NUREG/CR-6894PNNL-15346
Spent Fuel TransportationPackage Rlesponse to theCaldecott Tunnel Fire Scenario
Draft Report for Comment
Pacific Northwest National Laboratory
U.S. Nuclear Regulatory CommissionOffice of Nuclear Material Safety and SafeguardsWashington, DC 20555-0001
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NUREG/CR-6894PNNL-15346
Spent Fuel TransportationPackage Response to theCaldecott Tunnel Fire Scenario
Draft Report for Comment
Manuscript Completed: January 2006Date Published: February 2006
Prepared byH.E. Adkins, Jr., B.J. Koeppel, J. M. Cuta
Pacific Northwest National LaboratoryRichland, WA 99352
A. Hansen, NRC Project Manager
Prepared forSpent Fuel Project OfficeOffice of Nuclear Material Safety and SafeguardsU.S. Nuclear Regulatory CommissionWashington, DC 20555-0001Job Code J5167
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Allen HansenOWTN 13 D-13U.S. Nuclear Regulatory CommissionWashington, DC 20555-0001Phone: 301-415-1390E-mail: AGHanrc.govFax: 301-415-8555
ABSTRACT
On April 7, 1982, a tank truck and trailer carrying8,800 gallons of gasoline was involved in anaccident in the Caldecott Tunnel on State Route 24near Oakland, California. The tank traileroverturned and subsequently caught fire. Becausethis event is one of the most severe of the fivemajor highway tunnel fires involving shipments ofhazardous material that have occurred world widesince 1949, the United States Nuclear RegulatoryCommission (USNRC) selected it for analysis todetermine the possible regulatory implications ofsuch events for the transportation of spent nuclearfuel by truck.
The Fire Dynamics Simulator (FDS) codedeveloped and maintained by the National Instititeof Standards and Technology (NIST) was used todetermine the thermal environment in theCaldecott Tunnel during the fire. The FDS resultswere used to define boundary conditions for athermal transient model of a truck transport caskcontaining spent nuclear fuel. The NuclearAssurance Corporation (NAC) Legal WeightTruck (LWT) transportation cask was selected forthis evaluation, as it represents a typical truck(over-the-road) cask.
Detailed analysis of the response of the transportpackage to the fire was performed using theANSYSg computer code. The staff concluded thatsmall transportation casks similar to the NACLWT cask would probably experience degradationof some seals in this severe accident scenario. Themaximum temperatures predicted in the regions ofthe cask lid and the vent and drain ports exceed therated service temperature of the tetrafluoro-ethylene {TFE) or Vitong seals, making itpossible for a small release to occur due to CRUDthat might spall off the surfaces of the fuel rods.However, any release is expected to be very smalldue to a number of factors. These include (1) the
metallic lid seal does not exceed its rated servicetemperature and therefore can be assumed toremain intact, (2) the tight clearances maintainedby the lid closure bolts, (3) the low pressuredifferential between the cask interior and exterior,(4) the tendency for solid particles to plug smallclearance gaps and narrow convoluted flow pathssuch as the vent and drain ports, and (5) thetendency of CRUD particles to settle or plate outand consequently not be available for release.
USNRC staff evaluated the radiologicalconsequences of the package response to theCaldecott Tunnel fire. The results of thisevaluation strongly indicate that neither spentnuclear fuel (SNF) particles nor fission productswould be released from a spent fuel shipping caskinvolved in a severe tunnel fire such as theCaldecott Tunnel fire. The NAC LWT caskdesign analyzed for the Caldecott Tunnel firescenario does not reach internal temperatures thatcould result in rupture of the fuel cladding.Therefore, radioactive material (i.e., SNE particlesor fission products) would be retained within thefuel rods. The potential release calculated for theNAC LWT cask in this scenario indicates that anyrelease of CRUD from the cask would be verysmall - less than an A2 quantity (see footnote 3,Section 8)
iii
DISCLAIMER
This report was prepared as an account of work sponsored by an agency of the United States Government.Neither the United States Government nor any agency thereof, nor Battelle Memorial Institute, nor any oftheir employees, makes any warranty, express or implied, or assumes any legal liability or responsibilityfor the accuracy, completeness, or usefulness of any information, apparatus, product, or process disclosed,or represents that its use would not infringe privately owned rights.
CONTENTS
ABSTRACT................................................................................................................................................. .. iii
ABBREVIATIONS ............. xv
1 INTRODUCTION .1.1
2 CALDECOTT TUNNEL FIRE EVENT .2.1
3 NIST TUNNEL FIRE MODEL .3.1
4 TRANSPORTATION OF SPENT NUCLEAR FUEL .4.1
4. It NAC LWT Transport Package .4.1
5 ANALYSIS APPROACH .5.1
5.I1 Model of NAC LWT Transportation Package .5.2
5.2 NAC LWT Transportation Package within the Tunnel . .5.35.2.1 With ISO Container .5.45.2.2 Without ISO Container .5.5
5.3 NAC LWT Transportation Package Material Properties .5.6
6 ANALYSIS METHOD .. 6.1
6.1] Modeling Assumptions for Fire Transient .6.1
6.2 Boundary Conditions for Fire Transient .. 6.26.2.1 Boundary Temperatures from FDS Analysis .6.26.2.2 Convection Boundary Conditions .6.5
6.'3 Initial System Component Temperatures .6.9
6.4 Tunnel Fire Transient .6.12
7 ANALYSIS RESULTS .. 7.1
7.1 NAC LWT Package Response to Fire Transient .7.1
7.2 NAC LWT Package Short-Tenn Post-Fire Transient Response .7.4
7.3 NAC LWT Package Long-Term Post-Fire Transient Response .7.7
v
7.4 Summary of NAC LWT Package Peak Temperatures in Fire Transient ................................ 7.9
8 POTENTIAL CONSEQUENCES ....................................................... 8.1
8.1 Release Analysis ....................................................... 8.1
9 REFERENCES ....................................................... . 9.1
vi
FIGURES
2.1 Cross-section Diagram of Bore No. 3 of Caldecott Tunnel ............................................................. 2.1
3.1 Evolution of Tunnel Ceiling Centerline Temperatures Predicted in FDS Simulation of CaldecottTunnel Fire ...................................................................... 3.2
3.2 Evolution of Tunnel Wall Mid-line Temperatures Predicted in FDS Simulation of Caldecott TunnelFire ...................................................................... 3.2
3.3 Evolution of Tunnel Floor Centerline Temperatures Predicted in FDS Simulation of CaldecottTunnel Fire ...................................................................... 3.2
3.4 Evolution of Gas Velocity Profile near Tunnel Ceiling Centerline Predicted in FDS Simulation ofCaldecott Tunnel Fire ...................................................................... 3.3
3.5 Evolution of Gas Velocity Profile near Tunnel Mid-line Predicted in FDS Simulation of CaldecottTunnel Fire ...................................................................... 3.3
3.6 Evolution of Gas Velocity Profile near Tunnel Floor Centerline Predicted in FDS Simulation ofCa]decott Tunnel Fire ...................................................................... 3.3
3.7 Evolution of Gas Temperature Profile near Tunnel Ceiling Centerline Predicted in FDS Simulationof Caldecott Tunnel Fire ...................................................................... 3.3
3.8 Evolution of Gas Temperature Profile near Tunnel Mid-line Predicted in FDS Simulation ofCaldecott Tunnel Fire ...................................................................... 3.4
3.9 Evolution of Gas Temperature Profile near Tunnel Floor Centerline Predicted in FDS Simulation ofCa]decott Tunnel Fire ...................................................................... 3.4
3.10 Tunnel Surface Temperatures at Hottest Location Predicted for First Hour of FDS Simulation ofCaldecott Tunnel Fire ...................................................................... 3.6
3.11 Tunnel Air Temperatures at Hottest Location Predicted for First Hour of FDS Simulation ofCaldecott Tunnel Fire ...................................................................... 3.6
3.12 Tunnel Air Velocities at Hottest Location Predicted for First Hour of FDS Simulation of CaldecottTunnel Fire ...................................................................... 3.6
3.13 Peak Tunnel Surface Temperatures Predicted in 3-Hr FDS Simulation of Caldecott Tunnel Fire . 3.6
3.14 Peak Tunnel Gas Temperatures Predicted in 3-Hr FDS Simulation of Caldecott Tunnel Fire . 3.7
Vii
3.15 Peak Tunnel Gas Velocities Predicted in 3-Hr FDS Simulation of Caldecott Tunnel Fire ............. 3.7
4.1 NAC LWT Transport Package (without ISO Container) .4.1
4.2 NAC LWT Transport Package (with ISO Container) .4.1
5.1 ANSYS NAC LWT Cask Analysis Model Element Plot (with ISO Container) .5.1
5.2 ANSYS NAC LWT Cask Analysis Model Element Plot (without ISO Container) .5.1
5.3 Cross Section of NAC LWT Cask Model in ANSYS .5.2
5.4 NAC LWT Cask Geometry (without ISO Container) .5.3
5.5 NAC LWT Cask Geometry (with ISO Container) .5.4
5.6 Zones for Convection Computations Within the ISO Container .5.4
5.7 Zones for External Heat Transfer Between ISO Container and Tunnel .5.5
6.1 Peak Temperatures for Radiation Exchange During Fire Transient in Caldecott Tunnel .6.4
6.2 Peak Temperatures for Convection Heat Transfer During Fire Transient in Caldecott Tunnel . 6.4
6.3 Peak Velocities for Convection Heat Transfer During Fire Transient in Caldecott Tunnel . 6.4
6.4 Peak Temperatures for Radiation Exchange During Extended Transient in Caldecott Tunnel . 6.5
6.5 Peak Temperatures for Convection During Extended Transient in Caldecott Tunnel .6.5
6.6 Nusselt Number for Heat Transfer in Liquid Neutron Shield .6.8
6.7 Effective Conductivity of Liquid Neutron Shield Tank .6.8
6.8 Effective Conductivity of Liquid Neutron Shield Expansion Tank .6.8
6.9 LWT Cask (with ISO Container): Normal-Hot Condition Temperature Distribution (2.5 kW DecayHeat, 130F Ambient) .6.10
6.10 LWT Cask (with ISO Container): Normal Condition Temperature Distribution (2.5 kW Decay HeatLoad)6.............................................................................................................................................. 6.11
6.11 LWT Cask (without ISO Container): Normal Condition Temperature Distribution (2.5 kW DecayHeat Load) .6.11
viii
7.1 NAC LWT Cask (with ISO Container): Component Maximum Temperature Histories During FireTransient ........ .............................................................. 7.2
7.2 NAC LWT Cask (without ISO Container): Component Maximum Temperature Histories DuringFire Transient ...................................................................... 7.2
7.3 Lumped Fuel Assembly Temperature Distribution 0.7 hr into Transient ........................................ 7.4
7.4 NAC LWT Cask (with ISO Container): Maximum Temperature Histories for First 3 hours of FireTransient ...................................................................... 7.4
7.5 NAC LWT Cask (without ISO Container): Maximum Temperature Histories for First 3 hours ofFire Transient ....................................................................... 7.5
7.6 Maximum Predicted ISO Container Surface Temperature History Compared to NIST BoundaryCondition Temperatures ....................................................................... 7.5
7.7 Maximum Predicted Cask Outer Surface Temperature History for NAC LWT Cask without ISOContainer Compared to NIST Boundary Condition Temperatures ...................... ........................... 7.5
7.8 NAC LWT Cask (with ISO Container): Maximum Seal Temperature Histories for Drain/Vent Portsand Cask Lid During First 3 hours of Fire Transient ...................................................................... 7.6
7.9 NAC LWT Cask (without ISO Container): Maximum Seal Temperature Histories for Drain/VentPorts and Cask Lid During First 3 hours of Fire Transient .............................................................. 7.6
7.10 NAC LWT Cask (with ISO Container): Maximum Temperature Histories During 50-hourTransient ...................................................................... 7.8
7.11 NAC LWT Cask (without ISO Container): Maximum Temperature Histories During 50-hourTransient ...................................................................... 7.8
ix
TABLES
3.1 Air and Surface Temperatures Near Hottest Fire Location ............................................................ 3.4
3.2 Total Energy Flux Values Near Hottest Fire Location ................................................................ 3.4
3.3 Temperature and Energy Flux Values at Hottest Location in Tunnel ..................... ....................... 3.5
6.1 NAC LWT Component Temperatures at Various Decay Heat Loads .................... ....................... 6.10
6.2 NAC LWT Component Temperatures at 2.5 kW Decay Heat Load and 130'F Ambient ............. 6.11
6.3 NAC LWT Component Temperatures for 2.5 kW Decay Heat Load and 100'F Ambient ........... 6.12
7.1 NAC LWT Peak Component Temperatures During Fire Transient ................................................. 7.9
8.1 Assumptions Used for Release Estimate for NAC LWT Cask ......................................................... 8.2
8.2 Potential Release Estimate for NAC LWT Cask ..................................... ........................... 8.3
xi
APPENDIX
APPENDIX A - Material Properties for ANSYS Model of Legal Weight Truck Package ........... .......... A.1
xiii
ABBREVIATIONS
APDL ANSYSO Parametric Design Language
BWR Boiling Water Reactor
CFD Computational Fluid Dynamics
CRUD Chalk River Unknown Deposit (generic term for various residues deposited on fuel rodsurfaces, originally coined by Atomic Energy of Canada, Ltd. (AECL) to describe depositsobserved on fuel removed from the test reactor at Chalk River.)
FDS Fire Dynamics Simulator (code)
FEA Finite Element Analysis
ISO International Organization for Standardization (The International Organization forStandardization has decreed the use of the initials ISO for reference to the organization,regardless of the word order of the organization's name in any given language. Thisdefines a uniform acronym in all languages.)
LWT Legal Weight Truck
NIST National Institute of Standards and Technology
NTSB National Transportation Safety Board
OFA Optimized Fuel Assembly
PNNL Pacific Northwest National Laboratory
PWR Pressurized Water Reactor
SFPO USNRC Spent Fuel Project Office
SNF Spent nuclear fuel
TFE tetrafluoroethylene; generic term far polytetrafluoroethylene polymers such as Teflon®
USNRC United States Nuclear Regulatory Commission
xv
1 INTRODUCTION
Current NRC regulations specify that spentnuclear fuel shipping casks must be designed tosurvive exposure to a fire accident lasting at least30 minutes, with an average flame temperature ofno less than 14750F (802'C) [1]. The cask mustmaintain shielding and criticality control functionsthroughout the fire event and post-fire cool downin order to meet USNRC requirements.
On April 7, 1982, a tank truck towing a tank trailerand carrying 8,800 gallons (33,310 liters) ofgasoline was involved in an accident in theCaldecott Tunnel on State Route 24 near Oakland,California. The tank trailer overturned and caughtfire. Darmage to the tunnel ceiling indicated thatthe fire reached peak temperatures in the range of1880-19500 F (1026-10650 C) [2]. Based oncalculations for the amount of fuel present andavailable tunnel air flow, it is estimated that thegasoline fire burned for approximately 40 minutes,and temperatures above 14750F (802'C) may havebeen sustained for up to 35 minutes during thegasoline-fueled portion of the fire [3].
The severity of the Caldecott Tunnel fire hasraised questions about the performance of spentfuel casks licensed for transport by truck, shouldone be involved in such an accident. The staff ofthe USNRC Spent Fuel Project Office (SFPO) hasundertaken analyses to evaluate the impact thisevent could have had if it had involved a spentnuclear fuiel cask. As part of the investigationrelated to this accident, finite element analysis(FEA) evaluations were performed subjecting a
model of a typical truck transportation cask toexternal temperatures representing the predictedconditions of this fire.
Detailed temperature boundary conditions(including temperatures of the combustion gasesand the surrounding surfaces of the tunnel) wereobtained from fire simulations performed at theNational Institute of Standards and Technology(NIST). The purpose of this analysis of the spentfuel transport cask was to obtain an estimate of thetemperature response of the various components ofthe cask during and after the fire.
This report presents a detailed description of theanalysis, including boundary conditions, modelingapproach, and computational results. Section 2presents a brief description of the CaldecottTunnel fire based on the National TransportationSafety Board (NTSB) investigation of the event[2]. Section 3 describes the detailed temperatureboundary conditions obtained from the firesimulations performed by NIST. Section 4describes the NAC LWT spent fiuel transportationcask. Section 5 describes the analysis approachand the computational model developed for theanalysis. Section 6 presents the analysis method.Section 7 describes the results of the simulation,giving a detailed evaluation of the cask responseduring and after the fire. Section 8 provides ananalysis to determine the magnitude of anypotential release of radioactive material as aconsequence of the effects of the fire on the NACLWT transportation package.
1.1
2 CALDECOTT TUNNEL FIRE EVENT
This tunnel fire occurred shortly after midnight onApril 7, 1982 in Bore No. 3 of the CaldecottTunnel on State Route 24 near Oakland,California, as the result of an accident involving atank truck towing a tank trailer and carrying 8,800gallons (33,310 liters) of gasoline[2]. Bore No. 3of the Caldecott Tunnel is 3,371 ft (1027 m) long,with a two-lane roadway 28 ft (8.5 m) wide.Traffic is one-way from east to west, and theroadway has a 4% downgrade beginningapproximately 30 ft (9.1 m) into the tunnel.Figure 2.1 shows a diagram of the tunnel typicalcross-section. Vertical clearance between thetunnel ceiling and the roadbed is about 18 ft (5.5m) at the center, tapering to 17 ft (5.2 m) at theside walls. The tunnel width is approximately34.5 ft (10.5 m) between the sidewalls of the bore.The tunnel is actively ventilated by blowers with atotal capacity of 1.5 million cubic feet per minutethrough ducting above the tunnel ceiling.(However, the blowers were not operating at thetime of the fire.)
which varies in thickness from 6 ft (1.8 m) at thebottom to 2 ft (0.6 m) at the top. The wall surfaceis covered with 4.25-inch (10.8-cm) square greentiles. The ceiling between the roadway and theventilation ducting is 5.5-inch (14-cm) thickPortland cement concrete. Ventilation ports (5 ft x1 ft (1.5 m x 0.3 m)) covered with steel gratingsare spaced at 15-ft (4.6-m) intervals along bothsides of the ceiling for the full length of the tunnel.
In the accident, the tank truck and trailer collidedwith a stalled vehicle in the tunnel and wassubsequently struck by a bus. The tank traileroverturned and the entire vehicle (tanker andtrailer) came to rest approximately 1650 ft (503 in)from the west portal of the tunnel. Gasolinespilled onto the roadway from the damaged tanktrailer. A fire erupted, and within four minutes ofthe accident heavy black smoke began pouring outthe east portal of the tunnel. The tank truck, trailerand five other vehicles in the tunnel werecompletely destroyed by the fire, seven personswere killed, and the tunnel incurred major damage.
Based on NTSB evaluations of the fire debris andinterviews with emergency responders, theintensely hot gasoline-fueled portion of the fire isestimated to have lasted less than three-quarters ofan hour. Although vehicles within the tunnel werestill burning at 46 minutes after the start of thefire, firefighters in protective gear entered thetunnel to search for survivors and were able toapproach the location of the tanker truck.
At approximately 55 minutes after the start of thefire, the smoke had cleared sufficiently for theCalTrans supervisor to visually assess structuraldamage to the tunnel and authorize fire crews toenter the tunnel. After laying hoses from thenearest standpipe in the center bore of the tunnelthrough cross adits (a process that required
Figure 2.1. Cross-section Diagram of Bore No.3 of Caldecott Tunnel
The roadway pavement is Portland cementconcrete, as are the arched walls of the bore,
2.1
approximately 30 minutes), the fire crews beganfighting the fires due to the burning vehicles in thetunnel. Within approximately 25 minutes (about
2.7 hours after the time of the accident), these fireswere reported to be "under control".
2.2
3 NIST TUNNEL FIRE MODEL
Experts at the National Institute of Standards andTechnology (NIST) developed a model of theCaldecoti: Tunnel fire using the Fire DynamicsSimulator (FDS) code [3]1 and performed analysesto obtain predictions of the range of temperaturespresent in the tunnel during the fire event [4].FDS is a computational fluid dynamics (CFD)code that models combustion and flow of hotgases in fire environments. FDS solves the mass,momentum, and energy equations for a givencomputational grid, and is also able to construct avisual representation of smoke flow for the fire.Full details on the analyses performed by NIST areprovided in the report on the FDS analysis of theCaldecott Tunnel fire [4]. A brief description ofthe model and a summary of the results arepresented here.
To validate the FDS code for tunnel fireapplications, NIST developed fire models in FDS'based on the geometry and test conditions from aseries of fire experiments conducted by theFederal Highway Administration and ParsonsBrinkerhoff, Inc. as part of the Memorial TunnelFire Ventilation Test Program [5]. NIST modeleda 6.83x1(17 Btu/hr (20 MW) and a 1.71X 108 Btu/hr(50 MW) unventilated fire test from the MemorialTunnel Test Program, and achieved results usingFDS that were within 1000F (560 C) of therecorded data [3,6].
The NIST model of the Caldecott Tunnel for theFDS code consisted of the section of the tunnelthat experienced the most severe effects of the fire.The model extends from 1509 ft to 2297 ft (460 into 700 m) relative to the west portal of the tunnel,
IFormal publication of the FDS code documentationbegan in 2001 with Version 2. Continuing validationand development of the code led to Version 3 in 2002.Version 3 was used in the FDS analyses discussed inthis report.
with the fire location defined at 1673-1706 ft (510-520 in). The fire location corresponds to thelocation of the tank truck and trailer, which cameto rest with the front of the truck approximately1650 ft (503 m) from the west portal. From thefire center to approximately 2297 ft (700 in), thefire resulted in essentially uniform spalling of theconcrete on the tunnel walls and ceiling, theunderlying reinforcing steel was exposed, andthere was heat buckling of the steel ventilatoropening cover plates. The wall tiles and grout alsoshowed severe spalling in this region of the tunnel,and the fluorescent lighting fixtures andemergency phones were destroyed or damaged.
The computational grid for the tunnel fire modelconsisted of a fully three-dimensional (3-D)representation of this segment of the tunnel inorder to capture flame and gas behavior and theinteraction of the fire with the tunnel walls,ceiling, and floor. Based on boundary conditionsthat includes information on the available fuel andair sources, the FDS code calculates the energyrelease from the combustion process, the resultingflow of air and hot combustion gases, and local airand surface temperatures throughout the tunnel.
The FDS calculation simulated only the gasolinefire, and did not include the thermal energyreleased due to the burning vehicles. Compared tothe energy released by the gasoline fire, the energyreleased by the burning vehicles is negligible, andthese individual vehicle fires were located far fromthe hottest region in the tunnel. The tank truckitself was 328 ft (100 m) away from the hottestlocation in the tunnel during the fire. Of the fiveother vehicles destroyed in the fire, the closestvehicle was at least 223 ft (68 in) away from thehottest location in the tunnel, and the remainingvehicles were approximately 575 ft (175 in), and1224 ft (373 m) away.
3.1
Figures 3.1, 3.2, and 3.3 illustrate the modelresults, showing the evolution of the surfacetemperatures on the tunnel ceiling, walls and floorduring and immediately after the gasoline-fueledfire. Temperature profiles are shown along theaxial length of the portion of the tunnel included inthe model. These plots show the surfacetemperature profiles along the axial length of thetunnel at the ceiling centerline (see Figure 3.1),mid-way up the tunnel wall (see Figure 3.2), andat the centerline of the tunnel floor (see Figure 3.3)at various times during the fire transient.
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so e M 1 bd - - 1 _ M o ____|_E
Location In tunnel (meters. relative to west portal)
Figure 3.2. Evolution of Tunnel Wail Mid-huneTemperatures Predicted in FDSSimulation of Caldecott TunnelFire
a fiO Tat teea …tII
flT 2 e I TI I I
a fi~cTot2Si3 e£ -fierT tOeetm
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Location In tunnel (meters elv btowest w~portl1)
Figure 3.3. Evolution of Tunnel Faloord-inCeneTemperatures Predicted i
inFSSimulation of Caldecott TneTunnelFr
Tecae ,-of . th2is~e shift of th- --e hots loca-tio to a
- Ibt T int 3 l tte I I I t I _ I
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pocaao o tf cb usiongtses la he ar to Wst po
drivresi flow tEvolughthen ofTunnel, caorrynh
hea ofthe f nteirir e westanes lrenct
ff~~i FDS S or 1imulatio of Caldecot-t~ ~v ~r
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thecmaefcyton thfef aitro aenad fae thoempeatureaounnelg the lnth of the t . re.Expansion of combustion gases and heated airdrives air flow through the tunnel, carrying theheat of the fire from east to west along the lengthof the tunnel. The FDS code solves for thehydrodynamics of this flow, and calculates thethermal effect on the air and surface temperaturesalong the length of the tunnel. Figures 3.4, 3.5,
. . .
Location In tunnel (meters, relativ to West portal)
Figure 3.1. Evolution of Tunnel CeilingCenterline Temperatures Predictedin FDS Simulation of CaldecottTunnel Fire
The plots in Figures 3.1, 3.2, and 3.3 show that inthe first few minutes of the fire, the tunnel surfacetemperatures in the vicinity of the fire begin to riserapidly. Temperatures farther away down thelength of the tunnel (east of the fire location) beginto rise also, but initially these temperatures risemore slowly. After about the first five minutes ofthe fire, however, the tunnel surfaces to the east ofthe fire location are rising much more rapidly thanthe those near the fire. By the end of the fire, atabout 39-40 minutes elapsed time, the surfaces atabout 262-394 ft (80-120 m) to the east of the fire(i.e., at 1968-2099 ft (600-640 m) relative to thewest portal) are the hottest surfaces in the tunnel.
3.2
and 3.7 show the evolution of the predicted airflow velocities in the tunnel in the upper regionnear the ceiling, the mid-line region near thetunnel wall, and near the floor of the tunnel.
centerline, mid-way between the ceiling and floor.The evolving temperature profiles for the gasmoving at these velocities are shown in Figures3.7, 3.8, and 3.9.
la
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Location In tunnel (meters, relatve to west portal)
Figure 3.4. Evolution of Gas Velocity Profilenear Tunnel Ceiling CenterlinePredicted in FDS Simulation ofCaldecott Tunnel Fire
Figure 3.6. Evolution of Gas Velocity Profilenear Tunnel Floor CenterlinePredicted in FDS Simulation ofCaldecott Tunnel Fire
p a.
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a Ma Ya Sa sMo soo so a a s mo MS r a r
Location In tunnel (meters. relatie to west portal)
Figure 3.5. Evolution of Gas Velocity Profilenear Tunnel Mid-line Predicted inFDS Simulation of CaldecottTunnel Fire
The velocities in the plots in Figures 3.4 and 3.6for the air near the ceiling and floor of the tunnelare from nodes approximately 1 ft (0.3 m) fromtheir respective surfaces. The velocity in the mid-line region (in Figure 3.5) is along the tunnel
Figure 3.7. Evolution of Gas TemperatureProfile near Tunnel CeilingCenterline Predicted in FDSSimulation of Caldecott TunnelFire
The plots of the evolving tunnel surfacetemperatures in Figures 3.4, 3.5, and 3.6, and oftunnel air temperatures in Figures 3.7, 3.8, and 3.9show that the highest temperatures during the fire
3.3
do not occur at exactly the same location in theupper, middle, and lower regions of the tunnel.
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further 'downstream', at 2100 ft (640 in). Thepeak air temperature near the floor also occurs at2100 ft (640 in), but the peak floor temperatureoccurs at 2165 ft (660 in).
Table 3.1. Air and Surface Temperatures (IC)Near Hottest Fire Location
Location 580 600 620 640 660m (ft): (1903) (1969) (2034) (2100) (2165)
Upper air 1039 1 1040 985 929Mid-lineair 1016 1000 982 933Near-floor air 463 560 729 823 834
Ceilingcenterline 905 909 869 826Wall mid- |3
line 789 818 844 E850 839Floorcenterline 680 705 747 763
Location In tunnel (meters, relatiem to west portal)Do
Figure 3.8. Evolution of Gas TemperatureProfile near Tunnel Mid-linePredicted in FDS Simulation ofCaldecott Tunnel Fire
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Nevertheless, a "hottest location" must be definedin order to determine the boundary conditions thatwould be seen by a nuclear fuel shipping casksubjected to the extreme temperature conditions ofthe Caldecott Tunnel fire. The difficulty can beresolved by considering the energy output of thefire at a given location over time, as well as thetemperature history at that location. Table 3.2summarizes the heat flux values calculated withFDS for the locations of the hottest temperaturesalong the tunnel.
Table 3.2: Total Energy Flux Values Nearof Hottest Fire Location
W M s o em s e - te o on e a a
Location In tunnel (metere. relative to west portal)
Figure 3.9. Evolution of Gas TemperatureProfile near Tunnel Floor Center-line Predicted in FDS Simulation iCaldecott Tunnel Fire
enerqv flux (kW/hr-m')
Table 3.1 summarizes the peak temperaturespredicted with FDS for the upper, middle, andlower regions of the tunnel. The hottest locationnear the ceiling of the tunnel occurs at 1969 ft(600 in), the hottest air temperature at the mid-lineoccurs at 1903 ft (580 m), but the hottest mid-linewall temperature occurs 197 feet (60 meters)
Location 560 580 600 620 640 660
Im (ft): (1837) (1903) (1969) (2034) (2100) (2165)Ceilingcenterline 106 123 134 118 100 86Wallmid-line 61 78 89 95 96 89Floorcenterline 50 60 62 68 71 66
3.4
The highest heat flux from the fire occurs at 1969)ft (600 m) for the ceiling, and at 2100 ft (640 m)for the walls and floor. Based on the distributionof the heat output from the fire and the temper-ature distribution on the tunnel surfaces and in thetunnel air, the conditions at 2034 ft (620 m) givethe highest temperatures at the highest heat fluxvalues. Table 3.3 summarizes the conditions at2034 ft (620 m), and compares the temperatureand energy flux values to the peak values for theceiling, wall, and floor regions of the tunnel.
Table 3.3: Temperature and Energy FluxValues at Hottest Location inTunnel
Surface temperature (0C) % ofat 2034 ft (620 m) peak
Ceiling centerline 909 97%Wall mind-line 844 99%Floor centerline 747 97%
Energy flux (kW/hr-m 2)at 2034 ft (620 m)
Ceiling centerline 118 88%Wall mid-line 95 99%Floor centerline 68 96%
Air temperature (IC)at 2034 ft (620 m)
Ceiling centerline 1039 97%Wall mid-line 1000 96%Floor centerline 729 87%
temperatures at 2034 ft (620 m) are within 34%of their corresponding peak values, and althoughthe near-floor air temperature is only 87% of itspeak value at this location, at 1969 ft (600 m) it isonly 67% of the peak value. In terms of the effectof the fire conditions on a cylindrical package suchas a spent fuel transportation cask positionedwithin the tunnel, the conditions at 2034 ft (620 m)represent the best estimate of the "hottest location"in the tunnel, in that it maximizes the temperaturesand heat fluxes seen by all surfaces of the cask.
Figure 3.10 shows the temperatures of the tunnelceiling centerline, wall mid-line, and floorcenterline predicted with FDS for the first hour ofthe simulation at 2034 ft (620 m), defined as thehottest location in the tunnel during the gasoline-fueled portion of the fire transient. Figure 3.11shows the air temperatures during this time for theupper, middle, and lower regions of the tunnel at2034 ft (620 in). Figure 3.12 shows the predictedvelocities produced by the fire at the locations ofthe air temperatures shown in Figure 3.11. Thesevelocities are used to define the convective heattransfer conditions on the top, sides and bottom ofthe spent fuel cask during the fire. (Thetemperature-vs.-time and velocity-vs.-time valuesin these plots were smoothed to conservativelyremove the rapid stochastic variations typical offire dynamics, preserving only the major peaksand troughs defining the general physical behaviorof the simulated fire.)
Maximum gas temperatures calculated in the FDSmodel are on the order of 19650F (10740C). Themaximum tunnel surface temperatures arepredicted to be only about 1715'F (9350C) (seeFigure 3.10). Maximum air temperatures in theupper and middle regions of the tunnel arepredicted to exceed 18320 F (10000C) in the first 5to 6 minutes of the fire, and remain above thistemperature until the end of the gasoline-fueledportion of the fire (at approximately 40 minutes.)
Although none of the regions of the tunnel havetheir peak temperature or peak energy flux valuesexactly at 620 m, defining this point as the "hottestlocation" in the tunnel gives an overall energy fluxthat is within 99% of the peak energy flux, incombination with local tunnel surfacetemperatures that are within 1% to 3% of theirrespective peak values. The conditions at 2034 ft(620 in) are more severe than at 1969 ft (600 m),
which has slightly higher total energy flux andwall and ceiling temperatures, but much lowertunnel floor temperatures. Similarly, the air
3.5
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dropped to 154°F (68°C) or lower, and the tunnelsurface temperatures are less than 320°F (160°C)
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Figure 3.10. Tunnel Surface Temperatures atHottest Location Predicted forFirst Hour of FDS Simulation ofCaldecott Tunnel Fire
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Figure 3.11. Tunnel Air Temperatures atHottest Location Predicted forFirst Hour of FDS Simulation ofCaldecott Tunnel Fire
The FDS calculation was run out for a totaltransient time of three hours, including the 40-minute gasoline-fueled fire and a 2.3 hr cool-downperiod. Temperatures and velocities for the fullduration of the FDS calculation are shown inFigure 3.13 (for the tunnel surface temperatures),Figure 3.14 (for the tunnel air temperatures), andFigure 3.15 (for the tunnel air velocities.) By theend of this three-hour period, the tunnel airtemperatures predicted at the hottest location have
0.00 0.25 0.50 0.75 1.00 125 1.50 1.75 2.00 2.25 2.50 2.75 3.00
Elapsed lime (hours)
Figure 3.13. Peak Tunnel SurfaceTemperatures Predicted in 3-HourFDS Simulation of CaldecottTunnel Fire
3.6
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0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.00 2.25 2.50 2.75 3.00
Elapsed Time (hours)
0.00 0.25 0.50 0.75 1.00 1.25 1.50 1.75 2.80 2.25 2.50 2.75 3.00
Elapsed Tlme (hours)
Figure 3.14. Peak Tunnel Gas TemperaturesPredicted in 3-Hour FDSSimulation of Caldecott TunnelFire
Figure 3.15. Peak Tunnel Gas VelocitiesPredicted in 3-Hour FDSSimulation of Caldecott TunnelFire
3.7
4 TRANSPORTATION OF SPENT NUCLEAR FUEL
NRC regulations require spent fuel transportationpackages to be evaluated for a series ofhypothetical accident conditions that include afully engulfing fire with an average flametemperature of 14750F (802'C) for a period of 30minutes. The certification process must includeeither an open pool fire test or an analysis of thepackage for a fire exposure meeting these criteria.Packages must maintain shielding and criticalitycontrol functions throughout the sequence ofhypothetical accident conditions.
In this investigation, a typical spent nuclear fueltransportation package licensed for over-the-roadtransport by truck is subjected to boundaryconditions simulating the thermal conditions of theCaldecott. Tunnel fire, to determine the response ofthe package to these severe conditions. TheNuclear Assurance Corporation (NAC) LegalWeight Truck (LWT) transportation package wasselected for this analysis because it represents atypical package that can be transported by truck.A complete description of the package design andloading configurations can be found in thelicensing SAR [7]. A brief description of thisdesign is presented below.
4.1 NAC LWT Transport Package
The NAC LWT transportation package is certifiedto be carried on a standard tractor trailer truck. Itis typically shipped within an InternationalOrganization for Standardization (ISO) shippingcontainer. Figure 4.1 shows a picture of an LWTpackage on a flat-bed trailer with a personnelbarrier installed, but without an ISO container.Figure 4.2 shows an exterior view of the packagewithin an ISO container on a flat-bed trailer. Thispackage is designed to transport a variety ofcommercial and test reactor fuel types with widelyvarying maximum decay heat load specifications.
For this analysis, the cask was assumed to containa single PWR spent nuclear fuel assembly, with adecay heat load of 8,530 Btu/hr (2.5 kW). This isthe highest heat load the LWT package is rated forwith any spent fuel it is designed to carry (refer toAmendment 34 of the SAR [7]), and ensures aconservative thermal load for the package in thefire accident scenario.
Figure 4.1. NAC LWT Transport Package(without ISO container)
Figure 4.2. NAC LWT Transport Package(with ISO container)
The loaded package weighs approximately 52,000lb (23,586 kg). The containment boundaryprovided by the stainless steel package consists ofa bottom plate, outer shell, upper ring forging, andclosure lid. This cask has an additional outer
4.1
stainless steel shell to protect the containmentshell, and also to enclose the lead gamma shield.Neutron shielding is provided by a stainless steelneutron shield tank containing a water/ethyleneglycol mixture. An additional annular expansiontank for the mixture is provided, external to theshield tank. This component is strengthened
internally by a network of stainless steel stiffeners.Aluminum honeycomb impact limiters coveredwith an aluminum skin are attached to each end ofthe cask during transport. The entire package,including impact limiters, fits within an ISOcontainer, which is constructed of steel plate.
4.2
5 ANALYSIS APPROACH
In the analytical approach used to evaluate theresponse of the NAC LWT transportation cask tothe conditions of the Caldecott Tunnel fire, ahighly detailed three-dimensional (3-D) modelwas constructed. The ANSYS [8] general finiteelement analysis (FEA) package was selected forthis analysis, since it is a widely used analyticaltool for licensing analyses of spent nuclear fuelcasks. Using this approach, the model included allsignificant heat transfer paths within the cask andbetween the cask and the external environment.The computational model was subjected to thethermal environment of the tunnel during the firetransient using boundary conditions derived fromthe NIST simulation of the fire using FDScomputational fluid dynamics code.
The model of the NAC LWT cask constructed forANSYS consists of a detailed 3-D representationof a symmetric half-section of the spent fuelpackage and a complete cross section of thesurrounding tunnel wall. Because the cask can beshipped uncovered or enclosed in an ISO shippingcontainer, two models were constructed; one thatincluded the ISO container, and one that did not.For both cases, the cask is oriented horizontallywithin the tunnel. This orientation gives the caskor ISO container outer surface the maximumexposure to the highest temperatures in the fireenvironment. This includes exposure to the tunnelsurfaces for thermal radiation exchange and to theflow of hot gases generated by the fire, whichresults in significant convection heat transfer tothe package during the fire transient. A diagramof the package model (including the ISOcontaineri and part of the tunnel is shown inFigure 5. L. Figure 5.2 shows a similar diagram fbrthe analysis without the ISO container.
the fire, is omitted from the analysis. However,the cask is assumed to be located within the tunnelat a vertical height corresponding to the height ofthe flatbed. This assumption yields the minimumpossible distances for thermal radiation exchangewith the hottest surfaces in the tunnelenvironment, and exposes the cask to the hottestair temperatures in the tunnel.
Figure 5.1. ANSYS NAC LWT Cask AnalysisModel Element Plot (with ISO)
The flatbed of the truck, which would tend toshield the bottom of the cask from the effects of
Figure 5.2. ANSYS NAC LWT Cask AnalysisModel Element Plot (without ISO)
5.1
The model used 40,489 SOLID70 8-node brickelements and 4,776 SHELL57 4-node quadrilateralthermal elements to represent the structuralcomponents. A total of 7,165 SURF 152 elementswere used to include thermal radiation between theISO container surfaces and the tunnel, andconvection heat transfer at the ISO containersurfaces. Sixteen MATRIX50 elements were usedto model thermal radiation exchange betweensurfaces within the ISO container. The surfaceeffect elements were also used to generate solarinsolation loads for calculation of the initialsteady-state temperature distribution for the cask.
5.1 Model of NAC LWTTransportation Package
The model geometry for the internal componentsof the cask was developed from the vendor'sengineering drawings. The representation of thecask internal components was identical in bothcases considered, with and without the ISOcontainer enclosing the cask. The cask contains acylindrical solid aluminum basket that holds asingle fuel assembly. The helium gaps between thefuel and the basket, and between the basket andcask shell, were explicitly modeled with solidelements. The cask model cross section is shownin Figure 5.3.
The cask body is constructed of concentricstainless steel shells to provide structural supportand some gamma shielding. The innermost shellis surrounded by a layer of lead that acts as themain gamma shield. The outermost stainless steelshell is surrounded by an annular tank containing a56% solution of ethylene glycol and water whichacts as a neutron shield. The tank is contained byan outer stainless steel skin and an annular over-flow tank that extends approximately one-third theaxial length of the cask body.
All of these components were modeled using brickelements. The tank is constructed with sixteen
stainless steel support ribs connecting the skin tothe outer shell. These structures were modeledwith shell elements.
mAN
Figure 5.3. Cross Section of NAC LWT CaskModel in ANSYS
The cask bottom consists of a stainless steel base,a layer of lead shielding, and a steel cover. Theupper end of the cask is sealed with a stainlesssteel lid, as illustrated in Figure 5.4. Impactlimiters attached to each end of the cask consist ofan internal aluminum honeycomb structurecovered by an aluminum skin. The expansion tankto handle overflow of the liquid neutron shield hasan outer stainless steel skin.
In the ANSYS model, the cask is assumed to belocated relative to the tunnel surfaces at a levelcorresponding to the height it would be above thetunnel floor when sitting on the bed of the truck.For the analysis in which the cask is within an ISOcontainer, it is similarly assumed that the top ofthe ISO container is at a height corresponding tothe height of the container plus the height of thetruck bed.
All three possible modes of heat transfer (i.e.,conduction, convection, and radiation) werecarefully represented in the model for thermalenergy exchange between all of the components.
5.2
Conduction is handled inherently by the geometlyof the connections between the elements modelingeach cornponent, but convective and radiationmechanisms must be carefully implemented usingappropriate modeling options.
AMl.
Figure 5.4. NAC LWT Cask Geometry
Westinghouse 17x17 OFA fuel was used in thisevaluation. The fuel assembly was modeled withan effective conductivity determined using ahomogenization scheme similar to that presented.by Babney and Lotz [9], modified to include ahelium gap between the homogenized fuel regionand the fiuel basket. This yields a more realisticrepresentation of the temperature profile throughthe assencbly, and takes into account the effect ofthe non-uniform wall temperature distributionaround the assembly.
Axial conduction in the homogeneous fuel regionwas conservatively neglected in the fuel itself, andwas modeled in the cladding only, using theconductivity of Zircaloy modified by a weightingscheme based on the cross-sectional area. Theeffective density and heat capacity for the fuelregion was based on volumetric averages of theproperties of the helium cover gas, fuel rodcladding, and uranium oxide fuel pellets. Thedesign basis axial power profile from the SAR [7],which has a normalized peaking factor of 1.2, was
used to establish the volumetric heat generation of8,532 Btulhr (2.5 kW) over the active fuel lengthof the assembly.
The 0.225-inch (0.57-cm) gap filled with a heliumcover gas between the fuel and the basket wasmodeled with solid elements and used standardhelium thermal conductivity, density, and specificheat. Convection was ignored in this small gap.The 0.25-inch (0.64-cm) gap between the basketand the inner shell was modeled in the samemanner, assuming negligible convection. Gapsbetween the lead gamma shielding and cask innerand outer shells due to contraction of the lead afterpour were accounted for in the model bycomputing effective conductivities assuming boththermal radiation and conduction across the gap.Effective conductivities were also used to includethe effect of the Fiberfrax paper insulationbetween the lead and the steel cask shell.
Radiation interaction across helium-filledenclosures in the cask interior was modeled bycoating the surfaces of elements bordering theseregions with SHELL57 elements having specifiedemissive material properties. The SHELL57elements were then used to produce highly struc-tured AUX-12 generated MATRIX50 super-elements, each defined by an enclosure, and theAUX-12 hidden ray-tracing method was used tocompute view factors for each element in thesuperelement. A total of 10 MATRJX50superelements were defined to capture the thermalradiation interactions within the cask and canister.
5.2 NAC LWT TransportationPackage within the Tunnel
The presence or absence of the ISO container hasa significant effect on the environment seen bysurface of the LWT cask within the tunnel.Without the ISO container, the exterior surface ofthe cask is directly exposed to the tunnelenvironment during the fire. With the ISO
5.3
container, the exterior surface of the cask isshielded from direct interaction with the tunnelenvironment. Instead, the cask exchanges heatwith the inner surface of the ISO container, andonly the ISO container outer surface is directlyexposed to the tunnel environment. As a result,the two cases require somewhat different modelingapproaches to appropriately account for heattransfer at the cask surface.
Calculations for convection heat transfer on theexternal surface of the ISO container usedempirical relations for free convection over flatplates (see Section 6 for full details). Convectionat a surface was implemented using SURF 152elements. These elements are placed on theexterior surface of a body and communicate withthe designated sink temperature assigned to asingle node (called the "space node") to computethe heat flux.
5.2.1 With ISO Container
For the analysis with the cask enclosed in an ISOcontainer, the model illustrated in Figure 5.4 wasenclosed within additional elements modeling theISO container, as shown in Figure 5.5. For thelarge air volumes between the cask outer surfaceand the inner surface of the ISO container,conduction across the gaseous medium isnegligible, but significant convection currents willbe created by the buoyant forces due to the heatedsurfaces. Surfaces with unobstructed views ofother surfaces will also experience significantradiation exchange that is highly dependent on thesurface geometry. Therefore, heat exchangebetween the cask exterior and the containerinterior was modeled with internal free convectionand thermal radiation between interior surfaces.
Convective heat transfer rates between the outersurface of the cask and the inner surface of theISO container are expected to vary in differentregions, due to geometry considerations andvarying temperature gradients. This wasaccounted for in the model by dividing the volumeenclosed by the ISO container into 17 zones, asillustrated in Figure 5.6. A separate zone wasdefined on each end of the cask, three zones weredefined for the top, side, and bottom radialsurfaces of each impact limiter, and three similarzones were defined for the cask body along itsaxial length. The sink temperature for each zoneis computed as the average surface temperature ofthe participating cask surface elements and ISOcontainer inner surface elements for that zone.
Figure 5.5. NAC LWT Cask Geometry withinISO Container
Figure 5.6. Zones for Convection Computa-tions Within the ISO Container
5.4
A convective heat transfer coefficient is assignedto the cask and container elements based on thesurface geometry and the temperature differencebetween the surface and the local sink temperaturefor that zone (see Section 6.)
In addition to convection heat transfer at the casksurface, a total of five MATRIX50 superelementswere defined to capture the radiation interactionbetween the cask and interior surfaces of the ISC)container'. The heat exchange between thesesurfaces ;and the space node is computed byANSYS during the solution.
Convection and thermal radiation are also the twoavailable mechanisms for heat transfer from theexterior surface of the ISO container. In the fireanalysis, the initial temperature distribution isobtained from a steady-state calculation forboundary conditions specified by IOCFR71.71 [I],followed by a transient calculation representingthe fire.
During the fire, the sink node temperatures for theSURF 1 52 elements are set and the externalconvection coefficient is computed using a forcedconvection relation derived from gas temperaturesand velocities predicted in the NIST firesimulation. These results were obtained for thetop, side, and bottom of the tunnel, and applied tothree zones defined on the top, sides, and bottomof the ISO container, as illustrated in Figure 5.7.
Thermal radiation between the outer surface of dieISO container and the tunnel during and after thefire is incorporated by a MATRIX50 element, asdescribed above for radiation exchange betweensurfaces within the cask. The top, side, andbottom temperatures in the tunnel predicted in theNIST fire simulation with FDS are imposed asboundary conditions on the elements modeling thetunnel surfaces. Emissivity values of 1.0 for thetunnel surfaces and 0.9 for the ISO containerexterior surfaces were used, on the assumption that
these surfaces would be severely blackened duringthe fire due to the effect of sooting.
Figure 5.7. Zones for External Heat TransferBetween ISO Container andTunnel
5.2.2 Without ISO Container
For the analysis of the cask without an ISOcontainer, the cask model illustrated in Figure 5.4was connected directly to the tunnel environment.Calculations for convection heat transfer on theexternal surface of the cask were based onempirical relations for convection over cylinders(see Section 6 for full details). Convection at agiven surface was implemented using SURF 152elements, in essentially the same manner asdescribed above for the external surfaces of theISO container.
Similarly, radiation interaction between the caskouter surface and the tunnel was established bycoating all respective interacting surfaces withSHELL57 elements with specified emissivematerial properties. The SHELL57 elements werethen used to produce a highly structured AUX- 12generated MATRIX50 superelement.
The top, side, and bottom temperatures in thetunnel predicted in the NIST fire simulation with
5.5
FDS are imposed as boundary conditions on theelements modeling the tunnel surfaces. Emissivityvalues of 1.0 for the tunnel surfaces and 0.9 for theLWT cask exterior surfaces were used, on theassumption that these surfaces would be severelyblackened during the fire due to the effect ofsooting.
5.3 NAC LWT TransportationPackage Material Properties
The material thermal properties used in theanalytical model (with and without the ISOcontainer) were obtained from the vendor's SAR[7] and are listed in Appendix A. Somemodifications were made to the material propertiesto account for structural configuration changes andexpected effects of the fire. For the aluminumhoneycomb material, the significant void volumereduces the heat transfer capability compared tosolid material. The thermal conductivity assignedto the impact limiters was scaled by the ratio of thehoneycomb density to the solid aluminum density.
Modeling of the liquid neutron shield wascomplicated by the expectation that the 56%ethylene glycol liquid will exceed its boiling pointduring any fire simulation. This can be expectedto lead to tank rupture and vaporization of thecontents, which significantly affects the heattransfer behavior of the cask. Prior to rupture, theliquid in the tank is expected to sustain convectioncurrents due to temperature gradients through theliquid between the tank surfaces. After rupture,
empirical relations were used to obtain separateeffective conductivities for the shield tank andexpansion tank. (Refer to Section 6 for details oncorrelations used in this approach.)
The effective conductivity was determined as afunction of the average tank temperature and theradial temperature difference between the tankinner and outer surfaces. The material propertieswere updated between each time step during thetransient solution using ANSYSO ParametricDesign Language (APDL). The affected nodeswere assumed to consist of a 56% ethylene glycolsolution up to the point where the averagetemperature reached the mixture's boiling point of3500F (177-C).
When the average temperature in the tankexceeded the boiling point, it was assumed thatrupture occurred and the liquid was immediatelyvaporized. The effective conductivity was thencomputed using air as the medium. Thiscalculation was continued during the cool downperiod also. This formulation conservativelyneglects energy absorbed by the phase change(i.e., the heat of vaporization for the liquid), butmainly as a matter of convenience, since thiswould constitute a very small deduction from thetotal energy imparted to the cask. After rupture,thermal radiation exchange within the empty tankswas also activated using MATRIX50superelements.
5.6
6 ANALYSIS METHOD
Analyses have been performed by the NationalInstitute of Standards and Technology (NIST)using various assumptions related to the type offire that could have been sustained in the CaldecottTunnel. Results from the NIST analyses,including temperature and flow predictions for thepostulated fire and post-fire scenario, were used todevelop the boundary conditions applied to theANSYS model of the NAC LWT cask.
Section 6.1 lists the conservative assumptionsunderlying the analytical approach used. Section6.2 describes the boundary conditions derivedfrom the NIST simulation with FDS, and definestheir application to the ANSYS analysis of theNAC LWT cask. This includes temperatureboundary conditions and the approach used todefine convection and radiation heat transfer rates,and the methods used to account for materialdegradation during the fire. Section 6.3 describesthe initial steady-state conditions for the NACLWT cask model, with and without the ISOcontainer, at the beginning of the fire transient.Section 6.4 describes the procedure used for thetransient calculations.
6.1 Modeling Assumptions for FireTransient
A number of conservative assumptions were madein developing models and performing evaluationsof the thermal response of the NAC LWT spentfuel transport package to the Caldecott Tunnel firetransient. The assumptions of greatest impact arelisted below.
1) Boundary conditions were taken from thehottest location within the tunnel, which wasdetermined to be 2034 ft (620 in), which is 328ft (100 m) to the east (downstream) of thelocation of the fire, based on predictions of
peak gas temperatures in the lower, middle, andupper zones of the tunnel, and peak surfacetemperatures and energy fluxes on the tunnelfloor, walls, and ceiling.
2) The peak temperature values in each regionwere used to define boundary temperaturesover the entire region, rather than using thedetailed local temperature distributionspredicted in the FDS calculation. Thisapproach ensures a conservative estimate of theboundary temperatures, since the package doesnot see the peak temperatures on all surfaces,and in some cases may not see the peaktemperature on any surface. (For example, thetop of the package is not high enough to bedirectly exposed to the peak gas temperaturenear the top of the tunnel, but this value wasused as the ambient temperature for convectiveheat transfer to the upper surface of thepackage.)
3) The package cradle and the trailer bed wereomitted from the ANSYS model of the NACLWT package. These structures were neglectedbecause they could partially shield the packagefrom thermal radiation from the hot tunnelsurfaces or block convection heat transfer to thepackage due to the flow of hot gas generated bythe fire. This approach eliminated any potentialshielding of the package from thermal radiationand convection heat transfer from the tunnelenvironment.
4) During the simulated gasoline-fueled fire ( <0.7 hr) and the short-term post-fire cool downperiod (0.7 hr < t < 3.0 hr), it was assumed thatforced convection heat transfer at the outersurface of the package was due solely to airflow induced in the tunnel by the temperaturegradients of the fire. Convection heat transfer
6.1
rates were calculated using the gas velocities atthe locations of the peak gas temperatures, aspredicted in the NIST analysis with FDS. Thisapproach neglects the possible contribution ofadditional heat transfer from the package due tofree convection resulting from verticaltemperature gradients around the package.This boundary condition was switched to solelyfree convection after 3.0 hours, in theextrapolated extended cool down portion of thetransient. This conservatively neglects anyforced convection cooling of the packageduring the extended cool down period, whenthe gas velocities in the tunnel are predicted tohave dropped to relatively small values.
5) Attenuation of thermal radiation during the firedue to optical densification (i.e., smoke andparticulates from combustion and materialdegradation) was not taken into account in thetransient calculation. However, because thefire was reported to have produced thick blacksmoke, it was assumed that the outer surfacesof the package would 'see' the peak gastemperatures for thermal radiation exchange,rather than the tunnel surface temperatures.This provides a conservative treatment of heattransfer due to thermal radiation, since the FDScalculation predicted that the gas temperatureswould be higher than the tunnel surfacetemperatures during and shortly after the fire.For the analysis with the cask within an ISOcontainer, attenuation of thermal radiation wasalso neglected between the cask and innersurfaces of the ISO container. In these regions,radiation views were treated as clear andunobscured at all times during the transient.
6) Materials that would burn, boil off or meltduring the transient were assumed to remainintact during the fire. At the end of the fire, thethermal conductivity values for these materialswere reduced to that of air. The higher thermalconductivity values of the intact material tends
to maximize the heat input into the packageduring the fire. When these values are replacedwith the thermal conductivity of air, theaffected components present an added thermalbarrier to heat removal from the package afterthe fire. In addition, the energy absorbed bythese materials, due to latent heat of fusion orvaporization, was not subtracted from theenergy input to the package from the fire.
Given these assumptions and the extremelydetailed 3-D model of the spent fuel transportationpackage, the ANSYS analyses presented hereconstitute a conservative evaluation of theresponse of the NAC LWT cask to the CaldecottTunnel fire scenario. The boundary conditionsfrom the FDS simulation of the Caldecott Tunnelare presented in Sections 6.2 through 6.4.
6.2 Boundary Conditions for FireTransient
Boundary conditions from the NIST simulationwith FDS were selected from a locationapproximately 328 ft (100 m) downstream of thefire source. This location corresponds to thehottest gas temperatures and highest thermalenergy output of the fire (see the discussion inSection 3 and the plots in Figures 3.13, 3.14, and3.15.) Section 6.2.1 describes the tunnel surfacetemperatures and gas temperatures selected todefine the boundary conditions for the ANSYScalculation. Section 6.2.2 describes the heattransfer boundary conditions applied in theanalysis, based on the gas temperatures andassociated gas velocities.
6.2.1 Boundary Temperatures from FDSAnalysis
Peak tunnel surface temperatures, peak gastemperatures, and associated gas velocities overtime from the NIST simulation with FDS were
6.2
selected from a location approximately 328 ft (100m) downstream of the fire source. As a .conservative simplification of the finely detailedmeshing of the fluid nodes in the FDS simulation,the tunnel air volume was divided into threesections, consisting of an upper, middle and lowerregion. As an additional simplification, theanalysis neglected the shielding effect of caskmounting structures and the trailer bed supportingthe package above the roadway.
The regions were defined based on the geometryof the tunnel and the position of the packagewithin the tunnel. The upper region was definedas the tunnel volume extending from the tunnelceiling to 15.3 ft (4.7 m) above the tunnel floor.The middle region was defined as the volumeextending from 15.3 ft (4.7 m) to 1.0 ft (0.3 m)above the tunnel floor. The lower region wasdefined as the volume between the tunnel floorand 1.0 ft (0.3 m) above the tunnel floor. Thetunnel surfaces in the ANSYS model were dividedinto three: corresponding regions; ceiling, sidewalls, and floor. The upper region consists of theceiling and upper wall to 15.3 ft (4. 7 m) above thetunnel floor. The middle region consists of thetunnel wall from 15.3 ft (4.67 m) to 1.0 ft (0.3 in)
above the tunnel floor. The lower region consistsof the tunnel floor and up the wall to 1.0 ft (0.30m) above the tunnel floor.
Rather than tracking the local surface and gastemperatures, and gas velocities, predicted overthe fine mesh within each of these regions in thedetailed NIST simulations with FDS, the boundarytemperatures used in the ANSYS calculationswere defined by applying the peak temperatureand velocity values in a given region over theentire region. Within a given region, the predictedpeak tunnel surface temperature, peak gastemperature, and associated gas velocity as afunction of time were used to defined theboundary conditions for the entire region.
Using this conservative simplification, boundarytemperatures were specified for the top region,side region, and bottom region of the ANSYSmodel of the package within the tunnel. For theanalysis with the ISO container, the top regionconsists of the upper surface of the ISO container,the side region consists of the three verticalsurfaces of the half-section of symmetry of theISO container, and the bottom region of the modelconsists of the ISO container base.
For the analysis without the ISO container, the topregion consists of the upper 60-degree arc of the180-degree half-section of symmetry of the caskcircumference. The bottom region consists of thelower 30-degree arc of the cask circumference,and the side region consists of the 90-degree arcbetween the upper and lower region. Thisdivision was also applied to the impact limiters.
In clear air, the cask surfaces (without the ISOcontainer) or the ISO container surfaces would seethe tunnel surfaces for radiation exchange.However, during the fire portion of the transient,this view is obscured due to smoke and othercombustion gases filling the tunnel. This meansthat the package would see the gas temperaturerather than the wall temperature for radiation heattransfer. This is significant, since during the fireportion of the transient, the peak gas temperaturesfrom the upper and middle regions of the tunnelare generally 180-270'F (100-1500 C) above thepeak ceiling and wall surface temperatures (as canbe seen from Figures 3.10 and 3.11).
This is represented in the ANSYS simulations byspecifying the gas temperatures rather than thetunnel surface temperatures as the temperaturesseen by the package outer surfaces for thermalradiation heat transfer during the fire. After thefire, the smoke was reported to have cleared outfairly rapidly, so that in a relatively short time, thepackage surfaces would be expected to see thetunnel surfaces. This transition was modeled by
6.3
selecting the boundary temperature as the higherof the tunnel surface temperature or gastemperature for the given region.
troughs related to the general physical behavior ofthe simulated fire.
- nd Rf re12 0OO , I
0n
In practical terms, this means that the radiationboundary temperature for a region switches fromthe gas temperature to the tunnel surfacetemperature very shortly after the end of the fire.Figure 6.1 shows the boundary temperatures forthermal radiation for each region, which wereselected as the maximum of the gas temperature orthe surface temperature for the correspondingregion of the tunnel.
lenvd ffir |
19000.1800 0170016001500.-1400.
C 1300i 12r0
e 0,.0I YEgooo1700
pldn)-(tbo ) .......- (-Wsr) ...
0.00 0.50 1.0 D 1.50 2
Elapsed Mme (hours)LOO 2.50 3.00
Figure 6.2. Peak Temperatures for ConvectionHeat Transfer During FireTransient in Caldecott Tunnel
'_I:a '
'±
lend o firet
0.50 0.50 1.00 1.50
Elapsed lime (hours
I . .zoo 2.50 3.o0
i
I I
12 ......... L ..... a''''''' 1 ......... .............. I................
124. I L..4...........
........ ....... .. ................ ..............8I I4 ...............I............... ................ ..............i.Figure 6.1. Peak Temperatures for Radiation
Exchange During Fire Transient inCaldecott Tunnel
I, ....1.....-o , ............... ,.. ... & : .
000 0.0 100 1.00 zo0
Elapsed lime (hours)2.00 3oo
The boundary temperatures for convection heattransfer in each region is shown in Figure 6.2. Inall regions, this temperature is the correspondingpeak gas temperature from the NIST calculationwith FDS. The gas velocities used in each regionare also taken from the NIST calculation, at thelocation of the corresponding peak temperature.These velocities are shown for each region inFigure 6.3. These temperature-vs.-time andvelocity-vs.-time values used as boundaryconditions in the ANSYS calculation weresmoothed to conservatively remove the rapidstochastic variations typical of dynamic firebehavior, preserving only the major peaks and
Figure 6.3. Peak Velocities for ConvectionHeat Transfer During FireTransient in Caldecott Tunnel
The FDS analysis performed by NIST was carriedout for a 40-minute gasoline-fueled fire and 2.3-hour post-fire cool-down, for a total simulationduration of 3 hours. To determine the completetime and temperature response of the package, andexplore the effects of prolonged exposure to post-fire conditions in the tunnel, the ANSYS analysisextended the post-fire cool down to 50 hours.Tunnel surface and gas temperatures predicted
6.4
with FDS at 3 hours were extrapolated from 3hours out to 50 hours using a power function, torealistically simulate cool down of the tunnelenvironment.
Figures 6.4 and 6.5 show the boundarytemperatures for radiation and convection heattransfer, respectively, extrapolated from 3 hoursout to 50 hours. The extrapolation was performedby fitting a power function to the post-fire portionof each of the boundary temperature curves from.the FDS simulation, such that
Tn =an tba
whereTn= extrapolated boundary temperature of
region nan= leading coefficient of regression fit to
boundary temperature curve nbn= exponential coefficient of regression
fit to boundary temperature curve nt = elapsed time
le lend fir 51
convection heat transfer boundary at the packagesurface is switched from forced convection to freeconvection only. By 50 hours, the extrapolatedboundary conditions predict that the peak gastemperatures and surface temperatures in thetunnel will be back to the normal tunnel ambientair temperature of 680F (200C), and all boundarytemperatures are essentially constant.
2900 -- 'l l, l , .
17W ....... ...... ..... ..... ... .. ... .. ..... ..... ...... ........'SW ... .. '...... ...... '.......... ..... ,... tt..... - ( p I18o......0 i0 f i_(xp
''-(b(d.) .......
,- ,(botm.......C£
-12i ElD
0 5 10 15 20 25 30 35 40 46 50
Elapsed Time (hous)
Figure 6.5. Peak Temperatures for ConvectionDuring Extended Transient inCaldecott Tunnel
-(op) 1 ].(W.) . ..........
£120. 12
61II1I
10-
0 0 10 10 20 25 30 35 40 45 51'
Elapsed Tirme (hours)
6.2.2 Convection Boundary Conditions
The NIST analyses with FDS show that thethermal gradients created by the fire would resultin significant air flow past a body located in thetunnel downstream of the fire. This fire-forcedconvection would significantly affect heat transferaround the LWT cask or ISO container, and have astrong influence on the rate of increase of theoutermost surface temperatures of the package.The regional peak gas temperatures shown inFigure 6.2 and associated velocities shown inFigure 6.3 were used to define local time-dependent Nusselt number values on the surfacenodes corresponding to the upper, middle andlower regions of the package. The correspondingheat transfer coefficient is used to calculate thelocal convection heat at the package surface.
Figure 6.4. Peak Temperatures for RadiationExchange During ExtendedTransient in Caldecott Tunnel
By three hours into the transient (2.3 hr after theend of the simulated fire), the predicted gasvelocities for forced convection have dropped toless than 2 ft's (0.6 m/s). At that time, the
6.5
To maintain consistency between the two models,the same Nusselt number correlation was used todefine convection heat transfer at the cask surface(without the ISO) as for the analysis of thepackage within an ISO container. For both cases,the Nusselt number at the outer surface of thepackage was defined using the followingrelationships for gas flow over flat or slightlycurved surfaces at zero angle of attack [10];
for laminar flow (ReL < 500,000):
NUL = 0.665 Re' 2 Pr" 3
for turbulent flow (ReL > 500,000):
NUL = 0.032Re" Pr" 3
With the ISO container, the characteristic length,L, used in the ANSYS model to define the Nusseltnumber and Reynolds number for this applicationwas the horizontal ISO container wetted surfacelength (i.e., 240 inches). For the case without theISO container, the axial characteristic length wasdefined as 232 inches, based on the length of theexposed package body. A characteristic length of65 inches was used for the vertical surfaces of theimpact limiters on the ends of the cask.
The peak gas temperature predictions from theNIST analysis define the ambient sinktemperatures around the package during the firetransient and post-fire cool down period. TheNusselt number defines the rate of heat transferfrom the package, which is used in ANSYS tocalculate the local convection heat flux at the outersurfaces. Using the one of the above relationshipsfor Nusselt number (depending on the geometrybeing modeled and the hydrodynamics of the airflow), the code solves for local surfacetemperatures, T3, and calculates the convectioncomponent of the heat flux at the surface using theformula
q 11W= NUL L (Ts- Tarf)L
where k = thermal conductivity of ambient airL = characteristic lengthTs = cask surface temperature
TaLi = ambient external air temperature.
Separate boundary types were defined for the top,sides, and bottom surfaces of the package usingthe external air temperatures shown in Figure 6.2.The velocities in Figure 6.3 were used to definethe Nusselt number so that the boundary condi-tions on the cask would change with time.
By the end of 3 hours, the gas velocities predictedin the NIST calculation are down to 1 to 2 ftl/s (0.3to 0.6 m/s) or less (see Figure 3.15). Heat transferat the package surface for these flow conditions isa complex mixture of forced convection (due to airflow induced in the tunnel by the temperaturegradients of the fire) and free convection (drivenby the non-uniform circumferential temperaturesaround the package outer surface).
At velocities below about 3-5 ftWs (1 to 1.5 m/s),heat transfer rates predicted assuming forcedconvection are generally lower than heat transferrates due to natural convection for thetemperatures on and around the surface of thepackage. To avoid the modeling uncertaintiesassociated with mixed-mode heat transfer, forcedconvection only was assumed until the end of theNIST simulation, at 3 hours into the transient.From 3 hours to 50 hours, the heat transfer wasassumed to be natural convection only. Thecontribution of free convection at the packagesurface is ignored in the cool down from 0.7 to 3hours, and the contribution of forced convection isneglected in the cool down period from 3 to 50hours. This ensures a conservative treatment ofconvection heat transfer from the package surfaceduring the entire calculation.
6.6
For consistency, the natural or buoyancy-drivenconvective coefficients were those utilized todetermine the pre-fire component temperaturedistributions (i.e., Normal-Hot Conditions ofTransport, as defined in 10 CFR 71.71[l].) Theheat transfer coefficients were defined for theappropriate surface geometries using the followingrelationships [11,12,13]:
--for flow along a vertical plane or cylinder:
--laminar flow (104 < Grf Prf < 10)
h = 1.42(-)(L
--turbulent flow (Grf Prf > 109)
h =1.31(AT)"'
whereh = heat transfer coefficient, W/m2 oCAT = TW - TX, OC2, = surface or wall temperature, 0CZOO = ambient temperature, 0CL, = vertical or horizontal dimension, mGrf = Grashoff number of the gas at
film temperature; Tf = (TW + T/2Pry = Prandtl number of the gas at
film temperature
--for flow over a horizontal heated plate facingupward (cool side facing downward):
-- laminar flow (104 < Gry Prf < 09),
h~.2AT 1/4It =1.32(
-- turbulent flow (Gr, Prf > 109),
h =1.52(AT)11
--for laminar flow (104 < Gr1-Prf < 10) over aheated plate facing downward (cool side up):
ht =0.59(L
Definitions of material properties used to computeGrf1 Prf for use with these correlations were takenfrom Table A-3 of Kreith [13].
An empirical relationship for effectiveconductivity incorporating the effects of bothconduction and convection was used to determineheat exchange through the liquid neutron shield.In the SAR [7] analysis for the LWT cask, theeffective conductivity of the ethylene glycolmixture for conditions below 350TF wasdetermined using the correlation of Bucholz [14].This correlation defines the ratio of the effectiveconductivity to the actual thermal conductivity asequal to the Nusselt number, such that
keff = Nu = 0.135(Pr 2 Gr/(1.36+ pr))0.218kc
where keff = effective thermal conductivity ofmaterial in node
kc = thermal conductivity of motionlessfluid in node
Pr = Prandtl numberGr = Grashoff number.
--for flow over a horizontal cylinder:
--lamrrinar flow (104 < Grf-Prf <IO),
h AT14( )
where
a' = diameter, m
-- turbulent flow (Grf Prf > IO),
h =1.24(AT)" 3
6.7
The tunnel fire transient is outside the range ofapplicability of the Bucholz correlation, and ityields unrealistically large values for keff for theseconditions. An alternative correlation fromRaithby and Hollands [10], based on heat transferbetween concentric cylinders, was used in thisanalysis instead. This correlation producesreasonable values of kff, and the transientconditions are generally within its applicablerange. In this correlation, the Nusselt number isexpressed as
keff = Nu =0.386Dr(Pr/(0.861 + Pr))0 25Ra0.25
kc
where Rayleigh number (Ra = Pr*Gr) is based onthe temperature difference across the annular gap.
The dimensionless parameter Dr is defined:
D [ n(D0 /Di) 1Dr = 11~4 (I/D 315 + I/D"' )514]
where D. = annulus outer diameterD= annulus inner diameterd = width of annulus.
Figure 6.6 shows a plot of the Nusselt numberpredicted with these two correlations for the liquid(56% ethylene glycol and water mixture) in theneutron shield annulus.
Figures 6.7 and 6.8 show the effectiveconductivity for the annulus as a function of theaverage temperature and temperature differencefor the liquid neutron shield tank and expansiontank, respectively. (The sharp discontinuity in thecurves on both plots represents the abrupt phasechange assumed when the average temperature ofthe liquid reaches the boiling point of the ethyleneglycol and water mixture.) For low values of thetemperature difference, the results approach thosefor conduction-only conditions.
F, 6
5
4
zI
Z2
-- EbChdz (fro SAR) I I I'0 -- Ralthby & Rdands-l 975 -(GWr)…*
-V at Ratn C8p OrPera & D.W1it) I I 7 P
h -_tN.htrotSAR -4d -tF- - - -* RaithubY&H6olandsbrfr a..aIentdTs I 1
I I I I I I'D - - - - r T _- -T - - - 7 - - - 7 dTr-25 F -
0 L-__… - - _-…-I I I I I /
30 - - - - -- - - -…………- -- -- -t- -I I I I I I /
I , d7.1F I20 ---- --n-- i -- -- i-,-In-
I I 1 1 1I 1 10 dT,…
I I _I0_
11
1.E-02 1.E-03 1.E-04 1.E.0S 1.X.06 1.E+07 12.08 1.6E09 1.E.10
Pr-Gr (unitless)
Figure 6.6. Nusselt Number for Heat Transferin Liquid Neutron Shield
1.E-02,
1.E.01
I 1.E600
1.E-01
I .--dT.500I I I'-* dT:300
_ I I I I- dT.7II IdT.25
4 ---dT.100.
- -. 00.70
_ -
I I I II I I I
2 0 700 1200 1700
Awrage AnnuluS Temperature (F)
2700
Figure 6.7. Effective Conductivity of LiquidNeutron Shield Tank
1.E-02
.I I dT.500I -- dT . 300
-- dT - 2001.E-01 - - - 4- - - 4-- -- -- I-- - -- - --- --- -- dT .100
-- dT70. =| -- so.50
, _ I -- dT 25
t 1.E-00 - -- - - - - - - - - - - - - - - - - dT.10
I1.E01_ - - - - - - - - - - -- - - - - - -
1.E-02
200 700 1200 1700
Average Annulus Temperature (F)2200 2700
Figure 6.8. Effective Conductivity of LiquidNeutron Shield Expansion Tank
6.8
6.3 Initial System ComponentTemperatures
The hot normal conditions for transport were usedas initial conditions for these analyses. A heatgeneration rate equivalent to a decay heat load of8,530 Btu/hr (2.5 kW) was applied, withappropriate peaking factor, over the active fuelregion. For the analysis of the cask without anISO container, free convection at the packagesurface is handled by SURF 152 elements with aconstant heat transfer coefficient of 0.891 Btulh-*-OF (0.157 W/m2 -oC) and an ambienttemperature of 100I F (380C). For the analysiswith the ISO container, the natural convectioncorrelations for buoyancy-driven flow discussed inSection 6.2.2 were used to simulate convectionheat transfer between the outer surface of the caskand the inner surface of the ISO container, andbetween the outer surface of the ISO container andambient air.
For both analyses, with and without the ISOcontainer, solar insolation (i.e., radiation) isincorporated by using SURF 152 elements withheat generation on the outer surface of thepackage, at the rate specified in IOCFR71 [1]. Forpre-fire conditions, the emissivity of the casksurface or ISO container surface was specified at avalue representative of the local surface finish(e.g., 0.3 for bare stainless steel, 0.85 for paintedsurfaces.
The steady-state temperature distributionspredicted in the cask to define the initialconditions for the fire transient calculations wereverified by comparison with results reported in CieSAR [7]. Direct comparison is not possible,because the SAR [7] does not include anyanalytical cases similar to the detailed 3-D modelsused in this study. Because the main concern inanalyses -For normal transport conditions is todetermine a conservative rate of heat removal
from the cask, the applicant chose to perform aseries of highly conservative evaluations usingrelatively simple models to qualify the system forits Certificate of Compliance (CoC).
The most complex models presented in the SAR[7] involve simple 2-D ANSYS cross-sections inwhich the cutting plane includes the expansiontank as well as the neutron shield tank. Thisapproach does not allow axial heat flow out of theplane of the 2-D cross-section, and also assumesthat the decay heat load axial peak occurs on thatplane. This assumption places the spent nuclearfuel peak decay heat location under two concentrictanks filled with neutron shield material. Thisprovides conservatism for a steady-state analysis,since the expansion tank makes a longerconduction path over which to dissipate the decayheat. For the fire transient, however, theassumptions in this 2-D model would limit theheat input to the cask from the fire, and would notconstitute a conservative approach.
In the SAR [7], ANSYS cross-sectional modelswere also used to represent a 25-rod BWR basketassembly at 1.41 kW and a high burn-up PWRassembly at 2.1 kW. These models includeddetailed representation of the fuel pins, pin tubes,and can weldments with the pins resting on the pintubes via point contact. These models alsoincluded the ISO container, with solar insolationand 1000F (380 C) ambient temperature.
The design basis results presented in Amendment34 of the SAR [7] for a 2.5 kW PWR assemblyalso used a 2-D model of the cask. This is aHEATING5 model, with a 2-D axisymmetricrepresentation using effective diameters for thebasket and fuel assembly. This model neglects theISO container and impact limiters, and the 2-Dmodel cannot account for conduction andconvection at the assembly end cavities. Theambient temperature boundary condition for theseanalyses was specified as 130'F (54 0C).
6.9
The results reported for these three cases aresummarized in Table 6.1. As might be expected,the conservative 2-D ANSYS models predictrelatively high temperatures, compared to theresults obtained with the more detailedHEATING5 model. Of these three cases, only theHEATING5 analysis at 2.5 kW is sufficientlyclose to the initial steady state conditions assumedfor the fire transient to allow reasonablecomparisons to be made for verification of the 3-DANSYS model predictions.
Table 6.1. NAC LWT ComponentTemperatures at Various DecayHeat Loads
2.5 kW 141 4kW 2.1 kW0F O V0 (C) OFF(CC) F(C);(Table Table Table
Component 3A-2 171 _3.4-7 171) 3A-10 171)Fuel Cladding 472 (244) 358 (181) 671 (355)Aluminum PWR 276 (136) * 394 (201)InsertInner Shell 274(134) 249 (121) 385 (196)Gamma Shield 273 (134) 248 (120) 375 (191)Outer Cask Surface 229 (109) 185 (85) 308 (153)Neutron Shield 238 (114) 235 (113) 306 (152)Lid Seal 227 (108)Drain/VentPorts 231 (111 * *Impact LimitersISO Container* value not reported by applicantFor the purpose of this comparison, additionalcalculations were performed with the 3-D ANSYSmodel, with and without an ISO container, usingan ambient temperature boundary of 130'F (540C)at 2.5 kW decay heat load. (These calculationswere performed in addition to the cases at 100WF(380C) ambient temperature, which provided theinitial conditions for the fire transient calculation.)
Figure 6.9 shows the predicted temperaturedistribution from the ANSYS solution for this casewith the cask in an ISO container. Table 6.2presents detailed component temperature resultsobtained with the 3-D ANSYS model analyses,compared to the values published in the SAR [7]
for the HEATING5 model at this decay heat loadand ambient temperature boundary condition.
ANOYS a.0JUN 16 200515:50:30ODAL SOLUTION
STEP-ISUB -ITINE-.100E-03TEMP (AVG)RSYS-0P-SPGraph8USEFACSET-1Av~rf-44.tSPIN -1308M0 -432.855
130
148a..2186.7851
___ 224.642__243.571
262.499281.428
ER 300.3S6319.285
394.999:193.927432.855
Figure 6.9. LWT Cask (with ISO Container):Normal-Hot ConditionTemperature Distribution (2.5 kWDecay Heat, 1300F Ambient)
At first glance, the temperatures presented inTable 6.2 appear to show rather large differencesbetween the results obtained with the two models.The peak clad temperature predicted with theANSYS 3-D model is 4340 F (2230C), compared to4720F (2440C) reported in the SAR for theHEATING5 model [7]. Other componenttemperatures shown in the table are also lower forthe 3-D ANSYS model results, compared to thecorresponding SAR values. However, this is anexpected result, given the modeling differencesbetween the two cases. The 2-D cross-sectionrepresenting the cask in the HEATING5 modelshould result in more conservative predictedtemperatures, compared to the 3-D ANSYS model.
A more significant observation for the purposes ofthis comparison is that the differences in peakcomponent temperature between the twoapproaches are consistent. The radial temperaturedrop from the peak fuel cladding temperature tothe outer cask surface temperature is 2340 F(1300C) for the ANSYS 3-D model, compared to
6.10
the predicted temperature drop of 2430F (1350C)with the HEATING5 axisymmetric model. Thisclose agreement strongly suggests that bothmodels provide a similar representation of theradial heat transfer paths from the fuel cladding tothe environment. The differences in specifictemperature values predicted are due mainly todifferences in model complexity. The 3-Drepresentation in the ANSYS model accounts foraxial as well as radial heat transfer paths, whichthe 2-D HEATING5 model specifically excludes.
Table 6.2. NAC LWT ComponentTemperatures at 2.5 kW DecayHeat Load and 1301F Ambient
Curent SAR^ 0~>9 x; l SStudy,:Values
-ith4ISO (Table0 : .SYS) (3.4-24171) AT
Fuel Cladding 434 (223) 472 (244) 38 (21)
Aluminum PWRInsert 265 (129) 276 (136) 11(6)Inner Shell 228 (109) 274 (134) 46 (26)
Gamma Shield 227 (108) 273 (134) 46 (26)
Outer CaskSurface 200 (93) 229 (109) 29 (16)Neutron Shield 204 (96) 238 (114) 34 ([9)
Lid Seal 164 (73) 227 (108) 63 (:35)
Drain/Vent Ports 164 (73) 231 (111) 67 (37)
Impact Limiters Not167 (75) Modeled
SO Container 167 (75) Not
Modeled
conditions for these calculations were specified as1000F (380C) at 2.5 kW with solar isolation,corresponding to Normal Hot Conditions ofTransport as described in 10 CFR 71.71 [1].
aIsTs 8.0JUN 16 200515:s1:36NODAL SOLUTIONSTEP-ISUB -ITIME-. 180E-03TEMP (AWO)PSYS-0Po-.rtraphicsEFACET-1AVRES-Mat8187 -100SPEK -417.459
139.682is 9. 524179.36S199.206
__219.047__2386.888
cm 258.129EJ278.571
298.4i1
E 318. 253__ 38 094
3 193S3577.76397.6718417. 459
Figure 6.10. LWT Cask (with ISO Container):Normal Condition TemperatureDistribution (2.5 kW Decay Heat)
21±46.22.
34 -00
170.544
241.093276.366
427.459
Figure 6. L0 shows the temperature distribution forthe NAC LWT package within an ISO container,predicted with the ANSYS 3-D model for theinitial steady-state conditions before the firetransient. Figure 6.11 shows the temperaturedistribution predicted for the NAC LWT packagewithout an ISO container. The boundary
Figure 6.11. LWT Cask (without ISOContainer): Normal ConditionTemperature Distribution (2.5 kWDecay Heat)
The pre-fire steady-state component peaktemperatures predicted with the ANSYS 3-D
6.11
models are shown in Table 6.3. Thesetemperatures are somewhat lower than thosereported for the 3-D model in Table 6.2, due to thelower ambient boundary temperature, but thetemperature distribution is essentially identical.
Table 6.3. NAC LWT Pre-Fire ComponentTemperatures at 2.5 kW Decay
Heat Load and 100OF Ambientwithout ISO withlSO
(ANSYS) X (ANSYS)Component 0 F (CC) 0 F (6C)
Fuel Cladding 399 (204) 418 (214)luminum PWR
Insert 212 (100) 242 (117)
Inner Shell 172 (78) 205 (96)
Gamma Shield 171 (77) 204 (95)Outer CaskSurface 138 (59) 176 (80)
Neutron Shield 144 (62) 180 (82)
Lid Seal 121 (49) 138 (59)
DrainNent Ports 120 (49) 138 (59)
Impact Limiters 122 (50) 141 (61)
SO Container N/A 140 (60)
phase consists of the long-term post-fire cooldown, using extrapolated boundary conditions andextending from 3 hours out to 50 hours.
In the first phase of the calculation, the firetransient was initiated from the steady-stateconditions by setting the solar insolation to zero,adding the elements and appropriate thermalconnections comprising the model of the tunnel,and introducing the boundary conditionsrepresenting the fire. The transport package andtunnel surfaces were assign emissivities of 0.9 and1.0, respectively, to represent surfaces affected bysooting.
For the first phase of the transient (0 < t < 0.7 hr),during the intense, gasoline-fueled fire, a forcedconvection regime was assumed to exist on theexterior of the package, with the surface heattransfer coefficient calculated based on the gasvelocity predictions from the FDS analysisperformed by NIST. With the gas temperaturesfrom the NIST analysis defining the ambientboundary temperature, the convective heat flux atthe package surface could be determined in thesolution for the local surface temperature. Heattransfer due to thermal radiation was alsoincluded, with the source temperature for radiationexchange defined as the maximum of the tunnelwall temperature or tunnel gas temperature, toconservatively take into account the effects ofoptical densification due to smoke and othergasses released as a result of the fire.
As an additional conservatism to maximize theheat input to the package from the fire, thealuminum honeycomb impact limiters wereassumed to remain intact during the fire. The heatconduction paths into the cask provided by theimpact limiters were therefore maintained at thehigher value corresponding to the aluminumhoneycomb long after the predicted temperaturesindicated that this material would have beendestroyed or degraded by the fire. At the end of
6.4 Tunnel Fire Transient
The Caldecott Tunnel fire transient simulation forthe NAC LWT transport package consists of threephases. The transient calculation is initiated fromthe steady-state conditions described in Section6.3 (with or without the ISO container) for normalhot conditions, assuming insolation and 100'F(38 0C) ambient temperature, as per 1 OCFR71.71[1]. The first phase of the transient consists of theintense, gasoline-fueled fire, lasting approximately40 minutes. The second phase consists of theshort-term post-fire cool down, extending from theend of the fire (at 40 minutes) out the end of theNIST simulation with FDS, at 3 hours. The third
6.12
the fire, the properties of the nodes representingthis material were replaced with thermal propertiesof hot dry air.
Similarly, the ethylene glycol and water mixture inthe neutron shield tanks was assumed to remain inplace until the average temperature of this regionexceeded the boiling temperature of the liquid(350TF (1 770 C)). At that point, it was assumedthat the liquid was replaced by hot dry air.
In the second phase of the analysis, the post-firecool down from the end of the fire (at 40 minutes)to the end of the FDS simulation (at 3 hours), theconvective heat transfer at the package surfacewas assuned to consist of only forced convection,based on predicted gas velocities and temperatures
from the NIST analysis with FDS. In the thirdphase of the analysis, the post-fire cool-down wasextended from 3 hours out to 50 hours (49.3 hoursafter the end of the fire.) The boundary conditionsfor the additional 47 hours of the transient wereobtained from the temperatures and velocitiespredicted in the FDS analysis, extrapolated to 50hours using a power function (as discussed inSection 6. 1.) In this phase of the transient, theboundary condition at the package surface wasswitched from forced convection to freeconvection.
Results obtained using the ANSYS models of theNAC LWT cask (with and without an ISOcontainer) are discussed in Section 7.
6.13
7 ANALYSIS RESULTS
Due to temperature limits for the spent fuelcladding, closure seals, impact limiter materials,and neutron shield materials, these componentsare the most important elements to consider inevaluating the response of the transport systemsto the fire scenario. The peak clad temperaturelimit is important because the cladding is theprimary lission product containment boundaryfor the spent fuel. The temperature limit for theclosure seals is important because these sealsconstitute the outer-most containment boundaryfor the cask. The temperature limits for theneutron shield material and impact limiters areimportant because these materials are generallythe most vulnerable to damage or destructionduring the fire. The results of the analyses of theNAC LWT package are evaluated primarily inrelation to the peak predicted temperatures forthese components in the fire transient.
The ANSYS model of the NAC LWT packageconsists of 52,446 standard computationalelements and 16 superelements that are solvedeach time step. Calculations with this modelyield detailed temperature distributions that canbe analyzed to characterize the cask response tothe specified boundary conditions. The systemresponse predicted for the NAC LWT packagewith ANSYS for the fire transient conditions ispresented in the following three subsections, forthe three phases of the transient, as outlined inSection 6.4.
Section 7.1 presents the predicted response forthe first phase, which consists of the intensegasoline- fueled fire (i.e., the first 40 minutes ofthe transient.) Section 7.2 presents results forthe second phase of the transient, which consistsof the short-term post-fire cool down. Thisphase extends from the end of the fire (at 40minutes) to the end of the NIST simulation with
FDS (at 3 hours.) Section 7.3 presents resultsfor the third phase of the transient, whichconsists of the long-term post-fire cool downfrom 3 hours out to 50 hours, using boundaryconditions extrapolated from the cool downportion of the FDS simulation.
7.1 NAC LWT Package Responseto Fire Transient
Figure 7.1 shows the temperature response forthe NAC LWT cask and ISO container predictedwith ANSYS for the first hour of the transient.Figure 7.2 shows the temperature response forthe NAC LWT cask without an ISO containerfor the same boundary conditions. This timeinterval encompasses the intense gasoline-fueledfire, which lasted approximately 40 minutes,plus the first 20 minutes of the post-fire cooldown period.
For both cases, the temperature response is verynearly identical during this time interval.Without the ISO container, temperatures of out-board components (i.e., cask surface, vent/portseals, and impact limiters) rise somewhat fasterand reach slightly higher peak temperaturesduring the fire. However, the differences arerelatively small, and in both cases, the caskpackage exhibits essentially the same responseto the fire. In both cases, most components reachtheir peak temperature values during thisinterval, closely following the high boundarytemperatures during the fire and their rapiddecrease once the gasoline is consumed.
This behavior is due to the relatively lowthermal inertia of the package, because of itsrelatively small physical size. Direct conductionpaths into the cask are relatively short, and itssurface-to-volume ratio is relatively large.
7.1
Without the ISO container, thermal radiationheat transfer views include a large portion of thecask surface, due to the horizontal orientation ofthe package within the tunnel. Similarly, thesurfaces of the ISO container have essentiallyone-to-one views of the tunnel ceiling, walls,and floor.
1!
I-
l l l -Fua1l ie~r
law0 - - - - - - - -I_ - - - - - - - - - - -Cark Irmrer Surta-I ~-Cask Wuar Surr.s
1400 -9t -I - - - - - - - _ _-Drair~eM Sris
1200__I w; o__ -Erd LeadI / I / I I -ISO Caoarer
80 I-- ODO__--An - - - - - -
400 - - ---- --
o0 0.2 OA 4 O 0.8 1
Elapsed Thie (hours)
Figure 7.1. NAC LWT Cask (with ISOContainer): ComponentMaximum TemperatureHistories During Fire Transient
hr. just before the end of the fire. For the casewith the ISO container, the maximumtemperature on the exterior surface of the cask isonly 16940F (9230C), reached at 0.67 hr. This isbecause the ISO container acts as a thin thermalshield, protecting the cask surface from thedirect radiation view of the fire. The ISOcontainer itself reaches a peak temperature of17730F (9670C) at about 0.67 hr. This value ishigher than the peak temperature on the casksurface for this simulation, but is still somewhatlower than the peak temperature on theunshielded cask, in the case without the ISOcontainer.
These results show that the ISO container acts asa heat shield for the cask during the intensehigh-temperature portion of the fire, loweringthe peak temperature on the cask surface byabout 1590F (880C). Similarly, the maximumtemperature on the impact limiters is about18370F (10030C) without the ISO container,compared to 1714'F (9340C) obtained with theISO container, a difference of about 123OF(680C). This effect is also seen in the maximumtemperature for the drain and vent port seals.Without the ISO container, the nodesrepresenting this component reach a peak valueof 12870F (6970 C) by the end of the simulatedfire, compared to 10350 F (5570C) in the casewith the ISO container.
For both cases, with and without the ISOcontainer, the peak temperatures of the caskinner shell material, lid seal, and the lead gammashielding layers show a more gradual increaseduring the fire, and the temperatures of thesecomponents continue to rise after the end of thefire. At the end of the first hour of the transient,the peak temperature predicted for the cask innersurface has reached approximately 4000 F(2040() for the case with the ISO container, andis at about 6850F (3630() for the case withoutthe ISO container. In both cases, this
-Fud Oddog (Oku~) -Cask Or S.usce -Cask Our Suh r-k,,pad Limier -Lid Ceal -anlme Pot Sca5s0-Ck Cody Led -End Lead
1 r - r1400 I I- t/3~~|- -| l~- s-I
400
2 0 - - -- - - - -C S I I''/ j I I 1 ' I
0 51 .02 03 4 0.5 0.6 .7 08 09 1
Eapwd lme (bours)
Figure 7.2. NAC LWT Cask (without ISOContainer): ComponentMaximum TemperatureHistories During Fire Transient
Without the ISO container, the maximumtemperature on the exterior surface of the caskreaches a peak value of 1853 0F (1012CC) at 0.65
7.2
temperature is still rising at the end of the firsthour of the transient, which is approximately 20minutes after the end of the gasoline-fueled fire.
For the cask within the ISO container, thetemperature of the lead layer within the caskbody is predicted to reach a maximum of 870TF(4660C) at 0.9 hr elapsed time (i.e., about 15minutes after the end of the simulated fire.)Without the ISO container, this temperaturepeaks somewhat earlier, just after the end of thefire at 0.7 hr, at the somewhat highertemperature of 1031 F (5550C). Similarly, thetemperature of the lead layer in the cask bottomis predicted to reach a maximum of 952TF(511 C) at 1 hr elapsed time for the cask withinthe ISO container, while the peak is 1061TF(5720 C) for this component in the case with noISO container.
For both cases, these peak temperatures in thelead shielding are considerably above theestablished operating limit of 600TF (316'C)reported in the SAR [7] for this material. Thissuggests ithat there could be melting andslumping of the lead as a result of the fire.However, for the purposes of the thermalanalysis, the lead is assumed to remain intact, topreserve the good conduction path through thematerial and conservatively maximize heat inputto the cask during the transient. Also, as afurther conservatism, the energy that would beabsorbed by the material phase change in theprocess of melting the lead is not subtractedfrom the thermal load imposed on the cask bythe fire.
Because of the relatively low boiling point of theethylene glycol solution in the neutron shieldtank and overflow tank, the liquid is expected toboil off as part of the cask response to the firetransient, with or without the ISO container. Asdescribed in Section 5, the predictedtemperatures in the main tank and overflow tank
were monitored throughout the transient solutionto determine the predicted time of rupture andevaporation. Consistent with the standard fireanalysis included in the SAR [7], the tanks wereassumed to rupture when the predictedtemperature exceeds the ethylene glycol boilingpoint of 350'F (1770 C). As an additionalmeasure of conservatism, to further maximizethe heat input to the cask during the fire, tankrupture was assumed to occur only after theaverage ethylene glycol temperature exceeded350TF (1770C), rather than at the point when thepeak temperature reached this value.
Using this criterion, the ANSYS analysis for thecase with the ISO container predicts that theouter expansion tank would rupture atapproximately 13 minutes into the fire, and theinner tank would rupture at about 18 minutes.For the case without the ISO container, thistransition occurs slightly earlier, at about 10.5minutes for the outer tank and 13 minutes for theinner tank.
In both cases, basing the times of rupture for thetwo tanks on the average temperature rather thanthe peak temperature delays rupture to a slightlylater point in the transient than would bepredicted based on the peak temperature. Theeffect of this assumption is to increase the heatinput into the cask due to the fire, by extendingthe time interval that the relatively highconductivity ethylene glycol remains in thetanks. Following rupture, the effectiveconductivity of the expansion tank decreasessignificantly as a result of the expulsion of theethylene glycol volume, which is assumed to bereplaced with air. Cooling effects associatedwith this boiling process are neglected in theheat transfer solution. However, the calculationfully accounts for thermal radiation between thehot walls of the empty tanks.
7.3
The temperature response of the fuel cladding isthe slowest of all components in the cask, due tothe significant thermal inertia of the fuel, andbecause it has the longest heat transfer path tothe fire. For the case with the ISO container, thepredicted peak fuel cladding temperature hasincreased by only about 51F (2.80C), at the endof the gasoline-fueled fire. For the case withoutthe ISO container, the increase is slightlysmaller, only about 3.70F (20C). However, inboth cases, the rod surface temperatures areincreasing along the entire length of theassembly, as a result of the ends of the fuelassembly being exposed within the open cavitiesat the top and bottom of the cask. This isillustrated in Figure 7.3 for the calculation withthe ISO container. (The results for thecalculation without the ISO container arevirtually indistinguishable at this point in thetransient, and therefore are not shown in aseparate plot.)
IDE-2.6"
SW B 0E
I~10 i006: W:4
Mas 12W.8I
~~625
portion of the transient (i.e., during thesimulated fire) the peak fuel temperature occursat the center of the assembly (see Figure 7.3).
7.2 NAC LWT Package Short-Term Post-Fire TransientResponse
Figure 7.4 shows the peak temperaturespredicted for components of the cask within theISO container during the first three hours of theANSYS transient simulation. Figure 7.5 showsthe peak temperatures for these componentspredicted for the cask without the ISO container.In both cases, the cladding peak and averagetemperatures continue to rise after the simulatedfire, due to the severe temperature environmentin the tunnel. The ambient conditions in thetunnel immediately following the simulated firesignificantly retard the rate at which the fueldecay heat can be removed from the cask.
2000 -F -- Dronos- - , - Du-ration (NOTp~j
I Cfsk In.r Su£4ce1800 - -----……-- -- Ce k O.W S.1
1400…-l - - - - - - - -It. -Lid S.,J
0DmrakNed S.W..I Cask Body Led
10DO -- - -- - - -- - -End leed 4IScolft-a
-5-00… ---t--*--
20
0 3Elapsd Time (hours)
[L-1 Ndd* I-11, ntun CNA id 29M 2.520 11..5245
Figure 7.3. Lumped Fuel AssemblyTemperature Distribution 0.7 hrinto Transient
In both cases, the ends of the fuel rods directlysee the inner shell and ends of the cask. As theinner shell surrounding these ends heats up,thermal radiation exchange within the cavitiestransfers heat directly to the rod array. For this
Figure 7.4. NAC LWT Cask (with ISOContainer): MaximumTemperature Histories for First 3hours of Fire Transient
Once the simulated fire is over, however, thepredicted peak temperatures on outboardcomponents begin to drop rapidly. As noted inSection 7.1 for the calculations with and without
7.4
the ISO container, the cask outer surface, theimpact limiters, and drain and vent port sealsreach their peak temperatures by the end of thesimulated fire, and the lead gamma shieldingcomponents reach their peak temperatureswithin the first hour of the transient. The peaktemperatures on the cask inner surface and thelid seal reach their respective maximum valuesat a slightly later time; approximately 1.7 hrsinto the transient for the case with the ISOcontainer, and at about 1.3 hrs in the casewithout the ISO container.
2000I I I I -FNd aiding (Peak)
1800~~ - - - ~~~t - - - - - - - - - - Cask Imws thlecacZ , , , -Cask W. rS.%r~
1tiOO _ 4 _ 4_ i_; 4 Impact LWntRs
l4] -CekoF.*alt4'00 -vf- - - E --L- - - - I - - - - Z-c sEd en LR.a
C IIf I -Ed0FNISTTmesaI j
E 800 |t - - - -
100
of the fire. As a result of the low thermalinertial of this cask, peak temperatures in mostcomponents occur within about an hour after theend of the fire. The exception is the peakcladding temperature, which responds muchmore slowly to the adverse heat transferconditions imposed by the fire transient.
2000.
1800leow
1400C
C-i 1000
1800
400
200
* - -- - - - -ing T-r-at--…l -Top AOr T rn~peratr
- -4 - - Con-aine-
--------- I ------- -- --
- - - - - -- - - - - - -N- - - - - - - - - - - -- - -
- - - - - - - - - - - - - - - - - - - - --^I= =; Hi ~
04-jj0.00 3.00
Elapsed Time (hours)
I -
===7"I'1;; =1 -f ___ -V0.5 1.5
Elapsed TIme (hour)Zs :
Figure 7.6. Maximum Predicted ISOContainer Surface TemperatureHistory Compared to NISTBoundary ConditionTemperaturesFigure 7.5. NAC LWT Cask (without ISO
Container): MaximumTemperature Histories for First 3hours of Fire Transient
This behavior is in response to the rapidlydecreasing boundary temperatures, as illustratedby the ISO container peak temperature in Fig-ure 7.6, and the cask peak surface temperature(for the case without the ISO container) inFigure 7.7. These figures compare the predictedpeak temperature of the ISO container or thecask outer surface to the boundary temperaturesfor the tunnel ceiling and upper tunnel air fromthe NIST calculation with FDS.
Figures 7.4 and 7.5 show that in both cases (withand without the ISO container) the peaktemperatures for all cask components except thefuel begin to decrease at or shortly after the end
E
IA
-Callt; T-p-h-4- - - - - - -_Tp AW Tempshae
-C.* outar SW-- - - - - - - - -Ed f A.
EldI
- - - - - - - - - -- - - - - - - - - ------ ------- ----------
- - - - - - - - - -
- - - - - - - - - -
900.0
700.0
000.0
300.0
100.0 .- I
0 3
Elapeed Time (hour)
Figure 7.7. Maximum Predicted Cask OuterSurface Temperature History forNAC LWT Cask without ISOContainer Compared to NISTBoundary ConditionTemperatures
7.5
In both cases, the peak clad temperature is stillrising after three hours. However, it is predictedto be only about 4970F (2580C) by this time, forthe case with the ISO container. Without theISO container, the peak clad temperature ispredicted to be 4820F (2500 C) by the end ofthree hours.
For the case with the cask within the ISOcontainer, Figure 7.8 shows the maximumtemperature histories predicted for the seals inthe drain/vent ports and the lid for the first 3hours of the transient. Figure 7.9 shows themaximum temperatures of these components forthe case without the ISO container. (Thecalculated values were gathered by queryingnodes at the seals' locations, since the seals werenot explicitly represented in the model.) Thedrain and vent ports are sealed with TFE orViton® O-rings. The bolted lid is sealed withmetallic and TFE 0-ring seals. For the caskwithin the ISO container, the drain and vent portseals are predicted to reach a maximumtemperature of 10350F (5570C) by the end of thesimulated fire. Similarly, for the cask withoutan ISO container, these seals are predicted toreach a maximum temperature of 1287 0F(6970C) by the end of the fire. The lid seal ispredicted to reach 740'F (3930C) at 1.7 hrelapsed time for the cask within an ISOcontainer. Without an ISO container, the lid sealis predicted to reach 7950 F (4240C) at 1.33 hrelapsed time in the transient.
In both cases, the seal materials then graduallydecrease in temperature as the transient proceedsinto the post-fire cool down. The extreme rise intemperature during and immediately after thefire is due to the low thermal inertia of the NACLWT cask and the close proximity of the seals toexterior surfaces subject to thermal radiationdirectly from the tunnel environment, or fromthe inner surface of the ISO container.
a 600IA 400 -
200
MOI0
Figure 7.8.
Elapsed Time (hour)
NAC LWT Cask (with ISOContainer): Maximum SealTemperature Histories forDrain/Vent Ports and Cask LidDuring First 3 hours of FireTransient
1300
1100
2
700
300
-Ud Sea
- - --- -- Ed of R.
-- E0d d NIST TwangA
. _/-/t- L - - - - - - - -- -L - - - -
I II
-L- LD
Fk. D.1500. Post-F-, 0.050. (NIST Data)100
0 3Elapsed Time (hou)
Figure 7.9. NAC LWT Cask (without ISOContainer): Maximum SealTemperature Histories forDrain/Vent Ports and Cask LidDuring First 3 hours of FireTransient
With or without an ISO container, the maximumseal temperatures predicted in this transientexceed the maximum continuous-usetemperature limits of the drain and vent portseals used in this cask design. In the lid sealregion, the predicted maximum temperature is740'F (393 0C) for the case with the cask in an
7.6
ISO container, and is 7950F (4240C) without anISO container. Both are below the maximumcontinuous-use temperature limit of 800'F(4270C) JFor the metallic seals, but exceed themaximum continuous-use temperature limit ofof the drain and vent port seals. These limits are7350F (3910C) for TFE seals and 550'F (2880C)for the altnerative design Viton® seal. For thedrain and vent port seals, the predictedmaximum temperature values in both cases(10350F ( 5570C) with an ISO container, and12870F (6971C) without an ISO container), areseveral hundred degrees above the maximumcontinuous-use temperature limits for these sealmaterials.
Figures 7.8 and 7.9 show that in both cases, withand without an ISO container, the lid seal regionmaintains temperatures at or near the peaktemperature values for a relatively short timebefore beginning a steady decrease. Similarly,the maximum temperature values predicted ineach case for the drain and vent port regionclimbs very rapidly to the peak value, thensteadily decreases. This component is above themaximum continuous-use temperature limit forless than two hours. Since the noted limits forthe Viton®, TFE, and metallic 0-ring materialsare defined for continuous use, it is possible thatthe seals might survive these temperatureexcursions undamaged.
However, information is not available on therecommended short-term temperature limits forthese seals. Based on the continuous-usetemperature limits, the primary containmentbarrier of the NAC LWT is considered todegrade at the drain/vent ports and possibly atthe lid seal under the postulated conditions ofthis fire transient, with or without an ISOcontainer. An analysis evaluating the possibleradiological consequences of the NAC LWTcask responses to the Caldecott Tunnel fire ispresented in Section 8.
7.3 NAC LWT Package Long-Term Post-Fire TransientResponse
To evaluate the effects of prolonged exposure topost-fire conditions in the tunnel, thetemperatures predicted in the NIST analysiswere extrapolated from 3 hours to 50 hoursusing a power function in order to realisticallymodel the extended cool down of the tunnelenvironment. (See Section 6.2; the extrapolatedvalues are presented in Figures 6.4 and 6.5 forthe radiation and convection heat transferboundary conditions, respectively.) Thisconservative approach is equivalent to assumingthat the cask will be left in the tunnel up to twodays without any emergency responderintervention.
The external boundary conditions were extendedusing the conservative assumption of a purelyforced convection heat transfer regime for thefirst 3 hours of the simulation, then a purely freeconvection regime for the remainder of thecalculation (t > 3 hours). Figure 7.10 shows thetemperature response of various components ofthe cask for the long term transient calculation to50 hours, with the cask enclosed in an ISOcontainer. A similar plot is shown in Figure7.11 for the case of the cask without an ISOcontainer.
The maximum temperatures for mostcomponents were reached within a short timeafter the simulated fire, (see Sections 7.1 and7.2.) However, the predicted maximum fuelcladding temperature of 5580 F (2920C) for thecask within an ISO container is not reached untilabout 8 hours into the transient. Without an ISOcontainer, the peak clad temperature is reachedapproximately one hour sooner, at 7 hours intothe transient, and the maximum temperature issomewhat lower, at 5391F (2820C).
7.7
1500 _hs- |- ElrP'Ie0d Drtb S - -Fed Oaddig
~NIST I -Cask In-er Sfahce1600 _ _ --- ~- ' ______-CaskOuteSud.ee
1400 _ E _ _ -Impa __LUitrsI -Ud Seal
* 20_ _ _I _ I _ I _ I _ __,|Oain/VenliSeals0.120D _ - Di tSsl l l l-CtskBodyLrad
11000 Ed Lea-ISO C I IftI I
E i - -_ - -- - --- - - - - - -
I. ad I I I - l _ _
400-
2 0 I
Zircaloy clad spent nuclear fuel under accidentconditions [17].
200
18C
16C
14t
1100IE SC
SC
40
20
I I l I -FuePaaddng(Peak)00 1…-orC-an-S___ 1 = netfar Sedacr.
-Cask Outer SnaacI I I I I I -Im podU.itaas
00 rf - _- -- Iy Lead| I I I I I I -ira rclenrlartsan sI I I I I I C bir
I I I I IT -4.
00 L - L- - - - - - -________
l- l00 D.I.IS t
l
0 5 10 15 20 25 30
Elapsed Tm. (hours)
35 40 45 50
Figure 7.10. NAC LWT Cask (with ISOContainer): MaximumTemperature Histories During 50hour Transient
This difference is due to the effect of the ISOcontainer on the rate of heat removal from thecask in the post-fire cool down. The ISOcontainer shields the cask from the externalenvironment, slowing the rate of heat input tothe cask during the fire, and resulting in slightlylower peak temperatures on most of the caskcomponents, compared to the values predictedwithout the ISO container. However, after thefire, the ISO container slows the rate of heatremoval from the cask to the cooling tunnelenvironment. The unshielded cask, in the casewithout the ISO container, shows a slightlyfaster cool down, and does not reach as high avalue for the maximum peak claddingtemperature during the transient.
With or without the ISO container, the peak cladtemperature does not exceed the long-termstorage temperature limit of 7520F (400'C), andis far below the currently accepted short-termtemperature limie of 1058 0F (570'C) for
2 The short-term temperature limit of 10580F (570'C)is based on creep experiments performed on two fuelcladding test samples which remained undamagedwhen held at 10580F (570'C) for up to 30 and 71days [15]. This is a relatively conservative limit,
0 5 10 15 20 25 30 35Elapsed Mme (hours)
40 4a so
Figure 7.11. NAC LWT Cask (without ISOContainer): MaximumTemperature Histories During 50hour Transient
The plots in Figures 7.10 and 7.11 also showthat the NAC LWT cask is very close to a newsteady state for the extrapolated conditions in thetunnel at 50 hours. This behavior is consistentwith the lower thermal inertia of this cask,compared to the expected response of largermulti-assembly casks to severe fire transientconditions. The temperature distributions withinthe cask predicted for these two cases (with andwithout the ISO container) for the final steadystate differ somewhat from the temperaturespredicted for the initial conditions at the start ofthe transient.
The differences are the result of the changes inthe physical condition of the package after thefire, and the different boundary conditions forthe post-fire ambient environment of the tunnel.As a result of the fire, the liquid neutron shieldhas boiled away, the cask outer surfaces (or thesurfaces of the ISO container) have a much
since the temperature at which Zircaloy fuel rodsactually fail by burst rupture is approximately 1382 0F(750 0C)[16].
7.8
higher eraissivity due to sooting, and the impactlimiters have been damaged. In addition, theambient temperatures in the tunnel are lowerthan those assumed for Hot-Normal conditionsof transport, and there is no solar insolation.
7.4 Summary of NAC LWTPackage Peak Temperaturesin Fire Transient
Peak component temperatures over the entiretransient fire simulation predicted for the NACLWT cask (with and without the ISO container)are reported in Table 7.1. These results showthat for both cases, the lead shielding within thecask body is expected to reach temperatures thatfar exceed the established safe operating limit of6000F (316'C) [7] for this material. Thepredicted peak temperature of the lead shieldingfor the cask within the ISO container is 870'F(4660C) at 0.9 hr elapsed time, and the lead inthe bottom end of the cask is predicted to reach amaximum temperatures 9520F (511 C) at 1 hrelapsed time.
For the cask without the ISO container, the peaktemperatures in the lead shielding are slightlyhigher, and are reached at a slightly earlier timein the transient. The predicted peak temperaturein the cask lead shielding is 1031 'F (555°C),and the lead in the bottom end of the cask ispredicted to reach a maximum temperatures10610F ('572 0C). In both cases, the lead remainsfully contained within the steel cask body.However, melting and possible slumping of theshielding material within the steel containmentis expected.
In the severe conditions of this fire scenario, thealuminum honeycomb material of the impactlimiters mounted on the ends of the cask isexpected to reach temperatures that areapproximately 500-600'F (278-316'C) above its
commonly estimated melting temperature. Inthe case without the ISO container, the predictedpeak temperature on this component isapproximately 1237F (680C) hotter than in thecase with the ISO container. Without the ISOcontainer, the impact limiters are directlyexposed to the intense heat of the fire, ratherthan being shielded by the walls of the container.In either case, however, the impact limiterscannot reasonably be expected to remain intactafter the fire. However, these components arenot part of the cask structure, and are generallyexpected to be damaged or destroyed byaccident conditions. Their loss in the fire is notexpected to adversely affect the thermalperformance of the cask.
Table 7.1. NAC LWT Package PeakComponent Temperatures DuringFire Transient
withoutISO with ISO
(ANSYS) Time (ANSYS) TimeComponent 0F (0C) (hours) F (0C) (hours)FuelCladding 539 (282) 7.00 558 (292) 8.00AluminumPWR Insert 425 (218) 4.50 444 (229) 5.00Inner Shell 726 (386) 1.33 661 (349) 1.70LeadGammaShield 1031 (555) 0.70 870(466) 0.90Lead EndShield 1061 (572) 0.90 952 (511) 1.00Outer Shell 1853 (1012) 0.65 1694 (923) 0.67LiquidNeutronShield 1834 (902) 0.65 1656 (902) 0.65Lid Seal 795 (424) 1.33 740 (393) 1.70Drain/VentPorts 1287 (697) 0.67 1035 (557) 0.68ImpactLimiters 1837 (1003) 0.65 1714 (934) 0.65ISOContainer N/A 173(6) 0.65
7.9
8 POTENTIAL CONSEQUENCES
USNRC staff evaluated the potential for a releaseof radioactive material from the NAC LWTtransportation cask analyzed for the CaldecottTunnel fire scenario. The analysis indicates thatthe possibility of a release cannot be entirely ruledout for this cask because temperatures in the drainand vent port seal regions during the transientexceed the continuous-use temperature limit forthe TFE or Viton® seals. Although the cask lidpeak temperature remains significantly below thecontinuous-use temperature limit for its metallicseal, it exceeds the continuous-use temperaturelimit for its TFE seal.
Staff performed an analysis to determine themagnitude of any potential release. Based on thatanalysis (described below), it was determined thatany potential release from the NAC LWT caskwould be small-less than an A2 quantity.3 Thepotential release would not involve a release ofspent fuel or fission products, but could possiblyresult from CRUD spalling off the fuel rods.
8.1 Release Analysis
The thermal analyses for the NAC LWT cask(with and without an ISO container) show thatduring the Caldecott Tunnel fire scenario this caskdesign would maintain the single most importantbarrier (i.e., the fuel cladding) to prevent therelease of radioactive materials. The temperatureof the fuel cladding is conservatively predicted toreach 5580F (2920C) when the cask is enclosedwithin an ISO container. The predicted peak
3 An A2 quantity represents the threshold belowwhich an accident resistant package is not required.The acceptance requirement for Type B packages isthat they release less than an A2 quantity/week afterbeing subjected to the hypothetical accidentconditions in 10 CFR Part 71 [1].
cladding temperature is only 5390 F (2820C) whenit is assumed that the cask is not enclosed withinan ISO container. These predicted peaktemperatures are well below the long-termcladding temperature limit of 7520F (400'C) fornormal storage and transport conditions. Thesepeak temperatures are much lower than thecladding short-term temperature limit of 10580F(570'C), and far below its projected bursttemperature of 13820F (750'C).
The maximum temperatures predicted for the TFEseals used in the NAC LWT cask lid and the drainand vent ports approach or exceed the ratedcontinuous-use temperature limit of 7350F(3911C) for this material. The predictedtemperatures also exceed the safe operatingtemperature of 550"F (2880C) for the alternativedesign Vitong seals for the drain and vent ports.The maximum temperature predicted for the lid is740'F (3930C) for the cask within an ISOcontainer, and 7950F (4240C) without an ISOcontainer. The peak temperature of the vent anddrain port seals is predicted to reach 10350F(557 0C) for the cask within an ISO container, and1287 0F (6970C) without an ISO container.
Exceeding the service temperature of the seals onthe NAC LWrcask lid or vent and drain portsmeans that there is the potential for a release tooccur. Potential releases from the drain and ventports would be limited, however, by the narrow,convoluted flow paths of these structures.Potential releases through the lid seals would belimited by the presence of the undamaged metallicseal, and by the tight clearances of the close metal-to-metal contact between the lid and cask body.The close contact is maintained by the pre-loadcreated by the initial torque on the lid bolts.
8.1
Because the fuel cladding remains intact, it is notexpected that any radioactive material would bereleased from inside the fuel rods. Any release ofradioactive material from the cask would consistonly of CRUD particles that may flake off or spallfrom individual fuel rods.
The amount of releasable CRUD in the NAC LWIcask was estimated using data developed bySandia National Laboratory for analysis of CRUDcontribution to shipping cask containmentrequirements [18], and assuming the cask containsa PWR fuel assembly consisting of 289 fuel rods.An estimate of the maximum "spot" CRUDactivity shows that for 90% of PWR spent fuelrods the maximum activity is 20 pCi/cm2 or less[18, Table 1-15]. The ratio of the peak (i.e., themaximum "spot" CRUD activity) to averageconcentration on the rod surface varies by a factorof two for PWR fuel rods [18, Table I-12].
The CRUD activity estimates [18] are for newlydischarged spent nuclear fuel. This activity isexpected to decay by a factor of one-half for five-year cooled fuel, based on the decay rate for Co6w.This is a good approximation because 92% of theactivity for five-year cooled fuel comes from Co60.Based on this data, the average CRUD activity forfive-year-cooled PWR fuel rods is about 0.006curies per rod, based on a surface area of 1200 cm2
per rod. The average CRUD activity for a 17 x 17PWR assembly is therefore about 1.73 curies.
The amount of CRUD that could flake or spallfrom the surface of a PWR rod due to temper-atures calculated for the fuel rods in the thermalanalysis is estimated to be a maximum of 15%[18, Table I-10]. The major driving force formaterial release results from the increased gaspressure inside the cask due to increases in internaltemperature. The temperature change in the caskis bounded by the difference between themaximum gas temperature predicted during thefire transient and the gas temperature at the time
the cask is loaded. For this analysis, the loadingtemperature is defined as I 000F (380C), based onthe value reported in the SAR [7]. The maximumgas temperature is assumed to be the maximumpeak clad temperature predicted during thetransient. This yields a conservative estimate ofthe maximum possible temperature change.
A deposition factor of 0.90 was used to accountfor the deposition of CRUD particles on casksurfaces and fuel assemblies. This factor wasdeveloped as part of NRCsecurity assessments forspent nuclear fuel transport and storage casks, andis based on an analysis of the gravitational settlingof small particles. The value of 0.90 isconservative because it does not consider theeffects of particle conglomeration and plugging.It is also consistent with the values used in otherstudies [16]. The major assumptions used toestimate CRUD release are given in Table 8.1.
Table 8.1. Assumptions Used forRelease Estimate for NACLWT Cask
Parameter Assumed value
Number of Assemblies in Cask 1 PWR
Rods per Assembly 289Maximum "spot" CRUD Activity 2
on Fuel Rod 20__Ci/crn
Peak to axial average variation 2CRUD decay factor (5 yr)
ased on Co6) 0.5
Average surface area per rod 1200 cm2
Average CRUD Activity on PWRFuel Rod (5 yr cooled) 0.006 Ci
Average CRUD Activity on PWRAssembly (5 yr cooled) 1.73 Ci
Fraction of CRUD released dueo heating 0.15
Deposition Factor 0.90
8.2
To estimate the potential release from the NACLWT cask, a methodology similar to thatdeveloped at Sandia National Laboratory (forNUREG-6672 [16]) was used. This methodologywas developed for evaluation of the generic risksassociated with the transport of spent fuel by truckand rail from commercial power plants toproposed interim storage and disposal sites.
The potential release from the cask in this severefire accident can be calculated from the followingrelationship:
R T
where R = release (curies)Cl = amount of CRUD on fuel
assemblies (curies)S = fraction of CRUD released due
to heatingD) = deposition factorlp = peak internal temperature (0R)l' = initial internal temperature (0R)
clad temperature (5580 F (2920 C), compared to5390F (2820 C) without an ISO container), so thisvalue was used in determining the potential releaseestimate.
Table 8.2. Potential Release Estimate forNAC LWT Cask
'eInt al Pek Potentialtemperature temperature release
F (W) - PF (MR) (cunes)
100 (560) 558 (1018) 0.01
The potential release from the NAC LWT caskbased on five-year cooled fuel is estimated to beapproximately 0.01 curies of Co6w. Since the A2value for Co6o is 11 curies, the potential release isabout 0.001 of an A2 quantity (see footnote 2).Therefore, the potential radiological hazardassociated with an accident similar to theCaldecott Tunnel fire, if it were to involve a spentnuclear fuel cask in close proximity to the firesource, is quite small. The probability of such anoccurrence, based on tunnel accident frequency,flammable materials trucking accident statistics,and radioactive material shipment statistics, hasbeen estimated as one such accident every millionyears [19].
Table 8.2 shows the results obtained when thisequation is applied using the parameter valuesfrom Table 8.1 and the temperatures predicted forthe NAC LWT cask in this accident scenario. Theanalysis Ior the cask within an ISO containerresulted in the higher predicted maximum peak
8.3
9 REFERENCES
1. 1O C:FR 71. Jan. 1, 2003. Packaging andTransportation of Radioactive Material. Codeof Federal Regulations, U.S. NuclearRegulatory Commission, Washington D.C.
2. NTS]BIHAR-83/01. 1983. Multiple VehicleCollisions and Fire: Caldecott Tunnel, nearOakland, California, April 7, 1982. NationalTransportation Safety Board, Bureau ofAccident Investigation, Washington D.C.
3. McGrattan KB, HR Baum, RG Rehm,GP Fomey, JE Floyd, and S Hostikka.November 2001. Fire Dynamics Simulator(Version 2), User's Guide. NISTIR 6784,National Institute of Standards andTechnology, Gaithersburg, Maryland.
4. McGrattan KB. May 2005. NumericalSimulation of the Caldecott Tunnel Fire, April1982. NISTIR 7231, National Institute ofStandards and Technology, Gaithersburg,Maryland.
5. Bechtel/Parsons Brinkerhoff, Inc. November1995. Memorial Tunnel Fire Ventilation Tes.tProgram, Comprehensive Test Report,Prepared for Massachusetts HighwayDepartment and Federal HighwayAdministration.
6. McGrattan KB, HR Baum, RG Rehm,GP Forney, JE Floyd, and S Hostikka.November 2001. Fire Dynamics Simulator(Version 2), Technical Reference Guide.NISTIR 6783, National Institute of Standardsand Technology, Gaithersburg, Maryland.
2001, Rev. 33; Rev. 34. Nuclear AssuranceCorporation, Atlanta, Georgia.
8. ANSYS, Inc. 2003. "ANSYS Users Guidefor Revision 8.0," ANSYS, Inc., Canonsburg,Pennsylvania.4
9. Bahney RH III, TL Lotz. July 1996. SpentNuclear Fuel Effective Thermal ConductivityReport, BBAOOOOOO-01717-5705-00010 Rev.00. TRW Environmental Safety Systems,Inc., Fairfax, Virginia.
10. Guyer EC and DL Brownell, editors. 1989.Handbook of Applied Thermal Design.McGraw-Hill, Inc., New York, p. 1-42.
11. Kreith F and MS Bohn. 2001. Principles ofHeat Transfer, 6h Edition. Brooks/Cole,Pacific Grove, California.
12. Holman JP. 1986. Heat Transfer, 6 th Edition.McGraw-Hill, Inc.
13. Kreith F. 1976. Principles of Heat Transfer,
3 rd Edition. Intext Education Publishers,New York.
14. Bucholz JA. January 1983. Scoping DesignAnalyses for Optimized Shipping CasksContaining 1-, 2-, 3-, 5-, 7-, or 10-year oldPWRSpentFuel. ORNL/CSD/TM-149. OakRidge National Laboratory, Oak Ridge,Tennessee.
15. Johnson AB and ER Gilbert. September 1983.Technical Basis for Storage of Zircaloy-Clad
7. NRC Docket Number 71-9225. Legal WeightTruck Transport (LWT) Packaging SafetyAnalysis Report, July 2000, Rev. 2; September
4 ANSYS Release 10.0 (2005) was used for some of thegraphical post-processing, but all fire transientsimulations of the NAC LWT package were performedwith ANSYS Release 8.0.
9.1
Spent Fuel in Inert Gases, PNL-4835. PacificNorthwest Laboratory, Richland, Washington.
16. Sprung JL, DJ Ammerman, NL Breivik, RJDukart, and FL Kanipe. March 2000.Reexamination of Spent Fuel Shipment RiskEstimates, NUREG/CR-6672, Vol. 1(SAND2000-0234). Sandia NationalLaboratories, Albuquerque, New Mexico.
17. U.S. Nuclear Regulatory Commission.January 1997. "Standard Review Plan for DryCask Storage Systems." NUREG-1536,USNRC, Washington D.C.
Estimate of CRUD Contribution to ShippingCask Containment Requirements, SAND88-1358. Sandia National Laboratories,Albuquerque, New Mexico.
19. Larson DW, RT Reese, and EL Wilmot.January 1983. The Caldecott Tunnel FireThermal Environments, RegulatoryConsiderations and Probabilities, SAND-82-1949C;CONF-830528-8, Sandia NationalLaboratories, Albuquerque, New Mexico.Presented at 7th International Symposium onPackaging and Transportation of RadioactiveMaterials, 15 May 1983, New Orleans, LA.
18. Sandoval RP, RE Einziger, H Jordan, APMalinauskas, and WJ Mings. January 1991.
9.2
Appendix A
Material Properties for ANSYS Modelof Legal Weight Truck Package
Table A.1. 304 Stainless Steel
..Temperature ThermalConductivity Specific Heat'(°F- (tulhr-in-)0 Den1sity (Ibm/in 3 ) k(BtuIilnn-0 F) Description-..
70 0.7143 _ - 0.1141
212 0.7800 0.2888 0.1207392 0.8592 0.2872 0.1272 Used for cask
572 0.9333 0.2855 0.1320 body, cask lid,752 1.0042 0.2839 0.1356 spokes932 1.0717 0.2822 0.1385
1112 1.1375 0.2805 0.1412
Table A.2. 6061-T6 AluminumonductivtSecific: Hea
( .' (Btulh r- 'n t Den ( inbm- 0F) Description
32 9.7500
212 9.9167 0.0984 0.2140 Used for basket,
572 11.0833 . . IL 1, 2 skin
932 12.9167 _
Table A.3. 6061-T6 Aluminum Honeycomb
Tewperature Thermal'Conductivity S,'ecific Heat. (Btu/hr-in-.0 F) Density (bmin (Btubm- "Description
32 1.6965
212 1.7255 007156024Used for IL I= 1 250.017118056 0.214 (Honeycomb)572 1.9285(Hnyob
932 2.2475
Table A.4. 6061-T6 Aluminum Honeycomb
T peratr Thermal Conductivity SpecificHeat"'; 0F) (Btur-in-°F) Density (Ilbi (;tu/ /lbb m-0 F) d Des
32 1.4235212 1.4478 0.0144 0.214 Used for IL 2
572 1.6182 0. (Honeycomb)
932 1.8858 _
Table A.5. Helium
;Teperature,: ThermalCo'nductivity SpecificHeat.(0F) ; ' (Btuhr-inF) D t (b:/ie (Btullbm-°F) Description
200 0.00808 4.83E-06
400 0.00942 3.70E-06 1.24 Used for cask gap
600 0.01075 3.01E-06 and fuel gap
800 0.0115 2.52E-06
A.1
Table A.6. Chemical Copper LeadTemperature Thermal Conductivity Specific Heat
(OF) ,(Btu/hr-in-PF) -. Density (Ibm/in3) (Btullbm-F) i Description68 1.6651
209 1.6308400 1.526 0.3 0.06 Used for lead499 1.4111 regions581 1.2096630 1.0079
Table A.7. 56% Ethylene Glvcol SolutionAvg. : :,'Thermal
Temperature iConductivity 0Specific EHeat ^Densityi0 °F) S(Btu/hr-in-04) .(BtulbmnI~); (ibm/in3)50 0.0188 0.7405 0.039170 0.0187 0.7522 0.0389
100 0.0185 0.7696 0.0385150 0.0182 0.7979 0.0378200 0.0179 0.8255 0.0370250 0.0177 0.8522 0.0362260 0.0176 0.8575 0.0360270 0.0176 0.8627 0.0358280 0.0175 0.8679 0.0357290 0.0175 0.8731 0.0355300 0.0174 0.8782 0.0353310 0.0174 0.8833 0.0351320 0.0173 0.8884 0.0349330 0.0173 0.8934 0.0347340 0.0172 0.8984 0.0345350 0.0172 0.9034 0.0343
A.2
Table A.8. AirAvg. Thermal
'Tempenrture C Conducativit y Specific Heat Density
350 0.0017 0.2467 0.0000283450 0.0018 0.2494 0.0000252550 0.0020 0.2516 0.0000227650 0.0022 0.2533 0.0000206750 0.0023 0.2546 0.0000189850 0.0025 0.2556 0.0000175950 0.0026 0.2562 0.0000162
1050 0.0027 0.2566 0.00001521150 0.0029 0.2568 0.00001421250 0.0030 0.2570 0.00001341350 0.0031 0.2571 0.00001261450 0.0033 0.2571 0.00001201550 0.0034 0.2573 0.00001141650 0.0035 0.2576 0.00001081750 0.0036 0.2581 0.00001041850 0.0038 0.2589 0.00000991950 0.0039 0.2599 0.00000952050 0.0040 0.2614 0.0000091
Table A.9. Effective Conductivity for Liquid Neutron Shield with 1°F Temperature Gradient3 56% EthleneGcol Air
Avg Effective Cionductivity Effective'Conducivity Effective Conductivit Effective Conductivityperature NeutronShield& Epansin Tank Neutron kS ield~i-00 Expansion Tank
250 0.364 0.149 0.003 0.002
260 0.374 0.153 0.003 0.002
270 0.384 0.157 0.003 0.002
280 0.393 0.161 0.003 0.002
290 0.398 0.163 0.003 0.002
300 0.396 0.162 0.003 0.002
310 0.395 0.162 0.003 0.00232. 0.394 0.161 0.003 0.002
332 0.393 0.161 0.003 0.002
34(9 0.391 0.160 0.003 0.002
350 0.390 0.160 0.003 0.002
351 * __ 0.003 0.002
40(0 * __ 0.003 0.002
00 * . 0.003 0.002
60(0 _ 0.003 0.002
800 * 0.003 0.002
100C * 0.003 0.003
1200 0.003 0.003
150(C * 0.003 0.003
200C0 0.004 0.004
250CI * __ 0.004 0.004
A.3
Table A.10. Effective Conductivity for Liquid Neutron Shield with 10IF Temperature Gradient56% Ethylene Glycol Air
Avg. Effective Conductivity Effective Conductivity Effective Conductivity Effective ConductivityTemperature Neutron Shield Expansion Tank Neutron Shield Expansion Tank
5 0 i-F) (Btu/hr-in-° (Btu/hr-0n-65F) 0.2 00(Btur- 0.250 0.654 0.268 0.006 0.002260 0.673 0.276 0.006 0.002270 0.691 0.283 0.006 0.002280 0.704 0.288 0.006 0.002290 0.705 0.289 0.006 0.002300 0.703 0.288 0.006 0.002310 0.701 0.287 0.006 0.002320 0.699 0.286 0.006 0.002330 0.697 0.286 0.006 0.002340 0.695 0.285 0.006 0.002350 0.006 0.002351 0.006 0.002400 * * 0.006 0.002500 * * 0.006 0.002600 0.005 0.002700 0.005 0.002800 * * 0.005 0.002
1000 **0.005 0.0031200 * * 0.005 0.0031500 * * 0.004 0.0032000 * 0.004 0.0042500 0.004 0.004
A.4
Table A.11. Effective ConductivitY for Liquid Neutron Shield with 25IF Temperature Gradienti56%16 Ethyl eneGlycol Air
CAvg. Effective Conductivity Effective Conductivity EeCtive Conductivity Effcutive ductivitTemperature NeutrOn Shield Expanson Tank Neutron Shield 'Expansion Tank -
( 0)(uhrin-OF) (til h -ino0 F (Btix 4-n-0PF) (Btnlri-F250 0.840 0.344 0.008 0.003260 0.863 0.353 0.008 0.003270 0.882 0.361 0.008 0.00328(1 0.888 0.364 0.008 0.00329C0 0.885 0.363 0.007 0.00330CI 0.883 0.361 0.007 0.003310 0.880 0.360 0.007 0.00332(1 0.877 0.359 0.007 0.003330 0.875 0.358 0.007 0.00334CI 0.872 0.357 0.007 0.00335CI 0.007 0.003
351 * * 0.007 0.003
400 * * 0.007 0.003
50C * * 0.007 0.003
600 0.007 0.003
70CI * * 0.007 0.003
800 * * 0.006 0.003
1000 * * 0.006 0.003
1200 0.006 0.003
1500 * * 0.005 0.003
2000 0.005 0.004
2500 * * 0.005 0.004
A.5
Table A.12. Effective Conductivity for Liquid Neutron Shield with 50IF Temperature Gradient56% Eth Iene Glycol Air
Avg. Effective Conductivity Effective Conductivity tEffectve EffectfiveConductivityTemperature Neutron Shield Expansion Tank Conductivity Neutron Expansion Tank
(OF) (Btu/hr-in-OF) (Btu/hr-in-0F). Shield (Btu/hr-in-0 F) (Btulhr-in-F)^:250 1.061 0.434 0.009 0.004260 1.058 0.433 0.009 0.004270 1.055 0.432 0.009 0.004280 1.052 0.431 0.009 0.004290 1.049 0.430 0.009 0.004300 1.046 0.428 0.009 0.004310 1.043 0.427 0.009 0.004320 1.039 0.426 0.009 0.004330 0.009 0.004340 * * 0.009 0.004350 * * 0.009 0.004351 0.009 0.004400 0.009 0.003500 * * 0.008 0.003600 * * 0.008 0.003700 0.008 0.003800 0.008 0.003
1000 * * 0.007 0.0031200 0.007 0.0031500 * * 0.006 0.0032000 * * 0.006 0.0042500 0.006 0.004
A.6
Table A.13. Effective Conductivity for Liquid Neutron Shield with 70°F Temperature Gradient56%Eth Gle~y~o
Avg. Effective Conductivity Effective'Coinductivity EffectiveConductivity Effective ConductivityTemperatu~re ~Neutron Shield -Expansion Tank' Nur~ hel xaso Tank
(,, . e. 0F (Btu/hr-in-o ) (B' 0 r-n-0F (Bthlhr-in 0 r"(Btu/r-in-'- 0
25(1 1.151 0.471 0.010 0.004260 1.148 0.470 0.010 0.004270 1.144 0.469 0.010 0.00428Ci 1.141 0.467 0.010 0.00429CI 1.138 0.466 0.010 0.00430C 1.134 0.464 0.010 0.004310 1.131 0.463 0.010 0.00432Ci * * 0.010 0.004330 * * 0.010 0.004340* * 0.009 0.00435(* 0.009 0.004351 * * 0.009 0.004400 * * 0.009 0.00450CI * * 0.009 0.004600 * 0.009 0.004700 0.008 0.00380C0 * * 0.008 0.003
1000 * * 0.008 0.003120C * 0* 0.007 0.0031500 0.007 0.0032000 * * 0.006 0.0042500 * * 0.006 0.004
A.7
Table A.M4. Effective Conductivity for Liquid Neutron Shield with 100OF Temperature Gradient56% Ethi : ;e :iycol Air
Avg. Effective Conductivity Effective Conductivity Effective Conductivity Effective ConductivityTemperature Neutron Shield Expansion Tank Neutron Shield Expansion'Tank
(OF) (Btulhr-in-0F) (Btu/hr-in-0F) : (Btulhr-in-F) (Btu/hr-in-1F).250 1.253 0.513 0.011 0.004260 1.249 0.512 0.011 0.004270 1.245 0.510 0.011 0.004280 1.242 0.509 0.011 0.004290 1.238 0.507 0.011 0.004300 1.234 0.505 0.011 0.004310 * * 0.010 0.004320 * * 0.010 0.004330 0.010 0.004340 * * 0.010 0.004350 * * 0.010 0.004351 * * 0.010 0.004400 - * 0.010 0.004500 0.010 0.004600 * * 0.009 0.004700 * * 0.009 0.004800 * * 0.009 0.004
1000 * * 0.008 0.0031200 0.008 0.0031500 * * 0.008 0.0032000 * * 0.007 0.0042500 0.007 0.004
A.8
Table A.15. Effective Conductivity for Liquid Neutron Shield with 200°F TemperatureGradient
56% Eth lene Glccl Air -ecv E Effective Effective
-Ag. Conductivity' Condluctivit: C c ti CodctvtTemperature NeutronShield . Expansion Tank Neutron Shield Expansion Tank
.. u/Fri- 0F) : (Btui/hr-in- 0F) -,(Btu/hr-in-F) (Btu/hr-inu-°0 ;F
250 1.468 0.601 0.013 0.0052260 0.013 0.005270 * 0.013 0.005280 * * 0.013 0.005290 * 0.013 0.0053(10 * * 0.012 0.00531.0 * * 0.012 0.005320 * * 0.012 0.005330 0.012 0.0053'40 * * 0.012 0.005350 * * 0.012 0.005351 * 0.012 0.0054(0 0.012 0.005500 * 0.012 0.005600 * * 0.011 0.004700 * * 0.011 0.0048(10 * * 0.011 0.004
1000 * 0.010 0.00412CO * * 0.010 0.00415C(O * * 0.009 0.00420C0O * 0.008 0.00425C'0 * * 0.008 0.005
A.9
Table A.16. Effective Conductivity for Liquid Neutron Shield with 300IF TemperatureGradient
56% Ethylene Glycol AirEffective :.Effective Effective Effective
Avg. Conductivity Conductivity Conductivity ConductivityTemperature Neutron Shield Expansion Tank Neutron Shield Expansion Tank
: (0 f)' i/-' (Btu/hr-in- 0f) d: (Btutu/hr-in-0 (Btu'hr-in-°):250 * * 0.014 0.005260 * * 0.014 0.005270 * * 0.014 0.005280 0.014 0.005290 * * 0.014 0.005300 * * 0.014 0.005310 * * 0.014 0.005320 0.014 0.005330 * * 0.014 0.005340 * * 0.014 0.005350 * * 0.013 0.005351 0.013 0.005400 * 0.013 0.005500 * * 0.013 0.005600 * * 0.012 0.005700 0.012 0.005800 * * 0.012 0.0051000 * * 0.011 0.0041200 * * 0.011 0.0041500 * * 0.010 0.0042000 0.009 0.0042500 * * 0.009 0.005
A.10
Table A.17. Effective Conductivity for Liquid Neutron Shield with 500°F TemperatureGradient
K56% Ethylene Glyol AirEffetive Effective| Efective Effective
Avg. C, . ndutvit Conductivity ;..:.' Conductivity ConductivityTemperature Neut ;Shield 'Ex pansion Tanik', Neutron Shield Eansi' Tank
250 * * 0.016 0.006.260 * * 0.016 0.006270 * * 0.016 0.006:280 * * 0.016 0.006_290 * * 0.016 0.006:300 * * 0.015 0.006:310 * 0.015 0.006.320 * * 0.015 0.006330 0.015 0.006340 * 0.015 0.006:350 * * 0.015 0.006:351 * * 0.015 0.006400 0.015 0.006.500 * 0.014 0.0061500 * * 0.014 0.005'700 * * 0.014 0.005800 * * 0.013 0.005
1300 * 0.013 0.0051200 * * 0.012 0.0051.500 * * 0.011 0.0052000 * * 0.011 0.0042:500 * * 0.010 0.005
Table A.18. Emissivity Values for Radiation Heat Transfer- Emissiit 'Before Emissivi tI
CofmponentU0 M Fire During/Afterl FreCanister stainless steel 0.36 0.36Cask stainless steel 0.36 0.36Outer Neutron Shield 0.34 0.34Ihmer Neutron Shield _ 0.34 0.34
Basket stainless steel 0.36 0.36Fuel Clad zircaloy 0.8 0.8BEoral Plate aluminum clad 0.55 0.55Shell Interior stainless steel 0.36 0.36Cask Exterior stainless steel 0.85 0.9Tunnel/ISO various 0.9
A.11
NRC FORM 335 U.S. NUCLEAR REGULATORY COMMISSION AdREPORTNUMBER(9-2004) (Assigned by NRC, Add Vol,, Supp., Revs,NRCMD 3.7 fnd Addendum Numbers, i any.)
BIBLIOGRAPHIC DATA SHEET(So on #w fsew$rst NUREG/CR-6894
2. TITLE AND SUBTITLE .S DATE REPORT PUBLISHED
Spent Fuel Transportation Package Response to the Caldecot Tunnel Fire Scenario YEAR
Draft report for comment February 20064. FIN OR GRANT NUMBER
.___ J51675. AUTHOR(S) 6. TYPE OF REPORT
H. E. Adkins, Jr.B. J. Koeppel TechnicalJ. M. Cuta 7. PERIOD COVERED tIncSW Dams)
8. PERFORMINGORGANIZATION - NAME AND ADDRESS WNRCpdDiwsf, O~-osre2,on. U.& NuarRe u t Conrsn. andnwt Sndr1e cprover nram and ma a dressy
Pacific Northwest National LaboratoryRichland, WA 99:352
9., SPONSORING OGANIZATION - NAME AND ADDRESS (IINRC, tofA Ori~ RJ o~fC ogicr US. , ~ ~ >and tfu agdde s.)
Spent Fuel Project OfficeOffice of Nuclear Material Safety and SafeguardsU.S. Nuclear Regulatory CommissionWashington, D.C. 20555-0001
10. SUPPLEMENTARY NOTES
11.ABSTRACT 2owor;or l
On April 7, 1982, a tank truck and trailer carrying 8,800 gallons of gasoline was involved in an accident in the Caldecott tunnelon State Route 24 near Oakland, California. The tank trailer overturned and subsequently caught fire. The U.S. NuclearRegulatory Comrnission selected this accident for analysis 0 determine the possible regulatory implications.The staff concluded that small transportation casks similar to the NAC LWT cask would probably experience degradation ofsome seals in this severe accident scenario. However, any release is expected to be very small.
USNRC staff evaluated the radiological consequences of the package response to the Caldecolt tunnel fire. The results of thisevaluation strongly indicate that neither spent nuclear fuel (S>NF) particles nor fission products would be released from a spentfuel shipping cask involved in a severe tunnel fire such as the Caldecott Tunnel Fire. The NAC LWT cask design analyzed forthe Caldecott Tunnel fire scenario does not reach internal temperatures that could result in rupture of the fuel cladding.Therefore radioactive material (i.e., SNF particles or fission products) would be retained within the fuel rods. The potentialrelease calculated for the NAC LWT cask in this scenario indicates that any release of CRUD from the cask would be verysmall.
12. KEY WORDSIDESCRIIPTORS fLW words Or., zarttal WA sa3. AVAILASB fTY STATEMENT
Spent fuel transportation 4 unlimiTeCask thermal performance 14. SECING OLAs neaTsThermal analysis
unclassifiedTh Repo-m
unclassified15. NUMBER OF PAGES
16. PRICE
NfKC FORM 335 (9-2004) PRINTED ON RECYCLED PAPER
Federal Recycling Program
