Characterization of Thermal-Hydraulic 

and Ignition Phenomena in Prototypic, 

Full-Length Boiling Water Reactor Spent 

Fuel Pool Assemblies After a Postulated 

Complete Loss-of-Coolant Accident

Office of Nuclear Regulatory Research

NUREG/CR-7143

SAND-2007-2270
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infringe privately owned rights.
Characterization of Thermal-Hydraulic 

and Ignition Phenomena in Prototypic, 

Full-Length Boiling Water Reactor Spent 

Fuel Pool Assemblies After a Postulated 

Complete Loss-of-Coolant Accident

Manuscript Completed: October 2012

Date Published: March 2013

Prepared by

E. R. Lindgren and S. G. Durbin

Sandia National Laboratory

Albuquerque, NM 87185

G. A. Zigh, Technical Advisor

A. Velazquez-Lozada, Project Manager

NRC Job Code Y6758

Prepared for

Division of Systems Analysis

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

NUREG/CR-7143

SAND-2007-2270

 iii 

ABSTRACT 

The NRC regulates the operation of the civilian nuclear power plant fleet by establishing and 

enforcing regulatory requirements for their design, construction, and operation. To protect the 

health and safety of the public and environment, the NRC requires all nuclear power plants to 

have a spent fuel pool where the used reactor fuel assemblies are allowed to cool for a number of 

years before being moved to interim or permanent storage. Spent fuel pools (SFP) are robust 

structures with an extremely low likelihood of a complete loss of coolant under traditional 

accident scenarios. However, in the wake of the terrorist attacks of September 11, 2001, the SFP 

accident progression was reconsidered and reevaluated using best-estimate accident codes. 

 

In 2001, the NRC staff performed an evaluation of the potential accident risk in a SFP at 

decommissioning plants in the United States. This evaluation is documented on NUREG-1738, 

“Technical Study of Spent Fuel Pool Accident Risk at Decommissioning Nuclear Power Plants,” 

(ML010430066). The study described a modeling approach of a typical decommissioning plant 

with design assumptions and industry commitments, thermal-hydraulic analyses performed to 

evaluate spent fuel stored in the SFP at decommissioning plants, risk assessment of SFP 

accidents, consequence calculations and the implications for decommissioning regulatory 

requirements. Some assumptions in the accident progression were known to be necessarily 

conservative, especially the estimation of the fuel damage. Consequently, the NRC continued 

SFP accident research by applying best-estimate computer codes to predict the severe accident 

progression following various postulated accident initiators. These code studies identified 

various modeling and phenomenological uncertainties that prompted a need for experimental 

confirmation. The present experimental program was undertaken to address thermal-hydraulic 

issues associated with complete loss-of-coolant accidents in boiling water reactor SFPs. All of 

the experiments and numerical simulations described in this report were performed at Sandia 

National Laboratories (SNL) in Albuquerque, New Mexico. 

The objective of this project was to provide basic thermal-hydraulic data associated with a 

postulated SFP complete loss-of-coolant accident. The accident conditions of interest for the SFP 

were simulated in a full-scale prototypic fashion (electrically-heated, prototypic assemblies in a 

prototypic SFP rack) so that the experimental results closely represent actual fuel assembly 

responses. A major impetus for this work was to facilitate code validation (primarily MELCOR) 

and reduce modeling uncertainties within the code. 

The research summarized in this report achieved the stated objective and resolved several 

unexpected technical challenges related with thermocouple attachments and the choice of 

appropriate input power in the heated design test that would cause the ignition. The close 

coupling of the experimental and numerical programs allowed for rapid validation and 

improvement of the MELCOR whole pool calculations. Because of the success of this approach, 

this project will be used as a model for subsequent studies. 

 

 v 

CONTENTS 

ABSTRACT ................................................................................................iii 

LIST OF FIGURES ....................................................................................................................... ix

LIST OF TABLES ...................................................................................................................... xvii

EXECUTIVE SUMMARY ......................................................................................................... xix

ABBREVIATIONS/DEFINITIONS ........................................................................................... xxi

1 INTRODUCTION ..................................................................................................................1

1.1 Background ...................................................................................................................1 

1.2 Objectives and Scope of Testing ..................................................................................1 

1.3 Test Plan Synopsis ........................................................................................................1 

1.4 Report Organization ......................................................................................................3

2 HEATER DESIGN TEST 1 ...................................................................................................5

2.1 Objective .......................................................................................................................5

2.2 Experimental .................................................................................................................5

2.2.1 Electric Heater Design ...................................................................................5 

2.2.2 Initial Experimental Assembly .......................................................................7 

2.2.3 Final Experimental Assembly ......................................................................14 

2.2.4 Additional Sampling ...................................................................................15 

2.2.5 Initial Test Operation ...................................................................................16 

2.2.6 Initial Test Analysis .....................................................................................19 

2.2.7 Final Test Operation .....................................................................................19 

2.3 Data Analysis ..............................................................................................................23

2.3.1 Ignition Details .............................................................................................23

2.3.2 Smoke Sampling Results ..............................................................................31

2.4 Discussion ...................................................................................................................32

2.4.1 Time of Ignition and Burn Rate ...................................................................32

2.4.2 Post-Test MELCOR Analysis ......................................................................32

2.5 Technical Issues ..........................................................................................................33

2.5.1 TC Attachment .............................................................................................33

2.5.2 Molten Zirconium-Magnesium Oxide (Zr-MgO) Reaction .........................34

2.6 Conclusion ..................................................................................................................35

3 HEATER DESIGN TEST 2 .................................................................................................37

3.1 Background .................................................................................................................37

3.2 Objective .....................................................................................................................37

3.3 Electric Heaters ...........................................................................................................38

3.4 Experimental ...............................................................................................................38

3.4.1 Additional Sampling ....................................................................................47

3.4.2 Test Operation ..............................................................................................48

3.4.3 Ignition Details .............................................................................................50
 vi 

3.4.4 Post-Test MELCOR Analysis ......................................................................55

3.4.5 Post-Test Inspection .....................................................................................56

3.4.6 Smoke Sampling Results ..............................................................................61

3.5 Issues ...........................................................................................................................61

3.6 Conclusion ..................................................................................................................61

4 HYDRAULIC ANALYSIS OF THE SPENT FUEL POOL EXPERIMENT .....................63

4.1 Experimental Apparatus and Procedures ....................................................................63

4.2 Hydraulic Analysis .....................................................................................................66

4.2.1 Blocked Water Rod Results .........................................................................67

4.2.2 Unblocked Water Rod Results .....................................................................72

4.2.3 Water Rod Flow Rate ...................................................................................74

4.2.4 Bypass Hole Flow Rate ................................................................................78

4.3 Summary .....................................................................................................................79

4.3.1 Application of Experimental Results to a MELCOR Model .......................79

5 THERMAL−HYDRAULIC ANALYSIS OF THE SPENT FUEL POOL 

EXPERIMENT .....................................................................................................................83

5.1 Experimental Apparatus and Procedures ....................................................................83

5.1.1 Hardware Components .................................................................................83

5.1.2 Thermocouple Layout ..................................................................................83

5.1.3 Flow Measurements and Calibrations ..........................................................85

5.2 Modeling Efforts .........................................................................................................92

5.2.1 COBRA-SFS (Coolant Boiling in Rod Arrays-Spent Fuel Storage) ...........92

5.2.2 MELCOR – Severe Accident Analysis ........................................................94

5.3 Results .........................................................................................................................97

5.3.1 Transient Thermal Response ........................................................................97

5.3.2 “Steady State” Peak Cladding Temperature ...............................................102

5.3.3 Steady State Natural Induction Flow Rates ................................................102

5.4 Summary ...................................................................................................................103

6 SUMMARY OF THE 14 INCOLOY SHORT STACK THERMAL RADIATION 

EXPERIMENTS OF THE SFP PROJECT ........................................................................105

6.1 Assembly Design ......................................................................................................105

6.1.1 TC Layout ..................................................................................................106

6.2 MELCOR Analysis Methodology ............................................................................108

6.2.1 MELCOR Code ..........................................................................................108

6.2.2 Short 14 Assembly Model ........................................................................109

6.3 COBRA 14 Model ..................................................................................................112

6.4 Results .......................................................................................................................114

6.5 Model Sensitivity to Emissivity ................................................................................120

6.5.1 Additional COBRA Model Modifications .................................................128

7 FULL-SCALE ZIRCALOY PRE-IGNITION AND IGNITION RESULTS OF THE 

SFP EXPERIMENT ...........................................................................................................129

7.1 Apparatus ..................................................................................................................129

7.2 Pre-Ignition Testing and Analysis ............................................................................135

7.2.1 Comparison of Thermal Results .................................................................135

7.2.2 Comparison of Air Flow Results ................................................................136
 vii 

7.3 Ignition Test and Analysis ........................................................................................138

7.3.1 Thermal Data ..............................................................................................138

7.3.2 Oxygen Data ...............................................................................................139

7.3.3 Comparison of Air Flow Results ................................................................142

7.3.4 Comparison of Thermal Results .................................................................145

7.4 MELCOR Post-test Comparison Summary ..............................................................147

7.5 Discussion .................................................................................................................153

8 1×4 ZIRCALOY PRE-IGNITION AND IGNITION RESULTS OF THE SFP 

EXPERIMENT ...................................................................................................................155

8.1 Assembly Design ......................................................................................................155

8.1.1 General Construction ..................................................................................155

8.1.2 TC Layout ..................................................................................................158

8.1.3 Oxygen Monitors ........................................................................................161

8.2 Ignition Test Results .................................................................................................163

8.2.1 Input Parameters .........................................................................................163

8.2.2 Oxygen Profiles ..........................................................................................165

8.2.3 Axial Temperature Profiles ........................................................................167

8.2.4 Radial Temperature Gradients ...................................................................171

8.2.5 Post-Mortem of the Test Apparatus ...........................................................176

8.3 MELCOR Baseline Case Comparison ......................................................................177

8.3.1 Model Input Parameters .............................................................................177

8.3.2 Comparison of Results ...............................................................................180

8.4 MELCOR Sensitivity Study .....................................................................................184

8.4.1 Modification to Kinetics Model .................................................................185

8.4.2 Comparison of Results ...............................................................................187

8.5 Summary ...................................................................................................................191

9 REPORT SUMMARY .......................................................................................................193

9.1 Heater Design Tests ..................................................................................................193

9.2 Separate Effects Tests ...............................................................................................194

9.3 Integral Effects Tests ................................................................................................194

9.4 Key Findings .............................................................................................................195

9.5 Uncertainties and Potential Mitigating Factors ........................................................196

10 REFERENCES ...................................................................................................................197

APPENDIX A. SFP HEATER DESIGN TEST 1 DATA SUMMARY ................................... A-1

APPENDIX B. ERROR ANALYSIS .........................................................................................B-1

APPENDIX C. TSI HOT WIRE ANEMOMETER MEASUREMENTS ..................................C-1

APPENDIX D. ADDITIONAL DATA AND FINAL POST-TEST MELCOR RESULTS 

FOR THE FULL-SCALE ZIRCALOY IGNITION TEST ............................................... D-1

 ix 

LIST OF FIGURES 

Figure 1.1 Illustration of the closely-coupled relationship between the various testing 

elements and the supporting MELCOR modeling. ..................................................3 

Figure 2.1 Cumulative power distribution along a typical spent fuel rod and the electric 

heater rod. ................................................................................................................7 

Figure 2.2 Specific heat of electric heaters compared to spent nuclear fuel. ............................7 

Figure 2.3 Schematic of the heater design test. .........................................................................8 

Figure 2.4 Bundle of heater rods. ..............................................................................................9 

Figure 2.5 TC attachment detail. .............................................................................................11 

Figure 2.6 Quartz tube installed over heater bundle. ..............................................................11 

Figure 2.7 Stainless steel foil in place over quartz tube. .........................................................12 

Figure 2.8 Insulation and side-viewing slit. ............................................................................13 

Figure 2.9 Initial test setup. .....................................................................................................14 

Figure 2.10 More insulated final test setup. ..............................................................................15 

Figure 2.11 Temperature of air entering the bottom of the assembly and temperature of the 

air leaving the top of the assembly during the initial test operation. .....................16 

Figure 2.12 History of the electrical current supplied to each of the twelve heater rods 

during the initial test operation. .............................................................................17 

Figure 2.13 Axial temperature profile developed during the initial test operation. ..................17 

Figure 2.14 Radial temperature profiles developed during the initial test operation. ...............18 

Figure 2.15 Axial differential temperature change measured during the initial test 

operation. ...............................................................................................................18 

Figure 2.16 First derivative of axial temperature profile with major experimental events 

noted. ......................................................................................................................21

Figure 2.17 Close-up of chimney during ignition initiation. .....................................................22 

Figure 2.18 Smoke filled CYBL vessel during burn. ................................................................23 

Figure 2.19 Oxygen concentration of off gas during time of Zircaloy ignition. .......................24 

Figure 2.20 Heater currents during time of Zircaloy ignition. ..................................................24 

Figure 2.21 Axial temperature profile during time of ignition..................................................25 

Figure 2.22 First derivative of axial temperature profile during time of ignition. ....................25 

Figure 2.23 Radial temperature profile at the 36-in. elevation during ignition. ........................26 

Figure 2.24 First derivative of radial temperature profile at the 36-in. elevation during 

ignition. ..................................................................................................................26 

Figure 2.25 Radial temperature profile at the 24-in. elevation during ignition. ........................27 

Figure 2.26 First derivative of radial temperature profile at the 24-in. elevation during 

ignition. ..................................................................................................................27 

Figure 2.27 Radial temperature profile at the 12-in. elevation during ignition. ........................28 

Figure 2.28 First derivative of radial temperature profile at the 12-in. elevation during 

ignition. ..................................................................................................................28 

Figure 2.29 Temperature on the outside of insulation at the 42-in. level. .................................31 

Figure 2.30 Heater Design Test 1 PCT with and without post-breakaway oxidation 

kinetics. ..................................................................................................................33 

Figure 2.31 TC attachment options, (a) small hood and (b) full wrap. .....................................34 

Figure 2.32 Photo image of burnt heater rod cross sections, (a) lower rod segment end 

detail, and (b) upper rod segment end detail. .........................................................34 
 x 

Figure 2.33 Electron microprobe images of burnt heater rod section, (a) scanning electron 

microscopy (SEM) image, (b) Zirconium analysis, and (c) magnesium 

analysis. ..................................................................................................................35 

Figure 3.1 Schematic of the second heater design test. ..........................................................37 

Figure 3.2 Bundle of heater rods. ............................................................................................39 

Figure 3.3 TC attachment detail. .............................................................................................42 

Figure 3.4 Quartz tube installed over heater bundle. ..............................................................43 

Figure 3.5 Stainless steel foil in place over quartz tube and ceramic positioning block is 

in place. ..................................................................................................................44 

Figure 3.6 Installation of quartz wall TCs and light pipes into ceramic positioning block. ...45 

Figure 3.7 Detail of quartz tube wall TCs and light pipe installation. ....................................46 

Figure 3.8 Initial test setup for Heater Design Test 2. ............................................................47 

Figure 3.9 Heater currents during time of Zircaloy ignition with major experimental 

events noted. ..........................................................................................................50 

Figure 3.10 Voltage applied to heater rods. ..............................................................................50 

Figure 3.11 Oxygen concentration of off gas with major experimental events noted. .............51 

Figure 3.12 Axial temperature profile with major experimental events noted. ........................52 

Figure 3.13 First derivative of axial temperature profile with major experimental events 

noted. ......................................................................................................................52

Figure 3.14 Radial temperature profile at the 12-in., 24-in., and 36-in. elevations. .................53 

Figure 3.15 Axial temperature profile along quartz tube. .........................................................54 

Figure 3.16 Close-up of light pipes during burn. ......................................................................54 

Figure 3.17 Composite axial rod temperature profile using corrected wall temperature 

measurements. ........................................................................................................55 

Figure 3.18 Peak cladding temperature data and MELCOR calculation with and without 

post-breakaway oxidation kinetics. ........................................................................56 

Figure 3.19 Outer insulation jacket after burn. .........................................................................57 

Figure 3.20 Inlet gas temperature, exit gas temperature in chimney, and other 

miscellaneous temperatures. ..................................................................................58 

Figure 3.21 Post-test quartz tube and inner thermal radiation shield after removal of 

insulation. ...............................................................................................................59 

Figure 3.22 Detail of heater rod rubble. ....................................................................................60

Figure 3.23 Detail of heater rod rubble at the 36-in. level after bundle removal from test 

stand. ......................................................................................................................61

Figure 4.1 Experimental apparatus showing as-built port locations. ......................................64 

Figure 4.2 Pressure traces recorded during the blocked water rod testing for 

measurements across 2–8 and 8–17. ......................................................................65 

Figure 4.3 Acrylic sleeves are inserted over the top of the water rods to block the flow. ......66 

Figure 4.4 Pressure drop as a function of bundle velocity for the SFP assembly with 

blocked water rods. ................................................................................................68 

Figure 4.5 Hydraulic loss coefficients as a function of the upper, measured experimental 

flow range. .............................................................................................................69 

Figure 4.6 Cumulative pressure drop as a function of axial location for flow rates of 100, 

200, 300, 400, 500, and 600 slpm. .........................................................................72 

Figure 4.7 Pressure drop as a function of bundle velocity for the SFP assembly with 

unblocked water rods. ............................................................................................73 
 xi 

Figure 4.8 Total air flow rate through the assembly as a function of pressure drop across 

2–17 for blocked (closed) and unblocked (open) water rods. ................................75 

Figure 4.9 Fraction of flow rate through the water rods as a function of total flow rate as 

determined from the 2–17 pressure drop data. .......................................................76 

Figure 4.10 Pressure drop as a function of bundle velocity for the SFP assembly with 

unblocked water rods. ............................................................................................77 

Figure 4.11 Fraction of flow rate through the bypass holes as a function of total flow rate 

as determined from 1–B pressure drop data. .........................................................79 

Figure 5.1. Experimental mock fuel assembly showing internal, as-built (a) axial and 

(b) lateral thermocouple (TC) locations. ................................................................84 

Figure 5.2 Arrangement of insulation, radiation barriers, and external TCs. ..........................85 

Figure 5.3 Schematic showing the layout of the assembly inlet hot wires (hot wires 2 

and 4). ....................................................................................................................86 

Figure 5.4 Hot wire 4 velocity traverse across the assembly inlet pipe for input 

volumetric flow rates of 50 (red) and 125 slpm (blue). .........................................87 

Figure 5.5 Inlet hot wire calibrations for hot wire 2 (HW2) (diamonds) and hot wire 4 

(HW4) (squares) on 1/20/06 (closed) and 1/31/06 (open). ....................................88 

Figure 5.6 Schematic showing layout of the drain hole hot wire (HW1 and HW3). ..............88 

Figure 5.7 Drain hole hot wire calibrations for hot wire 1 (HW1) (diamonds) and hot 

wire 3 (HW3) (squares) on 1/20/06 (closed) and 1/31/06 (open). .........................89 

Figure 5.8. Schematic of the measurement of the pressure drop across the bypass holes for 

(a) calibration and (b) heated testing. ....................................................................90 

Figure 5.9. Annular flow rate as a function of the bypass pressure drop for the hydraulic 

assembly (diamonds), analytic solution (line), and the external bottom tie plate 

calibration (squares). ..............................................................................................91 

Figure 5.10. Schematic of the COBRA model layout. ...............................................................93 

Figure 5.11. Detail view of the COBRA bundle layout. ............................................................93 

Figure 5.12. Repeating patterns used in the MELCOR separate effects models. ......................95 

Figure 5.13. MELCOR nodalization of the full-length assembly experiment. ..........................97 

Figure 5.14 Transient thermal response at different axial locations for a power input of 

1370 W and closed bypass/open drains flow configuration. .................................98 

Figure 5.15 Transient thermal response at different axial locations for a power input of 

1370 W and open bypass/closed drains flow configuration. .................................99 

Figure 5.16. Comparison of the thermal transient response with the MELCOR and COBRA 

codes for a power input of 1370 W and open bypass/closed drains flow 

configuration. .......................................................................................................100 

Figure 5.17 Comparison of experimental (symbols) to MELCOR (solid lines) and 

COBRA (dashed lines) temperatures as a function of axial location at times of 

3 (), 6 (), 9 (), and 12 hours (). ...................................................................101 

Figure 5.18 Similar to the figure above for the open bypass/closed drains configuration. .....101 

Figure 5.19 PCT as a function of assembly input power for experimental (symbols), 

MELCOR (blue lines), and COBRA (red lines). .................................................102 

Figure 5.20 Volumetric flow rates as a function of assembly input power for experimental 

(symbols) and MELCOR (lines). .........................................................................103 

Figure 6.1 Schematic and pictures of the assembly design. ..................................................105 

Figure 6.2 Photographs showing the final insulated assembly. ............................................106 
 xii 

Figure 6.3 Overall TC layout. ...............................................................................................107 

Figure 6.4 Center assembly TC locations and naming conventions. ....................................107 

Figure 6.5 Peripheral assemblies TC locations and naming conventions. ............................108 

Figure 6.6 Schematic of the 14 configuration. ....................................................................109 

Figure 6.7 MELCOR nodalization of the full-length assembly experiment. ........................110 

Figure 6.8 Slab layout for the 14 assembly. ........................................................................113 

Figure 6.9 Individual rod array illustration. ..........................................................................114 

Figure 6.10 Symmetric peripheral bundle temperatures as a function of time for the 

powered test conducted on 3/21/06. .....................................................................115 

Figure 6.11 Power test 2/27/06 – 1.5 kW center assembly, 26-in. level. ................................116 

Figure 6.12 Power test 3/15/06 – 3 kW center assembly, 11-in. level. ...................................116 

Figure 6.13 Power test 3/15/06 – 3 kW center assembly, 26-in. level. ...................................117 

Figure 6.14 Power test 3/15/06 – 3 kW center assembly, 41-in. level. ...................................117 

Figure 6.15 Power test 3/21/06 – 3 kW center assembly, 11-in. level. ...................................118 

Figure 6.16 Power test 3/21/06 – 3 kW center assembly, 26-in. level. ...................................118 

Figure 6.17 Power test 3/21/06 – 3 kW center assembly, 41-in. level. ...................................119 

Figure 6.18 Power test 3/23/06 – 5 kW center assembly, 11-in. level. ...................................119 

Figure 6.19 Power test 3/23/06 – 5 kW center assembly, 26-in. level. ...................................120 

Figure 6.20 Power test 3/23/06 – 5 kW center assembly, 41-in. level. ...................................120 

Figure 6.21 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

11-in. level. ..........................................................................................................121 

Figure 6.22 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

26-in. level. ..........................................................................................................122 

Figure 6.23 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

41-in. level. ..........................................................................................................122 

Figure 6.24 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

11-in. level. ..........................................................................................................123 

Figure 6.25 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

26-in. level. ..........................................................................................................123 

Figure 6.26 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 

41-in. level. ..........................................................................................................124 

Figure 6.27 Effective two-surface emissivity as a function of the Zircaloy emissivity for 

stainless steel emissivities of 0.33, 0.38, and 0.43. ..............................................124 

Figure 6.28 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 

11-in. level. ..........................................................................................................125 

Figure 6.29 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 

26-in. level. ..........................................................................................................125 

Figure 6.30 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 

41-in. level. ..........................................................................................................126 

Figure 6.31 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 

for power test 3/21/06 – 3 kW center assembly, 11-in. level. .............................126 

Figure 6.32 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 

for power test 3/21/06 – 3 kW center assembly, 26-in. level. .............................127 

Figure 6.33 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 

for power test 3/21/06 – 3 kW center assembly, 41-in. level. .............................127 
 xiii 

Figure 6.34 Peripheral temperature profiles at 4.1 hours elapsed test time for the powered 

test conducted on 3/21/06. ...................................................................................128 

Figure 7.1 Schematic of externally mounted sensors. ...........................................................130 

Figure 7.2 In situ oxygen sensor seal and flow schematic. ...................................................131 

Figure 7.3 Plan view schematic of oxygen sensor installation. .............................................132 

Figure 7.4 Experimental mock fuel assembly showing internal, as-built (a) axial and 

(b) lateral TC locations. .......................................................................................133 

Figure 7.5 Schematic showing the layout of the assembly inlet hot wires 2 and 4. ..............134 

Figure 7.6 Inlet hot wire calibrations for hot wire 2 (diamonds) and hot wire 4 (squares). ..134 

Figure 7.7 Temperature as a function of time for the pre-ignition test with a power input 

of 2250 W. ...........................................................................................................135 

Figure 7.8 MELCOR predictions for temperature as a function of time for the preignition test with a power input of 2250 W. ........................................................136 

Figure 7.9 Flow rate and pressure drop across the bypass as a function of time for hot 

wire 2 (HW2) (blue), hot wire 4 (HW4) (red), MELCOR (green), and pressure 

gage (black). .........................................................................................................137 

Figure 7.10 Comparison of pre-ignition MELCOR predictions and experimentally 

measured assembly inlet flow rates as a function of bundle power. ....................138 

Figure 7.11 Temperature as a function of time at different axial levels during the ignition 

test. .......................................................................................................................139 

Figure 7.12 Signal and temperature response for the oxygen sensor located at z = 144 in.....140 

Figure 7.13 Comparison of the measured oxygen concentrations measured by the in situ

oxygen sensor and the flow-through oxygen monitor. ........................................141 

Figure 7.14 Measured oxygen concentration from the monitor devices. ................................142 

Figure 7.15 Comparison of MELCOR pre-test prediction with and without sensor leakage 

versus the measured assembly inlet flow rate. .....................................................143 

Figure 7.16 Comparison of MELCOR post-test prediction with additional sensor leakage 

versus the measured assembly inlet flow rate. .....................................................144 

Figure 7.17 Calculated leakage flow from the post-test enhanced leakage model. ................145 

Figure 7.18 Calculated temperature response from the posttest enhanced leakage model. ....146 

Figure 7.19 Measured median temperature response. .............................................................147 

Figure 7.20 Flow rate as a function of time for the experiment hot wires and the final posttest MELCOR model. ..........................................................................................148 

Figure 7.21 PCT for both the ignition experiment and the final post-test MELCOR model. .149 

Figure 7.22 Breakaway lifetime function as a variable of time in the MELCOR ignition 

model....................................................................................................................149 

Figure 7.23 Cladding temperatures at z = 24 in. for both experiment and MELCOR. ...........150 

Figure 7.24 Cladding temperatures at z = 48 in. for both experiment and MELCOR. ...........150 

Figure 7.25 Cladding temperatures at z = 72 in. for both experiment and MELCOR. ...........151 

Figure 7.26 Cladding temperatures at z = 96 in. for both experiment and MELCOR. ...........151 

Figure 7.27 Cladding temperatures at z = 119 in. for both experiment and MELCOR. .........152 

Figure 7.28 Cladding temperatures at z = 144 in. for both experiment and MELCOR. .........152 

Figure 7.29 Postmortem of the assembly after ignition. .........................................................153 

Figure 8.1 Left: Expanded schematic of the apparatus construction. Upper right: Photo of 

the center assembly under construction. Lower right: Photo of the 

undercarriage of the apparatus. ............................................................................156 
 xiv 

Figure 8.2 Flow diagram of mass flow controllers, air heaters, and assemblies. ..................157 

Figure 8.3 Photographs of the final assembly with insulation applied around (a) the 

perimeter and (b) the lower portion. ....................................................................158 

Figure 8.4 Overall type K (73K to 1523K temperature range) TC layout. ...........................159 

Figure 8.5 Center assembly TC locations and naming conventions. ....................................159 

Figure 8.6 TC locations and naming conventions in the peripheral assemblies. ..................160 

Figure 8.7 TC locations and naming conventions in the channel canisters and pool cells. ..161 

Figure 8.8 AMI Inc. Model 65 oxygen monitors. .................................................................162 

Figure 8.9 Oygen sampling locations for each assembly as shown in (a) the as-built 

apparatus and (b) a schematic with sampling tube penetrations indicated in 

red. .......................................................................................................................162

Figure 8.10 Center assembly input power for the ignition test. ..............................................164 

Figure 8.11 Air flow rates for the different channels inside the apparatus during the 

ignition test...........................................................................................................164 

Figure 8.12 Air inlet temperatures for the different channels inside the apparatus during 

the ignition test. ....................................................................................................165 

Figure 8.13 Response of all oxygen monitors as a function of time. ......................................166 

Figure 8.14 Overhead images of the test apparatus at elapsed test times of (a) 5 hours and 

(b) 7 hours. ...........................................................................................................167 

Figure 8.15 Temperature of the GX rod in the center assembly as a function of time for 

different axial locations. .......................................................................................168 

Figure 8.16 Temperature of the EZ and EQ rods in the west and east assemblies, 

respectively, as a function of time for different axial locations. ..........................169 

Figure 8.17 Detailed temperature of the peripheral assembly as a function of time at z = 

10.6 in. .................................................................................................................169 

Figure 8.18 Temperature across the central portion of the apparatus as a function of time 

at z = 10.6 in. ........................................................................................................172 

Figure 8.19 Temperature across the peripheral portion of the apparatus as a function of 

time at z = 10.6 in. ................................................................................................172 

Figure 8.20 Temperature across the central portion of the apparatus as a function of time 

at z = 25.6 in. ........................................................................................................173 

Figure 8.21 Temperature across the peripheral portion of the apparatus as a function of 

time at z = 25.6 in. ................................................................................................173 

Figure 8.22 Temperature across the central portion of the apparatus as a function of time 

at z = 40.6 in. ........................................................................................................174 

Figure 8.23 Temperature across the peripheral portion of the apparatus as a function of 

time at z = 40.6 in. ................................................................................................174 

Figure 8.24 Post-ignition test photographs of the apparatus with the insulation removed. ....176 

Figure 8.25 Post-ignition overhead photographs of (a) the apparatus and (b) the inlet air 

flow boxes. ...........................................................................................................177 

Figure 8.26 Comparison of air flow rates for the different channels inside the test 

apparatus and the MELCOR model. ....................................................................178 

Figure 8.27 Comparison of air inlet temperatures for the different channels inside the test 

apparatus and the MELCOR model. ....................................................................179 

Figure 8.28 Default emissivity model in MELCOR with high temperature correction. .........180 
 xv 

Figure 8.29 Comparison of oxygen concentrations for the different assemblies inside the 

test apparatus and the MELCOR model. .............................................................181 

Figure 8.30 Comparison of temperatures inside the center assembly and MELCOR model 

for z = 4 to 14 in. ..................................................................................................182 

Figure 8.31 Comparison of temperatures inside the center assembly and MELCOR model 

for z = 28 to 38 in. ................................................................................................183 

Figure 8.32 Comparison of temperatures inside the center assembly and MELCOR model 

for z = 37 to 47 in. ................................................................................................183 

Figure 8.33 Comparison of temperatures inside the peripheral assembly and MELCOR 

model....................................................................................................................184 

Figure 8.34 Comparison of CTE prediction with Kearns’ model for Zircaloy in the 

circumferential direction. .....................................................................................186 

Figure 8.35 Modification of the default MELCOR oxidation kinetics model. .......................186 

Figure 8.36 Average and PCT as a function of time in the center assembly at z = 37.6 in. ...187 

Figure 8.37 Comparison of oxygen concentrations for the different assemblies inside the 

test apparatus and the MELCOR sensitivity model. ............................................188 

Figure 8.38 Comparison of the maximum and average temperatures in the center assembly 

to the MELCOR baseline and sensitivity models. ...............................................189 

Figure 8.39 Comparison of the maximum and average temperatures in the peripheral 

canister to the MELCOR baseline and sensitivity models. ..................................189 

Figure 8.40 Comparison of the maximum and average temperatures in the peripheral 

assemblies to the MELCOR baseline and sensitivity models. .............................190 

Figure 8.41 Effect on temperature response in the full-scale Zircaloy MELCOR model to 

the modified oxidation kinetics model. ................................................................190 

 xvii 

LIST OF TABLES 

Table 1.1 Summary of the testing elements in the experimental program. .............................2 

Table 2.1 Composition of heater rods. .....................................................................................6 

Table 2.2 Correlation of assembly and rod power versus the required voltage. ......................6 

Table 2.3 Heater Design Test 1 instrument layout. ................................................................10 

Table 2.4 Chronology of events during experiment. .............................................................20 

Table 3.1 Composition of heater rods. ...................................................................................38 

Table 3.2 Heater Design Test 2 instrument layout. ................................................................40 

Table 3.3 Chronology of events during experiment. .............................................................49 

Table 4.1 Dimensions of assembly components in the 99 BWR. ........................................63 

Table 4.2 Summary of the flow areas and hydraulic diameters of the SFP experiment. .......66 

Table 4.3 SLAM and k coefficient analysis data for pressure drops between 2–8, 8–17, 

and 2–17 for blocked water rods. ...........................................................................68 

Table 4.4 SLAM and k values for different ranges of experimental air flow rate. ...................69 

Table 4.5 SLAM and k coefficients for the blocked water rod assembly. ................................70 

Table 4.6 Spacer SLAM and k coefficients calculated with the appropriate spacer 

hydraulic diameter and flow area. ..........................................................................70 

Table 4.7 Calculation of effective SLAM and k values for different span lengths. .................71 

Table 4.8 SLAM and k coefficient analysis data for pressure drops between 2–8, 8–17, 

and 2–17 for unblocked water rods. .......................................................................73 

Table 4.9 SLAM and k coefficients for the unblocked water rod assembly assuming all 

flow passes through the bundle. .............................................................................74 

Table 4.10 SLAM and k coefficients for the unblocked water rod assembly assuming equal 

air velocities in the bundle and water rods. ............................................................74 

Table 4.11 SLAM and k coefficient analysis data for pressure drops between 2–8, 8–17, 

and 2–17 for unblocked water rods. .......................................................................77 

Table 4.12 SLAM and k coefficients for the unblocked water rod assembly using the 

experimentally determined bundle flow rate (Qtot – Qwr). .....................................78 

Table 5.1 Dimensions of assembly components in the 99 BWR. ........................................83 

Table 5.2 SLAM and k analysis across 1–B and 17–B for Incoloy and stainless steel 

assemblies. .............................................................................................................85 

Table 5.3 Integrated flow rates from the hot wire velocity profiles. .....................................87 

Table 5.4 Summary of all thermal-hydraulic tests for the SFP experiment. ........................104 

Table 6.1 Summary of all 14 thermal radiation testing on the Incoloy short stack

assembly. ..............................................................................................................114 

Table 7.1 Elapsed experimental time to reach 15% oxygen concentration for the oxygen 

sensors and monitors. ...........................................................................................142 

Table 8.2 Elapsed experimental time to reach reduced oxygen concentrations for the 

oxygen monitors...................................................................................................166 

Table 8.3 Measured burn rate at different axial heights based on an ignition criterion of 

1220 K. .................................................................................................................170 

Table 8.4 Measured temperatures across the apparatus at 4.82 hours. ................................175 

Table 9.1 Summary of the testing elements in the experimental program. .........................193 

 xix 

EXECUTIVE SUMMARY 

Spent fuel pools (SFP) are robust structures with an extremely low likelihood of a complete loss 

of coolant under traditional accident scenarios. However, in the wake of the terrorist attacks of 

September 11, 2001, the SFP accident progression was reevaluated using best-estimate accident 

codes. In 2001, the United State Nuclear Regulatory Commission (NRC) staff performed an 

evaluation of the potential accident risk in a spent fuel pool (SFP) at decommissioning plants in 

the United States [1]. This evaluation is documented on NUREG-1738, “Technical Study of 

Spent Fuel Pool Accident Risk at Decommissioning Nuclear Power Plants,” (ML010430066). 

The best-estimate computer code studies identified various modeling and phenomenological 

uncertainties that prompted a need for experimental confirmation. This experimental program 

was undertaken to address thermal-hydraulic issues associated with complete loss-of-coolant 

accidents in boiling water reactor (BWR) SFPs. 

The objective of this project was to provide basic thermal-hydraulic data associated with a SFP 

complete loss-of-coolant accident. The accident conditions of interest for the SFP were simulated 

in a full-scale prototypic fashion (electrically-heated, prototypic assemblies in a prototypic SFP 

rack) so that the experimental results closely represent actual fuel assembly responses. A major 

impetus for this work was to facilitate code validation (primarily MELCOR) and reduce 

modeling uncertainties within the code. 

The Spent Fuel Pool Heatup and Propagation Phenomena project was conducted from April 

2004 until November 2006 over which time seven unique experimental apparatuses were tested 

at Sandia National Laboratories in Albuquerque, New Mexico. In addition to these experiments, 

extensive simulation efforts were undertaken with the MELCOR severe accident modeling code 

to understand and predict the behavior observed in the tests. The key findings from this 

integrated experimental and simulation program are: 

 Electrically heated spent fuel rod simulators can be fabricated with Zircaloy cladding to 

accurately represent the decay heat, thermal mass and Zircaloy reactivity of a prototypic 

spent fuel rod. 

 The measured form and friction loss coefficients of a prototypic BWR assembly were 

significantly different from generally accepted values. Use of the measured coefficients was 

vital for accuracy when calculating (with MELCOR) the naturally induced flow in a heated, 

prototypic BWR assembly. 

 Incorporation of “breakaway” Zircaloy oxidation kinetics into MELCOR was vital for 

accurately capturing the Zircaloy heat-up to ignition and oxygen consumption. 

 For the full length ignition test, the MELCOR model predicted the peak cladding 

temperature (PCT) of the assembly to within 40 K at all times and the time of ignition to 

within 5 minutes. 

 For the 1×4 ignition experiment, the standard MELCOR model predicted ignition in the 

center and peripheral assemblies to within 30 and 15 minutes, respectively. The error in 
 xx 

ignition timing between the simulations and experiment is approximately 10%. The 

difference in timing is likely due to the inability of the lumped parameter approach used in 

MELCOR to account for steep radial temperature gradients. 

 Post-mortem examination of the integral test assemblies revealed gross distortion of the 

pool rack and channel box, rubblization of the tubing bundle and accumulation of debris on 

the bottom tie plate that resulted in flow blockage. Flow blockage was also evident from 

molten aluminum (originating from Boral plates built into the pool rack) that collected on 

and below the bottom tie plates. 
 xxi 

ABBREVIATIONS/DEFINITIONS 

%LEL percent lower explosive limit 

%O2 percent oxygen 

ANL Argonne National Laboratory 

BE best estimate 

BWR boiling water reactor 

COBRA Coolant Boiling in Rod Arrays: thermal-hydraulic and radiative analysis 

code 

COBRA-SFS Coolant Boiling in Rod Arrays-Spent Fuel Storage 

COR CORE – fuel response model in MELCOR 

Cr chromium 

CTE coefficient of thermal expansion 

CVH control volume hydrodynamics model in MELCOR 

CYBL cylindrical boiling 

DAQ data acquisition 

EPA error propagation analysis 

Fe iron 

HS heat structure 

HW hot wire 

ID inside diameter 

IR infrared 

LEL lower explosive limit 

MELCOR severe accident analysis code 

MgO magnesium oxide 

Ni nickel 

NPT national pipe thread 

NRC Nuclear Regulatory Commission 

O2 oxygen 

OD outside diameter 

PCT peak cladding temperature 

PWR pressurized water reactor 

RADGEN view factor input generation code for COBRA 

RMS root mean square 

scfm standard cubic feet per minute (standard defined at 0oC and 1 atm) 

SEM scanning electron microscopy 

SFP spent fuel pool 

slpm standard liters per minute (standard defined at 0oC and 1 atm) 

Sn tin 

SNL Sandia National Laboratories 

TC thermocouple 

vol% volume percent 

wt% weight percent 

Zr Zirconium 

1 

1 INTRODUCTION 

1.1 Background 

In 2001, the United States Nuclear Regulatory Commission (NRC) staff performed an evaluation 

of the potential accident risk in a spent fuel pool (SFP) at decommissioning plants in the United 

States [1]. NUREG-1738 was prepared to provide a technical basis for decommissioning 

rulemaking for permanently shutdown nuclear power plants. The study described a modeling 

approach of a typical decommissioning plant with design assumptions and industry 

commitments, thermal-hydraulic analyses performed to evaluate spent fuel stored in the SFP at 

decommissioning plants, risk assessment of SFP accidents, consequence calculations and the 

implications for decommissioning regulatory requirements. It was known that some of the 

assumptions in the accident progression in NUREG-1738 were necessarily conservative, 

especially the estimation of the fuel damage. Furthermore, the NRC desired to expand the study 

to include accidents in the SFPs of operating power plants. Consequently, the NRC continued 

SFP accident research by applying best-estimate computer codes to predict the severe accident 

progression following various postulated accident initiators. The best-estimate computer code 

studies identified various modeling and phenomenological uncertainties that prompted a need for 

experimental confirmation [2]. The present experimental program was undertaken to address 

thermal-hydraulic issues associated with complete loss-of-coolant accidents in boiling water 

reactor (BWR) SFPs. 

1.2 Objectives and Scope of Testing 

The objective of this project was to provide basic thermal-hydraulic data associated with a SFP 

complete loss-of-coolant accident. The accident conditions of interest for the SFP were simulated 

in a full-scale prototypic fashion (electrically-heated, prototypic assemblies in a prototypic SFP 

rack) so that the experimental results closely represent actual fuel assembly responses. A major 

impetus for this work was to facilitate code validation (primarily MELCOR) and reduce 

questions about buoyancy driven flow, heat transfer modes, oxidation reaction kinetics and with 

the interpretation of the experimental results. It was necessary to simulate a cluster of assemblies 

to represent a higher decay (younger) assembly surrounded by older, lower power assemblies. 

Specifically, this program provided data and analysis confirming: (1) MELCOR modeling of 

inter-assembly radiant heat transfer, (2) flow resistance modeling and the natural convective flow 

induced in a fuel assembly as it heats up in air, (3) the potential for and nature of thermal 

transient (i.e., Zircaloy fire) propagation, and (4) mitigation strategies concerning fuel assembly 

management. 

1.3 Test Plan Synopsis 

Three configurations of fuel assemblies were utilized in the experimental testing. Initial tests 

were conducted in a small configuration in order to assess the performance and suitability of 

Zirconium clad electrically-heated spent fuel rod simulators. Two such tests were conducted in 

order to demonstrate that the heater design was capable of initiating a Zirconium fire. The second 

configuration examined a single, full-length highly prototypic fuel assembly (stainless steel, 

Zircaloy, and Incoloy versions) inside a prototypical pool rack cell. The stainless steel rod 

assembly was unheated and only used for the hydraulic characterization. The Incoloy fuel 

assemblies were used to conduct high temperature tests while minimizing complications of 

oxidizing Zircaloy surfaces. The Zircaloy version of the full-length fuel assembly was used for 
2 

tests that were taken to conditions of high temperature oxidation and ignition. The final 

configuration was five Zircaloy short (1/3 length) assemblies in a 33 pool rack. The short array 

of assemblies was designed to simulate a slice from the middle to upper portion of an array of 

full-length assemblies. This was accomplished in the final ignition test by accurately controlling 

both the flow rate and temperature of air introduced into the bottom of each partial assembly. 

Two types of testing were conducted: (1) separate effects tests and (2) integral effects tests. In 

the separate effects tests, the experiments were designed to investigate a specific heat transfer or 

flow phenomenon such as thermal radiative coupling or induced natural convective flow. These 

tests were non-destructive and involved some non-prototypic materials (e.g., stainless steel and 

Incoloy). This phase of testing involved conducting a greater number of less complicated tests. In 

the integral effects tests, all hydraulic flow and heat transfer phenomena were investigated 

simultaneously. These tests were specified with boundary conditions which led to the destruction 

of the experimental apparatus and thus involved a limited number of more complex experiments. 

Table 1.1 summarizes in chronological order the phased tests conducted as part of this 

experimental program. 

Table 1.1 Summary of the testing elements in the experimental program 

Description Purpose Assembly Rod material 

Heater Design Test electrical heater 

performance, preliminary 

data on zircaloy fire 

12 rod bundle Zircaloy 

Separate Effects Hydraulics – determine 

form loss and laminar 

friction coefficients 

Prototypic Stainless Steel 

Separate Effects Thermal hydraulics – 

Determine input 

conditions for partial 

length experiments 

Prototypic Incoloy 

Separate Effects Thermal radiation – 

radiation coupling in a 14 

arrangement 

Prototypic – Partial length Incoloy 

Integral Effects Axial Ignition – 

temperature profiles, 

induced flow, axial O2

profile, nature of fire 

Prototypic Zircaloy 

Integral Effects Radial Propagation – 

Determine nature of radial 

fire propagation 

Prototypic – Partial length Zircaloy 

A unique aspect of this project was the deliberate close coupling of the experiments with 

numerical analysis. The principal code used was the severe accident code MELCOR. At each 

step in the experimental program, MELCOR was used (1) as a tool for the experimental design, 

(2) for the pre-test results prediction, and (3) for post-test analysis of the calculated and measured 

responses. The post-test analysis helped identify important response parameters, which often led 

to improvements in the conduct of the next phase of testing and improvements in the modeling 

approach. The experimental and modeling findings and improvements from post-test analysis of 

the previous experimental program tests were subsequently used to design and predict the next 

phase of experiments, and so on. For example, the pressure drops measured in the hydraulic 

testing did not initially match the MELCOR predictions based on best-estimate hydraulic 
3 

parameters. When MELCOR was updated with the experimentally derived hydraulic parameters, 

the MELCOR predictions of the next thermal-hydraulic testing phase were in excellent 

agreement with the data. 

Figure 1.1 illustrates the closely coupled relationship between the various experiments and the 

MELCOR modeling. This figure also stresses the overall objective of this program to validate 

MELCOR with full-scale prototypic experimental measurements to reduce uncertainties in whole 

pool accident analyses. The phased, multi-scale test program and integrated relationships 

between the test program and the analytical modeling efforts were critical to the comprehensive 

understanding of the complex accident phenomena. 

Figure 1.1 Illustration of the closely-coupled relationship between the various testing elements and the 

supporting MELCOR modeling 

1.4 Report Organization 

This report is organized in a series of sections with each describing a particular experiment or 

phase of experiments. Section 2 describes the first heater design test where the design and 

physical properties of the heaters are described. Both the full oxygen test checkout test and 

Section 3 describe the second heater design test, which was conducted under reduced oxygen 

conditions. Section 4 describes the hydraulic characterization of the highly prototypic 99 BWR 

assembly. Next, Section 5 describes the thermal-hydraulic testing where the naturally induced 

flows inside a heated assembly are characterized. Section 6 describes the thermal radiative 

coupling experiments; Section 7 describes the ignition testing of a full-length Zircaloy assembly. 

Section 8 describes the ignition testing of the 14 array of short length Zircaloy assemblies. 

Finally, Section 9 summarizes the key findings. 

Heater Design Test 

Can the Zr be heated to 

ignition? 

Separate Effects Tests 

Integral Effects Tests 

MELCOR 

Calibrate hydraulics 

Model verification 

BWR SFP

Hydraulic Tests 

Determine form loss and 

laminar friction coefficients 

k’s and SLAM 

Hydraulic 

parameters 

Thermal Hydraulic Tests 

Determine natural draft flow 

characteristics and temperature 

profiles 

Test models 

Heater design 

Thermal Radiation Coupling 

Determine energy transfer to 

adjacent assemblies 

Test models 

Axial Ignition Characterization 

Full length Zr assembly 

Flow, temperature, 

O2 concentration, 

input parameters 

Radial Propagation 

Array of short length 

Zr assemblies 

Experimental 

design 

Induced flows 

Temp profiles 

Radial radiation 

coupling 

Whole Pool Analysis

Implications for fuel storage 

Pool arrangement 

Verify oxidation 

kinetics Prototypic 

ignition 

Estimate ignition 

test parameters 

Radial radiation 

effect on ignition 

in 1x4 

Experiment 

design 

Experiment 

design 

5 

2 HEATER DESIGN TEST 1 

2.1 Objective 

The main objective of the first heater design test was to verify that the heater design would 

produce high enough surface temperatures for long enough to ignite the Zircaloy cladding before 

the heater element fails. A secondary objective was to gain experience with monitoring a 

Zirconium fire and to identify important issues. 

The surrogate fuel assemblies used in this program consisted of a Zircaloy tubing jacket with a 

central Nichrome heater element surrounded by compacted magnesium oxide. The heaters were 

assembled by Watlow Electric Manufacturing Company (Watlow), a large manufacturer of 

industrial electric heaters. Watlow has considerable experience in manufacturing reliable heater 

rods. The Nichrome element melts when it reaches a temperature of about 1400C (1673 K). 

Typically, heaters of this type are rated for intermediate (~4 hours) duty at surface temperatures 

of 1000C (1273 K). This allows for a 400 K temperature drop between the Nichrome element 

and the surface of the heater jacket. The actual temperature drop between the jacket surface and 

the Nichrome element is highly dependent on the power output of the heater. The lower the 

power output, the lower the temperature drop, hence, the higher the temperature the heater jacket 

surface can reach before element failure. The power output of the heaters used in this study was 

at most 25 W/ft, which is considered very low (an order of magnitude lower than typical). Thus, 

surface temperatures of at least 1225C (1500 K) and perhaps as high as 1350C (1623 K) were 

expected before heater element failure. Depending upon the heat losses, the oxidation layer 

thickness, the ventilation characteristics (i.e., flow rate, inlet gas temperature, oxygen content), 

as well as other factors, 1225C could be above or below the temperature where rapid oxidation 

(i.e., ignition) occurs. 

2.2 Experimental 

2.2.1 Electric Heater Design 

The design of the Zircaloy jacketed heater was typical of those made by Watlow. Each heater 

was comprised of a central Nichrome heater element, compressed magnesium oxide powder, and 

an outer Zircaloy 2 jacket. The composition of the heater rods is listed in Table 2.1. The final 

magnesium oxide powder density was 2.720 g/cm3

 corresponding to a solids fraction of 0.759.
6 

Table 2.1 Composition of heater rods 

Component 

Nominal 

wt% 

Mass/rod 

(lb) 

Mass/rod 

(g) # Rods 

Total 

mass 

(g) 

Mwt 

(g/mol) 

Total 

moles 

(mol)

Zircaloy 2 0.368 

 Zr 98.40% 164.2513 12 1971.015 91.224 21.60632 

 Sn 1.30% 2.169986 12 26.03983 118.71 0.219357 

 Fe 0.18% 0.30046 12 3.605515 55.845 0.064563 

 Cr 0.10% 0.166922 12 2.003064 51.9961 0.038523 

 Ni 0.07% 0.116845 12 1.402145 58.6934 0.023889 

 100.05% 

MgO 0.377 171.0043 12 2052.052 40.3044 50.91385 

Nichrome 0.046 20.86525 12 250.383 

The diameter of the heater rods used in the heater design test bundle was smaller than prototypic 

due to the availability of Zircaloy tubing. The tubing used was 0.440-in. diameter, 0.028-in. wall 

Zircaloy 2 tubing for a boiling water reactor (BWR) assembly. This tubing was made into 48-in. 

long, 0.375-in. diameter heaters by Watlow in a process whereby the 0.440-in. tubing was drawn 

through a die that reduced the diameter to 0.375 in. and compressed the magnesium oxide 

powder considerably. Compression of the magnesium oxide powder is required to establish good 

thermal contact between the Nichrome heater element and the outer metallic sheath. 

Consideration was given to how well the electric heater would represent a spent fuel rod from a 

thermal perspective. The power produced by a spent fuel rod depends primarily on the age of the 

spent fuel rod, which is defined as the time since it was removed from the reactor. Table 2.2 

shows a typical BWR spent fuel assembly power for various ages along with the corresponding 

power per rod and linear foot . The electric heater rods were designed to produce 25 W/ft at 120 

volts. With this design, spent fuel assemblies with ages from as short as three days old to over 

two years old could easily be simulated. 

Table 2.2 Correlation of assembly and rod power versus the required voltage 

Time 

(days) 

Assembly 

peak power 

(kW) 

Total rod 

power 

(W) 

Linear rod 

power 

(W/ft) 

Voltage 

(V) 

3 23.93 323.40 25.87 122.07 

10 15.01 202.84 16.23 96.68 

100 5.17 69.80 5.58 56.72 

365 2.30 31.03 2.48 37.81 

730 1.33 17.91 1.43 28.73 

In order to simplify the heater rod fabrication, the power was distributed linearly along the 

heated length of the rod. A typical spent fuel rod produces a chopped cosine power distribution 

with 130% peak power at the center and 60% power at the ends. The difference between the 

cosine power distribution of a typical spent fuel rod and the linear power distribution of the 

electric heater rod is shown in Figure 2.1. The discrepancy is not great and can be accounted for 

in the model validation effort. 
7 

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 0.2 0.4 0.6 0.8 1

Elevation (Z/Zmax)

Cum Power (W/Wtot)

typical spent fuel rod

electric heater rod

Figure 2.1 Cumulative power distribution along a typical spent fuel rod and the electric heater rod 

Another important thermal property of the electric heater rod is the thermal mass. This is best 

characterized by the specific heat (on a volume basis) and is shown in Figure 2.2 for spent fuel, 

compacted magnesium oxide (based on the as-built compacted magnesium oxide density), and as 

aluminum oxide. The agreement between magnesium oxide and spent nuclear fuel is very close 

over the entire temperature range making the magnesium oxide an ideal surrogate. 

2.0E+06

2.2E+06

2.4E+06

2.6E+06

2.8E+06

3.0E+06

3.2E+06

3.4E+06

3.6E+06

3.8E+06

4.0E+06

0 500 1000 1500

Tem p (K)

rho Cp (J/m3/K)

Melcor spent fuel

MgO heater

Al2O3 heater

Figure 2.2 Specific heat of electric heaters compared to spent nuclear fuel 

2.2.2 Initial Experimental Assembly 

The experiment was constructed and operated in a cylindrical boiling (CYBL) vessel located in 

Sandia National Laboratories (SNL) in Albuquerque, New Mexico. Twelve 4-ft. long, 0.375-in. 

diameter heater rods with a pitch of 0.539 in., as shown in Figure 2.3, were assembled into a 

bundle (Figure 2.4). The heater rods were instrumented with 0.032-in. ungrounded Incoloy 

sheathed type K thermocouples (TCs) in a configuration as shown in Table 2.3. Table 2.3 also 
8 

lists other instrumentation used in the test and was collected by a computerized data acquisition 

system at nominally 1-second intervals. The TCs were attached to the Zircaloy tubing by spot 

welding a small piece of stainless steel foil across the tip, as shown in Figure 2.5. A 57-mm ID 

(2.244-in.) quartz tube was placed around the heater bundle, and the tube was surrounded with 

reflective foil (Figure 2.6 and Figure 2.7). Initially, the bundle was insulated with 3 in. of 

KaowoolTM. A small axial slit through the insulation and foil was made to permit viewing (see 

Figure 2.3 and Figure 2.8). However, in initial testing of this setup, the temperature rise rate was 

low indicating that the heat loss was too high. The initial test setup is shown in Figure 2.9. 

Plan View

Cross View

Side view 

 window 

Insulation

Heater 

rod

~2” channel 

Figure 2.3 Schematic of the heater design test 
9 

Figure 2.4 Bundle of heater rods 
10 

Table 2.3 Heater Design Test 1 instrument layout 

Heater Design Test Instrumentation 

Chan 

# Name Description 

Instrument 

type 

Max 

Volt Units 

0 D0 Heater Axial Elevation 0" K TC 

1 D6 Heater Axial Elevation 6" K TC 

2 D12 Heater Axial Elevation 12" K TC Key 

3 D18 Heater Axial Elevation 18" K TC A B 

4 D24 Heater Axial Elevation 24" K TC C D E F 5o SLIT 

5 D30 Heater Axial Elevation 30" K TC W X Y Z 

6 D36 Heater Axial Elevation 36" K TC G H 

7 D42 Heater Axial Elevation 42" K TC C 

8 D48 Heater Axial Elevation 48" K TC C 

9 W12 W Heater Elevation 12" K TC C 

10 X12 X Heater Elevation 12" K TC C 

11 Y12 Y Heater Elevation 12" K TC C 

12 Z12 Z Heater Elevation 12" K TC C 

13 W24 W Heater Elevation 24" K TC C 

14 X24 X Heater Elevation 24" K TC C 

15 Y24 Y Heater Elevation 24" K TC C 

16 Z24 Z Heater Elevation 24" K TC C 

17 W36 W Heater Elevation 36" K TC C 

18 X36 X Heater Elevation 36" K TC C 

19 Y36 Y Heater Elevation 36" K TC C 

20 Z36 Z Heater Elevation 36" K TC C 

21 F24 F Heater Elevation 42" K TC C 

22 INLET Air inlet temperature K TC C 

23 OVEN Oven temperature K TC C 

24 Chimney Chimney Air temperature K TC C 

25 A current Current of Heater A Current xducer 5 ampere 

26 B current Current of Heater B Current xducer 5 ampere 

27 C current Current of Heater C Current xducer 5 ampere 

28 D current Current of Heater D Current xducer 5 ampere 

29 E current Current of Heater E Current xducer 5 ampere 

30 F current Current of Heater F Current xducer 5 ampere 

31 G current Current of Heater G Current xducer 5 ampere 

32 H current Current of Heater H Current xducer 5 ampere 

33 W current Current of Heater W Current xducer 5 ampere 

34 X current Current of Heater X Current xducer 5 ampere 

35 Y current Current of Heater Y Current xducer 5 ampere 

36 Z current Current of Heater Z Current xducer 5 ampere 

37 Oxygen Oxygen in chimney Oxygen percent 

38* Air Flow Mass air flow into oven MKS Flow 10 slpm 

39** Out 42 Outer skin temp 42" K TC C 

*Not used 

**Final assembly only 
11 

Figure 2.5 TC attachment detail 

Figure 2.6 Quartz tube installed over heater bundle 
12 

Figure 2.7 Stainless steel foil in place over quartz tube 
13 

Figure 2.8 Insulation and side-viewing slit 
14 

Figure 2.9 Initial test setup 

2.2.3 Final Experimental Assembly 

After the initial heat test indicated a heat-up rate too low, the test was terminated, and the 

assembly was allowed to cool. It was decided to increase the insulation and abandon the sideviewing slit. The viewing slit was filled with insulation and a stainless steel foil wrap was placed 

around the initial 3-in. layer of insulation. An additional 3-in. wrap of insulation was added and a 

final stainless steel sheet metal wrap was added. In addition, the top plate and chimney were also 

insulated. The final test setup is shown in Figure 2.10. 
15 

Figure 2.10 More insulated final test setup 

2.2.4 Additional Sampling 

In addition to the data collected by the instrumentation shown in Table 2.3, three other sampling 

efforts collected data for this experiment. The first effort was to monitor the oxygen 

concentration of the exit gas, the second effort was the sampling of smoke particulates for metals 

analysis, and the third was video recording of the experiment from a number of perspectives. 

Oxygen: The oxygen concentration of the gas leaving the test bundle was monitored using an 

ITX (S/N: 0307010-286) portable atmosphere-monitoring instrument. The instrument measured 

both oxygen concentration and the percent lower explosive limit (%LEL). The sample was 

collected from the mid point of the chimney through stainless steel tubing that was routed 

through an ice bath and outside the CYBL vessel where the oxygen monitor was located. The 

sample was drawn using the instrument’s internal sample pump. The instrument’s sampling 

system included a small particulate filter that plugged shortly after smoke was emitted from the 

test bundle. The percent oxygen (%O2) and %LEL data were collected internally along with the 

time at 5-second intervals. 

Smoke: Health physics personnel installed two particulate sampling devices inside the CYBL 

vessel to analyze the total metal content of the smoke generated by the experiment. One device 

was located about 3 ft. above the chimney and the other was located about midway between the 

top of the chimney and the upper lip of the CYBL vessel. The sample filters were analyzed for 

the metals added to Zirconium to make the Zircaloy 2 alloy, which include tin, iron, nickel, and 

chromium. The samples were also analyzed for magnesium, a principal constituent of the 
16 

magnesium oxide insulation used inside the heaters. Table 2.1 presents the initial mass of each of 

these constituents that was initially present in the test bundle. 

Video: Video perspectives included a close-up looking down the chimney, a side view of the 

experimental setup, and a distant top view from the top opening of the CYBL vessel. 

2.2.5 Initial Test Operation 

The initial experimental setup was tested on December 9, 2004. The air preheater (and data 

collection system) was started at 11:06 a.m., and the air flow of 1.0 scfm was started at 11:13 

a.m. The preheat air temperature set point was 600 K (327C). The power to the heater rods was 

applied at 1 p.m. (114 minutes into experiment). The initial power was a total of 800 W, which 

corresponds to about 16.6 W/ft. At 2:24 p.m. (198 minutes) the total power was increased to 

1200 W or 25 W/ft. At 2:45 p.m., the power was terminated concluding the initial test. 

The data collected during this initial testing follows. Figure 2.11 shows the inlet and exit air 

temperatures. Figure 2.12 shows the current history of the test bundle. Figure 2.13 and Figure 

2.14 show the axial and radial temperature profiles. Figure 2.15 shows the rate of change of the 

axial temperature profile. 

Air Temp

0

50

100

150

200

250

300

350

400

450

500

0 50 100 150 200 250

time (min)

air temp (C)

inlet air temp

"exit air temp"

Figure 2.11 Temperature of air entering the bottom of the assembly and temperature of the air leaving 

the top of the assembly during the initial test operation 
17 

current history

0

250

500

750

1000

1250

1500

1750

2000

0 50 100 150 200 250

time (min)

current (mA)

A CURRENT

B CURRENT

D CURRENT

C CURRENT

E CURRENT

F CURRENT 

G CURRENT

W CURRENT

X CURRENT

Y CURRENT

Z CURRENT 

Figure 2.12 History of the electrical current supplied to each of the twelve heater rods during the initial 

test operation 

Axial Clad Temp Profile

0

100

200

300

400

500

600

700

800

900

1000

0 50 100 150 200 250

time (min)

Clad temp (C)

D48

D42

D36

D30

D24

D18

D12

D6

D0

Figure 2.13 Axial temperature profile developed during the initial test operation 
18 

Radial Clad Temp Profiles

0

100

200

300

400

500

600

700

800

900

1000

0 50 100 150 200 250

time (min)

Clad temp (C)

W36

X36

Y36

Z36

W24

X24

Y24

Z24

W12

X12

Y12

Z12

F 42

Figure 2.14 Radial temperature profiles developed during the initial test operation 

Axial Temp Rise on Rod D

-0.8

-0.6

-0.4

-0.2

0

0.2

0.4

0.6

0 2000 4000 6000 8000 10000 12000 14000 16000

time (sec)

Temp Change (K/s)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

Figure 2.15 Axial differential temperature change measured during the initial test operation 
19 

2.2.6 Initial Test Analysis 

In order to ensure ignition of the Zirconium cladding, the best available MELCOR model 

indicated the clad temperature needed to reach 1000C at a rate of at least 0.1 K/s. If the heat-up 

rate was not high enough, a Zirconium oxide layer could grow such that ignition would not be 

possible at the maximum power rating of the heaters. Figure 2.13 shows that the clad peak 

temperature was approaching 800C at 198 minutes into the test. Figure 2.15 shows that the heatup rate at this time (11,000 seconds) was only about 0.04 K/s, far below the 0.1 K/s needed. At 

this point, it was decided to increase the power to 1200 W. This increased the heat-up rate to 

0.12 K/s; but as the clad temperature approached 900C, the heat-up rate fell back to 0.04 K/s. 

Thus, ignition was not ensured so the test was terminated. The results of this initial testing 

clearly indicated that the heat losses were too great to ensure ignition of the heater rods. 

2.2.7 Final Test Operation 

The initial testing, though terminated, brought the Zirconium heater rods to sufficient 

temperature long enough to produce a significant oxide layer. This oxide layer would make 

igniting the rods more difficult than before. A number of measures were taken to ensure ignition 

during the final test even with this pre-established oxide layer. As described above, the insulation 

around the test bundle was more than doubled, and the side-viewing slit was abandoned. It was 

decided that the inlet air temperature would be increased to 750 K (477C) and that the heater 

power could be increased to 120% of rating without compromising the heaters. Finally, 

convective losses would be reduced and further oxide buildup minimized by reducing the air 

flow until sufficient temperatures were reached. 

Table 2.4 lists the chronology of events during the final test. Many of the main events are noted 

in Figure 2.16, which shows the temperature change of the axial profile. The data shown on this 

plot is the calculated first derivative of the temperature data over a moving 10-sample (nominally 

10-second) window. Presenting the temperature data as the first derivative most clearly denotes 

the temperature response to the events listed in Table 2.4. A complete record of all the data 

collected during the final test can be found in Appendix A, SFP2 Data Summary. 
20 

Table 2.4 Chronology of events during experiment 

Time 

(s) 

Event 

Value Event Description 

0 Inlet temperature set to 600 K, flow 1.0 scfm 

4813.5 Inlet temperature increased to 750 K 

6433.50 O2 data on (12:50 on O2 data file, 3 min difference) 

6485.70 800 W on heater rods (16.7 W/ft) 

7445.70 0.23 K/s 

8405.70 0.18 K/s 

9125.70 0.126 K/s 

9339.40 0.1 scfm air inlet (13:39:42 O2 time) 

9639.40 0.11 K/s 

9809.20 1200 W on heater rods (25 W/ft) 

10049.20 20.6 % O2

10169.20 0.17 K/s (20.5% O2) 

10889.20 20.4 %O2

11309.20 20.3 %O2

11549.20 20.2 %O2

11909.20 20 %O2

11969.20 19.9 %O2

12022.50 1440 W (94 min experiment time) 

12082.50 19.8 %O2

12142.50 19.7 %O2

12308.50 1 scfm air inlet IGNITION (14:28:02 O2 time) 

12628.50 8 %O2

12688.50 7.5 %O2 SMOKE (14:34 O2 time) 

12808.50 4.8 %O2 (filter may be plugged) 

12868.50 4.1 %O2 (restart timer) 

12988.50 2.8 %O2 (filter may be plugged) 

13948.50 20.5 %O2 (this sample w/o filter) 

15557.80 0 scfm air inlet (15:22:56 video time) 
21 

Axial Temp Change on Rod D

-0.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

5000 6000 7000 8000 9000 10000 11000 12000 13000 14000 15000 16000 17000

time (sec)

Temp Change (C/s)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

800W 0.1 scfm 1200W 1440W

 0.0 

 1.0 scfm scfm

Figure 2.16 First derivative of axial temperature profile with major experimental events noted 

The test was started on December 10, 2004, at 11 a.m. The inlet air flow was 1.0 scfm and 750 K 

(477C). Power was applied to the heater rods at a level of 800 W at about 12:50 p.m. At 1:30 

p.m., peak temperatures were approaching 700C, but heat rates were dropping to 0.1 K/s. At 

this time, the air flow rate was reduced to 0.1 scfm. This had only a minimal effect on the heat 

rate of the hotter rods; therefore, at about 1:45 p.m., the power was increased to 1200 W. At 

about 2:20 p.m., the peak temperatures were approaching 1000C, but the heat rate had dropped 

to less than 0.1 K/s. At this point, the power was increased to 1440 W, 120% of power rating. 

Five minutes later, with clad temperatures at 1000C, the air was increased back to 1 scfm. 

Ignition was immediately evident as a bright glow in the chimney, and at about 2:30 p.m., smoke 

was first seen (Figure 2.17). The burn continued until about 3:20 p.m., when the air flow was 

terminated and the test was allowed to cool off (Figure 2.18). 
22 

Figure 2.17 Close-up of chimney during ignition initiation 
23 

Figure 2.18 Smoke filled CYBL vessel during burn 

2.3 Data Analysis 

2.3.1 Ignition Details 

As shown in Figure 2.16, the temperature data after ignition and during the burn period is very 

complicated. Figure 2.19 through Figure 2.28 show 1000-second intervals of the data at the time 

of ignition for various locations in the bundle. From this detailed examination of the data, it can 

be determined when and where the ignition occurred and when the burn reached various 

thermocouples, as well as some insights gained into some complicating events that occurred 

during the burn. 
24 

Exit Gas

0

5

10

15

20

25

12000 12200 12400 12600 12800 13000

Time (sec)

Gas Conc (%)

O2(%)

LEL(%)

Air back up 

 to 1 cfm smoke

Figure 2.19 Oxygen concentration of off gas during time of Zircaloy ignition 

Heater Currents

1400

1450

1500

1550

1600

1650

1700

1750

1800

12000 12200 12400 12600 12800 13000

time (sec)

Current (mA)

A CURRENT

B CURRENT

C CURRENT

D CURRENT

E CURRENT

F CURRENT 

G CURRENT

H CURRENT

W CURRENT

X CURRENT

Y CURRENT

Z CURRENT 

Air back up 

 to 1 cfm smoke

Figure 2.20 Heater currents during time of Zircaloy ignition 
25 

Axial Temperatures on Rod D

0

200

400

600

800

1000

1200

11900 12000 12100 12200 12300 12400 12500 12600 12700 12800 12900

time (sec)

Temp (C)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

smoke 1440W 1.0 scfm

Figure 2.21 Axial temperature profile during time of ignition 

Axial Temp Change on Rod D

-0.6

-0.4

-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

11900 12000 12100 12200 12300 12400 12500 12600 12700 12800 12900

time (sec)

Temp Change (C/s)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

 1.0 scfm 1440W smoke

Figure 2.22 First derivative of axial temperature profile during time of ignition 
26 

Radial Temp Profile @ 36"

300

400

500

600

700

800

900

1000

1100

1200

1300

11700 11800 11900 12000 12100 12200 12300 12400 12500 12600 12700

Time (sec)

Temp (C)

W 36

X 36

Y 36

Z 36 

D 36

smoke 1.0 scfm

Figure 2.23 Radial temperature profile at the 36-in. elevation during ignition 

Temp Change @ 36"

-0.4

-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

11700 11800 11900 12000 12100 12200 12300 12400 12500 12600 12700

Time (sec)

Temp Change (C/s)

W 36

X 36

Y 36

Z 36 

D 36

1440W 1.0 scfm

Figure 2.24 First derivative of radial temperature profile at the 36-in. elevation during ignition 
27 

Radial Temp Profile @ 24"

300

400

500

600

700

800

900

1000

1100

1200

1300

12000 12100 12200 12300 12400 12500 12600 12700 12800 12900 13000

Time (sec)

Temp (C)

W 24 

X 24

Y 24

Z 24 

D 24 

Air back up 

 to 1 cfm

smoke

Figure 2.25 Radial temperature profile at the 24-in. elevation during ignition 

Temp Change @ 24"

-0.2

0

0.2

0.4

0.6

0.8

1

12000 12100 12200 12300 12400 12500 12600 12700 12800 12900 13000

Time (sec)

Temp Change (C/s)

W 24 

X 24

Y 24

Z 24 

D 24 

Air back up 

 to 1 cfm

smoke

Figure 2.26 First derivative of radial temperature profile at the 24-in. elevation during ignition 
28 

Radial Temp Profile @ 12"

0

200

400

600

800

1000

1200

1400

0 5000 10000 15000 20000

Time (sec)

Temp (C)

W 12

X 12

Y 12

Z 12

AIR INLET 

D 6

D 0

air off

power

 off

Figure 2.27 Radial temperature profile at the 12-in. elevation during ignition 

 Temp Change @ 12"

-0.5

-0.25

0

0.25

0.5

0.75

1

1.25

1.5

0 5000 10000 15000 20000

Time (sec)

Temp Change (C/s)

W 12

X 12

Y 12

Z 12

D 12

D 6

1 scfm air off

800 W

0.1 scfm

1200 W

1440 W

Figure 2.28 First derivative of radial temperature profile at the 12-in. elevation during ignition 

Time of Ignition: The time of ignition can be seen in the oxygen content measurement. Figure 

2.19 shows the oxygen content during this ignition period. As soon as the air flow was increased, 

the oxygen content decreased sharply, clearly indicating the onset of Zirconium ignition. Some 

time shortly after smoke was emitted from the test bundle, the filter on the oxygen sampling 

system plugged, and the data was compromised. It is not known what constituent in the off gas 

stream could have resulted in the lower explosive limit (%LEL) response. 
29 

Heater Short: Figure 2.20 shows the heater currents during the ignition and initial burn. The 

first complicating event occurred within 1 minute of returning the full air flow to the bundle 

when the heater shorted out and tripped the power breaker. This was unexpected, as heater 

failure was expected to occur as an open circuit when the heater element melted. The breaker 

was reset twice, but each time it tripped again within 20 seconds. 

It is postulated that the heaters shorted out when the Zircaloy cladding melted and penetrated the 

magnesium oxide insulation. This evidently greatly compromised the electrical insulating 

properties of the magnesium oxide and allowed a short from the heater element to the Zirconium 

cladding. A short to ground was provided by the TC sheaths, which were tied off to the 

electrically grounded steel test stand. The involvement of the TCs in the shorting process is also 

evident in the temperature data and will be discussed further. Electrically isolating the TC 

sheaths from ground may prevent the circuit breaker tripping problem and allow the heater to 

operate longer into the burn. 

Temperature Response: Figure 2.21 and Figure 2.22 show the axial temperature profile and the 

temperature change during the ignition period. Note that in general, this data shows that the 

temperatures were cooler at the bottom (e.g., D0 and D6) and hotter at the top (e.g., D42). The 

noted exception is D48, which requires some explanation. The heater rods were built with a 1.5-

in. no heat zone on each end. Thus, TC D48 at the top and D0 at the bottom are attached to the 

no heat zones and therefore measure lower temperatures than otherwise would be expected. The 

temperature of D0 is further lowered by its proximity of the relatively large thermal mass of the 

lower heater connection plate. 

Ignition and Burn Region: Ignition should be evident by a rapid increase in temperature leading 

to TC failure at which point the data becomes erratic. In most of the temperature data collected 

during this experiment, this expected behavior is not clearly evident. Instead, the temperature 

data drops sharply followed by TC failure. It is believed that this is due to a couple of reasons. 

The first is the shorting through the TC sheath discussed earlier. In Figure 2.21 , this effect is 

most clearly seen in the traces of TCs D18 and D24. The blips in the data correspond to the 

heater rod shorts. Note that the further down the bundle, the effect of the heater short is less 

evident. TCs D0, D6, and D12 show no effect of the short. TC D18 shows only a minor effect 

and recovers along its previous trend indicating the measurement was not compromised. TC D24 

is more seriously affected but recovers after the first two shorts. However, after the second short, 

the temperature indicated by D24 drops sharply, is then affected by the third short, and recovers 

again only to continue to drop precipitously and fail. TC D30 exhibits a similar behavior after the 

first heater short. TCs D36, D42, and D48 never recovered from the first short. Note that at the 

time of the first short, D42 was indicating a lower temperature than D36. Similar TC responses 

are found on the radial profiles at 36 in. and 24 in. as shown in Figure 2.23 and Figure 2.25, 

respectively. While the heater current spikes complicated the TC data, careful inspection does 

reveal important information. 

The cleanest TC failure due to Zirconium combustion for TC D18 is seen in Figure 2.21. As 

mentioned before, this TC was little affected by the heater power spikes. As the burn front 

reached the 18-in. level, the temperature indicated by this TC rose toward 1200C and then 

sharply dropped off and became erratic. The point of failure is more easily seen as occurring at 
30 

12,779 seconds on Figure 2.22, which shows the first derivative of temperature. The time 

between failure and obvious erratic behavior is about 1 minute. 

The next cleanest failure is seen on TC Z36 (as shown in Figure 2.23 ), one of the TCs making 

up the radial profile at the 36-in. level. Like D18, the temperature rose toward 1200C at 12,380 

seconds and then sharply fell off with about 1 minute between failure and erratic behavior. This 

occurred just after the air flow rate was turned up and just before the first heater power spike. 

Again, the temperature change plot shown in Figure 2.24 clearly demarks the time of Z36 failure. 

Figure 2.24 also shows some subtle but unusual behavior in the other TCs located at the 36-in. 

level. All the other TCs were cooling off prior to the air flow increase. There are no good 

physical explanations as to why or how the heater rod clad surface could begin to cool down 

even a small amount. This behavior is more likely due to loss of thermal contact with the 

cladding surface at the sensor tip or high temperature damage at a point below the sensor tip. 

Note that all of these TCs responded to the air flow increase as seen by the positive spike in the 

temperature change and then failed before the current spike. The temperatures indicated by these 

TCs were between 1000 and 1060C, quite a bit lower that that reached by Z36 before failure. 

This cooling behavior for TCs D48, D42, and D36 is also seen in Figure 2.22, but it is not seen at 

the 24-in. level in Figure 2.26 . 

Below Burn Region: The burn did not make it all the way down to the 12-in. level. Figure 2.27 

shows the temperature history for the 12-in. level along with the inlet air temperature and the 

TCs D0 and D6 for comparison. When the heater power was lost, the peak temperatures at the 

12-in. level were approaching 1000C, but Zirconium ignition did not occur. Rather, the rods 

began to cool down due to convective cooling of the inlet air. The same trends are evident on TC 

D0 but at lower temperatures. About 8 minutes after the heater power went off, the temperatures 

at the 12-in. level began to rise again due to the approach of the burn from above. The 

temperatures peaked about 20 minutes after the heater power was lost, resulting in the highest 

temperature detected, 1264C on Z12, but TC failure did not occur. Rather, the temperatures 

began to cool off at about the same rate as after the heater power was lost. The temperature 

change is shown in Figure 2.28. When the air was finally turned off, the temperatures at the 12-

in. level increased slightly due to the loss of convective cooling, supporting the conclusion that 

the TCs survived. The same trend is evident on TC D6. 

Bottom Plate Heat Loss: The temperature rise of TC D6 when the air flow was stopped is 

counter-intuitive when compared to the inlet air temperature. Because the air temperature was 

greater than D6 when the flow was stopped, it seems the temperature at D6 should have dropped 

faster, not risen. Looking at the temperature response of TC D0 provides the needed insight. The 

temperature at D0 was always over 100 lower than the inlet air temperature, and when the air 

flow stopped, the temperature here did drop at a faster rate. It seems the heat losses at the bottom 

plate of the bundle assembly were quite large. The temperature of the air entering the bundle was 

probably closer to the temperature indicated by TC D0 than by the inlet air temperature. 

Balance of Burn: After the approach of the burn to the 12-in. level at 13,350 seconds, it did not 

go out but returned to burn on the upper levels above 36 in., which were previously oxygen 

starved. This return is evident in the video records as a return of intense glow out the top of the 

chimney. No TC survived in the upper portion of the bundle to provide any temperature data, but 
31 

the intensity of the glow seems to indicate an even more intense burn. This is substantiated by 

TC Out 42 located on the outside of the insulation jacket at the 42-in. level (Figure 2.29). After 

the peak temperature was recorded at the 12-in. level, the temperature here rose dramatically. 

The intensity of the burn was subsiding when the air flow was terminated at 15,558 seconds. 

Data collection was terminated at 17,180 seconds. 

Outer Jacket Temp @ 42"

0

50

100

150

200

250

300

350

400

12000 13000 14000 15000 16000 17000

Time (sec)

Temp (C)

Out 42

Peak temp @ 12" air off

Figure 2.29 Temperature on the outside of insulation at the 42-in. level 

2.3.2 Smoke Sampling Results 

The most surprising product of this experiment was the analytical results of the smoke sampling. 

Magnesium was found instead of the tin that was expected. The sample taken close to the 

chimney collected more particulates than the sample taken further up in the CYBL vessel as was 

expected. The sample close to the chimney collected 580 g of magnesium (5.0 μg detection 

limit) but no tin (5.0 μg detection limit), nickel (1.0 μg detection), or iron (5.0 μg detection). 

Chromium was detected at a level of 2.1 μg, just above the detection limit of 2.0 μg, which 

leaves this result questionable. The other sample yielded 320 μg of magnesium and nothing else. 

The metals analysis provides no information on the speciation of the metals detected, but any 

magnesium sampled would most certainly be in the form of magnesium oxide. 
32 

2.4 Discussion 

2.4.1 Time of Ignition and Burn Rate 

The time of ignition can be set by the oxygen data as occurring when the air rate is increased to 

1.0 scfm at 12,300 seconds. The TCs in the upper portion of the bundle (D48, D42, D36, Y36, 

X36, and W36) exhibited a slight but unusual cooling trend 2 to 3 minutes prior to this. The first 

clear indication of ignition by TCs occurred at Z36 at 12,380 seconds. TC D30 failed about the 

same time indicating that ignition occurred between the 30-in. and 36-in. levels and 

compromised all TCs above this except Z36. The burn reached the TCs at the 24-in. level 

between 12,500 and 12,560 seconds. The apparent axial burn rate between the 30-in. and 24-in. 

level was 0.040 in/s. The burn clearly reached D18 at 12,778 seconds so the average axial burn 

rate slowed to 0.022 in/s. The burn made its closest approach to the 12-in. level at 13,350 

seconds, but its actual proximity is uncertain. Assuming the approach was close, the average 

axial burn rate was 0.010 in/s between levels 18 in. and 12 in. 

2.4.2 Post-Test MELCOR Analysis 

The Zirconium oxidation parameters used in the pre-test MELCOR modeling did not accurately 

explain the observed oxidation characteristics. The oxidation parameters used were thought to be 

the best estimate at the time. However, a few months prior to conducting Heater Design Test 1, a 

new Zircaloy 4 air oxidation kinetic study conducted at Argonne National Laboratory (ANL) 

was published [2]. Incorporation of this new kinetic information into MELCOR resulted in much 

better agreement with the experimental data. Figure 2.30 shows the peak cladding temperature 

(PCT) measured during the experiment along with the MELCOR prediction using the new 

kinetics. The new ANL kinetic study identified a breakaway phenomena in the oxidation of 

Zircaloy after which the kinetic rate of oxidation greatly increased. Thus, the study provides two 

oxidation kinetic parameters, one for pre-breakaway and one for post-breakaway, and empirical 

time-at-temperature criteria for the occurrence of breakaway. No mechanistic explanation of 

breakaway was identified. Incorporation of the pre- and post-breakaway kinetics into the 

MELCOR model resulted in excellent agreement with the data. The aborted initial attempt for 

conducting this experiment resulted in a thermal history too complex to test ANL’s breakaway 

criteria; therefore in Figure 2.30, the post-breakaway kinetics was triggered at 1350 K. The 

oxidation kinetic rate used in the pre-test modeling was midway between the pre- and postbreakaway kinetic rates. This resulted in the model over predicting the oxidation energy before 

breakaway and under predicting the oxidation energy after breakaway. 
33 

Peak Cladding Temperature for Heater Design Test #1

300

800

1300

1800

2300

2800

0 0.5 1 1.5 2 2.5 3

Time (hr)

Temperature (K)

Data PCT

w/Pre- and Post- Zr-4 Oxidation

Kinetics

Only Pre- Oxidation Kinetics

Heater Power Failure

Zr fire reaches bottom 

support plate and causes its 

Zr fire reaches bottom

Support plate and causes 

its failure 

Figure 2.30 Heater Design Test 1 PCT with and without post-breakaway oxidation kinetics 

2.5 Technical Issues 

A number of important issues regarding the experimental approach have been identified. These 

issues will be identified and the actions taken will be briefly discussed here. 

2.5.1 TC Attachment 

A contributing factor to the cooling trend shown by the TCs in the upper portion of the bundle 

may be due to loss of thermal contact between the TC tip and the heater rod surface. This could 

occur in at least two ways: 

1) The TC tips were held in contact with the Zircaloy surface by spot welding a stainless 

steel foil to the Zircaloy tube across the tip. Zirconium and iron form a Zirconium 

rich (88 wt% Zr) eutectic that melts at temperatures as low as 928C. If a portion of 

the spot weld melted at this temperature, thermal contact could be lost. 

2) The other possibility is that the formation of the Zirconium oxide layer eventually 

weakened the spot weld attachment. 

It was suspected that the TC tip attachment method was responsible for this problem. Figure 

2.31(a) shows how the TC was attached using a small hood of stainless steel foil. It is believed 

that this attachment method would be susceptible to either failure method described earlier. Both 

failure mechanisms could be prevented by wrapping the foil all the way around the rod and spot 

welding it to itself and the rod as shown in Figure 2.31(b). Thus, if the weld to the rod is lost, the 

weld to itself may hold. 
34 

Thermocouple

Metallic foil

Spot weld

Zircaloy clad

Heating element

Magnesium oxide (a) (b)

Figure 2.31 TC attachment options, (a) small hood and (b) full wrap 

2.5.2 Molten Zirconium-Magnesium Oxide (Zr-MgO) Reaction 

The predominance of magnesium in the smoke sampling filters strongly suggests there was a 

reaction between molten Zirconium and magnesium oxide. The likely reaction is 

 Zr(liq) + 2MgO => ZrO2 + 2Mg(gas) Reaction (1) 

where molten Zirconium reacts with the magnesium oxide packing and forms Zirconium oxide 

and magnesium gas. The reaction is endothermic, requiring about 350 kJ/mole, and the free 

energy of this reaction becomes negative (favored) above 1350C. The magnesium gas that is 

produced will enter the gas stream in the bundle and react exothermically (1450 kJ/mole) with 

oxygen wherever it is encountered, forming magnesium oxide. 

Cross sections of the burnt heater rods were examined to determine the extent of zirconium 

penetration. Figure 2.32 (a) and (b) show an end view of a heater rod segment taken from the 

lower and upper section of the test bundle, respectively. Many different crystalline phases are 

evident throughout the cross section. Note that there are three roughly concentric circular layers 

containing different solid materials. The outermost layer is whitish in color, the middle layer is 

dark grey and is thicker in the upper sample, and the inner central layer is lighter grey.

(a) (b) 

Figure 2.32 Photo image of burnt heater rod cross sections, (a) lower rod segment end detail, and 

(b) upper rod segment end detail 
35 

Figure 2.33 shows electron microprobe analyses of these three layers that indicate that the outer 

layer is pure Zirconium oxide as expected from the oxidation of the Zirconium metal in air. The 

middle grey is Zirconium oxide with a small fraction of magnesium. This layer is likely formed 

by the reaction of molten zirconium and magnesium oxide as shown in Reaction 1. The inner 

central region is still primarily magnesium oxide with a complex mix of minerals formed by the 

interaction of magnesium oxide with components of the molten Nichrome wire. There are small 

amounts of Zirconium associated with some of these minerals. This analysis helps to confirm 

that Reaction 1 was taking place. This analysis also shows that the Zirconium oxide formed by 

Reaction 1 is a small fraction of the Zirconium oxide formed by air oxidation. Given that the 

complicating reaction may not occur for an hour after ignition occurs, most if not all the 

objectives of the overall Spent Fuel Pool Heat-up and Propagation Experiment project can be 

achieved. It is recommended that no modifications to the basic heater design are needed to 

mitigate this issue. 

(a) (b) (c) 

Figure 2.33 Electron microprobe images of burnt heater rod section, (a) scanning electron microscopy 

(SEM) image, (b) Zirconium analysis, and (c) magnesium analysis 

2.6 Conclusion 

The main objective of the test was achieved. The Zircaloy clad was successfully ignited using the 

Watlow heater design. A number of issues were also identified. Issues of TC attachment and 

electrical grounding were identified and can likely be addressed through design improvements. 

An issue of high temperature Zirconium reaction with magnesium oxide used in the heater 

design was identified. The impact of this interaction on the suitability of the heater design was 

determined to be minimal and no design changes are required. 

37 

Plan view

Cross view

View windows 

and 

thermocouples 

Insulation

Heater 

rod

2.25” Quartz tube

6” 

12” 

18” 

24” 

30” 

36” 

42” 

Thermal 

radiation 

shields 

3 HEATER DESIGN TEST 2 

3.1 Background 

The first heater design test was successfully performed on December 17, 2004, and demonstrated 

that the Watlow heater design will result in the auto ignition of the Zircaloy cladding when 

heated under simulated spent fuel conditions in air. As documented in Section 2 a couple of 

issues with this experiment were identified including concerns about TC attachment and reaction 

of molten zirconium with the magnesium oxide used to fill the electric heater rods. 

3.2 Objective 

The objective of the second heater design test was to ignite the simulated fuel rod bundle under 

reduced oxygen conditions that might better represent the situation near the top of a full-length 

assembly. In a full-length assembly, oxidation of the Zirconium cladding in the lower, cooler 

region depletes the oxygen concentration of the air reaching the upper, hotter portion where 

ignition first occurs. The burn in this test was expected to be slower and cooler (making it easier 

to observe) and perhaps better represent the initial fire in a prototypic assembly. 

This experiment also served to test some instrumentation improvements suggested by the first 

test. In the first test, the type K TCs attached to the heater rods failed in a manner that made it 

difficult to determine with certainty where the burn front was located. One possible problem was 

detachment of the TC tip at high temperatures. To mitigate this possibility, an improved method 

for attaching the TCs to the heater rods was used (using a full wrap of shim). To provide more 

data on the location of the burn front, an array of 14 platinum TCs was added to the 

instrumentation. This array monitored the temperature axially along the outside surface of the 

quartz tube. Between each of these TCs was a 3-mm quartz light pipe, which was monitored 

visually and recorded on videotape (Figure 3.1). 

Figure 3.1 Schematic of the second heater design test 
38 

3.3 Electric Heaters 

The heater design was the same as that used in the first heater design test. The Zircaloy jacketed 

electric heaters were manufactured by Watlow Electric Manufacturing Company. The design of 

the heater was typical of those made by Watlow. Each heater was comprised of a central 

Nichrome heater element, compressed magnesium oxide powder, and an outer Zircaloy 2 jacket. 

The composition of the heater rods is shown in Table 3.1. The final magnesium oxide powder 

density was 2.720 g/cm3

, corresponding to a solids fraction of 0.759. 

Table 3.1 Composition of heater rods 

Component 

Nominal 

wt% 

Mass/rod 

(lb) 

Mass/rod 

(g) # Rods 

Total 

mass 

(g) 

Mwt 

(g/mol) 

Total 

moles 

(mol) 

Zircaloy 2 0.368 

 Zr 98.40% 164.2513 12 1971.015 91.224 21.60632 

 Sn 1.30% 2.169986 12 26.03983 118.71 0.219357 

 Fe 0.18% 0.30046 12 3.605515 55.845 0.064563 

 Cr 0.10% 0.166922 12 2.003064 51.9961 0.038523 

 Ni 0.07% 0.116845 12 1.402145 58.6934 0.023889 

 100.05% 

MgO 0.377 171.0043 12 2052.052 40.3044 50.91385 

Nichrome 0.046 20.86525 12 250.383 

The diameter of the heater rods used in the heater design test bundle was smaller than prototypic 

due to the availability of Zircaloy tubing. The tubing used was 0.440-in. diameter, 0.028-in. wall 

Zircaloy 2 tubing for BWR assembly. This tubing was made into 48-in. long, 0.375-in diameter 

heaters by Watlow in a process whereby the 0.440-in. tubing was drawn through a die that 

reduced the diameter to 0.375 in. and compressed the magnesium oxide powder considerably. 

Compression of the magnesium oxide powder is required to establish good thermal contact 

between the Nichrome heater element and the outer metallic sheath. 

3.4 Experimental 

The experiment was constructed and operated in the CYBL vessel located in Building 6585C in 

Technical Area III of SNL in Albuquerque, New Mexico. Twelve 4-ft long, 0.375-in. diameter 

heater rods with a pitch of 0.539 in. (as shown in Figure 3.1) were assembled into a bundle 

(Figure 3.2). The heater rods were instrumented with 0.032-in. ungrounded Inconel type K TCs 

in a configuration as shown in Table 3.2. Table 3.2 also lists other instrumentation used in the 

test and was collected by a computerized data acquisition system at nominally 2-second 

intervals. 
39 

Figure 3.2 Bundle of heater rods 
40 

Table 3.2 Heater Design Test 2 instrument layout 

Chan 

# Name Description 

Instrument 

type 

Max 

Volt Units 

0 D0 Heater Axial Elevation 0" K TC C 

1 D6 Heater Axial Elevation 6" K TC C 

2 D12 Heater Axial Elevation 12" K TC C 

3 D18 Heater Axial Elevation 18" K TC C 

4 D24 Heater Axial Elevation 24" K TC C 

5 D30 Heater Axial Elevation 30" K TC C 

6 D36 Heater Axial Elevation 36" K TC C 

7 D42 Heater Axial Elevation 42" K TC C 

8 D48 Heater Axial Elevation 48" K TC C 

9 W12 W Heater Elevation 12" K TC C 

10 X12 X Heater Elevation 12" K TC C 

11 Y12 Y Heater Elevation 12" K TC C 

12 Z12 Z Heater Elevation 12" K TC C 

13 W24 W Heater Elevation 24" K TC C 

14 X24 X Heater Elevation 24" K TC C 

15 Y24 Y Heater Elevation 24" K TC C 

16 Z24 Z Heater Elevation 24" K TC C 

17 W36 W Heater Elevation 36" K TC C 

18 X36 X Heater Elevation 36" K TC C 

19 Y36 Y Heater Elevation 36" K TC C Heater Nomenclature 

20 Z36 Z Heater Elevation 36" K TC C 

21 F24 F Heater Elevation 24" K TC C A B 

22 Oven Out Oven Outlet Temperature K TC C C D E F 3-mm quartz 

windows and type 

S/B TCs 23 Oven Wall Oven Wall Temperature K TC C W X Y Z 

24 Chimney Chimney Air Temperature K TC C G H 

25 Base Plate Base Plate Temperature K TC C 

26 Bottom Atm Gas Inlet Temperature K TC C 

27 Conn 21 In 21 Connector Temperature K TC C 

28 Out 12 Outer Insulation Temp 12" K TC C 

29 Out 36 Outer Insulation Temp 36" K TC C 

30 In 6 Inner Skin 6" S TC C 

31 In 9 Inner Skin 9" B TC C 

32 In 12 Inner Skin 12" S TC C 

33 In 15 Inner Skin 15" B TC C 

34 In 18 Inner Skin 18" S TC C 

35 In 21 Inner Skin 21" B TC C 

36 In 24 Inner Skin 24" S TC C 

37 In 27 Inner Skin 27" B TC C 

38 In 30 Inner Skin 30" S TC C 

39 In 33 Inner Skin 33" B TC C 

40 In 36 Inner Skin 36" S TC C 

41 In 39 Inner Skin 39" B TC C 

42 In 42 Inner Skin 42" S TC C 

43 In 45 Inner Skin 45" B TC C 

In6

In12

In9

In15

In18

0" 

6" 

12" 

18" 

24" 

30" 

36" 

42"

In21

In24

In27

In30

In33

In36

In42

In39

Win 43.5 

Win 40.5 

Win 37.5 

Win 34.5 

Win 31.5 

Win 28.5 

Win 25.5 

Win 22.5 

WIn 19.5 

Win 16.5 

Win 13.5 

Win 10.5 

Win 7.5 

In45
41 

Table 3.3 Heater Design Test 2 instrument layout (continued) 

Chan 

# Name Description 

Instrument 

type 

Max 

Volt Units 

44 A current Current of Heater A Current xducer 5 ampere 

45 B current Current of Heater B Current xducer 5 ampere 

46 C current Current of Heater C Current xducer 5 ampere 

47 D current Current of Heater D Current xducer 5 ampere 

48 E current Current of Heater E Current xducer 5 ampere 

49 F current Current of Heater F Current xducer 5 ampere 

50 G current Current of Heater G Current xducer 5 ampere 

51 H current Current of Heater H Current xducer 5 ampere 

52 W current Current of Heater W Current xducer 5 ampere 

53 X current Current of Heater X Current xducer 5 ampere 

54 Y current Current of Heater Y Current xducer 5 ampere 

55 Z current Current of Heater Z Current xducer 5 ampere 

56 Mass Flow Mass Air Flow into Oven MKS Flow 10 slpm 

57 Heater Volts Voltage Applied to Heaters Voltage xducer 10 volts 

Inner skin types B and S TCs are exposed junction with bead in contact with the quartz tube. 

Outer insulation TC is attached to the outer metal wrap opposite (approximate) the type S and B TCs. 

The connector TC is located on the mid point of the ceramic connectors at the end of the type B TC.

The TCs attachment method used to attach the TC to the Zircaloy tubing was slightly different 

than that used in the first heater test. In the first test, the TCs were attached by spot welding a 

small piece of stainless steel shim or foil across the tip. In the second heater test, the TC was 

attached using a full wrap of the stainless foil (Figure 3.3). The full wrap provides a more secure 

attachment that is less prone to failure due to the oxidation of the Zircaloy or the melting of the 

nickel-Zirconium eutectic. 
42 

Figure 3.3 TC attachment detail 

Note: The tips of the TCs are attached with a full wrap of shim stock. 

Another experimental improvement was the addition of a heater on the bottom brass base plate. 

In the first heater design test, it appeared that the thermal mass of this plate cooled the incoming 

gas flow such that the bottom heater rod temperature never approached the inlet air temperature. 

With this additional heater, the temperature of the base plate could be controlled to match the 

inlet air temperature. 

A 57-mm ID (2.244-in.) quartz tube was placed around the heater bundle (see Figure 3.4), and 

the tube was surrounded with reflective foil. The bundle was insulated with 5 in. of KaowoolTM. 

Two additional steel radiation shields were used. One was located on the outside of the 

insulation, and the other was located midway through the insulation. An axial array of 14 

platinum TCs (types S and B) was installed at 3-in. intervals in contact with the outside surface 

of the quartz tube. In between these TCs was a similarly installed quartz light pipe to serve as a 

visual indicator of the burn front location. The TCs and light pipes were held in place by a 

ceramic positioning block as pictured in Figure 3.5, Figure 3.6 and Figure 3.7. The final test 

setup is pictured in Figure 3.8. 
43 

Figure 3.4 Quartz tube installed over heater bundle 

Note: The mouse hole in the bottom of the quartz tube is used to route the TCs away from the bundle. 
44 

Figure 3.5 Stainless steel foil in place over quartz tube and ceramic positioning block is in place 
45 

Figure 3.6 Installation of quartz wall TCs and light pipes into ceramic positioning block 
46 

Figure 3.7 Detail of quartz tube wall TCs and light pipe installation 
47 

Figure 3.8 Initial test setup for Heater Design Test 2 

3.4.1 Additional Sampling 

In addition to the data collected by the instrumentation listed in Table 3.2, three other sampling 

efforts collected data for this experiment. The first effort was to monitor the oxygen 

concentration of the exit gas, the second effort was the sampling of smoke particulates for metals 

analysis, and the third was video recording of the experiment from a number of perspectives. 

Oxygen: The oxygen concentration of the gas leaving the test bundle was monitored using an 

ITX (S/N: 0307010-286) portable atmosphere-monitoring instrument. The instrument measured 

both oxygen concentration and the %LEL. The sample was collected from the mid point of the 
48 

chimney through a stainless steel tube that was routed through an ice bath and outside the CYBL 

vessel where the oxygen monitor was located. The sample was drawn using the instrument’s 

internal sample pump. The %O2 and %LEL data was collected internally along with the time at 

5-second intervals. 

Smoke: Health physics personnel installed three particulate sampling devices inside the CYBL 

vessel to analyze the total metal content (this provides no information on the chemical nature of 

the metal) of the smoke generated by the experiment. The sample filters were analyzed for the 

metals added to Zirconium to make the Zircaloy 2 alloy, which include tin, iron, nickel, and 

chromium. The samples were also analyzed for magnesium, a principal constituent of the 

magnesium oxide insulation used inside the heaters. Table 3.1 presents the initial mass of each of 

these constituents that was initially present in the test bundle. 

Video: Video perspectives included a close-up looking down the chimney, a side view of the 

experimental setup, and a distant top view from the top opening of the CYBL vessel. 

3.4.2 Test Operation 

The experiment was conducted on May 24, 2005. The chronology of significant events during 

the experiment is listed in Table 3.3. The air preheater (and data collection system) was started at 

11:58 a.m., and the air flow of 30.0 slpm was started 53.9 minutes into the test. The preheat air 

temperature set point was 600 K (327C). The base plate heater was powered at 61.8 minutes and 

collection of oxygen data began at 132.5 minutes. 
49 

Table 3.4 Chronology of events during experiment 

Clock 

Time 

Experiment 

Time (min) Event 

11:58:26 0.00 Start: data on, air heat on, 327 C set 

point 

12:52:18 53.90 Air flow on 30.0 slpm 

13:00:10 61.77 Base plate heat on 327 C set point 

14:10:56 132.54 O2 data on 

14:11:02 132.64 Heater power on 

14:18:20 139.94 Lower smoke samplers into CYBL 

14:28:34 150.17 Switch flow from air to 7.5% O2

14:50:28 172.07 Rod E shorts 

14:50:56 172.54 Rod E burns out 

14:51:10 172.77 Reset O2 valve, no sample until now 

15:12:00 193.60 Brief fire, rod H shorts 

15:12:22 193.97 Rod H burns out 

16:22:02 263.64 D48 fails 

16:23:18 264.90 D42 fails 

16:24:36 266.20 W, X, and Y36 fail 

16:25:08 266.74 O2 going to zero: ignition 

16:26:42 268.30 D36 fails 

16:26:44 268.34 Z36 fails 

16:28:32 270.14 All rod power off 

17:15:50 317.44 Gas flow off 

17:35:02 336.64 Data off 

Figure 3.9 shows the total power history and individual heater rod current histories, while Figure 

3.10 shows the applied voltage during the test. The power to the heater rods was applied at 132.6 

minutes. The total initial power was 975 W, which corresponds to about 20.3 W/ft. At 150.2 

minutes, the input gas was changed from air to 7.5 vol% O2 in nitrogen. At 172.1 minutes, heater 

rod E shorts and fails. The constant power control increased the current to the remaining heater 

rods, and the total power increased slightly to about 1000 W. This is 22.7 W/ft on the remaining 

heater rods. At 193.6 minutes, heater rod H shorts and fails. Again, the constant power control 

increased the current to the remaining heater rods. The total power increased slightly again to 

1050 W or 26.3 W/ft on the remaining heater rods. The total power was maintained at 1050 W 

until ignition was achieved at 266.7 minutes, shortly after which all heater rods failed. 
50 

Power and current history

600

800

1000

1200

1400

1600

1800

2000

0 50 100 150 200 250 300 350

time (min)

total power (W) or current (mA)

heater power

A CURRENT

B CURRENT

C CURRENT

D CURRENT

E CURRENT

F CURRENT 

G CURRENT

H CURRENT

W CURRENT

X CURRENT

Y CURRENT

Z CURRENT 

rod E shorts & fails rod H shorts & fails all power off

Figure 3.9 Heater currents during time of Zircaloy ignition with major experimental events noted 

Heater Volts

0

20

40

60

80

100

120

0 50 100 150 200 250 300 350

time (min)

Voltage (V)

Heater Volts

Figure 3.10 Voltage applied to heater rods 

3.4.3 Ignition Details 

The point of Zirconium ignition is most easily seen in the oxygen concentration data shown in 

Figure 3.11. At the point of ignition, the oxygen concentration drops sharply to zero. This 

occurred at 266.7 minutes into the test. Figure 3.12 and Figure 3.13 show the axial heater rod 

temperature profile and the first derivative of these temperatures, respectively. These plots show 

that D36 and D42 approached ignition in similar fashion, but D42 failed before D36. This 
51 

indicates that the fire ignited above D36 and below D42. Once ignition occurred, the TCs leading 

up to D42 (and D48) failed. Twelve minutes after D36 failed, the temperature at D30 peaked at 

1225C, the highest temperature measured by a type K TC. However, this TC did not fail; 

therefore, the burn front did not quite reach this level. Figure 3.14 shows the radial temperature 

profiles at the 12-in., 24-in., and 36-in. levels. Generally, the radial profiles are even at each 

level, with heater rod Z being a bit lower in temperature than the other heater rods. Note that 

heater rod Z was adjacent to the light pipe array, which could account for the slightly lower 

temperature. Ignition and TC failure is evident with all the TCs at the 36-in. level and not at the 

lower levels. Note that as soon as all the heater rods failed, the temperatures on the heater rods at 

the 24-in. level and below dropped sharply. Also, note that the TCs that did fail did not exhibit 

any of the peculiar behavior evident in the first heater design test indicating that the full wrap 

installation method worked as expected. 

Exit gas oxygen, LEL and flow rate

-5

0

5

10

15

20

25

30

35

0 50 100 150 200 250 300 350

time (min)

concentration (%) and flow (lpm)

O2(%)

LEL(%)

Mass FLOW

switch from air to 7.5% O2

O2 sample valve corrected

air on

flow off

ignition

Figure 3.11 Oxygen concentration of off gas with major experimental events noted 
52 

Axial rod temperature profile

0

200

400

600

800

1000

1200

1400

0 50 100 150 200 250 300 350

time (min)

Temperature (C)

D 48

D 42

D 36

D30

D 24 

D 18

D 12

D 6

D 0

air flow on

heater power on

Ignition

All rod power lost

Figure 3.12 Axial temperature profile with major experimental events noted 

Axial rod dT/dt

-20

-15

-10

-5

0

5

10

15

20

25

30

35

40

45

50

0 50 100 150 200 250 300 350

time (min)

dT/dt (C/min)

D 48

D 42

D 36

D30

D 24 

D 18

D 12

D 6

D 0

air flow on heater power on

rod E lost

rod H lost

Ignition All rod power lost

Figure 3.13 First derivative of axial temperature profile with major experimental events noted 
53 

Radial rod temperature profile 

0

200

400

600

800

1000

1200

0 50 100 150 200 250 300 350

time (min)

Temperature (C)

Z 36 

Y 36

X 36

W 36

D 36

Z 24 

Y 24

X 24

W 24 

D 24 

Z 12

Y 12

X 12

W 12

D 12

Figure 3.14 Radial temperature profile at the 12-in., 24-in., and 36-in. elevations 

Figure 3.15 shows the axial temperature profile along the outside surface of the quartz tube. 

Figure 3.16 shows the light pipes glowing during the burn. The platinum (types S and B) TCs 

operate to higher temperatures (1500 to 1600C) than the type K TCs used on the heater rods. 

Because these TCs are not located inside the quartz tube where the fire is located, they are much 

less prone to failure due to the fire. While this also means the temperature indicated does not 

accurately represent the temperature anywhere in the heater rod bundle and was generally lower 

than the temperature inside the heater rod bundle, the wall temperature measurements did 

provide an excellent indication of the location of the burn front. When all the heaters failed, all 

the temperatures at the 24-in. level and below dropped off quickly indicating that there was little 

oxidation energy being released at these levels. There was a 15-minute delay in the temperature 

drop at the 27-in. level, suggesting a small amount of oxidation energy release. The temperatures 

at the 30-in. level and above all increased significantly after the loss of the heater power, 

indicating significant energy input from Zirconium oxidation. 
54 

Axial temperature profile (quartz tube)

0

200

400

600

800

1000

1200

1400

1600

0 50 100 150 200 250 300 350

time (min)

Temperature (C)

In 45

In 42

In 39

In 36

In 33

In 30

In 27

In 24

In21

In18

In 15

In12

In 9

In 6

All rod power lost

heater power on

Ignition

Figure 3.15 Axial temperature profile along quartz tube 

Figure 3.16 Close-up of light pipes during burn 

From the data presented in Figure 3.15, it can be seen that TCs In 36 and In 39 heated up first 

and together indicating the burn initiated in this region. A few minutes later, TCs In 33 and In 42 

both sharply increased in temperature, indicating that the burn front moved both upward and 

downward 3 in. A few minutes later, the temperature at the 30-in. and 45-in. levels spiked 

suggesting further upward and downward spread of the oxidation front. The peak temperatures at 

the 30-in., 33-in., and 36-in. levels were all reached at about the same time; however, the peak 

temperature was not as high at the progressively lower levels. After the temperature at these 
55 

levels peaked, the temperature peaked at progressively higher levels, the next being at the 39-in. 

level, the 42-in. level, and finally the 45-in level. The highest peak temperature measured was 

1364C at the 42-in. level. 

As stated previously, the quartz wall temperature measurements were generally lower than the 

corresponding fuel rod temperatures. It is possible to correlate quartz wall temperature 

measurement with the measured fuel rod temperature measurements before the type K TC failure 

in order to estimate the offset of the wall temperature from the rod temperature. Figure 3.17 

shows the axial heater rod temperatures at the 36-in. and 42-in. levels. The higher temperature 

data starting at about 160 minutes is obtained from the quartz wall temperature measurements 

after adjusting the wall temperature to match the rod temperature shortly before the type K TCs 

failed. The peak rod temperature estimated in the rod bundle was about 1475C. This 

temperature is far short of the 1850C required to melt the Zircaloy cladding. The absence of any 

appreciable smoke during this experiment also suggests that the cladding did not melt, and the 

molten Zr-MgO reaction evident in the first heater design test did not occur. 

Composite Axial rod temperature profile

0

200

400

600

800

1000

1200

1400

1600

0 50 100 150 200 250 300 350

time (min)

Temperature (C)

D 42

D 36

In 42 corrected

In 36 corrected

Figure 3.17 Composite axial rod temperature profile using corrected wall temperature measurements 

3.4.4 Post-Test MELCOR Analysis 

Incorporation of the new ANL Zircaloy 4 air oxidation kinetic information [2] into MELCOR 

resulted in excellent prediction of the time of ignition. Figure 3.18 shows the estimated peak 

cladding temperature (see Figure 3.17) measured during the experiment along with the 

MELCOR prediction using the new kinetics. The new ANL kinetic study identified a breakaway 

phenomena in the oxidation of Zircaloy after which the kinetic rate of oxidation greatly 

increased. Thus, the study provides two oxidation kinetic parameters, one for pre-breakaway and 

one for post-breakaway and an empirical time-at-temperature criterion for the occurrence of 

breakaway. The time-at-temperature criterion was used to transition the MELCOR calculation 

from pre- to post-breakaway kinetics, which occurred at 250 minutes in excellent agreement with 
56 

the data. Also shown is the MELCOR prediction using only the pre-breakaway kinetics. Without 

the added oxidation energy, the rod temperature was predicted to drop sharply at 268 minutes 

when all the heater rods failed. 

Comparison of MELCOR with Data

0

200

400

600

800

1000

1200

1400

1600

0 50 100 150 200 250 300 350

time (min)

Peak Clad Temperature (C)

MELCOR 

MELCOR pre-breakaway

only

Data

Heater rod 

failure at 268 min

Breakaway

at 250 min

Heater rod power on

at 133 min

Figure 3.18 Peak cladding temperature data and MELCOR calculation with and without postbreakaway oxidation kinetics 

3.4.5 Post-Test Inspection 

Figure 3.19 shows the outer insulation jacket after the conclusion of the experiment. The labels 

show obvious signs of high temperature at and above the 36-in. level. Figure 3.20 shows the 

temperature of the outer jacket at the 12-in. and 36-in. levels. The peak jacket temperature 

measured at the 36-in. level was 200C. The condition of the labels at the 42-in. level suggests 

the surface temperature here reached a higher temperature. Figure 3.21 shows the inner quartz 

tube and inner thermal radiation reflector with the insulation removed. The upper portion of the 

bundle shows obvious signs of the high temperature burn. Figure 3.22 shows a detail of the 

heater rod rubble before the bundle was removed from the test stand. Figure 3.23 shows a detail 

of heater rod rubble at the 36-in. to 39-in. levels after bundle removal from the test stand. Note 

that much of the rubble is still held within the test bundle by intact Nichrome heater elements. 

Nichrome melts at about 1400C; therefore, the core of many of the heater rods did not reach this 

temperature. This is consistent with the estimate of a maximum clad temperature of 1475C. 

Because the heater rods were no longer heated once the cladding ignited, the cores of the heater 

rods were cooler than the burning surface. 
57 

Figure 3.19 Outer insulation jacket after burn 
58 

Gas & other miscellaneous temperatures

0

100

200

300

400

500

600

700

800

0 50 100 150 200 250 300 350

time (min)

Temperature (C)

OVEN Outlet

OVEN WALL

CHIMNEY

Base Plate

Bottom Atmosphere

B TC connecter

Outer Jacket 36

Outer Jacket 12

Figure 3.20 Inlet gas temperature, exit gas temperature in chimney, and other miscellaneous 

temperatures 
59 

Figure 3.21 Post-test quartz tube and inner thermal radiation shield after removal of insulation 
60 

Figure 3.22 Detail of heater rod rubble 
61 

Figure 3.23 Detail of heater rod rubble at the 36-in. level after bundle removal from test stand 

3.4.6 Smoke Sampling Results 

Total metals analysis of the three smoke sampling filters yielded no trace of any metals including 

magnesium, tin, iron, nickel, or chromium. This is consistent with the experimental observation 

that no smoke was generated during the experiment. 

3.5 Issues 

The main experimental issue identified by this experiment was the nature of the downward burn 

front advance. With the reduced oxygen expected at the burn ignition point, the burn front does 

not advance downward substantially once the heat input from the electric heaters is lost. With 

prototypic decay heated fuel rods, the heat input from the fuel rods would, of course, not be lost 

once ignition occurs and significant downward burn front migration would be expected. This is 

an experimental artifact that will be taken into account in the numerical model calibration phase 

of this project. 

3.6 Conclusion 

The main objective of the test was achieved. The Zircaloy clad was successfully ignited using the 

Watlow heater design under reduced oxygen conditions. As expected, the resulting burn was 
62 

slower and cooler than during the previous heater design test. The peak temperature of the clad 

was estimated to be 1475C. This is consistent with the observations that much of the Nichrome 

wire was still intact (melting point of 1400C), and no smoke was produced by the reaction of 

molten Zirconium (melting point of 1850C) and the magnesium oxide core. 

 
 

 63 

4 HYDRAULIC ANALYSIS OF THE SPENT FUEL POOL 

EXPERIMENT 

This report summarizes the findings of the SFP pressure drop experiments conducted between 

August 25 and October 12, 2005. The stated purpose of these investigations was to determine 

hydraulic coefficients, namely SLAM and k values, for the calibration of the MELCOR severe 

accident analysis code. The apparatus was operated in the laminar regime with Reynolds 

numbers spanning from 70 to 900, based on the bundle velocity and hydraulic diameter. 

4.1 Experimental Apparatus and Procedures 

Commercial components were purchased to create a prototypical BWR assembly including the 

top and bottom tie plates, spacers, water rods, channel box, and all related assembly hardware. 

Stainless steel conduit was substituted for the fuel rod pins for hydraulic testing. The diameter of 

the stainless steel rods was slightly smaller than prototypic pins, 1.1110-2 m versus 1.1210-2 m. 

The slightly simplified stainless steel mock fuel pins were fabricated based on drawings and 

physical examples supplied typical fuel type . The dimensions of the assembly components are 

listed in Table 4.1. 

Table 4.1 Dimensions of assembly components in the 99 BWR 

Description Lower (Full) Section Upper (Partial) Section 

Number of Pins 74 66 

Pin Diameter (m) 1.1110-2 1.1110-2

Pin Pitch (m) 1.4410-2 1.4410-2

Pin Separation (m) 3.2810-3 3.2810-3

Water Rod OD (main section) (m) 2.4910-2 2.4910-2

Water Rod ID (m) 2.3410-2 2.3410-2

Figure 4.1 shows the layout of the SFP pressure drop experimental assembly, including all 

available pressure port locations. Two Paroscientific Digiquartz differential pressure transducers 

(Model 1000-3D) were plumbed directly to the desired pressure ports. These pressure gauges use 

a highly sensitive quartz crystal to measure slight changes in differential pressure (resolution 

~0.02 Pa). For this report the pressure drop across two ports is reported as the low side port first 

followed by the high side port, e.g., the overall pressure drop across the entire assembly is 

reported as 1–B. 
 

 64 

Cross section 

above partial 

Port “B” 

162.875" 

Can 

bottom 

0.000" 

Can top 

reference 

101.078" Port 11 

159.823" Port 17

139.633" Port 15 

119.458" Port 13 

99.333" Port 10 

110.218" Port 12 (radial 22, 23) 

121.198" Port 14 

141.308" Port 16 

60.746" Port 9 

58.996" Port 8 

40.541" Port 7 

38.871" Port 6 

29.690" Port 5 (radial 18, 19, 20, 21) 

20.430" Port 4 

18.680" Port 3 

2.500" Port 2 

0.937" Port 1 

Digiquartz 

differential 

pressure 

sensor

Bypass holes 

– 2 places 

Figure 4.1 Experimental apparatus showing as-built port locations 

Measurements were recorded directly to the hard drive of a PC-based data acquisition system 

every 3 seconds using a LabView 7.1 interface. These measurements included the air flow rate 

through the assembly, ambient air temperature, ambient air pressure, and the assembly pressure 

drops. The procedure for collecting the data is detailed next. 
 

 65 

With each pressure transducer plumbed to two set port locations and with the air flow off, 

pressure drop measurements were recorded for a period of roughly 1 minute. These 

measurements were termed zero flow measurements and allowed for correction of any zero drift 

in the transducer. Next, the air flow was set to the desired rate with pressure drop readings 

subsequently acquired for 2 minutes. The air flow was then stopped, and zero flow 

measurements were again taken for 1 minute. This procedure was repeated for different air flow 

rates. The resulting pressure traces are shown in Figure 4.2. The pressure spikes evident during 

the reestablishment of flow are discarded before averaging for the pressure drops. Also, the slight 

zero drift of the transducer was corrected by subtracting the average of the zero flow 

measurements taken prior to and after each respective flow test. The zero corrections of the 

pressure drops were less than 0.15 N/m2

, which ocurred during an overall 1-B pressure drop 

measurement. 

-10

0

10

20

30

40

50

60

70

0 500 1000 1500 2000 2500 3000 3500 4000 4500

Time (sec)

Pressure (N/m 2

)

2-8 8-17

Figure 4.2 Pressure traces recorded during the blocked water rod testing for measurements 

across 2–8 and 8–17 

Experimental runs may be divided into two distinct groups, blocked and unblocked water rods. 

The bypass holes were blocked for all testing. To understand the influence of flow through the 

water rods, acrylic sleeves were inserted over the water rod exit holes for one series of 

experiments to force all of the air flow through the bundle (Figure 4.3). These experimental runs 

were then repeated using the apparatus with unblocked water rods. 
 

 66 

Figure 4.3 Acrylic sleeves are inserted over the top of the water rods to block the flow. Sleeve 

alignment on the water rods (a) and the sleeve cross-section (b) are shown above 

4.2 Hydraulic Analysis 

As discussed previously, the goal of this research was to determine the SLAM and k coefficients 

for use with the MELCOR code. The cross-sectional areas and hydraulic diameters used in the 

hydraulic analyses to follow are presented in Table 4.2. 

Table 4.2 Summary of the flow areas and hydraulic diameters of the SFP experiment. 

Description A (m2

) DH (m) 

Upper Bundle 1.0610-2 1.4110-2

Upper Bundle plus Water Rods 1.1410-2 1.4510-2

Upper Spacer 8.3510-3 3.5710-3

Lower Bundle 9.7910-3 1.1910-2

Lower Bundle plus Water Rods 1.0610-2 1.2410-2

Lower Spacer 7.3710-3 2.8910-3

Water Rod (1) 4.2910-4 2.4910-2

Curve fits to the pressure drop data were used to determine the SLAM and k coefficients of the 

assembly. The determination of these coefficients is discussed next. The major, or viscous, 

pressure loss is expressed in Equation 1. 















 















   2

ρ V

D

L P

2

bundle

H

major f 1 

Retrieval wire Adhesive tape 

(b)

Key 

Spacer 

Acrylic sleeve 

(a) 
 

 67 

The friction factor for laminar flow is written explicitly as 

 100 (bundle flow)

, where S 64 (pipe flow) S

LAM

LAM



  Re f 2 

Substituting for the Reynolds number yields 





 



 















   2

V μ

D

L P S bundle

2

H

major LAM 3 

The minor, or form, pressure drops across the assembly are given by 

 













    2

ρ V P k

2

bundle

minor 4 

Curve fits to pressure drop data are presented in the following format. In Equation 5, the 

quadratic term accounts for the minor losses and the linear term for the major losses. 

1 bundle

2 Ptotal  a2  Vbundle  a  V 5 

Because the total pressure drop is simply the sum of the major and minor pressure drops, the 

SLAM and k coefficients may now be determined explicitly. 

ρ

2 k

L μ

D S 2

2

2

H

LAM 1

a

a

 

















  



 6 

4.2.1 Blocked Water Rod Results 

The following is an example SLAM-k analysis of the curve fit to pressure drop data. Please refer 

to Figure 4.1 for the location of the pressure ports described next. The data in Figure 4.4 refer to 

the pressure drops across pressure ports 2–8, 8–17, and 2–17 for the assembly with blocked 

water rods (compiled from data file, “blocked3 full 8-17 partial 2-8.csv,” created October 4, 

2005). 

Note: The pressure drops are plotted versus the corresponding bundle velocity for each section. 

The hydraulic diameter and flow area of the lower portion of the assembly were used to calculate 

SLAM and k coefficients for the 2–17 data. Also, the pressure drop across 2–17 was not 

independently measured but is the summation of the pressure drops across 2–8 and 8–17. See 

Appendix B, Error Analysis, for details concerning the 95% confidence error bars on the data. 
 

 68 

y = 3.60E+00x2 + 5.87E+00x

R2

 = 9.99E-01

y = 9.33E+00x2

 + 2.29E+01x

R2

 = 9.99E-01

y = 1.24E+01x2

 + 2.83E+01x

R2

 = 9.99E-01

0

10

20

30

40

50

60

70

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

Vbundle (m/s)

 P (N/m 2

)

2-8 8-17 2-17

Figure 4.4 Pressure drop as a function of bundle velocity for the SFP assembly with blocked water rods 

The relevant information for the calculation of the SLAM and k coefficients is given in Table 4.3. 

Uncertainties in the SLAM and k coefficients listed in this report are taken as ±5 and ±1, 

respectively. See Appendix B, Error Analysis, for details concerning this uncertainty estimate. 

Table 4.3 SLAM and k coefficient analysis data for pressure drops between 

2–8, 8–17, and 2–17 for blocked water rods 

Pressure 

Drop L(m) A (m2

) DH (m) 

a1

(N·s/m3

) 

a2

(N·s2

/m4) SLAM  k 

# of 

spacers 

k per 

spacer 

2–8 1.43 0.0106 0.0141 5.87 3.60 88 7.3 2 3.7 

8–17 2.56 0.0098 0.0119 22.9 9.33 138 19 5 3.8 

2–17 4.00 0.0098 0.0119 28.3 12.4 109 25 7 3.6 

This analysis assumes air properties at local ambient conditions, typically  = 0.98 kg/m3

 and 

μ = 1.8510-5 N·s/m2

. Changes in air temperature and pressure are taken into account for 

measurements collected during different experimental runs. 

The influence of the experimental flow rate range is examined in Table 4.4. For this analysis, 

subsets of the entire data set were used to determine the curve fit coefficients. These results are 

also presented graphically in Figure 4.5. The hydraulic loss coefficients tend towards constant 

values as the experimental flow range is increased. The errors associated with these 

measurements are also more significant in the lower flow rate ranges. Therefore, the SLAM and k 

values determined from the largest data range are considered to be more applicable. 
 

 69 

Table 4.4 SLAM and k values for different ranges of experimental air flow rate 

Flow Rate 

Range (slpm) 

2–8 8–17 2–17 

SLAM  k SLAM  k SLAM  k 

50–200 46 20 99 47 75 64 

50–300 60 14 113 34 86 46 

50–400 75 10 126 25 99 33 

50–500 84 8.1 134 20 106 27 

50–600 88 7.3 138 19 109 25 

0

30

60

90

120

150

0 100 200 300 400 500 600 700

Upper Experimental Flow Rate Range (slpm)

SLAM, (k)

2-8 SLAM

2-8 Sum(k)

8-17 SLAM

8-17 Sum(k)

2-17 SLAM

2-17 Sum(k)

Figure 4.5 Hydraulic loss coefficients as a function of the upper, measured experimental flow range 

A summary of the SLAM and k coefficients for blocked water rod testing is shown in Table 4.5. 

These values were determined from the full experimental flow rate range of 50 to 600 slpm, or 

Reynolds numbers of 70 to 900 in the fully populated segment, respectively. The hydraulic 

diameter and flow area of the fully populated section was used to calculate the SLAM and k values 

for any span including the lower section, e.g., 1–B. Values of SLAM and k for segments 1–2, 4–6, 

and 6–7 are calculated using the characteristic hydraulics of both the partially and the fully 

populated sections. 
 

 70 

Table 4.5 SLAM and k coefficients for the blocked water rod assembly 

Description Segment A (m2

) L (m) DH (m) SLAM  k 

# of 

spacers 

k per 

spacer 

Long Spans 1–B 0.0098 4.19 0.0119 106 37 – – 

2–8 0.0106 1.44 0.0141 88 7.3 2 3.6 

8–17 0.0098 2.56 0.0119 138 19 5 3.8 

2–17 0.0098 4.00 0.0119 109 25 7 3.6 

4–8 0.0106 0.98 0.0141 85 4.0 1 4.0 

8–15 0.0098 2.05 0.0119 137 15 4 3.8 

4–15 0.0098 3.03 0.0119 111 18 5 3.6 

Individual 

Bundle Runs 

4–5 0.0106 0.24 0.0141 69 0.88 – – 

5–6 0.0106 0.23 0.0141 51 0.00 – – 

4–6 0.0106 0.47 0.0141 60 0.88 – – 

4–6 0.0098 0.47 0.0119 40 0.71 – – 

11–12 0.0098 0.23 0.0119 89 0.70 – – 

12–13 0.0098 0.23 0.0119 65 0.13 – – 

11–13 0.0098 0.47 0.0119 77 0.84 – – 

Spacers/Tie 

Plates 

1–2 0.0106 0.040 0.0141 49 0.42 – – 

1–2 0.0098 0.040 0.0119 33 0.36 – – 

6–7 0.0106 0.045 0.0141 408 3.1 1 3.1 

6–7 0.0098 0.045 0.0119 272 2.6 1 2.6 

13–14 0.0098 0.044 0.0119 729 3.2 1 3.2 

17–B 0.0098 0.13 0.0119 119 11 – – 

Next, the SLAM and k coefficients across 6–7 and 13–14 are recalculated using the appropriate 

spacer hydraulic diameter and flow area. As shown in Table 4.6, the values determined from this 

analysis are significantly reduced. 

Table 4.6 Spacer SLAM and k coefficients calculated with the appropriate spacer hydraulic diameter and 

flow area 

Segment A (m2

) L (m) DH (m) SLAM  k 

6–7 0.00835 0.045 0.00357 21 1.9 

13–14 0.00737 0.044 0.00289 32 1.8 

The values of SLAM through the spacers 6–7 and 13–14 are rather large compared to the other 

values in Table 4.5. These large values are due to the use of the bundle hydraulic diameter when 

calculating the SLAM coefficients. As a result, the SLAM values for segments including spacers, 

such as 1-B, have inflated values of SLAM when compared to a single section without spacers 

(11–13). 

The single section/spacer SLAM and k values can be manipulated to recreate the values observed 

in the multiple section/spacer data. To accomplish this, the values of SLAM must be weight 

averaged based on flow length. Equation 7 shows the general format for calculating the effective 

SLAM coefficient of an assembly span with an overall flow length of Ltot, “I” number of spacers, 

and “J” number of sections. 
 

 71 

tot

J

j 1

sect,j LAMsect,j

I

i 1

sp,i LAMsp,i

LAMeff L

L S L S

S















   



   

 7 

The loss coefficients will simply be the sum of the k values for the individual segments. The 

values for 4–6 and 6–7 are considered to be typical for a segment and spacer in the upper section. 

Similarly, 11–13 and 13–14 were used for the lower section. The top and bottom tie plate SLAM

and k coefficients are measured separately and are used to calculate the effective values for the 

1–B span. Two values are listed for segments 1–2, 4–6, and 6–7 in Table 4.5. These values 

denote the loss coefficients determined from the upper and lower assembly hydraulic areas and 

diameters. The first values listed in the table are used for spans including the upper portion of the 

assembly only. For spans including the lower assembly (e.g., 2–17), the values calculated with 

the lower assembly area and hydraulic diameter are used for segments in the partially populated 

region (for 6–7 SLAM = 272, k = 2.6). Table 4.7 gives the calculated effective SLAM and k 

coefficients based on the values from the individual sections/spacers. The SLAM values are within 

about 10% of each other. However, the effective k values are somewhat higher than the 

measured values. Discrepancies in these two sets are most likely due to the assumption that all 

segments and spacers have the same loss coefficient values throughout the assembly. 

Table 4.7 Calculation of effective SLAM and k values for different span lengths 

Pressure 

Drop 

Measured Effective 

SLAM  k SLAM  k 

1–B 106 37 104 39 

2–8 88 7.3 81 8.8 

8–17 138 19 133 20 

2–17 109 25 105 28 

4–8 85 4.0 75 4.9 

8–15 137 15 133 16 

4–15 111 18 106 20 

Finally, Figure 4.6 shows the cumulative pressure drop in the assembly as a function of axial 

location for various flow rates. The lines in the plot represent the pressure drop determined from 

Equations 3 and 4. The form loss and laminar friction coefficients were taken from Table 4.5. 

The dashed line at z = 2.7 m indicates the demarcation of the partially and fully populated 

sections. 
 

 72 

0

20

40

60

80

01234

z (m)

 P (N/m 2

)

Flow rate decreasing 

100 

200 

300 

400 

500 

600 

z 

A = 0.0098 m2

DH = 0.0119 m 

A = 0.0106 m2

DH = 0.0141 m 

Figure 4.6 Cumulative pressure drop as a function of axial location for flow rates of 100, 200, 300, 400, 

500, and 600 slpm 

The lines represent the analytically determined pressure drops using Equations 3 and 4 along 

with the values of SLAM and k listed in Table 4.5. 

4.2.2 Unblocked Water Rod Results 

The pressure drops across pressure ports 2–8, 8–17, and 2–17 are shown in Figure 4.7 in the SFP 

assembly with unblocked water rods. The pressure drop across 2–17 is again the summation of 

the measurements across 2–8 and 8–17. The pressure drops are slightly less than those presented 

in Figure 4.4 for a given velocity. (compiled from data files, “unblocked2 full 17-8 partial 8-

2.csv,” “unblocked3 full 8-17 partial 2-8.csv,” and “unblocked4 full 8-17 partial 2-8.csv,” 

created September 28, October 3, and October 12, 2005, respectively). Again, the pressure drops 

are plotted versus the corresponding section bundle velocity. The bundle velocities shown in this 

plot were calculated using the bundle area only and are discussed next. The hydraulic diameter 

and flow area of the lower portion of the assembly were used to calculate SLAM and k coefficients 

for the 2–17 data. 

ΔP (N/m2) 
 

 73 

y = 3.69E+00x2 + 5.33E+00x

R2

 = 9.93E-01

y = 9.47E+00x2

 + 2.07E+01x

R2

 = 9.98E-01

y = 1.26E+01x2

 + 2.56E+01x

R2

 = 9.98E-01

0

10

20

30

40

50

60

70

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

Vbundle (m/s)

 P (N/m 2

)

2-8 8-17 2-17

Figure 4.7 Pressure drop as a function of bundle velocity for the SFP assembly with unblocked water 

rods 

 Note: The bundle velocities shown in the plot were calculated for all the air flowing through the 

bundle. 

This analysis was performed to directly compare with the previous blocked runs. The values 

used for this analysis are shown in Table 4.8. The inertial losses appear to be similar, but the 

SLAM coefficient decreased somewhat from the blocked water rod values in Table 4.3. Again, the 

hydraulic diameter and flow area of the lower portion of the assembly were used to calculate 

SLAM and k coefficients for the 2–17 data. 

Table 4.8 SLAM and k coefficient analysis data for pressure drops between 2–8, 8–17, and 2–17 for 

unblocked water rods 

All flow is assumed to pass through the bundle for this analysis. 

Pressure 

Drop L(m) A (m2

) DH (m) 

a1

(N·s/m3

) 

a2

(N·s2

/m4) SLAM  k 

# of 

spacers 

k per 

spacer 

2–8 1.43 0.0106 0.0141 5.33 3.69 80 7.3 2 3.7 

8–17 2.56 0.0098 0.0119 20.7 9.47 125 19 5 3.8 

2–17 4.00 0.0098 0.0119 25.6 12.6 99 25 7 3.6 

A summary of the SLAM and k coefficients for unblocked water rod testing is shown in Table 4.9. 

Again, this analysis assumes all flow passes through the bundle. The hydraulic diameter and flow 

area of the fully populated section was used to calculate the SLAM and k values for any span 

including the lower section, e.g., 1–B 

ΔP (N/m2) 
 

 74 

Table 4.9 SLAM and k coefficients for the unblocked water rod assembly assuming all flow passes 

through the bundle 

Description Segment A (m2

) L (m) DH (m) SLAM  k 

# of 

spacers 

k per 

spacer 

Long Spans 1–B 0.0098 4.19 0.0119 96 37 – – 

2–8 0.0106 1.44 0.0141 80 7.3 2 3.7 

8–17 0.0098 2.56 0.0119 125 19 5 3.8 

2–17 0.0098 4.00 0.0119 99 25 7 3.6 

4–8 0.0106 0.98 0.0141 75 3.4 1 3.4 

8–15 0.0098 2.05 0.0119 130 14 4 3.5 

4–15 0.0098 3.03 0.0119 104 17 5 3.4 

Tie Plates 1–2 0.0106 0.04 0.0141 64 0.39 – – 

17–B 0.0098 0.13 0.0119 172 12 – – 

Next, the above analysis was repeated using the sum of the bundle and water rod areas to 

determine the bundle velocity. This approach sets the velocity of the air inside the water rods and 

the bundle to be equal. For this case, the fraction of flow through the water rods is 7.5 to 8.1% of 

the total flow rate for the upper and lower assembly sections, respectively. These values are 

simply the ratio of the water rod area to the total flow area (bundle plus water rod area) in the 

two sections. Table 4.10 lists a summary of the SLAM and k coefficients calculated from this 

analysis. The tie plates are omitted since the water rods have no flow area through these sections. 

The hydraulic diameter and flow area of the fully populated section was used to calculate the 

SLAM and k values for any span including the lower section. 

Table 4.10 SLAM and k coefficients for the unblocked water rod assembly assuming equal air velocities 

in the bundle and water rods 

Description Segment A (m2

) L (m) DH (m) SLAM  k 

# of 

spacers 

k per 

spacer 

Long Spans 1–B 0.0106 4.19 0.0124 113 43 – – 

2–8 0.0114 1.44 0.0145 92 8.6 2 4.3 

8–17 0.0106 2.56 0.0124 147 22 5 4.4 

2–17 0.0106 4.00 0.0124 117 30 7 4.3 

4–8 0.0114 0.98 0.0145 86 4.0 1 4.0 

8–15 0.0106 2.05 0.0124 154 17 4 4.3 

4–15 0.0106 3.03 0.0124 123 20 5 4.0 

4.2.3 Water Rod Flow Rate 

In the previous section, the air flow rate through the water rods was assumed to be based on flow 

area fractions. The water rod flow rate may be explicitly determined by comparing pressure drop 

data for both blocked and unblocked water rod assemblies. Figure 4.8 shows the total air flow 

rate through the assembly as a function of pressure drop across 2–17. This particular pressure 

drop was chosen because of the abundance of existing data across 2–17 and the fact that it spans 

the entire water rod. Power law fits were chosen to continuously represent the data based on a 

goodness of fit criteria. The flow rate through the bundle must be the same to create the same 

pressure drop. Therefore, the difference in the flow rate represented by the two curves is the flow 

rate through the water rods. 
 

 75 

y = 5.344E-04x7.823E-01

R2 = 9.997E-01

y = 5.811E-04x7.738E-01

R2 = 9.996E-01

0.0E+00

5.0E-03

1.0E-02

1.5E-02

0 10 20 30 40 50 60 70

P (N/m2

)

Qtot (m 3/s)

  Blocked

  Unblocked

Figure 4.8 Total air flow rate through the assembly as a function of pressure drop across 2–17 for 

blocked (closed) and unblocked (open) water rods 

Figure 4.9 shows the fraction of flow rate through the water rods as a function of total flow rate 

through the assembly. This curve was determined from the pressure drop data across 2–17. The 

experimental data span a range of 1.110-3 to 1.310-2 m3

/s. Initially, the flow rate through the 

water rods decreases sharply with increasing total flow rate. The flow fraction then approaches a 

value of approximately 5% at the highest total flow rate. 

Qunblocked - Qblocked = Qwr
 

 76 

0.00

0.02

0.04

0.06

0.08

0.10

0.0E+00 5.0E-03 1.0E-02 1.5E-02

Qtot (m3

/s)

Qwr / Q tot

Figure 4.9 Fraction of flow rate through the water rods as a function of total flow rate as determined 

from the 2–17 pressure drop data 

Now, the SLAM and k analysis shown in Section 4.2.2 may be repeated with experimentally 

determined water rod flow rates. Figure 4.10 shows the bundle pressure drop as a function of the 

bundle velocity calculated using the experimentally determined flow rate through the bundle 

(Qtot – Qwr). As expected, the bundle velocities are slightly less than those shown in Figure 4.7. 
 

 77 

y = 3.93E+00x2 + 5.79E+00x

R2

 = 9.93E-01

y = 1.00E+01x2

 + 2.24E+01x

R2

 = 9.98E-01

y = 1.34E+01x2

 + 2.78E+01x

R2 = 9.98E-01

0

10

20

30

40

50

60

70

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

Vbundle (m/s)

P (N/m 2

)

2-8 8-17 2-17

Figure 4.10 Pressure drop as a function of bundle velocity for the SFP assembly with unblocked water 

rods 

The bundle velocities shown in the plot were calculated using the experimentally determined 

bundle flow rate (Qtot – Qwr). 

The values used for the SLAM and k analysis of this pressure drop data are presented next in Table 

4.11. These values are quite similar to the blocked water rod values in Table 4.3. The hydraulic 

diameter and flow area of the lower portion of the assembly were used to calculate SLAM and k 

coefficients for the 2–17 data. 

Table 4.11 SLAM and k coefficient analysis data for pressure drops between 2–8, 8–17, and 2–17 for 

unblocked water rods 

 Note: Flow through the bundle (Qtot – Qwr) was determined experimentally for this analysis. 

Pressure 

Drop L(m) A (m2

) DH (m) 

a1

(N·s/m3

) 

a2

(N·s2

/m4) SLAM  k 

# of 

spacers 

k per 

spacer 

2–8 1.43 0.0106 0.0141 5.79 3.93 87 7.8 2 3.9 

8–17 2.56 0.0098 0.0119 22.4 10.0 135 20 5 4.0 

2–17 4.00 0.0098 0.0119 27.8 13.4 107 27 7 3.9 

Table 4.12 presents the results of the SLAM and k analysis using the experimentally determined 

bundle flow rate. These SLAM and k values are remarkably similar to the values obtained during 

the blocked water rod tests (Table 4.5). This agreement indicates that the flow rates through the 

water rods reported here are accurately portrayed over the experimental range of flow rates. The 

hydraulic diameter and flow area of the fully populated section was used to calculate the SLAM

and k values for any span including the lower section. 
 

 78 

Table 4.12 SLAM and k coefficients for the unblocked water rod assembly using the experimentally 

determined bundle flow rate (Qtot – Qwr) 

Description Segment A (m2

) L (m) DH (m) SLAM  k 

# of 

spacers 

k per 

spacer 

Long Spans 1–B 0.0098 4.19 0.0119 105 38 – – 

2–8 0.0106 1.44 0.0141 87 7.8 2 3.9 

8–17 0.0098 2.56 0.0119 135 20 5 4.0 

2–17 0.0098 4.00 0.0119 107 27 7 3.9 

4–8 0.0106 0.98 0.0141 81 3.6 1 3.6 

8–15 0.0098 2.05 0.0119 141 15 4 3.8 

4–15 0.0098 3.03 0.0119 113 18 5 3.6 

4.2.4 Bypass Hole Flow Rate 

Two bypass holes are located in the bottom tie plate (Figure 4.10). With the pool cell in place, 

these holes supply flow to the annular region between the canister and the pool cell. These holes 

were blocked for all testing referred to earlier in this report. To measure the flow rate through the 

bypass holes, the blockages were removed one at a time from these holes. The flow rate was then 

varied until the pressure drop across 1–B matched the blocked case. The difference in the flow 

rates was taken to be the flow rate through the bypass holes. These tests were conducted with the 

unblocked water rod configuration. 

Figure 4.11 shows the fraction of flow rate through the bypass holes as a function of total flow 

rate. The experimental range of flow rates was 2.210-3 to 5.410-3 m3

/s. For the prototypic 

situation of two open bypass holes, the fraction of total flow rate through these holes is 

considerable (up to 6.8% at the highest flow rate measured). These values represent a 

preliminary attempt to characterize the bypass flow rate in an isothermal assembly unbounded by 

a pool cell. Further investigations provided better insight into the flow across the bypass holes 

(See Figure 5.9). 
 

 79 

0.00

0.02

0.04

0.06

0.08

0.10

0.0E+00 2.0E-03 4.0E-03 6.0E-03

Qtot (m3

/s)

Qbyp/Q tot

1 hole 2 holes

Figure 4.11 Fraction of flow rate through the bypass holes as a function of total flow rate as determined 

from 1–B pressure drop data 

4.3 Summary 

Extensive hydraulic experiments have been conducted for a prototypic spent fuel assembly 

modeled after the prototypic 99 BWR to determine the appropriate SLAM and k coefficients for 

use with the MELCOR severe accident analysis code. The experimental range spanned the 

laminar region with volumetric flow rates of 1.110-3 to 1.310-2 m3

/s. The resulting Reynolds 

numbers based on the bundle velocity and hydraulic diameter were 70 to 900, respectively. 

These experiments included testing with the water rods in the assembly, both blocked and 

unblocked. The values best suited for use with MELCOR modeling the 99 prototypical BWR 

geometry are given in Table 4.9. 

4.3.1 Application of Experimental Results to a MELCOR Model 

MELCOR, like other control volume codes, includes constitutive relationships to specify form 

losses (i.e., minor losses) and wall friction losses (i.e., major or viscous) along a flow path as a 

hydraulic flow loss term to the momentum equation. Because a single MELCOR flow path may 

be used to represent a rather complicated flow path, the wall friction terms may be computed for 

one or more segments that are connected in series. The format of the user-specified input follows 

Equations 3 and 4. In the context of recent BWR SFP analysis applications, a BWR bundle may 

be divided into five control volumes with six connecting flow paths that represent the major and 

minor flow losses. Similarly, the interstitial bypass region may be analogously subdivided. 

As shown in Figure 4.1, the BWR fuel assembly contains seven grid spacers, upper and lower tie 

plates, full and partial rod regions, two water rods, and an inlet nozzle. It is generally not 
 

 80 

practical to include a detailed representation of all the geometry changes in the MELCOR model. 

In addition, the MELCOR code includes some hard-wired geometry models that further limit the 

modeling of the two large water rods inside the assembly and their associated flow. 

Consequently, modeling choices are required to represent the geometry of the BWR assembly. 

Relative to the hydraulic modeling in the MELCOR model, the experimental data are used in the 

following manner: 

 Based on the hydraulic impact of the fully populated versus the partial regions, the 

control volume boundaries were specified to span uniform geometry regions. Control 

volume boundaries were placed at the bottom of the lower tie plate, the transition from 

the fully populated rod region to the partial region, and at the top of the upper tie plate. 

By spatially dividing the two regions, the distinct flow loss effects can be extended to 

heated conditions, where the flow will accelerate along the length of the assembly. This 

division of hydraulic segments has been implemented into whole pool analyses. 

 The flow resistance in a flow path spans the region from cell-center of the lower control 

volume to cell-center of the upper control volume. Hence, the pressure in a given control 

volume represents the pressure in the center of the control volume. A single flow path 

may span more than one grid spacer and perhaps a tie plate. Segment data on a particular 

flow path defines the form and wall friction losses along the geometric regions 

encompassed in the flow path length. Several flow paths exhibited multiple flow 

segments due to geometry changes with the flow path and these net effects are calculated 

by the code. 

 For the initial application for BWR SFP analysis, the results from Table 4.9 were used, 

which include prototypical water rod flow effects. For the flow segments in the fully 

populated tube region, a SLAM of 125 was used and k losses of 3.8 were used for each 

spacer included in the range of the flow path. The flow area and hydraulic diameters were 

preserved. The total length across all flow paths in the fully populated region was 2.56 m 

and the total k was 19. Therefore, the total flow losses from 8–17 were exactly preserved, 

including the effect of flow within the water rods. Similarly, the partially populated tube 

region used a SLAM of 80 and k losses of 3.7 per spacer in the range of the flow path. The 

total length across all flow paths in the fully populated region was 1.44 m and the total k 

was 7.3. The flow path segments in the partial rod region used the larger flow area and 

hydraulic diameter as specified in 2–8 from Table 4.9. 

 Table 4.9 was used to specify the pressure drop across the upper tie plate (1–2), lower tie 

plates, and the inlet region (i.e., 17–B). Similar to above, the appropriate flow areas and 

hydraulic for these regions were used. 

 Finally, the results from Table 4.10 were used to specify the flow loss terms for the inlet 

nozzle leakage to the interstitial bypass region. The form losses in the MELCOR model 

were adjusted to match the measured total to bypass flow split. 

Until this hydraulic data became available, previous BWR SFP MELCOR analyses did not 

include the sophistication of separate flow resistances for the partial and fully populated regions 

of the BWR bundle. 

Other SLAM and k formulations found in this report are expected to be used for other analysis 

efforts. The following lists some on-going and expected future applications. 
 

 81 

 The results from the blocked water rods (Table 4.1 and Table 4.5) were used to estimate 

the flow resistance for a pressurized water reactor (PWR) assembly in a SFP rack. The 

PWR assembly has guide tubes, which may have some flow depending upon the internal 

storage configuration (i.e., various types of control rods and/or plugs may be used). 

However, in most circumstances, the guide tubes are not expected to have a significant 

flow. 

 The results in Table 4.6 may be used for more detailed scaling of BWR grid spacers to 

other types of grid spacers. They also help support the unexpected experimental finding 

of high flow losses in the grid spacer. 

 Using new modeling features in MELCOR, a water rod model was developed. The model 

uses explicit coupling to the assembly model and runs too slowly for “production” BWR 

SFP analysis. However, the water rod model and the water rod hydraulic data (Figure 

4.8) will be used in the experimental analysis to assess the impact of heating on the water 

rod hydraulics versus the faster running “production” model approach. The results from 

the blocked water rods (Table 4.3 and Table 4.5) were used to estimate the flow 

resistance in conjunction with the water rod model. 

Finally, it is worth noting that while the experimental analysis could not separately discern the 

contribution of wall friction distinct from form losses, the results analysis in the previous tables 

precisely satisfies the model input requirements for MELCOR, as well as other control volume 

codes, where the overall hydraulic losses are averaged across larger regions. The linear 

(i.e., SLAM) and quadratic (i.e., the k term) hydraulic loss coefficients are easily put into the 

MELCOR input format and will replicate the measured flow losses across the Reynolds number 

range of 70 to 900, including the overall nonlinear hydraulic effects from: 

 Flow development regions, 

 the grid spacer entrance, internal, and exit effects, 

 the inlet and exit flows to the water rods, and 

 the transition from the fully to partially populated tubes regions. 

 

 83 

5 THERMAL−HYDRAULIC ANALYSIS 

OF THE SPENT FUEL POOL EXPERIMENT 

This report summarizes the findings of the SFP thermal-hydraulic experiments conducted 

between January 16 and February 16, 2006. The stated purpose of these investigations was to 

determine the thermal and hydraulic response for the validation of both the MELCOR severe 

accident analysis code and the COBRA spent fuel storage code. Two annular flow configurations 

were considered, namely closed bypass and open drain holes and open bypass and closed drain 

holes. The power applied to the apparatus spanned 200 to 2500 W with inlet natural draft rates 

for the open bypass/closed drain configuration of 83 to 122 slpm, or Reynolds numbers of 100 to 

150 in the bundle. 

5.1 Experimental Apparatus and Procedures 

5.1.1 Hardware Components 

Commercial components were purchased to create the prototypic 99 BWR assembly including 

the top and bottom tie plates, spacers, water rods, channel box, and all related assembly 

hardware. Incoloy heater rods were substituted for the fuel rod pins for heated testing. The 

diameter of the Incoloy heaters was slightly smaller than prototypic pins, 1.0910-2 m versus 

1.1210-2 m. The slightly simplified Incoloy mock fuel pins were fabricated based on drawings 

and physical examples of a prototypic 9x9 BWR assembly. The dimensions of the assembly 

components are listed in Table 5.1. 

The entire fuel assembly was housed inside a stainless steel enclosure, simulating the rack walls 

of a SFP. This enclosure was then insulated with 0.15 m (6 in.) of Fiberfrax Durablanket® type S 

insulation. Two radiation barriers, consisting of stainless steel shim stock, were placed at 

0.075-m (3-in.) intervals. 

Table 5.1 Dimensions of assembly components in the 99 BWR 

Description Lower (Full) Section Upper (Partial) Section 

Number of Pins 74 66 

Pin Diameter (m) 1.0910-2 1.0910-2

Pin Pitch (m) 1.4410-2 1.4410-2

Pin Separation (m) 3.4810-3 3.4810-3

Water Rod OD (main section) (m) 2.4910-2 2.4910-2

Water Rod ID (m) 2.3410-2 2.3410-2

5.1.2 Thermocouple Layout 

A total of 135 TCs were used to characterize the thermal response of the apparatus. TCs were 

placed both internal and external to the bundle. Figure 5.1 shows schematically the location of all 

internal TCs. The TCs were named according to the two-axis alpha scheme followed by the axial 

height, as referenced from the top of the bottom tie plate. For example, the TC at CS119 was the 

third rod from the left and third rod from the top at the 119-in. axial level (Figure 5.1). 
 

 84 

Internal Thermocouples 

Cross section 

above partial 

z = 0"

Top of bottom tie

plate reference

Bypass holes 

– 2 places 

24" 

48" 

72" 

96" 

119" 

144" 

Radial array

24" spacing 

11 TC each level 

66 TC total 

Axial array A1 

6" spacing 

20 TCs 

Axial array A2 

12" spacing 

7 TCs 

Water rods inlet and exit 

4 TCs 

Total of 97 TCs 

a b c d e f g h i 

q

r 

s 

t 

u

v

x

y

z

24" and 96" levels 

48" and 119" levels 

a b c d e f g h i 

q

r 

s 

t 

u

v

x

y

z

(a) (b)

72" and 144" levels 

a b c d e f g h i 

q

r 

s 

t 

u

v

x

y

z

Figure 5.1. Experimental mock fuel assembly showing internal, as-built (a) axial and (b) lateral 

thermocouple (TC) locations 
 

 85 

External TCs were placed on the channel box, pool cell, and outer radiation barrier as shown in 

Figure 5.2. The arrangement of the insulation and radiation barriers is also illustrated. 

Figure 5.2 Arrangement of insulation, radiation barriers, and external TCs 

5.1.3 Flow Measurements and Calibrations 

Prior to heated testing, pressure drop measurements were acquired across the bottom tie plate 

(17–B) and the entire assembly (1–B). All hydraulic loss coefficients, SLAM and k, were within 

experimental uncertainty of previous hydraulic testing (stainless steel assembly) with the 

exception of form loss of the entire assembly (Table 5.2). The uncertainties in these data are 

uSLAM = ±5 and uk = ±1. 

Table 5.2 SLAM and k analysis across 1–B and 17–B for Incoloy and stainless steel assemblies. 

Note: All flow is assumed to pass through the bundle. 

Description Segment A (m2

) L (m) DH (m) SLAM  k 

Incoloy 1–B 0.0100 4.19 0.0124 98 34 

17–B 0.0100 0.13 0.0124 171 11 

Stainless 

Steel 

1–B 0.0098 4.19 0.0119 96 37 

17–B 0.0098 0.13 0.0119 172 12 

Two flow configurations were studied during the course of testing, referred to as open 

bypass/closed drains and closed bypass/open drains. By restricting the annular flow inlet to one 

pathway, the bundle and annular flow rates could be determined with much greater certainty. 

Five hot wire anemometers were used to characterize the flow at various locations in the 

apparatus. The assembly inlet was instrumented with two Omega FMA-900-V-R constant 

temperature hot wires as shown in Figure 5.3. For all heated testing, these probes were inserted a 

distance of 25.4 mm as measured from the pipe wall. 

Radiation 

barriers 

Thermocouples
 

 86 

Figure 5.3 Schematic showing the layout of the assembly inlet hot wires (hot wires 2 and 4) 

Figure 5.4 shows the velocity inside the assembly inlet pipe as a function of distance from the 

inner pipe wall and input flow rates of 50 and 125 slpm. The figure also shows the fixed location 

of the sensor head during heated testing, which is approximately 6.4 mm inside the protective 

sheath. These probes were positioned at this location to measure inside the bulk of the flow while 

minimizing blockages due to the insertion of the sensors. The integrated average values were 

used to verify the output of the hot wires with the input of the mass flow controllers. Table 5.3 

shows the comparison of the input flow rates versus the integrated average values. Typical errors 

were less than 7% with a maximum error of 9.39%. The volumetric flow rate measured with hot 

wire 4 was consistently high, perhaps due to the blockage introduced into the pipe during the 

insertion of the instrument. The hot wire represents a 4.5% area blockage when inserted to the 

center line of the pipe. 

101.6 

241.3

Anemometer 

ports – 2 places

A-A 

Section A-A 

25.4 

Traverse

direction

Hot wire 2

Hot wire 4

Fixed 

location 

25.4

6.4 

All dimensions in mm 

89 ID 

(3.5") 
 

 87 

0.0

0.1

0.2

0.3

0.4

0.5

0.00 0.01 0.02 0.03 0.04 0.05

Stage (m)

Velocity (m/s)

Centerline

Integrated 

average values 

Fixed probe 

location 

Q = 

50 125 slpm 

Figure 5.4 Hot wire 4 velocity traverse across the assembly inlet pipe for input volumetric flow rates of 

50 (red) and 125 slpm (blue) 

Table 5.3 Integrated flow rates from the hot wire velocity profiles 

Input flow 

rate (slpm) 

Integrated average 

velocity (m/s) 

Integrated flow rate 

(slpm)   %

Input

Integrated  Input

50 0.143 53.3 6.58 

75 0.220 82.0 9.39 

100 0.285 106 6.18 

125 0.358 133 6.56 

Figure 5.5 shows the calibrations for the inlet hot wires performed on January 20 and January 31, 

2006. These calibrations were used to determine the volumetric flow rate into the assembly inlet 

for the heated tests. The difference in outputs in the two hot wires is attributed to slight variations 

in positioning and instrument gains. The electronics of hot wire 2 were replaced prior to recalibration on January 31, 2006, due to failure of the original unit on January 26, 2006. This 

exchange caused an overall decrease in the instrument gain reflected by the change in the sensor 

output. 
 

 88 

HW2 - 1/20/06

y = 341.2x

R2 = 0.9986

HW4 - 1/20/06

y = 319.2x

R2

 = 0.9985

HW2 - 1/31/06

y = 435.0x

R2

 = 0.9993

HW4 - 1/31/06

y = 316.8x

R2

 = 0.9998

0

50

100

150

200

250

0 0.1 0.2 0.3 0.4 0.5 0.6

V (m/s)

Q (slpm)

HW2 - 1/20/06

HW4 - 1/20/06

HW2 - 1/31/06

HW4 - 1/31/06

Figure 5.5 Inlet hot wire calibrations for hot wire 2 (HW2) (diamonds) and hot wire 4 (HW4) (squares) 

on 1/20/06 (closed) and 1/31/06 (open) 

Figure 5.6 gives the location of the remaining two Omega FMA-900-V-R hot wires, hot wires 1 

and 3, in the drain hole entry pipes. Again, the protective sheath was inserted to a depth of 25.4 

mm from the inner pipe wall. The flow path through the drain holes into the annulus began 

through the drain entry pipes with an inner diameter of 40.6 mm (1.6 in.) followed by an abrupt 

contraction to a hole of 19.1-mm diameter. 

>DAQ System 

DAQ 

Input 

Hot wire 1 Hot wire 3

B-B

Section A-A

25.4

19.1 

A-A

Section B-B

40.6 

All dimensions in mm 

Figure 5.6 Schematic showing layout of the drain hole hot wire (HW1 and HW3) 

Sensor Output (-) 
 

 89 

Calibrations similar to before were performed for hot wires 1 and 3. Figure 5.7 shows the 

calibration curves for the drain hole hot wires conducted on January 20 and January 31, 2006. 

The electronics for hot wire 1 were damaged by the same electrical failure noted with hot wire 2 

on January 26, 2006. A substitute component was used for the calibration performed on January 

31, 2006, which resulted in an output signal of approximately half the original values for a given 

flow rate. 

HW3 - 1/20/06

y = 151.0x - 1.283

R2

 = 0.9964

HW1 - 1/20/06

y = 142.2x - 1.486

R2 = 0.9940

HW1 - 1/31/06

y = 246.8x + 1.414

R2

 = 0.9924

HW3 - 1/31/06

y = 156.2x - 1.308

R2 = 0.9947

0

20

40

60

80

100

120

0.0 0.1 0.2 0.3 0.4 0.5 0.6

Sensor Output (-)

Q (slpm)

HW1 - 1/20/06

HW3 - 1/20/06

HW1 - 1/31/06

HW3 - 1/31/06

Figure 5.7 Drain hole hot wire calibrations for hot wire 1 (HW1) (diamonds) and hot wire 3 (HW3) 

(squares) on 1/20/06 (closed) and 1/31/06 (open) 

Additional hot wire measurements were conducted to characterize the flow in the annulus. See 

Appendix C for further details. The flow in the annulus was shown to be nearly laminar and 

confirmed other measurements of the annular flow rate within instrument capabilities. 

The pressure drop across the bypass holes was measured to determine the flow rate in the 

annulus for testing with open bypass/closed drain holes (Figure 5.8). First, a bottom tie plate was 

externally calibrated to measure the pressure drop for different flow rates with known flows 

passing only through the bypass holes. For heated testing, the instrument was again configured to 

measure the pressure drop across the bypass. Q (slpm) 
 

 90 

Figure 5.8. Schematic of the measurement of the pressure drop across the bypass holes for 

(a) calibration and (b) heated testing 

Figure 5.9 shows the results of the bypass pressure drop calibrations from the external tie plate 

and previous hydraulic testing efforts. Also, the analytic solution for an infinite contraction and 

expansion is shown (kc = 0.5, ke = 1). The actual pressure drop indicates that the infinite limits of 

loss coefficients are slightly too high. A power law curve fit provides a reasonable estimate of 

the bypass volumetric flow rate as a function of pressure drop. 

#.# N/m2

Flow 

P 

#.# N/m2

Flow 

Open to 

ambient 

a) Calibration setup b) Experimental setup 

Flow 

blocked

P 
 

 91 

y = 4.032x0.5475

R2 = 0.9999

0

5

10

15

20

25

30

0 5 10 15 20

Bypass P (N/m2

)

Annular Flow Rate (slpm)

Hydraulic Assembly

Analytic - Hydraulic Losses

Bottom Tie Plate Calibration

Figure 5.9. Annular flow rate as a function of the bypass pressure drop for the hydraulic assembly 

(diamonds), analytic solution (line), and the external bottom tie plate calibration (squares) 
 

 92 

5.2 Modeling Efforts 

5.2.1 COBRA-SFS (Coolant Boiling in Rod Arrays-Spent Fuel Storage) 

COBRA-SFS is a lumped parameter finite difference code, written in FORTRAN, used to 

conduct thermal-hydraulic analysis of multi-assembly spent fuel storage. It is a fully implicit 

code which utilizes the Newton-Raphson technique to solve the equations of mass, momentum, 

and energy conservation. The code solves three-dimensional convective and conductive heat 

transfer and two-dimensional radiation heat transfer. The flow solution is limited to single-phase 

flow with or without buoyancy driven natural circulation. It can also determine flow and pressure 

fields in which the net flow is zero, simulating spent fuel stored in casks. COBRA-SFS has been 

validated and verified with experimental data for both steady state and transient calculations. 

The code is broken down into fluid channels, rods, solid nodes, radiation heat transfer, and 

boundary conditions. Both the fluid and solid thermal properties are temperature dependant. The 

fluid channels describe the flow geometry between the rods and other solid structures. The rods 

are solid nodes containing an inner core and outer cladding. The solid nodes define every other 

structure in the assembly such as the canister, pool cell, and storage cask. Radiation heat transfer 

is assumed to be gray and diffuse. The view factors between the rods and canister are generated 

from a supplementary program called RADGEN. This program generates view factors for a 

variety of rod arrangements. The view factors between other components in the assembly are 

manually calculated using Hottel’s crossed-string method. The boundary conditions are broken 

down into five groups: heat transfer, flow drag, plenum and outer boundary conditions, power 

and loading, and inlet/outlet conditions. All the information is entered into an input file in which 

an executable file of the FORTRAN code solves the model. 

5.2.1.1 Full Length Heated Assembly Model 

The COBRA model included relatively high lateral detail for the assembly components and rod 

array as shown in Figure 5.10 and Figure 5.11. Model details are summarized below. 

 Various powers with two flow arrangements: two open bypass holes or two open drain 

holes. 

 Specified steady state flow rate through both the annulus and bundle based on 

experimental values at 12 hours. 

 Eight partial rods included (102 in.): power across the entire axial length. 

 Internal flow assumed to be laminar (Nu = 4.36 bundle, Nu = 5.385 annulus). 

 Temperature dependant properties for materials and air. 

 Full length rods: 26 axial levels, 156-in. long, axial power from 0–144 in. 
 

 93 

Figure 5.10. Schematic of the COBRA model layout 

Figure 5.11. Detail view of the COBRA bundle layout 
 

 94 

5.2.2 MELCOR – Severe Accident Analysis 

The MELCOR 1.8.5 severe accident computer code with enhancements through Version RP was 

used to simulate the experimental SFP accident response. MELCOR is a full non-equilibrium, 

two-phase thermal-hydraulics code for the simulation of nuclear reactor severe accidents. It 

includes a full spectrum of models for simulation of reactor accidents progressing from design 

conditions to severe accidents with: 

 Comprehensive fuel degradation, 

 Steam and air oxidation with hydrogen production, 

 Fuel and structural component degradation, 

 Fission product release, transport, deposition, chemisorption, vapor condensation and 

evaporation, aerosol hygroscopic effects, iodine pool evolution, and gamma heating, 

 Core-concrete interactions, and 

 Simulation of containment engineering safety features (e.g., sprays, fan coolers, filters, 

and recombiners). 

Of particular relevance for application to the SNL experimental SFP tests, MELCOR has new 

fuel geometry models for boiling and pressurized water reactor SFP rack geometry. The 

specialized SFP geometry models include basic heat transfer models for convection, conduction, 

and radiation as well as a new breakaway oxidation kinetics model and fuel and rack degradation 

models. 

Version RP includes three recent modeling enhancements applicable to BWR SFP modeling: 

3) A new rack component, which permits better modeling of a SFP rack, 

4) A new oxidation kinetics model, and 

5) A simplified flow regime model. 

The new BWR SFP rack component permits proper radiative modeling of the SFP rack between 

groups of different assemblies. The new oxidation kinetics predicts the transition to breakaway 

oxidation in air environments on a node-by-node basis. The simplified flow regime model 

permitted simulation of liquid films draining down the BWR fuel assemblies during spray 

operation. 

5.2.2.1 Full Length Heated Assembly Model 

A MELCOR full-length assembly model was developed to analyze the temperature response to 

the SNL experimental SFP testing program. The model simulates a single assembly in a uniform 

pattern (Figure 5.12). For implementation into MELCOR, the uniform pattern model was 

represented by a single, high-powered assembly with a heavily insulated radial boundary (i.e., 

simulated with 6 in. of KaowoolTM insulation). 
 

 95 

14 

Pattern

Uniform 

Pattern

Checkerboard 

Pattern 

Recently discharged, high-powered assembly 

Low-powered assembly discharged many years earlier 

Figure 5.12. Repeating patterns used in the MELCOR separate effects models 

The interstitial area between the canister wall and the rack wall is modeled as the BWR bypass 

region. The bypass uses the same 9-control volume axial nodalization as the bundle region. A 

control volume represents the inlet tube connected to the bottom of the assembly. A flow path 

was included that represented the two tubes connected to the rack drain holes. The inlet tube, 

drain lines, and assembly exit were connected to a time-independent volume that represented the 

atmospheric conditions during the test (i.e., pressure and temperature). 

The hydraulic resistance was specified using the results from the experimental test program. The 

drain line and inlet tubes included standard losses for flow contractions, expansions, and wall 

friction. 

The BWR assembly canister is modeled with the MELCOR canister component. The rack walls 

are modeled using the new SFP rack component with stainless steel and aluminum (i.e., to 

represent Boral). MELCOR does not include an option to model the two large water rods in the 

assembly. Consequently, the water rod mass and surface area were included in the canister wall. 

The BWR assembly is represented by 14 axial levels in the COR package: 

 Level 1 is the pipe leading to the inlet of the assembly. 

 Level 2 is the base plate. 

 Level 3 is the inlet region between the inlet nozzle and the lower tie plate. 

 Level 4 is the start of the active fuel region. 

 Level 8 is the top of the partial rod active fuel region. 

 Level 9 represents the region of the plenum of partial rods. 

 Level 12 is the top of the full-length rod active fuel region. 

 Level 13 represents the plenum region of the full-length rods. 

 Level 14 represents the region between the upper tie plate and the top of the racks. 
 

 96 

The upper and lower tie plates are modeled as supporting plate structures, made of stainless steel. 

The 0.5-in. rack support plate at Level 2 is also modeled as a stainless steel supporting structure. 

MELCOR requires that the canister, water rods, and fuel cladding be specified as the same 

material. The default Zircaloy properties for these components were replaced with inconel. In 

reality, the experimental apparatus had a Zircaloy canister, Zircaloy water rods, and Incoloy 

heater rods. MELCOR’s Zircaloy oxidation kinetics model was disabled to prevent oxidation of 

the heater rods. The interior of the heater rods was modeled as compacted magnesium oxide. 

MELCOR’s gamma heating model was disabled to deposit all the power from the heater rods in 

the simulated fuel (i.e., the magnesium oxide). 

A constant power profile was used in each heater rod. Because eight partial heater rods did not 

extend the full length of the assembly, there was a step change in the overall, axial power profile 

at the top of the heated zone of the partial rods. 
 

 97 

CV-100

CV111 

CV119 

CV112 

CV113 

CV114 

CV115 

CV116 

CV117 

CV118 

Bundle 

CV211 C2119 CV212 CV213 CV214 CV215 CV216 CV217 CV218 Bypass 

CVH 

Nodalization

Core

Nodalization

Bottom 

tie plate 

Lower section, fully 

populated 

Upper section, 

partially 

populated 

Spacer 

Top tie 

plate 

Active 

heater 

zone 

4 

14 

5 

6 

7 

10 

8 

9 

11 

12 

13 

1 

2 

3 

Axial

Level 

Rack baseplate 

Inlet tube 

Inlet drain 

tubes (2)

Figure 5.13. MELCOR nodalization of the full-length assembly experiment 

5.3 Results 

5.3.1 Transient Thermal Response 

The transient thermal response of the apparatus at 0.610-m (24-in.) increments for an input 

power of 1370 W and closed bypass/open drains flow configuration is shown in Figure 5.14. The 

COR 
 

 98 

location of the peak temperature in the assembly increases upward with time until the top of the 

stack is the hottest at around 7 hours. 

300

400

500

600

700

0 2 4 6 8 10 12 14

Time (hrs)

Temperature (K)

24" 

48" 

72" 

96" 

119" 

144" 

Figure 5.14 Transient thermal response at different axial locations for a power input of 1370 W and 

closed bypass/open drains flow configuration 

The transient thermal response for an input power of 1370 W and open bypass/closed drains flow 

configuration in 0.610-m (24-in.) increments is shown Figure 5.15. This configuration is 

considered the more prototypic flow arrangement in an actual SFP. The location of the peak 

temperature in the assembly increases upward with time until the top is hottest at around 10 

hours. The peak temperature at 12 hours was 36 K hotter due to lower overall induced air flow 

rates and, thus, less heat convected out of the assembly. 
 

 99 

300

400

500

600

700

0 2 4 6 8 10 12 14

Time (hrs)

Temperature (K)

24" 

48" 

72" 

96" 

144" 

119" 

Figure 5.15 Transient thermal response at different axial locations for a power input of 1370 W and open 

bypass/closed drains flow configuration 

Figure 5.16 shows the comparisons of the MELCOR and COBRA codes to the 144-in. level 

temperature response for the test conducted with an input power of 1370 W and open 

bypass/closed drains configuration. The codes are within 3% error of the experimental values at 

all times. MELCOR predictions of temperature are somewhat better, although under predictive. 
 

 100 

300

400

500

600

700

0 2 4 6 8 10 12

Time (hrs)

Temperature (K)

Experimen

t

MELCOR

Experiment

MELCOR

COBRA

Figure 5.16. Comparison of the thermal transient response with the MELCOR and COBRA codes 

for a power input of 1370 W and open bypass/closed drains flow configuration 

Figure 5.17 shows a comparison of the experimental (symbols) and code results (lines) for an 

input power of 1370 W and a flow configuration of closed bypass/open drains. These 

temperature results are presented as a function of vertical location in the assembly at 3 to 12 

hours in increments of 3 hours. The code values are within 10% error over the entire assembly. 

However, these errors are within 6% for the top half of the assembly. 

Figure 5.18 is similar to the previous graph for the open bypass/closed drains configuration. 

Here, the codes predict temperatures within 8% error of the experimental results. The top half of 

the assembly is again more accurately portrayed with errors of less than 5%. 
 

 101 

300

400

500

600

700

0 24 48 72 96 120 144 168

z (in)

Temperature (K)

3 hrs

6 hrs

9 hrs

12 hrs

––––– MELCOR

–– –– COBRA

––––– MELCOR

–– –– COBRA

Figure 5.17 Comparison of experimental (symbols) to MELCOR (solid lines) and COBRA (dashed lines) 

temperatures as a function of axial location at times of 3 (), 6 (), 9 (), and 12 hours () 

 Note: The apparatus was configured with closed bypass/open drains and an input power of 1370 

W. 

300

400

500

600

700

0 24 48 72 96 120 144 168

z (in)

Temperature (K)

3 hrs

6 hrs

9 hrs

12 hrs

––––– MELCOR

–– –– COBRA

––––– MELCOR

–– –– COBRA

Figure 5.18 Similar to the figure above for the open bypass/closed drains configuration 
 

 102 

5.3.2 “Steady State” Peak Cladding Temperature 

Figure 5.19 shows the PCT of the assembly as a function of input power for test results 

(symbols), MELCOR (blue lines), and COBRA (red lines) in both flow configurations. 

MELCOR calculations were within 5% error for all cases. COBRA results for the closed 

bypass/open drains configuration was also within 5% error. However, errors of up to 15% were 

incurred while modeling the open bypass/closed drains cases. 

300

500

700

900

1100

0 500 1000 1500 2000 2500 3000

Assembly Power (Watts)

PCT (K)

Experiment

MELCOR

COBRA

Experiment

MELCOR

COBRA

Solid lines, Closed symbols - Open Bypass Experiment

Dashed lines, Open symbols - Open Drains

Figure 5.19 PCT as a function of assembly input power for experimental (symbols), MELCOR (blue 

lines), and COBRA (red lines) 

 Note: Data for both flow configurations are shown, open bypass/closed drains (solid lines, closed 

symbols) and closed bypass/open drains (dashed lines, open symbols). 

5.3.3 Steady State Natural Induction Flow Rates 

Figure 5.20 gives the naturally-induced volumetric flow rates in the assembly inlet and annulus 

as a function of input power. The red symbols and lines represent the open bypass/closed drains 

configuration; the blue entries represent the closed bypass/open drains configuration. Only 

predictions from MELCOR are shown, as flows were specified as inputs into COBRA. 

MELCOR results for the open bypass configuration are within experimental uncertainty for all 

but the lowest two input powers. The code tended to under predict the inlet flow rates for the 

open drains configuration but calculated annular flow rates to within uncertainty for both flow 

configurations. All code predictions were within 10% error. 
 

 103 

0

25

50

75

100

125

150

0 500 1000 1500 2000 2500 3000

Assembly Power (Watts)

Flow Rate (slpm)

Figure 5.20 Volumetric flow rates as a function of assembly input power forexperimental (symbols) and 

MELCOR (lines) 

 Note: The data represent the bundle (diamonds) and the annulus (squares) flow rates. Both flow 

configurations are shown, open bypass/closed drains (solid red lines, closed symbols) and closed 

bypass/open drains (dashed blue lines, open symbols). 

5.4 Summary 

Benchmark experiments were conducted with a highly prototypic mock BWR spent fuel 

assembly to measure thermal and hydraulic response. The two flow configurations studied were 

open bypass/closed drains and closed bypass/open drains. The open bypass configuration is 

considered the more prototypic flow setup. Peak cladding temperatures for both experiments and 

MELCOR were within 5% error at all assembly power inputs and flow configurations. Predicted 

temperatures from COBRA were within 5% and 15% error for the closed bypass/open drains and 

open bypass/closed drains configurations, respectively. Naturally-inducted flow rates were 

within 10% error over the entire power input range between tests and MELCOR. 

The overall agreement of the MELCOR calculations and the experimental data is attributed to 

the direct application of prototypic viscous and form loss coefficients determined in previous, 

unheated testing. 

Table 5.4 gives a summary of all heated testing of the apparatus. Measurements of note are listed 

for values at 12 hours of elapsed test time. 
 

 104 

Table 5.4 Summary of all thermal-hydraulic tests for the SFP experiment 

Test Configuration Assembly Temperatures (oC) Flow Rates (slpm)

Test Date

True RMS 

Power (W)

# of drain 

holes

# of bypass 

holes PCT

Max. Channel 

Temp.

Max. Pool 

Cell Temp. Inlet Annulus

01/24/06 202 0 2 99 91 84 83 9

01/26/06 402 0 2 158 144 133 106 11

02/01/06 902 0 2 282 266 250 119 14

01/17/06 1368 0 2 392 377 359 122 16

02/02/06 1801 0 2 468 441 435 118 17

01/19/06 2240 0 2 562 542 527 119 18

02/15/06 2505 0 2 589 568 557 110 17

01/23/06 200 2 0 101 92 85 69 42

01/25/06 403 2 0 150 138 125 94 53

01/16/06 1369 2 0 356 339 320 116 70

02/08/06 1812 2 0 440 423 407 113 72

01/18/06 2236 2 0 515 498 482 105 74

02/16/06 2505 2 0 532 516 502 99 75

01/27/06 1362 2 2 346 328 312 116 70

*Flow rates and temperatures reported at 12 hrs elapsed test time
 

 105 

6 SUMMARY OF THE 14 INCOLOY SHORT STACK THERMAL 

RADIATION EXPERIMENTS OF THE SFP PROJECT 

This report summarizes the findings of the SFP thermal radiation experiments conducted 

between February 27 and March 23, 2006. The stated purpose of these investigations was to 

determine the thermal response of a spent fuel in a 14 arrangement for the validation of both the 

MELCOR severe accident analysis code and the COBRA spent fuel storage code. The radiative 

heat transfer between assemblies was of particular interest. To this purpose, the experimental 

apparatus was specifically constructed to maximize the radiative heat transfer between a single, 

hot center assembly and four, cooler neighbors. 

6.1 Assembly Design 

The 14 Incoloy assembly was designed to study the radiative heat transfer of a centrally located 

hot bundle to its cooler neighbors (Figure 6.1). Bulk air flow was prevented during testing by 

blocking all flow paths into and out of the assembly. However, the apparatus was not leak tight 

and did allow volumetric expansion of the air as it heated. The entire assembly height was 

approximately 52 in., modeling about one-third of a full-length assembly. 

Figure 6.1 Schematic and pictures of the assembly design 

Note: These depictions do not show the insulation applied to the outside of the pool cell rack or to the corner cells. 
 

 106 

Commercial components were purchased to create the assembly, including two rod spacers, 

water rods, channel boxes, and the pool cell rack. The water rods and channel boxes were cut to 

size (52 in. high) from full-length prototypic components. Incoloy heater rods were substituted 

for the fuel rod pins for heated testing. The diameter of the Incoloy heaters was slightly smaller 

than prototypic pins, 1.0910-2 m versus 1.1210-2 m. The heater rods had unheated regions of 2 

in. at the top and bottom, resulting in a 48-in heated zone. The spacers were placed at 16 in. and 

36 in. as measured from the top of the bottom tie plate. 

Figure 6.2 shows the assembly after 6 in. of KaowoolTM insulation was applied to the exterior of 

the apparatus. The cables visible in the top and side views are the electrical feeds to the five 

mock fuel assemblies. The alumina pipes extending through the insulation in the top view 

connect to the water rods in each bundle. Cooling air was fed through the bottom of the 

apparatus into the water rods and out the alumina tubes at the conclusion of heated testing. 

Figure 6.2 Photographs showing the final insulated assembly 

6.1.1 TC Layout 

Figure 6.3 depicts the general TC layout of the 14 Incoloy assembly. Three main TC levels 

were placed at heights of 11, 26, and 41 in. as measured from the top of the bottom tie plate. The 

solid circles indicate heater rods with a TC attached. 
 

 107 

West

East

North

Top level 41" Middle level 26" Bottom level 11" 

South 

Figure 6.3 Overall TC layout 

Figure 6.4 provides the location and naming convention of the TCs in the center assembly. A list 

of all installed TCs is presented below the figure. The naming convention first identifies the 

bundle, followed by the rod and the axial height. An extra identifier (W) is included for TCs 

attached to water rods and appears after the bundle designation. 

Top 41" Middle 26" Bottom 11" 

C-WeU 41 

C-eX 41 

C-eY 41 

C-eZ 41 

C-fV 41 

C-fX 41 

C-fY 41 

C-WeU 26 

C-WdV 26 

C-aQ 26 

C-aR 26 

C-aS 26 

C-aT 26 

C-aU 26 

C-bR 26 

C-bS 26 

C-WeU 11 

C-aQ 11 

C-bQ 11 

C-bR 11 

C-cQ 11 

C-cR 11 

C-cS 11 

C-dQ 11 

C-dR 11 

C-dS 11 

C-dT 11 

C-eQ 11 

C-eR 11 

C-eS 11 

C-gX 11 

Additional 

Axial 

C-gX 3 

C-gX 19 

C-gX 33 

C-gX 49 

Q 

R 

S 

T 

U 

V 

X 

Y 

Z 

a b c d e f g h i 

Q

R

S

T

U

V

X

Y

Z

a b c d e f g h i 

Q

R

S

T

U

V

X

Y

Z

a b c d e f g h i 

C-bT 26 

C-bU 26 

C-cS 26 

C-cT 26 

C-cU 26 

C-cX 26 

C-dT 26 

C-gX 26 

C-fZ 41 

C-gX 41 

C-gY 41 

C-gZ 41 

C-hY 41 

C-hZ 41 

C-iZ 41 

North 

Figure 6.4 Center assembly TC locations and naming conventions 

Similar to Figure 6.4, Figure 6.5 presents the TCs located in the peripheral assemblies. Again, 

TCs attached to water rods are given the extra identifier (W) after the bundle identification. 
 

 108 

Q

R

S

T

U

V

X

Y

Z

Q

R

S

T

U

V

X

Y

Z

East Top 41" South Middle 26" West Bottom 11" 

E-WeU 41 

E-eQ 41 

E-eS 41 

E-eX 41 

E-eZ 41 

E-fR 41 

E-fV 41 

E-fY 41 

E-gQ 41 

E-gS 41 

E-gU 41 

E-gX 41 

E-gZ 41 

E-hR 41 

E-hT 41 

E-hV 41 

S-WeU 26 

S-WdT 26 

S-aQ 26 

S-aS 26 

S-aU 26 

S-bR 26 

S-bT 26 

S-cQ 26 

S-cS 26 

S-cU 26 

S-dR 26 

S-eQ 26 

S-eS 26 

S-eZ 26 

S-fR 26 

S-fT 26 

S-gQ 26 

S-gS 26 

S-gU 26 

S-gX 26 

S-hR 26 

S-hT 26 

S-iQ 26 

S-iS 26 

S-iU 26 

W-WeU 11 

W-aQ 11 

W-aS 11 

W-aU 11 

W-aX 11 

W-aZ 11 

W-bR 11 

W-bT 11 

W-bV 11 

W-bY 11 

W-cQ 11 

W-cS 11 

W-cU 11 

W-cX 11 

W-cZ 11 

W-dR 11 

E-hY 41 

E-iQ 41 

E-iS 41 

E-iU 41 

E-iX 41 

E-iZ 41 

Extra 

Axial 

E-eZ 3 

E-eZ 11 

E-eZ 19 

E-eZ 26 

E-eZ 33 

E-eZ 49 

W-dT 11 

W-dY 11 

W-eQ 11 

W-eS 11 

W-eX 11 

W-eZ 11 

Extra 

Axial 

W-eZ 3 

W-eZ 19 

W-eZ 26 

W-eZ 33 

W-eZ 41 

W-eZ 49 

Q 

R 

S 

T 

U 

V 

X 

Y 

Z 

a b c d e f g h i a b c d e f g h i a b c d e f g h i 

North 

Figure 6.5 Peripheral assemblies TC locations and naming conventions 

6.2 MELCOR Analysis Methodology 

6.2.1 MELCOR Code 

The MELCOR 1.8.5 severe accident computer code [3], with enhancements through Version RP, 

was used to simulate the experimental SFP accident response. MELCOR is a full nonequilibrium, two-phase thermal-hydraulics code for the simulation of nuclear reactor severe 

accidents. It includes a full spectrum of models for simulation of reactor accidents progressing 

from design conditions to severe accidents with 

1) Comprehensive fuel degradation, 

2) Steam and air oxidation with hydrogen production, 

3) Fuel and structural component degradation, 

4) Fission product release, transport, deposition, chemisorption, vapor condensation and 

evaporation, aerosol hygroscopic effects, iodine pool evolution, and gamma heating, 

5) Core-concrete interactions, and 

6) Simulation of containment engineering safety features (e.g., sprays, fan coolers, 

filters, and recombiners). 

Of particular relevance for application to the SNL experimental SFP tests, MELCOR has new 

fuel geometry models for boiling and PWR SFP rack geometry. The specialized SFP geometry 

models include basic heat transfer models for convection, conduction, and radiation, as well as a 

new breakaway oxidation kinetics model and fuel and rack degradation models. 
 

 109 

Version RP includes three recent modeling enhancements applicable to BWR SFP modeling: 

1) A new rack component, which permits better modeling of a SFP rack, 

2) A new oxidation kinetics model, and 

3) A simplified flow regime model. 

The new BWR SFP rack component permits proper radiative modeling of the SFP rack between 

groups of different assemblies. The new oxidation kinetics predicts the transition to breakaway 

oxidation in air environments on a node-by-node basis. The simplified flow regime model 

permitted simulation of liquid films draining down the BWR fuel assemblies during spray 

operation. 

6.2.2 Short 14 Assembly Model 

A MELCOR short-stack assembly model was developed to analyze the temperature response to 

the SNL experimental SFP testing program. The model simulates a portion of five assemblies in 

a 14 pattern (Figure 6.6). The fully populated rod region at the bottom of the assembly was 

represented in the radiation tests. For implementation into MELCOR, the 14 pattern was 

represented by a single, high-powered assembly in the first ring and four peripheral assemblies in 

a second ring. The second ring was surrounded with 6 in. of KaowoolTM insulation. 

Figure 6.6 Schematic of the 14 configuration 

Figure 6.7 shows the corresponding COR and control volume hydrodynamics (CVH) 

nodalizations. The BWR assembly is represented by 11 axial levels in the COR Package: 

 Level 1 is the tube leading to the inlet of the assembly (blocked). 

 Level 2 is the lower tie plate. 

 Level 3 is the start of the active fuel region. 

1x4 

Configuration

High decay heat 

center assembly

4 low decay heat 

peripheral assemblies

Insulated Insulated 
 

 110 

 Level 10 is the top of the full-length rod active fuel region. 

 Level 11 is the upper tie plate. 

Figure 6.7 MELCOR nodalization of the full-length assembly experiment 

COR 

Nodalization 

Active 

fuel 

height 

3 

11 

4 

5 

6 

9 

7 

8 

10 

1 

2 

Axial 

Level 

Bottom 

tie plate 

Fully populated 

 fuel section 

Spacer

Upper 

tieplate 

Spacer

CV-100 CV-200 

High decay heat 

center assembly 

4 low decay heat 

peripheral assemblies 

Ring 1 

Bypass 

Bundle

CVH 

Nodalization 

Ring 2 

Bypass 

Bundle

CV101 

CV111 

CV201 

CV211 

CV102 

CV112 

CV202 

CV212 

CV103 

CV113 

CV203 

CV213 

CV104 

CV114 

CV204 

CV214 

CV-300 
 

 111 

The upper and lower tie plates are modeled as supporting plate structures made of stainless steel. 

The 0.5-in. rack support plate at Level 2 is also modeled as a stainless steel supporting structure. 

While the model accounts for the thermal capacitance of these support structures, MELCOR 

treats the top and bottom surfaces as adiabatic. The actual apparatus was insulated on the top and 

bottom with 6 in. of KaowoolTM insulation. This difference may account for some end effects 

reflected in the experimental data but not in the model. 

In the heated region two COR cells were placed in each CVH volume. The interstitial area 

between the canister wall and the rack wall is modeled as the BWR bypass region. The bypass 

uses the same four-control volume axial nodalization as the channel region. At the bottom of 

each ring is a control volume that represents the inlet tube connected to the bottom of the 

assembly. However, the flow path representing the inlet tube was closed to represent the plugged 

inlet used in the radiation tests. The exits of the assembly were connected to a time-independent 

volume that represented the atmospheric conditions during the test (i.e., constant pressure and 

temperature). The gas in the assemblies was allowed to enter into CV-300 (an expansion 

plenum) as the assembly heated (i.e. the system did not pressurize just as in the actual apparatus). 

The hydraulic resistance was specified using the results from the SNL hydraulic test 

(Section 4.2.2). However, the gas was stagnant in the MELCOR model, except for the 

volumetric expansion as the gas heated. 

The BWR assembly canister was modeled with the MELCOR canister component. The center 

rack walls were modeled using the new SFP rack component with stainless steel and aluminum 

(i.e., to represent Boral). However, due to limitations in MELCOR to simultaneously represent 

the rack and the insulation, the 12-rack panels and surrounding insulation on the outside of the 

four peripheral assemblies were represented in MELCOR’s Heat Structure (HS) Package. The 

primary limitations of the heat structure rack/insulation model only occur in high temperature 

conditions (i.e., the HS Package does not model oxidation, melting, or relocation). However, the 

peripheral canisters will properly radiate to the heat structure rack/insulation model. The peak 

temperatures in the radiation tests were well below the rack melting temperature. 

MELCOR does not include an option to model the two large water rods in the assembly. 

Consequently, the water rod mass and surface area was included in the canister wall. 

The interiors of the heater rods were modeled as compacted magnesium oxide. MELCOR’s 

gamma heating model was disabled to deposit all the power from the heater rods in the simulated

fuel (i.e., the magnesium oxide). A constant, linear power profile was used in each heater rod. 

The emissivity of the stainless steel rack walls was modeled using the default model in 

MELCOR except where noted. The emissivity was a function of temperature, 

 = 0.042 + 0.000347 T 

The Zircaloy rod and canister emissivities were also modeled using MELCOR’s default model. 

The emissivity was a function of the oxide layer thickness, 
 

 112 

 = 0.325 + 0.1246x106 rox (for rox < 3.8810-6 m) 

 = 0.808642 – 50.0 rox (for rox > 3.8810-6 m) 

MELCOR’s steam emissivity model was effectively disabled because the assembly was filled 

with dry air. 

6.3 COBRA 14 Model 

 No-flow conditions with adiabatic surfaces at the top and bottom of model. 

 Radial heat loss through insulation only. 

 Density allowed to change with temperature. System does not pressurize. 

 Center assembly power 1500 W, peripheral assemblies have no power. 

 Internal flow assumed to be laminar (Nu = 4.36 Bundle, Nu = 5.383). 

 Eight axial levels, 48-in. long 

 Includes 33 rack arrangement with KaowoolTM insulation in the four corner cells and 

surrounding the entire assembly. 

 Temperature dependant properties for material and air. 

Figure 6.8 shows the overall slab layout of the COBRA model. The slab partitions were based on 

the geometry and composition of the actual hardware. 
 

 113 

Figure 6.8 Slab layout for the 14 assembly 

Figure 6.9 gives the detailed nodalization of the individual rod arrays. The yellow rods indicate 

the location of the partial rods in a full-scale assembly. Since COBRA only allows for a regular 

rod array, the water rods were modeled as the interstitial space bounded by the rods highlighted 

with red. 
 

 114 

Figure 6.9 Individual rod array illustration 

6.4 Results 

Table 6.1 is a summary of the tests conducted with the 14 Incoloy assembly. Testing focused on 

the high contrast heating, i.e., heating of only the center bundle. Power was terminated at 

maximum assembly temperatures of ~873 K to prevent melting of the Boral panels, which are 

primarily aluminum. The apparatus was then monitored as the center assembly cooled by 

radiating to the peripheral assemblies. All numerical results represent average bundle 

temperatures. 

Table 6.1 Summary of all 14 thermal radiation testing on the Incoloy short stack assembly 

Date Description Time at Power (hrs) Test Duration (hrs) 

2/27/2006 1.5 kW - Center only 12 12 

3/15/2006 3 kW - Center only 2.2 10.4 

3/21/2006 3 kW - Center only 4.1 6.8 

3/23/2006 5 kW - Center only 2 8.8 

Figure 6.10 shows symmetrically placed peripheral bundle temperatures as a function of time for 

the powered test conducted on March 21, 2006. Temperatures were within 10°C for all but the 

hottest location at the 41-in. level. Here, the maximum difference was 15°C. 
 

 115 

0

50

100

150

200

250

012345678

Time (hrs)

Temperature (oC)

E-EZ 11

W-EQ 11

W-EZ 26

S-IU 26

E-EZ 26

S-AU 26

W-EZ 41

E-EQ 41

Figure 6.10 Symmetric peripheral bundle temperatures as a function of time for the powered test 

conducted on 3/21/06 

Figure 6.11 through Figure 6.20 compare the experimental minimum (blue lines), maximum (red 

lines), and average bundle temperatures (black lines) to the predicted average bundle 

temperatures of COBRA (green dashed lines) and MELCOR (pink dashed lines). Discrepancies 

between the codes and experimental data are attributed to additional heat loss to the bottom and 

top surfaces, which is not treated in the numerical models. 
 

 116 

300

400

500

600

700

800

900

0 3 6 9 12

Time (hrs)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.11 Power test 2/27/06 – 1.5 kW center assembly, 26-in. level 

300

400

500

600

700

800

900

0369

Time (hrs)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.12 Power test 3/15/06 – 3 kW center assembly, 11-in. level 
 

 117 

300

400

500

600

700

800

900

0369

Time (hrs)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.13 Power test 3/15/06 – 3 kW center assembly, 26-in. level 

300

400

500

600

700

800

900

0369

Time (hrs)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.14 Power test 3/15/06 – 3 kW center assembly, 41-in. level 
 

 118 

250

350

450

550

650

750

850

950

02468

Time (hrs)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.15 Power test 3/21/06 – 3 kW center assembly, 11-in. level 

250

350

450

550

650

750

850

950

1050

02468

Time (hr)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.16 Power test 3/21/06 – 3 kW center assembly, 26-in. level 
 

 119 

250

350

450

550

650

750

850

950

1050

02468

Time (hr)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.17 Power test 3/21/06 – 3 kW center assembly, 41-in. level 

250

350

450

550

650

750

850

950

1050

0369

Time (hr)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.18 Power test 3/23/06 – 5 kW center assembly, 11-in. level 
 

 120 

250

350

450

550

650

750

850

950

1050

0369

Time (hr)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.19 Power test 3/23/06 – 5 kW center assembly, 26-in. level 

250

350

450

550

650

750

850

950

1050

0369

Time (hr)

Temperature (K)

Experiment: 

Maximum Minimum Average 

MELCOR 

COBRA 

Center 

Peripheral 

Figure 6.20 Power test 3/23/06 – 5 kW center assembly, 41-in. level 

6.5 Model Sensitivity to Emissivity 

Sensitivity to emissivity of the Zircaloy and stainless steel surfaces inside the 14 thermal 

radiation models is presented in the following pages. All reported values correspond to average 
 

 121 

bundle temperatures. Comparisons were made to the data collected during the powered test on 

March 21, 2006, 3 kW input power to the center assembly. The cases were chosen based on the 

following criteria: 

 MELCOR Baseline: εZirc = 0.325, εSS = 0.256 + 3.4710-4 (T-616.5) (where T is in units 

of °F) (Figure 6.21 through Figure 6.23). 

 COBRA Baseline: εZirc = 0.43, εSS = 0.43 (Figure 6.24 through Figure 6.26). 

 Maintain εeff: εZirc = 0.62, εSS = 0.33 (Figure 6.27 through Figure 6.30). 

 Set εZirc = 0.80, hold εSS = 0.43 (Figure 6.31 through Figure 6.33). 

Effective emissivity is defined for a two surface gray enclosure by the following formula: 

-1 1-ε 1-ε Zirc SS 1 ε = ++ eff ε F ε Zirc 12 SS

   

 

Here, F12 is the view factor between the Zircaloy channel box and stainless steel pool cell. 

Assuming F12 ≈ 1, simplifies the expression to the following: 

ε ε Zirc SS ε = eff ε -ε ε +ε Zirc Zirc SS SS

   

 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.21 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 11-in. level 

MELCOR Baseline Exp. 

MELCOR 
 

 122 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.22 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 26-in. level 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.23 MELCOR baseline emissivity for power test 3/21/06 – 3 kW center assembly, 41-in. level 
 

 123 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.24 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 11-in. level 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.25 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 26-in. level 

COBRA Baseline Exp. 

COBRA 

MELCOR 
 

 124 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.26 COBRA baseline emissivity for power test 3/21/06 – 3 kW center assembly, 41-in. level 

0.0

0.1

0.2

0.3

0.4

0.5

0 0.2 0.4 0.6 0.8 1

Zirc

Eff

SS2 = 0.33 0.38 0.43 = 0.33 0.38 0.43

Figure 6.27 Effective two-surface emissivity as a function of the Zircaloy emissivity for stainless steel 

emissivities of 0.33, 0.38, and 0.43 

 Note: The COBRA baseline was set at SS and Zirc = 0.43. 

COBRA 

Baseline 

εeff 
 

 125 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.28 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 11-in. level 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.29 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 26-in. level 

Maintain COBRA εeff Exp. 

COBRA 

MELCOR 
 

 126 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.30 Maintain effective emissivity for power test 3/21/06 – 3 kW center assembly, 41-in. level 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.31 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 for power test 

3/21/06 – 3 kW center assembly, 11-in. level 

Set εZirc = 0.80 

SS = 0.43 

Exp. 

COBRA 

MELCOR 
 

 127 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.32 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 for power test 

3/21/06 – 3 kW center assembly, 26-in. level 

300

400

500

600

700

800

900

1000

012345678

Time (hrs)

Temperature (K)

Figure 6.33 Increase Zircaloy emissivity to 0.80 and set stainless steel emissivity to 0.43 for power test 

3/21/06 – 3 kW center assembly, 41-in. level 
 

 128 

6.5.1 Additional COBRA Model Modifications 

Additional efforts were made to reconcile the differences in the COBRA model to the 

experimental data. In particular, the temperature profile in the peripheral assembly was 

examined. Figure 6.34 shows the evolution of the COBRA model along with two temperature 

profiles from the powered test from March 21, 2006. The temperature profiles were taken along 

the axis of highest temperature gradient. 

40

80

120

160

200

240

280

320

-1 0 1 2 3 4 5 6

x (in)

Temp. (oC)

Experiment

Experiment

COBRA

COBRA - emiss 0.33

COBRA - Revised

COBRA - WR

Peripheral at 26" level

Lateral temp. profile

3 kW @ 4.1 hrs

x

Center Assembly

Temp. Profile along Dashed Line

Figure 6.34 Peripheral temperature profiles at 4.1 hours elapsed test time for the powered test conducted 

on 3/21/06 

The red line shows the original COBRA baseline model. This model agrees well toward the 

hotter side of the assembly but over predicts the temperature at the opposite side. The next model 

iteration raised the Zircaloy emissivity and lowered the stainless steel emissivity to maintain the 

effective two-surface emissivity (green line) as discussed in the previous section. This change in 

emissivity had very little effect on the thermal response of the system. Next, the hydraulic loss of 

the two-rod spacers was added (pink line). This alteration hindered the convective cell which 

forms inside the peripheral assemblies. The result was elevated temperatures toward the side 

facing the center assembly and lower temperatures on the opposite side of the bundle. Finally, 

the model elements constituting the interior of the water rods were configured to prevent any 

axial flow (light blue line). This modification had minimal impact on the thermal response of the 

peripheral bundles. 
 

 129 

7 FULL-SCALE ZIRCALOY PRE-IGNITION AND IGNITION RESULTS 

OF THE SFP EXPERIMENT 

This report summarizes the findings of the SFP full-scale, pre-ignition and ignition experiments 

conducted between May 23 and June 8, 2006. The stated purpose of these investigations was to 

determine the thermal and hydraulic response for the validation of the MELCOR severe accident 

analysis code. The full-scale Zircaloy assembly was configured for the open bypass and closed 

drain holes flow arrangement. Pre-ignition testing was conducted to verify the thermal and 

hydraulic response of the assembly against previous heated experiments. These tests were 

conducted at power inputs of 1800 to 3000 W. The ignition test (5000 W) was conducted after 

feedback to MELCOR provided agreement with pre-ignition tests. The MELCOR pre-test 

predictions of the ignition experiment were accurate to within 40 K and 5 minutes of ignition. 

7.1 Apparatus 

The basic configuration of this apparatus was very similar to the full-length Incoloy assembly. 

Figure 7.1 shows the experimental apparatus before testing. A prominent feature of this setup 

was the side-mounted sensors, which were not included in the full-length Incoloy assembly. The 

largest of these instruments were the in situ oxygen sensors mounted at 2-ft. intervals along the 

extent of the assembly. Details of these sensors are provided later. The smallest sensors are TCs 

and light pipes that penetrated through the pool cell wall and monitored the Zircaloy canister 

wall. The type B and S TCs were placed roughly every 6 in. starting at z = 54 in., as measured 

from the top of the bottom tie plate. Silica light pipes were mounted between these TCs, again in 

6-in. increments. Finally, an infrared (IR) spectrometer was positioned at the z = 139.5-in. level 

to observe the Zircaloy channel box through a calcium fluoride window. 
 

 130 

Figure 7.1 Schematic of externally mounted sensors 

The Oxyfire high temperature in situ oxygen sensors were made by Marathon Sensors Inc., 

Cincinnati, Ohio. Figure 7.2 schematically shows the sensor flow paths and ceramic felt gasket 

seal location. The sensing element is at the tip of a hemispherically closed 3/8-in. diameter 

zirconia tube. The sensor exploits the phenomena that at high temperatures (>600°C), oxygen 

can readily diffuse through zirconia. Platinum coatings on the inside and outside surfaces of the 

zirconia tube tip allow the potential of the oxygen diffusion to be measured. The resulting 

potential follows the Nertz equation, 

 V = 0.0215 T ln(Cref/C) 

where 

 V is the potential produced in mV, 

 T is the sensor tip temperature in K, 

 Cref is the oxygen concentration in the reference gas, and 

 C is the measured oxygen concentration. 

Cross section 

above partial 

length rods 

24" 

48"

72" 

96"

119"

144" 

 O2 sensor 

 IR 

Spectrometer

 Light pipe

Thermocouple

0" 

Cross section 

of full bundle
 

 131 

Reference 

 air in 

Reference

air out 

Gas 

sample 

to flowthrough 

O2 monitor 

Small gas 

sample flow 

Zr canister 

wall 

Stainless steel 

pool cell wall 

Ceramic felt 

Pt O2 gasket 

sensor 

tip 

3/8” ZrO2

Ceramic 

sensor tube

Reference 

air purge 

tube 

Alumina 

protection 

tube

Pt/ZrO2 

sensor tip 

Ceramic felt

gasket 

Alumina 

protection 

tube 

1-1/4” NPT 

End 

view 

Figure 7.2 In situ oxygen sensor seal and flow schematic 

An integral component of the sensor is a type B TC located on the outside surface of the zirconia 

tube adjacent to the edge of the platinum coating. This TC provides a measure of the sensor tip 

temperature. 

As shown in Figure 7.2, the zirconia sensor tube is concentrically located inside an alumina 

protection tube. The Oxyfire sensors used were custom fabricated so that the protection tube 

ended 3/16 in. before the end of the sensor tip. This allowed for the sensor tip to protrude 

approximately 1/16 in. inside the Zircaloy channel box through a 3/4-in. diameter hole. The hole 

in the channel box was oversized to ensure that the sensor tip did not contact any metallic parts, 

which would result in shorting of the signal. A ceramic felt gasket was used to seal the gap 

between the end of the protection tube and the outer Zircaloy channel box wall. 

An additional port on the Oxyfire sensor allowed access to the annular gas space between the 

alumina protection tube and the zirconia sensor tube. This port allowed a small gas sample flow 

to be drawn from inside the Zircaloy canister and into an external flow-through type oxygen 

monitor providing a backup oxygen measurement. The oxygen monitors used were Model 65 

made by Advanced Micro Instruments Inc., Huntington Beach, California. Note that the oxygen 

sensor at the 24-in. level was damaged during installation and was abandoned in place. No 

oxygen measurements or samples were taken at this level. 
 

 132 

Figure 7.3 shows a plan view schematic of the Oxyfire sensor installation on the experimental 

assembly. The main support for the sensor is provided by a 1-1/2-in. stainless steel half-coupling 

welded to the pool cell. Large Swagelok fittings and 1-1/2-in. stainless steel tubing were used to 

extend the access port out past the insulation. The final Swagelok ferrule was compressed while 

exerting an axial force on the sensor. This helped compress the ceramic felt gasket between the 

end of the alumina protection tube and the Zircaloy canister wall. 

Insulation Outer wall 

Inner Zr wall 1-1/2” Tube and 

Swagelok fittings 

1-1/2” NPT half-coupling 

welded to pool cell 

1-1/2” to 1-1/4” 

Bell reducer 

Figure 7.3 Plan view schematic of oxygen sensor installation 

Figure 7.4 shows the general layout of all internal K-type TCs. The naming convention follows 

the dual-alpha character grid as labeled in the figure. The axial location is given after the rod 

location of the TC. An additional character W is added as a precursor to the TC name for those 

TCs attached to water rods. The TCs along an axial array also have an extra character A at the 

end of the name for distinction. 
 

 133 

Internal thermocouples 

Cross section 

above partial 

z = 0" 

Top of bottom tie 

plate reference 

Bypass holes 

– 2 places 

24" 

48" 

72" 

96" 

119" 

144" 

Radial array 

24" spacing 

11 TC each level 

66 TC total 

Axial array A1 

6" spacing 

20 TCs 

Axial array A2 

12” spacing 

7 TCs 

Water rods inlet and exit 

4 TCs 

Total of 97 TCs 

a b c d e f g h i 

q 

r 

s 

t 

u 

v 

x

y 

z 

24" and 96" 

levels 

48" and 119" levels 

a b c d e f g h i 

q 

r 

s 

t 

u 

v 

x

y 

z 

(a) (b) 

72" and 144" levels 

a b c d e f g h i 

q 

r 

s 

t 

u 

v 

x

y 

z 

Figure 7.4 Experimental mock fuel assembly showing internal, as-built (a) axial and (b) lateral TC 

locations 
 

 134 

Figure 7.5 gives the locations of the two hot wire anemometers placed inside the inlet of the 

assembly. These hot wires measure the induced flow rate of air into the apparatus. The unheated 

flow calibrations of these hot wires are shown in Figure 7.6. Here, the input volumetric flow rate 

into the assembly is plotted as a function of the measured velocities of the hot wires. 

101.6 

241.3

Anemometer 

ports – 2 places 

A-A 

Section A-A 

25.4

Hot wire 2 

Hot wire 4 

Fixed 

location

25.4

6.4 

All dimensions in mm 

89 ID 

(3.5") 

Figure 7.5 Schematic showing the layout of the assembly inlet hot wires 2 and 4 

y = 365.56x

R2

 = 0.9966

y = 339x

R2

 = 0.9984

0

40

80

120

160

200

0 0.1 0.2 0.3 0.4 0.5 0.6

V (m/s)

Q (slpm)

Hot Wire 2

Hot Wire 4

Figure 7.6 Inlet hot wire calibrations for hot wire 2 (diamonds) and hot wire 4 (squares) 
 

 135 

7.2 Pre-Ignition Testing and Analysis 

7.2.1 Comparison of Thermal Results 

Figure 7.7 gives the thermal response of the test assembly as a function of time for the preignition test with a power input of 2250 W. As with the previous Incoloy full-length testing, the 

assembly reached steady state near the bottom first and continued to heat at the top. The location 

of the hottest portion of the assembly also moved upward during heat-up. 

27

127

227

327

427

527

627

0 3 6 9 12

Time (hrs)

Temperature (K)

CS18-1A

CS36-1A

CS61-2A

GX78-3A

CS96-1

CS103-1A

CS119-2

CS132-2A

GX144-3

GX156-3A

300

400

500

600

700

800

900

Figure 7.7 Temperature as a function of time for the pre-ignition test with a power input of 2250 W 

Figure 7.8 shows the initial MELCOR model results for the same test as depicted in Figure 7.7. 

The MELCOR model was within 28 K at all times during the test. However, the predicted 

temperatures were consistently higher than those in the experiment were. This disparity was most 

likely due to the initial MELCOR model under predicting the induced flow rate, which is 

discussed next. 
 

 136 

300

400

500

600

700

800

900

0 3 6 9 12

Time (hr)

Temperature (K)

COR-TCL.104

COR-TCL.105

COR-TCL.106

COR-TCL.107

COR-TCL.108

COR-TCL.109

COR-TCL.110

COR-TCL.111

COR-TCL.112

COR-TCL.113

Figure 7.8 MELCOR predictions for temperature as a function of time for the pre-ignition test with a 

power input of 2250 W 

7.2.2 Comparison of Air Flow Results 

The induced flow rate into the assembly and bypass pressure drop as a function of time are 

shown in Figure 7.9 for the powered test of 2250 W. Here, the MELCOR model under predicts 

the flow rate by up to 13 slpm. The flow resistances in the model were based on the previous 

Incoloy experiments. However, the penetrations of the oxygen sensors into the bundle of the 

apparatus (Figure 7.2 and Figure 7.3) appeared to provide additional flow paths from the bundle 

to the annulus, despite being sealed with alumina felt. Attempts were made to resolve this issue 

on the apparatus and within MELCOR and are described later in this section. 
 

 137 

0

50

100

150

200

0 3 6 9 12

Time (hrs)

Q (slpm)

0

5

10

15

20



PBypass (N/m2

)

HW2

HW4

MELCOR

Pressure 1

Figure 7.9 Flow rate and pressure drop across the bypass as a function of time for hot wire 2 (HW2) 

(blue), hot wire 4 (HW4) (red), MELCOR (green), and pressure gage (black) 

Initial attempts to model the hydraulic behavior of the Zircaloy apparatus under predicted the 

inlet flow rates. Prior to the final ignition test, the trends from the pre-ignition testing showed a 

decreasing assembly flow rate with increasing assembly power. This was also shown in the 

Incoloy assembly but perhaps exaggerated in the MELCOR calculations. As shown in Figure 

7.10, the calculated trends from MELCOR using the best-estimate flow resistance trended with 

the original Incoloy correlation but were well below the new Zircaloy assembly data. 
 

 138 

Power versus Inlet Flow Rate

y = -0.0046x + 137

R2

 = 0.6627

y = -0.0167x + 147.96

R2 = 0.9973

0

20

40

60

80

100

120

140

0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000

Power (W)

Inlet Flow Rate (slpm)

Experiment - Zr Assembly

Experiment - Incoloy Assembly

MELCOR - Zr Assembly

MELCOR - Incoloy Assembly

MELCOR + Additional Leakage

Zr Assembly Trendlines 

to 5 kW

O2 sensor 

l k

Figure 7.10 Comparison of pre-ignition MELCOR predictions and experimentally measured assembly 

inlet flow rates as a function of bundle power 

The higher flow rate was attributed to leakage around the oxygen sensors. During pre-test repairs 

of the oxygen sensors, pressure drop measurements indicated leakages around the sensor 

penetrations through the canister. Recall that the oxygen sensors were placed at 2 ft., 4 ft., 6 ft., 8 

ft., 10 ft., and 12 ft. Following the 3-kW test, the MELCOR model was adjusted to match the 

higher measured flow in the 3-kW test (see red triangle at 3000 W in Figure 7.10). Rather than 

introduce small leaks at each sensor location, the bypass leakage through the nosepiece was 

increased until the overall leakage matched the 3-kW test. This enhanced (bypass) leakage (to 

the annulus) model was used for the 5-kW pre-test calculation. 

7.3 Ignition Test and Analysis 

7.3.1 Thermal Data 

The thermal response of the apparatus at different axial locations as a function of time is shown 

in Figure 7.11. The assembly first reached ignition between the 96- and 103-in. levels at ~7.3 

hours. The burn front then propagated downward through the assembly. Noise in the temperature 

readings above ~1000°C indicates failure of the type K TCs. 
 

 139 

0

200

400

600

800

1000

1200

1400

02468

Time (hrs)

Temp. (oC)

CS6-1A CS12-1A CS18-1A

CS24-1 CS30-1A CS36-1A

CS42-2A CS48-2 CS54-2A

CS61-2A GX66-3A GX72-3

GX78-3A GX84-3A CS90-1A

CS96-1 CS103-1A

Figure 7.11 Temperature as a function of time at different axial levels during the ignition test 

7.3.2 Oxygen Data 

Two independent oxygen measurements were made at each of the 2-ft. intervals from the 4-ft. 

elevation to the 12-ft. elevation. These measurements were made by an in situ sensor located just 

inside the bundle region and by a gas sample drawn into a flow-through oxygen monitor. Details 

are provided previously in the Apparatus section. Two additional gas samples were drawn from 

the top of the assembly with the oxygen content measured by two flow-through type monitors. 

One gas sample was taken from the bundle region and the other was taken from the annular 

region. 

Figure 7.12 shows the sensor signal and temperature response for the sensor located at the 144-

in. level. The sensor response is not valid for determining the oxygen concentration until the 

sensor temperature is greater than 600°C. Figure 7.13 compares the oxygen concentration at the 

144-in. level determined from the in situ sensor signal and temperature measurement, via the 

Nertz equation, with the corresponding response from the flow-through oxygen monitor. Both 

measurements closely agree on the timing of the precipitous drop in oxygen at 7.42 hours that is 

indicative of Zirconium combustion at or below this level. The close agreement on the ignition 

timing was also evident at each of the lower levels analyzed. After oxygen concentration 

dropped at the 144-in. level, the in-situ sensor consistently read near zero oxygen concentration 

while the flow-through monitor registered about 2% oxygen. This is attributed to small leaks of 

air into the gas sample stream drawn into the flow-through analyzer. The 210-mV signal 

generated by the in situ sensor can only be a result of oxygen molecules diffusing through the 
 

 140 

ceramic matrix of the sensing element. Therefore, false positive measurements are improbable 

leading the sensor manufacturer to state in the manual, “...basic sensor theory negates the 

possibility that the sensor will give an incorrect, low reading.” 

-150

-100

-50

0

50

100

150

200

250

0 2 4 6 8 10 12

Time (hr)

signal (mV)

0

100

200

300

400

500

600

700

800

Temperature ( C)

o

Oxygen 144" < 600 C

Oxygen 144" > 600 C

Oxygen Temp 144

Figure 7.12 Signal and temperature response for the oxygen sensor located at z = 144 in. 
 

 141 

0

5

10

15

20

25

0 2 4 6 8 10 12

Time (hr)

O2 Concentration (%)

144" Nertz O2 < 600C

144" Nertz O2 > 600C

144" O2 Monitor

Figure 7.13 Comparison of the measured oxygen concentrations measured by the in situ oxygen sensor 

and the flow-through oxygen monitor 

Figure 7.14 shows the response for all of the flow-through oxygen monitors. Table 7.1 gives the 

time at which an oxygen concentration of 15% was measured by each of the oxygen sensors and 

monitors. The oxygen measurement at the top bundle indicates that ignition occurred at 7.28 

hours. 
 

 142 

Oxygen Monitors

0

5

10

15

20

25

0369

Time (hr)

O2 Concentration (%)

O2 Top bundle

O2 144

O2 119

O2 96

O2 72

O2 48

Figure 7.14 Measured oxygen concentration from the monitor devices 

Table 7.1 Elapsed experimental time to reach 15% oxygen concentration for the oxygen sensors and 

monitors 

 Location 

O2 Sensor time 

to 15% O2 (hr) 

O2 Monitor time 

to 15% O2 (hr) 

Average time to 

15% O2 (hr) 

Apparent burn 

rate (in/hr) 

Top bundle - 7.277 7.277 - 

Top annulus - 7.419 7.419 - 

144 in 7.418 7.425 7.422 - 

119 7.436 7.427 7.432 - 

96 7.397 7.387 7.392 80.11 

72 7.684 7.699 7.692 51.15 

48 8.143 8.178 8.161 - 

- - - average 65.63 

7.3.3 Comparison of Air Flow Results 

The results of pre-test calculations with and without the added oxygen sensor leakage are shown 

in Figure 7.15. The green line shows the best-estimate total induced flow rate based on the 

previous full-length Incoloy experiments and SNL standards laboratory flow/pressure drop 

characterization of the nosepiece bypass holes. The black line shows the pre-test calculation at 5 

kW with the sensor leakage added to the bypass holes, which did not extrapolate well from 3 kW 

to 5 kW. Attempts to repair the oxygen sensor seals were made between the pre-ignition testing 

and the ignition test and may have increased the leakage. 
 

 143 

Both MELCOR models show a sharp downturn in flow rate at the time of ignition. This decrease 

is due to the rapid heating and expansion of the air in the assembly, producing a choking effect. 

The test data indicate an increase in the inlet flow rate immediately following ignition. This 

behavior suggests that the leakage into the annulus increased at this time. As the stainless steel 

pool cell melted away at the point of ignition, the pool rack and insulation slumped downward. 

This event pulled the oxygen sensors down as well, possibly corrupting the seal at the sensor tip. 

Attempts to refine the MELCOR model to more accurately portray the sensor leakages are 

discussed next. 

Assembly Flow Rate

Pre-test Calculation of 5 kW Full-length Zr Assembly Ignition Test

0

30

60

90

120

150

180

0369

Time (hr)

Flow Rate (slpm)

HW2 slpm 

HW4 slpm

MELCOR + Sensor Leakage

MELCOR, BE Leakage Higher annulus leakage tuned 

from 3 kW per-ignition test Best estimate 

Added 

Figure 7.15 Comparison of MELCOR pre-test prediction with and without sensor leakage versus the 

measured assembly inlet flow rate 

Individual leakage paths were introduced for the post-test calculation for each oxygen sensor at 

the 2-ft., 4-ft., 6-ft., 8-ft., 10-ft., and 12-ft. elevations. A leakage of 1.5×10-4 m2

/sensor was 

needed to match the measured flow data (Figure 7.16). This represents a ~0.070-in. annular gap 

around each sensor. 
 

 144 

Assembly Flow Rate

0

30

60

90

120

150

180

0369

Time (hr)

Flow Rate (slpm)

HW2 slpm 

 HW4 slpm

MELCOR

Figure 7.16 Comparison of MELCOR post-test prediction with additional sensor leakage versus the 

measured assembly inlet flow rate 

The measured assembly inlet flow rate achieved a steady value at ~4.7 hours. The peak measured 

temperature at 4.7 hours was 900 K. The temperature steadily rose thereafter until 6.9 hours 

when the temperature rise increased slightly near the ignition point and then dramatically at 

~7.3 hours as the breakaway phenomena occurred. Between 4.7 hours to ignition, oxidation 

became increasingly more important. However, the oxygen sensors and the MELCOR 

calculations did not measure a significant reduction in the oxygen concentration until close to 

ignition. Hence, the steady flow rate does not seem to be related to any thermal or chemical 

phenomena. This response remains unresolved, but the increased leakage MELCOR result tracks 

the overall magnitude of the flow reasonably during this period but not the flattening trend 

(Figure 7.16). 

A sudden increase in flow was measured at 7.59 hours. The most probable cause was the 

catastrophic failure of the oxygen sensor seal to the canister at the time of ignition. To address 

this issue, the MELCOR leakage flow area was increased by a factor of 10 at all sensor locations 

at 7.56 hours (results shown in Figure 7.17) to represent this assertion. Consequently, instead of 

decrease in flow rate at the start of ignition (see pre-test result in Figure 7.15), the flow increased 

to almost the value measured in the experiment. Examination of the calculated flows showed the 

additional leakage flow came primarily from the 2-ft. elevation (Figure 7.17). There were smaller 

outflows from the assembly to the annulus at the 4-ft., 6-ft., and 12-ft. elevations. Near the 

ignition front (8 ft. and 10 ft.), the leakage was inward, which fed the oxidation reaction. Hence, 

a larger opening(s) of the oxygen sensor(s) could have caused the sharp increase in flow rate. 

However, the MELCOR model suggests that the most dramatic impact is near the bottom of the 

assembly (especially, the 2-ft elevation). At 7.59 hours, the measured ignition front was between 

the 96-in. and 72-in. elevations, which are significantly above the bottom of the assembly. 

Consequently, a few residual questions remain: 
 

 145 

 Through numerous sensitivity calculations, replication of the magnitude or duration of 

the high flow period was not duplicated. The flow rate increased to >180 slpm from 7.7 

to 8 hours. This increase could be the result of many factors, but apparently, there was 

little to no blockage. 

 Based on the MELCOR calculations, the most significant impact to the inlet flow rate 

comes from the lowest sensor. The failed oxygen sensor seal was most likely near the 

ignition front (72 in. to 96 in.). However, the postulated loss of oxygen sensor seals 

seems reasonable given the observed slumping of the pool cell and insulation as 

discussed earlier in this section. 

Annulus Leakage Flow

-0.0004

-0.0002

0

0.0002

0.0004

0.0006

0.0008

0.001

0.0012

0.0014

0.0016

0369

Time (hr)

Mass Flow (kg/s)

2-ft

4-ft

6-ft

8-ft

10-ft

12-ft

Figure 7.17 Calculated leakage flow from the post-test enhanced leakage model 

7.3.4 Comparison of Thermal Results 

The MELCOR model predicted the same burn rate as in the experiment. The measured ignition 

front started at ~96 in. at 7.3 hours and moved to 42 in. by 8.3 hours, or 54-in/hr. The MELCOR 

ignition front started calculation in the 117-in. to 131-in. cells at 7.5 hours and moved to the 77-

in. to 96-in. cells by 8.5 hours, or 54-in/hr. Here the entire length of the nodes is considered to be 

burning; thus, the MELCOR model burned between the 131-in. and 77-in. levels in the stated 

timeframe. 

 Here much smaller nodes in MELCOR would have better tracked the ignition front. 

Furthermore, more energy is required to ignite a 19-in. node versus a continuously 

downward moving front. 

The measured ignition front started at ~96 in. at 7.3 hours versus 117 in. to 131 in. at 7.5 hours in 

the MELCOR calculation. Possible reasons for this difference are presented next. 
 

 146 

 The initial breakaway oxidation lifetime value cannot be input in MELCOR. When the 

assembly flow rate was increased, the ignition timing increased from 7.3 to 7.5 hours. 

There was a small oxidation layer (i.e., estimated maximum of 15 μm at the top of the 

assembly). Consequently, some pre-oxidation effects could have effected the location of 

ignition. 

 As shown in Figure 7.18 and Figure 7.19, the calculated and measured pre-ignition 

temperature responses were relatively close across a broad axial height. Hence, subtle 

local factors may have influenced the exact ignition location. Again, the MELCOR 

breakaway correlation seems to accurately predict the physics governing the experiment. 

In the final comparison calculations, an offset was added to the lifetime function to move 

ignition from 7.5 hours to 7.3 hours. 

 Sensitivity calculations were performed. When the initial oxide layer above the partial 

rods was increased to 24 μm, the difference between the temperatures above and below 

the top of the partial rods was lessened (i.e., <20 K at the time of ignition). However, the 

117-in. to 131-in. nodes still transitioned to breakaway first. 

 In the MELCOR model, the magnesium oxide and cladding in the unheated plenum at the 

top of the partial rods is accurately modeled. However, it is not possible to only power 

the full-length rods and not the partial rods in this region. For a sensitivity calculation, the 

power distribution was changed to also heat the plenums above the partial rods. The 

results did not change the location of ignition. 

Axial Cladding Temperature Profile

300

500

700

900

1100

1300

1500

1700

1900

2100

2300

0123456789

Time (hr)

Temperature (K)

TCL 109 (96 - 102 in)

TCL 110 (102 - 117 in)

TCL 111 (117 - 131 in)

MELCOR ignited at 117"-131" first 

but 96"-102" was not far behind.

Figure 7.18 Calculated temperature response from the posttest enhanced leakage model 
 

 147 

Median Axial Temperature Data from Full-Length Zr Assembly Test

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

24 in - Median

48 in - Median

72 in - Median

96 in - Median

119 in - Median

144 in - Median

The test ignitioned between 96" to 

102" but 119" was also close to 

ignition. i iti

Figure 7.19 Measured median temperature response 

Due to the complex nature of breakaway and perhaps stochastic or other factors, the location of 

ignition is not accurately predicted by MELCOR. The role of nitrogen appears critical to the 

onset of breakaway oxidation in which nitrogen actually enhances the magnitude of the oxidation 

rate versus pure oxygen. SEM of frozen cladding samples just prior to the onset of breakaway 

shows a very erratic and complex grain matrix that has been disrupted by the nitrogen. 

Finally, the assembly heat balance in MELCOR for this test/configuration leads to higher 

temperatures than in the test near the top portion of the apparatus. Hence, ignition in the model at 

117 to 131 in. was inherent of the range of sensitivities explored. However, in real SFP 

assemblies there will be an axial power profile that will skew the power toward the center of the 

assembly. The model cannot capture the nuances of the test apparatus such as the heat balance 

near the top of the partial rods, scratches in rods near the grid spacer, or some localized power 

variances that led to ignition between 96 in. and 102 in. 

7.4 MELCOR Post-test Comparison Summary 

Changes from the pre-test prediction are summarized by the following statements: 

 The leakage rate was increased between the canister and the annulus to the equivalent of 

1.5×10-4 m2

/sensor. 

 A factor of 10× further increase in leakage rates was assumed at all sensor locations at 

7.59 hours. 

 A small adjustment to the breakaway lifetime correlation was applied to move ignition 

from 7.5 hours to 7.3 hours. The inlet flow rate as a function of time is plotted in Figure 

7.20 for both the experiment and the final post-test MELCOR model, which reflects the 

changes outlined earlier. 
 

 148 

Assembly Flow Rate

0

30

60

90

120

150

180

0369

Time (hr)

Flow Rate (slpm)

HW2 slpm 

 HW4 slpm

MELCOR

Figure 7.20 Flow rate as a function of time for the experiment hot wires and the final post-test MELCOR 

model 

The following are some various additional observations: 

 Calculated power failure occurred at PCT = 1445 K and 7.31 hours (almost identical to 

measured failure at 1444 K and 7.31 hours). The maximum temperature along the 

assembly is shown in Figure 7.21 for the experiment and the final post-test MELCOR 

model. 

 Interestingly, the heaters failed at 7.31 hours at a PCT of 1444 K (almost the identical 

point as Heater Design Test 2 and the best-estimate value for the full-length Zircaloy 

ignition test). 

 Calculations were terminated at 9 hours. In the experiment, the argon purge was started at 

8.7 hours. The assembly inlet flow damper was closed a few minutes before. 

 MELCOR was primarily burning downward at cell 108, but cell 111 (117 in. to 131 in.) 

was also slowly burning at the end of the calculation (Figure 7.22). 

 As discussed earlier, the ignition front started lower in the experiment. Hence, no ignition 

occurred in the calculation at the lower levels (see Figure 7.23 and Figure 7.24). The 58-

in. to 77-in. cells (107) were rapidly approaching ignition at the end of the calculation 

(Figure 7.25). 

 The pre-ignition heat-up rates at the 48-in. to 119-in. elevations were in good agreement 

with the data (see Figure 7.24 through Figure 7.27). The 24-in. data is also well bounded 

by the two MELCOR cells around this location. 
 

 149 

 The pre-ignition temperatures at 144 in. were below the calculation (Figure 7.28). The 

heat loss appears to have been under estimated in the calculation. 

Maximum Cladding Temperature

300

500

700

900

1100

1300

1500

1700

1900

2100

2300

0 2 4 6 8 10 12

Time (hr)

Temperature (C)

Data Max

MELCOR

Figure 7.21 PCT for both the ignition experiment and the final post-test MELCOR model 

Breakaway Lifetime Function

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0123456789

Time (hr)

Lifetime (-)

Level 104 (0 - 19 in)

Level 105 (19 - 38 in)

Level 106 (38 - 58 in)

Level 107 (58 - 77 in)

Level 108 (77- 96 in)

Level 109 (96 - 102 in)

Level 110 (102 - 117 in)

Level 111 (117 - 131 in)

Level 112 (131 - 145 in)

Level 113 (145 - 156 in)

Figure 7.22 Breakaway lifetime function as a variable of time in the MELCOR ignition model Temperature (K) 
 

 150 

Axial Cladding Temperature Profile

24" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

24 in - Max

24 in - Median

24 in - Min

TCL 105 (19 - 38 in)

TCL 104 (0 - 19 in)

MELCOR

MELCOR

Figure 7.23 Cladding temperatures at z = 24 in. for both experiment and MELCOR 

Axial Cladding Temperature Profile

48" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

48 in - Max

48 in - Median

48 in - Min

TCL 106 (38 - 58 in) MELCOR

Figure 7.24 Cladding temperatures at z = 48 in. for both experiment and MELCOR 
 

 151 

Axial Cladding Temperature Profile

72" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

72 in - Max

72 in - Median

72 in - Min

TCL 107 (58 - 77 in) MELCOR

Figure 7.25 Cladding temperatures at z = 72 in. for both experiment and MELCOR 

Axial Cladding Temperature Profile

96" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

96 in - Max

96 in - Median

96 in - Min

TCL 109 (96 - 102 in)

TCL 108 (77- 96 in) MELCOR

MELCOR

Figure 7.26 Cladding temperatures at z = 96 in. for both experiment and MELCOR 
 

 152 

Axial Cladding Temperature Profile

119" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

119 in - Max

119 in - Median

119 in - Min

TCL 111 (117 - 131 in)

TCL 110 (102 - 117 in)

MELCOR

MELCOR

Figure 7.27 Cladding temperatures at z = 119 in. for both experiment and MELCOR 

Axial Cladding Temperature Profile

144" Level

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

144 in - Max

144 in - Median

144 in - Min

TCL 112 (131 - 145 in) MELCOR

Figure 7.28 Cladding temperatures at z = 144 in. for both experiment and MELCOR 
 

 153 

7.5 Discussion 

The initial pre-test MELCOR model predicted the PCT of the assembly to within 40 K at all 

times. The model also predicted the time of ignition to within 5 minutes. However, the model 

under predicted the assembly inlet flow rate by up to 20 slpm. This difference was likely due to 

leakage around the ceramic felt used to seal the oxygen sensors to the channel can. 

The post-test MELCOR model accounted for the leakages as flow areas of 1.5×10-4 m2

/sensor 

(or ~0.070-in. annular gap around the seal). These leakage areas were increased to 1.5×10-3

m2

/sensor at ignition in order to duplicate the probable catastrophic failure of the sensor seals. In 

addition, a small adjustment was made to the breakaway lifetime correlation to move the ignition 

time to the experimentally observed 7.3 hours. 

Finally, Figure 7.29 shows photos of the assembly after ignition. The top left shows the top of 

the assembly. The pool cell (dark gray) has slumped downward from its initial height, which 

matched that of the channel can (whitish gray). The bottom left image is of the top 4 ft. of the 

assembly as viewed from below. The damage to the pool cell is clearly evident. The two photos 

on the right show the internal damage to the rod bundle in the vicinity of the initial ignition. 

Figure 7.29 Postmortem of the assembly after ignition 

 

 155 

8 1×4 ZIRCALOY PRE-IGNITION AND IGNITION RESULTS OF THE 

SFP EXPERIMENT 

This report summarizes the findings of the SFP heated radial experiments conducted between 

August 1 and August 31, 2006. The stated purpose of these investigations was to validate the 

MELCOR severe accident analysis code. The apparatus consisted of five shortened assemblies, 

recreating 1.22-m (4-ft.) sections of a SFP in a 1×4 arrangement. For this testing, only the center 

assembly was electrically heated up to 4.87 kW. Air flow rates and inlet temperatures were 

controlled at the inlets of the center and peripheral assemblies separately. 

8.1 Assembly Design 

The 14 Zircaloy assembly was designed to study the potential propagation of a cladding fire 

from a centrally located hot bundle to its cooler neighbors. The assembly segments were located 

in an analogously shortened but otherwise highly prototypic 3×3 pool rack fabricated by Holtec 

Inc. This piece of hardware included prototypic Boral neutron absorption plates between cells. 

The Boral plates consisted of boron impregnated aluminum. Air flow and temperature were 

prescribed using metered flow rates into several air heaters. The entire assembly height was 

approximately 52 in. with a heated length of 48 in. The apparatus simulated the top 48 in. of the 

fully-populated bundle in a full-length assembly . The simulation of this bundle region was 

chosen based on the fact the full length Zircaloy experiment ignited at the top of the fully 

populated bundle region and burned downward. 

8.1.1 General Construction 

Figure 8.1 shows different views of the experimental apparatus. An exploded schematic is 

pictured on the left with details of the center assembly construction highlighted. The actual 

center assembly is shown under construction in the upper right photo. A completed peripheral 

assembly stands behind the center assembly. The lower photo depicts the flow arrangement 

underneath the apparatus. Copper tubing supplied the air ovens (MHI Inc., Model LTA750), 

which then delivered heated air to the apparatus. 

Details of the flow arrangements are provided in Figure 8.2. Two flow paths are formed in each 

assembly, the bundle and the annular region formed by the pool rack and the channel box. Due to 

the higher flow rate in the center assembly bundle, two air ovens were employed to provide the 

heated air. Flow rates were controlled by a series of mass flow controllers (MKS Instruments 

Inc. Model 1559A). The air to the four peripheral annuluses was fed by two air ovens. After 

these ovens, the flow of each air heater was split into half with each stream entering a single 

peripheral annulus. 

Photographs of the completed apparatus are shown in Figure 8.3. The apparatus was surrounded 

by 6 in. of Fiberfrax Durablanket® type S insulation with stainless steel shim at the 3- and 6-in. 

insulation levels serving as radiation barriers (Figure 8.3a). The corner pool cells of the rack 

were also filled with insulation. Finally, the air lines going from the air ovens to the apparatus 

were also insulated (Figure 8.3b). 
 

 156 

Figure 8.1 Left: Expanded schematic of the apparatus construction. Upper right: Photo of the center 

assembly under construction. Lower right: Photo of the undercarriage of the apparatus 

Center 

assembly 

Inlet flow 

box 

Air heaters 

3×3 Holtec 

pool cell rack 
 

 157 

Figure 8.2 Flow diagram of mass flow controllers, air heaters, and assemblies AH 1A 

East South Center North West AH 6 AH 7 AH 8 AH 2 AH 3 AH 4 AH 5 

MF 6

MF 7

MF 8 

MF 1

MF 2

MF 3 

MF 4

MF 5

Air 

Manifold 

in 

Mass Flow 

Controller Air Heater(s) Destination

1 1A and 1B Center bundle

2 2 Center annulus

3 3 North bundle

4 4

North and West 

annulus

5 5 West bundle

6 6 East bundle

7 7

South and East 

annulus

8 8 South bundle

Bundle

Annulus

AH 1B 

North

West

South 

East

Bundle

Flow cross section
 

 158 

Figure 8.3 Photographs of the final assembly with insulation applied around (a) the perimeter and 

(b) the lower portion 

 Note: The corner cells were filled with insulation. 

8.1.2 TC Layout 

The type K TC (73 K to 1523 K temperature range) configuration for the ignition test was based 

on the 14 Incoloy testing experiences (Section 6.1.1). Figure 4.14 depicts the general TC layout 

of the 14 Zircaloy assembly. Three main TC levels were placed at heights of 10.6, 25.6, and 

40.6 in. as measured from the top of the bottom tie plate. The solid circles indicate heater rods 

with a TC attached. 

Figure 8.5 provides the location and naming convention of the TCs in the center assembly. A list 

of all installed TCs is presented below the figure. The naming convention first identifies the 

bundle, followed by the rod and the axial height. An extra identifier (W) is included for TCs 

attached to water rods and appears after the bundle designation. 
 

 159 

Figure 8.4 Overall type K (73K to 1523K temperature range) TC layout 

Figure 8.5 Center assembly TC locations and naming conventions 

Similar to Figure 8.5, Figure 8.6 presents the TCs located in the peripheral assemblies. Again, 

TCs attached to water rods are given the extra identifier (W) after the bundle identification. 

In addition to the type K TCs, seven type B TCs (500K to 1923K) were used to monitor the 

temperature of the gas exiting the top of each assembly. The tip of each TC was positioned just 

Top 40.6" Middle 25.6" Bottom 10.6" 

C-WeU 40.6 

C-eX 40.6 

C-eY 40.6 

C-eZ 40.6 

C-fV 40.6 

C-fX 40.6 

C-fY 40.6 

C-WeU 25.6 

C-WdV 25.6 

C-aQ 25.6 

C-aR 25.6 

C-aS 25.6 

C-aT 25.6 

C-aU 25.6 

C-bR 25.6 

C-bS 25.6 

C-WeU 10.6 

C-aQ 10.6 

C-bQ 10.6 

C-bR 10.6 

C-cQ 10.6 

C-cR 10.6 

C-cS 10.6 

C-dQ 10.6 

C-dR 10.6 

C-dS 10.6 

C-dT 10.6 

C-eQ 10.6 

C-eR 10.6 

C-eS 10.6 

C-gX 10.6 

Additional 

Axial 

C-gX 1.6 

C-gX 4.6 

C-gX 7.6 

C-gX 13.6 

C-gX 16.6 

C-gX 19.6 

C-gX 22.6 

C-gX 28.6 

C-gX 31.6 

C-gX 34.6 

C-gX 37.6 

C-gX 43.6 

C-gX 46.6 

C-gX 49.6

Q 

R 

S 

T 

U 

V 

X 

Y 

Z 

a b c d e f g h i 

Q

R

S

T

U

V

X

Y

Z

a b c d e f g h i 

Q

R

S

T

U

V

X

Y

Z

a b c d e f g h i 

C-bT 25.6 

C-bU 25.6 

C-cS 25.6 

C-cT 25.6 

C-cU 25.6 

C-cX 25.6 

C-dT 25.6 

C-gX 25.6 

C-fZ 40.6 

C-gX 40.6 

C-gY 40.6 

C-gZ 40.6 

C-hY 40.6 

C-hZ 40.6 

C-iZ 40.6 

North 

West

East

North

Top level 40.6" Middle level 25.6" Bottom level 10.6" 

South 
 

 160 

below the top tie plate. Each peripheral assembly had one type B TC above the bundle, the 

center assembly had two type B TCs above the bundle and one type B TC was located in the 

center assembly annulus. 

Figure 8.6 TC locations and naming conventions in the peripheral assemblies 

Finally, the locations of the TCs placed on the channel canisters and pool cell walls are identified 

in Figure 8.7. These TCs are labeled as either channel (Cha-) or pool cell (P-), followed by the 

assembly identifier (e.g. C for center, E for east), then the direction of the mounting face, next 

the axial height, and finally the corresponding rod location nearest the TC (if applicable). The 

annular region in Figure 8.7 has been exaggerated in order to show the locations of these TCs 

distinctly. 

Q

R

S

T

U

V

X

Y

Z

East Top 40.6" South Middle 25.6"

E-WeU 40.6 

E-eQ 40.6 

E-eS 40.6 

E-eX 40.6 

E-eZ 40.6 

E-fR 40.6 

E-fV 40.6 

E-fY 40.6 

E-gQ 40.6 

E-gS 40.6 

E-gU 40.6 

E-gX 40.6 

E-gZ 40.6 

E-hR 40.6 

E-hT 40.6 

E-hV 40.6 

S-WeU 25.6 

S-WdT 25.6 

S-aQ 25.6 

S-aS 25.6 

S-aU 25.6 

S-bR 25.6 

S-bT 25.6 

S-cQ 25.6 

S-cS 25.6 

S-cU 25.6 

S-dR 25.6 

S-eQ 25.6 

S-eS 25.6 

S-eZ 25.6 

S-fR 25.6 

S-fT 25.6 

S-gQ 25.6

S-gU 25.6 

S-gX 25.6 

S-hR 25.6 

S-hT 25.6 

S-iQ 25.6 

S-iS 25.6 

S-iU 25.6 

Extra 

Axial 

S-cU 1.6 

S-cU 7.6 

S-cU 13.6 

S-cU 19.6 

S-cU 31.6 

S-cU 37.6 

S-cU 43.6 

S-cU 49.6 

W-WeU 10.6 

W-aQ 10.6 

W-aS 10.6 

W-aU 10.6 

W-aX 10.6 

W-aZ 10.6 

W-bR 10.6 

W-bT 10.6 

W-bV 10.6 

W-bY 10.6 

W-cQ 10.6 

W-cS 10.6 

W-cU 10.6 

W-cX 10.6 

W-cZ 10.6 

W-dR 10.6 

E-hY 40.6 

E-iQ 40.6 

E-iS 40.6 

E-iU 40.6 

E-iX 40.6 

E-iZ 40.6 

Extra 

Axial 

E-eQ 4.6 

E-eQ 10.6 

E-eQ 16.6 

E-eQ 22.6 

E-eQ 28.6 

E-eQ 34.6 

E-eQ 46.6 

W-dY 10.6 

W-eQ 10.6 

W-eS 10.6 

W-eX 10.6 

W-eZ 10.6 

Extra 

Axial 

W-eZ 1.6 

W-eZ 7.6 

W-eZ 13.6 

W-eZ 19.6 

W-eZ 31.6 

W-eZ 37.6 

W-eZ 43.6 

W-eZ 49.6 

Q 

R 

S 

T 

U 

V 

X 

Y 

Z 

a b c d e f g h i a b c d e f g h i 

Q

R

S

T

U

V

X

Y

Z

West Bottom 10.6"

a b c d e f g h i 

North 

Q

R

S

T

U

V

X

Y

Z

North Axial 

a b c d e f g h i 

Extra 

Axial 

N-gU 4.6 

N-gU 10.6 

N-gU 16.6 

N-gU 22.6 

N-gU 28.6 

N-gU 34.6 

N-gU 40.6 

N-gU 46.6 
 

 161 

Figure 8.7 TC locations and naming conventions in the channel canisters and pool cells 

8.1.3 Oxygen Monitors 

Oxygen concentration was monitored at the top of each of the five assemblies. The oxygen 

monitors used were Model 65 made by Advanced Micro Instruments Inc., Huntington Beach, 

California. Figure 8.8 shows the installation of the five monitors. 

Figure 8.9 shows (a) photographic detail of the as-built sample locations and (b) a diagram with 

the sample location in each assembly denoted as a red circle. A small gas sample was drawn 

from just below the top tie plate through a ceramic tube into stainless steel tubing and through 

the oxygen monitor. The sampling locations were centrally located representing a bulk or 

average value. 

Periph. Cell 

 

P-C-W10.6EQ 

P-C-W10.6CQ 

P-C-W10.6AQ 

P-W-E10.6EZ 

P-W-E10.6CZ 

P-W-E10.6AZ 

P-W-S10.6AU 

P-W-W10.6EQ 

P-W-W10.6CQ 

P-W-W10.6AQ 

Corner Cell 

P-SW-N10.6IS 

P-SW-N10.6IU 

P-SW-N10.6IX 

Misc 

PSW corn10.6 

P-N-N 10.6 

West 

East

South North

Middle level 25.6" Bottom level 10.6" 

Channel Can 

Cha CW-10.6-EQ 

Cha CW-10.6-CQ 

Cha CW-10.6-AQ 

Cha WE-10.6-EZ 

Cha WE-10.6-CZ 

Cha WE-10.6-AZ 

Cha WS-10.6-AR 

Cha WS-10.6-AU 

Cha WS-10.6-AX 

Cha WW-10.6-EQ 

Cha WW-10.6-CQ 

Cha WW-10.6-AQ 

Axial array 

ChaNS-4.6-GU 

ChaNS-10.6-GU 

ChaNS-16.6-GU 

Channel Can 

Cha CS-25.6-AQ 

Cha CS-25.6-AS 

Cha CS-25.6-AU 

Cha SN-25.6-IU 

Cha SN-25.6-IS 

Cha SN-25.6-IQ 

Cha SW-25.6-GQ 

Cha SW-25.6-EQ 

Cha SW-25.6-CQ 

Cha SS-25.6-AQ 

Cha SS-25.6-AS 

Cha SS-25.6-AU 

Axial array 

ChaNS-22.6-GU 

ChaNS-28.6-GU 

Channel Can 

Cha CE-40.6-EZ 

Cha CE-40.6-GZ 

Cha CE-40.6-IZ 

Cha EW-40.6-IQ 

Cha EW-40.6-GQ 

Cha EW-40.6-EQ 

Cha EN-40.6-IX 

Cha EN-40.6-IU 

Cha EN-40.6-IS 

Cha EE-40.6-EZ 

Cha EE-40.6-GZ 

Cha EE-40.6-IZ 

Axial array 

ChaNS-34.6-GU 

ChaNS-40.6-GU 

ChaNS-46.6-GU 

Periph. Cell 

P-C-S25.6AU 

P-S-N25.6IU 

P-S-W25.6EQ 

P-S-S25.6AQ 

P-S-S25.6AS 

P-S-S25.6AU 

Corner Cell 

P-SW-E25.6EZ 

Misc 

PSE corn25.6 

Periph. Cell 

 

P-C-E40.6EZ 

P-C-E40.6GZ 

P-C-E40.6IZ 

P-E-W40.6IQ 

P-E-W40.6GQ 

P-E-W40.6EQ 

P-E-N40.6IU 

P-E-E40.6EZ 

P-E-E40.6GZ 

P-E-E40.6IZ 

Corner Cell 

P-NE-S40.6AX 

P-NE-S40.6AU 

P-NE-S40.6AS 

Misc 

PNE corn40.6 

P-N-N 40.6 

Top level 40.6" 
 

 162 

Figure 8.8 AMI Inc. Model 65 oxygen monitors 

Figure 8.9 Oxygen sampling locations for each assembly as shown in (a) the as-built apparatus and (b) a 

schematic with sampling tube penetrations indicated in red 
 

 163 

8.2 Ignition Test Results 

8.2.1 Input Parameters 

The air flow rate, temperature, and oxygen concentration input parameters for this experiment 

were chosen to mimic flow conditions entering the top 48 in. of the fully populated sections of 

five full-length assemblies arranged in a 1×4 pattern. This experimental representation does not 

directly simulate the lower 52 in. of fully populated sections or the upper 58 in. of partially 

populated sections of a full length assemblies. The center assembly represented a 17.5-day-old 

assembly (11.8 kW) and the four peripheral assemblies represented the average age of 

assemblies in the pool (0.2 kW). As shown in Figure 8.10, the center assembly was powered at 

4.87 kW which represents the desired fully populated section of a full-length 11.8 kW assembly 

with a 1.2 peaking factor. For experimental simplicity, the rod heaters in the four peripheral 

assemblies were not powered. The representative power for each peripheral assembly was only 

0.08 kW, which was lower than the capability of the power controller and therefore considered 

insignificant.

The magnitude of the other dynamically controlled input parameters were chosen based on 

previous experimental results (i.e., full-length Incoloy and Zircaloy assemblies) and MELCOR 

modeling with some additional experimental constraints. The axial oxygen concentration results 

from the full-length Zircaloy assembly ignition test (see Section 7.3.2) clearly showed that no 

oxygen depletion occurred 24 in. below the burn front meaning that atmospheric air could be 

used as the feed gas. As the burn front approached the bottom tie plate of the apparatus, the 

assumption of no oxygen depletion begins to deviate from the prototypic situation since a full 

length assembly would extend 52 in. below the test apparatus.

Figure 8.11 shows the flow rates for the different channels inside the apparatus imposed during 

the ignition test. Figure 8.12 shows the inlet temperatures for the different channels inside the 

apparatus during the ignition test. The flows were precisely controlled to the desired levels by 

programming the mass flow controllers with the LabView data acquisition program. At the time 

of ignition in the center assembly at about 4.9 hours, the test flow program was manually 

interrupted to begin a programmed ramp down in flow rate. The inlet temperatures were 

controlled more coarsely. The temperature controllers used only allowed for a prescribed linear 

ramp to a set temperature. The elevated temperatures shown in Figure 8.12 after 6 hours were 

due to the burn front reaching the bottom tie plate and influencing the inlet air temperature 

measurements. 
 

 164 

0

1000

2000

3000

4000

5000

6000

012345678

Time (hrs)

Power (W)

Figure 8.10 Center assembly input power for the ignition test 

Note: Average power during the heated time was 4.87 kW. 

0

20

40

60

80

100

120

012345678

Time (hrs)

Flow Rate (slpm)

FlowCentBund FlowCentAnul FlowNorthBund FlowSouthbund

FlowEastBund FlowWestBund Flow N&W Anul Flow S&E Anul

2x Peripheral 

annuluses 

Peripheral 

bundles 

Figure 8.11 Air flow rates for the different channels inside the apparatus during the ignition test 

Note: the flow into the peripheral annuluses was split into two after passing through the air oven. 
 

 165 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

Air Cent Bund

Air Anul Cent

Air N Bund

Air S Bund

Air E Bund

Air W Bund

Air Anul Periph

Figure 8.12 Air inlet temperatures for the different channels inside the apparatus during the ignition test 

8.2.2 Oxygen Profiles 

Figure 8.13 shows the response for all of the flow-through oxygen monitors except the north 

assembly. The oxygen monitor for the north assembly failed during installation and is not shown. 

Table 8.2 gives the time at which an oxygen concentration of 20%, 15%, and 5% was measured 

by each of the monitors considered. Also shown is the time to propagate from the center 

assembly to the peripheral assemblies based on the time to reach 15% oxygen. These oxygen 

measurements indicate that ignition occurred first in the center assembly at 4.83 hours. Figure 

8.14a shows the overhead view of the apparatus at 5 hours elapsed test time. The west peripheral 

assembly was the first to ignite at 6.73 hours, 1.9 hours after the center assembly. The east and 

south peripheral assemblies were not far behind, igniting about 6 minutes later at 6.84 and 6.87 

hours or about two hours after the center assembly. 

Although oxygen concentration first decreased peripherally in the west assembly, the oxygen 

consumption rates in the west were lower than in the south and east assemblies. The oxygen 

concentrations dropped much steeper and to lower levels in the south and east assemblies 

indicating a faster radial propagation. Video evidence confirms these trends. The video shows 

the west brightened first followed by the south, east, and then north assemblies (see Figure 

8.14b). Smoke, which indicates temperatures high enough to melt the Zircaloy, was emitted first 

from the east assembly followed by the south, west, and then north assemblies. Post-mortem 

inspection indicated that the south and east peripheral assemblies were the most heavily 

damaged, as discussed in more detail later. 
 

 166 

0

5

10

15

20

25

012345678

Time (hrs)

Oxygen (%)

Oxygen Center

Oxygen South

Oxygen East

Oxygen West

Figure 8.13 Response of all oxygen monitors as a function of time 

Table 8.2 Elapsed experimental time to reach reduced oxygen concentrations for the oxygen monitors 

 Location 

O2 Monitor time 

to 20% O2 

(hr) 

O2 Monitor time 

to 15% O2

(hr) 

O2 Monitor time 

to 5% O2

(hr) 

Radial propagation 

time for 15% O2

(hr) 

Center 3.39 4.83 4.92 -- 

West 5.88 6.73 6.96 1.90 

South 6.12 6.87 6.91 2.05 

East 6.22 6.84 6.86 2.01 
 

 167 

(a) (b) 

Figure 8.14 Overhead images of the test apparatus at elapsed test times of (a) 5 hours and (b) 7 hours 

8.2.3 Axial Temperature Profiles 

Ignition occurred in the center of the center assembly at 4.83 hours (see Figure 8.15) at levels 

31.6, 34.6, and 37.6 in. The burn front moved downward in the center assembly at an overall 

rate of 26 in/hr. The corresponding peak temperature in the peripheral assemblies increased 

correspondingly, but self-sustaining ignition did not occur until after the burn front in the center 

assembly reached the bottom tie plate (see Figure 8.16). Based on data from the oxygen 

monitors, the west peripheral assembly was the first to ignite at 6.72 hours. Ignition occurred 

uniformly across the 10.6-in. level of the west assembly by 6.9 hours. At this time all 

thermocouples failed, but the burn continued according to video evidence in all assemblies for 

over 33 hours despite the termination of all air flows and failed attempts to quench the fire with 

argon gas. 

Oxygen concentration is shown along with the axial temperature plots in Figure 8.15 and Figure 

8.16. For the center assembly shown in Figure 8.15, the oxygen concentration begins to drop 

below 20% when the hottest level measured reaches 1058 K. The oxygen concentration drops 

sharply to 15% when the hottest level rises rapidly to 1300 K suggesting these two criteria as 

indicators for ignition in the center assembly. The radial gradient in the center assembly near the 

ignition point is fairly shallow, on the order of 100 K (see Figure 8.22) so Zircaloy temperatures 

in the range of 1200 K to 1300 K are needed to drop the oxygen concentration to low levels at 

the top of the assembly. 

Figure 8.16 shows the temperature histories for the axial array in the east and west peripheral 

assemblies. The axial array is on the rod closest to the center assembly as shown in the inset 

diagram. The temperature of this rod increased sharply in response to the high temperature from 

the burn in the center assembly. As the burn front proceeded downward in the center assembly, 

the corresponding temperature on the axial array in the peripheral assembly peaked to higher and 

higher temperatures. However, self-sustaining ignition did not occur even though the 

temperatures were peaking at 1200 K to 1300 K. These temperatures decreased once the burn 

front in the center assembly passed by. As will be discussed in some more detail below, this is 

due to the steep radial temperature profile present in the peripheral assemblies. The additional 

oxidation energy was not sufficient to overcome the radiative heat loss when cool neighboring 

rods within the assembly were present. Once the radial temperature gradients were flattened 
 

 168 

significantly, ignition did occur. Ignition occurred simultaneously at the 1.6-in., 4.6-in., and 7.6-

in. levels at 6.8 hours and then burned upwards. Once ignition happened at the bottom of the 

assembly, all the TCs failed so it is not possible to provide any burn rate information. 

For the east and west peripheral assemblies shown in Figure 8.16, the situation is complicated by 

the steep radial temperature gradients. Figure 8.17 shows a detail of the west peripheral assembly 

along with the oxygen concentration near the time of the assembly ignition. When the hottest rod 

closest to the center assembly (W-EZ-10.6) first reaches 1300 K, the coolest rod farthest from 

the center assembly (W-EQ-10.6) is just over 700 K or almost 600 K cooler. Only the oxygen in 

the air passing near the hottest rods is consumed. The oxygen in the air flowing through the 

cooler regions bypasses the hottest rods and keeps the oxygen levels elevated when measured at 

the top. When the rod temperatures at the cooler side of a peripheral assembly approach 1200 K, 

oxygen concentration then begins to drop sharply. 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

0

3

6

9

12

15

18

21

Oxygen (%)

C-GX 1.6

C-GX 4.6

C-GX 7.6

C-GX 10.6

C-GX 13.6

C-GX 16.6

C-GX 19.6

C-GX 22.6

C-GX 25.6

C-GX 28.6

C-GX 31.6

C-GX 34.6

C-GX 37.6

C-GX 40.6

C-GX 43.6

C-GX 46.6

C-GX 49.6

O2 Center

TC location 

in red

South North

East 

West 

Figure 8.15 Temperature of the GX rod in the center assembly as a function of time for different axial 

locations 
 

 169 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

0

3

6

9

12

15

18

21

Oxygen (%)

W-EZ 1.6

E-EQ 4.6

W-EZ 7.6

W-EZ 10.6

W-EZ 13.6

E-EQ 16.6

W-EZ 19.6

E-EQ 22.6

W-EZ 25.6

E-EQ 28.6

W-EZ 31.6

E-EQ 34.6

W-EZ 37.6

E-EQ 40.6

W-EZ 43.6

E-EQ 46.6

W-EZ 49.6

Oxygen East

Oxygen West

TC locations 

in red

South North

East 

West

Figure 8.16 Temperature of the EZ and EQ rods in the west and east assemblies, respectively, as a 

function of time for different axial locations 

300

500

700

900

1100

1300

1500

1700

5 5.5 6 6.5 7 7.5 8

Time (hrs)

Temperature (K)

0

3

6

9

12

15

18

21

Oxygen (%)

Cha WE-10.6-EZ

W-EZ 10.6

W-EX 10.6

W-WEU 10.6

W-ES 10.6

W-EQ 10.6

Cha WW-10.6-EQ

Oxygen West

TC locations 

in red

South 

North 

East 

West 

Figure 8.17 Detailed temperature of the peripheral assembly as a function of time at z = 10.6 in. 
 

 170 

When ignition occurred in the center assembly at 4.83 hours (see Figure 8.15), the temperatures 

at levels 31.6, 34.6, and 37.6 in. increased sharply while the rest of the rod temperatures drop 

significantly. The drop in temperature was due to the expected heater rod failure once the rods 

ignited. At the time of ignition, the power input from Zircaloy oxidation is greater than twice that 

of the decay power supplied by the heaters therefore the impact of the heater rod failure is 

minimal, especially in the peripheral assemblies. The three locations that increased in 

temperature upon heater rod failure indicate the axial level where ignition occurred. The 

temperature at 34.6 in. increased at the fastest rate indicating this level was closest to the 

initiation point. The closest measurement points above and below this point, 37.6 and 31.6 in. 

respectively, increased at a slightly slower but equal rate. This suggests that the ignition initiation 

occurred across the 6-in. axial level from 31.6 to 37.6 in. The burn front then moved downward. 

Table 8.3 summarizes the time of ignition and apparent axial burn rate for all the axial levels. 

Comparison was made at 1220 K because this is the highest temperature available for all the 

axial levels. The burn front moved downward at an overall rate of 26 in/hr, which is half that 

measured in the full-length ignition test. This decrease in burn rate is most likely due to the 

transmission of energy to the colder peripheral assemblies via heat transfer. 

As the burn front progressed downward, it is not likely that all of the Zirconium was consumed. 

The unconsumed Zirconium did not continue to oxidize due to lack of oxygen. When the burn 

front reached the bottom tie plate, it is likely that the burn would stall at this level until all of the 

Zirconium is consumed and then progress back upward burning the remaining unconsumed 

Zirconium. Since all the TCs were lost by this time, there is no data to measure the upward burn 

rate, however, video evidence shows that the center assembly burned long after the front reached 

the bottom. 

Table 8.3 Measured burn rate at different axial heights based on an ignition criterion of 1220 K 

Axial location 

(in) 

Time to 1220K 

(hr) 

Downward burn rate 

(in/hr) 

34.6 4.820 -- 

31.6 4.880 50.0 

28.6 5.020 21.4 

25.6 5.145 24.0 

22.6 5.275 23.1 

19.6 5.453 16.9 

16.6 5.623 17.6 

13.6 5.705 36.6 

10.6 5.806 29.7 

7.6 5.896 33.3 

4.6 6.055 18.9 

1.6 6.090 85.7 

 Overall 26.0 
 

 171 

8.2.4 Radial Temperature Gradients 

Figure 8.18 through Figure 8.23 show the temperature history of the central and peripheral 

assemblies at each of the three radial temperature arrays located at axial positions 10.6, 25.6, and 

40.6 in. The series of temperatures shown parametrically span the apparatus radially at each of 

the axial levels as shown in the inset diagram. For each of the center assembly plots, the 

temperatures span from the center of the water rods, across three heater rods, the center channel 

box, pool rack, Boral neutron absorption plate, the peripheral channel box, and the inner most 

peripheral heater rod. In the peripheral plots, the temperatures span from the Boral plate across 

the peripheral channel box, two heater rods, the center of the water rods, two more heater rods, 

the back side of the channel box, and the back side of the peripheral pool cell. Note that the last 

three temperatures shown in the center assembly plots are the same as the first three temperatures 

shown in the peripheral plots. 

Figure 8.19, Figure 8.21, and Figure 8.23 show the radial temperature histories in the peripheral 

assemblies at the 10.6-in., 25.6-in., and 40.6-in. levels, respectively. Ignition is only evident at 

the 10.6-in. level because once this level burned all the TCs leading to the two upper levels were 

lost. The temperatures in the peripheral assemblies responded to the burn front in the center 

assembly with heating progressing from the face adjacent to the center assembly and moving 

inward toward the back wall. Figure 8.21 shows the radial temperature histories at the 25.6-in. 

level. The burn front in the center assembly reached this level at 5.15 hours. The temperature of 

the canister and the first row of heater rods reached temperatures of 1200 to 1300 K and the 

oxygen concentration was still near 20%; however, the oxidation reaction was quenched once the 

burn front passed. The oxidation reaction could not be sustained because of radiative heat loss to 

the back half of the peripheral assembly where the heater rods and canister were only about 

650 K. 

Figure 8.19 shows the radial temperature histories at the 10.6-in. level in the west peripheral 

assembly. The burn front in the center assembly reached this level at 5.8 hours. The temperatures 

of the channel box and first row of heater rods reached 1300 K, but oxidation was not sustained 

because the back half of the assembly was still less than 700 K. The center burn front moved 

down past the 10.6-in. level and reached the bottom tie plate shortly after 6.1 hours. The burn 

front then likely stalled until all of the Zirconium was consumed and then started back upwards. 

This stationary burn front encouraged radial propagation to the peripheral assemblies. Based on 

data from the oxygen monitors, the west peripheral assembly was the first to ignite at 6.72 hours 

and ignition occurred uniformly across the 10.6-in. level of the west assembly at 6.9 hours. The 

timing of the radial propagation is not represented prototypically in this experiment due to the 

shortened assembly length. Based on the observed behavior of this experiment and the previous 

integral test, radial propagation in prototypic full length assemblies would likely have occurred 

only after the center assembly burn front reached the bottom tie plate and returned upwards to 

consume the remaining Zircaloy. This extended burn length could add at least 2 hours to the time 

of radial propagation in a prototypic situation. 
 

 172 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

C-WEU 10.6

C-ES 10.6

C-ER 10.6

C-EQ 10.6

ChaCW-10.6-EQ

P-C-W10.6EQ

P-W-E10.6EZ

Cha WE-10.6-EZ

W-EZ 10.6

South North 

East

West

TC locations in red 

Figure 8.18 Temperature across the central portion of the apparatus as a function of time at z = 10.6 in. 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

P-W-E10.6EZ

Cha WE-10.6-EZ

W-EZ 10.6

W-EX 10.6

W-WEU 10.6

W-ES 10.6

W-EQ 10.6

Cha WW-10.6-EQ

P-W-W10.6EQ

South North 

East 

West 

TC locations in red 

Figure 8.19 Temperature across the peripheral portion of the apparatus as a function of time at z = 10.6 in. 
 

 173 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

C-WEU 25.6

C- CU 25.6

C-BU 25.6

C-AU 25.6

Cha CS-25.6-AU

P-C-S25.6AU

P-S-N25.6IU

Cha SN-25.6-IU

S-IU 25.6

South North 

East 

West 

TC locations in red 

Figure 8.20 Temperature across the central portion of the apparatus as a function of time at z = 25.6 in. 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

P-S-N25.6IU

Cha SN-25.6-IU

S-IU 25.6

S-GU 25.6

S-WEU 25.6

S-CU 25.6

S-AU 25.6

Cha SS-25.6-AU

P-S-S25.6AU

South North 

East 

West 

TC locations in 

Figure 8.21 Temperature across the peripheral portion of the apparatus as a function of time at z = 25.6 in. 
 

 174 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

C-WEU 40.6

C-EX 40.6

C-EY 40.6

C-EZ 40.6

Cha CE-40.6-EZ

P-C-E40.6EZ

P-E-W40.6EQ

Cha EW-40.6-EQ

E-EQ 40.6

TC locations in red 

South North 

East 

West

Figure 8.22 Temperature across the central portion of the apparatus as a function of time at z = 40.6 in. 

300

500

700

900

1100

1300

1500

1700

012345678

Time (hrs)

Temperature (K)

P-E-W40.6EQ

Cha EW-40.6-EQ

E-EQ 40.6

E-ES 40.6

E-WEU 40.6

E-EX 40.6

E-EZ 40.6

Cha EE-40.6-EZ

P-E-E40.6GZ

TC locations in red 

South North

East 

West

Figure 8.23 Temperature across the peripheral portion of the apparatus as a function of time at z = 40.6 in. 
 

 175 

Table 8.4 summarizes the temperature and temperature drop between these radial locations at the 

time of initial ignition at 4.83 hours. The largest temperature drops (on the order of 100 K) were 

across the annular gaps between the pool cells and the adjacent channel box walls. The 

temperature drop across the pool cell Boral plate varied considerably from 19 K at the 10.6-in. 

level to 94 K at the 25.6-in. level. The radial temperature gradients are the smallest, and in some 

cases reversed on the back side of the peripheral assemblies. Thermal radiation and conduction 

results in the back pool cell and channel box walls reaching higher temperatures sooner than the 

adjacent heater rods. 

Table 8.4 Measured temperatures across the apparatus at 4.82 hours 

Axial Level => 

Temperature and radial temperature drop at 4.82 hours 

10.6 in. 25.6 in. 40.6 in. 

Radial Location 

 

 center peripheral 

Temp 

(K) 

ΔT 

(K) 

Temp 

(K) 

ΔT 

(K) 

Temp 

(K) 

ΔT 

(K) 

Center water rod 1018 -7 1170 6 1228 29 

Center rod 1 1025 7 1164 24 1199 41 

Center rod 2 1018 35 1140 47 1158 47 

Center rod 3 983 44 1093 55 1111 54 

Center channel 939 92 1038 61 1057 111 

Pool cell - center side 847 19 977 94 946 57 

Pool cell - peripheral side 828 98 883 95 889 104 

Peripheral channel - center side 730 67 788 45 785 59 

Peripheral rod 4 663 35 743 88 726 91 

Peripheral rod 5 628 6 655 30 635 41 

Peripheral water rod 622 24 625 44 594 82 

Peripheral rod 6 598 30 581 35 512 25 

Peripheral rod 7 568 8 546 -17 487 -25 

Peripheral channel - back side 560 20 563 5 512 3 

Pool cell - back side 540 558 509 

The axial temperature profiles placed initial ignition in the center assembly at the 34.6-in. level. 

This level is located between the 25.6-in. and 40.6-in. radial temperature array levels shown in 

Figure 8.20 and Figure 8.22, respectively, which help with determining the radial location of the 

initial ignition. At all three of the radial levels, the central water rod temperature and the adjacent 

heater rod temperatures tracked together as the hottest. Figure 8.22 shows the temperatures along 

TC=> 1 2 3 4 5 6 7 water 

channel 

box 

pool 

water 
 

 176 

the radial segment across the center assembly at the 40.6-in. level. Just after the time of ignition, 

the water rod temperature increased above that of the adjacent heater rod. Figure 8.20 shows the 

temperatures along the radial segment across the center assembly at the 25.6-in. level. After 

ignition the heater rod power was lost, and the water rod and adjacent heater rod temperatures 

dropped. However, the heater rod temperature dropped more, leaving the water rod as the hottest 

location. When the burn front reached this level, the water rod was first to ignite at 5.06 hours, 

almost 4 minutes before the adjacent heater rod. This strongly suggests that the point of ignition 

in the center assembly was in the very center at the water rods at the 34.6-in. level. By the time 

the burn front reached the 10.6-in. level the radial temperature gradients had flattened. Here, the 

heater rods and water rods ignited simultaneously as shown in Figure 8.18. 

8.2.5 Post-Mortem of the Test Apparatus 

Figure 8.24 shows two photographs of the exterior of the apparatus after the ignition test with the 

insulation removed. On the left, the extensive damage to the south and east peripheral assemblies 

is clearly evident. The south and southeast cells have been breached at the middle axial level. 

Indeed, the entire pool rack shows signs of warpage and buckling. Note the exposed channel 

cans at the top of the apparatus. The original construction placed the top of the assemblies and 

pool rack at the same level (see Figure 8.3a). The photo on the right depicts the damage incurred 

in the north and west assemblies. Similar to the south and east, the pool rack shows signs of 

melting and geometric failure. 

 

Figure 8.24 Post-ignition test photographs of the apparatus with the insulation removed 

South East North West 

Exposed 

channel box 

Exposed 

channel box 
 

 177 

Figure 8.25a shows an overhead of the apparatus and further demonstrates the distortion caused 

during the ignition. Water rods in the center and east assemblies unmistakably protrude through 

the top tie plate indicating the slumping of the other structural components (bundle, channel box 

and pool rack). Figure 8.25b shows the air inlet flow boxes after their removal from the 

apparatus. The south assembly burnt through the bottom flow plate and deposited rod debris into 

the flow box. The east assembly deposited approximately 1 to 1.5 in. of aluminum from the 

Boral panels into its flow box. Remarkably, post-ignition inspection of all flow paths in the 

bundles and annuluses did not reveal any catastrophic blockages. 

(a) (b) 

Figure 8.25 Post-ignition overhead photographs of (a) the apparatus and (b) the inlet air flow boxes 

8.3 MELCOR Baseline Case Comparison 

The MELCOR baseline model was built from the previous 1×4 efforts (see Section 6.2.1). For 

this case the flow paths were open, allowing air flow rates and temperatures to be defined as inlet 

conditions. The nodalizations described in the earlier section are accurate. 

Recall that MELCOR does not account for radial temperature gradients in the fuel assembly tube 

bundle (as discussed in Section 6.2.2). In spite of this limitation, the experiments described in 

Section 6 demonstrate that MELCOR can adequately represent the radative coupling between the 

center and peripheral assemblies especially with respect to the impact on the center assembly. 

The primary objective of MELCOR whole pool calculations is to accurately predict when the 

first assembly ignites; therefore, accurately representing the center assembly is of paramount 

importance. 

8.3.1 Model Input Parameters 

The model input parameters were chosen based on earlier modeling and testing efforts. Figure 

8.26 shows the inlet air flow rates during the ignition test and the corresponding MELCOR 

model values. Again, the flow rate shown for the peripheral annuluses is twice the flow rate 

entering a single annulus. Measured and modeled values are all within experimental error. At the 

point of ignition in the center assembly the test flow program was manually interrupted to begin 

a ramp down in flow rate. 

North 
 

 178 

0

20

40

60

80

100

120

01234567

Time (hrs)

Flow Rate (slpm)

MELCOR_CentBund MELCOR_CentAnul MELCOR PeriphBund

MELCOR_PeriphAnul FlowCentBund FlowCentAnul

FlowPeriphBund FlowPeriphAnul

Figure 8.26 Comparison of air flow rates for the different channels inside the test apparatus and the 

MELCOR model 

Inlet air temperatures for the ignition test and the MELCOR model are plotted in Figure 8.27. 

The model values are consistent with the measured temperatures up until the ignition front 

neared the bottom of the apparatus. At this time, the TCs registered the heat-up of the stainless 

steel base plate. The air temperature measurement of the heated air into the north and west 

annuluses is not shown. This measurement was anomalously much higher than its counterpart in 

the south and east annuluses or bundles and was therefore disregarded. Peripheral annulus 

temperatures in the MELCOR model were assumed to track with the south and east annulus air 

temperature measurement until they reached the average temperature of the peripheral bundle air 

temperature. From this point on, the MELCOR peripheral annulus air temperature tracked with 

the MELCOR peripheral bundle air temperature. 
 

 179 

300

500

700

900

1100

1300

1500

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Time (hr)

Temperature (K)

MELCOR CentBund MELCOR CentAnul MELCOR PeriphBund MELCOR PeriphAnul

Air Cent Bund Air Anul Cent Air N Bund Air S Bund

Air E Bund Air W Bund Air Anul Periph

Figure 8.27 Comparison of air inlet temperatures for the different channels inside the test apparatus and 

the MELCOR model 

The default emissivity model for Zircaloy in MELCOR is shown graphically in Figure 8.28 [3]. 

The emissivity is assumed to be a function of oxide layer thickness, increasing sharply from 

0.325 for unoxidized Zircaloy to a nearly constant value of ~0.8 for oxide thicknesses greater 

than 3.88 m. For Zircaloy surfaces above a temperature of 1500 K, the emissivity is modified 

by a multiplicative factor decreasing the emissivity back to the initial value of 0.325 at ~1800 K. 
 

 180 

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

0 10 20 30 40 50

Oxide Layer Thickness (mm)

Emissivity (-)

MELCOR Default

MELCOR High Temp Model

1500 1600 1700 1800 1900

Temperature (K)

Figure 8.28 Default emissivity model in MELCOR with high temperature correction 

8.3.2 Comparison of Results 

Figure 8.29 gives the measured and predicted oxygen concentrations in the center and peripheral 

assemblies. Based on a 15% concentration level, the MELCOR model overestimated the time to 

reach ignition in the center by 31 minutes. The ignition of the peripheral assemblies was 

predicted to within 3 and 11 minutes for the west and south, respectively. 

(m) 
 

 181 

0

5

10

15

20

25

012345678

Time (hr)

Oxygen (%)

MELCOR Center MELCOR Periph

Oxygen Center Oxygen South

Oxygen East Oxygen West

Figure 8.29 Comparison of oxygen concentrations for the different assemblies inside the test apparatus 

and the MELCOR model 

Figure 8.30 shows the measured and modeled temperatures in the center assembly between z = 4 

to 14 in. as a function of time. Note that the center GX rod was about 30 K hotter than the 

average of radial temperatures and 20 K cooler than PCT (C-DT and C-FV) at the same axial 

level. Therefore, the MELCOR model conservatively represents a temperature hotter than the 

average rod temperature. This trend was also evident in the 1x4 Incoloy experiments discussed 

in Section 6. The model heat-up matched the experiment within 25 K for the first 4 hours. The 

apparatus then heated considerably faster than predicted. A possible explanation involves an 

increase in the reaction kinetics at temperatures around 1090 K and is discussed in a later 

sensitivity section. Heater failure occurred at 4.83 and 5.43 hours for the experiment and 

prediction, respectively. The MELCOR criteria for heater rod failure was a PCT of 1444 K. 

Figure 8.31 compares the model and experiment temperatures in the center assembly for axial 

locations bounding the initial location of ignition in the test apparatus. Again, the initial heat-up 

compares well with the experiment heating at a faster rate above temperatures of about 1090 K. 

The data indicates that ignition occurred at 4.82 hours at the 34.6” level. The baseline MELCOR 

model accurately predicted the ignition location to be between 31” and 37” and the ignition time 

was 36 minutes later than the experiment. In agreement with the data, MELCOR also predicted 

the burn front would move downward. 

The temperatures from the upper section of both model and experiment are plotted in Figure 

8.32. These results are similar to those discussed earlier. The model captures the beginning heatup accurately but under predicts the temperatures from 4 hours elapsed time to ignition. 
 

 182 

MELCOR utilizes a lumped thermal analysis that inherently characterizes each axial level in any 

given assembly with a single average temperature. This approach represents the center assembly 

relatively well because it is radially symmetric and the thermal gradients are relatively small. 

Due to asymmetry and steep thermal gradients, the lumped approach does not characterize the 

peripheral assemblies as well. Figure 8.33 gives the temperatures in the peripheral assembly for 

the MELCOR model and experiment. The west EZ rod is situated just inside the channel can on 

the center-facing side and represents the hottest location in the peripheral bundle. The MELCOR 

model predicted lower temperatures during the heat-up, which is to be expected since the single 

temperature MELCOR uses to characterize the tube bundle is more indicative of an average 

temperature. The base case MELCOR model predicted the timing of the peripheral assembly 

ignition closely, being only about 10 minutes early. However, the location of the peripheral 

assembly ignition and the direction of the burn front movement were not well represented. The 

base case MELCOR model predicted the peripheral assembly would ignite at the 19” to 25” level 

and burn downward while the data showed that the peripherals ignited simultaneously at the 1.6-

in., 4.6-in., and 7.6-in. levels at 6.8 hours and then burned upwards (see Section 8.2.3). Overall, 

the base case MELCOR representation of the center assembly ignition was very good. The 

representation of the peripheral assemblies was not as good due to the limitation of 

characterizing the tube bundle with a single temperature thus ignoring the complications of steep 

radial temperature gradients. 

300

500

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1100

1300

1500

1700

012345678

Time (hr)

Temperature (K)

MELCOR 7" - 13"

MELCOR 1" - 7"

C-GX 13.6

C-GX 7.6

C-GX 4.6

Figure 8.30 Comparison of temperatures inside the center assembly and MELCOR model for z = 4 to 14 in. 
 

 183 

300

500

700

900

1100

1300

1500

1700

1900

012345678

Time (hr)

Temperature (K)

MELCOR 31" - 37"

MELCOR 25" - 31"

C-GX 37.6

C-GX 34.6

C-GX 31.6

C-GX 28.6

Figure 8.31 Comparison of temperatures inside the center assembly and MELCOR model for z = 28 to 

38 in. 

300

500

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900

1100

1300

1500

1700

012345678

Time (hr)

Temperature (K)

MELCOR 43" - 49"

MELCOR 37" - 43"

C-GX 46.6

C-GX 43.6

C-GX 37.6

Figure 8.32 Comparison of temperatures inside the center assembly and MELCOR model for 

z = 37 to 47 in. 
 

 184 

300

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900

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1300

1500

1700

012345678

Time (hr)

Temperature (K)

MELCOR 1" - 7"

MELCOR 7" - 13"

MELCOR 13" - 19"

MELCOR 19" - 25"

MELCOR 25" - 31"

MELCOR 31" - 37"

MELCOR 37" - 43"

MELCOR 43" - 49"

W-EZ 7.6

W-EZ 13.6

W-EZ 19.6

W-EZ 25.6

W-EZ 31.6

W-EZ 37.6

W-EZ 43.6

W-EZ 49.6

Figure 8.33 Comparison of temperatures inside the peripheral assembly and MELCOR model 

8.4 MELCOR Sensitivity Study 

While the representation of the center assembly ignition by the base case MELCOR model was 

relatively good, the time to ignition was slightly over predicted. Since this is not conservative, a 

sensitivity study was conducted to determine which parameters could explain the discrepancy. In 

the hour before ignition, radiation was the most important heat transfer mechanism and 

zirconium oxidation was increasing in importance as an energy input. The two parameters 

considered were the emissivity of the zirconium oxide surfaces, which effects the radiation heat 

transfer, and the Zircaloy oxidation kinetics transition to breakaway, which effects the energy 

input from zirconium oxidation. Although these two parameters are not the only possible 

explanation for the differences between the MELCOR simulations and the test results, they do 

appear to be the variables with the greatest influence to reconcile the MELCOR results. 

As mentioned in the previous section, an increased heat-up rate was observed during the ignition 

test at temperatures around 1090 K. Adjustment of the zirconium oxide emissivity could be used 

to bring the MELCOR calculation in agreement with the increased heat-up rate observed; 

however, an aphysically low emissivitiy (~0.10) was required. Furthermore, the agreement 

between the MELCOR oxygen and the oxygen data was not significantly improved. Therefore, 

error in the zirconium oxide emissivity is not considered a viable explanation for MELCOR 

missing the increased heat-up rate after 1090 K and thereby slightly over predicting the time to 

ignition. 

Modification of the transition from pre- to post-breakaway Zircaloy oxidation kinetics brought 

the MELCOR calculation into agreement with both the increased heat-up rate and the oxygen 

concentration history observed. There is also great significance to temperatures around 1090K 
 

 185 

because this happened to be coincident with an alpha to beta phase change in Zircaloy that 

mechanistically could result in an oxidation kinetic rate change. The remainder of this section 

details the changes made to the baseline model to explore this sensitivity. 

8.4.1 Modification to Kinetics Model 

As previously asserted, the increased oxidation is assumed to be linked to the alpha to beta phase 

change occurring at ~1090 K. Figure 8.34 shows the circumferential coefficient of thermal 

expansion (CTE) for Zircaloy as a function of temperature [4]. The MATPRO manual further 

lists the transition for the phase change occurring from 1083 to 1244 K. The sharp change in the 

circumferential, as well as the axial, CTE leads surprisingly to decrease in volume as the material 

changes from the alpha to beta phases. This volume change is hypothesized to provide a 

mechanism for increased oxygen diffusion through the outer crust of Zirconium oxide via 

fracture and rupture of this external surface. 

Figure 8.35 shows the pre- and post-breakaway oxidation correlations from the ANL Zircaloy 

oxidation study [2]. For the default MELCOR model, the oxidation reaction follows the prebreakaway correlation until an auxiliary time-at-temperature function reaches a preset value. The 

reaction rate then jumps to the post-breakaway correlation. Note that in the ANL study, the 

kinetic rates determined were based on pre- and post reaction weight measurements and the 

reaction was conducted under isothermal conditions. Furthermore, only one temperature 

considered (1173 K) was above 1090K. This means that the ANL study provides no information 

on how the kinetics transition from pre- to post-breakaway and would not have detected any 

effects resulting from the alpha to beta phase change. 

As shown in Figure 8.35, the MELCOR oxidation correlations were modified for this study to 

ramp up to 5.3 times the pre-breakaway rate, from 1065 to 1115 K. The same time-attemperature function from the default model is still tracked and determines the shift from the 

modified kinetics curve to the post-breakaway correlation. The bounding temperature values of 

1065 to 1115 K were chosen based on the phase change temperature (1090 K) and the difference 

between the peak and average temperatures (50 K or ± 25 K) in the center assembly near the 

location of the initial ignition (see Figure 8.36). The reaction rate increase of 5.3 times was 

chosen to best match the heat-up rate after 1090K as well as the time of ignition. Although 

empirical, this modification is not considered to be unreasonable and is fully consistent with the 

ANL kinetics. 
 

 186 

Figure 8.34 Comparison of CTE prediction with Kearns’ model for Zircaloy in the circumferential 

direction 

 Note: Reproduced from MATPRO manual [4]. 

1.E-10

1.E-08

1.E-06

1.E-04

1.E-02

5 6 7 8 9 10 11 12 13 14 15

10000/T

k(T) [kg2/m4-s]

ANL Zr-4 Pre-Breakaway - Eqn. 6.8

ANL Zr-4 Post-Breakaway - Eqn. 6.9

Pre-breakaway (1090 K Phase Change)

Transition from low 

temperature to high 

temperature pre-breakaway 

oxidation kinetics

1065 

1115 

ANL Zr-4

Post-Breakaway 

Oxidation Correlation

ANL Zr-4

Pre-Breakaway 

Oxidation 

Figure 8.35 Modification of the default MELCOR oxidation kinetics model 
 

 187 

300

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700

900

1100

1300

1500

1700

0123456

Time (hrs)

Temperature (K)

Average at 38 in.

PCT at 38 in. Ignition and

thermocouple failure

dT/dt = 98 K/hr

dT/dt = 53 K/hr

dT/dt = 50 K/hr

dT/dt = 85 K/hr

First rod to 1090 K

at 3.423 hrs and

z = 37.6"

T = 50 K

Figure 8.36 Average and PCT as a function of time in the center assembly at z = 37.6 in. 

8.4.2 Comparison of Results 

Figure 8.37 compares the oxygen monitor results with the predicted values from the oxidation 

sensitivity case. The depletion in the center assembly was modeled to within 1 minute based on a 

15% criterion. The sensitivity case also captured the moderate decrease in oxygen concentration 

from 3 to 4.83 hours. The peripheral assemblies ignited just over 1 hour faster in the simulations. 

This discrepancy is likely due to the inability of the model to capture the radial temperature 

gradient in the peripheral assembly. 
 

 188 

0

5

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15

20

25

012345678

Time (hr)

Oxygen (%)

MELCOR O2 Sens - Center

MELCOR O2 Sens - Periph

Oxygen Center

Oxygen South

Oxygen East

Oxygen West

Figure 8.37 Comparison of oxygen concentrations for the different assemblies inside the test apparatus 

and the MELCOR sensitivity model 

Figure 8.38 shows the thermal response of both MELCOR models and two data series, the axial 

maximum of all radial PCTs and the axial maximum of all radially-averaged temperatures. The 

MELCOR baseline and sensitivity models track with the maximum of the PCTs up until 3 hours, 

which coincides with model temperatures of 1065 K. At times past 3 hours, the sensitivity model 

begins to deviate from the baseline and falls between the two data curves. The time of ignition 

was again predicted to within 1 minute, here based on a criterion of 1220 K. 

Similar to Figure 8.38, the thermal response of the peripheral assemblies channel box and bundle 

are shown in Figure 8.39 and Figure 8.40, respectively. The models trend with the axial 

maximum of the radially-averaged temperatures until ignition in the center assembly, or 4.83 

hours. The sensitivity model then follows the general trend of the lower data series but ignites 1 

hour earlier than measured. This difference in ignition times is attributed to the lack of the model 

to fully appreciate the radial temperature gradient. Therefore, the model could not capture the 

quenching effect of radiative heat loss within the peripheral assemblies on the oxidation reaction. 

Finally, the modified kinetics model was applied to the full-scale ignition test to determine its 

applicability to other situations. Figure 8.41 shows the thermal response of the full-scale 

apparatus and both models. Using the modified kinetics leads to an ignition time within 10 

minutes of the measured response, or a 2.3% error. The impact of the sensitivity case appears to 

have minimal impact on the overall response of the full-scale ignition test. 
 

 189 

300

500

700

900

1100

1300

1500

1700

0123456

Time (hr)

Temperature (K)

MELCOR

Baseline

MELCOR O2

Sensitivity

Data - max of

maximums

Data - max of

averages

Figure 8.38 Comparison of the maximum and average temperatures in the center assembly to the 

MELCOR baseline and sensitivity models 

300

500

700

900

1100

1300

1500

1700

01234567

Time (hr)

Temperature (K)

MELCOR

Baseline

MELCOR O2

Sensitivity

Data - max of

maximums

Data - max of

averages

 

Figure 8.39 Comparison of the maximum and average temperatures in the peripheral canister to the 

MELCOR baseline and sensitivity models 
 

 190 

300

500

700

900

1100

1300

1500

1700

01234567

Time (hr)

Temperature (K)

MELCOR

Baseline

MELCOR O2

Sensitivity

Data - max of

maximums

Data - max of

averages

Figure 8.40 Comparison of the maximum and average temperatures in the peripheral assemblies to the 

MELCOR baseline and sensitivity models

300

500

700

900

1100

1300

1500

1700

012345678

Time (hr)

Temperature (K)

Data Max

MELCOR Baseline

MELCOR O2 Sensitivity

Figure 8.41 Effect on temperature response in the full-scale Zircaloy MELCOR model to the modified 

oxidation kinetics model 
 

 191 

8.5 Summary 

The 1×4 ignition experiment represents the second integral effects test of the Zircaloy ignition 

phenomena and propagation program. This test utilized an array of five electrically-heated 

prototypic partial-length assembly segments. In addition, the assemblies were constructed of 

Zircaloy-clad heater rods and allowed for flow through the bundle and annular regions of each 

assembly. The objective of the test was to study the thermal history, ignition, and radial burn 

propagation behavior of a central heated assembly when adjacent to cold neighbors. The 

experimental approach was designed to represent a one-third axial piece of prototypic assemblies 

located in the upper portion of the fully populated bundle region. Each assembly was supplied 

with metered, heated air. The results of the full-length ignition test and MELCOR modeling were 

used to determine the flow and temperature histories to impose as boundary and inlet conditions. 

The center assembly was heated at 4.87 kW, which is the full-length equivalent of 17.5 day old 

fuel. This hot assembly was surrounded by four peripheral assemblies representing fuel at the 

average pool background power, or 0.2 kW. The center assembly ignited at the 34.6-in. level 

after 4.83 hours and burned down to the bottom tie plate by 6.1 hours. The peripheral assembly 

ignited near bottom at 6.8 hours and burned upwards. Post-mortem inspection revealed severe 

damage to both the pool rack and fuel assemblies. 

The initial MELCOR model estimated the time to ignition in the center assembly within 31 

minutes based on a 15% oxygen concentration criterion. Propagation into the peripheral 

assemblies was predicted to within 3 and 11 minutes for the west and south, respectively. The 

model heat-up matched the experiment within 25 K for the first 4 hours. The apparatus then 

heated considerably faster than predicted, perhaps due to a change in the oxidation kinetics. 

A sensitivity case was created in MELCOR to study the effect of an increased reaction rate at 

temperatures of a known alpha to beta phase change in Zircaloy. Many other parameters were 

considered and while consideration of the reaction kinetic parameter is reasonable it is not the 

only parameter that could be used to explain the small discrepancy in the MELCOR’s prediction 

of the center ignition time. The default MELCOR reaction kinetics were increased from the prebreakaway correlation by a factor of 5.3 times over the temperature range of 1065 to 1115 K. 

The kinetic rate increase was chosen such that the ignition in the center assembly was predicted 

to within 1 minute based on criteria of both 15% oxygen concentration and temperature of 1220 

K. The sensitivity model predicted propagation of ignition in the peripheral assemblies 1 hour 

earlier than measured. This difference in ignition propagation time is attributed to the lumped 

analysis approach inherent to MELCOR that results in a lack of the model to fully capture the 

radial temperature gradients in the peripheral bundles. As a result, the model could not capture 

the quenching effect of radiative heat loss within the peripheral assemblies on the oxidation 

reaction. 

 

 193 

9 REPORT SUMMARY 

The Spent Fuel Pool Heatup and Propagation Phenomena Experimental program was conducted 

from April 2004 until November 2006; over which time seven experimental apparatuses were 

tested. The objective of this project was to provide basic thermal-hydraulic data associated with 

a SFP complete loss-of-coolant accident. The accident conditions of interest for the SFP were 

simulated in a full-scale prototypic fashion (electrically-heated, prototypic assemblies in a 

prototypic SFP rack) so that the experimental results closely represent actual fuel assembly 

responses. A major impetus for this work was to facilitate code validation (primarily MELCOR) 

and reduce questions associated with interpretation of the experimental results. 

Table 9.1 summarizes the phased experimental approach employed to study Zircaloy fires in 

prototypic spent fuel assemblies. Three basic types of experiments were used to study these 

accident situations. As a proof of concept, two heater design tests were first performed to 

determine the suitability of the electrically-heated, Zircaloy-clad spent fuel rod simulators. Next, 

three separate effects tests were conducted to study and understand specific phenomena 

independently. Finally, two prototypic assemblies were heated to ignition in the integral effects 

test series.

Table 9.1 Summary of the testing elements in the experimental program 

Description Purpose Assembly Rod material 

Heater Design Test electrical heater performance, 

preliminary data on Zircaloy fire 

12 rod bundles Zircaloy 

Separate Effects Hydraulics – Determine form loss and 

laminar friction coefficients 

Prototypic Stainless Steel 

Separate Effects Thermal hydraulics – Measure naturally 

induced buoyancy driven flow rates. 

Prototypic Incoloy 

Separate Effects Thermal radiation – Radiation coupling in 

a 14 arrangement. 

Prototypic – Partial length Incoloy 

Integral Effects Axial ignition – Temperature profiles, 

induced flow, axial O2 profile, nature of 

fire. 

Prototypic Zircaloy 

Integral Effects Radial propagation – Determine nature of 

radial fire propagation. 

Prototypic – Partial length Zircaloy 

9.1 Heater Design Tests 

The main objectives of the two heater design tests were to prove that electrical heater rods could 

simulate spent fuel under accident conditions. The internal Nichrome heating element was able 

to successfully bring the Zircaloy cladding up to air ignition temperature before failure. In 

addition, the thermal mass of the magnesium oxide insulation packing was in excellent 

agreement with that spent fuel over a wide temperature range. As a result, the initial heat-up rates 

closely represented the heating of spent fuel rods. 
 

 194 

9.2 Separate Effects Tests 

For the first separate effects test series, a highly prototypic BWR assembly was constructed 

using prototypic components and stainless steel rods. The commercial components included the 

top and bottom tie plates, seven spacers, two water rods, the channel box, and all the nuts, 

springs, and washers used in the assembly process. The fuel rods were constructed of 7/16-in. 

stainless steel tubing that was within 0.002 in. of the diameter of a prototypic fuel rod. The pins 

were machined from stainless steel based on vender drawings and welded to the tubing ends. 

Extensive hydraulic characterization was conducted to determine the appropriate viscous and 

form loss coefficients for use with the MELCOR severe accident analysis code. The parameters 

determined were different from previously used and accepted textbook parameters. The 

experimental range spanned the laminar region with volumetric flow rates of 1.110-3 to 1.310-2

m3

/s. The resulting Reynolds numbers based on the bundle velocity and hydraulic diameter were 

70 to 900, respectively. These experiments included testing with the water rods in the assembly, 

both blocked and unblocked. 

For the second separate effects testing series, benchmark experiments were conducted with a 

highly prototypic electrically-heated BWR spent fuel assembly to measure thermal and hydraulic 

response. Peak cladding temperatures for both experiments and MELCOR simulations were 

within 5% error for all flow configurations and for all assembly power inputs, which ranged from 

200 W to 2500 W. Naturally-inducted flow rates were within 10% error over the entire power 

input range between tests and MELCOR. The overall excellent agreement of the MELCOR 

calculations with the experimental data is attributed to the direct application of prototypic 

viscous and form loss coefficients determined in previous, unheated hydraulic characterization. 

For the third separate effects testing series, an array of five electrically-heated prototypic partial 

length assembly segments were used to study thermal radiation effects. The experimental 

approach was designed to represent a one-third axial length from a grouping of five prototypic 

assemblies. The grouping arrangement studied was a central heated assembly with an adjacent 

assembly on each side. The assembly segments were located in an analogously shortened but 

otherwise highly prototypic 33 pool rack fabricated by Holtec Inc. Flow into the bottom and out 

of the top of each assembly was blocked in order to minimize convection and maximize the 

significance of thermal radiation heat transfer. The center assembly was rapidly heated to 

temperatures of 900 K and abruptly stopped. The temperature histories of the center and 

peripheral assemblies were analyzed with two codes, MELCOR and COBRA. Both codes were 

in very good agreement with the experimental data. MELCOR predicted the center assembly 

response better than the peripheral assembly response. This is due to assumptions built into 

MELCOR that allow axial temperature gradients in the fuel assemblies but no radial temperature 

gradients. Since prediction of initial ignition in the center assembly is of primary importance, the 

small deficiency in the peripheral assembly is acceptable. 

9.3 Integral Effects Tests 

In the first integral effects test, a single, highly-prototypic full length assembly was allowed to 

heat up until the Zircaloy-clad heater rods ignited. The objective was to measure the thermal 

history leading to ignition, the axial burn front advance rate, the buoyancy driven flow leading 

up to and during the burn, and the axial oxygen concentration profile. The results of these 

measurements were used to set boundary conditions in the second integral effects test. Ignition 
 

 195 

occurred after 7.3 hours of heating near the top of the full populated bundle region 96 in. above 

the bottom tie plate and 60 in. below the top tie plate. The MELCOR model predicted the peak 

cladding temperature of the assembly to within 40 K at all times and the ignition time to within 5 

minutes. 

The 1×4 ignition experiment represents the second integral effects test of the Zircaloy ignition 

phenomena and propagation program. This test utilized an array of five prototypic partial length 

assembly segments. The assemblies were constructed of Zircaloy-clad heater rods and allowed 

for flow through the bundle and annular regions of each assembly. The objective of the test was 

to study the thermal history, ignition, and radial burn propagation behavior of a central heated 

assembly when adjacent to cold (unheated) neighbors. The experimental approach was designed 

to represent a one-third length axial piece of prototypic assemblies located in the upper portion 

of the fully populated bundle region. Each assembly was supplied with metered, heated air. The 

results of the full-length ignition test and MELCOR modeling were used to determine the flow 

and temperature histories to impose as boundary and inlet conditions. The center assembly 

ignited at the 34.6-in. level after 4.83 hours and burned down to the bottom tie plate by 6.1 

hours. The peripheral assemblies ignited near bottom at 6.8 hours and burned upwards. Initial 

MELCOR modeling simulated the time to ignition in the center assembly to within 0.5 hours. 

Sensitivity of the model to the default oxidation kinetics correlations was also investigated, 

resulting in code predictions within 1 minute of the measured ignition in the center assembly. 

9.4 Key Findings 

The Spent Fuel Pool Heatup and Propagation Phenomena project was conducted from April 

2004 until November 2006; over which time seven unique experimental apparatuses were tested 

at Sandia National Laboratories in Albuquerque, New Mexico. In addition to these experiments, 

extensive simulation efforts were undertaken with the MELCOR severe accident modeling code 

to understand and predict the behavior observed in the tests. The key findings from this 

integrated experimental and simulation program are: 

 Electrically heated spent fuel rod simulators can be fabricated with Zircaloy cladding to 

accurately represent the decay heat, thermal mass and Zircaloy reactivity of a prototypic 

spent fuel rod. 

 The measured form and friction loss coefficients of a prototypic BWR assembly were 

significantly different from generally accepted values. Use of the measured coefficients was 

vital for accuracy when calculating (with MELCOR) the naturally induced flow in a heated, 

prototypic BWR assembly. 

 Incorporation of “breakaway” Zircaloy oxidation kinetics into MELCOR was vital for 

accurately capturing the Zircaloy heat-up to ignition and oxygen consumption. 

 For the full length ignition test, the MELCOR model predicted the peak cladding 

temperature (PCT) of the assembly to within 40 K at all times and the time of ignition to 

within 5 minutes. 

 For the 1×4 ignition experiment, the standard MELCOR model predicted ignition in the 

center and peripheral assemblies to within 30 and 15 minutes, respectively. The error in 
 

 196 

ignition timing between the simulations and experiment is approximately 10%. The 

difference in timing is likely due to the inability of the lumped parameter approach used in 

MELCOR to account for steep radial temperature gradients. 

 Post-mortem examination of the integral test assemblies revealed gross distortion of the 

pool rack and channel box, rubblization of the tubing bundle and accumulation of debris on 

the bottom tie plate that resulted in flow blockage. Flow blockage was also evident from 

molten aluminum (originating from Boral plates built into the pool rack) that collected on 

and below the bottom tie plates. 

9.5 Uncertainties and Potential Mitigating Factors 

In the course of this research, some uncertainties and approaches to mitigating the postulated 

accident were noted. These potential mitigating factors were not further evaluated as they were 

outside the study’s scope. 

 Characterization of a BWR assembly with the channel box removed was considered as a 

possible mitigation strategy. Removal of the BWR canister could increase the naturally 

induced flow rate and thus improve the coolability of the assembly. This hypothesis was 

noted on this research but was not quantified or confirmed. 

 The independent flow paths evaluated as part of this study (i.e. bundle, water rods, and 

bypass) were quantified by indirect methods. Direct measurements using techniques such 

as laser Doppler anemometry were not used, so this increased uncertainty in the test 

results. The partitioning of flow is of much greater importance in a PWR assembly, 

where the flow can exit the fuel bundle into the annular region along the entire length of 

the assembly. 

 The presence of both viscous and form loss in pure bundle runs indicates possible entry 

or exit effects from the spacers. Direct flow velocity measurements in the bundle region 

to confirm the presence of these disturbances or other unexpected flow structures were 

not made. 

 Water sprays may represent a possible mitigation approach for a complete-loss-of-coolant 

accident. If the supply of makeup water is not adequate to keep the pool from draining, 

spraying the available water onto the assemblies may be effective cooling strategy. 

Currently, the amount of water needed and the best delivery approach is uncertain. 

 Flooding the pool with an inert gas is another possible mitigation strategy to prevent or 

halt the combustion of Zircaloy components. The inert gas could displace the oxygen and 

thereby shut down the oxidation reaction. Cooling would still need to be considered to 

avoid releases due to ballooning of the fuel cladding and other high temperature effects. 
 

 197 

10 REFERENCES 

1. NUREG-1738, “Technical Study of Spent Fuel Pool Accident Risk at Decommissioning 

Nuclear Power Plants,” U.S. Nuclear Regulatory Commission, Washington, DC, February 

2001. 

2. NUREG/CR-6846, “Air Oxidation Kinetics for Zr-Based Alloys,” prepared by Argonne 

National Laboratory, U.S. Nuclear Regulatory Commission, Washington, DC, July 2004. 

3. NUREG/CR-6119, “MELCOR Computer Code Manuals, Reference Manual, Version 1.8.5, 

Vol. 2, Rev. 2,” U.S. Nuclear Regulatory Commission, Washington, DC, September 2001. 

4. NUREG/CR-6150, “SCDAP/RELAP5/MOD 3.2 Code Manual,” U.S. Nuclear Regulatory 

Commission, Washington, DC, July 1998. 

 

 A-1 

APPENDIX A. SFP HEATER DESIGN TEST 1 DATA SUMMARY 

Heater Currents

0

500

1000

1500

2000

5000 7000 9000 11000 13000 15000

time (sec)

Current (mA)

A CURRENT

B CURRENT

C CURRENT

D CURRENT

E CURRENT

F CURRENT 

G CURRENT

H CURRENT

W CURRENT

X CURRENT

Y CURRENT

Z CURRENT 

Figure A1 

Air Temperatures

0

100

200

300

400

500

600

700

0 5000 10000 15000 20000

time (sec)

temp (C)

OVEN WALL

AIR INLET 

CHIMNEY

Figure A2 
 

 A-2 

Exit Gas

0

5

10

15

20

25

6000 8000 10000 12000 14000 16000 18000

Time (sec)

Gas Conc (%)

O2(%)

LEL(%)

Figure A3 

Axial Temperatures on Rod D

-200

0

200

400

600

800

1000

1200

1400

1600

0 5000 10000 15000 20000

time (sec)

Temp (C)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

CHIMNEY

Figure A4 
 

 A-3 

Radial Temp Profile @ 12"

0

200

400

600

800

1000

1200

1400

0 5000 10000 15000 20000

Time (sec)

Temp (C)

W 12

X 12

Y 12

Z 12

Figure A5 

Radial Temp Profile @ 24"

-200

0

200

400

600

800

1000

1200

1400

1600

0 5000 10000 15000 20000

Time (sec)

Temp (C)

W 24 

X 24

Y 24

Z 24 

Figure A6 
 

 A-4 

Radial Temp Profile @ 36"

-200

0

200

400

600

800

1000

1200

1400

1600

0 5000 10000 15000 20000

Time (sec)

Temp (C)

W 36

X 36

Y 36

Z 36 

Figure A7 

Radial Temp Profile @ 42"

-200

0

200

400

600

800

1000

1200

1400

0 5000 10000 15000 20000

Time (sec)

Temp (C)

F 42 

Out 42

Figure A8 
 

 A-5 

Axial Temp Change on Rod D

-0.2

-0.1

0

0.1

0.2

0.3

0.4

0.5

5000 6000 7000 8000 9000 10000 11000 12000 13000 14000 15000 16000 17000

time (sec)

Temp Change (C/s)

D 48

D 42

D 36

D 30

D 24 

D 18

D 12

D 6

D 0

800W 0.1 scfm 1200W 1440W

 0.0 

 1.0 scfm scfm

Figure A9 

 

 B-1 

APPENDIX B. ERROR ANALYSIS 

The error and uncertainty inherent to an experimental result are critical to the accurate 

interpretation of the data. Therefore, the uncertainties in the experimental measurements are 

estimated in this section. Results of this analysis are given, followed by a general description of 

the method used and a brief explanation of the source of each reported measurement uncertainty. 

The overall standard uncertainty of an indirect measurement y, dependent on N indirect 

measurements xi, is defined in Equation 1. The standard uncertainty associated with an indirect 

measurement is analogous to the standard deviation of a statistical population. 

 

















 

N

i

i

i

u

x

y u

1

2

2 1

Here, u is used to define the standard uncertainty of a measurement. 

The expanded uncertainty U is reported in this appendix and defines the bounds that include 95% 

of the possible data. The expanded uncertainty is usually defined as some multiple of the 

standard uncertainty. Equation 2 shows the definition of the expanded uncertainty as used in the 

following sections. 

U  2  u 2

B.1 Uncertainty in Bundle Velocity for the Apparatus 

The uncertainty in the bundle velocity was determined using error propagation analysis (EPA) 

for the blocked water rod measurements. Errors associated with auxiliary flows through the 

bypass and water rods are not considered. However, the magnitude of these flows is small 

compared with the highest, measured bundle flow rate. The bundle velocity was determined from 

Equation 3 in which Qi is the volumetric flow rate in slpm for each flow controller, Abundle is the 

bundle area, R is gas constant for air, T is the ambient air temperature, and P is the ambient air 

pressure. The first term in the equation represents the conversion from slpm to kg/s. 

  



 



  



















   P

R T

A

Q Q

min

s 60

ρ l 0.001m

V

bundle

1 2 STP

3

bundle 3

Equation 4 gives the relation between the overall uncertainty of Vbundle and the contributions 

from the measurement uncertainties of Q, , and P. 

 

   

1 2

bundle

bundle

2 2 2

1 2 2 3 QQ T

STP 2

V 2 2

bundle 12 12

2 2 P A

bundle

RT RT Q QR

0.001m ρ PP P l

s 60 A min Q Q RT Q Q RT

P PA

uu u

u

u u

                                            

    

4 
 

 B-2 

Table B.1 summarizes the values used to determine the overall uncertainty of the bundle 

velocity. The overall uncertainty in Vbundle was found for the two highest volumetric flow rates, 

Q1 = 300 slpm and Q2 = 300 slpm, at a standard ambient condition of T = 298 K, P = 83,400 Pa, 

and Abundle = 9.8×10-3 m2

. The standard uncertainty was determined to be Vbundle u = 0.01 m/s. The 

uncertainty was most affected by volumetric flow rate, Q, which contributed 80% of the overall 

uncertainty. 

Table B. 1 Measurement uncertainties and intermediate calculations for Vbundle

Measurement, xi Standard Uncertainty, ui Influence Coefficient  

i

bundle

x

V



 Section 

Containing 

Explanation 

Volumetric Flow 

Rate, Q1

3.0 slpm 

P

RT B.1.1 

Volumetric Flow 

Rate, Q2

3.0 slpm 

P

RT B.1.1 

Ambient Air 

Temperature, 

1.1 K  

P

R 2 Q 1 Q  B.1.2 

Ambient Air 

Pressure, P 

110 Pa  

2 P

RT 2 Q 1 Q  B.1.3 

Bundle Hydraulic 

Area, Abundle

3.4 × 10-5 m2

 

2

bundle

Q Q RT 1 2

PA

 B.1.4 

B.1.1 Uncertainty in Volumetric Flow Rate Q 

The volumetric flow rate was controlled with two MKS volumetric flow controllers operated in 

parallel (Model # 1559A-24174). The uncertainty of the volumetric flow rate was determined 

from the stated manufacturer’s upper uncertainty of 1% of full scale. The uncertainties in flow 

rate were 3 slpm for flow controllers 1 and 2. The value shown in Table B.1 represents these 

standard uncertainties associated with the volumetric flow rate. 

B.1.2 Uncertainty in Ambient Air Temperature 

The air temperature was measured with a standard k-type TC. The standard uncertainty for this 

type of TC is uT = 1.1 K. 

B.1.3 Uncertainty in Ambient Air Pressure 

The air pressure was measured with a Setra Systems barometer (Model 276). The uncertainty of 

the ambient air pressure was taken from the manufacturer’s calibration sheet, which indicated an 

uncertainty in the instrument of ±0.1% of full scale (110,000 Pa). Therefore, the standard 

uncertainty in the pressure reading is uP = 110 Pa. 

B.1.4 Uncertainty in Bundle Hydraulic Error 

The inner dimension of the channel box was measured to within ± 0.127 mm (0.005 in). This 

tolerance leads to a standard uncertainty of 3.4 × 10-5 m2

 in the hydraulic area. The uncertainty in 
 

 B-3 

the inner channel box dimension dominates the uncertainty in the hydraulic cross-sectional area, 

including variations of ± 0.0254 mm (0.001 in) in the outer diameter of the rods. 

B.2 Uncertainty in Pressure Drop Measurements 

The manufacturer of the Digiquartz pressure transducers used in these experiments lists a static 

error band of ±0.02% of full scale. This error band includes repeatability, hysteresis, and 

conformance. Furthermore, these error bands consider the zero-drift of the instrument over 

periods of up to 14 years. Conversations with the manufacturer indicate the experimental 

procedure followed for these investigations, namely the zero flow measurements to correct any 

zero drift and the relatively short experimental data collection times (~ 2 minutes), should place 

the uncertainty in any pressure data closer to the resolution of the instrument, or 1 part per 

million of full scale. The observed noise level in the zero flow measurements was approximately 

twice the instrument resolution. This noise-based uncertainty is approximately 0.04 Pa, which is 

smaller than the plotted symbols in this report. Any spread in the data outside this range is 

believed to be caused by positioning errors associated with the installation and removal of the 

ported canister between experimental runs. 

B.3 Uncertainty in SLAM and k Coefficients 

The following procedure was adopted to determine the uncertainty in the SLAM and k 

coefficients. Because the greatest experimental uncertainty comes from the bundle velocity, the 

influence of the velocity on the quadratic curve fits was examined. The pressure drops across 2–8 

and 8–17 for the blocked water rod assembly were curve fit as a function of Vbundle ± Vbundle u . 

Figure B. 1 shows the resulting curve fits to the pressure drop data across 8–17. Using these 

curve fit coefficients, the error associated with the SLAM and and k coefficients may now be 

determined. This procedure was also followed for the derived pressure drop across 2–17. 
 

 B-4 

y = 9.68E+00x2 + 2.21E+01x

R2

 = 9.99E-01

y = 8.95E+00x2

 + 2.37E+01x

R2

 = 1.00E+00

0

10

20

30

40

50

60

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4

Vbundle (m/s)

 P (N/m 2

)

Figure B. 1 Pressure drop across 8–17 as a function of bundle velocity 

The two dashed curves represent the quadratic fits to the data shifted by ± Vbundle u . 

Table B. 2 summarizes the SLAM and k coefficients determined from this error analysis. The 

hydraulic diameter and bundle velocity used for 2–17 was the same as 8–17. 

Table B. 2 SLAM and k coefficients showing the effect of the uncertainty in bundle velocity 

Fitting velocity (m/s) 

2–8 8–17 2–17 

SLAM  k SLAM  k SLAM  k 

Vbundle 88 7.3 138 19.0 109 25.3 

Vbundle+ Vbundle u 92 7.1 143 18.2 113 24.3 

Vbundle- Vbundle u 84 7.6 133 19.7 105 26.3 

The uncertainty for SLAM and k coefficients appears to be dependent on bundle location of the 

pressure data. The maximum differences in the SLAM and k coefficients in Table B. 2 are taken to 

be the uncertainty, SLAM u = 5 and uk = 1. These values represent the conservative limit. 

B.4 Uncertainty in Thermal-Hydraulic Measurements 

The uncertainty in volumetric flow rate was determined using EPA. The flow rates were 

determined from Equation 5 in which Qi is the volumetric flow rate in slpm for each flow 

controller. 
 

 B-5 

Q = Q +Q tot 1 2 5

Equation 6 gives the relation between the overall uncertainty of Qtot and the contributions from 

the measurement uncertainties of Qi. 

  tot 1 2

2 2 2

Q QQ uuu       6

Table B. 3 summarizes the values used to determine the overall uncertainty of the volumetric 

flow rates. The overall uncertainty in Qtot was found to be 4.24 slpm. For situations where only 

one flow controller was used, the uncertainty was simply the stated manufacturer limit of 1% of 

full scale, or 3 slpm. 

Table B. 3 Measurement uncertainties and intermediate calculations for Qtot

Measurement, xi

Standard 

Uncertainty, ui

Influence Coefficient 

Qtot 

i x



 Section Containing 

Explanation 

Volumetric Flow 

Rate, Q1

3 slpm 1 A.1.1 

Volumetric Flow 

Rate, Q2

3 slpm 1 A.1.1 

B.4.1 Uncertainty in Omega Hot Wires 

The 95% uncertainty in the Omega FMA-900-V-R hot wires was stated by the manufacturer as 

3% of full scale. This translates to an expanded uncertainty of 0.01524 m/s, or 3 sfpm. 

B.4.2 Uncertainty in Measured Temperatures 

The standard limits of error of the k-type TCs are quoted by the manufacturer as ± 1.1 K. This 

uncertainty in temperature is smaller than the data symbols and does not display on the graphs 

presented in the text. 

 

 C-1 

APPENDIX C. TSI HOT WIRE ANEMOMETER MEASUREMENTS 

In addition to the integrated hot wires 1 and 3 (HW1 and HW3), a standalone TSI constant 

temperature hot wire system (Model IFA 300) was instrumented as a backup measurement to 

determine the flow rate in the annulus. During heated tests, this instrument was located at z = 

0.305 m (12 in.) inside the annulus formed by the channel box and pool cell. The lower threshold 

of this instrument was 0.2 m/s as determined from the lowest achievable external calibration 

velocity, described next. 

The single sensor, end flow probes (Model 1210-20) were externally calibrated prior to insertion 

into the assembly annulus. Figure C.1 shows the TSI calibration equipment (Model 1129). The 

probes are placed above the nozzle, which is capable of flows ranging between 0.2 and 5 m/s. 

Typical calibrations for these experiments are taken between 0.2 and 1 m/s. 

The system records ambient pressure and temperature during the calibration. Once the 

calibration file was saved to the dedicated TSI system, the probe was placed into the assembly 

annulus. The pressure and temperature at the probe were input into the system during data 

acquisition. These values were referenced back to the original calibration file with appropriate 

temperature and pressure corrections applied. The pressure was taken to be local ambient, and 

the temperature was measured with a TC placed next to the probe in the annulus (Figure C.2). 

The probe was mounted to a linear stage to enable traverses of the annulus. However, the probe 

was placed in the middle of the annulus for most of the testing. 

Figure C.1 Photographs of the TSI calibration equipment 
 

 C-2 

Figure C.2 Schematic showing the arrangement of the TSI hot wire in the annulus 

Figure C.3 gives the response of the TSI anemometer to a given volumetric annular flow rate for 

calibrations performed on January 20 and January 31, 2006. For these calibrations the probe was 

located in the center of the annulus. The two data sets are within the experimental uncertainty, 

which is dominated by uncertainty in the flow rate. 

TSI 1/31/06

y = 4.01E-03x2

 + 1.14E+00x + 2.46E+01

R2 = 1.00E+00

TSI 1/20/06

y = 3.55E-03x2

 + 1.15E+00x + 2.29E+01

R2 = 1.00E+00

-20

0

20

40

60

80

100

120

-40 -20 0 20 40 60 80

Velocity (cm/s)

QAnnulus (slpm)

TSI 1/20/06

TSI 1/31/06

Figure C.3 In situ calibration of the TSI hot wire located in the center of the annulus performed on 

1/20/06 (blue diamonds) and 1/31/06 (red squares) 

Traverse 

direction 

Pool cell wall 

Channel box wall 

TSI hot wire

k-type TC 

To TSI IFA 300 

Linear stage 

“V” block 
 

 C-3 

Figure C.4 shows a velocity traverse across the annulus with the TSI hot wire probe. The solid 

line depicts the corresponding laminar velocity profile for a volumetric flow rate of 72 slpm 

passing through the annulus. This traverse was collected on January 17, 2006, during a heated 

test with 1370 W applied to the apparatus in the closed bypass/open drains configuration. The 

data were collected for this traverse at approximately 608 minutes elapsed test time. Typical 

experimental uncertainties are shown on the graph. Recall that calibrations were limited to a 

lower operating threshold of 20 cm/s. The flow through the annulus appears to be laminar in 

nature. 

0

5

10

15

20

25

30

35

40

-0.006 -0.004 -0.002 0.000 0.002 0.004 0.006

x Annulus (m)

U Annulus (cm/s)

Figure C.4 Velocity traverse of the TSI hot wire across the annulus 

Note: The solid line depicts the corresponding analytical laminar velocity profile. 

D-1 

APPENDIX D. ADDITIONAL DATA AND FINAL POST-TEST 

MELCOR RESULTS FOR THE FULL-SCALE ZIRCALOY 

IGNITION TEST 

The following graphs represent additional data and final post-test MELCOR results from 

the full-scale ignition test not already presented in the report. 

Heater Power

0

1

2

3

4

5

6

0123456789

Time (hr)

Power (kW)

Heater Power

Cir. 12 "A" Power

Heaters fail at 7.31 hrs

PCT = 1444 K (96")

Pave = 4.975 kW

Figure D. 1 Input power for the experiment and the final post-test MELCOR model as a 

function of time 

Oxide Layer Thickness

0

100

200

300

400

500

600

0369

Time (hr)

Thickness (microns)

Lev 4

Lev 5

Lev 6

Lev 7

Lev 8

Lev 9

Lev 10

Lev 11

Lev 12

Lev 13

Figure D. 2 Cumulative oxide layer predictions from MELCOR at different axial levels 
D-2 

Bundle Oxygen Concentration 

5000 Watt Full Zr Test

0.00

0.05

0.10

0.15

0.20

0.25

0 2 4 6 8 10 12

Time (hr)

Mole Fraction (-)

CVH-X.4.101

CVH-X.4.102

CVH-X.4.103

CVH-X.4.104

CVH-X.4.105

CVH-X.4.106

CVH-X.4.107

CVH-X.4.108

CVH-X.4.109

Figure D. 3 Oxygen concentration at different axial levels within the MELCOR model 

Axial Cladding Temperature Profile

Individual Rod Data from 103" to 114"

300

500

700

900

1100

1300

1500

0123456789

Time (hr)

Temperature (K)

CS103-1A

CS108-1A

AU108-1

CS114-2A

TCL 110 (102 - 117 in)

Figure D. 4 Cladding temperatures at z = 103 to 114 in. for both experiment and MELCOR 
D-3 

0

40

80

120

160

200

02468

Time (hrs)

Q (slpm)

0

4

8

12

16

20



PBypass (N/m2

)

Hot Wire 2

Hot Wire 4

Pressure 1

Figure D. 5 Volumetric flow rate and bypass pressure drop as a function of time 




UNITED STATES 

NUCLEAR REGULATORY COMMISSION

WASHINGTON, DC 20555-0001

--------------------

OFFICIAL BUSINESS
NUREG/CR-7143 Characterization of Thermal-Hydraulic and Ignition Phenomena in Prototypic, 

Full-Length Boiling Water Reactor Spent Fuel Pool Assemblies After a 

Postulated Complete Loss-of-Coolant Accident