by the same method, and certified j to the same criteria as the production lot poison. De sample coupon is of a thickness similar to the poison used within the storage system and not less than 5 ,
I by 15 inches on a side. Figure 4-19 shows a typical coupon. Each poison specimen is encased in a stainless steel jacket of an austenitic stainless steel alloy j identical to that used in the storage system, formed so as to encase the poison material and fix it in a position and with tolerances similar to the design used ,
I for the storage system. The jacket is closed by tack welding in such a manner as to retain its form throughout the test period and still allow rapid and easy opening without causing mechanical damage to the poison specimen contained within. ,
he jacket permits wetting and venting of the specimen similar to the actual rack /
g environment.
i 4.8.3 Specimen Evaluation Af ter the removal of the jacketed poison specimen from the cell at a designated time, a careful evaluation of that specimen will be made to determine its actual condition as well as its apparent durability for continued function. Immediately af ter the removal, the specimen and jacket section will be visually examined for any effects of environmental exposure. Specific attention will be directed to the examination of the stainless steel jacket for any evidence of physical degradation. Functional evaluation of the poison asterial will be accomplished by the following measurements:
o A neutron radiograph of the poison specimen aids in the determination of the maintenance of uniformity of the boron distribution.
o Neutron attenuation measurements will allow evaluation of the continued nuclear effectiveness of the poison. Consideration will be given in the analysis of the attenuation measurements to the level of accuracy of such measurements, as indicated by the degree of repeatability normally observed by the testing agency.
o A measurement of the hardness of the poison asterial will establish the continuance of physical and structural durability. Se hardness acceptability criterion requires that the specimen hardness will not
~
be less than hardness listed in the qualifying test document for laboratory test specimen irradiated to 10 reds. The actual hardness measurement will be made after the specimen has been withdrawn from the pool and allowed to air dry for not less than 48 hours5.555556e-4 days <br />0.0133 hours <br />7.936508e-5 weeks <br />1.8264e-5 months <br /> to allow for a meaningful correlation with the pre-irradiated sample. ;
0077L/00111 4 -32 Revision 1 3
o Measurement of the length, the width, and tha overage thickness I and comparison with the pre-exposure data vill indicate i l
dimensional stability within the variation range reported in the
~
Boraflex laboratory test reports.
In the event of any observed deterioration of the coupon that could affect the poison function, an insediate inspection of the " poison panels" in the rack will be performed. =The NRC will be advised immediately if the inspection indicates degradation of the poison asterial.
4.9 REFERENCES
FOR SECTION 4
- 1. Nuclear Regulatory Commission, " Standard Review Plan for the Review of Saf ety Analysis Reports for Nuclear Power Plants", NUREG-0800, Revision
- 1. July 1981.
- 2. ASME Boiler & Pressure Vessel Code,Section III, Subsection NF (1983 Edition up to and including Summer 1984 Addenda).
- 3. USNRC Regulatory Guide 1.29, " Seismic Design Classification," Rev 3, 1978.
- 4. " Friction Coefficients of Water Lubricated Stainless Steels for a Spent l Fuel Rack Facility," Prof. Ernest Rabinowicz, MIT, a report for Boston -
Edison Company, 1976.
- 5. USNRC Regulatory Guide 1.92, " Combining Modal Responses and Spatial Components in Seismic Response Analysis," Rev 1, February 1976.
N; 6. "The Component Element Method in Dynamics with Application to l
b Earthquake and Vehicle Engineering," S Levy and J P D Wilkinson, McGraw Hill, 1976. .
- 7. " Dynamics of Structures," R W Clough and J Pension, McGraw Hill (1975).
- 8. " Mechanical Design of Heat Exchangers and Pressure Vessel Components,"
Chapter 16, K P Singh and A I Soler, Arcturus Publishers, Inc., 1984.
- 9. R J Fritz, "The Effects of Liquids on the Dynamic Motions of Innersed Solids," Journal of Engineering for Industry, Trans. of the ASME, February 1972, pp 167-172.
- 10. " Dynamic Coupling in a closely Spaced Two-Body Systen Vibrating in Liquid Medius: The Case of Fuel Racks," K P Singh and A I Soler, 3rd International Conference on Nuclear Power Safety, Keswick, England, May 1982.
- 11. " Flow Induced Vibration," R D Blevens, VonNostrand (1977).
- 12. " Fluid Mechanics," M C Potter and J F Foss, Ronald Press, p. 459 (1975).
- 13. St Lucie Plant Unit I, Updated Final Safety Analysis Report, Docket No 50-335.
GT 4-33 0077L/0011L
Tcble 4-1 BORAFLEX EXPERIENCE FOR HIGH DENSITY RACKS o Site Plant Type NRC Docket No.
Point Beach 1 & 2 PWR 50-226 & 301-Nine Mile Point 1 BWR 50-220 Oconee 1 & 2 PWR 50-269 & 270 Prairie Island 1 & 2 PWR 50-282 & 306 Calvert Cliffs 2 PWR 50-318 BWR 50-254 & 265 Quad Cities 1 & 2 Watts Bar 1 & 2 PWR .
50-390 & 391 Waterford 3 PWR 50-382 Fermi 2 BWR 50-341 H B Robinson 2 PWR 50-261 River Send 1 -
BWR 50-458 Rancho Seco 1 PWR 50-312 !
1 Nine Mile Point 2 BWR 50-410 .
l 50-400 l Shearon Harris 1 PWR l Millstone 3 - PWR 50-423 ]
Grand Gulf 1 BWR 50-416 Oyater Creek BWR 50-219 V C Summer PWR 50-395 Diablo Canyon 1 & 2 PWR 50-275 & 323 .
Byron Units 1 & 2 PWR 50-454 & 455 CT 4-34 0077L/0011L
' !lll I o
0 6 7 7 5 8 4 0 0 0 0 5 2 9
F 1 2 1 1 1 3 1 S
s _
s _
e t t t t t t r t e
t e e e e a e e l L l l l l l l S t e0 e2 _
e L e6 4 e0 5 e0 e8 7 e8 8 8 5 0 rr A 5 5 1 1G2 8 1 1G1 9 0 4G2 8 1 1G2 8
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4 0 ._
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P t
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AC oe L 8l 3l l 4l 9l 2l 4l _
MN i t L 3 e6 5 0 e6 3 e6 4 e6 0e6 2 e6 2 e2 _
1 MI U
se A 3 9 6 5 6 4 3 1 6 7 G2 6 5 0 t e
sr W - G1 - G2 6G2 - G1 G2 - G1 p S E 2
R ec _
SA rn po 6 4 9 7 7 9 1 8 4
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T E m u t t t t t t t = o p MP m e e e a e 9a a )
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M e
0r I T 8 3 3 2 0 1 3 0tF I
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e L
A 6, e 06 3 8 8 3 8 8 2 3 5 8 2 0 F rre pae r
t r W 8G2 2G1 1G1 1G2 3G2 2G1 1G2 smeS eoh(
S e eCS b 1 0 1 8 6 4 3 r r ee 8 4 5 9 7 5 7 t eeo uR F Sttt eec e S 1 2 2 1 1 1 1
_ if rrra xo
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7t 9t 3 t 3t 3t 5t 0t beooy nn M 3a 7 a 3 a 5 e 0e 7 a 0e 9, e l 9, e l 3, e l 1, e l 4, e l 3, e l 8, e l RCCt T e A 9 1 4 4 8 0 9 3 5 0 eeef M 198 163 195 195 297 2G2 2G5 t tt a aaaS e mma e ) iii=
gS n2 t tt n( o lllF i i1 UUUS d et a sc4 l 1 v
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TABLE 4-3
/'
's STRESS / STRAIN
SUMMARY
FOR LINER AND ANCHORS Strain or Disp 1 Stress or Force Actual Allowable SF Remark Actual Allowable SF
-11.1 No Buckling -- -3.847x10-4 -2.0x10-3 5.20 Constraint-Liner in/in Induced Load kai due to in/in hydrostatic pressure.
Evaluation not required -- 0.025 in 0.0625 in' 2.5 Constraint-Liner Induced Load )
Anchor by Code.
1.525 2.03 1.33 Evaluation not required -
Horizontal kips /in by Code. Seismic Load kips /in (N-3)
NOTE: Actual stress and strain are based on Factored (abnormal)
Load condition and allowables are based on 5arvice (ncrual)
Load condition where these terms are as defined in Paragraph CC-3220 of ACI-ASME Section III Division 2.
6 I
4 i
I I
I l
0077L/0011L j 4-36 l
l
TABLE 4-4 SOIL BEARING SIRESSES (KSF)
Corner Location D + L + SSE N-E 4.1 S-E. 1.7 N-W 12.0 5-W 7.9 NOTE: Allowable soil bearing attass is 12.0 KSF l- ~t l
l i
l
. (
0 4-37 0077L/0011L-
. = - _
TABLE 4-5 STABILITY SAFETY FACTORS :
Earthquake Overturning Sliding Loading Type ' D + L + SSE. D + L + SSE l SSE(N-S) + Vert. Up 7.82 3.10 SSE(E-W) + Vert. Up 4.59 3.31 .
1 i
\
l l
O O 4-38 0077t/0011L
~
l Table 4-6 j
DEGREES OF FREEDOM Displacement Rotation -
Location ug uy ug 6x Sy 8:
(Node) p3 q4 q5 q6 1 p1 p2 1* Point 18 is assumed fixed to base at XB , YB, Z-0 2 Point 2 is assumed attached to rigid rack at the top most point.
2* p7 ' p8 Pi = qi (t) + U i(t)
Other r p9, p10 Battling I l pli, p12 Node points 3* , 4* , 5*
Masses l pl3, pl4 l l wp15, pl6
/~T u
4 1
i CT 0077L/0011L 4 -39 Revision 1
Table 4-7 NUMBERING SYSTEM FOR GAP ELEMENIS AND FRICTION ELEMENTS O
Nonlinear Springs (Cap Elements) (64 total) l I.
Node Location Description Number 1 Support S1 Z compression only element 2 Support S2 Z compression only element ;
3 Support S3 Z compression only element l 4 Support S4 Z compression only element ;
5 2 ,2
- X rack / fuel assembly ispect element 6 2 ,2
- X rack / fuel assembly ispect element ;
7- 2 ,2
- Y rack / fuel assembly impact element i 8 2 ,2
- Y rack / fuel assembly impact element !
l 9-24 other rattling masses 25-44 Bottom cross-section Inter-rack impact elements l of rack (around edge) _
45 44 Top cross section Inter-rack impact elements l of rack (around edge ~) Inter-rack impact elements II. Friction Elements (16 total)
Number Node Location Description 3
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 53 X direction friction 6 Support S3 Y direction friction 7 Support S4 . X direction friction l 8 Support S4 Y direction friction l 9 S1 X Slab soment 10 S1 Y Slab soment 11 S2 X Slab soment ;
12 S2 Y Slab soment 13 53 X Slab soment 14 S3 Y Slab soment 15 54 . X Slab soment 16 S4 Y Slab soment-l l
l 0077L/0011L 4-40 Revision 1 l
I u.___._____._ .. . . _ _ . . . _ _ _ _ _ _ _ _ _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ _ _ . _ _ _ _ _ . _ _ _ _ _ _ _ _ ____.__o
Table 4-8 RACK MATERIAL DATA Young's Yield Ultimate Strength Strength Modulus E (psi) Sy (psi) Su (psi)
Material 65,000 l 304L Stainless 27.9 x 106 23,150 ,
Steel i ASME Table Table f Section III Table I-4.0 I-2 .2 I-3.2 {
Reference l GD O
0077L/0011L 4-41 Revision 1
Table 4-9 l
ADJUSTABLE HEIGHT SUPPORT MAIIRIAL DATA
( (Reference Temperature = 1500F)
Young's Yield Ultimate Modulus 104 Strength Strength Material (psi) ksi kai 23.15 65 l Upper part 27.9 g
(Female)
S A-240-304L 101.04 140 l Lower Part 27.9 (Male)
SA-564430 (age hardened to 11000F) 0077L/0011L 4-42 Revision 1
l TABLE 4-10 BOUNDING VALUES FOR STRESS FACIORS Stress Factors
- R R3 R4 R5 R6 R2 Run No. (2 Upper values for rack base - lower values for upper part of the support feet) l a a a a a SSE a I a E E F F
- = .8, fun l l
Module B2 I
l a a a a j SSE a, a, a a a a l
- = .2, full a a l I Module B2 a a a a a a SSE 10 cells I a i E F p= .8, a l
loaded Module B2 ]
l l
d
- The terms a, b, e and d imply the stress factors Ri (i = 1, 2 .. 6) are bounded by the following limiting values:
l I
- /
a: 1.0 j
b: 1.5 c: 1.75 d: 2 ,
l l
l l
l l O' 0077L/0011L 4-43 Revision 1
TABLE 4-10 (Cont'd) l BOUNDING VALUES For STRESS FAC70RS 1
O Stress Factors
- Rg. R3 R4 R5 R6 R1 Run No. (upper values for rack base - lower values for upper part of the support feet) .g
.1 a a a a a SSE a i i a a 5 5 F * . 8, Module H1 full a a a a a SSE a I I I i E F
- = .8 Nodule Hi Positive x-half loaded a a a- a a a SSE
- " .2 , a i a i I i
. Module Hi Positive x-half loaded 0 4 44 Revision 1 0077L/0011L i
TABLE 4-10 (Cont'd)
BOUNDING VALUES FOR SIRESS FAC70RS O ,
1 Stress Factors
- R1 R R3 R4 R5 R6 Run No.
(2 Upper values for rack base - lower values for upper part of the support feet) 1 a a a a a SSE a I a I I y = .2,10 cells a a loaded Module B2 l
a a a a a SSE a I 2 a E E
- = .8, Positive a x-half loaded Module B2 i
a a a a a a SSE y= .2, Positive a I a i I I x-half loaded Module B2 l
y SSE a a a a a a Full i i E E c c
- = .8, Module G1 D~
0077L/0011L 4-45 Revision 1
m , -__
i l
.j IABLE 4-10 (Cont'd) l BOUNDING VALf1ES FOR SIRESS FACTORS Stress Factors
- R1 Rg R3 R4 R5 R6 Run No. (Upper values for rack base - lower values for upper part of the support feet) a a a a a SSE a i i a i a p = .2 a Module C1 Fall a a s a a a SSE P = .2 a a E 5
- c c Module C1 Positive x half loaded a a a a a a SSE
- = .2, Module C1 i a i i I i-Positive x '
- half loaded a a, a a a a SSE l a a a 5 5 1 P * .2 a Module C1 10 Cells loaded l l
1 l
l O 0077L/0011L 4-46 Revision 1
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5.0 COST / BENEFIT AND ENVIRONMENTAL ASSESSMENT i
5.1 COST / BENEFIT AND THERMAL ASSESSMENT The cost / benefit of the chosen reracking alteration is demonstrated in b the following sections.
5.1.1 Need for Increased Storage Capacity
- a. FPL currently has no contractual arrangements with any fuel reprocessing facilities. l FPL executed three contracts with the Department of Energy (DOE) on June 16, 1983 pursuant to the Nuclear Waste Policy Act of 1982, but the disposal facilities are not expected to be available for spent fuel any earlier than 2003.
l
- b. Table 5-1 includes a proposed refueling schedule for St. Lucie Unit 1 and the expected number of fuel assemblies that will be transferred into the spent fuel pool at each refueling until the l total existing capacity is reached and it becomes impossible to conduct a full refueling past cycle 9-10 shutdown in 1991. At l
present the licensed capacity of Unit 1 is 728 storage cells. All calculations in the table for loss of full core reserve (FCR) are !
based on the number of licensed total cells in the pool. The table is then continued assuming the installation of 1706 replacement ;
cells and loss of full core reserve is projected in the year 2009.
- c. The St. Lucie Unit 1 spent fuel pool is expected to contain 529 spent fuel assemblies at the time of reracking.
I%-
bU- d. Adoption of this proposed spent fuel storage expansion would not necessarily extend the time period that spent fuel assemblies would be stored on site. Spent fuel will be sent offsite for final disposition under existing legislation, but the government facility l is not expected to be available until after 2003. l
- e. The estimated date when.the spent fuel pool will be filled with the proposed increase in storage capucity is provided in Table 5-1. In addition to the fuel assemblies, six storage locations are occupied by non-fuel equipment (e.g. dummy fuel assembly, trash basket, etc.).
f 5.1.2 Estimated Costs Total construct. ion cost associcted with the proposed modification is 8 million dollars. This figure includes the cost of designing and fabricating the spent fuel racks; engineering costs; and installation and support costs at the site.
O 5-1 0078L/0011L
5.1.3 Consideration of Alternatives
- a. There are no operational commercial reprocessing facilities available for FPL's needs in the United States, nor are there l espected to be any in the foreseeable future.
- b. At the present time, fhety are no existing available-independent spent fuel storage. f a :ili '.ies. While plans are being formulated by DOE for construction of a spent' fuel repositu y per the Nuclear Waste Policy Act of 1982, this facility is not expected to be available to accept spent fuel any earlier than 2003.
- c. At present, FPL has no license to transship fuel between facilities.
St. Lucie Unit 1 lost full core reserve' capacity upon startup of cycle 7 in 1987. Permanent transfer of St. Lucie Unit 1 spent fuel to other facilities would only compound storage problems there and is not a viable option.
- d. Estimates for costs of replacement power were calculated based on the'last official rate of return. The assumption was made that the unit could be operated without asintaining full core reserve, thus cycle 09-10 in 1990 would be the last refueling possible with existing storage capacity. Table 5-2 indicates the average yearly fuel cost increases for St. Lucie Unit 1 af ter three years of shutdown. Plant shutdown would place a heavy financial burden on Florida residec*s within FPL's service area and cannot be justified.
5.1.4 Resources Consitted Reracking of the spent fuel pools will not result in any irreversible and i irretrievable commitments of water, land, and air resources. The land area now used for the spent fuel pools vill be used more efficiently by safely increasing the density of fuel storage.
The materials used for new rack fabrication are discussed in Section 4.7.1.
These materials are not expected to significantly foreclose alternatives available with respect to any other licensing actions designed to improve the possible shortage of spent fuel storage capacity. 1 5.1.5 Thermal Impact on the Environment Section 3.2 considered the following: the additional heat load and the anticipated maximum temperature of water in the SFP that would result from the proposed expansion, the resulting increase in evaporation rates, t're additional heat load on component and/or plant cooling water systesis, and whether there will be any significant increase in'the amount of heat released to the environment. As discussed in Section 3.2, the proposed increase in storage capacity will result in an insignificant impact on the environment.
1 l
c O' 5-2 0078L/0011L
1 1
5.2 RADIOLOGICAL EVALUATION 5.2.1 Solid Radwaste l currently, resins sre generated by the SFP purification systes..No Current l O, frequency of resin change out is approximately once per year.
significant ,
increase in volume of solid radioactive wastes.is expected due to the new racks. It -is estimated' that a sinimal amount of additional resins will be generated by the spent fuel pool cleanup systen during reracking. The most recent isotopic analysis of the spent fuel pool resin is present in Table 5-7.
Operating plant experience with high density fuel storage has not indicated '
any ser' suable increase in the solid radioactive wastes generated by the I incre sod fuel storage capability.
5.7.2 Gaseous Releases Table 5-3 summarises the FHB Gaseous releases in 1985 and 1986. No significant increaseo are expected as a result of the raracking.
5.2.3 Personnel Erposure
- a. The range of values for recent (October 1985 refueling) gassa isotopic analyses of spent fuel pool water is shown on Table 5-4.
- b. Operating erperience shows dose rates of less than 10 ares / hour either at the edge or above the center of the spent fuel pools regardless of the quantity of fuel stored. This is not erpected to change with the proposed raracking because radiation levels above the pool are due primarily to radioactivity in the water,
/'~ k- which esperience shows to return to a level of equilibrium.
Stored spent fuel is so well shielded by the water above the
' fuel that dose rates at the top of the pool from this s'urce o are negligible.
- c. There have been negligible concentrations of airborne radioactivity from the spent fuel pools. Operating plant experience with dense fuel storage has shown no noticeable increases in airborne radioactivity above the spent fuel pool or at the site boundary. Recent spent fuel pool airborne radioactivity is depicted in Table 5-8. No significant increases are erpected. f rom more dense storage. l
- d. As stated in Section 5.2.1, raracking and utilization of the new
- racks will result in no significant increase in the radweste generated by the spent fuel pool cleanup system. This is '
because operating experience has shown that with high density storage racks, there is no significant increase in the radioactivity levels in the spent fuel pool water, and no significant increase in the annual man-rea due to the increased l
fuel storage, including the changing of spent fuel' pool cooling systes resins and filters. )
O 5-3 0078L/0011L
e, Most .of the corrosion products " crud" associated with spant fuel storage is released soou ofter.fual is ratoved from tha reactor. Once fuel is placed into the pool storage positions, additional crud contribution is minimal.
The highest possible water level is maintained in the spent fuel Should pool to keep exposure as low as reasonably achievable.
crud buildup ever be detected on the spent fuel pool walls around the pool edge, it could easily be washed down.
f.
There is no sccess underneath the spent fuel pool. During normal operation, the radiation dose rate around the outside of the poo1~could increase locally up to .53 area per hour should freshly discharged fuel be located in the cells adjacent to the pool liner. This dose rate will decrease to below .25 ares per hour after approximately 25 days. The depth of the water above the fuel is sufficient so there will be no sessurable increase in dose rates above the pool due to radiation emitted directly from the fuel.
Operating experience has shown a negligible increase in man-rea due to the increased fuel storage with high density racks. Therefore, a negligible ,
)
increase in the annual man-res is expected at St. Lucie as a result of the increased storage capacity of che spent fuel pools with the higher density l storage racks.
The existing St. Lucie or Turkey Point health physics program did not have to be modified as a result of the previous increase in scorage of spent fuel. It is not anticipated that the health physics program will need to be modified for this increase in storage. capability.-
N
! 5.2.4 Radiation Protection During Re-Rack Activities 5.2.4.1 General Description of Protective Measures The radiation protection aspects of the spent fuel' pool modification are the !
I responsibility of the Plant Health Physicist, who is assisted by his staff, with the support of the Corporate. Health Physicist and his staff. Gamma radiation levels in the pool area are constantly monitored by the station Area l Radiation Monitoring Systes, which has a high level alarm feature.
Additionally, periodic radiation and contamination surveys are conducted in work areas as necessary. Where there is a potential for significant airborne radionuclides concentrations, continuous air samplers can be used in addition to periodic grab sampling . Personnel working in radiologically controlled areas will wear protective clothing and when required by work area conditions respiratory protective equipment, as required by the applicable Radiation Work Permit (RWP). Personnel sonitoring equipment is assigned to and worn by all personnel in the work area. At a minimus, this aquipment consists of a thermoluminescent dosimeter (TLD) and self-reading pocket dosimeter.
Additional personaal sonitoring equipment, such as estremity badges, are utilized as required.
Contamination control measures are used to protect persons from internal exposures to radioactive material and to prevent the spread of contamination. l Work, personnel traffic, and the movement of asterial and equipment-in and out of the area are controlled so as to minimize contamination problems. Ma.terial o 5-4 0078L/0011L o 4 h
end equipaint will be sonitored cad appropriately decontaninsted and/or wrapped prior to removal from the spent fuel pool ares. 2he station radiation protection staff will closely monitor and control all aspects of the. work so that personnel exposures, both internal and external, are maintained as low as O reasonably achievable (ALARA).
Water levels in the spent fuel pool will be maintained to provide adequate shielding from the direct radiation of the spent fuel. prior to rack re placement , the spent fuel pool cleanup system will be operated to reduce the activity of the pool water to as low a level as can be practically achieved.
5.2.4.2 Anticipated Exposures During Reracking Table 5-5 is a summary of expected exposures for each phase of the Unit 1 .
reracking operation. These estimates are made based on the proposed i installation plan, including fuel transfers, the use of long-handled tools, and the onsite decontamination cleanup and packaging of the old storage racks. Also, current pool radioactivity levels were conservatively increased in calculating these exposures. The total occupations 1 exposure for the Unit l 1 reracking operation is conservatively estimated to be between 10 and 15 l i
person-rea. See Table 5-5. .
5.2.5 Rack Disposal i 7he spent fuel storage rack modules that will be removed from the spent fuel pool weigh between 30,000 and 42,000 pounds each. The total weight of these racks is approximately 222 tons and the racks occupy a total uncompacted volume of approximately 11,900 ft 3 . They will-be cleaned of loose
- contamination, packaged and. shipped to a licensed radioactive waste processing facility.
i{
Shipping containers will meet the requirements of DOT regulations pertaining to radioactive waste shipments, including limitations with respect to the waste surface dose and radionuclides activity distribution. Shipping .
containers will be certified to meet all requirements for a strong tight package. The maximum weight of a loaded shipping container will be in accordance with the American Association of State Highway and Transportation Officials (AASH70). Trucks and drivers used for rack and waste transportation will have all permits and qualifications required by the Federal DCT and the DOT for each State through which the truck will pass.
I At the waste processing facility, the racks will be decontaminated to the mariaua extent possible. Remaining portions of the racks and contaminated waste generated from decontamination will be buried at a licensed radioactive waste burial site. In preparing non decontaminable waste for shipment and subsequent burial, volume r' eduction methodologies will be employed such as compaction, combining metallic materials with " soft wante" to minimize void space, and super compsetion where feasible.
O 0078L/0011L 5 -5 Revision 1
- 5. 3 ACCIDENT EVALUATION 3.3.1 Spent Fuel Handling Accidents 5.3.1.1 Fuel Assembly Drop Analysis For a drop on top of the rack, the fuel assembly will come to rest horizontally on top of the rack with a miniaua separation distance.from the fuel of more than 12 inches, suf ficient to preclude neutron coupling. Maximum espected deformation under seismic or accident conditions will not reduce the minimum spacing between the' dropped assembly and the stored fuel assemblies to less than 12 inches. Consequently, fuel assembly drop sceidents will not result in a significant increase in reactivity due to the separation distance. Furthermore, soluble boron in the pool water would outstantially reduce the reactivity and assure that the true reactivity is always less than the limiting value for any conceivable dropped fuel accident.
As discussed in Section 4.1.2.2, the proposed spent fuel pool modifications will not increase the radiological consequences of fuel handling accidents prev $nysly evaluated in Section 9.1.4.3 of the St Lucie Unit 1 Updated FSARl21 5.3.1.2 Cask Drop Analysis 5.3.1.2.1 Cask Handling As discussed in Subsection 9.1.4.3 of the St Lucie Unit 1 Updated FSAR(2),
limit switches prevent movement of the cask beyond the spent fuel cask storage area in the northeast corner of the fuel pool; prevent interference of the N' cask crane bridge, trolley, and hoist with fuel racks or building structures; and restrict vertical lift of the cask to an elevation sufficient to gain entry to the Fuel Handling Building. The rarack progras does'not alter the ]
cask handling procedures described in Updated FSAR Sectson 9.1. h cask etion 5.1.1 handlingcranemeetsthedesignandoperattunalrequirementsof(p/.
of NUP.EG-0612, " Control of Heavy Loads at Nuclear Power Plants" 5.3.1.2.2 Radiological Consequences For the calculation of radiological consequences potentially resulting from a cask drop accident, two cases were evaluated regarding the number of fuel assemblies that are assumed to suffer a loss of integrityt l Case I One-third of a core is placed in the spent fuel pool each year during refueling fer the next 23 years, until the pool is filled. The number of assemblies damaged is equal to the number offloaded during a normal refueling plus the ,
remainder of the pool filled with discharged assemblies from previous refuelings.
O 5-6 0078L/0011L
Csse II: One-third cf a core is placed in tha spent fuel pool e:ch year during refueling for tha azzt 20 years. Following the i 21st year of operation, the entire core is removed f ros the reactor and placed into the pool, which fills the pool. The number of assemblies damaged is equal to a full-core offload plus the remainder of the pool filled with discharged assemblies from previous refuelings. 1 The model for calculating the thyroid and whole-body exclusion area boundary doses incorporates the co f rvativeassumptionsspecifginStandardReview Plan (SRP) Section 15.7.5 and Regulatory Guide 1.25 with the exception that a 1.0 Radial Peaking Factor (RPF) is utilized. An RPF of 1.65 as specified in Regulatory Guide 1.25 is intended to represent the highest ,
burnup fuel assembly. While this value say be appropriate for the analysis of l r
a postulated accident involving a single assembly, it is grossly overconservative when applied to an analysis of a normal refueling batch or a full core whose fuel assemblies have various exposure histories. An RPF of 1.0 has been selected as being more representative for the off-load of one or 4
l more regions from the core and has been applied to each aerembly in the present analysis. The use of a 1.0 RPF for the calculation of cask drop radiological consequeness St. Lucie Unit 1 plant.50)has been previously submitted to the NRC for FPL's The core inventory used in the analysis of the dropped spent fuel cask is given by the St. Lucie Unit 1 Updated FSAR Table 15.4.1-1c. As indicated in the St. Lucie Unit 1 Updated FSAR Table 15.4.1-4 and the St. Lucie Unit 1 Updated FSAR Subsection 2.3.4.3, the 0-2 hour exclusion area boundary (EAB) X/Q value of 8.55 x 10-5 eee/m3 is used for the analysis. The i
results of the analysis demonstrate that by retaining the required decay time of spent fuel in the pool to be the minimum times imposed in Technical Specification 3.9.14 prior to moving a epent fuel cask into the spent fuel
~
l
$ )1-
~
pool, the potential offsite doses are less than 10 percent' of 10 CFR Part 100 limits even if all the assemblies in a full pool are damaged and no credit is i
taken for filtration. For a decay time in the spent fuel pool of either 1180 hours0.0137 days <br />0.328 hours <br />0.00195 weeks <br />4.4899e-4 months <br /> (Case I described above) or a decay time of 1490 hours0.0172 days <br />0.414 hours <br />0.00246 weeks <br />5.66945e-4 months <br /> (Case II), the EAB thyroid dose, which is governing, is approximately 15 res. The SRP 15.7.5 acceptance criterion for these analyses (25 percent of 10 CFR 100) is 75 res.
The whole-body doses calculated fqr Case I and Case II for the corresponding decay times are less than 0.1 res, compared to the SRP 15.7.5 acceptance criterion of 6 res. Accordingly, the Technical Specification 3.9.14 decay time requirements prior to cask handling operations are acceptable. Thislis conservative, since not all spent fuel storage modules located in the pool are susceptible to impact from any single cask drop. Thus, the proposed spent fuel pool modifications do not increase the radiological consequences of the cask drop accident previously evaluated.
5.3.1.2.3 overhead Cranes Except for the area described in Section 5.3.1.2.1, the spent fuel cask erane is not capable of traveling over or into the vicinity of the spent fuel pool.
A complete cask crane component description, cask handling description, and ,
cask crane design evaluation ere provided in Updated FSAR Section 9.1 and will i not be affected as a result of the rarack program.
O 5-7 0078L/0011L
5.3.1.2.4 Acceptability The accident aspects of review establish acceptability with respect.to Sections 5.3.1.2.1 and 5.3.1.2.2 of this report.
' Technical Specification requirements for spent fuel decay time prior to moving a spent fuel cask into the spent fuel pool containing freshly-discharged fuel assemblies result in potential offsite doses less than 10 percent of 10 CFR I Part 100 limits should a dropped cask strike the stored fuel assemblies.
5.3.1.3 Abnormal Location of a Fuel Assembly The abnormal location of a fresh unirradiated fuel assembly of 4.5% enrichment could, in the absence of soluble poison, result in exceeding the design reactivity. limitation (k Thir could occur if the assembly were j to be either positioned off of 0.95).utside and adjacent to a storage rack module or inadvertently loaded into a Region 2 acorage cell, with the latter condition producing the larger positive reactivity increment. Soluble poison, however, is present in the spent fuel pool water (for which credit is permitted under these conditions) and would maintain the reactivity substantially less than the design limitation.
The largest reactivity increase occurs for accidentally placing a new fuel assembly into a Region 2 storage cell with all other cells fully loaded.
Under this condition, the presence of 500 ppa soluble boron assures that the infinite multiplication factor would.not exceed the design basis reactivity.
With the normal concentration of soluble poison present (1720 ppa boron), k.
is normally less than 0.80 and will not be critical even if Region 2 were to be fully loaded with f resh fuel of 4.5% enrichment. Administrative procedures will be used to confirm and assure the continued presence of soluble poison in 1... the spent fuel pool water.
5.3.1.4 New Fuel Storage in Region 2 In a confituing calculation, it was determined that a checkerboard storage pattern in Region 2 would allow new fuel assemblies cf 4.5% enrichment to be safely accommodated without exceeding the limiting 0.95 k ff value. In this l checkerboard loading pattern, the. fuel assemblies are loc,ated on a diagonal array with alternate storage cells empty of any fuel.
The Monte Carlo calculation (AMPX-KENO) resulted in a k. of 0.8084 + 0.0085.
With a one sided K-factor (13) for 95% probability at a 95% confidence level and a 4 k of 0.0125 for uncertainties (Table 3-1 for Region 2), the maximua k . is 0.857, which is substantially less than the 0.95 limiting value. Thus, Region 2 may be safely used for the temporary storage of new fuel assemblies ;
provided the storage configuration is restricted to the checkerboard pattern with alternate storage locations empty of fuel.
5.3.2 Fuel Decay Technical Specification 3.9.14 requires decay time for freshly discharged fuel prior to movement of the cask into the pool. As a result, with the increased storage capacity, the radiological consequences of a c'ask drop are less than 10 percent of the requirements of 10 CFR Part 100.
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5-8 0078L/0011L
5.3.3 Loads over Spant Feel Administrative procedures and Tech Spec 3.9.7 limits the maximum weight of )
loads that say be transported over spent fuel.
5.3.4 Temperature and Water Density Effects The moderator temperature coefficient of reactivity in both regions is l nessgive; a moderator temperature of 40C, with a water density of 1.0 g/cm , was assumed for the raference designs, which assures that the true reactivity will always be lower, regardless of temperature.
Temperature effects on reactivity have been calculated and the results are ,
shown in Table 5-6. Introducing voids in the water internal to the storage l
cell (to simulate boiling) decreased reactivity. Since, at saturation j i
temperature, there is no significant thermal driving force, voids due to boiling will not occur in the outer (flux-trap) water regir,n of Region 1.
5.3.5 Conclusions Since the spent fuel cask will not be handled over er in the vicinity of spent fuel as discussed in Section 5.3.1.2.1, the proposed modification does not result in a significant increase in the probability of the cask drop accident previoggy evaluated in the R Lucie Updated FSAR or Safety Evaluation ReportW1 Furthermore, as shown in Section 5.3.1.2.2, by requiring a '
j minimum decay time for spent fuel prior to moving a spent fuel cask into the spent fuel pool, the potential offsite doses are less than 10 percent of 10 CFR Part 100 lisits should a dropped cask strike the stored fuel assemblies.
The proposed spent fuel pool modifications do not increase the radiological consequences of a cask drop accident previously evaluated.
Since there will be a negligible change in radiological conditions due to the l increased storage capacity of the spent fuel pool, no change is anticipated in l
I the radiation protection program. In addition, the environmental consequences of a postulated fuel handlint accident in the spent fuel pool, described in Updated FSAR Section 15.0, remain unchanged. Therefore, there will be no changeorgpacttoanypreviousdeterminationsoftheFinalEnvironment Based on the foregoing, the proposed amendments will not statement .
significantly affect the quality of the human environment; therefore, under 10 l CFR 51, issuance of a negative declaration is appropriate.
l i
O 5-9 0078L/0011L
j
5.4 REFERENCES
' FOR SECTION 5
- 1. PROMOD III Computer Code, Version 22.8, Energy Management Associates.
- 2. St Lucie Plant - Unit 1, Updated Final Safety Analysis Report, O Docket No. 50-335.
- 3. Nuclear Regulatory Commission, " Control of Neavy Loads at Nuclear ]
Power Plants", NUREG-0612, July 1980.
- 4. Nuclear Regulatory Consission, Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants, NUREG-0800, Revision 1, July 1981.
- 5. Nuclear Regulatory Commission, " Assumption 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," Regulatory Guide 1.25, March 1972.
- 6. St Lucie Plant Unit 1, Final Safety Analysis Report, Section 9.1, Docket No. 50-335.
- 7. St Lucie Plant - Unit 1, Technical Specifications, Facility Operating License DPR-67.
- 8. NUREG 0575, " Final Environmental Impact Statement on Handling and Storage of Spent Light Water Power Reactor Fual," Vol 1-3 USNRC August 1979.
- 9. St Lucie Plant - Unit 1. Safety Evaluation Report. Licenses DPR-Docket Nos. 50-335. !
I
- 10. St Lucie Plant - Unit 1, Final Environmental Statement, Docket No.
50-335.
k 5-10 0078L/0011L
TABLE 5-1 l NUCLEAR FUEL DISCHARGE INFORMATION
^-O- ST LUCIE UNIT 1 Cumulative Total Number of of Spent Fuel-Cycle . Shutdown Assemblies Assemblies No Dates Discharged in the Pool 3/28/78 60 60 01 02 3/31/79 68 128 3/15/80 88 216 03
- 04 9/9/81 64 280 2/26/83 372 05 06 10/21/85 92(1) 73 445 07 2/7/87 84 5298 728 CURRENTLY INSTALLED / USABLE CELLS ACTUAL CYCLE INFORMATION THROUGH CYCLE SEVEN, PROJECTED THEREAFTER 08 10/8/88 72 601 09 3/16/90 80 681 10 10/1/91 64 745 3/15/93 -
.68 813 6 11 877 i
12 10/1/94 64 ,
13 3/15/96 68 945' I 14 10/1/97 64 1009 15 3/15/99 64 1077 16 10/1/2000 64 1141 17 3/15/02 68 1209 18 10/1/03 64 1273 19 3/15/05 - 68 1341 20 10/1/06 64 1405 21 3/15/08 68 1473 22 10/1/09 64 1537(2) 23 3/15/10** 68 1605 24 10/1/11 64 1669 25 3/15/13 68 1737 26 10/1/14 64 1801 END OF LIFE- 3/1/16 217 Final Offload 2018 l
- FULL CORE RESERVE (FCR) LOST AT 511 CELLS WITH CURRENT RACKS; RERACK REQUIRED' TO REGAIN FCR
l (2) FCR LOST AT 2489 CELLS WITH POISONED RERACK (ASSUMES 1706 AVAILABLE Ca STORAGE I4 CATIONS) 5-11 0078L/0011L ,
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TA3LE 5-3 GASEOUS RELEASES
~s%p/, FROM FUEL HANDLING BUILDING i
1985 Radionuclides Curies Xe-133 64.9 Xe-135 9.9 4 I-131 2.74 E-4 (
Kr-85m 1.0 Kr-87 7.94 E-1 Kr-88 1.5 l J 1986 Radionuclides Curies Xe-133 20.1 Xe-135 7.5 I-131 1.45 E-4 Kr-85m 1.0 Kr-87 1.3
'* - Kr-88 1.3 1-133 9.88 E-5 l
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5-13 0078L/0011L
TABLE 5-4 GAMMA ISOTOPIC ANALYSIS SPENT FUEL POOL WATER O.
ACTIVITY RADIONUCLIDES 4.90 E-4 pCi/mi )
Co-58 i Co-60 6.70 E-4 #Ci/a1 Cs-134 5.80 E-4 #Ci/mi Cs-137 9.32 E-4 #Ci/ai H-3 7 70 E-2 #C1/mi-l 1
I 1
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O 5-14 0078L/0011L=
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- TABLE 5-5 i
ANTICIPATED DOSES DURING RERACKING O _
Dose (man-rea)
Project l
'l 0.5 Spent Fuel Movement Install Rack Removal Crane 0.1 and Remove Crane 1 Decontaminate Racks Underwater 2.5 f and Remove free. Pool Decontamination of Racks in 2.0 Cask Washdown Bidg and Crate for Shipment Install New Racks 2.5 Repair Equipment 0.5 Measurement of New Racks and 1.0 Du.my Assembly Testing Health Physics Goverage and Sarveys 2.0
( 1.0 General Entry and Inspection Fuel Pool Floor vacuum and 2.0 Filter Disposal TOTAL 14.1 .
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5-15 0078L/0011L
m-TABLE 5-6
]
EFFECT OF TEMPD.ATURE AND VOID ON CAICULATED -
REACTIVITY OF STORAGE RACK Incremental Reactivity Change, dk j Case Region 2 Region 1 0
4C Reference Raference-l 200C -0.0014 -0.0009 5000 -0.0054 -0.0032 800C -0.0109 -0.0062 1200C -0.0216 -0.0117 1200 + 20% void -0.0767 -0.0446 L
I l
1 l
l O'
5-16 0078L/0011L o - _ _ _ _ _ _ _ _ _ _ _ _ - - _ _ - . _ _ _ _ --
d TABLE 5-7 SPENT FUEL POOL PURIFI' CATION SYSTEM RADIONUCLIDES ANALYSIS REPORT RESIN ACTIVITY RADIONUCLIDES ACTIVITY NON-TRANSURANIC pCi/cm 3 Co-58 56.93 23.25 .;
Cs-137 i Cs-134 17.06 Co-60 6.63 I-131 3.37 cc-136- 1.09 Mn-54 0.44 C-14 6.6'3E-3 Tc-99 1.63E-4 I-129 1.95E-6 H-3 7.6E-2 ;
l Sr-90 3.95E-3 Ni-63 1.39 Fe-55 9.68E-2
.. TRANSURANIC nCi/gm Pu-239, 240 1.39E-5 Pu-241 2.52E-3 Ca-242 9.28E-6 TRU* 3.71E-5 l Resin Volume = 35 ft 3 or 0.991 m3
- Other alpha-emitting transuranic nuclides with half-lives greater than 5 years.
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5-17 0078L/001LL
TA3LE 5-8 SPENT FUEL POOL AIR. BORNE ACTIVITY RADIONUCLIDES ANALYSIS REPORT CONCENTRATION l
RADIONUCLIDES yCi/mi
< 3.0E-11 As-110m l Ar-41 < 9.5E-11 l Ba-139 <1.2E-10
< 7. 8E-11 Ba-140 < 8.5E-11 Ba-141 I ce-139 <2.SE-11 Ce-144 < 1. 5E-10 Ce-141
< 3.6E-11 Co-57 <1.7E-11 Co-58
- 3.6E-11 Co-60
- 4.4E-11 Cr-51
- 2.0E-10 Ca-134
- 2.7E-11 Ca-137
< 3.7E-11 Ca-138 < 2.8E-11 Fe-59 < 4.0E-11 I-131 1.0E-9 I-132 < 3.7E-11 C"' I-133 = 2.0E-11
' - I-134 < 2. 7E-11 I-135' < 8.0E-11 Kr-85 =c5.2E-9
<2.7E-11 Kr-85a Kr-87 <4.3E-11 La-140 < 2.0E-11 I La-142 *9.9E-11 Mn-54
< 2.5E-11 Mo-99
- 1.4E-10 Nb-95 < 2.4E-11 Nb-97 < 3.4E-11
" < " = Lower level of D.;tection and was not detected in the sample.
I-131 and Xe-133 were detected in the sample. Fuel movement and fuel reconstitution was in progress during the sample collection. These are normally not detected and their lower levels of detection are I-331 = 3.0E-11 Ci/mi and Xe-133 = 9.0E-11 pCi/al.
5-18 0078L/0011L
TABLE 5-8 (Cont'd)
SPENT FUEL POOL AIRBORNE ACTIVITY RADIONUCLIDES ANALYSIS REPORT
_ RADIONUCLIDES CONCENTRATION
- C1/mi N;-239 .-
Rb-88 <6.4E-11 Rb-89 <1. 9E-10 Ru-103 <6.7E-11 Ru -106 < 2. 4E-11 Sb-124 <1.8E-10 Sb 125 *2.6E-11 Sa-113 < 5. 5E-11 Sr-85
- 1. 7E-11 Sr-91 < 2. 3E-11 Sr-92 < 7.1E-11 Te-99a <1. 7E-11 Te-132 < 2. 2E-11 I Xe-131a <2.3E-11 Xe-133 <1 1E Xe-133s 3.4E-10 i
g" Xe-135 < 2. 0E-10
~
Xe-138 <1. 3E-10 Y-88 <1. 0E-10 Y-91ts <2.3E-11 l 2r-65 < 3. 7E-11 Zr-95 < 5. 4E-11 f,di Zr-97 1
< 3. 9E-11 i y
< 1.9E-11 a 1
'+ i
< " = Lower level of Detection ce in and was not dete t d the sample.
.I-131 and Xe-133 were detected in the sample.
I-131=r 3.0E-11normally not edetected and their lower f{\
level pCi/a1 and Xe-133 as collection. These are 9.0E-11 pCi/al.s of detection are
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\ 5-16
! 0078L/0011L
\ _ _ _ _ _ _ _ _ _ - _ _