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NUREG/CR-7250, Thermal-Hydraulic Experiments Using a Dry Cask Simulator
	ML18310A140
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Issue date:	10/31/2018
From:	Durbin S G, Lindgren E R, Marshall S; Office of Nuclear Regulatory Research, Sandia
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	Meyd, Donald
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NUREG/CR-7250 Thermal-Hydraulic Experiments Using Dry Cask Simulator Office of Nuclear Regulatory Research

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NUREG/CR-7250 Thermal-Hydraulic Experiments Using Dry Cask Simulator Manuscript Complet ed: October 2018 Date Published: October 2018 Prepared by:

S. G. Durbin

E. R. Lindgren Sandia National Laboratories Albuquerque, NM 87185 Shawn Marshall, NRC Project Manager Office of Nuclear Regulatory Research

iii ABSTRACT A series of well-controlled tests were conducted using a single, prototypic-geometry boiling water reactor (BWR) fuel assembly inside of a pressure vessel and enclosure to mimic the thermal-hydraulic responses of both aboveground and belowground dry storage casks. This simplified test assembly was shown to have similarity with prototypic systems through dimensional analysis.

The data were collected over a broad parameter set including simulated decay power and internal helium pressure. These data were collected and documented with the intent to be used for validation exercises with thermal-hydraulic codes and computational fluid dynamics simulations. The primary values of interest, air mass flow rate and peak cladding temperature, and their uncertainties are highlighted in this report.

v TABLE OF CONTENTS ABSTRACT ......................................................................................................................

.................... iiiTABLE OF CONTENTS ........................................................................................................................ vLIST OF FIGU RES ...............................................................................................................

............... viiLIST OF TA BLES ................................................................................................................

................ xiEXECUTIVE SU MMARY.............................................................................................................

...... xiiiABBREVIATIONS AN D ACRONYMS...............................................................................................

xv1 INTRODUC TION ................................................................................................................

............. 1-11.1 Objectiv e .................................................................................................................

................. 1-21.2 Prev ious Stu dies ..........................................................................................................

............ 1-21.2.1 Smal l Scale, S ingle As sembly ......................................................................................

1-21.2.2 Full-Scale , Multi-Assemb ly ............................................................................................

1-21.2.3 Unique ness of Dry Cas k Simulator

...............................................................................

1-42 APPARATUS AND PROCEDURES ..............................................................................................

2-12.1 Gene ra l Construction

......................................................................................................

......... 2-12.2 De sign of the Heated Fuel Bundle

..........................................................................................

2-42.3 In strumentation

...........................................................................................................

............. 2-62.3.1 Thermocouples (TCs) ...................................................................................................

2-62.3.2 Pres sure Vessel

.......................................................................................................... 2-152.3.4 Ho t Wire Anemometer s ............................................................................................... 2-182.4 Air Ma ss Flow Rate

................................................................................................................ 2-182.4.1 Flow Straightening ......................................................................................................

2-192.4.2 Aboveground Air F low Meas urement .........................................................................

2-192.4.3 Belowground Air Flo w Measurem ent .......................................................................... 2-222.5 Cro ss-Wind Testing

........................................................................................................

....... 2-243 ABOVEGROU ND RESULTS

.........................................................................................................

3-13.1 Ste ady State Analyses

.....................................................................................................

....... 3-13.1.1 Peak C lad ding Temperatur e and Air Mass Flow Rate .................................................

3-13.1.2 Two-Dimension al Temperature Contours

....................................................................

3-33.1.3 Transverse Temperature Profiles including t he TC Lance ...........................................

3-53.1.4 Summary Dat a Table s ..................................................................................................

3-63.2 Tr ansient Analyses

........................................................................................................

.......... 3-83.2.1 Transient Re sponse of TC La nce and Corresponding Cladding

...............................

3-104 BELOWGROU ND RESULTS

.........................................................................................................

4-14.1 Ste ady State Analyses

.....................................................................................................

....... 4-14.1.1 Peak C lad ding Temperatur e and Air Mass Flow Rate .................................................

4-14.1.2 Two-Dimensional Velocity Conto urs .............................................................................

4-34.1.3 Transvers e Temperature Profiles I ncluding t he TC Lance ..........................................

4-44.1.4 Summary Dat a Table s ..................................................................................................

4-54.2 Tr ansient Analyses

........................................................................................................

.......... 4-84.2.1 Transient Re sponse of TC La nce and Corresponding Cladding

.................................

4-94.3 Cro ss-Wind Analyses

.......................................................................................................

..... 4-112.3.

......................................................................................................... 2-1 vi 5

SUMMARY

.....................................................................................................................

................. 5-16 REFERENC ES ..................................................................................................................

.............. 6-1APPENDIX A ERRO R ANALYSIS

...................................................................................................

A-1APPENDIX B CHANNEL LIST FROM ABOVEGROUND TESTING

.............................................

B-1APPENDIX C DIMENSIO NAL ANALYSES

.....................................................................................

C-1APPENDIX D VERIFICATION OF HO T WIRE A NEMOMET ERS ..................................................

D-1APPENDIX E THERMOCOUPLE LANCE ANOM ALY ...................................................................

E-1 vii LIST OF FIGURES Figure 1-1Typical vertical aboveground storage cask system. ................................................. 1-1Figure 1-2Typical vertical belowground storage cask system. ................................................. 1-1Figure 2-1General design showing the plan view (upper left), the internal helium flow (lower left), and the external air flow for the aboveground (middle) and belowground configurations (right).

........................................................................... 2-2Figure 2-2Carbon steel pressure vessel. .................................................................................. 2-3Figure 2-3CYBL facility housing the aboveground version of the BWR cask simulator. .......... 2-4Figure 2-4Typical 99 BWR components used to construct the test assembly including top tie plate (upper left), bottom tie plate (bottom left) and channel box and spacers assembled onto the water rods (right). ....................................................... 2-5Figure 2-5Typical TC attachment to heater rod. ....................................................................... 2-6Figure 2-6Experimental BWR assembly showing as-built a) axial and b) lateral thermocouple locations. ............................................................................................ 2-7Figure 2-7Definition of coordinate references in test apparatus. .............................................. 2-8Figure 2-8BWR channel box showing thermocouple locations. ............................................... 2-9Figure 2-9Storage basket showing thermocouple locations. .................................................. 2-10Figure 2-10Pressure vessel showing thermocouple locations. ................................................. 2-11Figure 2-11Ducting for aboveground configuration showing thermocouple locations.

............. 2-12Figure 2-12 Ducting for belowground configuration showing thermocouple locations. ............. 2-13Figure 2-13Location of thermocouples for gas temperature measurements at elevations of 1.219, 2.438, 3.658 m (48, 96, and 144 in.). ...................................................... 2-14Figure 2-14TC elevations for the TC lance. .............................................................................. 2-15Figure 2-15Power control system and test circuits. .................................................................. 2-17Figure 2-16Schematic of the instrumentation panel for voltage, current and power measurements. ........................................................................................................ 2-17Figure 2-17Photographs of the two types of hot wire anemometer tips. .................................. 2-18Figure 2-18Photograph of the honeycomb element used for flow straightening. ..................... 2-19Figure 2-19Aboveground configuration showing the location of the hot wire anemometer. .... 2-20Figure 2-20Mass flow rate as a function of hot wire output for forced flow. ............................. 2-20Figure 2-21Schematic showing the location of the inlet duct profiles for aboveground testing. .....................................................................................................................

2-21Figure 2-22Diagram showing the integration scheme for the calculation of air mass flow rate for the aboveground configuration. .................................................................. 2-21Figure 2-23Natural-to-forced flow correlation. ........................................................................... 2-22Figure 2-24Location of air flow measurement instrumentation for the belowground configuration. ........................................................................................................... 2-23Figure 2-25Radial positioning of the hot wire anemometers for belowground testing. ............ 2-23Figure 2-26Diagram showing the integration scheme for the calculation of air mass flow rate for the belowground configuration. .................................................................. 2-24Figure 2-27Layout of the cask simulator and wind machine for cross-wind testing. ................ 2-25Figure 2-28Schematic showing the local coordinates of the wind machine. ............................ 2-25Figure 2-29Velocity contours of the wind machine for maximum cross-wind. .......................... 2-26Figure 2-30Correlation of the two-dimensional, integrated average velocity (W2D, avg) to the average of the three fixed hot wire anemometers (W3-Pt, avg). ........................... 2-26Figure 3-1Steady state peak cladding temperature as a function of power. ............................ 3-1Figure 3-2Steady state air flow rate as a function of power. ..................................................... 3-2Figure 3-3Steady state peak cladding temperature as a function of absolute internal vessel pressure. ........................................................................................................ 3-2 viii Figure 3-4Steady state air mass flow rate as a function of absolute internal vessel pressure. ....................................................................................................................

3-3Figure 3-5Steady state temperature contours for 5.0 kW at different internal helium pressures. .................................................................................................................. 3

-4Figure 3-6Steady state temperature contours for 0.5 kW at different internal vessel pressures. .................................................................................................................. 3

-4Figure 3-7Steady state transverse temperature profile at z = 3.023 m (119 in.) for the test conducted at 5.0 kW and 800 kPa helium. ........................................................ 3-5Figure 3-8Steady state transverse temperature profile at z = 3.023 m (119 in.) for the test conducted at 0.5 kW and 0.3 kPa air. ................................................................ 3-6Figure 3-9Peak cladding temperature as a function of time for tests conducted at 800 kPa helium. ................................................................................................................ 3-9Figure 3-10Total air mass flow rate as a function of time for tests conducted at 800 kPa helium. .......................................................................................................................

3-9Figure 3-11Time to reach steady state as a function of power for the various vessel pressures tested. ..................................................................................................... 3-10Figure 3-12Comparison of TC lance and cladding temperatures at z = 3.023 m (119 in.)

as a function of time for the test conducted at 5.0 kW and 800 kPa helium. ......... 3-11Figure 3-13Comparison of TC lance and cladding temperatures at z = 3.023 m (119 in.) as a function of time for the test conducted at 0.5 kW and 0.3 kPa air. ................. 3-11Figure 4-1Steady state peak cladding temperature as a function of power. ............................ 4-1Figure 4-2Steady state air mass flow rate in the inlet annulus as a function of power. ........... 4-2Figure 4-3Steady state peak cladding temperature as a function of absolute internal vessel pressure. ........................................................................................................ 4-2Figure 4-4Steady state air mass flow rate in the inlet annulus as a function of absolute internal vessel pressure. ........................................................................................... 4-3Figure 4-5Steady state velocity contours for 5.0 kW at different internal helium pressures. .................................................................................................................. 4

-3Figure 4-6Steady state velocity contours for 0.5 kW at different internal vessel pressures. .................................................................................................................. 4

-4Figure 4-7Steady state transverse temperature profile at z = 3.023 m (119 in.) for the test conducted at 5.0 kW and 800 kPa helium. ........................................................ 4-5Figure 4-8Steady state transverse temperature profile at z = 3.023 m (119 in.) for the test conducted at 0.5 kW and 0.3 kPa air. ................................................................ 4-5Figure 4-9Peak cladding temperature as a function of time for tests conducted at 800 kPa helium. ................................................................................................................ 4-8Figure 4-10Total air mass flow rate as a function of time for tests conducted at 800 kPa helium. .......................................................................................................................

4-9Figure 4-11Time to reach steady state as a function of power for the various vessel pressures tested. ....................................................................................................... 4-9Figure 4-12Comparison of TC lance and cladding temperatures at z = 3.023 m (119 in.) as a function of time for the test conducted at 5.0 kW and 800 kPa helium. ......... 4-10Figure 4-13Comparison of TC lance and cladding temperatures at z = 3.023 m (119 in.) as a function of time for the test conducted at 0.5 kW and 0.3 kPa air. ................. 4-11Figure 4-14Normalized air mass flow rates as a function of cross-wind speed for 1.0 kW tests. ........................................................................................................................

4-12Figure 4-15Normalized air mass flow rates as a function of cross-wind speed for 2.5 kW tests. ........................................................................................................................

4-13Figure 4-16Normalized air mass flow rates as a function of cross-wind speed for 5.0 kW tests. ........................................................................................................................

4-13 ix Figure 4-17Normalized air mass flow rates as a function of cross-wind speed for 100 kPa tests. ................................................................................................................. 4-14Figure 4-18Normalized air mass flow rates as a function of cross-wind speed for 800 kPa tests. ................................................................................................................. 4-14Figure 4-19Orientation of the wind machine and test assembly. .............................................. 4-15Figure 4-20Velocity contours for 5.0 kW and 100 kPa at different cross-wind speeds. ........... 4-15

xi LIST OF TABLES Table 2-1Dimensions of assembly components in the 99 BWR. .......................................... 2-5Table 2-2List of proposed equipment for power control. ....................................................... 2-18Table 3-1Steady state results for the primary assembly measurements at 0.3 kPa air. ......... 3-6Table 3-2Steady state results for the primary assembly measurements at 100 kPa helium. .......................................................................................................................

3-7Table 3-3Steady state results for the primary assembly measurements at 450 kPa helium. .......................................................................................................................

3-7Table 3-4Steady state results for the primary assembly measurements at 800 kPa helium. .......................................................................................................................

3-8Table 4-1Steady state results for the primary assembly measurements at 0.3 kPa air. ......... 4-6Table 4-2Steady state results for the primary assembly measurements at 100 kPa helium. .......................................................................................................................

4-6Table 4-3Steady state results for the primary assembly measurements at 450 kPa helium. .......................................................................................................................

4-7Table 4-4Steady state results for the primary assembly measurements at 800 kPa helium. .......................................................................................................................

4-7Table 4-5Rise in peak cladding temperature attributed to cross-wind conditions. ................ 4-11

xiii EXECUTIVE

SUMMARY

The thermal performance of commercial nuclear spent fuel dry storage casks is evaluated through detailed numerical analysis. These modeling efforts are completed by the vendor to demonstrate performance and regulatory compliance. The calculations are then independently verified by the Nuclear Regulatory Commission (NRC). Carefully measured data sets generated from testing of full sized casks or smaller cask analogs are widely recognized as vital for validating these models. Recent advances in dry storage cask designs have significantly increased the maximum thermal load allowed in a cask in part by increasing the efficiency of internal conduction pathways and by increasing the internal convection through greater canister helium pressure. These same canistered cask systems rely on ventilation between the canister and the overpack to convect heat away from the canister to the environment for both aboveground and belowground configurations. While several testing programs have been previously conducted, these earlier validation attempts did not capture the effects of elevated helium pressures or accurately portray the external convection of aboveground and belowground canistered dry cask systems. The purpose of this investigation was to produce validation-quality data that can be used to test the validity of the modeling presently used to determine cladding temperatures in modern vertical

dry casks. These cladding temperatures are critical to evaluate cladding integrity th roughout the storage cycle. To produce these data sets under well-controlled boundary conditions, the dry cask simulator (DCS) was built to study the thermal-hydraulic response of fuel under a variety of heat loads, internal vessel pressures, and external configurations. An existing electrically heated but otherwise prototypic BWR Incoloy-clad test assembly was deployed inside of a representative storage basket and cylindrical pressure vessel that represents a vertical canister system. The symmetric single assembly geometry with well-controlled boundary conditions simplified interpretation of results. Two different arrangements of ducting were used to mimic conditions for aboveground and belowground storage configurations for vertical, dry cask systems with canisters. Transverse and axial temperature profiles were measured throughout the test assembly. The induced air mass flow rate was measured for both the aboveground and belowground configurations. In addition, the impact of cross-wind conditions on the belowground configuration was quantified. Over 40 unique data sets were collected and analyzed for these efforts. Fourteen data sets for the aboveground configuration were recorded for powers and internal pressures ranging from 0.5 to 5.0 kW and 0.3 to 800 kPa absolute, respectively. Similarly, fourteen data sets were logged for the belowground configuration starting at ambient conditions and concluding with thermal-hydraulic steady state. Over thirteen tests were conducted using a custom-built wind machine.

The results documented in this report highlight a small, but representative, subset of the available data from this test series. This addition to the dry cask experimental database signifies a substantial addition of first-of-a-kind, high-fidelity transient and steady-state thermal-hydraulic data sets suitable for CFD model validation.

xv ABBREVIATIONS AND ACRONYMS ANSI American National Standards Institute BWR boiling water reactor DAQ data acquisition DCS Dry Cask Simulator DOE Department of Energy EPRI Electric Power Research Institute FCRD Fuel Cycle Research and Development MSB multi-assembly sealed basket NRC Nuclear Regulatory Commission PCT peak cladding temperature PID proportional-integral-differential controller PWR pressurized water reactor SCR silicon controlled rectifier SNF spent nuclear fuel SNL Sandia National Laboratories TCthermocouple VCC ventilated concrete cask

1-11 INTRODUCTION The thermal performance of commercial nuclear spent fuel dry storage casks is evaluated through detailed analytical modeling. These modeling efforts are performed by the vendor to demonstrate the performance and regulatory compliance and are independently verified by the Nuclear Regulatory Commission (NRC). Most commercial dry casks in use today store the fuel in an aboveground configuration, although belowground storage has grown in recent years. Both horizontally and vertically oriented aboveground dry cask systems are currently in use. Figure 1-1 shows a diagram for a typical vertical aboveground system. Cooling of the assemblies located inside the sealed canister is enhanced by the induced flow of air drawn in the bottom of the cask and exiting out the top of the cask. Figure 1-1 Typical Vertical Aboveground Storage Cask System Figure 1-2 shows a diagram for a typical, vertical belowground system. For belowground configurations air is drawn in from the top periphery and channeled to the bottom where it then flows upward along the wall of the canister and exits out the top center of the cask. Figure 1-2 Typical Vertical Belowground Storage Cask System Source:www.nrc.gov/readingrm/doccollections/factsheets/storagespent fuelfs.htmlSource:www.holtecinternational.com/productsandservices/wasteandfuelmanagement/historm/

1-2 Carefully measured data sets generated from testing of full sized casks or smaller cask analogs are widely recognized as vital for validating design and performance models. Numerous studies have been previously conducted [Bates, 1986; Dziadosz and Moore, 1986; Irino et al., 1987; McKinnon et al.,1986]. Recent advances in dry storage cask designs have significantly increased the maximum thermal load allowed in a cask in part by increasing the efficiency of internal conduction pathways and by increasing the internal convection through greater canister helium pressure. These vertical, canistered cask systems rely on ventilation between the canister and the overpack to convect heat away from the canister to the environment for both above and belowground configurations. While several testing programs have been previously conducted, these earlier validation attempts did not capture the effects of elevated helium pressures or accurately portray the external convection of aboveground and belowground canistered dry cask systems. Thus, the enhanced performance of modern dry storage casks cannot be fully validated using previous studies. 1.1 Objective The purpose of this investigation was to produce a data set with a detailed error analysis (see Appendix A) that can be used to test the validity of the modeling presently used to determine cladding temperatures in modern vertical dry casks, which are used to evaluate cladding integrity throughout the storage cycle. To produce these data sets under well-controlled boundary conditions, the dry cask simulator (DCS) was built to study the thermal-hydraulic response of fuel under a variety of heat loads, internal vessel pressures, and external configurations. The results documented in this report highlight a small, but representative, subset of the available data from this test series. To illustrate the breadth of the data sets collected for each test, an example channel list for the data acquisition system (DAQ) can be found in Appendix B. In addition, the results generated in this test series supplement thermal data collected as part of the High Burnup Dry Storage Cask Project [EPRI, 2014]. A shortened version of the thermal lance design deployed in the Cask Project was installed in the DCS. The installation of this lance in the DCS assembly allowed the measurement of temperatures inside of a "guide tube" structure and direct comparisons with fuel cladding. 1.2 Previous Studies 1.2.1 Small Scale, Single Assembly Two single assembly investigations were documented in the mid-1980s [Bates, 1986; Irino et al., 1987]. Both included electrically heated 1515 pressurized water reactor (PWR) assemblies with thermocouples installed to directly measure the surface temperature of the cladding. In Bates (1986) the electrically heated assembly was instrumented with 57 TCs distributed over 7 axial levels. In Irino et al. (1987) the electrically heated assembly was instrumented with 92 TCs distributed over 4 axial levels. In Bates (1986) a single irradiated 1515 PWR assembly was also studied using 105 thermocouples distributed equally into each of the fifteen guide tubes at seven axial levels. All experiments were limited to one atmosphere helium or air, and all imposed a constant temperature boundary condition on the outer cask wall in an attempt to achieve prototypic storage temperatures in the fuel assembly bundle. 1.2.2 Full-Scale, Multi-Assembly Several full-scale multi-assembly cask studies were also documented in the mid-1980s to early 1990s, one for a BWR cask with unconsolidated fuel assemblies [McKinnon et al., 1986] and the 1-3 others for PWR casks with both consolidated and unconsolidated fuel [Dziadosz et al., 1986; McKinnon et al., 1987; Creer et al., 1987; McKinnon et al.,1989; McKinnon et al., 1992]. Only in the most recent study was a ventilated cask design tested. In all studies the cask were studied with internal atmospheres ranging from vacuum up to 150 kPa (21.8 psia) using air, nitrogen, or helium. In the first study [McKinnon et al., 1986], 28 or 52 BWR assemblies with a total heat load of 9 or 15 kW, respectively, were contained in REA 2023 prototype steel-lead-steel cask with a water-glycol neutron shield. Thirty-eight TCs were installed on the cask interior. Twenty-four of those were installed in direct contact with the center rod in 7 assemblies at up to 7 different elevations. Twelve were installed on the basket at 3 different elevations. Two TCs were installed in direct contact with a fuel rod located on the center outer face of an assembly. The cask was tested in a vertical and horizontal orientation with atmospheres of vacuum or nitrogen at 145 kPa (21 psia) average or helium at 152 kPa (22 psia) average. In the earliest full scale PWR cask study [Dziadosz et al., 1986], twenty-one PWR assemblies with a total heat load of 28 kW were contained in a Castor-V/21 cast iron/graphite cask with polyethylene rod neutron shielding. The interior of the cask was instrumented with sixty thermocouples deployed on ten lances located in eight guide tubes and two basket void spaces.

Two of the assembly lances were installed into the center assembly. Note, with the use of TC lances inside of the assembly guide tubes; no direct fuel-cladding temperatures were measured. The cask was tested in a vertical and horizontal orientation with atmospheres of vacuum or nitrogen at 57 kPa (8.3 psia) or helium at 52 kPa (7.5 psia). A relatively low total heat load of 12.6 kW was tested in a Westinghouse MC-10 cask with 24 PWR assemblies [McKinnon et al., 1987]. The MC-10 has a forged steel body and distinctive vertical carbon steel heat transfer fins around the outer circumference. The outer surface of the cask was instrumented with 34 thermocouples. The interior of the cask was instrumented with 54 thermocouples deployed on 9 TC lances in 7 fuel assembly guide tubes and 2 basket void spaces. The cask was tested in a vertical and horizontal orientation and interior atmosphere was either a vacuum or 150 kPa (21.8 psia) helium or air. A pair of studies using the same TN-24 cask was tested with 24 PWR assemblies with 20.5 kW total output [Creer et al., 1987] or 24 consolidated fuel canisters with 23 kW total output [McKinnon et al.,1989]. The TN-24P has a forged steel body surrounded by a resin layer for neutron shielding. The resin layer is covered by a smooth steel outer shell. The TN-24P is a prototype version of the standard TN-24 cask with differences in the cask body thickness, basket material and neutron shield structure. The TN-24P also incorporates 14 thermocouples into the basket structure. In both studies the fuel was instrumented with 9 TC lances with 6 TCs per lance, 7 in fuel guide tubes and 2 in simulated guide tubes in basket void spaces. The outside surface was instrumented with 35 TCs in the unconsolidated fuel study [Creer et al., 1987] and 27 TCs in the consolidated fuel study [McKinnon et al., 1989]. In both studies the cask was tested in a vertical and horizontal orientation with the interior atmosphere as either a vacuum or 150 kPa (21.8 psia) helium or air. A seventh test was conducted in the consolidated fuel study [McKinnon

et al.,1989] for a horizontal orientation under vacuum, with insulated ends to simulate impact limiters. None of the previous studies discussed so far included or accounted for internal ventilation of the cask. Both of the single assembly investigations imposed constant temperature boundary conditions [Bates, 1986; Irino et al., 1987], and four full-scale cask studies discussed so far 1-4[Dziadosz et a l., 1986; McKinnon et al., 1987; Creer et al., 1987; McKinnon et al.,1989] considered externally cooled cask designs. In only one previous study was a ventilated cask design considered, and this cask was the VSC-17 [McKinnon et al., 1992]. The VSC-17 cask system consists of a ventilated concrete cask (VCC) and a removable multi-assembly sealed basket (MSB). The VCC is steel lined and incorporates four inlet vents to the outside neat the bottom and four outlet vents near the top.

When the MSB is placed inside the VCC, an annular gap is formed and the vents allow air to be drawn in from the bottom through the annular gap and out the top vents. The lid on the MSB is a specially designed bolted closure that seals the basket interior and closes off the top of the cask above the top vents. The VSC-17 is a specially designed test version (holding 17 PWR assemblies) of the commercial VSC-24 cask (holding 24 PWR assemblies). The VSC-17 is smaller and lighter and incorporates the bolted lid to facilitate testing. The VSC-24 is larger and utilizes a welded lid canister for containing the spent fuel assemblies. In the investigation of the VSC-17 cask, 17 consolidated PWR fuel canisters with a total heat load of 14.9 kW were utilized. The cask system was instrumented with 98 thermocouples. Forty-two of these were deployed on 7 TC lances with 6 TCs each. Six lances were installed in the fuel canisters and one was installed in a basket void space. Nine TCs were located on the outer MSB wall and 9 TCs were located on the inner VCC liner. Ten TCs were embedded in the VCC concrete wall. One TC was located at each vent inlet and outlet. Thirteen TCs were located on the outer cask surface and weather cover. Testing consisted of six runs, all in a vertical orientation. In four of the tests the MSB was filled with helium at an average pressure of 95 kPa (13.8 psia). The vents were either all unblocked, or the inlets were half blocked, or the inlets were fully blocked, or both the inlets and outlets were fully blocked. The other two runs were with unblocked vents and 84 kPa (12.2 psia) nitrogen or vacuum. 1.2.3 Uniqueness of Dry Cask Simulator This investigation differed from previous studies in several significant ways. Principle among these was that the canister pressure vessel was tested with helium pressures up to 800 kPa and assembly powers up to 5.0 kW until a steady state temperature profile was established. During the apparatus heating, the helium pressure was controlled to be constant to within +/-0.3 kPa (0.044 psi). Additionally, ventilated design boundary conditions for aboveground and belowground configurations were explicitly simulated. The present study also differs from previous studies in terms of experimental approach. Rather than striving to achieve prototypic peak clad temperatures by artificially imposing a temperature boundary condition on the canister wall, this study represented the physics of near-prototypic boundary conditions.

2-1 2 APPARATUS AND PROCEDURES This chapter describes th e various subsystems, construction , and methods used for this testing. The test apparatus desig n was guid ed by an attempt to match critical dimensionles s grou ps with prototypic systems as reasonably as possible, namely Re ynol ds , Rayleigh, and Nusselt nu m bers. The dimensional anal yses revealed that a sc aling distortion in simulated assembly power would b e nec ess a ry to more clos ely match the thermal-hydraulic response of a full-sized spent fuel storage cask. This need fo r additi onal decay heat is reas ona ble given the higher external surface-area-to-volume ratio of a single-assembly arrange ment as in the DCS compared to a modern canister with up to 89 ass emblies. A mo re rigorous treatmen t of the test apparatus d e sign was recorded and is available for furthe r details [Durbin, et al., 2016], and a summary of the dimensional analyses is provided in Appendi x C. Each phas e of experimental apparatus design and implement ation was al so guid ed by extensive, meticulous c omputational fluid d ynamics (CFD) modeling that is not explicit ly detailed in this repo rt. A brief description and e xample of modeling results may be fou nd in Zigh , et a l., 20 17. As an example, these models provi ded information o n the flo w profile development a nd thermal gradi ents th at were critical to the optimization of flow straightening a nd hot wire anemometer placements.

2.1 General

Construction The gene ral design details are show n in Figure 2-1. An existing electrically heat ed but otherwise protot ypic BWR Incoloy-clad test assembly was deployed insi de of a representative storage basket and cylindrical pre ssure vessel that represents the canist er. The symmetric single- assembly geometry wi th well-controlled boundary conditions simplified interpretation of results. Various configurat ions of oute r concen tric ducting were used to mimic conditions for aboveground and belowground storage configurations of vertical, dr y-cask system s with canist ers. Radial and axial temperature profiles we re measured for a wide range of decay power and canister pressu res. Of particular inte res t was the evaluation of the effect of increased helium pressure on heat load fo r both the aboveground and belowground configurations.

T he effect of wind spee d was also measured for the belowground configurat ion. Externally, air-mass flo w rates were calculated fr om measurements of t he induce d ai r velocities in the ex ternal ducting.

2-2Figure 2-1 General Design Showing the Plan View (upper left), the Internal Helium Flow (lower left), and the External Air Flow for the Aboveground (middle) and Belowground Configurations (right)

Fi gure 2-2 s hows t he ma jor c ar bo n s te el co mpo nen ts u s ed t o fa br ica te th e pr ess ur e v esse l. Th e 4.5 7 2 m (1 8 0 i n.) l o n g v e rt i c a l test s e c t i on was made fr o m 0.254 m (10 i n.) Sche d u l e 4 0 p i p e welded to Class 300 fla nges. The 0.35 6 0.2 5 4 m (1 4 1 0 i n.) S c h e d u l e 4 0 reduc i n g t e e w a s n e e d ed to fac i l ita te t h e ro u tin g o f o v e r 15 0 th e rm o cou p les (TC s) thr o ugh the p re s s u re ve s s e l. B l i nd f lan ge s w i t h thr eaded acce s s p or t s for T C a n d po wer lead pass-t h r o u g h s were b o lted to t he to p o f th e ve r t ic al t es t st an d s ec t i o n an d t he sid e s o f the red uc in g t ee. T he m ax imu m al lo w abl e working p ressur e w a s 2,4 0 0 k Pa at 4 0 0 C. B a r stock t a b s w e re w e l d ed ins id e t h e 0.25 4 m (10 i n.) fl ang e on th e te e to s up port t he t est ass emb ly and on th e to p o f th e t est s ect io n to al low a n ins ula t e d t op bo u n da r y con d i t ion. 10 i n. S c h. 40 p i pe ID = 10.02 in. MAWP = 24 bar at 400 C Channel Box "Basket Cell" "Canister" Abovegroun d Belowgroun dInternal Helium Flow Patterns Top of AssemblyBottom of AssemblyInstrumentation N eutral lea d Induced air flows Outside of shells insulated "Hot" electrical lead 2-3 Figure 2-2 Carbon Steel Pressure Vessel The test configurations were assembled and operated inside of the Cylindrical Boiling (CYBL) test facility, which is the same facility used for earlier fuel assembly studies [Lindgren and Durbin, 2007]. CYBL is a large stainless steel containment vessel repurposed from earlier flooded- containment/core-retention studies sponsored by DOE. Since then, CYBL has served as an excellent general-use engineered barrier for the isolation of high-energy tests. The outer vessel is 5.1 m in diameter and 8.4 m tall (16.7 ft. in diameter and 27.6 feet tall) and constructed with 9.5 mm (0.375 in.) thick stainless steel walls. Figure 2-3 shows a scaled diagram of the CYBL facility with the aboveground version of the test DCS inside. Reducing Tee (Instrument Well) 4.572 m (Test Section) 2-4 Figur e 2-3 CYBL Facility Housing the Aboveground Ver sion of th e BWR Cask Si mulator 2.2 Design of the Heated Fuel Bundle The highly pr ototypic fuel assembly was modeled after a 9 9 BWR fuel assembly.

Commercial components we re purchased to create the assembly, including the top and bottom tie plates, spacers, water rods, channel box , and all rela ted assembly hardware (se e Figure 2-4). Incoloy heater rods were su bstituted for the fuel rod pi ns for heated te sting. Due to fabrication constraints, the diameter of the Incolo y heaters was slightly s m aller than prototypic pins , 10.9 mm versus 11.2 mm. The slightly si mplified Incolo y mock fuel pins we re fabric ated based on drawings and physical e xamples from the nucl ear component supplier. The di mensions of th e assembly components are listed b elow in Table 2-1.

2-5 Tabl e 2-1 Dimensions of Assembly Components in the 99 BWR Description Lower (Full) Section Upper (Partial) Section Number of pins 74 66 Pin diameter (mm) 10.9 10.9 Pin pitch (mm) 14.4 14.4 Pin separation (mm) 3.48 3.48 Water rod OD (main section) (mm) 24.9 24.9 Water rod ID (mm) 23.4 23.4 Nominal channel box ID (mm) 134 134 Nominal channel box OD (mm) 139 139 Figur e 2-4 Typical 99 BWR Components Used to Construct the Test Assembly Including Top Tie Plate (upper left), Bottom Tie Plate (bottom left) and Channel Box and Spacers Assembled Onto the Water Rods (right) The thermocouples used are ungrounded-junction, Type K, with an Incoloy-sheath diameter of 0.762 mm (0.030 in.) held in intimate contact with the cladding by a thin Nichrome shim. This shim is spot welded to the cladding as shown in Figure 2-5. The TC attachment method allows the direct measurement of the cladding temperature.

2-6 Figure 2-5 Typical TC Attachment to Heater Rod 2.3 Instrumentation The test apparatus was instrumented with thermocouples (TCs) for temperature measurements, pressure transducers to monitor the internal vessel pressure, and hot wire anemometers for flow velocity measurement in the exterior ducting. Volumetric flow controllers were used to calibrate the hot wire probes. Voltage, amperage, and electrical power transducers were used for monitoring the electrical energy input to the test assembly. Ninety-seven thermocouples were previously installed on the BWR test assembly. Details of the BWR test assembly and TC locations are described elsewhere [Lindgren and Durbin, 2007].

Additional thermocouples were installed on the other major components of the test apparatus, such as the channel box, storage basket, canister wall, and exterior air ducting. TC placement on these components is designed to correspond with the existing TC placement in the BWR assembly. Hot wire anemometers were chosen to measure the inlet flow rate because this type of instrument is sensitive and robust while introducing almost no unrecoverable flow losses. Due to the nature of the hot wire measurements, best results are achieved when the probe is placed in an isothermal, unheated gas flow. 2.3.1 Thermocouples (TCs) 2.3.1.1 BWR Assembly TC locations The existing electrically-heated, prototypic BWR Incoloy-clad test assembly was previously instrumented with thermocouples in a layout shown in Figure 2-6. The assembly TCs are arranged in axial and radial arrays. The axial cross-section is depicted in Figure 2-6a, and radial cross-sections are shown in Figure 2-6b. The axial array A1 has TCs nominally spaced every 0.152 m (6 in.), starting from the top of the bottom tie plate (z o = 0 reference plane). Axial array A2 has TCs nominally spaced every 0.305 m (12 in.), and the radial arrays are nominally spaced every 0.610 m (24 in.). The spacings are referred to as nominal due to a deviation at the 3.023 m 2-7 (119 in.) elevation, resulting from interference by a spacer. Note that the TCs in the axial array intersect with the radial arrays. Figure 2-6 Experimental BWR Assembly Showing As-Built a) Axial and b) Lateral Thermocouple Locations Internal Thermocouples Cross section above partial rods z o = 0 Top of bottom tie plate Bypass holes - 2 24 48 72 96 119 144 Radial Array 24 in. spacing 9 TC each level 54 TC total Axial array A1 6 in. spacing 26 TCs Axial array A2 12 in. spacing 13 TCs Water rods inlet and exit 4 TCs Total of 97 TCs a b c d e f g h i q r s

t u v x

y z 24 & 96 48 & 119 a b c d e f g h i q r s

t u

v x y z (a) (b) 72 & 144 a b c d e f g h i q r s

t u

v x

y z Key for radial cross sections Axial array A1, 6 in. spacing Axial array A2, 12 in. spacing Radial array on rods, 24 in. spacing Radial array on water rods Partial rod locations TC lance location (Ends at 106 in. level) TC lance locations f all dimensions are in inches unless otherwise notedin.fm1443.6581193.023962.438 721.829 481.219 240.610 Quadrant 2 3 1 4 x y N E S W 2-8 Based on the need to optimally balance the TC routing through the assembly, the axial and radial array TCs were distributed among three separate quadrants, relying on the assumption of axial symmetry. Also shown in Figure 2-6 is the location of the TC lance (for more details see Section 2.3.1.8). The quadrant for the lance deployment was chosen to minimize the possibility of damaging any of the previously installed TCs. The TC spacing on the lance matched the elevation of the TCs in the upper portion of the A1 and A2 axial arrays and the radial array at 3.023 m (119 in.) and 3.658 m (144 in.) elevations. Figure 2-7 shows the definition of the reference coordinate system. The reference origin is defined as being in the center of the top of the bottom tie plate. The x-axis is positive in the direction of Quadrant 4 and negative in the direction of Quadrant 2. The y-axis is positive in the direction of Quadrant 3 and negative in the direction of Quadrant 1.

Figur e 2-7 Definiti on of Coor dinate Ref erences in T est Ap paratus 2.3.1.2 BWR Channel B ox T C L ocations The BW R channel box was instrumented with 25 TCs as depicted in Fig ure 2-8. Twenty-one of the TC s were on the cha nnel faces, three were on the corners and one was on the pe destal. The TCs on the faces of the channe l box were nomin ally located at lxl, lyl =

0.069, 0 m (2.704, 0 in.) or lxl, lyl = 0, 0.069 m (0, 2.704 in.), depending on th e quadrant in which they we re placed. TCs on the corners were located at lxl, lyl = 0.

065, 0.065 m (2.564, 2.564 in.). The refere nce plane, z o , was measured from the top of the bottom tie pla te, the sa me a s the BW R assembly.

Multiple T Cs on different faces at a given elevatio n were available to c heck the axial symmet ry assumption at 0.610 m (24 in.) intervals, starting at the z = 0.610 m (24 in.) elevation.

x y z Bottom tie plate N E S W 2-9 Figur e 2-8 BWR Channel Box Sho wing Thermocouple Lo cations 2.3.1.3 Storage B asket T C L ocations The stora ge basket was i nstrumented with 26 T Cs as depicted in Figure 2-9. Twenty-two of the TCs we re on the basket faces at the s ame positions as on the c hannel bo x, four were on the corners (the corner T C a t the 4.1 91 m (165 in.) level did not corresp ond to a chan nel box TC) and one was on the basket face at the elevat ion of the pedes tal. TCs located o n the bask et faces were located at lxl, lyl = 0, 0.089 m (0, 3.5 in.) and lxl, lyl = 0.08 9, 0 m (3.5, 0 in.). TCs on the corners were located at lxl, lyl = 0.08 3, 0.083 m (3.281 , 3.281 in.). The referenc e plane , z o , was measured from the to p of the botto m tie plate.

N E S W 2-10 Figur e 2-9 Storage B asket Sh owing Thermocouple Locations 2.3.1.4 Press ure Ve ssel TC Locations The pressu re vessel was instru mented with 27 TCs as depic ted in Figur e 2-10. Twenty-four of the TCs we re ali gned with the TCs on th e storage ba sket faces and three were alig ned with the T Cs on the storage basket corners.

TC s aligne d with the stora ge basket faces we re located at lxl, lyl = 0, 0.137 m (0, 5.375 in.) and lxl, lyl = 0.137, 0 m (5.375, 0 in.). TCs aligned with the storage

basket corners were located at lxl, lyl =

0.097 , 0.097 m (3.801, 3.801 in.). The re ference plane, z o , was measured fro m the top of th e botto m tie plate.

N E S W 2-11 Figur e 2-10 Pressure Vessel Showing Thermocouple Lo cations 2.3.1.5 Aboveground Configurati on Ducting TC L ocations The concentric air-flow duct for the abovegroun d configuration was instrumented with 27 thermocouples depicted i n Fig ure 2-11. Twen ty-four of the TC s were alig ned with th e TCs on the channe l box and storage basket faces; three were aligned with the corners. The face-aligned TCs were located at lxl, lyl = 0, 0.233 m (0, 9.16 4 in.) and lxl, lyl =

0.233, 0 m (9.164, 0 in.). The corner-align ed TCs were locate d at lxl, lyl = 0.165, 0.165 m (6.480, 6.4 80 in.). Th e reference plane, z o , wa s measured from the top of th e bottom tie plate. N E S W 2-12 Figur e 2-11 Ducti ng for Aboveground Configuration Showing Thermocouple Locations 2.3.1.6 Belowground Config urati on Ducting TC L ocations The concentric air-flow duct for the belowgro und configuration was instrumented with 24 thermocouples depicted i n Fig ure 2-12. Twenty-one of the TCs were align ed with the TCs on the channe l box and storage basket faces; three were aligned with the corners. The face-aligned TCs were nomin ally located at lxl, lyl =

0, 0.316 m (0, 12.42 7 in.) and lxl, lyl =

0.316, 0 m (12.427 , 0 in.). The corner-align ed TCs we re nominally loca ted at lxl, lyl =

0.223, 0.2 23 m (8.7 87, 8.78 7 in.). The referenc e plan e, z o , was measured from the top of the bottom tie pla te. N E S W 2-13 Figure 2-12 Ducting for Belowgrou nd Configuratio n Showi ng Thermocoupl e Locations 2.3.1.7 Gas Temp erature TC Locations Up to 37 T Cs were us ed to measure the temperat ure of the gas flo wing in the va rious regions of the test apparatus at three different elevati ons, as depicted in Figu re 2-13. For the aboveground

configuration testi ng, the outer most gas TC s were installed but the ou ter shell (s hell 2) was not in place. Th e center re gion sho wn in red de notes heliu m flowing upwa rd while it was heated insid e the assembly and storage basket.

Moving outward, the region sho wn in orange depicts heliu m flowi ng downwa rd as it c ooled alon g the inner pre ssure vessel wall. A total of 17 T C s were used for gas temp erature measurements inside the pressu re vess el. Mo re TCs were us ed at the upper

two ele vations where higher temper ature and temperature gradients were measured.

Moving further outward the region sh own in gree n is ai r moving upwa rd as it heated al ong the outer pressu re vesse l wall. The outer most region, sho wn in bl ue, is c ool air flowing downward in

the belowground configurati on. Fo r the aboveground configurati on, the outer blu e re gion was open to ambi ent. The narro w yellow region on the outsid e o f e ach of the concentric air ducts represents a 6 mm (0.2 5 in.) thick layer of high te mperature insulation.

N E S W 2-14 Figure 2-13 Location of Thermocouples for Gas Temperature Measurements at Elevations of 1.219, 2.438, 3.658 m (48, 96, and 144 in.) 2.3.1.8 Thermocouple Lance A custom TC lance was deployed in the upper portion of the test assembly above a partial length rod, as illustrated previously in Figure 2-6. Design details of the lance are shown in Figure 2-14. The design provided for a pressure boundary along the outer surface of the lance, with a pressure seal at a penetration in the top flange using standard tube fittings. The lance was made by the same fabricator using the same process and materials as the TC lances that were used in the full-scale High Burnup Dry Storage Cask Research and Development Project [EPRI, 2014]. The TC spacing was designed to correspond with TCs installed on the test assembly heater-rod cladding to provide a direct comparison between the two measurements. Direct comparisons between TC lance and corresponding clad-temperature measurements will aid in the interpretation of the TC lance data generated during the High Burnup Cask Project.

N E S W 2-15 Figure 2-14 TC Elevations for the TC Lance 2.3.2 Pressure Vessel Two high-accuracy, 0 to 3447 kPa (0 to 500 psia), absolute-pressure transducers (OMEGA PX409-500A5V-XL) were installed in the lower reducing tee for redundancy. The experimental uncertainty associated with these gauges is +/-0.03% of full scale, or +/-1.0 kPa (0.15 psi). At least one of these transducers was operational for each heated test. For testing below atmospheric pressure, a dedicated vacuum transducer 0 to 100 kPa (0 to 14.5 psia) absolute (OMEGA PXM409-001BV10V) was used in place of the higher-range absolute-pressure transducers. All penetrations and fittings were selected for the apparatus to have helium leak rates of 1E-6 std.

cm 3/s or better at 100 kPa. In addition, spiral-wound gaskets capable of leak rates of better than 1E-7 std. cm 3/s were used to form the seals at each flange. The ANSI N14.5 leak rate of 1E-4 std. cm 3/s [ANSI, 2014] would result in an observable pressure drop of 0.03 kPa (4E-3 psi) after a one week period, which is far below the experimental uncertainty of 1.0 kPa (0.15 psi). Leaks in the as-built apparatus were identified and repaired as best as possible. Ultimately, a small leak All dimensions in inches 2-16 path of undetermined origin remained, and a positive pressure control system was implemented to maintain pressure as described next. Under subatmospheric (0.3 kPa) conditions, the system leak path resulted in air infiltrating the pressure vessel. Therefore, the residual gas composition for 0.3 kPa testing was air, not helium. 2.3.2.1 Pressure Control A helium pressure control system was implemented using the high-accuracy, absolute-pressure transducers, three low-flow needle valves, and three positive-shutoff actuator valves under control of the LabView DAC system. Two actuator valves (vent) controlled helium flow out of the vessel, and the third valve (fill) controlled helium flow into the vessel. As the vessel heated up, the expanding helium was vented out the first actuator and needle valve to maintain a constant pressure. A second vent valve (overflow) activated if the vessel continued to pressurize. As steady state was reached, the small helium leak slowly reduced the helium pressure, at which point the control system opened the third actuator valve (fill) to allow a small helium flow through the third needle valve. Overall, the pressure control system maintained the helium pressure constant to +/-0.3 kPa (0.044 psi). For the subatmospheric tests, the pressure control system was not utilized. A vacuum pump was used to evacuate the vessel, and the ultimate vacuum achieved was a balance between the vacuum pump and the small amount of air leaking into the vessel. 2.3.2.2 Pressure Vessel Internal Volume Measurement The pressure vessel was pressurized with air in a manner that allowed the measurement of the as-built total internal volume. The pressure vessel was first pressurized to 100 kPa (14.5 psia). The pressure vessel was then slowly pressurized to 200 kPa (29.0 psia) with a high-accuracy 0 to 5 liters-per-minute flow controller (OMEGA FMA 2606A-TOT-HIGH ACCURACY). A high-accuracy, 0 to 3447 kPa (500 psia), absolute-pressure transducer (OMEGA PX409-500A5V-XL) was used to monitor the transient fill progression. The transient mass flow and pressure data were used to determine the total internal volume to be 252.0 liters, with an uncertainty of +/-2.6 liters. 2.3.3 Power Control A diagram of the test assembly power control system is shown in Figure 2-15, and the details inside the instrument panel are shown in Figure 2-16. The electrical voltage and current delivered to the test assembly heaters was controlled by a silicon controlled rectifier (SCR) to maintain a constant power. The data acquisition (DAQ) system provided a power setpoint to a PID controller that sent a control signal to the SCR based on the power measurement. The power, voltage, and current measurements were collected by the DAQ. The details of the instrumentation used to control and measure the electrical power are provided in Table 2-2.

2-17 Figure 2-15 Power Control System and Test Circuits Figure 2-16 Schematic of the Instrumentation Panel for Voltage, Current and Power Measurements Watt Transducer Voltage Transducer Voltage Signal Neutral Signals to DAQ Power Feedback Signal Current Transducer Current Signal ~5.0 kW @ 60 VAC 2-18 Table 2-2 List of Proposed Equipment for Power Control DescriptionManufacturer Mo delAC Watt Transducer Ohio Semitronics PC5-001DY230 AC Voltage Transducer Ohio Semitronics AVTR-001D AC Current Transducer Ohio Semitronics ACTR-005DY06 PID Controller Watlow Electric Manufacturing PM6C1FJ1RAAAASCR Power Controller Watlow Electric Manufacturing PC91-F25A-1000 2.3.4 Hot Wire Anemometers The hot wire anemometers used for this testing were TSI models 8475 and 8455. The sensor tip details are shown in Figure 2-17. For scale, the largest shaft diameter shown was 6.4 mm (0.25 in.). The sensing element of the model 8455 is protected inside of an open cage and is sensitive to flows down to 0.13 m/s (25 ft/min), with a fast response time of 0.2 seconds. The sensing element of the model 8475 is the ball at the tip, which results in sensitivity to flows down to 0.05 m/s (10 ft/min) but with a much larger response time of 5 seconds. Hot wire anemometers were chosen to measure the inlet flow rate because this type of instrument is sensitive and robust, while introducing almost no unrecoverable pressure loss. Due to the nature of the hot wire measurement, for best results the probes were placed in the gas flow at the flow inlets where temperature and thermal gradients were minimal. Figure 2-17 Photographs of the Two Types of Hot Wire Anemometer Tips 2.4 Air Mass Flow Rate The methods for determining the induced air flow in the aboveground and belowground configurations were similar but have some distinct differences. Both methods used hot wire anemometers to measure inlet air velocity and subsequently calculate an overall air-mass flow rate. For the aboveground configuration, the hot wires were fixed in the center of the inlet ducts and subjected to known mass flow rates of air using mass-flow controllers during a series of pre-test measurements. The output of the hot wires was then correlated to the forced mass flow rate input. Additionally, a velocity profile was measured along the short dimension of the center of the inlet during steady state operation of each heated, buoyancy-driven (natural) test. A mass flow rate was calculated from these velocity profiles and provided a correction correlation between the natural-to-forced flow data.

2-19For the belowground configuration, forced flow calibration in the annulus between Shell 1 and Shell 2 was not possible. The mass flow was determined by integrating the velocity profiles of multiple hot wire anemometers positioned around the annulus. For belowground testing, eight hotwires were mounted on motorized stages (Velmex Stage XN10-0040-M02-71, Motor PK245-01AA) at equidistant positions. The data acquisition computer communicated with the stage controller (Velmex Controller VXM-4) to identify and verify hot wire positioning. An additional four hot wires were added to one half of the Shell 1 and Shell 2 annulus for belowground, cross-wind testing to more accurately measure the effect of larger velocity gradients. 2.4.1 Flow Straightening To obtain the most stable and repeatable measurements possible, a honeycomb element was inserted into the inlets of both the aboveground and belowground configurations. This honeycomb served to align the flow in the desired direction and reduce any flow disturbances on the hot wire measurements. As shown in Figure 2-18, a plastic honeycomb element was chosen with a cell diameter, wall thickness, and flow length of 3.8, 0.1, and 51.6 mm (0.150, 0.004, and 2.030 in.), respectively. This type of flow straightening element was found to provide the greatest reduction in hot wire fluctuations while introducing the smallest pressure drop to the system. The effective, frictional coefficient for this honeycomb material was found to be D = 2.7E6 m

-2 for porous media in CFD simulations.

Figur e 2-18 Photograph of the Hon eycomb Element Us ed for Flow Str aightening

2.4.2 Abovegrou

nd Air Flow Measurement The in let and hot wi re arrangement for the aboveg round configuration is s hown in Figure 2-19.

Four rectangular ducts wi th as-bu ilt cross sectional dimensions of 0.22 9 m (9.03 in.) by 0.100 m (3.94 in.) convey ed the i nlet flow into the simulated cask.

One TSI Model 8475 a nd three TSI Model 8455 hot wire anemometers were used for these tests.

Hot wire anemometers were located 0.22 9 m (9.00 in.) downstream from the i nlet of e a ch duct along the centerline of flow. 51.6 Circular Cells 3.8 twall = 0.1 All dimensions in mm

2-20 Figure 2-19 Aboveground Configuration Showing the Location of the Hot Wire Anemometer 2.4.2.1 Forced Flow Correlation The outputs of the hot wire anemometers were correlated using metered, forced flow. Air flow was metered into each of the inlet ducts individually, and the response of each anemometer in the center of the inlet recorded for a range of flow rates as shown in Figure 2-20. A least-squares regression was used to define the linear coefficients to convert the hot wire anemometer output to mass flow rate during heated testing. Figure 2-20 Mass Flow Rate as a Function of Hot Wire Output for Forced Flow Honeycomb flow straightener Hot wire anemometer 0.229 m .

2-21 2.4.2.2 Inlet Duct Flow Profiles Velocity profiles were collected across the short dimension (0.100 m) at the end of each powered test. The profiles were measured with the hot wire anemometer along the x-axis of the duct at 0.229 m (9.00 in.) from the duct entrance as shown in Figure 2-21. Figure 2-21 Schematic Showing the Location of the Inlet Duct Profiles for Aboveground Testing These velocity profiles were integrated to determine the relationship of the air-mass flow rate during heated, buoyancy-driven testing to that measured during the forced flow testing. The integrated, natural air-mass flow rate is given in Equation 2.1. Here, the reference density is defined by the standard conditions for the TSI hot wires, or ref = 1.2 kg/m 3 at 21.1 °C and 101.4 kPa. The area for each measurement is given by the product of the profile step size, x, and the width of the inlet duct (W = 0.229 m). Figure 2-22 gives a visual representation of the integration scheme. 2.1 Figure 2-22 Diagram Showing the Integration Scheme for the Calculation of Air Mass Flow Rate for the Aboveground Configuration W x y x w 1 w N x z y Profiles along dashed line 2-222.4.2.3 Natural-to-Forced Flow Correlation Air-mass flow rates from the natural (integrated profiles) and forced (mass flow controller) methods were compared after testing. Recall, flow velocity data was collected with the hot wires centrally located in the ducts during general testing and was converted to mass flow rate using the pre-test forced flow correlations. Velocity profiles were recorded only at the end of each heated test when steady state was achieved. This comparison, as shown in Figure 2-23, revealed that the natural air-mass flow rate was less than that indicated from the forced-flow correlation by a factor of 0.9344. Therefore, the two correlations are applied successively to the hot wire voltage to obtain the best estimate of air mass flow rate. Comparisons of velocity profiles revealed that the boundary layer for the natural flow was larger than the forced flow case. This difference corresponded to the lower observed mass flow rate for natural conditions.

Figur e 2-23 Natural-To-Forced Flow Corr elation 2.4.3 Belowgrou nd Ai r Flow Measurement The in let and hot wi re arrangement f or the belowground configuration is s hown in Figure 2-24.

Velocity prof iles were col l ected across the annular ga p defined by sh ell 1 and shell 2 d uring heated testing at z = 0.508 m (20.00 in.) or 3.336 m (131.37 in.) from the bottom of the inlet duct.

The profil es were meas ured from the inner surface of s hell 2 to the ou ter surface of the ins ula tion attac h ed to shell 1 as shown in Figure 2-24.

. .

2-23 Figure 2-24 Location of Air Flow Measurement Instrumentation for the Belowground Configuration Figure 2-25 shows the radial positioning for the hot wire anemometers for the both phases of the belowground testing. The first arrangement with eight equally-spaced hot wires was used for powered testing without cross-wind. Four additional hot wires were added in the second configuration along one half of the annulus to measure larger velocity gradients than possible with 45° spacing. Figure 2-25 Radial Positioning of the Hot Wire Anemometers for Belowground Testing Profiles along dashed line 0.508 3.238 Honeycomb flow straightener N E S W 0.606 Hot wires z Air inlet Air outlet All dimensions in meters Hot wire ports -8 plcs.Hot wire ports -12 plcs. (Cross-wind) 22.5° Automated traverses in annulus 45° N S E W Cross-wind

2-24 The velocity profiles from the hot wires were integrated to calculate the air mass flow rate during heated, buoyancy-driven testing. The integrated, natural air-mass flow rate is given in Equation 2.2. Again, the reference density is defined by the standard conditions for the TSI hot wires, or ref = 1.2 kg/m 3 at 21.1 °C and 101.4 kPa. The area for each measurement is given by the product of the radius, r, profile step size, r, and the arc angle in radians, . The arc angle for a given hot wire is assumed to bisect the azimuths formed between the index hot wire and the nearest hot wires. The first index is defined as the hot wire identifier. The second index denotes the radial position. Figure 2-26 gives a visual representation of the integration scheme. Verification tests were conducted to determine the accuracy of determining the air mass flow rate through velocity measurements and integration as discussed in Appendix D.

2.2 Figure

2-26 Diagram Showing the Integration Scheme for the Calculation of Air Mass Flow Rate for the Belowground Configuration 2.5 Cross-Wind Testing A wind machine was fabricated and installed in the CYBL vessel to study the effect of a continuous cross-wind on the thermal and hydraulic response of the system. This wind machine consisted of three air-driven blowers connected to a specially fabricated duct with outlet dimensions of 1.295 0.762 m (51.0 30.0 in.). The duct served two purposes. First, it redirected the flow from a vertical orientation to a horizontal direction via a long-sweep elbow. Second, the duct allowed the insertion of flow straightening elements to make the air velocity at the outlet as uniform as reasonably achievable. The top and bottom of the wind machine duct outlet were installed approximately 0.12 m (4.625 in.) above the DCS air outlet and 0.18 m (7.25 in.) below the DCS air inlet, respectively. The distance between the outer edge of the DCS air inlet and the duct outlet was 0.17 m (6.75 in.). The wind machine was centered side-to-side on the DCS assembly with the duct extending 0.13 m (5.25 in.) on either side of the DCS air inlet. Figure 2-27 shows the position of the wind machine relative to the assembly. A local coordinate system for the wind machine is defined in Figure 2-28.

w1,1 w 1, N r w2,1 w 2, N 1 HW 1 HW 2 2 HW 3 3/2 w3,1 w 3, N HW M M/2 r w M , N w M ,1 2-25Figure 2-27 Layout of the Cask Simulator and Wind Machine for Cross-Wind Testing Figur e 2-28 Schemati c Showing th e Local Coordinates of t he Wind Ma chine Hot wire measurements were taken across the wind machine outlet to determine wind speed and unifo rmity. Prior to heated testing, h o t wire measuremen ts were taken fo r three differe nt wind speeds at 45 re gularly s paced locations.

Figur e 2-29 sho ws the velocity contours of o ne such effort near the upp er range of achievab le wind spe eds (W 2D, avg = 5.2 m/s {11.6 mp h}). For heated cross-wind testing, two-dimensiona l mapping was not possible. Therefore, ho t wire a nemometers were fixed at thre e locations as shown in Figure 2-

29. Figure 2-30 gi ves the correlation betwe en the i ntegrated av erage velocity (W 2D, avg) and th e average of the three hot wires (W 3-Pt, avg). This correlation was applied to the 3-point avera ge to provide an estimate of the average wind speed at the ou tlet of th e wind m achine for heated testi ng. 0.17 0.18 0.12 All dimensions in meters x y z z y x N S E W Origin at center of the face of the duct outlet

2-26 Figure 2-29 Velocity Contours of the Wind Machine for Maximum Cross-Wind Note: The fixed positions of the hot wires used for the 3-point average wind speed are marked in the figure.

Figure 2-30 Correlation of the Two-Dimensional, Integrated Average Velocity (W2D, avg) to the Average of the Three Fixed Hot Wire Anemometers (W3-Pt, avg) w (m/s) Locations for 3-Point Averaging (Fixed Hot Wire Positions)

3-13 ABOVEGROUND RESULTS 3.1 Steady State Analyses A total of fourteen tests were conducted, where the apparatus achieved steady state for various assembly powers and pressures. The power levels tested were 0.5, 1.0, 2.5, and 5.0 kW. The vessel pressures tested were vacuum (0.3 kPa), 100, 450, and 800 kPa absolute. A scaling analysis [Durbin, et al., 2016] showed that elevated powers up to 5.0 kW were warranted to drive the induced air flow to prototypic levels. The criterion for steady state was considered met when the first derivative with respect to time of any given TC in the test apparatus was 0.3 K/h. The steady state values reported here represent the average of data collected between the "start of steady state" and the end of the test. 3.1.1 Peak Cladding Temperature and Air Mass Flow Rate Figure 3-1 and Figure 3-2 present the steady state data as peak cladding temperature (PCT) and total induced air flow rate, respectively, as a function of power for each vessel pressure tested. Figure 3-3 and Figure 3-4 present the same PCT and flow data but as a function of vessel pressure for each power tested. Generally, the peak temperatures and induced air flow both increased significantly with power level and decreased slightly with helium pressure. The notable exception was that the peak cladding temperature increased significantly as the vessel pressure was decreased from 100 kPa absolute helium to 0.3 kPa absolute air. Recall that subatmospheric testing resulted in a vessel gas composition of air due to the leak path discussed in Section 2.3.2. Figure 3-1 Steady State Peak Cladding Temperature as a Function of Power 3-2 Figure 3-2 Steady State Air Flow Rate as a Function of Power Figure 3-3 Steady State Peak Cladding Temperature as a Function of Absolute Internal Vessel Pressure 3-3 Figure 3-4 Steady State Air Mass Flow Rate as a Function of Absolute Internal Vessel Pressure 3.1.2 Two-Dimensional Temperature Contours Figure 3-5 shows 2-D temperature contour plots from the center of the assembly through the basket, pressure vessel, shell 1, and ambient for the high-power tests (5.0 kW) at the three helium pressures tested (100, 450, and 800 kPa absolute). Figure 3-6 shows 2-D temperature contour plots for the low power tests (0.5 kW) at the four vessel pressures tested (0.3, 100, 450 and 800 kPa absolute). For both power levels, the peak temperatures decreased with increasing vessel pressure. The location of the PCT also shifted from ~1/3 of the assembly height to near the top of the assembly for vessel pressures of 0.3 to 800 kPa, respectively.

3-4 Figure 3-5 Steady State Temperature Contours for 5.0 kW at Different Internal Helium Pressures Figure 3-6 Steady State Temperature Contours for 0.5 kW at Different Internal Vessel Pressures P = 100 kPa P = 450 kPa P = 800 kPa Temp. (K) P = 100 kPa P = 450 kPa P = 800 kPa Temp. (K) P = 0.3 kPa

3-5 3.1.3 Transverse Temperature Profiles including the TC Lance Figure 3-7 shows the steady state transverse temperature profile at the z = 3.023 m elevation for the 5.0 kW and 800 kPa aboveground case. Figure 3-8 shows a similar steady-state transverse temperature profile at the 3.023 m elevation for the 0.5 kW and 800 kPa case. The TC lance was located at y = -0.042 m. The assembly TCs for comparison with the TC lance were located starting at x = 0 m and continued along the negative x-direction. Assuming symmetry, the lance is plotted on the x-axis. The TC lance was in good agreement with the interpolated temperature of the two closest assembly TCs. As received and installed, the lance TCs above the 3.023 m (119 in.) elevation exhibited anomalous behavior during some tests as discussed in detail in Appendix E. TC lance data for the 3.023 m (119 in.) elevation is presented because no anomalous behavior was evident. A modification was made to the TC lance that eliminated the anomalous behavior for the affected TCs shortly before cross-wind testing of the belowground configuration, which was the last phase of testing. The behavior of the TCs at the 3.023 m (119 in.) elevation and below was not impacted by the modification. Figure 3-7 Steady State Transverse Temperature Profile at z = 3.023 m (119 in.) for the test Conducted at 5.0 kW and 800 kPa Helium x y 3-6 Figure 3-8 Steady State Transverse Temperature Profile at z = 3.023 m (119 in.) for the Test Conducted at 0.5 kW and 0.3 kPa Air 3.1.4 Summary Data Tables The steady-state value of the peak temperature for each region of the test apparatus is presented in the following summary tables. Table 3-1 through Table 3-4 present these peak temperatures and corresponding location along with the measured power, ambient temperature, and induced air mass flow rate for each power level tested at a given vessel pressure. The corresponding minimum and maximum values over the steady-state measurement period are also presented. Table 3-1 Steady State Results for the Primary Assembly Measurements at 0.3 kPa Air Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4924584043613283122992.53E-02Max0.5104594053623303153032.87E-02Min0.4724564033613283112962.17E-02LocationDT_2_48Channel_4_48Basket_3_72PV_2_108S1_2_119AllAssembly TotalAverage1.0045494704063513233013.51E-02Max1.0415504714073523243033.84E-02Min0.9345494704063513222993.14E-02LocationDT_1_24Channel_4_48Basket_3_72PV_1_96S1_2_119AllAssembly Total 1 0.5 x y 3-7 Table 3-2 Steady State Results for the Primary Assembly Measurements at 100 kPa Helium Table 3-3 Steady State Results for the Primary Assembly Measurements at 450 kPa Helium Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Ambient (K)Air Flow Rate (kg/s)Average0.5043763593443283122982.64E-02Max0.5253763593443283123002.88E-02Min0.4823753593443283112962.44E-02LocationFV_3_72Channel_4_72Basket_4_96PV_2-3_119S1_2_119AllAssembly TotalAverage1.0014344053783503212993.53E-02Max1.0174354053793503213013.75E-02Min0.9854344043783493212983.21E-02LocationFV_3_72Channel_4_72Basket_3_72PV_2-3_119S1_2_119AllAssembly TotalAverage2.4935705114614033483005.31E-02Max2.5165705114614033483025.61E-02Min2.4715705114604023472985.02E-02LocationDT_2_48Channel_3_60Basket_3_72PV_2-3_119S1_2_119AllAssembly TotalAverage5.0107156305544673873016.89E-02Max5.0397166315554683893057.21E-02Min4.9697146285534663852996.54E-02LocationDT_2_48Channel_4_48Basket_3_72PV_2-3_119S1_2_119AllAssembly Total0.5 12.5 5 Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Ambient (K)Air Flow Rate (kg/s)Average0.5133673533413263112962.41E-02Max0.5293673533413273122992.66E-02Min0.4893673523403263102932.07E-02LocationFV_3_144Channel_2_119Basket_3_132PV_2-3_119S1_4_159AllAssembly TotalAverage1.0474263993773513232993.28E-02Max1.0734273993773513243023.63E-02Min1.0184253973763503222952.82E-02LocationFV_3_144Channel_2_119Basket_3_132PV_3_144S1_4_159AllAssembly TotalAverage2.4915454944514013463004.76E-02Max2.5515464954524023483035.06E-02Min2.4565434924493993452994.52E-02LocationDT_1_96Channel_2_119Basket_2_108PV_2-3_119S1_3_132AllAssembly TotalAverage4.9726896125474653842996.55E-02Max5.0306906135484663863026.87E-02Min4.9106896115474643832976.16E-02LocationDT_1_96Channel_1_84Basket_2_108PV_2-3_119S1_2_119AllAssembly Total2.5 5 10.5 3-8 Table 3-4 Steady State Results for the Primary Assembly Measurements at 800 kPa Helium 3.2 Transient Analyses Figure 3-9 and Figure 3-10 show the peak cladding temperature and total assembly air mass flow rate for each power tested at 800 kPa absolute helium pressure. The air flow rate data was smoothed over a fifteen-minute moving window for clarity of presentation. Ninety-five percent uncertainties are also presented for select data points, 1% of reading for temperature (+/-7 K maximum) and +/-1.5E-3 kg/s for flow rate.

Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4993593473383293122982.21E-02Max0.5163593473383293122992.43E-02Min0.4843583473383293122961.91E-02LocationFV_3_144Channel_3_144Basket_4_159PV_1_156S1_4_159AllAssembly TotalAverage0.9854103883743563232973.10E-02Max1.0584103893743563243003.48E-02Min0.9674103883733553232942.72E-02LocationFV_3_144Channel_3_144Basket_4_159PV_4_159S1_4_159AllAssembly TotalAverage2.5035214774444083492984.69E-02Max2.5475214774444093503034.92E-02Min2.4445214774434083492964.39E-02LocationFV_3_144Channel_3_144Basket_4_159PV_4_159S1_4_159AllAssembly TotalAverage4.9976595905334663873006.26E-02Max5.0216595905334673873036.60E-02Min4.9566585895324663872995.99E-02LocationFV_3_144Channel_3_144Basket_3_144PV_4_159S1_4_159AllAssembly Total 50.5 12.5 3-9 Figure 3-9 Peak Cladding Temperature as a Function of Time for Tests Conducted at 800 kPa Helium Figure 3-10 Total Air Mass Flow Rate as a Function of Time for Tests Conducted at 800 kPa Helium Steady state conditions were reached in about 15 hours1.736111e-4 days 0.00417 hours 2.480159e-5 weeks 5.7075e-6 months . Figure 3-11 shows the time required to reach steady state as a function of power for the various test pressures. The time to steady state was independent of power and helium pressure for the 450 kPa and 800 kPa cases. For the 100 kPa helium pressure tests there was a slight dependence on power with 13 hours1.50463e-4 days 0.00361 hours 2.149471e-5 weeks 4.9465e-6 months required at 5.0 3-10 kW and 18 hours2.083333e-4 days 0.005 hours 2.97619e-5 weeks 6.849e-6 months required at 0.5 kW. The vacuum tests were the most sensitive to power, with up to 31 hours3.587963e-4 days 0.00861 hours 5.125661e-5 weeks 1.17955e-5 months required to reach steady state in the 0.5 kW case. Figure 3-11 Time to Reach Steady State as a Function of Power for the Various Vessel Pressures Tested 3.2.1 Transient Response of TC Lance and Corresponding Cladding Figure 3-12 shows the temperature of the TC lance and adjacent cladding TCs (assuming symmetry) as a function of time at the 3.023 m elevation for the 5.0 kW and 800 kPa case. Figure 3-13 shows the temperature of the TC lance and adjacent cladding TCs at the same elevation fo r the 0.5 kW and 0.3 kPa case. Ninety-five percent uncertainties are also presented for select datapoints as 1% of reading for temperature (+/-7 K maximum). The transient response of the TC lance and the adjacent cladding TCs were similar. The temperature indicated by the lance TC was roughly midway between the adjacent clad TCs. The good agreement provided validation that th e TC lance provides an accurate indication of nearby cladding temperatures. Again, TC lance datafor the 3.023 m (119 in.) location is presented because no anomalous behavior was evident at this elevation.

3-11 Figure 3-12 Comparison of TC Lance and Cladding Temperatures at z = 3.023 m (119 in.) as a function of Time for the Test Conducted at 5.0 kW and 800 kPa Helium Figure 3-13 Comparison of TC Lance and Cladding Temperatures at z = 3.023 m (119 in.) as a Function of Time for the Test Conducted at 0.5 kW and 0.3 kPa Air

4-14 BELOWGROUND RESULTS 4.1 Steady State Analyses A total of fourteen tests were conducted, where the apparatus achieved steady state for various assembly powers and vessel pressures. The power levels tested were 0.5, 1.0, 2.5 and 5.0 kW. The vessel pressures tested were vacuum (0.3 kPa), 100, 450 and 800 kPa absolute. A scaling analysis [Durbin, et al., 2016] showed that elevated powers up to 5.0 kW were warranted to drive the induced air flow to prototypic levels. Again, a summary of these dimensional analyses is provided in Appendix C. The criterion for steady state was considered met when the first derivative with respect to time of any given TC in the test apparatus was 0.3 K/h. The steady state values reported here represent the average of data collected between the "start of steady state" and the end of the test. 4.1.1 Peak Cladding Temperature and Air Mass Flow Rate Figure 4-1 and Figure 4-2 present the steady-state data as peak cladding temperature (PCT) and integrated air-mass flow rate in the inlet annulus, respectively, as a function of power for each vessel pressure tested. Figure 4-3 and Figure 4-4 present the same PCT and mass flow rate data but as a function of vessel pressure for each power tested. As in the aboveground configuration, the peak temperatures and induced air mass flow rate for the belowground configuration both increased significantly with power level and decreased slightly with helium pressure. The notable exception was that the peak cladding temperature increased significantly as the vessel pressure was decreased from 100 kPa absolute helium to 0.3 kPa absolute air. Recall that subatmospheric testing resulted in a vessel gas composition of air due to the leak path discussed in Section 2.3.2. Figure 4-1 Steady State Peak Cladding Temperature as a Function of Power 4-2 Figure 4-2 Steady State Air Mass Flow Rate in the Inlet Annulus as a Function of Power Figure 4-3 Steady State Peak Cladding Temperature as a Function of Absolute Internal Vessel Pressure 4-3 Figure 4-4 Steady State Air Mass Flow Rate in the Inlet Annulus as a Function of Absolute Internal Vessel Pressure 4.1.2 Two-Dimensional Velocity Contours Figure 4-5 shows 2-D velocity contour plots in the inlet annulus of the assembly for the high-power tests (5.0 kW) at the three helium pressures tested (100, 450, and 800 kPa absolute). As shown in Figure 4-5, the honeycomb flow straightening element was installed in two "C" pieces creating two seams. Because of the installation method, the honeycomb was likely compressed, especially at the seams. A deficit in the flow is observable in the velocity contour plots, particularly at these seams, indicating non-ideal behavior in the flow straightening. Figure 4-6 shows 2-D velocity contour plots for the low power tests (0.5 kW) at the four vessel pressures tested (0.3, 100, 450, and 800 kPa absolute). Figure 4-5 Steady State Velocity Contours for 5.0 kW at Different Internal Helium Pressures P = 100 kPa P = 450 kPa P = 800 kPa Velocit y (m/s) = 6.99E-2 kg/s = 6.51E-2 kg/s = 6.11E-2 kg/s Honeycomb seams 4-4 Figure 4-6 Steady State Velocity Contours for 0.5 kW at Different Internal Vessel Pressures 4.1.3 Transverse Temperature Profiles Including the TC Lance Figure 4-7 shows the steady state transverse temperature profile at the z = 3.023 m elevation for the 5.0 kW and 800 kPa belowground case. Figure 4-8 shows a similar steady state transverse temperature profile at the 3.023 m elevation for the 0.5 kW and 800 kPa case. The TC lance was

located at y = -0.042 m. The assembly TCs for comparison with the TC lance were located starting at x = 0 m and continued along the negative x-direction. Assuming symmetry, the lance is plotted on the x-axis. The TC lance was in good agreement with the interpolated temperature of the two closest assembly TCs. As received and installed, the lance TCs above the 3.023 m (119 in.) elevation exhibited anomalous behavior during some tests as discussed in detail in Appendix E. TC lance data for the 3.023 m (119 in.) elevation is presented because no anomalous behavior was evident. A modification was made to the TC lance that eliminated the anomalous behavior for the affected TCs shortly before cross-wind testing of the belowground configuration, which was the last phase of testing. The behavior of the TCs at the 3.023 m (119 in.) elevation and below was not impacted by the modification. P = 100 kPa P = 450 kPa P = 800 kPa Velocity (m/s)

P = 0.3 kPa

= 3.63E-2 kg/s = 2.64E-2 kg/s = 2.24E-2 kg/s = 2.18E-2 kg/s 4-5 Figure 4-7 Steady State Transverse Temperature Profile at z = 3.023 m (119 in.) for the Test Conducted at 5.0 kW and 800 kPa Helium Figure 4-8 Steady State Transverse Temperature Profile at z = 3.023 m (119 in.) for the Test Conducted at 0.5 kW and 0.3 kPa Air 4.1.4 Summary Data Tables The steady-state value of the peak temperature for each region of the test apparatus is presented in the following summary tables. Table 4-1 through Table 4-4 present these peak temperatures and corresponding location along with the measured power, ambient temperature, and induced air x y x y 4-6 flow rate for each power level tested at a given vessel pressure. The corresponding minimum and maximum values over the steady-state measurement period are also presented. Table 4-1 Steady State Results for the Primary Assembly Measurements at 0.3 kPa Air Table 4-2 Steady State Results for the Primary Assembly Measurements at 100 kPa Helium Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Shell 2 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4984544033623293133012972.59E-02Max0.5244554033633303143032992.73E-02 Min0.4684514003603273113002952.46E-02LocationDT_2_48Channel_4_48Basket_3_72PV_4_72S1_4_119S2_4_48AllIntegrated TotalAverage0.9965384664063523233042983.63E-02Max1.0405394664063523253073003.67E-02Min0.9565374654063513233032963.54E-02LocationDT_1_24Channel_4_48Basket_3_72PV_1_84S1_2_119S2_4_48AllIntegrated Total 1 0.5Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Shell 2 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4983743583433273102992952.64E-02Max0.5233743583433273113012962.67E-02Min0.4713733573433273102992942.61E-02LocationFV_3_72Channel_4_72Basket_3_72PV_4_72S1_4_119S2_4_48AllIntegrated TotalAverage0.9964334033783493213012953.61E-02Max1.0284334043783493213012973.65E-02 Min0.9674324033773493213002933.58E-02LocationFV_3_72Channel_3_60Basket_3_72PV_4_72S1_2_119S2_4_48AllIntegrated TotalAverage2.4945635084594033493052965.33E-02Max2.5455645084604033493062975.35E-02 Min2.4465635074594033493052955.29E-02LocationDT_2_48Channel_3_60Basket_3_72PV_3-4_72S1_2_119S2_2_48AllIntegrated TotalAverage4.9947046245564733943132966.99E-02Max5.0367046255564743953142987.04E-02 Min4.9547036245564723933122956.94E-02LocationDT_2_48Channel_3_60Basket_3_72PV_3-4_72S1_2_119S2_4_96AllIntegrated Total0.5 12.5 5 4-7 Table 4-3 Steady State Results for the Primary Assembly Measurements at 450 kPa Helium Table 4-4 Steady State Results for the Primary Assembly Measurements at 800 kPa Helium Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Shell 2 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4983663513393253092982942.24E-02Max0.5263663523393253092992972.33E-02Min0.4693653513383243092982922.14E-02LocationDT_2_119Channel_2_119Basket_4_119PV_2-3_119S1_2_119S2_4_48AllIntegrated TotalAverage0.9994203943723473203002963.21E-02Max1.0294203953723483213032973.25E-02 Min0.9674203943713473193002943.12E-02LocationDT_2_119Channel_2_119Basket_4_119PV_2-3_119S1_2_119S2_4_96AllIntegrated TotalAverage2.4945464944534023493072984.88E-02Max2.5385464954534033513093004.93E-02 Min2.4475454944524013493072964.85E-02LocationDT_1_96Channel_2_108Basket_2_108PV_2-3_119S1_2_119S2_4_96AllIntegrated TotalAverage4.9946896125474663893122966.51E-02Max5.0306896125484663903132986.57E-02 Min4.9336896125474653893112936.42E-02LocationFV_3_72Channel_4_72Basket_2_108PV_2_108S1_2_119S2_1_96AllIntegrated Total2.5 5 10.5Nominal Power (kW)Power (kW)Cladding (K)Channel (K)Basket (K)Vessel (K)Shell 1 (K)Shell 2 (K)Ambient (K)Air Flow Rate (kg/s)Average0.4983633513413303143033002.18E-02Max0.5233643513413303153053022.26E-02Min0.4683633503403293133032992.06E-02LocationFV_3_144Channel_3_144Basket_3_144PV_1_156S1_4_119S2_3_72AllIntegrated TotalAverage0.9994063843673493203012963.06E-02Max1.0384063843673493203032983.11E-02 Min0.9644053843673493193002943.01E-02LocationFV_3_144Channel_3_144Basket_3_144PV_1_156S1_1_144S2_4_96AllIntegrated TotalAverage2.4945244794434043503103004.57E-02Max2.5465254794434043513123024.62E-02 Min2.4305244794434033493092994.51E-02LocationFV_3_144Channel_3_144Basket_3_144PV_1_156S1_1_144S2_4_96AllIntegrated TotalAverage4.9946615915314653893132976.11E-02Max5.0656625925324663903163006.16E-02 Min4.8796615915304643883122966.08E-02LocationDT_2_119Channel_2_119Basket_2_108PV_2-3_119S1_2_119S2_4_96AllIntegrated Total 50.5 12.5 4-84.2 Transient Analyses Figure 4-9 and Figure 4-10 show the peak cladding temperature and total air mass flow rate for each power tested at 800 kPa absolute helium pressure. The integrated results from the air velocity profiles were converted to calculate the total air-mass flow rate in the inlet annulus. Ninety-five percent uncertainties are also presented for select data points, 1% of reading for temperature (+/-7 K maximum) and +/-1.1E-3 kg/s for mass flow rate. On average, the pressurized belowground configurations took a few hours longer to reach steady state than the corresponding aboveground configurations requiring about 17 hours1.967593e-4 days 0.00472 hours 2.810847e-5 weeks 6.4685e-6 months . Figure 4-11 shows the time required to reach steady state as a function of power for the various test pressures. The time to steady state was independent of power and helium pressures, except for the vacuum case. For the 100 kPa helium pressure tests, there was a slight dependence on power, with 13 hours1.50463e-4 days 0.00361 hours 2.149471e-5 weeks 4.9465e-6 months required at 5.0 kW and 18 hours2.083333e-4 days 0.005 hours 2.97619e-5 weeks 6.849e-6 months required at 0.5 kW. The vacuum tests were the most sensitive to power, with up to 27 hours3.125e-4 days 0.0075 hours 4.464286e-5 weeks 1.02735e-5 months required to reach steady state in the 0.5 kW

case. Figure 4-9 Peak Cladding Temperature as a Function of Time for Tests Conducted at 800 kPa Helium 4-9 Figure 4-10 Total Air Mass Flow Rate as a Function of Time for Tests Conducted at 800 kPa Helium Figure 4-11 Time to Reach Steady State as a Function of Power for the Various Vessel Pressures Tested 4.2.1 Transient Response of TC Lance and Corresponding Cladding Figure 4-12 shows the temperature of the TC lance and adjacent cladding TCs (assuming symmetry) as a function of time at the 3.023 m elevation for the 5.0 kW and 800 kPa case. Figure 4-10 4-13 shows the temperature of the TC lance and adjacent cladding TCs at the same elevation fo r the 0.5 kW and 0.3 kPa case. Ninety-five percent uncertainties are also presented for select datapoints as 1% of reading for temperature (+/-7 K maximum). The transient response of the TC la nce and the adjacent cladding TCs were similar. The temperature indicated by the lance TC was roughly midway between the adjacent clad TCs. The good agreement provided validation that th e TC lance gives an accurate indication of nearby cladding temperatures. Again, TC lance data for the 3.023 m (119 in.) location is presented because no anomalous behavior was evident at thiselevati on.Figure 4-12 Comparison of TC Lance and Cladding Temperatures at z = 3.023 m (119 in.) as a Function of Time for the Test Conducted at 5.0 kW and 800 kPa Helium 4-11 Figure 4-13 Comparison of TC Lance and Cladding Temperatures at z = 3.023 m (119 in.) as a Function of Time for the Test Conducted at 0.5 kW and 0.3 kPa Air 4.3 Cross-Wind Analyses Two types of cross-wind tests were conducted. In both types of tests, the apparatus was first allowed to reach thermal steady-state for the given test conditions and zero cross-wind. For constant cross-wind testing, the wind machine was then started and wind speed was maintained for 12 to 18 hours2.083333e-4 days 0.005 hours 2.97619e-5 weeks 6.849e-6 months . A limited number of these extended duration tests were conducted. In all cases the rise in PCT attributed to the cross-wind was small and within the experimental error of the temperature measurement. Table 4-5 shows the temperature rise attributed to the cross-wind for each of these cases. Table 4-5 Rise in Peak Cladding Temperature Attributed to Cross-Wind Conditions At the higher wind speeds, the compressor was not able to run for these extended periods. During these tests the induced air-mass flow rate obtained 95% or greater of the steady state value almost immediately. For the second type of cross-wind testing, the wind speed was changed at one hour intervals to more efficiently probe the effect of cross-wind speed on the induced air flow rate. Thermal steady-state was not reestablished. The effect of cross-wind velocity (from 0.5 to 5.4 m/s) on the induced air flow rate was measured for three powers (1.0 kW, 2.5 kW, and 5.0 kW) and three helium pressures (100 kPa, 450 kPa and 800 kPa). Figure 4-14 to Power (kW)Pressure (kPa)Cross-Wind (m/s)PCT (K) (kg/s) / o1.01001.30.22.62E-020.711.01002.70.62.06E-020.561.01005.31.72.38E-020.655.01001.41.75.79E-020.815.01002.73.74.50E-020.635.01005.35.84.02E-020.56 4-12 Figure 4-18 present the normalized air-mass flow rate as a function of cross-wind velocity for the various test cases. As the wind speed increased from zero, the normalized air-mass flow rapidly dropped to a minimum of between 0.5 to 0.6 at a cross-wind speed between 2.5 and 5.0 m/s and then slowly increased as the cross-wind speed was increased further. Error bars are included on every other data point for enhanced clarity. As the applied power increased, the error in the normalized air-mass flow rate decreased noticeably. The error did not change noticeably with helium pressure. Figure 4-14 Normalized Air-Mass Flow Rates as a Function of Cross-Wind Speed for 1.0 kW Tests 4-13 Figure 4-15 Normalized Air-Mass Flow Rates as a Function of Cross-Wind Speed for 2.5 kW Tests Figure 4-16 Normalized Air-Mass Flow Rates as a Function of Cross-Wind Speed for 5.0 kW Tests 4-14 Figure 4-17 Normalized Air-Mass Flow Rates as a Function of Cross-Wind Speed for 100 kPa Tests Figure 4-18 Normalized Air Mass Flow Rates as a Function of Cross-Wind Speed for 800 kPa Tests Figure 4-20 shows velocity contours for the induced air flow in the annulus between shell 1 and shell 2 for the 5.0 kW and 100 kPa test at various cross-wind speeds. The wind was imposed on the top, or North side, of the image as indicated by the arrows in Figure 4-19. At zero cross-wind, 4-15 the contours were not azimuthally symmetric with higher velocities in the Northeast and Southwest quadrants. The asymmetry was likely due to flow restrictions at the seam of the two halves of the honeycomb flow straightener located at the Northwest and Southeast quadrants. For a cross-wind speed of 1.3 m/s (3.0 mph), the azimuthal symmetry was improved. At a cross-wind speed of 2.7 m/s (6.0 mph), the induced air-flow velocity was enhanced on the windward side and nearly stagnant on the leeward side. The contrast between the induced air flow velocity on the windward and the leeward sides was diminished at 5.3 m/s (11.8 mph). Figure 4-19 Orientation of the Wind Machine and Test Assembly Figure 4-20 Velocity Contours for 5.0 kW and 100 kPa at Different Cross-Wind Speeds x y N S E W 1.3 m/s (3.0 mph) 2.7 m/s (6.0 mph) 5.3 m/s (11.8 mph)

Velocit y (m/s) Cross-Wind = 0 m/s = 0.072 kg/s = 0.057 kg/s = 0.045 kg/s = 0.042 kg/s

5-15

SUMMARY

A test apparatus simulating a modern dry cask was successfully constructed and operated to produce first-of-a-kind, high-fidelity transient and steady-state thermal-hydraulic data sets suitable for CFD model validation. An existing electrically heated but otherwise prototypic BWR Incoloy-clad test assembly was deployed inside of a representative storage basket and cylindrical pressure vessel that represented the canister. Simulated decay power was scaled to mimic the desired range of prototypic dimensionless groups. One unique aspect of the test apparatus was the capability to pressurize the simulated canister to a wide range of pressures, from sub-atmospheric (0.3 kPa) to the upper range of prototypic values (800 kPa). Test configurations for both vertical aboveground and belowground storage cask systems were tested. A wind machine was used to test the effect of wind speed on the peak cladding temperature and induced air mass flow rate in the belowground configuration. Cladding temperatures were measured with 0.762 mm (0.030 in.) diameter Type K thermocouples installed in direct contact with the Incoloy heater cladding. The induced air-mass flow rate was determined by integrating velocity profiles measured with hot wire anemometers that impose negligible pressure drop. A total of fourteen tests were conducted with the apparatus in the aboveground configuration. Similarly, fourteen tests were conducted with the apparatus in the belowground configuration. For these twenty-eight tests, the assembly was operated from initial, ambient conditions to thermal-hydraulic steady state for each unique combination of assembly power and vessel pressure. The power levels tested were 0.5, 1.0, 2.5, and 5.0 kW. The vessel pressures tested were vacuum (0.3 kPa), 100, 450, and 800 kPa absolute. A previous scaling analysis showed that elevated powers up to 5.0 kW were warranted to drive the induced air flow to prototypic levels. Over thirteen tests were conducted with the wind machine and the apparatus in the belowground configuration. The effect of cross-wind velocity (from 0.5 to 5.4 m/s) on the induced air mass flow rate was measured for three powers (1.0 kW, 2.5 kW, and 5.0 kW) and three helium pressures (100 kPa, 450 kPa, and 800 kPa). The performance of the aboveground and belowground storage cask configurations were relatively similar, as expected. All steady state peak temperatures and induced air mass flow rates increased with increasing assembly power. Peak cladding temperatures decreased with increasing internal helium pressure for a given assembly power, indicating increased internal convection. In addition, the location of the PCT moved from near the top of the assembly to ~1/3 the height of the assembly for the highest (800 kPa absolute) to the lowest (0.3 kPa absolute) pressure studied, respectively. This shift in PCT location is consistent with convective heat transfer increasing with internal helium pressure. The highest average steady state PCT achieved was 715 K for 5.0 kW and 100 kPa helium pressure. This temperature was in the range of the NRC limits for allowable PCT of 673 K for normal operation and 843 K for off-normal operation [US NRC, 2003]. For the cross-wind test series, as the wind speed increased from zero, the normalized air mass flow rate rapidly dropped to a minimum of between 0.5 to 0.6 at a cross-wind speed between 2.5 and 5.0 m/s and then slowly increased as the cross-wind speed increased further. Over 40 unique data sets were collected and analyzed for these efforts. The results documented in this report highlight a small, but representative, subset of the available data. This addition to the experimental database signifies a substantial addition of first-of-a-kind, high-fidelity transient and steady-state thermal-hydraulic data sets suitable for CFD model validation.

6-16 REFERENCES

[1]ANSI, American National Standards Institute, "American National Standard for Radioactive Materials - Leakage Tests on Packages for Shipment," ANSI N14.5-20 14, June 2 014.[2]Bates, J.M., "Single PWR Spent Fuel Assembly Heat Transfer Data for Computer Code Evaluations," Pacific Northwest Laboratory, Richland, Washington, PNL-5571, Janu ary 1986.[3]Creer, J.M., T.E. Michener, M.A. McKinnon, J.E. Tanner, E.R. Gilbert, R.L. Goodma n,"The TN-24P PWR Spent Fuel Storage Cask: Testing and Analyses", EPRI NP-5 128 Proj. 2406-4, PNL-6054, Pacific Northwest Laboratory, Richland, Washington, Apri l 1987.[4]Durbin, S.G., E.R. Lindgren, A. Zigh, and J. Solis, "Description of Dry Cask Simulato r for Measuring Internal and External Thermal-Hydraulic Performance," SAND2016-0176C, Trans. Am. Nucl. Soc., New Orleans, LA, June 2016.

[5]Dziadosz, D., E.V. Moore, J.M. Creer, R.A. McCann, M.A. McKinnon, J.E. Tanner, E.

R.Gilbert, R.L. Goodman, D.H. Schoonen, M Jensen, and C. Mullen, "The Castor-V/2 1 PWR Spent-Fuel Storage Cask: Testing and Analyses," Electrical Power ResearchInstitute, EPRI NP-4887, Project 2406-4, PNL-5917, Pacific Northwest Laboratory,Richland, Washington, November 1986.

[6]EPRI, Electric Power Research Institute, "High Burnup Dry Storage Cask Research andDevelopment Project: Final Test Plan," Contract No.: DE-NE-0000593, February 2014.

[7]Irino, M., M. Oohashi, T. Irie, and T. Nishikawa, "Study on Surface Temperatures ofFuel Pins in Spent Fuel Dry Shipping/Storage Casks," IAEA-SM-286/139P, in Proceedings of Packaging and Transportation of Radioactive Materials (PATRAM '86)

, Volume 2, p. 585, International Atomic Energy Agency Vienna, 1987.

[8]Lindgren, E.R. and S.G. Durbin, "Characterization of Thermal-Hydraulic and IgnitionPhenomena in Prototypic, Full-Length Boiling Water Reactor Spent Fuel Po ol Assemblies after a Complete Loss-of-Coolant Accident", SAND2007-2270, SandiaNational Laboratories, Albuquerque, New Mexico, April 2007.

[9]McKinnon, M.A., J.W. Doman, J.E. Tanner, R.J. Guenther, J.M. Creer and C.E. Ki ng,"BWR Spent Fuel Storage Cask Performance Test, Volume 1, Cask Hand ling Experience and Decay Heat, Heat Transfer, and Shielding Data," PNL-5777 Vol. 1,Pacific Northwest Laboratory, Richland Washington, February 1986.

[10]McKinnon, M.A., J.M. Creer, C. L. Wheeler , J.E. Tanner, E.R. Gilbert, R.L. Goodma n, D.P. Batala, D.A. Dziadosz, E.V. Moore, D.H. Schoonen, M.F. Jensen, and J.H.Browder, "The MC-10 PWR Spent Fuel Storage Cask: Testing and Anal ysis," EPRI NP-5268, PNL-6139, Pacific Northwest Laboratory, Richland, Washington, July 1987.

[11]McKinnon, M.A., TE Michener, M.F. Jensen, G.R. Rodman, "Testing and Analyses ofthe TN-24P Spent Fuel Dry Storage Cask Loaded with Consolidated Fuel", EPRI NP-6191 Project 2813-16, PNL-6631, Pacific Northwest Laboratory, Richland, Washington,February 1989.

[12]McKinnon, M.A., R.E. Dodge, R.C. Schmitt, L.E. Eslinger, & G. Dineen,, "Performanc e Testing and Analyses of the VSC-17 Ventilated Concrete Cask", EPRI-TR-100305,Electric Power Research Institute, Palo Alto, California, May 1992.

6-2[13]Nakos, J.T., "Uncertainty Analysis of Thermocouple Measurements Used in Normal an d Abnormal Thermal Environment Experiments at Sandia's Radiant Heat Facility and Lurance Canyon Burn Site," SAND2004-1023, Sandia National Laboratorie s, Albuquerque, New Mexico, April 2004.

[14]US NRC, "Cladding Considerations for the Transportation and Storage of Spent Fuel

", Interim Staff Guidance-11 Rev. 3 (2003).

[15]Zigh, A., S. Gonzalez, J. Solis, S.G. Durbin, and E.R. Lindgren, "Validation of the Computational Fluid Dynamics Method using the Aboveground Configuration of th e Dry Cask Simulator," SAND2017-6104C, Trans. Am. Nucl. Soc., San Francisco, CA, June 2017.

A-1 APPENDIX A ERROR ANALYSIS The uncertainty and error inherent to an experimental result are critical to the accurate interpretation of the data. Therefore, the uncertainties in the experimental measurements are estimated in this section. Results of this analysis are given, followed by a general description of the method used and a brief explanation of the source of each reported measurement uncertainty. The overall standard uncertainty of an indirect measurement y, dependent on N indirect measurements x i, is defined in Equation A-1. The standard uncertainty associated with an indirect measurement is analogous to the standard deviation of a statistical population. N i i i u x y u 1 2 2 A-1 Here, u is used to define the standard uncertainty of a measurement. The expanded uncertainty, U, is reported in this appendix and defines the bounds that include 95% of the possible data. The expanded uncertainty is assumed to be defined as the product of the standard uncertainty and the Student's t-value. Unless otherwise stated, all uncertainty measurements are assumed to be based on a Student's t-distribution with no fewer than 30 measurements. The associated t-value for 95% intervals is 2.0 for 29 degrees of freedom. Therefore, Equation A-2 shows the definition of the expanded uncertainty as used in the following sections for a 95% confidence interval.

U = t value u A-2 Table A-1 summarizes the expanded uncertainty for each measurement used in this report. Table A-1 Summary of the Expanded Uncertainty Determined for each Measurement Measurement, xUnitsExpanded Uncertainty, U xPeak clad temperature K7.0E+00Ambient temperature K3.0E+00Ambient pressurekPa, abs1.1E-01Helium pressurekPa, abs1.0E+00VacuumkPa, abs3.0E-01Voltage V3.8E-01Current A3.8E-01PowerkW7.5E-02Forced air mass flow ratekg/s5.9E-04Induced air mass flow rate (aboveground)kg/s1.5E-03Induced air mass flow rate (belowground)kg/s1.1E-03Induced air mass flow rate (cross-wind)kg/s1.3E-03Normalized air mass flow rate, /o-5.6 E-0 2Cross-wind speedm/s4.9E-02 A-2A.1 Temperature MeasurementsA.1.1 Uncertainty in Clad Temperature Measureme nt Clad temperature was measured with a standard k-type TC. The expanded uncertainty for this type of TC is U T = 1% of the reading in Kelvin [Nakos, 2004]. The maximum peak clad temperature reading was 716 K for the aboveground 5.0 kW, 100 kPa helium test. The maximum expanded uncertainty for the cladding temperature is U PCT = +/-7.0 K. A.1.2 Uncertainty in Ambient Air TemperatureThe air temperature was measured with a standard k-type TC. The expanded uncertainty for this type of TC is U T = 1% of the reading in Kelvin [Nakos, 2004]. The maximum ambient temperature reading was 305 K for the aboveground 5.0 kW, 100 kPa helium test. The maximum expanded uncertainty for the ambient temperature is UT-amb = +/-3.0 K. A.2 Pressure MeasurementsA.2.1 Uncertainty in Ambient Air Press ure The air pressure was measured with a Setra Systems barometer (Model 276). The uncertainty of the ambient air pressure was taken from the manufacturer's calibration sheet, which indicated an expanded uncertainty in the instrument of +/-0.1% of full scale (110 kPa). Therefore, the expanded uncertainty in the pressure reading is UP-atm = +/-0.11 kPa. A.2.2 Uncertainty in Helium Vessel PressureThe helium pressure was measured using an Omega model PX409-500A5V-XL, 0 to 3447 kPa (500 psia), pressure transducer. The resolution of the transducer allowed the pressure control system described in Section 2.3.2.1 to maintain the pressure constant to +/-0.3 kPa (0.044 psi).

However, with the "-XL" accuracy identifier the linearity deviates +/-0.03% from the best straight line, which at full scale is +/-1.0 kPa (+/-0.15 psi). Therefore, the expanded uncertainty is UP-He = +/-1.0 kPa. A.2.3 Uncertainty in Air Vessel PressureThe residual air pressure was measured using an Omega model PXM409-001BV10V, 0 to 100 kPa absolute (0 to 14.5 psia), pressure transducer. The linearity deviates +/-0.08% from the best straight line, which at full scale is +/-0.08 kPa (+/-0.012 psi). However, the span and zero shift for temperature compensation are each +/-0.5%, which for full scale is +/-0.5 kPa (+/-0.073 psi). The geometric mean of these three expanded uncertainties is +/-0.3%, or +/-0.3 kPa (+/-0.044 psi). This value of 0.3 kPa absolute was assumed to be the smallest determinable pressure under vacuum conditions. Therefore, all vacuum tests are reported as 0.3 kPa, even though the gage typically read less than this value. A.3 Uncertainty in Electrical MeasurementsThe voltage, current, and power supplied to the internal spent fuel assembly heater rods were measured by Ohio Semitronics, Inc. instrumentation. The voltage was monitored by a model AVTR-001D voltmeter with an expanded uncertainty of UVolt = +/-0.38 V. The current was monitored by a model ACTR-005DY06 current meter with an expanded uncertainty of UAmp =

A-3+/-0.38 A. The power was monitored with a model PC5-001DY230 Watt meter with an expanded uncertainty of UWatt = +/-0.075 kW. A.4 Flow MeasurementsThe methodology for determining the induced air flow in the aboveground and belowground configurations was different. As described in detail in Section 2.4.2 for the aboveground configuration, correlation of the hot wires in the inlet ducts was performed by imposing a known mass flow rate of air through the ducting with the hot wires held in a fixed location and then implementing a small correction based on velocity profile measurement and integrating to a total mass flow for the buoyancy driven flows. For the belowground configuration described in detail in Section 2.4.3, a forced flow correlation in the annulus between Shell 1 and Shell 2 was not possible, so the mass flow was determined by integrating eight velocity profiles (twelve for cases with wind). A.4.1 Aboveground ConfigurationA.4.1.1 Uncertainty in Air Mass Flow ControllersThe air flow was controlled using an OMEGA FMA-2623A, 0 to 3000 slpm (or 5.92E-2 kg/s at the standard conditions of 25 °C and 101.4 kPa), mass flow controller. The maximum expanded uncertainty is +/-1.0% of full scale at full flow or +/-5.9E-4 kg/s. A.4.1.2 Uncertainty in Hot Wire Anemometer MeasurementsThe parameter values needed to determine the induced air flow from the hot wire measurements are listed in Table A-2 and Table A-3 along with the parameter's expanded uncertainty, influence coefficient, and contribution to the error. V TSI is the voltage output of the TSI Model 8455 hot wire anemometer. The expanded uncertainty is given by the manufacturer as +/-0.025 m/s for the ambient temperatures encountered. The full-scale voltage output is 10 V, so the expanded error in the voltage output is +/-0.25 V. Standard conditions for the TSI hotwire are 21.1 °C and 101.4 kPa. The primary calibration of the hot wires was performed by metering a measured flow of air with the hot wire centered in the duct at the position indicated in Figure 2-19. Figure 2-20 shows the forced flow calibration curve for the TSI Model 8455 hot wire located in a fixed position in the center of an inlet duct as shown in Figure 2-21, along with the equation for the best linear through the data. The constant linear fit coefficient, a TSI,0, is -8.0E-04 kg/s, with an expanded error of 9.0E-05 kg/s based on the fit of the linear correlation. The first order linear fit coefficient, aTSI,1, is 2.8E-03 kg/s/V, with an expanded uncertainty of 1.8E-05 kg/s/V. An additional correlation was needed to relate the naturally induced flow to the metered forced flow. After each powered test during steady state, the hot wire was traversed across the narrow dimension of the duct, as shown in Figure 2.21, to generate a velocity profile. The profile was integrated across the area of the duct to calculate the total naturally induce flow. Figure 2-23 shows the correlation between the more direct measurements of the naturally induced flow-based on the velocity profile measurement made only at the end of the test and the less direct measurement based on the forced flow correlation with the hot wire in the fixed location maintained throughout the ~24 hour transient to steady-state. The correlation coefficient, C corr, is 0.9344, with an expanded uncertainty of 1.3E-2 based on a t-value of 2.2 for the 12 data points used to define the correlation. The mass flow in each duct is determined with an expanded error of +/-7.4E-04 kg/s. The error in the hot wire air velocity measurement contributed 80% of the error, followed by the natural-flow to forced-flow correlation, which contributed 15% of the error.

A-4 Table A-2 Parameters Values and Uncertainty Analysis for a Single Hotwire Measurement in the Aboveground Configuration Table A-3 outlines the calculation of the total mass flow from the four ducts. The expanded error in the total air mass flow of U = +/-1.5e-03 kg/s. Table A-3 Uncertainty Analysis for Combining Multiple Hotwire Measurements into a Total Induced Flow Rate in the Aboveground Configuration A.4.2 Belowground Configuration (Annular Gap)The details for the determination of the total induced air mass flow rate in the belowground configuration are given in Section 2.4.3. In the belowground configuration, a forced-flow correlation in the annulus between Shell 1 and Shell 2 was not possible, so the mass flow was determined by integrating eight velocity profiles. Separate verification tests were conducted to determine the accuracy of deriving the air mass flow rate from velocity measurements and integration as discussed in Appendix D The temperature of the air flow in the annular gap was up to 41°C, which raises the expanded error of the measurement to +/-0.051 m/s. This value of +/-0.051 m/s includes the standard instrument uncertainty of +/-0.025 m/s (2.5% of full scale) and +/-0.026 m/s (0.2% of full scale per °C above 28 °C). However, the velocity gradient between the different profiles at the same radial location introduces an uncertainty greater than the instrument uncertainty. This uncertainty may be conceptualized as the potential error introduced by using a centrally measured velocity to calculate the mass flow rate across a small but finite area. This gradient-based uncertainty was estimated for all hot wires for three different test conditions (1 kW and 100 kPa; 2.5 kW and 450 kPa; 5 kW and 800 kPa). The root mean square of all gradient-based uncertainties was found to

be U V = +/-0.085 m/s, which exceeds the instrument uncertainty. For the purposes of this uncertainty analysis and the cross-wind uncertainty analysis to follow, this value of +/-0.085 m/s is adopted. Hotwire air-velocity measurements were made at fourteen equidistant locations across the annular gap. The integration process involves calculation of an associated flow area for each velocity measurement. Table A-4 presents the pertinent inputs for the calculation along with the expanded uncertainty, influence coefficient, and contribution. The expanded uncertainty in the Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(/x i)/])Contribution V TSIV8.0E+002.5E-013.2E-020.80 aTSI, 0kg/s-8.0E-049.0E-054.1E-030.01 aTSI, 1(kg/s)/V2.8E-031.8E-056.7E-030.03 Ccor r--9.3E-011.3E-021.4E-020.15kg/s2.0E-027.4E-043.6E-021.00Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(/x i)/])Contribution1kg/s2.0E-027.4E-049.0E-030.252kg/s2.0E-027.4E-049.0E-030.253kg/s2.0E-027.4E-049.0E-030.254kg/s2.0E-027.4E-049.0E-030.25kg/s8.2E-021.5E-031.8E-021.00 A-5flow area for each air velocity measurement is +/-2.4E-05 m

2. Table A-5 presents a representative integration calculation to determine the mass flow and expanded uncertainty for one of the eight hotwires. Table A-4 Representative Calculation to Estimate the Expanded Error of Flow Area Determination Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(A/x i)/A])Contributionrm3.1E-016.4E-032.0E-021.00rm4.8E-035.0E-065.2E-040.00/2--1.3E-01------A m 21.2E-032.4E-052.0E-021.00 A-6Table A-6 presents the calculation of the total air mass flow and expanded uncertainty based on all eight hotwires. The expanded error for the total air mass flow determination in the belowground configuration is U = +/-1.1E-03 kg/s. Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(i/x i)/i])Contribution vi,1m/s3.1E-018.5E-021.1E-02 0.06Ai,1 m 27.8E-032.4E-051.3E-04 0.00 vi,2m/s4.8E-018.5E-021.3E-02 0.09Ai,2 m 29.1E-032.4E-052.0E-04 0.00 vi,3m/s6.1E-018.5E-021.3E-02 0.09Ai,3 m 29.0E-032.4E-052.5E-04 0.00 vi,4m/s6.0E-018.5E-021.3E-02 0.08Ai,4 m 28.9E-032.4E-052.5E-04 0.00 vi,5m/s6.4E-018.5E-021.3E-02 0.08Ai,5 m 28.7E-032.4E-052.6E-04 0.00 vi,6m/s6.1E-018.5E-021.3E-02 0.08Ai,6 m 28.6E-032.4E-052.5E-04 0.00 vi,7m/s6.0E-018.5E-021.2E-02 0.08Ai,7 m 28.4E-032.4E-052.5E-04 0.00 vi,8m/s5.7E-018.5E-021.2E-02 0.07Ai,8 m 28.3E-032.4E-052.4E-04 0.00 vi,9m/s5.5E-018.5E-021.2E-02 0.07Ai,9 m 28.1E-032.4E-052.3E-04 0.00 vi,10m/s5.2E-018.5E-021.2E-02 0.07Ai,10 m 28.0E-032.4E-052.1E-04 0.00 vi,11m/s4.8E-018.5E-021.2E-02 0.07Ai,11 m 27.8E-032.4E-052.0E-04 0.00 vi,12m/s4.0E-018.5E-021.1E-02 0.06Ai,12 m 27.7E-032.4E-051.6E-04 0.00 vi,13m/s3.6E-018.5E-021.1E-02 0.06Ai,13 m 27.6E-032.4E-051.5E-04 0.00 vi,14m/s2.5E-018.5E-028.9E-03 0.04Ai,14 m 26.1E-032.4E-051.0E-04 0.00Refkg/m 31.2E+00------ikg/s8.7E-033.9E-044.5E-02 1.00Table A-5 Representative Integration Calculation to Determine the Mass Flow and Expanded Error for One of the Eight Hotwires A-7 Table A-6 Calculation of the Total Mass Flow and Expanded Error from the Eight Hotwires used in the Belowground Configuration A.4.3 Cross-Wind ConfigurationThe determination of the total mass flow of air for the belowground configuration with cross-wind was similar to the belowground configuration except twelve hot wires were used as described in detail in Section 2.5. Table A-4 and Table A-5 are applicable. Table A-7 shows the calculation using twelve hotwires. Using the twelve hotwires the expanded error for the total air mass flow determination in the belowground configuration is U = +/-1.3E-03 kg/s. Table A-7 Calculation of the Total Mass Flow and Expanded Error from the Twelve Hotwires used in the Cross-Wind Configuration The effect of cross-wind was evaluated using a normalized flow variable, /o, defined as the air mass flow with wind divided by the mass flow without wind under the same conditions. The expanded uncertainties for /o are presented in Table A-8 for various test conditions. Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(/x i)/])Contribution1kg/s8.7E-033.9E-045.6E-030.122kg/s1.1E-025.2E-047.4E-030.213kg/s8.8E-033.9E-045.6E-030.124kg/s7.5E-033.4E-044.8E-030.095kg/s9.6E-034.3E-046.1E-030.146kg/s9.6E-034.3E-046.1E-030.147kg/s9.0E-034.1E-045.8E-030.138kg/s5.5E-032.5E-043.5E-030.05kg/s7.0E-021.1E-031.6E-021.00Measurement, x iUnitsValueExpanded uncertainty, U iInfluence coefficient (U i*[(/x i)/])Contribution1kg/s6.8E-033.9E-045.4E-030.102kg/s5.6E-033.2E-044.5E-030.073kg/s5.8E-033.4E-044.7E-030.074kg/s4.7E-032.7E-043.8E-030.055kg/s4.4E-032.6E-043.6E-030.046kg/s4.5E-032.6E-043.6E-030.047kg/s3.8E-032.2E-043.1E-030.038kg/s4.2E-032.4E-043.3E-030.049kg/s7.2E-034.1E-045.8E-030.1110kg/s9.8E-035.6E-047.8E-030.2011kg/s9.3E-035.4E-047.5E-030.1912kg/s5.6E-033.2E-044.5E-030.07kg/s7.2E-021.3E-031.7E-021.00 A-8 Table A-8 Expanded Uncertainties in Normalized Mass Flow, /o, for Various Conditions Tested A.4.3.1 Cross-Wind VelocityThe area-weighted average cross-wind velocity was determined using the same type TSI Model 8455 hot wire anemometers fixed at three locations shown in Figure 2-29. As discussed in Section 2.5, the average of the three fixed hotwires was correlated with the area weighted average of 45 regularly spaced points. The standard error about the best straight line was

+/-0.0113 m/s. Using the t-value of 4.3 for the three data-point correlation, the expanded error for the area weighted cross-wind velocity is Uwind = +/-0.049 m/s. ConditionsExpanded uncertainty, U i5 kW, 100 kPa2.5E-025 kW, 800 kPa2.8E-022.5 kW, 100 kPa3.3E-022.5 kW, 800 kPa3.8E-021.0 kW, 100 kPa4.8E-021.0 kW, 800 kPa5.6E-02 B-1 APPENDIX B CHANNEL LIST FROM ABOVEGROUND TESTING The results presented in the body of the test report describe the most important quantities as determined by the authors. This presentation represents a fraction of the information collected from the test assembly. Table B-1 gives the complete channel list for the aboveground configuration as an example to the reader of the extent of the available data. Table B-1 Channel List for Aboveground Configuration Testing SlotChannelTC#InstrumentNomenclatureInstrumentTypeSlotChannelTC#InstrumentNomenclatureInstrumentType101WDVINType"K"TC2033FV72_3Type"K"TC112WDVOUTType"K"TC2134FV144_3Type"K"TC123WFTINType"K"TC2235CS6_1AType"K"TC134WFTOUTType"K"TC2336CS12_1AType"K"TC145WEU24Type"K"TC2437CS18_1AType"K"TC156WEU48Type"K"TC2538CS24_1Type"K"TC167WEU72Type"K"TC2639CS30_1AType"K"TC178WEU96Type"K"TC2740CS36_1AType"K"TC189No_DataType"K"TC2841CS42_2AType"K"TC1910WEU144Type"K"TC2942CS48_2Type"K"TC11011WDV24_1Type"K"TC21043CS54_2AType"K"TC11112WDV96_1Type"K"TC21144CS61_2AType"K"TC11213WFT48_2AType"K"TC21245CS90_1AType"K"TC11314WFT72_3AType"K"TC21346CS96_1Type"K"TC11415WFT119_2AType"K"TC21447CS103_1AType"K"TC11516WFT144_3AType"K"TC21548CS108_1AType"K"TC11617DT24_1Type"K"TC21649CS114_2AType"K"TC11718DT48_2Type"K"TC21750CS119_2Type"K"TC11819DT96_1Type"K"TC21851CS126_2AType"K"TC11920DT119_2Type"K"TC21952CS132_2AType"K"TC12021CU24_1Type"K"TC22053No_DataType"K"TC12122CU96_1Type"K"TC22154GX72_3Type"K"TC12223ES48_2Type"K"TC22255GX78_3AType"K"TC12324ES119_2Type"K"TC22356GX84_3AType"K"TC12425CX24_1Type"K"TC22457GX138_3AType"K"TC12526CX96_1Type"K"TC22558GX144_3Type"K"TC12627GS48_2Type"K"TC22659GX150_3AType"K"TC12728GS72_3Type"K"TC22760GX156_3AType"K"TC12829GS119_2Type"K"TC22861AQ24_1Type"K"TC12930GS144_3Type"K"TC22962AQ48_2Type"K"TC13031GU72_3Type"K"TC23063AQ96_1Type"K"TC13132GU144_3Type"K"TC23164AQ119_2Type"K"TC B-2SlotChannelTC#InstrumentNomenclatureInstrumentTypeSlotChannelTC#InstrumentNomenclatureInstrumentType3065AS24_1Type"K"TC50129g96_CB_2.9_1Type"K"TC3166AS96_1Type"K"TC51130g96_CB_2.9_1SType"K"TC3267No_DataType"K"TC52131g144_CB_2.9_1Type"K"TC3368No_DataType"K"TC53132g144_CB_2.9_1SType"K"TC3469No_DataType"K"TC54133g144_CB_4.0_34Type"K"TC3570AU96_1Type"K"TC55134g144_CB_2.9_3Type"K"TC3671AU108_1Type"K"TC56135g144_CB_2.9_3SType"K"TC3772No_DataType"K"TC57136Basket_Int_12_1Type"K"TC3873AX96_1Type"K"TC58137Basket_(5.5)_4Type"K"TC3974AZ24_1Type"K"TC59138Basket_0_4Type"K"TC31075AZ96_1Type"K"TC510139Basket_12_1Type"K"TC31176CQ48_2Type"K"TC511140Basket_24_1Type"K"TC31277CQ119_2Type"K"TC512141Basket_24_4Type"K"TC31378EQ48_2Type"K"TC513142Basket_24_41Type"K"TC31479EQ60_2Type"K"TC514143Basket_36_2Type"K"TC31580EQ119_2Type"K"TC515144Basket_48_2Type"K"TC31681EQ132_2Type"K"TC516145Basket_48_4Type"K"TC31782GQ48_2Type"K"TC517146Basket_60_3Type"K"TC31883GQ119_2Type"K"TC518147Basket_72_3Type"K"TC 31984IQ48_2Type"K"TC519148Basket_72_4Type"K"TC32085IQ72_3Type"K"TC520149Basket_72_3 4T y p e"K"TC32186IQ119_2Type"K"TC521150Basket_84_1Type"K"TC32287IQ144_3Type"K"TC522151Basket_96_1Type"K"TC32388IS72_3Type"K"TC523152Basket_96_4Type"K"TC32489IS144_3Type"K"TC524153Basket_108_2Type"K"TC32590IU72_3Type"K"TC525154Basket_119_2Type"K"TC32691IU84_3Type"K"TC526155Basket_119_4Type"K"TC32792IU144_3Type"K"TC527156Basket_119_2 3T y p e"K"TC32893IU156_3Type"K"TC528157Basket_132_3Type"K"TC32994IX72_3Type"K"TC529158Basket_144_3Type"K"TC33095IX144_3Type"K"TC530159Basket_144_4Type"K"TC33196IZ72_3Type"K"TC531160Basket_156_1Type"K"TCSlotChannelTC#InstrumentNomenclatureInstrumentTypeSlotChannelTC#InstrumentNomenclatureInstrumentType4097IZ144_3Type"K"TC60161Basket_159_4Type"K"TC4198Instr_Well_LeadsType"K"TC61162Basket_165_41Type"K"TC4299Instr_Well_IntType"K"TC62163Basket_Int_156_1Type"K"TC43100Pedestal_BaseType"K"TC63164g(7.6)_BV_3.5_2Type"K"TC44101Pedestal_(5.5)_4Type"K"TC64165g48_BV_4.3_4Type"K"TC45102Channel_0_4Type"K"TC65166g48_BV_4.8_34Type"K"TC 46103Channel_12_1Type"K"TC66167g72_BV_4.3_2Type"K"TC47104Channel_24_1Type"K"TC67168g96_BV_4.8_4 1T y p e"K"TC48105Channel_24_4Type"K"TC68169g96_BV_3.8_1Type"K"TC49106Channel_24_4 1T y p e"K"TC69170g96_BV_4.3_1Type"K"TC410107Channel_36_2Type"K"TC610171g96_BV_4.8_1Type"K"TC411108Channel_48_2Type"K"TC611172g144_BV_4.3_1Type"K"TC412109Channel_48_4Type"K"TC612173g144_BV_4.3_1SType"K"TC413110Channel_60_3Type"K"TC613174g144_BV_4.8_3 4T y p e"K"TC414111Channel_72_3Type"K"TC614175g144_B V_3.8_3Type"K"TC415112Channel_72_4Type"K"TC615176g144_BV_4.3_3Type"K"TC416113Channel_72_3 4T y p e"K"TC616177g144_BV_4.8_3Type"K"TC417114Channel_84_1Type"K"TC617178g167_BV_3.5_3Type"K"TC418115Channel_96_1Type"K"TC618179g167_BV_3.5_1SType"K"TC419116Channel_96_4Type"K"TC619180PV_Int_12_1Type"K"TC420117Channel_108_2Type"K"TC620181PV_0_4Type"K"TC421118Channel_119_2Type"K"TC621182PV_12_1Type"K"TC422119Channel_119_4Type"K"TC622183PV_24_1Type"K"TC423120Channel_119_2 3T y p e"K"TC623184PV_24_4Type"K"TC424121Channel_132_3Type"K"TC624185PV_24_4 1T y p e"K"TC425122Channel_144_3Type"K"TC625186PV_36_2Type"K"TC426123Channel_144_4Type"K"TC626187PV_48_2Type"K"TC427124Channel_156_1Type"K"TC627188PV_48_4Type"K"TC428125Channel_159_4Type"K"TC628189PV_60_3Type"K"TC429126g48_CB_2.9_4Type"K"TC629190PV_72_3Type"K"TC430127g72_CB_2.9_2Type"K"TC630191PV_72_4Type"K"TC431128g96_CB_4.0_4 1T y p e"K"TC631192PV_72_3 4T y p e"K"TC B-3SlotChannelTC#InstrumentNomenclatureInstrumentTypeSlotChannelTC#InstrumentNomenclatureInstrumentType70193PV_84_1Type"K"TC90257g96_S1S2_10.8_4Type"K"TC71194PV_96_1Type"K"TC91258g144_S1S2_10.8_34SType"K"TC72195PV_96_4Type"K"TC92259g144_S1S2_10.8_3Type"K"TC73196PV_108_2Type"K"TC93260S2_0_4Type"K"TC74197PV_119_2Type"K"TC94261S2_12_1Type"K"TC75198PV_119_3Type"K"TC95262S2_24_1 4T y p e"K"TC76199PV_119_4Type"K"TC96263S2_24_1Type"K"TC77200PV_119_2 3T y p e"K"TC97264S2_24_4Type"K"TC78201PV_132_3Type"K"TC98265S2_36_2Type"K"TC79202PV_144_1Type"K"TC99266S2_48_2Type"K"TC710203PV_144_3Type"K"TC910267S2_48_4Type"K"TC711204PV_144_4Type"K"TC911268S2_60_3Type"K"TC712205PV_156_1Type"K"TC912269S2_72_3 4T y p e"K"TC713206PV_159_4Type"K"TC913270S2_72_3Type"K"TC714207PV_165_4Type"K"TC914271S2_72_4Type"K"TC715208PV_Int_156_1Type"K"TC915272S2_84_1Type"K"TC716209g48_VS1_5.6_4Type"K"TC916273S2_96_1Type"K"TC717210g48_VS1_6.4_4Type"K"TC917274S2_96_4Type"K"TC718211g48_VS1_7.2_4Type"K"TC918275S2_108_2Type"K"TC 719212g48_VS1_8.1_4Type"K"TC919276S2_119_2 3T y p e"K"TC720213g48_VS1_7.2_3 4T y p e"K"TC920277S2_119_2Type"K"TC721214g96_VS1_5.6_1Type"K"TC921278S2_119_3Type"K"TC722215g96_VS1_6.4_1SType"K"TC922279S2_119_4Type"K"TC723216g96_VS1_7.2_1Type"K"TC923280S2_132_3Type"K"TC724217g96_VS1_8.1_1SType"K"TC924281S2_144_1Type"K"TC725218g96_VS1_7.2_4 1T y p e"K"TC925282S2_144_3Type"K"TC726219g96_VS1_7.2_4Type"K"TC926283S2_144_4Type"K"TC727220g144_VS1_7.2_34Type"K"TC927284Lance_108Type"K"TC728221g144_VS1_7.2_3Type"K"TC928285Lance_114Type"K"TC729222S1_0_4Type"K"TC929286Lance_119Type"K"TC730223S1_12_1Type"K"TC930287Lance_126Type"K"TC731224S1_24_1 4T y p e"K"TC931288Lance_132Type"K"TCSlotChannelTC#InstrumentNomenclatureInstrumentTypeSlotChannelTC#InstrumentNomenclatureInstrumentType80225S1_24_1Type"K"TC100289Lance_138Type"K"TC81226S1_24_4Type"K"TC101290Lance_144Type"K"TC82227S1_36_2Type"K"TC102291Lance_150Type"K"TC83228S1_48_2Type"K"TC103292Lance_156Type"K"TC84229S1_48_4Type"K"TC104293S1_96_1_InsType"K"TC85230S1_60_3Type"K"TC105294S1_96_4_InsType"K"TC86231S1_72_3 4T y p 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C-1 APPENDIX C DIMENSIONAL ANALYSES C.1 ProcedureThe dimensional analyses were conducted in two parts, one that considers helium flow internal to the pressure vessel and another that considers the external air flow (see Figure 2-1). For the internal analysis, the modified Rayleigh number (Ra*H) based on the channel height (H) is defined in Equation C-1, where g is acceleration due to gravity, is the thermal expansion coefficient, q is the uniform surface heat flux, is the thermal diffusivity, is the kinematic viscosity and k is the thermal conductivity. A simple correlation for the Nusselt number (Nu H) in a channel with uniform heating on one side and equivalent, uniform cooling on the other side is given in Equation C-2 [Bejan, 1995]. In these equations, the channel height is given as H and the hydraulic diameter of the helium downcomer is listed as DH, Down. The modified Rayleigh was chosen for these analyses because for these pre-test calculations the heat flux was easily estimable, but the temperature difference between the heated surfaces and the gas was not available.

kH"qg Ra*H 4C-19192340Down,H*H H D HRa.Nu C-2C.2 ResultsC.2.1 Internal AnalysisThe results of the internal analysis for the aboveground DCS at low and high power and the aboveground prototypic cask are presented in Table C-1. Again, this internal analysis relates to the helium flow and heat transfer inside the spent fuel and the downcomer in the pressure vessel (i.e. canister). The average helium-mass flow rate and velocity, Reynolds number, modified Rayleigh number, and the Nusselt number for the prototypic cask compare favorably with the DCS operated at low power.

C-2 Table C-1 Comparison of Internal Dimensionless Groups for the DCS and Dry Cask Systems with Helium at 700 kPa Parameter Aboveground DCS DCS Cask Power (W) 500 5,000 36,900 He (kg/s) 1.3E-3 1.8E-3 2.1E-2 DH, Down (m) 0.053 0.053 0.14 Wavg (m/s)0.061 0.126 0.078 ReDown 170 190 250

H R a 3.1E11 5.9E11 4.6E11 H N u 200 230 200 C.2.2 External AnalysisFor the external analysis, the hydraulic diameter of the air-flow channel is substituted for the channel height. This substitution yields a channel-based, modified Rayleigh number, as given in Equation C-3. Again, this external analysis relates to the air flow and heat transfer in the annulus formed by the pressure vessel (i.e. canister) and the overpack. A Nusselt number correlation for a channel with uniform heat on one side and insulated on the other side is given in Equation C-4 [Kaminski and Jensen, 2005]. Again, the channel height is listed as H. However, the hydraulic diameter listed in these equations is defined by the annular air channel between the canister and the first shell, or "overpack".

kD"qg Ra H*D H 4C-32152512 24HDRa.HDRa Nu H*D H*D D H H H C-4 Results of the external analysis are presented in Table C-2. The average air flow velocity, Reynolds number, modified Rayleigh number, and the Nusselt number for the prototypic cask compare favorably with the DSC operated at high power.

C-3 Tabl e C-2 Comparison of External Dimensionless Groups for the DCS and Dry Cask Systems with Helium at 700 kPa Parameter Aboveground DCS DCS Cask Power (W) 500 5,000 36,900 Air (kg/s) 0.039 0.083 0.350 D H (m) 0.184 0.184 0.096 Wavg (m/s) 0.37 0.76 1.26 Re3,700 7,100 6,100 H*D R a 2.7E8 2.7E9 2.3E8 H D Nu 16 26 14 C.3 SummaryDimensional analyses indicate that the anticipated ranges of relevant dimensionless groups (Reynolds, Modified Rayleigh, and Nusselt numbers) bracket or closely approach prototypic values for both the aboveground and belowground configurations. While designed to match prototypic values, the expected test matrix will include values that exceed currently acceptable values for decay heat, internal helium pressure, and peak cladding temperatures to gain more insight into the underlying behavior of the system. C.4 ReferencesA. BEJAN, Convection Hea t Transfe r, 2 nd Ed., John Wiley and Sons, (1 995).D.A. KAMINSKI an d M.K. JENSEN, Introduction to The rmal and Fl uids Engineering, JohnWiley a nd Sons, (2005).