Text

Spent Fuel Rack Criticality Analysis Considering Boraflex Gaps & Shrinkage
	ML20149H159
Person / Time
Site:	Byron, Braidwood
Issue date:	06/30/1994
From:	Srinilta S; COMMONWEALTH EDISON CO.
To:	;
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ML20149H153	List: ML20149H149; ML20149H155; ML20149H159;
References
	NUDOCS 9411180099
	Download: ML20149H159 (43)
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Byron and Braidwood Spent Fuel !

Rack Criticality Analysis Considering Boraflex Gaps and Shrinkage l

l June 1994 W.D. Newmyer H.R. Lam J.W. Miller MJ. Hone Verified: b ' b' N' S. Srinilta Co esign B Approved: ,k[ , /d4&4(

C. R. Savage, Manager Core Design B 9411180099 941107 PDR ADOCK 05000454 p PDR

TABLE OF CONTENTS 1.0 Introduction............................................................................................................I 1.1 Design Description.. . .. . . . . . . .. . .. ... . .. .1 1.2 Design Criteria. .. . ... ....2 2.0 a n a l y I i ca l M et h od s ................................................... .. ...................................... . 3 2.1 Criticality Calculation Methodology . . . . . . . . ....... . . . .. 3 2.2 Reactivity Equivalencing for Bumup and IFBA Credit ... .... . . . . .......4 2.3 Boraflex Shrinkage And Gap Methodology.. . . . . . . . . . . . . . . . . . . . . . . . ..5 3.0 Criticality Analysis of Region 1 Spent Fuel Raeks .............................................. 7 3.1 Reactivity Calculations.. ..... . . . . . .. . . . . . . . . ..... . .7 3.2 IFB A Credit Reactivity Equivalencing. . .. . . . .. .. .. . . . .. .9 3.2.1 IFBA Requirement Determination.. . . . . . . . . . . . . . .. ls 3.2.2 Infinite Multiplication Factor.. . . . . . .. .. ... . . . . . .11 3.3 Sensitivity Analysis and Soluble Boron Worth... . .. . . .. . . . . . . .12 4.0 Criticality Analysis of Region 2 Spent Fuel Raeks .............................................13 4.1 Reactivity Calculations. . . . .. ... .. . . . . . . . . . ... . .... . .13 4.: Burnup Credit Reactivity Equivalencing. . . . .... . . . . . . . . . . . . . . . . .... ..16 4.3 Sensitivity Analysis and Soluble Baron Worth. . .. .. ..... . .. . . . . . . . . . . . . . . . ..17 5.0 Discussion of Post ula t ed A cci dents.................... .................... ............................. ! 8 6.0 S u mma ry of C riticality R esults ............... ........ .................. .................... ...... 19 H i b l i o g ra p h y .. .... .... .... ........ .. .......... .... .. .... .. .. ........ .... .. .. .. .. .. .... .......... ....... .. 3 8 Byron and Braidwood Spent Ftiel Racks i

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LIST OF TABLES Table 1. Fuel Parameters Employed in the Criticality Analysis. ... .. . . . .. . .20 Table 2. Benchmark Critical Experiments... . .. .. . . . . . . . . . . . . . .. . . .. 21 Table 3. Benchmark Critical Experiments PHOENIX Comparison.. .. . . . . 22 Table 4 Data for U Metal and UO2 Critical Experiments (Part 1 of 2) . .. ... 23 Table 4. Data for U Metal and UO2Critical Experiments (Part 2 of 2) .. .... . .. 24 Table 5. Comparison of PHOENIX Isotopics Predictions to Yankee Core 5 Measurements.. . ... . . . . . . . . . .. . . . . . . . . . . 25 Table 6. Byron and Braidwood Region 1 Spent Fuel Rack Ke rr Summary.. . . . 26 Table 7. Byron and Braidwood Region 1 Spent Fuel Rack IFBA Requirement.. . . 27 Table 8. Byron and Braidwood Region 2 Spent Fuel Rack Keff Summary. .28 Table 9. Byron and Braidwood Region 2 Spent Fuel Rack Minimum Burnup Requirement.. .. . . . . .. .. . . . . . . . . . 29 i

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Byron and Braidwood Spent Fuel Racks ii ,

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LIST OF FIGURES Figure 1. Byron and Braidwood Spent Fuel Pool Layout... . . . . . ....30 Figure 2. Byron and Braidwood Region 1 Spent Fuel Rack Storage Cell. .. . . .31 Figure 3. Byron and Braidwood Region 2 Spent Fuel Rack Storage Cell. . . . . . . .32 Figure 4. Byron and Braidwood Region 1 Spent Fuel Rack ,

IFB A Requirement...... . . .. .. . . .. . . . . . . . . .. . 33 Figure 5. Byron and Braidwood Region 1 Spent Fuel Rack Reactivity Sensitivity. .. . . . . . . . . . . . . . .34 Figure 6. Byron and Braidwood Region 1 and 2 Spent Fuel Rack Soluble Boron Worth . . ... ... .. . . . . . . . . . . . . . . . . . . . . . . . . . . . .35 Figure 7. Byron and Braidwood Region 2 Spent Fuel Rack Burnup Credit... .. ... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. . 36 Figure 8. Byron and Braidwood Region 2 Spent Fuel Rack Reactivity Sensitivity. . . . . . . . . . . . . . . . . . . . . . . . . . . . . .. .. .. . 37 t

Byron and Braidwood Spent Fuel Racks iii

1.0 Introduction This report presents the results of a criticality analysis of the Commonwealth Edison Byron and Braidwood spent fuel storage racks with consideration of Botaflex shrinkage and gaps.

The spent fuel storage rack designs considered herein are existing arrays of fuel racks, previously qualified for storage of various 17x17 fuel assembly types with maximum enrichments up to 4.25 w/o U 235 ,

In this analysi.s, each of the two unique storage racks in the Byron and Braidwood spent fuel pools will be manalyzed with consideration of Boraflex shrinkage and gap development. To provide for ,

future fuel management flexibility, storage limits will be developed to allow storage of '

Westinghouse 17x17 OFA fuel with nominal enrichments up to and including 5.0 w/o by employing credit for Integral Fuel Burnable Absorbers (IFBA) and accumulated fuel assembly burnup.

The following storage configurations and enrichment limits am considered in this analysis:

Spent Fuel Racks Storage of 17x17 OFA fuel assemblies with nominal enrichments up to Region 1 4.2 w/o U 235 utilizing all available storage cells. Fresh and burned fuel assemblies with higher initial nominal enrichments up to 5.0 w/o U235 can also be stored in these racks provided a minimum number of ;

IFB As are present in each fuel assembly. IFBAs consist of neutron l absorbing material applied as a thin ZrB2 coating on the outside of the UO 2fuel pellet. As a result, the neutron absorbing materialis a non-removable or integral part of the fuel assembly once it is manufactured.

Spent Fuel Packs Storage of Westinghouse 17x17 OFA assemblies utilizing all available Region 2 storage cells. The Westinghouse 17x17 OFA fuel assemblies must i have an initial enrichment no greater than 1.60 w/o U235 (nominal) or satisfy a minimum burnup requirement.

The Byron and Braidv. nod spent fuel rack analyses are based on maintaining Ke g 5 0.95 for [

storage of 17x17 OFA fuel assemblies under full water density conditions.

1.1 Design Description The Byron and Braidwood spent fuel storage rack layout is depicted in Figure 1 on page 30. The spent fuel rack storage cells for Region 1 and Region 2 are shown in Figure 2 on page 31 and Figure 3 on page 32, respectively, with nominal dimensions provided on each figure.

The fuel parameters relevant to this analysis are given in Table 1 on page 20. With the simplifying assumptions employed in this analysis (no grids, sleeves, axial blankets, etc.), the various types of Westinghouse 17x17 OFA fuel (V5, V+, and P+) are beneficial in terms of extending bumup capability and improving fuel reliability, but do not contribute to any c

i Byron and Braidwood Spent Fuel Racks 1

meaningfulincrease in the basic assembly reactivity. Therefore, future fuel assembly upgrades do not require a criticality analysis if the fuel parameters specified in Table 1 continue to remain bounding.

1.2 Design Criteria Criticality of fuel assemblies in a fuel storage rack is prevented by the design of the rack which limits fuel assembly interaction. This is done by fixing the minimum separation between fuel assemblies and inserting neutron poison between them.

The design basis for preventing criticality outside the reactor is that, including uncertainties, there is a 95 percent probability at a 95 percent confidence level that the effective neutron multiplication factor, K g, e of the fuel assembly array will be less than 0.95 as recommended by ANSI 57.2-1983 and NRC guidanceU )

Byron and Braidwood Spent Fuel Racks 2

2.0 Analytical Methods 2.1 Criticality Calculation Methodology The criticality calculation method and cross section values are verified by comparison with critical experiment data for fuel assemblies similar to those for which the racks are designed. This benchmarking data is sufficiently diverse to establish that the method bias and uncertainty will apply to rack conditions which include strong neutron absorbers, large water gaps and low moderator densities.

The design method which insures the criticality safety of fuel assemblies in the fuel storage rack uses the AMPX(23) system of codes for cross-section generation and KENO Va(4) for reactivity determination.

The 227 energy group cross-section library that is the common starting point for all cross-sections used for the benchmarks of KENO Va and the KENO Va storage rack calculations is generated from ENDF/B-V(2) data. The NITAWL(3) program includes, in this library, the self-shielded resonance cross-sections that are appropriate for each particular geometry. The Nordheim Integral Treatment is used. Energy and spatial weighting of cross-sections is performed by the XSDRNPM(3) program which is a one-dimensional S transpon theory code. These multigroup cross-section sets are then used as input to KENO Va l ) which is a three dimensional Monte Carlo theory program designed for reactivity calculations.

KENO Va Monte Carlo calculations are always performed with sufficient neutron histories to assure convergence. A typical KENO Va Monte Carlo calculation involves more than 60.000 neutron histories which is significantly more than the default of 30,000. To assure adequate convergence, the KENO Va edits which show Average Ke g per Generation Run and Average Ke g by Generation Skipped are examined. These edits provide a visual inspection on the overall convergence of the KENO Va Monte Carlo results.

A set of 44 critical experiments (5.6.7.8.9) has been analyzed using the above method to demonstrate its applicability to criticality analysis and to establish the method bias and uncertainty. The benchmark experiments cover a wide range of geometries, materials, and enrichments, ranging from relatively low enriched (2.35,2.46, and 4.31 w/o), water moderated, oxide fuel arrays separated by various materials (B4 C, aluminum, steel, water, etc.) that simulate LWR fuel shipping and storage conditions to dry, harder spectrum, uranium metal cylinder arrays at high enrichments (93.2 w/o) with various interspersed materials (Plexiglass and air). Comparison with these experiments demonstrates the wide range of applicability of the method. Table 2 on page 21 summarizes these experiments.

1 The highly enriched benchmarks show that the criticality code sequence can correctly predict the

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reactivity of a hard spectrum environment, such as the optimum moderation condition often !

considered in fresh rack and shipping cask analyses. However, the results of the 12 highly l enriched benchmarks are not incorporated into the criticality method bias because the enrichments are well above any encountered in commercial nuclear powcr applications. Basing the method bias solely on the 32 low emiched benchmarks results in a more appropriate and more conservative bias.

Byron and Braidwood Spent Fuel Racks 3 ,

The 32 low enriched, water moderated experiments result in an average KENO Va Ke rrof 0.9930.

Comparison with the average measured experimental Ke rrof 1.0007 results in a method bias of 0.0077. The standard deviation of the bias value is 0.00136 AK. The 95/95 one-sided tolerance limit factor for 32 values is 2.20. Thus, there is a 95 percent probability with a 95 percent confidence level that the uncertainty in reactivity, due to the method, is not greater than 0.0030 AK. This KENO Va bias and uncertainty are consistent voth the previous Westinghouse bias and uncenainty calculated for KENO IV(10)

Material and construction tolerance reactivity effects and reactivity sensitivities are determined using the transpon theory computer code, PHOENIX (ll) PHOENIX is a depletable, two-dimensional, multigroup, discrete ordinates, transpon theory code which utilizes a 42 energy group nuclear data library.

The agreement betyven reactivities computed with PHOENIX and the results of 81 critical benchmark experiments is summarized in Table 3 on page 22. Key parameters describing each of l the 81 experiments are given in Table 4 on page 23. These reactivity comparisons again show l good agreement between experiment and PHOENIX calculations. i 2.2 Reactivity Equivalencing for Burnup and IFBA Credit !

Storage of spent fuel assemblies with initial enrichments higher than that allowed by the methodology described in Section 2.1 is achievable by means of the concept of reactivity is predicated upon the reactivity decrease associated with equivalencing. Reactivity fuel depletion or the eddition of IFBA(equivalencing2) fuel rods. A series of reactivity calculations i performed to generat. a set of enrichment-burnup or enrichment-IFBA ordered pairs which all yield an equivalent Ke rr when the fuelis stored in the Byron and Braidwood spent fuel racks. The data points on the reactivity equivalence curve are generated with the transpon theory computer code, PHOENIX.

A study was done to examine fuel reactivity as a function of time following reactor shutdown.

Fission product decay was accounted for using CINDER (13) CINDER is a point-depletion computer code used to determine fission product activities. The fission products were permitted to decay for 30 years after shutdown. The fuel reactivity was found to reach a maximum at approximately 100 hours0.00116 days 0.0278 hours 1.653439e-4 weeks 3.805e-5 months after shutdown. At this time, the major fission product poison, Xe135, has nearly completely decayed away. Funhermore, the fuel reactivity was found to decrease continuously from 100 hours0.00116 days 0.0278 hours 1.653439e-4 weeks 3.805e-5 months to 30 years following shutdown. Therefore, the most reactive time for a fuel assembly after shutdown of the reactor can be conservatively approximated by re;aoving the Xe 135 The PHOENIX code has been validated by comparisons with experiments where the isotopic fuel composition has been examined following reactor shutdown. In addition, an extensive set of benchmark critical experiments has been analyzed with PHOENIX. Comparisons between measured and predicted uranium and plutonium isotopic fuel compositions are shown in Table 5 on page 25. The measurements were made on fuel discharged from Yankee Core 5(14) The data in Table 5 on page 25 shows that the agreement between PHOENIX predictions and measured I isotopic compositions is good.

Uncertainties associated with the burnap and IFB A dependent reactivities computed with Byron and Braidwood Spent Fuel Racks 4

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PHOENIX are accounted for in the development of the individual reactivity equivalence limits.

For burnup credit, an uncertainty is applied to the PHOENIX calculational results which starts at zero for zero burnup and increases linearly with burnup, passing through 0.01 AK at 30,000 l MWD /MTU. This bias is considered to be very conservative and is based on consideration of the i

good agreement between PHOENIX predictions and measurements and on conservative estimates of fuel assembly reactivity variances with depletion history. For IFBA credit applications, an uncertainty of approximately 10% of the total number of IFBA rods is accounted for in the ,

development of the IFBA requirements. Additionalinformation conceming the specific .

uncertainties included in each of the Byron and Braidwood burnup credit and IFBA credit limits is provided in the individual sections of this report. .

2.3 Boraflex Shrinkage And Gap Methodology ;

As a result of blackness testing measurements performed at other storage rack facilities, the i presence of shrinkage and gaps in some of the Boraflex absorber panels has been noted. The effects of Boraflex shrinkage and gaps will be considered in both the Region 1 and Region 2 spent fuel rack criticality evaluations performed for this report.

Previous generic studies of Boraflex shrinkage and reactivity effects have been performed 05) for ;

storage rack geometries which resemble the Byron and Braidwood spent fuel racks. The results ;

of these studies! (and experience gained in performing similar studies for other rack geometries) !

indicate that: ;

.When absorber panel shrinkage occurs evenly and uniformly (equal pullback is experienced at both ends and the panel remains axially centered and intact), meaningful increases in rack .

reactivity will not occur until more than 7.0 inches of total active fuel length is exposed (3.5 i inches on each end). Assuming a conservative 4% shrinkage scenario, combined top and bot- :

tom fuel exposure will reach 10.08 inches given the initial Byron and Braidwood Region 1 l Boraflex panel length of 139.5 inches and 5.76 inches given the initial Byron and Braidwood Region 2 Boraflex panellength of 144 inches. For this level of uniform top and bottom expo- l sure, generic study data indicates that reactivity will increase by about 0.01 AK for Region 1 l

and will not increase for Region 2.

.When absorber panel shrinkage occurs all at one end, experience has shown that the reactivity impact will remain approximately constant even when an identical length of exposure is 3 added to the opposite end. For the one-end scenario, generic data indicates that reactivity will increase by well over 0.06 AK when 4% uniform, one-end shrmkage is assumed in the Region I racks and by 0.02 AK in the Region 2 racks.

.When absorber panel shrinkage is assumed to result in the formation of a single large gap in every panel, and all panel gaps are conservatively positioned at the vertical centerline of the active fuel, generic study data indicates that reactivity will increase dramatically once a gap size of 1 inch has been exceeded. For an assumed 4% shrinkage at Byron and Braidwood, the data indicates that reactivity will increase by more than 0.06 AK if all shrinkage is modeled as a single, large (5.58 inch in Region 1 and 5.76 inch in Region 2) gap at the centerline.

These generic study results indicate that Boraflex shrinkage and gap formation will result in ]

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1. Note: The genene data in Reference 14 does not include the effect of Boral insens.

i Byron and Braidwood Spent Fuel Racks 5 l

6 extremely large reactivity impacts for the conservative scenarios of single-end exposure and mid- !

plane gap development. Accommodating this level ofimpact in the Byron and Braidwood spent i fuel rack limits would cause an unreasonable and unacceptable loss of enrichment storage 1 capability. Therefore, a conservative, but more realistic treatment of shrinkage and gap formation will be considered in this criticality evaluation.

To conservatively bound the current and future development of shrinkage and gaps, the following assumptions will be employed in the criticality evaluations performed for each of the Byron and Braidwood storage regions which utilize Boraflex absorbers:

1. All absorber panels will be modeled with 47c width shrinkage.

2. All absorber panels will be modeled with 47c length shrinkage (5.58 inches in Region I and 5.76 inches in Region 2) which will be assumed to occur either uniformly (where the panel remains intact over its entire length) or non-uniformly (where a conservative, single 4 inch gap develops somewhere along the panel length).

3. For those panels which are modeled with a gap, the remainder of the 47c length shrinkage not accounted for by the single 4 inch gap will be conservatively applied as bottom or top end shrinkage.

4. Gaps will be distributed randomly with respect to axial position for the absorber panels which are modeled with gaps.

5. Determination of which panels experience shrinkage and which experience gaps will be ,

based on random selection. Several scenarios will be considered to cover the complete spec-trum of shrinkage and gap combinations:

1007c of the panels experience nonuniform shrinkage (random gaps).

507c of the panels experience nonuniform shrinkage (random gaps) and the remaining 507c of panels experience uniform shrinkage (pullback) from the bottom-end.

50% of the panels experience nonuniform shrinkage (random gaps) and the remaining 507c of panels experience uniform shrinkage (pullback) from the top-end.

1007c of the panels experience uniform shrinkage (pullback) from the bottom-end.

100% of the panels experience uniform shrinkage (pullback) from the top-end.

6. A criticality model which simulates 16 storage cells and 64 individual absorber panels for Region 1 and 16 storage cells and 32 individual absorber panels for Region 2 will be employed to provide sufficient problem size and flexibility for considering gaps and shrinkage on a random basis.

7. All absorber material which is lost to shrinkage of gaps will be conservatively removed from

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the model. In reality, the absorber material is not lost - it is simply repositioned by shrinkage to the remaining intact areas of the panel.

The above assumptions are conservative and bounding with respect to the upper bound values for shrinkage and gaps recommended by EPRI.

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3.0 Criticality Analysis of Region 1 Spent Fuel Racks I This section describes the analytical techniques and models employed to perform the criticality analysis and reactivity equivalencing evaluations for the Byron and Braidwood Region I spent fuel storage racks.

Section 3.1 describes the reactivity calculations performed for Region 1 with the nominal enrichment up to 4.20 w/o U 235 Section 3.2 describes the analysis which allows for storage of assemblies with nominal enrichments above 4.20 w/o U 235 and up to 5.00 w/o U 235 by taking credit for Integral Fuel Burnable Absorbers (IFB As). Section 3.3 presents the results of calculations performed to show the reactivity sensitivity of variations in enrichment, center-to-center spacing, and Boraflex loading.

3.1 Reactivity Calculations Ta show that storage of bumed and fresh 17x17 OFA fuel assemblies in the Region I spent fuel racks satisfies the 0.95 Ke g criticality acceptance criteria, KENO is used to establish a nominal reference reactivity and PHOENIX is used to assess the effects of material and construction tolerance variations. The nominal temperature range of 50*F to 140*F is considered in the analysis. A final 95/95 Keg is developed by statistically combining the individual tolerance impacts with the calculational and methodology uncertainties and summing this term with the nominal KENO reference reactivity.

The following assumptions are used to develop the nominal case KENO model for storage of fuel assemblies in the Byron and Braidwood Region I spent fuel rack:

1. The fuel assembly parameters relevant to the criticality analysis are based on the Westing-house 17x17 OFA design (see Table 1 on page 28 for fuel parameters).

2. All fuel assemblies contain uranium dioxide at a nominal enrichment of 4.20 w/o over the entire length of each rod.

3. The fuel pellets are modeled assuming nominal values for theoretical density and dishing frac-tion.

4. No credit is taken for any natural or reduced enrichment axial blankets.

5. No credit is taken for any U 234 or U 236 in the fuel, nor is any credit taken for the buildup of fission product poison material.

6. No credit is taken for any spacer grids or spacer sleeves.

7. No credit is taken for any burnable absorber in the fuel rods.

8. The moderator is pure water (no boron) at a temperature of 68'F. A limiting value of 1.0 gm/cm3is used for the density of water to conservatively bound the range of normal (50*F to 140*F) spent fuel pool water temperatures.

9. The array is infinite in lateral (x and y) extent and finite in axial (vertical) extent.

10. All available storage cells are loaded with fuel assemblies.

I1. Nominal Boraflex poison plate dimensions for width, thickness and length are assumed.

Byron and Braidwood Spent Fuel Racks 7

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12. Boral inserts are modeled at the worst case thickness of 0.079 inches.

To conservatively evaluate the effects of Boraflex shrinkage and gap development, the methodology described in Section 2.3 is employed. Five shrinkage / gap scenarios are examined to cover the spectrum of shrinkage-to-gap ratios from 1007c gaps and 07c shrinkage through 07c gap and 1007c shrinkage. Assignment of which panels have gaps or shrinkage, and the axial location of the gap is based on random selection.

With the above assumptions, the KENO calculation for the nominal case results in a Ke g of 0.9232 with a 95 percent probability /95 percent confidence level uncertainty of 0.0024 AK. This K en is the nominal reactivity assuming no Boraflex gaps or shrinkage.

KENO calculation for the worst case of Boraflex gaps and shrinkage resulted in no increase in the l K eg. This is due to the presence of Boral inserts which are present next to the areas where !

Boraflex gaps and shrinkage are modeled. The Ke g for the nominal case without Boraflex gaps and shrinkage will therefore be used as the reference reactivity for the Region I storage configuration.

Calculational and methodology biases must be considered in the final Ke g summation prior to comparing against the 0.95 Ke g limit. The following biases are included:

Methodology: As discussed in Section 2 of this report, benchmarking of the Westinghouse KENO Va methodology resulted in a method bias of 0.0077 AK.

IO B Self Shielding: To correct for the modeling assumption that individual BIOatoms are homogeneously distributed within the absorber mateial (versus clustered about each B4 C parti-cle), a bias of 0.0011 AK is applied.

Watec Temperature: To account for the effect of the normal range of spent fuel pool water temperatures (50*F to 140*F) on water cross section properties, a reactivity bias of 0.0011 AK '

is applied. The reactivity effect of spent fuel pool water temperatme on water density was con-sidered in assumption 8 above.

To evaluate the reactivity effects of possible variations in material characteristics arid mechanical /

construction dimensions PHOENIX perturbation calculations are performed. For the Byron and Braidwood Region I spent fuel rack configuration UO 2material tolerances are considered along with construction tolerances related to the cell I.D., cell pitch, stainless steel thickness, and Boraflex poison panels. Uncertainties associated with calculation and methodology accuracy are also considered in the statistical summation of uncertainty components.

The following tolerance and uncertainty components are considered in the total uncertainty .

statistical summation:

U 235 Enrichment: The standard DOE enrichment tolerance of 0.05 w/o U 235 about the nom-inal 4.20 w/o U 233 reference enrichment was evaluated with PHOENIX and resulted in a reac-tivity increase of 0.0022 AK.

UO 2Density: A 2.07c variation about the nominal 95% reference theoretical density was evaluated with PHOENIX and resulted in a reactivity increase of 0.0027 AK.

Fuel Pellet Dishing: A variation in fuel pellet dishing fraction from 0.07c to 2.07c (about the nominal 1.2117c reference value) was evaluated with PHOENIX and resulted in a reactivity Byron and Braidwood Spent Fuel Racks 8

increase of 0.0017 AK.

Storage Cell I.D.: The 0.032 inch tolerance about the nominal 8.85 inch reference cell 1.D.was evaluated with PHOENIX and resulted in a reactivity increase of 0.0001 AK.

Storage Cell Pitch : The 0.05 inch tolerance about the nominal 10.32 inch reference cell pitch in the north / south direction was evaluated with PHOENIX and resulted in a reactivity increase of 0.0007 AK. The 0.05 inch tolerance about the nominal 10.42 inch reference cell pitch in the east / west direction was evaluated with PHOENIX and resulted in a reactivity increase of 0.0008 AK.

Stainless Steel Wall Thickness: The 0.005 inch tolerance about the nominal 0.06 inch refer-ence stainless steel wall thickness was evaluated with PHOENIX and resulted in a reactivity increase of 0.0003 AK.

Horaflex Absorber Width: The 0.0625 inch tolerance about the nominal 7.75 inch Boraflex panel width was evaluated with PHOENIX and resulted in a reactivity increase of 0.0002 AK.

Horaflex Absorber Thickness: The 0.007 inch tolerance about the nominal 0.075 inch Boraflex panel thickness was evaluated with PHOENIX and resulted in a reactivity increase of 0.0003 AK.

Horaflex Hm Loading: The 0.0017 gm/cm tolerance 2 about the nominal 0.0238 gm/cm 2 Boraflex Bm loading was evaluated with PHOENDC and resulted in a reactivity increase of 0.0014 AK.

Assembly Position: The KENO reference reactivity calculation assumes fuel assemblies are synunetrically positioned within the storage cells since experience has shown that centered fuel assemblies yield equal or more conservative results in rack K err than non-centered (asymmet-ric) positioning. Therefore, no reactivity uncertainty needs to be applied for this tolerance since the most reactive configuration is considered in the calculation of the reference K err.

Calculation Uncertainty: The KENO calculation for the nominal reference reactivity resulted in a Kerr with a 95 percent probability /95 percent confidence level uncertainty of 0.0024 AK.

Methodology Uncertainty: As discussed in Section 2 of this report, comparison against benchmark experiments showed that the 95 percent probability /95 percent confidence uncer-tainty in reactivity, due to method,is not greater than 0.0030 AK.

The maximum K err for the Byron and Braidwood alternating rows storage configuration is developed by adding the calculational and methodology biases and the statistical sum of independent uncertainties to the KENO reference reactivity. The summation is shown in Table 6 on page 26 and results in a maximum Ke rrof 0.9389.

Since K err is less than 0.95 including uncertainties at a 95/95 probability / confidence level, the acceptance criteria for criticality is met for storage of 17x17 OFA fuel assemblies with nominal enrichment up to 4.2 w/o U 235 in the Byron and Braidwood Region 1 spent fuel racks.

3.2 IFBA Credit Reactivity Equivalencing Storage of fuel assemblies with nominal enrichments greater than 4.20 w/o U 235 in the Region 1 spent fuel storage racks is achievable by means of the concept of reactivity equivalencing. The concept of reactivity equivalencing is predicated upon the reactivity decrease associated with the Byron and Braidwood Spent Fuel Racks 9

addition of Integral Fuel Burnable Absorbers (IFBA)o2h IFBAs consist of neutron absorbing material applied as a thin ZrB2 coating on the outside of the UO2 fuel pellet. As a result. the neutron absorbing material is a non-removable or integral part of the fuel assembly once it is manufactured. l l

Two analytical techniques are used to establish the cr' icality criteria for the storage of IFB A fuel in the fuel storage rack. The first method uses reactiuty equivalencing to establish the poison material loading required to meet the criticality limits. The poison material considered in this analysis is a zirconium diboride (ZrB )2 coating manufactured by Westinghouse. The second method uses the fuel assembly infinite multiplication factor to establish a reference reactivity. The reference reactivity point is compared to the fuel assembly peak reactivity to determine its acceptability for storage in the fuel racks.

3.2.1 IFBA Requirement Determination A series of reactivity calculations are performed to generate a set of IFBA rod number versus enrichment ordered pairs which all yield the equivalent Ke rrwhen the fuelis stored in the Region I spent fuel racks. The following assumptions were used for the IFBA rod assemblies in the PHOENIX models:

1. The fuel assembly parameters relevant to the criticality analysis are based on the Westing-house 17x17 OFA design (see Table 1 o.n page 20 for fuel parameters).

2. The fuel assembly is modeled at its most reactive point in life.

3. The fuel pellets are modeled assuming nominal values for theoretical density and dishing frac-tion.

4. No credit ic taken for any natural enrichment or reduced enrichment axial blankets.

5. No crec'it is taken for any U 23 .1 or U 236 in the fuel, nor is any credit taken for the buildup of fission product poison material.

6. No credit is taken for any spacer grids or spacer sleeves.

7. The IFBA absorber material is a zirconium diboride (ZrB 2) coating on the fuel pellet. Each IFB A rod has a nominal poison material loading of 1.50 milligrams Bi oper inch, which is the minimum standard loading offered bv Westinghouse for 17x17 OFA fuel assemblies.

8. The IFB A B oi loading is reduced by 57c to conservatively account for manufacturing toler-ances and then by an additional 257c to conservatively model a minimum poison length of 108 inches.

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9. The moderator is pure water (no boron) at a temperature of 68'F with a density of 1.0 gm/cm

10. The array is infinite in lateral (x and y) and axial (vertical) extent. This precludes any neutron leakage from the array.

Figure 4 on page 33 shows the constant Kerrcontour generated for the Region 1 spent fuel racks.

Note the endpoint at 0 IFB A rods where the nominal enrichment is 4.20 w/o and at 64(lX) IFB A rods where the nominal enrichment is 5.00 w/o. The integretation of the endpoint data is as follows: the reactivity of the fuel rack array when filled with fuel assemblies enriched to a

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235 nominal 5.00 w/o U with each containing 64(1.0X) IFB A rods is equivalent to the reactivity of Byron and Braidwood Spent Fuel Racks 10

the rack when filled with fuel assemblies enriched to a nominal 4.20 w/o and containing no IFB As. The data in Figure 4 on page 33 is also provided on Table 7 on page 27 for both 1.0X and 2.0X IFBA rods. ,

l it is important to recognize that the curve in Figure 4 on page 33 is based on reactivity ;

equivalence calculations for the specific enrichment and IFBA combinations in actual rack ;

geometry (and not just on simple comparisons of individual fuel assembly infinite multiplication factors). In this way, the environment of the storage rack and its influence on assembly reactivity is implicitly considered.

The IFBA requirements of Figure 4 on page 33 were developed based on the standard IFBA patterns used by Westinghouse. However, since the worth of individual IFB A rods can change depending on position within the assembly (due to local variations in thermal flux), studies were performed to evaluate this effect and a conservative reactivity margin was included in the development of the IFBA requirement to account for this effect. This assures that the IFBA requirement remains valid at intermediate enrichments where standard IFBA patterns may not be available. In addition, to conservatively account for calculational uncertainties, the IFB A requirements of Figure 4 on page 33 also include a conservatism of approximately 107c on the total number of IFB A rods at the 5.00 w/o end (i.e., about 6 extra IFB A rods for a 5.00 w/o fuel ,

assembly).

Additional IFBA credit calculations were performed to examine the reactivity effects of higher IFB A linear B io loadings (1.5X and 2.0X). These calculations confum that assembly reactivity remains constant provided the net Bi o material per assembly is preserved. Therefore, with higher IFB A B 10loadings, the required number of IFB A rods per assembly can be reduced by the ratio of the higher loading to the nominal 1.0X loading. For example, using 2.0X IFB A in 5.00 w/o fuel assemblies allows a reduction in the IFB A rod requirernent from 64 IFB A rods per assembly to 32 IFBA rods per assembly (64 divided by the ratio 2.0X/1.0X).

3.2.2 Infinite Multiplication Factor The infinite multiplication factor, K ,is used as a reference criticality reactivity point, and offers an alternative method for determining the acceptability of fuel assembly storage in the Region I spent fuel racks. The reference K is determined for a nominal fresh 4.20 w/o fuel assembly.

The fuel assembly K calculations are performed using the Westinghouse licensed core design code PHOENIX-PM. The following assumptions were used to develop the infinite multiplication facter model:

1. The Westinghouse 17x17 OFA fuel assembly was analyzed (see Table 1 on page 20 for fuel parameters). The fuel assembly is modeled at its most reactive point in life and no credit is taken for any burnable absorbers in the assembly.

235

2. All fuel rods contain uranium dioxide at a nominal enrichment of 4.20 w/o U over the entire length of each rod.

3. The fuel array model is based on a unit assembly configuration (infinite in the lateral and axial extent)in Byron and Braidwood reactor geometry (no rack).

Byron and Braidwood Spent Fuel Racks 11

4. The moderator is pure water (no baron) at a temperature of 68' F with a density of 1.0 gm/

cm 3 Calculation of the infinite multiplication factor for the Westinghouse 17x17 OFA fuel assembly in the Byron and Braidwood core geometry resulted in a reference K, of 1.470. This includes a 1(7c AK reactivity bias to conservatively account for calculational uncertainties. This bias is consistent with the standard conservatism included in the Byron and Braidwood core design refueling shutdown margin calculations.

For IFBA credit, all 17x17 fuel assemblies placed in the Region I spent fuel racks must comply with the enrichment-IFBA requirements of Figure 4 on page 33 or have a reference K less than or equal to 1.470. By meeting either of these conditions, the maximum rack reactivity will then be less than 0.95, as shown in Section 3.1.

3.3 Sensitivity Analysis and Soluble Boron Worth To show the dependence of Keg on fuel and storage cells parameters as requested by the NRC W ,

the variation of the K en with respect to the following parameters was developed using the PHOENIX computer code:

1. Fuel enrichment, with a 0.50 w/o U 235 delta about the nominal case enrichment.

2. Center-to-center spacing of storage cells, with a 0.50 inch delta about the nominal case center-to-center spacing.

IO

3. Boraflex B loading, 2 with a 0.01 gm/cm delta about the nominal case Boraflex B io loading.

Results of the sensitivity analysis are shown in Figure 5 on page 34.

PHOENIX calculations were also perfmmed to evaluate the reactivity benefits of soluble boron for the Region i spent fuel storage configuration. Results of these calculations are provided in Figure 6 on page 35. As the cune shows, the presence of soluble boron in the Byron and Braidwood spent fuel pool provides substantial reactivity margin.

Byron and Braidwood Spent Fuel Racks 12

I l

4.0 Criticality Analysis of Region 2 Spent Fuel Racks !

This section describes the analytical techniques and models employed to perform the criticality analysis and reactivity equivalencing evaluations for the Byron and Braidwood Region 2 spent ,

fuel storage racks.

Section 4.1 describes the reactivity calculations performed for Region 5 2 with Westinghouse ,

17x17 OFA assemblies at nominal enrichments up to 1.60 w/o U Section 4.2 describes the analysis which allows for storage of Westinghouse 17x17 OFA above 1.60 w/o and up to 5.0 w/o U 235 with minimum burnup requirements. Section 4.3 presents the results of calculations performed to show the reactivity sensitivity of variations in enrichment, center-to-center spacing, and Boraflex loading.

4.1 Reactivity Calculations To show that storage of burned and fresh Westinghouse 17xl7 OFA fuel assemblies in the Region 2 spent fuel racks satisfies the 0.95 Ke g criticality acceptance criteria, KENO is used to establish a nominal reference reactivity and PHOENIX is used to assess the effects of material and construction tolerance variations. The nominal temperature range of 50*F to 140*F is considered in the analysis. A final 95/95 Kerr is developed by statistically combining the individual tolerance ;

impacts with the calculational and methodology uncertainties and summing this term with the nominal KENO reference reactivity.

The following assumptions are used to develop the nominal case KENO model for storage of fuel assemblies in the Byron and Braidwood Region 2 spent fuel rack:

1. The fuel assembly parameters relevant to the criticality analysis are based on the Westing-house 17x17 OFA design (see Table 1 on page 20 for fuel parameters).

2. All fuel assemblies contain uranium dioxide at a nominal enrichment of 1.60 w/o over the ,

entire length of each rod.

3. The fuel pellets are modeled assuming nominal values for theoretical density and dishing frac-tion. !

4. No credit is taken for any natural or reduced enrichment axial blankets. ,

234 236

5. No credit is taken for any U or U in the fuel, nor is any credit taken for the buildup of fission product poison material.

6. No credit is taken for any spacer grids or spacer sleeves.

7. No credit is taken for any burnable absorber in the fuel rods.

8. The moderator is pure water (no boron) at a temperature of 68'F. A limiting value of 1.0 gm/

cm3is used for the density of water to conservatively bound the range of normal (50*F to ;

140*F) spent fuel pool water temperatures.

9. The array is infinite in lateral (x and y) extent and finite in axial (vertical) extent.

10. All available storage cells are loaded with fuel assemblies.

11. Nominal Boraflex poison plate dimensions for width, thickness and length are assumed.

i l

13 ;

Byron and Braidwood Spent Fuel Racks

To conservatively evaluate the effects of Boratlex shrinkage and gap development, the methodology described in Section 2.3 is employed. Five shrinkage / gap scenarios are examined to cover the spectrum of shrinkage-to-gap ratios from 1007c gaps and 07c shrinkage through 07c gap !

and 1007c shrinkage. Assignment of which panels have gaps or shrinkage, and the axiallocation l of the gap is based on random selection.

With the above assumptions, the KENO calculation for the nominal case results in a Ke g of ,

0.9113 with a 95 percent probability /95 percent confidence level uncertainty of 0.0018 AK. l This K en is the nominal reactivity assuming no Boraflex gaps or shrinkage.

Three gap and shrinkage cases were considered for the worst case of Boraflex gaps and shrinkage.

The results are shown below with 95/95 uncertainty:

1007c of the panels experience uniform shrinkage (pullback) from the bottom end assum- )

ing the Boraflex panel starts 6.0625" from the bottom of the rack. l KENO K eg = 0.9371 0.0018 AK :

50% of the panels experience nonuniform shrinkage (random gaps) and the remaining 507c of the panels experience uniform shrinkage (pullback) from the bottom end assuming the Boratlex panel stans 6.0625" from the bottom of the rack.

KENO Ke n = 0.9206 0.0018 AK 1007c of the panels experience uniform shrinkage (pullback) from the bottom end assum-ing the Boraflex panel starts 5.0625" from the bottom of the rack.

KENO K en = 0.9244 0.0018 AK The KENO results show the highest K eg for the case assuming the Boraflex panels start 6.0625" from the bottom of the rack. The KENO result for the case assuming the Boraflex panels start 5.0625" from the bottom of the racks shows a more favorable Ke g. Since uniform shrinkage (pullback) for the bottom end of every Boraflex panel is an overly conservative estimate of the occurance of Boraflex gaps and shrinkage,it is more realistic and still conservative to use the 50%

nonuniform shrinkage and 507c uniform shrinkage K eg as the worst case of realistic Boraflex gaps and shrinkage. This K eg will be used as the reference reactivity for the Region 2 storage contiguration.

Calculational and methodology biases must be considered in the final Ke g summation prior to comparing against the 0.95 Ke n limit. The following biases are included:

Methodology: As discussed in Section 2 of this report, benchmarking of the Westinghouse KENO Va methodology n >uited in a method bias of 0.0077 AK.

HMSelf Shielding: To correct for the modeling assumption that individual BWatoms are ;

homogeneously distributed within the absorber material (versus clustered about each B4 C par- ;

I ticle), a bias of 0.0026 AK is applied.

Water Temperature: To account for the effect of the normal range of spent fuel pool water temperatures (50*F to 140*F) on water cross section properties, a reactivity bias of 0.0020 AK is applied. The reactivity effect of spent fuel pool water temperature on water density was con-sidered in the above assumption.

To evaluate the reactivity effects of possible variations in material characteristics and mechanical /

Byron and Braidwood Spent Fuel Racks 14 l

construction dimensions, PHOENIX perturbation calculations are performed. For the Byron and Braidwood Region 2 spent fuel rack configuration. UO 2material tolerances are considered along with construction tolerances related to the cell 1.D., cell pitch, stainless steel thickness, and Boraflex poison panels. Uncertainties associated with calculation and methodology accuracy are also considered in the statistical summation of uncenainty components.

The following tolerance and uncertainty components are considered in the total uncertainty statistical summation:

U 235 Enrichment: The standard DOE enrichment tolerance of 0.05 w/o U 235 about the nom-inal 1.60 w/o U 235 reference enrichment was evaluated with PHOENIX and resulted in a reac-tivity increase of 0.0104 AK.

UO 2Density: A 2.07c variation about the nominal 95% reference theoretical density was evaluated with PHOENIX and resulted in a reactivity increase of 0.0037 AK.

Fuel Pellet Dishing: A variation in fuel pellet dishing fraction from 0.07c to 2.07c (about the nominal 1.2117e reference value) was evaluated with PHOENIX and resulted in a reactivity increase of 0.0022 AK.

Storage Cell I.D.: The 0.032 inch tolerance about the nominal 8.85 inch reference cell I.D.was evaluated with PHOENIX and resulted in a reactivity increase of 0.0012 AK.

Storage Cell Pitch : The MLO21/-0.059 inch tolerance about the nominal 9.011 inch reference cell pitch was evaluated with PHOENIX and resulted in a reactivity increase of 0.0008 AK.

Stainless Steel Wall Thickness: The 0.005 inch tolerance about the nominal 0.06 inch refer-ence stainless steel wall thickness was evaluated with PHOENIX and resulted in a reactivity increase of 0.0004 AK.

Horaflex Absorber Width: The 0.0625 inch tolerance about the nominal 7.25 inch Boraflex panel width was evaluated with PHOENIX and resulted in a reactivity increase of 0.0010 AK.

Bornflex AbsorberThickness:The 0.007 inch tolerance about the nominal 0.041 inch Boraflex panel thickness was evaluated with PHOENIX and resulted in a reactivity increase of 0.0002 AK.

2 2 Horaflex B 101oading: The 0.0009 gm/cm tolerance about the nominal 0.0130 gm/cra Boraflex B io loading was evaluated with PHOENIX and resulted in a reactivity increase of 0.0028 AK.

Assembly Position: The KENO reference reactivity calculation assumes fuel assemblies are symmetrically positioned within the storage cells since experience has shown that centered fuel assemblies yield equal or more conservative results in rack K err than non-centered (asymmet-ric) positioning. Therefore, no reactivity uncertainty needs to be applied for this tolerance since the most reactive configuration is considered in the calculation of the reference Ke rr.

Calculation Uncertainty: The KENO calculation foi the nominal reference reactivity resulted in a K ert with a 95 percent probability /95 percent confidence level uncertainty ofi0.0018 AK.

Methodology Uncertainty: As discussed in Section 2 of this report. comparison against benchmark experiments showed that the 95 percent probability /95 percent confidence uncer-tainty in reactivity, due to method. is not greater than 0.0030 AK.

Byron and Braidwood Spent Fuel Racks 15

[

The maximum K.g for the Byron and Braidwood Recion 2 spent fuel storage centiguration is t j developed by adding the calculational and methodology biases and the statistical sum of independent uncertainties to the KENO reference reactivity. The summation is shown in Table 8 on page 28 and results in a maximum Ke g of 0.9449.

Since Keg is less than 0.95 including uncertainties at a 95/95 probability /contidence level, the acceptance criteria for criticality is met for storage of Westinghouse 17xl7 OFA fuel assemblies l 235 with nominal enrichments up to 1.60 w/o U in the Byron and Braidwood Region 2 spent fuel l racks.

l 4.2 Burnup Credit Reactivity Equivalencing Storage of burned fuel assemblies in the Byron and Braidwood Region 2 spent fuel racks is achievable by means of the concept of reactivity equivalencing. The concept of reactivity equivalencing is predicated upon the reactivity decrease associated with fuel depletion or the addition of IFBA fuel rods. For burnup credit, a series of reactivity calculations are performed to generate a set of enrichment-fuel assembly discharge burnup ordered pairs which all yield an equivalent K eg when stored in the spent fuel storage racks.

Figure 7 on page 36 dows the constant K eg contour generated for the Byron and Braidwood Region 2 spent fuel racks. This curve represents combinations of fuel enrichment and discharge burnup which yield the same rack multiplication factor (K eg) as the rack loaded with 235 Westinghouse 17x17 OFA fresh fuel (zero burnup) at 1.60 w/o U Note in Figure 7 on page 36, the endpoints are O MWD /MTU where the Westinghouse 17x17 OFA enrichment is 1.60 w/o, and 45089 MWD /MTU where the Westinghouse 17x17 OFA enrichment is 5.00 w/o. The interpretation of the endpoint data is as follows: the reactivity of the spent fuel rack containing Westinghouse 17x17 OFA 5.00 w/o fuel at 45089 MWD /MTU is equivalent to the reactivity of the rack containing Westinghouse 17x17 OFA 1.60 w/o fresh fuel.

The burnup credit curve shown in Figure 7 on page 36 includes a reactivity uncertainty of 0.0150 AK, consistent with the minimum burnup requirement of 45089 MWD /MTU for Westinghouse 17x17 OFA at 5.00 w/o.

It is important to recognize that the curve in Figure 7 on page 36 is based on calculations of constant rack reactivity. In this way, the environment of the storage rack and its influence on assembly reactivity is implicitly considered. For convenience, the data from Figure 7 on page 36 is also provided in Table 9 on page 29. Use of linear interpolation between the tabulated values is acceptable since the curve shown in Figure 7 on page 36 is linear in between the tabulated points.

The effect of axial bumup distribution on assembly reactivity has been considered in the development of the Byron and Braidwood Region 2 burnup credit limit. Previous evaluations have been performed to quantify axial bumup reactivity effects and to confirm that the reactivity equivalencing methodology described in Section 2.2 results in calculations of conservative burnup credit limits". The previous evaluations show that axial burnup effects can cause assembly reactivity to increase at burnup-enrichment combinations which are well beyond those calculated for the Byron and Braidwood Region 2 burnup credit limit. Therefore. additional accounting of axial burnup distribution effects in the Byron and Braidwood Region 2 burnup credit limit is not necessary.

Byron and Braidwood Spent Fuel Racks (Rev. 2) 16

4.3 Sensitivity Analysis and Soluble Boron Worth To show the dependence of Kdf on fuel and storage cells parameters as requested by the NRC W ,

the variation of the Keff with respect to the following parameters was developed using the PHOENIX computer code:

1. Fuel enrichment, with a 0.50 w/o U 235 delta about the nominal case enrichment.

2. Center-to-center spacing of storage cells, with a +0.50/-0.0425 inch delta about the nominal case center-to-center spacing.

i

3. Boraflex B 10 loading, with a 0.01 gm/cm2delta about the nominal case Boraflex B o loading.

Results of the sensitivity analysis are shown in Figure 8 on page 37.

PHOENIX calculations were also performed to evaluate the reactivity benefits of soluble boron for the Region 2 spent fuel storage configuration. Results of these calculations are provided in Figure 6 on page 35. As the curve shows, the presence of soluble boron in the Byron and Braidwood Region 2 spent fuel pool provides substantial reactivity margin.

l l

l 1

l Byron and Braidwood Spent Fuel Racks 17 I

l

5.0 Discussion of Postulated Accidents Most accident conditions will not result in an increase in Keff of the rack. Examples are:

Fuel assembly drop The rack structure pertinent for criticality is not excessively deformed on top of rack and the dropped assembly which comes to rest horizontally on top of the rack has sufficient water separating it from the active fuel height of stored assemblies to preclude neutronic interaction.

Fuel assembly drop Design of the spent fuel racks is such that it precludes the insertion of a between rack fuel assembly between rack modules.

modules Loss of cooling Reactivity decreases since loss of cooling causes an increase in systems temperature, which causes a decrease in water density, which results in decreased reactivity.

However, two accident can be postulated which would increase reactivity beyond the analyzed condition. One such postulated accident would be a fuel assembly misload into a position for which the restrictions on location, enrichment, or bumup are not satisfied. To very conservatively estimate the reactivity impacts of such an occurrence in the spent fuel racks, the impact of loading a fresh assembly at 4.2 w/o U 235 in the middle of a 5x5 array of Region 2 spent fuel rack cells 235 with fresh assemblies at 1.60 w/o U fuel assemblies was determined. The reactivity increase associated with this misloading is less than 0.0438 AK.

A second accident which could result in increased reactivity would be a "cooldown" event during which the pool temperature would drop below 50*F. Calculations show that if the Region I spent fuel pool water temperature was to decrease from 50*F to 32*F, reactivity could increase by about 0.0011 AK, and if the Region 2 spent fuel pool water temperature was to decrease from 50*F to 32*F, reactivity could increase by 0.0020 AK ,

For occurrences of any of the above postulated accidents, the double contingency principle of ANSI /ANS 8,1-1983 can be applied. This states that one is not required to assume two unlikely, independent, concurrent events to ensure protection agamst a criticality accident. Thus, for these postulated accident conditions, the presence of soluble boron in the storage pool water can be assumed as a realistic initial condition since not assuming its presence would be a second unlikely event.

The worth of soluble boron in the Byron and Braidwood spent fuel pool has been calculated with PHOENIX and is shown in Figure 6 on page 35. As the curves show, the presence of soluble baron m the pool water reduces rack reactivity significantly and is more than sufficient to offset the positive reactivity impacts of any of the postulated accidents. To bound the 0.0438 AK reactivity increase from the most limiting accident in the spent fuel racks, it is estimated that 300 ppm of soluble boron is required.

Therefore should a postulated accident occur which causes a reactivity increase in the Byron and Braidwood spent fuel racks, Kerr ill w be maintained less than or equal to 0.95 due to the presence of at least 300 ppm of soluble boron in the spent fuel pool water.

Byron and Braidwood Spent Fuel Racks 18

, 4 6.0 Summary of Criticality Results l For the storage of fuel assemblies in the spent fuel storage racks, the acceptance criteria for . -

criticality requires the effective neutron multiplication factor, K g, e to be less than or equal to 0.95, including uncenainties, under all conditions.

This report shows that the acceptance criteria for criticality is met for the Byron and Braidwood Fresh Fuel Storage Racks for the storage of Westinghouse 17x17 OFA fuel assemblies and for the Byron and Braidwood Spent Fuel Storage Racks for the storage of 17x17 fuel assemblies with the following configurations and enrichment limits:

Spent Fuel Racks Storage of 17x17 OFA fuel assemblies with nominal enrichments up to 4.2 w/o U 235 utilizing all available storage cells. Fresh and burned fuel Region 1 assemblies with higher initial enrichments up to 5.0 w/o U 235 can also be stored in these racks provided a minimum number of IFB As are present in each fuel assembly. IFB As consist of neutron absorbing material applied as a thin ZrB2 coating on the outside of the UO2 fuel .

pellet. As a result, the neutron absorbing material is a non-removable or integral part of the fuel assembly once it is manufactured.

Spent Fuel Racks Storage of Westinghouse 17x17 OFA assemblies utilizing all available Region 2 storage cells. The Westinghouse 17x17 OFA fuel assemblies must have an initial enrichment no greater than 1.60 w/o U235 (nominal) or satisfy a minimum burnup requirement. ,

The analytical methods employed herein conform with ANSI N18.2-1973, " Nuclear Safety Criteria for the Design of Stationary Pressurized Water Reactor Plants," Section 5.7 Fuel Handling System: ANSI 57.2-1983, " Design Objectives for 13VR Spent Fuel Storage Facilities at i Nuclear Power Stations," Se-ction 6.4.2: ANSI N16.9-1975, " Validation of Calculational Methods for Nuclear Criticality Safety"; and the NRC Standard Review Plan, Section 9.1.2, " Spent Fuel ;

S torage".

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Byron and Braidwood Spent Fuel Racks 19 .

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, 1 Table 1. Fuel Parameters Employed in the Criticality Analysis i Parameter W 17x17 OFA )

Number of Fuel Rods per Assembly 264 Rod Zirc-4 Clad O.D. (inch) 0.3600 Clad Thickness (inch) 0.0225 Fuel Pellet O.D.(inch) 0.3088 Fuel Pellet Density (% of Theoretical) 95 Fuel Pellet Dishing Factor (%) 1.211 Rod Pitch (inch) 0.496 Number of Zirc-4 Guide Tubes 24 Guide Tube O.D. (inch) 0.474 Guide Tube Thickness (inch) 0.016 Number of Instrument Tubes 1 Instrument Tube O.D. (inch) 0.474 Instrument Tube Thickness (inch) 0.016 l

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Byron and Braidwood Spent Fuel Racks 20 i

y Table 2. llenchmark Critical Experiments

-4 O "'*' " " " # " ##' D General Description Reflector Separating Material Measured Kg ca Number U2 (w/o) (ppm) (K,n +/- One Sigma)

E i uo2 Kml Lattisc 2.46 w ater water U 16 4 U2 O W 34 / +/- U.10234 m 2 UO2Rod Lattice 2.46 water water 1037 1 Otol 0 99161 +/- 0 (0193 m 3 UO2Rod Lattice 2.46 water water 764 1.0 m 0 0.99459 +/- 0 (U194 h5:

4 5

UO Rod Lattice UO Rod Lattice 2.46 2 46 water water B4C pms B4C pins 0

0 0.9999 100 x) 0 98766 +/- 0 u)217 0 98838 +/- 0111221 6 6 UO Rod Lattice 246 water B4C pins 0 1.0097 I 00132 +/- O m221

& 7 8

UO2Rod Lattice UO2Rod Lattice 2.46 2.46 water water B4C rins B4C pins 0

0 0 9998 0 99570 + /- O H)225 g 1 0083 0.99905 +/- O m210 e 9 UO2Rod Lattice 2 46 water water 0 1(U30 0.99660 +/- O m299 g to UO Rod Lattice 2.46 water water 143 1 (XX)1 0 99707 +/- 0 m199 1I UO Rod Lattice 2.46 w ater stainless steet 514 1Omo 0.99862 +/- 0 00203 r! 12 UO Rod Lattice 2.46 water stainless steel 217 1.011X) 0.99411 +/- 0 00207 E

~~

13 UO Rod Lattice 2.46 water borated aluminum 15 11XXU 0.99229 +/- 0 9)218 14 UO Rod Lattice 2.46 water borated aluminum 92 10001 0.98847 +/- O m208 f

o 15 16 UO2Rod Lattice UO Rod Lattice 2.46 2.46 water borated aluminum ivrated aluminum 395 0.9998 0 9M24 +/- O m205 water 121 1.0001 0.9M68 +/- 0.01209 vT 17 UO Rod Lattice 2.46 water borated aluminum 487 11XUO 0 99521 +/- 01U195 18 UO Rod Lattice 2.46 water borated aluminum 197 I (KU2 0 99203 +/. 0 00211 19 UO Rod Lattice 2.46 water lurated aluminum 634 1.0102 0 98924 +/- O m201 20 UO2Rod Lattice 2.46 water borated aluminum 320 IIX03 0.99461 4/- O m197 21 UO2Rod Lattice 2.46 water h> rated aluminum 72 0.9997 0 98700 +/- O m220 22 UO2Rod Lattice 235 water borated aluminum 0 100tX) 0.99347 +/- 0 (X)128 23 UO Rod Lattice 2 35 water stainless steel 0 11XXX) 0.99566 +/- 0 011116 24 UOf Rod Lattice 235 water water 0 1.00lX) 0 99785 +/- 0 Oil 239 25 UO2Rod Lattice 235 water stainless steel 0 1.G WIO 0.98964 +/- O m240 26 UOzRod Lattice 235 water borated aluminum 0 11XX10 0.98841 +/-Om234 27 U0 2Rod Lattice 235 water BC4 0 1 (XX10 0.99015 +/- 0 0)2 31 28 UO2Rod Lattice 431 water stainless stee! 0 1.0 x 0 0.99063 +/- 0 (U247 29 UO2Rod Lattice 431 water water 0 1.0000 0.98986 +/-0(U228 30 UO2Rod Lattice 431 water stainless steel 0 1.0(X X) 1.00011 +/- 0.00248 31 UO2Rod Lattice 431 water borated aluminum 0 1 (XXX) 1 (XU70 +/- 010254 32 UO2Rod Lattice 431 water lurated aluminum 0 1.0000 11XX)88 +/-01U253 33 U-metal Cyhnders 93.2 bare air 0 l uxo 0 98997 +/- O m257 34 U-metal Cylinders 93.2 bare air 0 11X UO 0.99815 +/- 0 91242 35 U. metal Cylinders 93.2 bare air 0 11XX10 0.99250 +/- 0 00230 36 U-metal Cylinders 93.2 bare air 0 1.0ax) 0.99288 +/- 0 00247 37 U-metal Cylinders 93.2 bare air 0 1.0ax) 0 99869 +/-() 00235 38 U-metal Cylinders 93.2 bare air 0 100tX) 0.99796 4/- 0.m236 39 U-metal Cylinders 93.2 bare plexiglass 0 1.0(n U 0.99799 +/- O m261 40 U-metal Cylinders 93.2 paraffin plexiglass 0 11KUO IJXU61 +/- 0 m265 41 U-metal Cylmders 93.2 bare plexiglass 0 11XXX) 0.99961 +/- 0 m243 42 U-metal Cylinders 93.2 paraffin plexiglass 0 11XXX) 1.0!O54 +/- 0 m272 43 U-metal Cylinders 93.2 paraffin plexiglos 0 100tU 11X1171 +/- O m246 44 U-metal Cylinders 93.2 paraffin plexiglass 0 11XUO 11X)375 +/-Om274 tJ

~

1

. \

Table 3. Ilenchmark Critical Experiments PIIOENIX Comparison Description of Number of PHOENIX K eg Using Experiments Experiments Experiment Buckling UO2 l Al clad 14 0.9947 SS clad 19 0.9944 l l

Borated H2O 7 0.9940 Subtotal 40 0.9944 U-Metal Al clad 41 1.0012 TOTAL 81 0.9978 l

1 l

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Byron and Braidwood Spent Fuel Racks 22

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y-a Table 4. Data for U Metal and UO2 Critical Experiments (Part I of 2)

O "#' ""

Case Cell A/O II 20/U Fuel Density Clad OD Clad Thickness

$ Number Type Un2 Ratio (g/cc) (cm) (cm) a (cm) (cm) tc : tiexa 1.3 z6 luz /33 63263 Aluminum 1.6916 .usitu z.itou ou y 2 Ilexa 1328 3.95 7.53 15265 Aluminum 16916 .07110 2.3590 00 E 3 Ilexa 1328 4 95 7.53 1.5255 Aluminum 1.6916 .07110 2.5120 00 5: 4 Ilexa 1328 3 92 7.52 .9855 Aluminum 1.1506 .07110 1.5580 00 6 5 llexa 1328 4.89 7.52 .9855 Aluminum 1.1506 .07110 1.6520 00

8. 6 Ilexa 1328 2 88 10.53 .9728 Aluminum 1.1506 .07110 1.5580 Of m 7 Ilexa 1328 3.58 10.53 .9728 Aluminum 1.1506 .07110 1.6520 0.0 c 8 Ilexa 1328 4.83 10.53 .9728 Aluminum 1.1506 .07110 1.8060 00

@ 9 Square 2.734 2.18 10.18 .7620 SS-3N .8594 .01085 1.0287 0.0 10 Square 2.734 2 92 10.18 .7620 55 304 B594 .04085 1.lG49 0.0

[ 11 Square 2.734 3.86 10 18 .7620 SS-304 .8594 .at085 1.1938 00 o

~

12 Square 2.734 7.02 10.18 .7620 SS 304 .8594 .at085 14554 0.0 13 Square 2.734 8.49 10.18 .7620 SS-3G4 .8594 .N085 1.5621 0.0

[ 14 Square 2.734 1038 10.18 .7620 SS-3N .8594 .G1085 1.6891 00 p 15 Square 2.734 2.50 10.18 .7620 SS-304 .B594 .04085 1.0617 00 v 16 Square 2.734 4.51 10.18 .7620 SS-3G4 .8594 .at085 1.2522 00 17 Square 3.745 2.50 10.27 .7544 SS-3M .8600 .04060 1.0617 0.0 18 Square 3.745 4.51 10 37 .7544 SS 304 .8600 .01060 1.2522 00 19 Square 3.745 4.51 10 37 .7544 SS-304 .8600 .G4060 12522 0.0 20 Square 3.745 4.51 1037 .7544 55-304 .8600 .N060 1.2522 4560 21 Square 3.745 4.51 10 37 .7548 55-304 .8600 .G4060 1.2522 7090 22 Square 3.745 4.51 10 37 .7544 SS-304 .8600 .G1060 1.2522 1260 0 23 Square 3.745 4.51 10 37 .7544 S5 304 .8600 .04060 1.2522 1334.0 24 Square 3.745 4.51 1037 .7541 SS-304 .8600 .04060 1.2522 1477.0 25 Square 4.069 2.55 9.46 1.1278 SS-301 1.2090 .04060 1.5113 00 26 Square A069 2.55 9.46 1.1278 S5-3G1 1.2090 .01060 1.5113 3392.0 27 Square 4.069 2.14 9.46 1.1278 55-304 1.2090 .01060 1.4500 0.0 28 Square 2.490 2St 10.24 1.0297 Aluminum 1.2060 .08130 1.5113 00 29 Square 3.037 241 9.28 1.1268 SS-3G1 1.1701 .07163 1.5550 0.0 30 Square 3.037 8.16 9.28 1.1268 SS-304 1.2701 .07163 2.1980 00 31 Square 4.069 2.59 9.45 1.1268 SS-304 1.2701 .07163 1.5550 00 32 Square 4.069 3.53 9.45 1.1268 SSJ108 1.2701 .07163 1.6840 00 33 Square 4.069 8.02 9 45 1.1268 SS-304 1.2701 .07163 2.1980 00 34 Square 4.069 9.90 9.45 1.1268 SS-304 1.2701 .07163 23810 00 35 Square 2.490 2 84 10 24 1.0297 Aluminum 1.2060 .08130 1.5113 1677.0 36 Ilexa 2.096 2.06 1038 1.5240 Aluminum 1.6916 .07112 2.1737 0.0 37 Ilexa 2.096 3.09 1038 1.5240 Aluminum 1.6916 .07112 2.4052 00 38 Ilexa 2.096 4.12 1038 1.5240 Aluminum 1.6916 .07112 2.6162 00 39 Ilexa 2.096 6.14 1038 1.5240 Aluminum 1.6916 .07112 2.9891 0.0 40 flexa 2.096 8.20 1038 1.5240 Aluminum 1.6916 .07112 3.3255 00 41 Ilexa 1.307 1.01 18.90 1.5240 Aluminum 1.6916 .07112 2.1742 00 42 Ilexa I J07 1.51 18.90 1.5240 Aluminum 1.6916 .07112 2.4054 00 N ,

w-- -

. ~ . . . -. -

y Table 4. Data for U Metal ami UO2 Critical Experiments (Part 2 of 2) a O #'

Case Cell A,0 il2 0/U Fuel Density Clad OD Lattn e 1%h U"2 * * *'*' *"*' "*" ""

$ Numler Tyle Ratio (g/cc) (cm) um)

o. (cm) (cm)

T 4) liexa 1.30/ i U4 15.YU l.3240 Alummum 1.6916 .U4112 26162 UU U 44 Ilex a 1.307 3 01 18 90 1.5240 Aluminum I6916 .07112 2.9896 00 E 45 llexa 1.307 4 02 18.90 1.5240 Aluminum 1.6916 .07112 3.3249 00 s: 46 Ilexa 1.160 1.01 18.90 1.5240 Aluminum 1.6916 .07112 2 1742 00 0 47 Ilexa 1.160 1." ' 18.90 1.5240 Aluminum 14916 07112 2.4054 0.0

8. 48 Ilexa 1.160 2 01 18 90 1.5240 Alummum I.6916 .07112 2 6162 00 m 49 Ilexa 1.160 3.01 18.90 1.5240 Aluminum 1.6916 .07112 2.9896 00 c 50 llexa 1.160 4.02 18.90 1.5240 Aluminum I.6916 .07112 3.3249 00

@ Si llexa 1.m0 1 01 18 90 1.5240 Aluminum I 6916 .07112 2.1742 00 52 Ilexa 1.mo 1.51 18.90 1.5240 Aluminum 1.6916 07112 2.4054 00

[

n 53 54 Ilexa 11exa 1n10 1.010 2.02 3.01 18 90 18.90 1.5240 1.5240 Aluminum Aluminum 1.6916 .07112 2.6162 00

~ 1.6916 .07112 2.9896 00 55 llexa 1n10 4.02 18 90 L5240 Aluminum 1 6916 .07112 3 3249 00 h 56 Ilexa 1.307 1.00 18.90 .9830 Aluminum 1.1506 .07112 1.4412 00 0 57 Ilexa 1.307 1.52 18 90 .9830 Aluminum 1.1506 .07112 1.5926 00 IA 58 Ilexa 1.307 2 02 18.90 .9830 Aluminum 1.1506 .07112 1.7247 00 59 Ilexa 1307 3 02 18.90 .9830 Atuminum 1.1506 .07112 1.9609 00 60 llexa 1.307 4 02 18.90 .9830 Aluminum 1.1506 .07112 2.1742 00 61 Ilexa 1.160 1.52 18 90 .9830 Aluminum 1.1506 .07112 1.5926 00 62 Ilexa 1.160 2.02 18.90 .9830 Aluminum 1.1506 .07112 I.7247 00 63 Ilexa 1.160 3.02 18.90 .9830 Aluminum 1.1506 .07112 1.9609 00 64 11exa 1.160 4.02 18 90 .9830 Aluminum I.1506 .07112 2.1742 00 65 llexa I.160 1.00 18.90 .9830 Aluminum I.1506 .07112 1.4412 00 66 Ilexa 1.160 1.52 18.90 .9830 Aluminum 1.1506 .07112 1.5926 00 67 Ilexa 1.160 2.02 18 90 .9830 Aluminum 1.1506 .07112 1.7247 00 68 Ilexa 1.160 3.02 18.90 .9830 Aluminum 1.1506 .07112 1.9609 00 69 Ilexa 1.160 4.02 18.90 .9830 Aluminum I.1506 .07112 2.1742 00 70 llexa 1.040 1.33 18.90 19.050 Aluminum 2.0574 .07620 2.8687 0.0 71 Ilexa 1.mo 1.58 18.90 19.050 Aluminum 2.0574 .07620 3.0086 00 72 Ilexa 1.040 1.83 18.90 19.050 Aluminum 2.0574 .07620 3.1425 00 73 Ilexa 1.fM0 2.33 18.90 19.050 Aluminum 2.0574 .07620 3.3942 00 74 Ilexa 1n10 2.83 18.90 19.050 Aluminum 2 0574 .07620 3.6284 00 75 llexa 1n10 3.81 18.90 19 050 Aluminum 2.0574 .07620 4.0566 00 76 Ilexa 1.310 2.02 18.88 1.5240 Aluminum 16916 .07112 2.6160 00 77 Ilexa 1.310 3.01 18.88 1.5240 Aluminum 1.6916 .07112 2.9900 00 78 Ilexa 1.159 2 02 18.88 1.5240 Aluminum 1.6916 .07112 2.6160 00 79 Ilexa 1.159 3.01 18.88 1.5240 Aluminum 1.6916 .07112 2.9900 00 80 llexa 1.312 2.03 18.88 .9830 Aluminum 1.1506 .07112 1.7250 0.0 81 Ilexa 1.312 3.02 18.88 .9830 Aluminum 1.1506 .07112 1.9610 00

Table 5. Comparison of PHOENIX Isotopics Predictions to Yankee Core 5 Measurements Quantity (Atom Ratio) % Difference U 235 /U -0.67 U 236/U -0.28 U 238/U

-0.03 239 Pu /U +3.27 Pu 240/U +3.63 241 Pu /U -7.01 242 Pu /U -0.20 Pu 239/U238 +3.24 Mass (Pu/U) +1.41 FISS-Puff 0T-Pu -0.02 f

Byron and Braidwood Spent Fuel Racks 25

~ - .

1 l

Table 6. Byron and Braidwood Region 1 Spent Fuel Rack Ke n Summary l

AK Ker !

1 Nominal KENO Reference Reactivity: 0.9232 !

Calculational & Methodology Biases:

Methodology (Benchmark) Bias +0.0077 B io Particle Self-Shielding Bias +0.0011 !

Pool Temperature Bias (50*F - 140*F) +0.0011 3 TOTAL Bias 40.0099 ,

Best Estimate Nominal K,g: 0.9331 j Tolerances & Uncertainties:

UO2Enrichment Tolerance +0.0022 UO2Density Tolerance +0.0028 Fuel Pellet Dishing Variation +0.0016 Cell Inner Diameter +0.0001 l

Cell Pitch North / South +0.0007 i Cell Pitch East / West +0.0008 Stainless Steel Thickness +0.0003 Borafiex Thickness +0.0003 Boraflex Width +0.000' B10Loading 40.0014 ,

Calculational Uncertainty (95/95) +0.0024 Methodology Bias Uncertainty (95/95) +0.0030 TOTAL Uncertainty (statistical) 40.0057 ;

Final K eg Including Uncertainties & Tolerances: 0.9389 [

i i

Byron and Braidwood Spent Fuel Racks 26

i Table 7. Byron and Braidwood Region i Spent Fuel Rack IFB A Requirement l l

1 1.0X (1.50 mgAn) WH A 2.0X (3.M mgAn) WB A Enrichment (w/o)

RodsIn Assembly Rods In Assembly 4.20 0 0 4.40 16 8 4.60 32 16 4.80 48 24 5.00 64 32 Byron and Braidwood Spent Fuel Racks 27

~

Table 8. Byron and Braidwood Region 2 Spent Fuel Rack Ke gSummary ,

l AK Kg e

Nominal KENO Reference Reactivity: 0.9206 Calculational & Methodology Biases: )

Methodology (Benchmark) Bias +0.0077 B10 Particle Self-Shielding Bias +0.0026 Pool Temperature Bias (50*F - 140*F) +0.0020 TOTAL Bias +0.0123 Best-Estimate Nominal K,g: 0.9329 Tolerances & Uncertainties:

UO2EnrichmentTolerance +0.0104 UO2DensityTolerance +0.0037 Fuel Pellet Dishing Variation +0.0022 Cell Inner Diameter +0.0012 Cell Pitch +0.0008 Stainless Steel Thickness +0.0004 l

Boraflex Thickness +0.0002 Boraflex Width +0.0010 B io Loading 40.0028 Calculational Uncertainty (95/95) +0.0018 Methodology Bias Uncertainty (95/95) +0.0030 TOTAL Uncertainty (statistical) +0.0120 Final K,g Including Uncertainties & Tolerances: 0.9449

)

l Byron and Braidwood Spent Fuel Racks 28

)

+

Table 9. Ilyron and Ilraidwood Region 2 Spent Fuel Rack Minimum Ilurnup Requirement Nominal Enrichment Iturnup -

(w/o) (51WD/N1TU) 1.60 0 1.80 4500 2.00 8316 2.20 11500 2.40 14300 2.60 16890 2.80 19500 3.00 22080 3.20 24400 3.40 27000 3.60 29300 3.80 31700 4.00 34026 4.20 36300 4.40 38500 4.60 40800 4.80 43000 5.00 45089 l

l l

I 29 1 Byron and Braidwood Spent Fuel Racks (Rev. 2)

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Figure 1. Byron and Braidwood Spent Fuel Pool Layout Byron and Braidwood Spent Fuel Racks 30

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10.42" i 0.05" ,

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Figure 2. Byron and Braidwood Region 1 Spent Fuel Rack Storage Cell Byron and Braidwood Spent Fuel Racks 31 l

l

! t.ATTICE SPac:xc I

! 9.011" +0.021"/-0.059" !

I I I: :i ih= s.sso o.ozz sax r.o. M i i.I ii m:c --

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Figure 3. nyron and Braidwood Region 2 Spent Fuel Rack Storage Cell Byron and Braidwood Spent Fuel Racks 32

64

/

-. /

/

56 /

/

ACCEPTABLE /

48 /

/

, /

/

3 40 /

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8 /

2 /

$. 32 - /

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E / /

x / ,/

$ 24 / '

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1.0X IFBA Loading 8

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2.0X IFBA Loadiny j ,

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0 4.20 4.40 4.60 4.80 5.00 235 U Enrichment (w/o)

Figure 4. Ilyron and Braidwtx>d Region 1 Spent Fuel Rack IFBA Requirement Byron ar.d Braidwood Spent Fuel Racks 33

1 1

0.04 1 I

i l

0.03 ' = . ,

i 1

t 0'02 "- !

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i

-0.04 i

-0.50 0.00 0.50 ;

Delta Enrichment (w/o) ;

e

-0.50 ,

0.00 0.50 l Delta Center-to-Center Spacing (in) l

)

-0.01 0.00 0.01 ;

Delta B 10 Loading (gm-B10/cm2) !

i Figure 5. Hyron and Hraidwood Region 1 Spent Fuel Rack Reactivity Sensitivity i i

i Byron and Braidwood Spent Fuel Racks 34 l

i i

0.00

\,

.\ :

.s (

. N

-0.05

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.2 I; .- x' .

m E ".

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-0.20 .

Region 1 .

--- Region 2 -

-0.25 0.30 I O 500 1000 1500 2000 Soluble Boron Concentration (ppm)

Figure 6. Hyron and Braidwood Region I and 2 Spent Fuel Rack Soluble Boron Worth l

l i

I E

Byron and Braidwood Spent Fuel ~bcks 35

i

, \

I l

1 1

50000

/

J 40000 ACCEPTABLE '

/

S e

/

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2

~

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/,

UNACCEPTABLE ::

a )

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0 4.0 4.5 5.0 1.5 2.0 2.5 3.0 3.5 U 235 Enrichment (w/o)

Figure 7. Ilyron and Braidwood Region 2 Spent Fuel Rack Burnup Credit 36 Byron and Braidwood Spent Fuel Racks (Rev. 2)

l 0.10 l l

/ l 7

/

0.05 ' '

x /

'N /

N /

2 N._ /

0.00 !

9. / 'MC ..... . __

15 / ~' ~

~-.Q ......_

g j 1,y ;

w >

y /

s

= '

/

-0.05 N /

$ / c

+

7

/ Enrichment

-0.10 / Center-to-Center

/ B-10 Loading

/

-0.15

-0.50 0.00 0.50 Delta Enrichment (w/o)

-0.50 0.00 0.50 Delta Center-to-Center Spacing (in)

-0.01 0.00 0.01 10 10 Delta B Loading (gm-B /cm2) i

~

Figure 8. Hyrun and Braidwood Region 2 Spent Fuel Rack Reactivity Sensitivity Byron and Braidwood Spent Fuel Racks 37 i

1 1

Bibliography I

1. Nuclear Regulatory Commission, Letter to All Power Reactor Licensees from B. K. Grimes, OT Position for Review and Acceptance of Spent Fuel Storage and Handling Applications, j April 14,1978.

l

2. W. E. Ford Ill, CSRL-V: Processed ENDFIB-V 227-Neutron-Group and Pointwise Cross-l Section Librariesfor Criticality Safety. Reactor and Shielding Studies, ORNLICSDffM-160, June 1982.

3. N. M. Greene Ah!PX: A Afadular Code Systemfor Generating Coupled hfultigroup Neutron-Gamma Librariesfrom ENDFIB, ORNLffM-3706, March 1976.

4. L. M. Petrie and N. F. Landers, KENO Va--An Improved Afonte Carlo Criticality Program With Supergrouping, NUREG/CR-0200, December 1984.

5. M. N. Baldwin, Critical Experiments Supporting Close Proximity Water Storage of Power Reactor Fuel, BAW-1484-7, July 1979.

6. S. R. Bierman and E. D. Clayton, Criticality Separation Between Subtritical Clusters of 2.35 wt%235U Enriched UO2 Rods in Water with Fixed Neutron Poisons, PNL-2438, October 1977.

7. S. R. Bierman and E. D. Clayton, Criticality Separation Between Subcritical Clusters of 4.29 wt%235U Enriched UO2 Rods in Water with Fixed Neutron Poisons, PNL-2615, August 1979.

8. S. R. Bierman and E. D. Clayton, Criticality Experiments with Subtritical Clusters of 2.35 wt% and 4.31 wt% 235U Enriched UO2 Rods in Water at a Water-to-Fuel Volume Ratio of

/.6, PNL-3314, July 1980.

9. J. T. Thomas, Critical Three-Dimen sional Arrays of U(93.2) hietal Cylinders, Nuclcar Science and Engineering, Volume 52, pages 350-359,1973.

10. D. E. Mueller, W. A. Boyd, and M. W. Fecaau (Westinghouse NFD), Qualification of KENO Calculations with ENDFIB-V Cross Sections, American Nuclear Society Transac-tions, Volume 56, pages 321-323, June 1988.

11. Nguyen, T. Q. et. al., Qualification of the PHOENIX-PIANC Nuclear Design Systemfor Pressurized Water Reactor Cores. WCAP-11596-P-A, June 1988 (Westinghouse Propri-etary).

12. Davidson, S.L., et al, VANTAGE 5 Fuel Assembly Reference Core Report, Addendum 1, WCAP-10444 P-A, March 1986.

13. England, T. R.. CINDER - A One-Point Depletion and Fission Product Program, WAPD-TM-334, August 1962.

Byron and Braidwood Spent Fuel Racks 38

< o ,

14. Melchan, J. B., Yankee Core Evaluation Program FinalReport, WCAP-3017-6094. January f 1971. ,

15. W. A. Boyd and D. E. Mueller (Westinghouse NFD), Effects ofPoison Panel Shrinkage and ;

Gaps on FuelStorage Rack Reactivity, American Nuclear Society Transactions. Volume 56, i pages 323-324, June 1988.

16. W. A. Boyd and M. W. Fecteau (Westinghouse NFD), Effect ofAxial Burnup on Fuel Storage Rack Burnup Credit Reactivity, American Nuclear Society Transactions, Volume 62, Pages 328-329, November 1990.

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Byron and Braidwood Spent Fuel Racks 39 j

ML20149H159

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