ML19350C109

Spent Fuel Pool Mod for Increased Storage Capacity, Licensing Rept on High Density Spent Fuel Racks

ML19350C109
Person / Time
Site:	Quad Cities
Issue date:	03/31/1981
From:	Singh K, Sarah Turner BLACK & VEATCH, COMMONWEALTH EDISON CO., EECBLAVS, JOSEPH OAT CORP.
To:
Shared Package
ML19350C104	List: ML19350C103 ML19350C105 ML19350C109
References
OAT-J-2443, TM-600, NUDOCS 8103300487
Download: ML19350C109 (86)

Text

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I LICENSING REPORT ON i HIGH DENSITY SPENT FUEL RACKS iI FOR I

QUAD CITIES PROJECT OAT J-2443

I COMMONWEALTH EDISON P.O. #249405 l

I BY i

l K.P. SINGH, Ph.D. '

VICE PRESIDENT, ENGINEERING JOSEPH OAT CORPORATION '-

AND l SrANtEr TURNEa, Pa.D.

VICE PRESIDENT SOUTHERN SCIENCE APPLICATIONS, INC.

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REPORT TM-600 I 8/d30oo987

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TABLE OF CONTENTS Page

1. PREFACE 1.1 2 2.1 1.GENERALARRANGEMENT

3. RACK CONSTRUCTION 3.1 References to Section 3 3.3

4. NUCLEAR CRITICALITY ANALYSIS 4.1 4.1 Design Bases. 4.1 4.2 Geometric and Calculational Models 4.3 I 4.2.1 Reference Fuel Assembly 4.2.2 Alternative Fuel Assembly Designs 4.2.3 Calculational Models 4.3 4.3 4.7 4.2.3.1 Analytical Methods 4.7 I 4.2.3.2 Calculational Bias and Uncertainty 4.2.3.3 Trend Analysis 4.2.4 Reference Fuel Storage Cell 4.9 4.9 4.10 4.3 Reference Subcriticality and Mechanical Tolerance Variations 4.11 I 4.3.1 4.3.2 4.3.3 Nominal Case Alternative Fuel Assemblies Boron Loading Variation 4.11 4.11 4.11 4.3.4 Storage Cell Lattice Pitch Variation 4.13 I 4.3.5 4.3.6 Stainless Steel Thickness Variations Fuel Enrichment and Density Variation 4.13 4.15 4.15 1 4.3.7 Boraflex Width Tolerance Variation 4.3.8 Effect of Zirconium Fuel Channel 4.15

! 4.4 Abnormal and Accident Conditions 4.17 4.4.1 Fuel Assembly Positioning in Storage Rack 4.17 4.4.2 Effect of Zirconium Flow Channel Distortion 4.17 4.4.3 Temperature and Water Density Effects 4.18 l 4.4.4 Abnormal Positioning of Fuel Assembly

Outside Storage Rack 4.18 l 4.4.5 Missing Absorber Plate 4.19 4.4.6 Dropped Fuel Asserbly Accident 4.19

, 4.4.7 Fuel Rack Lateral Movemeat 4.21 4.5 Summary 4.22 l

l 3 References 4.24 50 HYDRO-THERMAL CONSIDERATIONS 5.1 5.1 Heat Generation Calculations (Later) 5.1 I

5.2 Analysis of Pool Thermal-Hydraulics 5.2 i

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I I References 5.4

6. SEISMIC 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 6.4 6.2.3 Fluid Coupling 6.5 6.2.4 Damping 6.6 5 6.2.5 Impact 6.7 6.3 Stress Analysis 6.8 I 6.3.1 Stiffness Characteristics 6.3.2 Combined Stresses and Corner Displacements 6.8 6.9 6.4 Time Integration of the Equations of Motion 6.11 6.5 Structural Acceptance Criteria 6.13 References 6.24

7. MISCELLANEOUS ANALYSES 7.1

8. RADIOCHEMICAL ANALYSES (Later) 8.1 8.1 Objective and Assumptions 8.1 8.2 Method of Analysis 8.3 I References 8.4

9. POOL STRUCTURAL CALCULATIONS (Later)

10. NEUTBnN ABSORBER MATERIAL 10.1 l -

10.1 Chemical Composition 10.1 10.2 Physical Properties 10.2

) g 11. IN-SERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABS R ING MATERIAL 11.1 5

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1. PREFACE Quad-Cities Nuclear Power Station consists of two (2) generating units (Unit 1 and Unit 2), each with a General Electric BWR-3 Reactor.

The station is owned by Commonwealth Edison Company (65%) and Iowa-Illinois Gas and Electric Company (25%), and is operated by Common-wealth Edison Company. The two utilities share the electrical output in proportion to the ownership.

The station will loose full core discharge capability at the Fall 1981 Unit 2 refueling outage. A limited number of unused fuel I racks of the original design are available to be installed, which could extend the date for loss of full core discharge capability to the Fall of 1982. No further extension is currently possible.

I Commonwealth Edison Company, in its function as operator, prt-poses to increase the spent fuel storage capacity by replacing the I present spent fuel storage racks with new high density storage racks.

This modification will include use of a neutron absorber material in the racks, as an increase of Keff from 0.90 to 0.95.

The specification for design, construction and quality assur-ance of the high density racks, was prepared by Quadrex, a San Jose based company. The mechanical design, seismic analysis, hydrothermal analysis and other related calculations as well as fabrication of the hardware will be performed by Joseph Oat Corporation. Joseph Oat Cor-I poration, based in Camden, N.J., possesses ASME Code stamps for Section III, Class 1, 2, 3 and MC pressure vessels and components. Oat also has a qualified Q.A. program in accordance with 10CFR 50 Appendix B.

Southern Science Applications, Inc. of Dunedon, Florida, in serving as a consultant to Joseph Oat Corporation in the arm s of criticality analysis and other radionuclide evaluations. Southern Science has a qualified Q.A. System audited and accepted by the Q.A.

division of Joseph Oat Corporation.

I Consulting support on the overall effort is provided by NUS Corporation of Rockville, Maryland.

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2. GENERAL ARRANGEMENT Quad-Cities Units 1 and 2 each possess fuel storage pools 33' wide x 41' long. The high density racks described in this Report are engineered to achieve the dual objective of maximum protection iI against structural loadings (such as ground motion) and maximization of available storage locations. Greater width t< height aspect I ratio provides greater margin against rigid body tipping.

modules are made as large as possible within the constraints of trans-Hence, the j

portation and site handling capabilities.

Quad-Cities Unit 1 pool will contain high density fuel racks in i j 8 different module sizes. The module types are labelled A through H in Figure 2.1, which also shows their relative placement. There will '

be a total of 3,657 storage locations in the Quad-Cities Unit 1 pool.

Quad-Cities Unit 2 pool will contain 3,913 storage cells arranged in 20 racks. There will be 7 different module types in this pool.

Figure 2.2 shows the module placement details with respect to the pool inside boundary.

l Table 2.1 gives the detailed module data, e.g., weight, quantity  ;

and number of storage locations.

All modules are free standing, i.e., they are not anchored to the pool floor or connected to the pool wall through snubbers or lateral I i restraints. The minimum gap between adjacent racks is 3.5" at all locations. Sufficient gap is also maintained between the modules and the pool walls. The minimum gap between the fuel pool wall and '

rack modules is 7h" (ref. Figure 2.1) . Adequate clearance from other -

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pool resident hardware is also provided. Due to the gaps pro-vided, the possibility of inter-rack impact, or rack collision with other pool hardware during the postulated ground motion events is precluded.

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a TABLE 2.1 MODULE DATA  ;

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5 l CELLS CELLS APPROX.

NO. O F, N.S. E.W. WEIGllT

Type QUANTITY CELLS / MODULE DIRECTION DIRECTION LBS/ MODULE CO*1MENTS .

1

! A 12 210 15 14 27,000 Rectangular Modules !

I B 6 196 14 14 25,500 Square Modules

! C 8 182 14 13 17,000 Rectanqular Modules D 2 169 13 13 22,000 Square Modules i

g E 4 120 15 8 15,500 Rectangular Modules ,

" 1' 2 256 16 16 33,300 Square Modules i G 4 224 16 14 29,000 Rect. angular Modules l

II 1 192 16 12 25,000 Rectangular Modules l l

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I I 210 CELLS 210 CELLS CE LS 196 CLLLS 224 CELLS -

,182 (IS x 14) (15 x 14 ) l14 'X l ) (14 x 14) (16 x 14) -

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3. RACK CONSTRUCTION The racks will be constructed from ASTM 240-304, austenitic steel sheet material, ASTM 204-304 austenitic steel plate material, and ASTM 182-F304 austenitic steel forging material. Boraflex, a patented brand name product of Bisco

• will serve as the nuetron absorber material.

The detailed radiological properties of Broaflex may be found in Sec-tion 4 and Section 10.

A typical module will contain storage cells which have 6" minimum I internal cross-sectional opening. These cells will be straight to within 10.125". These dimensions ensure that fuel assemblies with maximum permissible channel deformations can be inserted into the storage cells.

Figure 3.1 shows a horizontal cross-section of an array of 3 x 3 I cells. The cells provide a smooth and continuous surface for lateral contact with the fuel assembly. The construction of the rack modules

, may best be described by exposing the basic building blocks of this design, namely the " cruciform", " ell" and " tee" elements, shown in Figure 3.2. The cruciform element is made of 4 angular sub-elements, "A" (Figure 3.3) with the neutron absorber material tightly sandwiched between the stainless sheets. The cruciform assembly I has 5" high stainless strips, which ensure against slippage of the material downwards due to gravitation loads or operating

" poison" conditions. The fabrication procedure leads to 100% surface contact (in macroscopic sense) between the poison and the stainless sheets.

The top of the cruciform is also end welded using a spacer strip as shown in Figure 3.4. Skip welding at the top ensures proper venting of the sandwiched space in the cruciform spokes.

The " ell" and " tee" elements are constructed similarly using angu-I lar sub-element "B", and flat sub-element "C" (Figure 3.5). Having fabricated the required quantities of the " cruciform", " tees" and

" ells", the assembly is performed in a specially designed fixture I which serves the function of maintaining dimensional accuracy while

• Bisco, a Division of Brand, Inc., 1420 Renaissance Drive, Park Ridge, Illinois 3.1

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I welding all the contiguous spokes of all elements using fillet welds.

Figure 3.6 shows the fillet welds in a 4x4 array. In this manner, the cells are produced which are bonded to each other along their long I edges, thus in effect, forming an " egg-crate".

The bottom ends of the cell walls are welded to the base plate which has 5.25" diameter holes concentric with cell center lines.

Machined sleeve elements are positioned in the base plate and attached ,

I to the base plate through circular fillet welds (Figure 3.7). The conical machined surface on the sleeve provides a contoured seating surface for the " nose" of the fuel assembly. Thus, the contact stresses ,

at the fuel assembly nose bearing surface are minimized.

The central hole in the sleeve provides the coolant flow path for heat transport from the fuel assembly cladding. Lateral holes in the I cell walls (Figure 3.7) provide the redundant flow path in the un-likely event that the main coolant flow path is clogged.

The rack assembly is typically supported on four plate type sup-ports. The supports elevate the module base plate 6.5" above the pool floor level, thus creating the water plenum for coolant flow.

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l REFERENCES TO SECTICN 3 I 1. Oat drawing D-7070, " Fuel Cell Details", Joseph Oat Corporation, Jan, 1981.

2. Oat drawing D-7071, " Elements of Cell", Joseph Oat Corporation, I

Jan., 1981.  !

l I 3. Oat drawing D-7072, " General Arrangement and Assembly", Joseph Oat Corporation, Jan., 1981.

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4. Oat drawing D-7073, " Fixed Support Detail", Joseph Oat Corporation, f Jan., 1981.

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4. NUCLEAR CRITICALITY ANALYSIS 4.1 Desian Bases I The spent fuel storage racks are designed to assure that a k eff equal to or less than 0.95 is maintained with the racks fully loaded with fuel of the highest anticipated reactivity and flooded with unborated water at a temperature corresponding to the highest reactivity. The maximum calculated reactivity includes a margin for uncertainty in reactivity calculations and in mechanical tolerances, statistically combined, such that the true keff "i11 I be equal to or less than 0.95 with a 95% probability at a 95%

confidence level.

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

e General Design Criterion 62 - Prevention of Criticality in e uel Storage and Handling.

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

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NRC Standard Review Plan, Section 3.8.4 and 9.1.2 as they apply to spent fuel racks.

Regulatory Guide 3.41, Validation of Calculational l g Method for Nuclear Criticality Safety (and related l

g ANSI N16.9-1975).

e ANSI N210-1976, Design Objectives for Light Water Reactor Spent Fuel Storage Facilities at Nuclear

' Power Plants.

e ANSI N18.2-1973, Nuclear Safety Criteria for the lI Design of Stationary Pressurized Water Reactor Plants.

The design basis fuel assembly is an 8 x 8 array of fuel rods (BWR type) containing UO 2 at a maximum uniform enrichment of 3.2% U-235 by weight, corresponding to 15.49 grams U-235 per axial I

4.1

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centimeter of fuel assembly. Fuel assemblies containing gado-linium burnable poison or assemblies of other configurations or enrichments, e.g., 7 x 7 array, may also be safely accommodated in I the spent fuel storage racks provided the maximum reactivity is less than or equal to the reactivity of the design basis fuel assembly.

To assure the true reactivity will always be less than the calcu-I lated reactivity, the following conservative assumptions were made:

I e Moderator is pure,unborated water at a temperature corresponding to the highest reactivity.

e Lattice of storage racks is infinite in all di.ec-tions; i.e., no credit is taken for axial or radial neutron leakage.

e Neutron absorption in minor structural members is neglected; i.e., spacers and Inconel springs are replaced by water.

e Pure ::irconium is used for cladding and flow channel; l I i.e., higher neutron absorption of alloying materials in Zircaloy is neglected.

e The spent fuel storage rack will accomodate, with I the required subcriticality, fuel assemblies with maximum expected distortion of the Zr flow channel.

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I 4.2 4.2.1 Geometric and Calculational Models Reference Fuel Assembly I

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I The reference design basis fuel asserbly, illustrated in Fig. 4.1, l

I is an 8 x 8 array of fuel rods with two of the central rods re-placed by zirconium " water-rods." The square Zircaloy channel surrounding the fuel is 0.080 inches thick and 5.433 inches outside dimension. A maximum uniform U-235 enrichment of 3.2%  !

U-235 by weight was assumed as the design basic corresponding to an average axial loading of 15.49 grams U-235 cer axial centimeter in each fuel assembly.

The maximum expected distortion of the Zr channel is illustrated f in Fig. 4.2, which is taken from GE Specification 22A5866, Rev. O.

Since the curved surface of the bulged Zr channel cannot be ade-quately described in the two-dimensional computer codea used for  :

1 I analysis, an approximation, preserving the Zr thickness and weight, was necessary as shown by the dotted lines in Fig. 4.1. This l

should represent a reasonable approximation, since the reactivity effect due to bulging of the Zr channel is small (see Section 4.3.8).

4.2.2 Alternative Fuel Assembly Designs The spent fuel storage racks are also intended to accommodate fuel assemblies consisting of a 7 x 7 array and an 8 x B array, both i

l containing fuel with enrichment less than 3.2% U-235 by weight.

l Specifications for these fuel assemblies and for the reference fuel assembly are listed in Table 4-1, which permits comparisons lg3 1

of the designs. The reactivity of the alternative fuel assemblies is lower than the reactivity of the design basis fuel assembly, primarily because of the lower enrich =ents in the alternative assemblies. Consequently, the design basis fuel assembly is the assembly of highest anticipated reactivity and is the limiting case.

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Table 4-1 FUEL ASSEMBLY DESIGN SPECIFICATIONS Fuel Assembly Jesignation 8 x 8R '

(Reference) 7 x 7/7 x 7R 8x3 r i

Fuel Rod Data Outside diameter, in. 0.483 0.563 0.493 Cladding thickness, in. 0.032 0.032/0.037 0.03, t Cladding material Zr-2 Zr-2 Zr-2 (

Pellet density,iT.D. 95 93/95 95 I l Pellet diameter, in. 0.410 0.488/0.477 0.416  !

Enrichment, wt% U-235 3.2* 2.12/2.30 2.62 Grams U-235 per axial em 15.49 11.25/12.46 13.26  :

I Water Red Data Outside diameter, in. 0.591 -

0.493 i Wall thickness 0.030 -

0.034 i Material Zr-2 -

Zr-2 Number per assembly 2 none 1 l

lI Fuel Assembly Data  !

Number of fuel rods 62 49 63 Fuel rod pitch, in. 0.640 0.738 0.640 l Fuel channel outside dimension, in. 5.438 5.438 5.438 I

• Fuel channel wall l i thickness, in. 0.000 0.080 0.080 Fuel channel material Zr-4 Zr-4 Zr-4 I

I 6 inches of natural uranium at both ends of fuel rod. ,

I 4.6 i

- - - - , , - - , , , _ , _ ~ _ _ , _ _ _ _ , . , _

I

, i lI l lg I m In the present analysis, gadolinium burnable poison is not included in the fuel. However, the spent fuel storage racks can I safely accommodate gadolinium-bearing fuel of higher U-235 enrich-ment than that specified for the design basis, provided the reac-tivity of the fuel assembly is less than or equal to that of the reference design basis fuel assembly. For comparison purposes, j the calculated reactivity (by AMPX-KENO, see Section 4.2.3 below) of the design basis fuel assembly on a 6.00-inch lattice spacing I is 1.362 20.004 (le) with unborated water in the standard reactor  :

core geometry (see Fig. 4.3).

I 4.2.3 Calculational Models l 4.2.3.1 Analytical Methods ,

(

I Nuclear criticality analyses of the high density spent fuel storage rack were performed with the Jc4PX -KENO l 2 computer package, using the 123-group XSDRN cross-section set and the NITAWL sub-  !

routine for U-238 resonance shielding effects (Nordheim integral treatment). A*4PX-KENO has been extensively benchmarked against a number of critical experiments (e.g., Refs. 3, 4, and 5).

I For investigation of small reactivity effects (e.g. , mechanical I tolerances), a four-group diffusion / blackness theory method of analysis (NULIF-CNRM-PDQ7 ) was used (Ref. 5) to calculate small I

incremental reactivity changes. This model has been used pre-viously with good results and is normally used only to evaluate trends and small incremental reactivity effects that would other-wise be lost in the KENO statistical variation. Where possible, i

trends calculc+.ed by AMPX-KENO and by diffusion / blackness theory l were compared and found to be in good agreement, well within the f statistical uncertainty of KENO calculations.

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DIM. INCHE S 0.032 0.410 l

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O 493 R S l T l U 0.380 l 0.591 l 0 531 l 0.030 l V I

1 Fig. 4.3 Typical c7:e configuration of BWR-type fuel assemblics as used in the Guad Cities reactor.

4.8

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4.2.3.2 Calculational Bias and Uncertainty 5

l Results of benchmark calculations on a series of critical j experiments indicate a calculational bias of 0, with an uncer-l tainty of 0.0123 (95% probability at a 95% confidence level).

In addition, a small correction in the calculational bias is necessary to account for the slightly larger gap thickness l (1.1 inches) between fuel assemblies in the Quad Cities spent 1

i fuel rack compared to the corresponding thickness (0.644 inches)

]

in the benchmark critical experiments. Based upon the correla-tion developed in Ref. 5, the correction for water-gap thickness in the Quad Cities spent fuel storage rack is -0.0036 ak (under-prediction). Thus, the net calculational bias is 0.0036 ! 0.0123, including the effect of the water-gap thickness.

4.2.3.3 Trend Analysis Trend analysis of benchmark calculations on critical experiments

- with varying boron content in the absorber plate between fuel assemblies indicates a tendency to overpredict k eff with higher reactivity worth of the boron absorber. In the Quad Cities spent fuel rack, the boron worth is about 40% ak, or s2.7 times the highest boron worth (15.9% ak) in the critical experiments analyzed in Ref. 5. Based upon extrapolation of the trend analysis, AMPX- l KENO calculations of the Quad Cities rack would be expected to I overpredict k gg by an estimated 3.1% ak, including allowance for water-gap thickness. Statistically combining the standard devia- -

5 tion of the regression analysis (t0.0027, la) and a typical standard deviation of the KENO variation of the mean ( 0.005, 10),

I the maximum uncertainty would be :0.0116, including a one-sided 6

tolerance factor of 2.03 (951 probability at a 95% confidence level) for an assumed 60 generations in a KENO calculation. Thus, to the extent extrapolation of the linear regression analysis is valid, the AMPX-KENO calculation of the Quad Cities rack will bc l

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4.9 f I

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high (overprediction) by 0.031 0.012 2k, or a minimum overpre-l diction of 0.019 ak including calculational uncertainty. Although extrapolation of the regression trend much beyond the range of the measurements may be questionable, the analysis does indicate that AMPX-KENO calculations would be expected to overpredict k eff when strong boron absorbers are present. No credit is taken for the expected overprediction other than to indicate an additional level of conservatism in the criticality analysis of the Quad Cities spent fuel storage rack.

. 4.2.4 Reference Fuel Storage Cell l

l The nominal spent fuel storage cell model used in the criticality analyses is shown in Fig. 4.1. The rack is composed of Boraflex absorber material sandwiched between two 0.075-inch s tainless-steel plates. The fuel assemblies are centrally located in each storage cell on a ncminal lattice spacing of 6.22 inches. For two-dimensional X-Y analysis , a zero current (reflecting) boundary l condition was applied in the axial direction and at the center-line through the Boraflex absorber on all four sides of the cell, I

effectively creating an infinite array of storage cells. The Boraflex absorber has a nominal thickness of 0.070 inches and a 2

nominal B-10 areal density of 0.01728 grams B-10 per em .

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_ _ . . - _ . . . _ _ . _ _ . . . . _ . - . _.n___________._

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I I 4.3 Reference Subcriticality and Mechanical Tolerance Variations 4.3.1 Nominal Case (8 x 8 Fuel Assembly of 3.2 wt% U-235)

Under normal conditions, with nominal dimensions, the calculated k is 0.9155 0.0036 (la with 140 generations). For a one-sided tolerance factor of 1.879, corresponding to 95% probability at a l 95% confidence limit with 140 generations, the maximum deviation j of k is 0.0067.

4.3.2 Alternative Fuel Assemblies The alternative 8 x 8 fuel assembly of 2.62 wt% U-235 enrichment I will have an appreciably lower reactivity than the reference 3.2% enriched assembly, because of the lower enrichment. For the 7 x 7 assembly at an assumed enrichment of 2. 8 wt% U-23 5, AMPX-KENO calculations with nominal dimensions yielded a k eff f 0.890 0.005, which is substantially less than that of the ref-erence 8 x 8 fuel assembly. For the enrichments indicated in Table 4-1, the reactivity would be even lower. Thus, the reference I 8 x 8 assembly, with 3.2% U-235 enrichment, is the limiting case.

4.3.3 Boron Loading Variation The Boraflex absorber plate is nominally 0.070 inches thick with 2

a B-10 areal density of 0.01728 g/cm . Manufacturing tolerance limits are 10% in both thickness and boron ccntent. This assures that,at any point where the minimum boron loading (0.0155 5 g B-10/cm )

and minimum Boraflex thickness (0.063 inch) may coincide, the boron areal density will not be less than 0.014 g B-10/cm .

I Calculations were made of k gg with variations in Boraflex absorber loading and thickness. Results of these calculations, given in Fig. 4.4, indicate that the k can be described by the following I

4.11

W M M M M M W M M M) W K mW W W WM M O.96

-In k m = 0.0 6 494 In(B-lO) + 0. 3 519

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.010 . 0 11 .012 .013 .014 .015 .016 .017 .018 .019 .02 0 B-l O , gms/cm 2 Fig. 4.4 Log-log plot of calculated k g versus b-10 loading.

I

.I regression fit (least squares) to the data over the range of B-10 I

I loading from 0.010 to 10.020 g/cm'.

9 2

-In k gg = 0.06494 in (B-10, g/cm ) + 0.3519 l

Within the precision of the calculations, this relationship indi-cates that the 10% tolerance linit on either boron content or l Boraflex thickness results in the same incremental reactivity l change of 0.0063 ak. The trend calculated both by AMPX-KENO and l 8 by diffusion / blackness theory is the same within the analytical l l uncertainty.

I 4.3.4 Storage Cell Lattice Pitch variation The design storage cell lattice spacing between fuel assemblies is 6.220 inches. For manufacturing tolerances of +0.125 or -0.000 inches, increasing the lattice pitch from the minimum 6.220 inches to 6.345 inches (maximum tolerance) reduces reactivity by

0. 0113 ! 0. 006 ak , as calculated by AMPX-KENO or by 0.0094 ak calculated by diffusion / blackness theory. Thus, the nominal case

• • "i'i'" '" ""S**' 'ere ""d '"* "" ' " "*^ 'i"i'Y ' ^ ^"i" I

g pitch increase is negative. A larger increase in lattice pitch j

! produces in even larger negative effect. Results of calculations at several lattice spacings and boron loadings are shown in f Fig. 4.5 in terms of the overall fuel region volume fraction in the spent fuel storage cell (0.6775 for the nominal design).

4.3.5 Stainless Steel Thickness variations I The nominal stainless-steel thickness is 0.075 inches. The I reactivity ef fect of the expected stainless-steel thickness toler-ance variation (10.002 inches) was calculated to be diffusion / blackness theory, since the reactivity increment is too 0.0005 ok by small to be calculated by AMPX-KENO.

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4.14 E

'I I 4.3.6 Fuel Enrichment and Density Variation 1 The design basis enrichment, 3.2 wtt U-235 or 15.49 grams U-235 per axial centimeter in each fuel assembly, defines the f uel I of highes t anticipated reactivity.

ment would result in reduced reactivity. Calculations of the Reductions in U-235 enrich-sensitivity to small enrichment variations by diffusion / blackness theory yielded an average coefficient of 0.0075 ak per 0.1 wt%

U-235 in the range from 3.1 to 3.3 wt % U-235.

Calculations made with the UO 2 fuel density reduced from the maximum of 10.41 g/cm to 10.25 g/cm indicate that the storage rack k is reduced by 0.0002 ak (diffusion / blackness theory).

Thus, in the expected range of UO 2 densities, the reactivity effect is negligible.

4.3.7 Boraflex Width Tolerance Variation The calculational model (Fig. 4.1) uses a Boraflex blade width of 5.86 inches. This is conservative since in the final design of the storage cell, the minimum Boraflex absorber width is nominally  ;

I 5.91 inches, including tolerances. The calculational model thus results in the highest reactivity (0.9155 :0.0036), and the greater width of the actual absorber would further decrease reactivity.

4.3.8 Effect of Zirconium Fuel Channel Elimination of the zirconium fuel channel results in a small decrease in reactivity (-0.0035 ak) as calculated by diffusion /

blackness theory. More significant is a small positive reactivity I effect resulting from bulging of the zirconium channel, which moves the channel wall outward toward the Boraflex absorber. For the maximum expected bulging (to 5.925 inches outside dimc.sion) uni-formly throughout the assembly, an incremental reactivity of I +0.0039 ak would result. as calculated by diffusion / blackness

'I a

4.15

I 1 I theory using the approximate geometric model for the flow channel indicated by the dotted lines in Fig. 4.1. Since actual bulging of the flow channel would not be the maximum everywhere in all assemblies, the reactivity effect can be statistically ccmbined with the reactivity eff ect of other mechanical deviations.

Fuel assembly bowing yields a negative reactivity effect and is treated under abnormal conditions (Section 4.4 below).

l I 4.3.8 Summary of Statistical Variations Calculated reactivity increments from mechanical and fabrication tolerances are summarized in Table 4-2.

Table 4-2 CALCULATED STATISTICAL VARIATIONS IN REACTIVITY (MECHANICAL) l T Case Tolerance Incremental Reactivity, ok Boron concentration 10% ;0.0063 B Boraflex thickness 10% 70.0063 Zero to negative  !

Lattice pitch [00.000 inch SS tolerance 0.002 inch 10.0005 Channel bulge 0.49 inch max +0.0039

• Zero to negative Fuel enrichment and density Boraflex width ** Zero to negative I Statistical average (Root-mean-square of positive increment) 0.0097 '

I

• Design basis enrichment of 3.2% U-235 by weight is the upper i **

limit.

Boraflex width used is conservatively less than the minimum width expected including tolerances.

.i

,.16 E

I l 4.4 Abnormal and Accident Conditions l

I Although credit is permitted for absorptic, by other abscrbers under abnormal conditions, the follcwing evaluations were made I without any additional absorber material in the spent fuel storage pool. Tc the extent any additional absorbers may be present in I the realistic case, the following analyses are even more conserva-tive.

4.4.1 Fuel Assembly Positioning in Storage Rack The fuel assembly is normally located in the center of the storage rack cell with bottom fittings that mechanically prevent lateral movement of the fuel assemblies. Nevertheless, calculations were made with adjacent fuel assemblies (each c sumed to be located on one side of its cell with the zirconium fr- 1 channel touching the g

E SS-Boraflex plate) creating an infinite series of two-assembly clusters separated only by the SS-Boraflex plate. For this case, the calculated reactivity was slightly less than the. nominal design case (by 0.0011 Ak). Calculations were also made with the fuel assembly moved into the corner of the storage rack cell (four-assembly cluster at closest approach), resulting in an even L

larger negative reactivity effect (calculated decrease in kgg of s0.01). With the zirconium fuel channel removed, the reactivity effect of off-se' fuel assenblies is even more negative. Thus, I the nominal ca:.

of the storage i rith the fuel assembly pcsitioned in the center

. ell, yields the maximum reactivity.

4.4.2 Effect of Zirconium Fuel Channel Distortion Consequencer of bulging of the zirconium fuel channel have been treated as a statistical deviation in Section 4.3.7 above. Bowing of the zirconium channel (including fuel rods, see Fig. 4. 2) results in a negative reactivity ef f ect analagous to that of positioning the fuel asserbly toward one side of the s torage cell I

!I 4.17 1I l

I Thus, bewing will result as described in Section 4.4.1 above.

in a reduction in reactivity. ,

i 4.4.3 Temperature and Water Density Effects Decreasing temperature from the nominal 68 F to 39 F (maximum water density) is calculated to increase reactivity by 0.0007 ok, as indicated in Table 4-3 (reactivity effects calculated by i

diffusion / blackness theory). Increasing the water temperature or introducing voids (to simulate boiling) decreases reac tivi ty ,

j as shown in the table.

Table 4-3 EFFECT OF TEMPERATURE AND VOID ON CALCULATED REACTIVITY OF STORAGE RACK Case #oo Comment I 39 F +0.0007 Maximum water density 68 F 0 Reference 104 F -0.004 p (H O) = 0.992 2

176 F -0.013 o(H'0) = 0.972 o -0.020 p(H O) = 0.958 B 212 F 2 212 F with 50% void -0.175 Simulates boiling 4.4.4 Abnormal Positioning of Fuel Assembly Outside Storage Rack Since the storage rack criticality calculations were made assuming an infinite array of storage cells, positioning a fuel assembly outside and adjacent to the actual rack cannot add reactivity; such positioning would result in a keff that is lower than the k I calculated for the infinite array. This has been confirmed by two-dimensional PDQ analysis of finite racks with a new fuel I element positioned outside and adjacent to the rack.

i 1

, 4.18

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I  !

I 4.4.5 Missing Absorber Plato l 2

Should a Beraflex absorber plate be missing frce between fuel as s emblies - i . e . , postulated to be lost by an undefined acci-dent - the reactivity will be slightly higher than the reference case. Calculations performed in two dimensions (PDQ7) indicate I the largest reactivity increment is less than +0.0031 ak due to the loss of a single plate. Because of mesh size limitations in l

PDO7, symmetry considerations (with reflective boundary condi-tions) effectively resulted in the loss of an absorber plate from one side of every 15 storage cells. Thus, the calculated incre-mental reactivity addition due to the loss of an absorber plate should be conservative.

4.4.6 Dropped Fuel Assembly Accident A postulated fuel asserrbly drop into and to the bottom of a storage rack cell results in a configuration that is that of the nominal cell configuration; therefore, a drop will not result in  !

a reactivity greater than that of the nominal design case. l To investigate the possible reactivity eff ects of other postulated drop accidents, calculations were made for unpoisoned assemblies i

I separated only by water. Figure 4.6 shows the results of these  ;

calculations. From these data, the reactivity (k ) will be less than 0.95 for any spacing greater than N8 inches. For a straight drop on top of the rack, an inclined drop or a fuel assembly ,

lying horizcntally on the top of the rack, the minimum separation distance is s9 inches. Maximum expected deformation under seis-mic or accident conditions (see Sections 6 and 7) will not reduce the minimum spacing to less than 8 inches. In addition, the upper 6 inches of fuel is natural uranium oxide, which affords a  !

further effective separation from the higher-enriched active fuel l in the storage racks. Finally, a three-dimensional PDQ analysis, j with a new fuel element immediately above the active fuel in the

)

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4.20 i

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storage rack (neglecting structural material) confirms tha t the reactivity is less than that of the design basis infinite array.

I Fuel assembly drop accidents will not result in an increase in reactivity above that calculated for the infinite nominal design storage rack.

4.4.7 Fuel Rack La teral .'lovement Normally, the individual racks in the spent fuel pool are separa-ted by a water-gap of 1 to 2 inches. For finite fuel racks, this separation would reduce the actual maximum reactivity of i

the racks. Should lateral motion of a fuel rack occur, for what-

! ever reason, closing the gap between racks, the reactivity would, in the limit, only approach the limiting reactivity of the reference infinite array.

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I 4.21 I

I 4.5 Sum: nary I The criticality analyses of the spent fuel storage rack under I normal and abnormal conditions are summarized in Table 4-4.

Table 4-4 SU: DIARY CF CRITICALITY CALCULATIONS i

U' Case 03 00 Comment Normal Conditions k reference 0.9155 Section 4.3.1 I

Calculational bias +0.0036 Includes gap correction l

I Uncertainties Bias Calculational 0.0123 20.0067 Section 4.2.3 Section 4.3.1 I Mechanical 10.0097 0.0170 Section 4.3.8, Table 4-2 S ta tis tical combina tion Total 0.9191 0.0170 Maximum k 0.9361 Abnormal and Accident Conditions Decreased temperature +0.0007 Maximum water density negative l I Increased temperature or void Fuel element positioning negative Fuel channel bowing negative Lost / missing absorber plate +0.0031 Conservative Fuel handling accident negligible l

Lateral rack movement negligible I Thus, ak mum k of 0.936 is conservatively estimated to be the maxi-under the wors t combination of calculational and mechani-cal uncertainties (normal conditions), with a 951 probability at a

, I I 4.22 I

iI i

I 955 confidence level. Under the wors t combination of abnormal and accident conditions, the maximum k g could be as much as 0.940.

Removal of the zirconium fuel channel from all assemblies would reduce the maximum k to 0.933 (normal conditions). Ifthe I trend toward overprediction with bcron worth (Section 4.2.3.3) is valid, the maximum expected k under normal conditions would be 0.905.

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REFERENCES I 1. Green, Lucious, Petrie, Ford, White, Wright, PSR-63/AMPX-1 (code package) , AMPX Modular Code Sys tem for Generating Coupled Multigroup Neutron-Gamma Libraries From ENDF/B, i

ORNL-TM-3706, Oak Ridge National Laboratory, March 1976.

2. L. M. Petrie and N. F. Cross, KENO-IV, An Improved Monte Carlo Criticality Program, ORNL-4938, Oak Ridge National Laboratory, November 1975.

3. S. R. Bierman et al., Critical Separation Between Subcriti-cal Clusters of 4.29 wtt U235 Enriched UO2 Rods in Water I with Fixed Neutron Poisons, NUREG/CR-0073, Battelle Pacific Northwest Laboratoryies, May 1978, with errata sheet issued by the USNRC Augus t 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.

5. S. E. Turner and M. K. Gurley, Benchmark Calculations for Spent Fuel Storage Racks, Report SSA-127, Southern Science Applications, Inc., July 1980.

6. M. G. Natrella, Experimental S tatis tics , National Bureau of I S tandards , Handbook 91, August 1963.

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5. HYDRO-THERMAL CONSIDERATIONS A central objective in the design of the high density fuel rack is to ensure adequate cooling of the fuel assembly cladding. In the following, a brief synopsis of the design basis, method of analysis and computed results is given.  ;

5.1 Heat Generation Calculations:

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5.2 Analysis of Pool Thermal-liydraulics In order to determine an upper bound on the maximum fuel I cladding temperature, a series of conservative assumptions are made. The most important assumptions are listed below:

a. As stated above, the fuel pool will contain spent fuel with varying " time-after-shut-down" t s. Since the heat emission falls off rapidly with increasing t s, it is obviously conservative to assume that all fue' assen-blies are fresh (t s = .100 hours0.00116 days <br />0.0278 hours <br />1.653439e-4 weeks <br />3.805e-5 months <br />), and they all have had I four years of operating time in the Reactor (Ref. 1).

The heat emission rate of each fuel assembly is assumed to be equal, and it can be computed from Ref. (2).

b. As shown in Figures 2.1 and 2.2, the modules occupy an irregular floor space in the pool. For purposes of the hydrothermal analysis, a circle circumscribing the actual rack floor space is drawn. It is further assumed that the cylinder with this circle as its base is packed with fuel assemblies at the nominal pitch of 6.22". (Figure 5.1) .

c. The downcomer space around the rack module group varies, as shown in Figure 5.1. The minimum downcomer gap (715 ")

available in the pool is assumed to be the total gap avail-able around the . aalized cylindrical rack; thus, the maximum resistance to downward flow is incorporated into I the analysis.

No downcomer flow is assumed to exist between the rack I

d.

modules.

I In this manner, a conservative idealized model for the rack assemblage is devised. The water flow is axisym-metric about the vertical axis of the circular rack assemblage, and thus, the flow is two dimensional (axi-syrr.etric three dimensional) . The governing equation to characterize the flow field in the 'ol can now be written. It is shown in Ref. (4) that the resulting integral equation can be solved for the lower plenum velocity field (in the radial direction) and axial 5.2

I I

I velocity (in-cell velocity field), by using the method of collocation. It should be added here that the hydro-dynamic loss coefficients which enter into the formula-I tion of the integral equation are also taken from well recognized sources in the literature; and wherever dis-crepancies in reported values exist, the conservative values are consistently used.

After the axial velocity field is evaluated, it is a straight-forward matter to compute the fuel assembly cladding temperature. The knowledge of the overall flow field enables pinpointing the storage location with the minimum axial flow (i.e: maximum water outlet temperature).

I This is called the most " choked" location. It is recog-nized that these storage locations, where rack modula 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" loca-I tion. Knowing the global plenum velocity field, the revised axial flow through this choked cell can be cal-culated by solving the bernoulli's equation for the flow circuit through this cell. Thus, an absolute upper bound on the water exit temperature and maximum fuel cladding temperature is obtained. It is believed that in view of the preceding assumption, the temperatures calculated in this manner over-estimate the temperature rise that will actually be obtained in the pool.

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I "

IDE ALIZ ED OU T LIN E OF POOL BOUNDARY ,

^ '

IDE ALIZE D OUTLINE RACK ASSE BLY OF R ACK AS SEMBLY ACTU AL OUTLINE /

OF POOL f I 3

-/ /

/

/ /// /'

RACK ASSEMBLY

/l/ / '

/ .

ASS UMED ADDED

{f, ,) FUEL ASSEMBLIES no. s.1 RACK S P A C E E N V E.L PIN G C Y LIN D ER

( QU AD-CITIES STATI ON UNIT I 82)

I

I REFERENCES TO SECTION 5 I 1. FSAR, Quad-Cities, Section 10, Auxiliary and Emergency Systems.

2. Regulatory Commission, Standard Review Plan, Branch Technical I U.S.

Position. APCS 8 9-2 Rev. 1, Nov. 1975.

3. "A Method for Hydro-Thermal Analysis of High Density Fuel Racks",

Oat Standard Document #20, Rev. 1, (1980).

I I

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l l

I I

I I

I 1 I I

5.5

I I

6. SEISMIC ANALYSIS 6.1 Analysis Outline The spent fuel storage racks are seismic category I equip-ment. Thus, in accordance with Ref. (1), they are required to remain functional during and after an SSE (Safe Shutdown Earth-quake). As noted previosly, 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 I (which corresponds to greatest rack inertia), or it may be par-tially loaded so as to produce naimum geometric eccentricity in the structure. The coefficient of friction, p, between the supports and pool floor is another indeterminate factor. According to Rabinowic z [ 2] , the results of 199 tests performed on austenitic stainless plates submerged in water show a mean value of p to I be .503 with a standard deviation of 0.125. The upper and lower bounds (p I 2a) are thus 0.753 and .253, respectively. Two sep-arate analyses are performed for this rack assembly with values I of p equal to 0.2 (lower limit) , and 0.8, respectively.

the following six separate analyses are performed:

In summary,

1. Fully loaded rack (all storage locations occupied);

p = 0.8 (u = coefficient of friction)

2. Fully loaded rack.. u = 0.2

3. Half loaded rack to produce maximum geometric asymetry about I the major dimension of the rectangular rack u = 0.8

4. Half loaded rack to produce maximum geometric asymmetry about the major dimension of the rectangular rack u = 0.2

5. Half loaded rack to produce maximum loading asymmetry about a diagonal y = 0.8

6. Half loaded rack to produce maximum loading asymmetry about a diagonal u = 0.2 f.1

I I The method of analysis employed is the Time History method. The ground acceleration coincidently in three directions is specified by the owner of the power plant.

The object of the seismic analysis is to deternine the struc-tural response (stresses, deformation, rigid body notion, etc.)

I due to simultaneous application of the three orthogonal excita-tions. Thus, recourse to approximate statistical summation tech-niques such as SRSS method (Ref. 3) is avoided and the dependabil-ity of computed results is ensured.

The siesmic analysis is performed in four steps; namely (i) Development of non-linear dynamic model consisting of beam, gaps, spring, damper and inertia coupling ele-I ments; (ii) Derivation and computation of element stiffnesses using a sophisticated elastostatic model; (iii) Layout the equations of motion, decouple these equations and solve them using the " component element time inte-gration" procedure (Ref. 4). Determine nodal forces.

- (iv) Compute the detailed stress field in rack structure using the detailed elastostatic model from the nodal forces calculated in Step III above. Determine if the stress and displacement limits (given in Section 6.5)

I are satisfied.

A brief description of the dynamic model now follows.

'I lI l

lI I 6.2

I I 6.2 Fuel Rack - Fuel Assembly Model 6.2.1 Assumptions

a. The fuel rack metal structure is represented by five lumped masses connected by appropriate elas-tic springs. (Refer to Figure 6.1).

b. The fuel assemblies are represented by five lumped I masses located, relative to the rack, in a manner which simulates either full or partially filled conditions.

c. The fuel rack base is considered as a rigid body supported at four points.

d. The rack base supports may slide or lift off the pool floor, I e. The pool floor is assumed to have a known ground acceloration in three orthogonal directions.

f. Fluid coupling between rack and assemblies, and between rack and adjacent racks is simulated by introducing appropriate inertial coupling into the system kinetic energy.

I g. Potential impacts between rack and assemblies is accounted for by appropriate spring gap connectors between masses involved.

h. Fluid damping between rack and assemblies, and between rack and adjacent rack is simulated by inclusion of appropriate equivalent linear damping.

i. The supports are modeled as rigid beams for dynamic analysis. The bcttom of the support legs is attach-ed to a frictional spring as described in Section 6.2.2. The elastic properties of the support I

g B beams are derived and used in the nnal computa-tions to determine support leg stresses.

j. The effect of sloshing is shown to be negligible and is hence neglected. It is to be noted that the top of the rack is over 20' below the free water surface.

I 6.3

I I 6.2.2 Model Description The absolute degrees of freedom associated with each of the mass locations i, i* is as follows (Figure 6.1).

I LOCATION ux DISPLACEMENT uy uz 6x ROTA ION y Oz (NODE)

I 1 P1 P2 P3 94 95 96 1* Point is assumed fixed to base at XB,YB,Z=0 2 P7 Pg q11 gl2 I 2* P 8

P 10 9 9 1 13 15 17 18 3* P y4 P 16 4 E P 19 21 923 924 4* P E 20 22 l

5 P P P 9 29 9 30 9 31 25 27 32 L

5* P E 28 I 26 Thus, there are 32 degrees of freedom in the system.

l Note that elastic motion of the rack in extension is represented by generalized coordinates P3 and P32-i I This is due to the rela'tively high axial rigidity of the rack. Torsional motion of the rack relative to its base is governed by q31' I A schematic description of the rack supports is given I in Figure 6.2. The members joining nodes 1 to 2, 2 to 3, etc., are beam elements with deflection due to bend-ing and shear capability (Ref. 4, pp 156-161). The elements of the stiffness matrix of these beam ele-ments are readily computed if the effective flexure 1I

I I modulus, torsion modulus, etc. for the rack struc-ture are known. These coefficients follow from the elastostatic model as described later. The node points i* (i = 1,2 5) denote the cumulative I

mass for all the fuel assemblies distributed at 5 elevations. Referring to G.E. specification (Ref. 5), the bending and torsional stiffnesses of the fuel assembly (channeled or unchanneled) are sev-eral orders of magnitude smaller than the rack beam elements. Hence, it is reasonable to neglect the spring elements joining these lumped masses. In order to demonstrate that fuel assembly structural springs can be disregarded to produce conservative results, the case (refer to Section 6.1) which yields maximum rack primary stress is also run with beam I - springs connecting fuel assembly lumped masses.

results are available in Ref. (7). The nodes i The are located at X = XB,Y=YB in the global c:oordinate system shown in Figure 6.1. Y The coordinates (XB' B) are determined by the center-of-mass of the set of fuel assemblies. For a completely loaded rack XB*YB = 0.

6.2.3 Fluid Coupling An effect of some significance requiring careful I modeling is the so-called " fluid coupling effect".

If one body of mass my vibrates adjacent to another body (mass m ), and both bodies are submerged in a 2

frictionless fluid medium, then the Newton's equation of motion for the two bodies have the form (mi + M11) X1-M12 X2 = applied forces on mass m1(6.1)

-M 21 1+ I"2 + M22} 2 = applied forces on mass M2 M11, M12, M21 and M22 are fluid coupling coefficients which depend on the shapes of the two bodies, their relative disposition; etc. Fritz (5) gives data for I M g3 for various body shape and arrangements. It is to 6.5

I I

be noted that the form of Eq. (6.1) indicates that effect of the fluid is to add a certain amount of mass to the body (M 11 to body 1), and an external force which is proportional to the acceleration of the adjacent body (mass m2). Thus, the acceleration I of one body affects the force field on another.

force is a strong function of the inter-body gap, This reaching large values for very small gaps. This I inertial coupling is called fluid coupling.

an important effect in rack dynamics.

It has 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 as described in detail in Reference (6). The fluid coupling is be-I tween nodes i and i (i = 2, 3 ... 5) in Figure 6.1.

Furthermore, nodal masses i are coupled to the refer-ence frame through inertial coupling coefficients.

Finally, virtual mass is included in vertical direc-tion vibration equations of the rack; and virtual inertia is added to the governing equations corres-ponding to rotational degrees of freedom, such as q4, q5' 9 6 ' 911, etc.

6.2.4 Damping In reality, damping to the rack motion arises from material hysteresis (material damping), relative inter-component motion in structures (structural damping),

! and fluid drag effects (fluid damping). (Ref. 17)

Only fluid damping is included in the analysis. The jI fluid damping acts on the i nodal masses, as well as on i nodal masses. The equivalent values of linear l

j dampers for various types of motions are derived in Ref. (7). An analysis of rack stresses in the ab-sence of fluid damping is also performed to obtain an understanding of the contribution of damping in abating stresses and displacements.

6.6

e

! 6.2.5 Impact The fuel assembly nodes i will impace the corres-lI ponding structural mass node i. To simulate this impact, 4 impact springs around each fuel assembly node are provided (Figure 6.3). The fluid dampers are also provided in para:lel to the springs. The spring constant of the springs is equal to the local stiffness of the vertical panel computed by evaluat-ing the deflection of a 6" diameter circular plate

(.075") uniformly loaded and built in around the edge. The spring constant calculated in this manner I should provide an upper bound on the local stiffness of the structure.

A brief description of the elastostatic model now follows.

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6.7

I 6.3 Stress Analysis 6.3.1 Stiffness Characteristics The fuel rack is a multi-cell, folded-plate struc-ture which has what is colloquially called an " egg-crate" configuration. This type of construction is I very similar to the so-called " stressed-skin" con-struction of ribs spars and cover plates which are widely used in aircraft construction. Techniques developed in the field of aircraft structural analy-sis are utilized herein to find the stresses and de-formations in such structures. These methods have been thoroughly tested and their reliability has been documented in a number of well-known publications (e.g. Ref. 8 thru 12).

Figure 6.4 shows two cross-sections of the fuel rack which is modeled as a rectangular network of plates interconnected along nodal lines shown as points in i

Fig. 1-A. An arbitrary load with components Fx ,

i Fy i , Fz acts as an arbitrary elevation on one of the nodal lines. We find the displacements and stresses due to such a typical load according to I the stressed skin model as follows: ,

The torsional deformations are solved for by using I the classical theory of torsion for multi-celled, thin-walled cross-sections (Ref. 13).

The bending deofmration is found by using the theory of shear flow (Ref. 12) wherein all axial stresses are carried by the effective flanges (or stringers) formed by the intersections of the plates and all transverse shears are carried by the plates modeled as shear panels.

From a knowledge of the shear flows, the bending and torsional deformations, it is possible to provide a set of influence functions or the following section properties for the fuel rack as a whole:

6.8

I l (EI)eq = Bending rigidity (in two places)

(GJ)eq = Torsional rigidity (AE)eq = Extensional rigidity k = Shear deformation coefficient s

such properties are used for the dynamic analysis of seismic loads. The detailed equations are documented in Ref. (7).

I 6.3.2 Combined Stresses and Corner Displacements The cross-sectional properties and the Timoshenko shear correction factor calculated in the previous section are red into a dynamic analysis of the sys-I tme shown in Figure 6.5 with a specified ground mo-tion simulating earthquake loading. From the dynamic analysis, the stress resultants (Fx, Fy, Fz, Mx, My, M)z acting as shown in Figure 6.6 are computed for a large number of times t = at, 2 at ... etc, at a selected number of cross sections. The displacements (Ux, Uy, Ug ) at selected nodal points on the z axis I are also provided by the dynamic analysis as well as rotations (0 x, O y, Og) of the cross-sections at the I nodes.

Figure 6.7 shows a typical sub-division of the struc-ture into elements, nodes and sections. The stresses are calculated at all sections and the displacements at all four corners of the rack are calculated at

these elevations.

Since o varies linearly over the cross-section and achleves its extreme values at one of the four cor-ners of the rack, the shear stresses due to torsion-al loads (M z) achieve their extreme values near the middle of each side. The shear stresses due to lateral forces (Fx, F )ywill achieve their extreme values at the center of the cross section or at the I

6.9 L

I middle of each side (sae Ref. 7). Thus, candidates for the most critical point on any section will be the points labelled 1, 2 ... 9 in Figure 6.8. The expression for the combined stress and kinematic dis-placement for each of these points is written out.

I Similarly the stresses in the support legs are eval-uated.

An Oat proprietary computer program "EGELAST" com-putes the stresses at the candidate points in each level. It sorts out the most stressed location in space as well as time. The highest stress, and maxi-mum kinematic displacement are thus readily found.

i t

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I I eo

6.4 Time Integration of the Equations of Motion Having assembled the structural model, the dynamic equa-tions of motion corresponding to each degree of freedom can be written by using Newton's second law of motion; or using Lagrange's equation. For example, the motion of Node 2 in x-direction (governed by the generalized coordinate p7) is written as follows:

The inertial mass is "21 + A211 + B211 where m 21 is the mass of node 2 for x-directional motion.

A 211 is the fluid coupling mass due to interaction with I node 2 .

B 211 is the fluid coupling mass due to interaction of node~

2 with the reference frame (interaction between adjacent racks).

Hence Newton's law gives (m21 + A211 + B211) [7 + A212 E8 + B212 u = 02 where Q2 represents all the beam spring and damper forces I on node 2, and A of node 2 ; and B 212 is the cross term fluid coupling effect 212 is the cross term fluid coupling effect of the adjacent racks. 'u' represents the ground accelera-tion.

Let i 97=P7-u i.e: q7 is the relative displacement of node 2 in x-direc-with respect to the ground. Substituting in the above l equation, and rearranging, we have (m21+ A211+ B211) 5'7 + A212 D'8 = 02- ("21 + ^211 + B211

+A212 + B212 I" Similar equation for each one of the 32 degrees of freedom can be written out. The system of equations can be repre-sented in matrix notation as:

6.11 L

I I

[M] {'q'} =

[Q] + {Gl where the vector [Q] is a function of nodal displacement and velocities, and {G} depends on the coupling inertias and the ground acceleration.

~

Pre-multiplying above equation by [M) renders the re-sulting equations uncoupled in mass.

We have:

~

{'q') = [M] ~ [Q] + [M] {G)

This equation set is mass uncoupled, displacement coupled; and is ideally suited for numerical solution using the cen-tral difference scheme. The computer program developed I by G.E. and described in Ref. (4) performs this task in an efficient manner. This computer program, named "DYNAHIS" in Oat's computer program library is documented in Ref.

(4), and also internally at Oat.

Having determined the internal forces as a function of time, the computer program "EGELAST" computes the detailed stress and displacement field for the rack structure as described in the preceding section.

lI lI i

iI e.12 g

L

I I i Structural Acceptance Criteria I 6.5 There are two sets of criteria to be satisfied by the rack modules:

(a) Kinematic Criteria: This criterion seeks to ensure that adjacent racks will not impact during SSE (condition E' in Ref. 14), assuming the lower bound value of the pool surface I friction coefficient. It is further required that the fac-tors of safety against tilting specified in Raf. (15) are met (1.5 for OBE, 1.1 for SSE).

(b) Stress Limits: The stress limits of the ASME Code.

with latest Addendum (1) Criteria : Section III, Sub-Section NF, 1980 EditionAwere chosen to be met, since this Code provides the most consistent set of limits for various stress types, and various loading conditions. The following loading cases (taken out of the set specified in Ref. (14) are meaningful.

~

SRP Designation ASME Designation (i) D+L Level A (normal condition) 1 (ii) D+L+E Level B (upset condition)

(iii) D+L+T g No ASME Designation. Primary mem-I brane plus bending stress required to be limited to lesser of 2 S t and Siu I (iv) D+L+T + E No ASME Designation. Stress limit same as (iii) above No ASME Designation. Stress limit (v) D+L+Ta+E same as above (vi) D+L+T a + E' Level D (faulted condition) where D: Dead weight induced stresses L: Live load induced stresses E: O.B.E. (Time history loading)

E': S.S.E.

I To: Stresses due to assymmetric heat emission from the fuel assemblies Thermal stresses due to postulated high energy pipe Ta:

break vield stress of the material, S u: ultimate rtress g .S,:

I 6.13

I I The conditions T 3 and T o cause local thermal stresses to be produced. The worst situation will be obtained when an isolated stor-I age location has a fuel assembly which is generating heat at the maxi-mum postulated rate. The surrounding storage locations are assumed to contain no fuel. Furthermore, the loaded storage location is assumed to have unchanneled fuel. Thus, the heated water makes unobstructed contact with the inside of the storage walls thereby producing maxi-

.. mum possible temperature difference between the adjacent cells. The

' secondary stresses thus produced are limited to the body of the rack.

1 i.e., the support legs do not experience the secondary (thermal)

(2) Basic Data: The following data on the physical proper-ties of the rack material are obtained from the ASME Code,Section III, appendices.

TABLE 6.1 PHYSICAL PROPERTY DATA Property : Young's Yield Ultimate Allowable l l Modules Strength Strength Stress i E S S S y

Value 28.3x10 25 KSI 71 KSI 17.8 Psi KSI Reference ~ Table Table Table Table I-6.0 I-2.2 I-3.2 I-7.2 l

_ _.' i (3) Stress limits for normal, upset and faulted conditions:

The following limits are obtained from NF-3230 in conjunc-I tion with Appendix XVII as modified by the USNRC Regulatory Guide 1.124.

(3.1) Normal and upset conditions (level A or level B).

I (i) Allowable stress in tension on a net section =

F= .6S y or F t

• t Ft is equivalent to primary membrane stresses tEvaluated at 2000F. This temperature is higher than the pool water bulk temperature under any of the loading conditions under considera-tion.

I e.14

,g

I I (ii) On the gross section, allowable stress in shear is F = 0.4 S v y

= (0.4)(25000) = 10,000 Psi (iii) Allowable stress in compression, Fa a

, 3_(h) / 2C c S

v kl d-(5)+

J 37kl) - (7) 8C c @_

I where I kl 7 = the largest effective slenderness ratio

= = 147.81 Ce ( )

Y Substituting ntunbers, we obtain, for both 7upport leg and " egg crate" region:

Fa = 15000 Psi

! (iv) Maximum bending stress at the outermost fiber due to flexure about one plane of symmetry:

F b= .60S y = 15000 Psi (v) Combined flexure and compression:

l C my f by g f a, , C mx b fx, 3y lE Fa D x Fbx D y r by where l

1 I f:a Direct compressive streas in the section fbx: Maximum flexural stress x-axis f Maximum flexural stress y-axis l

by:

Cmx =C my = 0.85 Dx = 1 Ffx I

g e.1e .

I I

I D y =1-f where 2

E 7,x e , 12n I 23(ki b)2 rb (vi) Combined flexure and compression (or tension)

I I

a

.6 S y

+h+

rbx by 5 1,0 The above requirement should be met for both direct I tension or compression case.

I (3.2) Faul';ed Condition:

F-1370 (SectionII$,AppendixF), states that the limits for the faulted condition are I 12 (

S

) times the corresponding limits for normal condition. Thus the multiplica-tion factor is 0 )

= 2.0 Factor = (1. 2) (1 l

(3.3) Thermal Stresses:

l There are no stress limits for thermal (self-limit-ing) stresses in Class 3-NF Structures for linear j type supports.

l

! However, the range of primary and secondary stress intensity is required to be limited to 3 Sm in the manner of class 1 components. Sm is the allow-I able stress intensity of the rack material at the maximum operating temperature.

I I

e 1e

.I

I l I d l Z da 3 h . ~l ) ~l

/

C O U PLIN G ELEMENTS I TYPICAL FU EL ASS EM B LY 3 GROUP M AS S g

H TYPICAL FUEL l RACK MASS 2 FUEL R ACK B A SE l =

2 Ay, =

/

1 / 1 i

" Ax "' Ys l Y l ,

1p 7

• 2 l , ,

, m 3 .

8

' I h

/S S '

l FUEL R ACK SUPPORT l

( X X8, YB - LOCATION OF CE N TROI D O F FU EL l ROD GROUP M ASSES - RELATIVE TO CENTER OF FU EL R A C K g

ni = UNIT VECTORS I FI G. 6.1 DYN A MIC M O DEL l s.17 -

I I -

CELLS MODULE l

B ASE PL ATE }~ ~ ~ ^l7~' i I ) J :l

<ms) ;5 EEnd67_:JE-iribvi l u

.y ,

y d I y --

d= i,,

~

6/2 1 V r 7 h

, l/' " d I v 2 n

xxxxgxxxxxxxxy 8

i = 1 5 '/ e4 S Q. -

I -

JANNN NN , l\\\\\NL I

h

/ h

/ /

I ,, d ' [Z g i 3 '/4 - -

S e.

g/ s/

I //L i J /l

/ "

l (T Y P.)

F _VN\\\\\1 NNNNNWF_~

I i I era. 6.2 SUPPORT

'I 6.18

_.q I

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I I Y U

IMPACT SPRIN G S I

I ,

E n$l{S d .

=:

,,e7 ,_

l DAMPERS I RIGID FR AM E l

,I l

= X

}

I F l G. 6.3 l M PACT SPRI N GS AND F LUI D DAMPERS g

I e .1e

I I

I I ya j, F I B L  ;  ; B Fx I "X .

(a) TOP VIEW g

l Zn I u , 'Fz .

lI = Fh l (b) AXIAL CROSS S ECTI ON ( B-B )

i

/,usin u////

I.

FIG.6.4 (a) HO-RIZONTAL CROSS

'g SECTION OF RACK

'I (b) VERTICAL CROSS SECTION OF RACK

,I I -

I

,I e.20

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CELL Z(W) 3 l WALLS

!g

,% :::ffC  :::

/d^^^^2^f= s= n, c l a 4,

C' a:NxC7 c

'I - -

I , y(V)

A RIGIO PLATE

g j I s~/

i I

p- p s..x ,

p,'

wQ F//

= x cu)

I

/

a' SUPPORTS 1 l n uz FIG.6 5

'g a j uy 6rF a ,

(

'I g/ y,=ux

=

,C I A e

,I I F I G. 6.6

I 6.21

I I

I

'I I hZ NODEI I /y E L.I l S E C. ! -- -_.-

N O D E 2 ---- EL.2 SEC.2  : --

g N O D E 3 --.- EL3 l S E C. 3 - --

NODE 4 - :

e ag EC.4 -- .

. , E L.5

, - ,a - S EC.5

. NODE 5 J-e__ g '-

S E C. 6 g 4 - -. -

EL7- ~

gR O OT O F R'A CK' , -EL.8 y

ASEC 8 -

S E C. 9 N O. O F E L E M E N T S = 8 l

N O. OF S EC TI'ON S' =9 N O. 'O F N O'D E S =5

.I l F I G. 6.7 I

e.22 g

I l l

I l

I I

.y b @

I 8

e g @x I @

I ' ,

I @ @

= 0 -

I I

I FI G. 6 8 I

I I .

g e.23

References to Section 6

1. Regulatory Guide 1.29, Seismic Design Classification, Rev. 2, Feb. 1976.

2. " Friction Coefficients of Water Lubricated Stainless Steels for a Spent Fuel Rack Facility", by Prof. Ernest Rabinowicz, M.I.T., a report for Boston Edison Company.

3. Regulatory Guide 1.92, Combining Modal Rasponses and Spatial Components in Seismic Response Analysis, Rev. 1, Feb. 1976.

Regulatory Guide 1.61, Damping Values for Seismic Design of Nuclear Power Plants, Oct. 1973.

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. General Electric specification 22A5866, Rev. 1, Appendix II, Fuel Assembly Structural Characteristics.

6. R.J. Fritz, "The Effect of Liquids on the Dynamic Motions of Immersed Solids", Journal of Engineering"for Industry, Trans.

of the ASME, Feb. 1972, pp. 167-173.

7. " Seismic Analysis of Quad Cities High Density Racks, Oat report ,

No. (later) .

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, Eaglewood Cliffs, N.J., 1966.

. 11. J.S. Przemienicki, Theory of Matrix Structural Analysis, McGraw-Hill, N.Y., 1966.

12. P. Kuhn, Stresses in Aircraft and Shell Structures, McGraw-Hill, N.Y., 1956.

13. S.P. Timoshenko and J.N. Goodier,' Theory of Elasticity, McGraw-Hill, N.Y., 1970, Chap. 10.

, I 6.24 y

l 14. U.S. Nuclear Regulatory Commission, Standard Review Plan, NUREG-75/087, Section 3.8.4.

l

15. SRP NUREG-75/087, Section 3.8.5.

16. NRC Regulatory Guide 1.124.

f I '

I 1 i

l

<I l

'I I

I .

,I .

'E I

I I

I I

I .

- - - ._ . _ . _ _ . . _ . . _ _ _ . _ _ _ _ . _ _ . _ . _ _ _ _ _ . _ _ _ _ . _ _ . . __ ___,_. . _ _ _ _ _ _ _ _ _ . _. _ .i

I I 7. MISCELLANEOUS ANALYSES In addition to the ground motion analyses, the following mechanical loads are analyzed:

a. Dropped Fuel Accident I A fuel assembly (weight - 600 lbs.) dropping from 36" above a storage location and impacting the base. Local failure of the I base plate is acceptable; however, a substantial impact with the pool liner is not allowed. The sub-criticality of the adjacent fuel assemblies is not be violated.

b. Dropped Fuel Accident II One fuel assembly dropped from 36" above the rack and hits top of rack. Permanent deformation of the rack is allowed but is required to ce limited to the top region such that the rack cross-sectional geometry at the level of the top of the active fuel hnd below) is not alter.ed.

c. Jammed Fuel Handling Equipment and Horizontal Force I A 2000 lb. uplift force and 1000 lb. horizontal force applied at the top of rack at the " weakest" storage location. The force is assumed to be applied on one wall of the storage cell boundary as an upward shear force. The damage, if any, is re-quired to ba limited to the region above the top of the active fuel.

The above loading conditions are analyzed to determine an upper bound on the plastic deformation zones. It is shown that the plastic deformation is limited to the rack structure well removed from the active fuel regions. Thus the suberiticality of the fuel arrays is not modified or violated.

l l

l I

I I

7.1

I I 10. NEUTRON ABSORBER MATERIAL The material utilized for neutron attenuation in the racks is Boraflex; a proprietary product of Bisco, a Division of Brand Industrial Services. This material is available in sheet form I which facilitates easy handling and a close control of lateral dimensions during fabrication. This material has found wide-spread acceptance due to its durability, and a remarkable reten-tion of physical and mechanical properties when subject to high or low flux irradiation under typical fuel pool environments.

A brief resume'of the established information on this material is given in the following:

10.1 Chemical Composition The elemental composition of the Boraflex proposed can be divided into two catagories, the polymeric matrix system and the boron carbide power. The element.al composition of I each to the nearest 0.5 wt. % is listed below:

TABLE I

- Elemental Composition of Boraflex Components by Weight %

ELEMENT POLYMER B4C Silicon 411 -

Oxygen 37%

Hydrogen 4.5% -

l Carbon 17.5% 23.5%

l Boron - 76%

l Iron; soluble borons - 0.5%

1 2 The minimum B loading is 0.014 grams /cm ac a nominal thick-ness of .070". The criteria suggests a formulation based on 42 wt. % boron carbide to assure that the specified B con-tent is exceeded at the minimum acceptable manufacturing tolerance thickness (I 10% typical, t .010" maximum). The l elemental content of Boraflex based on this formulation would l be as follows:

I 10.1

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I I TABLE II Elemental Composition of Boraflex Containing 42 wt. %BC 3 (by wt. %)

Silicone 24.0%

0xygen 21.5%

Hydrogen 2.5%

Carbon 20.0%

Boron 32.0%

Iron, soluble boron - trace I Note that the isotopic B 10 content expressed as wt. % of total boron is typically 18.0 1 .4.

I 10.2 Physical Properties Boraflex has been extensively tested for physical and mechanical characteristics when subjected to high and low rate irradiation while contained in air, deionized water I or borated water environments. Careful laboratory data on ,

neutron attenuation, elemental boron leaching, residual activity, gas generation, etc. were also taken and docu-mented. Bisco report 748-10-1 contains detailed description of the procedures and recorded results. It is shown that the exposure of boraflex in air to 2.81 x 10 rads gamma from a spent fuel source results in no significant physical I changes nor in the generation of any gas. Irradiation to the level 1.03 x 10 11 rads gamma w:.th a substantial con-current neutron flux in air, deionized water, and borated water environments causes some increase in hardness and tensile strength of boraflex. During that irradiation a certain amount of gas is generated but beyond the level of 0 rads gamma it drops off considerably. The rate of gas 1 x 10 generati. n is found to be greater when B 4 C is irradiated in deionized or borated water in absence of boraflex, thus confirming the function of boraflex polymer as our escapsulant which mitigates the interaction between boron carbide and the environment. Vent holes are provided on top of each f ~

storage cell compartment to eliminate gas entrapment.

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10.2

I I Measurements of the specimen width, thickness, weight, speci-

, fic gravity at pre- and post-irradiation stages indicated minuscule variation in these quantities.

Experiments also show that neither irradiation, environment or boraflex composition has any discernible effect on the neutron transmission of boraflex. Tests also prove that boraflex does not possess leachable halogens that may be I extracted ints the pool environment in the presence of radia-tion. Similar conclusions are reached regarding leaching of elemental boron out of boraflex.

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10.3

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I 11. IN-SERVICE SURVEILLANCE PROGRAM FOR BORAFLEX NEUTRON ABSORBING MATERIAL 11.1 Program Intent A sampling plan to verify the integrity of the neutron ab-sorber material employed in the high density fuel racks in the long-term environment is describr.d in this section.

The program is intended to be conducted in a manner which allowa access to representative absorber material samples without dis-rupting the integrity of the fuel storage system. The program is I tailored to evaluate the material in normal use mode, and to forecast future changes using the data base developed.

11.2 Description of Specimens The absorber material, henceforth referred to as " poison" used in the surveillance program must be representative of the mater-ial used within the storage system. It must be of the same composi-tion, 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 I than 4" x 4" on a side. Figure 1 showed a typical coup'on. Ee'h poison specimen must be encased in a stainless stell 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 toler-ances similar to that designed into the storage system. The jacket would be closed by tack welding in such a manner as to retain its form throughout the use period yet allow rapid and easy opening I without contributing mechanical damage to the poison specimen con-tained within.

11.3 Test The test conditions represent the vented conditions of the cruciform elements. The samples will be located adjacent to the fuel racks and suspended from the spent fuel pool wall. Eighteen I (18) test samples are to be fabricated in accordance with Figure 1 and installed in the pool when the racks are installed.

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The procedure for fabrication and testing of samples I shall be as follows:

a. Samples shall be cut to size and carefully weighed in milligrams;

b. Length, width and average thickness of each specimen to be measured and recorded; 4

Il . c. Samples shall be fabricated in accordance with Figure 1 and installed in pool; I d. Two samples shall be removed at each time instant per the schedule shown in Table 1.

I 11.5 Specimen Evaluation After removal of the jacketed poison specimen from the fuel pool at the designated time, a careful evaluation of that speci-men will be made to determine its actual condition as well as its apparent durability for continued function. Separation of the poison from the stainless steel specimen jacket must be performed carefully to avoid mechanically damaging the poison. specimen. Immediately upon removal, the specimen and jacket section should be visually examined for any effects of environmental exposure. Specific attention should be directed to the examination of the stainless steel jacket for evi-I dence of physical degradation. Functional evaluation of the poison material is accomplished by the following measurements:

a. A neutron radiograph of the poison specimen will allow for a determination of the maintenance of uniformity of the boron distribution;

b. Neutron attenuation measurements of the specimen made in a fashion consistent with that described in the Poison Material Qualifying Test Data will, by comparing vi.th I

the attenuation of preirradiated poison as listed in that document, allow evaluation of the continuing nuclear effectiveness of the poison. Consideration must be given in the analysis of the attenuation measurements for the level of accuracy of such measurements as indicated by

'9 11.2

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the degree of repeatability normally observed by the testing agency;

c. A measurement of the hardness of the poison material will establish the continuance of physical and structural durability. Hardness acceptability criterion requires that the specimen hardness will not exceed the hardness I listed in the qualifying test document for lab test specimen irradiated to 10 11 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, width and average thick-ness and comparison with the pre-exposure data will in-dicate dimensional stability within the variation range reported in the Boraflex laboratory test reports.

A detailed procedure paraphrasing the spirit of this program is prepared for step-by-step execution of the test procedure and interpretation of the test data.

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11.3

I I TABLE 1 l I Date Installed I

' " INITIAL FINAL WEIGHT PIT WEIGHT WEIGijT CHANGE PENETRATION E SCHEDULE (mg/Cm 2 _Yr) (mg/Cm'-Yr) (mg/cm -Yr) mil /Yr

~E 2 90 day 1r l

I' 4 180 day 1P 5

I 6 1 year Mr 7

8 5 year 1F 9

10 10 year 1r 11 12 15 year 3r l

13 I

l 14 20 year 1r 15 16 30 year 1r 17 I

18 40 year v I

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