ML20079P405
| ML20079P405 | |
| Person / Time | |
|---|---|
| Site: | Summer  |
| Issue date: | 12/31/1983 |
| From: | SOUTH CAROLINA ELECTRIC & GAS CO. |
| To: | |
| Shared Package | |
| ML20079P400 | List: |
| References | |
| NUDOCS 8401310209 | |
| Download: ML20079P405 (190) | |
Text
-
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LICENSING REPORT ON NEW FUEL STORAGE RACKS FOR THE VIRGIL C.
SUMMER NUCLEAR STATION NRC DOCKET NO. 50-395 SOUTH CAROLINA ELECTRIC AND GAS COMPANY COLUMBIA, SOUTH CAROLINA 29218 DECEMBER 1983 4
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i O 8401310209 840123 PDR ADOCK 05000395 4
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TABLE OF CONTENTS Page SECTION 1 - NEW FUEL STORAGE RACKS I
l.0 New Fuel Storage Racks 1-1 LIST OF FIGURES I
Page SECTION 1 Fig. 1.1 New Fuel Storage Rack 1-2 l
Arrangement I
1.2 Reactivity Effect of Low 1-3 Density Moderator in New Fuel Storage Rack i
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x 1.0 New Fuel Storage Racks The new fuel storage racks for the V.
C.
Summer Nuclear Station consist of two arrays of storage cells, each containing 30 cells in a 2 x 15 pattern.
Figure 1.1 illustrates the new fuel storage cell arrangement and shows the geometry used in the criticality analysis.
Each storage cell consists of a 0.075-in.-
thick stainless-steel box (9.0-in.
inside dimension) with a
minimum of 21 in, between cell centerlines.
Although fuel is normally stored in the dry condition, the criticality analysis considered flooding with
- clean, unborated water ranging in density from 1.0 to hypothetical, low values (e.g.,
fog, mist, or foam).
Preliminary survey calculations with diffusion theory suggested a
second maximum in reactivity peaking at a
hypothetical water density of ~10%-15%.
r~s Since diffusion theory is known to be inadequate in very dry U
lattices, three-dimensional AMPX-KENO calculations were used in the low-moderator-density region to define the maximum reactivity under optimum moderating conditions.
For these calculations, the array of fuel storage cells (using quarter-core geometry as indi-cated in Pig.
1.1) was assumed to be reflected by full-density j
water on the outer boundaries and on both top and bottom of the array.
Low-density water was used within the storage boxes and l
between the array of storage cells.
I l
(
Figure 1.2 shows the calculated reactivities as a function l
of moderator density within and between the storage cells.
These i
calculations indicate a low-density maximum reactivity of ~ 0.72 occurring at a water density of 0.10 g/cm3 This low-density maximum reactivity, however, is less than that (k 0.915) for
=
the fully flooded condition.
In either event, the reactivity is substantially less than the limiting reactivity of 0.98 specified O
in SRP 9.1.1.
Hence, it is concluded that unirradiated fuel of 0
4.3% enrichment may be safely stored in the new fuel racks of the V.
C. Summer Nuclear Station.
1-1
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1.1 New fuel storage rack arrangement.
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.9 1.0 W AT E R DENSITY, gms / cc Fig.
1.2 Reactivity effect of low density moderator in new fuel storage rack.
LICENSING REPORT O
ON HIGH-DENSITY SPENT FUEL RACKS FOR VIRGIL C. SUMMER NUCLEAR STATION NRC DOCKET NO. 50-395 i
O SOUTH CAROLINA ELECTRIC & GAS COLUMBIA, SOUTH CAROLINA 29218
.-DECEMBER, 1983 O
()
TABLE OF CONTENTS Page SECTION 1 - INTRODUCTION 1.0 Introduction 1-1 SECTION 2 - GENERAL ARRhNGEMENT 2.0 General Arrangement 2-1 SECTION 3 - RACK CONSTRUCTION 3.1 Fabrication Details 3-1 3.1.1 Regions 1 and 2 3.1.2 Region 3 3.2 Codes, Standards and Practices for 3-7 the Spent Fuel Pool Modification SECTION 4 - NUCLEAR CRITICALITY ANALYSIS 4.1 Design Bases 4-1 V('T Summary of Criticality Analyses 4-3 4.2 4.2.1 Normal Operating Conditions 4-3 4.2.2 Abnormal and Accident Conditions 4-3 4.3 Reference Fuel Storage Cell 4-9 4.3.1 Reference Fuel Assembly 4-9 4.3.2 Region 1 Storage Rack 4-10 4.3.3 Region 2 Storage Cells 4-10 4.3.4 Region 3 Storage Cells 4-10 l
4.4 Analytical Methodclogy 4-15 4.'4.1 Reference Analytical Method 4-15 and Bias 4.4.2 Manuf acturing Tolerances and 4-16 Uncertainties 4.4.3 Fuel Burnup Calculations 4-16 11
TABLE OF CONTENTS [ continued]
Page
'e 4.4.4 Long-Term Decay 4-18 4.4.5 Effect of Axial Burnup Distribution 4-20 4.5 Reference Subcriticality and Mechanical Tolerance Variations 4-26 4.5.1 Nominal Design Cases 4-26 4.5.2 Boron Loading Variation 4-27 4.5.3 Storage Cell Lattice Pitch 4-28 Variations 4.5.3.1 Inner Water Thickness Variations 4-28 4.5.3.2 Outer (Flux-Trap) Water 4-28 Thickness Variation 4.5.4 Stainless Steel Thickness variations 4-29 4.5.5 Fuel Enrichment and Density Variation 4-29 w
4.5.6 Boraflex Width Tolerance Variation 4-30 (w/
)
4.5.7 Eccentric Positioning of Fuel Assembly in Storage Rack 4-30 4.6 Abnormal and Accident Conditions 4-31 4.6.1 Temperature and Water Density Effects 4-31 4.6.2 Dropped Fuel Assembly Accident 4-32 4.6.3 Abnormal Positioning of Fuel Assembly Outside Storage Rack 4-32 4.6.4 Lataral Rack Movement 4-33 SECTION 5 - THGRMAL-HYDRAULIC CONSIDERATIONS 5-1 5.1 Decay Heat Calculations for the Spent 5-1 Fuel Cooling 5.1.1 Basis 5-1 5.1.2 Model Description 5-3 5.1.3 Decay Heat Calculation Results 5-6
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TABLE OF CONTENTS (Continued) 7-V) 4 5.2 Thermal Hydraulic Analysis for Spent Fuel 5-10 Cooling 5.2.1 Basis 5-10 5.2.2 Model Description 5-11 5.2.3 Results 5-13 REFERENCES TO SECTION 5 5-16 SECTION 6 - STRUCTURAL ANALYSIS 6-1 6.1 Analysis Outline 6-1 6.2 Fuel Rack - Fuel Assembly Model 6-3 6.2.1 Assumptions 6-3 6.2.2 Model Description 6-5 6.2.3 Fluid Coupling 6-6 6.2.4 Damping 6-7 6.2.5 Impact 6-8 6.2.6 Assembly of the Dynamic Model 6-8 6.3 SLress Analysis 6-12 6.3.1 Stiffness Characteristics 6-12 6.3.2 Combined Stresses and Corner Displacements 6-13 l
6.4 Time Integration of the Equations of Motion 6-14 6.5 Structural Acceptance Criteria 6-17 6.6 Results 6-22 6.7 Summary of Mechanical Analyses 6-24 REFERENCES TO SECTION 6 6-27 SECTION 7 - SPENT FUEL POOL FLOOR STRUCTURAL ANALYSIS 7.1 Introduction 7-1 7.2 Analysis Methods 7-1 O
iv
4
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TABLE OF CONTENTS
'u (Continued) 7.'3 Assumptions 7-2 7.4 Load Combinations 7-2 7.5 Results 7-3 7.6 Conclusions 7-4 SECTION 8 - ENVIRONMENTAL EVALUATION 8.1 Summary 8-1 8.2 Characteristics of Stored Fuel 8-2 8.3 Related Industry Experience 8-3 8.4 V.C. Summer Operating Experience 8-4 8.5 Spent Fuel Pool Cooling.and Clean-Up System (FPCC) 8-4 8.6 Fuel Pool Radiation Shie.lding 8-6 8.7 Radiological Consequences 8-6 8.8 Re-racking Operation 8-7 8.9 Conclusions 3-7 References to Section 8 8-9 SECTION 9 INSERVICE SURVEILLANCE PROGRAM FOR 10-1 BORAFLEX NEUTRON ABSORBING MATERIAL 9.1 Program Intent 10-1 9.2 Description of Specimens 10-1
~
9.3 Test 10-2 9.4 Specimen Evaluation 10-2 SECTION 10 COST / BENEFIT ASSESSMENT 10-1 10.1 Specific Needs for Spent Fuel Storage 10-1 O
V
O TABLE OF CONTEt1TS (Continued) 10.2 Cost of Spent Fuel Storage 10-2 10.3 Alternatives to Spent Fuel Storage 10-2 10.4 Resource Commitments 10-4 REFERENCES TO SECTION 10 10-5 SECTION 11 Quality Assurance Program 11-1 11.1 Introduction 11-1 11.2 General 11-1 11.3 System Highli,ghts 11-1 11.4 Summary 11-3 0
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LIST OF FIGURES Page SECTION 2 Fig. 2.1 Module Layout 2-5 SECTION 3 Fig. 3.lA 3x3 Typical Array (Poisoned Cells) Region 1 3-10 3.lB 3x3 Typical Array (Poisoned Cells) Region 2 3-11 3.2(a) Small Angular Subelement 3-11 3.2(b) Large Angular Subelement 3-11 3.3 Basic Cell Element 3-12 3.4A Welding Sequence of Composite Boxes 3-14 3.4B Typical Cell Elevation 3-15 3.5 Adjustable Support 3-16 3.6 Fixed Support 3-17 3.7 3x3 Typical Array (Unpoisoned Cells)
Region 3 3-18 SECTION 4 Fig. 4.1 Limiting discharge fuel burnup for Region 2 storage rack for fuel of various initial enrichments 4-7 4.2 Limiting discharge fuel bt:rnup for Region 3 storage rack for fuel of various initial enrichments 4-8 4.3 Configuration of Region 1 spent fuel i
l storage cell 4-12 1
4.4 Configuration of Region 2 spent fuel storage cell 4-13 4.5 Configuration of Region 3 spent fuel storage cell 4-14 4.6 Decrease in K with fuel burnup under oo I
hot operating conditions 4-22 4.7 Long-term change in infinite multiplication j
factor (K o) of 4.3% enriched fuel burned to o
20 Mwd /kgU 4-23 V1A i
,_-m 4
LIST OF FIGURES (continued}
Page
()
4.8 Long-term change in infinite multiplication factor (koo) for 4.3% enriched fuel burnup 42 Mwd /kgU 4-24 4.9 Relative axial burnup distribution from NUREG/CR-0722 4-25 4.10 Infinite multiplication factor (koo) of 4.3%
enriched fuel assemblies separated only by 4-34 water SECTION 5 Fig. 5.1.1 (a) Pool Bulk Temperature, Normal Discharge 5-17 (b) Pool bulk Temperature, Normal Discharge 5-18 (c) Power Discharge, Normal Discharge 5-19 (d) Power Discharge, Normal Discharge 5-20 5.1.2 (a) Pool Bulk Temperature, Full Core 5-21 Discharge (b) Pool Bulk Temperature, Full Core
(}
Discharge 5-22 (c) Power Discharge, Full Core Discharge 5-23 (d) Power Discharge, Full Core Discharge 5-24 5.1.3 (a) Pool Bulk Temperature, Normal Discharge with loss of 1 SFPHX 5-25 (b) Pool Bulk Temperature, Normal Discharge with loss of 1 SFPHX 5-26 l
(c) Power Discharge, Normal Discharge with loss of 1 SFPHX 5-27 l
(d) Power Discharge, Normal Discharge with l
. loss of 1 SFPHX 5-28 l
1 5.1.4 (a) Pool Bulk Temperature, 1/3 Core Discharge with loss of 2 SFPHX 5-29 (b) Pool Bulk Temperature, 1/3 Core Discharge with loss of 1 SFPHX 5-30 l
O viii i
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LIST OF CIGURES (continued) 0 (c) Power Discharge, 1/3 Core Discharge 5-31 with Loss of 1 SFPHX (d) Power Discharge, 1/3 Core Discharge 5-32 with Loss of 1 SFPHX 5.2.1 Idealization of Rack Assembly 5-33 5.2.2 Thermal Chimney Flow Model 5-34 SECTION 6 Fig. 6.1 Dynamic Model 6-29 6.2 Impact Springs and Fluid Dampers 6-30 6.3 Spring Mass Simulation.or Two Dimensional Motion 6-31 6.4(a) Horizontal Cross Section of Rack 6-32 6.4(b) Vertical Cross Section of Rack 6-32 6.5 Dynamic Model 6-33 6.6 Stress Resultants Orientation 3-34 6.7 Subdivision of a Typical Rack 6-35 6.8 Finite Elements Model Crost Section 6-36 LIST OF FIGURES (continued) 6.9 Horizontal OBE, 4 percent Damping 6-36 6-10 Hotizontal Y, OBE, 4 percent Damping 6-38 6-11 Vertical OBE, 4 percent damping 6-39 1
O
E LIST OF FIGURES
()
(continued)
SE'CTION 7 Spent Fuel Pool Structural Analysis Fig. 7.1 Plan on Spent Fuel Pool 7-7 7.2 Overall Foundation Plan 7-8 7.3 Section Looking West through Spent Fuel Pool 7-9 7.4 Section Looking North through Spent Fuel Pool 7-10 SECTION 9 Fig. 9-1 Test Coupon 9-5 O
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C' LIST OF TABLES N'
s^ s Page I
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SECTION~1
'd Table,1.L
'V.
C.
Summer Nuclear Station Assembly
- - Discharges (Tentati've Schedule) 1-3
' '. - ~ _
SECTION 2~
s
~ Table'2.1 Design' Nodule Data 2-3 Table 2.2 Module Dat'a 2-4 SECTf0N 3 sg s
Table 3.1-Boraflex Experience-for High Density Racks 3-2 SECTION 4 Tabre 4.1 Summar of Criticality Safety Analyses 4-4
's 4.2 Reactivity Effects of Abnormal and Accident Conditions 4-5 4'.3*
Fuel Assembly Design Specifications 4-9
-4.h
'C6mparison of Cold, Clean Reactivities
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-.Ca cu ated at 20 Mwd /kgU and 42 Mwd /kgU 4-17 l
l 4.S's Long-Term Changes in Reactivity 4-19 4.6 Effect of Temperature and Void on
[-
Calcuiated Reactivity of Storage Rack 4-31 SECTION 5 s
s.
Table 5.1.1 List of 'ases Analyzed 5-7 C
5.1)2'Ma'himum Pool Bulk Temperature, t,
Coincident
' Total Power Q1 and Coincident Specific Power for the Hottest Assembly 5-8 5.1 3 Time (Hrs) to Boiling and Boiling Vaporization Rate From the Instant All
Cooling is Lost 5-9 5.2.'l Maximum Local Pool Water Temperature and Local Fuel Cladding Temperature at Instance of Maximum Pool Bulk Temperature 5-14
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LIST OF TABLE (Continued) 5.2.2 Pool and Maximum Cladding Temperature At the 5-15 Instance Fuel Assembly Transfer Begins SECTION 6 Table 6.1 Degrees of Freedom 6-5 6.2 Numbaring System for Springs, Gap Elements, 6-10 Friction Elements 6.3 Physical Property Data 6-19 6.4 Support Material Data 6-19 6.5 File DSCLOl Module A (lixil), Coef =
.8, 6-28.1 Full Rack 6-28.2 SECTION 7 Table 7.1 Caisson Evaluation-Required vs. Minimum Available Capacity 7-5 7.2 Required vs. Available Capacity In Spent Fuel Pool Walls and Slab 7-6 Section 9 Table 9.1 Time Schedule for Removing Coupons 9-4 I
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INTRODUCTION d
The purpose of this report is to describe the design, fabrication, and safety analysis of High Density Spent Fuel Storage Racks produced by Joseph Oat Corporation for Virgil C.
Summer Nuclear Station.
V.C.
Summer Nuclear Station is co-owned by South Carolina Electric and Gas Company and South Carolina Public Service Authority, with the former serving as the principal owner and operating agent.
The plant is located in Jenkinsville, South Carolina, approximately 26 miles northwest of Columbia, South Carolina.
V.C.
Summer is a single unit Pressurized Water Reactor (Westinghouse design) with a design capacity of 900 Mwt(e).
The reactor core contains 157 fuel assemblies rated to produce 2775 thermal megawatts.
At present there are no stored fuel assemblies in the pool.
The power station is slated to go into commercial operation in January, 1984.
The plant is currently licensed for the storage of 682 spent fuel assemblies at a maximum of 3 5% enrichment.
As shown in Table 1.1, the storage pool will lose full core discharge capability in the year 1994.
The proposed pool storage densification will equip the pool with 1276 storage locations.
Table 1.1 indicates that the proposed reracking of the pool will provide adequate storage with full core discharge capability well into the next century (c.
2008).
Table 1.1 is based on a conservatively estimated 18 month fuel cycle.
Current trends 1
towards extdnded burn-up and higher enrichment would further extend the tiime span of on-site storage.
The racks proposed herein are of free
- standing, self supporting variety.
The principal materials of construction are ASTM SA240, Type 304 for the storage locations and "BORAFLEX", a O
patented product of BISCO (a division of Brand, Inc.).
O 1-1
~ _ _ _. _
The specifications for design, construction and quality
. assurance for the high-density spent fuel storage racks were prepared by South Carolina Electric. and Gas Company.
The mechanical design, seismic / structural canalysis, thermal-hydraulic
- analysis, and other related calculations as well as the fabrication of the hardware are performed by Joseph Oat I
Corporation.
Southern Science, a division of Black & Veatch, is serving as a consultant to Joseph Oat Corporation in the area of criticality analysis.
Gilbert Commonwealth of
- Reading, Pennsylvania is pavidi.ng.the analytical appport in the areas of pool slab / wall qualification, and radiological considerations.
The analyses performed by Joseph Oat Corporation in conjunction with Gilbert Commonwealth and Black and Veatch demonstrate that acceptable margins of safety exist with respect to appropriate NRC. and ASME acceptance criteria. - A cost-benefit comparison of several potential spent fuel disposition. alternatives indicate that reracking of the V.C.
Summer Pool is the lowest risk and most cost-effective alternative and that neither the reracking i
.O o9er eiem eer twe i=cre ea e=-
tee eer se ef irr dietee 1
material pose an increased hazard to the Plant Staff or the public.
The following sections provide a synopsis of the design, 1
fabrication, neutronics
- analysis, thermal / hydraulic
- analysis, structural analysis, accident analysis, environmental analysis l
and cost-benefit appraisal of the High Density Spent Fuel Racks.
(
In particular, the integrity of the rack structure under the I
specified combinations of inertial, seismic, and mechanical loads and thermal gradient per NUREG-0800 is demonstrated.
Also included-are concise descriptions of the rack In-service Surveillance Program and the Joseph Oat Corporation Quality l
Assurance Program.
The Joseph Oat Corporation Quality Assurance Program has been reviewed and found acceptable for engineered fabrication of ASME Section III Class 1, 2,
3 and MC components by both ASME and NRC.
1-2
Table 1.1 V.C. Summer Nuclear Station Fuel Assembly Discharge (Tentative Schedule)
Remaining Remaining Storage Storage Total Discharged Capability with-Capability Refueling Discharge Assemblies in Spent Fuel out Proposed with Proposed Date Assembly Pool Following Refueling Expansion Expansion Fall 1984 44 44 638 1232 i
~
Fall 1985 72 116 566 1160 Spring 1987 68 184 498 1092 Fall 1988 72 256 426 1020 Spring 1990 68 324 358 952 Fall 1991 72 396 286 880 i
Spring 1993 68 464 218 812 Fall 1994 72 536 146*
740 Spring 1996 68 604 78 672 Fall 1997 72 676 6**
600 Spring 1999 68 744 532 Fall 2000 72 816 460 Spring 2002 68 884 392 Fall 2003 72 956 320 i
Spring 2005 68 1024 252 Fall 2006 72 1096 180 l
112*
Spring 2008 68 1164 l
Fall 2009 72 1236 40**
Full core discharge capability lost (157 assemblies)
- Normal discharge capability lost (= 72 assemblies) t 1-3
o 2.
GENERAL ARRANGEMENT b
The high density spent fuel racks consist of individual cells with 8.85" x 8.85" (nominal) square cross section, each of which accommodates a single PWR (Westinghouse) fuel assembly.
The cells are arranged in modules of varying size. A total of 1276 cells are arranged in 11 distinct modules in 3
regions.
Region 1
is designated for storage of freshly discharged fuel assemblies with enrichments up to 4.3 weight percent U-235.
The cells in Region 2 are reserved for accommodating fuel assemblies with initial enrichments of 4.3 weight percent U-235 and a minimum burnup of 20,000 MWD /MTU.
The remaining cells, i.e.
Region 3 cells, are capable of accommodating fuel assemblies with initial enrichments of 4.3 weight percent U-235 and a
minimum burnup of 42,000 MWD /MTU.
Fig. 2.1 shows the arrangement of the rack modules in the i
V.C. Surtmer pool in the 3 regions described above.
A 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 2
capabilities.
As shown in Fig.
2.1, there are 11 discrete modules arranged l
l in the fuel pool at 1-7/8" minimum inter-modulo gap.
Table 2.1 gives the relevant design data on each region.
The modules in the three regions are of 4 different types.
Table 2.2 summarizes the l
physical data for each module type.
lo 21 L
The modules a,re not anchored to the pool floor, to each other, O
or to the pool walls.
A minimum gap of 1-7/8" is provided between the modules to ensure that kinematic movements of the modules during the Plant Des'ign Basis Earthquake will not cause inter-module impact. Adequate clearcnce with other pool hardware, eg. light fixtures, etc. is also provided.
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Table 2.1 Design Data 10 Cell Pitch diin 6 B Flux Trap Region (nominal)
Leading Gap (nominal) 2 1
10.4025"
.022 gm/cm 1.1605 1
2 2
10.4025" x 10.1875"
.0015 gm/cm 1.2605 x 1.0455 3
10.116" unpoisoned 1.086 O
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Table 2.2 Module Data Approximate Region Module
' IMeliv.1er Array Weight No.
Type Ouantity module:
Size (lb/ module) 1 A
2 121 llxll 36300 2
B 1
99 llx9 28300 3
C 5
121 llxil 25500 3
D 3
110 llx10 23100 O
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FI G. 2 1 MODULE LAYOUT
3.
RACK CONSTRUCTION
- 3. 1 Fabrication Details:
3.1.1 Recions 1 and 2:
i The rack module is fabraicated from ASTM 240-304 austenitic l
stainless steel sheet and plate material, and ASTM A182-Type F304 forging material.
The weld filler material utilized in body welds is ASME SFA-5.9 Type 308. Boraflex, a patented brand name product of BISCO
- serves as the neutron absorber material. The detailed radiological properties of Boraflex may be found in Section 4.
The l
experience list of Boraflex is given in Table 3.1.
4 A typical module contains storage' cells which have an 8.85" nominal cross sectional opening.
This dimension ensures that fuel assemblies with maximum expected axial bow can be inserted and
)
removed from the storage cells without any damage to the fuel assemblies or the rack modules.
j Figs. 3.l A arid 3. lB show a horizontal cross section of a 3x3 array.
The cells provide a smooth and continuous surface for lateral contact with the fuel assembly.
The anatomy of the rack modules is best exposed by describing the basic building blocks of l
the design, namely (a)
Internal square tube (b)
Neutron Absorber Plate (Boraflex) envelope angular i
elements (c).' Angular structural element I
(d)
Base plate (e)
Support assembly (f)
Top lead-in Os.
- BISCO, a Division of Brand, Inc.,
1420 Renaissance Drive, Park Ridge, Illinois l
3-1 l
l
)
Table 3.1 BORAFLEX EXPERIENCE FOR HIGH DENSITY RACKS Plant NRC Licensing Site Type Docket #
Status Point Beach -l & 2 PWR 50-226 & 301 Issued Nine Mile Point - 1 BWR 50-220 Issued Oconee 1 & 2 FIIR
'50-%9 & 270 Issued Prairie Island 1& 2 PWR 50-282 & 306 Issued Calvert Cliffs - 2 PWR 50-318 Issued
- Quad Cities - 1 & 2 BWR 50-254 & 265 Issued Midland - 1 & 2 PWR 50-329 & 330 Pending Watts Bar 1 & 2 PWR 50-390 & 391 Pending Waterford - 3 PWR 50-382 Pending-
- Fermi - 2 BWR 50-341 Issued H.B.
Robinson - 2 PWR 50-261 Pending River Bend - 1 BWR 50-458 Pending
- Rancho Seco - 1 PWR 50-312 Pending Nine Mile Point - 2 BWR 50-410 To be applied for Shearon Harris - 1 PWR 50-400 To be applied for Millstone - 3 PWR 50-423 To be applied for
- Grand Gulf
-1 BWR 50-416 Pending
- Oyster Creek BWR 50-219 Pending
- Joseph Oat Corporation fabricated racks.
3-2
..-,._...,,,_,,.m.
.y.,,
_,._7
(3 V
(a) Internal' square tube:
This element provides the lateral bearing surface to the fuel assembly.
It is fabricated by joining two formed channels using a
controlled seam welding operation.
The weld penetration in the seam welded zone is required to be 90% minimum.
This element is 8.85" square (nominal) cross section x 169" long.
(b) Neutron Absorber Plate (Boraflex) envelope angular elements:
Boraflex surrounds the square tube on all four sides over a length of 138" wi.!oh completely envelopes the active fuel length except the top 3 and bottom 3 inches.
(c) Angular structural elements:
Two angular subelements, illustrated in Fig. 3.2 (a) and (b) comprise the structural support gridwork for the fuel racks.
One set of large and small angular subelements is placed around the square tube with the Neutron Absorber interposed in-between, as shown in the cross section in Fig.
3.3.
The fillet welds indicated in Fig.
3.3 are made while the l
angular subelements exert a contact pressure on the neutron absorber sheets in the welding
- fixture, thereby ensuring a continuous surf ace. contact in a macroscopic sense, between the constituent elements bf the sandwich.
As shown in Fig.
3.4B, bottom spacer sheets (also made from ASTM 240-304 material) i 3-3 l
position the horaflex shee ts in the vertical g
(
direction.
The top of the argular sub-elements is welded to the square tube using a suitable spacer.
In this manner a
composite box assembly is fabricated.
An ar. ray of composite box assemblies welded as indic:ated in Figs. 3.lA and 3.lB form the
" honey-comb" griioneer'k ef cells which harnesses the structural s treangtih of all sheet and plate type members in an edf,fiefte nt manner.
Figure 3.4.A illustrates a ty,gma:ft weiding sequence to obtain a honey comb construtetloc. The array of composite boxes has overall be ndiimp,, torsional and axial rigidities which are an order of magnitude greater than configurations utilizing grid bar type of construction.
(d)
Base Plate:
The base plate is a 5/8" thick plate type member pg V
which has 5" diameter holes concentrically located with respect to the internal square tube.
These I
holes provide the primary path for coolant flow.
Secondary flow paths are available between adjacent cells via the lateral flow holes (1.0" diameter) near the root of the " honey-comb" (Figure 3.4B).
The honey comb is welded to the base plate with 1/8" fillet welds.
(e)
Support Assembly:
7
. Each module has 4 support legs.
One support leg is of fixed height,(Fig.
3.6) the other three are adjustable in length to enable leveling of the rack.
The variable height support assembly consists of a flat-footed spindle which rides into an internally threaded cylindrical member.
The f
cylindrical member is attached to the underside of
('
the base i
3-4
f'()'
pl. ate through double fille.t and partial penetration weld.
The baste of the flat-footed spindle sits on the pool floor. Leveling of the rack modules is accomplished b>y turning the hex sprocket in the spindle using :a loog aun (approximately 16' long) hex head wre ncth..
Fig 3.5 shows a vertical cross section of the addustabIe support assembly.
The supports elevate the module base plate approximately.' 6 1/4" -above pool
- floor, thus creating the water plenum for coolant flow.
The lateral holes in the cylindtical member provide the coolant entry path leading into the bottom of the storage locations.
(f)
Top Lead-In:
Contiguous walls of adjacent cells are connected by a
suitably designed lead-in for fuel assembly v.-
insertion.
These lead-in joints also aid in reducing the lateral deflection of the inner square tube due to the impact of fuel assemblies during g
the ground motion (postulated seismic motion specified in the FSAR). This construction procedure leads to natural venting locations for the inter-cell space where the neutron absorber material is located. The fabrication of the rack modules is performed under a
strict quality assurance system suitable for ASME Section III, i
. Class 1,
2 and 3 manufacturing which has been in place at Joseph Oat Corporation for over 10 years.
O 3-5
O 3.1.2 Region 3:
f G
The rack modules in Reg ion 3 are fabricated from the same material as that used for Regions 1
and 2
- modules, i.e.
ASTM-240-304 austenitic stain'less steel and ASTM A182 Type F-304 forging material.
No neutron.nbsorber material is used.
A typical module also ocetains storage cells which have an 8.85" nominal cross-sectional opening.
Fig. 3.7 shows a horizontal cross-section of a 3x3 array.
The rack construction varies from that for Regions 1 and 2 in as much as the internal square tube and the neutron absorber plate are eliminated.
Hence the basic components of this design is as follows:
(a)
Angular structural element (b)
Baseplate (c)
Support Assembly
/m(j (d)
Top Lead-in 4
In this construction, two angular structural elements form the cell of an 8.85" nominal cross-sectional opening in addition to functioning as part of the structural support gridwork as illustrated in Fig.
3.7.
The fillet welds for these unpoisoned cells are also shown in Fig. 3.7.
l The baseplate and support assemblies are exactly the same as those described for Regions 1 and 2.
Contiguous walls of adjacent cells are also connected by a suitably designed lead-in for fuel assembly insertion.
l l
I 3-6
3.2 CODES, STANDARDS, AND PRA:CTICES FOR THE SPENT FUEL POOL O
(y MODIFICATION The following are the public domain codes, standards, and practices to which the fuel storage racks are
- designed, constructed and assembled, and/or pool structur.e analyzed.
Additional problem-specific references related to detailed analyses are given at the end of each section and at the beginning of the section for Section 4.
I.
Design Codes (a)
AISC Manual of Steel Construction, 8th edition (1980)
(b)
ANSI N210-1976 Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Stations.
(c)
American Society of Mechanical Engineers (ASME), Boiler & Pressure Vessel Code,Section III, 1980 Edition up to and including Winter 1982 addenda. (Subsection NF)
(d)
ASNT-TC-1A June,
- 1980, American Society for i
Nondestructive Testing (Recommended Practice for Personnel Qualifications)
II. Material Codes l
(a)
American Society for Testing and Materials (ASTM) Standards - A240.
(b)
American Society of Mechanical Engineers (ASME), Boiler & Pressure Vessel Code, Section
~
gV II, Parts A and C,
1980 Edition up to and including Summer 1983 addenda.
3-7
III. Welding Codes (a)
ASME Boiler and Pressure Vessel Code, Section IX-Welding and Brazing Qualifications, 1980 Edition up to and including Summer 1983 addenda.
IV. Quality Assurance, Cleanliness, Packaging, Shipping, Receiving, Storage, and Handling Requirements (a)
ANSI 45.2.2, Packaging,
- Shipping, Receiving, Storage and Ilandling of Items for Nuclear Power Plants.
(b)
ANSI 45.2.1, Cleaning of Fluid Systems and Associated Components During Construction Phase of Nuclear Power Plants.
(c)
ASME Boiler and Pressure Vessel,Section V,
Non-destructive Examination, 1980
- Edition, including Summer 1983.
(d)
ANSI-N16.1-75 Nuclear Criticality Safety Operations with fissionable materials outside reactors.
(e)
ANSI-N16.9-75 Validation of Calculation Methods for Nuclear Criticality Safety.
(f)
ANSI-N45.2.11-1914 Quality Assurance Requirements for the Design of Nuclear Power Plants.
V.
Other References (a)
NRC Regulatory Guides, division 1,
regulatory guides 1.13, 1.29, 1.31, 1.61, 1.71, 1.85, 1.92, and 1.124 (revisions as applicable).
3-8
,--i
---,--v.--
,-,,.--.m..e,.e..-y
&w
-m.--e.r
.+>--a
,-e---
m ------ --..
-., ~
._ =.
i; (b)
General Design Criteria for Nuclear Power Plants, Code of Federal Regulations, Title 10, 9
Part 50, Appendix A (GDC Nos.
1, 2, 61, 62, and 63).
(c)
NUREG-08GO, Standard Review Plan (1981).
(d)
"NRC Position for Review and Acceptance of Spent Fuel Storage and Handling Applications,"
dated April 14, 1978, and the modifications to this document of January 18, 1979.
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DEPEND ON REGION FIGURE 3 2 (b) LARGE ANGULAR SUBELEMENT O
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O aesion 3 3 -18
s
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,^ 4.
NUCLEAR CRITICALITY ANALYSIS
> e s
~
"4.1 Design Bases s
The high density spent fuel storage racks for the V.
C.
Summeb5 Station 1 are designed to assure that a k gg equal to or e
less than 0.95 is maintained with the racks fully loaded with fuel of the highest; anticipated reactivity in each of three regions and flooded [vith unborated water at a temperature corre-spo'nding to the. highest reactivity.
The maximum calculated re-
~
activity i n c l u d,e s. a margin for uncertainty in reactivity cal-culations and in ' mechanical tolerances, statistically combined,
'such that the; true k gg will be equal to or less than 0.95 with a e
95% probability at a 95% confidence level.
Applicable codes, standards and regulations, or pertinent sections thereof, include the following.
Prevention of Criticality
. 1.
e General Design Criterion 62 in Fuel Storage and Handling.
s e
USNRC Standard Review Plan, NUREG-0800, Section 9.1.2,
^ Spent Fuel Storage.
e ' 'USNRC Regulatory Guide 1.13, Spent Fuel Storage Facility Design Basis (proposed), December 1981.
e USNRC letter of April 14,
- 1978, to all Power Reactor OT Position for Review and Acceptance of Licensees-Spent Fuel Storage and Handling Applications, including
, modification letter dated January 18, 1979.
'O'SNRC R,e'gulatory Guide 3.41, Validation of Calculational e
Method for Nuclear Criticality Safety (and related ANSI N16.9-1975).
ANSI / ASS-57.2-1983, Design Requirements for Light Water
~
o Reactor ^ Spent Fuel Storage Facilities at Nuclear Power Plants.
m tut 4-1
~.
e
,1*
g p,
m
,...-m
pd e
ANSI N210-1976, Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear Power Plants.
e ANSI N18.2-1973, Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants.
To assure the true reactivity will always be less than the calculated reactivity, the following conservative assumptions were made.
o Moderator is pure, unborated water at a temperature cor-responding to the highest reactivity.
e Lattice of storage racks is assumed infinite in all di-rections; i.e.,
no credit is taken for axial or radial neutron leakage (except in the assessment of certain abnormal / accident conditions where leakage is inherent).
e Neutron absorption in minor structural members is neglected; i.e.,
spacer grids are replaced by water.
r
The design basis fuel assembly is a 17 x 17 array of fuel at a maximum initial rods (Westinghouse design) containing UO2 enrichment of 4.3% U-235 by weight, corresponding to 54.30 grams U-235 per axial centimeter of fuel assembly.
Three independent regions are provided in the spent fuel storage
- pool, with separate criteria defining the highest anticipated reactivity in each of the three regions as follows.
e Region 1 is designed to accommodate new unirradiated fuel with a maximum enrichment of 4.3 wt.% U-235, or spent fuel regardless of the discharge fuel burnup.
e Region I2 is designed to accommodate spent fuel of 4.3 wt.% U-235 initial enrichment, which has accumulated a minimum burnup of 20,000 Mwd /mtU.
Region 2 will also safely accept fuel of lower discharge fuel burnup pro-vided the initial enrichment is correspondingly lower.
e Region 3 is designed to accommodate fuel of 4.3 wt.%
U-235 initial enrichment which has accumulated a minimum
('
burnup of 42,000 Mwd /mtU.
Region 3 will also safely
's accept fuel of lower discharge fuel burnup provided the initial enrichment is correspondingly lower.
4-2
4.2 Summary of Criticality Analyses 4.2.1 Normal Operating Conditions The criticality analyses of each of the three separate regions of the 9 pent fuel storage pool described above are sum-marized in Table 4.1 for the anticipated normal storage condi-tions.
The calculated maximum reactivity in Regions 2 and 3,
which includes a 0.015 Ak extra additive allowance for uncer-tainty in burnup calculations, provides an additional margin of 1% ak or more below the limiting value of 0.95.
As cooling time increases in long-term storage, decay of Pu-241 results in sig-nificant decrease in reactivity, which will provide an increasing subcriticality margin and tend to further compensate for any uncertainty in depletion calculations.
Spacing between the three different rack modules is sufficient to preclude adverse inter-p action between modules.
Regions 2 and 3 can accommodate fuel of lower discharge fuel burnup provided the initial enrichment is correspondingly lower.
Figures 4.1 and 4.2 illustrate, as a function of the initial fuel enrienment, the minimum acceptable burnup which yields the maximum reactivity given in Table 4.1 for Regions 2 and 3,
respectively.
These curves are intended to be incorpo-rated in the Technical Specifications supplemented with appro-priate administrative procedures to assure verified burnup as specified in draf t Regulatory Guide 1.13, Revision 2.
4.2.2 Abnormal and Accident Conditions Although credit for soluble poison normally present in the pool water is permitted under abnormal / accident conditions, most abnormal or accident conditions will not result in exceeding the limiting reactivity of 0.95 even in the absence of soluble poi-son.
One accident condition that could potentially exceed the 4-3
(~)
Table 4.1 Summary of Criticality Safety Analyses LJ Region 1 Region 2 Region 3 Minimum burnup with 4.3%
0 20 Mwd /kgU 42 Mwd /kgU initial fuel enrichment Temperature assumed 40*F 68'F 150*F for analysis Reference k, 0.9323 0.9024 0.9168 (AMPX-KENO)
Calculational bias 0
0 0
Uncertainties Bias
- 0.0030
- 0.0030 i0.0030 Calculational
- 0.0038
- 0.0039
- 0.0043 B-10 concentration T0.0017 T0.0019 NA Boraflex thickness 70.0032 T0.0121 NA Boraflex width T0.0004 70.0005 NA
()
Inner box dimension 0.0008 70.0013 T0.0033 Flux-trap water gap 70.0059 T0.0071 70.0031 SS tolerance
- 0.0006
- 0.0006 10.0040 Fuel enrichment
- 0.0020
- 0.0020 0.0020 Fuel density t0.0023 t0.0023
- 0.0023 Statistical
- 0.0090
- 0.0154
- 0.008'6 combination Total 0.9323 0.9024 0.9168
- 0.0090 A 0.0154
- 0.0086 Eccentric assembly negative
+0.0032
+0.0012 l
position l
Extra allowance for NA 0.015 0.015 burnup uncertainty l
Maximum reactivity 0.941 0.936 0.942 l
l 4-4
[
c limiting reactivity is the inadvertent loading of a new fuel assembly (4.3% enrichment) into Region 2 or Region 3 storage cells, with the simultaneous occurrence of a loss of all soluble l
poison.
Administrative procedures will be necessary to preclude the possibility of simultaneous occurrence of these ' two inde-pendent accident conditions.
Effects on reactivity of other abnormal and accident condi-tions evaluated are summarized in Table 4.2.
Table 4. 2 Reactivity Effects of Abnormal and Accident Conditions Accident / Abnormal Conditions Reactivity Effect Temperature increase Negative in Regions 1 and 2; positive in Region 3 Void (boiling @ 24 8
- F)
Negative in all regions Assembly outside rack Negligible Assembly lying on top of rack Negligible Lateral rack module movement Negligible of the abnormal / accident conditions cited above, a positive reactivity effect results only from an increase in temperature above the nominal maximum pool temperature of 150*F assumed for criticality evaluation of the Region 3 storage rack.
Tempe r-atures above 1R0*F are considered accident conditions, in which case the soluble poison would maintain reactivity at an accepta-bly low value.
Nevertheless, in the absence of soluble poison, the increase in rea.ctivity is calculated to be only
- 0. 01 A k at 248'F (approximate boiling temperature of the bulk coolant at the submerged depth of the fuel racks).
Thus, even in the simul-taneous occurrence of two independent accident conditions, the 4-5
O(_/
maximum reactivity does not exceed the limiting value of 0.95.
At 248*F, voids resulting from boiling have a negative reactivity effect.
O 1
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INITI A L EN RIC HM E NT, wt.% U-235 a
Fig. 4.1 Limiting discharge fuel burnup for Region 2 storage rack for fuel of various initial enrichments.
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INITI AL ENRI C H M ENT, wt. % U-235 kJ Fig. 4.2 Limiting discharge fuel burnup for Region 3 storage rack for fuel of various initial enrichments.
4-8
x.)
4.3 Reference Fuel Storage Cell 4.3.1 Reference Fuel Assembly The reference design basis fuel assembly, illustrated in Fig.
4.3, is a 17 x 17 array of fuel rods with 21 rods replaced by 20 control rod guide tubes and one instrument thimble.
Table 4.3 summarizes the fuel assembly design specifications and the expected range of significant parameters.
Table 4.3 Fuel Assennly Design Specifications Fuel Rod Data Outside dimension, in.
0.374 Cladding thickness, in.
0.0225 Cladding material Zr-4
/"N Pellet diameter, in.
0.3225
(_),
Dishing factor 0.012 UO2 density, % T. D.
95
- 2 3
UO2 stack density, g/cm 10.296
- 0.2172 Enrichment, wt. % U-235 4.3
- 0.05 Fuel Assembly Data Number of fuel rods 264 (17 x 17 array)
Fuel rod pitch, in.
0.496 Control rod guide tube Number 20 0.D.,
i n, 0.482 Thickness, in.
0.016 Material Zr-4 Instrument thimble Number 1
0.D.,
in.
0.482 Thickness, in.
0.016 Material Zr-4 O
u-23s 1o a1=2 g/ axial cm of assembly 54.30
- 1.78 4-9
4.3.2 Region 1 Storage Rack The nominal spent fuel storage cell used for the criticality analyses of Region 1 storage cells is shown in Fig.
4.3..
The rack is composed of Boraflex absorber material sandwiched between a 0.049-in.
inner stainless-steel box and a
0.065-in.
outer stainless-steel box.
The fuel assemblies are centrally located in each storage cell on a nominal lattice spacing of 10.4025
- 0.0625 in.
Stainless-steel tabs connect one storage cell box to another in a rigid structure and define an outer water space j
between boxes.
' Itis outer water space constitutes a flux-trap i
between the two Boraflex absorber plates that are essentially opaque (black) to thermal neutrons.
The Boraflex absorber has a thickness of 0.082
- 0.007 in, and a nominal B-10 areal density 2
of 0.0265 grams per cm,
4.3.3 Region 2 Storage Cells Figure 4.4 illustrates the storage cell design used for the Region 2 storage cells.
In Region 2,
the rectangular storage cells are located on a lattice spacing of 10.4025
- 0.0625 in. in one direction and 10.1875
- 0.0625 in. in the other direction.
Boraflex absorber material of 0.032-in.
thickness, sandwiched between 0.049-in. and 0.065-in. stainless-steel plates, consti-tutes the walls of each storage cell.
The outer water-gap flux-I trap is 1.0455
- 0.0625 in. in one dimension and 1.2605
- 0.0625 in. in the other direction, as indicated in Fig.
4.4.
For fuel of 4.3 wt.% U-2,35 initial enrichment burned to 20 Mwd /kgU, the 2
nominal design B-10 areal density is 0.002 g/cm,
4.3.4 Region 3 Storage Cells l
l l
Region 3 storage cells, designed for fuel of 4.3 wt.% U-235 initial enrichment burned to 42 Mwd /kgU, is unpoisoned, other than the 0.090-in. thick stainless-steel plates forming the walls 4-10
l.
j' of the storage cell.
These cells, shown in Fig. 4.5, are loccted l
on a lattice spacing of 10.116
- 0.032 in. defining a 1.086 i 0.032-in. water gap between the steel walls.
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4-12 1
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i 1
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9.142 d
.0625
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h 10.1875 2.0625
?'
(PlTCH)
( N OT TO SCALE)
,U Fig. 4.4 Configuration of Region 2 spent fuel storage cell.
4-13
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l (NOT TO SCALE) r' k.)N Fig. 4.5 Configuration of Region 3 spent fuel storage cell.
l 4-14 L
()
4.4 Analytical Methodology 4.4.1 Reference Analytical Method and Bias The reference nuclear criticality analyses of the high den-sity spent fuel storage rack were performed with the AMPX -KENO 2 I
computer package, using the 123-group GAM-THERMOS cross-section set and the NITAWL subroutine for U-238 resonance shielding effects (Nordheim integral treatment).
AMPX-KENO has been ex-tensively benchmarked against a number of critical experiments (e.g., Refs. 3, 4, 5, and 6), including those4,6 most representa-tive of spent fuel storage racks.
In the geometric model used in KENO, each fuel rod and cladding were described explicitly.
For two-dimensional X-Y analysis, a zero current (white albedo) boundary condition was applied in the axial direction and at the centerline through the
(
outer water space (flux-trap) on all four sides of the cell, effectively creating an infinite array of storage cells.
Results of the benchmark calculationr 6 on a series of criti-cal experiments indicate a calculational bias of 0, with an un-certainty of
- 0.003 (95%
probability at a
95%
confidence level).
In addition, a small correction in the calculational bias might be necessary to account for the internal water-gap thickness (0.418 in.) between rack walls and the fuel assemblies in the V.
C. Summer spent fuel rack compared to the corresponding I
thickness (0.644 in.)
in the benchmark critical experiments.
Based upon the ' correlation developed in Ref.
6, the correction for water-gap thickness in the V.
C.
Summer spent fuel storage rack indicates a small overprediction of ~0.002 Ak.
For conser-vatism, the overprediction is neglected and the net calculational l
bias is taken as 0.000
- 0.003, including the effect of the water-gap thickness.
- O l
4-15 o
qV 4.4.2 Manufacturing Tolerances and Uncertainties For investigation of small reactivity effects due to manu-facturing tolerances, the CASMO computer code 7 was used to calcu-late small incremental reactivity changes that would otherwise be lost in the normal statistical variation associated with Monte Carlo techniques (i.e., KENO).
CASMO 4.s a two-dimensional trans-port-theory code, based on capture probabilities, th at allows an explicit description of each fuel pin in an assembly, as well as an approximate description of the storage cell geometry.
Geomet-rical approximations necessary in CASMO include the following.
e The outer stainless-steel plates and the connecting tabs were homogenized with the water in the flux-trap region.
e The Boraflex absorber plate was necessarily described as 8.948 in, wide, rather than the actual 8.45-in width.
x e
In Region 2,
the rectangular cell geometry was V
represented as a square of equal area.
Despite these approximations, reactivities calculated by CASMO were within 0.0014 ak of the reference AMPX-KENO calcu-lation for Regions 1 and 2.
For Region 3, the CASMO-calculated reactivity was 0.0057 ak above the corresponding AMPX-KENO calcu-lation.
Thus, the differential reactivity calculated by CASMO should provide reliable estimates of the uncertainties associated with manufacturing tolerances.
4.4.3 Fuel Burnup Calculations Fuel burnup calculations were performed primarily by the CASMO code.
However, to enhance the credibility of the burnup calculations (in lieu of critical experiments), the CASMO results 8
were independently confirmed by calculations with the EPRI-CELL and NULIF9 codes.
Figure 4.6 compares results of the three O
V independent methods of burnup analysis under reactor operating 4-16
h"%
99 conditions.
Agreement between CASMO and EPRI-CELL is very good (within 0.002 Ak at 20 Mwd /kgU and 42 Mwd /kgU), although re-activities calculated by NULIF were somewhat lower, probably due to differences in treatment of temperature effects and resonance capture.
For additional information, burnup-dependent reactivi-ties extracted from a CHEETAH-P calculation 10 for a comparable reactor system with essentially the sa:ae core nuclear properties are also shown on Fig.
4.6.
In addition to good agreement in depletion calculations, of equal importance for storage rack criticality analyses are re-activity comparisons under conditions more representative of fuel to be stored in the racks (cold, xenon-free).
Table 4.4 compares the cold, xenon-free reactivities calculated, at 68'F, by the three codes at the two reference fuel burnups.
V Table 4.4 Ccmparison of Cold, Clean Reactivities Calculated at 20 Mwd /kgU and 42 Mwd /kgU Burnup k, Xe-Free @ 68'F Calculational Method
@ 20 Mwd /kgU e 42 Mwd /kgU CASMO 1.2597 1.0743
[
EPRI-CELL
- 1.2612 1.0747*
NULIF 1.2625 1.0789 r
l 1
- At maximum reactivity during long-term decay (see Section l
4.4.4).
1 OV l
l 4-17 i
l
g()
No definitive method exists for determining the uncertainty in burnup-dependent reactivity calculations.
All three codes discussed above have been used with good accuracy to follow reactivity changes in operating reactors.
CASMO has been exten-sively benchmarkedll against cold, clean critical experiments (including plutonium-bearing fuel),
Monte Carlo calculations, reactor operations, and heavy-element concentrations in ll of eleven criti-irradiated fuel.
In particular, the analyses cal experiments with plutonium-bearing fuel showed an average k gg of 1.002
- 0.011 (95%/95%), showing adequate treatment of e
the plutonium nuclides.
With fuel of 42 Mwd /kgU burnup, the total reactivity worth of the fission products (excluding Xe and Sm) is estimated to be 12% ak.
Assuming the fission product reactivity worth is accurate to
- S%,
the uncertainty would be *0.006 ok.
Statisti-cally combined with the uncertainty derived from the analysis of (m
\\
Pu-bearing critical experiments, the total uncertainty becomes
- 0.0125.
However, due to the possible existence of a small positive reactivity increment from the axial distribution in burnup (see Section 4.4.5 below),
the uncertainty has been increased to 0.015 ak in Region 3 (42 Mwd /kg fuel) and treated as an additive term for the present evaluation, rather than being combined statistically with other uncertainties.
This is believed to be a conservative estimate, particularly in view of the substantial negative reactivity contribution from aged fuel as discussed in Section 4.4.4.
Although the uncertainty at lower l-burnup (e.g.,
20 Mwd /kgU) would normally be expected to be less, 1
this refinement has not been made and the same uncertainty (0.015 ak) has been assigned to Region 2.
4.4.4 Long-Term Decay q
Since the fuel racks in Regions 2 and 3 are intended to con-d tain spent fuel for long periods of time, calculations were made 4-18
~ _, _ _ _
7s using EPRI-CELL (which incorporates the CINDER code
- as a sub-routine) to follow the long-term changes in reactivity of spent fuel over a 30-year period.
Figures 4.7 and 4.8 show these changes for fuel burned to 20 Mwd /kg and 42 Mwd /kg, respec-tively.
Early in the decay period, xenon grows in and sub-sequently decays, with the reactivity (k.) reaching a maximum at about 10 days after reactor shutdown.
( This maximum value is listed in Table 4.4 above.)
For longer storage periods, the decay of Pu-241 (13-year half-life) again reduces reactivity, as shown in Figs. 4.7 and 4.8.
The reference design criticality safety calculations do not take credit for this long-term reduc-tion in reactivity, although this effect woMld afford an increasing subcriticality margin in Regions 2 and 3 of the spent fuel storage pool.
For illustrative purposes, Table 4.5 lists the long-term reduction in reactivity below the maximum used for the reference criticality safety evaluation.
pV Table 4.5 Long-Term Changes in Reactivity ak from Maximum Reactivity Storage Time, years
@ 20 Mwd /kg
@ 42 Mwd /kg 0.5
-0.0016
-0.005 1.0
-0.0032
-0.011 4.0
-0.0098
-0.034 10.0
-0.0192
-0.060 20.0
-0.0283
-0.086 30.0
-0.0335
-0.101
- CINDER tracks the decay and burnup dependence of 179 distinct N(d fission products.
4-19
To realize the advantage of the long-term reduction in reactivity (especially in Region 3),
it will be necessary to develop a specified loading pattern for spent fuel and rely upon administrative procedures to assure that the more recently-dis-charged assemblies are separated from each other and interspersed
' among lean reactive (more aged) spent fuel assemblies.
4.4.5 Effect of Axial Burnup Distribution Initially, fuel loaded into the reactor will burn with a slightly. skewed cosine power distribution.
As burnun progresses, 1
the burnup distribution will tend to flatten, becoming more highly burned in the central regions than in the upper and lower ends.
.This effect may be clearly seen in the curves compiled in Ref. 12.
At high burnup, the more reactive fuel (less than average burned) occurs near the ends of the fuel assembly where leakage is expected to reduce the reactivity worth.
Con-sequently, it is anticipated that distributed burnup fuel assem-blies would exhibit a slightly lower reactivity than that calcu-lated for the average burnup.
Based on calculated burnup dis-tributions at an average of ~ 20 Mwd /kgU, the distributed burnup case shows a lower reactivity than the uniform average case, as l
expected, in one-dimensional axial calculations.
As burnup pro-I gresses, the distribution, to some extent, tends to be self-regu-1 l
lating as controlled by the axial power distribution, precluding I
the existence of large regions of significantly lower burnup (unless locally perturbed by the long-term insertion of control rods).
Although the axial burnup distribution is normally expected to have slightly less raactivity than the uniform average burnup l3 in a fuel assembly case, the experimentally determined burnup removed from the H.
B.
Robinson reactor (similar to the V.
C.
Summer Station) was used as a basis for evaluating the potential reactivity effect of the axial burnup distribution in spent 4-20
(D V
fuel.
Figure 4.9 shows the reported gamma scan (proportional to burnup) together with burnups measured on selected pellets based upon Nd-148 mass-spectrometric determinations.
This distribution indicates that the upper 1.6 feet of the assembly and the lower 1.2 feet have burnups less than the average.
Based upon the distribution in Fig.
4.9, one-dimensional diffusion theory calculations
( CASMO-derived,
homogenized diffusion constants for fuel of different burnups in Region 3 cells),
the distributed burnup case showed a slightly higher reactivity
(+0.003 ak) than the reference reactivity for the averaged burnup.
Although this may be unique to the particular Robinson fuel
- assembly, or to experimental uncertainty in determining the actual fuel burnup, the results suggest there may be a possibility of a small positive reactivity effect due to the axial distribution in burnup.
Because of the self-regulating nature of the power distribution during in-core operation, any positive reactivity ef fect from distributed burnup cannot be very large.
However, the potential for a positive reactivity effect was considered in establishing an overall reactivity uncertainty in burnup calculations, as discussed in Section 4.4.3 above.
- Reference 14 suggests that the average assembly power may be ~26 Mwd /kgU rather than the 27.4 Mwd /kgU implied in Fig.4.9 which would eliminate the apparent excess reactivity attributed to the distributed burnup effect.
4-21
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10 YEARS A FTER R E M OVAL 'F RO M REACTOR Fig. 4.7 Long-term change in infinite multiplication factor (km) of 4.3% enriched fuel burned to 20 Mwd /kgu.
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10 YEARS AFTER R EM OVAL F ROM REACTOR Fig. 4.8 Long-term chanije in infinite multiplication factor (k o) for 4.31, enriched fuel o
burned to 42 Mwd /kgu.
C b
C)
O O
DISTANCE FROM BOTTOM OF ROD (f t )
O I
2 3
4 5
6 7
8 9
10 11 12 13 l
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i DISTANCE FROM BOTTOM OF ROD (m)
Fig. 4.9 Relative axial burnup distribution from NUREG/CR-0722.
f l
o k'
4.5 Reference Subcriticality and Mechanical Tolerance variations 4.5.1 Nominal Design Cases Under normal conditions, with nominal dimensions, the calcu-lated k, for 4.3% enriched fuel in Region 1 is 0.9323
- 0.0021 (lo with 300 generations of 500 neutrons each) for the nominal 15 case.
For a one-sided tolerance factor of 1.799, corresponding to 95% probability at a 95% confidence limit with 300 genera-tions, the maximum deviation of k,,
is *0.0038.
The corresponding k,, from the CASMO calculation is 0.9337, which, despite the necessary geometrical approximations in
- CASMO, provides addi-tional confidence in the validity of the reference AMPX-KENO calculation.
In Region 2,
CASMO calculations, with fuel burned to 20
('
Mwd /kgU in the reference design storage cell at 68'F, yielded a
(
k,,
of 0.9010.
Iterative CASMO calculations with fresh fuel of varying enrichments resulted in an enrichment of 2.30 wt.% U-235 yielding the same k,,
value.
AMPX-KENO calculations were then made on fresh fuel of 2.30% enrichment (to compensate for the geometric approximations in CASMO), yielding a k,
of 0.9024
- 0.0039 (55% probability at a 95% confidence level).
Subse-l quently, iterative burnup and storage cell calculations were made f
with CASMO for fuel of varying initial enrichments (3.6%, 3.0%,
and 2.5%), in each case searching for the burnup which gave the same k,,
as the reference fuel at 20 Mwd /kgU.
These converged burnup values aie those shown in Fig.
4.1.
At the design basis burnup (20 Mwd /kgU), the sensitivity to burnup is calculated to i
be 0.0066 Ak per Mwd /kgU.
{
A similar procedure was used for the Region 3 storage cells at a reference temperature of 150* F, giving a k,,
from the CASMO
(]m calculation of 0.9225.
Iterative CASMO calculations determined an equivalent enrichment of 1.42% U-235 for this same value of l
4-26 l
/7 reactivity.
This' enrichment, in an AMPX-KENO calculation, gave n t*1 k,, of 0.9168 t 0.0043 (95% probability at a 95% confidence level) as the reference design reactivity for the Region 3 storage cells.
Iterative CASMO calculations, as described above (at initial ^ enrichments of 3.6%, 3.0%, 2.5%, and 1.8% E) determined the-burnup values presented in Fig. 4.2 for various initial en-richments.
At the design basis burnup of 42 Mwd /kgU, the calcu-lated sensitivity to burnup is 0.0074 Ak per Mwd /kgU.
4.5.2 Boron Loading Variation The Boraflex absorber plates used in Region 1 storage cells are nominally 0.082 in.
thick, with a B-10 areal density of 0.0265 g/cm2 Independent manufacturing tolerance limits are
- 0.007 in.'in thickness and *0.00245 g/cm2 in B-10 content.
This assures
- that, at any point where the minimum boron loading 2
(0.02405 grams B-10/cm ) and minimum Boraflex thickness (0.075 fl in.) may coincide, the boron areal density will not be less than 0.022 grams B-10/cm2 CASMO calculations indicate that these tolerance limits result in an incremental reactivity change (un-certainty) of *0.0017 Ak for boron content and
- 0.0032 for Bora-flex thickness variations.
In Region 2,,
the Boraflex absorber plates are nominally 0.032 in. thick with a B-10 concentration corresponding to an areal density of 0.0020 4%.
For an independent thickness tol-erance of *0.007 in., the minimum B-10 areal density at any point where the minimum thickness (0.025 in.) and minimum concentration g/cm ),'might coincide, the areal density will not be 2
(0.00192 less than 0.0015 grams-B-10/cm2 CASMO calculations show an incremental reactivity effect (uncertainty) of *0.0019 ak for B-10 concentration and 0.0121 ak for Boraflex thickness.
Boraflex poison sheets are not used in the Region 3 storage cells.
4-27
4.5.3 Storage Cell Lattice Pitch variations he design storage cell lattice spacing between fuel assem-blies ranges from 10.1875 in Region 3 to 10.4025 in Region 1.
An increase in storage cell lattice spacing may or may not reduce reactivity depending upon other dimensional changes that may be associated with the increase in lattice spacing.
Decreasing lattice spacing by decreasing the outer (flux-trap) water thick-
~
ness always increases reactivity, although decreasing the inner water thickness (between the fuel and the inner stainless-steel box) may result in a small increase or decrease in reactivity.
We reactivity ef fect of the outer (flux-trap) water thickness, however, is usually more significant.
Both of these effects have been evaluated for the independent design tolerances in each of the three regions.
4.5.3.1 Inner Water Wickness variations The inner stainless-steel box dimension, 8.850 i 0.032 in.
for all three regions, defines the inner water thickness between the fuel and the inside box.
For the stated tolerance limit, the calculated uncertainty in reactivity is *0.0008 Ak in Region 1,
- 0.0013 Ak in Region 2, and *0.0033 Ak in Region 3.
In Region 1, l
k, increases as the inner stainless-steel box dimension (and de-
'rivative lattice spacing) increases, while in Regions 2 and 3,
increasing the box dimension reduces reactivity.
4.5.3.2 Outer ( Flux-Trap) Water Wickness Variation In Region 1, the design outer (flux-trap) water thickness is l
1.1605 0.0625 in.,
which results in an uncertainty of
- 0.0059 Ak due to the tolerance in flux-trap water thickness, j
assuming the water thickness is simultaneously reduced on all l
four sides.
Region 2 is rectangular with a flux-trap water 3
thickness of 1.2605
- 0.0625 in.
in one direction and 1.0455
- s 4-28
w _
(3 0.06'25 in. ~.in the other direction.
Assuming the thickness
~
y) tolerance 'is, simultaneously applied on all four sides, the un-certainty in re' activity is *0.0071.
For Region 3,
which uses only a sin'gle steel box and no Boraflex, the tolerance in gap waterr, thickness results in a
reactivity uncertainty of
- 0.0031 ak.
Since the fabrication tolerance on each of the four
~
sides is statistically independent, the actual reactivity uncertainties ; would be approximately one-fourth of the values shown, although the more conservative values have been used in the-criticality. evaluation.
4.5.4 Stainless-Steel Thickness Variations The nominal stainless-steel thickness is 0.049 in. for the inner box and 0.065 in. for the outer box.
The maximum positive reactivity effect of the expected stainless-steel thickness tolerance variation (*0.004 in.) was calculated by CASMO to be
.Q
(/
- 0.0006 ak.
-In Region 3, the 0.005-in. tolerance on the single stainless-steel box wall has a larger effect and is calculated to result in a *0.0040 ak uncertainty.
4.5.5 Fuel Enrichment and Density Variation The design maximum enrichment is 4.30 0.05 wt.% U-235.
Calculations of the sensitivity to small enrichment variations by CASMO yielded a coefficient of 0.0041 ak per 0.1 wt.% U-235 in l
Region 1 at the design enrichment.
For the tolerance on U-235 l
enrichment of
.*0.05 in wt.%,
the uncertainty on k,
is i0.0020 A k.
'Ihe same value has been assumed for Regions 2 and 3.
Calculations were made with the UO fuel density in Region 2
t 1 rsnging from a minimum of 93% theoretical density to a maximum i
value of 97% theoretical density.
For the mid-range value (95%
r I
p T.D.) used for the reference design calculations, the uncertainty I
in reactivity is *0.0023 ak over the maximum range of UO2 den-sities expected.
This uncertainty is assumed to apply in all three regione,.
4-29
4.5.6 Boraflex Width Tolerance Variation C/
The reference storage cell design for Regions 1 and 2 ( Figs.
4.1 and 4.2) uses a Boraflex blade width of 8.45
- 0.0625 in.
A positive increment in reactivity occurs for a decrease'in Bora-flex absorber width.
For the width tolerance of -0.0625 in., the maximum calculated reactivity increment is +0.0004 ok in Region 1 and +0.0005 in Region 2.
Increasing the Boraflex width decreases reactivity.
4.5.7 Eccentric Positioning of Fuel Assembly in Storage Rack The fuel assembly is normally located in the center of the storage rack cell with bottom fittings and spacers that mechan-ically limit lateral movement of the fuel assemblies.
Neverthe-less, calculations were made with the fuel assemblies moved into the corner of the storage rack cell (four-assembly cluster at h) closest approach).
These calculations resulted in a negative u./
36 with diffusion coef-reactivity effect for Region 1 using PDQp7 ficients generated by NULIF.
For this region, fuel assembly bowing will produce a small negative reactivity ef fect locally, and the nominal case, with the fuel assemblies positioned in the center of the storage rack cell, yields the maximum reactivity.
Similar calculations were performed for Regions 2 and 3 (four-assembly cluster at closest approach) yielding a positive reac-tivity increment of +0.0032 in Region 2 and +0.0012 in Region 3.
For conservatism, these potential reactivity increments were l
added to the reference calculated k,, although eccentric posi-tioning (if any) would be expected to be randomly distributed l
(statistically) throughout the rack.
l
- his calculational approach was necessary since the reactivity
'(#
T effects are too small to be calculated by KENO, and CASMO geome-try does not permit eccentric positioning of a fuel assembly.
l 4-30
4.6 Abnormal and Accident Conditions 4.6.1 Temperature and Water Density Ef fects The temperature coefficient of reactivity in Region 1 is negative and a temperature of 40*F, corresponding to a water density of 1.0, was arsumed for the reference design.
In Region 2 the maximum reactivity occurred at 68'F, and in Region 3 the temperature coef ficient of reactivity is positive in the temper-ature range to which the racks are normally exposed.
For this reason, a design basis temperature of 150*F was assumed for the criticality evaluation in Region 3.
Temperatures above 150* F are considered accident conditions and credit for the soluble poison actually present would maintain a low reactivity.
Temperature effects on reactivity have been calculated and the results are shown in Table' 4.6.
Introducing voids in the v
water internal to the storage cell (to simulate boiling) de-creased reactivity, as shown in the table.
Voids due to boiling will not occur in the outer (flux-trap) water region.
Table 4.6 Ef fect of Temperature and Void on Calculated Reactivity of Storage Rack Incremental Reactivity Change, ak Case Region 1 Region 2 Region 3 40*F Reference
-0.0002 68*F (20*C)
-0.0021 Reference
-0.0082 104
- F (40 *C)
-0.0055
-0.0002 150*F (65.56*C)
Reference 248*F (120*C)
-0.0205
-0.0049
+0.0108 248'F with 20% void
-0.1111
-0.0564
-0.0015 P
4-31
(q m,/
With soluble poison present, the temperature coefficients of reactivity would be expected to differ from those inferred by the data in Table 4. 6.
However, the reactivities would also be lower at all temperatures, and the data in Table 4.6 is pertinent to the higher-reactivity unborated case.
4.6.2 Dropped Fuel Assembly Accident To investigate the possible reactivity effect of a postu-lated fuel assembly drop accident, calculations were made for unpoisoned assemblies separated only by water.
Figure 4.10 shows the results of these calculations.
From these data, the reactiv-ity (k ) will be less than 0.95 for any water-gap spacing greater than
~6 in. in the absence of any absorber material other than water between assemblies.
For a drop on top of the rack, the fuel assembly will come to rest horizontally on top of the rack with a minimum separation distance of >12 in.
Maximum expected gS
\\l deformation under seismic or accident conditions (see Sections 6 and 7) will not reduce the minimum spacing between fuel assem-blies to less than 12 in.
Consequently, fuel assembly drop accidents will not result in an increase in reactivity above that calculated for the infinite nominal design storage rack.
4.6.3 Abnormal Positioning of Fuel Assembly Outside Storage Rack If a fresh fuel assembly of the highest initial reactivity were to be positioned outside and adjacent to a fuel rack, the reactivity could potentially be increased by a percent or more.
If the fuel element had accumulated some burnup prior to dis-charge, the reactivity increase would be less.
The fuel racks, however, are designed so that the space between the fuel rack and the pool wall (<5 in.) is not sufficient to permit a fuel assem-bly being abnormally positioned outside a fuel rack.
Similarly, t
(.)
t
(/
the spacing between rack modules is too small to permit inserting 4-32
nC an extraneous ascembly in these positions.
The Region 3 cells facing the cask area represent the only nodule faces where an extraneous assembly might conceivably be positioned.
To preclude this, those modules facing the cask area will have mechanical stcps to assure that a minimum space of 8 in. will be maintained between - the rack module and a potential extraneous assembly.
From Fig. 4.10, this spacing is more than adequate to maintain k less than 0.95.
Furthermore, soluble boron is normally present in the spent fuel pool (for which credit is permitted in this condition) and would reduce the maximum k to substantially less than 0.95.
Therefore, it is concluded that the abnormal posi-tioning of a fuel assembly outside and immediately adjacent to the storage rack is not a credible occurrence.
4.6.4 Lateral Rack Movement Lateral motion of the rack modules under seismic conditions g.g V
could alter the spacing between rack modules.
However, the lat-eral motion is not of sufficient magnitude to reduce the spacing
-to less than the nominal spacing of the flux-trap water gaps in the reference storage cell.
In addition, soluble boron would substantially reduce the k, under the postulated conditions.
e 9
4-33
/'
g i
1
's )
_2-l.30 t_._.
-~~
l.25 -
--- _ __4 l.20 ----
t' _
~_
8
.as 1.15 -__[g7i:g_.
,z--- --]- - - -
~E r:(1_- r_.1g
~
3 4
x g_
_=g-g 4
~
V)
W l.10 - Ca.
. g.
z Z
z o 1.05-
.;g - ---- _. -, _ _ _. _. _
4= 3. g; -- --. _
H g-
= _ _ _ _ _ _ _ _.
- _ _ __ =.
g u< l. 0 0 --
z.
W 2
0.98-O.96-
~
0.94-0.92-0.90
~ ~ ~ '
~~~' ~ ^ ~ ~
~^
~ - ~
~ ~ ', ^ ~ ~
' - - ~
~
O 2
4 6
8 10 12 W ATER G AP BETWEEN FUEL ASSEMBLIES, Inches n
(
Fig. 4.10 Infinite multiplication factor (b ) of 4.3% enriched fuel assemblies separated only by water.
4-34
REFERENCES 1.
Green, Lucious, Petrie, Ford, White, Wright, PSR-63/AMPX-1 (code package), AMPX Modular Code System for Generating coupled Multigroup Neutron-Gamma Libraries from ENDF/B, ORNL-TM-3706, Oak Ridge National Laboratory, March 1976.
2..
L.
M.
Petrie and N.
F.
Cross, KENO-IV, An Improved Monte Carlo Criticality Program, ORNL-4938, Oak Ridge National Laboratory, November 1975.
1 3.
S.
R.
Bierman et al.,
Critic [h5 Subcritical Clusters of 4.29 wt% U Enriched 002 Rods in
' Water with Fixed Neutron Poisons, NUREG/CR-0073, Battelle Pacific Northwest. Laboratoriec, May 1978, with errata sheet issued by the USNRC August 14, 1979.
4.
M.
N.
Baldwin et al.,
Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel, BAW-1484-7, The Babcock & Wilcox Company, July 1979.
l S.
R.
M. Westfall and J. R. Knight, Scale System Cross-Section Validation with Shipping-Cask Critical Experiments, ANS Transactions, Vol. 33, p. 368, November 1979.
6.
S.
E.
Turner and M.
K.
- Gurley, Evaluation of AMPX-KENO Benchmark Calculations for High Density Spent Fuel Storage 1
- Racks, Nuclear Science and Engineering, 80(2):
230-237, February 1982.
7.
A.
Ahlin and M.
- Edenius, CASMO
-A Fast Transport Theory Depletion Code for LWR Analysis, ANS Transactions, Vol. 26,
- p. 604, 1977.
CASMO-2E Nuclear Fuel Assembly Analysis, Application Users Manual, Rev. A, Control Data Corporation, 1982.
8.
W.
J.
Eich, Advanced Recycle Methodology Program, CEM-3, Electric Power Research Institute, 1976.
9.
W.
A.
Wittkcipf, NULIF Neutron Spectrum Generator, Few-Group Cons tint Generator and Fuel Depletion Code, BAW-426, The Babcock & Wilcox Company, August 1976.
10.
NAI Modified
- LEOPARD, Rev.
2, NAI Report 71-13 (Proprietary), Nuclear Associates International Corporation, December 10, 1973.
f.
" CHEETAH-P "
report module within the LEAHS Nuclear Fuel Management and Analysis Package, Publication No. 84004100 (Proprietary), Nuclear Associates International Corporation, d
July 1974.
(
REFERENCES ( Continued )
11.
E.
E.
- Pilat, Methods For the Analysis of Boiling Water Reactors (Lattice Physics),
YAEC-1232, Yankee Atomic Electric Co., December 1980.
M.
Edenius et al.,
CASMO Benchmark Report, Studsvik/RF-78/6293, Aktiebolaget Atomenergi, March 1978 12.
H.
- Richings, Some Notes on PWR (W)
Power Distribution Probabilities For LOCA Probabilistic
- Analyses, NRC Memorandum to P. S. Check dated July 5,1977.
13.
R.
A.
Lorenz et al., Fission Product Release From Highly Irradiated LWR Fuel, NUREG/CR-0722, February 1980.
14.
S.
J.
Dagbjartsson et al., Axial Gas Flow in Irradiated PWR Fuel
- Rods, TREE-NUREG-ll58, EG&G
- Idaho, Inc.,
September 1977.
15.
M.
G.
Natrella, Experimental Statistics National Bureau of Standards, Handbook 91, August 1963.
16.
W.
R.
Cadwell, PDQ-7 Reference Manual, WAPD-TM-678, Bettis p
y Atomic Power Laboratory, January 1967.
1 i
A 5.
THERMAL-HYDRAULIC CONSIDERATIONS j
b A central objective in the design of the high-density fuel rack is to ensure adequate cooling of the fuel assembly cladding.
In the following, a brief synopsis of the design basis, the method of analysis, and computed results are given.
Similar analysis has been used in previous licensing reports on high density spent fuel racks for Fermi II (Cocket 50-341), Quad Cities I and II (Dockets 50-254 and 50-265), Rancho Seco (Docket 50-312), Grand Gulf Vnit 1 'Itocket 50-416),
and Oyster Creek (Docket 50-219).
5.1 Decay Heat Calculations for the Spent Fuel This report section covers requirement III.l.5(2) of the NRC "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications" issued on April 14, 1978.
This requirement states that calculations for the amount of thermal energy removed by the spent fuel cooling system shall be made in accordance with Branch Technical Position APCSB 9-2
" Residual Decay Energy for Light Water Reactors for Long Term Cooling"1 The calculations contained herein have been made in accordance with this requirement.
5.1.1 Basis The V.C.
Summer reactor is rated at 2775 Megawatt-Thermal (MWT).
The core contains 157 fuel assemblies.
Thus, the average operating pdwer per fuel assembly, Po, is 17.675 MW.
The fuel assemblies are assumed to be removed from the reactor after a 1000 EFPD (Effective full power days) in accordance with the plant system diagram. The fuel discharge can be made in one of the following two modes:
(i)
Normal discharge - Mode (i)
()v (ii) Full Core discharge - Mode (ii) 5-1
As shown in Table 1.1, the anticipated fuel batch size for normal
[]
discharge may vary from 68 to 72.
- However, for analysis purposes, we assume the batch size to be 80 fuel assemblies.
The fuel transfer begins after 144 hours0.00167 days <br />0.04 hours <br />2.380952e-4 weeks <br />5.4792e-5 months <br /> of cool-off time in the reactor (time after shut down).
It is assumed that the time period of discharge of this batch is 27 hours3.125e-4 days <br />0.0075 hours <br />4.464286e-5 weeks <br />1.02735e-5 months <br />. (Three assemblies transferred to the pool per hour). The cooling system consists of two seismic category I spent fuel cooling circuits.
The bulk temperature analysis assumes the typical *~ operating condition of two spent fuel. pool coolets working in parallel.
For comparison purposes, the bulk temperatures corresponding to batch sizes of 1/3 rd core (53 assemblies) and 76 assemblies are also calculated.
- However, for these cases the normal' condition of only one cooler in operating condition is analyzed and reported herein.
Mode (ii) corresponds to a full core discharge (157 assemblies).
Eul.1 core off-load condition implies that the 4
reactor core has no remaining fuel. It is assumed that the total time period for the discharge'of one full core is 52 hours6.018519e-4 days <br />0.0144 hours <br />8.597884e-5 weeks <br />1.9786e-5 months <br /> (after 144 hours0.00167 days <br />0.04 hours <br />2.380952e-4 weeks <br />5.4792e-5 months <br /> of shut down time in the reactor).
The dischatge rate to the pool is assumed to be continuous and uniform. The bulk temperature analysis assumes two spent fuel pool coolers working in parallel.
I j
The water inventory in the reactor cavity cooled by the l
RHR heat exchanger exchanges heat with the fuel pool water mass i
through the refueling canal.
This important source of heat t
removal is also neglected in the analysis.
Thus, the results obtained for both modes (i) and (ii) are highly conservative.
l Normal condition refers to normal di'scharge assuming only one cooler in operating condition;
- however, Arpically two
/~N
()
coolers are available.
l L
5-2 y
-g, w -
w---
m-
~ ~
(~h In the, following, all relevant perforraance data for the
\\_/
spent fuel pool heat exchangers is given.
a.
Spent Fuel PoDT Heat Exchanger:
Type Tube and shell Quantity 2
Performance data o Heat transferred 14.02 x 106 Btu /hr Tube Side o Fluid flow 1,800 gpm
~
o Pool water inlet temperature 135*F o Outlet temperature, 119.2*F Shell Side o Fluid flow 1800 gpm o Coolant inlet temperature 105'F
()
o Outlet temperature 120.8'F o Touling Eactor 0.0005 The above data enables complete characterization of the thermal performance of the fuel pool heat exchanger.
5.1.2 Model Description Reference (l) is utilized to compute the heat dissipation requirements in the pool.
The total decay power consists of," fission products decay" and " heavy element decay".
l Total decay-power P for a fuel assembly is given as a linear l
function of Po and an exponential function of to and ts+
i.e.: P = Po f(to,ts) whero
()
P=
linear function of Po 5-3 l
1
p P,o= average operating power per fuel assembly V
to= cumulative exposure time of the fuel assembly in the reactor t = Time elapsed since reactor shutdown s
The uncertainty factor K,
which occurs in the functional relationship f
(to,ts) is set equal to 0.1 for t
> 107 s
sec in the interest ' of conservatism.
Furthermore, the operating power Po is taken equal to the rated power, even though the reactor may be operaMng at less than its rated power during most of the period of exposure of the batch of fuel assemblies. Finally, the computations and results reported here are based on the discharge taking place when the inventory of fuel in the pool will be at its maximum resulting in an upper bound on the computed decay heat rate.
O Having determined the heat dissipation rate, the next task is to evaluate the time-temperature history of the pool water.
Table 5.1.1 identifies the loading cases examined.
The pool bulk temperature time history is determined using the first law of thermodynamics (conservation of energy).
A number of simplifying assumptions are made wnich render the analysis conservative.
The principal ones are:
1.
The cooling water temperature in the fuel pool cooler is based on the maximum postulated values given in the FSAR.2 2.
The heat exchangers are assumed to have maximum fouling.
Thus, the temperature effectiveness, s,
for the heat exchangers utilized in the analysis are the lowest postulated values:
S=
0.526 for 5-4
fuel pool
< coolers, S
is calculated from heat qV exchanger technical data sheets.No heat loss is assumed to take place through the concrete floor.
4.
No credit is taken for the improvement in the film coefficients of the heat exchangers as the.
operating temperature rises.
- Thus, the film coefficient used in the computations are lower bounds.
5.
No credit is taken for evaporation of the pool water.
The basic energy conservation relationship for the pool heat exchanger system yields:
Ct
= 0 03 (5.1.2) 1 dT D0 where Ct:
Thermal capacitance of stored water in the l
- pool, t:
Temperature of pool water at time, t 0 :-
Heat generation rate due to stored fuel 1
l assemblies in the pool.
01 is a
known i
function of
- time, T
from the preceding section.
02:
Heat removed in the fuel pool cooler.
03:
Heat removed in the RHR heat exchanger (03=0 if RHR is not used).
5-5 i
l
The pool has total water inventory of 38162 cubic feet phen all racks are in pl. ace in the pool and every storage location is occupied.
5.1.3 Decay Heat Calculatfion Results:
The calculations were performed for the pool disregarding the additional thermal capacity and cooling system available in the transfer channel, and the reactor cavity.
For a specified coolant init.t temperature and flow
- rate, the quantities 02 and 03 are shown to be linear function of t
in a recent paper by Singh (3).
As stated 0,
is an exponential function of T.
Thus Equation
- earlier, 1
(5.i.2) can be integrated to determine t directly as a function of T.
The results are plotted in Figures (5.1.1) to (5.1.4).
The results show that the pool water never approaches the boiling f) point under the most adverse conditions.
These figures also give v
01 as a function of T.
Pour plots are generated for each case.
The first and third plots for eacli case shows temperature and power generation respectively for a period extending from t=
2r where T is the total time of fuel transfer.
The 0 + r =
n n
second and fourth plots show the same quantities (i.e.
temperature and power generation respectively) over a
long period.
The long-term plots are produced to indicate the required operating time for the heat exchangers.
Summari.ed results are given in Table 5.1.2.
~
Finally, computations are made to determine the time interval to boiling after all heat dissipation paths are lost. Computations are made for each case under the following two assumptions:
(i)
All cooling sources lost at the instant pool bulk temperature reaches the maximum value.
f's a
5-6
O O
O TABLE 5.1.1 LIST OF CASES ANALYZED i
- i i
j Case No.
Condition No. of No. of No. of Total Time Cool off time 4
fuel spent fuel RHR's to transfer before transfer assemblies pool HXS in-service fuel into begins, hrs.
discharged the pool N
the hrs.
I 1
Normal discharge 80 2
0 27 144 2
Full core 157 2
0 52 144 discharge l
3, Normal discharge 76 1
0 25 144 4
Normal discharge 53 1
0 18 144
~
l 1
4 l
l
O O
O TABLE 5.1.2 MAXIMUM POOL BULK TEMPERATURE t, COINCIDENT TOTAL POWER Q1 and COINCIDENT SPECIFIC POWER FOR THE HOTTEST ASSEMBLY Case No. of Time Maximum Coincident Coincident 01x10-6 Notes No.
Assemblies to transfer pool bulk time (sinct specific BTU / hour fuel into temp.*F initiation power q, pool, hrs.
of fuel BTU /sec.
transfer, hrs.
1 80 27 123.3 34 49.77 17.018 Maximum normal batch size 2
157 52 136.7 58 47.37 29.454 Full coro offload 3:
76 25 139.7 40 49.13 -
16.123
' Normal condition 4
53 18 131.2 34 49.77 12.179 Normal condition.
1/3 rd core off-load e
^
o.
O
~
O TABLE 5.1.3 t-TIME (Hrs) TO BOILING AND BOILING VAPORIZATION RATE FROM THE INSTANT ALL COOLING IS LOST Case No.
, CONDITION 1 CONDITION 2
.Loss of Cooling at maximum Loss of Cooling at-maximum pool bulk temperature power discharge rate Time (Hrs)
Vap. Rate Time (Hrs)
~
P Vap. Rate Ib./hr.
lb./hr.
1 13 17135 12 17380 Yu 2
6 30050 6
30360 3
10 16332 11
' 16738 4
16 12230 16 12553 4
4 l
q l
l
rx (ii) All cooling paths lost at the instant the heat j
i'v dissipation power reaches its maximum value in the pool.
Results are summixixed in Table 5.1.3.
Table 5.1.3 gives the bulk boiling
- v. api..riiz at ion rate for both cases at the instant the boiling commences.
This rate will decrease with time due te reduced heat generation in the fuel.
5.2 Thermal-Mydraulics Analyses for Spent Fuel Cooling This report section covers requirement III.l.5(3) of the NRC "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications" issued on April 14, 1978.
Conservative methods have'been used to calculate the maximum fuel cladding temperature as required therein.
- Also, it has been determined that nucleate boiling or voiding of coolant on the p
surface of the fuel rods does not oc. cur.
v
5.2.1 Basis
In order to determine an upper bound on the maximum fuel cladding temperature, a series of conservative assumptions are mada.
The most important assumptions are listed below:
a.
As stated above, the fuel pool will contain spent fuel with varying
" time-after-shutdown" (ts).
Since the heat emission falls off rapidly with
~
increasing ts, it is obviously conservative to assume that all fuel assemblies are fresh (ts 144 hours0.00167 days <br />0.04 hours <br />2.380952e-4 weeks <br />5.4792e-5 months <br />) and they all have had 24000
=
hours of operating time in the reactor.
The heat emission rate of each fuel assembly is assumed to
+
be equal.2 O
5-10
b.
As shown in Figures 2.1 in Section 2, the modules (o) occupy an irregular floor space in the pool.
For purposes of the hydrothermal analysis, a circle circumscribing the actual tack floor space is drawn.
It is further assumed that the cylinder with this circle as its base is packed with fuel assemblies at the nominal pitch of 10.405 inches (see Figure 5.2.1).
c.
The downcomer space around the rack module group varies, as shown in Figure 5.2.1.
The nominal downcomer gap available in the pool is assumed to be the total gap available around the idealized cylindrical rack; thus, the maximum resistance to downward flow is incorporated into the analysis.
d.
No downcomer flow is assumed to exist between the rack modules.
O
(.
5.2.2 Model Description In this manner, a conservative idealized model for the rack assemblage is devised.
The water flow is axisymmetric about the vertical axis of the circular rack assemblage, and thus, the flow is two-dimensional (axisymmetric three-dimensional). Fig.
5.2.2 shows a typical " flow chimney" rendering of the thermal hydraulics model. The governing equation to characterize the flow field in the pool can now be written.
The resulting integral equation can, be solved for the lower plenum velocity field (in the radial' direction) and axial velocity (in-cell velocity field), by using the method of collocation.
It should be added that the hydrodynamic loss coefficients which enter into the formulation of the integral equation are also taken from well-recognized sources 4 and wherever disc.repancies in reported values exist, the conservative values are consistently used.
O
()
Reference [5] gives the details of mathematical analysis used in this solution process.
5-11
(Q";
Af ter,the axial vr/locity field is evaluated, it is a straight-forward matter to compute the fuel assembly cladding temperature.
The knowledge of the overall flow field enables pinpointing the storage locatiion with the minimum axial flow (i.e. maximum water outlet temIperature).
This is called the most
" choked" location.
It is recognized that some storage locations, where rack module supports are located, have some additional hydraulic resistance not encountered in other cells.
In order to find an upper bound on the temperature in such a cell, it is assumed that it is located at the most
" choked" location.
Knowing the global plenum velocity field, the revised axial flow through this choked cell can be calculated by solving the Bernoulli's equation for the flow circuit through this cell.
Thus, an absolute upper bound on t,he water exit temperature and maximum fuel cladding temperature is obtained.
It is believed that in view of the aforementi6ned assumptions, the temperatures calculated in this manner overestimate the temperature rise that
[]
will actually occur in the pool.
Tha maximum pool bulk temperature t is computed in Section 5.1.3 and reported in Table 5.1.2.
The corresponding average power output from the hottest fuel assembly, q is also reported in that table.
The maximum radial peaking factor is 1.55 for the V.C.
Summer installation.
Thus, it is conservative to assume that the maximum specific power of a fuel assembly is given by 9A " 9 Or where ar = 1.55 The maximum temperature rise of pool water in the most disadvantageously placed fuel assembly is given in Table 5.2.1 for all loading cases.
Having determined the maximum " local" water temperature in the pool, it is now possible to determine the maximum fuel cladding temperature.
It is conservatively assumed that the total peaking factor aT is 2.32.
Thus, a fuel rod can produce 2.32 times the average heat emission rate over a 5-12
small length.
The axial bert. dissipation in a rod is known to a
reach a maximum in the central region, and taper off at its two extremities.
For the sake o.f stided conservatism it is assumed that the peak heat emisslan ocesrs at the top where the local water temperature also reaches its maximum.
Furthermore, no credit is taken for axial conduction of heat along the rod.
The.
highly conservative model thus constructed leads to simple algebraic equations wh.ich directly give the maximum local cladding temperature, tc.
5.2.3 Results
Table 5.2.1 gives the maximum local cladding temperature, t at the instant the pool bulk temperature has c,
attained its maximum value.
It is quite possible, however, that the peak cladding temperature occurs at the instant of maximum value of qA, i.e.,
at the instant when the fuel assembly is first placed in a storage location.
Table 5.2.2 gives the maximum local cladding temperature at t =
0.
It is to be noted that there are wide margins to local boiling in all cases.
The local boiling temperature near the top of the fuel cladding is 240*F.
Furthermore, the cladding temperature must be somewhat higher than the boiling temperature to initiate and sustain nucleate boiling.
The above considerations indicate that a j
comfortable margin against the initiation of localized boiling l
exists in all cases.
e O
5-13 l
O O
O TABLE 5.2.1 MAXIMUM LOCAL POOL WATER TEMPERATURE AND LOCAL FUEL i
t l'
CLADDING TEMPERATURE l
AT INSTANCE OF MAXIMUM POOL BULK TEMPERATURE 1
1 Case No.
Max. Local Pool Maximum Coincident Local Case Waterremperature 'F Cladding Temperature *F Identified i
1 155.5 184.0 80 Assemblies q
Cooling Mode A 2
139.8 167.1 157 Assemblies Cooling Mode A Y
5 3
171.6 199.8 76 Assembil0h Cooling Mode B 1
4 163.4 191.9 53 Assemblies Cooling Mode B i
e
- Cooling Mode A means 2 spent fuel pool coolers (SFPHX).
1 Cooling Mode B means only 1 SFPHX in operation.
i i
i l
8 i
]
1
O TABLE 5.2.2 POOL A14DiMAXIMUM CLADDING.. TEMPERATURE AT THE INSTANCE FUEL ASSEMBLY TRANSFER BEGINS Case No.
Cladding Coincident Pool Temp.
'F
- Temp,
- F Bulk Local i
1 or 2 172.8 108 142.0 3 or 4 175.7 110.9 145.0 i
i O
5-15
+ -
REFERENCIS TO SECTION 5 1.
NUREG 0800 U.S. Nuclear Regulatory Commission, Standard Review Plan, Branch Technfmal Position, ASB 9-2, Rev. 2, July 1981.
2.
FSAR, V.C. Summer Nuclear Station.
3.
Journal of Heat Transfer, Transactions of the ASME August,1981, Vol. 103, "Some Fundmental Relationships for Tubular Heat Exchanger Thermal Performance," K.P. Singh.
4.
General Electric Corporation, R&D Data Books, " Heat Transfer and Fluid Flow, 1974 and updates.
5.
4th National Congress of the ASME, "A Method for Computing the Maximum Water Temperature in a Fuel Pool Containing Spent Nuclear Fuel", paper 83-NE-7, Portland, Oregon (June 1983).
O O
5 I
5-16 i
8 i
S-oo PEAK VALUE = 123.3 F AT 34 HRS.
m.
/
^8
- a..o CDN-Ed" 8
CASE 1 gC o
Ido Number of Assemblies 80
=
hk Time of Discharge 27 hrs.
=
[
Spent Fuel Pool Heat Exchangers=
2 R.H.R. Heat Exchanger 0
=
CLo w$
Z F--
- 8 l
4 i
a-8
?
4 o
1.00 1'0.00 2'0.00 3'0.00 3'0.00 5'O.00 6'O.00 1
TIME (HOURS)
FIG. 5.1.1 (a) P0OL RULK TEMcERATURE; NORf4AL DISCHARGE l
5-17
O 8
i S!-
o 9
PEAK VALUE = 123.3 F AT 34 HRS.
0 5-
/
o E".
-o CD84-O
=i WO mf Number of Assemblies 80
=
p Time of Discharge 27 hrs.
=
Spent Fuel Pool Heat Exchangers =
2 g
. R.H.R. Heat Exchanger 0
=
z4 w.
W*
8 i
S-8 4
o 1.00 2'.00
/.00 6'.00 8'.00 1'0.00 i2.00 TIME (DAYS)
O FIG. 5.11(b) P0OL BULK TEMPERATURE; NORMAL DISCHARGE 5-18 l
[
i O
o o,
=
o9oo.
"e PEAK VALUE = 17.25 x 10 BTU /HR.
6
,4 AT 27 HRS.
S oo E'o re.
Nw
"")
6--
ED CASE 1 g
Number of Assemblies 80
=
4 Time of Discharge 27 hrs.
=
E S.F.P. Ileat Exchangers 2
=
- C I
R.H.R. lient Exchangers 0
o
=
Uo gr3 -
o$-
sw K
u.1Zo O.
Q.O-o
%.00 1'.00 2'.00 3'.00 4'O.00 5'.00 6'0.00 0
O 0
O TIME (HOURS)
O
~
FIG. 5.1.1 (c)
POWER DISCliARGE; NORMAL DISCllARGE 5-19 5
l 0
o o
o' w.
N o9oo.
6 e
PEAK VALUE = 17.25 x 10 BTU /HR. AT 27 HRS.
- o no mg Ze Na.
D O
a CASE 1 e
E Number of Assemblies 80
=
4 "E
Time of Discharge 27
=
o C)o m-S.F.P. Heat Exchangers 2
=
go.
8D c
R.H.R. Heat Exchangers 0
=
K La No O.
11.0-o
%.00 2'.00 4'.00 s'.oo s'. co l'o.oo t'2.00 TIME (DAYS)
O FIG. 5.1.1 (d) POWER DISCHARGE; NORMAL DISCHARGE 5-20
O 0
WM, VALUE = 136.65 F AT 58 HRS.
8 8.
w 8
4 2-
. CASE 2 d
i Number of Assemblies 157 g
=
CD$-
Time of Discharge 52
=
W Q
Spent Fuel Pool Heat Exchangers=
2 o
R.H.R. Heat Exchanger 0
=
mo O.
x F"
4 EWo CLo Ef w.
F-
- i o
2 O
l O,
l m
x O
1.00 2'O.00
/0 00 6'0.00 8'O.00 i00.00 i20.00 TIME (HOURS) l O l
FIG. 5.1.2 (a) POOL RULK TEMPERATURE; FULL CORE DISCHARGE i
l I
s-21 x
s
'\\
)-
O 0
PEAK VALUE = 136.65 F AT 58 HRS.
i c-8 i
C-E.
gQ CASE 2 o
Number of Assemblies
= 157 Time of Discharge 52
=
o SFP Heat Exchanger 2
=
Og.
R.H.R. Heat Exchanger 0
=
CC Wo 0.0 Ef w.
8 i
8 N
1.00 4'.00 8'.00 i2.00 i6.00 2'0.00 2'4.00 TIME (DAYS)
O FIG. 5.1.2 (b) POOL BULK TEMPERATURE; FULL CORE DISCHARGE 5-22
O
)
o PEAK VALUE = 29.77 x 10 BTU /IIR.
O, AT 52 ilRS.
o S~
o?og-CASE 2 vg Number of Assemblies 157
=
Time of Discliarge 52
=
3o%*
S.F.P. Heat Exchangers 2
=
o.
N O8 R.H.R. Heat Exchangers 0
=
~3 O
80 l--
a.
Ow-C.D W
<Cx8 o.
(fjo HO-O Wwza O.
O-E-o
%.00 2'O.00 4'0.00 6'0.00 8'0.00 1'00.00 1'20.00 TIME (HOURS)
O FIG. 5.1.2 (c)
POWER DISCHARGE; FULL CORE DISCHARGE 5-23
- i. - - - -.
O o
6 PEAK VALUE = 29.77 x 10 BTU /liR. AT S2 ilRS.
8.
m 8
8.
.N
'O
,c*oo E*d.
NW CASE 2
"")
O bg Number of Assemblies 157
=
gg" Time of Discharge S2
=
S.F.P. ficat Exchangers 2
=
E R.II.R. licat Exchangers 0
=
4x8 o.
U2O HO-O W
LAJ28 O.
a.8-l o9
- b.co 4'.00 e'. 00 i2.00 t's.oo 2'o.co 2'4.00 TIME (DAYS)
O FIG. 5.1.2 (d) Pt 3ER DISCilARGE: FULL CORE DISCf!ARGE l
l l
5-24 l
l
O PEAK VALUE = 139.7 F at 40 hrs.
{
o
.o4w oo d
CASE 3 w"
Number of Assemblies
= 76 o
Time of Discharge
= 25 hrs.
LL.o Spent Fuel Pool Heat Exchar1gers
= 1 CDQ-R.H. R. Heat Exchanger
= 0 O
w o
W *.
E ns O cu.
B" 4
KWo (1,,o Ed LLitu.
F-
- o N.
w oo w
1.00 i0.00 2'0.00 3'0.00 4'O.00 5'0.00 6'O.00 TIME (HOURS)
O FIG. 5.1.3(a) Pool Bulk Temperature; Normal Discharge with Loss of 1 SFPHX.
5-25
O PEAK VALUE = 1.FJ.7% at 40 hrs.
o k.
[
w o
E.
e o
CASE 3 m*
11.o w"
Number of Assemblies w
=
76 o
Time of Discharge
=
25 hrs.
Spent Fuel Pool Heat Exchangers
=
1 m
R.H.R. Heat Exchanger
=
0 Em Dns.
4W 11Jo lo Z'o ww.
F--
- o b.
o w_.
4.00 a'.co 4'.00 s'. co s'. co io.co t'a.co TIME (DAYS)
FIG. S.1.3 (b) Pool Bulk Temperature; Normal Discharge With Loss of 1 SFPHX l
5-26
O O
9o'-
N o9oo.
m N 6
PEAK VALUE = 16.5875 x 10 BTU /HR. at 25 hrs.
Ow*oo Wg 1 40 N*.
3 s
D--
G1S.
CASE 3 OS Number of Assemblies
=
76 Time of Discharge 25 hrs.
=
Io Spent Fuel Pool Heat Exchanger
=
1 Uo U3@*
R.H.R. Heat Exchanger
=
0 H
Q W
tdx8 O.
CLO-oo
.00 1'0.00 2'0.00 3'0.00
/0.00 5'0.00 6'0.00 TIME (HOURS) i FIG. 5.1.3 (c) Power Discharge; Normal Discharge With l
Loss of 1 SFPHX.
5-27
_s s--
m.
s.s O
i o
E.
N o9oo.
e tu Oy PEAK VALUE = 16.5875 x 10 BTU /hr.
oo E*
".3 F-O E"o o.
Do CASE 3 m-CD E
4To Number of Assemblies
=
76 CJo 25 hrs.
U3
- Time of Discharge
=
H@~
Q Spent Fuel Pool Heat Exchangers
=
1
=
0 g
R.H.R. Heat Exchanger tux8 o.
EE' o
%.00 2'.00 4'.00 6'.00 8'.00 l'O.00 i2.00 l
TIME (DAYS) l i
FIG. 5.1.3 (d) Power Discharge; Normal Discharge With loss of 1 SFPHX.
5-28 9
-9 p9
..v--.
3w.,.
_pg
,m e
o.
l
\\
g PCAK VALUE = 131.2 F at 34 hrs.
C-8 4
2-CME 4 E,9 A
m.
Number of Assemblies 53
=
Time of Discharge 18 hrs.
=
~,
Spent Fuel Pool Heat Exchangers 1
=
10 9 R.H.R. Heat Exchanger 0
=
Go Dat.
F*
4 T
LILio D_ o E5 tu._
l--
- 8 4
U~
8 1
i
.00
[0.00 2'0.00 3'0.00
/0.00 5'0.00 6'O.00 TIME (HOURS)
O'-
FIG. 5.1.4 (a) Pool Bulk Temperature,1/3 Core Discharge with Loss of 1 SFPHX 5-29
O 8
E PEAK VALUE = 131.2 F at 34 hrs.
0 w
a i
2-o E*
W N.
w" CASE 4 9
O o
Number of Assemblies 53
=
f$-
Time of Discharge 18 hrs.
=
4 Spent Fuel Pool Heat Exchangers 1
=
EWo R.H.P,. Heat Exchanger 0
=
Q.o Zf w
t--
- 8 t
l 8
5 1.00 2'.00
/.00 6'.00 8'.00 i0.00 i2.00 TIME (DAYS)
O
~
FIG. 5.1.4 (b) Pool Bulk Temperature; 1/3 Core Discharge With Loss of 1 SFPHX 5 30
i t
8 a
0" 6
PEAK VALUE = 12.55 x 10 BTU /HP.. at 18 hrs.
8
.-o
,e*o o.
i gI8 CASE 4 Nw.
"3 i
Number of Assemblies 53 CD
=
~g Time of Discharge 18 hrs.
=
O ',
Spent Fuel Pool Heat Exchanger 1
o
=
Ede CD R.H.R. Heat Exchanger 0
=
E 4To Oo in M@-
O E
tu38 o.
Q-$-
8 i
.00 1'0.00 2'0.00 3'0.00 4'0.00 5'O.00 6'O.00 TIME (HOURS)
O FIG. 5.1.4 (c) Power Discharge; l/3 Core Discharge With Loss of 1 SFPHX 5 31
O 8
i g..
6 PEAK 1RteE s'at.55 x 10 BTU /hr.
. E.
o
,1*o Y'
x8 Nw.
3 CASE 4 F-O e@
Nuinber of Assemblies 53
=
O*g.
Time of Discharge 18 hrs.
=
CD Spent Fuel Pool Heat Exchangers 1
=
E<
R.H.R. Heat Exchanger 0
=
Zo Oo U3
- H@-
O E
LLJa 3o O-E8" 8
6
%.00 2'.00 4'.00 6'.00 S'.00 1'0.00 i2.00 TIME (DAYS)
O FIG. 5.1.4 (d) Power Discharge; 1/3 Core Discharge With Loss of 1 SFPHX 5-32
Actual id e a liz e d O wtlin e Outline of NOf Rack. Assembly Rack Assembly N.
ff/
v
//////////8,/'
k-m g
~
Actual Ou l.ine idealized Assumed Added
- =. y
~. I
.... 11..
o FIG. 5.2.1 IDEALIZATION OF RACK ASSEMBLY 5-33
l Water Assumed At The Pool Bulk Temperature
-+=
r'
'S J L a
/~7
/
__/
TouTA"
/
/
r_.3 po O
" 1 9
"I 1
o w<
2 3
3' 4
J O
w S
O Heat Addition 3
V 9
u.2 d
3 8
/
e, T \\N A
/
/
=
( 27
/
/
/
)
/
O
,.-s'
~
FIG. 5.2.2 THERMAL CHIMNEY FLOW MODEL 5-34
6.
STRUCTURAL ANALYSIS p
The purpose of this section 12 to demonstrate the structural adequacy of the spent fuel rack design under normal and accident loading conditions.
The method of analysis presented herein is similar to that previously used in the Licensing Repor's on High Density Fuel Racks for Fermi II (Docket 50-341.), Quad Cities I and II (Dockets 50-254 and 50-265), Rancho Seco (Docket No. 50-312),
Grand Gulf Unit 1 (Docket No. 50-415) and Oyster Creek (Docket No.
50-219). The results show that the high density spent fuel racks are structurally adequate to resist the postulated stress combinations associated with normal and accident conditions.
6.1 Analysis outline:
The spent fuel storage racks are Seismic Category I
equipment. Thus, they are required to remain functional during and after an SSE (Safe Shutdown Earthquake).1 As noted previously, these racks are neither anchored to the pool floor, nor are they
^
attached to the side walls.
The individual rack modules are not interconnected.
Furthermore, a particular rack may be completely loaded with fuel assemblies (which corresponds to greatest rack inertia), or it may be partially loaded so as to produce maximum geometric eccentricity in the structure.
The coefficient of
- friction, u,
between the supports and pool floor is another indeterminate factor.
According to Rabinowicz,2 the results of 199 tests performed on austenitic stainless steel plates submerged in water show a mean value of p to be 0.503 with a standard deviation of 0.125.
The upper and lower bounds (72o) are thus 0.753 and 0.2f3, respectively.
Two separate analyses are performed for this rack' assembly with values of p equal to 0.2 (lower limit) and 0.8 (upper limit) respectively. Initially, the following four separate analyses are performed on the largest rack module (Module A in Table 2.2).
1.
Fully loaded rack (all storage locations occupied),
p = 0.8 ( p = coefficient of friction).
2.
Fully loaded rack, p= 0.2.
6-1
O" 3.
Empty rack, y = 0.8.
4.
Empty rack, u = 0.2.
B.ased on the results of these runs, additional analyses are performed.
The actual studies performed for the different rack modules are summarized in Section 6.6.
The method of analysis employed is the time history me hod.
The pool slab acceleration data are developed from the original pool floor response spectra.
The object of the seismic analysis is to determine the structural response (stresses, deformation, rigid body motion, etc.)
due to simultaneous application of the three orthogonal excitations.
Thus, recourse to approximate statistical summation techniques such as " Square-Root-of-the-Sum-of-the-Squares" method 3 is avoided and the dependability of computed results is ensured.
The seismic analysis is performed in four steps; namely 1.
Dey~elopment of nonlinear dynamic model consisting of
- beam, gap,
2.
. Derivation and computation of element stiffnesses using a sophisticated elastostatic model.
O 6-2
l l
O 3.
Layout of the equations of
- motion, and inertial decoupling and solution of the equations using the
" component element time integration" procedure 4,5 to determine nodal and element forces and displacements of nodes.
r 4.
Computation of the detailed stress field in the rack structure, using the detailed elastostatic model, from the nodal forces calculated in Step III above.
Determine H 'the stress and displacement limits,given in Section 6.5, are satisfied.
A brief description of the dynamic model follows.
6.2 Fuel Rack - Fuel Assembly Model 6.2.1 Assumptions O
- a. The fuel rack metal structure is represented by five lumped masses connected by appropriate elastic springs as shown in Figure 6.1.
The spring rates simulate the elastic behavior of the fuel rack as a beamlike structure.
- b. The fuel assemblies are represented by five lumped masses located, relative to the rack, in a manner which simulates either fully or partially loaded conditions.
- c. The local flexibility of the rack-support interface is modeled conservatively in the analysis.
- d. The rack base support may slide or lift off the pool floor.
O 6-3
C)
- e. The pool floor is assumed to have a known time history of ground accelerations along the three orthogonal directions.
- f. Fluid coupling between rack and assemblies, and between rack and adjacent racks is simulated by intradiacing appropriate inertial coupling into the system kinetic energy.
- g. Potential impacts between rack and assemblies are accounted for by appropriate spring gap connectors between masses involved.
- h. Fluid damping between rack and assemblies, and between rack and adjecent rack is conservatively neglected.
s
- i. The supports are modeled as extensional elements for dynamic analysis.
The bottom of a support leg is attached to a
frictional spring as described in Section 6.2.2.
The cross section properties of the support beams are derived and used in the final computations to determine support leg stresses.
- j. The effect of sloshing can be shown to be negligible I
-at the bottom of a pool and is hence neglected.
l i
o a
l 6-4 L
6.2.2 Model Description The absolute degrees of freedom associated with each of the mass locations i, i* are as follows (see Figure 6.1):
Table 6.1 Degrees of Freedom Location Dicplacement Rotation (Node) ux uy ug 0
O o
x y
g 1
P1 P2 93 94 95 96 1*
Point is assumed fixed to base'at X 'Y,2=0 B
B 2
p7 99 911 912 2*
P8 P10 3
P13 P15 917 918 3*
P14 P16 4
P19 P21 923 924 4*
P20 P22 l
5 P25 P27 P32 929 930 431 5*
p26 P28
- O i
i 6-5 l
c
Thus, there are 3a degrees of 2.reedom in the system.
Note that elastic motion of the rack in extension is represented by generalized coordinates P3 and P32 This is due to the relatively high axial rigidity of the rack.
Torsional motion of the rack relative to its base is governed by q31-The members joining nodes 1 to 2, 2 to 3, etc., are the beam elements with deflection capability due to bending and shese (see Reference 4,pp.
156-161.).
The elements of the stiffness matrix of these beam elements are readily computed if the effective flexure modulus, torsion modulus, etc., for the rack structure are known.
These coefficients follow from the elastostatic model as described later.
The nodal points i (i =
1,2..
- 5) denote the fuel rack mass at the 5
elevations.
The node points i*
(i*
=
1,2..
- 5) denote the cumulative mass for all the fuel assemblies distributed at 5 elevations.
The element stiffnesses of the fuel assembly are obtained from the structural properties of the OCNS fuel assemblies. The nodeu i* are located at x =
XBa Y
YB in the global coorditnate system shown in Figure 6.1.
6.2.3 Fluid Coupling An effect of some significance requiring careful modeling is the so-called " fluid coupling effect."
If one body of mass mi vibrates adjacent to another body (mass m2), and both bodies are submerged in a
frictionless fluid
- medium, then the Newton's G
O 6-6
I V
equation of motion for the two bodies have the form (mi + M11) XI-M12 X2 = applied forces on mass mi "M21 X1 + (m2 + M22) X2 = applied forces on mass m2
- M11, M12e
- M21, and M22 are fluid coupling coefficients which depend on the shape of the two bodies, their relative disposition, etc.
Fritz6 gives data for Mij for various body shapes and arrangements.
It is to be noted that the above equation indicates that effect of the fluid is to add a certain amount of mass to the body (Mil to body 1), and an external force which is proportional to the acceleration of the adjacent body (mass m2)*
- Thus, the acceleration of one body affects the force field on another.
This force is a strong function of the interbody gap, reaching large values forms very small gaps.
This inertial coupling is called fluid coupling.
It has an important effect in (j
rack dynamics.
The lateral motion of a fuel assembly inside the storage location will encounter this effect.
So will the motion of a rack adjacent to another rack.
These effects are included in the equations of motion.
The fluid coupling is between nodes i and i* ti 2,3..
5) in Figure 6.1.
Furthermore, nodal manses i
=
contain coupling terms which model the ef fect of fluid in the gaps between adjacent racks.
Finally, fluid virtual mass is included in vertical direction vibration equations of the rack; virtual inertia is added to the governing equations corresponding to rotational degrees of freedom, such as g4, qi, q6' 911, etc.
6.2.4 Damping In reality, damping of the rack motion arises from material hysteresis (material damping), relative intercomponent motion in structures (structural damping),
and fluid drag effects (fluid
/~'i V
damping).
The fluid damping acts on the i and i* nodal masses.
In 6-7
,Q the analysis, a raaximum of 41 structural damping is imposed on elements of the rack structure during SSE seismic simulations.
Actual structural damping values used in the analysis are provided in Table 6.4.
This is in accordance with the FSAR and NRC guidelines 7
Material and fluid damping are conservatively neglected.
6.2.5 Impact The fuel assemb'ly* nodes 1*
will impact the corresponding structural mass node i.
To simulate this impact, 4 impact springs around each fuel assembly node are provided (see Figure 6.2).
The fluid dampers are also provided in parallel with the springs.
The spring constant of the impact springs is assumed equal to the local stiffness of the vertical panel computed by evaluating the peak deflection of a six inch diameter circular plate subject to a specified uniform pressure, and built in around the edge. The spring constant calculated in this manner should provide an upper bound on the local stiffnesses of the structure.
6.2.6 Assembly of the Dynamic Model The dynamic model of the rack, rack base plus supports, and internal fuel assemblies, is modeled for the general three dimensional (3-D) motion simulation, by five lumped masses and inertial nodes for the rack, base, and supports, and by five lumped masses for the assemblage of fuel assemblies.
To simulate the connectivity and the elasticity of the configuration, a total of 18 linear spring. dampers, 20 nonlinear gap elements, and 16 nonlinear friction elements are used.
A summary of spring-damper, gap, and friction elements with their connectivity and purpose is presented in Table 6.2.
i If we restrict the simulation model to two dimensions (one O
horizontal motion plus vertical
- motion, for example) for the V
purposes of model clarification only, then a descriptive model of the simulated structure which includes all necessary spring, gap, 6-8
I
(
and friction eleme,nts is shown in Figure 6.3.
The beam springs, t)
Kar KB at each level, which represent a rack segment treated as a structural beam,4 are locatd in Table 6.2 as linear springs 2,
3,
6 ',
7, 10, 14, and 15.
The extensional spring, K
which e,
simulates the lowest elastic motion of the rack in extension relative to the rack base, is given by linear spring 18 in Table 6.2.
The remaining spring-dampers either have zero coefficients (fluid damping is neglected),
or do not enter into the two-dimensional (2-D) motion shown in Figure 6.3.
The rack mass and inertia, active in rack bending, is apportioned to the five levels of rack mass; the rack mass active for vertical motions is apportiioned to locations 1 and 5 in the ratio 2 to 1.
The mass and inertia of the rack base and the support legs is concentrated at node 1.
The impacts between fuel assemblies and rack show up in the gap elements, having local stiffness K1, in Figure 6.3.
In Table 6.2, these elements are gap elements 3,
4, 7,
8, 15, 16, 19 and
- 20. The support leg spring rates K6 are modelled by elements 9 and 10 in Table 6.2 for the 2-D case.
Note that the local elasticity of the concrete floor is included in K6 To simulate sliding potential, friction elements 2 plus 8 and 4 plus 6 (Table 6.2) are shown in Figure 6.3.
The local spring rates Kg reflect the lateral elasticity of the support legs.
Finally, the support l
rotational friction springs KR, reflect the rotational elasticity I
of the foundation.
The nonlinearity of these springs (friction elements 9 plus 15 and 11 plus 13 in Table 6.2) reflects the edging l
limitation imposed on the base of the rack support legs.
For the 3-D simulation, carried out in detail for this analysis, additional springs and support elements (listed in Table 6.2),
are included in the model.
Coupling between the two horizontal seismic motions is provided by the offset of the fuel assembly group centroid which causes the rotation of the entire rack.
The potential exists for the assemblage to be supported on 1 6-9
Table 6.2 Numbering System for Springs, Gap Elements, Friction Elements I.
Spring Dampars (18 total)
Number Node Location Description 1
1-2 X-Z rack shear spring 2
1-2 Y-Z rack shear 3
1-2 Y-Z rack bending spring 4
1-2 X-Z rack bending 5
2-3 X-Z rack shear 6
2-3 Y-Z 7
2-3 Y-Z s
rack bending 8
2-3 X-Z 9
3-4 X-Z rack shear 10 3-4 Y-Z 11 3-4 Y-2 rack bending 12 3-4 X-Z 13 4-5 X-Z rack shear 14 4-5 Y-Z 15 4-5 Y-Z rack bending 16 4-5 X-2 17 17-5 Rack torsion spring 18 1-5 z rack extensional spring O
l 6-10
-.-a
Tabic 6.2 (continued) e'}
's /
II.
Nonlinear Springs (Gap P'9ments) (20 total)
Number Node Location Description 1
2,2*
X rack / fuel assembly impact spring 2
2,2*
X rack / fuel assembly impact 3
2,2*
Y rack / fuel assembly impact 4
2,2*
Y rack / fuel assembly impact 5
3,3*
X rack / fuel assembly impact 6
3,3*
1 rack / fuel assembly impact 7
3,3*
Y rack / fuel assembly impact 8
3,3*
Y rack / fuel assemlby impact 9
Support S1 2 compression spring 10 Support S2 2 compression spring 11 Support S3 2 compression spring 12 Support S4 Z compression spring 13 4,4*
X rack / fuel assembly impact spring 14 4,4*
X rack / fuel assembly impact spring 15 4,4*
Y rack / fuel assembly impact spring 16 4,4*
Y rack / fuel assembly impact spring 17 5,5*
X rack / fuel assembly impact spring 18 5,5*
X rack / fuel assembly impact spring (s) 19 5,5*
Y rack / fuel assembly impact spring 20 5,5*
Y rack / fuel assembly impact spring III. Friction Elements (16 total)
Number Node Location Description 1
Support S1 X direction support friction 2
Support S1 Y direction friction 3
Support S2 X direction friction 4
Support S2 Y direction friction 5
Support S3 X direction friction 6
Support S3 Y direction friction 7
Support S4 X direction friction 8
Support S4 Y direction friction 9
S1 X Floor Moment 10 S1 Y Floor Moment 11 E2 X Floor Moment 12 S2 Y Floor Moment 13 J3 X Floor Moment 14 S3 Y Floor Moment 15 S4 X Floor Moment 16 S4 Y Floor Moment i
V 6-11
to 4 rack suppor,ts during any instant of a complex 3-D seismic event.
All of these potentia 3 events may be simulated during a 3-D motion and have been observed in the results.
A brief description ot the elastostatic model now follows.
This detailed model is used to obtain overall beam stiffness formulae for the rack dynamic model, and to determine detailed stress distributions in the rack from a knowledge of the results of the time history analysis.
6.3 Stress Analysis 6.3.1 Stiffness Characteristics:
4
's a multicell, folded-plate structure i
The fuel rack which has what is colloquially called a
" honey-comb" configuration.
This type of construction is very similar to the so-called " stressed-skin" construction of ribs, spars, and cover plates which are widely used in aircraf t construction. Techniques developed in the field of aircraft structural analysis are utilized herein to find the stresses and deformations in such structures.
These methods have been thoroughly tested and their reliability has been dccumented in a number of publications.8-12 Figure 6.4 shows two cross-sections of the fuel rack which is j.
modeled as a'
rectangular network of plates interconnected along l
nodal lines shown as points in Figure 6.1.
An arbitrary load with l
components
- Fxi, F i, Fzi acts at an arbitrary elevation on y
one of the nodal lines.
We find the displacements and stresses due to such a typical load according to the stressed-skin model as follows.
.The torsional deformations are solved for by using the classical theory of torsion for multicelled, thin-walled, cress sections.13 O
l l
6-12 i
_~
f
/"
(\\
The bending deformation ?s found by using the theory of shear flow 12 wherein all axial stresse s are carried by the effective flanges (or stringrrs) formed by the intersections of the plates and all transverse shears are carried by the plates modeled as t
shear panels.
From a knowledge of the shear flows, the bending and torsional deformations, it is possible to provide a
set of influence functions or the Tollowing section properties for the fuel rack as a whole:
(EI)eq Bending rigidity (in two places)
=
(GJ)eq Torsional rigidity
=
(AE)egs Extensional rigidity
=
ks Shear deformation coefficient
=
such properties are used for the dynamic analysis of seismic loads and serve to establish values for the spring rates of the elastic beam elements representing each rack section.
6.3.2 Combined Stresses and Corner Displacements The cross-sectional properties and the Timoshenko shear correction factor calculated in the previous section are fed into a dynamic analysis of the system shown in Figure 6.5, with a
specifieil ground motion simulating earthquake loading.
From the dynamic
- analysis, the stress resultants (Fx, Fy, Pz, Mx,
~
y, Mz) act. as shown in Figure 6.6 are computed for a large M
number of times t =
at, 2At, etc., at a selected number of cross sections.
The displacements (Ux, U
Uz) at selected nodal y,
points on the - z axis are also provided by the dynamic analysis as well as the rotations (O
e, e) of the cross sections at x,
y z
the nodes.-
6-13
oV Figure 6.7 shows a typicud subdivision of the structure into elbments, nodes, and sectLuns.
The stresses are calculated at all sections and the displacements at all four corners of the racks are calculated at these elevations.
Since the ax'ial stress varies linearly over the cross section and achieves its extreme values at one of the four corners of the
- rack, the shear.stres ses due to torsional loads (Mz) achieve their extreme values near the middle of each side.
The shear stresses due to lateral forces (F
Fy) will achieve their x,
extreme values at the center of the cross section or at the middle of each side.
Thus, candidates for the most critical point on any section will be the points labelled 1 through 9 in Figure 6.8.
The expression for the combined stress and kinematic displacement for each of thece points is written out.
Similarly, the stresses in the support legs are evaluated.
U A
validated Joseph Oat Corporation proprietary computer program "EGELAST"i computes the stresses at the candidate points at each level.
It sorts out the most stressed location in space as well as time.
The highest stress and maximum kinematic displacements are thus readily found.
6.4 Time Integration of the Equations of Motion llav ing assembled the structural model, the dynamic equations of motion corresponding to each degree of freedom can be written by using Newton's I second law of motion or using Lagrange's equation.
For example, the motion of node 2 in y-direction (governed by the generalized coordinate pg) is written as follows:
i t This code has been previously utilized in licensing of similar racks for Fermi II (Docket No.
50-341), Quad Cities I and II g
(Docket Nos.
50-254 and 265),
and Rancho Seco (Docket No.
50-312).
l l
6-14 l
l
.n_
s A
s s
(")
The inertial mass,is:
3*22 + A211 + B211 where m22 is the mass of node 2 for y-directional motion.
A211 is the fluid coupling mass due to interaction with node 2*,
s and s
3 B2Il is the fluid coupling mass due to interaction of node 2 with the reference frame (interaction between adjacent racks).
t Hence, Newton's law gives-s (m22 + A211 + B211) P9 + A212 P10 + B212 u = 09 where 09 represents all the beam' spring and damper forces on node
._, 2, and A212-is the cross term fluid coupling effect of node 2*;
B212 iS'-the cross term fluid coupling effect of the adjacent racks, end u represents ground motion.
x Let 99 = 99 - u 910
- P10 - u That is, q9, is' t,he relative displacement' of node 2 in x-direction f
with, respect' :to the ground.
Substituting in the above equation, l
and rearranging, we have:
s (m22'+ A21'l + B211) 99 + A212 Si.0 09 - (m22 +
l A211 + B211 + A212 + B212 ). -
~.. -
l 6-15 l
(G9 A similar equation for each one of the 32 degrees of freedom can be written.
The system of equations can be represented in matrix notation as:
(M) {'q'} = (0) + {G}
where the vector (0) is a function of nodal displacements and velocities, and {G} depends on the coupling inertia and the ground acceleration.
Premultiplying above equation by (M ]-1 renders the resulting equations uncoupled in mass.
We have:
{**} = [g]-1 [o] + [g]-1 {c}
The generalized force 0,
which contains the effects 9
of all spring elements acting on node 2 in the " direction" of coordinate q9 (the relative displacement of node 2
in the y direction), can easily be obtained from a free body analysis of node 2.
For example, in the 2-D model shown in Figure 6.3, contributions to 09 are obtained from the two shear springs of the rack structure, and the two impact springs which couple node 2*
and node 2.
Since each of these four spring elements contain couplings with other component deformations through the spring force-deformation relations, considerable static coupling of the complete set of equations results.
The level of static coupling of the equations further increases when 3-D motions are considered due to the inclusion of rack torsion and general fuel assembly group centroid effect.
For example, referring to Figure 6.3, and Table 6.1, a 2-D simulation introduces static coupling between coordinates 2,9 and 15 in the expression for 0; this coupling comes from the shear 9
springs simulating the rack elasticity which have constitutive relations of the form F
=K (99 - 92)
,K (415 - 49)
Further, the s
s impact springs introduce two additional forces having constitutive equations of the form P'
= Ky (qg - ql0)
., Of course, at any instant, these forces may be zero if the local gap is open.
The local gap depends on the current value of q9 - ql0 6-16
[]
It should be, noted that in the numerical simulations run to verify structural integrity during a seismic event all elements of the fuel assemblies are a.ssamed to move in phase.
This will provide maxinium impact force level, and hence induce additional conservatism in the time history analysis.
This equation set is mass uncoupled, displacement coupled, and is ideally suited for numerical solution using the central difference scheme.
The computer program named "DYNAHIS"t, developed by General Electric Company and further enhanced by Joseph Oat Corporation, performs this task in an efficient manner.
Having determined the internal forces as a function of time, the computer program "EGELAST"i computes the detailed stress and displacement fields for the rack structure as described in the preceding section.
6.5 Structural Acceptance Criteria There are two sets of criteria to be satisfied by the rack modules:
(a) Kinematic Criterion:
This criterion seeks to ensure that adjacent racks will not impact during SSE (condition E'14) assuming the lower bound value of the pool floor surface friction coefficient. It is further required that the factors of safety against tiltingl5 are met (1.5 for OBE, 1.1 for SSE).
(b) Stress Limits (l[The stress limits of the ASME Code,Section III, Subsection NP, 1983 Edition were chosen to be met, since t These codes has been previously utilized in licensing of similar racks for Fermi II (Docket No. 50-341), Quad Cities I and II (Docket Nos.
50-254 and 265),
and Rancho Seco (Docket No.
c f
50-312).
6-17
this Code provides the most consistent set of limits for various stress types, and various loading conditions.
The following loading combinations are applicable (ref.
14, Sec. 3.8.4).
Load Combination Acceptance Limit D+L Level A service limits D + L + To D + L + To + E D + L + Ta + E Level B service limits D + L + To + Pg D + L + Ta + E' Level D service limits D + L + Fd The functional capability of the fuel racks should be demonstrated where D=
Dead weight induced stresses L=
Live load induced stresses; in this case stresses these are developed during lifting.
Fd:
Force caused by the accidental drop of the heaviest load from the maximum possible height C'
Pg:
Upward force on the racks caused by postulated stuck fuel assembly E:
Operating Basis Earthquake E':
Differential temperature induced loads (normal or upset condition)
Ta:
Differential temperature induced loads (abnormal design conditions)
The conditions Ta and To cause local thermal stresses to be produced.
The worst situation will be obtained when an isolated storage location has a fuel assembly which is generating heat at the maximum postulated rate.
The surrounding storage locations are assumed to contain no fuel.
The heated water makes unobstructed j
l contact with the inside of the storage walls, thereby producing maximum possible temperature difference between the adjacent cells.
The secondary stresses thus produced are limited to the O
l
(,/
body of the rack; that is, the support legs do not experience the secondary (thermal) stresses.
o l
6-18
(2) Basic Data:
'Ite following data on the physical properties of the rack material are obtained from the ASME Codes,Section III, appendices.
Table 6.3 Physical Property Data
- Property Young's Yield Ultimate Modulus Strength Strength
@ 2000F
@2000F
@2000F E
S s
y u
value 28.3 x 106 25 KSI 71 KSI psi Section III Table Table Table Reference I-6.0 I-2.2 I-3.2
- Evaluated at 2000F.
This temperature is higher than the pool water bulk temperature under any of the loading conditions under consideration.
Table 6.4 b
Support Material Data (Table 6.4 gives the equivalent data for the support materials.)
Young's Modulus Yield Strength Material at 200'F at 200*F 6
1 ASTM 479-521800 27.5 x 10 50,000 psi 2 SA564-630 (hardened at 27.6 x 106 125,000 1075'F)
(3.1)
Normal and upset conditions (level A or level B):
(i) Allowable stress in tension on a net section =Ft =0.6 Sy or Ft =(0.6) (25000) =15000 psi (rack material)
Ft is equivalent to primary membrane stresses 30,000 psi (for support Ft (.6) (50,000) a feet)
(ii) on the gross section, allowable stress in shear is Fy = 0.4 S (0.y) (25000) = 10000 psi (main 4
=
rack body)
Ft= (.4)(50,000) =20,000 psi (forsupport feet)
(
6-19
,(iii) Allowable stress in compression, Fa (1 - (
)
2C 8
c y
F
=
((5)
(3 (
)
8Cc] - [(
8C
)
+
c where 2
q,J (2x E)
J Y
Substituting
- numbers, we
- obtain, for both support leg and " honey-comb" region:
Fa = 15000 psi (main rack body)
Fa = 30,000 psi (support legs)
(iv)
Maximum bending stress at the outermost fiber due to flexure about one plane of symmetry:
Pb = 0.60 Sy = 15000 psi (rack body)
Fa = 30,000 psi (support feet)
(v) Combined flexure and compression:
f f
C,x bx C,y by F
DF DF a
x bx y by where f:
Direct compressive stress in the a
section.
f x:
Maximum flexural stress along x-axis b
fby:
Maximum flexural stress along y-axis Cmx = Cmy = 0.85 i
O 6-20
f O
a o
=1-x pi ex f"
D
=1-Y pi ey where 2
12w E p.
ex 2
kl 23 ( b)
- b (vi) Combined flexure and compression (or tension)
I E
E a
bx by
,p
< l.0 0.6 S F
P y
bx by The above requirement should be met for both direct tension or compression case.
(3.2)
Faulted Condition:
F-1370 (section III, Appendix F),
states that the lim,ts for the faulted condition are 1.2 (S /Ft) times the corresponding limits for y
normal condition.
- Thus, the multiplication 1
factor is 5000 (1.2)
= 2.0 Factor
=
15000 i
O 6-21
6.6 Results
.I3 Q;
Figures 6.9, 6.10, and 6.11 show the pool slab motion in horizontal x,
horizontal y,
and vertical directions, respectively.
These plots correspond to the Operating Basis Earthquake.
The corresponding SS8 time history motions are conservatively assumed to be 1.62 times the OBE values.
Since there are several rack module configurations (Fig. 2.1) it was decided to make an exhaustive analysis of one rack type.
We note that module A is an above-average size module, and hence will produce above-average floor reaction and support stress levels. Therefore, module A is chosen for performing extensive analyses.
The support locations for module A are taken' to be three cells in-board of the edge of the rack along one side.
This assumption will maximize rack displacement and ro ta t' ion.
Appropriate simulations are also carried out for other limiting rack p
\\
geometrics (e.g.
tipping study for rack with low cross section to height aspect
- ratio, stress evaluations for the heaviest module, etc.).
To determine the magnitude of structural dampers, free lateral vibration plots of the top of rack A (in X and Y directions) for fully loaded and empty conditions were developed. The dominant natural frequency of vibration thus evaluated enables computations of the linear structural dampers. The percentage structural damping for SSE condition is assumed to be 4%
and modifications to the l
stiffness matrix to incorporate damping is based on the dominantg-f requency of 10 cps.
Having determined the damper i
characte'ristic data, the dynamic analysis of the rack module is performed using the computer program DYNAHIS.
A complete synopsis of the analysis of module A subject to the safe shutdown earthquake motions is presented in the-photocopied computer print-outs labelled as Table 6.5.
Table 6-22
.----e w
w
(mj 6.5 gives
,the maximum values of stress factors (Ri 1,2,3,4,5,6).
The values given in the tables are the (i
=
maximum values in time and space (all sections of the rack).
The various stress factors are listed below for convenience of reference.
R:
Ratio of tensile stress on a
net section to its 1
allowable OBE value R2:
Ratio of gross shear on a net section to its allowable OBE value R3:
Ratio of net compressive ' stress to its allowable OBE value for the section Rg:
Ratio of maximum bending stress in one plane to its allowable value in OBE R:
Combined flexure and compressive factor S
R6:
Combined flexure and tension (or' compression) factor O()
The allowable value of Ri (i 1,2,3,4,5,6) is 1 for OBE
=
condition, and is 2 for SSE condition (see Section 6.5).
The displacement and stress tables given herein are for the SSE condition.
It is noted that the maximum displacements are less than the limiting value for inter-rack impact.
The maximum stress factors (Ri) are well below limiting value for SSE condition for all sections.
Five plots (Figs. 6.12 through 6.15) show the variation j
of support reactions with time.
Figure 6.12 shows the t
l variation of the total rack support reaction (sum of four feet) with time.
Figures 6-13 through 6.15 show the reaction variation in individual feet.
The plots vividly portray the module motion.
We note that support feet never lose contact with the ground.
However, the module displacements are not infinitesimal.
This is due to the in-board location of b
v support feet assumed in the analysis.
Subsequent analyses 6-23 L
,~
(]
with support feet located at the outermost corner cell locations produce substantially lower stresses and displacements.
Seismic simulations for the tipping conditions are carried out by increasing the horizontal SSE accelerations by 50*.15 The calculations indicate that the rack remains stable, and the gross movement remains within the limit of small motion theory. Thus the rack module is seen to satisfy both kinematic and stress criteric with large margins of safety.
Analysis of welded joints in the rack also show large margins of safety.
6.7 Summary of Mechanical Analyses:
The mathematical model constructed to determine the impact velocity of the above falling objects is based on several conservative assumptions, such as
- 1. The virtual mass of the body is conservatively assumed to be equal to its displaced fluid mass.
17 Evidence in the literature indicates that the virtual mass can be many times higher.
- 2. The minimum frontal area is used for evaluating drag coefficient.
- 3. The drag coefficients utilized in the analysis are l8 lower bound values reported in the literature 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 drag coefficient.
O-6-24
-(O
- 4. The,f alling bodies are assumed to be rigid for the j
purposes of impact stress calculation on the rack.
The solution of itbe immersed body motion problem is found analyticaMy.. The impact velocity thus computed is used to determite (Ae maximum stress generated due to stress wave ggation.
e With this model, the fodlowing analyses are performed.
(i)
Dropped Fuel Accident I A fuel assembly (weight - 1616 pounds with control rod assembly) is dropped from 36 inches above the module which impacts, the base.
Local piercing of the base p' late is not found to occur.
Direct impact with the pool liner does not occur.
The subcriticality of the adjacent fuel assemblies is
(
not violated.
(ii) Dropped Fuel Accident II One fuel assembly dropping from 36 inches above the rack and hitting the top of the rack.
Permanent deformation of the rack is found to be limited to the top region such that the rack cross-sectional geometry at the level of the top of the active fuel (and below) is not altered.
l (iii) Jammed Fuel-Handling Equipment and Horizontal Force A
4400-pound uplift force and a
1100-pound horizontal force are applied at the top of the rack at the " weakest" storage location; the force is l
assumed to be applied on one wall of the storage cell boundary as an upward shear force. Th'e - damage 6-25
is found to be limited to the region above the top of the active fuel.
i These analyses prove that the rack modules are engineered to prearide maximum safety against all postulated abnormal and accident conditions.
l O
O G
6-26
REFERENCES TO SECT'ON 6 I
1, USNRC Regulat( y Guide 1,29, " Seismic Design Classification,"
Rev. 3, 1978.
2.
" Friction Coefficients &E Water Lubricated Stainless Steels for a Spent Fuel Rack Facility," by Prof. Ernest Rabinowicz, MIT, a report for Boston Edison Company, 1976.
3.
U.S. Nuclear Regulatory Commission, Regulatory Guide 1.92,
" Combining Modal Responses and Spatial Components in Seismic Response Analysis," Rev. 1, February 1976.
4.
"The Component Element Method in Dynamics with Application to Earthquake and Vehicle Engineering" by S.
Levy and J.P.D.
Wilkinson, McGraw Hill, 1976.
5.
" Dynamics of Structures" R.W. Clough & J. Penzien, McGraw Hill (1975).
6.
R.J.
Fritz, "The Effects of Liquids on the Dynamic Motions of Immersed Solids," Journal of Engineering for Industry, Trans. of the ASME, February 1972, pp. 167-172.
7.
USNRC Regulatory Guide 1.61, Damping Values for Seismic f-)
Design of Nuclear Power Plants, 1973.
v 8.
J.T. Oden, " Mechanics of Elastic Structures," McGraw Hill, N.Y.,
1967.
9.
R.M. Rivello, " Theory and Analysis of Flight Structures,"
McGraw-Hill, N.Y.,
1969.
10.
M.F. Rubinstein, " Matrix Computer Analysis of Structures,"
Prentice-Hall, Englewood Cliffs, N.J.,
1966.
11.
J.S.
Przemienicki, " Theory of Matrix Structural Analysis,"
McGraw-Hill, N.Y.,
1966.
12.
P.
Kuhn, " Stresses in Aircraft and Shell Structures,"
i l
McGraw-Hill, N..Y.,
1956.
13.
S.P. Timoshenko and J.Ne Goodier, " Theory of Elasticity,"
McGraw-Hill, N.Y.,
1970, Chapter 10.
14.
U.S. Nuclear Regu.latory Commission, Standard Review Plan, NUREG-0800 (1981).
15.
U.S. Nuclear Regulatory Commission, Standard Review Plan, Section 3.8.5, Rev.
1, 1981.
O 6-27 1
i 16.
U.S. Nuclear' Regulatory Cermission, Regulatory Guide 1.124,
" Design Limits and Loadixq Combinations for Class 1 Linear-Type Component so;vperts, November 1976.
17.
" Flow Induced Vibratioi" by R.D.
Blevins, VonNostrant (1977).
18.
" Fluid Mechanics" by M.C. Pot te r arid J. F. Foss, Ronald press,
- p. 459 (1975).
O e
9 O
O 6-28
y
[
O[\\
U U
h TABLE 6.5
-o FILE DSCLO1 MODULE A (11 X 11)
COEF.=
.8, FULL RACK I,
' ' ~ ~ - ' '
~ ~ ~ ~ ~
~ ~'
l
~'
'WED, DEC 14E"1983,'11 45'A4 ~~~~' PROGRAM EGELAST - ~ ~~ 7..
~7
'j
'1
~~
~ ~ ~ ~
8*
"1
.Y SCL A2,COFs.S. sos 4% 810HZ,DATAzDSCLO1,11X11 FULL RACK $$E QUAKE
.(L INPUT _ PAP 4 METERS NO. OF NODES (NMMNOD) s 5
q y
Wo. OF ELEMENTS (MUMEL)s 12 PHIFT_0PT10h.(IOP_T)___m_.0._ ___
O X-HALFWIDTH (A2) a 5.713E+01
-+
Y-HALFWIOTH (B2) s 5.713E+01 E L E w. _ M L Fl.W.T H_MZ 1_ s_.2. 12.0ff_q1 m
'd 0
STDESS C0r.FFICIF%TS FOR RACKt
'1 CfX _a_P.900E-03 CFYa P 900F-03 QFZ _s 2.970E-0) 2.702E-04 CMX a 1.406E-04 CMY a 1.406E-04 CTX z 2.702E-04 CTY z n
0 STRFSS__COrFFIC1FNIS FOR SUPPORTS:
' ~ ' '
p.
CFX2 = 9.990E-02 CFY2 = 9.990F-02 CFZ2 s 4.440E-02 u
CMX2 a 1.693F-02 CMY2 = 1.693E-02 CTX2 x 0.000E+00 CTY2s 0.000E+00 of 0
STRESS COEFFICIFNTS FOR SUPP0HT BOTTOM O
STRESS COEFFICIENTS FOR SUPPORTS:
CFX7 = 1,791E-01 CFY2 m 1.791E-01 CFZ2 = 7.960E-02 C*X2 a 1.590E-01 CMY2 = 1.590E-01 CTX2 = 0.000E+00 CTY2s 0.000E+00 e
l 0
STATIC STRESS COEFFICIENTS CCXP,CCXM,CCYP,CCYM,CXSH,CYSHs 0.000F+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00 0.000E+00
~
~ ~
A FTT, FAT,FVT,FOTs 10000.0 18000.0 12000.0 14000.0 E70.rA0,Fv0,Fs0=
30000.0 30000.0 20000.0 30000.0
=
FTR, FAB,FVR,FBas 75000.0 75000.0 55000.0 75000.0 r
SECTION h0 OF HOST (JROOT) = 5
_ TOT A k_Mo, _G E S ECT I O N S_( N U M S EC ) s_,13._,.
4 l
1 TABLE OF wnXIMAX COUIVAI.ENT STRESS l
1 l
Q
-EECT. _jtIME_EGINT WAX,_SEO,__
DJR,STRFSS X=PE40 SFR Y-RENO STRESYLAT.SHERR YLAT.SHFAR NET J_ HEAR) 9 0
NO.
NO.
NO.
(SEMxux)
(S0)
(59%)
(59Y)
(TX)
(TY)
( TP. )
l "l"g
___1 404 __6___2.19AE+03
-6.665E+00__1.627E+01
-7.209E+01
-2.OR3E+02
-4.81er+01
-1.099r+03 2
404 8
4.4R1F+03
-1.333F+01 7.816E+01
-3.005E+02
-4.602E+02
-1.352R+02
-2.241E+03 3
404 8
6.552V+03
-t.909F+01 2.117E+02
-6.603E+02
-6.026E*02
-2.603E+02
-3.274r+03 4
401 8
9.506E+03
-2.666E+01_ _4_,196E+02
-1.095E+03
-6 017E+02
-3.549E+02
-4.247R+03 3
5 404 8
8.513E+03
-2.666E+01 5.194F+02
-1.325E+03
-6.857F+02
-3.549F+n2
-4.247E+03 6
774 8
3.1SdE+04
-1.447E+04 1.010E+e4 1.051E+04 9.9266+03
-0.2dEE+03 9.926E+03 7
396 2
3.528E+04
-1.31Sv+04
-1. 297E.?04 1,12pr+04 1 06dF+04 1.213E+0a
-1.610E+C4 8
746 1
3.556E+04
-1.256E+04
-1.357E+04
-1.097E+04
-1.029E+04 1.271F+04 1.626E+04
't 9
453 4
4.494E+04
-1.392F+04 1.632E+04
-1.281E+04
-1.202F+04
-1.716E+04
-2.063E+04 t _4.2 F + 0_1_ _;1,7 A O F + 0 4 1.66%F+04 2.415E+04 R
i"I to 7]4 1
5,602E+04_
-2,59 3 &;+ 0 4
-2._4 2 6 Ff,0J 4
11 396 4
6.263E+04
-2.35HE+04 1.115E+03
-5.17eE+01
-1.a07F+04
-2.174E+04
-2.696E+04 12 746 3
6.262E+04
-2.251E+04 1.101E+03 6.012F+02 1.94SE+04
-2.279E+04 2.915E+04
h TABLE s (C( _ nuid)
DUTSCL01.DATAR&Clf. ENGR THU, DEC 15, 1F93, 3:24 P1 PACE 10
..1%_
_.453_ _2.__.7.R1.0E+04
-1 496E+04
.t. 320F+03 7_929F+02 2.199r+04 3.077t+04
-3.699E+04 6
1 CONF >ITIONS 4MEU X-DISPLACE 8ENT OT A CO R PIE R TSMAXI4AX Np0E T i VE 3 4 X. CO P N ERJEN.Ty r) t 0 4_ta_ _.C_E N TR O I C A [.
TQP,51 QN A L NO.
X-DISP.
X-DISP.
Y-DISP.
ANGLE
- 1. __74$ _.S.611F*0)
_.-7.J46 M l_.__.5.01RE-0J
-2.600E-01 2
745 7.142F-01
~5.657E-01 4.478E-01
-2.600E-03 3
745 5.61RE-01
-4.161E-01 3.ojog-03
-2.550E-03 4
_744 4.132E-01
-2.659C-01 3.7Y/f=01
-2,5798;-03 (j '1 i
5 697 2.724E-01
-1.147E-01 2.7406-01
-2.760E-03 3.,:
~.
J.'
- 8 cQ8(DlTJQR$_ gB Enf Y-DISPLACEMFNT AT COR NM._1.S M A X I tt 4 X e
.g 400E TI"E MAX. CORNER CENTRQt0AL CENTHOIDAL TORSIONAL
~
l
MD-Y-f!5P.
7-0I50 1.0 T EP_,
A.NG(g _
j' 1
744 9.312E-01
-4.190E-01
-8.107E-01
-2.110E-03
(,
2 R19 7.463E-01
-1.1c7F-01 6.972E-01
-8.Suag-04 l'
1_ _B 2 0_____6. 2 7 2 E - 01_._- 1. 215 f-01__5. 7 R 7 E- 01___- 6. 4 9 5 5-0 4
';j 4
755
- 5. 2 5 4 t.-01
-1.230E-01 3.962E-01
-2.314E-03
{9 5
754 4.39eE-01
-8.450E-02 3.0446-01
-2.366E-01 h.
STRUCTUPA* ACCEPTANCF-AS9E 7:F
- ~
(
SECTION t.UMSEP 1
m #1 31=__.901AT_ TIME _1085 _R2=.. 012dT_TJ'E,.__317 R 1=_. ? 0 EA T___IIM.E 757 y
R4s 0 h T T T I P* E 763 R5=
010AT TIME 756 R6=
.011AT TIME 756 s " " -- "
SECTIOM Nlt h 8 FR 2
co g'1
- g :
R 1 =_ 0 0 2 A T_ T f " F,,j o R 5_,,4 7 =,__. OJ O A T _TJ P E 75^
R3=
037AT TT"F' 757 R4=
026AT T18E 763 45=
.041AT TIVE 7S6 R6=
.04 ALT TIME 756 c-SFCTION PU"AFP 3
R1=_.003AT_ TIME 1045...P2=.
042AT TIMF 756 R3=
086AT 7 t " g_7ji7 R4=
059A7 TIME 764 R5=
.fs90AT TINF 756 R6=
.106AT TIME 756 SECTIO 4 Nilu W R 4
r 8.1= _ 008AT.TI"E__10a5_S2=__
0474T_TTPE 756 P3=
1.97AT_ TIME 757-94=
.10047 Ti6E 764 R5=
.150AT TIME 397 R6=
176AT TIME 397
%ECT104 MM4FR 5
"i 91 =_. 0 0 4 A I T 1
- E._10 5 5 _R 2 =.__. 017 A T_ T I " E__7 %_ _ 3 3 =_.17 9 4I_T I M E 757 l
9 R4=
.122AT TT"E 7C6 R5=
.182AT TIME 397 R6=
.214AT TI"E 397
~ " - '
6l SFCTION NU43ER 6
91= _.447AT T!"F_
775_.R2=__.24 RAT ___T18E.__774 R33 3._4 2 4 LT I'! E 779 j
94=
3504T TI"E 774 P5=
.895AT T1PP.
774 R6=
9684T T18E 774
_.__Ris _. 449AT TIAE. 397 SECTION N U WiER 7
.R2=__.364AT_ TIME __)d3 93=_._.47dAT.TI4C 756 17 R4=
.51447 TIvE 383 R5=
925AT TIME 396 R6= 1.011Af TIPE 396
/ #1 SECTION NU"9ER 8
d R1=_.(19AT,Tj"E_746.,87=_. 32A4T_TJ"f_,_757 p ) =_._,.4 6 3 A T _I T 4 F._,.7 5 7 R4=
.395AT TIME 707 PS=
913AT TIME 746 k6= 1.000AT TIME 746 SECTION NU"9FP 9
R1 =. _. 4 6 4 A T T I M E._ _4 5 3 _. ft 2 =_. 4 2 9 A T_I I M C_4 5 3 R3=
_ ft l l A T _TI ME._ $.5 3
- 4=
429AT *T"P 152 R5= 1.0004T TI."E 453 R6=
.209AT TIME 453
,"g SECTION NU"PER 10 R1=__'.349AT T1"r__775 _ p2= _.167AT_TTPF..77_4_ P)= _.0} RAT TJ y _ 7.74 R4=
033AT TI"E 666 R5=
378AT TT"E 775 R6=
.38 3 AT TTME - 775 SFCTION ffU"BER 11 g-
O t
lE Z.
3 s
g
/
CO U PLIN G ELEMENTS v
44 TYPIC AL FU E L A S S EM B LY 3
GROUP M AS S H
TYPIC AL FUEL RACK MAS S O
2<
FUEL R AC K B A S E 2
AY
=
p
[
l Ax
_L ljwe j
y y
Y 7
/- + - %
- B i
/
i 6
i j
a I
i h
,h FUEL R ACK SUPPORT
/
/ ni X
X8, YB - LOC ATION OF CE N TROI D O F FU EL ROD G ROUP M ASSES - REL ATIVE TO O
CENTER OF FU EL R A C K Di = UNIT VECTORS FIG.
6.1 D Y N A M I C MODEL
__________ - 2 9_ _ _ _ _ _ _ _ _
)
6
O IMPACT d
SPRIN G S
/
n
~
T
[H MASS O
m i*
ut ]
?
F LUID DAMPERS RIGID FR A M E l
i X
1 O
F I G.
6.2 i M PA C T S P R I N G S AND F LUI D D A M PERS s-so
M 5
K, K (typ.)
3 a 4
4 Seismic l Hl MAkW Motions E
M.
~' '
Z
~
~ > > -
\\
Fuel Assembly Group Lumped Mass y
(Typ.)
y s/
$ 5 WW s
Rack Lamped.
3 Mass & Inertia For Hothostal O,
Motions (Typ.)
%s y
I b WN Kr 2
K, (Typ.)
I l
m a
ig-i 4-- TB +
K h
g[
s K
'/
$wv efy-
/t
/
s, K'
l Tt K n A,
O l
Figure 6.3 Spring Mass Simulation For Two Dimensional Motion 6-31 I
O
~
i I, Fy y, j B
L
- 8 u Fx y
(a) TOP VIEW i
z o I
O lPF i
=x
( 3) AXlAL CROSS S ECTION ( B-B )
y u u,,,,,,,
FIG.
6.4 (a) HORIZONTAL CROSS
' SECTION OF RACK
~
(b) VERTICAL CROSS SECTION OF RACK O
6-32
U CELL z(W)
WALLS t
^ l l E A s s'
/ A vlAA)0(PC
' ^$(>(VW W n
b= NyC a :N x Cy
/ $
~
~
C A,-
g, y(y)
A RIGID PLATE g
j rCF o <n, y
e=e
,=
A=-dj, Q
A-e i
/
e
' SUPPORTS
~
FI G.
6.5 oynamic uoeei nMz M
jy a
Y "
fz
/
C A
B O
F G.
6.6 s tress aesuitant s o ri e n t a t io n 6-33
aZ N O DE.1
/
g E L. I SEC.I ~
/
N O D E 2 --*
-EL2 SEC.2 ~
NODE 3 ~
E L,3 O
S E C. 3 ---
N O D E 4 a
- E L.4 f--
S EC 4 -
'/y E L.5
[S E C,6 S E C. 5 ---
/
4_ v'g. NODES g_ yp g
-EL.8 E L,7 -
ROOT OF R ACK
~
S E C. 8 S E C,9 N O.
F ELEMENTS = 8 N O. OF S EC Tl ON S~
=9 N O. 0F N O D ES
=5 O
F G.
6.7 S U B DIVISIO N O F A TYPIC A L R A C K 6-34
4 0
I f-
.m e
~
b
@y O
g g
0
=
=
Fl G.
6.8 FINITE ELEMENTS MODEL CROSS-SECTIGN O
6-35
O O
O
~ V. C. Summer Fuel Pool 3-Horizontal X go-m z
f f
I a
o II I
) l' 1
I i
fh f
/
ll e
i t
li ga i
)
l
< ?-
i O
i0.00 2'.00 4'.00 6'.00 8'.00 l'O.00 i2.00 i'4.00 1'6.00 TIME (SEC)
FIG. 6.9
O O
O V.
C.
Summer Fuel Pool i
Horizontal Y
O 25-
\\
t m
o I
l 1
I Ho I
(
f l
Af i
pi
<tf-1 l
l '
n w
i i
l }
}
gj ll I
.~
on J
Sp:-
l f
0.00 2'.00 4'.00 6'.00 8'.00 l'0.00 1'2.00 1'4.00 l'6.00 TIME (SEC) i FIG. 6.10
O o
O V.
C.
Summer Fuel Pool Vertical
@d-m 6
8
~
ih
\\l
(
I l
\\ \\ \\
\\
t
\\
\\
a u
i
?
5
'o 00 2'.00 4'.00 s'. oo 8 00 s'o.co sh.co s'4.co s's.co FIG. 6.11
SUM OF SUPPORT REACTIONS -
SSE 3
E j
TOTAL PERIOD = 15 SEC.
7
}
e-I I
)
d I'
)
i o
e i
i 0)
^
- i com
);
{$
dy-f f
I I
Z 1
l l
o i
+
j 4
g
-l-
- -(
l 2
g W
i
.e m
T l
0
\\
)
l l
E,.
i N
'O.000 Sb.
ibO.
ih0.
2h0.
2h0.
3b0.
350.
400.
TIME V.
C.
SUMMER FUEL POOL FIG. 6.12
O O
O SUPPORT LEG i - SSE
?
i i
TOTAL PERIDE
=
15 SECONDS 7-
\\
7-
'o l
w f
i i
I 1
S*-
L f
l go a
l 1
\\
b l
(
5 1
H&m g
i
[
I
/
m-L -
- - -V<
tu i
i i
l m
f rf.
I 1
m I
'O.000 56.
100.
150.
200.
250.
3'00.
350.
400.
TIME!
FIG. 6.13 V.
C.
SUMMER FUEL POOL
l
.O:
O O
SUPPORT LEG 2 - SSE - TOTAL PERIOD = 15 SEC.
o e
I i
i E_
a
_.o 1
,1*
o>
I 1
EN j!
(
f 1
l j
gy-l
)
l l
h Y f l-o j
l k
l j
07 t
i i
i W
E l
)
l E,.
I E
j
'0.000 5b.
ibO.
ISO.
260.
250.
3h0.
330.
400.
TIME FIG. 6.14
.V.
C.
SUMMER FUEL POOL
I
\\l
- O 0
0
?
4 l;
l g
0
(
5 C
l 3
E S
1 5
I 0
1 0
3
=
D W
O 1
i 0
I 5
R
(
2 I
E q
P L
O L
O
.E O A
0 T
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SUMMER FUEL POOL
7.
SPENT FUEL POOL STRUCTURAL ANALYSIS
(~)
7.1 Introduction The high density rack modules for long term fuel storage described in Sections 2 and 3 are located in the Spent Fuel Pool of the Fuel Handling Building shown on Figure 7.1.
The Spent Fuel Pool structure is a,
reinforced concrete structure supported on ca.issons down to competent rock and is integrated with the remainder of the building.
Figure 7.2 shows a plan of caisson layout, and Figures 7.3-7.4 show cross sections through the structure.
The pool walls and slab are 6' "d
thick and the caissons are 3'-0" and 4'-0" in diameter.
It has been demonstrated that the existing Spent Fuel Pool structure and foundations maintain their structural integrity for all postulated loading conditions for the new high density racks.
In particular, the requirements of NUREG-0800, Standard Review Plan Section 3.8.4 and the PSAR have been met.
7.2 Analysis Methods The structural analysis methods used for the existing Spent Fuel Pool are described in the FSAR Section 3.8 and the dynamic analysis is in Section 3.7.
Investigation of the existing design and analysis indicates that reasonable margins existed in the walls, slab, and caissons.
This condition allows a simplified conservative approach to be taken for analyzing the structure for increased loads from the new racks and spent fuel.
The models used for the analysis were idealized to produce upper bound results for tension and compression conditions in the structure at any given location.
The walls and slab were analyzed as a continuous two-dimensional frame supported by the caissons and surrounding 7-1
integral concretq floors.
The design of the caissons was re-evaluated to include the effects of the additional loads from the new racks and spent fuel.
)
7.3 Assumptions 1.
The loading used to qualify the pool structure and caissons assumes that all racks are fully loaded and that the loads are evenly distributed.
2.
All caisson supports to the pool slab are considered pinned for analyzing the slab and walls;
- however, moments in the caissons from horizontal forces are considered in evaluating the caisson design.
3.
The seismic forces from the racks are represented by equivalent static loads on the pool slab.
An equivalent static load equal to two times the weight of the racks is used for seismic vertical loads for the Safe Shutdown Earthquake (SSE) condition.
4.
The Maximum Operating Temperature (To) of spent fuel pool is assumed to be 150*.
In section 5, the maximum water temperature is shown to be limited to 137'F (Fig.
5.1.2(a),
p.
5-21) for full core off-load condition.
Even in the normal condition of one (out of two) spent fuel pool cooling loops in operating condition, the maximum temperature is limited to 140*F (Fig. 5.1.3(a),
- p. 5-25).
7.4 Load Combinations In compliance with USNRC Standard Revipw Plan section 3.8.4 the following load combinations were evaluated for the reinforced O
ce crete see e re 1 eeet. trectere =
e.
7-2
()
U = 1.4,D + 1.7L + l JE U= (0.75) (1.4D + 1 7L + 1.9E + 1.7To) l U = D + L + To + E U = D + L + Ta + 1.25E U= (0.75) (1.4D + }.7L + 1.7To)
U = D + L + Ta Where D = Dead Load (Including Hydrostatic)
L
= Live load E
= Operating Basis Earthquake l
E
= Operating Temperature Ta = Accident Temperature,
Note:
The spent fuel pool cooling system is seismic category I
which allows accident temperature (Ta) to be
()
replaced by Operating temperature To in accordance with Standard Review Plan 9.1.3.
7.5 Results 1.
Caisson The results of the caisson evaluation are shown on Table 7.1.
The capacity of the caisson is dependent upon the interaction curve of Bending Moments and Axial Fo r,ce s.
The results show that all the affected caissons have sufficient capacity to sustain the loading for the new rack conditions with further margin available.
f 2.
Walls and Slab l
-The results of the wall and slab evaluation are shown on 7_
l (-)
Table 7.2.
The capacity of the walls and slab is 7-3
p dependent upon the interaction curve of Bending Moments and Membrane Forces.
The results show that the walls and slab have suf ficient capacity to sustain the loading from the new rack conditions.
Considerable reduction in bending due to thermal gradient is possible if a
cracked-section analysis is performed.
However, except in the case of slab north / south with Bending causing tension in the bottom (total bending 352 Kip ft), it was not necessary to utilize reduction due to a cracked section 1n order to show sufficient capacity exists.
7.6 Conclusions Review of the rack seismic analysis results in Section 6 shows that assumption 3
in Section '7.3 of this report was conservative.
Assumption 4,
that the maximum operating temperature To will not exceed 150*F, is confirmed by the cooling analysis result of Section 5.
The results of the circulation of the Spent Fuel Pool walls, slab and caissons show that adequate margin exists in the structure to meet the requirements of NUREG-0800, Standard Review Plan Section 3.8.4 and FSAR requirements for the high density rack modules.
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Table 7.1 Celsson Evaluation - Required vs. Minimum Avellable Capacity i
Affected Celsson REQUIRED CAPACITY AVAILABLE CAPACITY Alloweble bending on Celsson #
Dia.(FT) Compression (KIPS)
Moment (KIP-FT)
Compression (K_lPS)
Moment (KIP-FT)
Rack Sock _ets (KIPS)
~
]
7 4.0 2525 939 2699 1600 3817 8
4.0 2508 939 2699 1600 3817 9
4.0 2400 939 2699 1600 3230 y
10 4.0 2591 939 2699 1600 3230 Ut 16 4.0 1160 939 2699 1600 3230 17 3.0 976 657 1615 675 2257 18 4.0 1804 939 2547 1755 3230 19 3.0 1044 657 1615 675 2257 I
20 3.0 951 657 1615 675 2257 21 4.0 1764 939 2547 1755 3817 4
i 1
In all cases the governing load combination is 1.4D + 1.7L + 1.9E I
4 1
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O Table 7.2t Required vs. Avellable Capacity In Spent Fuel Pool Wells and Slab REQUIRED CAPACITY AVAILABLE CAPACITY Bending Membrane Shear Bending Membrane Shear WORSE CASE LOCATION GOVERNING LOAD COMBINATION (KIP-FT) _ KIPS)
(KIPS) (KIP-FT)
(KlPS)
(KIPS)
(
WALLS Bending Tension on Outside Face (0.75)(1.4D + 1.7L + 1.7T )
913 53.6(Comp) 930 160HComp)
O Bending Tension on inside Face 1.4D + 1.7L + 1.9E 304 23.8(Comp) 449 1665(Comp) y (0.75) (1.4D + 1.7L+1.9E+1.TT )
45.5 77,4 O
m SLAB-NORTH / SOUTH Bending Tension on Bottom (0.75)(1.4D + 1.7L+ 1.9E + 1.7T )
552 42.1(Tens) 67.8 899 168(Tens) 74.8 O
Bending Tension on Top 1.40 + 1.7L + 1.9E 200 26.6(Tens) 444 16d(Tens)
SLAB-EAST / WEST Bending Tension on Bottom (0.75)(1.40 + 1.7L + 1.0E+ 1.7T )
988 617.2(Comp) 1176 1669tComp)
O Bending Tension on Tcp 1.40 + 1.7L + 1.9E 171 3.9(Tens) 61.0 463 168(Tens) 74.6 i Only liml' Ing values of shears and moments are reported.
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m 8.0 ENVIRONMENTAL EVALUATION 8.1 Summary Installation of High Density Spent Fuel Storage Racks at V.C.
Summer Nuclear Station (VCSNS) will increase the licensed storage capacity of the spent fuel from 682 to a maximum of 1276 assemblies.
Radiological consequences of expanding the capacity have been evaluated with thesobjective of determining if there is significant additional on site or off site radiological impact relative to that previously reviewed and evaluated l.
In addition, radiological impact to operating personnel has been evaluated to ensure that exposures remain As Low As Is Reasonably Achievable (ALARA).
The decay heat loading and the radiological burden to the spent fuel pool water are determined almost entirely by refueling V
operations.
The frequency of refueling operations and the conduct of refueling are independent of the increased capacity of the storage pool, except that the increased capacity will reduce fuel movement and allow continued normal operation.
Since the fuel assemblies which will utilize the bulk of the storage capacity (and will ultimately fill all incremental capacity above that of the existing design) are aged, their contribution to either the peak decay-heat load or the increased radiological impact, in terms of increased dose, is negligibly small.
A study 2
performed by the NRC supports this conclusion.
Consequently, the increase irr the storage capacity of the spent fuel pool will neither significantly alter the operating characteristics of the current pool nor result in a measurable change in impact on the environment.
'A 8-1 1
em 8.2 Characteristics of Stored Puel Because of radioactive diecay, the heat generation rate and the intensity of gamma radiaticn from the spent fuel assemblies decreases substantially with cooling time.
After a cooling time of about 4 years,3 the decay heat generation rate is less than 2%
of the rate at 7 days--the nominal time at which depleted fuel assemblies are transferred to the spent fuel pool.
The intensity of gamma radiation is very nearly proportional to the decay heat and decreases with cooling time in a similar manner.
~The bulk of heat loading is due to freshly discharged fuel; moreover, aged fuel contributes very little to the total heat load.
Therefore, it is not expected that this expansion will significantly increase the thermal dissipation to the environment.
Since the ir. tensity of gamma radiation follows the decline in decay heat generation rate, it is similarly concluded
(~
that there would be no significant increase in gamma radiation A
due to the expanded storage.
It is important to note that the aged fuel in the expanded storage capacity will not contain significant amounts of radioactive iodine or short-lived gaseous fission products, since these would have decayed during the refueling period.
The Krypton-85 which might escape from defective fuel assemblies has 2
been shown to do so quickly (i.e.
within a short time after discharge from the core).
Further, the residual Krypton-85 will l
be contained within the fuel pellet matrix and hence any leakage would occur at very low rates 2 Cesium 134/137 2, is strongly bound within the fuel pellet matrix and its dissolution rate in water is extremely small. Any cesium dissolved in the pool water is easily controllable in the clean up system (demineralizer-ion exchanger resin bed) 2 Thus the planned storage expansion will not significantly increase the release of gaseous radionuclides.
l O l
l 8-2 i
8.3 Related Industry Experience
[,)
V Experience with storing spent fuel underwater has been substantial,3,5 These references show that the pool water 2
- activity, normally
- low, during refueling periods experiences a
small increase which decays rapidly with time.
Typical concentrations 4 of radionuclides in spent fuel pool water range f rom 10-4uci/ml or less to 10-2uci/ml with the higher value associated with refueling operations.
References 2 and 4 also state that the increase in pool water activity during refueling can be attributed to:
a.
dislodging (sloughing off) of corrosion products on the fuel assembly during transfer and handling operations b.
the possible short-term exposure of fuel pellets to pool water via a cladding defect, and O.
U c.
mixing of the spent fuel pool water with the higher activity reaccor coolant.
Upon cessation of the refueling operations the fuel pool water and the reactor coolant system would be isolated from each
- other, thereby terminating transport of corrosion products from the Reactor Coolant System.
Thus, deposition of crud is a function of refueling operations and is not impacted by the expanded storage.
Furthermore, it has been shown6 that release of fission products from failed fuel decreases rapidly after s hutdown to essentially negligible levels.
The fuel that have very low pellets are made of inert UO2 solubility in water and the propensity for corrosion of the cladding (Zircaloy 2) at spent fuel pool water 2
temperatures is virtually nil,4 Thus the only mechanism available for the release of the gaseous g
V 8.3
fission products is dlif f usion through the U02 Pellet.
\\
It has been shown that at low water temperatures
(<l50*F) the dif f usion coefficient is extremely small7 Therefore, thie small increase in activity of the spent fuel pooh water is due to either crud transport, fission prcxtucts release, or cross flow from the reactor coolant system and is only a function of refueling operations.
It is reasonable to assume that the increased capacity of the spent fuel pool will reduce fuel handling operations, thereby reducing the probability of increased pool water activity due to crud dislodging.
- Thus, the expansion of fuel pool storage capacity will not cause a significant increase in dose either on site or of,f site.
The corrosion properties of irradiated Zircalloy-2 cladding have been reviewed in References 2 and 5, and the conclusion is drawn that the corrosion of the cladding in spent fuel pool water is negligibly small.
The minor incremental heating of pool water, due to the expansion of storage capacity, is far too small to materially affect the corrosion properties of Zircalloy-2 cladding.
8.4 V.C.
Summer Operating Experience At present there are no stored fuel assemblies in the V.C. Summer fuel pool.
8.5 Spent Fuel Pool Cooling and Cleanup System (FPCC)
It has been shown previously (Section 5 and V.C.
Summer FSAR that the cooling system at the V.C.
Summer Nuclear Station is adequate to handle the expected heat loads and maintain the temperature peaks within acceptable limits.
It has been shown in' 8-4
(~)
\\/
Section 5 that the small increase in heat load due to the storage capacity expansion.will neither significantly increase the thermal dissipation to the environment nor increase the propensity for corrosion of the cladding.
It has also been shown that the crud deposition in the spent fuel pool water occurs during refueling outages and that the planned expansion will not increase crud deposition.
The fuel pool clean-up system (filter and demineralizer) is designed to maintain fuel pool water clarity and is operated and maintained in accordance with V.C.
Summer Nuclear Station operation procedures.
The clean-up system takes a surface skim from the fuel pool an cleans it through a process of filtration and demineralization to prevent crud build-up on the fuel pool walls at the water-to-air interface.
(~N The spent fuel pool water is ampled and analyzed periodically to confirm proper operation of the pool clean-up system.
The frequency of filter and resin replacement is determined primarily by requirements for water clarity rather than the loading of fission product radionuclides.
The fuel pool demineralizer contains 54 cubic feet of mixed bed bead type resin.
The cation resin is a strongly acidic, highly cross linked, sulfonated styrene-divinyl benzene cepolymer.
The anion resin is a strongly basic, quaternary ammonium-poly (styrene-divinyl benzene) resin.
A resin probe energizes a local status light when the demineralizer is empty.
Local pressure ^ gauges on the discharge of the spent fuel purification pump, upstream of the demineralizer and on the inlet of the spent fuel purification filters downstream of the demineralizer, are available for monitoring under administrative control to determine demineralizer resin status.
O v
8-5
)
The fuel pool filter cartridge consists of epoxy impregnated cellulose fiber media and stainless steel hardware, with integral seals.
The filters are monitored for pressure drop with local pressure gauges on inlet and outlet, and for flow with a local flow indicator located downstream of the filters.
The present expected annual quantity of solid radwaste generated by the Spent Fuel Pool Purification System is about 54 ft3 The SFP modification is not expected to result in a significantly higher quantity of solid radwaste.
8.6 Fuel Pool Radiation Shielding Radiation shielding for the spent fuel pool is provided by six foot thick concrete pool walls and by approximately 24 feet of water above the spent fuel storage racks. A thrae-dimensional shielding analysis was performed on the spent fuel pool assuming the pool is filled to capacity with the
(
proposed storage densification arrangement.
This analysis shows that radiation dose levels will be less than 1 mr/hr on the outside of the pool walls and at the pool surface from the stored spent fuel.
This radiation level meets the V.
C.
Summer Nuclear Station design radiation zoning for the fuel handling building.
Thus, it is concluded that the proposed fuel storage modification vill not appreciably increase radiation levels or personnel doses in the fuel handling building and surrounding areas above current estimates.
8.7 Radiological Consequences As stated earlier and confirmed by other studies 2,4,5,6,8,9, it can be shown that there will be no significant increase in activity due to Krypton-85, Cesium 134/137 or crud j
buildup on pool walls.
It is concluded that the incremental impact from the release of either volatile fission products or
()
crud with the expanded capacity of the spent fuel pool will be negligibly small.
8-6
8.8 Reracking Operation O
The existing spent fuel racl;s, in the spent fuel pool, are to be removed prior to the discharge of any spent fuel from the core.
These racks have not been exposed to spent fuel and are only minimally contaminated.
Therefore, it is concluded that significant radiation dose to individuals involved in the reracking is not anticipated.
8.9 Conclusions Based upon the industry experience and evaluations discussed in previous sections, the following conclusions are made.
I l
l o
Minor -increases in radiological burden to the pool water, if any, can be adequately handled by the fuel pool clean up system (filter and demineralizer),
thereby maintaining the radionuclide concentration in the water at an acceptably low level.
(
o No appreciable increase in solid radioactive wastes (i.e.
filter media and demineralizer resin) is anticipated.
o No increase in release of radioactive gases is expected, since any long-lived inert radioactive gas potentially available for release (i.e.,
Kr-85) will
~
have. leaked from the fuel either in the reactor core during operation or during the first few months of residence in the pool. Further, Vol.
1, Ref.
3, (pp.
4-16) has shown airborne activity to be considerably lower than that allowable by Table 1 of 10CFR Part 20, Appendix B.
Therefore, the planned expansion will not significantly increase the release of radioactive gases.
I 8-7
o The existing spent fuel pool cooling system will keep l
the pool water temperature at an acceptable level (see Section 5 - Thermal Hydraulic Considerations).
o The existing radiation protection monitoring systems and program are adequate to detect and warn of any unexpected abnormal increasas in radiation level. This provides sufficient assurance that personnel exposures can be maintained As I.ow As is Reasonably Achievable.
o since the re-racking operations will be performed prior to any spent fuel being placed in the fuel pool, the existing racks are expected to be only minimally contaminated.
Hence removal and disposal of the existing racks will have only minor radiological impact.
o Expanding the storage capacity of the spent fuel pool will not significantly increase the onsite or offsite radiological impact above. that of the~ currently authorized storage capacity, nor is any significant increase in environmental radiological or non-radiological impact anticipated.
e O
O 8-8
1 l
l REFERENCES TO SECTION 8 Cs 1.
"FSAR", V.C. Summer Nuclear Station, Docket No. 50-395.
2.
" Handling and Storage of Spent Light Water Power Reactor Fuel, Vol.
1, Executive Summary and Text, USNRC August 1979.
3.
NUREG 0800, USNRC Standard Review Plan - Branch Technical Position ASB9-2, Rev.
2, July 1981.
4.
A.B.
- Johnson, Jr.,
" Behavior of Spent Nuclear Fuel in Water Pool Storages-BNWL-2256, September 1977.
5.
J.R. Weeks, " Corrosion of Materials in Spent Fuel Storage Pools", BNL-NUREG-2021, July 1977.
6.
J.M.
- Wright,
" Expected Air and Water Activities in the Fuel Storage Canal",
WAPD-PWR-CP 1723, (with addendum) undated.
7.
ANS 5.4 Proposed Standard,
" Method for Calculating the Fractional Release of Volatile Fission Products from Oxide Fuel", American Nuclear Society, issued for review 1981.
O 8.
" Licensing Report on High Density Spent Fuel Racks for U
Ouad Cities Units 1 and 2,",
Docket Nos.
50-254 and 50-265, Commonwealth Edison Company, June 1981.
9.
" Licensing Report for High Density Spent Fuel Storage Racks", Rancho Seco Nuclear Generating Station, Sacramento Municipal Utilities
- District, Docket No.
50-312, June 1982.
i l
e 8-9 O
(
9.
INSERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING MATERIAL 9.1 Program Intent:
A sampling program to verify the integrity of the neutron absorber material employed in the high-density fuel racks in the long-term environment is described in this section.
The program is intended to be conducted in a manner which allows access 'to the representative absorber material samples without disrupting the integrity of the entire fuel storage system.
The program is tailored to evaluate the material in normal use mode, and to forecast future changes using the data base developed.
9.2 Description of Specimens:
The absorber material, henceforth referred to as " poison",
D used in the surveillance program must be representative of the d
material used within the storage system.
It must be of the same composition, produced by the same method, and certified to the same criteria as the production lot poison.
The sample coupon must be of similar thickness as the poison used within the storage system and not less than 5 3/4' x3 inches on a side.
Figure 9.1 shows a typical coupon.
Each poison specimen must be encased in a
stainless steel jacket of an identical alloy to that used in the storage system, formed so as to encase the poison material and fix it in a position and with tolerances similar to that design used for the storage system.
The jacket has to be closed by tack l
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.
The jacket should permit wetting and venting of the specimen similar to the actual rack environment.
t
' O 9-1
(~}
9.3 Test:
The test conditions represent the vented conditions of the box elements.
The samples are to be located in one of the cells of a rack. Eighteen test samples are to be fabricated in accordance with Figure 9.1 and installed in the pool when the racks are installed.
The procedure for fabrication and testing of samples is as given below:
- a. The samples should be cut to size and weighed carefully in milligrams'.
- b. The length, width, and the average thickness of each specimen is to be measured and recorded.
- c. The samples should be fabricated in accordance with Figure 9.1 and installed in a cell (Region I).
- d. Two samples should be rethoved at each time interval according to the schedule shown in Table 9.1.
/7 9.4 specimen Evaluation:
V After the removal of the jacketed poison specimen from the cell at a designated time, a careful evaluation of that specimen should be made to determine its actual condition as well as its apparent durability for continued function.
Separation of the poison frcm the stainless steel specimen jacket must be performed carefully to avoid mechanical damage to the poison specimen.
Immediately after the removal, the specimen and jacket section should visually be examined for any effects of environmental exposure.
Specific attention should be directed to the examination of the stainless steel jacket for any evidence of physical degradation.." Functional evaluation of the poison material can be accomplished by the following measurements:
,O
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9-2
l V)
- a. A ne,utron radiograph of the poison specimen aids in
(~
the determination of the maintenance of uniformity of the boron distribation.
- b. Neutron attensation measurements will allow evaluation of the continued nuclear ef fectiveness of the poison.
Consideration must be given, in the analysis of the attenuation measurements, for the level of accuracy of such measurements as indicated by the degree of repeatability normally observed by the testing agency.
- c. A measurement of the hardness of the poison material will establish the continuance of physical and structural durabiity.
'The hardness acceptability criterion requires that the specimen hardness will not exceed the hardness listed in the qualifying test document for laboratory test specimen irradiated to 1011 rads.
The actual hardness measurement should be made after the specimen has been withdrawn from the pool and allowed to air dry for not less than 48 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 preirradiated sample.
- d. Measurement of the
- length, the
- width, and the average thickness and comparison with the pre-exposure data will indicate dimensional stability within the variation range reported in the Boraflex
.laborat[ory test reports.
A detailed procedure paraphrasing the intent of this program will be prepared for step-by-step execution of the test procedure and interpretation of the test data.
OV 9-3
TABLE 9.1 Time Schedule for Removing Coupons Date Installed INITIAL FINAL WEIGHT PIT WEIGHT WCIGHT CHANGE PENETRATION 2
2 2
SCHEDULE (mg/Cm -Yr)
(mg/Cm -Yr)
(mg/Cm -Yr) mil /Yr 1
90 day "
2 3
4 180 day "
5 l '
6 1 Year O,
8 5 Yearl '
9 10 10 Yearl 11 l
12 15 Year 13 14 20 Year
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15 U
16 30 Year 17 18 40 YearI' 4
O TIME SCHEDULE FOR REMOVING ' COUPONS
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9-4
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FIG. 9.1 -
TEST COUPON 9-5
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10.0 COST / BENEFIT ASSESSMENT A cost / benefit assessment has been prepared in accordance with the requirements of reference 1 Section V,
Part 1.
The purpose of the assessment is to demonstrate that the installation of high-density spent fuel storage racks is the most advantageous means of handling spent
- fuel, considering the needs of our customers for a dependable source of electric power.
The material is presented to satisfy the NRC's need for information; it is the position of SCE&G that no environmental impact statement need be prepared in support of the request, because there will be no significant impact on the human environment.
NRC precedent establishes that alternatives and economic costs need not be discussed when there is no significant environmental impact.
- However, for the sake of completeness, alternatives to re-racking, for additional spent fuel storage capacity, are discussed in Section 10.3.
10.1 Specific Needs for Spent Fuel Storage Disposal of V.C. Summer spent nuclear fuel is scheduled to be carried out by the Department of Energy in or after 1998 in accordance with Public Law 97-425; Nuclear Waste Policy Act of 1982. As V.C.
Summer spent fuel may not be accorded a high priority under the DOE program, SCE&G is seeking to provide a spent fuel storage capacity to support approximately twenty-five years of nominal operation.
No other contractual arrangements exist for the interim
- storage or reprocessing of spent fuel from V.C.
Summer Nuclear Station; therefore, increased storage capacity in the V.C.
Summer fuel pool is the only viable option under ' consideration.
Table 1.1, the fuel discharge schedule, indicates that with the high density spent fuel racks, loss of full core discharge capability (FCDC) will occur in 2008.
O 10-1
L 10.2 Cost of Spen't Fuel Storage The design and manufacture of the spent fuel storage racks will be undertaken by the organizations described in Section 1.
It is expected that the total project cost will be between 1.2 and 1.4 million dollars.
10.3 Alternatives to" Spent Fuel Storaga South Carolina Electric & Gas has considered the various alternatives to the proposed onsite spent fuel storage.
These alternatives:are as follows:
N o
Shipment of fuel to a reprocessing or independent spent fuel storage / disposal facility O
No) commercial spent fuel reprocessing facilities are
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presently operating in the United States.
SCPSA and SCE&G have made contractual arrangements whereby spent nuclear fuel and/or high level nuclear waste will be s
accepted and disposed of by the U.S.
Department of Energy; but such services are not expected to be available before 1998. The V.C.
Summer Nuclear Station existing spent fuel st.orage capacity will not provide full core discharge capability beyond 1993. Spent fuel acceptance and disposal by the Department of Energy is
- not, therefore, an alternative to increased on-site pool storage capacity.
o shipment of fuel to another reactor sitie Shipment of V.C.
Summer Nuclear Station fuel to another reactor site bv l
10-2
could pro, vide short tera relief to the storage capacity V
problem.
However, transshipment of spent fuel merely serves to transfer the problem to another site and does not result in any additional net long-term storage capacity.
Accordingly, SCE&G doen not consider the I
transshipment of spent fuel to be an appropriate alternative to high-density spent fuel storage at the site.
o Not completing the reactor plant /not operating the plant after the current spent fuel storace capacity is exhausted As indicated in NUREG-0575, " Final Environmental Impact Statement on Handling and Storage of Spent Light Water Power Reactor Fuel," the replacement of nuclear power by coal-generating capacity would cause excess mortality to rise from 0.59-1.70 to 15-120 per year for 0.8 GWY(e).
Based on these facts, not operating the plant or shutting down the plant after exhaustion of spent fuel discharge capacity are not viable alternatives to high density storage in the spent fuel pool. The prospective 1983 expenditure of approximately
$1.4 million for the high density racks is small compared to the estimated value of replacement power equivalent to the plant's energy output: approximately S9 million per month in 1983 and between $18.1 and
$22.7 million per month in 1990-1991.
subhect The of the comparative economics. associated with various spent fuel options is the subject of Chapter 6 of NUREG-0575.
Although the material presented is generic, it is of value in comparing the costs of the varioup options. Of the options presented in Chapter 6 of NUREG-05752, high-density spent fuel storage at the site is the most economic option at $18 per 10-3
[a]
KgU.
The price of "Away From Reactor (AFR)" fuel storage, if available, would be $115 per EgU.
This corresponds to 0.5 mill /Kwh from a 1000 MWe power reactor for AFR storage. The marginal cost per KgU of high density spent fuel racks for V.C.
Summer Nuclear Station is SS.
10.4 Resource Commitments The expansion of the V.C.
Summer Nuclear Station spent fuel storage capacity will require the following primary resources:
o Stainless steel - 290,600 pounds o
Boraflex neutron absorber 6900 pounds of which 3200 pounds is Boron Carbide (BqC) powder.
The requirement for stainless steel represents a small fraction of the total domestic production for 1983.3 Although the fraction of domestic production of BgC, required for the fabrication, is somewhat higher than that for stainless steel, it is unlikely that the commitment of EqC to this project will affect other alternatives.
Experience has shown that the production of BqC is highly variable and depends on need, but could easily be expanded to accommodate additional domestic needs.
l i
O 10 4
... - ~
REFERENCES TO SECTION 10 t
1.
B.K. Grimes, "OT Position for Review and Acceptance of Spent Fuel Storages and Handling Applications," April 14, 1978.
i 2..
NUREG-0575, " Handling and Storage of Spent Light Water Power Reactor Fuel", Vol. 1-3, USNRC, August, 1979.
3.
" Mineral Facts and Problems," Bureau of Mines Bulletin 671, 1980.-
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10-5
b 11.
QUALITY ASSURANCE PROGRAM w/
11.1 Introduction This chapter provides a general description of the Quality Assut~ance Program that is implemented to assure that the quality objectives of the contract specification are met.
11.2 General The Quality Assurance Program used on this project is based upon the system' described in Joseph Oat's Nuclear Quality Assurance Manual.
This system is designed to provide a flexible, but a highly controlled system for the design, manufacture and testing of customized components in accordance with various
- Codes, specifications, and regulatory requirements. The Joseph Oat Nuclear Quality Assurance Program has been accepted by ASME and found to be adequate by NRC audit team.
The philosophy behind Oat's Quality Assurance System is that it shall provide for all controls.necessary to fulfill the contract requirements with sufficient simplicity to make it functional on a day to day basis.
As this system is applied to most of the contracts which Joseph Oat obtains, implementation of it is almost second nature to Oat's personnel.
The system readily adapts to dif ferent designs and component configurations, making possible the coastruction of many varied forms of equipment.
The highlights of this system, as addressed in the following paragraphs, provide an overview of the system and how it has been applied to the customer specifications and regulations.
11.3 System Highlights:
The design control is organized to provide for careful review of all contract requirements to extract each individual design and quality criteria.
These criteria are trans}ated into design and quality control documents customized to the contract requirements
/)
and completely reviewed and approved by responsible personnel.
y 11-1
The system for contrdt of purchased material entails generating detailed descript3ws of each individual item of material along with speci fications for any special requirements
\\
such as impact
- testing, corrosion
- testing, monitoring, or witnessing of chemical analysis, provision of overcheck specimens, special treatments or conditioning of material, source inspection, and provision of documentation of performance of any of the above.
Material receipt inspection includes a complete check of all material and its -documentation.
Upon acceptance, each item of material is individually listed on a control sheet issued once a week to assure that only accepted material goes into fabrication.
The fabrication control system provides that a shop traveler is prepared for each subassembly and assembly in each contract.
The traveler is generated specifically to provide step by step instructions for fabrication, inspection,
- testing, cleaning, packaging, etc. which address all standard and special requirements of the contract specifications.
Special attention is given to deployment of fabrication sequence and inspection steps to preclude the possibility of missing poison sheets or incorrect sheets (incorrect B10 loading).
Due to the tendency of contract specifications to require special examination techniques or test procedures, all nondestructive examination procedures and test procedures are custom written to apply to each given component within a contract.
The system provides f o r.
qualificatior, and written certification of personnel performing quality related activities including nondestructive examination and fabrication inspection, welding, engineering, production supervision and auditing.
Other requirements of a solid quality control system are fully covered as specified in the Quality Assurance Manual including 11-2
document control,, control of measuring and test equipment, control of nonconforming material and parts, corrective action auditing and other areas as specified.
11.4 Summary:
Joseph Oat Corporation's Quality Assurance System provides the full measure of quality assurance required by the contract.
All special requirements of the specifications are covered including source inspection of material and witnessing of material testing by the
- Engineer, furnishing of material certifications and test reports within five days of shipment, and obtaining verification of qualification testing of poison materials.
Oat has a long history of providing excellent quality control over a wide range of equipment types such as the high density fuel racks.
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O 11-3
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