Spent Fuel Pool Project Phase I:

Pre-Ignition and Ignition Testing of a Single 

Commercial 17x17 Pressurized Water Reactor 

Spent Fuel Assembly under Complete Loss 

of Coolant Accident Conditions 

Office of Nuclear Regulatory Research

NUREG/CR-7215

Spent Fuel Pool Project Phase I: 

Pre-Ignition and Ignition Testing of a Single 

Commercial 17x17 Pressurized Water Reactor 

Spent Fuel Assembly under Complete Loss 

of Coolant Accident Conditions

Manuscript Completed: July 2013

Date Published: April 2016

Prepared by:

S.G. Durbin, E.R. Lindgren and A.S. Goldmann 

Sandia National Laboratories

Albuquerque, NM 87185

M. Zavisca, Z. Yuan, R. Karimi, A. Krall, and M. Khatib-Rahbar 

Energy Research, Inc.

Rockville, MD 20847-2034

G. A. Zigh, NRC Technical Monitor

A. Velazquez-Lozada, NRC Project Manager

S. Gonzalez, NRC Project Manager

NRC Job Code N6777

Office of Nuclear Regulatory Research

NUREG/CR-7215

 

iii

ABSTRACT

The US Nuclear Regulatory Commission (NRC), in collaboration with the Organization for 

Economic Corporation and Development (OECD), and 12 international partners, conducted an 

experimental program to obtain experimental data for the characterization of hydraulic and 

ignition phenomena of prototypic light water reactor fuel assemblies in a spent fuel pool under 

complete loss of coolant accidents for validation of the MELCOR severe accident computer 

code. MELCOR is a fully integrated, engineering-level computer code whose primary purpose 

is to model the progression of postulated accidents in light water reactors as well as non-reactor 

systems (e.g., spent fuel pool and dry cask). The experimental program, were conducted in two 

phases at Sandia National Laboratories. The first phase, described in this NUREG, focused on 

axial heating and zirconium fire propagation in a single 17x17 PWR fuel assembly. The results 

from the first phase of the experiments demonstrate that the MELCOR computer code can 

accurately simulate ignition timing and burn propagation in a single 17x17 PWR assembly under 

complete loss of coolant conditions.

v

CONTENTS

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

LIST OF FIGURES......................................................................................... ...........vii

LIST OF TABLES......................................................................................................................xi

EXECUTIVE SUMMARY .........................................................................................................xiii

ABBREVIATIONS ..................................................................................................................xvii

1 INTRODUCTION ............................................................................................................1-1

2 Experimental Apparatus and Procedures ...................................................................2-1

2.1 Flow Measurements and Calibrations ...............................................................2-1

2.1.1 Assembly Hydraulics .............................................................................. 2-1

2.1.2 Hot Wire Anemometer at the Assembly Inlet........................................... 2-3

2.1.3 Laser Doppler Anemometry Measurements............................................ 2-4

2.2 Thermocouple Layout .......................................................................................2-7

2.3 General Construction ........................................................................................2-8

2.3.1 Power Control......................................................................................... 2-8

2.3.2 Instrumentation External to the Assembly............................................. 2-10

3 Cell 1 Pre-Ignition Testing and Analysis .....................................................................3-1

3.1 Cell 1 – 1.0 kW Results.....................................................................................3-3

3.2 Cell 1 – 3.5 kW Results.....................................................................................3-7

4 Cell 2 Pre-Ignition Testing and Analysis .....................................................................4-1

4.1 Cell 2 – 1.0 kW Results.....................................................................................4-3

4.2 Cell 2 – 3.5 kW Results.....................................................................................4-6

4.3 Internal Assembly Mass Flow ...........................................................................4-8

4.3.1 Forced Flow Calibrations ........................................................................ 4-9

4.3.2 Naturally Induced Velocity Profiles........................................................ 4-10

5 Cell 2 Ignition Testing and Analysis ............................................................................5-1

5.1 Ignition Test and MELCOR results....................................................................5-1

5.1.1 Thermal-Hydraulic Comparisons............................................................. 5-1

5.1.2 Exhaust Stream Analyses....................................................................... 5-5

5.1.3 Post-Test Examination and Deconstruction ............................................ 5-8

5.2 MELCOR Sensitivity Studies...........................................................................5-10

5.2.1 Axial Nodalization Sensitivity (24A_1R) ................................................ 5-10

5.2.2 Radial Nodalization Sensitivity (12A_2R).............................................. 5-12

5.2.3 Combined Axial and Radial Nodalization Sensitivity (24A_2R) ............. 5-13

6 SUMMARY .....................................................................................................................6-1

7 REFERENCES ...............................................................................................................7-1

APPENDIX A – ERROR PROPAGATION ANALYSIS ...........................................................A-1

APPENDIX B – OXIDATION OF ZIRCONIUM........................................................................B-1

APPENDIX C – MELCOR MODEL DESCRIPTION................................................................C-1

vii

LIST OF FIGURES

Figure 1.1 Various components in a typical 17×17 PWR fuel assembly .............................1-1

Figure 2.1 Overall pressure drop as a function of average air velocity in the 

assembly for the 221.3 mm storage cell (Cell 1) ...............................................2-2

Figure 2.2 Overall pressure drop as a function of average air velocity in the 

assembly for the 223.4 mm storage cell (Cell 2) ...............................................2-2

Figure 2.3 Schematic showing the layout of the inlet hot wire ............................................2-3

Figure 2.4 Flow rate as a function of the TSI Model 8455 hot wire 

anemometer voltage .........................................................................................2-4

Figure 2.5 Dual-beam backscatter LDA system components and principles ......................2-5

Figure 2.6 Schematic of the LDA system for measuring velocity profiles in the 

PWR 17×17 assembly ......................................................................................2-6

Figure 2.7 Photographs showing the a) overall LDA setup, b) close-up of the 

probe head and optical window, and c) flow inlet and seed source ...................2-6

Figure 2.8 Thermocouple layout for the Phase I test assembly ..........................................2-7

Figure 2.9 Detail view showing TC attachment to a fuel rod...............................................2-8

Figure 2.10 Power control system and test circuit................................................................2-9

Figure 2.11 Schematic of the instrumentation panel ............................................................2-9

Figure 2.12 PWR test assembly and external instrumentation ...........................................2-11

Figure 2.13 Schematic of the assembly exhaust sampling system.....................................2-12

Figure 3.1 Comparison of the average test (blue diamonds) and MELCOR (pink) 

temperatures during pre-ignition testing for Cell 1 at z = 3.150 m (124 in.)

and t = 12 hrs....................................................................................................3-2

Figure 3.2 Comparison of the test (blue diamonds) and MELCOR (pink) flow rates 

during pre-ignition testing for Cell 1 at t = 12 hrs ...............................................3-3

Figure 3.3 Cell 1 bundle temperatures as a function of time at 1.0 kW simulated 

decay power for the average test temperature (blue diamonds) and 

MELCOR (pink) ................................................................................................3-4

Figure 3.4 Induced flow rate within the Cell 1 assembly at a simulated decay 

power of 1.0 kW for the test (blue solid) and MELCOR (pink) ...........................3-5

Figure 3.5 Cell 1 bundle temperatures as a function of axial height in the assembly 

at 1.0 kW simulated decay power and 12 hours elapsed for the test 

maximum (red squares), average (blue diamonds), minimum (green 

triangles), and MELCOR temperatures (pink) ...................................................3-6

Figure 3.6 Cell 1 bundle temperatures as a function of axial height in the assembly 

at 1.0 kW simulated decay power and 20 hours elapsed for the test 

maximum (red squares), average (blue diamonds), minimum (green 

triangles), and MELCOR temperatures (pink) ...................................................3-6

Figure 3.7 Cell 1 bundle temperatures as a function of time at 3.5 kW simulated 

decay power for the average test temperature (blue diamonds) and 

MELCOR (pink) ................................................................................................3-7

Figure 3.8 Induced flow rate within the Cell 1 assembly at a simulated decay power 

of 3.5 kW for the test (blue solid) and MELCOR (pink)......................................3-8

Figure 3.9 Bundle temperatures as a function of axial height in the assembly at 

3.5 kW simulated decay power and 12 hours elapsed for the maximum 

test temperature (red squares), average test temperature (blue 

diamonds), and MELCOR (pink) .......................................................................3-8

Figure 3.10 Contours of temperature within the assembly at the midplane for an 

input power of 3.5 kW .......................................................................................3-9
viii

Figure 3.11 Contours of temperature within the assembly at z = 3.150 m (124 in.) 

level for an input power of 3.5 kW ...................................................................3-10

Figure 4.1 Comparison of the average test (blue diamonds) and MELCOR (pink) 

temperatures during pre-ignition testing for Cell 2 at z = 3.150 m (124 in.) 

and t = 12 hrs....................................................................................................4-2

Figure 4.2 Comparison of the test (blue diamonds) and MELCOR (pink) flow 

rates during pre-ignition testing for Cell 2 at t = 12 hrs ......................................4-3

Figure 4.3 Cell 2 bundle temperatures as a function of time at 1.0 kW simulated 

decay power for the average test temperature (blue diamonds) and 

MELCOR (pink) ................................................................................................4-4

Figure 4.4 Cell 2 bundle temperatures as a function of axial height in the assembly 

at 1.0 kW simulated decay power and 12 hours elapsed for the maximum 

test (red squares), average test (blue diamonds), minimum test (green 

triangles) and MELCOR (pink) ..........................................................................4-5

Figure 4.5 Cell 2 bundle temperatures as a function of axial height in the assembly 

at 1.0 kW simulated decay power and 20 hours elapsed for the maximum 

test (red squares), average test (blue diamonds), minimum test (green 

triangles) and MELCOR (pink) ..........................................................................4-5

Figure 4.6 Induced flow rate within the Cell 2 assembly at a simulated decay 

power of 1.0 kW for the test (blue solid) and MELCOR (pink) ...........................4-6

Figure 4.7 Cell 2 bundle temperatures as a function of time at 3.5 kW simulated 

decay power for the average test (blue diamonds) and MELCOR (pink)...........4-7

Figure 4.8 Bundle temperatures as a function of axial height in the assembly at 

3.5 kW simulated decay power and 12 hours elapsed for the maximum 

test (red squares), average test (blue diamonds), minimum test (green 

triangles), and MELCOR (pink) .........................................................................4-7

Figure 4.9 Induced flow rate within the Cell 2 assembly at a simulated decay 

power of 3.5 kW for the test (blue solid) and MELCOR (pink) ...........................4-8

Figure 4.10 Normalized mass flux profiles within the fuel assembly for measured 

flows of 150 (red squares), 186 (blue diamonds), and 225 slpm (green 

triangles) and for simulated flows of 120 (open pink squares) and 

240 slpm (open light-blue diamonds) ..............................................................4-10

Figure 4.11 Normalized mass flux profiles within the fuel assembly of the naturally 

induced flows at 0.5 kW for measured (red squares) and simulated 

(open blue diamonds) .....................................................................................4-11

Figure 4.12 Normalized mass flux profiles within the fuel assembly of the naturally 

induced flows at 1 kW for measured (red squares) and simulated 

(open blue diamonds) .....................................................................................4-11

Figure 4.13 Bundle and annulus mass flow fraction at z = 0.442 m (17.4 in.) within 

the assembly for the bundle from the LDA line average (red squares), 

CFD line average (blue diamonds), and CFD area average (black circles) 

and for the annulus from the LDA line average (open red squares), CFD 

line average (open blue diamonds), and CFD area average (open 

black circles)...................................................................................................4-12

Figure 4.14 Bundle and annulus mass flow fraction within the assembly for the 

bundle forced flow (red squares), 0.5 kW (blue diamonds), 1 kW (green 

triangles), and 3.5 kW (purple circles) and for the annulus forced flow 

(open pink squares), 0.5 kW (open light-blue diamonds), 1 kW (open 

light-green triangles), and 3.5 kW (open light-purple circles)...........................4-13
ix

Figure 5.1 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated 

decay power for the maximum test temperature (red squares), average 

test temperature (blue diamonds), and MELCOR 12A_1R (pink dashed) .........5-2

Figure 5.2 Bundle temperatures as a function of axial height in the assembly at 

5.0 kW simulated decay power and 12 hours elapsed for the test maximum 

(red squares), test average (blue diamonds), test minimum (green 

triangles), and MELCOR 12A_1R (pink dashed)...............................................5-3

Figure 5.3 Bundle temperatures as a function of axial height in the assembly at 

5.0 kW simulated decay power and 10 hours elapsed for the test maximum 

(red squares), test average (blue diamonds), test minimum (green 

triangles), and MELCOR 12A_1R (pink dashed)...............................................5-4

Figure 5.4 Induced flow rate within the Cell 2 assembly at a simulated decay power 

of 5.0 kW for the test (blue solid) and MELCOR 12A_1R (pink dashed)............5-4

Figure 5.5 Exit oxygen concentration as a function of time for assembly bundle (red 

squares), assembly annulus (blue diamonds), and MELCOR 12A_1R 

(pink dashed)....................................................................................................5-5

Figure 5.6 Overhead photographs of the assembly exit and exhaust sampling lines 

showing a) the initial configuration, b) immediately prior to the break in the 

bundle sample line, and c) immediately after the break in the bundle 

sample line........................................................................................................5-5

Figure 5.7 Mole fraction of nitrogen, oxygen, and argon as a function of time during 

the ignition test..................................................................................................5-6

Figure 5.8 Ratio of nitrogen to argon in the exhaust stream and assembly flow rate 

as a function of time..........................................................................................5-8

Figure 5.9 Deconstruction photographs of the SFP Phase I test assembly ........................5-9

Figure 5.10 Photographs of the SFP Phase I fuel bundle after the ignition test..................5-10

Figure 5.11 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated 

decay power for the maximum test temperature (red squares), average 

test temperature (blue diamonds), and MELCOR 24A_1R (pink dashed) .......5-11

Figure 5.12 Bundle temperatures as a function of axial height in the assembly at 

5.0 kW simulated decay power and 10 hours elapsed for the test 

maximum (red squares), test average (blue diamonds), test minimum 

(green triangles), and MELCOR 24A_1R (pink dashed)..................................5-11

Figure 5.13 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated 

decay power for the maximum test temperature (red squares), average 

test temperature (blue diamonds), MELCOR 12A_2R[1] (pink long 

dashed), and MELCOR 12A_2R[2] (orange short dashed) .............................5-12

Figure 5.14 Bundle temperatures as a function of axial height in the assembly at 

5.0 kW simulated decay power and 10 hours elapsed for the test maximum 

(red squares), test average (blue diamonds), test minimum (green 

triangles), MELCOR 12A_2R[1] (pink long dashed), and MELCOR 

12A_2R[2] (orange short dashed)...................................................................5-13

Figure 5.15 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated 

decay power for the maximum test temperature (red squares), average 

test temperature (blue diamonds), MELCOR 24A_2R[1] (pink long 

dashed), and MELCOR 24A_2R[2] (orange short dashed) .............................5-14

Figure 5.16 Bundle temperatures as a function of axial height in the assembly at 

5.0 kW simulated decay power and 10 hours elapsed for the test maximum 

(red squares), test average (blue diamonds), test minimum (green 

triangles), MELCOR 24A_2R[1] (pink long dashed), and MELCOR 

24A_2R[2] (orange short dashed)..........................................................5-14
x

Figure A.1 Flow rate as a function of hot wire anemometer voltage..............................A-6

Figure B.1 Comparison of the MELCOR breakaway timing fit to Zr-4 data (Ref. 10).........B-3

Figure C.1 Nodalization of PWR spent fuel assembly (Cell 2 in Phase I).......................C-2

Figure C.2 Heat structures for the rack wall (Cell 2 in Phase I)....................................C-3

Figure C.3 Representation of the core cell nodes...............................................................C-4
xi

LIST OF TABLES

Table 2.1 Summary of assembly hydraulics for storage Cell 1 and Cell 2.........................2-3

Table 2.2 Detailed list of the LDA system components .....................................................2-5

Table 2.3 List of equipment used for power control.........................................................2-10

Table 3.1 Summary of pre-ignition testing results for Cell 1 at z = 3.150 m (124 in.).........3-1

Table 4.1 Summary of pre-ignition testing results for Cell 2 at z = 3.150 m (124 in.).........4-1

Table A.1 Measurement uncertainties and intermediate calculations for wo .....................A-2

Table A.2 Measurement uncertainties and intermediate calculations for Aassembly .............A-2

Table A.3 Measurement uncertainties and intermediate calculations for SLAM. .................A-4

Table A.4 Measurement uncertainties and intermediate calculations for Σk. ....................A-4

Table A.5 Measurement uncertainties and intermediate calculations for DH.....................A-4

Table A.6 Measurement uncertainties and intermediate calculations for Pwet. ..................A-5

Table A.7 Measurement uncertainties and intermediate calculations for air density, ρ. ....A-5

Table A.8 Measurement uncertainties and intermediate calculations for assembly flow 

rate from the TSI hot wire.................................................................................A-6

Table A.9 Measurement uncertainties and intermediate calculations for assembly flow 

rate from the Dwyer 2 hot wire. ........................................................................A-6

Table B.1 MELCOR Fit of the Timings for Transition from Pre-Breakaway to PostBreakaway Oxidation Reaction Kinetics Zircaloy-4 in the ANL Experiments 

(Ref. 10)...........................................................................................................B-4

xiii

EXECUTIVE SUMMARY

In 2001, an evaluation of the potential accident risk in a spent fuel pool (SFP) at a US nuclear 

plant was performed. NUREG-1738, “Technical Study of Spent Fuel Pool Accident Risk at 

Decommissioning Nuclear Power Plants,” describes a modeling approach for a typical 

decommissioning plant with design assumptions and industry commitments, the thermal 

hydraulic analyses performed to evaluate spent fuel stored in the SFP, the risk assessment of 

SFP accidents, the consequence calculations, and the implications for decommissioning 

regulatory requirements. Some of the assumptions in the accident progression were known to 

be necessarily conservative, especially the estimation of the fuel damage. The SFP accident 

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

progression following various postulated initiators. These studies identified various modeling 

and phenomenological uncertainties that prompted a need for experimental confirmation.

Previous analyses have shown that fuel assemblies can ignite and propagate in a SFP during a 

complete loss of coolant scenario. Hence, qualified data obtained in representative fuel 

configurations were needed to confirm these results. In 2003, the NRC undertook an 

experimental program to address thermal-hydraulic conditions and zirconium fire propagation 

during a complete loss of coolant event in a boiling water reactor spent fuel pool. These 

experiments are summarized in NUREG/CR-7143. The experiments showed that the measured 

form and friction loss coefficients of a prototypic BWR assembly were significantly different from 

generally accepted values and that the use of the measured coefficients was vital for accuracy 

when calculating (with MELCOR) the naturally induced air flow rate. The incorporation of 

“breakaway” Zircaloy oxidation kinetics into MELCOR was also necessary to accurately capture

the Zircaloy heat-up to ignition and oxygen consumption.

In May 2009, the NRC, in collaboration with the Organization for Economic Cooperation and 

Development/Nuclear Energy Agency (OECD/NEA) and 12 international partners signed an 

agreement called the “OECD/NEA Sandia Fuel Project - An Experimental Programme and 

Related Analyses for the Characterization of Hydraulic and Ignition Phenomena of Prototypic 

Water Reactor Fuel Assemblies.” This program defined an experimental test matrix, 

experimental conditions and parameters to be examined with the overall objective being to 

perform a highly detailed thermal-hydraulic characterization of a full-length commercial fuel 

assembly mockup to provide data for the direct validation of severe accident computer codes 

(i.e., MELCOR). This experimental work was valuable to the OECD/CSNI community to 

facilitate severe accident code validation (ATHLET-CD, ASTEC, Ansys Fluent, DRACCAR, and 

MELCOR) and reduce modeling uncertainties within the codes. Further, it provided a 

collaborative exchange of information on severe accident modeling which resulted in increased 

knowledge to participating members that can be used in their regulatory programs and future 

research.
xiv

The study was conducted in two phases; Phase I (documented in this NUREG), focused on 

axial heating and burn propagation in a single assembly and Phase II (documented in 

NUREG/CR-7216) focused on axial and radial heating, burn propagation and the effects of fuel 

rod ballooning. Phase II consisted of a single heated assembly surrounded by four unheated 

assemblies representing a 1x4 loading pattern.

A single test assembly of a full-length commercial 17x17 PWR fuel bundle was constructed 

using prototypic, commercial components with heater rods made from zirconium alloy tubing 

supplied by an industrial vendor. The fuel assembly contains the core skeleton including eleven 

spacers permanently attached to twenty-five guide tubes and 264 fuel rods which pass through 

the spacers and are held captive in the assembly by the top and bottom nozzles. The assembly 

was placed into two different size pool racks during the testing. The test assembly was 

completely insulated to model boundary conditions representing a “hot neighbor” loading 

pattern. The test assembly was fully instrumented including hot wire anemometers (flow rate), 

oxygen sensors, gas chromatograph (Ar and N2 quantification), quartz light pipes (visual 

observation), laser Doppler anemometer (velocity), pressure transducers and thermocouples.

Separate effects tests were performed to investigate the assembly hydraulic response. For 

these tests, the assembly was unheated, and flow was forced into the assembly covering the 

expected range of flow rates. Tests were performed and these values were computed for both 

pool rack cell sizes. The experimental data for flow rate and pressure drop was used to 

compute both the frictional and inertial flow resistance coefficients SLAM and Σk (along with their 

associated uncertainties).

Pre-ignition tests were conducted using a uniform axial power profile with electrically heated 

rods to simulate decay powers from 0.5 to 3.5 kW (in steps of 0.5 kW). These tests were also 

performed with both pool rack cell sizes. The key parameters for these non-destructive tests 

are temperatures throughout the fuel assembly and inlet mass flow rate.

The ignition test was conducted at a simulated decay power of 5.0 kW and using the slightly 

larger pool rack cell size. The power is equivalent to an offload age of approximately 17 months 

(assuming fuel burn up of 45 GWd/MTHM). Ignition of the Zircaloy within the assembly was first 

indicated at an elapsed test time of 12 hours 40 minutes. Power to the assembly was lost 

shortly after ignition occurred.

During the ignition test, a residual gas analyzer (RGA) was used to monitor the amount of 

nitrogen and argon exiting the top of the assembly. The ratio of nitrogen to argon was used to 

determine if nitrogen was being consumed by reaction with zirconium. While encountering 

some sampling difficulties, the RGA successfully measured a significant amount of nitrogen 

consumption at the start of ignition and during the burn front progression to the bottom of the 

assembly. The single point calibration with air was used for samples containing significant 

concentrations of oxygen. The analysis of two additional calibration gases allowed adjustment 

of the air calibration for samples that were devoid of oxygen. During the burn phase, all of the 

oxygen was removed from the air drawn into the assembly converting 14 percent of the initial 
xv

zirconium to ZrO2. Depending on which background ratio was assumed, between 20 percent to 

40 percent of the zirconium in the assembly was converted to ZrN. These results indicate that 

the hot oxygen starved environment remaining after the passage of the burn front is ideal for 

significant zirconium nitride formation.

A MELCOR model representing the test assembly was developed consisting of a stack of 10 

control volumes and 17 axial core levels in a single radial ring (12 axial levels along the heated 

length of the rods). This MELCOR model was used to simulate both the pre-ignition and ignition 

tests.

For the pre-ignition tests with the smaller pool rack cell size (Cell 1), the MELCOR model with 

the base SLAM and Σk parameters over predicted the induced air inlet mass flow rate, resulting in 

a slight under prediction of the assembly temperatures. To improve the results, the value of 

SLAM used was increased from 132 to 145 (at the top end of the uncertainty range). This 

resulted in the mass flow rate predictions falling within the experimental data uncertainty range 

for all power levels and improving the temperature response so that MELCOR is within 19K 

(34°F) of the measured test data. For the pre-ignition tests with the larger pool rack cell size 

(Cell 2), the MELCOR model with the base SLAM and Σk parameters shows excellent agreement 

for both induced air inlet mass flow rate and temperatures. MELCOR calculated temperatures 

are within 7K (13°F) of the experimental data for all power levels. The MELCOR computed 

mass flow rates were within the experimental data uncertainty range for all power levels. 

For the 5.0 kW ignition test, MELCOR shows excellent agreement for the peak cladding 

temperature in the assembly, and accurately captures the ignition time. Just prior to the ignition, 

MELCOR was over predicting the temperature throughout the assembly. This was traced to 

MELCOR over predicting the oxidation rate. Earlier in the transient, MELCOR was accurately 

predicting the peak cladding temperature in the assembly, however, had a tendency to over 

predict the temperature in the lower portion of the assembly. The MELCOR predicted mass 

flow rate shows excellent agreement to the experimental test data up until the time of ignition, 

after which time, the MELCOR mass flow rate shows significant oscillations. Overall, MELCOR 

is capable of accurately predicting the induced mass flow rate, peak cladding temperature and 

ignition time for a single assembly in a spent fuel pool under complete loss of coolant 

conditions. Additional MELCOR analyses were performed to investigate axial and radial 

nodalization. The addition of nodes in both the axial and radial directions did not result in 

changes to the overall results, implying that there are a sufficient number of nodes in the base 

case model.

xvii

ABBREVIATIONS

ANL Argonne National Laboratory

BSA burst signal analyzer

BWR Boiling Water Reactor

CFD computational fluid dynamics

COR MELCOR Core Package

CVH MELCOR Control Volume Hydrodynamics Package

DAQ data acquisition

DH hydraulic diameter

ID inside diameter or dimension

LDA laser Doppler anemometer

LOCA Loss of Coolant Accident

MELCOR severe accident analysis code

NEA Nuclear Energy Agency

NRC Nuclear Regulatory Commission

NS Non-supporting Structure

OECD Organization for Economic Co-operation and Development

PCT peak cladding temperature

PID Proportional-integral-derivative

PWR pressurized water reactor

RGA residual gas analyzer

SCR silicon-controlled rectifier

SFP spent fuel pool

SFPP Spent Fuel Pool Project

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

SNL Sandia National Laboratories

TC Thermocouple

1-1

1 INTRODUCTION

The testing in Phase I of the spent fuel pool (SFP) project focused on axial heating and burn 

propagation. The near-prototypic test assembly was constructed to represent a commercial 

17×17 PWR fuel bundle. The various components comprising a typical 17×17 PWR assembly 

are illustrated in Figure 1.1. The main structural component of the assembly is the core 

skeleton, which consists of eleven spacers permanently attached to twenty-five guide tubes. 

The 264 fuel rods pass through the spacers and are held captive in the assembly by the top and 

bottom nozzles.

Spacer

Bottom 

nozzle

Guide 

tubes Top 

nozzle

Spacer

Figure 1.1 Various components in a typical 17×17 PWR fuel assembly

The single full-length, electrically heated PWR assembly was fabricated using prototypic, 

commercial 17×17 PWR components and 9.50 mm (0.374 in.) heater rods made from 11.18 

mm (0.44 in.) Zircaloy-2 tubing supplied by an industrial vendor. The as-built heater rod 

diameter of 9.50 mm (0.374 in.) is slightly smaller than the typical PWR design value of 9.53 

mm (0.375 in.) but is not expected to significantly affect system hydraulics. The heater rods 

were manufactured by a commercial vendor using the same fuel rod simulator design and 

Zircaloy-2 material that was highly successful in the BWR study (Ref. 1). An oxidation kinetics

study was performed that showed no difference between the Zircaloy-2 used to make the 

heaters and Zircaloy-4 used in prototypic PWR fuel rods (Ref. 2). The spent fuel rod simulators 

for Phase I had a linear power profile and a maximum output of 15.8 W/m (4.8 W/ft), which was 

three times greater than expected to produce ignition.

Two storage cell sizes were fabricated for study in Phase I testing. Cell 1 had an inner 

dimension (ID) of 221.3 mm (8.7 in.) and was chosen to match the middle-sized cell in the PWR 

hydraulic characterization study (Ref. 3). Cell 2 had an inner dimension of 223.4 mm (8.8 in.) 

and was chosen to match the cell size in the pool rack to be used in Phase II testing. A series 

of eight pre-ignition tests were conducted for each cell to determine the thermal-hydraulic 

response of the mock-spent fuel assembly. In addition, MELCOR simulations were performed 

for each of these pre-ignition tests for both cells and the ignition test of Cell 2. MELCOR 

sensitivity analyses were also conducted to examine the effect of using more refined

discretization.

2-1

2 EXPERIMENTAL APPARATUS AND PROCEDURES

2.1 Flow Measurements and Calibrations

2.1.1 Assembly Hydraulics

A single pressure port was installed near the bottom of the storage cell to allow isothermal 

pressure drop measurements to determine the hydraulic loss coefficients of the apparatus with 

unheated rods. The location of this pressure port is detailed in Section 2.3.2. Figure 2.1 shows 

the overall assembly pressure drop of the Cell 1 (221.3 mm ID) configuration as a function of 

average air velocity (with the guide tubes plugged). The SLAM and Σk as determined from the 

quadratic fit of the data are 132 and 29.2, respectively. Figure 2.2 shows two simultaneous 

measurements of the overall assembly pressure drop for the Cell 2 (223.4 mm ID) configuration 

as a function of average air velocity. The SLAM and Σk, as determined from the quadratic fit of 

the data, are 146 and 24.8, respectively. Appendix A summarizes calculations and definitions 

for SLAM and Σk. Thermocouples were not installed during these forced flow hydraulic tests.

The dependence of the hydraulic parameters on the storage cell size differs from that 

determined during the PWR hydraulic characterization study (Ref. 3). The previous hydraulic 

characterization was conducted with a prototypically ideal PWR assembly. As the storage cell 

size increased, the SLAM decreased significantly and Σk decreased minimally with a net result 

that hydraulic resistance decreased. With the Phase I PWR assembly, the SLAM and Σk are 

comparable within the calculated uncertainties, as the cell size did not change significantly. 

Figures 2.1 and 2.2 show the measured pressure drop as a function of velocity based on 

separate measurements (ΔP1 and ΔP2). Comparing Figure 2.1 and Figure 2.2 at high velocity 

(1.6 m/s (5.25 ft/s)), the pressure drop with Cell 2 is only slightly lower than with Cell 1. At 0.8 

m/s (2.62 ft/s) and lower, there is little discernible difference between the pressure drop in Cell 1 

and Cell 2. The velocity range of interest for complete loss-of-coolant conditions is on the order 

of 0 to 0.25 m/s (0 to 0.82 ft/s).

The uncertainties in the hydraulic measurements are detailed in Appendix A, Error Propagation 

Analysis. Unless otherwise noted, the uncertainties quoted throughout this report refer to the 

expanded uncertainties, which give the 95 percent confidence level that the measurement lies 

within the cited bounds. The uncertainty in the average assembly velocity is ±0.078 m/s

(±0.256 ft/s). The uncertainty in the pressure measurement is ±0.1 N/m2 (±1.45·10-5 psi). The 

estimated uncertainty in the flow loss coefficients is ±13 and ±1.4 for SLAM and Σk, respectively. 

Table 2.1 presents a summary of the hydraulics for both Cell 1 and 2. The as-built, average 

diameter of the simulated fuel rods was 9.5 mm (0.374 in.). The presence of the thermocouples 

was considered in calculating the flow area and hydraulic diameter. The average number of 

thermocouples at any given height in the assembly is 60, each with an outer diameter of 0.79 

mm (0.031 in.). The decrease to the flow area with the inclusion of the thermocouples is 

approximately 0.1 percent of the non-instrumented flow area. These thermocouples were 

bundled into four groups that added an estimated 41 mm (1.6 in.) of wetted perimeter, or 0.4

percent of the total wetted area.

Major differences between the hydraulic assembly and the heated assembly are the presence of 

thermocouples and a slight randomly oriented curvature present in the heater rods in the heated 

assembly. Variation in the spacing between adjacent rods is evident in the velocity profile 

measurements discussed in Section 4.3. The heater rods to be used in the Phase II testing will 

be mechanically straightened in an effort to make them more ideally prototypic. 
2-2

Additional hydraulic characterization will be conducted on the center heated assembly as well 

as the unheated peripheral assemblies to examine the issues of rod straightness and presence 

of thermocouples.

Figure 2.1 Overall pressure drop as a function of average air velocity in the assembly for the 

221.3 mm storage cell (Cell 1)

Figure 2.2 Overall pressure drop as a function of average air velocity in the assembly for the 

223.4 mm storage cell (Cell 2)
2-3

Table 2.1 Summary of assembly hydraulics for storage Cell 1 and Cell 2

Description Cell 1 Cell 2

Inner Dimension (mm) 221.3 223.4

Flow Area (m2) 0.0273 0.0283

Hydraulic diameter, DH

(mm) 11.2 11.6

SLAM 132 146

Σk 29.2 24.8

2.1.2 Hot Wire Anemometer at the Assembly Inlet

A hot wire anemometer was placed in the 154 mm (6.065 in.) ID pipe that defines the inlet to the 

test assembly. A hot wire anemometer was chosen to measure the inlet flow rate because this 

type of instrument is sensitive and robust while introducing almost no unrecoverable pressure 

loss. The placement of the hot wire is shown in Figure 2.3. A TSI Model 8455 hot wire

anemometer was used for these tests. A honeycomb element was added to the inlet entrance 

to reduce the influence of any air flow disturbances within the experimental enclosure on the hot 

wire measurements.

Figure 2.3 Schematic showing the layout of the inlet hot wire

A series of unheated calibration runs were performed to calibrate the output of the hot wire 

anemometer. Air flows were metered into the bottom of the assembly via the inlet pipe and the 

response of the anemometer was recorded for each flow rate. A least-squares regression was 

performed for each pool cell configuration as shown in Figure 2.4. These linear coefficients 

were used to determine the indicated flow rate from the hot wire anemometer during heated 

testing. Although the data for the two pool cells give different calibration coefficients, the 

apparent flow rates from both calibration curves are well within the uncertainty of the 

measurement of ±12 slpm, especially for flow rates in the observed range of testing (< 250 

slpm).

914

305

A-A

Section A-A

TSI

Fixed 

location

54

154 ID 

(6.065 in.)

TSI Anemometer

12

Honeycomb element

13.5

All dimensions in mm
2-4

Figure 2.4 Flow rate as a function of the TSI Model 8455 hot wire anemometer voltage

2.1.3 Laser Doppler Anemometry Measurements

Laser Doppler anemometry (LDA) is a non-intrusive, optical technique used to measure the 

instantaneous velocity in a flow field at the intersection of two coherent laser beams. The most 

common method of LDA used at present is the dual-beam anemometer system. Typically, a 

single laser beam is split into two mutually coherent polarized light waves, which intersect to 

form a spheroid-shaped region called the measuring volume. Particles passing through the 

measuring volume with a given velocity scatter light from the light beams to produce the LDA 

signal. For the investigations detailed herein, a dual-beam system operating in backscatter 

mode as shown in Figure 2.5 was used. The photodetector was a photomultiplier (PM), and a 

Bragg cell was used to introduce a reference frequency, allowing the measurement of near-zero 

velocities. The signal was then processed in a burst analyzer and sent to a PC-based data 

acquisition system.

LDA measurements were initially planned for the Cell 1 configuration only. However, failure of 

the burst signal analyzer (BSA) during Cell 1 efforts forced these measurements to be collected 

during Cell 2 pre-ignition testing. The data collected are of the same nature for Cell 2 as it 

would have been for Cell 1 with only a slightly larger cell ID. This change to the original test 

plan was necessary to prevent further delay to the project and does not significantly affect the 

technical content.
2-5

Figure 2.5 Dual-beam backscatter LDA system components and principles

The average velocity profiles were measured in these experiments by a single-component laserDoppler anemometer (Dantec 1-D FlowExplorer). This LDA system is composed of the 

FlowExplorer probe head, a photomultiplier, a burst analyzer, a motorized stage and controller 

unit, a PC-based data acquisition (DAQ) system, and data processing software. These 

components are listed in Table 2.2.

Table 2.2 Detailed list of the LDA system components

Description Manufacturer Serial Number

FlowExplorer probe 

head

Dantec 

Dynamics 0115

BSA F60 – Burst 

analyzer

Dantec 

Dynamics 437

Photomultiplier Dantec 

Dynamics 119

1-D Stage Isel Automation 505

1-D Stage controller Isel Automation 502

DAQ – PC Dell F51KYD1

Figure 2.6 shows the layout of the test components for LDA measurements. The LDA probe is 

mounted externally to the PWR assembly on a motorized stage. The laser beams pass through 

the optical window into the assembly and measures the velocity at the intersection of the 

beams. In this manner, the local velocity can be measured across the assembly in between rod 

banks. All measurements in this report were collected between the rod banks identified in 

Figure 2.6 at z = 0.442 m (17.4 in.). Figure 2.7 gives three photographs of the LDA setup. 

These photographs depict a measurement just inside the optical window. The rubber coupling 

shown in Figure 2.7c was connected to the flow conditioner for pressure drop measurements, 

hot wire calibrations, and forced flow velocity profile baselines. Further details of the flow 

conditioner, flow metering system, and LDA may be found in the PWR hydraulic 

characterization study (Ref. 3).

t (measured)

Signal

Time

Measuring volume Backscattered light

Flow with particles

dfr (known)

Bragg Lase

cell

Processor

Photomultiplier
2-6

Figure 2.6 Schematic of the LDA system for measuring velocity profiles in the PWR 17×17 

assembly

Figure 2.7 Photographs showing the a) overall LDA setup, b) close-up of the probe head and 

optical window, and c) flow inlet and seed source

Direction of 1-D 

automatic stage travel

Optical window

Velocity 

profile along 

dashed line

z

y

x

LDA probe head

a)

b)

c)

6 in. rubber 

coupling

Flow 

conditioner / 

Seed source
2-7

2.2 Thermocouple Layout

Thermocouples (TCs) were installed throughout the assembly to measure thermal response 

during heat up. The TC layout was unchanged between the Cell 1 and the Cell 2

configurations. A total of 131 TCs were installed on the guide tubes and heater rods. The top 

11 TCs, representing those at z = 3.759 and 3.934 m (148 and 154.875 in.), exited the top of 

the assembly. The remainders of TCs were routed through the bottom of the assembly. All of 

Phase I testing utilized these internal TCs. Figure 2.8 shows the layout and identification 

scheme of these TCs. An additional 60 TCs were placed on the storage cell (30) and the

thermal radiation barrier (30). The thermocouples were placed 15.24 cm (6 in.) apart along the 

storage cell and thermal radiation barrier. These TCs were designated by their physical 

orientation, e.g. Pool_N_F_27” was a pool TC on the North face at an elevation of 0.686 m (27 

in.) and Rad_SE_C_52” was a radiation barrier TC on the Southeast corner at an elevation of 

1.321 m (52 in.). Other TCs were placed to measure temperatures of interest such as flow inlet 

and ambient temperature.

Figure 2.8 Thermocouple layout for the Phase I test assembly

Note: Thermocouples were installed using US customary units. The data file headers and TC 

identifiers give positioning in inches, e.g. J_9_124” is installed on the middle instrument tube at z

= 124 in. (3.150 m).

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

Axial Levels = 148, 100, 52, 4*

* - “A” TCs installed at 4.875

142, 94, 46 136, 88, 40 130, 83, 34

A B C D E F G H J K L M N O P Q R

A9, A13, A17, C9, C12,

D14, F9, F12, J9

D14, J15 J15, J17, M15, N17,

O14, R17

O14, P9, SYM D4

A B C D E F G H J K L M N O P Q R A B C D E F G H J K L M N O P Q R A B C D E F G H J K L M N O P Q R

Axial Levels = 124, 76, 27 118, 69, 22 110, 64, 16 154.875, 106, 58, 10

J9, M6, M9, O4, P6,

P9, R1, R5, R9

J3, O4 A1, D4, E1, F3,

J1, J3, Sym E17

D4, C9

A B C D E F G H J K L M N O P Q R A B C D E F G H J K L M N O P Q R A B C D E F G H J K L M N O P Q R A B C D E F G H J K L M N O P Q R

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

High-density axial array, 15.25 cm Medium-density axial array, 30.5 cm Low-density axial array, 61 cm 

Guide tube Fuel rod 

y

x
2-8

TCs were attached to the assembly by spot welding Nichrome ribbon over the thermocouple 

sheath as shown in Figure 2.9. An additional piece of Nichrome was attached a short distance 

away from the tip of the TC to provide strain relief. All TCs used in this testing were 

ungrounded, K-type with a 0.813 mm (0.032 in.) Super Omegaclad XL sheath diameter. The 

majority of the TCs used were Omega Engineering part number TJ192-CAXL-032U-192-

SMPW-M.

Figure 2.9 Detail view showing TC attachment to a fuel rod

2.3 General Construction

2.3.1 Power Control

Figure 2.10 shows the configuration of the power control system for the SFP Phase I test series. 

The data acquisition (DAQ) system generates a power set point based on user input via a 

LabVIEW graphical user interface. This set point signal is relayed to a proportional-integralderivative (PID) controller. The PID controller determines an appropriate power control signal 

by comparing the power set point to the feedback signal from the system Watt transducer. The 

silicon controlled rectifier (SCR) power controller receives the control signal and allows the 

prescribed electrical power into the resistive load of the test assembly.
2-9

Figure 2.10 Power control system and test circuit

The internal layout of the instrumentation panel is shown in Figure 2.11. The transducers 

measure power, voltage, and current applied to the assembly. The signal from the Watt 

transducer is output to the DAQ and the PID controller as a feedback signal. The components 

used in this testing are listed in Table 2.3.

Figure 2.11 Schematic of the instrumentation panel

120 VAC SCR

Source

Test Assembly

~15 kW @ 120 VAC

Power Set Point

Instrumentation 

Signals

Current Signal

(5:200)

Instrumentation

Panel

8

EZ

PID

Voltage 

Signal

Power 

Feedback

Signal

Power Control Signal

Neutral

AC 

Power

DAQ System

Watt

Transducer

Voltage

Transducer

Voltage 

Signal

Current

Signal

Neutral

Signals 

to DAQ

Power 

Feedback 

Signal

Current

Transducer
2-10

Table 2.3 List of equipment used for power control

Description Manufacturer Model

AC Watt Transducer Ohio Semitronics PC5-001D

AC Voltage Transducer Ohio Semitronics 3VTR-001D

AC Current Transducer Ohio Semitronics 3CTR-010D

PID Controller Watlow Electric Manufacturing PM6C1FJ1RAAAAA

SCR Power Controller Watlow Electric Manufacturing PC91-F25A-1000

2.3.2 Instrumentation External to the Assembly

Figure 2.12 shows the location of external instrumentation. All dimensions are referenced from 

the bottom of the bottom nozzle (z = 0). A single pressure port was located at z = 0.011 m 

(0.43 in.) to calculate the overall assembly pressure drop using atmospheric pressure as the 

reference pressure. These pressure drops were used to determine the SLAM and Σk of the asbuilt assembly. The laser-Doppler anemometer (LDA) was located mid-bundle between the 

bottom nozzle/debris catcher and the first spacer. Also, quartz light pipes were placed 

approximately every 0.3048 m (1 ft) for visually observing the burn front. Three sample tubes 

were placed at the exit of the assembly. An exhaust air temperature was measured using a TC 

inserted into one of the tubes with the exit air drawn across the TC junction. The other bundle 

and annulus sample tubes supplied the two oxygen concentration transmitters (Advanced Micro 

Instruments, Model 65, Part 6ANA0056). An additional slip stream from the bundle sample tube 

was diverted to the residual gas analyzer (RGA) to measure relative concentrations of oxygen 

and nitrogen to the argon present in the assembly exhaust.

A schematic of the sampling system is provided in Figure 2.13. Two vacuum pumps were used 

to draw the samples from the top of the assembly. The main vacuum pump was used to pull 

relatively high volumetric flow rates through the sample tubes in order to reduce residence time 

of the samples. The secondary vacuum pump was used to pull slip stream samples for the 

oxygen concentration transmitters and the RGA. The RGA had a resident vacuum system 

internal to the instrument to provide a final sample stream.
2-11

Figure 2.12 PWR test assembly and external instrumentation

LDA

Optical 

windows

Beam dump

0.442

0.305 – 13 places (Light pipes 

for external burn front tracking)

0.011

Pressure port for overall 

assembly ∆P 

t

z 0.616

All dimensions in meters Exhaust TC

Bundle O2

Annulus O2

0.222
2-12

O2 Annulus

O2 Bundle

Exhaust TC

RGA

1.1 cfm

Secondary

Vacuum 

Pump

1.1cfm

Main Vacuum 

O Pump 2 sensors

1⁄4 in. Cu tubing

Filters

Vent to Vessel

Vent to Vessel

Vacuum Gauge

O2 Vessel

.1-1 scfh

Rotameter

.1-1 scfh

Rotameter

.2-2.5 scfh

Rotameter

Figure 2.13 Schematic of the assembly exhaust sampling system

The amounts of nitrogen, oxygen, and argon in the exhaust stream directly above the test 

bundle were measured with a Pfeiffer RGA (Model GSD 300T). This instrument employed a 

heated quartz capillary for sampling air at ambient pressures (83 kPa (12 psi)). The standard 

instrument configuration contains a tungsten filament for ionizing the sample and a quadruple

mass spectrometer for detection. A single ion (amu 28, 32, 40) was used to monitor for each 

gas. Data was recorded at 10 second intervals over the duration of the test. Before the test 

started, the instrument was tuned to reduce signal drift. Data were recorded in terms of detector 

current and then converted to a mole or volume fraction based on the analysis of ambient air, 

which served as a single point calibration gas. After initial evaluation of the ignition test data, 

two additional calibration gases were analyzed using the RGA. The first calibration gas was 

1.22 percent argon in nitrogen. The other calibration gas was 4.76 percent argon in nitrogen. 

These additional calibration gases spanned ranges of interest as determined during post-test 

analyses of the RGA data.

The detector current for each of the three major gas species was normalized by the sum of the 

three detector currents. Since the three gas species ionized with different efficiency, correction

factors are required to adjust the current ratios into volume fraction of each of the gas species. 

Assuming ambient air is 0.78084 N2, 0.20946, O2 and 0.00934 Ar. The correction factors 

determined on the day of the ignition test were 0.957, 1.153, and 2.401, respectively. 

Samples were drawn through ceramic tubes at the top of the assembly at two independent 

locations, one from the center of the bundle just above the electrical bus plate and the other 

from the annular region between the storage cell and the outer perimeter of the bundle just 

below the electrical bus plate. The bus plate was located 0.173 m (6.82 in.) below the top of the 

storage cell. The sample for the RGA analysis was initially configured to be drawn from the 

center bundle location. Later, an additional sampling point was added at the annulus.
3-1

3 CELL 1 PRE-IGNITION TESTING AND ANALYSIS

The tests were performed using a uniform power profile with electrically heated rods. The 

pre-ignition tests were performed for seven different power levels, varying from 0.5 kW to 

3.5 kW. The power was maintained constant during the tests. All tests were performed for a 

period of 12 hours, except the two 1 kW tests, which were conducted for a period of 20 hours.

A 0.1524 m (6 in.) layer of Kaowool insulation was installed around the cell and wrapped with a 

0.9 mm (0.036 in.) thick sheet of stainless steel. In the pre-ignition tests, the assembly was 

cooled by natural circulation. During the tests, the pressure and air inlet temperature were 

maintained at the ambient conditions of Albuquerque, NM.

A summary of the pre-ignition testing results for Cell 1 are presented in Table 3.1. This table 

gives the average temperatures and flow rates at 3, 6, 9, and 12 hours for z = 3.150 m (124 in.). 

The minimum and maximum temperatures of the bundle at each time are also shown.

The 1.0 kW test was run twice, as shown in the second and third columns of test entries in

Table 3.1, to examine repeatability. The temperatures were within 3 K (5.4°F) at all positions 

and times, which is slightly greater than the experimental uncertainty of ±2.2 K (±4.0°F). The air 

flow rates were within 12 slpm for the two measurements, which is equal to the experimental 

uncertainty of ± 12 slpm.

Table 3.1 Summary of pre-ignition testing results for Cell 1 at z = 3.150 m (124 in.)

True RMS Power (W) 500 1000 1000 1500 2000 2500 3000 3500

Test Date 10/27/2010 10/12/2010 11/08/2010 10/22/2010 10/21/2010 10/20/2010 10/19/2010 10/18/2010

Flow Rate (slpm) 56 116 128 145 164 175 176 184

Min. Temp. (K) 320 343 344 365 392 411 438 455

Max. Temp. (K) 327 359 359 388 422 445 479 503

Avg. Temp. (K) 324 352 353 379 409 430 461 482

Pool Cell Avg. Temp. (K) 316 336 336 356 380 398 425 437

Rad. Barrier Avg. Temp. (K) 302 301 296 297 298 298 301 301

Ambient Avg. Temp. (K) 301 300 295 295 296 296 297 298

Flow Rate (slpm) 114 155 156 174 183 183 184 181

Min. Temp. (K) 343 386 388 429 474 513 559 594

Max. Temp. (K) 354 410 410 459 510 552 600 645

Avg. Temp. (K) 349 399 400 445 493 534 581 621

Pool Cell Avg. Temp. (K) 339 379 380 421 465 503 550 580

Rad. Barrier Avg. Temp. (K) 300 302 299 300 304 306 310 312

Ambient Avg. Temp. (K) 298 299 296 295 298 298 299 300

Flow Rate (slpm) 138 172 165 181 182 182 176 170

Min. Temp. (K) 363 425 426 483 545 599 661 718

Max. Temp. (K) 377 452 450 515 579 636 698 756

Avg. Temp. (K) 371 440 439 500 563 619 682 739

Pool Cell Avg. Temp. (K) 359 418 420 477 538 593 656 708

Rad. Barrier Avg. Temp. (K) 300 304 303 304 309 316 321 324

Ambient Avg. Temp. (K) 297 299 297 295 297 300 300 300

Flow Rate (slpm) 149 175 165 181 179 173 168 158

Min. Temp. (K) 381 458 457 526 598 665 738 820

Max. Temp. (K) 396 487 484 559 634 702 775 850

Avg. Temp. (K) 390 474 472 544 617 685 759 838

Pool Cell Avg. Temp. (K) 377 451 453 522 594 662 736 809

Rad. Barrier Avg. Temp. (K) 301 307 307 309 315 323 331 338

Ambient Avg. Temp. (K) 296 299 298 297 297 300 301 300

Bundle

Bundle

3 hrs

6 hrs

Bundle

Bundle

9 hrs

12 hrs
3-2

A MELCOR input model was developed containing 10 control volumes and 17 axial core levels 

(12 heated) in a single radial ring. Details about the MELCOR input model can be found in 

Appendix C. MELCOR version 1.86.yv.3084 was used for the analyses.

The experimental data of Cell 1 was compared with two MELCOR models using the values of 

132 for SLAM (as shown in Table 2.1) and a SLAM of 145 (which includes the uncertainty of +13). 

A value of 30 for Σk was used in both cases. Figure 3.1 shows the comparison of temperature 

for the pre-ignition tests at t = 12 hrs. The model agrees with the measured temperatures within 

19 K (34°F) for the SLAM of 145 and within 35 K (63°F) for the SLAM of 132 case for all powers.

Figure 3.2 shows the comparison of the airflow rates for the pre-ignition tests and MELCOR at 

t = 12 hrs. The model agrees with the measured flow rate to within 8 slpm for the SLAM of 145

and 30 slpm for the SLAM of 132 case for all powers. This flow rate includes the air flow through 

the bundle and the annulus. The increased MELCOR flow rate leads directly to the model 

under predicting the assembly temperature. As shown in Figures 3.1 and 3.2, temperature and 

mass flow predictions for a SLAM value of 145 agrees favorably with experimental data as 

compared to the predictions using SLAM of 132. As such, a SLAM value of 145 is recommended 

for Cell 1.

350

400

450

500

550

600

650

700

750

800

850

0 500 1000 1500 2000 2500 3000 3500 4000

Temperature (K)

Power (W)

Test - Avg

MELCOR, SLam 145

MELCOR, SLam 132

Figure 3.1 Comparison of the average test (blue diamonds) and MELCOR (pink) temperatures 

during pre-ignition testing for Cell 1 at z = 3.150 m (124 in.) and t = 12 hrs
3-3

100

125

150

175

200

225

250

0 500 1000 1500 2000 2500 3000 3500 4000

Flow Rate (slpm)

Power (W)

Test - Avg

MELCOR, SLam 145

MELCOR, SLam 132

Figure 3.2 Comparison of the test (blue diamonds) and MELCOR (pink) flow rates during preignition testing for Cell 1 at t = 12 hrs

While eight pre-ignition tests representing seven simulated decay powers were conducted for 

the Phase I – Cell 1 configuration, only powers of 1.0 and 3.5 kW are discussed in more detail in 

the following sections of the report. These tests were chosen because they bracket the 

experimental results. The 1.0 kW test was selected over the 0.5 kW test because the 1.0 kW 

test was conducted to 20 hours duration and represents the closest test result to steady state 

achieved in these studies.

3.1 Cell 1 – 1.0 kW Results

The peak, average test temperatures used to compare with MELCOR were calculated using the 

logic shown in Equation 3.1. The average data were determined primarily at the radial arrays 

with nine TCs, z = 4, 27, 52, 76, 100, 124, and 148 in. levels. In order to gain more axial 

information, the average values at the z = 124 and 148 in. levels were used in conjunction with 

the limited TC data at z = 130, 136, 142, and 154.875 in. levels to extrapolate average 

temperatures at these intermediate levels.

avg ( ( ( ))) =const. T ( ) max avg T , for = 4, 27, 52, 76, 100, 124, 130, 136, 

 142, 148 and 154.875 in. levels z t = zt z 3.1

Figure 3.3 shows the Cell 1 bundle temperatures at 1.0 kW simulated decay power for the peak, 

average test and MELCOR temperatures. The experiment and MELCOR are within 5 K (9°F)

for the SLAM of 145 and within 6 K (11°F) for the SLAM of 132 case at t = 12 hrs. MELCOR under 

predicts the peak cladding temperature (PCT) by 13 K (23°F) for the SLAM of 145 case and by 

30 K (54°F) for the SLAM of 132 case at t = 20 hrs.
3-4

300

325

350

375

400

425

450

475

500

525

550

0 2 4 6 8 10 12 14 16 18 20

Temperature (K)

t (hrs)

MELCOR, SLam 145

MELCOR, SLam 132

Test-Avg

Figure 3.3 Cell 1 bundle temperatures as a function of time at 1.0 kW simulated decay power 

for the average test temperature (blue diamonds) and MELCOR (pink)

Figure 3.4 compares the transient induced flow rate of the Cell 1 configuration with an applied 

test power of 1.0 kW for the test and MELCOR results. The experimental enclosure was 

opened at approximately 13 hours elapsed test time to allow free access to both hot wire and 

laser Doppler anemometers. Traverses with an additional hot wire anemometer, which are not 

shown in this report, were conducted to measure the velocity profile inside the inlet pipe. In 

addition, alignment of the LDA and insertion of neutral density filters as required for operations

were performed during this time period. MELCOR over predicts the observed flow data for a 

SLAM of 132 and match favorably for a SLAM of 145 case. For that reason, a SLAM value of 145 is 

recommended for Cell 1.
3-5

0

25

50

75

100

125

150

175

200

225

0 2 4 6 8 10 12 14 16 18 20 22

Flow Rate (slpm)

Time (hrs)

Test

MELCOR, SLam 145

MELCOR, SLam 132

Figure 3.4 Induced flow rate within the Cell 1 assembly at a simulated decay power of 1.0 kW 

for the test (blue solid) and MELCOR (pink)

Figure 3.5 compares the Cell 1 maximum, average, and minimum test temperatures in the 

bundle as a function of axial height with MELCOR values at 12 hours elapsed time. MELCOR 

under predicts the average temperature in the top part of the assembly but over predicts the 

temperature in the lower part. The axial location of the MELCOR calculated PCT is accurate to 

within the discretization of the MELCOR model.

Similar to Figure 3.5, Figure 3.6 shows the test and MELCOR temperatures at t = 20 hours. 

Again, MELCOR over predicts the average temperature of the assembly in the lower part and 

under predicts the temperature in the upper part. The location of PCT is predicted to within less 

than one node in the MELCOR results for both cases. Figures 3.5 and 3.6 show that an SLAM of 

145 predicted the axial temperature profile more accurately than an SLAM of 132. Figures 3.4 

and 3.5 attest further that SLAM of 145 for Cell 1 is more appropriate.
3-6

300

350

400

450

500

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Avg

Test - Min

MELCOR, Slam 145

MELCOR, SLam 132

Figure 3.5 Cell 1 bundle temperatures as a function of axial height in the assembly at 1.0 kW 

simulated decay power and 12 hours elapsed for the test maximum (red squares), 

average (blue diamonds), minimum (green triangles), and MELCOR temperatures 

(pink)

300

350

400

450

500

550

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Avg

Test - Min

MELCOR, Slam 145

MELCOR, SLam 132

Figure 3.6 Cell 1 bundle temperatures as a function of axial height in the assembly at 1.0 kW 

simulated decay power and 20 hours elapsed for the test maximum (red squares), 

average (blue diamonds), minimum (green triangles), and MELCOR temperatures 

(pink)
3-7

3.2 Cell 1 – 3.5 kW Results

Figure 3.7 shows the Cell 1 maximum bundle temperatures at 3.5 kW heater power for the 

peak, average test and MELCOR temperatures. Experiment and MELCOR peak cladding 

temperatures agree to within 40 K (72°F) for SLAM of 132 and within 30 K (54°F) for SLAM of 145 

case at all times.

300

350

400

450

500

550

600

650

700

750

800

850

900

0 1 2 3 4 5 6 7 8 9 10 11 12

Temperature (K)

Time (hrs)

Test-Avg

MELCOR, SLam 145

MELCOR, SLam 132

Figure 3.7 Cell 1 bundle temperatures as a function of time at 3.5 kW simulated decay power 

for the average test temperature (blue diamonds) and MELCOR (pink)

Figure 3.8 compares the transient, induced flow rate through the assembly in the Cell 1 

configuration for the 3.5 kW test and MELCOR simulation. Disturbances to the flow rates 

between 2 and 3 hours were likely due to personnel entry into the vessel to diagnose lack of 

signal from auxiliary instrumentation. MELCOR over predicts the flow rate for a SLAM value of 

132 case match favorably the data for a SLAM of 145 case for times greater than t > 3 hrs. For 

that reason a SLAM value of 145 is also recommended for the 3.5 kW test of Cell 1.

Figure 3.9 compares the Cell 1 maximum, average and minimum test temperatures in the 

bundle as a function of axial height with MELCOR at 12 hours elapsed time. MELCOR under 

predicts the average temperature in the upper part of the assembly but over predicts the 

temperature in the lower part in both cases. The axial location of the MELCOR predicted PCT 

is accurate to within the nodalization of MELCOR.

The pre-ignition analysis for Cell 1 demonstrated that MELCOR accurately modeled the flow 

rate and peak cladding temperature using a Σk value of 30 and a SLAM value of 145.
3-8

0

25

50

75

100

125

150

175

200

225

250

0 1 2 3 4 5 6 7 8 9 10 11 12

Flow Rate (slpm)

Time (hrs)

Test

MELCOR, Slam 145

MELCOR, Slam 132

Figure 3.8 Induced flow rate within the Cell 1 assembly at a simulated decay power of 3.5 kW 

for the test (blue solid) and MELCOR (pink)

300

400

500

600

700

800

900

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Min

Test - Avg

MELCOR, SLam 145

MELCOR, SLam 132

Figure 3.9 Bundle temperatures as a function of axial height in the assembly at 3.5 kW 

simulated decay power and 12 hours elapsed for the maximum test temperature 

(red squares), average test temperature (blue diamonds), and MELCOR (pink)
3-9

Figure 3.10 shows contours of temperature within the test apparatus at the midplane for the 

3.5 kW test at t = 3, 6, 9, and 12 hours. The data represented in this plot reflect all available 

thermocouple data at the midplane of the apparatus with interpolation in between measurement 

locations. The midplane of the apparatus is defined at y = 0 m (see lower right assembly crosssection in Figure 2.8). A graphic of the fuel assembly and the surrounding insulation have been 

overlaid onto the plot. The insulation is denoted with diagonal hatch marks. Temperature data 

is not available for z > 3.934 m (154.875 in.) and z < 0.102 m (4 in.). The PCT was observed at 

z = 3.454 m (136 in.), which is also shown in Figure 3.7 The duration of this test was 12 hours.

Figure 3.10 Contours of temperature within the assembly at the midplane for an input power of 

3.5 kW

(t = 3 hrs) (t = 6 hrs) (t = 9 hrs) (t = 12 
3-10

Figure 3.11 shows contours of temperature within the test apparatus at z = 3.150 m (124 in.) 

and t = 3, 6, 9, and 12 hours for the 3.5 kW test. Similar to Figure 3.10, this plot reflects all 

available thermocouple data at z = 3.150 m (124 in.) of the apparatus with interpolation in 

between measurement locations. A cross-sectional graphic of the fuel assembly and the 

surrounding insulation have been overlaid onto the plot. The insulation is denoted with diagonal 

hatch marks. The plot at t = 12 hours clearly demonstrates the transverse thermal gradient in 

the assembly. The temperature difference across the assembly is approximately 30 K (54°F)

and is also documented in Table 3.1. 

Figure 3.11 Contours of temperature within the assembly at z = 3.150 m (124 in.) level for an 

input power of 3.5 kW

t = 3 hrs t = 6 hrs

t = 9 hrs t = 12 

h
4-1

4 CELL 2 PRE-IGNITION TESTING AND ANALYSIS

A summary of the pre-ignition testing results for Cell 2 are presented in Table 4.1. This table 

gives the average temperatures and flow rates at 3, 6, 9, and 12 hours for z = 3.150 m (124 in.). 

The minimum and maximum temperatures of the bundle at each time are also shown.

The 1.0 kW test was repeated to examine repeatability as shown in the second and third 

columns of test entries in Table 4.1. The temperatures during these repeatability tests were 

within 3 K (5.4°F) at all positions and times, slightly exceeding the experimental uncertainty of 

±2.2 K (±4.0°F). The air flow rates were within 9 slpm for the two measurements, which is 

better than the experimental uncertainty of ±12 slpm.

Table 4.1 Summary of pre-ignition testing results for Cell 2 at z = 3.150 m (124 in.)

True RMS Power (W) 500 1000 1000 1500 2000 2500 3000 3500

Test Date 2/21/2011 12/07/2010 02/16/2011 01/26/2011 01/12/2011 01/11/2011 01/06/2011 01/05/2011

Flow Rate (slpm) 97 123 114 167 171 186 193 204

Min. Temp. (K) 322 345 347 371 386 415 428 457

Max. Temp. (K) 330 361 361 393 413 449 467 504

Avg. Temp. (K) 327 354 355 383 402 434 450 482

Pool Cell Avg. Temp. (K) 317 337 340 359 372 400 412 441

Rad. Barrier Avg. Temp. (K) 298 299 300 299 304 303 304 301

Ambient Avg. Temp. (K) 297 298 299 298 302 300 301 299

Flow Rate (slpm) 128 182 175 197 209 209 215 211

Min. Temp. (K) 344 387 390 432 468 514 549 594

Max. Temp. (K) 356 411 412 463 504 555 591 640

Avg. Temp. (K) 351 400 401 448 486 535 571 619

Pool Cell Avg. Temp. (K) 339 380 382 423 457 504 537 585

Rad. Barrier Avg. Temp. (K) 299 298 301 302 306 310 305 308

Ambient Avg. Temp. (K) 298 295 299 298 300 302 299 298

Flow Rate (slpm) 146 189 188 208 213 206 205 197

Min. Temp. (K) 364 424 427 484 537 598 647 706

Max. Temp. (K) 378 450 452 517 573 637 686 748

Avg. Temp. (K) 372 437 440 501 556 618 668 729

Pool Cell Avg. Temp. (K) 359 417 420 477 529 590 639 700

Rad. Barrier Avg. Temp. (K) 301 301 303 306 310 317 316 320

Ambient Avg. Temp. (K) 298 295 297 298 299 301 299 298

Flow Rate (slpm) 154 193 197 205 214 200 201 194

Min. Temp. (K) 382 451 454 519 583 652 712 781

Max. Temp. (K) 398 482 484 557 625 695 755 823

Avg. Temp. (K) 390 467 469 538 604 675 734 804

Pool Cell Avg. Temp. (K) 376 446 449 513 578 648 708 778

Rad. Barrier Avg. Temp. (K) 302 303 306 309 314 323 325 330

Ambient Avg. Temp. (K) 297 296 298 297 298 302 299 298

Bundle

Bundle

3 hrs

6 hrs

9 hrs

12 hrs

Bundle

Bundle

On Table 2.1, the value for SLAM is 146 for Cell 2. However, based on fundamentals of fluid 

mechanics, this value should be lower than the value for Cell 1 because the flow area of Cell 2 

is larger. For that reason, the experimental data of Cell 2 was compared with two MELCOR 

models using the values of 146 for SLAM as shown in Table 2.1 and a SLAM of 133 including the 

uncertainty of -13. On the other hand, on Table 2.1, the value of Σk is 25 for Cell 2. Based on 

fluid mechanics fundamentals this value should be higher because the flow experiences a larger 

expansion after spacer grids. After comparing MELCOR results using the values of 25 and 30 

for Σk (Figures 4.1 and 4.2), no significant differences were observed and for that reason a 

value of 30 for Σk was used for Cell 2 MELCOR analysis. As observed in Figure 4.1, the test 

and MELCOR model results agree within 16 K (29°F) for a SLAM of 133 and within 7 K (13°F) for 
4-2

a SLAM of 146 at all powers. Similarly, Figure 4.2 shows the comparison of the airflow rates for 

the pre-ignition tests and MELCOR at t = 12 hrs. The model agrees with the measured flow rate

within 20 slpm for a SLAM of 133 and within 10 slpm for a SLAM of 146 for all powers.

350

400

450

500

550

600

650

700

750

800

850

0 500 1000 1500 2000 2500 3000 3500 4000

Temperature (K)

Power (W)

Test - Avg

MELCOR, SLam 146, Σk 25

MELCOR, SLam 146, Σk 30

MELCOR, SLam 133, Σk 30

Figure 4.1 Comparison of the average test (blue diamonds) and MELCOR (pink) temperatures 

during pre-ignition testing for Cell 2 at z = 3.150 m (124 in.) and t = 12 hrs
4-3

100

125

150

175

200

225

250

0 500 1000 1500 2000 2500 3000 3500 4000

Flow Rate (slpm)

Power (W)

Test - Avg

MELCOR, SLam 146, Σk 25

MELCOR, SLam 146, Σk 30

MELCOR, SLam 133, Σk 30

Figure 4.2 Comparison of the test (blue diamonds) and MELCOR (pink) flow rates during preignition testing for Cell 2 at t = 12 hrs

While eight pre-ignition tests representing seven simulated decay powers were conducted for 

the Phase I – Cell 2 configuration, only powers of 1.0 and 3.5 kW will be discussed in more 

detail in the following sections of the report. These tests were chosen because they bracket the 

experimental results. The 1.0 kW test was selected over the 0.5 kW test because the 1.0 kW 

test was conducted for 20 hours and represented the closest test result similar to a steady state 

condition achieved in these studies.

4.1 Cell 2 – 1.0 kW Results

Figure 4.3 shows the Cell 2 maximum bundle temperatures at 1.0 kW simulated decay power 

for the average test and MELCOR temperatures. The average experiment and MELCOR peak 

cladding temperatures agree to within 20 K (36°F) for a SLAM of 133 and within 10 K (18°F) for a 

SLAM of 146.
4-4

300

325

350

375

400

425

450

475

500

525

0 2 4 6 8 10 12 14 16 18 20

Temperature (K)

t (hrs)

MELCOR, SLam 146

MELCOR, SLam 133

Test-Avg

Figure 4.3 Cell 2 bundle temperatures as a function of time at 1.0 kW simulated decay power 

for the average test temperature (blue diamonds) and MELCOR (pink)

Figure 4.4 compares the Cell 2 maximum, average, and minimum 1.0 kW test temperatures in 

the bundle as a function of axial height with the MELCOR values at 12 hours. The test values 

are those defined in the maximum operator on the right hand side of Equation 3.1 for 

t = 12 hours. MELCOR under predicted the average temperature in the top part of the 

assembly but over predicted the temperature in the lower part. The axial location of the 

MELCOR predicted PCT is accurate to within the discretization of the MELCOR model. Similar 

to Figure 4.4, Figure 4.5 shows the test and MELCOR temperatures at t = 20 hours. Again, 

MELCOR over predicts the average temperature of the assembly in the lower part and under 

predicts the temperature in the upper part. The location of PCT is predicted to within less than 

one node in the MELCOR results for both cases. Figures 4.4 and 4.5 show that an SLAM of 146

predicted the axial temperature profile more accurately than an SLAM of 133.
4-5

300

350

400

450

500

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Avg

Test - Min

MELCOR, Slam 146

MELCOR, SLam 133

Figure 4.4 Cell 2 bundle temperatures as a function of axial height in the assembly at 1.0 kW 

simulated decay power and 12 hours elapsed for the maximum test (red squares), 

average test (blue diamonds), minimum test (green triangles) and MELCOR (pink)

300

350

400

450

500

550

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Avg

Test - Min

MELCOR, Slam 146

MELCOR, SLam 133

Figure 4.5 Cell 2 bundle temperatures as a function of axial height in the assembly at 1.0 kW 

simulated decay power and 20 hours elapsed for the maximum test (red squares), 

average test (blue diamonds), minimum test (green triangles) and MELCOR (pink)
4-6

Figure 4.6 compares the transient induced flow rate of the Cell 2 configuration with a power of 

1.0 kW for the test and MELCOR results. The experimental enclosure was opened at 

approximately 13 hours elapsed test time to allow free access to the laser Doppler anemometer. 

The sharp increase and subsequent decrease at 13 hours was due to the experimental 

enclosure equilibrating to the building temperature. The data and simulation agree within the 

experimental uncertainty of 12 slpm at all times in both cases for the MELCOR simulation using 

SLAM of 146, with the exception of the disturbance at t = 13 hours as previously described. 

MELCOR predictions using SLAM of 133 over predicted the mass flow rare as shown in Figure 

4.6.

0

25

50

75

100

125

150

175

200

225

0 2 4 6 8 10 12 14 16 18 20 22

Flow Rate (slpm)

Time (hrs)

Test

MELCOR, SLam 146

MELCOR, SLam 133

Figure 4.6 Induced flow rate within the Cell 2 assembly at a simulated decay power of 1.0 kW 

for the test (blue solid) and MELCOR (pink)

4.2 Cell 2 – 3.5 kW Results

Figure 4.7 shows the Cell 2 bundle temperatures at 3.5 kW heater power for the average test 

and MELCOR temperatures. Experiment and MELCOR peak cladding temperatures agree to 

within 28 K (50°F) for an SLAM of 133 and within 18 K (32°F) for an SLAM of 146 at all times.

Figure 4.8 compares the Cell 2 maximum, average, and minimum 3.5 kW test temperatures in 

the bundle as a function of axial height with two MELCOR simulations at 12 hours elapsed time. 

The test values are those defined in the maximum operator on the right hand side of Equation 

3.1 for t = 12 hours. Again, MELCOR simulations over predicted the average temperature in the 

lower part of the bundle and under predicted the temperature in the upper part of the bundle for 

both cases. The axial location of the MELCOR PCT is accurate to within the nodalization of the 

model.
4-7

300

350

400

450

500

550

600

650

700

750

800

850

900

0 1 2 3 4 5 6 7 8 9 10 11 12

Temperature (K)

Time (hrs)

Test-Avg

MELCOR, SLam 146

MELCOR, SLam 133

Figure 4.7 Cell 2 bundle temperatures as a function of time at 3.5 kW simulated decay power 

for the average test (blue diamonds) and MELCOR (pink)

300

400

500

600

700

800

900

0 0.5 1 1.5 2 2.5 3 3.5 4

Temperature (K)

z (m)

Test - Max

Test - Min

Test - Avg

MELCOR, SLam 146

MELCOR, SLam 133

Figure 4.8 Bundle temperatures as a function of axial height in the assembly at 3.5 kW 

simulated decay power and 12 hours elapsed for the maximum test (red squares), 

average test (blue diamonds), minimum test (green triangles), and MELCOR (pink)
4-8

Figure 4.9 compares the transient induced flow rate through the assembly in the Cell 2

configuration for the 3.5 kW test and MELCOR simulations. MELCOR predicted the mass flow 

rate favorably within the experimental error of 12 slpm for SLAM values between 133 and 145.

0

25

50

75

100

125

150

175

200

225

250

275

300

0 1 2 3 4 5 6 7 8 9 10 11 12

Flow Rate (slpm)

Time (hrs)

Test

MELCOR, Slam 146

MELCOR, Slam 133

Figure 4.9 Induced flow rate within the Cell 2 assembly at a simulated decay power of 3.5 kW 

for the test (blue solid) and MELCOR (pink)

4.3 Internal Assembly Mass Flow

The behavior of the internal flow in the assembly was explored with a laser Doppler 

anemometer (LDA) and with computational fluid dynamics (CFD). These measurements and 

simulations were for both forced (unheated) and natural circulation (heated) cases. In order to 

account for differences in density for the heated cases, the results in this section are presented 

as mass flux profiles and mass flow fractions.

A series of average velocity profiles were collected with the LDA along the x-axis in the 

assembly at z = 0.442 m (1.45 ft) and y = -0.0189 m (-0.74 in.) (See Figure 2.6) for both forced 

and naturally induced flows. These profiles were measured between the inside of the optical 

window to slightly beyond the assembly center. The forced flow data were collected to establish 

a baseline with which to compare the free convection generated velocity profiles. Three forced 

flows of 150, 186, and 225 slpm were chosen because they bracketed the expected flow rates 

for the heated tests. Velocity profiles were measured for heated tests at applied powers of 0.5

and 1 kW. These naturally induced flow profiles were collected after the assembly had been 

under power for 12 hours. The change in the measured inlet flow rate between 12 and 20 hours 

was less than 3 percent of the flow rate at 12 hours for both powered tests. The local air density 

in the assembly was estimated from the interpolated cladding and pool cell temperatures along 

the x-axis at z = 0.442 m (1.45 ft) and y = -0.0189 m (-0.74 in.).
4-9

The three-dimensional CFD simulations were performed on a 1/8th wedge of the PWR fuel 

assembly, taking advantage of symmetry conditions in the cross-sectional geometry. ANSYS 

Fluent version 14.0.0 was used for all calculations. The spacers and nozzles were explicitly 

modeled as flow restrictions with simplified geometric representations, i.e. sheet metal 

construction without turning vanes. The problem solution was obtained using a steady-state, 

laminar solver. The inlet was defined with a uniform velocity condition for forced flow cases and 

with a pressure boundary condition for naturally induced flow cases. The outlet was defined 

with a pressure boundary condition for all cases. Thermal radiation was modeled by the 

discrete ordinates method. The momentum, energy, and discrete ordinates equations were 

solved with 2nd order up-wind discretization scheme. The continuity equation was linked to the 

momentum equation through the Pressure Staggering Option.

4.3.1 Forced Flow Calibrations

Figure 4.10 shows the three normalized mass flux profiles collected during the forced flow 

baselines. The force flow calibration used an unheated assembly. This plot also shows two 

profiles produced from computational fluid dynamics (CFD) simulations. The local mass flux, 

density multiplied by velocity, is normalized by the uniform, average bundle velocity and inlet air 

density. The inlet air density for both the LDA and CFD normalizations is defined as ρo = 0.986 

kg/m3 (0.062 lb/ft3

) at To = 297 K (75°F) and Po = 84 kPa (12.2 psi). The x-dimension is 

normalized by the fuel rod pitch, P. Again, all profiles were collected between the rod banks 

identified in Figure 2.6 at z = 0.442 m (17.4 in.). The three normalized profiles exhibit nearly 

identical behavior as might be expected. Significant variations occur at locations where 

measurements were more difficult due to the presence of a solid surface such as a fuel pin or 

the optical window. Alignment of the LDA apparatus was performed to center the measurement 

volume between the rod banks depicted in Figure 2.6. Unlike the CFD profiles, the amplitudes 

of the measured maxima and minima are not equal due to random orientations of the heater 

rods, which had a slight curvature in their shape imparted from the manufacturing process. The 

rods at x/P = 2 and 6 were particularly distorted, nearly preventing penetration to the center of 

the assembly. However, the integrated line averages of all the LDA (dashed black lines) and 

CFD (dashed gray lines) profiles were within about 1percent of each other in both the bundle (0 

≤ x/P ≤ 8) and in the annulus (x/P > 8). These integrated line averages of forced flows are 

repeated in subsequent graphs of heated flux profiles for reference.
4-10

Figure 4.10 Normalized mass flux profiles within the fuel assembly for measured flows of 150 

(red squares), 186 (blue diamonds), and 225 slpm (green triangles) and for 

simulated flows of 120 (open pink squares) and 240 slpm (open light-blue 

diamonds)

4.3.2 Naturally Induced Velocity Profiles

Figure 4.11 gives the measured and simulated normalized mass flux profiles of the naturally 

induced flow for an assembly power of 0.5 kW. Again, the long dashed lines represent the 

integral averages of each profile from 0 ≤ x/P ≤ 8 for the bundle and x/P > 8 for the annulus. 

The black and gray dashed lines represent the integrated line average of the forced flow cases 

for the LDA and CFD, respectively. The integral averages indicate that the bundle flow 

increases and the annular flow decreases for the naturally induced flows compared to the forced 

flow values. The LDA and CFD 0.5 kW profiles show decreases in the annulus of 10 and 5

percent of the integral line averages compared to the forced cases, respectively. Conversely,

the bundle mass fluxes increased by 3 and 8 percent over the forced integral average for the 

LDA and CFD results, respectively.

Figure 4.12 gives the normalized mass flux profiles at an assembly power of 1 kW for naturally 

induced flow as a function of x/P at 12 hours. The forced flow integral averages are repeated in 

the plot as the black dashed (LDA) and gray dashed (CFD) lines. The integrated line averages 

of the heated mass fluxes in the bundle are 5 and 10 percent higher than the forced cases for 

the LDA and CFD results, respectively. Comparison of the heated and forced annulus line 

averages reveals decreases of 11 and 8 percent for the LDA and CFD results, respectively.
4-11

Figure 4.11 Normalized mass flux profiles within the fuel assembly of the naturally induced 

flows at 0.5 kW for measured (red squares) and simulated (open blue diamonds)

Figure 4.12 Normalized mass flux profiles within the fuel assembly of the naturally induced 

flows at 1 kW for measured (red squares) and simulated (open blue diamonds)
4-12

By multiplying the average, normalized mass fluxes in the bundle and annulus by their 

respective areas, the mass flow fraction may be estimated. This technique will knowingly 

overestimate the flows especially in the bundle because the mass fluxes are estimated along 

lines of maximum value, whereas the flow field is two-dimensional. However, this line-based 

estimation is the only method available for the LDA data. The CFD data provides both this line 

average value and the actual mass flows in the bundle and annulus. Figure 4.13 shows the 

mass flow fractions for the LDA and CFD line average technique as a function of power. The 

area-based, or actual, mass flow fractions from the CFD results are also provided. As 

expected, the line-based mass flow fractions in the bundle overestimate the area-based value 

by up to 0.033. However, the line average mass flow fractions for the LDA and CFD results 

agree to within 0.007. All the data suggest that the mass flow fraction in the bundle is 

increasing with assembly power at this axial location. The CFD results are explored next to 

determine the evolution of the mass flow fractions in the bundle and annulus as the flow passes 

up the assembly.

Figure 4.13 Bundle and annulus mass flow fraction at z = 0.442 m (17.4 in.) within the assembly 

for the bundle from the LDA line average (red squares), CFD line average (blue 

diamonds), and CFD area average (black circles) and for the annulus from the LDA 

line average (open red squares), CFD line average (open blue diamonds), and CFD 

area average (open black circles)

The area-average mass flow fractions in the bundle and annulus from the CFD simulations are 

next plotted as a function of the z-coordinate in Figure 4.14. The limited number of points 

shown in the graph are at the computational inlet and exit as well as two mid-bundle locations. 

The effects of the flow area contraction at spacers are not depicted in the graph but were 

captured in the CFD results. For reference, the bundle and annulus flow areas represent about 

75 and 25 percent of the total flow area in a bundle section, respectively. As expected, the 

forced mass flows partition according to the flow area ratios at the inlet where a uniform velocity 

condition is imposed. The forced flow appears to asymptote to approximately 0.63 of the flow in 
4-13

the bundle and 0.37 in the annulus. The heated cases show an inlet bundle mass flow fraction 

of 0.64, which then quickly increases as the flow moves up the bundle to z = 0.442 m (1.45 ft). 

The amount of this increase appears to be dependent on the assembly power. The bundle 

mass flow fraction then decreases as the flow moves further up the assembly, approaching the 

forced flow value. Finally, the bundle flow fraction at the exit of all the CFD simulations shows a 

decrease to approximately 0.6 due to the close proximity of a spacer at z = 3.89 m (12.8 ft). 

The annular mass flow fraction mirrors the behavior in the bundle.

Figure 4.14 Bundle and annulus mass flow fraction within the assembly for the bundle forced 

flow (red squares), 0.5 kW (blue diamonds), 1 kW (green triangles), and 3.5 kW 

(purple circles) and for the annulus forced flow (open pink squares), 0.5 kW (open 

light-blue diamonds), 1 kW (open light-green triangles), and 3.5 kW (open lightpurple circles)

5-1

5 CELL 2 IGNITION TESTING AND ANALYSIS

Phase I of the Spent Fuel Pool Project concluded with the ignition of the test assembly in the 

Cell 2 configuration. This test was conducted between the dates of February 28 and March 1, 

2011. The test was started at 20:00:25 Mountain Standard Time. The applied power to the 

assembly was 5.0 kW, which represents an assembly approximately 17 months after offload. 

The power to the assembly was lost when ignition occurred. Ignition of the Zircaloy within the 

assembly was first indicated at the z = 3.302 m level (130 in.) at an elapsed test time of 

12.66 hours based on a thermal criterion of 1,200 K (1,700°F).

5.1 Ignition Test and MELCOR results

The baseline MELCOR model used for Phase I predictions represents the bundle as a series of 

twelve axial nodes along the heated length. Similar to a grid convergence study, this baseline 

model was adapted to include more axial nodes and additional rings to represent the bundle. 

An axial nodalization refinement of twice as many nodes was applied to the bundle. The radial 

refinement consisted of repartitioning the single ring into two rings with roughly equal cross 

sectional areas and therefore equal bundle masses.

The MELCOR calculations were all performed with current best practice modeling parameters 

and sensitivity coefficients for the analysis of severe accidents, with the exception of radiative 

exchange factors. These exchange factors were adjusted based on the discretization of the 

model. Exchange factors should be based on standard expressions for simple geometries, 

where possible, or on experimental data or detailed radiation calculations for complicated 

geometries involving intervening surfaces, such as for radiation between “representative” 

structures in cells containing a number of similar structures (e.g. fuel rod bundles). In absence 

of this information, they should be treated parametrically to examine the effects of radiation on 

the course of a calculation. The values for exchange factors are user defined input and can 

range from zero to unity. The baseline values for the MELCOR model are 0.1 for both axial and 

radial exchange factors. The baseline model is therefore denoted as 12A_1R (0.1A, 0.1R) for 

12 axial nodes, 1 radial ring, 0.1 axial exchange factor, and 0.1 radial exchange factor. 

Sensitivity analyses for the radial exchange factor was not considered, as it is ignored by 

MELCOR in a single ring model.

5.1.1 Thermal-Hydraulic Comparisons 

Figure 5.1 shows temperature as a function of time for the test maximum, test average (at 

maximum temperature elevation), and peak cladding temperature of the MELCOR single ring, 

twelve axial node model. The MELCOR results are again denoted by the number of axial nodes 

and number of rings, along with the applied axial and radial radiative exchange factors (e.g. 

12A_1R (0.1A, 0.1R) denotes the single ring, twelve node model with axial and radial exchange 

factors of 0.1). The test values are truncated beyond the time of ignition due to loss of 

instrumentation and noise in the remaining TCs. The MELCOR post-test prediction of ignition 

time was within one percent of the observed test ignition time.
5-2

Figure 5.1 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated decay power 

for the maximum test temperature (red squares), average test temperature (blue 

diamonds), and MELCOR 12A_1R (pink dashed)

Figure 5.2 shows the axial temperature distribution in the assembly prior to ignition at 

t = 12 hours for the test maximum, test average, test minimum, and the MELCOR post-test 

prediction with 5.0 kW input power. The model over predicts the temperature for all locations in 

the assembly. The MELCOR PCT is within 60 K (108°F) of the test maximum, and the 

prediction of the PCT location is accurate to within the nodal resolution of the model. However, 

the onset of oxygen consumption in the MELCOR model occurs prior to what was shown by the 

experiment, as seen in Figure 5.5. The MELCOR model and experiment PCT are not expected 

to match at 12 hours because the higher-than-measured oxidation in the MELCOR model is 

forcing the temperature higher. Therefore, the axial temperature profiles at 10 hours were 

compared when MELCOR and the experiment were in more agreement with respect to oxygen 

consumption.
5-3

Figure 5.2 Bundle temperatures as a function of axial height in the assembly at 5.0 kW 

simulated decay power and 12 hours elapsed for the test maximum (red squares), 

test average (blue diamonds), test minimum (green triangles), and MELCOR 

12A_1R (pink dashed)

Figure 5.3 shows the axial temperature distribution in the assembly prior to ignition at 

t = 10 hours for the test maximum, test average, test minimum, and MELCOR post-test 

prediction with 5.0 kW input power. The model over predicts the temperature for all locations in 

the assembly but is within 28 K (50°F) of the test maximum. The prediction of PCT location is 

accurate to within the nodal resolution of the model.

The measured and predicted induced flow rates for the ignition test are shown in Figure 5.4. 

The MELCOR results are within the experimental uncertainty until the point of ignition, as 

mentioned earlier.

As observed in the analysis of ignition test, the MELCOR input model of 12 axial nodes and 1 

radial node, captured the ignition time accurately, within 1 percent.
5-4

Figure 5.3 Bundle temperatures as a function of axial height in the assembly at 5.0 kW 

simulated decay power and 10 hours elapsed for the test maximum (red squares), 

test average (blue diamonds), test minimum (green triangles), and MELCOR 

12A_1R (pink dashed)

Figure 5.4 Induced flow rate within the Cell 2 assembly at a simulated decay power of 5.0 kW 

for the test (blue solid) and MELCOR 12A_1R (pink dashed)
5-5

5.1.2 Exhaust Stream Analyses

5.1.2.1 Oxygen Concentration Monitoring

Figure 5.5 gives the bundle, annulus, and MELCOR oxygen concentrations in the assembly as 

a function of time. The ceramic, bundle sample tube became caught in the top bus plate and 

was shattered at 13.2 hours as the assembly shifted and slumped. The bundle O2 sensor and 

RGA slipstream sample lines were reconnected to the main annulus sample line at 13.4 hours. 

The annulus O2 sensor was abandoned after 13.4 hours. Figure 5.6 shows the state of the 

sample lines at various times of interest throughout the test.

Figure 5.5 Exit oxygen concentration as a function of time for assembly bundle (red squares), 

assembly annulus (blue diamonds), and MELCOR 12A_1R (pink dashed)

Figure 5.6 Overhead photographs of the assembly exit and exhaust sampling lines showing a) 

the initial configuration, b) immediately prior to the break in the bundle sample line, 

and c) immediately after the break in the bundle sample line

a) b) c)

Bundle O2 

sample 

Bundle 

tube prior 

to failure

Bundle 

tube failure
5-6

5.1.2.2 Residual Gas Analyzer

Figure 5.7 shows the mole or volume fraction of nitrogen, oxygen and argon as determined by 

the RGA analysis. The RGA analysis was started eleven hours into the test, which was almost 

two hours before ignition. Near the end of the test at 17.7 hours, ambient air was sampled for 

calibration. After initial evaluation of the data, additional calibration of the RGA was performed 

to test the validity of the air calibration for the samples that were devoid of oxygen. Two 

calibration gases were used. The first contained 1.22 percent argon in nitrogen, which 

represents the argon concentration naturally in air after all the oxygen is removed. The other 

calibration gas contained 4.76 percent argon in nitrogen, which represents the maximum argon 

concentration apparently measured, based on the single air calibration. The apparent argon 

concentration in the 1.22 percent and 4.76 percent calibration gases was determined to be 1.66 

percent and 7.85 percent respectively based on the air calibration. A linear correction 

correlation was used to adjust the argon and nitrogen concentrations for the samples collected 

during the burn phase when the oxygen concentration was essentially zero. The single point air 

calibration was solely used up until this point when significant oxygen was present.

Figure 5.7 Mole fraction of nitrogen, oxygen, and argon as a function of time during the 

ignition test

At about 12.66 hours, the oxygen concentration began to drop sharply indicating ignition had 

begun. Shortly after ignition at 13.2 hours, a mishap occurred when the ceramic tube on the top 

bus plate snapped (See Figure 5.6). Due to the break in the sampling tube, the RGA sample 

location changed to just above the top of the storage cell. This resulted in the dilution of the 

sample with ambient air. At 13.4 hours, the bundle sample was abandoned and the RGA was 

configured to draw sample from the annulus sample location.

Ignition Tube Break

Annulus Sampled
5-7

Once the sample was drawn from the annulus, the oxygen concentration dropped to zero due to 

the zirconium oxidation reaction. The removal of oxygen from the gas stream increased the 

nitrogen concentration significantly to 0.978 but not enough to be explained solely by the 

absence of oxygen, which would be 0.988. However, the argon concentration increased by a 

factor of 1.7 times greater than can be explained by the absence of oxygen. These combined 

observations are consistent with significant nitrogen removal by zirconium nitride formation. 

When nitrogen is not being consumed, the ratio of the nitrogen to argon in atmospheric air 

should remain constant at 83.6, termed here the background ratio. However, due to previous 

testing of the argon purge system, the concentration of argon inside the cylindrical boiling facility

vessel was artificially elevated above ambient to about 1.2 percent before the ignition test. The 

corresponding nitrogen to argon ratio was measured at 66.7. Thus, the background ratio during 

the ignition test was between 66.7 and 83.6. The uncertainty in the background ratio 

corresponds to uncertainty in the estimated nitrogen removal and ZrN formation. 

Since argon is inert, any change in this ratio should be due to a change in the relative amount of 

nitrogen present. A ratio below the background value signifies nitrogen removal. Figure 5.8

shows the nitrogen to argon concentration ratio. After the sample point was switched to the 

annulus, the measured nitrogen-to-argon ratio was 48.6, indicating that between 27 percent and 

42 percent (depending on the background ratio assumed) of the nitrogen drawn into the 

assembly was converted to ZrN during this burn phase of the experiment. Based on this 

nitrogen consumption and the measured inlet flow rate, the initial zirconium in the assembly 

converted to ZrN during this period was estimated to be between 20 and 40 percent. The 

amount of zirconium oxide formed was estimated with greater certainty based on the measured 

flow rate and assuming total oxygen consumption. During this first pass of the burn front 14

percent ± 3 percent of the initial zirconium was converted to ZrO2. Also plotted in Figure 5.8 is 

the total induced flow. During the burn phase, the induced flow correlates with the nitrogen to 

argon ratio. When the flow is high the ratio increases, signifying that more nitrogen makes it out 

of the assembly. When the flow is low the ratio is low signifying that less nitrogen makes it out 

the top of the assembly. The correlation of the nitrogen argon ratio with the independent 

induced flow measurement provides some confidence in the relative accuracy of the method. 

The absolute accuracy is still under scrutiny. The single point calibration with air inherently 

assumes a linear detector response with changes in concentrations. This assumption may not 

be valid, especially for argon, which underwent a multi-fold increase in concentration. These 

results indicate that the hot oxygen starved environment left after the passage of the burn front 

is ideal for significant zirconium nitride formation.
5-8

Figure 5.8 Ratio of nitrogen to argon in the exhaust stream and assembly flow rate as a 

function of time

5.1.3 Post-Test Examination and Deconstruction

The SFP Phase I test assembly continued to react for several days after the first ignition. The 

temperature within the assembly was low enough after about one week to allow for examination 

and removal of the test materials from the vessel at the Cylindrical Boiling facility. As expected, 

the internal geometry of the fuel bundle and storage pool cell were heavily distorted due to the 

severe thermal environment caused by the ignition of the zirconium. Figure 5.9 shows a 

sequence of photographs throughout the deconstruction process. These photographs are 

labeled alphabetically to denote chronology. The outer thermal radiation barrier was removed 

first, followed by the high-temperature insulation. The top ~0.6 m (24 in.) portion of the fuel 

bundle was detached and lifted from the assembly as shown in Figure 5.9 (b) and (c). Although 

heavily oxidized, the stainless steel pool cell remains structurally intact in this segment. This 

segment of the fuel bundle represents the most preserved piece and also coincides with the fuel 

located above the original ignition location. The fuel located below this location was largely 

unstable and disintegrated upon removal of the insulation. The stainless steel pool cell in this 

lower portion either had relocated further down the assembly or was so heavily oxidized as to 

be no longer structurally viable.

Figure 5.10 shows photographs from various perspectives of the Phase I fuel bundle. Figure 

5.10 (a) is the view through the LDA port with the insulation and quartz window removed. Slag 

from the pool cell is evident in the lower portion of the port. Figure 5.10 (b) and (c) show the 

initial point of ignition, z = 3.302 m (130 in.), before and after the removal of the top section, 

respectively. The top ~0.6 m (24 in.) portion of the fuel bundle is shown in Figure 5.10 (d). 

Figure 5.10 (e) and ( f ) show the lower part of the fuel bundle from different angles.

Ignition

Annulus Sampled
5-9

Figure 5.9 Deconstruction photographs of the SFP Phase I test assembly

b)

c)

d)

e)

f)

g)

h)

a)
5-10

Figure 5.10 Photographs of the SFP Phase I fuel bundle after the ignition test

5.2 MELCOR Sensitivity Studies

5.2.1 Axial Nodalization Sensitivity (24A_1R)

Figure 5.11 shows temperature as a function of time for the test maximum, test average at 

maximum temperature elevation, and peak cladding temperature of the MELCOR single ring, 

twenty-four axial node model for the 5 kW ignition case. The axial radiative exchange factors of 

the single ring, twenty-four axial node model had to be increased compared to the twelve node 

model to account for the finer nodalization. The axial radiative exchange factor adjustment is 

made because the nodes are shorter, which effectively increases the node-to-node view factor. 

The MELCOR results for this model were again within one percent of the observed ignition test 

time.

Figure 5.12 shows the axial temperature distribution in the assembly prior to ignition at 

t = 10 hours for the test maximum, test average, test minimum, and the MELCOR 24A_1R 

model. The model over predicts the temperature for all locations in the assembly. The 

MELCOR PCT is within 35 K (63°F) of the test maximum, and the prediction of PCT location is 

accurate to within the nodal resolution of the model.

a) b) c)

d) e) f)
5-11

Figure 5.11 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated decay power 

for the maximum test temperature (red squares), average test temperature (blue 

diamonds), and MELCOR 24A_1R (pink dashed)

Figure 5.12 Bundle temperatures as a function of axial height in the assembly at 5.0 kW 

simulated decay power and 10 hours elapsed for the test maximum (red squares), 

test average (blue diamonds), test minimum (green triangles), and MELCOR 

24A_1R (pink dashed)
5-12

5.2.2 Radial Nodalization Sensitivity (12A_2R)

Figure 5.13 shows temperature as a function of time for the test maximum, test average, and 

peak cladding temperature of the MELCOR two ring, twelve axial node model. The assembly 

was divided in two concentric rings in the MELCOR model. The inner ring represented 132 

heated rods and the outer ring represented the remaining 132 heated rods and the rack. The 

temperature of the first, inner ring is denoted as 12A_2R[1] in the legend. The second, outer 

ring is denoted as 12A_2R[2]. This convention is repeated throughout the remainder of this 

report. The two ring model radial radiative exchange factors were increased to unity. The 

increase to unity was made because a large fraction of the rods in ring 1 are expected to 

radiatively interact in ring 2. If a smaller exchange factor was used, ring 1 temperatures would 

be unrealistically high. The MELCOR time to ignition was again within one percent of the 

observed ignition test time.

Figure 5.13 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated decay power 

for the maximum test temperature (red squares), average test temperature (blue 

diamonds), MELCOR 12A_2R[1] (pink long dashed), and MELCOR 12A_2R[2] 

(orange short dashed)

Figure 5.14 shows the axial temperature distribution in the assembly prior to ignition at 

t = 10 hours for the test maximum, test average, test minimum, and both MELCOR rings. The 

model over predicts the temperature for all locations in the assembly for ring 1. The MELCOR 

PCT is within 30 K (54°F) of the test maximum, and the prediction of PCT location is accurate to 

within the nodal resolution of the model.
5-13

Figure 5.14 Bundle temperatures as a function of axial height in the assembly at 5.0 kW 

simulated decay power and 10 hours elapsed for the test maximum (red squares), 

test average (blue diamonds), test minimum (green triangles), MELCOR 12A_2R[1] 

(pink long dashed), and MELCOR 12A_2R[2] (orange short dashed)

5.2.3 Combined Axial and Radial Nodalization Sensitivity (24A_2R)

Figure 5.15 shows temperature as a function of time for the test maximum, test average, and 

peak cladding temperature of the MELCOR two ring, twenty-four axial node model. The axial 

exchange factor had to be increased in the two ring model, similarly to the 24A_1R model. The 

MELCOR result was within two percent of the observed ignition test time.

Figure 5.16 shows the axial temperature distribution in the assembly prior to ignition at 

t = 10 hours for the test maximum, test average, test minimum, and the MELCOR model. The 

model over predicts the temperature for all locations in the assembly for ring 1. The MELCOR 

PCT is within 40 K (72°F) of the test maximum, and the prediction of PCT location is accurate to 

within the nodal resolution of the model.

The results of the sensitivity study demonstrate that increasing the core nodalization beyond the 

baseline model does not noticeably improve accuracy for the purposes of the SFP Phase I 

modeling efforts. Given the treatment of the thermal system as a lumped mass within 

MELCOR, refined nodalizations are not expected to necessarily improve the agreement with the 

test data as long as the original discretization is reasonably resolved. Furthermore, user inputs 

such as axial and radial exchange factors require re-examination with each node refinement.
5-14

Figure 5.15 Cell 2 bundle temperatures as a function of time at 5.0 kW simulated decay power 

for the maximum test temperature (red squares), average test temperature (blue 

diamonds), MELCOR 24A_2R[1] (pink long dashed), and MELCOR 24A_2R[2] 

(orange short dashed)

Figure 5.16 Bundle temperatures as a function of axial height in the assembly at 5.0 kW 

simulated decay power and 10 hours elapsed for the test maximum (red squares), 

test average (blue diamonds), test minimum (green triangles), MELCOR 24A_2R[1] 

(pink long dashed), and MELCOR 24A_2R[2] (orange short dashed).
6-1

6 SUMMARY

Testing of a single, full-scale 17×17 PWR mock fuel assembly for Phase I of the Spent Fuel 

Pool Project has been completed. The near-prototypic fuel assembly was constructed to 

measure the thermal-hydraulic response of spent fuel under complete loss-of-coolant accident 

scenarios. Testing included a pre-ignition series for Cell 1 and Cell 2 and concluded with a final, 

destructive ignition experiment for Cell 2. The assembly in the Cell 1 configuration, ID = 

221.3 mm (8.71 in.) and in the Cell 2 configuration, ID = 223.4 mm (8.80 in.), were tested in air 

environments below ignition temperatures for simulated decay powers of 0.5, 1.0, 1.5, 2.0, 2.5, 

3.0, and 3.5 kW. The final ignition test was conducted only for Cell 2 at a simulated decay 

power of 5.0 kW, or equivalent offload age of approximately 17 months.

The ignition test was conducted with an applied power of 5.0 kW. The time to ignition was 

captured by MELCOR to within one percent of the observed time. The MELCOR maximum 

temperature was within 60 K (108°F) of the maximum test temperature for times prior to ignition. 

Nodalization refinements to the baseline MELCOR model were performed to examine sensitivity 

of the model to increased discretization of the axial dimension and the effect of additional rings 

to capture transverse thermal gradients. Both axial and radial grid convergence efforts did not 

yield improvements to the predictions of temperature or time to ignition. Through sensitivity 

calculation for Cell 1 and Cell 2 cases, it was concluded that a Σk value of 30 matches very well 

the data for both Cell 1 and Cell 2. On the other hand, a SLAM value of 145 matches very well 

the data for Cell 1 while SLAM values between 146 and 133 match the data for Cell 2. In the 

ignition test, the MELCOR input model of 12 axial nodes and 1 radial node, captured the ignition 

time accurately, within 1 percent.

Laser Doppler anemometer (LDA) measurements and computational fluid dynamics (CFD) 

simulations were performed for both heated and unheated cases. These data indicate that flow 

is drawn preferentially into the bundle from the annulus for naturally induced flows as compared 

to forced flows at z = 0.442 m (1.45 ft). Based on the integrated line average mass fluxes of the 

LDA data, the bundle mass flow fraction increased by 0.030 and 0.034 over the forced flow case 

for 0.5 and 1 kW, respectively. Further analysis of the CFD data reveals that the bundle mass 

flow fraction for naturally induced flows initially increases in the lower part of the assembly and 

then decreases as the flow moves up the assembly. The highest simulated power case of 3.5 

kW indicated that the mass flow fractions approach those in the forced flow case for z ≥ 1.5 m

(4.9 ft).

During the ignition test, a residual gas analyzer (RGA) was used to monitor the amount of 

nitrogen and argon exiting the top of the assembly. The ratio of nitrogen to argon was used to 

determine if nitrogen was being consumed by reaction with zirconium. While encountering 

some sampling difficulties, the RGA successfully measured a significant amount of nitrogen 

consumption at the start of ignition and during the burn front progression to the bottom of the 

assembly. The single point calibration with air was used for samples containing significant 

concentrations of oxygen. The analysis of two additional calibration gases allowed adjustment 

of the air calibration for samples that were devoid of oxygen. 
6-2

During the burn phase, all of the oxygen was removed from the air drawn into the assembly 

converting 14 percent of the initial zirconium to ZrO2. Depending on which background ratio 

was assumed, between 20 percent to 40 percent of the zirconium in the assembly was 

converted to ZrN. These results indicate that the hot oxygen starved environment remaining 

after the passage of the burn front is ideal for significant zirconium nitride formation.
7-1

7 REFERENCES

1. E.R. Lindgren, and S.G. Durbin, “Characterization of Thermal-Hydraulic and Ignition 

Phenomena in Prototypic, Full-Length Boiling Water Reactor Spent Fuel Pool Assemblies 

after a Complete Loss-of-Coolant Accident,” NUREG/CR 7143, April 2013.

2. M. Steinbrück, J. Jung, and M. Walter, “Separate-Effects Tests on the Investigation of HighTemperature Oxidation Behavior and Mechanical Properties of Zircaloy-2 to Be Used in the 

SFP PWR Tests,” Karlsruher Institute for Technology Letter Report, 2010.

3. S.G. Durbin, and E.R. Lindgren, “Laminar Hydraulic Analysis of a Commercial Pressurized 

Water Reactor Fuel Assembly,” NUREG/CR 7144, March 2013.

4. J. Cardoni, “MELCOR Model for an Experimental 17×17 Spent Fuel PWR Assembly,” 

SAND2010-8249, November 2010.

5. F. P. Incropera et al., "Fundamentals of Heat and Mass Transfer," 4th edition, John Wiley & 

Sons, Inc. (1998).

A-1

APPENDIX A – ERROR PROPAGATION 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 A.1. The standard uncertainty associated with an 

indirect measurement is analogous to the standard deviation of a statistical population.

∑= 















∂

∂ =

N

i

i

i

u

x

y u

1

2

2 A.1

Here, u is used to define the standard uncertainty of a measurement.

The expanded uncertainty U is reported in this appendix and defines the bounds that include 95

percent of the possible data. The expanded uncertainty is usually defined as some multiple of 

the standard uncertainty. For purposes of this report, all uncertainty measurements are 

assumed to behave according to a normal distribution. Therefore, Equation A.2 shows the 

definition of the expanded uncertainty as used in the following sections for a 95 percent

confidence interval, where the factor of 1.96 has been rounded to a value of 2.

U u = ⋅ 2 A.2

A.1 Uncertainty in Average Bundle Velocity

The uncertainty in the bundle velocity was determined using error propagation analysis (EPA) 

for the blocked guide tube measurements in the Cell 2 configuration (ID = 223.4 mm (8.80 in.)). 

The assembly velocity was determined from Equation A.3 in which Qi is the volumetric flow rate 

in slpm for each flow controller, Aassembly is the cross sectional area of the assembly, R is gas 

constant for air, T is the inlet air temperature, and P is the ambient air pressure. The first term 

in the equation represents the conversion from slpm to kg/s.

8

3

STP 1

o

assembly

m Q 0.001ρ l R T

60 s A P min

i

i w =

  ⋅   ⋅ =          

∑

A.3
A-2

Equation A.4 gives the relation between the overall uncertainty of wo and the contributions from 

the measurement uncertainties of Aassembly, Q, Τ, and P. The uncertainty in these fundamental 

measurements are defined in Sections A.1.1 through A.1.4, respectively.

o

assembly

2 2

8

tot 2 3 Q T

STP 1 assembly assembly 2

2 2

tot tot

2 2 P A

assembly assembly

RT Q R

0.001ρ

m P A P A l

s 60 min Q RT Q RT

P A P A

i

i

w

u u

u

u u

=

          +   ⋅   ⋅ ⋅    =  

       + +   

   ⋅ ⋅     

∑

A.4

Table A.1 summarizes the values used to determine the overall uncertainty of the assembly

velocity. The overall uncertainty in wo was found for the highest volumetric flow rate achieved 

during testing, Qtot = 2,300 slpm, at a typical ambient condition of T = 298 K (77°F), P = 83.4 Pa 

(12.1 psi), and Aassembly = 0.0283 m2 (0.305 ft2

). The standard uncertainty was determined to be 

wo u = 0.039 m/s (0.128 ft/s). The uncertainty was dominated by the assembly cross-sectional 

area (Aassembly) and the air temperature (T) contributing 96 and 3 percent of the overall 

uncertainty, respectively. The remainder was due to the uncertainty in the flow rate and 

atmospheric pressure.

Table A.1 Measurement uncertainties and intermediate calculations for wo

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂w o/∂x i)/w o]) Contribution

Q1 slpm 100 0.5 0.000 0.00

Q2 slpm 200 1.0 0.000 0.00

Q3 slpm 300 1.5 0.001 0.00

Q4 slpm 300 1.5 0.001 0.00

Q5 slpm 300 1.5 0.001 0.00

Q6 slpm 300 1.5 0.001 0.00

Q7 slpm 400 2.0 0.001 0.00

Q8 slpm 400 2.0 0.001 0.00

T K 298 1.1 0.004 0.03

P N/m2 83400 55 0.001 0.00

Aassembly m2 0.0283 6.0E-04 0.021 0.96

w o m/s 1.793 0.039 0.0216 1.00

A.1.1 Uncertainty in Assembly Cross Sectional Area

The inner dimension of the storage cell was measured to within ±1.3 mm (±0.05 in.). The outer 

diameter of the simulated fuel rods was measured to within ±0.04 mm (±0.0015 in.). The outer 

diameter of the guide tubes was measured to within ±0.03 mm (±0.001 in.). This tolerance 

leads to a maximum standard uncertainty of 6·10-4 m2 (2·10-3 ft) in the hydraulic area.

Table A.2 Measurement uncertainties and intermediate calculations for Aassembly

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂Aassembly/∂x i)/AassemContribution

ODpin m 0.0095 0.00004 0.005 0.06

ODGT m 0.0122 0.00003 0.000 0.00

IDpool m 0.2234 0.00130 0.021 0.94

Aassembly m2 0.0283 0.00060 0.021 1.00
A-3

A.1.2 Uncertainty in Volumetric Flow Rate Q

The volumetric flow rate was controlled with eight MKS volumetric flow controllers operated in 

parallel (Model # 1559A-24174). The expanded uncertainty of the volumetric flow rate was 

determined from the stated manufacturer’s upper uncertainty of 1 percent of full scale. The 

standard uncertainties in flow rate were 0.5 slpm for flow controller 1, 1 slpm for flow controller 

2, 1.5 slpm for flow controllers 3 through 6, and 2 slpm for flow controllers 7 and 8. The value 

shown in Table A.1 represents these standard uncertainties associated with the volumetric flow 

rate. 

A.1.3 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 (2.0°F).

A.1.4 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 expanded uncertainty in the instrument of ±0.1 percent of full scale (110 kPa (16 psi)). 

Therefore, the standard uncertainty in the pressure reading is uP = 55 Pa (0.008 psi).

A.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 percent of full scale, or ±4.1 N/m2 (6·10-4 psi). This error band includes 

repeatability, hysteresis, and conformance. Furthermore, these error bands consider the zerodrift 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 largest observed zero drifts in the zero 

flow measurements were less than 0.92 N/m2 (1.3·10-4 psi), which is smaller than the plotted 

symbols in this report.

A.3 Uncertainty in SLAM and Σk Coefficients

The uncertainties in the SLAM and Σk coefficients were evaluated from the hydraulic equations 

that describe their behavior and the curve fit coefficients determined from the pressure drop 

measurements. Curve fits to pressure drop data have been presented to capture the form 

losses (quadratic term) and the friction losses (linear term).

2 P 2 1 o o ∆= ⋅ + ⋅ aw aw A.5

The equations for SLAM and Σk are written as the following.

2

H

LAM 1

D S 2

Lμ

a

  = ⋅   ⋅  

A.6

2 2 k

ρ

⋅a Σ = A.7
A-4

The standard uncertainties for the flow loss coefficients were determined using these equations 

for the Cell 2 configuration. The calculation of the standard uncertainties for SLAM and Σk are 

outlined in Table A.3 and Table A.4, respectively. The expanded uncertainties for SLAM and Σk

are ±13 and 1.4, respectively. Explanations of the uncertainties in the underlying 

measurements are given in the following sections.

Table A.3 Measurement uncertainties and intermediate calculations for SLAM

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂SLAM/∂x i)/SLAM]) Contribution

a1 s/m 41 0.4 0.011 0.06

DH m 0.0116 0.0002 0.043 0.94

L m 4.0472 0.003 0.001 0.00

SLAM -- 146 6.5 0.044 1.00

Table A.4 Measurement uncertainties and intermediate calculations for Σk

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂Σk/∂x i)/Σk]) Contribution

a2 s

2

/m2 12.5 0.3 0.027 0.930

ρ kg/m3 0.976 0.0037 0.007 0.070

Σk -- 26 0.7 0.0283 1.000

A.3.1 Uncertainty in Pressure Drop Curve Fit Coefficients

The pressure drop through the assembly was measured for various bundle velocities as shown 

in Figure 2.2. Curve fits to these data were used to determine the quadratic and linear 

coefficients. Standard statistical treatment within Excel was used to determine the standard 

error associated with each coefficient. The standard errors were ua1 = 0.4 and ua2 = 0.3.

A.3.2 Uncertainty in Hydraulic Diameter

The hydraulic diameter is determined from the wetted perimeter and the cross sectional flow 

area (Section A.1.1). Table A.5 outlines the determination of the uncertainty in the hydraulic 

diameter.

Table A.5 Measurement uncertainties and intermediate calculations for DH

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂DH/∂x i)/DH]) Contribution

Aassembly m2 0.0283 0.0006 0.021 0.98

Pwet m 9.7715 0.0321 0.003 0.02

DH m 0.0116 0.0002 0.021 1.00

A.3.2.1 Uncertainty in the Wetted Perimeter

The uncertainty in the wetted perimeter was estimated using the known uncertainties in the 

geometry of the assembly. Table A.6 shows the quantities used to determine the uncertainty in 

the wetted perimeter. The uncertainty is completely determined by the variation in the heater 

pin diameter.
A-5

Table A.6 Measurement uncertainties and intermediate calculations for Pwet

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂Pwet/∂x i)/Pwet]) Contribution

ODpin m 0.0095 0.00004 0.003 0.97

ODGT m 0.0122 0.00003 0.000 0.00

IDpool m 0.2234 0.00130 0.001 0.03

Pwet m 9.772 0.032 0.003 1.00

A.3.3 Uncertainty in Hydraulic Length

The length of the bundle was measured to within ±6.4 mm (0.25 in.). This tolerance was taken 

to be the expanded uncertainty. Therefore, the standard uncertainty reported in the table is half 

this value.

A.3.4 Uncertainty in Air Density

The density of the air flowing through the assembly was determined using the ideal gas law and 

measurements of the air flow temperature (Section A.1.3) and ambient pressure (Section A.1.4). 

The resulting uncertainty in the determination of the air density is summarized in Table A.7. The 

compressibility of air at test conditions is 0.9996, validating the use of the ideal gas law.

Table A.7 Measurement uncertainties and intermediate calculations for air density, ρ

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂ρ/∂x i)/ρ]) Contribution

T K 298.0 1.1 0.004 0.969

P N/m2 83400 55 0.001 0.031

ρ kg/m3 0.976 0.004 0.004 1.000

A.4 Uncertainty in Hot Wire Anemometer Measurements

Two types of hot wire anemometers were initially installed for measurement of the naturally 

induced flow rate through the assembly. These included the Dwyer Model 641RM and the TSI 

Model 8455. Only the TSI measurements were included in this report. Details of the two Dwyer 

hot wires are included in the Cell 2 Quick Look report. Because the data from the Dwyer hot 

wires is still included in the ignition data files, estimates of the uncertainty in the Dwyer hot wires 

is included in this error propagation analysis.

Figure A.1 shows the calibration curves for all three hot wires in the Cell 2 configuration. The 

uncertainties in the instrument voltages were determined from manufacturer specifications (see 

Sections A.4.1 and A.4.2). The uncertainties in the curve fit coefficients are estimated in 

Section B.4.3.

The standard uncertainties in the TSI and Dwyer hot wires were determined to be ±6 and

±9 slpm, respectively. Therefore the 95 percent confidence interval in these measurements 

were ±12 and ±18 slpm for the TSI and Dwyer hot wires, respectively. Although the uncertainty 

calculations for the Dwyer hot wire was calculated only using Dwyer 2, the uncertainty in 

Dwyer 1 is assumed to be equivalent to the Dwyer 2 values.
A-6

Figure A.1 Flow rate as a function of hot wire anemometer voltage

Table A.8 Measurement uncertainties and intermediate calculations for assembly flow rate 

from the TSI hot wire

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂QTSI/∂x i)/QTSI]) Contribution

VTSI V 2 0.05 0.029 0.797

aTSI, 0 slpm -28.5 1.9 0.011 0.117

aTSI, 1 slpm/V 101.4 0.8 0.010 0.086

QTSI slpm 175 6 0.032 1.000

Table A.9 Measurement uncertainties and intermediate calculations for assembly flow rate 

from the Dwyer 2 hot wire

Measurement, x i Units Value Standard uncertainty, u i Influence coefficient (u i·[(∂QDwyer2/∂x i)/QDwyer2]) Contribution

VDwyer2 V 1.19 0.08 0.049 0.962

aDwyer2, 0 slpm 40.2 1.3 0.008 0.023

aDwyer2, 1 slpm/V 113.0 0.9 0.006 0.014

QDwyer2 slpm 175 9 0.049 1.000

A.4.1 Uncertainty in TSI Hot Wire Anemometer

The standard uncertainty of the TSI Model 8455 hot wire was stated by the manufacturer as 

±0.5 percent of full scale. For a full scale output of 10 V, this translates to a standard

uncertainty of 0.05 V.
A-7

A.4.2 Uncertainty in Dwyer Hot Wire Anemometers

The expanded uncertainty of the Dwyer Model 641RM hot wire was stated by the manufacturer 

as ±3 percent of full scale. For a full scale output of 5 V, this translates to a standard

uncertainty of 0.075 V.

A.4.3 Uncertainty in Hot Wire Curve Fit Coefficients

The outputs of the hot wire anemometers were measured for various bundle flow rates as 

shown in Figure A.1. Least-squares curve fits to these data were used to determine the slope 

and intercept. Standard statistical treatment within Excel was used to determine the standard 

error associated with each coefficient. The standard errors for the TSI curve fit coefficients were

uTSI, 1 = 0.8 slpm/V and uTSI, 0 = 1.9 slpm. The standard errors for the Dwyer 2 hot wire curve fit 

coefficients were uDwyer2, 1 = 0.9 slpm/V and uDwyer2, 0 = 1.3 slpm.

B-1

APPENDIX B – OXIDATION OF ZIRCONIUM

The oxidation of zirconium alloys at high temperatures in steam, oxygen and nitrogen containing 

environments has been the subject of numerous studies. The temperature range of interest 

spans from 773 to 1,673 K (932 to 2,552°F). The pertinent reactions and the energy released at 

1,500 K (2,240°F) are listed below (Ref. 1):

Zr + O2 → ZrO2 ΔHrxn = -1,094.6 kJ/mol Zr B.1

Zr + 2H2O → ZrO2 + H2 ΔHrxn = -529.1 kJ/mol Zr B.2

Zr + 1⁄2 N2 → ZrN ΔHrxn = -365.4 kJ/mol Zr B.3

ZrN + O2 → ZrO2 + 1⁄2 N2 ΔHrxn = -729.4 kJ/mol Zr B.4

B.1 General Oxidation (Ref. 2)

At elevated temperatures, zirconium readily reacts with oxygen to form a zirconium dioxide 

product layer by reaction (Equation B.1). The reaction is highly exothermic. On a fresh 

zirconium surface the initial reaction rate is rapid but as soon as a product layer is formed, the 

reaction rate is controlled by the diffusion of oxygen through the growing product layer. 

Provided there is ample supply of oxygen at the surface and the oxide layer remains dense and 

intact, the oxygen reaction rate is solid state diffusion limited and well described by parabolic 

reaction kinetics. Parabolic kinetics describes diffusion controlled situations where the reaction 

rate decreases as the product layer thickens. With solid state diffusion the driving force is the 

concentration gradient in the solid phase, no porosity is required. This is in contrast to gas 

phase diffusion through a porous product layer. The zirconium oxidation rate is not dependent 

on the gaseous oxygen concentration or even the oxygen bearing species, thus pure oxygen 

(Equation B.1) and pure steam (Equation B.2) react at nearly the same rate and form nearly 

identical oxide layers. Under these conditions, at the gas/oxide interface the surface oxide is 

fully stoichiometric ZrO2. At the inner oxide/metal interface, the oxide is substoichiometric 

(~ZrO1.9) which drives the solid-state diffusion through the oxide layer. However, if the bulk 

gaseous oxidant concentration drops to low levels or if, at very high temperatures, the oxidation 

rate becomes very high, gaseous diffusion of oxidant to the surface may control the oxidation 

rate (Ref. 3). This situation is called oxygen/steam starvation. As a consequence, net oxygen 

diffusion from the oxide to the metal causes the formation of substoichiometric oxide scale and 

may even result in the complete dissolution of the oxide layer into the remaining metal phase. 

B.2 Breakaway

As long as the growing product layer remains dense and intact, zirconium oxidation follows 

parabolic kinetics. However, eventually (and repeatedly) the product layer defects and “breaks 

away” from the inner metal interface and the reaction rate increases significantly. Several 

mechanisms seem to result in the breakaway phenomena. Often cited (Refs. 2, 4, 5, 6, 7) is the 

tetragonal to monoclinic phase transformation in the zirconia (ZrO2) product layer at 

temperatures below 1,323 K (1,922°F) resulting in a 3 to 5 percent volume change (Refs. 8, 9). 

This phase transformation is clearly an important initiator for breakaway kinetics and the 

breakaway phenomena is not observed during isothermal tests at 1,373 K (2,012°F) and above. 

The mechanism leading to breakaway may also cause radial cracking along columnar grain 

boundaries to relieve accumulated compressive stresses (Ref. 7). The radial cracks provide 

direct access for the gas phase to reach fresh metal. 
B-2

Additional mechanisms come into play when nitrogen is present in the gas phase. Above 

1,073 K (1,472°F) where nitriding becomes important, breakaway occurs much sooner when 

nitrogen is present. Steinbrueck (Ref. 6) measured the time to breakaway for Zr alloys in 

oxygen-Ar, synthetic air - Ar and steam-Ar mixtures. The time to breakaway generally 

decreases with increasing temperature due to the faster oxidation kinetics with a "local 

maximum" at around 1,073 K (1,472°F) caused by the alpha-beta phase transition of the metal 

connected with increased plasticity. Duriez (Ref. 7) measured the time to breakaway for bare 

Zr-4 in air. The time to breakaway was about 450 minutes at 873 K (1,112°F), 100 minutes at 

973 K (1,292°F), 30 minutes at 1,073 K (1,472°F) and 20 minutes at 1,173 K (1,652°F). At 

1,073 K (1,472°F) and 1,173 K (1,652°F) the time to breakaway is an order of magnitude faster 

when nitrogen is present. Duriez (Ref. 7) proposes a mechanism where the oxygen is rapidly 

consumed in the bottom of radial cracks producing a local oxygen starved, nearly pure nitrogen 

atmosphere. The nitrogen reacts with zirconium to produce ZrN in a spot wise manner. As the 

oxidation front moves inward the ZrN islands become embedded in oxide and eventually 

oxidize. The oxidation of ZrN produces a 46 percent volume increase, which submits the oxide 

layer to local high stresses causing further cracking, and the scenario repeats itself. The oxide 

formed by this process is more porous and less protecting than the oxide layer formed initially.

B.3 MELCOR Kinetic Treatment

The oxidation of zirconium is modeled in MELCOR as limited by both solid-state diffusion of 

oxygen through the oxide layer (assumed parabolic) and gaseous diffusion of steam or oxygen 

to the oxide surface. Both rates are calculated and the slower is used to define the reaction 

rate.

The zirconium oxidation kinetics used in MELCOR is based on the recent kinetics study 

conducted at Argonne National Laboratory (ANL) (Ref. 10). Air oxidation of bare Zr-4 was 

observed at temperatures as low as 773 K (932°F). Test specimens were held at a constant 

temperature and the weight gain due to oxidation and nitride formation was measured as a 

function of time. The initial reaction rates were described by parabolic kinetics expected for 

solid state oxygen diffusion. However, the rate of oxidation increased after some period of time 

and persisted for the remainder of the test. The higher “breakaway” kinetic rate was also found 

to follow a parabolic rate law. 

MELCOR incorporates both the pre- and post-breakaway kinetic correlations developed by 

Natesan and Soppet for bare Zr-4 (Ref. 10). The kinetic correlations found in Natesan and 

Soppet are given in units of mass of oxygen consumed while the kinetic correlation used by 

MELCOR are in units of mass of zirconium reacted, so an appropriate unit conversion is 

required.

A breakaway lifetime function is used to transition the model from pre- to post-breakaway 

kinetics. The model calculates an oxidation “lifetime” value for Zircaloy components in each cell 

using the local Zircaloy cladding temperature. Figure B.1 shows the breakaway timing data 

from the ANL tests. As the specimen temperature increased, the amount of time until 

breakaway became shorter.

For implementation into MELCOR, the ANL data was curve fit as follows.

( ) 0

t dt LF

τ T

′ = ∫ B.5
B-3

where, τ(T) is

τ(T) =10PLOX B.6

and PLOX

PLOX = -12.528 ⋅ log10 T + 42.038 B.7

A comparison of the data with Equation B.7 is shown in Figure B.1. For reference, specific 

values of the breakaway timing are provided in Table B.1.

Figure B.1 Comparison of the MELCOR breakaway timing fit to Zr-4 data (Ref. 10)

The MELCOR breakaway oxidation model calculates the lifetime function at every node in the 

MELCOR model with Zircaloy cladding. The oxidation kinetics linearly transitions from the prebreakaway correlation at LF = 1 to post-breakaway kinetics at LF = 1.25. Hence, only nodes

that have exceeded the lifetime function will have the higher post-breakaway oxidation kinetics.
B-4

Table B.1 MELCOR Fit of the Timings for Transition from Pre-Breakaway to Post-Breakaway 

Oxidation Reaction Kinetics Zircaloy-4 in the ANL Experiments (Ref. 10)

Specimen 

Temperature

Breakaway Timing (Eqn. 

B.6)

ANL Data Used in Curve 

Fit

400°C (673 K) 1,125 hr (Extrapolated) -

450°C (723 K) 458 hr (Extrapolated) -

500°C (773 K) 198 hr 144 hr

550°C (823 K) 90 hr -

600°C (873 K) 43 hr 64 and 81hr

650°C (923 K) 22 hr -

700°C (973 K) 11 hr 4.8 hr

750°C (1,023 K) 5.9 hr -

800°C (1,073 K) 3.3 hr 1.7 hr and 5.8 hr

850°C (1,123 K) 1.8 hr -

900°C (1,173 K) 1.1 hr 1 hr and 1.4 hr

References

1. I. Barin, Thermodynamic Data of Pure Substances, VCH Verlagsgsellshaft mbG,

Weinheim, and VCH Publishers, NewYork,1995.

2. M. Steinbrueck, “Prototypical experiments relating to air oxidation of Zircaloy-4 at high 

temperatures”, J. Nucl. Mat. 392 (2009) 531-544.

3. D.R. Olander, “Materials chemistry and transport modeling for severe accident analyses 

in light-water reactors. I: External cladding oxidation”, Nuclear Engineering and Design 

148 (1994) 253-271. 

4. J.H. Baek, and Y.H. Jeong, “Breakaway phenomenon of Zr-based alloys during a hightemperature oxidation”, J. Nucl. Mat. 372 (2008) 152-159.

5. C. Duriez, M. Steinbrueck, D. Ohai, T. Meleg, J. Birchley, T. Haste, “Separate-effect 

tests on zirconium cladding degradation in air ingress situations”, Nucl. Eng. Des. 239 

(2009) 244-253

6. M. Steinbrueck, “Oxidation of Zirconium Alloys in Oxygen at High Temperatures up to 

1600 oC”, Oxid. Met. (2008) 70:317-329.

7. C. Duriez, T. Dupont, B. Schmet, and F. Enoch, “Zircaloy-4 and M5 high temperature 

oxidation and nitriding in air”, J. Nucl. Mat. 380 (2008) 30-45.

8. R.N. Patil, and E.C. Subbarao, "Monoclinic-tetragonal phase transition in zirconia: 

mechanism, pretransformation and coexistence”, Acta Crystallographica Section A 26, 

535 (1970). 
B-5

9. J. Chevalier, L. Gremillard, A.V. Virkar, and D.R. Clarke., “The tetragonal-monoclinic 

transformation in zirconia: Lessons learned and future trends”, J. Am. Ceram. Soc. 92 

(2009), 1901-1920.

10. K. Natesan, and W.K. Soppet, “Air Oxidation Kinetics for Zr-Based Alloys”, NUREG/CR6846, ANL-03/32, Argonne National Laboratory, June 2004.

C-1

APPENDIX C – MELCOR MODEL DESCRIPTION

Spatial representation of the assembly in both the CVH and COR packages is essentially onedimensional, consisting of a stack of 10 control volumes and 17 axial core levels in a single 

radial ring (12 axial levels along the heated length of the rods), as shown in Figures C.1 through 

C.3.

Heater rod and other pool cell dimensions and masses were modeled based on data obtained 

from the facility description. As a simplification in the base case model, the mass of Nichrome 

resistive heating wire was excluded from consideration (i.e., the "fuel" component in the COR 

package was represented purely as the correct mass of MgO), and the small mass of upper 

plate and nozzle was also excluded.

The steel rack surrounding the pool cell was represented by mass allocated to the RK 

component available in recent versions of MELCOR for core models of type BWR-SFP and 

PWR-SFP. The layer of steel-wrapped Kaowool insulation surrounding the cell was 

represented as a radial boundary heat structure, which will thermally interact with the cell by 

means of thermal radiation heat transfer.

Spatial variation is accounted for only in the vertical direction (i.e., there is only one “ring” in the 

nodalization). Figures C.1 through C.3 show the nodalizations. The COR package uses 17 

cells, with 12 of them lying in the heated region. Ten control volumes are defined: one for each 

of the free spaces above and below the assembly (CV020 and CV010), one for the region to the 

sides of the assembly (CV030), and eight for the assembly (CV101 – CV108). The rack wall is 

accounted for by MELCOR heat structures. Other masses are represented by COR package 

components.
C-2

010

101

102

103

104

105

106

107

108

“Fuel” rod = 

3.918 

m

heated length = 

3.613 

m

0.254 

m

0.0 m

0.1651 m

0.3214 m

0.9236 m

1.5258 m

2.1280 m

2.7302 m

3.3324 m

3.9346 m

4.1884 m

020

“fuel” rod 5.0 m

upper “pool”

bottom “pool”

0.051 

m

005

109

108

107

106

105

104

103

102

101

002

CV30

Figure C.1 Nodalization of PWR spent fuel assembly (Cell 2 in Phase I)
C-3

el = 0.0 m

= 0.1524 m

= 0.1651 m

= 0.3214 m

= 0.6225 m

= 0.9236 m

= 1.2247 m

= 1.5258 m

= 1.8269 m

= 2.1280 m

= 2.4291 m

= 2.7302 m

= 3.0313 m

= 3.3324 m

= 3.6335 m

= 3.9346 m

= 4.0463 m

= 4.1884 m

HS 10100

HS 10200

10300

10400

10500

10600

10700

10800

10900

11000

11100

11200

11300

11400

11500

11600

HS 11700

insulation

stainless steel

Figure C.2 Heat structures for the rack wall (Cell 2 in Phase I)
C-4

Ax 1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

010

101

102

103

104

105

106

107

108

Fuel rod = 

3.918 

m

Heated length = 

3.613 

m

0.254 

m

Axial level CV number Elevation (m)

0.0

0.1524

0.1651

0.3214

0.6225

0.9236

1.2247

1.5258

1.8269

2.1280

2.4291

2.7302

3.0313

3.3324

3.6335

3.9346

4.0463

4.1884

Figure C.3 Representation of the core cell nodes






NUREG/CR-7215 Spent Fuel Pool Project Phase I: Pre-Ignition and Ignition Testing of a Single Commercial 17x17 Pressurized 

Water Reactor Spent Fuel Assembly under Complete Loss of Coolant Accident Conditions