Text

Rev 8 to LaSalle Station Unit 2,Spent Fuel Storage Capacity Mod Sar
	ML20214U923
Person / Time
Site:	LaSalle
Issue date:	08/31/1986
From:	; COMMONWEALTH EDISON CO.
To:	;
Shared Package
ML20214U898	List: ML20214U897; ML20214U903; ML20214U923;
References
	8601--84, 8601--84-R8, 8601-00-0084, 8601-00-0084-R08, NUDOCS 8610020026
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{{#Wiki_filter:. COMMONWEALTH EDISON LASALLE STATION UNIT 2 SPENT FUEL STORAGE CAPACITY MODIFICATION SAFETY ANALYSIS REPORT APRIL 1986 8601-00-0084 Revision 8 - AUGUST 1986 0 8610020026 060719 PDR ADOCK 05000374 p PDFt

LA SALLE SPENT FUEL STORAGE MODIFICATION TABLE OF CONTENTS Pages 1.0 Introduction 1-1 2.0 Requirements for Thermal-Hydraulic Analysis 2-1 2.1 Introduction 2-1 2.2 Methods and Assumptions used in Thermal-2-1 Hydraulic Analysis l l 2.3 Approach to Thermal-Hydraulic Analysis 2-6 2.4 Detailed Analysis and Results 2-7 2.5 Conclusions 2-20 2.6 References 2-21 3.0 Criticality Analysis 3-1 3.1 Analytical Technique 3-1 3.2 Calculational Approach 3-5 3.3 Evaluation of Criticality Safety 3-6 3.4 Tolerances and Uncertainties 3-7 3.5 Accident Analysis 3-9 3.6 Design Conservatisms 3-10 3.7 New Fuel Designs 3-11 1 l 3.8 References, Tables and Figures 3-13 4.0 Seismic Analysis 4-1 4.1 References 4-1 i 4.2 Introduction 4-3 4.3 Analysis Objectives 4-12

l. i l' a 5.0 Mechanical Analysis . 5-1 [ 4 3 5.1 Summary 5-1 i 5.2 Description of New Spent Fuel Racks 5-1 5.3 Mechanical Analysis 5-7 J 5.4 References 5-18 I t e i .j. 4 1 t 4 i e a + } l .., =.. - -

1.0 Introduction Commonwealth Edison is. currently acquiring high density spent fuel racks to replace the racks supplied by the NSSS supplier of a low density design. This Safety Analysis is provided to support Commonwealth Edison's request for NRC review and approval of new spent fuel racks for La Salle County Station Unit 2. There are two spent fuel pools at La Salle Station, the existing racks in each of these pools have 1080 storage cells. In the 1989-90 time frame, they will no longer have Full Core Discharge Reserve. Replacing these racks with high density racks Figure 1 1 will extend the storage capacity for each plant into the year 2000. The high density neutron absorber racks are seismically designed with no non-load bearing parasitic structures. The basic rack consists of precision made boxes welded together in rows and columns-to form a highly damped honeycomb structure'. The neutron absorbing material is trapped between box walls and enclosed on all sides. This Safety Analysis complies with the NRC position paper "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, dated April 14, 1978, as amended by the NRC letter dated January 18, 1979. The storage racks proposed for La Salle are similar to the Nine Mile Point 2 racks which-were recently licensed. The purpose of this report is to provide a description and the technical information necessary for evaluation of-the safety aspects for the proposed modification to the La Salle Unit 2 spent fuel storage pool. 1-1

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2.0 Requirements for Thermal-Hydraulic Analysis 2.1 Introduction The scope of the analysis covers aspects of the Sargent and Lundy Specification T-3758 and includes the following: Computation of decay heat loads for the spent fuel a. pool in accordance with the NRC's SRP 9.1.3, Branch Technical Position APCSB 9-2 and ANSI /ANS 57.2. b. Independent verification of the adequacy of the heat exchangers supplied by Yuba Industries and furnished by Sargent and Lundy. Determination of pool bulk temperatures for design c. base decay heat loads and cooling conditions. d. Temperature changes and heat-up rates for-the loss of spent fuel pool cooling accident under normal refueling-and full core offload conditions. e. Recirculation flow characteristics in the hottest and average spent fuel assemblies to determine local coolant temparature changes and peak clad temperatures. A local path and an under-rack path are~ examined. f. Investigation of gamma heating in the fuel box con-taining a fuel assembly to assure adequate coolant flow exists. g. Conductive cooling of intercellular water gaps. h. Determination of the temperature distributions in the fuel box, poison, fuel box interface. 2.2 Methods and Assumptions used in the Thermal-Hydraulic Analysis The methods used in the thermal-hydraulic report of the spent fuel ~ pool (Reference 2.6.2) involve relatively uncomplicated correla-tions for friction factors, loss coefficients, and heat transfer coefficients that make a detailed computer analysis unnecessary. Further simplifying but conservative assumptions reduce the mathematical complexity to the point where hand calculations or programmable calculators are all that are required. 2-1

The design critoria used for the thermal-hydraulic analysis of the spent fuel pool for La Salle County

Station - Unit 2 are in accordance with the NRC "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications", issued April 14 1978.

Addit-ional conditions are given by Sargent and Lundy Engineers in Addendum 2 and Specification No. T-3758 for the Spent Fuel and Special Storage Racks, July 17, 1985. Based on the NRC Position Paper and the Sargent and Lundy Specifications, the following are established as design bases or requirements: a. Decay heat loads for a full pool are to be determined in accordance with the NRC Branch Technical Position APCSB 9-2, " Residual Decay Energy for Light Water Reactors for Long-Term Cooling", Section 9.2.5-8a of the Standard Review Plan. (This version is superceded by ASB 9-2 but is identical in form. The initials merely reflect the ~ branch change to Auxiliary Systems Branch.) Full pool decay heat loads and temperatures are. computed for the following cases:

1. A normal refueling discharge of 240 fuel assemblies cooled 7 days after reactor shutdown (DARS) @ 4FA/HR.

The remainder of the pool is filled with normal refueling discharges cooled 18 months each. The pool maximum tem-peratrue is limited to 120 F. A pool containing 3120 fuel assemblies is assumed. This allows for a final full core off-load of 764 assemblies; This leaves 189 cells empty because not enough room remains for a complete normal refueling.

2. An equilibrium core of 764 fuel assemblies discharged 7 DARS @ 4FA/HR and 30 days after the last refueling.

A pool containing 3884 fuel assemblies is assumed. The pool bulk temperature is limited to 150*F. e 2-2

3. Sama as caso 2, except that the dischargo occurs 90 days after the last refueling.

It is normal for-one spent fuel cooling loop to be in operation for the design bases decay heat loads.

b. To ensure - that adequate time exists for an alternate cooling method to be implemented in the event of a loss of spent fuel pool cooling capability accident, the heat-up rate is calculated and the time required for pool boiling to occur is determined.

c. Coolant flow rates, temperature increases, and peak clad temperatures are determined for worst case condit-ions (i.e. high pressure drops and low heat transfer conditions for channeled or unchanneled fuel assemblies, high bundle decay heat, etc.) to verify that boiling shall not occur.

d. The effect of gamma heating in the fuel box and intercell spaces between fuel assemblies is analyzed.

Gamma heating shall not cause boiling in these positions. Adequate flow must be established to preclude the possibility of trapping air or steam anywhere in the. fuel racks.

e. Coolant flow paths and sparger locations affecting the analysis shall be identified.

As noted in the design bases, conservative assumptions are employed for evaluations of all coolant and clad temperatures. Some additional assumptions used for the thermal and hydraulic analysis of the spent fuel pool are as follows:

a. In determining the pool bulk temperatures, only one (of two) cooling loop is assumed to be operational.

A 5% heat exchanger tube blockage is also assumed.

b. The thermal inertias of the concrete walls and the coolant and piping outside the pool boundaries are neglected in the transient heat-up analysis.

2-3

c. The pool surface is not assumed to mix to a lower. pool bulk temperature in the heat-up analysis following the

. loss of spent fuel pool cooling accident.

d. All decay energy is assumed to be absorbed in the fuel and surrounding coolant for the hot assembly or natural cir-culation analysis.

(In reality, some gamma radiation will be absorbed in the adjacent fuel boxes and poison).

e. The gamma decay heat absorbed.in the fuel box wall is taken to be proportional to the mass densities of the materials in the spent fuel pool.

(In reality, most of the gamma radiation never leaves the fuel assembly due to strong uranium attenuation.) Gamma heating proportional to the mass fraction is roughly equivalent to the assumption of uniform gamma flux in the repeating unit cell.

f. A circulation flow path from the East wall or downcomer to a position along the West wall is assumed for the hottest assembly.

This derates.the flow to the hottest assembly since flow down the three remaining walls is also possible.

g. Worst case combinations with the fuel assembly channeled (for high bundle pressure drop) and unchanneled (for lower clad heat transfer) are assumed in the natural circulation analysis for the hottest fuel assembly.

h. In gamma heating of the fuel box walls, the fuel assembly is assumed to be channeled.

4

i. The hottest assembly is assumed to generate 7 x 10 BTU /hr, more than 1.5 times higher than the decay heat generated by the average spent fuel bundle in the hot batch of spent fuel for each of the three cases analyzed.

Sinusoidal heat flux distributions force the clad hot spot factor above 2.4 for all' cases. 2-4

j. The dominant pressura drops are over estimated by, factors of 1.5 to 2.0 in most cases.

k. Material properties (e.g. thermal conductivities, densities, and specific heats) are generally assumed to be independent of temperature and are evaluated at some specified (average, inlet, or surface) temperature.

The major areas of concern in the thermal and hydraulic analysis are verification that the clad and coolant temperatures do not become high enough to cause boiling. In the event of the loss of spent fuel pool cooling accident, the heat-up rate must be slow enough to allow an alternate coolant system to be con-nected and operating before pool boiling occurs. O t e 2-5

2. 3 Approach to Thermal-Hydraulic Analysis In this analysis, the decay heat rates for the spent fuel pool, (SFP) are calculated for the normal refueling and emergency core off-load conditions.

Limits on the spent fuel pool-bulk temperatures are calculated for the specified mass fl'ow rates and design bases decay. heat loads for the condition of one heat exchanger operational. The spent fuel pool heat-up. rate and time until pool boiling following the loss of spent fuel pool cooling accident was then computed. Two recirculation paths are identified. The first is a local path involving the hottest spent fuel assembly and an adja-cent long-term cooled spent fuel assembly is considered. The second is a more complex path with under-the-rack flow. The coolant temperature increases and maximum clad temperatures are calculated for the hottest fuel assemblies for normal refueling and full core off-load conditions. ~ Gamma heating of a fuel cell box containing a fuel assembly and poison " slabs" is consid,ered. Temperature profiles in the box, poison, fuel box interface are then found. i f 4 l r 2-6

2.4 Dntailed Analysis and Results

2. 4.1 overview In this section, we present an analysis overview for the calculation summaries-that follow.

Decay heat frac-tions are computed according to NRC position standards. Total heat loads for the normal refueling and full core off-load conditions are then calculated. Heat exchanger adequacy is verified and the mass flow rates and coolant temperatures are calculated for the three design cases. The thermal inertia of the spent fuel pool (SFP) is computed. Heat-up rates and the time taken for the pool water to reach 200 F and 212 F following a loss of spent fuel pool cooling accident are found. Make-up rates at pool boiling are also determined in this section. Natural circulation cooling analyses are performed. A local recirculation path and a more complete under-rack path are considered. Clad and coolant temperature dis-tributions are specif'ied in these worst case analyses. Gamma heating of the fuel box walls and poison " Slabs" adjacent to the hottest assembly is investigated. The temperature distributions in the stainless steel wall at the fuel box and poison-fuel box interface are deter-mined. Flow blockage for the hottest fuel assembly is considered. 2.4.2 Decay Heat Loads For The Spent Fuel Pool The NRC Branch Technical Position APCSB 9-2 (or ASB 9-2) is used to compute the decay heat fractions for the La Salle County Station - Unit 2 spent fuel pool -(SFP). For 7 cooling times greater than 10 sec. (116 days), APCSB 9-2 does not specify a fission product decay uncertainty factor, but SRP Section 9.1.3 recommends a value of 0.1 for times 7 10, secs. and is used here. l 2-7 . ~.

Based on APCSB 9-2 and the three cases outlined in the design bases, the decay heat fractions and the SFP decay heat loads are as follows: Case 1: Normal Refueling, 7 days after reactor' shutdown (DARS) 240 FA; 7 DARS @ 4 ASSY5/HR 10.02 x 106 BTU /hr remainder of the SFP 3.96 x 106 BTU /hr Total heat load 13.98 x 106 BTU /hr Case 2: Full Core Off-Load, 7 DARS, 30 days in-reactor 764 FA; 7 DARS @ 4 ASSY$/HR 22.82 x 106 BTU /hr 240 FA; 30 DARS 4.78 x 106 BTU /hr remainder of the SFP

4. 01 x 106 BTU /hr Total heat load 31.61 x 106 BTU /hr Case 3: Full Core Off-Load, 7 DARS, 90 days in-reactor 764 FA; 7 DARS @ 4 ASSYS/HR 25.12 x 106 BTU /hr.

240 FA; 90 DARS 2.96 x 106 BTU /hr remainder of the SFP 3.87 x 106 BTU /hr Total heat load

31. 95 x 106 BTU /hr The reactor is a BWR with an eighteen month equilibrium refueling cycle (1/3 core refueling discharge) operating at a rated 3323 MWth.

A burn-up of 1 MWD /MTU is approx-imately equivalent to 1 full power hour. The racks will provide storage for 4073 fuel assemblies, equivalent to 5.33 cores. Allowing for the full core off-load capacity, 13 refueling discharges (or 20 years. of spent fuel) can be safely accommodated before spent fuel relocation becomes necessary. 9 2-8

As applied here, a 25% margin of conservatism is expected through the use of APCSB 9-2 and the pool histories used.

This margin allows for uncertainties of approximately 10% (for curve fitting)., 5% (for fission yields), and 10% (for PU uncertainties that are likely to introduce additional credits). 2.4.3 Spent Fuel Pool Heat Exchanger Design Adequacy and Pool Bulk Temperatures. This analysis determines the limits on the spent fuel pool bulk temperature and is divided into three parts. In Part 1, the Yuba Industries, Inc. heat exchanger (HX) for the LSCS2 spent fuel pool is checked for consistency. A reference to the data sheets or specifications for the Yuba Industries, Inc. HX can be found in Reference 2.6.17. The overall heat transfer coefficient, U, is given as: UClean = 553 BTU /h~r ft2=p U TU/hr ft2ep Dirty All fluid properties are evaluated at a mean temperature of 110*F. The mass flow rates for the one tube pass /one shell pass HX are: m = 1.5 x 106 lbm/hr - shell side (SFP) h m = 2.0 x 106 lbm/hr - tube side (CW) h The HX effectiveness under these conditions is &= 0.388 In Part 2, the value of U is calculated explicitly by accounting for all thermal resistances expected between the shell and tube fluids. The wall resistance is determined by the tube thermal conductivity. The film coefficients ,at the inner and outer tube surf aces are estimated using a Dittus-Boelter correlation. 2-9

In Part 3, the off-design heat exchanger temperatures are determined for normal refueling and full core off-load conditions. For a conservative margin, we assume 5% of the heat exchanger tubes are plugged. It is normal for one loop to be operational. The three design cases are presented below: TABLE 2.4.3-1 SPENT FUEL POOL AND HEAT EXCHANGER TEMPERATURE LIMITS NO. OF T CONDITIONS HX's HIN HOUT HAVG COUT gy= 13. 98 x 106 BTU /hr Normal Refueling 7 DARS @ 4 ASSY'S/HR 1 119.6 110.3 114.9 102.0 q2= 31. 61 x 106 BTU /hr Full Core Off-Load 7 DARS @ 4 ASSY'S/HR, 30 days in reactor 1 149.4 128.3 138.8 110.8 q3 = 31. 95 x 106 BTU /hr Full Core Off-Load 7 DARS @ 4 ASSY'S/HR, 90 days in reactor 1 150.0 128.6 139.3 111.0 NOTES: All temperatures in F. T = peak pool bulk temperature (HX inlet - tubes) HIN T = p ol average temperature HAVG T = sparger discharge temperature (HX outlet - tubes) HOUT-T = cooling water outlet temperature COUT T = 95 F = cooling water inlet temperature CIN Under normal refueling conditions, THIN (120*F 7 DARS with one HX in-operation. Spec. T-3758 limits this value to 120*F, in agreement with the value tabulated here. Since 5% tube plugging is assumed in all cases the heat loads 2-10

1 are considered conservatively high (by 25% to cover uncer-tainties), all temperatures are extreme temperatures for the given conditions of cooling time (DARS) and discharge type (i.e. normal refueling or full core of f-load). With full core off-load 7 DARS, 30 days in reactor and 90 days in reactor, the peak pool bulk temperature of 149.4 F and 150*F are less than or at the spec. limit (at 7 DARS) of 150*F. I i 2.4.4 Pool Thermal Inertia and Heat-Up Rates. In this analysis, the spent. fuel _ pool (SFP) heat-up rate and time until pool boiling following the loss of ' spent fuel pool cooling accident is considered. The SFP thermal inertia for La Salle County Unit 2 is found. A full pool 2 is assumed and is conservative since the UO2 and Zircaloy in the fuel assembly have a lower thermal inertia than the water these materials displace. For simplicity (and al'so-conservatism) we take pc = 60 BTU /ft

F for both water p

and stainless steel. Only the water and materials within the liner boundaries are considered. Materials that possess some thermal inertia that are neglected include: primary and secondary water and piping exterior to the liner plates, the cask pit, the transfer canal and gates, and the concrete walls of the spent fuel pool. The total SFP thermal inertia under these simplifying but conserva-tive assumptions is 6 I = 2.48 x 10 BTU /*F p with water and stainless steel accounting for more than 91% of this. i f In the heat-up analysis, the initial pool temperature is l taken to be T - the inlet HX temperature. Pool mixing HIN ,to a lower pool temperature is neglected. Following the loss of SFP cooling, the pool heat-up rates, times to -11 2

1 l reach 200'F (and 212*F), boil-off rates, and elevation changes at boiling are determined. These are tabulated on the following page. TABLE 2.4.4-1 RESULTS OF HEAT-UP ANALYSIS FOLLOWING THE LOSS OF SPENT FUEL POOL COOLING LA SALLE UNIT 2 BTUt Conditions q T 2 Lt(to 200*F/ g 'lbmq Q (GPM) [1 N , hr 3 hr' 212 F) UI' d L Normal Refuelin9 13.98x106 5.62 14.2/16.3 hrs 14,400 28.9 .21 7 DARS Full Core Off-Load 31.61x106 12.72 3.93/4.87 hrs' 32,500 65.3 .47 7 DARS, 30 days in reactor Full Core Off-Load 31.95x106 12.8 3.90/4.84 hrs 32,900 6 6 '. 0' .48 7 DARS, 90 days in reactor Definitions: DARS = Days after reactor shutdown g = total SFP decay heat load (full pool assumed) T = heat-up rate following loss of SFP cooling at = time to reach 200*F/212 F from T Of I HIN cooling loop 6 = boil-off rate at 212'F, 14.7 psia. 0 = make-up rate (at y = 8. 3 lbm/ gal) h = rate of change of pool elevation at boiling The above at's are considered long enough for an alternate cooling source to become operational following the loss of SFP cooling. 2-12

2. 4.5 Natural Circulation Cooling of the Spent Fuel..

In this analysis, two recirculation paths are identified which are the natural circulation cooling of the LaSalle County Station Unit 2 spent fuel assemblies. A local path where coolant is convectively driven up the hottest assembly and down a " cold" assembly is studied first. A second path flowing under the spent fuel racks, up the hot assemblies, into the mixing region above the racks, and finally down the West wall of the pool to complete the path is then modeled and analyzed. Apart from the estimation of the coolant inlet temperatures to the hot batch of spent fuel, these flow paths are decoupled from the cooling loop (s) and SFP heat exchanger (s). The following conservative assumptions are used in this analysis: 1. The. hottest assembly decay heat is taken to be 4 7.0 x 10 BTU /hr. This hottest assembly generates no less than 1.5 times more heat than the average assembly from the most recent spent fuel discharged for the three respective cases: Case 1: Normal refueling, 7 DARS @ 4 ASSY'S/HR Case 2: Full Core Off-Loading, 7 DARS @ 4 ASSY'S/ HR, 30 days in reactor Case 3: Full Core Off-Loading, 7 DARS @ 4 ASSY'S/ HR, 90 days in reactor A sinusoidal heat flux distribution at the clad surface is the assumed, so that the hot spot factor exceeds the average by a factor of1f/2 = 1.57. 2. The remainder of the pool is filled with 763 (full 4~ core less one) assemblies, each generating 4.5 x 10 BTU /hr. 2-13

3. The hot batch of spent fuel is located at the West wall so that the under-rack path length and distance from the operating sparger is maximized. 4. The fuel assembly pressure drop (versus flow) is accu-rately correlated with available experimental data. The data is for the channeled fuel assembly at 100"F, 34 . psia and will be worst case since 100'F is a represen-tative low temperature (high viscosity) and the channeled assembly has a smaller flow area than the unchanneled assembly. As a function of the volume flow rate Q in the fuel assembly, 1.564 p = 973 (Q) is the fuel assembly pressure drop equation, here p is in psf and Q in cfs. 5. Fortheunitcellgeometryofthefuelassembly,kiE 8x8,.483" rods,.640" pitch), the Nusselt number for the (laminar) flow in the spent fuel assembly is 5.7. The clad heat transfer coefficient is evaluated on the basis of an unchanneled fuel assembly (lowest h) and found to be 32' BTU /hr ft2*F. For the channeled assembly, it would be closer to 45. 6. For the under-rack flow path (path 2), the losses calculated include: a. West wall friction losses b. 90' turn losses c. Under-rack friction losses d. Expansion and contraction losses at the rack supports e. Fuel bundle losses f. Branching - Momentum losses e 2-14

Flow branching / momentum losses are typically small (and recoverable) in' comparison to the total losses. The dominant loss (d) - the expansion and contraction losses contribute more than 50.6% to the total flow loss. These were overestimated by a factor of 10%. _ 7. The driving pressure (caused by water density varia-tions) is given by 9L / 2_' apd g 2C o .sp: 4 p where 8= I fh p }7 is the thermal coefficient I of expansion for the water-only a mild function of pressure. L is the heated (active) l'ength (12.5 f t) and q is the decay heat rate. Fluid properties are evalu-ated at 125*F, a low temperature for the core of f-load case. At 125'F, S = 2.6 x 10~4'F and at 150*F, S= ~ 3.1 x 10~4'F-1 The driving pressure is then derated by approximately 20%. 8. For the local path, the fuel assembly inlet temperature is taken to be the hottest pool bulk temperature (THIN from Section 2 4 3). Since considerable mixing can occur in path 2 and cold water is discharged under the rack, T is taken to be the pool average temperature (i.e. IN T from Section 2.4.3. ggyg A summary of significant results that apply to all three cases '(for the purposes of determining peak coolant and clad temperatures) are as follows: 2-15

TABLE 2.4.5-1 RESULTS RELATED TO PEAK COOLANT AND CLAD TEMPERATURES Path 1 Path 2 volumeflow$atein 8 ft /sec .011 .0084 the hottest assembly (GPM) (4.8) (3.5) Total pressure drop = APd ft 2.9 3.5 ~ (3. 2 under rack) Coolant AT - *F 29 38 Position of clad hot spot - ft 7.8 8.2 Difference (T -TIN)- F 52 59 cladmax Following assumption 8, the peak clad and coolant temperatures T and T are as tabulated below: cladmax OUT TABLE 2.4.5-2 PEAK CLAD AND COOLANT TEMPERATURES Path 1 Path 2 Conditions TOUT ( F) Tcladmax( F) TOUT ( F) Tcladmax( F) Case 1 - Normal Refueling 148.6 171.8 158 178.8 7 DARS,4 ASSYS/HR Cane,2 - Full-Core Off-Load 178.6 201.6 188 208.8 I 7 DARS,,30 days ,r-ronc+qr Case 3 Full Core Qff-Load 178.6 201.6 183 208.8 inbhef8E5r2 ^ NOTE: T a 240'F at top of racks ( 25 psia) sat T a 245'F at peak clad T (~30 psia) sat All cases are significantly subcooled and void fractions are negligible. Since the actual flow paths will be complicated combinations of local and under-rack paths, the temperatures will not exceed those indicated. The indicated temperatures are much lower than the 700-800*F which is the lower bound for low temperature ( sensitization of austenitic stainless steel. 2-16

2. 4. 6 Gamma Heating of the Fuel Box Walls, Poison, and Inter-cellular Water.

In this analysis, gamma heating of the fuel box walls, poison, and intercellular water is investigated. Fission product decay accounts for virtually all residual heat in the spent fuel pool with minimum cooling times t 5 7 days. At this time, a realistic but upper limit s on the gamma fraction is 0.62, based on the primary reference for the NRC position standard APCSB 9-2. A typical 1 MeV electron will travel approximately 0.016" in the UO fuel. Thus, all B electrons will be stopped 2 in the fuel or surrounding clad and coolant. A typical, but higher than average, 1 MeV gamma ray has a mean free path of approximately 0.56" in the UO fuel - comparable 2 to the pellet diameter.410". Therefore, the fuel will not stop all the gamma radiati6n emitted by the decaying fission products. It will then be conservative to assume the following when estimating the energy deposition in the fuel box, poison, and intercellular water: 1. The fuel box is located within an infinite array of hottest assemblies - each generating 7 x 10" BTU /hr, 62% of which is gamma. 2. The gamma energy absorbed in a unit cell comprised of one fuel assembly, one fuel box, and 4 "1/2" Poison " slabs" is proportional to a given material's mass fraction. This is roughly equivalent to the assumption of uniform y-flux since p/p is approximately constant for all materials at a given gamma energy. 2-17

3. Water in the fuel box must remove gamma heat due to energy depositionc in the 0.090" thick fuel box wall, the. fuel box water itself, and the poison slabs adjacent to it. The flow rate between the fuel can and the cell wall is determined by equating the driving head to the loss head. The coolant temperature corresponding to this flow rate is then computed and is seen to be less than the boiling temper-ature of the coolant. Intercellular water is free to flow between the fuel assembly channel and the fuel box boundary. the flow areas defined by two.281" diameter nozzle holes in each fuel assembly are large enough to remove heat in the 0.090" box walls and poison " slabs" so that boiling will not. occur. 2.4.7 Stainless Steel Temperatures. In this calcuation, temperature distributions in the stainless steel box walls are determined. Heat source terms (due to y-irradiation) follow the same conservative assumptions used in Section 2.4.6. Channeled 5uel is also assumed and the resulting geometry contains water gaps adjacent to the box wall interfaces. The maximum temperature of the stainless steel is seen to be less than the boiling temperature of the coolant. e 2-18

One dimensional heat conduction in the series of " slabs" defined by the fuel assembly channels, water gaps, stainless steel box walls, and poison slabs is modeled using the steady-state, 1-D, heat conduction equation: h (T -Tc)

9

II s where q is the volumetric heat generation rate in the " slab" and T is the surface temperature of the " slab". s Denoting the coolant temperatures in the fuel box as T ' stainless c steel, poison,and coolant were all discovered not to exceed T by c more than 59*F. Since convection is neglected and all q are conservatively high, the 59 F variation above T is conservative. c In Table 2.4.5-2, T ex eeds T N*** OUT) by 2-F,-thus cladmax c all points in the racks are subcooled by at least 20 F, thus film boiling of the intercellular water will not occur. The maximum. temperature gradient occurs across the fuel box interface and is modeled using the steady-state, 1-D, heat conduction equations stated previously. The maximum temperature difference is limited to less than 1 F, and is of a magnitude which should produce negligible thermal stresses in the stainless steel box. o 2-19

2.5 Conclusions 4 The detailed thermal-hydraulic analyses described in Sections 2.4 through 2.4.7 address the concerns, intent, and design bases of the NRC's Position Paper "OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications" and the Sargent and Lundy Specification Number T-3758 plus Addendum 1. A Based on these analyses, our professional staff has concluded that the spent fuel pool for La Salle County Station - Unit 2 can be adequately cooled in accordance with the suggested regulatory standards of the Nuclear Regulatory Commission and. ] comply with specifications outlined by Sargent and Lundy l Engineering Corporation. 1 9 l 1 i i i l I 4 t I i s l 2-20

2.6 References 1. Sargent and Lundy Specification T-3758 and Addendum 1, Spent Fuel and Special Storage Racks for LaSalle County Station - Unit 2, Commonwealth Edison Company, Project No. 7043-73, July 17, 1985. 2. U.S. Tool & Die, Inc. " Thermal-Hydraulic Report: Spent Fuel Storage Racks: LaSalle County Station - Unit 2, "8601-00-0083, Prepared for Sargent and Lundy Engineers, April, 1986. 3. Arya, A. P., Fundamentals of Nuclear Physics, Allyn and Bacon Inc., Boston, MA, 1966. 4. Telecon, Ray Linder (U.S. Tool & Die) and Bright Wong (Commonwealth Edison Co. ), March 11, 1986 (8601-00-0050). 5. UST&D Drawing 8601-1, Rev. 1, February 18, 1986 " Plan Arrangement of Spent Fuel Storage Racks." 6.

Dailey, J.

W. and D. R. F. Harleman, Fluid Dynamics, Addison-Wesley Publishing Co., Reading, MA, 1966. 7. Eckert, E. R. G. and R. M. Drake, Heat and Mass Transfer, 2nd Edition, McGraw-Hill, New York, NY, 1959. 8. El-Wakil, Nuclear Heat Transport, International Textbook Company, New York, NY, 1971. 9. General Electric Technical Paper 22A5866 Rev. 1.

10. Kays, W. M. and A.

L. London, Compact Heat Exchangers, 2nd Edition, McGraw-Hill, New York, NY, 1964. ll.,Kreith, F., Principles of Heat Transfer, 2nd Edition, International Textbook Company, Scranton, PA, 1968. 2-21

12. Baumeister and Marks, Standard Handbook for Mechanical Engineers, 7th Edition, McGraw-Hill, New York, NY, 1967.

13. NRC Branch Technical Position APCSB 9-2, " Residual Decay Energy for. Light Water Reactors for Long-Term Cooling,"

Standard Review Plan, Section 9.2.5-8, 1975.

14. NRC Branch Technical Position ASB 9-2,

" Residual Decay Energy for Light Water Reactors for Long-Term Cooling," Standard Review Plan, Section 9.2.5-8a, Rev. 1, 1978.

15. Perkins, J.

R. and R. W. King, " Energy Release from the Decay of Fission Products," Nuclear Science and Engineering

Vol, 3,

1958, p. 726.

16. Yuba Heat Transfer Division, Yuba Industries, Inc.,

QN-223-73, La Salle County Station - Unit 2, Commonwealth Edison Co., August 21, 1973.

17. Streeter, Victor, ed., Handbook of Fluid Dynamics, McGraw-Hill, New York,'NY, 1961.

18. Lederer, C.

M., J. M. Hollander, and I. Perlman, Table of the Isotopes, 6th Edition, J. Wiley and Sons, New York, NY, 1967.

19. Chemical Engineering / Desk Book Issue / April 14, 1969.

20. Giedt, Principles of Engineering Heat Transfer, D. Van Nostrand Co.,

1957. 9 e 2-22

w 3.0 CRITICALITY ANALYSIS The following discussion summarizes the design of the spent fuel racks with respect to criticality safety. The analytical techniques described here are similar to those used to successf ully license spent fuel racks for several other plants, the most recent being those for Point Beach Units 1 and 2 and Nine Mile Point Unit 1. 3.1 ANALYTICAL TECHNIOUE U ) computer program was used to generate macroscopic cross The LEOPARD sections for input to four energy group dif fusion theery calculations which are performed with the PDQ-7 I ) program. LEOPARD calculates the neutron energy spectrum over the entire. energy range from thermal up to ' ~~ 10 Mev and determines averaged cross sections over appropriate energy The fundamental methods used in the LEOPARD program are those groups. used in the MUFTI 3) and SOF0CATE "Iprograms which were developed under the Naval Reacter Program and thus are well founded and extensively tested techniques. In addition, Westinghouse Electric Corporation, the developers of the original LEOPARD program, demonstrated the accuracy of these methods by extensive analysis of measured critical assemblies consisting of slightly enriched UO fuel rods.(5) 2 in addition, Pickard, Lowe and Garrick, Inc. (PLG) has made a number of improvements to the LEOPARD program to increase its accuracy for the calculation of reactivities in systems which contain significant amounts of plutonium mixed with U0. PLG has tested the accuracy of these 2 modifications by analyzing a series of UO and Pu0 -UO critical 2 2 2 l experiments. These benchmarking analyses not only demonstrate the improvements obtained for the analysis of Pu0 -UO systems but also 2 2 l demonstrate that these modifications have not adversely af fected the accuracy of the PLG-modified LEOPARD program for calculations of slightly enriched 00 systems. 7 l l l i l 3-1

The U0 critical experiments chosen for benchmarking include variations 2 in H 0/UO v lume ratios. U-235 enrichments, pellet diameters and' 2 2 cladding materials. Although the LEOPARD model also accurately calculates the reactivity effects of soluble boron, these experiments have not been included in the LEOPARD benchmarking criticals since the spent fuel pool calculations do not involve soluble boron. Neutron leakage was represented by using measured buckling input to infinite lattice LEOPARD calculations to represent the critical assembly. A sununary of the results is shown in Table 3.1-1 for the 27 measured criticals chosen as being directly applicable for bench-marking the LEOPARD model for generating group average cross section for spent fuel rack criticality calculations. The average calculated k,ff is 0.9979 and the standard deviation from this average is 0.0080 ak. Reference 5 raised questions concerning the accuracy of the measured buckling reported for the experiments number 12 through 19. If these data are excluded, the average calculated k,gg for the remaining 19 experiments is 1.0006 with a standard deviation from this value of 0.0063 ak. In all of these experiments there are significant uncertainties in the measured bucklings which are necessary inputs to the LEOPARD analysis. These uncertainties are the same order of magnitude as the indicated errors in the LEOPARD results, and therefore a more definitive set of experimental data is used to establish the accuracy of the combined LEOPARD /PDQ-7 model used for the criticality analysis of the spent fuel racks. 1 The PDQ series of programs have been extensively developed and tested over a period of 25 years, and the current version, PDQ-7, is an accurate and reliable model for calculating the subcritical margin of the prcposed spent fuel rack arrangement. This code or a mathematically equivalent method is used by all the U.S. suppliers of light water reactor cores and i reload fuel. In addition, this code has received extensive utilization in the U.S. Naval Reactor Program. As a specific demonstration of the accuracy of the calculational model used for.the spent fuel rack calculations, the combined LEOPARD /PDQ-7 model has been used to calculate fourteen measured just critical 3-2 ,.-._-,e- ,.-...,,.,,n-- ,---.r

assemblies. The criticals are high neutron leakage systems with a large variation in H 0/UO y 2 2 lume ratio and include parameters in the sar'ne range as those applicable to the spent fuel rack design. Experiments l including solubl'e boron are included in this demonstration since the ability of PDQ-7 to calculate neutron leakage ef fects is of primary interest. The use of soluble boron allows changes in the neutron leakage of the assembly while maintaining a uniform lattice and thus allows a better test of tne accuracy of the model. Furthermore, it eliminates the error associated with the measured bucklings, which is inherent in the LEOPARD benchmarks, thus permitting determinations of the actual calculational uncertainty which must be accounted for in the spent fuel rack criticality analysis. These combination LEOPARD /PDQ-7 calculations result in a calculated average k,77 of 0.9928 with a standard deviation about this value of 0.0012 ak. These results, as shown in Table 3.1-2, demonstrate that the proposed LEOPARD /PDQ-7 calculational model can calculate the reactivity of the proposed spent fuel rack arrangements with an accuracy' of better than 0.010 ak at the 95 percent confidence level. The cross sections for the Bor*aflex* neutron absorbing material which is an integral part of the design are calculated using fundamental techniques that have been successfully applied in the past to thin heavily absorbing mediums such as control rods.0 0) The procedure is straightforward and is comprised of several well defined steps: 1. The Boraflex* sheets are associated with the stainless steel and water areas exterior to the fuel twdle to define a one-dimensional slab geometry representing Sh orN tr material volume fractions. An equivalent LEOPARD cylindr"r si e.-iimensional geometry is used to obtain a first estimate of the spatial and energy self-shielded cross 10 section for the B in the Boraflex*. 3-3

l 2. Using the energy averaged cicroscopic cross sections fron 1., integral transport theory is applied in slab geometry using They's method for calculating flux depressions and shielding factors to 10 determine an appropriate B number density. This approach is similar to that of Amouyal and Benoist. 3. The self-shielded number densities calculated in Step 2 are again input to LEOPARD to obtain corrected microscopic B10 cross sections. 4. Blackness theory is applied to obtain macroscopic cross sections which will produce the requ, ired boundary conditions at the surface of the Boraflex* sheets. In addition to the fourteen critical assemblies in Table 3.1-2, the LEOPARD /PDQ model was used to calculate the k,gg for twelve additional critical assemblies, seven of which incorporated thin, heavily-absorbing materials for which the procedure just described was used to determine the macroscopic cross section. These twelve criticals were performed by Battelle Pacific Northwest Laboratories specifically for the purpose of providing benchmark critical experiments in support of spent fuel criticality analysis. They are described in detail in Reference 17. The results of these critical experiments are summarized in Table 3.1-3. The first seven of these twelve experiments include fixed neutron poison absorber plates, and the average k,ff calculated for these just critical experiments was 0.9935, with a standard deviation around this value of 0.0007 Ak. The other five critical experiments in this series do not include absorber plates and the average k,7f calculated for these just critical assemblies was 0.9944, with a standard deviation around this.value of 0.0007 ak. The overall average k,f f calculated for these twelve just critical assemblies was 0.9939, with a standard deviation around this value of 0.0008 ak. o 3-4

g applicability of the methods described above.for use in calculating the subtritical margin of these fuel storage rack designs, and demonstrates that the accuracy of better than 0.010 ak at the 95 percent confide'nce level established for the LEOPARD /PDQ-7 model applies equally well to designs incorpora' ting fixed neutron absorbers for which blackness theory is used to calculate the macroscopic cross sections. As a result of this approach to separately benchmark both the cross sections and the diffusion theory calculations against applicable critical assemblies, the " transport theory correction f actor" is implicitly included in the derived calculational uncertainty factor. 3.2 CALCULATIONAL APPROACH The PDQ-7 program is used in the final predictions of the reactivity of the spent fuel storage racks. The calculations are performed in four energy groups and take into account ~all the significant geometric details of the fuel assemblies, fuel boxes and major structural components'. The-geometry used for most of the calculations is a basic cell representing one-quarter of the area of a repeating array of stainless steel boxes. The specific geometry of this basic cell is shown in Figure 3.2-1. The calculational approach is to use the basic cell to calculate the reactivity of an infir.ite array of uniform spent fuel racks and to account for any deviations of the actual spent fuel rack array from this assumed infinite array as perturbations on the calculated reactivity of the basic cell. The effects of manufacturing tolerances, as well as thermal uncertainties, including f uel and water temperature and density variations, are also trea',ed as perturbations on the calculated reactivity of the basic cell. The fuel assemblies used for this analysis are the General Electric 8x8 design for which data are provided in Table 3.2-1. The fuel bundles were assumed to be unirradiated with an enrichment of 3.416 weight percent l. l 3-5 l L

U-235 over the entire slightly enriched section of the bundle. This is equivalent to a loading of 16.52 gm of U-235 per axial cm of the slightly enriched section of the fuel bundle. All of the calculations were performed at a uniform pool temperature of 68'F, except when the reactivity effects of pool teeperature were taken into account as a perturbation on the basic cell calculations. 3.3 EVALUATION OF CRITICALITY SAFETY _ For the average rack cell, which is defined as the basic cell for analysis purposes, the pitch wil.1 be 6.255 inch, and the ka of this cell is.9179. Figure 3.3-1 presents the spent fuel storage rack reactivity as a function of fuel bundle enrichment for the basic cell geometry. The reactivity of the basic cell as a function of B10 loading in the Boraflex* is shown in Figure 3.3-2. The 8 loading which was used 10 for the criticality analysis was the minimum loading to be incorporated into the design. This corresponds to a B10 loading of.020 grams per square centimeter of cross sectional area in a nominal thickness of .075 inch. A detailed one-dimensional axial model, which utilized flux weighted cross sections from planar cell representations of the rack, was used to calculate the net reactivity offect of axial neutron leakage including the natural uranium end sections of the assembly and the semi-circula'r cutouts of the Boraflex sheets. The resulting net effect 15 an increase in k= of.0026 ok. As derived in Section 3.1, the combined LEOPARD /.000 model bias to be added is.0061 ak. If Ir channels are stored on the fuel bundles, the calculated k of the rack is increased by +.0045 ak. 3-6

As shown in Table 3.3-1, the net effect of all the calculational biases is.0132 Ak, which therefore increases the basic cell k= to.9311,. 3.4 TOLERANCES AND UNCERTAINTIES There are also a number of tolerances and uncertainties which result in perturbations which must be considered in the criticality analysis. The reactivity ef fects of all such positive perturbations are then combined statistically in accordance with Referencc 18 to determine a single reactivity perturbation which is added to the calculated basic cell multiplication factor (including biases) to determine the final conservative evaluation of the spent fuel rack maximum possible multiplication factor. ~The tolerance on the Boraflex thickness is i.010 inch, but the minimum 10fc,2 for any thickness within this tolerance. loading is.020 gm B Calculations demonstrate the k= is largest for the minimum Boraflex thickness of.068 inch, and the resulting pertubation to the basic cell - is.0006 ak. The worst case in terms of manufacturing tolerances from a reactivity perspective is represented by the min'imum fuel box inside dimension. The k for this minimum inside dimension spent fuel storage rack cell at 68'F for a fuel assembly enrichment of 3.416 weight percent U-235 is .9197. Since the k= of the basic cell is.9179, the perturbation in k= due to tolerances on fuel box cell dimensions is.0018 ak, which corresponds to a sensitivity af.072 ak/ inch. To determine a conservative evaluation of the' reactivity effects of thickness variations in the stainless steel structural materials, all stainless steel members were assumed to be at the most reactivity limiting thicknesses allowed by the tolerances. The reactivity is highest for the maximum stainless steel thickness and the k. of the resulting basic cell is.9180. Therefore, the perturbation due to stainless steel tolerances is.0001 ak. 3-7

a The reactivity of the spent fuel storage rack.was evaluated for the effect of manufacturing tolerances on 002 density. The reference' cell is based on the fuel design value of 955 theoretical density. The worst case of a UO density of 96% theoretical density was examined. 2 The resulting perturbation to the basic cell was determined to be.0014 Ak due to an increase in pellet density from 955 to 965 of theoretical density. With regard to fuel position uncertainties within the fuel boxes, calculations confirm the fact that the fuel assemblies when center located in their most reactive positions within the fuel boxes. The fuel bur.dles in a two by two array of storage locations were moved off center in such a direction as to place all four fuel assemblies uniformly in closer proximity to one another. These calculations confirmed that tha maximum k= is obtained with each fuel bundle centered in its storage position. Based on the results of the calculational model benchmarking described in Section 3.1, the ko uncertainty in the model, which corresponds to a 95/95 confidence statement, is.0022 ak. The reactivity of the basic cell as a function of temperature is shown in Figure 3.3-3. With a maximum pool temperature of 200*F, the k= is less than the vqlue at 68'F by.0215 ak. This is to be expected for this design which incorporates a heavy loading of 8 10 as a neutron poison and indicates that the lower temperature conditions produce the higher spent f uel ~ storage rack k=. Since the reference case was based on a temperature of 68*F, which is clearly conservative, no additional reactivity effcct needs to be added to account for temperature. The sensitivity of the spent fuel rack multiplication factor to variations in the water density throtghout the pool is illustrated in Figure 3.3-4. Again, the effect.of the heavy BIO loading is to produce the most reactive conditions at full water density. A suma'ry of the perturbations to the basic cell reactivity calculations is shown in Table 3.3-1. As shown in this table, the total reactivity perturbation to be added to the biased basic cell reactivity to account for tolerances and uncertainties is.0032 ak. 3-8

Therefore, the conservatively calculated reactivity of the spent fuel rack fully loaded with unirradiated bundles with 3.416 weight perc'ent U-235 and 'no burnable poison is.9343for a pool temperature of 68'F including conservative allowances for manufacturing and calculational uncertainties. 3.5 ACCIDENT ANALYSIS i The fuel racks are designed to prevent a dropped fuel bundle from penetrating and occupying a position other than a normal fuel storage location. The only positive effect of such a bundle on the reactivity of the rack would be by virtue of a reduction in axial neutron leakage from the rack. Since there is approximately 13 inches between the top of the active fuel and the top of the boxes comprising the rack, a dropped fuel bundle will be neutronically decoupled from the fuel in the rack and would not have any measurable effect on the reported maximum possible reactivity of the spent fuel storage rack. The lattice of the fuel bundles results in an undermoderated configuration; so that any crushing or compaction of the fuel bundles would tend to reduce the neutron multiplication factor of the spent fuel pool. Theref' ore",' deformations resulting from the dropping of heavy objects into the f uel pool or f rom the ef fects of earthquakes or tornadoes will not produce a criticality accident. The reactivity ef fect of a fresh fuel assembly located adjacent to the fully loaded spent fuel storage rack has been evaluated for all postulated locations other than normal fuel storage locations. The model used to evaluate the maximum effect of a fuel bundle, which is accidentally mislocated directly. adjacent to the outer row of fuel storage racks, consisted of a 3 box by 6 box section of the fuel rack I with the extra bundle located directly on the centerline of the 3 box array. It is obvious that the centerline location selected, which is directly in line with the centerline of the stored fuel bundle, is the position which results in maximum reactivity. The maximum perturbation i associated with such an accident has been calculated to be.0098 ak. For this accident, even assuming the worst case k= of.9343, the spent fuel storage rack design assures that the multiplication factor is less than 0.95. l

Because of the well founded, conservative technique used for determination of the infinite multiplication factor, there is more than reasonable assurance that this spent fuel rack design will not cause a significant hazard to the public health and safety resulting from criticality considerations. 3.6 DESIGN CONSERVATISMS BEST ESTIMATE CALCULATIONS When the fuel assembly is represented by an explicit fuel pin distribution of selected U-235 enrichments typical.of the General Electric Company's intra-assembly fuel pin arrangement (which produces a bundle slightly enriched section enrichment of 3.416 weight percent U-235), the ke is calculated to be.9054. This.is less than the kan of the basic cell which utilizes a single average enrichment in all fuel pins, and therefore, the perturba-tion to be applied to account for the more realistic explicit multi-enrichment fuel pin distribution is .01254k. Doubling the number of mesh points used to represent the basic cell geometry results in a reduction in ka-of .0004 as shown in Table 3.3-1. As discussed previously, the basic cell calculations make the conservative assumption that all fuel bundles are unirradiated and contain no burnable poisons. Calculations, verified by reactor operation, show that with the burnable poison loadings required for a fuel bundle initial enrichment of 3.416 w/o, the maximum possible bundle kneis at'least 0.030ak less than the initial ka of the bundle with no burnable poison. In addition, the spacer grids in the fuel bundle reduce the calculated kan by 0.00294 k. These effects would reduce the calculated km of the basic cell by at least 0.04584k. Thus, the actual maximum possible multiplication factor of the spent fuel racks when completely filled with unirradiated fuel of 3.416 w/o U-235 and under the worst accident condition is less than.8983. )

- ~.. - 3.7 NEW FUEL DESIGNS ~ I It is anticipated that in the future new reload fuel designs will be developed, and the initial enrichments of these designs may be greater than the 3.416 w/o loading analyzed herein. However, the maximum reactivity of such higher enrichment designs is limited i by the existing reactor control systems which must provide assurance that the specified reactor shutdown margin will be maintained at all times for all reload fuel designs. 1 As a practical matter Gd 023 burnable poison containing fuel rods are incorporated in all BWR reload fuel designs, and for the higher enrichment designs, increased Gd 0 2 3 concentrations and/or larger numbers of burnable poison containing fuel rods are incorporated in such designs. To provide a criterion to establish that it is safe to store higher enrichment designs, a generic calculation model was evaluated. An average assembly enrichment of 4.25 w/o U-235 was selected for this generic evaluation which corresponds to a maximum average planar enrichment of 4.558 w/0 U-235 in the fuel assembly assuming i natural uranium in the top and bottom six inches. Using this enrichment, cell calculations were performed corresponding to different concentrations of Gd 0 2 3 in burnable poison containing i fuel' rods. The results, shown in Figure 3.3-5, allow the deter-mination of the storage rack k. as a function of the fuel assembly

k. in the reactor lattice which was varied by changing the burnabic poison concentration.

Figure 3.3-5 may be used to determine the resulting k. of the LaSalle Spent Fuel Storage Racks when loaded with fresh fuel with an initial enrichment of 4.558 w/0 (for the maximum planar average 4 enrichment of the buadle) as a function of the maximum k. achieved i by the bundle during its irradiction in the reactor core. The k. j of the 'uel assembly is evaluated for the reactor core geometry when loaded with channeled fuel bundles and all control rods withdrawn at 68'F. The upper limit curve is based on the k of f 3-11 3 -,-m-v,-,,- y,www m-y m-n ,,+,w- ,--e -,m,-+ r_,--enen_~, .-en-n-m-, ,,w_.e-

the combined fuel assembly plus zircaloy channel and thq appropriate amount of interassembly water associated with the fuel assembly. The lower limit curve is based on the k of the fuel assembly only - excluding the zircaloy channel and associated external water chan-nels. The limit curves are conservative for cpplication to initial enrichment fuel less than 4.558 w/o. As an example assume a new fuel design has a marimum in-reacter k of 1.35. Using Figure 3.2-5 the corrcsponding fuel rac!t k would be 0.903 for the unchannelled fuel and 0.920 for the channelled m fuel. Thus the fuel design could be safely stored in the fuel rack, because the fuel rack k. is less than the 0.95 limit, both with and without the channel. a 0 e l 3-12 i ~, - --. - _. ,,---,w n

n-.-,-..

--.n-, -n

3.8 REFERENCES

1. R. F. Barry,' 'LEOPARO--A Spectrum Dependent Non-Spatial Oepletion Code for the IBM-1094," WCAP-3269, September 1963. 2. W. R. Caldwell, "P00-7 Reference Manual," WAPD-TM-678, January 1967. 3. H. Bohl, E. Gelbard and G. Ryan, "MUFT-4--Fast Neutron Spectrum Code for the IBM-740," WADP-TM-72 July 1957. 4. H. Amster and R. Suarez, "The Calculation of Thermal Constants Averaged Over a Wigner-Wilkins Flux Spectrum: Description of the SOFOCATE Code," WAPD-TM-39, January 1957. 5. L. E. Strawbridge and R. F. Barry, " Criticality Calculations for Unifor1a Water-Moderated lattices," Nuclear Science and Engineering, 23, 58, 1965. 6. "Large Closed-Cycle Water Reactor Research and Development Program Progress Report for the Period January 1 - March 31,1965." WCAP-3269-12. i 7. " List of Equipment and Apparatus at WREC " Westinghouse Reactor Evaluation Center, February 1967. 8. W. L. Orr, H. I. Sternberg, P. Deramaix, R. H. Chastain, L. Sinder and A. J. Impink, 'Saxton Plutonium Program, Nuclear Design of the Saxton Partial Plutonium Core," WCAP-3385-51,. December 1965. (Also EURAEC-1490) i 9. R. O. Leamer, W. L. Orr, R. L. Stover, E. G. Taylor, J. P. Tobin and A. Bukair, "Pu0 -UO Fueled Critical Experiments," WCAP-3726-1, 2 2 i July 1967. i 3-13' .--,..----,,--,-...-,.,,-.--,w-._ - ~. -.w.m

.-,_m-,,e,..-,

,w-,..m,ry-.-,m,,-,

10. A. F. Henry, "A Theoretical Method for Determining the Worth of Control Rods,'WAPO-218, August 1959.

l

11. P. W. Davis 5n, et al., ' Yankee Critical Experiments Measurements on Lattices of Stainless Steel Clad Slightly Enriched Uranium Dioxide Fuel Rods in Light Water,' YAEC-94 Westinghouse Atomic Power Division (1959).

l

12. V. E. Grob and P. W. Davison, et al., " Multi-Region Reactor Lattice Studies - Results of Critical Experiments in Loose Lattices of U0 Rods in H 0," WCAP-1412, Westinghouse Atomic Power Division (1960).

2 2

13. W. J. Eich and W. P. Kovacik, " Reactivity and Neutron Flux Studies in Mult1 region Loaded Core," WCAP-1433, Westinghouse Atomic Power Division (1961).

14. W. J. Eich, Personal Connunication (1963).

15. T. C. Engelder, et al., Measurement and Analysis of Uniform Lattices of Slightly Enriched U0 Moderated by D 0-H Mixtures,"

2 2 2 8AW-1273, the Babcock & Wilcox Company (1963).

10. A. L. MacKinney and R. M. Ball. " Reactivity Measurements on Unptrturteed, Slightly Enriched Uranium Dioxide Lattices " 8AW-1199, the Babcock & Wilcox Company (1960).

17. Batte11e Pacific Northwest Laboratories, " Critical Separation Between Suberitical Clusters of 2.35 WtX 235-U Enriched UO Rods in Water-2 with Fixed Neutron Poisons," PNL-2438.

)

18. 'OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications," U.S. NRC, April 14, 1978.

e 3-14 l 4

TABLE 3.1-1. StHiARY OF LEOPARD RESULTS FOR MEASURE 0 CRITICAL 5 l I Case ** Referdnce Enrichment H 0/U sity 0"CIII"I U*IC"I * "' 0 er of er ck s Number Number (aton 5) V lume g,jg,3) (c,) ) (,) ( ) ,-2 k ) g i 1 11 2.734 2.13 10.18 0.7620 0.8594 0.04085 1.0287 40.75 1.0015 2 11 2.734 2.93 10.18 0.7620 0.8594 0.04085 1.1049 53.23 1.0052 l 3 11 2.734 3.80 10.18 0.7620 0.8594 0.04085 1.1938 63.28 1.0043 4 12 2.734 7.02 10.18 0.7620 0.8594 0.04085 1.4554 65.64 1.0098 5 12 2.734 8.49 10.18 0.7620 0.8594 0.04085 1.5621 60.07 1.0118 6 12 2.734 10.13 10.18 0.7620 0.8594 0.04085 1.6891 52.92 1.0072 7 13 2.734 2.50 10.18 0.7620 0.8594 0.04085 1.0617 47.5 1.0008 ( 8 13 2.734 4.51 10.18 0.7620 0.8594 0.04085 1.2522 68.8 O.9987 9 13 3.745 2.50 10.37 0.7544 0.8600 0.0406' 1.0617 68.3 1.0010 10 13 3.745 4.51 10.37 0.7544 0.8600 0.0406 1.2522 95.1 1.0025 i 11 14 3.745 4.51 10.37 0.7544 0.8600 0.0406 1.2522 95.68 1.0009 { Y 12 15 4.099 2.55 9.46 1.1278 1.2090 0.0406 1.5113 88.0 0.9889 i

13 15 4.099 2.14 9.46 1.1278 1.2090 0.0406 1.450 79.0 0.9830 H

l 14 16 4.099 2.59 9.45 1.1268 1.2701 0.07163 1.555 69.25 0.9999 1 15 16 4.069 3.53 9.45 1.1268 1.2701 0.07163 1.684 85.52 0.9958 16 16 4.069 8.02 9.45 1.1268 1.2701 0.07163 2.198 92.84 1.0040 i 17 16 4.069 9.90 9.45 1.1268 1.2701 0.07163 2.381

91. 79 0.9872 18 16 3.037 2.64 9.28 1.1268 1.2701 0.07163 1.555 50.75 0.9946 4

19 16 3.037 8.10 9.28 1.1268 1.2701 0.07163 2.198 68.81 0.9809 1 20 8 0.714* 1.68 9.52 0.8570 0.9931 0.0592 1.3208 ' 108.8 0.9912 i 21 8 0.714* 2.17 9.52 0.8570 0.9931 0.0592 .1.4224 121.5 1.0029 i 22 8 0.714* 4.70 9.52 0.8570 0.9931 0.0592 1.8669 159.6 0.9944 i 23 6 0.714* 10.76 9.52 0.8570 0.9931 0.0592 2.6416 128.4 1.0008 24 9 0.729* 1.11 9.35 1.2827 1.4427 0.0800 1.7526 89.1 0.9902 j 25 9 0.729* 3.49 9.35 1.2827 1.4427 0.0800 2.4785 104.72 1.0055 1 26 9 0.729* 3.49 9.55 1.2827 1.4427 0.0800 2.4785 79.5 0.9948 j 27 9 0.729* 1.54 9.35 1.2327 1.4427 0.0800 1.9050 90.0 0.9878 i 4 1ese are Pu02 in Natural UO - 2 !ases 1 though 19 are with stainless steel clad, Cases 20 through 27 are zircaloy. i w

/ TABLE 3.1-2. IESTINGHOUSE UO2 2R-4 C1.A0 CYLINORICAL CORE CRITICAL EXPERIMENTS. PITCM CRITICAL NO* k'If EXPERIMENT CONCENTRATION SUCKING FOR FUEL REGICBf (IN.) OF PINS (LEOPAA0/P0Q-) (p) 20 PARD W 3 1 0.600 0 .008793 489.4 19.021 0.9912 2 0.690 0 .009725 317.0 17.605 0.9941 3 0.848 0 .000637 251.6 19.276 0.9927 4 0.976 0 .006454 293.0 23.935 0.9935 5 0.600 306. .007177 659.9 22.088 0.9927 6 0.600 536.4 .006244 807.2 24.429 0.9937 7 0.600 727.7 .005572 950.2 26.504 0.9940 8 0.600 104. .000165 546.3 20.097 0.9919 9 0.600 210. .007599 607.1 21.106 0.9917 10 0.600 330. .007601 669.5 22.240 0.9916 11 0.600 446. .006661 735.3 23.315 0.9909 12 0.600 657.1 .005809 895.4 ' 25.727 0.9944 13 0.840 104. .007320 321.0 21.772 0.9938 14 0.848 218. .006073 420.5 24.919 0.9925 0.9928 Mean 0.0012 5td NOTES: (a) Feel Reelon Data Enrichment = 2.719 w/o U-235 (b) Thickness of water reflector is that required to attain Fuel Density. 10.41 g/an3 total radius of 50 on for model. Pellet Density = 0.20 la Clad IR = 0.2027 in Clad OR = 0.23415 in (c) Sf =.00527 cm-2 D \\ 3-16

TABLE 3.1-3. BATTELLE FIXED NEUTRON POISON CRITICALSIIII length No. of' Absorber Absorber 8 U"C' Cise Times Assemblies k,pp Type Thickness Width

In Array Cluster of Clusters LEOPARD /PJQ 020 20x17 3

Boral .713 m .645 cm 6.34 m 0.9932 017 22.21x168 3 = a 5.22 cm 0.9944 002 20x18.88+ 1 = 2.732 cm 0.9925 028 20x16 3 S.S. 485 cm .645 m 6.88 m 0.9946 027 20x16 3 S.5. .302 cm 7.43 ca 0.9935 032 20:17 3 S.S. 1.1w/o B .298 cm .645 m 7.56 cm 0.9933 038 20x17 3 S.5. 1.6d/o B T.36 m 0.9931 0028 20x18.075 1 None 0.9956 015 20xl7 3 11.92 cm 0.9942 013 20:16 3 8.39 m 0.9945 = 022 20x15 3 021 20x14 3 6.39 ca 0.9933 4.46 m 0.9946 9 e { 3 .L *7

s TABLE 3.1-3 (continued) BATTELLE FIXED NEUTRON POISON CRITICALSII I STATISTICAL

SUMMARY

Series Number Mean k gg e a Boral 3 0.9934 0.0008 S.S. 2 0.9941 0.0006 S.S. (Borated) 2 0.9932 0.0001 Fixed Poison Total 7 0.9935 0.0007 Non-Poison Total 5 0.9944 0.0007 Overall 12 0.9939 0.0008

This is in units of pitch (Pitch = 2.032 cm).

x Center assembly was 20x16 and the other two were elongated at 22.21x16. + 20x18.88 was one assembly with a boral sheet on two sides. Fuel region data: Enrichment = 2.35 w/o, pellet radius = 0.5588 cm, Clad OR =.635 cm, Wall thickness .0762 cm, Pitch = 2.032 cm. = 0 e e 3-18

TABLE 3.2-1. FUEL BUNDLE CHARACTERISTICS ** Fuel Bundle Geometry 8x8 Active Fuel Height (in.) 150. Rod Pitch (in.) 0.64 Assembly Cross Section* (in.) 5.12 Fuel Rods Material sintered 002 Pellet Diameter (in.) 0.410 Pellet Immersion Density (%TD) 95.0 Cladding Material Zr-2 Outside Diameter (in.) 0.483 Thickness 0.032 Water Rod Material Zr-2 Outside Diameter (in.) 0.591 Thickness 0.030 Spacers Material Zr-4 with Inconel X-750 Springs Number per bundle 7 Fuel Channel Material Zr-4 Inside Dimension (in.) 5.278 Wall Thickness (in.) 0.10

Fuel assembly is comprised of 64 - 0.64 in. x 0.64 in. cells:

62 fuel rod cells and two water rod cells.

- Licensing Topical Report, General Electric Standard Application for Reactor Fuel," NE00-240ll-A-4, 82NE0057 Class I, January 1982.

k 3-19

i Figura 3.2-1 BASIC RACK CELL GEOMETRY AND DIMENSIONS ^ (ALL DIMENSIONS IN INCHES) A J A A A ~ / k D Q G K O M NEd / N / 5 9 @/ / wM e z v i G. l l l /\\W / / /l - v MATERIALS 1. Homogenized Fuel Pin Cells 2. Explicit Water 3 Explicit Stainless Steel 4. Homogenized Water Pod Cell S. Explicit Boraflex' e / 3-20

TABLE 3.3-1.

SUMMARY

OF PERTURBATIONS TO THE MULTIPLICATION FACTOR OF THE 8ASIC CELL Description ak Elfeet k. Basic Cell at 68*F, 3.416 w/o U-235 in .g179 enriched cgnter section of bundle,.020 gm B-10 cm in 0.075 inch Boraflex Calculational Biases Axial effects including natural uranium +.0026 ends and Boraflex cutouts LEOPARD /PDQ Model Bias +.0061 Zr Channel Stored on Fuel Assembly +.0045 Basic Cell including Biases .9311 Tolerances and Uncertainties Minimum Boraflex thickness +.0006 Minimum pitch (i.e., box to box) +.0018 Tolerance on SS box thickness +.0001 Maximum pellet density +.0014 Fuel position uncertainty 0 Calculational uncertainty (ks) +.0022 Total Uncertainty (statistical) +.0032 Extra Assembly Accident (next to rack) +.0098 Maximum, including worst accident .9441 Design Conservatisms Explicit multi-enrichment pin distribution .0125 Mesh spacing effect .0004 Burnable Absorbers (poison) .0300 Spacer Grids .0029 Best estimate .89 3 e 3-21 l [

? } TABLE 3.2-2. FUEL BUNDLE VOLUME FRACTIONS (68F) i .l 4 1 U02 PELLETS .3123 4 i Zr-2 CLADDING =.1113 He =.0138 i HO =.5626 2 i 4 i f I I l l 4 t l e i i t 3-22 \\

~ O Figure 3.2-2 BASIC RACK CELL GEOMETRY AND DIMENSIONS (All Dimensions in Inches) Zr Channel on Fuel Bundle n d h h h / h \\ / oh N QN s v h U$N N y / uR @f N / a w xN'N'N'NNN'@N'xz y ~ \\' y l (5 ~ N N'N'N'N / / / $/ / / / b y MATERIALS 1. Ilomogenized Fuel Pin Cells 2. Explicit Water 3. Explicit Stainless Steel 4. liomogeni::ed Water Rod Cell 5. Explicit Boraflex 6. Zr-II 0 ( 1. 2 9 /1. 0 0 )~ 2 3-23

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~. - PrVE FIGURE'3.3-2 INFINITE MULTIPLICATION FACTOR VS BIO LOADING IN BORAFLEX (.075 inch thick Boraflex) i,, I J.! !. ! i1 i: ! l ! I i i ' I I i i l ~ i l I i i i i i- 'l 'l I i i ir .930 + r. ~ } .i-i jf .j. e i i

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FIGURE 3.3-3 INFINITE MULTIPLICATION FACTOR VS TEMPERATURE

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lE' p

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4.0 SEISMIC ANALYSIS 4.1 Introduction i 4.1.1 Scope of Seismic Analysis The scope of the seismic analysis covers aspects of Sargent & Lundy Specification T-3758, Art'icles 305.4 and 305.8, and includes the following: a. A detailed dynamic analysis utilizing the time-history method for three statistically independent components of dynamic motion. (Art. 305.4) b. Artificial acceleration time histories'were generated by Design Decisions, Inc., and used in the analysis. (Art. 305.5) c. Damping values of 2% and 4% were used for the OBE and SSE, respectively. (Art. 305.6) d. Forces from sliding and overturning due to dynamic loads were found. (Art. 305.7) e. Coefficient of friction values of 0.2 and 0.8 were used in the analysis. (Art. 305.8) 4 4.1.2 Aspects of NRC Standard Review Plan (Ref. 2) include: a. "When the time history analysis method is employed for seismic analysis, two types of analysis are generally performed depending on the complexity of the problem: (1) To obtain maximum responses due to each of the three components of the earthquake motion: in this case the method for combining the three-dimensional effects is identical to that described in item 6a, (which is quoted in the next paragraph, b.) except that the maximum responses are calculated using the time history method instead of the. spectrum method..." (Ref. 2,3.7.2,II, 6b.) b. "When the response spectra method is adopted for seismic analysis, the maximum structural responses due to each of the three components of earthquake motion thould be combined by taking the square root of the sum of the square of the maximum codirectional responses caused by each of the three components of earthquake motion at a particular point of the structure or of the mathematical model. (Ref. 2,3.7.2,II,6a.) I c. "The use of equivalent static load factors as vertical response loads for the seismic design.. is acceptable only if it can be justified that the structure is rigid in the vertical direction. The criterion for rigidity is that the lowest frequency in the vertical direction is more than 33 cps." (Ref. 2,3.7.2,II, 10.) 4-1

4.1.3 Analysis Objectives The objectives.of this seismic analysis are to determine the following during OBE and SSE seismic events: 1. The maximum loads imposed on the fuel storage racks. The following loads are found: a. The maximum forces at the gaps for all assemblies. b. Forces at gaps on individual fuel assemblies. c. Maximum vertical and horizontal forces at rack base for each direction of motion. d. Forces on individual pedestals. i e. Maximum resultant vertical and horizontal i forces at base. 2. The maximum distance the racks will slide and/or lift off. The following are addressed: a. Elastic displacements (horizontal movement). b. Sliding displacements (horizontal movement). c. Liftoff (vertical movement). Summaries of the results are shown in Tables 4.2.1-1,-2 4 -3. In each table the results are grouped in 5 sets which are explained on page 4-10. f The results of this analysis will be used in the mechanical analysis to evaluate the structural integrity of the racks-when subjected to these loads and to determine that the racks will not impact pool walls, spargers, or any appurtenances. 4.1. 4 Scope of this report The loadings considered in this report are: 1. Deadweight of the fuel storage racks and the fuel assemblies with and without channels. 2. Submerged weights of the fuel storage racks and j the fuel assemblies with channels. 3. Seismic loading, both OLE and SSE, as provided by the acceleration time history of the pool floor, the time history data includes seismic, chugging, j and safety relief valve (SRV) discharge. 4-2

Horizontal responses to the seismic motion of 'the ground are obtained by evaluating the loadings for two difference boundary conditions, as follows: 1. The horizontal motion is restrained by a horizontal force equal to 0.2 times the normal force. This is the minimum anticipated friction factor between the rack pedestals and the floor, (Ref. 7). These results give the 1 maximum distance the racks will move during a seismic event. 2. Differential motion between the pedestal and the floor is prevented. This is done by placing a spring between the rack and a fixed point. The horizontal spring constants are determined by combining the flexibility of the pedestal, base, and the boxes. This represents the 0.8 friction condition. Analyses performed on similar racks have shown that a coefficient of friction of about 0.5 is sufficient to keep a rack from sliding appreciably. A run for a 240-cell rack subjected to an East-West SSE event with a friction factor of 0.8 is included in this report. A maximum displacement of 0.117 inch occurred. For this case it is considered to be more realistic to replace the friction element with a spring which represents the horizontal flexibility between the bottoms of the pedestals and the center of the rack. 4.1. 5 Methods of Analysis The vertical seismic analysis was performed using the equivalent static response spectra method. This may be used if the vertical natural frequency is greater than 33 HZ, (Ref. 2,3.7.2,II.10) accelerations of 0.46g for OBE and 0.669 for SSE were taken from the revised response spectra curves of Ref. 1. These values were applied to the deadweight to obtain the total vertical forces. The vertical reaction loads i were combined with the horizontal seismic loads using the square root sum of the squares method as specified in Ref. 2. The horizontal seismic analyses was done using the time history method of analyses in conjunction with time history data supplied by Design Decisions, Inc. The time histories were generated in accordance with Ref. 1, Article 305.5. This method accounts for the non-linearities inherent in the spent j fuel storage racks, which include: i } 1. Fuel-to-rack wall impacts 2. Rack sliding 3. Vertical impact due to rack tipping 1 43

The time history analysis was done using a spe'cial purpose computer program "RACKOE". This program was developed specifically to analyze fuel storage rack behavior resulting from seismic disturbances. The program solves the equations of motion explicitly using Euler's Extrapolation Formula. A formal presentation of the algorithm and program verification is provided in Appendix E. The nose of the fuel is considered to rest at the bottom of the cell's center with a simple support (hinge) between the fuel and the rack. There is a gap between the fuel and the rack wall along the sides. The gaps between the fuel and the box walls can close, thereby causing impact with the rack walls. The space between the fuel and the rack wall is filled with water. As the fuel and the box wall move relative to each other, hydrodynamic forces are set up due to the acceleration of the water. These forces are exerted on the fuel and rack structure, and tend to mitigate impact forces. Methods described by Fritz (Ref. 3), Dong (Ref. 4) and Stokey et al (Ref. 5) are used to quantify these hydrodynamic forces. The effects of hydrodynamic forces between the rack and the pool wall are included in the analyses. Damping values used for this analysis are taken from Regulatory Guide 1.61, (Ref. 6). The rack boxes are weldt.d together with " diameter fusion welds. When the welds are stressed there will be some localized deformation. The damping values are between those for welded steel and bolted steel structures. In the interest of conservatism the lower values for welded steel structures are used, i.e., 2% and 4% of critical for OBE and SSE, respectively. Friction between the rack and the pool floor is handled by a special friction element. The normal force on this element is the force in the pedestal springs which, due to rack tdpping, can be greater than the deadweight of the rack. Details of "RACKOE" and how the various parameters are used in the computer code are contained in Appendices C, D and E,' as follows: Appendix C - Outlines the methods used for the fluid analysis and demonstrate how the fluid effects are incorporated. Appendix D - Documents experimental verification of fluid effects. Appendix E - Describes and provides verification for the non-linear impact analysis. 'I ine-histories of the pool floor accelerations are used as input. Damping is described in Attachment E-l. Sliding capabilities of "RACKOE" are described in Attachment E-2. 4-4

4.1. 6 Models for Seismic Analysis A finite element representation of the rack with fuel assemblies is shown in rigure 4.1.6.-l. ()Representsmassnodes 1 Base (horizontal) 2-6 Rack (horizontal) 7-11 Fuel Assy (horizontal) 12 Rotary Inertia 13 Vertical mass of rack plus contents ~~ Represents flexible elements 1-5 Rack 6-10 Fuel Assy 11 Horizontal restraint 12,13 Vertical supports 14,15 Vertical supports l L. Represents gap el'ements: flexibility and local damping between fuel and rack walls. ' MHrw = Hydrodynamic mass (rack to wall) ,, I Mhrf = Hydrodynamic mass (rack to fuel) The following runs have been made and the important results shown in Tables 4.2.1.1-1,-2, and -3. Preliminary runs showed that the channeled fuel yielded higher forces than the unchanneled, so that only results for channeled fuel are tabulated for 0.2 friction factor. For unchanneled fuel, only the results for a 240-cell rack fuel are included. Channeled fuel (180-cell and 240-cell racks): Coefficient of friction = 0.2 OBE: North-South and East-West SSE: North-South and East-West Non-sliding (coefficient of friction = 0.8) OBE: North-South and East-West SSE: North-South and East-West Unchanneled fuel (240-cell rack): Coefficient of friction = 0.2 SSE: North-South and East-West Non-sliding (coefficient of friction = 0.8) OBE: North-South East-West SSE: North-South East-West 4-5

4.L7 Equipment Description The general arrangement of spent fuel storage racks for the La Salle County Station - Unit 2, is shown in _Ref. 18. The' array is composed of.20 modules designed to hold 4073 fuel assemblies. The modules have capacities varying from 174.to 240 assemblies each. A 240-cell rack is shown in Ref. 19. Each module is composed of vertical stainless steel thin-wall rectangular boxes welded together with local fusion welds. All boxes are about 6" x 6" x 168" long. Each box will hold one fuel assembly. A fuel support plate-is welded to the bottom of each fuel box. This plate has a chamfered hole which provides lateral location for the bottom of the fuel assembly and provides the coolant passage for the fuel. Each rack is supported on four vertically adjustable pedestals, which are located at the corners of the rack. There is an additional pedestal at the center, the primary purpose of which is to raise the natural frequency of vibra-tions in the vertical direction. The center pedestal is not adjustable. A rack will be installed by first lowering it onto the center pedestal, and then adjusting the corner pedestals. Figures 4.1.7.-l and -2 _are plan and elevation views, respectively, of a cell, showing.the poison wall design. l These are used in finding the rack weight. 4. 1. 8 Material Properties The following values are used in the seismic analysis: Material Density Young's Shear Modulus Modulus (LB/CU.FT) (PSI) (PSI) 304 SS 501.0 28.0E6 10.77E6 Zircaloy 409.0 13.0E6 5.00E6 Concrete N/A 3.5E6 N/A Water 62.4 Boraflex 104.4 e 46

0 0 .,m. s.w -m + DI R ECTIC N CF LDAD / l l' 5 10 l 5 h\\ (Q 9 l 4 [b, e ' @ $H rn 9 P 4 @P9 87l - - - 0 - . 4----r s o, ;, 7 .@ h /Vrh d/ si g) jFvers 8 (M,_ _ T ' h, 9 12)) F herir f Fver FIGURE 4.1.6-1 9 e ~ 4-7 ,m._ ~..

FIGURE 4.1.7-1

6. 2 S*5 NeM. Pi7cN 4

7 Ski + N6 poisott thrH x/39* Lou 6 4(;yo ARCA RwVcnoN hw 6 NorcdES) (,020 9r/cm ' 8 ~/0) w ~ R + &M, ,,/U,,,,,,,,v ',, i /,,,, 6Q I q 7 c ' & M ' "f'F W ' t " J W 'e v '? n e r o ? W eeeW'wa s 3 .q a l ~ -(.o75 cms. R.isoN)h .oce sw. . oa 2 su. (090 nom. KkLL) n . 086 Min'. ~ . 0 9 'l /. W. ,[ N' s 4-(d000 Abd S UsRE-)..E202Mw' f Q w 6.t.20,41Ax. y o l At/SI? AGE /%1ANUFA CTUfcE D Box Citrs/O~r t. -m e-q 6.I80 t.020 Sc?vARE z V ) y, w / } L AVERAGE AcRoss RAISELLAREAs As MAMUFACTURED 6 2 6~f 2 02 5 SQL/ARE) v u b / e eb 4 W s c ~ s e qy i 1<, 3 48 l

FIGURE 4.1.7-2 g liil T({_,Af-JOCAL FVDM/ WCLD n J g 6 FAms, asen w'At.L. 525 m A y t; I i a I I I 1, N .8 8 R. ~ v pq rYP -i N t I i l I I l ^ n- - l , y [ e.56 e ,g T, zi I h 4]! NorcNe~S i b Q N{ f 3 <hACEs 'e i d M bI BORAFL E X \\ \\ 1 poisod g" N AAA TEPIAL I f' ~', Y G _ 020 p<.lctr? Bf b ( u l d S e m I Q h TO TA L hi?EA l l ~ W

17. 3 h c.,

2/o fsr A MsTcHrs l \\ y 729 l L4 t [ T \\ e%T 3 d i 1 -s I l 1 h (h I 0 4 4 1 S t n i g bj t Y 4. D N s l V M l,N a u y y b.-. d y t g r-u i l 1 l 4-9

4.2 Analysis Objectives To determine the following during OBE and SSE events: 1. The maximum loads imposed on the fuel storage racks. 2. The maximum distance the racks will slide and/or lift off. 4. 2.1. Summary of Results Tabulated results are grouped and identified by " sets" numbered 1 through 5 as shown on Tables 4. 2.1.-1,-2 and -3. The values in each set are explained below. Set #1 - Maximum forces (KIPS). These are taken directly from the computer output " Summary for Complete Time History". They are the maximum values during the period of the time history. For the gap elements each value.is the maximum force acting between the fuel and the rack wall at the gap position. The vertical pedestal forces represent the maximum force on two pedestals. The horizontal force is the maximum horizontal restraining force acting on the base of the rack. Set #2 - Loads on individual fuel assemblies and pedestals. The maximum contact forces are the forces of Set #1 divided by the number of fuel assemblies in the rack. The pedestal forces are the forces of Set #1 divided by two. This is because liftoff occurs during each seismic event shifting the total reaction to two pedestals. Set #3 - Maximum forces at the base of the rack. The Fvert values, NS, EW and VT, are the values of Set #1 minus the submerged weight. The Fhoriz values are taken directly from Phz in Set #1. The SRSS values are calculates using: VERTICAL: SRSS = SWT + [(Fns)^+ (Few)^ + (Fvt)* HORIZONTAL: SRSS = f(Fhz)*ns + (Fhz )' ew WHERE: SWT = SUBMERGED' WEIGHT Fns = (Fvert)ns = (Fvt)ns - BF Few = (Fvert) ew = (Fvt) ew - BF ' Fvt = (Fvert)vt = (g f actor) *DWT 4-10

For conservatism all forces are considered to act on,two pedestals, since when the maximum forces occur on a pair of pedestals the other two pedestals may have lifted off the ground. The values for Set #4 are one half the values in Set #3. Set #5 - Horizontal and Vertical Movement of the Base (Inches) Elastic - The amount the base will deform as a result of the internal flexibility of the rack when the pedestals are restrained from horizontal motion. Sliding - The amount the base will move when the bottom element of the rack is considered to be rigid and a 0.2 friction factor is used to restrain movement in the hori-zontal direction. { Liftoff - The maximum distance the pedestal will move vertically off the floor during the seismic event. The values *DWT, BF and SWT are the deadweight of the dry rack and fuel, the buoyant force that acts on these, and the i submerged weight of the rack, respectively. They are taken from Section 4, Summary of Weights. The friction forces are the maximum horizontal forces developed at the base of the rack when the friction factor of 0.2. The maximum forces on one pedestal for the seismic disturbances are for the 240-cell rack. The following values are SRSS, sums taken f rom Table 4. 2.1-2. Vertical (lbs) Horizontal (lbs) OBE 431,700 92,522 SSE 497,687 118,660 The maximum values for other outputs are: SSE Elastic Displacement: 0.503 in. 240-cell rack East-West Sliding Displacement: 1.888 in. 240-cell rack (Unchan) North-South Liftoff

0.964 in. 240-cell rack North-South OBE Elastic Displacement: 0.428 in. 240-cell rack (Unchan)

East-West Sliding Displacement: 0.106 in. 180-cell rack East-West i Liftoff 0.502 in. 240-cell rack (Unchan) East-West l l l t ( 4 11

TABLE '4. 2,1-1

SUMMARY

OF RESULTS FOR 180 CELL RACK - CHANNELED, (FILE SUMISO)

SET C1 - MAX. FORCES (KIPS)----------------------

A~ GAP ELEMENT # PEDESTAk D;R EVT 1 2 3 4 5 Fvt Fhz NS OBE 0.0 9.1 !?,7 26.9 4?.1 4S5.0 6?.0 EW OBE' O.0 0,2 10.3 18.8 40.5 307,4 101.0 NS SSE 7.2 40.1 59,8 66,8 80.1 625,5 06.2 EW SSE 6.0 39,4 58.2 63,7 77,1 506.6 152.3

SET #2 - FORCES ON INDIVIDUAL F/A'S (LBS) AND PEDESTALS (HIPS)---

NS OBE 0, 51, 109, 149, 273, 242,5 34,5 EW OBE 0 1,

57. 104, 225, 193.7 50.5 NS SSE 40, 223, 332, 371 445.

312.8 43.1 EW SSE 33, 219, 323, 3'_74

428, 253,3 76.1

-SET #3 - MAX. FORCES AT EAIE (LES) -SET #5-POVEMENT AT BAIE (IN5) Fver t F ric e i: ELASTIC S_ID1 % LIFTOF.: ilS 7.'LE ?!9,650, 9., c i.: 0. - - -- --------- e. S+ 7 0.035 0.217 EU GrE / 01, c 41- . s 1, 0.:.. - -- -------- - :t. 1. 1 0,10' 0,127 VT CEE 66,550 e SRS5 574.:10 2 2,32C 14S SSE 499,1 0, Le, 2 00, -- ----------- U. 41 3 0.720 0.626 EW 55E 3E O, 2-10.

5 2, 0 0 0. - - - - ---------

181 0,870 0.602 UT SSE 94.034, O. SRSS 760,810. 175,000. -5ET #4 - MAX, F0RCEb O4 PEDESTAL (LE5)-- NS OPE 179,320, 34,500, EW 05E 130,520 $0,500 DWT=144,671 LB3, EF = 1E.320 LDS. VT OBE 33,270 O. SWT=126,352 LES, SRSS 237,450, 41,160 FRICTION FORCEE-(LE) t 0.2 FACTOR NS SSE 249,575. 43,100 NSOBE = 26,710 EWOBE = 25,430 EW SSE 190,100, 76,150. N3SSE = 36,760 EWSSE = 25,270. VT SSE 47,018 O. SRSS 380,405, 87,500, 4-12. h .-g. wr .-r ?

.~ TABLE 4,2:.1-2 SUF. MARY OF RESULTS FOR 240 CELL RACK - CHANNELED. (FILE SUM 240)

SET

1 MAX, FORCES (KIPS)- -

AT GAP ELEF.ENT# PEDESTAL DIR EVT 1 2 3 4 5 Fvt Fh: NS OEE 0.0 0.0 0.0 8.7 39.5 553.7 121.1 EW CEE 4.6 37.7 47.7 57.4

96. 6 736.5 140.0 NS 55E 4.8 45.3 72.3 E5.9 110.0 809.2 159.2 EW SSE 6.E 42.4 66.9 77.3 104.7 675.3 176.0

SET #2 - FCRCEE ON INO1VIDUAL F/A'S (LES) AND PEEESTALS ( t. I PS ) ---

NS OEE 0. O. U. '.: 5.. 165. 279.4 60.6 EW ODE 19 157 199 2.v. 401. 25.6.3 70.0 NS SSE 20, 129 301 3MB. 455'. 404.6 79.6 EW SSI 2 './. 177 "77 L 2.. 406. J37.9 55.0 -SET v 3 - r-M , _ FE E1: 4.- E-i.7.i (LLI: SET v5-C. 2.9E.,T AT I.,? E (Ir45. F, +r t T r. o r 1 : Eu45 TIC SL'DI.'.G LIF~CrF NS CEE 09 :,5'. 4. ! 1,10 0 -- - - ---- ---- -0. 3 :' 4 0.0 M 0.137 EW CEE 5.6,384

0, 0 0 0. ------------- 0. 4 0 0 0.054 0.270 VT OCE S3540 O.

SRSS 663,427 185,100 NS 55E 641,084 159,200,-------------0.400 1.021 0.964 EW SSE 507,484 17/., 0 0 0, ------------- 0. 5 0 3 1.651 0, 6S2 VT SIE 125,121 O. r

e....,..

.. /. m..g.

:.c_

v. ..s. -1.ET .4 - f14/. F I-::Ci_1. GN F ZL'LI.Tet t I: )-- N ; 0:.E . i t. 2 'i.. .0. 50.$. EW ODE 2 34, l i. 2 70,e:30 DWT=192,495 LES. Dr = 24,578 LES. VT 02E 44,274 O SWT=168,116 LES. SRBS 431.700 L,522. FRICTION FORCES (LC) e 0.2 FACTCR L i. SSE 12c,540. '4.r Oe, NSOSE = 54,270 EWCBE = 33,620 Ew ? i.E

51.642.

S_. 000 G5SSE = 36.520 EWL5E = 34,640 VT SIE

62,550 O.

SRSS 497,687 118,'660 A-13 ,,.--.3m, .e '~

TABLE 4.2.1-3 S'JMMARY OF RESULTS FOR 240 CELL RAOr< - UNCHANNELEC. ( FILE' SU.*.240U)

SET 61 - MAX. FORCES (KIPS)----------------------

AT GAP ELEt1ENTtt PEDESTAL DIR EVT 1 2 3 4 5 Fvt Fh: NS ODE 13.6 44.4 49.0 50.4 50.5 542,0 100.7 EW OLE 5.1 35.B 47.5 45.4 52.4 620.6 149.7 NS SSE 21.9 55.5 65.3 50.5 63.3 714.5 152.7 E J 5 5E 29.5 53.E: 65.2 59.6 62.1 733.4 123.8

SET #2 - FOPCES ON I f L I VI D:.' AL F / A

S (LDS) A.ND FEDESTALS r: 1 F 5 ) ---

NS OBE 57, 105 204 210 210 271.0 50.4 EW 03E 21, 149, 198 107 2:C. 5;0.3 74.9 NS SSE 91 231, 272 223, 295 357.2 76.4 EW c.5Z 120, 24t, 280 1*?. ./t.9 %6. 7 64.4 -SET #3 - MAX, FORCE. Ai I41. ( i. 21., -? ET

t. 5-M.. E. ;ErwT A~

DL5E (INS. Fver t 8 *.cr 1: Et.AST C SLIL:t.G _:FTOFF N3 02E 390,070, 10 :$,700 - - - -------- :5.252 0.356 EW 03E E.4,670 2 4 9, 7 00, - --- --------- 0. 4 6 4 0.502 VT GEE 79.752 C 5 :35 766,390 133,4: 3. Ni. SLE 012,5'O. 1 1 2, 70 0, - - - - ---------0. 2 4

1. Si. i' O 75 ?

EW 5?E 5'*1,470. . 0. 5 00 ------------0 . 0 6 3 1.E76 U.732 VT SEE 11 49:. O.- SRDS v61,108c. 7 *, e -SET #4 - MAX. F.';;C ES ON :TDE'.TAL I5.-- te 3 C DE .95.035 50,35(, EW ODE 204,335 74 E.5e DWT=173,374 LE.S, DF = 21,433 LDS. VT ODE 3?,076. O. $WT=151,920 L35. SRSS 330,440 90,210 FRICTION FO.CES (LB) t 0.2 FACTOR NS SSE 231,285 76,550 NSSSE = 56640 .e EW SEE 290,735, 64,400 EWSSE = 69340. VT SSE 56,547 O. NOTE: UNCHANNELLED FUEL WAS NOT RUN FOR TFE OBE SRSS 434,400, 9>,889 CONDITION. 4-14 l l i . - -. - ~.

4.3 References 1. Specification T-3758, Spent Fuel and Special Storage Racks, La Salle County Station - Units 1 and 2, Commonwealth Edison Company, Project No. 7043-73. 2. U.S. Nuclear Regulatory Commission, Standard Review Plan 3.7.2 " Seismic System Analysis," Revision 1, July, 1981. 3.

Fritz, R.J.,

"The Effects of Liquids on the Dynamic Motions of Immersed Solids," ASME February, 1972. 4.

Dong, R.R.,

" Effective Mass and Damping of Submerged Structures," UCRL-52342, L.L.L., April, 1978. 5.

Stokey, W.J.,

Scavuzzo, R.J.

and Radke, E.E., " Dynamic Fluid Structure Coupling of Rectangular Modules in Rectangular Pools," ASME Special Publication PVP-39, 1979. 6. Regulatory Guide 1.61, " Damping Values for Seismic Design of Nuclear Power Plants," October, 1973. 7. Rabinowicz, E. " Friction Coefficients of Water-Lubricated Stainless Steels for a Spent Fuel Rack Facility", Study performed for Boston Edison Co., November, 1976. 8. ASME Boiler and Pressure Vessels, NUCLEAR VESSELS, Section III, 1980 edition. 9. G.E. Technical Paper 22A5866,.Rev. Dec. 26, 1979. Appendix II, FUEL ASSEMBLY STRUCTURAL CHARACTERISTICS. 10. R.D.

Blevins, Ph.D., FORMULAS FOR NATURAL FREQUENCY AND MODE SHAPE, Van Nostrand_Reinhold Co.. New York, N.Y.,

1959. 11. S. Timoshenko, S.Woinowsky-krieger, THEORY OF PLATES AND SHELLS, McGraw-Hill Book Co., Inc., New York, 1959. 12. R.J. Roark, FORMULAS FOR STRESS AND STRAIN, McGraw-Hill Book Co., New York 3rd Edition, 1954. 13. C.M. Harris and C.E. Crede, Editors, SHOCK AND VIBRATION HANDBOOK, McGraw-Hill Book Co., New York, 2nd Edition, 1976. ,14. R. Szilard, THEORY AND ANALYSIS OF PLATES, Prentice-Hall, Inc., Englewood Cliffs, N. J., 1974. 4-15

15. G.E. Drawing No. 829E431 Fuel Bundle. 16. G.E. Drawing No. 829E276 Channel. 17. Boraflex Neutron Shielding Material, BISCO REPORT No. N-38, Bisco Products, Inc., 1420 Renaissance Drive, Park Ridge, IL 60068. 18. UST&D Drawing No. 8601 General arrangement of ~ storage racks. 19. UST&D Drawing No. 8601-3, 240 cell rack. e 4-16

5.0 Mechanical Analysis 5.1 Summary The new spent fuel racks are free standing poison wall design. These spent fuel storage racks provide smooth full length square storage cells of stainless steel in a welded honeycomb structure. Each storage cell, except on the periphery of the complete array, is bordered on all four sides by Boraflex neutron absorbing poison sheets sandwiched between adjacent cell walls. Each rack is provided with five pedestals. Only the four corner pedestals are analyzed to take the rack loads and are screw adjustable for rack leveling. UST&D provides the appropriate tool to make these adjustments from the surface through the cells of the pedestals. The male thread is chrome plated to Federal Specification QQ-C-320B Class 2a (.0002 to.0005 thickness). Our experience shows that this is the protection needed to prevent thread seizure during rack leveling, especially since the rack weight must be supported on the installation crane while pedestal adjustments are being made. After leveling operations are complete, the pedestal thread is no longer a moving part. The-height of the bottom of the rack above the pool floor, resulting from the necessary vertical dimension of the pedestal structure, provides adequate underneath space for cooling water flow. The individual racks are sized to fit into the space avail-able in the pool. Where rack pedestals coincide with leak chan-nels or floor liner seams, there will be floor plates under these pedestals which are grooved on the underside to span the trenches or floor liner seams and transfer loads to acceptable floor areas. The rack sizes are dimensionally compatible with overland trucking constraints on width and height. 5.2 Description of New Spent Fuel Racks 5.2.1 Module Construction The La Salle rack design is a honeycomb configuration of identical stainless steel boxes with sheet Boraflex poison material captured between the side walls of all adjacent boxes. To provide the space for the poison sheet between boxes, a double row of matching flat round raised areas are coined into the side walls of all boxes. The raised dimension of these locally formed areas on each box wall is half the thickness of-the poison sheet. The boxes are fused together at all these local areas. The poison sheets are scalloped along their edges to clear these raised areas. 5-1 4 _,,_,m

The poison sheets are axially centered on the active fuel region of the stored spent fuel assemblies. They are physically captured, as stated above, between adjacent box walls and within the double row of raised areas on the box walls. In addition each is effectively contained axially by a narrow sheet of stainless steel positioned at the bottom of the poison and welded across one of the two adjacent box walls. This sheet is the same thickness or less than the poison material. Between adjacent racks each poison sheet is held in place on each outside box of one of the racks by a thin cover of stainless steel which is welded intermittently, all around, to the box wall. All of the poison in this design concept is unsealed. Each fuel storage cell has a welded-in bottom plate to support the stored fuel assembly. It has a chamfered central hole to accept the tapered nose of the fuel assembly and provide for the cooling wster flow. The storage cell material is.090" thick stainless steel. The sheet Boraflex poison material is nominally 10 .075" thick with a probable B loading of.020 gm. per sq. cm. The thin stainless steel cover sheets are.025" thick. MATERIALS 1. Poison Design a. Spent Fuel Boxes, Channels, Stainless Steel Poison Cladding, Bottom Plates, Type 304 Pedestals, Shims, Lead-In Guides, and Rack Appurtenances b. Poison Material Boraflex (See Appendix B) 2. Special Tools 1 a. Rack Lifting Stainless Steel Type 304 b. Spreader Bars for Rack Stainless Steel Horizontal Lift and Rack Type 304 Upending c. Vertical Spreader Stainless Steel Type 304 d. Cable slings and Straps Commercial as required l 5-2

5.2.2 Rack Fabrication The fabrication and' assembly of the Spent Fuel Storage Rack is manufactured in strict adherence to in-house controls and tight tolerance, as well as those described within the client specification requirements. Our expertise, as well as experience, has enabled us to control growth, shrinkage and distortions by means of close tolerance manufacturing and special fixturing. A majority of our parts and sub-assemblies are machined to tight tolerance, thereby preventing, wherever possible, undesirable accumulation of tolerance and unacceptable deviation from required alignment. Finally, our experience has proven that the added time incurred by close tolerance machining is offset by the relative ease in assembly and final function testing. The following are typical requirements used in fabrication and assembly, as well as general informatioh used in our manu-facturing: The fuel boxes and water boxes are formed in channel sections, fusion welded continuously full length, and constructed out of two (2) 4 full length sheets. Each box (fuel, water) is constructed out of two (2) sheets (to form channel section) continuous over their total height.- All interior surfaces of the fuel boxes shall be smooth (125 AA microinches). All weld beads shall be flush with interior surface. There shall be no protrusions inside the fuel boxes that will interfere with insertion or withdrawal of fuel assemblies or to result in marring, scoring, or other damage to the fuel assemblies. Racks shall be fabricated by welding unless otherwise agreed by Commonwealth Edison. The fabricated racks shall comply with all requirements of the Specification and the CECO. approved drawings. Racks shall be fully assembled into their final module size and checked for compliance with dimen-sional requirements, alignment, and clear-ance in our shop prior to shipment to the jobsite. Any errors or discrepancies dis-covered shall be corrected prior to shipment. Fabrication and installation of the racks shall conform to the general requirements of Sub-section NF of Section III, Division I, of the ASME Boiler and Pressure Vessel Code for Class 3 component supports. 5-3

Austenitic stainless steel shall be supplied in the solution-annealed condition. Except as required and allowed during welding, austenitic stainless steel material t shall not be heated above 350' F unless it is subsequently given a full solution annealing as the material Supplier's recom-mended temperature and holding period followed by water-quenching from the annealing temper-ature. Cutting, forming, welding and handling of materials shall be performed under the supervision of experienced personnel quali-fied to work with the specified materials. Flame cutting is not permitted. Contact with dissimilar metals, low-melting point metals, chlorides, halogens, sulfur, phosphorus, and other potentially harmful materials shall b'e avoided. Drawing dimensions, geometric requirements, and tolerances shall be interpreted and defined in accordance with ANSI Y14.5. Projections such as weld reinforements will not be permitted on the inside surfaces or envelope of the storage cell. Corners shall be rounded. Sharp edges shall be deburred or chamfered. Weld splatter, chips, burrs, and other foreign matter shall be removed from the base metal in their entirety and i the affected surface restored to a 125-micro-inch (maximum) finish. Abrasive blast-cleaning is not permitted. The inside of fuel boxes shall have a commercial 2B finish, except where the fusion welding of the longitudinal seam weld is applied. Reworked interior surfaces that may come _ / M. /-/NM" in contact with the fuel assemblies shall have a surf ace roughness not exceeding 125 AA microinches. (Reworked interior surfaces /5 07 f for the failed-fuel and miscellaneous parts racks may have roughnesses not exceeding 7- 250 AA microinches). The interface between the lead-in guides and the fuel storage boxes shall be blended to provide a smooth entrance and exit for the fuel assemblies. 5-4

All interior edges of the fuel boxes, lateral and vertical, which may contact the fuel assemblies, shall be finished to a 1/R" minimum radius and chamfer. If they are chamfered, blending of the intersection edges is required. Threads shall be clean with all burrs and ragged edges removed. All male threads will be chrome plated. Each spent fuel storage rack assembly shall be prominently marked by means of Vibratool or stainless steel identifi-cation plate welded to the rack. All welding procedures and welder quali- -fications shall be in accordance with the ASME Boiler and Pressure Vessel Code, Section IX, and/or applicable sections of AWS Section Dl-1, Structural Welding Code. 5.2.3 Quality Assurance UST&D will provide documentary evidence of material trace-ability, inspections, and tests for each rack / shipping piece in accordance with our Approved Quality Assurance Program. This-documentation will be submitted with each shipment, of finished product, to the job site. I Quality Assurance Plan Outline This document describes.the Quality Assurance Program Plan which U. S. Tool & Die, Inc.-(UST&D) follows to control the activities required for the design, procurement, fabrication, inspection, packaging and shipping of nuclear' power plant struct-ural components involving Seismic Category I structures, systems, and components, including their foundations and supports. Classification of structures, systems, and components shall be as follows: Modification to existing structures, systems and components shall be designated the same seismic classification as the existing system. New structures, systems and components shall be designated a seismic classification in accordance'with the guidelines in the current edition of the USNRC Regulation Guide 1.29. Defined are the management systems, procedures, and controls established to assure conformance with applicable section of Nuclear Regulatory Commission Regulation 10CFR50, Appendix B, Quality Assurance Criteria'for Nuclear Plants; American National 5-5

Standards Institute N45.2, " Quality Assurance Program' for Nuclear Power Plants"; American Society of Mechanical Engineers; Boiler and Pressure Vessel Code, Section III, Nuclear Power Plant Componenta and applicable Nuclear Reguistory Commission Guides. This plan is subject to continuous review and shall be revised when required to reflect latest changes in Quality Assurance Policy and Planning for fulfillmel.t of Contract Requirements. Definitions of terms used in this plan are consistent with those in ANSI N45.2, Section 1.4. The UST&D-Quality Assurance Program main objectives are: To establish and implement methods for controlling design and procurement activities, and analytical studies for assurance that requirements for design, safety, and materials, and fabrica-tion are defined and correctly translated into design documents, procedures and instructions. To establish and employ selection, classification, and identification practices for materials, components, parts, and processes to be used for controlling and verifying quality of items and services throughout the project and specify' trace-ability, quality, procurement, and certification required for the components covered by this document. To establish and implement planning and control procedures required to assure that design, fabrication, processing, inspection and test activities conform to the latest drawings, specifications, procedures and instructions approved by Client, if applicable, or Engineering as required. It is the endeaver of UST&D to design, procure, fabricate and deliver quality equipment in accordance with contract require-j ments by systematic planning, controlled execution, and effective quality assurance supervision. The desired quality will be achieved by providing clearly defined design documents, thorough checking, complete analysis, adequate inspection coverage for proper process control, proper l auditing of design and fabrication adequacy, and adequate records to document product quality. Control of quality shall be implemented by adherence to l policies and procedures described in this Quality Assurance Program Plan. l UST&D recognizes the Client's right to audit this Quality Assurance Program as necessary to comply with the Client's Obligation under the applicable codes, standards, and regulations. Necossary documentation and records shall be furnished the Client j on a timely basis. The Client shall normally schedule audits 30 days in advance or in accordance with contractual requirements. 5-6 l 1

UST&D iq to design, procure, fabricato, certify, and deliver nuclear power plant structural components in accordance with the contract ordering data. The principal function of Quality Assurance is to assure that an effective Quality Assurance Program is working. Quality Assurance (QA) monitors, by review and audit, the task of each project. The significant tasks selected for moni-toring are based upon QA assessment of program work scope and contract objectives. QA reserves audit authority to propose and implement hold points. Auditing is the technique utilized on all projects covered by this QA Program Plan. QA has authority to identify problem areas, initiate and recommend corrective or preventative action for resolution of problems at all levels of Project Management, Engineering, Procurement, Manufacturing, and Quality Control organizations. The Quality Assurance organization has the vested authority to prevent non-conforming work or components from being incorporated into the final product. QA will be implemented by all Managers. Assistance will be provided by QA to consult, train and indoctrinate personnel, as required. Significant failures and malfunctions will be reported to the Client through Project Management. Thorough analysis will be conducted to determine the cause and appropriate action required to prevent recurrence. The failure analysis will be formally reported to the Client with pertinent information pro-vided as required. Action will be implemented expeditiously to resolve all problem areas. Documentation of all procurement, quality control, and ' quality assurance records will be kept on file with trace-ability to components being reviewed or inspected. 5.3 Mechanical Analysis 5.3.1 Basis for Analysis The effect of earthquake on the free standing Spent Fuel Storage Racks is addressed in the Seismic Analysis (Ref. 2). Section 4.0 explains the method and the computer program which requires earthquake input in the form of acceleration time histories. SRV and chugging forces are included. Loads and displacement values, horizontal and vertical, are generated. The loads from the Seismic Analysis are used in the Mechanical Analysis for rack stress calculations. These loads ~ are also used to determine the acceptability of the pool floor loadings which in turn are to be used in the determination of the pool structural integrity. 5-7

5.3.2 Dropped Fuel Bundle Analysos These analyses consider both the straight drop and slant drop of a fuel bundle. These analyses are supported by actual physical tests perfor=cd on the rack sections. The analytical results were confirmed by the tests. 5.3.3 Results of Dropped Fuel Bundle Analyses The results of these analyses confirm that 'he suberit-t icality array is maintained in the event of this accident. 5.3.4 Summary of Results The detailed structural analysis shows that the spent fuel storage equipment meets the requirements of the speci-fication. Stresses in the internal rack welds are computed due to both deadweight and seismic loads. Seismic loads are the peak reactions from the seismic analysis. Seismic stresses are calculated using the SRSS method and are combined with the deadweight stresses. The analysis of any external loads such as the drop of a fuel bundle are supported by physical testing. The result of any of these accidents would not be catastrophic and would-not destroy the suberitical array in the pool. Any resulting damage to the racks would be minimal. The maximum stresses for the rack and rack components are listed in Table 5.3.4.-l along with the allowables, and factors of safety. The Level B allowable stresses are 77% greater than the Level A allowable stresses and the Level D allowable stresses are 167% greater than the Level A allowable stresses. 5.3.4.1 Methods of Analysis The analyses use the loads developed in the Seismic Analysis Report (Reference 2) and applies them to high stress portions of the rack and supporting structure. Stresses (axial, shear, bending and torsion) are then evaluated and compared to allowables to determine structural adequacy. 5-8

TABLE. 5. 3. 4.-l STRESS

SUMMARY

ALLOW-STRESS STRESS ABLE SAFETY COMPONENT TYPE LOAD (KSI) (KSI) FACTOR RACK WELDS 240 B-B OBE 13.60 29.06 2.14 INTERNAL RACK WELD SIIEAR SSE 19.15 29.06 1.52 (Fusion) 240 C-C 20.10 29.06 1.44 INTERNAL RACK WELD SIIEAR SSE 25.38 29.06 1.14 (Fusion) 180 C-C-SSE 17.86 29.06 1.63 INTERNIST RACK WELD SliEAR (Fusion) PEDESTALS RACK BOTTOM PLATE TO BOX OBE 22.50 29.07 1.29 WALL (0.09" x 0.125" Weld) S1! EAR SSE 27.50 29.07 1.05 PEDESTAL TOP PLATE TO RACK OBE 19.62-29.07 1.48 BOTTOM PLATE (k" Weld) SIIEAR SSE 25.11 29.07 1.15-1" T!!ICK PLI,TE BENDING SSE 12.04 31.8 2.64 T11 READ (EXTERNAL) OBE 8.84 9.87 1.11 S!! EAR SSE 10.21 10.73 1.05 TIIREAD (INTERNAL) OBE 7.22 9.87 1.36 SilEAR SSE 8.34 10.73 1.28 1 o 5-9 C

5.3.4.2 Dapign Criterin According to the equipment specification, Reference 1, the design and analysis of the spent fuel and special storage racks shall be performed in accordance with the requirements of the ASME Boiler and Pressure Vessel Code,.Section III, Division 1, Subsection NP (Ref.9). The racks shall be classified as ANS Safety Class 3, ASME Code Class 3 component supports, and Seismic Category I structures. The analysis is to be performed to include anticipated loadings such as seismic forces, pool hydrodynamic forces (SRV, CHUGGING, CONDENSATION, OSCILLATION), pool slosh forces, impact forces from accident and abnormal occurrences, thermal forces, applied loads and deadloads. The accident and abnormal forces shall include cases (a), (b), (d), (f), [h), and (i) of paragraph 6. 4. 2.1. 3 of ANSI /ANS 57. 2 (Ref. 10). These cases are: (a) Tipping or falling of a spent fuel asscmbly, (d) Fuel drop accidents, (f) Horizontal movement of fuel before complete removal f rom ra ck, (h) Objects that may fall onto the stored assemblies, and (i) Missib generated by failure of rotating s machinery or generated by natural phenomena as described in facility SAR. Per paragraph 305.12 of Reference 1, the loads are to com-bined in accordance with Appendix D of Section 3.8.4-II.3 of the Standard Review Plan - (Reference ll) with certain modifications. This appendix stipulates that spent fuel pool racks are to be designed to criteria for Subsection NF for Class 3 component supports. The load combina~tions are listed in Table 5.3.4.2-1 The abbreviations and their significance in the analysis of spent fuel storage racks are as follows:

a. D -- Dead ~ Load including the rack deadweight and the entrained pool water.

b. L -- Live Loads including fuel assemblies and fuel channels.

i j

c. T -- Temperature effects and loads during normal o

operating or shutdown conditions. The Thermal-Hydraulic Analysis, Reference 5, i concludes that the temperature gradient across the rack structure, due to differential l heating between a full and an empty cell, is negligible, as is the temperature gradient l through the thickness of the cell walls (less than 2"F). Additionally, the pool tempera-ture after a full core discharge is shown to be approximately 120*F(with heat exchangers l operational). This temperature (120 F) will, i therefore, be used for evaluating material i properties for the load combinations con-taining this term. { 5-10

- TABLE 5. 3. 4. 2-1 4 i LOAD COMBINATION ACCEPTANCE LIMIT i I -D + L Level A' service limits 2. D+L+To D+L+To+E Level B service limits D+L+Ta+E i D+L+To+Pf l 1 D + L'+ T + E' Level D service limits D+L+F The functional capability d l of the racks should be demonstrated. {- i 1 NOTE: The provisions of NF-3231.1 (Reference 9) of the American Society of Mechanical Engineers, Boiler and Pressure Vessel Code, Section III, Division 1, shall be amended by the requirements'of Paragraphs c.2.3 i and 4 of Regulatory Guide 1.124 (Reference 12) entitled " Design Limits and Load Combinations for l Class 1 Linear-Type Component Supports". i i l 4 5-11 l

d. Ta-- Temparature effecta et the highest temparcture asaocicted with the postulated ebnormal design conditiona.

In the case of a spent fuel storage pool, this would consist of a full core discharge with the heat exchanger inoperable. The Thermal-Hydraulic Analysis concludes that the Lc Salle County Station - Unit 2 Pool will be heated to the boiling point in ev4 hours. Therefore, the maximum temperature of the racks will be assumed to be 240'F, the saturation temperature under 23' of water.

e. E -- Operating basis earthquake (OBE), replaced by UCBV per 305.12a of Reference 1 where UCBV is the loads generated by the service level B artificial acceleration time history.

f. E'-- Safe shutdown earthquake (SSE), replaced by FCBV per 305.12b of Reference 1 where FCBV is the loads generated by service level C artificial acceleration time history.

g. P -- Upward force on rack caused by postulated g

structural fuel assembly. Reference Page 3.8 of the Specification, " External Loadings" gives 1200 lb. uplift force due to fuel handling equipment jamming or maloperation.

h. F -- Loads due to accidental drop of heaviest d

load from maximum possible height. From NF-1201 (Reference 9), the racks and pedestals would be considered as linear-type supports. NF-3143 states that linear-type supports shall be analyzed based en the maximum stress theory when elastic analysis is performed in accordance with the rules of NP-3300. This section states that when linear clastic analysis is performed, the allowable stresses are deter-mined in NF-3320 which provides stress limits for design and Level A service limits in Section NP-3322. For Level B and Level D, stress limits are to be increased by one-third over the factors shown in Table NF-3523.2-1. Bearing type stress limits are cxcluded from rules for Level D service. The Level A service stress limits are summarized in Table 5.3.4.2-2 The factors that need to be applied to the Level A limits to obtain the Level B and Level D service limits are given in Table 5.3.4.2-3. a 5-12 )

's e TABLE 5.3.4.2-2 LEVEL A SERVICE LIMITS PER SUBSECTION NF STRESS CONDITION ALLOWABLE REFERENCE STRESS * -NF-s 'a. Tension on net section 0.6S 3322.1 (a) (1) Y (except for pin connected members) b. Tension on net section 0.45S 3322.1 (a) (2) Y of pin connected members c. . Shear on gross section

0. 4S,

3322.1 (b) (1) } d. Shear on effective throat 24 KSI (120'F) Table NF-3324.5 (a)-1 of fillet welds 21 KSI (240'F) e. Bending of solid rcund 0.75S 3322.1 (d) (2) and square bars and Y solid rectangular sections bent about their weaker axis f. Tension on extreme fibers 0.60S 3322.1 (d) (5) (a) Y of hot rolled or built-up members in bending g. Compression bearing on 0.9S 3322.1 (f) (1) milled surfaces and pins Y h. Compression bearing on 1.5S 3322.1 (f) (3) bolt projected area 1. Threads 0.62Su 3324.6 (a) (2) (a) (1) 5 j. Shear on Weld Base Material 24 KSI (120'F) Table NF-3324.5 (d) -2 21 KSI (240'F) S and S at temperature u I 5-13 1

TABLE

5. 3. 4. 2-3 ELASTIC ANALYSIS STRESS CATEGORIES AND STRESS LIMIT FACTORS FOR CLASS 3

LINEAR TYPE SUPPORTS DESIGNED BY ANALYSIS COMPONEN Stress Umst factors for Loaang Ca'egories [Ncte (1)l Serwce Levet Serv <e Le.et Service Level Stress Category A 8 ($bte (2)) D INote (3)l Primary Stretses (Note ($4 r, = 1.0 r. =l77 r, =2 66 r, = 1.0 r, = /.7 7. Note ta >l 4 2 66 thone m) - K, = 1.0 t. = I.7 7 e K,, = 2 6 4 but stress s but streu s si, of ("t<al 8acf critical bu-6s6ng stress bucaleng stress Prunary Pius $econdary Evaluauon as reew Stresws (Note (6)! for tn.: evaluation.ared for crocai b.,ckl.ng for a!! Ic ad.ng catepr.es. Tne requerement cf t%s $utmarticle snasa be Pean $tresws Evaluation not reau. red. NCVE NCLATU RE: K, = stress 1.mit factoe appi. cable to the Design allowatne tewie and Dending stresws f, = stress IWW1 factor aD()pcafle to tit Destgi allowable sheae stresys K., = stress limet fador ACDinCable to the Ce5 gn allcWafle compreus.e an.4 and beqoing stmWs tC orterm.ne Ducki s u e N0f E $: (1) Contici of debemat on es not ensured by tnew stren 1. nut factm we reo.f*o t> De go $ cec t cat on, oeiermat.on centrol muu cor%d* red wparat*'y. i. (2) F,, F,. and K,

1.0 for desegn of snubte t g

(3) $taess shall not encerG C 73.. (a) ( shall Pot esce=10 a2(. ($) IC' $r*.'ce Lev *ls A. 6, C, anJ 0, Stre.sP' ehto tv ***!rarrt et fm e t W.irePM J'O N0* mit ;ns of a' pre 89a'e ?tr*s#, f s&ng sPal' te cDn o* Iitremal st'essas n.then tw k%N*t as M*'d D) f*4 J#2. Il r*PC 9J! t* *% 2 a#*C e t,) 7' 20, J' I, at I"i* f.83 #* Eor UMe le** A.4DJ 0, D"A*T D8J' SPf C#" !!'**' *s V.all b* lirico LC J f a"3'

* 's' e *** * ** k-I e

5 14

5.3.4.4 M".tcrial Propertico The. storage racks and their supports are fabricated from stainless steel, ASTM-A240, Type 304 plate. Properties are ? and T (240'F). Values are evaluated at both T9.(120 F)2 2 and I-3.2 *f ASME, Section III, Division taken from Tables 1 1, Appendix I. 120*F 240*F* Yield Stress 28.8 KSI 24.0 KSI ~ Ultimate Stress 74.0 KSI 69.2 KSI 6 Modulus of Elasticity 28.0 x 10 PSI 27.4 x 10 pg7 6 Shear Modulus 11.0 x 10 PSI 10.6 x 10 PSI Density 500 PCF 499 PCF Values of 120'F and 240*F are linearly interpolated between the values given for 100*F and 200*F, and 300*F.

Eased on no boiling in FA's under 23' water.

5.3.4.5 Allowable Stresses for Service Load Conditions The allowable stresses for the loading conditions from Table 5.3.4.2-1 are calculated using the limits specified in Table 5.3.4. 2-2 and the material properties from section 5.3.4.4. Table 5.3.4.5-1 gives the allowable stresses for 3 categories. (1) Level A-normal, (2) Level B-OBE, and (3) Level D-SSE. All use the matcrial properties at 240'F. This is slightly conscrva-tive but experience has shown that these combinations are less severe and never control the design. e 5-15 t

TABLE 5.3.4.5-1 ALLOWABLE STRESSES IN KSI (1000 lb./in2) 240'F TYPE OF STRESS LEVEL A LEVEL B LEVEL D a. Tension - membrane 14.4 25.5 38.3 b. Tension - not section at holes 10.8 19.1 28.7 c. Shear - gross section 9.6 17.0 25.5 d. Shear - weld base material 21.0 29.06 29.06 e. Shear - fillet weld throat 21.0 29.06 29.06 f. Shear fusion weld 21.0 29.06 29.06 g. Bending - tension / compression 18.0 31.8 47.9 on solid sections h. Bending - tension / compression 14.4 25.5 38.3 on rolled or built up sections i. Compression - bearing 21.6 38.2 N/A j. Compression - bolt bearing N/A k. Shear (Threads)

8.58 9.87 10.7,3
LEVEL A = 1.0 I

B = 1.15 - For Bolts See Table 3225.2-1, Subsection NF I C = 1.25 e 5-16

5.3.4.6 Equipm:nt Dateription The Spent Fuel Storage equipment consists of 20, poison wall design racks. The complete pool layout is shown in Figure 1.1. The rack array is free standing. There is no gap between the racks as they butt against each other in all four directions. A 5.3 inch gap is present between the West wall of the racks and the West wall of the pool and a 3.0 inch gap is present between the South wall of the racks and the South wall of the ~ pool. All rack components and appurtenances are fabricated from 304 stainless steel. The pool layout of fuel storage and box dimensions are all in agreement with the criticality analysis. Each rack is provided with four screw adjustable pedestals and one non-adjustable pedestal welded to the bottom of the rack. These pedestals transmit the rack weight and vertical seismic loads to the pool floor. 5.3.4.7 Conclusion As a result of the Structural Analysis the following conclusions result: 5.1 The spent fuel storage racks are structurally safe and will maintain a suberitical array during all credible storage conditions. 5.2 The racks are structurally safe for full pool storage of spent fuel subassemblies. 5.3 The rack fusion weld stresses are conservatively analyzed and acceptable, based on adequate calculated factors of sa'fety, the smallest of which is 1.14. 5.4 The suberitical array is maintained for all specified external loading conditions. These include straight and inclined drop of a fuel bundle. l l 5-17 l l l

5.4 Rsferrncta 1. Specification for Fuel Storage Racks, La Salle County Station - Unit 2, Commonwealth Edison Specification No. T-3758, Dated 7-17-85. 2. Seismic Analysis of the La Salle County Station - Unit 2, Phase I, Spent Fuel Storage Racks. (8601-00-0082 April 1986) 3. Criticality Analysis for La Salle County Station - Phase I, Spent Fuel Storage Racks. (8601-00-0007 April 1986) 4. Spent Fuel Storage Racks Fuel Box Crush Tests Maximum Density Rack Design Typical PWR Fuel, Report and Calculation 90 7F 17. 5. Thermal and Hydraulic Analysis Report, Spent Fuel Storage Pool, La Salle County Station - Unit 2. (8601-00-0083 April 1986) 6. Blodgett, " Design of Welded Structures," The James F. Lincoln Arc Welding Foundation, 1966. 7. Rabinowicz, E., " Friction Coefficients of Water-Lubricated Stainless Steels for a Spent Fuel Rack Facility". Study performed for Boston Edison Co., November 1976. 8. Roarke & Young, " Formulas for Stress and Strain," Fifth Edition, McGraw-Hill Book Co., 1975. i l 5-18 _-}}

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