ML15075A340

NUH32PHB-0400, Revision 2, Benchmarking of the Ansys Model of the OS200FC Transfer Cask

ML15075A340
Person / Time
Site:	Calvert Cliffs
Issue date:	03/10/2015
From:	AREVA
To:	Office of Nuclear Material Safety and Safeguards
Shared Package
ML15075A350	List: ML15075A339 ML15075A340 ML15075A341 ML15075A342 ML15075A344 ML15075A345 ML15075A346 ML15075A347 ML15075A348 ML15075A349
References
NUH32PHB-0400, Rev. 2
Download: ML15075A340 (44)

Text

ENCLOSURE 10 NUH32PHB-0400, Revision 2, Benchmarking of the ANSYS Model of the OS200FC Transfer Cask Calvert Cliffs Nuclear Power Plant March 10, 2015

CONTROLLED COPY E-281 A Form 3.2-1 Calculation Cover Sheet Calculation No.:

Revision No.:

NUH32PHB-0400 2

Page: 1 of 43 A R EVA Revision 8 DCR NO (if applicable): NUH32PHB-018 PROJECT NAME: NUHOMS 32PHB System PROJECT NO: 10955 CLIENT: CENG - Calvert Cliff Nuclear Power Plant (CCNPP)

CALCULATION TITLE:

Benchmarking of the ANSYS Model of the OS200FC Transfer Cask

SUMMARY

DESCRIPTION:

1) Calculation Summary This calculation benchmarks the thermal analyses of the OS200FC Transfer Cask loaded with 32PTH1 DSC using ANSYS against the calculation performed in [7] using Thermal Desktop and SINDA/FLUINT when forced air circulation is used to improve the thermal performance of the system.

2) Storage Media Description Secure network server initially, then redundant tape backup If original issue, is licensing review per TIP 3.5 required?

Yes El No ED (explain below) Licensing Review No.:

This calculation is prepared to support a Site Specific License Application by CCNPP that will be reviewed and approved by the NRC. Therefore, a 10CFR72.48 licensing review per TIP 3.5 is not applicable.

Software Utilized (subject to test requirements of TIP 3.3): Version:

ANSYS 10.0 Calculation is complete: Digitally signed by VENIGALLA Venkata Date: 2015.03.03 14:21:36

-05'00' Originator Name and Signature: Venkata Venigalla Date:

Calculation has been checked for consistency, completeness and correctness:

11: Digitally signed by LIU SI14:45:20

.Date: 2015.03.03

-05'00' Checker Name and Signature: Hui Liu Date:

I Calculation is approved for use:

PATEL Girish 2.S.4.45=T1I2D8D41399567441 7FCF, InPATEL G rih Project Engineer Name and Signature: Girish Patel Date:201S.03.03 16:1 :06 -05'00'

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 2of43 REVISION

SUMMARY

AFFECTED AFFECTED REV. DESCRIPTION PAGES Computational 1/O 0 Initial Issue All All 1 Update the title for Table 5-6 1,2 4 and None 21 2 The temperature term Tj is corrected to T in Tables 4-5 and 1,2, 6,12 None 4-6 in response to RAI 6-11 from NRC. and 13

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AREVA Page: 3 of 43 TABLE OF CONTENTS Page 1.0 P u rp o s e ............................................................................................................................. 5 2.0 References ........................................................................................................................ 6 3.0 Assumptions and Conservatism .................................................................................... 7 3.1 OS200FC Mass Flow Rate Model ....................................................................... 7 3.2 OS200FC TC Model ............................................................................................ 7 4 .0 D e s ign Inp ut ...................................................................................................................... 9 4.1 Benchmarking Cases ........................................................................................... 9 4.2 Major Dimensions in the OS200FC TC Model ...................................................... 9 4.3 Thermal Properties of Materials in the OS200FC TC Model ................................. 9 4.4 Surface Properties of Materials ........................................................................... 14 4.5 Design Criteria ................................................................................................... 14 5.0 Methodology .................................................................................................................... 15 5.1 Flow Rate Model ............................................................................................... 16 5.2 OS200FC TC Model .......................................................................................... 19 6.0 Results and Discussion ............................................................................................. 32 7.0 Conclusion ...................................................................................................................... 39 8.0 Listing of Computer Files ........................................................................................... 40 APPENDIX A Total Heat Transfer Coefficients ................................................................. 42

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 4 of 43 LIST OF TABLES Page Table 4-1 Design Load Cases for 32PTHl DSC in OS200FC TC ................................... 9 Table 4-2 List of Materials in the OS200FC TC Model .................................................. 10 Table 4-3 Thermal Conductivity of Spacer Disc ............................................................. 11 Table 4-4 Thermal Conductivity of Slide Rail ............................................................... 11 Table 4-5 Air Thermal Properties [3] ............................................................................ 12 Table 4-6 Helium Thermal Conductivity [3] ................................ 13 Table 5-1 TC/DSC Annulus Hydraulic Diameter and Flow Area Calculation .................. 17 Table 5-2 List of the Friction Factors in the Mass Flow Model ...................................... 17 Table 5-3 Mass Flow Rates Along Each Annular Segment .......................................... 19 Table 5-4 D ecay Heat Flux ........................................................................................ .. 20 Table 5-5 S o lar Heat F lux ........................................................................................... .. 20 Table 5-6 Distance between 32PHB Canister and TC Centerline ................................. 21 Table 5-7 Heat Transfer Coefficients in the DSC/TC Annulus for Forced Air Flow ..... 23 Table 5-8 Heat Transfer Coefficients in the DSC/TC Annulus for Forced Air Flow ..... 24 Table 6-1 Steady State Operations with FC, (40.8 kW), [°F] ......................................... 32 Table 6-2 Steady State Operations with FC, (31.2 kW), [OF] ......................................... 33 Table 6-3 Heat Removed by Forced Cooling ................................................................ 34 Table 7-1 Maximum Differences between OS200FC TC Thermal Analysis using ANSYS and SINDA/FLUINT ........................................................................ 39 Table 8-1 List of G eom etry Files .................................................................................... 40 Table 8-2 Summary of ANSYS Runs ............................................................................ 40 Table 8-3 Associated Files and Macros ........................................................................ 41 LIST OF FIGURES Page Figure 5-1 Location of 32PTH1 DSC within OS200 FC TC ............................................. 25 Figure 5-2 Finite Element Mesh of Flow Rate Model with FLUID116 Elements .............. 26 Figure 5-3 Finite Element Model of OS200FC TC with 32PTH1 DSC ............................ 27 Figure 5-4 OS200FC TC Finite Element Model, Components ....................................... 28 Figure 5-5 Gaps in OS200FC Transfer Cask Model ..................................................... 29 Figure 5-6 Typical Decay Heat and Insolance Boundary Conditions .............................. 30 Figure 5-7 Typical Convection and Radiation Boundary Conditions ............................... 31 Figure 6-1 Temperature Distributions 32PTH1 DSC Shell @ 40.8 kW .............................. 35 Figure 6-2 Temperature Distributions OS200 TC @ 40.8 kW ............................................. 36 Figure 6-3 Temperature Distributions 32PTH1 DSC Shell @ 31.2 kW .............................. 37 Figure 6-4 Temperature Distributions OS200 TC @ 31.2 kW ............................................. 38

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 5 of 43 1.0 PURPOSE Thermal Desktop@ and SINDA/FLUINT were used in the previous applications such as Amendment 8 and Amendment 10 to NUHOMS UFSAR [2] to simulate the forced air cooling option for transfer casks. The SINDA/FLUINT models for OS197FC and OS200FC transfer casks (TCs) were studied extensively and approved by NRC as documented in the SER to Amendment 10 [2]. In order to use ANSYS to simulate the forced air cooling option for the CCNPP-FC TC, the ANSYS model is validated in this calculation by benchmarking it against the Thermal Desktop and SINDA/FLUINT model. The OS200FC TC loaded with 32PTH1 DSC and heat loads of 31.2 kW and 40.8 kW are considered for the benchmarking to envelope the conditions expected for the CCNPP-FC TC with a maximum heat load of 29.6 kW.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 6 of 43

2.0 REFERENCES

1 U.S. Code of Federal Regulations, Part 71, Title 10, "Packaging and Transportation of Radioactive Material".

2 Updated Final Safety Analysis Report for the Standardized NUHOMS Horizontal Modular Storage System for Irradiated Nuclear Fuel, NUH-003, Rev. 11.

3 Rohsenow, Hartnett, Cho, "Handbook of Heat Transfer", 3 rd Edition, 1998.

4 Rohsenow, Hartnett, Ganic, "Handbook of Heat Transfer Fundamentals", 2 nd Edition, 1985.

5 ASME Boiler and Pressure Vessel Code,Section II, Part D, "Material Properties", 1998 Edition through 2000 Addenda.

6 ANSYS computer code and On-Line User's Manuals, Version 10.0.

7 Calculation, "Thermal Analysis of OS200 Transfer Cask Loaded with 32PTH1 DSC",

Transnuclear, Inc., Calculation No. NUH32PTH1-0450, Rev. 1.

8 NOT USED.

9 Calculation, "Thermal Analysis of MP197HB Transport Cask for Normal Conditions of Transport", Transnuclear, Inc., Calculation No. MP197HB-0401, Rev. 2.

10 Henninger, J. H., "Solar Absorptance and Thermal Emittance of Some Common Spacecraft Thermal-Control Coatings," NASA Scientific and Technical Information Branch, NASA Reference Publication 1121, 1984.

A Calculation Calculation Revision No.:

No.: NUH32PHB-0400 2

AR EVA Page: 7of43 3.0 ASSUMPTIONS AND CONSERVATISM All the assumption and conservatism considered in this calculation are the same as those described in [7] for the purpose of benchmarking. These assumptions and conservatism are listed in the following sections.

3.1 OS200FC Mass Flow Rate Model The annulus between the DSC shell and the TC inner liner are divided into parallel, individual segments along the DSC axis. No circumferential air flow is considered between the parallel segments. Since the presence of circumferential flow will tend to exchange hotter air in the narrower segments of the annulus with cooler air in the wider segments of the annulus, ignoring the potential for circumferential flow will yield conservative temperature estimates for the peak temperatures on the DSC shell and TC inner liner [7].

Based on [7], an air flow rate of 450 cfm is considered for forced air cooling. To evaluate the air flow rate in each of the parallel segments, a constant pressure boundary condition is applied at the inlet such that the total mass flow rate at the outlet is equal to the total airflow rate of 450 cfm. Since the pressure drop through the annulus between the DSC shell and TC inner shell is the major factor controlling the amount of air flow rate in each segment, the mass flow rate model considers only the annulus over the length of the DSC to determine the mass flow rate through each segment.

3.2 OS200FC TC Model Heat load is simulated by heat generation distributed uniformly over the basket length on the homogenized region. The fuel basket is assumed to be centered within the DSC cavity. As such, a 1.25 inch long helium filled gap is assumed between the top and bottom of the fuel basket region and the DSC's closure lid and bottom end plug. An approximate 13.25 inch long cask spacer is used to position the short DSC within the TC cavity. For the purposes of this calculation the spacer is modeled as 0.75-inch thick inner and outer cylinders enclosed with 0.75-inch thick top and bottom plates. The inner cylinder is assumed to have a 22 inch ID,while the outer cylinder is assumed to have a 66.5 inch ID[7].

For regions between the canister support rails (orientation 150 to 1800) convection from forced air flow is ignored due to the narrowness of the gap between the DSC and TC inner liner.

For the transfer operation in horizontal orientation, the lower halves of the TC cylindrical surfaces are not exposed to insolance. No solar heat flux is considered over these surfaces. To remove any uncertainty about the solar impact on the vertical surfaces, the entire surface areas of vertical surfaces are considered for application of the solar heat flux.

No convection is considered for the exterior vertical surface of the liquid filled neutron shield shell.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 8 of 43 When the TC is in the horizontal orientation, the DSC is supported by the four canister rails depicted in Figure 2-2 of [7]. The thermal resistance between the DSC and the canister rails is assumed to be approximately 2.7 Btu/hr-in 2 -OF [7]. The material properties are listed in Section 4.3.

The grapple ring is not modeled in the current analysis; it is conservatively replaced with air.

Radiation heat exchange is considered between the DSC/Spacer and the cask inner liner, between the lead gap and structural shell and also between the fuel basket and the DSC top and bottom end cover plates by using the ANSYS [6] AUX12 processor.

The following gaps are considered in the OS200FC transfer cask model:

a 0.037" radial gap between the gamma shield and the structural shell [7].

0.037" radial gap assumed between the gamma shield and the structural shell is based on the calculated uniform lead gap at room temperature for OS200FC TC in [7]. This assumption is conservative since the lead gap reaches its maximum dimension of 0.037" at room temperature and decreases to 0" at the lead melt point.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 9of 43 4.0 DESIGN INPUT 4.1 Benchmarking Cases The following cases shown in Table 4-1 are analyzed in this calculation to benchmark the thermal performance of the OS200FC TC with 32PTH1 DSC with forced air flow using ANSYS

[6]. These benchmarking cases envelope the conditions expected for CCNPP-FC TC with 29.6 kW heat load with forced cooling. Design load cases such as the vertical loading where- the forced cooling is not used are not considered in this evaluation.

Table 4-1 Design Load Cases for 32PTHI DSC in OS200FC TC Applicable Conditions Air Flow Ambient Decay Temperature Insolation Heat Case Description (cfm) (°F) Max('7 (kW) 1 Normal Hot 450 106 x 40.8 2 Normal Hot 450 106 x 31.2 Notes:

(1) Insolation in accordance with 10CFR71.71(c) (1) [1].

4.2 Major Dimensions in the OS200FC TC Model Major dimension of 32PTH1 DSC used in the OS200FC TC model are the same as those used in [7].

4.3 Thermal Properties of Materials in the OS200FC TC Model Materials used in the OS200FC TC model are listed in Table 4-2 and are the same as those used in [7].

The material properties for SA240 Type 304, ASTM B29 Lead and NS-3 are listed in Table 3-1 of [7] and the same are used in this analysis. The material properties for the water filled neutron shield are listed in Table 3-4 of [7] for radial directions and Table 3-3 of [7] for axial directions.

The material properties for the homogenized fuel basket are listed in Table 3-2 of [7].

The effective conductivity values used in this calculation are noted in Table 4-3 through Table 4-4. The axial effective conductivity for the the spacer are calculated using the methodology described in Section 5.3 of [9].

The heat transfer coefficients for the forced air flow over the DSC/TC annulus are calculated using the same correlations described in [7],Section 4.2 and are presented in MassFlow ConvCoeff 32PTH1_31 kW.xls for the 31.2 kW heat loads and MassFlowConvCoeff_32PTH1_41 kW.xls for the 40.8 kW heat loads as noted in Table 8-3.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 10 of 43 Table 4-2 List of Materials in the OS200FC TC Model Component Mat # in ANSYS Model . Material DSC Shell 1 SA240-Type 304 DSC Homogenized Bottom Plates 2,3 SA240-Type 304(l)

DSC Outer Top Cover Plate 4 SA240-Type 304 DSC Top Shield Plug 5 SA240-Type 304(1' DSC Helium 33 Helium DSC Homogenized Fuel Basket 41 Eff. Conductivity See Table 3-2 of [7]

TC Inner Shell 7 SA240-Type 304 TC Gamma Shield 8 Lead TC Structural Shell 9 SA240-Type 304 TC Top NS-3 / Top Cover 10 SA240-Type 304 TC Top Cask Lid 11 XM-1 9 TC Inner/Outer Spacer Disc 13 5A240-Type 304 (Cylindrical Shell)

TC Inner/Outer Spacer Disc (Top 14 Effective Conductivity of SA240-Type and Bottom Plates) 304 with 0.0625" Axial Gap TC Outer Shell 17 SA240-Type 304 TC NS-3 Top and Bottom 18 NS-3 Eff. Conductivity, See Table 3-4 of [7]

TC Neutron Shield - 1 (top & bottom) 21 (Shied Se 1 & 17)

(Shield Sections 1 & 17)

Eff. Conductivity, See Table 3-4 of [7]

(Shield Sections 2 to 13)

TC Neutron Shield - 3 23 Eff. Conductivity, See Table 3-4 of [7]

(Shield Sections 14 to 16)

TC Cask Slide Rail 43 Effective Properties (See Table 4-4)

TC FLUID116 Flow Elements 31 --

TC LINK34 Convection Elements 61-75 --

TC LINK34 Convection Elements (At Entrance though wedges)

TC LINK34 Convection Elements 82 (At Exit thought Top Cask Lid)

TC Air 32 Air TC Gamma Shield Air Gap 34 Eff. Conductivity, See Table 4-1 of [7]

TC DSC/Cask Annulus Air Gap 36 Air Note: (1) Assumed as SA240 Type 304 for consistency with [7]

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 11 of 43 Table 4-3 Thermal Conductivity of Spacer Disc Top and bottom plates of Spacer Plate thickness = 0.75 in Gap thickness = 0.0625 in Temp k SS304 [7] Temp k air k air [3] k axial k radial (F) (Btu/hr-in-F) (K) (W/mn-K) (Btu/hr-in-F). (Btu/hr-in-F) (Btu/hr-in-F) 70 0.717 294.4 0.0257 0.0012 0.015 0.717 100 0.725 311.1 0.0269 0.0013 0.015 0.725 200 0.775 366.7 0.0308 0.0015 0.017 0.775 300 0.817 422.2 0.0345 0.0017 0.019 0.817 400 0.867 477.8 0.0381 0.0018 0.021 0.867 500 0.908 533.3 0.0415 0.0020 0.023 0.908 600 0.942 588.9 0.0449 0.0022 0.025 0.942 700 0.983 644.4 0.0482 0.0023 0.027 0.983 800 1.025 700.0 0.0514 0.0025 0.029 1.025 900 1.058 755.6 0.0545 0.0026 0.031 1.058 1,000 1.092 811.1 0.0576 0.0028 0.032 1.092 Table 4-4 Thermal Conductivity of Slide Rail Thickness I 0.12 I in Contact Resistance [7] 2.7 Btu/hr-in 2 -0F Temp k SS304 [5] k eff (OF) (Btu/hr-in-0 F) (Btu/hr-in-0 F) 70 0.717 0.223 100 0.725 0.224 200 0.775 0.228 300 0.817 0.232 400 0.867 0.236 500 0.908 0.239 600 0.942 0.241 700 0.983 0.244 800 1.025 0.246 900 1.058 0.248 1000 1.092 0.250

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 12 of 43 Table 4-5 Air Thermal Properties [3]

Temperature Thermal conductivity Temperature Thermal conductivity (K) (W/m-K) (OF) (Btu/hr-in-°F) 200 0.01822 -100 0.0009 250 0.02228 -10 0.0011 300 0.02607 80 0.0013 400 0.03304 260 0.0016 500 0.03948 440 0:0019 600 0.04557 620 0.0022 800 0.05698 980 0.0027 1000 0.06721 1340 0.0032 The above data are calculated based on the following polynomial function from [3].

k = -C, T' for conductivity in(W/m-K) and T in (K)

For250 < T < 1050 K CO -2.2765010E-03 Cl 1.2598485E-04 C2 -1.4815235E-07 C3 1.7355064E-10 C4 -1.0666570E-13 C5 2.4766304E-17 Specific heat, viscosity, density and Prandtl number of air are used to calculate heat transfer coefficients in APPENDIX A based on the following data from [3].

C,, =XA, T' for specific heat in (kJ/kg-K) and T in (K)

_________________For 250 <T < 1050 K AO 0.1 03409E+1 Al j-0.2848870E-3 A2 j0.781681 8E-6 A3 j-0.4970786E-9 A4 10.1 077024E-1 2

,a =X-B, T' for viscosity (N-s/m 2)x 106 and T in (K)

For 250 < T < 600 K For600 < T < 1050 K BO -9.8601E-1 BO 4.8856745 B1 9.080125E-2 B1 5.43232E-2 B2 -1.17635575E-4 B2 -2.4261775E-5 B3 1.2349703E-7 B3 7.9306E-9 B4 -5.7971299E-11 B4 -1.10398E-12 p = PI RT for density (kg/m 3) with P=1 01.3 kPa; R = 0.287040 kJ/kg-K; T = air temp in (K)

Pr = Cp/plk Prandtl number

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 13 of 43 Table 4-6 Helium Thermal Conductivity [3]

Temperature Thermal conductivity Temperature Thermal conductivity (K) (W/m-K) (OF) (Btu/hr-in-°F) 300 0.1499 80 0.0072 400 0.1795 260 0.0086 500 0.2115 440 0.0102 600 0.2466 620 0.0119 800 0.3073 980 0.0148 1000 0.3622 1340 0.0174 1050 0.3757 1430 0.0181 The above data are calculated based on the following polynomial function from [3]

k = C, T' for conductivity in (W/m-K) and T in (K)

For 300 < T < 500 K for 500< T < 1050 K CO -7.761491E-03 CO -9.0656E-02 Cl 8.66192033E-04 Cl 9.37593087E-04 C2 -1.5559338E-06 C2 -9.13347535E-07 C3 1.40150565E-09 C3 5.55037072E-10 C4 0.OE+00 C4 -1.26457196E-13

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 14 of 43 4.4 Surface Properties of Materials All the surface properties used in this calculation are the same as those described in Section 3 of [7] and are described below for reference.

An emissivity of 0.587 is assumed for the exterior surfaces of the 32PTH1 DSC, the inner shell of the TC, the exterior surface of the liquid neutron shield, and the stainless steel skin enclosing the NS-3 material at the top and bottom of the TC. For conservatism, an emissivity of 0.46 is assumed for the machined stainless steel surfaces of the top and bottom forgings of the TC [7].

An emissivity value of 0.6 is used for the inner surface of the structural shell to account for the expected surface oxidation that will occur during the lead pour process [7].

An emissivity of 0.587 is assumed for the radiation exchange between the fuel basket and the DSC inner end plates.

Solar absorptance values of 0.39 and 0.47 are given in [10] for rolled and machined stainless steel plates, respectively. For conservatism, it is assumed that the absorptivity and the emissivity of stainless steel are equal in this calculation. Solar absorptivity of 0.587 is used for

.the exposed stainless steel surfaces.

4.5 Design Criteria The following criteria are considered for the maximum differences between ANSYS and SINDA/FLUINT models for the benchmarking purposes.:

(1) +/-5 0F for the maximum DSC shell temperature.

(2) +/-5% for the heat removed by forced convection cooling.

(3) +/-1+OF for the air exit temperature.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 15 of 43 5.0 METHODOLOGY The NUHOMS OS200FC TC contains design provisions for the use of forced air circulation to improve its thermal performance. The system will consist of redundant, industrial grade pressure blowers and power systems, ducting, etc. When operating, the fan system is expected to generate a flow rate of 450 cfm or greater, which will be ducted to the location of the ram access cover at the bottom of the cask.

The following are the steps to determine the maximum steady state temperatures of the DSC/ITC components with the forced convection using ANSYS:

1. Assume a ATair for initial runs, Calculate Texit and Tavg based on the initial guess and the air properties based on Tavg.

Where, T1., =T,,mb +AT.,,

T,.g = (TYab + T,,,) / 2 Ta.,b = 106 0 F

2. Run Flow Rate Model described in Section 5.1 iteratively based on average properties of air calculated in previous step to compute the air mass flow rate in each DSC/TC annulus segment. (Run ID: "FlowRate_32PTH1_31 kW for 31.2 kW and "FlowRate_32PTH 1_41kW' for 40.8 kW load cases as listed in Table 8-2)

3. Determine the heat transfer coefficients within the annulus based on the mass flow rates computed in Step 2 (see worksheet "Hcdata" for mass flow rate in Ibm/hr and Hccalc for convection coefficients in "MassFlow ConvCoeff 32PTH1 31kW.xls" for the 31.2 kW load case and "MassFlowConvCoeff_32PTHl_41 kW.xls" for the 40.8 kW load case as noted in Table 8-3)

4. Run Thermal Model (Run ID: "TR_32PTHI1_31" for 31.2 kW and "TR_32PTHI1_41" for 40.8 kW load cases as listed in Table 8-2) described in Section 5.2 based on mass flow rates and heat transfer coefficients calculated in Step 2 and Step 3.

5. Calculate Texit, Tavg, and ATair based on results from Thermal Model in Step 4.

6. Ifdifference between assumed ATair in Step 1 and calculated ATair in Step 5 is less than 1°F, stop iterations, otherwise proceed to Step 7.

7. Rerun the Flow Rate Model described in Section 5.1 and Step 2 with air properties based on Tavg from Step 5.

8. Ifdifferences between air mass flow rates in each DSC/TC annulus segment from Step 7 and Step 2 are less than 0.1 Ibm/hr, stop iterations, otherwise proceed to Step 9.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AREVA Page: 16 of 43

9. Repeat Steps 4 to 9 until the solution converges.

5.1 Flow Rate Model The forced air enters the TC from the ram access opening and the airflow turns and enters the ten (10) flow paths formed by the 1.0" thick wedge segments welded to the TC's bottom. After the forced air exits from the flow paths formed by the wedge segments, the airflow turns and

• flows in the annulus between the DSC and the TC's inner liner. Given the gap between the DSC and TC varies with circumferential position, plus variances in the heating of the air, the airflow will distribute itself around the circumference of the DSC/TC inner liner, until an equal pressure drop is achieved everywhere.

For the purposes of this calculation, each half of the annulus is divided into 19 angular segments as shown in Table 5-1 with 00 at the top of the normally horizontal TC and 1800 at the bottom.

The mass flow rate along each of the 19 angular segments is calculated using the Flow Rate Model. The mass flow rates obtained from this model are used as input to the thermal model of the DSC/TC described in Section 5.2.

The 19 annular segments for forced air flow are modeled using FLUID116 elements with their length equal to the length of the DSC. The potential for circumferential airflow is conservatively ignored as discussed in Section 3.1. The flow area and hydraulic diameter for each annular segment are calculated based on the position of the DSC within the TC cavity. The determination of the gap between the DSC and the TC inner liner as a function of circumferential position was made considering a DSC shell outer diameter of 69.75 inches, a TC inner liner inner diameter of 70.5 inches, and two 0.120-inches thick slide rails that are located 120 from the centerline of the TC. The second set of sliding rails are conservatively ignored. Table 5-1 presents the calculation basis for the gap between the TC and DSC and the associated hydraulic diameter and air flow area.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 17 of 43 Table 5-1 TC/DSC Annulus Hydraulic Diameter and Flow Area Calculation Cask '"X" Cask "Y" DSC "X" DSC "Y" Hydraulic Angle Location Location Location Location Diameter Flow Area Section Anale Segment (degrees) (in) (in) (in) (in) (m) (m) 1 -5 5 0 0.000 35.250 0.000 34.614 0.0323 2.479E-03 2 5 15 10 6.121 34.714 6.011 34.092 0.0321 2.464E-03 3 15 25 20 12.056 33.124 11.844 32.541 0.0315 2.419E-03 4 25 35 30 17.625 30.527 17.324 30.007 0.0305 .2.345E-03 5 35 45 40 22.658 27.003 22.289 26.563 0.0292 2.245E-03 6 45 55 50 27.003 22.658 26.587 22.309 0.0276 2.121E-03 7 55 65 60 30.527 17.625 30.090 17.372 0.0257 1.978E-03 8 65 75 70 33.124 12.056 32.688 11.897 0.0236 1.819E-03 9 75 85 80 34.714 6.121 34.301 6.048 0.0214 1.648E-03 10 85 95 90 35.250 0.000 34.875 0.000 0.0191 1.473E-03 11 95 105 100 34.714 -6.121 34.390 -6.064 0.0167 1.296E-03 12 105 115 110 33.124 -12.056 32.856 -11.958 0.0145 1.125E-03 13 115 125 120 30.527 -17.625 30.316 -17.503 0.0124 9.637E-04 14 125 135 130 27.003 -22.658 26.844 -22.525 0.0105 8.177E-04 15 135 145 140 22.658 -27.003 22.546 -26.869 0.0089 6.914E-04 16 145 155 150 17.625 -30.527 17.551 -30.398 0.0076 5.888E-04 17 155 165 160 12.056 -33.124 12.012 -33.002 0.0066 5.131E-04 18 165 175 170 6.121 -34.714 6.101 -34.598 0.0060 4.667E-04 19 175 185 180 0.000 -35.250 0.000 -35.136 0.0058 4.510E-04 The friction factor along the length of the DSC/cask annulus is calculated as:

f = (1.58

• InRe - 3.28)- 2 [Section 4.2.2.2 of 7]

Table 5-2 lists the friction factors as a function of Reynolds numbers.

Table 5-2 List of the Friction Factors in the Mass Flow Model Re f 4*f Re f 4*f 1 0.093 0.372 1500 0.015 0.058 100 0.063 0.250 1750 0.014 0.055 200 0.039 0.154 2000 0.013 0.052 300 0.030 0.122 3000 0.011 0.046 400 0.026 0.105 4000 0.010 0.041 500 0.023 0.094 5000 0.010 0.039 600 0.021 0.086 6000 0.009 0.037 800 0.019 0.075 8000 0.008 0.034 1000 0.017 0.069 12500 0.007 0.030 1250 0.016 0.063 22500 0.006 0.025 The areas, hydraulic diameters, and friction factors calculated for the 19 annular segments are applied as real constants to the FLUID1 16 elements. The friction factors are applied using the TB,FCON command as function of temperature and Reynolds number.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 18 of 43 The total mass flow rate based on 450 cfm discharge from the fan for the 40.8 kW load case is calculated as follows:

m = 450 cfm

• Peiraveragetemp

=450* (0.3048)3

• 0.986 kg 60 s

=0.2094 kg / s

Where, m = Total Mass Flow Rate, (kg/s)

Pair ,averageep= Density of air based on Average Air Temperature (kg/M 3 )

The air density in the above equation is calculated based on a pressure drop of 6 inches water gauge taken from calculations presented in Section 4.2.2.1 of [7] and an initial air exit temperature of 2730 F for the 40.8 kW heat load case. The final air exit temperature is determined iteratively through the steps shown in Section 5.0. The same evaluation for 31.2 kW heat load case and an initial air exit temperature of 239 0 F results in a total mass flow rate of 0.2151 kg/s.

The forced air introduced in the annular gap between the DSC and the cask distributes itself based upon the flow area and hydraulic diameter. The Flow Rate Model computes the air flow rate in each annular segment based on achieving an equal pressure drop over any segments of the annulus. The Flow Rate Model for determining the mass flow rates is shown in Figure 5-2.

A constant volumetric airflow rate of 450 cfm is assumed to evaluate the air mass flow rate in each of the parallel segments. A constant pressure is applied at the inlet of the air flow into the DSC/TC annulus and the mass flow at the outlet is computed for the flow along the 19 annular segments. The pressure at the inlet is iteratively changed until the total mass flow rate at outlet of the 19 annular segments is equal to total mass flow rate of 0.2094 kg/s and 0.2151 kg/s for the 40.8 and 31.2 kW load cases respectively.

The mass flow rates obtained for each of the 19 angular segments for use in the OS200FC TC thermal model along with the hydraulic diameters and flow areas are presented in Table 5-3 for the 40.8 kW and 31.2 kW load cases.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 19 of 43 Table 5-3 Mass Flow Rates Along Each Annular Segment 40.8 kW 31.2 kW Massflo Massflo Hydraulic Flow w w Diameter Area Section (Ibm/hr) (Ibm/hr) (in) (in2) 1 100.3 102.9 1.27 3.84 2 99.3 101.8 1.26 3.82 3 96.3 98.7 1.24 3.75 4 91.3 93.7 1.20 3.63 5 84.3 86.9 1.15 3.48 6 76.1 78.3 1.09 3.29 7 67.2 69.0 1.01 3.07 8 57.8 59.5 0.93 2.82 9 48.8 50.1 0.84 2.56 10 40.2 41.3 0.75 2.28 11 31.8 32.7 0.66 2.01 12 24.6 25.4 0.57 1.74 13 18.6 19.1 0.49 1.49 14 13.8 14.2 0.41 1.27 15 10.1 10.5 0.35 1.07 16 7.5 7.8 0.30 0.91 17 5.8 6.0 0.26 0.80 18 4.8 5.0 0.24 0.72 19 4.5 4.7 0.23 0.70 5.2 OS200FC TC Model A half-symmetric, three-dimensional finite element model of OS200FC TC loaded with 32PTH1 DSC simulating forced air flow is developed using ANSYS Version 10.0 [6] to provide the maximum component temperatures for the benchmarking purposes.

The model contains the cask shells, cask bottom plate, cask lid, canister shell, spacer and canister end plates with a homogenized basket.

SOLID70 elements are used to model the components including the gaseous gaps. Surface elements SURF152 are used for applying the insolation boundary conditions. Radiation between the homogenized fuel basket and the canister inner cover plates, along the gap between canister and TC inner liner and also along the gap between the gamma shield and structural shell is modeled using the AUX12 processor with SHELL57 elements used to compute the form factors.

Decay heat load is applied as a uniform volumetric heat generated throughout the homogenized region of the basket. The volumetric heat generation rate is calculated as:

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 20 of 43 Q

qz(DI/2) 2 Lb q"= Volumetric Heat Generation Rate (Btu/hr-in 3)

Q = decay heat load (Btu/hr) (to convert from kW multiply by 3412.3)

Di = Canister inner Diameter (in)

Lb = Basket length (in)

The applied decay heat values in the model are listed in Table 5-4 Table 5-4 Decay Heat Flux Heat Load Heat Load Di Lb Decay heat flux (kW) (Btu/hr) (in) (in) (Btu/hr-in 2 )

31.2 106464 0.1770 68.75 162 40.8 139222 0.2315 The insolance is applied as a heat flux over the TC outer surfaces using average insolence values from 10CFR71 [1]. The insolance values are averaged over 12 hours1.388889e-4 days <br />0.00333 hours <br />1.984127e-5 weeks <br />4.566e-6 months <br /> and multiplied by the surface absorptivity factor to calculate the solar heat flux. The solar heat flux values used in OS200FC TC model are summarized in Table 5-5.

Table 5-5 Solar Heat Flux Surface Material Shape Insolance Solar Total solar heat flux over 12 hrs [1] Absorptivity (1) averaged over 12 hrs (gcal/cm 2 ) (Btu/hr-in 2 )

Stainless Steel Curved 400 0.587 (2) 0.501 Flat vertical 200 0.587 (2) 0.250 Notes:

(1) See Section 4.4 for surface properties.

(2) Solar absorptivity of stainless steel is taken equal to its emissivity.

Convection and radiation heat transfer from the cask outer surfaces are combined together as total heat transfer coefficients. The total heat transfer coefficients are calculated using free convection correlations from Rohsenow Handbook [3] and are incorporated in the model using ANSYS macros. These correlations are described in APPENDIX A. The ANSYS macros used in this calculation are listed in Section 8.0.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 21 of 43 During transfer when the cask in a horizontal orientation, the canister shell rests on two slide rails in the TC. These rails are flat stainless steel plates welded to the inner shell of the TC. The thickness of the slide rail is 0.12".

The angle between the lower rail and the vertical plane is 12 degree. Considering this configuration shown in Figure 5-1, the distance between the centerline of DSC and centerline of the cask are calculated as follows.

R 22 = R 12 + X2 - 2 R, x Cos(a)

With R, = Di, TC/2 - trail R2= Do, DSC / 2 a = 120 x = Distance between the canister and TC centerlines Di, TC = Inner diameter of TC Do, DSC = Canister outer diameter trail = cask slide rail thickness = 0.12" The calculated value for x is listed in Table 5-6. In the ANSYS model, the canister is shifted down by the amount of x in the Cartesian y-direction within the TC cavity.

Table 5-6 Distance between 32PHB Canister and TC Centerline DSC Type Di,TC Do,DSC R1 R2 a x (in) (in) (in) (in) (degree) (in) 32PTH1 70.5 69.75 35.13 34.88 12 0.26 Forced air circulation through the annulus of the DSC/TC is modeled using the FLUID1 16 and LINK34 elements. The FLUID1 16 element models the forced air flow along the axial length of the DSC/cask annulus by conducting heat and transmitting the fluid between its nodes, whereas the LINK34 elements model the convection from the DSC/TC surfaces due to the forced air flow.

The FLUID1 16 elements are modeled such that they are connected to the LINK34 convection elements.

The mass flow rates obtained from the Flow Rate Model described in Section 5.1 for each of the annular segments from 00 to 1500 are applied to the FLUID116 elements using the "SFE,,,hflux" command.

Based on the mass flow rates obtained for each of the annular segments from 00 to 1500, the convection heat transfer coefficients for the DSC/TC annulus are computed using the correlations for flow within ducts and pipes. The convection heat transfer coefficients are computed as a function of the local hydraulic diameter, the Reynolds number, and the

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 22 of 43 thermophysical properties of air. These convection heat transfer coefficients are applied to the LINK34 elements using the mpdata,hf/mp,hf commands.

The correlations for the convection coefficients are identical to those in [7] and are taken from equations 7, 43, 44, 45, 57, and 57a from Chapter 7 of [4] as follows:

For 0.5 < Pr < 2000 and 104 < Re < 5x1 06:

Nu hDh Re x Pr xf/2 k 1.07 +12.7(Pr2/ 3 - 1)(f/2)0 5 Re- VxpxDh I-t 2

f = (1.58x In Re-3.28)-

For 0.5 < Pr < 2000 and 3000 < Re < 104:

hcD Nu-= h - (Re-1000)xPrxf/2 2 3 0 5 k 1.0 + 12.7(Pr / - 1)(f/2)

For 0.5 < Pr < 2000 and 0 < Re < 3000:

Nu = hCDh = 2.035 x (x" )- 1/ 3) -0.7, for x*< 0.01 k

Nu = 2.035 x (x* )-(13) -0.2, for 0.01 < x* < 0.06 Nu = 3.657 + 0.0998/x*, for x* > 0.06 Where:

Nu = Nusselt number hc = convection coefficient Dh = hydraulic diameter k = thermal conductivity of fluid at film temperature V = flow velocity p = density of fluid at the film temperature

= dynamic viscosity Pr = Prandtl number f = friction factorRe = Reynolds number x* = entry length factor = x/Re/Dh/Pr x = length of duct/pipe Forced convection is omitted conservatively and conduction is assumed in the region between the canister support rails (i.e., approximately 1500 to 1800) due to the narrowness of the gap

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 23 of 43 between the DSC and the TC inner liner. Based on the above correlations and the mass flow rates from Section 5.1 the heat transfer coefficients for the annular segments from 0° to 150° are calculated and are presented in Table 5-7 and Table 5-8 for the 40.8 and 31.2 load cases, respectively.

The material properties used in the OS200FC TC model are listed in Section 4.0.

The geometry of the OS200FC TC Model is shown in Figure 5-3 through Figure 5-5.

Typical boundary conditions for the Thermal Model of OS200FC TC are shown in Figure 5-6 through Figure 5-7.

The OS200FC transfer cask with 31.2 kW and 40.8 kW heat loads and forced cooling are analyzed under steady-state conditions using SINDA/FLUINT in [7]. Therefore, steady-state analyses are performed using ANSYS to compare the maximum component temperatures for the 32PTHl DSC in an OS200FC transfer cask with 31.2 kW and 40.8 kW heat loads with forced cooling.

Table 5-7 Heat Transfer Coefficients in the DSC/TC Annulus for Forced Air Flow (40.8 kW Load Case)

Heat Transfer Coefficients (Btu/hr-in 2 -OF)

Temp Section Section Section Section Section Section Section Section (OF) Wedge(1) 1 2 3 4 5 6 7 8 110 0.027 0.027 0.027 0.027 0.026 0.025 0.024 0.023 210 0.028 0.028 0.028 0.028 0.027 0.026 0.025 0.024 310 0.004 0.029 0.029 0.029 0.028 0.028 0.027 0.025 0.024 410 0.030 0.030 0.029 0.029 0.028 0.027 0.026 0.024 510 0.030 0.030 0.030 0.029 0.029 0.028 0.026 0.024 Heat Transfer Coefficients (Btu/hr-in 2 -°F*i

... . .. ... .. C. o ... ... .. ( .....t r. ... in2-F Temp Section Section Section Section Section Section Section Exit at Top (OF) 9 10 11 12 13 14 15 Lid (1) 110 0.022 0.020 0.009 0.009 0.011 0.012 0.014 210 0.022 0.020 0.010 0.011 0.012 0.014 0.016 310 0.022 0.010 0.011 0.012 0.013 0.015 0.018 0.016 410 0.022 0.011 0.012 0.013 0.015 0.017 0.020 510 0.011 0.012 0.013 0.014 0.016 0.018 0.022 Notes:

(1) The lowest heat transfer coefficient is used for the Wedge and Exit at Top for conservatism

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 24 of 43 Table 5-8 Heat Transfer Coefficients in the DSCITC Annulus for Forced Air Flow (31.2 kW Load Case)

Heat Transfer Coefficients (Btu/hr-in 2-°F)

Temp Section Section Section Section Section Section Section Section (OF) Wedge(l) 1 2 3 4 5 6 7 8 110 0.028 0.028 0.028 0.027 0.027 0.026 0.025 0.024 210 0.029 0.029 0.028 0.028 0.028 0.027 0.026 0.024 310 0.004 0.030 0.029 0.029 0.029 0.028 0.027 0.026 0.025 410 0.030 0.030 0.030 0.030 0.029 0.028 0.027 0.025 510 0.031 0.031 0.031 0.030 0.029 0.028 0.027 0.025 Heat Transfer Coefficients (Btu/hr-in2-°F H eat. ... ns.... . ............ \ u....... 2. -OF)

Temp Section Section Section Section Section Section Section Exit at Top (OF) 9 10 11 12 13 14 15 Lid (1 110 0.022 0.021 0.009 0.009 0.011 0.012 0.014 210 0.023 0.021 0.010 0.011 0.012 0.014 0.016 310 0.023 0.010 0.011 0.012 0.013 0.015 0.018 0.016 410 0.023 0.011 0.012 0.013 0.015 0.017 0.020 510 0.011 0.012 0.013 0.014 0.016 0.018 0.022 Notes:

(1) The lowest heat transfer coefficient is used for the Wedge and Exit at Top for conservatism

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 25 of 43 TC_

R = Di, TC /2 - trail R 2 = D, DSC / 2 Angle 120 _

IDSC~CL Slide Rail Figure 5-1 Location of 32PTH1 DSC within OS200 FC TC

Calculation No.: NUH32PHB-0400 Revision No.: 2 Page: 26 of 43 The mass flow rate along each of the 19 annular A constant segments is pressure is obtained at the applied at the outlet for use in inlet, such that the OS200FC total mass flow thermal model.

rate is 0.2096 kg/s and 0.2147 kg/s for 40.8 and 31.2 kW load cases.

Figure 5-2 Finite Element Mesh of Flow Rate Model with FLUID116 Elements

Calculation No.: NUH32PHB-0400 Revision No.: 2 Page: 27 of 43 Gamma Shield Neutron Shield Structural Shell Cask Lid

/

Bottom /

NS-3 Spacer Homogenized Fuel Top Forging Bottom Forging Basket Figure 5-3 Finite Element Model of OS200FC TC with 32PTH1 DSC

Calculation No.: NUH32PHB-O400 Revision NO.: 2

_ Page: 28 of 43

/

Outer Shell Exit Nodes Coupled to FLUIDI 16 Elements I

I I/

/

Air Flow Inlet I

Fluid11 6 and Link34 Elements Snenl Shell Nodes, Fixed at 0

106 F Helium Gap DSC Shell Shield Plug

/ X TOP Cover Plate Homogenized Fu/el Basket Helium Gap Homogenized Bottomn Cover Plates Figure 5.4 OS200FC TC Finite Element Model, Components

Calculation No.: NUH32PHB-0400 Revision No.: 2 Page: 29 of 43

/

0.037" radial gap between structural shell and gamma shield Figure 5-5 Gaps in OS200FC Transfer Cask Model

32 HB

,/

4 NU n No:o .NUH32PHBo400 2

lc ula tio NO.:

CaRevision 30 of 43 Page'

.01967

.03934

.05901

.078681

.098351

.118021

.137691

-157361

.177031 ne

\ e eratio Re gi o -

uMetri . . G en sk etRegion Vol t . a d Bake Uniform ug h o u m ogeni z e Thr o Ho Conditions Boundary A Insolance A1SSYS 10.*A HPLU

.250469 If

.3061.9~

.3 33958

.361788 cziL -389618

.417448 Conditions Boundary and Insolance 5.6 Decay Heat Typical Figure

,/

5. Conditin Fi'gure Typical ConvectiOn and FMaciation SOundary

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 32 of 43 6.0 RESULTS AND DISCUSSION Table 6-1 and Table 6-2 present a summary of the maximum/average temperatures for the various components of the OS200FC TC obtained using the methodology described in Section 5.0 using ANSYS and presents a comparison to the temperatures obtained using SINDA/FLUINT in [7] for the design cases noted in Table 4-1.

Table 6-1 Steady State Operations with FC, (40.8 kW), [°F]

Temperature (OF) AT (OF)

Normal Hot Normal Hot TANSYS - TSINDA Component ANSYS SINDA/FLUINT

[7]

Max. DSC Shell 435 431 +4 Inner Liner 347 339 +8 Gamma Shield 339 333 +6 Structural Shell 288 283 +5 Neutron Shield, Max. /Avg. 283/216 278/210 +5/+6 Neutron Shield Outer Skin 271 267 +4 Forced Air, Inlet / Exit 106/273 106/275 0/-2 Bulk Average NS-3 207(1) 206 +1 Closure Lid 263 272 -9 Top Forging 292 299 -7 Bottom Forging 178 169 +9 Notes:

(1) The Bulk Average NS-3 is the maximum of the bulk average temperatures obtained for NS-3 at top and bottom.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 33 of 43 Table 6-2 Steady State Operations with FC, (31.2 kW), [OF]

Temperature (*F) AT ('F)

Normal Hot Normal Hot TANSYS - TSINDA Component ANSYS SINDA/FLUINT

[7]

Max. DSC Shell 373 370 +3 Inner Liner 299 293 +6 Gamma Shield 294 289 +5 Structural Shell 251 247 +4 Neutron Shield, Max. /Avg. 247/195 243/192 +4/+3 Neutron Shield Outer Skin 237 235 +2 Forced Air, Inlet / Exit 106 /239 106/243 0/-4 Bulk Average NS-3 191(1) 188 +3 Closure Lid 234 236 -2 Top Forging 258 263 -5 Bottom Forging 166 156 +10 Notes:

(1) The Bulk Average NS-3 is the maximum of the bulk average temperatures obtained for NS-3 at top and bottom.

As seen from the data presented in Table 6-1 and Table 6-2, the maximum temperature difference between the DSC shell temperatures obtained from the ANSYS and SINDA/FLUINT calculation in [7] is +31F for the 32PTH1 DSC with 31.2 kW and +40F for the 32PTH1 DSC with 40.8 kW. Further, the maximum difference in the average exit air temperature is -40 F for both 40.8 kW load case and -30 F for 31.2 kW load case.

In addition, the maximum difference between the exit air temperatures assumed in Section 5.1 (273 0 F and 239°F for the 40.8 and 31.2 kW load cases respectively) and exit air temperatures obtained from the thermal model as presented in Table 6-1 and Table 6-2 (273.40 F and 239.30 F for the 40.8 and 31.2 kW load cases respectively) is -0.40 F. Therefore, no further iterations are required.

The maximum heat removed by forced convection from the OS200FC TC is summarized in Table 6-3 for methodologies using ANSYS and SINDA/FLUINT. The maximum difference is

-3% less heat removed by forced cooling for the 31.2 kW 32PTH1 DSC using ANSYS methodology.

As seen from the results presented in Table 6-1 and Table 6-2, the maximum temperatures for the closure lid and the top forging are lower by -90F and -70 F, respectively for the 40.8 kW 32PTH1 DSC and -20 F and -50F, respectively for the 31.2 kW 32PTH1 DSC. The SINDA/FLUINT model used in [7] models the convection between the DSC end plates, interior surfaces of the top forging and the interior surfaces of the top closure lid. In the current ANSYS

A Calculation Calculation No.:

Revision No.:

Page:

NUH32PHB-0400 2

34 of 43 AREVA analyses, the convection in the region between the DSC top end plate, interior surfaces of the top forging and the interior surfaces of the top closure lid is not modeled and heat transfer through conduction is assumed. This results in a lower heat being removed from the DSC outer surface due to forced convection and is conservative for the calculation of the fuel cladding and DSC top cover plate temperatures. However this also results in reduced heat transfer to the top forging and closure lid resulting in lower temperatures, but since the temperatures of these components are significantly below the code limits of 800'F [7], there is no effect on the thermal performance of the-OS200 TC.

Table 6-3 Heat Removed by Forced Cooling Mass%

Texit Tinitiai Tavg Tavg Cp Flow m.cp.AT Difference in QFC Heat Methodology (OF) (OF) (OF) (K) (J/kg-K) (kg) (kW)

Load 31.2 ANSYS 239 106 173 351 1011 0.213 15.91 -3 SINDA/FLUINT[7] 243 106 175 352 1011 0.213 16.39 1 40.8 ANSYS 273 106 190 361 1012 0.208 19.53 -1 SINDA/FLUINT[7] 275 106 191 361 1012 0.208 19.76 The temperature profiles for the OS200FC and the 32PTH1 DSC with forced convection using ANSYS and SINDA/FLUINT [7] are shown in Figure 6-1 and Figure 6-2 for the 40.8 kW heat load and Figure 6-3 and Figure 6-4 for 31.2 kW heat load, respectively. The temperature profiles of the 32PTH1 DSC Shell and the OS200FC TC noted above from ANSYS and SINDA/FLUINT are similar. As seen from figures, the maximum temperature on the DSC Shell occurs towards the top of the DSC in both ANSYS and SINDA/FLUINT due to the forced air flow, which enters the OS200FC TC at the bottom and exits at the cask top lid.

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 Page: 36 of 43 AN NOW2 2009 09:56:15 PLOT N. 3 NODAL SOLUTION SUB =4 SIEP=1 SMN =132.25 SM =346.549 132.25 I 156.061 179. 812 203. 683 227.494 251.305 ED275.116 298. 927 322. 738 346.549 ANSYS Node > 344. 4 344. 4 32, 9 301,5 280 258,.5

.... 237. 2

21. 6 194.2 172. 7 151.3, (129.8 TemDer0.ture IF], Tine = 0 hr SINDNFLUINT [7]

Figure 6-2 Temperature Distributions OS200 TC @ 40.8 kW

I Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 39 of 43

7.0 CONCLUSION

Based on the results presented in Section 6.0, thermal analyses of OS200FC TC with forced air flow and 32PTH1 DSC for 31.2 and 40.8 kW using ANSYS satisfies all the criteria described in Section 4.5. Table 7-1 summarizes the maximum differences between the ANSYS and SINDA/FLUINT for the thermal analyses of OS200FC with 32PTH1 DSC for 31.2 and 40.8 kW heat loads.

Table 7-1 Maximum Differences between OS200FC TC Thermal Analysis using ANSYS and SINDA/FLUINT Max AT (F)

Component TANSYS - TSINDA Max. DSC Shell +4 Forced Air, Exit -4 Difference in QFC (1- QFC, SINDA / QFC, Ansys)

Max Difference in Heat Removed by Forced Cooling As seen from Table 7-1, the maximum difference between the DSC shell temperatures is within

+50F, between the air exit Temperature is within +/-10°F, and between the heat removed by forced convection cooling is within +/-5%. Hence all the criteria set in Section 4.5 for the thermal analyses of OS200FC/32PTH1 DSC with forced air flow are satisfied. In conclusion, the ANSYS model is validated to use for the thermal evaluation of CCNPP-FC TC equipped with forced air convection cooling.

A Calculation Calculation No.:

Revision No.:

Page:

NUH32PHB-0400 2

40 of 43 AREVA 8.0 LISTING OF COMPUTER FILES All the runs are performed using ANSYS version 10.0 [6] with operating system "Linux RedHat ES 5.1", and CPU "Opteron 275 DC 2.2 GHz" / "Xeon 5160 DC 3.0 GHz".

A list of the files to create the finite element model of OS200FC with 32PTH1 DSC is shown in Table 8-1.

Table 8-1 List of Geometry Files File Name Description Date/Time (Input and Output)

Macro to create geometry of OS200FC 10/08/2009 0S200_32PTH1 with 32PTH1 DSC 05:01 PM A summary of ANSYS runs is shown in Table 8-2.

Table 8-2 Summary of ANSYS Runs Date / Time Run Name Description for Output File 31.2 kW Load Cases Flow Rate Model to determine the mass 10/12/2009 FlowRate_32PTH1_31kW flow rates for 31.2 kW Heat Load. 01:07 PM TR_32PTH1_31 OS200FC with 32PTH1 DSC and 11/18/2009 Forced Convection- 31.2 kW 11:18 AM Flow Rate Model to determine the mass 10/20/2009 FlowRate_32PTH1_41kW flow rates for 40.8 kW Heat Load. 01:47 PM OS200FC with 32PTH1 DSC and 11/02/2009 TR_32PTHl_41 Forced Convection- 40.8 kW 09:56 AM

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 41 of 43 ANSYS macros, and associated files used in this calculation are shown in Table 8-3.

Table 8-3 Associated Files and Macros File Name Description Date / Time Total heat transfer coefficients for 06/16/08 HCL_0S200.MAC horizontal cylindrical surface 09:26 AM Total heat transfer coefficients for 06/16/08 VPL OS200.MAC vertical flat surface 09:32 AM Material properties for OS200FC TC 09/30/09 Mat_0S200+32PTH1.inp and 32PTH1 DSC 10:58 AM 11/18/2009 MassFlowConvCoeff_32PTH1_31kW.xls Spreadsheet for calculating the hydraulic diameters, friction factors, 11:41 AM mass flow rates and heat transfer 11/18/2009 MassFlowConvCoeff_32PTH 1_41 kW.xls coefficients 11:42 AM

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 42 of 43 APPENDIX A TOTAL HEAT TRANSFER COEFFICIENTS Total heat transfer coefficient, ht, is used to combine the convection and radiation heat transfer together.

ht = hr + hc

Where, hr = radiation heat transfer coefficient (Btu/hr-in 2-°F) hc = free convection heat transfer coefficient (Btu/hr-in 2 _oF)

The radiation heat transfer coefficient, hr, is given by the equation:

h, F 12 (T[ Tiibi Tu,4--Tamh I Btu/hr-in 2 -OF

Where, 6 = surface emissivity F 12 = view factor from surface 1 to ambient = 1 2 4 a = 0.1714 x10-8 Btu/hr-ft -R Tw = surface temperature (OR)

Tamb = ambient temperature (OR)

Surface emissivity values are discussed in Section 4.4.

The following equations from Rohsenow handbook [7], Section 4.4 are used to calculate the free convection coefficients.

Horizontal Cylinders:

Ra=GrPr ; Gr= g 8 (T, -T.)D' 2 V,

-~2 L

{

Nu = 0.60 +

0.387RaD91V66 1+ (0.559/Pr) / ]8/27 {

h, Nu k D

with g = gravitational constant =9.81 (m/s 2 )

0 = expansion coefficient = 1/T (1/K)

T = absolute temperature (K) v = kinematic viscosity (m2/s)

Calculation No.: NUH32PHB-0400 Calculation Revision No.: 2 AR EVA Page: 43 of 43 D = diameter of the horizontal cylinder (in) k = air conductivity (W/m-K)

The above correlations are incorporated in ANSYS model via macro" HCLOS200.MAC" listed in Section 8.0.

Vertical Flat Surfaces:

Gr= g 8 (Tv 2 T.)

Ra=GrPr NuL = 2.8 with In(1+ 2.8/NUT )

NUT = /4 Ra114 N C, w 4[ 0.503 with Q =[(1 + (0.492/Pr) 91 16 )49 1

I Nut = CVRa1/ 3 with Cv = 0.13Pr 0 22 (1+ 0.61 Pr'0 81)°42 Nu =[(Nu,)m + (Nu,)m I /M with m=6 for 1<Ra < 1012

= Nu k L

with g = gravitational constant =9.81 (m/s 2 )

13 = expansion coefficient = 1/T (1/K)

T = absolute temperature (K) v = kinematic viscosity (m2/s)

L = height of the vertical surface (in) k = air conductivity (W/m-K)

The above correlations are incorporated in ANSYS model via macro" VPLOS200.MAC" listed in Section 8.0.