Spent Fuel Pool Project Phase II: Pre-Ignition and Ignition Testing of a 1x4 

Commercial 17x17 Pressurized Water Reactor 

Spent Fuel Assemblies under Complete Loss 

of Coolant Accident Conditions 

 

Office of Nuclear Regulatory Research

NUREG/CR-7216

Spent Fuel Pool Project Phase II: Pre-Ignition and Ignition Testing of a 1x4 

Commercial 17x17 Pressurized Water Reactor 

Spent Fuel Assemblies under Complete Loss

of Coolant Accident Conditions

Manuscript Completed: September 2013

Date Published: April 2016

Prepared by:

S.G. Durbin, E.R. Lindgren and L. Humphries

Sandia National Laboratories

Albuquerque, NM 87185

Z. Yuan, M. Zavisca and M. Khatib-Rahbar

Energy Research, Inc.

Rockville, MD 20847-2034

R. Beaton

U.S. Nuclear Regulatory Commission

Washington, DC 20555-0001

G. A. Zigh, NRC Technical Monitor

R. Beaton, NRC Project Manager

S. Gonzalez, NRC Project Manager

NRC Job Code N6777

Office of Nuclear Regulatory Research

NUREG/CR-7216

 

iii

ABSTRACT

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

Economic Co-operation 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 was conducted in two 

phases at Sandia National Laboratories. The first phase focused on axial heating and zirconium 

fire propagation in a single 17×17 PWR fuel assembly. The second phase consisted of preignition and ignition tests of five prototypic PWR full-length 17×17 fuel assemblies to simulate a 

complete loss-of-coolant-accident in a spent fuel pool. The assemblies were placed in a 1×4 

configuration where the center fuel assembly was electrically heated and the four surrounding 

assemblies were unheated. Two of the unheated assemblies contained pressurized rods to 

study ballooning effects.

v

TABLE OF CONTENTS

1.1 Objective .................................................................................................................... 1-1

1.2 Testing Outline........................................................................................................... 1-1

2.1 General Construction................................................................................................. 2-1

2.1.1 Assembly Layout............................................................................................ 2-4

2.1.2 Dimensions of the Test Assembly ................................................................. 2-5

2.2 Design of the Heated Fuel Bundle ............................................................................ 2-6

2.2.1 Heater Design ................................................................................................ 2-6

2.2.2 Heater Rod Electrical Connections ............................................................... 2-7

2.3 Design of the Pressurized Peripheral Bundles........................................................ 2-12

2.3.1 Welding Techniques .................................................................................... 2-13

2.3.2 Argon Fill Gas .............................................................................................. 2-15

2.4 Instrumentation ........................................................................................................ 2-16

2.4.1 Center Bundle Thermocouples.................................................................... 2-16

2.4.2 Peripheral Bundle Internal Instrumentation................................................. 2-17

2.4.3 Inlet Instrumentation .................................................................................... 2-19

2.4.4 External Instrumentation.............................................................................. 2-20

2.4.5 Pressure Loss Measurements..................................................................... 2-22

2.4.6 Exhaust Gas Analysis.................................................................................. 2-25

4.1 General Thermal-Hydraulic Response...................................................................... 4-1

4.2 Burn Propagation....................................................................................................... 4-3

4.2.1 Axial Propagation........................................................................................... 4-3

4.2.2 Transverse Propagation ................................................................................ 4-6

4.3 Fuel Clad Ballooning.................................................................................................. 4-7

4.4 Exhaust Gas Analysis................................................................................................ 4-8

4.5 Post-Test Examination............................................................................................. 4-11

5.1 Introduction ................................................................................................................ 5-1

5.2 MELCOR Models ....................................................................................................... 5-1

5.2.1 9-Ring Model.................................................................................................. 5-3

5.2.2 2-Ring Model.................................................................................................. 5-8

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

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

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

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

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

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

2 APPARATUS AND PROCEDURES...............................................................................2-1

3 PRE-IGNITION TEST RESULTS....................................................................................3-1

4 IGNITION TEST RESULTS ............................................................................................4-1

5 MELCOR MODELING RESULTS...................................................................................5-1
vi

5.2.3 Oxidation Kinetics ........................................................................................ 5-11

5.3 Code Improvements................................................................................................. 5-12

5.4 Pre-Ignition Tests..................................................................................................... 5-14

5.4.1 1 kW pre-ignition.......................................................................................... 5-14

5.4.2 8 kW pre-ignition test ................................................................................... 5-15

5.5 Ignition Test.............................................................................................................. 5-24

5.6 MELCOR Sensitivities ............................................................................................. 5-34

5.6.1 Hydraulic Loss Coefficients ......................................................................... 5-34

5.6.2 Transition from pre- to post-breakaway oxidation kinetics.......................... 5-42

6.1 Phase II Test Results................................................................................................. 6-1

6.2 MELCOR Results....................................................................................................... 6-3

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

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

LIST OF FIGURES

Figure 1.1 Decay heat of a PWR 17×17 assembly as a function of time from 

discharge ..........................................................................................................1-2

Figure 1.2 Approximate pool loading scheme for a 1×4 test layout with an 

insulated boundary condition ............................................................................1-3

Figure 2.1 Various components in a typical 17×17 PWR fuel assembly.............................2-1

Figure 2.2 Thermal mass per unit length comparison of spent fuel and MgO fuel 

rod simulators ...................................................................................................2-2

Figure 2.3 Comparison of thermal conductivity for UO2, spent fuel, and MgO fuel 

rod simulators ...................................................................................................2-3

Figure 2.4 Layout of the Phase II test assembly.................................................................2-5

Figure 2.5 Dimensions of the as-built pool rack and the design insulation thickness..........2-6

Figure 2.6 Design of the electrically heated fuel simulators................................................2-7

Figure 2.7 Design of the top electrical bus plate.................................................................2-8

Figure 2.8 Assembly detail of the top electrical bus plate design .......................................2-8

Figure 2.9 Detail drawing of the top of the assembly showing two heater rods and 

a single guide tube............................................................................................2-9

Figure 2.10 Cross-sectional detail drawing of the top electrical connection from the 

power feed into the heater electrical pins ........................................................2-10

Figure 2.11 Detail drawing of the bottom of the assembly showing the electrical 

connection of a single heater rod to the bottom nozzle (neutral) .....................2-10

Figure 2.12 Power control system and test circuits ............................................................2-11

Figure 2.13 Schematic of the instrumentation panel for “C” circuit .....................................2-12

Figure 2.14 Schematic of the internal geometry of a Phase II peripheral rod and 

spent fuel ........................................................................................................2-13

Figure 2.15 Top end plug weld setup.................................................................................2-13

Figure 2.16 Loading of fuel rod with MgO surrogate fuel and stainless steel 

plenum spring .................................................................................................2-13

Figure 2.17 Bottom end plug weld setup............................................................................2-14

Figure 2.18 Schematic of the closure weld fixture..............................................................2-14

Figure 2.19 Close-up of gas fill process and top end plug weld location ............................2-15

Figure 2.20 Top end plug closure weld ..............................................................................2-15

Figure 2.21 Thermocouple layout for the Phase II center, heated fuel bundle....................2-17

Figure 2.22 Layout of internal instrumentation for the unpressurized, peripheral 

fuel bundles ....................................................................................................2-18

Figure 2.23 Layout of internal instrumentation for the pressurized, peripheral 

fuel bundles ....................................................................................................2-19

Figure 2.24 Layout of instrumentation in the inlets of the assemblies.................................2-19

Figure 2.25 Details of the positioning of the hot wire anemometers and pool rack 

pressure ports.................................................................................................2-20

Figure 2.26 TC layout on the external faces of the pool rack .............................................2-21

Figure 2.27 TC and light pipe layout on the exterior of the thermal radiation barrier...........2-21

Figure 2.28 Definition of the equivalent cross-flow paths at the bottom of the 

pool rack .........................................................................................................2-22

Figure 2.29 Overall pressure drop as a function of average air velocity in the 

assembly for all peripheral assemblies............................................................2-23

Figure 2.30 Overall pressure drop as a function of average air velocity in the 

assembly for the center assembly...................................................................2-24
viii

Figure 2.31 Pressure drop between the center and peripheral assemblies as a 

function of average cross-flow velocity............................................................2-25

Figure 2.32 Schematic showing planned locations of gas sampling lines...........................2-26

Figure 2.33 Schematic of gas sample conditioning system................................................2-27

Figure 3.1 Peak cladding temperatures in the center and peripheral bundles during 

Phase II pre-ignition testing at z = 3.150 m [124 in] and t = 12 h.......................3-3

Figure 3.2 Flow rates in the center and peripheral bundles during Phase II 

pre-ignition testing at t = 12 h............................................................................3-3

Figure 3.3 Peak cladding temperatures as a function of time at 1.0 kW simulated 

decay power for the center and peripheral bundles...........................................3-5

Figure 3.4 Induced flow rates as a function of time at 1.0 kW simulated decay 

power for the center and peripheral bundles .....................................................3-5

Figure 3.5 Peak cladding temperatures as a function of time at 4.0 kW simulated 

decay power for the center and peripheral bundles...........................................3-6

Figure 3.6 Induced flow rates as a function of time at 4.0 kW simulated decay 

power for the center and peripheral bundles .....................................................3-6

Figure 3.7 Peak cladding temperatures as a function of time at 8.0 kW simulated 

decay power for the center and peripheral bundles...........................................3-7

Figure 3.8 Induced flow rates as a function of time at 8.0 kW simulated decay 

power for the center and peripheral bundles .....................................................3-7

Figure 4.1 Peak cladding temperature in the center and peripheral assemblies 

during the ignition test.......................................................................................4-1

Figure 4.2 Measured flow rates in the center and peripheral assemblies during the 

ignition test........................................................................................................4-2

Figure 4.3 Oxygen concentrations of the center and peripheral assemblies during 

the ignition test..................................................................................................4-3

Figure 4.4 Temperature in the middle of the center assembly for different 

axial heights......................................................................................................4-4

Figure 4.5 Maximum temperatures in the peripheral assemblies of the rods 

closest from the center assembly for different axial heights ..............................4-4

Figure 4.6 Temperature contours at the midplane of the assembly and different 

times during the ignition test .............................................................................4-5

Figure 4.7 Maximum temperatures in the peripheral assemblies of the rods closest 

and furthest from the center assembly along the plane of symmetry.................4-6

Figure 4.8 Temperature contours throughout the assembly at z = 2.540 m [100 in.] 

and different times during the ignition test.........................................................4-7

Figure 4.9 Indicated pressures in the peripheral assemblies of the rods closest 

(J-1) and furthest (J-17) from the center assembly along the plane 

of symmetry ......................................................................................................4-8

Figure 4.10 Photograph of ballooned and oxidized sections of cladding from the 

SFP Phase II ignition test..................................................................................4-8

Figure 4.11 Mole fraction concentration of major and minor constituents of 

sample gas .......................................................................................................4-9

Figure 4.12 Normalized argon and helium ratios for nitrogen and oxygen..........................4-10

Figure 4.13 Photographs of the test assembly during various stages of deconstruction.....4-12

Figure 5.1 CVH and COR nodalization of PWR test assemblies in 1 by 4 

configurations (side view)..................................................................................5-7

Figure 5.2 CVH and COR Nodalization of PWR test assemblies in 1 by 4 

configuration (top view).....................................................................................5-8

Figure 5.3 MELCOR radial layout of 2-ring model..............................................................5-9

Figure 5.4 MELCOR 2-ring model nodalization................................................................5-10
ix

Figure 5.5 MELCOR 2-ring 9-rod model of the PWR SFP Experiment.............................5-10

Figure 5.6 Comparison of the MELCOR breakaway timing fit to Zr-4 data. ......................5-12

Figure 5.7 Cladding temperatures in peripheral assemblies for 15 kW ignition test 

at 6 hours........................................................................................................5-13

Figure 5.8 Center assembly peak cladding temperature for the 1 kW pre-ignition test .....5-16

Figure 5.9 Peripheral assembly peak cladding temperature for the 1 kW 

pre-ignition test ...............................................................................................5-16

Figure 5.10 Central assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours ...............................................................................................5-17

Figure 5.11 Peripheral assembly temperature versus elevation for the 1 kW 

pre-ignition test at 24 hours.............................................................................5-17

Figure 5.12 Central assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours ...............................................................................................5-18

Figure 5.13 Peripheral assembly temperature versus elevation for the 1 kW 

pre-ignition test at 24 hours.............................................................................5-18

Figure 5.14 Central assembly volumetric flow rate for the 1 kW pre-ignition test................5-19

Figure 5.15 Total peripheral assembly volumetric flow rate for the 1 kW 

pre-ignition test ...............................................................................................5-19

Figure 5.16 Center assembly peak cladding temperature for the 8 kW 

pre-ignition test ...............................................................................................5-20

Figure 5.17 Peripheral assembly peak cladding temperature for the 8 kW 

pre-ignition test ...............................................................................................5-20

Figure 5.18 Central assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours ...............................................................................................5-21

Figure 5.19 Peripheral assembly temperature versus elevation for the 8 kW 

pre-ignition test at 12 hours.............................................................................5-21

Figure 5.20 Central assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours ...............................................................................................5-22

Figure 5.21 Peripheral assembly temperature versus elevation for the 8 kW 

pre-ignition test at 12 hours.............................................................................5-22

Figure 5.22 Central assembly volumetric flow rate for the 8 kW pre-ignition test................5-23

Figure 5.23 Total peripheral assembly volumetric flow rate for the 8 kW 

pre-ignition test ...............................................................................................5-23

Figure 5.24 Peak cladding temperature in the center assembly for the 15 kW 

ignition test......................................................................................................5-27

Figure 5.25 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test......................................................................................................5-27

Figure 5.26 Peak cladding temperature in the center assembly for the 15 kW 

ignition test......................................................................................................5-28

Figure 5.27 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test......................................................................................................5-28

Figure 5.28 Central assembly ignition propagation for the 15 kW ignition test....................5-29

Figure 5.29 Peripheral assembly ignition propagation for the 15 kW ignition test...............5-29

Figure 5.30 Central assembly temperature versus elevation for the 15 kW ignition 

test at 5 hours .................................................................................................5-30

Figure 5.31 Peripheral assembly temperature versus elevation for the 15 kW ignition 

test at 5 hours .................................................................................................5-30

Figure 5.32 Central assembly temperature versus elevation for the 15 kW ignition 

test at 6 hours .................................................................................................5-31

Figure 5.33 Peripheral assembly temperature versus elevation for the 15 kW ignition 

test at 6 hours .................................................................................................5-31
x

Figure 5.34 Radial temperature distribution at 1.93 m [76 in] for the 15 kW ignition 

test at 5 hours .................................................................................................5-32

Figure 5.35 Radial temperature distribution at 3.45 m [136 in] for the 15 kW ignition 

test at 5 hours .................................................................................................5-32

Figure 5.36 Central assembly volumetric flow rate for the 15 kW ignition test....................5-33

Figure 5.37 Peripheral assembly volumetric flow rate for the 15 kW ignition test ...............5-33

Figure 5.38 CPU times for the MELCOR simulations of the 15 kW ignition test .................5-34

Figure 5.39 Flow rate in central assembly for 1 kW pre-ignition test using the 

2-ring model....................................................................................................5-36

Figure 5.40 Flow rate in peripheral assemblies for 1 kW pre-ignition test using the 

2-ring model....................................................................................................5-36

Figure 5.41 PCT in center assembly for 1 kW pre-ignition test using the 2-ring model.......5-37

Figure 5.42 PCT in peripheral assemblies for 1 kW pre-ignition test using the 

2-ring model....................................................................................................5-37

Figure 5.43 Flow rate in central assembly for 8 kW pre-ignition test using the 

2-ring model....................................................................................................5-38

Figure 5.44 Flow rate in peripheral assemblies for 8 kW pre-ignition test using the 

2-ring model....................................................................................................5-38

Figure 5.45 PCT in center assembly for 8 kW pre-ignition test using the 2-ring model.......5-39

Figure 5.46 PCT in peripheral assemblies for 8 kW pre-ignition test using the 

2-ring model....................................................................................................5-39

Figure 5.47 Flow rate in central assembly for 15 kW ignition test using the 

2-ring model....................................................................................................5-40

Figure 5.48 Flow rate in peripheral assemblies for 15 kW ignition test using the 

2-ring model....................................................................................................5-40

Figure 5.49 PCT in center assembly for 15 kW ignition test using the 2-ring model ...........5-41

Figure 5.50 PCT in peripheral assemblies for 15 kW ignition test using the 2-ring 

model..............................................................................................................5-41

Figure 5.51 Modification of the default MELCOR oxidation kinetics model.........................5-44

Figure 5.52 Peak cladding temperature in the center assembly for the 15 kW 

ignition test......................................................................................................5-45

Figure 5.53 Peak cladding temperature in the center assembly for the 15 kW 

ignition test......................................................................................................5-45

Figure 5.54 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test......................................................................................................5-46

Figure 5.55 Total oxidation heat generation rate in core for the 15 kW ignition test............5-46

Figure 5.56 Center assembly peak cladding temperature DT/dt for the 15 kW 

ignition test......................................................................................................5-47

Figure 5.57 Central assembly ignition propagation for the 15 kW ignition test....................5-47

Figure 5.58 Peripheral assemblies ignition propagation for the 15 kW ignition test ............5-48
xi

LIST OF TABLES

Table 2.1 Masses of components in the SFP project Phase II test assembly....................2-4

Table 2.2 List of equipment used for power control.........................................................2-12

Table 2.3 Hydraulic dimensions for the center assembly, average peripheral 

assembly, and the cross-flow path ..................................................................2-23

Table 2.4 Hydraulic loss coefficients of the SFP Phase II apparatus...............................2-25

Table 2.5 List of RGA calibration gases..........................................................................2-28

Table 3.1 Summary of pre-ignition testing results for Phase II at z = 3.150 m 

(124 in.) ............................................................................................................3-2

Table 5.1 Summary of main facility design parameters for the Phase II tests....................5-2

Table 5.2 Radiation view factors used in the MELCOR model..........................................5-6

Table 5.3 Ignition times and locations for the 15 kW ignition test. ...................................5-26

Table 5.4 Ignition times for the 15 kW ignition test..........................................................5-44

Table 6.1 Summary events during the SFP Phase II ignition test......................................6-2

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 Organisation for Economic Co-operation and 

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

agreement called the “OECD/NEA SFP 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.

The study was conducted in two phases; Phase I focused on axial heating and burn propagation 

in a single assembly and Phase II (documented in this report) 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 1×4 loading pattern. 

These experiments, completed in June 2012, were performed at Sandia National Laboratories.

In this second experimental phase, electrically resistive heaters with Zircaloy cladding were 

placed in the center cell of a 3×3 pool rack and surrounded by four unheated fuel assemblies. 

This arrangement imitated the situation of a recently offloaded assembly surrounded by much 

older and thus lower decay heat assemblies. Two of the peripheral fuel assemblies were 

constructed with pressurized rods that ballooned when sufficiently heated, and the other two 
xiv

fuel assemblies were constructed with vented rods that did not balloon. Nine pre-ignition tests 

were conducted with powers ranging from 1 kW to 15 kW, representing assemblies from as 

young as 3 months after offload to as old as 8 years, assuming 45 GWd/MTHM burnup. The 

pre-ignition tests were conducted for 12 hours or until a peak cladding temperature of 900 K

[1,160°F] was reached in the center assembly. Tests conducted at powers greater than 8 kW 

were terminated before 6 hours. Tests at 1 kW and 4 kW were conducted for 24 hours in 

attempts at achieving steady state. As a measure of repeatability, the 4 kW test was performed 

twice, once for 12 hours and again for 24 hours.

For the ignition test, the total applied power to the center assembly was 15 kW, which 

represented an assembly with a 45 GWd/MTHM burnup approximately 3 months after offload. 

All five fuel assemblies were completely consumed as a result of the Zircaloy cladding fire. The 

thermal-hydraulic behavior of the fuel assemblies was monitored during the ignition test as well 

as the timing of the ballooning of the pressurized rods. Finally, the depletion of oxygen and 

nitrogen in the exhaust stream were also directly measured.

Ignition of the Zircaloy was first observed in the center fuel assembly at the z = 3.302 m [130 in]

level and at an elapsed test time of 6.31 hours based on a thermal criterion of 1,200 K

[1,700°F]. The cladding fire also propagated transversely into the peripheral assemblies starting 

at 7.08 hours and at the z = 3.150 m [124 in] level. The ignition had progressed across the 

entire cross section of all of the peripheral assemblies by 8.74 hours. The ballooning events in 

the two assemblies occurred within approximately 3 minutes of each other at a given location 

despite the North and South assemblies being filled to pressures of 6.2 MPa [900 psi] and 

5.2 MPa [750 psi], respectively. The thermal-hydraulic behavior of assemblies with ballooned 

rods did not appear to be discernible from the unpressurized assemblies.

The gases measured at the exhaust were nitrogen, oxygen, argon, water, carbon dioxide, and 

helium. The analysis shows two distinct stages of oxidation. The primary oxidation stage was 

very energetic and continued for 11 hours after ignition. The second stage continued for 

88 hours after the first stage. During the first stage all oxygen entering the assembly was 

consumed by the zirconium oxidation and a substantial portion of the nitrogen was consumed 

by production of the nitride from much (if not most) of the remaining unoxidized zirconium. 

During the second stage, all of the oxygen entering the assembly was consumed by the 

oxidation of zirconium nitride as indicated by the release of nitrogen gas.

The additional release of energy from the production of zirconium nitride, along with the 

significant reduction in convective energy removal during the primary oxidation stage are 

phenomena to be considered when assessing the propagation of heat and fire to adjacent 

assemblies during a loss of water inventory accident in SFPs.

Several MELCOR models were developed to simulate the PWR spent fuel assembly 

experiment. These models include both 2-ring and 9-ring variants. The 9-ring model, where 

assemblies are subdivided by rows of fuel rods, provides a more detailed modeling of the 

experiment to be used in both assessing the MELCOR models as well as the limitations 

inherent in the 2-ring model. A 2-ring model is preferred as any reasonable full-scale modeling 

of a reactor plant spent fuel pool would be simulated with such a relatively coarse grid due to 

computer performance and model complexity issues. The MELCOR models were used to 

simulate both the pre-ignition and ignition tests.
xv

The conditions of this experiment challenge the MELCOR approach to representing an 

assembly (or collection of assemblies) with a single MELCOR node (or ring) consisting of a 

single lumped parameter representing all fuel rods in that node. In particular, a single 

representative temperature for the fuel rods in the unheated assembly is insufficient to capture 

the large temperature gradients and the dependence of oxidation rates on the local clad 

temperature. Consequently, the MELCOR code was modified to allow improved temperature 

refinement within a single radial ring by providing a means for the user to specify multiple fuel 

rod types within that ring. Note that this work is a proof of concept and that the models are 

currently valid up until material begins to relocate. Further model development and refinement 

is necessary prior to these modifications becoming part of an official release version of 

MELCOR.

The resulting analysis shows that MELCOR is capable of predicting the conditions of the 

OECD/NEA SFP experiment with either a 9-ring or 2-ring multiple rod model. These models 

show good agreement to important parameters such as PCT, ignition time and ignition 

propagation. The new 2-ring multiple rod model has also been compared with the 9-ring model 

and comparisons show that the new modeling is able to predict a similar response but with 

fewer COR rings. As a result, code performance is improved and the input model is greatly 

simplified.

xvii

ABBREVIATIONS

ANL Argonne National Laboratory

BWR boiling water reactor

CFD computational fluid dynamics

DAQ data acquisition

DH hydraulic diameter

GWd gigawatt-days

ID inside diameter

IFM intermediate fluid mixer

MELCOR severe accident analysis code

MTHM metric ton of heavy metal

NEA Nuclear Energy Agency

NRC Nuclear Regulatory Commission

OD outside diameter

OECD Organisation for Economic Co-operation and Development

PCT peak cladding temperature

PWR pressurized water reactor

PID proportional-integral-derivative

RGA residual gas analyzer

SCR silicon controlled rectifier

SEM secondary electron multiplier

SFP spent fuel pool

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

SNL Sandia National Laboratories

TC thermocouple

1-1

1 INTRODUCTION

1.1 Objective

The objective of the Phase II testing was to perform a highly detailed thermal-hydraulic 

characterization of five full-length, commercial 17×17 pressurized water reactor (PWR) fuel 

assembly mock-ups to provide data for the direct validation of MELCOR or other appropriate 

severe accident codes. Phase I testing concluded with ignition of a single PWR assembly in 

12.7 hours using 5 kW power. The present PWR characterization is similar to that conducted 

for the previous boiling water reactor (BWR) study and will lead to a full-scale multi-PWR fire 

test where the zirconium alloy cladding is heated in air to ignition and the fire propagates to 

adjacent assemblies (Ref. 1). This report details the findings from the Phase II testing of the 

Spent Fuel Pool (SFP) Project.

1.2 Testing Outline

The SFP project was conducted in two phases. The Phase I test series was conducted using a 

single PWR assembly and has been concluded (Ref. 1). Phase II focused on transverse

heating and burn propagation in five full-length assemblies. The fuel assemblies were arranged 

in a pool rack with the heated assembly in the center cell. The four unheated peripheral fuel 

assemblies shared a cell wall with the center assembly, representing older spent fuel. All mock 

fuel assemblies were constructed with zirconium alloy cladding and prototypic structural 

components. The center assembly was constructed with electrically resistive heaters. The 

thermal mass of the compacted MgO powder, used to make the electric heater, is an excellent 

match to spent fuel as demonstrated in the previous BWR study. The peripheral assemblies 

were loaded with MgO pellets in order to closely match the thermal mass of spent fuel. Two of 

the four peripheral assemblies were pressurized with argon to simulate ballooning of the fuel 

clad during the ignition test.

A series of pre-ignition tests were conducted to build a database of the thermal-hydraulic 

response of fuel below ignition temperatures. Figure 1.1 shows the decay power of a PWR fuel 

bundle as a function of time from discharge. These values were calculated using the 

normalized decay heats provided in Ade and Gauld (Ref. 2). Several burnups are shown in the 

graph to orient the reader, but a value of 45 GWd/MTHM was chosen as the reference burnup 

because it approximates the current industry average (Ref. 3). The range of the pre-ignition 

tests represents fuel aged between 2.8 months to 8 years after discharge for the reference 

burnup of 45 GWd/MTHM, or 15.0 to 1.0 kW decay heats, respectively. The ignition test was 

conducted at a decay heat of 15.0 kW.
1-2

Figure 1.1 Decay heat of a PWR 17×17 assembly as a function of time from discharge

The baseline, testing parameters for Phase II are summarized below.

• Characteristic pool cell size – 224.2 and 222.7 mm [8.83 and 8.77 in] average, as-built

inner dimensions of the center and peripheral assemblies, respectively

• Pre-ignition test powers – 1 to 15 kW (Simulated 96 to 2.8 months from offload for 

45 GWd/MTHM, respectively)

• Ignition test power – 15 kW electrical input (Simulated 2.8 months from offload for

45 GWd/MTHM)

The ignition test determined the location in the center assembly where ignition first occurred and 

the nature of the burn propagation throughout the neighboring fuel assemblies. As shown in 

Figure 1.2, the unpowered peripheral assemblies experimentally represented a “cold neighbor” 

situation and demonstrated the potential of zirconium fires initiated in the higher decay power 

assemblies to spread through the remainder of a fully populated pool. The remainder of this 

report outlines the details of the Phase II pre-ignition and ignition test results.

Pre-ignition Tests 

Ignition Test 
1-3

Figure 1.2 Approximate pool loading scheme for a 1×4 test layout with an insulated 

boundary condition

1×4 Pattern Repeated Phase II Layout

Insulated

Boundary

1 Heated Center 

Assembly

4 Unpowered 

Surrounding 

Assemblies

2-1

Spacer

2 APPARATUS AND PROCEDURES

2.1 General Construction

The testing in the Spent Fuel Pool (SFP) Project Phase II focused on the nature of the 

zirconium fire in the five full-length fuel bundles and specifically the propagation of the fire from 

the center to the peripheral bundles. The test assembly prototypically represented five

commercial 17×17 PWR fuel bundles. The various components comprising a typical 17×17 

PWR assembly are illustrated in Figure 2.1. 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.

Figure 2.1 Various components in a typical 17×17 PWR fuel assembly

The four peripheral PWR assemblies were fabricated using prototypic, commercial 17×17 PWR 

components including a skeleton, top and bottom nozzles, 9.50 mm [0.374 in] OD by 0.57 mm 

[0.0225 in] wall tubing, top and bottom end plugs, and plenum springs. The center heated PWR 

assembly was fabricated using prototypic, commercial 17×17 PWR components and 9.53 mm 

[0.375 in] OD heater rods made from 11.18 mm [0.44 in] zirconium alloy tubing supplied by an 

industrial vendor. The wall thickness of the Zircaloy-2 cladding was approximately 0.71 mm 

[0.028 in] and was relatively unchanged by the heater fabrication process. The spent fuel rod 

simulators for Phase II had a linear power profile and a maximum output of 31.1 W/m [9.5 W/ft], 

which was twice that expected to produce ignition.

An important attribute of the mock fuel designs was the fact that the thermal mass of the 

magnesium oxide (MgO) powder used to simulate the spent fuel was virtually the same as spent 

fuel over the entire temperature range of interest (as shown in Figure 2.2). The thermal mass 

was estimated from the measured solid fraction of the MgO used in the heater rods and 

peripheral rods and the specific heat data for MgO compiled in Touloukian Vol. 5, 1970 (Ref. 4). 

These spent fuel simulators therefore heated at approximately the same rate and stored the 

same amount of thermal energy as prototypic spent fuel rods. The magnesium oxide in the 

heated rods was constructed by compacting MgO powder around a coil of electrically resistive 

Nichrome wire. This compacted ceramic powder also forms the electrical insulation between 

the central heating element and the Zircaloy-2 cladding. The peripheral assemblies were 

constructed using sintered MgO pellets, which were inserted into Zircaloy-4 cladding. The 

distinction between the two types of mock spent fuel were due to differences in the as-built 

densities of the MgO inside the rods. The symbols in the plot for the Phase II heater and 

peripheral rods represent the measured, average value for each design.

Bottom 

nozzle

Guide 

tubes Top 

nozzle
2-2

Figure 2.2 Thermal mass per unit length comparison of spent fuel and MgO fuel rod 

simulators

Next, the thermal conductivity of the compacted MgO powder was compared to the thermal 

conductivity of spent fuel and fresh UO2 fuel. The thermal conductivity is of lesser importance 

due to the short conduction path between the heater element and the outer cladding and the 

slow heat rate of the simulated accident. Figure 2.3 compares the estimated thermal 

conductivity as a function of temperature for the compacted MgO used in the heater rods, the 

UO2 used by MELCOR, and spent fuel from FRAPCON-3. The spent fuel estimate was based 

on corrections applied to the MELCOR UO2 data as described in NUREG/CR-6534 Vol. 2. The 

thermal conductivity of compacted MgO powder is a function of the degree of compaction. The 

thermal conductivity of the compacted MgO used in the heater rods was estimated based on the 

measured solid fraction of MgO and extrapolation of data compiled in Touloukian Vol. 2, 1970. 

At low temperatures the thermal conductivity of the compacted MgO is about a third of the 

estimated thermal conductivity of spent fuel but as the temperature increases the comparison 

converges. At high temperatures the thermal conductivity of the compacted MgO is only 15 

percent lower than spent fuel or fresh UO2 fuel.
2-3

Figure 2.3 Comparison of thermal conductivity for UO2, spent fuel, and MgO fuel rod 

simulators

Table 2.1 gives the individual and cumulative masses of all components within the SFP Phase II 

test assembly. Major sub-assemblies include the pool rack, PWR fuel skeletons, peripheral 

rods, and heater rods. The mass of thermocouples within the assembly is neglected. Note that 

the mass of Zircaloy in the center and peripheral assembly differs due to the different wall 

thickness of the Zr-4 tubing used in the peripherals and the Zr-2 tubing used to make the heater 

rods in the center assembly.
2-4

Table 2.1 Masses of components in the SFP project Phase II test assembly

Part Name Material Quantity Individual Part Weight (kg) Total Part Weight (kg)

3×3 Pool Rack

Tie bar Stainless steel 32 0.041 1.31

Neutron absorber Aluminum 1100 16 5.77 92.32

Internal sheathing Stainless steel 16 6.70 107.19

Pool cell Stainless steel 5 57.35 286.74

Shim stock Stainless steel 8 0.866 6.93

Filler panel Stainless steel 4 16.63 66.51

Subtotal 561.00

Skeleton

Skeleton* ZIRLO 1 32.67 32.67

Bottom nozzle Stainless steel 1 5.26 5.26

Top nozzle Stainless steel 1 5.71 5.71

Debris catcher ZIRLO 1 1.43 1.43

Bottom bolts Stainless steel 24 0.007 0.18

Guide tube inserts Stainless steel 24 0.004 0.09

Subtotal 45.35

Peripheral Rods

Cladding Zr-4 1056 0.406 428.92

Top end plug Zr-4 1056 0.003 3.33

Bottom end plug Zr-4 1056 0.008 8.62

Plenum spring Stainless steel 1056 0.006 6.29

MgO ceramic MgO 1056 0.512 541.01

Subtotal 988.16

Heater Rods

Cladding Zr-2 264 0.539 142.20

MgO 142.25" heated + 2" unheated MgO 264 0.504 133.18

MgO 10" upper unheated "plenum" MgO 264 0.036 9.48

Subtotal 284.86

Phase II Assembly Total 1879.37

*

 Includes 24 guide tubes, 1 instrumentation tube, 7 full spacers, and 3 IFM spacers

2.1.1 Assembly Layout

Figure 2.4 shows the layout of the Phase II fuel bundles with and without the pool rack in place. 

The heated assembly was placed in the center pool cell and was surrounded by unheated 

assemblies. Two of these peripheral assemblies were constructed with pressurized fuel rods to 

simulate ballooning of the cladding during the ignition test.
2-5

Figure 2.4 Layout of the Phase II test assembly

2.1.2 Dimensions of the Test Assembly

The Holtec pool rack for 17×17 PWR fuel assemblies incorporates pool cells with an inner 

dimension of 224.5 mm [8.84 in]. Therefore, a 224.5 mm [8.84 in] pool cell was the design 

basis for the Phase II ignition test. As with the Phase I pool cells, the Phase II pool rack was

constructed of 1.91 mm [0.075 in] thick stainless steel material. The actual inner dimensions of 

each of the pool cells are shown in Figure 2.5. The inner dimension of the center cell, used in 

this experiment was 224.2 mm [8.83 in], which is nearly identical to the 224.5 mm [8.84 in]

Holtec design. The as-built peripheral cells are slightly smaller than the design with an average 

inner dimension of 222.7 mm [8.77 in], or less than a one percent difference. The insulation 

scheme for Phase II is also illustrated in Figure 2.5. The corner cells were filled with high 

temperature insulation. The entire assembly was surrounded by approximately 152.4 mm [6 in]

of the same high temperature insulation. A stainless steel thermal radiation barrier was installed 

around the external insulation.

Pressurized

Unpressurized

Unpowered

Heated
2-6

-

Figure 2.5 Dimensions of the as-built pool rack and the design insulation thickness

2.2 Design of the Heated Fuel Bundle

2.2.1 Heater Design

The design for the Phase II heater rods is shown in Figure 2.6. This design is identical to those 

used in Phase I except that linear power density was doubled for Phase II. The heated and 

unheated lengths for the fuel rod simulator and a PWR fuel rod are compared for reference. As 

shown in Detail A, the lower, unheated length of the heater included a steatite standoff 15.9 mm 

[0.625 in] for electrical isolation and an internal power introduction length 50.8 mm [2 in]. The 

height of the axial heated zone with respect to the top of the bottom nozzle was preserved as 

closely as possible between the heater rod 3.680 m [144.875 in] and the reference PWR fuel 

rod 3.683 m [145 in]. Due to the necessary electrical connection at the top, the heater rod 

design was extended 0.066 m [2.60 in] above the prototypic length. Details of the top and 

bottom electrical connections are given in Section 2.2.2.

North

South

West East

222.6

222.5

222.6

222.9 224.2

Corner cells 

filled w/ 

insulation

152.4

Insulation

(All dimensions in mm)
2-7

Figure 2.6 Design of the electrically heated fuel simulators

Note: The vertical dimension in the plan view has been scaled 6:1 to show the heated and 

unheated zones. PWR fuel layout shown for reference.

2.2.2 Heater Rod Electrical Connections

Introducing electrical power into the assembly presented several engineering challenges. An 

electrical current was fed into the assembly at the top of the assembly by applying voltage, up to 

120 VAC, across the heater rods. The bottom electrical connections terminate in the bottom 

nozzle, which was tied to the circuit neutral leg. Electrically isolating the heater cladding, guide 

tubes, and pool cell was crucial to prevent a short circuit in the power loop.

The upper electrical bus plate and connections were specifically designed to survive at higher 

temperatures. Figure 2.7 shows a rendering of the design of the top electrical bus plates. The 

bus plates were cut from three pairs of nickel-copper alloy 400 (Monel) plates and replaced the 

top nozzle in the center fuel bundle. The bus plates were electrically divided into three zones of 

88 heater rods each to accommodate the higher power requirements of the Phase II experiment 

and facilitated installation of the bus plates onto the heater rod power posts. To insure even 

distribution of power to the heater rods, the power was introduced into each zone via three 

6.4 mm [0.25 in] Monel threaded rods as shown in Figure 2.7.
2-8

Figure 2.7 Design of the top electrical bus plate

Each bus plate zone was comprised of a pair of upper and lower plates. The heater rod 

connection collars for the 88 heater rods were open on the same quadrant corner on a given 

plate. The connection collars on the mating plate were open on the opposite quadrant. When 

installed, the 88 heater rod connection pins passed through the plate in the larger interstitial 

area between four pin connection collars. Then, the plate was slid diagonally until all of the 

heater rod pins were captured in the connection collars. The top and bottom plates were moved 

in opposite directions so once in position each heater rod pin has 360° contact with the 

connection collar as shown in Figure 2.8. The Monel threaded rods are truncated in this 

rendering. For the actual installation, the all-thread rods extended approximately 1 m above the 

bus plates before transitioning to a traditional copper power lead.

Figure 2.8 Assembly detail of the top electrical bus plate design

The top of the assembly is represented schematically in Figure 2.9. The top electrical bus 

plates were left out of the drawing to avoid confusion. Figure 2.9 shows the top of two heater 

rods and a guide tube in cross section. Steatite standoffs and an alumina sphere isolated the 

cladding and guide tubes from electrical contact with the top bus plates. The additional length 

of the heater rods over the prototypical value was required to place the electrical connection 
2-9

above the top of the guide tubes. A detailed drawing of the power connection, including the top 

bus plates, is shown in Figure 2.10.

Figure 2.9 Detail drawing of the top of the assembly showing two heater rods and a 

single guide tube

The top and bottom plates were held in place by the threaded Monel power rods as shown in 

Figure 2.10. On the lower plates, the power connection locations were threaded to receive the 

power rod. The power connection locations on the top plate were slightly oversized, through 

holes. The power rod was passed through the upper plate and threaded into the lower plate. A 

Monel nut and washer was tightened down onto the top plate to lock the plate together.

The bottom PWR nozzle served as the electrical bus for the neutral connections. Blind holes 

machined into the bottom nozzle accepted the connection pin of each heater rod as shown in 

Figure 2.11. High temperature silver contact grease was placed in each hole to insure good 

connectivity.
2-10

Figure 2.10 Cross-sectional detail drawing of the top electrical connection from the 

power feed into the heater electrical pins

Figure 2.11 Detail drawing of the bottom of the assembly showing the electrical 

connection of a single heater rod to the bottom nozzle (neutral)

Figure 2.12 shows the configuration of the power control system for the SFP Phase II test 

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

a LabVIEW graphical user interface. This set point signal was relayed to a proportional-integralderivative (PID) controller. The PID controller determined 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 received the control signal and allowed the 

prescribed electrical power into the resistive load of the test assembly.
2-11

Figure 2.12 Power control system and test circuits

Note: Power instrumentation and control shown only for “C” circuit. Actual installation includes 

instrumentation and PID control for all three circuits.

The internal layout of the instrumentation panel is shown in Figure 2.13. The transducers 

measured Watts, voltage, and current applied to the assembly. The signal from the Watt 

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

components used in this testing are listed in Table 2.2.

120 VAC

Source

Test Assembly

~30 kW @ 120 VAC

Instrumentation 

Signals

Current Signal

(5:200)

Instrumentation

Panel

8

EZ

PID

Voltage 

Signal

Power 

Feedback

Signal

Power Control Signal

Neutral

Power Set Point DAQ System

“A” circuit

“B” circuit

“C” circuit

SCR
2-12

Figure 2.13 Schematic of the instrumentation panel for “C” circuit

Table 2.2 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 ManufacturingPM6C1FJ1RAAAAA

SCR Power Controller Watlow Electric Manufacturing PC91-F25A-1000

2.3 Design of the Pressurized Peripheral Bundles

A significant technical challenge in Phase II of the SFP Project was the manufacture and 

installation of pressurized fuel simulators in two of four peripheral assemblies. Welding 

techniques were developed for the attachment of the prototypic end plugs to the rod cladding. A 

total of 528 pressurized fuel rods were manufactured for installation in two peripheral 

assemblies for Phase II pre-ignition and ignition testing. These efforts used prototypic test 

articles with Zircaloy-4 (Zr-4) cladding and end plugs.

An initial series of small-scale rodlets were produced to evaluate the acceptability of welds 

between the fuel end plugs and the cladding and were documented in the Phase II test plan

(Ref. 19). After the qualification of these welds during the pilot study, a production run of fulllength rods was commissioned. These rods contained magnesium oxide (MgO) pellets and a 

stainless steel spring installed inside of the cladding to simulate the thermal mass of spent 

nuclear fuel. Figure 2.14 compares the design of the Phase II peripheral rods with spent fuel. 

Due to a higher gas volume fraction in the simulated fuel region, the fill pressure of the Phase II 

peripheral rods was compensated in order to mimic the pressure of spent fuel at the time of 

cladding rupture.

Watt

Transducer

Voltage

Transducer

Voltage 

Signal

Current

Signal

Neutral

Signals 

to DAQ

Power 

Feedback 

Signal

Current

Transducer
2-13

Figure 2.14 Schematic of the internal geometry of a Phase II peripheral rod and spent 

fuel

2.3.1 Welding Techniques

2.3.1.1 Orbital Welds

Two orbital welds were required between the Zr-4 cladding and fuel end plugs, one at the top 

and another at the bottom of the fuel rod. Figure 2.15 shows a cross-sectional view of the setup 

for welding the top end plug onto the Zr-4 fuel rod. Prior to the actual orbital welding procedure, 

both rod ends were reamed and all parts were thoroughly cleaned. External shield gas was 

provided by the orbital welding equipment.

Figure 2.15 Top end plug weld setup

With the top end plug welded, the argon purge setup was removed. The rod was then loaded 

with a stainless steel spring and MgO ceramic pellets as shown in Figure 2.16. Again, the 

design plenum spacing was 203.2 mm (8 in.).

Figure 2.16 Loading of fuel rod with MgO surrogate fuel and stainless steel plenum 

spring

Figure 2.17 shows the reversal of purging system in which the top portion of the fuel rod was 

inserted and secured into the purging apparatus. After purging the interior of the rod with argon, 

3.66 m

“Fuel” length

203 mm

Plenum

A

A

B

B

Phase 2 Peripheral Rod

Spent Fuel

9.50 mm OD

0.11 mm A-A

B-B

17.7% gas 

volume 

fraction

8.36 mm ID MgO

Spent 

Fuel

44% gas volume fraction (plenum)

82.3% gas volume fraction (plenum)

g

56% gas

volume fraction

Teflon Ferrules Top End 

Plug

Orbital Weld

Argon
2-14

the bottom end plug was then inserted and welded to the Zr-4 cladding. The Phase II peripheral 

fuel rod was then ready for final pressurization and closure welding.

Figure 2.17 Bottom end plug weld setup

2.3.1.2 Closure Weld

The 0.64 mm [0.025 in] through hole in the top end plug was used to pressurize the rods and 

was then welded closed. Figure 2.18 shows the closure weld fixture (CWF). A TIG electrode 

was inserted and secured at a predetermined depth such that when the fuel rod was inserted a 

small gap existed between the electrode tip and top end plug surface. The assembled fuel rod 

was inserted until the chamfered portion of the top end plug was seated firmly against the 

chamfered opening of the hex bushing as shown in Figure 2.19.

Figure 2.18 Schematic of the closure weld fixture

Figure 2.19 shows a close up of the flow path of the fill gas. The hex bushing had two 

machined slots on opposite sides to accommodate gas flow. The hole in the top end plug 

provided the final opening for the pressurizing gas to pass through allowing purging and 

pressurization.

Bottom End 

Plug

Orbital Weld

Pressurizing Gas

Ceramic 

insulator

Sealing 

gland

Fusion 

weld Sealing gland

TIG electrode

Argon
2-15

Figure 2.19 Close-up of gas fill process and top end plug weld location

After all electrical connections for the welder were configured, the final closure weld of the top 

end plug was made under pressure. Figure 2.20 shows a cross-sectional rendering and a photo 

of an actual welded top end plug. At this point the rod was fully pressurized and sealed.

 

Figure 2.20 Top end plug closure weld

2.3.2 Argon Fill Gas

Initial planning called for the use of helium to pressurize the Phase II rods. Initial attempts to 

implement a helium fill system for use with the SFP Phase II pressurized rods were 

unsuccessful in making the closure weld. Discussions with welding experts at Sandia National 

Laboratories (SNL) indicated that the high pressure helium gas was likely preventing sufficient 

ionization between the TIG welding tip and the top plug of the fuel element. For reference the 

first ionization energy of helium is approximately 1.6 times higher than argon. This ionization 

was crucial for forming and sustaining an electrical arc between the weld tip and the work piece. 

SNL was unable to determine from the fuel vendor their method of pressurization and closure. 

It is possible that the vendor utilizes some other technology such as laser or resistance welding 

that is not affected by the increased cover gas pressure.

As an alternative, argon was used as a substitute gas. The thermal mass of argon in a 

pressurized rod is theoretically identical to that of a helium-filled rod. Although helium is almost 

one order higher with regards to thermal conductivity, the effect on the thermal response of the 

Phase II pressurized rods, which is on the order of seconds, was negligible compared to the 

heat up times during testing of hours. Furthermore, a significant fraction of the gas inside spent 

fuel rods is comprised of the fission product gases Xe and Kr which would lower the effective 

thermal conductivity of the helium gas mixture.

Pressurizing Gas
2-16

The SFP Phase II South and North pressurized assemblies were filled to initial charging 

pressures of 5.2 MPa [750 psi] and 6.2 MPa [900 psi], respectively. These values represent a

range of expected peak pressures typical of spent fuel in industry under similar accident 

conditions.

2.4 Instrumentation

The instrumentation of the SFP Phase II assembly was based on the designs proposed in the 

Phase II Test Plan (Ref. 19). The instrumentation included hot wire anemometers, oxygen 

sensors, residual gas analyzer (RGA), strain gauges, and thermocouples (TCs). The following 

sections give corrections for the as-built instrumentation and also document details not 

previously discussed.

2.4.1 Center Bundle Thermocouples

Approximately 160 TCs were located within the center bundle. The TCs were attached to the 

heaters and guide tubes by strapping the tip of the TCs with a small piece Nichrome shim stock 

to the rod. All TCs to be 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.

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

identifiers give positioning in inches, e.g. C_J-9_124 was installed in the center bundle on the 

middle instrument tube at z = 3.150 m [124 in]. Ignition in the PWR occurred near the top of the 

assembly. Therefore, most of the bundle TCs were routed out the bottom of the assembly. TCs 

located at or above z = 3.30 m [130 in] were routed out the top of the assembly.

Figure 2.21 shows the TC layout of the center bundle. This updated schematic represents a 90° 

counterclockwise rotation compared to the Test Plan. A majority of the TCs were installed on 

the outside surface of the guide tubes before any of the heater rods were installed. TCs were 

also installed on some outer row heater rods. Red denotes the location of high density arrays 

where TCs were located at 15.25 cm [6 in] intervals. Yellow denotes the location of medium 

density arrays where TCs were located at 30.5 cm [12 in] intervals. Green denotes the location 

of low density arrays where TCs were located at 61 cm [24 in] intervals. The A-1 location in the 

center fuel bundle was oriented to the Northwest corner of the test apparatus.
2-17

Figure 2.21 Thermocouple layout for the Phase II center, heated fuel bundle

Note: The center bundle was installed with rod A1 oriented nearest the Northwest corner of the 

center pool cell.

2.4.2 Peripheral Bundle Internal Instrumentation

The as-built installation for the TCs and strain gauges in the peripheral assemblies is the same 

as documented in the Test Plan but are repeated here for clarity. Similar to the center 

assembly, the peripheral assemblies were instrumented with TCs at 15.25 cm [6 in] spacing. 

Unlike the center assembly, which was ideally symmetric about both the local x- and y-axes, the 

peripheral assemblies are symmetric only along a single axis. Therefore, the TCs were placed 

throughout these assemblies to primarily measure the thermal gradient from the center-facing 

cell wall to the outer rack wall. As in the center bundle, most of the TCs were routed through 

the bottom of the fuel bundles. However, TCs located at or above z = 3.30 m [130 in] were 

routed out the tops of the bundles.

Figure 2.22 shows the layout of the thermocouples in the unpressurized, peripheral assemblies. 

An extra designator was added to distinguish the global orientation of each bundle, e.g. “E” for 

East. An example TC designator is E_U_J-9_124 for a TC installed in the center of the East, 

unpressurized bundle at z = 3.150 m [124 in].

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

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

1

2

3

4

5

6

7

8

9

10

11

12

13

14

15

16

17

Axial Levels = C-148,

C-100, C-52, C-4*

* - “A” TCs installed at 4.875

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

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

J9, J12, J15, J17, M12, M15,

N17, O14, R17, SYM C9

J9, O14, P9 J9, O14, P6, P9,

R1, R5, R9, SYM J15

J3, J9, O4, SYM D14

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 = C-124,

C-76, C-27

C-118, C-69, C-22 C-110, C-64, C-16 C-154.875, C-106,

C-58, C-10

A1, E1, F3, F6, J1, J3,

J6, J9, SYM P6, SYM P9

C9, D4, J9 A9, A13, A17, C9, C12, D14,

J9, SYM J1, SYM R13

D14, J9, J15

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

Guide tube Fuel rod simulator

y

x
2-18

Similarly, Figure 2.23 gives the layout of the internal instrumentation in the two pressurized, 

peripheral bundles. No TCs were attached to the mock fuel pins because welding to the 

pressurized rods was prohibited. In addition to TCs, twenty-four strain gauges (Vishay MicroMeasurements Model CEA-03-062UQ-350) were attached to the pressurized fuel pins as shown 

in the left diagram of Figure 2.23 in a 3-wire, quarter-bridge configuration. Twenty-three of the 

gauges were actively monitoring the strain in the cladding due to the internal pressure of the 

rods. These strain gauges were attached directly to the cladding in between the debris catcher 

elements between the z = 0.105 to 0.111 m [4.125 to 4.375 in] levels. The remaining strain 

gauge was attached to the nearby debris catcher as a control to quantify the apparent strain due 

to temperature changes experienced by the gauges.

Figure 2.22 Layout of internal instrumentation for the unpressurized, peripheral fuel 

bundles

Axial Levels = W-4.875, W-100,

E-27, E-124

W-52, W-148,

E-76

W-10, W-16, W-22, W-58, W-64, W-69,

W-106, W-110, W-118, W-154.875

E-34, E-40, E-46, E-83, E-88,

E-94, E-130, E-136, E-142

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

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

J1, J3, J6, J9, J12, 

J15, J17, O4, O14, P9, 

R1, R5, R9, R13, R17

J1, J3, J6, J9,

J12, J15, J17

Guide tube Fuel rod simulator

Thermocouple Center-facing wall

J1, J3, J6, J9, J12, 

J15, J17, D4, D14, C9, 

A1, A5, A9, A13, A17
2-19

Figure 2.23 Layout of internal instrumentation for the pressurized, peripheral fuel 

bundles

2.4.3 Inlet Instrumentation

The inlet of the fuel assemblies were instrumented with hot wire anemometers, TCs, and 

pressure ports. Figure 2.24 shows an isometric view of the lower test stand and two side views 

of a center and peripheral pipe. The top of each inlet pipe was located at z = -38 mm [1.5 in], 

where z = 0 mm [0 in] at the bottom of the fuel assemblies by definition.

Figure 2.24 Layout of instrumentation in the inlets of the assemblies

305

241

127

241

z = -38

TSI Hot 

Wires

Inlet TCs

Pressure

Ports

Peripherals

All dimensions in mm

Center

Axial Levels = S-4.875, S-100,

N-27, N-124

* Single strain gauges at 4.25

Double strain gauges at 4.125 and 4.375

S-52, S-148,

N-76

S-10, S-16, S-22, S-58, S-64, S-69,

S-106, S-110, S-118, S-154.875

N-34, N-40, N-46, N-83, N-88,

N-94, N-130, N-136, N-142

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

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

C9, D4, D14, J3,

J6, J9, J12, J15

J3, J6, J9, J12,

J15, O4, O14, P9

J3, J6, J9,

J12, J15

Guide tube Fuel rod simulator

Thermocouple Two strain gauges Single strain gauge Center-facing wall
2-20

A total of six hot wire anemometers were placed in each of the 154 mm [6.065 in] ID pipe that 

defines the inlet to the five test assemblies. One was located in the inlet pipe of each of the 

peripheral assemblies and two were located in the inlet pipe of the center assembly. Hot wire 

anemometers were chosen to measure the inlet flow rate because this type of instrument is 

sensitive and robust while introducing almost no unrecoverable pressure loss. A typical 

placement of the hot wire is shown in Figure 2.25. TSI Model 8455 hot wire anemometers were 

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. Also, the flow encountered a slight contraction of 127 mm [5 in] as it passed

from the inlet pipe through the base plate. This diameter of 127 mm [5 in] was chosen based on 

the designs of commercial pool racks.

Figure 2.25 also shows the placement of the four pressure ports on the outside perimeter of the 

pool rack to the peripheral fuel assemblies. These pressure ports were centered in the middle 

of the fuel assemblies at z = 11 mm [0.43 in]. The center assembly was not accessible and 

therefore did not have a pressure port on the pool wall. All pressure drop measurements for the 

center fuel assembly were collected using the pressure port located in the inlet pipe as shown in 

Figure 2.24.

Figure 2.25 Details of the positioning of the hot wire anemometers and pool rack 

pressure ports

2.4.4 External Instrumentation

A number of thermocouples (TCs) were attached to the outside of the pool rack. Figure 2.26

shows the placement of these TCs on each face of the pool rack. The installation was 

staggered around the pool rack in 0.305 m [12 in] increments. The TCs were placed primarily in 

the center of the face of the peripheral cell, which was recorded as Pool_Face_Axial-Height_0. 

Arrays of TCs at the same axial level were placed to record the thermal gradient from the center 

of the peripheral cell to the corner of the pool rack. These TCs were recorded as 

Pool_Face_Axial-Height_Transverse-Location, where the three additional transverse locations 

of interest were the peripheral-corner cell interface 111.8 mm [4.4 in], the center of the corner 

cell 226.1 mm [8.9 in], and the corner of the pool rack 340.4 mm [13.4 in].

914

305

A-A

Section A-A

TSI

Fixed 

location

64.3

154 ID 

(6.065 in.)

TSI anemometer

13.5

127

11

All dimensions in mm

z

Perimeter 

of pool rack

38

51

Peripheral

fuel assembly

Honeycomb element 

(3.2 mm cell ID)
2-21

Thermocouples were also attached to the outside of the stainless steel thermal radiation barrier 

as shown in Figure 2.27. Similar to the pool rack, these TCs were installed in staggered fashion 

in 0.305 m [12 in] increments. All these TCs were placed either centered on a peripheral pool 

cell face or in the corner of the assembly. Quartz light pipes (3 mm [0.12 in] diameter) fully 

penetrated through the radiation barrier and insulation to the pool rack on both the North and 

East faces at 0.305 m [12 in] increments as a means to optically observe the location of the 

ignition front.

Figure 2.26 TC layout on the external faces of the pool rack

Figure 2.27 TC and light pipe layout on the exterior of the thermal radiation barrier

0.61

0.61

0.61

0.61

0.61

0.61

South West

z

North East

0.61

0.61

0.61

0.61

0.61

0.61

0.305

TC

All measurements in meters

z 0.61

0.61

0.61

0.61

0.61

0.61

North East South West

0.305 0.61

0.61

0.61

0.61

0.61

0.61

0.305

TC Light Pipe TC and Light Pipe

All measurements

in meters
2-22

2.4.5 Pressure Loss Measurements

A series of isothermal, forced flow calibrations were conducted to characterize the hydraulics of 

the Phase II fuel assemblies. These results of these calibrations include the flow loss 

coefficients for the center and peripheral assemblies. As discussed in the previous section, all 

pressure measurements in the center assembly were collected via a pressure port installed on 

the inlet pipe. The pressure drop measurements in the peripheral assemblies were taken using

pressure ports installed in the inlet pipe and in the pool rack wall at z = 0.011 m [0.43 in]. 

Cross-flow paths between the center and peripheral assemblies at the bottom of the pool rack 

were discovered as part of these analyses and are characterized for implementation in modeling 

efforts.

Figure 2.28 shows a cross section of the as-built test apparatus with the fuel bundles omitted. 

The cross-flow paths are interpreted for these analyses as four holes at the bottom of the pool 

rack. These flow paths were a result of an imperfect seal at the interface of the pool rack and 

the test stand.

Figure 2.28 Definition of the equivalent cross-flow paths at the bottom of the pool rack

The hydraulics for the center and an average peripheral assembly are given in Table 2.3. The 

assumed hydraulics for the cross-flow paths are also defined. The cross-flow paths may 

alternatively be conceptualized as a small slit with dimensions of 1.2 × 51 mm [0.046 × 2 in]. 

These flow paths were not part of the original design but were instead an artifact of the 

apparatus construction.

North

South

West East

222.6

222.5

222.6

222.9 224.2

Corner cells 

filled w/ 

insulation

(All dimensions in mm)

Equiv. 

diam. = 8.7

4 places
2-23

Table 2.3 Hydraulic dimensions for the center assembly, average peripheral 

assembly, and the cross-flow path

Description Center Peripheral Cross-Flow Path

Inner Dimension 224.2 mm [8.83 in] 222.65 mm [8.77 in] 8.7 mm [0.34 in]

Flow Area 0.0287 m2 [44.49 in2] 0.0283 m2 [43.87 in2] 5.89E-5 m2 [0.0913 in2]

Hydraulic diameter, DH 11.7 mm [0.46 in] 11.6 mm [0.46 in] 8.7 mm [0.34 in]

Two pressure drops were recorded in the peripheral assemblies. The first pressure drop was 

between the bottom of the pool rack at z = 0.011 m [0.43 in] and atmosphere. The second 

pressure drop was between a pressure port located z = -0.165 m [6.5 in] from the bottom of the 

fuel assemblies in the 0.154 m [6.065 in] diameter inlet pipe and atmosphere. Between this 

lower pressure port and the one installed on the pool rack, the flow experiences a slight 

contraction followed immediately by an expansion as it passes through a 0.127 m [5 in] hole in 

the test stand plate with a thickness of 0.013 m [0.5 in].

Figure 2.29 shows the pressure drop in all four peripheral assemblies as a function of average 

assembly air velocity for the pressure port in the pool rack and inlet pipe. Because of the 

additional form losses from passing through the hole in the test stand plate, the pressure drop is 

slightly higher in the inlet pipe than just inside the assembly. These results account for the 

cross-flow leaving the peripheral assemblies into the center assembly.

Figure 2.29 Overall pressure drop as a function of average air velocity in the assembly 

for all peripheral assemblies
2-24

Figure 2.30 shows the pressure drop for the center assembly as a function of air velocity. The 

pressure measurement for the center assembly was only possible for the inlet pipe because the 

pool rack wall of the center cell was not accessible. Again, the effects of the cross-flow were 

incorporated into these results.

Figure 2.30 Overall pressure drop as a function of average air velocity in the assembly 

for the center assembly

The pressure drop across the center and peripheral assemblies is shown as a function of crossflow velocity in Figure 2.31. This velocity was defined as the air velocity across a single crossflow path. All four cross-flow paths were assumed to be equal for these analyses. The 

pressure drop across the center and peripheral pool cells was assumed to be completely from 

contraction and expansion.
2-25

Figure 2.31 Pressure drop between the center and peripheral assemblies as a function 

of average cross-flow velocity

Table 2.4 summarizes the hydraulic loss coefficients for the Phase II apparatus. The friction 

and form losses appear slightly higher in the center assembly than in the peripheral assemblies. 

To compare with previous results, the appropriate difference in Σk from the peripheral 

measurements, pipe to pool, should be subtracted from the corresponding center Σk. 

Therefore, the form losses in the center assembly for the overall fuel bundle pressure drop are 

68 and 31.7 for upper flow ranges of 300 and 600 slpm, respectively.

Table 2.4 Hydraulic loss coefficients of the SFP Phase II apparatus

Center Peripheral Cross-Flow

Pipe Pool Pipe Center-Peripheral

Upper Flow Rate (slpm) 300 600 738 300 508 300 508 16

Upper velocity (m/s) 0.238 0.476 0.560 0.238 0.400 0.238 0.400 5.904

SLAM 119 129 126 113 118 113 118 --

Σk 73.0 37.9 43.6 47.5 29.1 52.5 35.3 1.5

2.4.6 Exhaust Gas Analysis

2.4.6.1 Oxygen Sensors

Oxygen concentrations were continuously monitored (Advanced Micro Instruments, Model 65, 

Part 6ANA0056) at six locations at the top of the assemblies as shown in Figure 2.32. All 

sample lines consisted of stainless steel tubing and were placed below the top bus plates (or 

top nozzles in the peripherals), which was located 0.173 m [6.82 in] below the top of the pool 

rack. The center assembly was sampled at two locations, one near the center of the bundle 

(avoiding power leads) and the other from the annulus region between the power bus plate and 
2-26

the storage cell wall. Each of the four peripheral assembles were sampled from the center of 

the bundle just below the top nozzle. In addition to the oxygen sensor, the exhaust gas in the 

center bundle was also monitored by a residual gas analyzer, which is detailed in the next 

section.

Figure 2.32 Schematic showing planned locations of gas sampling lines

2.4.6.2 Residual Gas Analyzer

Under the appropriate conditions, nitrogen will react with zirconium to form zirconium nitride

(ZrN). Steinbrück (2009) has shown that the zirconium-nitrogen reaction occurs readily when 

oxygen is absent in the gas phase and simultaneously oxygen is present in the zirconium metal 

phase. In particular it was determined that oxygen-stabilized α-Zr(O) was the form of zirconium 

most reactive to nitrogen. After ignition in the present experiment, the conditions required for 

rapid nitriding were ideally present above the burn front as the burn front moved downward.

Nitrogen reactions within the burning center assembly were monitored in order to determine the 

degree and timing of zirconium nitride formation. Tracking nitrogen reactions required

simultaneous measurement of nitrogen and ambient noble gas concentrations. The noble 

gases behaved as inert tracers. If nitrogen was consumed to form zirconium nitride, the ratio of 

nitrogen to the noble gases decreased. When zirconium nitride was oxidized, nitrogen was 

released and the ratio of nitrogen to the noble gases increased. The primary noble gas 

considered was argon, but helium was also monitored as a backup. The ballooning peripheral 

rods were pressurized with ultra-high purity argon so when the rods ruptured the background 

concentration of argon in the CYBL vessel was altered. As a mitigation, fresh air was

introduced at the bottom of the CYBL vessel and the top floor of the CYBL building was actively 

ventilated. The total inventory of argon in the 528 pressurized rods was about half the amount 

of argon initially present in the air contained inside the CYBL vessel.

The sampling and analysis of the gas stream exiting the center assembly was performed with 

an in-house built sample conditioning system coupled to a mass spectrometer (Pfieffer Model 

GSD 300T). With the Pfeiffer GSD 300T RGA, gases were ionized by a tungsten filament. 

These ions were filtered by the quadrapole mass spectrometer and then detected and amplified 

by the secondary electron multiplier (SEM). The SEM produced an electric current for each 

unique ion species, which was converted into gas concentrations and gas concentration ratios.

Center bundle

Center annulus

Sampling line

RGA
2-27

The gases of interest and the ion masses monitored were helium (4.0), water (18.0), nitrogen 

(28.0 and 14.0), oxygen (32.0 and 16.0), argon (40.0) and carbon dioxide (44.0). The analysis 

of these eight ions was recorded every 10 seconds. At two hour intervals, ambient air was 

sampled for a few minutes for calibration purposes.

The sample was collected through a stainless steel tube near the center of the center assembly 

just below the bundle bus plate as depicted in Figure 2.32. Again, the bus plate was located 

0.173 m [6.82 in] below the top of the storage cell. The sample conditioning system shown in 

Figure 2.33 was designed to continuously pull the gas sample at ambient pressure (84.1 kPa 

[12.2 psia]) from the top section of the center test assembly with a small vacuum pump (GAST 

model DOA-107-AA). The sample was passed through one of two parallel in-line filters 

(Swagelok model SS-4TF-05) housing a 0.5 micron filter cup (Swagelok model SS-4F-K4-05) 

and then into a capillary line for analysis by mass spectrometry. The pressure drop across the 

filters was monitored with a 13.8 kPa [2 psi] differential pressure gauge. The gas pressure at 

the capillary tip was kept constant at 76.19 kPa [11.05 psia] with an electronic backpressure 

controller (Proportion Air model DQPV1TFEE020BXL). As long as the pressure drop across the 

filter was less than 6.8 kPa [1.0 psi] the back pressure controller was able to maintain a 

constant sample pressure. Block valves permitted isolation of a filter for replacement without 

compromising the sample gas analysis. A 3-way solenoid valve (ASCO model EF8320G202), 

which was controlled through the RGA software, was used to periodically switch from the gas 

sample stream to an ambient air stream . The ambient air stream was also regulated to 

76.19 kPa [11.05 psia] to serve as a calibration gas. It was drawn from inside the CYBL 

building, but outside the CYBL vessel.

Figure 2.33 Schematic of gas sample conditioning system

When significant oxygen was present in the sample before ignition and after all oxidation 

reactions were complete, the detector current was converted to a mole or volume fraction based 

on the analysis of ambient air. When oxygen was absent after ignition the detector output was 

converted to mole or volume fraction based on the analysis of the five calibration gases 

0.5 μm

Filters

Pressure controlled sample

Mass Spectrometer

Back

pressure 

controller

 

 

Vacuum 

pump

Vent to 

atmosphere

3-way control valve

Ambient

calibration air

By-pass needle valve

DP

gauge

Particle laden sample

84.1 kPa (12.2 psia)

~1.0 slpm

0.5 μm

filters Vent to

atmosphere

By-pass needle valve

3-way control valve

Vacuum

pump

Back 

pressure 

controller Pressure controlled sample

Mass Spectrometer

Ambient 

calibration air
2-28

described in Table 2.5. The calibration gases were designed to span nitrogen to argon ratios of 

27 to 140 which, when divided by the nitrogen to argon ratio found in ambient air correspond to 

normalized N2/Ar of 0.33 to 1.67. Normalized N2/Ar less than unity signifies nitrogen removal; 

normalized N2/Ar greater than unity signifies nitrogen release. During the Phase I ignition test, 

the normalized nitrogen to argon ratio was found to range from 0.7 to 0.4.

Table 2.5 List of RGA calibration gases

Calibration Gas Ar (%) Balance N2 (%) N2/Ar Normalized N2/Ar

1 3.54 96.46 27 0.33

2 1.76 98.24 55 0.67

3 1.18 98.82 84 1.00

4 0.89 99.11 112 1.33

5 0.71 99.29 140 1.67
3-1

3 PRE-IGNITION TEST RESULTS

A summary of Phase II pre-ignition testing is presented in Table 3.1. Only tests that extended to 

12 hours are shown in the table. Three additional pre-ignition tests were conducted at powers 

of 10.0, 12.0, and 15.0 kW but did not exceed 6 hours duration. This table gives the average 

temperatures and flow rates at 3, 6, 9, and 12 hours for the center and peripheral bundles at 

z = 3.150 m [124 in]. The minimum and maximum temperatures of the center and peripheral 

bundles at each time are also shown. Figure 3.1 shows the maximum temperature in the center 

and peripheral assemblies at the 3.150 m [124 in] level and a time of 12 hours as a function of 

applied power. As expected, the peak temperature monotonically increased in all assemblies 

with higher applied power. Figure 3.2 shows the induced flow at 12 hours in the center and 

peripheral assemblies as a function of applied power. The value shown for the peripherals in 

this plot and in Table 3.1 represent an average across all the peripheral assemblies. The flow in 

the center assembly increased with power at low powers, peaked at 4.0 kW then decreased at 

higher powers. The flow in the peripheral assemblies increased with power until 8.0 kW, which 

may be the peak flow. The flows in the center and peripheral assemblies appear to converge at 

higher powers.

The 4.0 kW test was repeated to examine repeatability as shown in the third and fourth columns 

of test entries in Table 3.1, as well as in Figure 3.1 and Figure 3.2. The maximum and average 

temperatures in the center bundle were within 3 K [5.4°F] at all positions and times, which is 

slightly greater than the experimental uncertainty of ±2.2 K [±3.96°F]. The minimum 

temperature in the center bundle was as much as 8 K [14.4°F] higher in the second test. In 

addition, the temperatures in the peripheral assemblies were as much as 14 K [25.2°F] higher in 

the repeat test. These thermal differences are likely due to the accumulation of oxide on both 

the Zircaloy-2 cladding in the center bundle and on the center stainless steel pool cell, which 

continued to accrue between the initial 4.0 kW test on 5/9/12 and the second test on 5/21/12. 

The thermal record in the peripheral fuel bundles did not show temperatures high enough to 

indicate significant oxide layer growth during the entirety of pre-ignition testing. However, 

increases in the oxide layers on the center bundle and center pool cell surfaces would 

effectively increase the emissivity of the heated surfaces and therefore increase the radiative 

heat transfer between the center and peripheral fuel bundles. With the exception of the 

peripheral measurements at 3 hours, the air flow rates were within 10 slpm for the two 

measurements, which is better than the experimental uncertainty of ±12 slpm. The 

measurements in the peripherals at 3 hours are considered outliers because the hot wires were 

effectively unresponsive below flow rates of approximately 50 slpm.
3-2

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

True RMS Power (kW) 1.0 2.0 4.0 4.0 6.0 8.0

Test Date 5/14/2012 5/7/2012 5/9/2012 5/21/2012 5/4/2012 5/18/2012

Center Flow Rate (slpm) 109 162 193 185 191 193

Peripheral Flow Rate (slpm) 0 25 37 2 62 73

Max. Temp. (K) 361 414 515 515 618 684

Avg. Temp. (K) 350 392 474 475 556 620

Min. Temp. (K) 335 364 429 434 490 556

Max. Temp. (K) 306 311 325 326 350 358

Avg. Temp. (K) 300 297 298 300 312 312

Min. Temp. (K) 298 294 291 294 301 298

Pool Cell Avg. Temp. (K) 299 295 294 296 305 302

Rad. Barrier Avg. Temp. (K) 298 293 294 299 300 294

Ambient Avg. Temp. (K) 297 292 294 299 300 294

Center Flow Rate (slpm) 140 182 198 191 185 185

Peripheral Flow Rate (slpm) 28 54 79 71 102 117

Max. Temp. (K) 404 487 621 618 741 812

Avg. Temp. (K) 382 450 567 567 672 745

Min. Temp. (K) 359 410 507 516 600 674

Max. Temp. (K) 317 334 362 371 406 462

Avg. Temp. (K) 304 305 314 318 340 356

Min. Temp. (K) 299 296 296 299 310 311

Pool Cell Avg. Temp. (K) 301 299 303 305 321 322

Rad. Barrier Avg. Temp. (K) 298 294 296 300 302 297

Ambient Avg. Temp. (K) 298 294 296 300 301 296

Center Flow Rate (slpm) 154 187 197 191 191 180

Peripheral Flow Rate (slpm) 49 70 103 97 132 141

Max. Temp. (K) 431 526 668 665 784 859

Avg. Temp. (K) 404 485 614 614 720 794

Min. Temp. (K) 377 440 553 560 652 727

Max. Temp. (K) 328 353 398 411 468 546

Avg. Temp. (K) 309 316 334 343 376 416

Min. Temp. (K) 302 302 309 313 333 345

Pool Cell Avg. Temp. (K) 304 306 317 321 344 359

Rad. Barrier Avg. Temp. (K) 299 295 298 301 302 300

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

Center Flow Rate (slpm) 159 187 194 191 189 175

Peripheral Flow Rate (slpm) 56 83 118 115 148 150

Max. Temp. (K) 449 547 690 687 802 883

Avg. Temp. (K) 419 504 637 637 742 822

Min. Temp. (K) 390 456 579 583 676 758

Max. Temp. (K) 336 368 427 440 503 595

Avg. Temp. (K) 316 328 355 369 410 472

Min. Temp. (K) 306 310 326 332 362 395

Pool Cell Avg. Temp. (K) 309 314 332 340 371 403

Rad. Barrier Avg. Temp. (K) 299 297 301 303 304 305

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

12 h

Center

Peripheral

9 h

Center

Peripheral

Center

Peripheral

3 h

6 h

Center

Peripheral
3-3

Figure 3.1 Peak cladding temperatures in the center and peripheral bundles during 

Phase II pre-ignition testing at z = 3.150 m [124 in] and t = 12 h

Figure 3.2 Flow rates in the center and peripheral bundles during Phase II pre-ignition 

testing at t = 12 h
3-4

The peak cladding temperature (PCT) and induced flow rate are respectively shown as a 

function of time in Figure 3.3 and Figure 3.4 for the center and peripheral assemblies operated 

at 1.0 kW simulated decay power. After 24 hours the PCT and flow in the center assembly 

appeared to reach steady state. After 24 hours the PCT and flow in the peripheral assemblies 

was still rising slowly. The flow in the four peripheral assemblies was measured independently 

and tracked in a nearly identical fashion.

Figure 3.5 and Figure 3.6 show respectively the PCT and induced flow as a function of time in 

the center and peripheral assemblies when powered at 4.0 kW. After 24 hours, the PCT in both 

the center and peripheral assemblies was still rising slowly while the flow in both had reached 

steady state. The 4.0 kW case was tested twice, once for 12 hours and again for 24 hours. For 

the 12-hour test conducted 12 days earlier than the 24-hour test, the peak cladding temperature 

in the center assembly agreed closely with the 24-hour test and ended only slightly higher. 

However, the peak cladding temperature in the peripheral assembly was noticeably lower at all 

times for the 12-hour test. This is consistent with increased radiative coupling between the 

center assembly and the inner pool rack walls as a result of increased emissivity produced by 

repeated heating during testing. The induced flow in both the center and peripheral assemblies 

for all times was greater during the 12-hour test. For clarity in the graph, the induced flow for 

the peripheral assemblies in Figure 3.6 is shown as the average of the four peripheral 

assemblies.

The PCT and induced flow for 8.0 kW simulated decay power are shown in Figure 3.7 and 

Figure 3.8, respectively. After 12 hours the PCT in both the center and peripheral assemblies 

was still increasing. At the same time, the induced flow was dropping in the center assembly 

and rising in the peripheral assemblies in convergent fashion. The four peripheral flow 

measurements were within the experimental error.
3-5

Figure 3.3 Peak cladding temperatures as a function of time at 1.0 kW simulated 

decay power for the center and peripheral bundles

Figure 3.4 Induced flow rates as a function of time at 1.0 kW simulated decay power 

for the center and peripheral bundles
3-6

Figure 3.5 Peak cladding temperatures as a function of time at 4.0 kW simulated 

decay power for the center and peripheral bundles

Figure 3.6 Induced flow rates as a function of time at 4.0 kW simulated decay power 

for the center and peripheral bundles
3-7

Figure 3.7 Peak cladding temperatures as a function of time at 8.0 kW simulated 

decay power for the center and peripheral bundles

Figure 3.8 Induced flow rates as a function of time at 8.0 kW simulated decay power 

for the center and peripheral bundles

4-1

4 IGNITION TEST RESULTS

Phase II of the OECD/NEA SFP Project concluded with the ignition of the test assembly on 

June 6, 2012. The test was started at 00:01:42 Mountain Daylight Saving Time. The total 

applied power to the center fuel assembly was 15.0 kW, which represents an assembly with a 

45 GWd burnup approximately 3 months after offload. The center assembly was surrounded on 

four sides by unpowered, peripheral assemblies. Ignition of the Zircaloy within the center 

assembly was first observed at the z = 3.302 m [130 in] level and at an elapsed test time of 

6.31 hours based on a thermal criterion of 1,200 K [1,700°F]. The first ballooning event in the 

pressurized, peripheral assemblies occurred at 6.67 hours. Based on the same criterion of 

1,200 K [1,700°F], the Zircaloy fire first propagated into the peripheral assemblies at an elapsed 

test time of 7.08 hours at the z = 3.150 m [124 in] level.

4.1 General Thermal-Hydraulic Response

Figure 4.1 gives the peak cladding temperature as a function of time for the center and 

peripheral assemblies. Both data series have been truncated due to the failure of the 

thermocouples at elevated temperatures. The center assembly first reached ignition based on a 

temperature of 1200 K [1,700°F] at 6.31 hours. The peripheral assemblies first registered 

ignition temperatures at 7.08 hours.

Figure 4.1 Peak cladding temperature in the center and peripheral assemblies during 

the ignition test

Figure 4.2 shows the measured flow rates in the center and peripheral assemblies as a function 

of time during the ignition test. The flow rates in the peripheral assemblies did not register 

accurately prior to about 1.4 hours due to the relative insensitivity of the hot wire anemometers 

at low flow rates. The first major disturbance in the flow occurs at approximately 6.7 hours and 

corresponds to the loss of electrical power in the center assembly. A sharp decrease in the 
4-2

peripheral assembly flow rates was recorded after 8.7 hours, which coincides with the time at 

which the ignition front had progressed across the entire cross section of the peripheral 

assemblies. Debris was observed exiting the bottom of the inlet pipes and may be the cause of 

the erratic signals at times greater than t = 10 h.

Figure 4.2 Measured flow rates in the center and peripheral assemblies during the 

ignition test

Figure 4.3 shows the oxygen concentration as measured in the center bundle, center annulus, 

and individual peripheral assemblies. The sampling pump was activated at 5.7 hours, after 

measurable oxygen depletion had already begun. The depletion of oxygen measured in the 

bundle of the center assembly corresponds with the ignition time of 6.31 hours. The peripherals 

began registering significant oxygen consumption at times greater than approximately 8 hours, 

which is when the burn front had neared the middle of the peripheral assemblies. None of the 

sensors registered complete depletion of oxygen, although the RGA does indicate nearcomplete consumption of oxygen immediately after ignition. Therefore, the O2 concentrations 

as measured by the oxygen sensors are considered unreliable after the sample registers less 

than 8 percent. The non-zero response of the oxygen sensors after ignition is not understood, 

but air entrainment at the top of the fuel assemblies into the sample lines is the most likely 

explanation. Evidence of air entrainment was also observed in the RGA samples.
4-3

Figure 4.3 Oxygen concentrations of the center and peripheral assemblies during the 

ignition test

4.2 Burn Propagation

4.2.1 Axial Propagation

The Zircaloy ignition started in the middle of the center assembly at the z = 3.302 m [130 in]

level. The burn front briefly propagated both upward and downward from this level but failed to 

reach the uppermost portion of the center assembly before becoming oxygen starved. The 

ignition front primarily swept down the assembly toward the source of oxygen at the assembly 

inlet. The burn progressed at an average rate of 0.457 m/h [18 in/h]. Figure 4.4 shows the 

temperature in the middle of the center assembly at different axial heights as a function of time.

Similarly, Figure 4.5 shows the axial propagation of the Zircaloy ignition in East and West 

assemblies. The cladding fire started at z = 3.150 m [124 in] level (not shown) and propagated 

almost exclusively downward. The fuel above this initial level in the peripheral assemblies 

exhibited heavy oxidation as evidenced during the post-test inspection but did not exhibit a 

thermal excursion typically associated with breakaway kinetics during the 13 hour test. As 

expected, the peripheral average burn rate of 0.466 m/h [18.35 in/h] nearly matches that of the 

center assembly.

Figure 4.6 depicts the manner in which the cladding fire spread through the fuel bundles. 

Initially, the burn front propagated preferentially downward in the center assembly. At late times 

in the test (t > 10 h), the burn fronts in the center and peripheral merged and appear as a planar 

front.
4-4

Figure 4.4 Temperature in the middle of the center assembly for different axial heights

Figure 4.5 Maximum temperatures in the peripheral assemblies of the rods closest 

from the center assembly for different axial heights

Center 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
4-5

Figure 4.6 Temperature contours at the midplane of the assembly and different times 

during the ignition test

t = 6 h t = 6.5 h t = 7 h t = 7.5 h t = 8 h

t = 8.5 h t = 9 h t = 9.5 h t = 10 h t = 12 h
4-6

4.2.2 Transverse Propagation

The cladding fire also propagated transversely into the peripheral assemblies starting at 

t = 7.08 h. Figure 4.7 shows the maximum temperatures of the instrumented rods closest to the 

center assembly in the unpressurized (J-1) and pressurized assemblies (J-3) and the furthest 

from the center assembly in the unpressurized (J-17) and pressurized assemblies (J-15). The 

ignition had progressed across the entire cross section of all of the peripheral assemblies by 

t = 8.74 h. Examination of these results reveals a degree of asymmetry in the propagation into 

the peripheral assemblies. The East J-3 rod reached the ignition criterion 0.26 hours after the 

North J-3 rod. Similarly, the difference in time between the rods further from the center 

assembly show a 0.27 hours spread between the first to ignite (South J-15) and the last to reach 

1,200 K [1,700°F] (North J-15).

Figure 4.7 Maximum temperatures in the peripheral assemblies of the rods closest 

and furthest from the center assembly along the plane of symmetry

Figure 4.8 gives temperature contours across the center and peripheral fuel assemblies at the 

z = 2.540 m [100 in] level at different times of interest. This figure further demonstrates the 

nature of the burn propagation in the transverse direction. These contours were constructed 

with all available thermocouple data at this axial level with interpolation between measurement 

locations. These data were reflected across the diagonal plane of symmetry to represent onefourth of the assembly geometry for visualization purposes and thus are artificially symmetric. 

As described earlier, the ignition front is known to have propagated asymmetrically into the 

peripheral assemblies by times of up to 0.27 hours but is ignored in these plots. Thermocouples 

that exceeded 1,480 K [2,204°F] were consumed but were assumed to remain at a temperature 

above this threshold for times after their destruction in the plots. The plots show that the 

cladding fire first propagated preferentially into the peripheral assemblies near the midplane and 

proceeded to burn from the center facing rod bank to the outermost rod bank in a nearly planar 

fashion.

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

Center
4-7

Figure 4.8 Temperature contours throughout the assembly at z = 2.540 m [100 in] and 

different times during the ignition test

4.3 Fuel Clad Ballooning

Figure 4.9 shows the indicated pressure in the rods closest (J-1) and furthest (J-17) from the 

center assembly. The North and South assemblies were initially filled to 6.2 MPa [900 psi] and 

5.2 MPa [750 psi], respectively. The ballooning events in the two assemblies occurred within 

approximately 3 minutes of each other at a given location, regardless of the initial pressure. 

Ballooning did not likely occur at a single axial level but rather between z = 2.692 and 3.302 m 

[106 and 130 in] based on PCT data at the time each rod ballooned. The thermal-hydraulic 

behavior of assemblies with ballooned rods did not appear to be discernible from the 

unpressurized assemblies.

A photograph of post-test ballooned segments of fuel are shown in Figure 4.10. The stacked 

arrangement in the photo should not be perceived by the reader as a dual ballooning event on 

the same rod. Rather, each remnant represents the ballooning of a single rod. Due to the 

chaotic nature of the debris field, the originating location of these samples could not be 

determined. The variety of samples do show a spectrum of ballooning events ranging from 

extreme tears in the cladding to localized gashes. All the samples show the additional effects of 

slumping, oxidation, and spallation. The surrogate fuel is still visible in some samples such as 

in the lower-middle artifact.

t = 6 h 7.436 h 7.75 h

8 h 8.052 h 8.152 h

8.271 h 8.472 h 8.6 h
4-8

Figure 4.9 Indicated pressures in the peripheral assemblies of the rods closest (J-1) 

and furthest (J-17) from the center assembly along the plane of symmetry

Figure 4.10 Photograph of ballooned and oxidized sections of cladding from the SFP 

Phase II ignition test

4.4 Exhaust Gas Analysis

A residual gas mass spectrometer (RGA) was used to monitor the concentrations of the gas 

constituents exiting the center assembly over a 130 hour period. The gas constituents 

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

Center
4-9

monitored were nitrogen, oxygen, argon, water, carbon dioxide and helium. The gas sample 

stream was conditioned by filtering particulates and the sample pressure was held constant by 

an electronic back pressure controller. Ambient air was automatically sampled at two hour 

intervals to correct changes in the instrument sensitivity, which drifted downward by a factor of 

about four for all gases. Calibration gases containing certified levels of argon with the balance 

nitrogen were used to correct the argon and nitrogen concentrations for the periods when 

oxygen was absent from the gas sample.

Figure 4.11 shows the sample gas composition on a dry basis as a function of time over the 

entire 130 hour test period. Also shown on Figure 4.11 is the total of all mole fractions that 

should be unity. Any deviation from unity is a measure of the total error in the analytical 

scheme. From time zero until ignition at 6.3 hours the oxygen concentration was at ambient 

levels. At ignition the oxygen concentration dropped sharply to near zero. From about 10 hours 

until about 16 hours the oxygen concentration was unstable apparently due to air infiltration into 

the gas sample. Shortly after ignition occurred in the center assembly the pressurized rods in 

the North and South peripheral assemblies began to balloon. A series of ruptures occurred 

from about 6.6 hours until 8.5 hours. Since the pressurized rods were filled with argon, each 

rupture caused a spike in the argon concentration. At about 7.5 hours the CYBL vessel 

extraction fan was energized which began purging the vessel with fresh air. The excess argon 

was slowly removed and the argon concentration returned to ambient levels by about 30 hours. 

At about 64 hours the oxygen level began to rise. By 105 hours all gas concentrations had 

returned to ambient levels.

Figure 4.11 Mole fraction concentration of major and minor constituents of sample gas

Since argon and helium are inert noble gases and do not react with zirconium, the concentration 

of either of these gases relative to the reactive gas species provides a measure of a reactive 

gas removal and subsequent release. Figure 4.12 shows the ratio of nitrogen and oxygen to 
4-10

argon (and helium for the case of nitrogen) normalized by the ratios found in ambient air. A 

normalized argon ratio of unity signifies no change from ambient air. Normalized ratios that are 

less than unity signify removal of a gas species. Normalized ratios greater than unity signify the 

release of a gas species. Due to the excess argon release during pressurized rod ballooning 

and the subsequent slow removal of this excess, the argon concentration between 6.5 and 

30 hours is elevated above ambient. The increased argon concentration artificially reduces the 

argon ratios. During this period, the helium ratio is more relevant even though the very low 

concentration is near the detection limit of the instrument and results in a noisy signal.

Figure 4.12 Normalized argon and helium ratios for nitrogen and oxygen

The O2/Ar plot closely mirrors the oxygen concentration plot in the previous figure. At 

6.31 hours when ignition occurs, the O2/Ar ratio drops sharply to near zero where it remains until 

just after 60 hours. Air infiltration into the gas sample between 10 and 18 hours is evident. 

Between 60 and 80 hours, the O2/Ar ratio rises rapidly to near ambient conditions. The return to 

ambient levels is complete at 105 hours signifying the end of all oxidation reactions.

Under the appropriate conditions, nitrogen will react with zirconium to form zirconium nitride 

(ZrN). Steinbrück (2009) has shown that the zirconium-nitrogen reaction occurs readily when 

oxygen is absent in the gas phase and simultaneously oxygen is present in the zirconium metal 

phase. In particular it was determined that oxygen-stabilized α-Zr(O) was the form of zirconium 

most reactive to nitrogen. The ratio of nitrogen to argon and helium was used to monitor the 

reaction of nitrogen with zirconium. The analysis shows two distinct stages of oxidation over the 

course of the test. The primary oxidation stage was very energetic and continued for over 

11 hours after ignition as defined by the transition of the nitrogen to helium ratio from less-than 

unity to greater-than unity. During this stage, all of the oxygen entering the assembly was 

consumed by the oxidation of zirconium as described in Equation 4.1. Also, a substantial 

portion of the nitrogen was consumed by production of the nitride from much, if not most, of the 

unoxidized zirconium remaining above the burn front as given in Equation 4.2.
4-11

Zr + O2 → ZrO2 ΔHrxn = -1,094.8 kJ/mol Zr 4.1

Zr + 1⁄2 N2 → ZrN ΔHrxn = -365.4 kJ/mol Zr 4.2

The nitride reaction releases about a third of the energy of complete oxidation; however, the 

mass of zirconium converted to nitride may be two to three times greater than the mass 

converted to oxide during the primary oxidation stage. The kinetics of Equation 4.2 has been 

studied by Steinbrück and Jung (2011). A strong reaction rate dependence on the oxygen 

content of the oxygen-stabilized -Zr(O) metal phase was found such that an oxygen content 

increase from 2 wt. percent to 6.5 wt. percent increased the reaction rate at 1,100°C [2,012°F] 

by nearly two orders of magnitude. The reaction rate for 6.5 wt percent O metal was found to 

exhibit a complex temperature dependence where the rate increased for temperatures up to 

1,200°C [2,192°F] but then decreased for higher temperatures.

All reactions consumed gaseous reactants so the flow exiting the assembly was substantially 

reduced, which may have allowed air infiltration into the exhaust gas sample. The additional 

release of energy from the wholesale production of the nitride coupled with the significant 

reduction in convective energy removal are important phenomena to consider when assessing 

the propensity of an ignited assembly to propagate the fire to adjacent assemblies.

During the second oxidation stage, all of the oxygen entering the assembly was consumed by 

the oxidation of zirconium nitride as indicated by the release of nitrogen gas in Equation 4.3.

ZrN + O2 → ZrO2 + 1⁄2 N2 ΔHrxn = -729.4 kJ/mol Zr 4.3

This stage continued for 88 hours with the end defined by the return of all gas normalized ratios 

to unity. The lower energy release of zirconium nitride oxidation (two thirds of complete 

zirconium oxidation) coupled with increased convective energy removal resulted in a less 

energetic burn during the second stage. The oxidation kinetics given in Equation 4.3 have been 

studied by Krusin-Elbaum and Wittmer (1983), where a parabolic dependence of oxide growth 

on time was observed which indicates a diffusion limited reaction.

4.5 Post-Test Examination

Figure 4.13 shows photographs of the test assembly from different perspectives during the 

deconstruction activities. Image (a) shows the top ~0.61 m [24 in] of the assembly with the 

insulation removed from the assembly. The North peripheral cell shows loss of structural 

integrity of the pool cell towards the bottom of the photo. Image (b) is a top-down perspective of 

the assembly after the top ~0.61 m [24 in] of the assembly was removed. The Zircaloy 

components were nearly all transformed to ceramic by the fire and were highly brittle. No 

evidence of the center pool rack walls remained. The East assembly is shown with the outer 

pool cell wall removed in image (c). The stainless steel plenum spring is seen protruding 

through the fuel to the right of the image. The top nozzles of the East and North assemblies are 

shown in images (d and e), respectively. Accumulation of MgO at the top of the assemblies was 

more pronounced in the North and West (not shown) than in the East and South (not shown). 

Finally, a piece of vitrified insulation from a corner cell is pictured in image (f). The hightemperature insulation in the corner cells relocated to the bottom of the assembly and fused into 

large solid blocks.
4-12

Figure 4.13 Photographs of the test assembly during various stages of deconstruction

a)

b)

c) d)

e)

f)

North Bundle

East Bundle East Bundle

Vitrified Insulation

Top-Down of Debris Field

Before Deconstruction
5-1

5 MELCOR MODELING RESULTS

The OECD/NEA SFP Project Phase II experiments present a challenging modeling scenario 

where the central heated assembly can ignite, leading to an intense heat flux to the surrounding 

unheated assemblies, resulting in a large temperature gradient across the peripheral 

assemblies. These experiments are used to validate the spent fuel pool models that are part of 

the MELCOR code capabilities, including an assessment of the air oxidation models. The

MELCOR input models used are described in Section 5.2. A developmental version of 

MELCOR based on version 2.1.5287 was used for these simulations. Code modifications are 

described in Section 5.3. Simulation results for the pre-ignition and ignition tests are described 

in Sections 5.4 and 5.5 respectively. Sensitivity studies are described in Section 5.6.

5.1 Introduction

For this analysis, 2-ring and 9-ring MELCOR models were developed to simulate the PWR 

spent fuel assembly experiments performed at Sandia National Laboratories. A 2-ring model is 

the preferred approach for a reasonable model of an actual spent fuel pool. The 2-ring model

simulation using a relatively coarse nodalization improves code run time and decreases model 

complexities in comparison to the 9-ring model. In the 9-ring model, assemblies are subdivided 

by rows of fuel rods which provide a more detailed modeling of the experiment to be used in

assessing the MELCOR models. The 9-ring model was developed by Energy Research Inc.

and the 2-ring model was developed at SNL based on the 9-ring model. It is important to 

compare the detailed 9-ring calculation with a 2-ring calculation that utilizes identical radiation 

coupling between the spent fuel rack and the fuel rods in both the central and peripheral 

assemblies. The input models used are described in more detail in Section 5.2.

In order to more accurately predict the temperature response with the 2-ring model, code 

improvements were made to more adequately represent the intense thermal gradient between 

the central heated assembly and the unheated (peripheral) assemblies, as well as the 

temperature gradient between rods in the peripheral assemblies. The improvements involve the 

addition of multiple rods within a single core ring. These code changes, described in 

Section 5.3, are still developmental and have not been fully implemented into the code but were 

made as a proof of concept. In particular, the current implementation of these models does not 

account for core degradation properly and consequently, the calculations performed are only 

sufficient in calculating up to the point of ignition in the test. For the purpose of this analysis, 

this is acceptable given the test data is generally no longer available after ignition due to failures 

of the thermocouples at elevated temperatures. Additional work and refinement of these 

models is anticipated prior to these new features being incorporated in a release version of 

MELCOR. Comparisons of results predicted by the 2-ring and 9-ring models are discussed in 

Sections 5.4 and 5.5.

5.2 MELCOR Models

Three MELCOR models were developed to simulate Phase II of the OECD/NEA SFP 

experiments. These included a 2-ring model, a 2-ring multiple rod model and a 9-ring model. 

Table 5.1 presents a summary of the main design parameters for the five assemblies in the 

1 by 4 configuration in Phase II of the tests.
5-2

Table 5.1 Summary of main facility design parameters for the Phase II tests

Parameter 

Group Description Units Value

Overall Five full-length assemblies in 3×3 pool rack with 

heated assembly in the center cell - -

Power and time duration for nine pre-ignition 

tests

kW 

(h)

1 (24), 2 (24), 4 

(12), 4 (24), 6 

(12), 8 (12), 10 

(4), 12 (4) and 

15 (3)

Power for ignition test kW 15

Fuel Assembly “Fuel” (MgO) outer diameter mm 8.19

Heater rod cladding outer diameter mm 9.53

Control rod guide tube outer diameter mm 12.19

Number of heater rods per assembly - 264

Number of control rod guide tubes per assembly - 24

Number of instrument tubes per assembly - 1

Number of spacers in assembly - 11

Heater rod filler material - MgO

Heated rod axial length m 3.613

Fuel rod & tube axial length m 3.918

Total mass of steel cladding per assembly kg 5.748

Total mass of MgO in center assembly kg 142.66

Total mass of Zry in center assembly kg 142.2

Total mass of MgO in four peripheral assemblies kg 541.01

Total mass of cladding in four peripheral 

assemblies kg 440.87

3×3 Pool Rack Storage cell inside width (center cell/peripheral 

cells) mm 224.2/222.7

Flow area (center/peripheral) m2 0.0286/0.0279

Equivalent hydraulic diameter (center/peripheral) mm 11.74/11.46

Storage cell rack thickness mm 1.91

Insulation thickness mm 152.4

Number of pool cells - 5

Total mass of 3×3 pool rack kg 561

Others Major loss coefficient (SLAM) in the central 

assembly for 300 slpm upper flow range - 119*

Total form loss coefficient (Σk) in the central 

assembly for 300 slpm upper flow range - 68*

Major loss coefficient (SLAM) in the peripheral 

assemblies for 300 slpm upper flow range - 113*

Total form loss coefficient (Σk) in the peripheral 

assemblies for 300 slpm upper flow range - 47.5*

* Derived SLAM and Σk values based on the low flow range pressure drop parameters (see Table 

2.4.) Note that Σk = 68 for the center assembly eliminates the form losses incurred from 

passing through the inlet and test stand.
5-3

5.2.1 9-Ring Model

The 9-ring model has spatial representation of the five assemblies in a 1 by 4 configuration and 

is fundamentally limited by the MELCOR core (COR) package's restriction to two-dimensional

geometry (i.e., axial and radial nodalization with no azimuthal component). In the model, 17 

axial COR package levels were present in each of 9 radial rings, as shown in Figures 5.1 and 

5.2. In COR axial levels 3 through 15, four COR cells (rings) were assigned to each axial level 

of the heated center assembly, whereas five COR rings describe the peripheral assemblies as a 

whole. Since in MELCOR version 1.8.6 (which was used early in the project) the total number 

of control volumes in the core and the lower plenum regions is limited to 99, each control 

volume (CV) was assigned to two axial levels in the lower region of the assemblies while each 

CV corresponded one-to-one to each axial level in the upper region of the assemblies. In 

MELCOR 2.1, which was used for this project, the limit of 99 CVs is no longer applied which 

allows development of larger more complex models, however, the model noding was not 

changed to take advantage of this as it was not expected to improve the results. COR axial 

levels 16 and 17 (above the top of the heated portion of the rods) shared a single level of nine 

control volumes, and axial levels 1 and 2 occupied the assembly inlet (a single level of two 

control volumes, one for the center and one for the peripheral assemblies).

Note that the nine COR rings describe cross-sectional regions of the test assemblies in 

increasing distance from the middle of the center assembly; the four peripheral assemblies were 

lumped together, with no distinction possible, for example, between the North and South 

assemblies. The steel-sheathed Kaowool insulation layer surrounding the peripheral 

assemblies was modeled as a series of MELCOR heat structures.

A total of 94 control volumes (CV) were used to represent the five assemblies in the 1 by 4 

configuration (i.e., a vertical stack of 10 CVs in each of the nine COR rings, in addition to two 

inlet, one outlet, and one downcomer volume), which are denoted by the numbers in black color 

at the center of each cell in Figure 5.1. A total of 173 flow paths, denoted by the numbers in 

blue along with arrows, interconnected these control volumes, with vertical flow paths, and cross 

flow paths between adjacent rings of each assembly. There was no cross-flow path through the 

rack walls between the central and the peripheral assemblies, except for a leakage flow path at 

the bottom of the rack, representing an imperfect seal at the interface of the rack and the test 

stand. There were in total 153 COR cells, which are denoted by the numbers in red at the top 

on the right side of each cell (note that in some cases, above the heated length of the rods or 

the lower region of the assemblies, one CV is denoted as containing two COR cells). Air enters 

the assemblies from the inlet pipes at the bottom and flows upwards through the assemblies as 

a result of natural buoyancy forces as the assemblies are heated. The exits of the assemblies 

were connected to a time-independent control volume representing atmospheric conditions 

(CV020). Hydraulic loss factor values derived from experimental measurements (see 

Section 2.4.5) for the Phase II tests were used in this study.

In the COR package nodalization, the flow area and number of heated fuel rods in the center 

assembly were divided into four rings, while five rings model the four peripheral assembly 

unheated rods in increasing distance from the center. The flow area and the number of fuel 

rods varied with each ring, such that the radial nodalization became finer where larger 

temperature gradients were expected. In particular, rings 4 and 5 (i.e., adjacent to the rack wall) 

consisted of a single column of fuel rods. In order to preserve the rectangular geometry of the 

peripheral assemblies, rings 5 to 9 were modeled such that the distance from ring 4 to ring 9 

was the same as the inner width of the peripheral assemblies. In the vertical direction, the core 

region consisted of 12 axial levels in the heated region (i.e., from core axial level 4 to axial level 

15, colored in pink for the center assembly as seen in Figure 5.1), corresponding to about 0.3 m
5-4

[1 ft] length of heater rod per axial level; two axial levels above the top of the heated region to 

model the unheated plenum region and outlet nozzles (COR levels 16 and 17); and three axial 

levels below the heated region to model the inlet nozzles (COR level 3), the base-plate (COR 

level 2), and the pipe under the base-plate (COR level 1), respectively.

Heater rod, pool cell, and other dimensions and masses were modeled based on data obtained 

from the Phase II test plan (Ref. 20). The experimental tests used compacted MgO as rod filler 

that was uniformly heated by a central Nichrome wire to simulate UO2 fuel; therefore, the 

thermo-physical properties of UO2 fuel were replaced in the input model by MgO properties. As 

a simplification in the 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 

mass of MgO); this approximation was found in the case of Reference 8 to result in negligible 

differences in model output. Zircaloy mass in the cladding was specified as cladding 

component mass. The mass from the 24 guide tubes, one instrumentation tube, and 11 

assembly grid spacers per assembly was specified as Zircaloy non-supporting structure (NS) 

component mass, whereas the base-plate was specified as a supporting structure component 

(SS) made of stainless steel. The relatively small masses of the upper plate and nozzle were 

included in the input model, but represented as MELCOR heat structures. These heat 

structures were defined as constituting the upper axial boundary of the “core” as required by the 

MELCOR COR package.

The steel racks surrounding the pool cells were represented by mass allocated to the new rack 

(RK) component for core model of type SFP-PWR. The layer of steel-wrapped Kaowool 

insulation surrounding the peripheral assemblies and at the corners of the 1 by 4 cell 

configuration was represented as a vertical stack of COR radial boundary heat structures, which 

thermally interacted with the peripheral assemblies by means of radiative and conductive heat 

transfer.

In order to represent the actual rectangular geometry of the test assemblies, component surface 

areas and coolant flow areas were calculated from the real geometry (e.g., the boundary 

surface area for each ring of rings 5 to 9 was a product of the width of a test assembly multiplied 

by 4 peripheral assemblies) rather than by first deriving a fictitious and approximate cylindrical 

model geometry.

The default radial model of thermal radiation in MELCOR for PWR spent fuel pools consist of 

intra-cell radiation between cladding and rack components in each cell, in addition to inter-cell 

radiation between each rack component and the cladding in the adjacent ring. Such a model 

was not best suited to the nodalization used in the present study, since MELCOR would (for 

example) calculate no direct thermal communication between heater rod surfaces in different 

rings of one assembly. Therefore, in order to correctly reflect inter-cell radiation exchange, the 

conventional radial inter-cell radiation view factor defined by input parameter FCELR was set to 

zero; view factor-surface area pairs were directly input between specific pairs of components 

(e.g., ring 5 cladding to ring 6 cladding) using the optional CORHTR cards. Calculation of these 

inter-cell component-to-component view factors in the radial direction were performed with 

ANSYS FLUENT (Refs. 17 and 18) and are based upon the component geometry.

ANSYS FLUENT can compute the view factors for radiation problems and save them to a file 

when the surface-to-surface (S2S) radiation model is invoked. The surface-to-surface radiation 

model can be used to account for the radiation exchange in an enclosure of gray-diffuse 

surfaces. The energy exchange between two surfaces depends in part on their size, separation 

distance, and orientation. These parameters are accounted for by a geometric function called a 
5-5

“view factor”. When using the S2S radiation model, ANSYS FLUENT provides two methods for 

computing view factors: the ray tracing method (which is selected by default) and the hemicube 

method. The hemicube method is available only for 3D and axisymmetric cases and ray tracing 

method is only available when Face to Face is selected for basis. The hemicube method uses a 

differential area-to-area method and calculates the view factors on a row-by-row basis. The 

view factors calculated from the differential areas are summed to provide the view factor for the 

whole surface. While the hemicube method projects radiating surfaces onto a hemicube, the 

ray tracing method instead traces rays through the centers of every hemicube face to determine 

which surfaces are visible through that face. The ray tracing method does not subdivide the 

faces as can be done when using the hemicube method. In the current calculation, the ray 

tracing method was used to obtain the view factors between surfaces. The approximation was 

made that the radiative heat transfer cannot travel further than the next adjacent component. 

For example, the fuel rods in ring 1 can only radiate to the fuel rods in ring 2. In reality, some 

small fraction of the radiation (~0.3 percent) would go directly from ring 1 to ring 3, ring 4 and 

the rack.

The true projection of each visible face onto the hemicube can be accurately accounted for by 

using a finite-resolution hemicube. As described previously, the faces are projected onto a 

hemicube. Because of the finite resolution of the hemicube, the projected areas and resulting 

view factors may be overestimated or underestimated. This type of error is referred to as 

aliasing effects and can be reduced by increasing the value of the resolution. The value of the 

resolution can be adjusted by the user in order to reduce the impact of aliasing effects, as 

described previously. Different resolution values were used to eliminate any errors from aliasing 

effects. In addition to the default resolution value of 10, values of 100 and 1,000 were used. 

Minimal differences were noticed between the view factor values for these resolutions. As such, 

the final view factor values were free from aliasing effects error.

For the 9-ring model, the COR_HTR input table is used to specify radiative heat transfer paths 

between rods and from the racks to the rods. Note that the user manual defines the radiation 

constant supplied by the user, VFA, as the product of the view factor and the surface area used 

in calculating the radiant heat flow. However, an optional field following this factor can be 

supplied by the user to indicate that VFA is the view factor and the surface area is taken from 

the intact area of the ‘from’ radiating component. Therefore, VFA represents the view factor for 

the ‘from’ component. Furthermore, the outer surface area is determined from ASCELR. When 

radiation occurs to or from the RK component, the inner surface area is calculated as the total 

intact rack surface area (inside and outside) and subtracting the outer surface area (ASCELR). 

A summary of all radiation view factors defined in the 9-ring MELCOR models is provided in 

Table 5.2.

The outer edge of the peripheral assemblies in the Phase II tests (i.e., the regions farthest from 

the center assembly) were expected to experience relatively low temperatures. Therefore, 

unlike conditions pertaining in an actual reactor core, thermal conduction at the radial boundary 

may be significant in comparison to radiative heat losses. This factor was accounted for by 

using the optional COR_BCP card to add a thermal conduction path from the cladding of ring 9 

to the Kaowool insulation heat structure. Note that MELCOR modeling of the rack component 

does not permit such a conductive path to be specified directly from the rack to the insulation, 

necessitating the use of cladding temperature as a surrogate (since the gradient at that location 

is expected to be small).
5-6

Although ballooning occurred during the Phase II experiment, MELCOR does not include the 

option to model ballooning of fuel cladding. Consequently, the current input model does not 

reflect the fact that two peripheral assemblies were pressurized with noble gas (argon) to 

simulate ballooning of fuel cladding.

In addition, because the fuel region is fundamentally restricted to a two-dimensional 

representation in MELCOR (i.e., variation of properties allowed in only the vertical and radial 

directions), influence of the heat loss through the rack walls of the lateral sides of the peripheral 

assemblies on the temperature (an azimuthal variation) cannot be properly modeled. 

Nevertheless, the heat losses through the side walls of the peripheral assemblies were 

expected to be small due to the presence of thick insulation at the corners of the 1 by 4 cell 

configuration. The four corners of the 3 by 3 array were packed with Kaowool insulation.

Blockage of flow paths by particulate debris was not explicitly modeled in this input, since the 

thermocouples were expected to be damaged prior to debris relocation, and therefore code 

predictions past that time were not useful for validation.

The derived values of SLAM and Σk were based on the flow measurements for Phase II 

apparatus, which were used in the present calculations. The selected values of SLAM and Σk in 

the center assembly are 119 and 68, respectively, and the values of SLAM and Σk in the 

peripheral assemblies are 113 and 47.5, respectively. In order to correctly predict the heat 

transfer in the laminar free convection regime, the MELCOR correlations for forced convection 

and turbulent free convection were de-activated. Only the MELCOR correlation for laminar free 

convection was used for these calculations since the flows have very low Reynolds numbers.

Table 5.2 Radiation view factors used in the MELCOR model

From To Radiation View Factor

Rods in ring 1 Rods in ring 2 0.322

Rods in ring 2 Rods in ring 3 0.215

Rods in ring 3 Rods in ring 4 0.264

Rods in ring 4 Center Assembly Rack 0.392

Center Assembly Rack Rods in ring 5 1.000

Rods in ring 5 Rods in ring 6 0.398

Rods in ring 6 Rods in ring 7 0.269

Rods in ring 7 Rods in ring 8 0.211

Rods in ring 8 Rods in ring 9 0.138

Rods in ring 9 Peripheral Assembly Rack 0.072
5-7

20 (Exit Section)

10 (Inlet Section) 11 (Inlet Section)

005

30

003

002

511

510

509

513

512

511

510

509

508

507

506

505

504

561

560

559

611

610

609

613

612

611

610

609

608

607

606

605

604

661

660

659

711

710

709

511

510

509

507

505

503

713

712

711

710

709

708

707

706

705

704

711

710

709

707

705

703

761

760

759

861

860

859

811

810

809

911

910

909

811

810

809

807

805

803

911

910

909

907

905

903

813

812

811

810

809

808

807

806

805

804

913

912

911

910

909

908

907

906

905

904

102 202 302 402

101 201 301 401

502 602 702 802

501 601 701 801

902

901

611

610

609

607

605

603

111

110

109

113

112

111

110

109

108

107

106

105

104

111

110

109

107

105

103

161

160

159

211

210

209

213

212

211

210

209

208

207

206

205

204

211

210

209

207

205

203

261

260

259

311

310

309

313

312

311

310

309

308

307

306

305

304

311

310

309

307

305

303

361

360

359

411

410

409

413

412

411

410

409

408

407

406

405

404

411

410

409

407

405

403

112 212 312 412

112 212 312 412

113 213 313 413

113 213 313 413

101 201 301 401

101 201 301 401

114 214 314 414

114 214 314 414

162

163

164 264 364

114

115

214

116

117

215

314

216

217

315

316

317

414

415

416

417

263 363

262 362

501 601 701 801 901

503 603 703 803 903

514 614 714 814

514 614 714 814 914

914

516 564 664 764 864

517

616

617 716

717 816817 916

917

513 613 713 813

513 613 713 813 913

913

515 615 715 815 915

563 663 763 863

512 612 712 812

512 612 712 812 912

912

514 614 714 814 914

562 662 762 862

115 215 315 415 515 615 715 815 915

103 203 303 403

151 251 351

103 153

203 253303 353

403

105 155

205 255

305 355405

107 157

207 257

307 357407 507 557607 657707 757 857 807 907

505 555605 655705 755 855 805 905

503 553603 653703 753 853 803 903

501 601 701 801 901 551 651 751 851

899

Heated 

region

Un-heated 

region

103

103

105

105

106

CV number FL number

COR cell 

number

Figure 5.1 CVH and COR nodalization of PWR test assemblies in 1 by 4 configurations 

(side view).
5-8

Center Assembly Peripheral Assemblies

R3 R5 R7 R8 R9

Insulation layer 

modeled as heat 

structure

Rack-insulation 

conduction with air 

gap

Radiation view 

factors clad-clad, 

rack-rack

R1 R2 R4 R6

Figure 5.2 CVH and COR Nodalization of PWR test assemblies in 1 by 4 configuration 

(top view).

5.2.2 2-Ring Model

SNL converted the 9-ring model described above into a 2-ring model. Radially, there are two 

COR rings with the inner ring corresponding to the heated assemblies and the four peripheral 

assemblies modeled with a single outer ring (see Figure 5.3). The outer unheated assemblies 

are modeled in one ring having cross-sectional area equal to the sum of the cross-sectional 

area of all four unheated assemblies. Similarly, flow area, fuel mass, clad mass, control rod 

mass, and rack masses are conserved. The four inner rack walls that separate the heated 

assembly from the unheated assemblies is modeled as a rack component located in the inner 

ring and the remaining rack walls are located in the outer ring. The 2-ring model uses a total of 

33 CVH control volumes and 34 flow paths as seen in Figure 5.4.

Two 2-ring input models were developed and assessed against the experimental data. The first 

model has only a single fuel rod type for each ring, resulting in a single temperature in that ring 

(at each axial node). This is the only modeling capability in the current released version of 

MELCOR. Consequently, MELCOR can calculate only an average fuel rod temperature for the 

heated assemblies and a single average for the four unheated assemblies. The second model, 

referred to as the 2-ring multiple rod model, utilizes new modeling features that were added to 

correct for current code limitations. This model allows multiple fuel rod types within each ring to 

better account for the temperature gradient that was observed in the experiment.

Multiple fuel rod components are modeled for both the inner assembly (4 groups) and the outer 

assemblies (5 groups) as seen in Figure 5.5. When the multiple rod/ring model is invoked, the 

nodalization is similar to the 9-ring model described above where now rod groups in the 2-ring 

multiple rod model correspond to rings in the 9-ring model.
5-9

While the intent is to have the 2-ring multiple rod model as similar as possible to the 9-ring 

model, it is recognized that there are still some modeling differences. One important difference 

between the two models is that each COR cell in the 9-ring model has a separate CVH control 

volume, while in the 2-ring multiple rod model, all fuel rod types for a given COR cell share the 

same CVH control volume. As an example, consider the four fuel rod types in the new 2-ring 

multiple rod model. While each fuel rod can have a unique temperature (at a given COR cell), 

all of these rods will see the same air temperature and velocity. In the 9-ring model, each of the 

associated four fuel rods can see a different air temperature and velocity. In the peripheral 

region, this can lead to natural circulation flow patterns that cannot be obtained in a 2-ring 

model where there is only a single flow path through the assemblies. As an example, one such 

flow pattern that has been observed in the MELCOR runs is in the peripheral region where air 

flow is upwards in ring 5 (closest to the center heated assembly) and downward in ring 9 

(farthest from the center heated assembly). These natural circulation flow patterns affect the 

heat transfer early in the transient when convective heat transfer is still the dominate heat 

transfer mode and is responsible for some of the differences seen between the 2-ring multiple 

rod and 9-ring models.

The 9-ring and 2-ring multiple rod model use different approaches for specifying and calculating 

the radiant heat exchange between components. For the 9-ring model, view factors are 

specified on the COR_HTR records as user defined heat transfer paths and radiant heat 

transfer is calculated in subroutine COR_HTR. For the 2-ring multiple rod model, view factors 

are specified on the COR_ROD_VF records and radiant heat transfer is calculated from the 

COR_CORRAD subroutine. The same view factors that were applied in the 9-ring model were 

used in the 2-ring multiple rod model.

Experiment Geometry MELCOR Geometry

Figure 5.3 MELCOR radial layout of 2-ring model
5-10

Figure 5.4 MELCOR 2-ring model nodalization

Assembly geometry

Identification of fuel rod 

components

MELCOR representation

Figure 5.5 MELCOR 2-ring 9-rod Model of the PWR SFP Experiment

Ring 1 –

Heated 

Assembly

Ring 2 –

Un-heated 

assemblies
5-11

5.2.3 Oxidation Kinetics

The oxidation kinetics of zirconium in air is governed by the formation of high density zirconium 

nitride, resulting in the development of cracks that expose additional surface area for reaction 

with oxygen (i.e., breakaway). In addition, separate effect tests have shown the potential for 

strong breakup of the oxide crust under conditions of oxidation in pure air, resulting in enhanced 

oxidation rates, and therefore, to temperature excursions at the elevation of the breakup and 

following starvation of the oxidant. In the experiment, the crust breakup at different axial and 

radial nodes of the assembly was the starting point for the fuel heatup leading to ignition of the 

zirconium fire. Generally, ignition started in the upper region of the fuel assembly and 

subsequently, oxidation propagated downward into the assembly.

In MELCOR, metal oxidation is calculated using standard parabolic kinetics, with appropriate 

rate constant expressions for Zircaloy and steel, limited by gaseous diffusion considerations if 

necessary. The default treatment of breakaway is based on a simplified air oxidation model 

which is coupled with a breakaway model that has been based on an empirical fit to Argonne 

National Laboratory (ANL) experimental data (Ref. 13) as seen in Figure 5.6. The calculation of 

the breakaway oxidation reaction using a cumulative lifetime fraction (LF) function is defined in 

Equation 5.1.

( ) ( ) ∫ =

t

o T

dt LF t

τ

5.1

The MELCOR breakaway oxidation model calculates the lifetime function at every node in the 

model with Zircaloy cladding. At a critical value of the lifetime fraction (LF=1.0), the transition to 

breakaway is assumed to begin. The transition to post-breakaway oxidation kinetics is 

completed when LF=1.2. The instantaneous breakaway lifetime τ(T) (in seconds) as a function 

of temperature T (in K) is given in Equation 5.2.

( ) 12.528 log 42.038 10 10− + = T

τ T 5.2

The coefficients in Equation 5.2 were obtained by log-linear regression on experimental data 

obtained in tests performed at ANL and are necessarily subject to some uncertainty due to 

measurement error and to scatter in the data. As seen in Figure 5.6, there is a rather large 

uncertainty in the breakaway time. Variations within the ANL experimental data were assumed

to be ±0.626, i.e. the 95 percent confidence level as determined by the product of a two-sided 

t-student distribution (2.365) and the standard error of the curve fit (0.265). As a result, a lower 

and upper bound were obtained on the second term of the exponent in Equation 5.2 between 

41.412 and 42.664, respectively.
5-12

Figure 5.6 Comparison of the MELCOR breakaway timing fit to Zr-4 data.

5.3 Code Improvements

MELCOR Version 2.1.5287 was used as the starting point for the code modifications. This 

version includes the rack component in the COR package for modeling the spent fuel pool rack 

as well as the Sandia breakaway oxidation kinetics model allowing the code to predict the 

transition to breakaway oxidation in an air environment on a node-by-node basis. Furthermore, 

while it was not developed as part of this project, this code version contains the Paul Scherrer 

Institute (PSI) air oxidation models for breakaway (Ref. 15). It should be noted that the 

MELCOR COR package is the set of routines that models the components found in a reactor 

core. Because of similarities and flexibility in the code modeling, MELCOR can also be used for 

modeling the spent fuel pool assemblies and the user input distinguishes this configuration from 

an reactor core.

The conditions of this experiment challenge the MELCOR approach to representing an 

assembly (or collection of assemblies) with a single MELCOR node (or ring) consisting of a 

single lumped parameter representing all fuel rods in that node. In particular, a single 

representative temperature for the fuel rods in the unheated assembly is insufficient to capture 

the large temperature gradients and the dependence of oxidation rates on the local clad 

temperature. This is more readily understood when examining the temperatures measured in 

the unheated peripheral assemblies as shown in Figure 5.7 and the temperature contours as 

seen in Figure 4.8. Prior to ignition, the temperature of fuel rods near the interface with the 

heated assembly is almost 400 K [720°F] higher than it is for rods at the outer boundary of the 

assembly. The peak temperatures adjacent to the heated assembly are about 700 K [800°F] 

while the average temperature (which would be modeled by MELCOR using a 2-ring model) is 

only 450 K [350°F]. This large temperature gradient is particularly significant given that 

oxidation is highly temperature dependent and the oxidation reaction generates heat that then 

contributes further to the burn. For example, the oxidation rate at 700 K [800°F] is several 

orders of magnitude greater than it is at the average temperature of 550 K [530°F]. If MELCOR 
5-13

were modeling the assembly with a single node (as is done in the 2-ring model), the average 

cell temperature would be used in calculating the oxidation rate in that ring. This would 

significantly underestimate the oxidation that occurs in the unheated assembly and would 

therefore also underestimate the associated heat generation.

Consequently, the MELCOR code was modified to allow improved temperature refinement 

within a single radial ring by providing a means for the user to specify multiple fuel rod types 

within that ring. As modified, the user can now specify up to five fuel rod types in each radial 

ring. For example, in the peripheral region, a single fuel rod type can be used to represent the 

row of rods closest to the rack that divides the center assembly from the peripheral assemblies. 

Additional rod types can be defined with coarser nodalization for rods as they move closer to the 

outer rack.

The input for these new fuel rod types (currently limited to five) is rather simple. The user 

specifies the total fuel rod masses, control rod masses, and areas for a ring, just as before. 

However, new inputs are used to specify mass fractions for each fuel type, mass fractions for 

each control rod type, and view factors connecting fuel rods.

The models are currently strictly valid up until material begins to relocate which is sufficient for 

the present experimental validation of the code. This is acceptable given the test data is 

generally unavailable after ignition due to failures of the thermocouples at elevated 

temperatures. Further model development and refinement are necessary to allow relocation of 

different fuel rod types for spent fuel pool accident progression analysis and before becoming 

part of an official release version of MELCOR.

Figure 5.7 Cladding temperatures in peripheral assemblies for 15 kW ignition test at 

6 hours
5-14

5.4 Pre-Ignition Tests

As described previously, there were nine pre-ignition tests performed at power levels ranging 

from 1 kW to 15 kW. MELCOR runs were made for all power levels using each of the three 

MELCOR models described in Section 5.2. The three MELCOR models include: 1) 2-ring, 2) 

2-ring multiple rod/ring and 3) 9-ring. Modeling assumptions were as similar as possible for 

each calculation, though it is recognized that there may still be small modeling differences. 

Given the number of MELCOR simulations and the fact that most show the same phenomena, 

not all runs will be described and only selected results will be presented to demonstrate different 

topics of interest.

5.4.1 1 kW pre-ignition

The 1 kW pre-ignition test was run for 24 hours in an attempt to achieve steady state conditions. 

Figures 5.8 and 5.9 show the peak cladding temperature in the central and peripheral 

assemblies respectively for both 2-ring models and the 9-ring model. While close to steady 

state, the temperature is still increasing slightly at 24 hours in both the experiment and 

MELCOR simulations. For the experimental data, the PCT is computed as the maximum 

temperature observed at all locations in the assembly and may move from fuel rod to fuel rod as 

the transient progresses. For MELCOR, the PCT is a misnomer and is actually an average for a 

group of rods. For example, in the 9-ring model ring 1 contains 24 fuel rods (see Figure 5.2) 

that all must have the same temperature given they are all lumped together into a single rod 

component. By definition, the average temperature in a group of rods will be lower than the 

peak, so it is expected that the MELCOR PCT will generally be lower than the experimental data 

PCT.

In the center heated assembly, the MELCOR results show good agreement with the peak 

cladding temperature. In the peripheral assemblies, the 2-ring model shows poor agreement 

with the experimental data, while the other models show reasonable agreement. In all the 

MELCOR results, the heat from the center assembly rods is transferred to the rack by 

convection and radiation, then to the peripheral assembly rods by radiation only. The lack of 

convective heat transfer from the center assembly rack to the peripheral rods is a known 

limitation of the MELCOR code. Given the initial small temperature difference between the 

center assembly rods, rack and peripheral rods, the initial radiation heat transfer is close to 

zero. With little to no radiation and no convective heat transfer, the MELCOR prediction of PCT 

in the peripheral rods initially lags that of the test data. Once the center assembly rods and rack 

heat up, radiation heat transfer begins to dominate and the MELCOR calculations (2-ring 

multiple rod and 9-ring) show the PCT surpass that of the test data and generally show good 

agreement. However, the 2-ring model shows the PCT lagging significantly behind the test 

data. The peripheral temperature response in the 2-ring model lags that of the other two 

MELCOR models due to the fact that all of the peripheral rods must be heated up as opposed to 

only the first row of rods in the 2-ring multiple rod and 9-ring models.

Figures 5.10 and 5.11 show the central and peripheral temperature as a function of elevation for 

the 2-ring and 2-ring multiple rod models at 24 hours. While the MELCOR models show good 

agreement for the peak temperatures, both the 2-ring and 2-ring multiple rod model in general 

over predict the temperatures in the central heated assembly. The 2-ring multiple rod model 

shows a temperature difference of ~15 K [27°F] between the hottest and coldest rod in the 

center assembly, while the test data shows a difference of ~60 K [108°F]. In general, all the 

MELCOR results show that the temperatures in the center assembly are more tightly coupled 

together than observed in the test data. Examination of the test data show that the coldest rods 

in the center assembly are located in and around the corners. Given that the two-dimensional 
5-15

MELCOR model does not allow for azimuthal variations, there are no rods in the MELCOR 

model which are in “corner” locations. Figures 5.12 and 5.13 show the central and peripheral 

temperatures as a function of elevation for the 9-ring model. These results are similar to the 

2-ring multiple rod model.

The induced volumetric flow rate is shown in Figures 5.14 and 5.15 for the center and peripheral 

assemblies respectively. All three of the MELCOR calculations show the flow rate is over 

predicted in the center assembly. In the peripheral assemblies, the MELCOR calculations show 

the flow rate is under predicted. Note that while not shown in the Figures, the uncertainity in the 

experimental data is ±12 slpm for each assembly. Also note that the hot wire anemometers 

were generally unresponsive below 50 slpm/assembly. The differences in flow rate are mainly 

due to the use of constant loss coefficients SLAM and Σk. Sensitivities on the use of constant 

versus variable loss coefficients are described in more detail in Section 5.6.1.

5.4.2 8 kW pre-ignition test

The 8 kW test was run for 12 hours and stopped prior to the center assembly PCT reaching 

900 K [1,160°F]. Figures 5.16 and 5.17 show the peak cladding temperature in the central and 

peripheral assemblies respectively for both 2-ring models and the 9-ring model. In the center 

heated assembly, the MELCOR results for the 2-ring multiple rod and 9-ring model show 

excellent agreement with the peak cladding temperature. As expected, the 2-ring model 

somewhat underpredicts the peak cladding temperature. In the peripheral assemblies all 

MELCOR runs show a delayed increase in temperature similar to that in the 1 kW test.

Figures 5.18 and 5.19 show the central and peripheral temperature as a function of elevation for 

the 2-ring and 2-ring multiple rod models at 12 hours. In general, the 2-ring multiple rod model 

over predicts the temperatures in the center assembly similar to the 1 kW case. In the 

peripheral region, the 2-ring model temperature is in good agreement with the experimental 

average temperature. The 2-ring multiple rod model shows a good temperature distribution 

between the inner and outer most rods and somewhat overpredicts the temperatures relative to 

the data. Figures 5.20 and 5.21 show the central and peripheral temperature as a function of 

elevation for the 9-ring model. These results are similar to the 2-ring multiple rod model as 

expected.

The induced volumetric flow rate is shown in Figures 5.22 and 5.23 for the center and peripheral 

assemblies respectively. All of the MELCOR calculations show the flow rate is over predicted in 

the center assembly. However, the MELCOR flow rates in the center assembly for the 8 kW 

test show better agreement than the results from the 1 kW test. In general, the MELCOR center 

assembly calculated flow rates better match the experimental data as the power level increases. 

In the peripheral assemblies, the MELCOR calculations show the flow rate is initially under 

predicted, then over predicted. Note that the experimental data shows an initial flow rate just 

under 200 slpm which is believed to be in error as the flow was zero at the start of the 

experiment. The hot wire anemometers were generally unresponsive/errant below 

50 slpm/assembly (200 slpm total for the four peripheral assemblies). In the peripheral 

assemblies, above 200 slpm (total for all four assemblies), the experimental data for this test 

shows the same overall trends and falls in line with the other pre-ignition tests measured flow 

rate, giving some confidence that the data is within the ±12 slpm/assembly instrument 

uncertainty. Similar to the 1 kW test described above, the differences in flow are mainly due to 

the use of constant loss coefficients SLAM and Σk. Sensitivities on the use of constant versus 

variable loss coefficients are described in more detail in Section 5.6.1.
5-16

Figure 5.8 Center assembly peak cladding temperature for the 1 kW pre-ignition test

Figure 5.9 Peripheral assembly peak cladding temperature for the 1 kW pre-ignition 

test
5-17

Figure 5.10 Central assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours

Figure 5.11 Peripheral assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours
5-18

Figure 5.12 Central assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours

Figure 5.13 Peripheral assembly temperature versus elevation for the 1 kW pre-ignition 

test at 24 hours
5-19

Figure 5.14 Central assembly volumetric flow rate for the 1 kW pre-ignition test

Figure 5.15 Total peripheral assembly volumetric flow rate for the 1 kW pre-ignition test
5-20

Figure 5.16 Center assembly peak cladding temperature for the 8 kW pre-ignition test

Figure 5.17 Peripheral assembly peak cladding temperature for the 8 kW pre-ignition 

test
5-21

Figure 5.18 Central assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours

Figure 5.19 Peripheral assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours
5-22

Figure 5.20 Central assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours

Figure 5.21 Peripheral assembly temperature versus elevation for the 8 kW pre-ignition 

test at 12 hours
5-23

Figure 5.22 Central assembly volumetric flow rate for the 8 kW pre-ignition test

Figure 5.23 Total peripheral assembly volumetric flow rate for the 8 kW pre-ignition test
5-24

5.5 Ignition Test

The ignition test was conducted with a total applied power of 15 kW to the center fuel assembly. 

The first several hours of the ignition test, where the PCT is below 900 K [1,160°F], show 

results/trends similar to that of the pre-ignition tests described above. MELCOR calculations 

were performed for each of the three models described above which include: 1) 2-ring, 2) 2-ring 

multiple rod/ring and 3) 9-ring. Comparisons were made between the models and experimental 

data (when available) in Figures 5.24 through 5.38. Since the focus is on the new 2-ring 

multiple rod model, plots for the 2-ring and 9-ring models will only be shown in a few selected 

figures. In addition to the three models described above, a second calculation was performed 

with the 2-ring multiple rod model where the best fit curve to the ANL data was replaced with the 

lower uncertainty curve (as seen in Figure 5.6) to take into consideration the known uncertainty 

in this parameter. Further sensitivities on oxidation kinetics are described in Section 5.6.2.

Figures 5.24 and 5.25 show the peak cladding temperature response in the center and 

peripheral assemblies respectively. The 2-ring model, which uses a single rod/radial ring, 

shows the lowest temperatures and shows ignition is significantly delayed over the data and 

other models as expected. The 2-ring multiple rod and 9-ring models show good agreement 

with PCT until 1,025 K [1,385°F], at which point the experiment continues to heat up while the 

MELCOR results reach a quasi-steady condition. This quasi-steady condition remains until the

time-at-temperature condition is met and the model switches from pre- to post-breakaway 

oxidation kinetics. Once this occurs, the additional heat supplied from the oxidation reaction 

increases the PCT leading to fuel rod ignition. It is understood that ignition does not occur at a 

specific temperature, however, for convenience and in order to make comparisons, it is defined 

in this report as when the temperature exceeds 1,200 K [1,700°F]. In the peripheral region, 

ignition for the 2-ring multiple rod and 9-ring models is delayed similar to that seen in the center 

assembly, however, the 2-ring model did not show ignition during the 12 hour time period shown 

in Figure 5.25 but does ignite prior to 14 hours.

While the intent is to have the 2-ring multiple rod model as similar as possible to the 9-ring 

model, it is recognized that there are still some modeling differences as described above in 

Section 5.2.2. These differences result in the 9-ring model retaining more heat in the center 

assembly thus an earlier ignition time. Similar to that seen in the pre-ignition tests, the 9-ring 

model temperature in the peripheral region lags that of the 2-ring multiple rod model.

Figures 5.26 and 5.27 show the PCT in the center and peripheral assemblies for the 2-ring 

multiple rod model with the uncertainty in the ANL data taken into consideration. As seen in 

these figures, when the lower uncertainty is used the MELCOR computed ignition occurs prior 

to the experiment. This implies the measured time to ignition is predicted by the ANL 

breakaway data within uncertainty. A summary of the ignition times and first location of igntion

are shown in Table 5.3.

Ignition propagation for the center heated assembly and unheated peripheral assemblies is 

shown in Figures 5.28 and 5.29 respectively. In the center assembly the experimental data fall 

within the 2-ring multiple rod model with the ANL best fit and lower uncertainty curves. While 

ignition occurs earlier when using the lower uncertainty curve, the ignition propagates downward 

at a rate similar to the test data. In addition, the 2-ring multiple rod model shows good 

agreement for the elevation of the initial ignition in the center assembly. For about the first half 

hour after ignition, the burn front proceeded downward somewhat faster than in the test and 

results in the first occurrence of ignition in the peripheral assemblies at a much lower elevation 

than in the test data. There is more uncertainty in the peripheral region as ignition times vary 

from assembly to assembly (i.e., north, south, east and west) as seen in Figure 5.29.
5-25

Figures 5.30 and 5.31 show the temperature versus elevation at 5 hours. Note that at this time, 

cases with the ANL data best fit and lower uncertainty correlations are the same, so only the 

calculation using the best fit is shown. The MELCOR calculation shows the peak temperatures 

are in excellent agreement with the test data, however, the average MELCOR temperature is 

larger than the test data as the experiment still shows a wide range of temperatures at any 

given elevation. In the peripheral region, MELCOR is in excellent agreement with the test data, 

although the temperature of the rods closest to the center assembly rack are over predicted. 

Figures 5.32 and 5.33 show the temperature versus elevation at 6 hours, which is just prior to 

ignition in the experiment. Overall, the MELCOR calculation using the ANL data best fit shows 

good agreement to the peak temperature in the center assembly. However, as seen in 

Figure 5.32, at an elevation of 3.3 m [130 in], at least one rod has begun to increase in 

temperature and is near 1,100 K [1,520°F]. In the experiment, a single rod can heat up to the 

point of ignition and generate additional heat from the oxidation kinetics reaction, while in the 

MELCOR model, a minimum of 24 rods (smallest group of rods in this model) must undergo the 

same process. This is a known limitation of system level codes which use the lumped 

parameter approach.

Figures 5.34 and 5.35 show the radial temperature profile at 5 hours and two different 

elevations (1.93 m [76 in] and 3.45 m [136 in]). The experimental data presented here is that 

which is available in a straight line outward from the centerline of the center assembly towards 

each of the four peripheral assemblies (North, South, East and West). As such, it does not 

show temperatures that are from the corner rods in the assemblies. Prior to ignition, the 

MELCOR radial temperature profile is in excellent agreement with the test data. As previously 

described, the MELCOR temperatures in the center assembly are closely coupled together and 

show a relatively flat profile, while the data shows a somewhat larger decrease from assembly 

centerline outward. In the peripheral region, the MELCOR temperature closest to the center

assembly rack is higher than the test data, otherwise is in excellent agreement. Overall, 

MELCOR is in good agreement predicting the radial temperature profile through the assemblies.

The volumetric flow rates for the center and peripheral assemblies are shown in Figures 5.36

and 5.37. In the central assembly, the flow is initially over predicted, however, by the time of 

ignition, the MELCOR model shows excellent agreement with the test data. In the peripheral 

region, the models initially under predict the flow, then over predict the flow after ~3 hours. As 

noted above for the pre-ignition tests, the differences in flow are mainly due to the use of 

constant loss coefficients SLAM and ∑k. Sensitivities on the use of constant versus variable loss 

coefficients are described in more detail in Section 5.6.1.

The effect of nodalization (specifically the total number of cells used) on calculation 

performance can be significant. This calculation was run with the three models where the 

maximum time step was reduced to 0.2 seconds. This time step size was chosen because it 

meant that all calculations followed the same time history to eliminate the problem of excessive 

time step reduction in any particular model. Therefore meaningful comparisons can be made in 

the plots of total CPU time as seen in Figure 5.38. Initially, all calculations follow a linear trend, 

indicating that time-step reduction is not occurring in any of the models. Eventually, some of the 

models (particularly the 9-ring model) begin to show some time-step reduction and therefore the 

slope in the line increases. Note that the additional fuel rod types that are added to the 2-ring 

model has only a minor impact on the calculation time. This model offers an acceptable tradeoff to the modeler in terms of improving the fidelity of the calculation to predict the burn 

propagation from the heated assembly into the unheated assemblies. However, when 

increasing the number of COR rings to nine, the calculation time is significantly larger. This 

would be a less acceptable trade-off for the modeler. It is also expected that with more control 
5-26

volumes and flow paths, the CV package could have more numerical problems, leading to 

additional time-step reductions.

Another reason for using the 2-ring multiple rod model is the simplicity of the input model. The 

input model is nearly identical to the 2-ring model with a few new input records for defining mass 

fractions for fuel rod types, mass fractions for control rod types, and view factors. The inputs for 

the 9-ring model (COR_KFU, COR_KCL, COR_ KNS, COR_SS, COR_CIT, COR_EDR, 

CR_RDR, COR_BFA, COR_SA, and COR_RSA) are nine times larger than for the 2-ring 

model. This makes maintenance of an input model much more complex. For the 2-ring model, 

it is possible to renodalize within a ring by changing at most three input tables. For the 9-ring 

model, each of these input tables would need to be modified.

Table 5.3 Ignition times and locations for the 15 kW ignition test

Description Time of ignition* Location of Ignition**

Center Peripheral Center Peripheral

Experiment 6.31 hours 7.08 hours 3.302 m 

[130.0 in]

3.150 m 

[124.0 in]

2-ring, ANL data best fit correlation 11.6 hours 13.8 hours 3.182 m 

[125.3 in]

2.279 m 

[89.7 in]

2-ring multiple rod, ANL data best fit 

correlation 10.1 hours 10.9 hours 3.182 m [125.3 in]

2.580 m 

[101.6 in]

2-ring multiple rod, ANL data lower 

uncertainty correlation 5.62 hours 6.67 hours 3.182 m [125.3 in]

1.978 m 

[77.9 in]

9-ring, ANL data best fit correlation 8.90 hours 10.7 hours 3.483 m 

[137.1 in]

2.279 m 

[89.7 in]

* Ignition defined when temperature exceeds 1,200 K [1,700°F].

** Location of MELCOR is at node center where ignition first occurred.
5-27

Figure 5.24 Peak cladding temperature in the center assembly for the 15 kW ignition 

test

Figure 5.25 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test
5-28

Figure 5.26 Peak cladding temperature in the center assembly for the 15 kW ignition 

test

Figure 5.27 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test
5-29

Figure 5.28 Central assembly ignition propagation for the 15 kW ignition test

Figure 5.29 Peripheral assembly ignition propagation for the 15 kW ignition test
5-30

Figure 5.30 Central assembly temperature versus elevation for the 15 kW ignition test 

at 5 hours

Figure 5.31 Peripheral assembly temperature versus elevation for the 15 kW ignition 

test at 5 hours
5-31

Figure 5.32 Central assembly temperature versus elevation for the 15 kW ignition test 

at 6 hours

Figure 5.33 Peripheral assembly temperature versus elevation for the 15 kW ignition 

test at 6 hours
5-32

Figure 5.34 Radial temperature distribution at 1.93 m [76 in] for the 15 kW ignition test 

at 5 hours

Figure 5.35 Radial temperature distribution at 3.45 m [136 in] for the 15 kW ignition test 

at 5 hours
5-33

Figure 5.36 Central assembly volumetric flow rate for the 15 kW ignition test

Figure 5.37 Peripheral assembly volumetric flow rate for the 15 kW ignition test
5-34

Figure 5.38 CPU times for the MELCOR simulations of the 15 kW ignition test

5.6 MELCOR Sensitivities

Two sensitivity studies were performed for simulation of the Phase II tests. These sensitivity 

cases were intended to investigate the effects on key figures-of-merit of the following 

phenomenological or modeling issues: 1) hydraulic flow loss coefficients SLAM and Σk and 2) 

transition from pre- to post-breakaway oxidation kinetics.

5.6.1 Hydraulic Loss Coefficients

The hydraulic loss coefficients, namely the frictional loss coefficient (SLAM) and the inertial loss 

coefficient (Σk) values were determined experimentally as described in NUREG/CR-7144 

(Ref. 14). To determine these values, flow was forced through a fuel assembly at a given rate, 

and the pressure drop through the assembly was measured. From this data, equations for the 

loss coefficients were determined (Equations 5.3 and 5.4) where a1 and a2 are determined from 

curve fits to the ΔP data versus velocity data, Dh is the hydraulic diameter, L is the length over 

which the ΔP is measured, μ is the viscosity and ρ is the density.

















⋅ = ⋅ ⋅ L μ

D S a h

LAM

2

1 2 5.3

ρ

2 2 a k ⋅ Σ = 5.4
5-35

For the base runs, these loss coefficients are then input into the MELCOR model as constants

over the entire length of the fuel assembly. However, as parameter SLAM is a function of 

viscosity and Σk is a function of density, both of these parameters will change significantly 

during the transient. As temperature increases, SLAM decreases and Σk increases. These two 

changes together represent a situation with competing effects. As temperature increases, the 

value of SLAM decreases which leads to an increase in flow, while Σk increases leading to a 

decrease in flow. However, the magnitude of the changes are not equal and are not necessarily 

linear with respect to temperature. Given that these values are input as constants, it is not 

expected that the MELCOR computed flow will match the experimental data over the entire 

transient.

Sensitivity runs were performed with the 2-ring model where control functions were used to 

compute SLAM and Σk at every flow path in the assemblies. Note that the starting point model 

used for these sensitivities is not the final 2-ring model presented above, rather it is an earlier 

version.

As seen in Figures 5.45 through 5.56 the results of these sensitivities show that the flow rate is 

in much better agreement with the test data when using variable loss coefficients in both the 

lower power non-ignition tests as well as the 15 kW ignition test. While sensitivities were 

performed for all power levels used in the experiments, the figures only show results for the 

1 kW and 8 kW pre-ignition tests and the 15 kW ignition test. For each of the three power 

levels, the figures show flow in the central and peripheral assemblies as well as the peak 

cladding temperature in both central and peripheral assemblies. With the exception of the 1 kW 

pre-ignition test, the temperature response did not noticeably change given the change in flow 

rate. This is due to the fact that the experiments are dominated by radiation heat transfer which 

is not dependent on flow. At very low powers, convective heat transfer is important, and 

induced flow rates will make a difference in the results.

Computing the loss coefficients at each flow path takes eight control functions. To compute all 

the loss coefficients in just the center assembly of the 2-ring model requires 120 control 

functions. If this method were used for the 9-ring model, it would require over one thousand 

control functions, leading to an overly complicated and difficult to maintain model.

The resulting flow through the peripheral region did not show the same improvement as the 

center assembly with the use of variable loss coefficients. One possible reason for this is that 

the loss coefficients were based on tests with forced flow where the flow was somewhat uniform 

across the unheated assembly. In the case of the experiments, the heatup of the assembly is 

asymmetrical and flow would initially start only on the side that is being heated by the center 

assembly. The loss coefficients for this situation would be different than a case where flow was 

uniform across the assembly. If convective heat transfer is expected to be a significant factor, 

as would be the case in lower powered events, then a more thorough attempt at matching the 

experimental flow should be investigated.
5-36

Figure 5.39 Flow rate in central assembly for 1 kW pre-ignition test using the 2-ring 

model

Figure 5.40 Flow rate in peripheral assemblies for 1 kW pre-ignition test using the 

2-ring model
5-37

Figure 5.41 PCT in center assembly for 1 kW pre-ignition test using the 2-ring model

Figure 5.42 PCT in peripheral assemblies for 1 kW pre-ignition test using the 2-ring 

model
5-38

Figure 5.43 Flow rate in central assembly for 8 kW pre-ignition test using the 2-ring 

model

Figure 5.44 Flow rate in peripheral assemblies for 8 kW pre-ignition test using the 

2-ring model
5-39

Figure 5.45 PCT in center assembly for 8 kW pre-ignition test using the 2-ring model

Figure 5.46 PCT in peripheral assemblies for 8 kW pre-ignition test using the 2-ring 

model
5-40

Figure 5.47 Flow rate in central assembly for 15 kW ignition test using the 2-ring model

Figure 5.48 Flow rate in peripheral assemblies for 15 kW ignition test using the 2-ring 

model
5-41

Figure 5.49 PCT in center assembly for 15 kW ignition test using the 2-ring model

Figure 5.50 PCT in peripheral assemblies for 15 kW ignition test using the 2-ring model
5-42

5.6.2 Transition from pre- to post-breakaway oxidation kinetics

As discussed in Section 5.2.3, the coefficients in Equation 5.2 for the breakaway lifetime as a 

function of temperature were obtained by log-linear regression on experimental data obtained in 

tests performed at ANL and were subject to some uncertainty due to measurement error and to 

scatter in the data (See Figure 5.6). In order to understand the experimental results and assess 

the effect of the uncertainty, two sensitivity calculations were performed. The first is the use of a 

different oxidation model and the second is through the use of both low and high temperature 

pre-breakaway oxidation kinetics.

For the first sensitivity, an independent oxidation kinetics model was used. This model, which 

was recently added to the MELCOR code, was developed at the Paul Scherrer Institute and is 

referred to as the PSI oxidation kinetics model in this report. The PSI oxidation kinetics model 

was based on multiple separate and integral effects tests and is described in detail in 

Reference 15. While this model uses pre-breakaway parabolic kinetics similar to the default 

MELCOR model, it also includes the use of oxide thickness, linear kinetics, accelerated linear 

kinetics and a transition phase.

In the previous BWR spent fuel pool analysis (Ref. 18) it was discovered that better agreement 

to the test data could be made with use of both low and high temperature pre-breakaway 

oxidation kinetics correlations (see Section 8.4 of Reference 18). A sensitivity run was made 

where the high temperature oxidation rate was defined as 5.3 times that of the base (low 

temperature) oxidation rate. The factor of 5.3 is what was used in Reference 18. A transition 

from the low to high temperature oxidation rates was set to occur over the range 1,000 K 

[1,340°F] to 1,100 K [1,520°F]. This transition range is somewhat different than that used in the 

previous BWR analysis, as the values selected here were based on the PWR experimental data 

which showed a larger difference between PCT and average temperature than the BWR data. 

The transition region and new high temperature pre-breakaway kinetics are shown in 

Figure 5.51.

All sensitivities were made using the 2-ring multiple rod model. Figures 5.52 through 5.58 

present results from four cases: 1) default oxidation kinetics using ANL data best fit coefficients, 

2) default oxidation kinetics using ANL data lower uncertainty coefficients, 3) PSI oxidation 

model and 4) low and high temperature pre-breakaway oxidation kinetics.

Figures 5.52 and 5.53 present the peak cladding temperature in the center assembly at two 

different scales. As seen in these figures, the case using the ANL data best fit correlation 

approaches ~1,050 K [1,430°F] and reaches a quasi-steady condition where the total heat 

generated (core power and oxidation heat) is balanced by the convective and radiative heat 

losses. This results in a significant delay in ignition timing compared to the experiment. The 

system remains at this quasi-steady condition until the time-at-temperature is reached and 

breakaway occurs, which significantly increases the oxidation heat generated and leads to 

ignition. From Figure 5.6, it is seen that at 1,050 K [1,430°F] the time to breakaway is 

approximately 4 hours. In the case where the ANL data lower uncertainty curve is used, the 

time to breakaway (at the same temperature) is only 1 hour and demonstrates the large 

uncertainty in the breakaway data. When the lower uncertainty curve is used, the default 

oxidation kinetics model predicts ignition ~45 minutes before the experiment.
5-43

In addition to the difference in timing of breakaway (based on the best fit and lower uncertainty 

correlations) it appears that there is too little oxidation heat generated in the default oxidation 

model. In order to see the effect of additional oxidation heat, a high-temperature oxidation 

correlation was used (as described above). When using the low and high temperature oxidation 

correlations, the PCT does not reach a quasi-steady condition and increases at a rate slightly 

larger than the test data and shows ignition time is in excellent agreement with the experiment. 

The PSI model shows good agreement with the test data and predicts ignition about 30 minutes 

earlier than the experiment.

Figure 5.54 shows the peak cladding temperature in the peripheral region. Overall, these 

results are consistent with the center assembly and show delays from 30 to 60 minutes between 

center assembly ignition and peripheral assembly ignition. The PSI model adds more oxidation 

heat than the other models which results in a shorter time (30 minutes) between center and 

peripheral ignition. Table 5.4 shows ignition times for the sensitivity cases.

Figure 5.55 shows the total oxidation heat generation rate in the core. This figure shows that 

once breakaway occurs, significantly more oxidation heat is generated, leading to ignition. 

Figure 5.56 shows the change in center assembly peak cladding temperature over time (DT/dt). 

This figure shows that all the MELCOR cases show good agreement with the test data up 

through ~4.5 hours. At this time, all MELCOR cases and the experimental data show a PCT 

around 1,025 K [1,385°F]. In the case with the default oxidation kinetics and use of ANL best fit 

correlation, the DT/dt drops to zero while the data shows DT/dt levelling off around 50 K/hr 

[90°F/hr] and then increasing. In the case where the ANL lower uncertainty correlation is used, 

a sharp increase is seen just after 5 hours when breakaway occurs. The PSI model shows a 

more gradual increase after 5 hours and begins increasing faster than the data. The low and 

high temperature pre-breakaway oxidation kinetics case shows a much more gradual increase

similar to the test data, however, it occurs earlier than the test data.

Figures 5.57 and 5.58 show the ignition propagation in the center and peripheral assemblies 

respectively. As seen in these figures the case with the low and high temperature prebreakaway oxidation kinetics shows excellent agreement with the experiment. The default 

kinetics with ANL lower uncertainty correlation and the PSI model show similar results where 

ignition occurs earlier than the experiment. The case with default oxidation kinetics and the 

ANL best fit correlation show ignition is delayed over the experiment. In all cases, the rate of 

ignition propagation downward is in excellent agreement with the test data.

Overall, oxidation kinetics plays a significant role in this experiment. The default correlation in 

MELCOR tends to under predict the oxidation heat generated around 1,050 K [1,430°F] which 

results in a delay in breakaway. A sensitivity case with the ANL lower uncertainty correlation 

shows that the experimental data falls within the uncertainty range of the ANL data. The results 

indicate that there is reasonable uncertainty in the breakaway lifetime correlation that can yield 

a range of ignition times for this test spanning at least several hours. The use of low and high 

temperature pre-breakaway oxidation kinetics correlations resulted in excellent agreement to 

the experimental test data. The use of the PSI oxidation model resulted in good agreement with 

the test data showing ignition about 30 minutes early. Overall, the oxidation kinetics and the 

breakaway lifetime correlation are very important determinants of the predicted test results.
5-44

Table 5.4 Ignition times for the 15 kW ignition test

Description Time of ignition*

Center Peripheral

Experiment 6.31 hours 7.08 hours

2-ring multiple rod, ANL data best fit correlation 10.1 hours 10.9 hours

2-ring multiple rod, ANL data lower uncertainty correlation 5.62 hours 6.67 hours

2-ring multiple rod, PSI oxidation model 5.83 hours 6.38 hours

2-ring multiple rod, low and high temperature pre-breakaway 

oxidation kinetics 6.28 hours 7.27 hours

* Ignition defined when temperature exceeds 1,200 K [1,700°F].

Figure 5.51 Modification of the default MELCOR oxidation kinetics model
5-45

Figure 5.52 Peak cladding temperature in the center assembly for the 15 kW ignition 

test

Figure 5.53 Peak cladding temperature in the center assembly for the 15 kW ignition 

test
5-46

Figure 5.54 Peak cladding temperature in the peripheral assemblies for the 15 kW 

ignition test

Figure 5.55 Total oxidation heat generation rate in core for the 15 kW ignition test
5-47

Figure 5.56 Center assembly peak cladding temperature DT/dt for the 15 kW ignition 

test

Figure 5.57 Central assembly ignition propagation for the 15 kW ignition test
5-48

Figure 5.58 Peripheral assemblies ignition propagation for the 15 kW ignition test
6-1

6 SUMMARY

6.1 Phase II Test Results

Nine pre-ignition test were conducted with powers ranging from 1 kW to 15 kW which 

represented assemblies from as old as 96 months to as young as 3 months after offload 

(assuming 45 GWd/MTHM burnup). The pre-ignition tests were conducted for 12 hours or until 

the peak cladding temperature of 900 K [1,160°F] was reached in the center assembly. Tests 

conducted at powers greater than 8 kW were terminated before 12 hours. Tests at 1 kW and 

4 kW were conducted for 24 hours in attempts at achieving steady state. As a measure of 

repeatability, the 4 kW power was tested twice, once for 12 hours and again for 24 hours.

The second stage of the OECD/NEA Spent Fuel Pool Project successfully concluded with the 

ignition of the Phase II test assembly on June 6, 2012. The total applied power to the center 

assembly was 15 kW, which represented an assembly with a 45 GWd burnup approximately 

3 months after offload. The center assembly was surrounded on four sides by unpowered, 

peripheral assemblies. This configuration represented a “cold neighbor” configuration in a spent 

fuel pool. All five fuel assemblies were completely consumed as a result of the Zircaloy 

cladding fire. In addition, all of the pressurized fuel elements in the two pressurized, peripheral 

fuel bundles ballooned as a result of exposure to temperatures experienced during the ignition 

test.

Ignition of the Zircaloy was first observed in the center fuel assembly at the z = 3.302 m [130 in]

level and at an elapsed test time of 6.31 hours based on a thermal criterion of 1,200 K

[1,700°F]. The fire initially burned both upward and downward from this level but failed to reach 

the uppermost portion of the center assembly before becoming oxygen starved. The ignition 

front primarily swept down the fuel assemblies toward the most abundant source of oxygen at 

the assembly inlet. The burn progressed at an average rate of 0.457 m/hr [18 in/hr]. The 

cladding fire also propagated transversely into the peripheral assemblies starting at t = 7.08 h

and at the z = 3.150 m [124 in] level. The ignition had progressed across the entire cross 

section of all of the peripheral assemblies by t = 8.74 hr. Examination of the individual 

peripheral assemblies revealed a degree of asymmetry in the propagation into the peripheral 

assemblies by up to 0.27 hours. This asymmetry is not ascribed to the ballooning events in the 

pressurized assemblies but is rather most likely due to the limitations of construction and the 

chaotic nature of the Zircaloy fire.

The ballooning events in the two assemblies occurred within approximately 3 minutes of each 

other at a given location despite the North and South assemblies being filled to different 

pressures, 6.2 MPa [900 psi] and 5.2 MPa [750 psi] respectively. Ballooning did not likely occur 

at a single axial level but rather between z = 2.692 and 3.302 m [106 and 130 in] based on PCT 

data at the time each rod ballooned. The thermal-hydraulic behavior of assemblies with 

ballooned rods did not appear to be discernible from the unpressurized assemblies.

A residual gas mass spectrometer was used to monitor the concentrations of the gas 

constituents exiting the center assembly over a 130 hour period. The gas constituents 

monitored were nitrogen, oxygen, argon, water, carbon dioxide, and helium. Since argon and 

helium are inert noble gases and do not react with zirconium, the concentration of these gases 

relative to the reactive gas species provides a measure of the removal and release of the 

reactive gases. The analysis shows two distinct stages of oxidation over the course of the test. 

The primary oxidation stage was very energetic and continued for 11 hours after ignition. 

During this stage all the oxygen entering the assembly was consumed by oxidation of zirconium 
6-2

and a substantial portion of the nitrogen was consumed by production of the nitride from much 

(if not most) of the remaining unoxidized zirconium. During the 88 hour long second oxidation 

stage, all of the oxygen entering the assembly was consumed by the oxidation of zirconium 

nitride as indicated by the release of nitrogen gas. The additional release of energy from the 

wholesale production of zirconium nitride coupled with the significant reduction in convective 

energy removal during the primary oxidation stage are important phenomena to consider when 

assessing the propensity of an ignited assembly to propagate ignition to adjacent assemblies.

Finally, Table 6.1 lists a summary of events during the SFP Project Phase II ignition test. This 

list is not exhaustive but is intended to highlight the most significant events as judged by the 

authors.

Table 6.1 Summary events during the SFP Phase II ignition test

Local 

Time 

(MDT)

Elapsed 

Test Time 

(h)

Event

6/6/12 

0:00 -0.02 Data file started

6/6/12 

0:01 0.00 Test start, >95% of 15 kW

6/6/12 

6:19 6.31 Max center temperature = 1,200 K [1,700°F], ignition at z = 3.302 m (C_J-9_130)

6/6/12 

6:32 6.54 [O2] in center assembly measured by RGA dropped below 10%

6/6/12 

6:41 6.67 First pressurized rod ballooned

6/6/12 

6:41 6.68 Heater rod power lost

6/6/12 

6:41 6.82 Transverse propagation across center complete at z = 3.150 m (C_A-1_124)

6/6/12 

7:05 7.08 Max peripheral temperature = 1,200 K [1,700°F], ignition at z =3.150 m (N_J-3_124)

6/6/12 

7:35 7.50 CYBL vessel exhaust fan activated

6/6/12 

7:42 7.70 Nitrogen consumption by zirconium nitride (ZrN) formation began

6/6/12 

8:13 8.22 [O2] in South assembly dropped below 10%

6/6/12 

8:32 8.53 [O2] in East assembly dropped below 10%

6/6/12 

8:33 8.54 Last pressurized rod ballooned

6/6/12 

8:36 8.60 [O2] in West assembly dropped below 10%

6/6/12 

8:45 8.74 Transverse propagation across peripherals complete at z = 2.388 m (N_J-1_94)

6/6/12 

8:46 8.77 [O2] in North assembly drops below 10%
6-3

Table 6.1 Summary events during the SFP Phase II ignition test (Cont’d)

Local 

Time 

(MDT)

Elapsed 

Test Time 

(h)

Event

6/6/12 

12:18 12.29 Axial propagation in center to z = 0.559 m (C_J-9_22)

6/6/12 

12:54 12.91 Data acquisition failure

6/6/12 

17:24 17.40 Net nitrogen consumption ends, nitrogen release by ZrN oxidation dominates

6/8/12 

22:43 70.70 [O2] measured by RGA rises above 10%, O2 consumption and N2release ending

6/10/12 

4:01 100.00 All oxidation complete

6.2 MELCOR Results

Three MELCOR models were developed to simulate the OECD/NEA SFP Phase II experiments. 

These models included a 9-ring model, a 2-ring model and a 2-ring model with multiple 

rods/ring. The ignition test is characterized by the conditions of ignition in a central fuel 

assembly and the propagation of the burn to peripheral unheated assemblies. However, the 

2-ring model used with the current release version of MELCOR is unable to capture this 

phenomenon accurately because it cannot account for the large temperature gradient across 

the fuel assemblies that was observed in the experiments and necessary to predict radiant heat 

transfer and local ignition conditions. Code modifications were made to MELCOR 

Version 2.1.5287 which increase the number of fuel rod components available in a single ring 

and allow the user to simulate the test with much greater fidelity. MELCOR was successfully 

validated and this analysis shows that MELCOR is capable of predicting the conditions of the 

OECD/NEA SFP experiment with either a 9-ring or the 2-ring multiple rod model. The 2-ring 

multiple rod model (which was the focus of this analysis) shows good agreement to important 

parameters such as PCT, ignition time and ignition propagation when compared to the 

experimental test data when uncertainties in the oxidation kinetics are considered. When using 

the default oxidation kinetics model, better agreement to the experimental test data was found 

when using the lower uncertainty curve for breakaway timing over the best fit curve.

The new 2-ring multiple rod model has also been compared to the 9-ring model and 

comparisons show that while not identical, the new modeling is able to predict a similar 

response. The detailed 9-ring model did show slightly better agreement to the test data in terms 

of ignition timing in the center assembly than the 2-ring multiple rod model when using the code 

defaults for oxidation kinetics. As a result of the new code modifications that allow multiple 

rods/ring, code performance is improved and the 2-ring multiple rod input model is greatly 

simplified over the 9-ring model. It should be noted that the code modifications described are a 

work in progress and the models are currently strictly valid up until material begins to relocate. 

Further model development and refinement is necessary before these changes become part of 

a release version of MELCOR.

Two sensitivity studies were performed for simulation of the Phase II tests to investigate the 

effects on key figures-of-merit. The sensitivities on hydraulic loss coefficients showed that the 

MELCOR computed flow rate can better match the experimental data when variable loss 

coefficients are used, rather than constant values. These sensitivities also showed that the flow 
6-4

rate had little difference on fuel rod temperatures in all but the lowest powered case (1 kW preignition test). For situations where radiation heat transfer is the dominate heat transfer mode, 

the use of constant hydraulic loss coefficients is acceptable. For situations where convective 

heat transfer is dominate, variable hydraulic loss coefficients are recommended.

The sensitivities on pre- to post-breakaway oxidation kinetics showed that the results are highly 

sensitive to both oxidation kinetics and the transition to breakaway. Results from the 

sensitivities show that the predicted time of ignition can vary up to several hours based on the 

selection of individual models and selected modelling parameters. When using the default 

oxidation kinetics and default lifetime model (best fit curve in Figure 5.6), ignition times were 

delayed up to several hours, however, when using the lower uncertainty curve with the lifetime 

model, ignition times occurred about 45 minutes prior to that of the experiment. The PSI 

oxidation model was found to be more conservative than the default MELCOR model leading to 

ignition occurring 30 minutes prior to the experiment. Excellent agreement in ignition timing was 

obtained when using both low and high temperature pre-breakaway oxidation kinetics where the 

high temperature rate correlation was increased by a factor of 5.3 over the low temperature 

correlation as was done in the previous BWR study.
7-1

7 REFERENCES

1. NUREG/CR-7215, “Spent Fuel Project Phase I: Pre-Ignition and Ignition Testing of a Single 

Commercial 17×17 Pressurized Water Reactor Spent Fuel Assembly under Complete Loss 

of Coolant Accident Conditions,” U.S. Nuclear Regulatory Commission, Washington, DC, 

March 2016.

2. B.J. Ade and I.C. Gauld, 2011, “Decay Heat Calculations for PWR and BWR Assemblies 

Fueled with Uranium and Plutonium Mixed Oxide Fuel Using SCALE,” ORNL/TM-2011/290.

3. International Atomic Energy Agency, 2007, “Current Trends in Nuclear Fuel for Power 

Reactors,” 51st IAEA General Conference.

4. Y.S. Touloukian, and E.H. Buyco, Thermophysical Properties of Matter. Vol 5: Specific 

Heat of Nonmetallic Solids, IFI/Plenum, New York – Washington, 1970, pp. 142-144, Table 

42, Curve 10.

5. G.A. Berna, C.E. Beyer, K.L. Davis, and D.D. Lanning, “FRAPCON-3: A Computer Code for 

the Calculation of Steady-State, Thermal-Mechanical Behavior of Oxide Fuel Rods for High 

Burnup”, NUREG/CR-6534 Vol. 2, PNNL-11513 v.2, 1997, pp. 2.19-2.21.

6. Y.S. Touloukian, R.W. Powell, C.Y. Ho, and P.G. Klemens, Thermophysical Properties of 

Matter. Vol 2: Thermal Conductivity Nonmetallic Solids, IFI/Plenum, New York –

Washington, 1970, pp. 159-165, Table 27, Curves 62 to 70.

7. G. Zigh, and A. Velázquez-Lozada, “Experiments on Ignition of Zirconium-Alloy in a 

Prototypical Pressurized-Water Reactor Single Fuel Assembly in a Spent Fuel Pool During a 

Complete Draindown” J. Nuclear Engineering, Submitted for Publication (2013)

8. M. Steinbrück, “Prototypical experiments relating to air oxidation of Zirconium-4 at high 

temperatures,” J. Nucl. Mater. 392, 531-544 (2009).

9. M. Steinbrück, and M. Jung, “High-temperature reaction of α-Zr(O) with nitrogen,” 

Proceedings of ICAPP 2011, Nice, France, May 2-5, 2011, Paper 11395.

10. L. Krusin-Elbaum, and M. Wittmer, “Oxidation kinetics of ZrN thin films,” Thin Solid Films, 

107 (1983) 111-117.

11. NUREG/CR-6119, “MELCOR Computer Code Manuals, Vol. 1: Primer and Users’ Guide, 

Version 1.8.6,” U.S. Nuclear Regulatory Commission, Washington, DC, September 2001.

12. F.P. Incropera, and P. DeWitt, Fundamentals of Heat and Mass Transfer, John Wiley & 

Sons (1998).

13. NUREG/CR-6846, “Air Oxidation Kinetics for Zr-Based Alloys,” U.S. Nuclear Regulatory 

Commission, Washington, DC, July 2004.

14. NUREG/CR-7144, “Laminar Hydraulic Analysis of a Commercial Pressurized Water 

Reactor Fuel Assembly,” U.S. Nuclear Regulatory Commission, Washington, DC, 

January 2013.
7-2

15. Jonathan Birchley and Leticia Fernandez-Moguel, "Simulation of air oxidation during a 

reactor accident sequence: Part 1 – Phenomenology and model development," Annals of 

Nuclear Energy 40 (2012), pp. 163–170.

16. ANSYS FLUENT User Guide Version 14.5, ANSYS, Inc., Canonsburg, PA, 2012.

17. ANSYS FLUENT Theory Guide Version 14.5, ANSYS, Inc., Canonsburg, PA, 2012.

18. NUREG/CR-7143, “Characterization of Thermal-Hydraulic and Ignition Phenomena in 

Prototypic, Full-Length Boiling Water Reactor Spent Fuel Pool Assemblies After a 

Postulated Complete Loss-of-Coolant Accident,” U.S. Nuclear Regulatory Commission, 

Washington, DC, March 2013.

19. S.G. Durbin, and E.R. Lindgren, “Characterization of Hydraulic and Ignition Phenomena 

of Pressurized Water Reactor Fuel Assemblies: Phase II Test Plan”, SANDIA report 

SAND2012-1447, March 2012.
A-1

APPENDIX A – EXHAUST GAS ANALYSIS DETAILS

A.1 Introduction

The sampling and analysis of the gas stream exiting the center assembly was performed with 

an in-house built sample conditioning system coupled to a mass spectrometer (Pfieffer Model 

GSD 300T). With the Pfeiffer GSD 300T RGA, gases were ionized by a tungsten filament. 

These ions were filtered by the quadrapole mass spectrometer and then detected and amplified 

by the secondary electron multiplier (SEM). The SEM produced an electric current for each 

unique ion species, which was converted into gas concentrations and gas concentration ratios.

The gases of interest and the ion masses monitored were helium (4.0), water (18.0), nitrogen 

(28.0 and 14.0), oxygen (32.0 and 16.0), argon (40.0) and carbon dioxide (44.0). The analysis 

of these eight ions was recorded every 10 seconds.

A.2 Instrument Calibration

The responses for all the ions analyzed drifted due to temperature changes, filament age, 

instrument usage since refurbishment, and changes in baseline offset. In order to quantify 

instrument drift and signal deterioration, ambient air was sampled every two hours. During this 

two-hour sample gas period, measurements (cycles) were made about every 10 seconds for a 

total of 688 cycles. After two hours, a baseline offset calibration was performed and a threeway valve was activated to change the input gas to ambient air, which was sampled for about 

three minutes or a total of 18 cycles.

A.2.1 Baseline Drift Correction

The baseline offset calibration feature was used to reset the baseline value for signal noise to a 

standard current value, which was determined by averaging the signal in a region of the mass 

spectrum that typically has no current signal from the sample. Sample currents for all the 

different gas species were measured relative to this low baseline value. The offset calibration 

was performed every two hours, just before the analysis of ambient air. For the calibration, the 

collection tubing, fixtures, and fittings were purged with ambient air for two minutes. After the 

purge, the offset calibration was performed followed by the ambient air measurement. Figure 

A.1 shows the raw current for each ion analyzed as a function of time. The baseline offset 

calibration resulted in significant discontinuity of the currents before and after calibration for the 

minor components, helium and carbon dioxide. The discontinuity reflects the amount of 

baseline drift that occurred over the 2-hour sample analysis period. The direction and 

magnitude of the discontinuity changed over a 24 hour diurnal cycle suggesting the cause is 

due to small temperature changes in the gas sample ambient temperature. The baseline drift 

was corrected by assuming that the current for a given species should be the same before and 

after each calibration cycle. The required offset was added proportionally to each data point 

based on the time since the last calibration. The actual correction also included a 30 point 

moving average smoothing. The adjustment assumed a linear drift in baseline correction 

between calibration points. The effect of applying this treatment to the data was to smooth 

ridges caused by recalibration of the instrument and to set responses relative to the level of 

baseline noise. Figure A.2 shows the raw ion currents after the baseline offset correction and 

smoothing. The correction is applied to all species for all times except between about 5.3 hours 

and 11.2 hours when only helium is corrected. Over this period the uncorrected raw data for all 

species except helium are used so that the step changes in concentration due to rapid gas 

reaction with zirconium, argon release during rod ballooning and air infiltration into the gas 

sample are preserved unaltered.
A-2

A.2.2 Sensitivity Drift Correction

Figure A.2 also shows that a downward drift in overall sensitivity is evident for all ions analyzed. 

The gas sample consisted of ambient air near the beginning of the test (time zero) and at the 

end of the test (time 130 hours). The difference between the initial and final current for each ion 

is a measure of sensitivity drift over the test period. After 130 hours of analysis, the total drift for 

argon was about a factor of four.

1.0E-14

1.0E-13

1.0E-12

1.0E-11

1.0E-10

1.0E-09

1.0E-08

0 20 40 60 80 100 120

Raw Current (Amps)

Time (hr)

28 amu N2

32 amu O2

14 amu N

16 amu O

18 amu H2O

40 amu Ar

44 amu CO2

4 amu He

Figure A.1 Raw currents for all ions monitored.

Time (h)
A-3

1.0E-14

1.0E-13

1.0E-12

1.0E-11

1.0E-10

1.0E-09

1.0E-08

0 20 40 60 80 100 120

Smoothed Current (Amps)

Time (hr)

Baseline Corrected

28 amu N2

32 amu O2

14 amu N

16 amu O

18 amu H2O

40 amu Ar

44 amu CO2

4 amu He

Figure A.2 Baseline corrected currents for all ions.

The baseline corrected currents were used to determine the responses for all the major 

components of air including, helium, water, nitrogen, oxygen, argon and carbon dioxide. The 

two ions measured for nitrogen and oxygen were summed to yield a single current for these two 

gas species. The current value of each air component was divided by the sum of all currents on 

a dry basis (water is excluded) to determine the response ratio for each gas.

The response ratios determined from the analysis of ambient air every two hours served as a 

running calibration for determining the molar or volumetric fraction of each gas species in the 

sample gas. The accepted concentration of a given species in ambient air (see e.g. 

http://en.wikipedia.org/wiki/Atmosphere_of_Earth) was divided by the corresponding response 

ratio derived from the analysis of ambient air to generate a calibration factor. The concentration 

of a given species was determined by multiplying the response ratio generated by the analysis 

of the unknown sample gas by the corresponding calibration factor generated by the analysis of 

ambient air.

Each calibration period yielded slightly different correction values for equivalent gases. The 

difference between one value and the next was attributed partly to instrument drift and ambient 

temperature changes. A single two-hour sample analysis period was adjoined by two 

calibration periods. The two sets of correction factors were applied to the response ratios for a 

set of 688 sample gas measurements. The multiplier used to make the correction was 

calculated by linearly weighting the two sets of correction factors relative to the placement of the 

measurement with the calibration interval.

Time (h)
A-4

A.2.3 Calibration Correction for Low Oxygen

Over the entire test period, the oxygen content of the gas mixture exiting the top of the 

assemblies varied between normal ambient air and very low oxygen levels. The RGA response 

for nitrogen and argon (the primary gas species besides oxygen) were different when oxygen 

was part of the mixture versus the response when oxygen was mostly absent. The use of air as 

a calibration gas was appropriate when the oxygen concentration was above 11 percent; 

however, using an air calibration when the oxygen levels were below 11 percent resulted in a 

nitrogen concentration error of greater than 2 percent. Calibration factor correction relationships 

were derived for nitrogen and argon when oxygen was mostly absent in the sample gas. Five 

calibration gases described in Table A.1 were employed to derive these calibration curves.

Table A.1 Calibration gases for the RGA

Calibration 

gas Ar 

Balance 

N2 N2/Ar

Normalized 

N2/Ar

1 0.71% 99.29% 139.8 1.67

2 0.89% 99.11% 111.4 1.33

3 1.18% 98.82% 83.7 1.00

4 1.76% 98.24% 55.8 0.67

5 3.54% 96.46% 27.2 0.33

The five certified calibration gases were sampled the day before the ignition test began and the 

day after sampling was terminated. The response ratios for the calibration gases were 

evaluated against the correction factors for nitrogen and argon derived from using ambient air 

as a calibration gas. This was done to determine the apparent measured concentration of 

nitrogen and argon in low-oxygen conditions from using ambient air as a standard. The 

measured concentrations of nitrogen and argon from the pre- and post- test calibrations were 

averaged. The resulting correlations between the average measured concentration and the 

certified concentration are shown in Figures A.3 and A.4. It was important to establish this 

relationship because ambient air was used as a standard for calibration of gases in the sample 

stream during periods when there was very little oxygen present.
A-5

The correction correlations were used to correct the measured nitrogen and argon 

concentrations when the oxygen concentration in the sample stream was less than 11 percent. 

Figure A.5 shows the sample gas composition as a function of time over the entire 130 hour test 

period.

0.00

0.01

0.02

0.03

0.04

0.05

0.06

0.0

0.2

0.4

0.6

0.8

1.0

0 20 40 60 80 100 120

Minor Gas Composition (Mole Fraction Ar, CO2 )

Major Gas Composition (Mole Fraction N2, O2 )

Time (hr)

Total

N2

O2

Total

N2

O2

CO2

Ar

Argon 

release

during

ballooning

Ar

N2

CO2

Ar

O2

Excess argon 

slowly removed

Air 

infiltration

Figure A.5 Mole fraction concentration of major and minor constituents of sample gas.

Also shown on Figure A.5 is the total of all mole fractions, which should be unity. Any deviation 

from unity is a measure of the total error in the analytical scheme. From time zero until ignition 

Figure A.3 Calibration correlation for argon. Figure A.4 Calibration correlation for nitrogen.

Time (h)
A-6

at 6.3 hours and again from about hour 80 to the end of sampling, oxygen concentrations were 

at ambient levels and the total was unity. From about 20 to 70 hours the oxygen concentration 

was near zero and the calibration gas correction was applied. The total was about 1.007 or 0.7 

percent above unity. From about 10 hours until about 16 hours and again from about 69 to 75 

hours the oxygen concentration was unstable or in transition and the total was low by up to 2.5 

percent. After ignition occurred in the center assembly at 6.3 hours, the pressurized rods in the 

North and South peripheral assemblies ruptured. A series of ruptures occurred from about 7 

hours until 8.3 hours. Since the pressurized rods were filled with argon, each rupture caused a 

spike in the argon concentration. At about 7.5 hours the CYBL vessel extraction fan was 

energized, which began purging the vessel with fresh air. The excess argon was slowly 

removed and argon concentration did not return to ambient levels until about 30 hours.

A.3 Data Analysis: Argon and Helium Ratios

Since argon is inert and does not react with zirconium, the concentration of argon relative to 

other reactive gas species was used as a measure of their removal and subsequent release. 

Figure A.6 shows the argon ratios for nitrogen, oxygen and, carbon dioxide normalized by the 

ratios found in ambient air. A normalized argon ratio of unity signifies no change from ambient 

air. Normalized argon ratios that are less than unity signify removal of a gas species. 

Normalized argon ratios greater than unity signify the release of a gas species. However, due 

to the argon release during pressurized rod ballooning and the subsequent slow removal of this

excess argon the argon concentration between 6.5 and 30 hours is elevated above ambient. 

The increased argon concentration artificially reduces the argon ratios. In anticipation of this 

problem, helium was also measured. No certified calibration gas were available for developing 

a correction correlation for low oxygen conditions so the helium concentration was adjusted 

using the correlation developed for argon. The helium concentration in air is very low and near 

the detection limit of the instrument resulting in a noisy signal. Nonetheless meaningful data 

was obtained that is not influenced by the argon released during rod ballooning. Also shown in 

Figure A.6 is the reduced N2/He ratio.

The O2/Ar plot closely mirrors the oxygen concentration plot in the previous figure. At 6.3 hours 

when ignition occurs, the O2/Ar ratio drops sharply to near zero where it remains until just after 

60 hours. Air infiltration into the gas sample between 10 and 18 hours is evident. Between 60 

and 80 hours, the O2/Ar ratio raises to near ambient conditions. The return to ambient levels is 

complete at 105 hours signifying the end of all oxidation reactions.

The N2/Ar, N2/He and CO2/Ar plots reveal two distinct oxidation stages and are discussed in the 

subsections below.

A.3.1 Stage 1 Oxidation

During the initial or primary oxidation stage and immediately following ignition at 6.3 hours, the 

ratios drop below unity indicating consumption of the gases. The Stage 1 reactions all involve 

the oxidation of zirconium metal. During this period the oxygen is consumed by the direct 

oxidation of zirconium metal:

Zr + O2 → ZrO2 A.1

The N2 ratios drop below unity signifying that nitrogen gas entering the center assembly is partly 

consumed by the zirconium nitride formation reaction:

Zr + 1⁄2 N2 → ZrN A.2
A-7

The carbon dioxide argon ratio also drops, suggesting that CO2 is consumed by this zirconium 

carbide formation reaction:

2Zr + CO2 → ZrC + ZrO2 A.3

Note that during the primary oxidation stage, the oxidation reaction consumed a gaseous 

reactant and produce solid products with no gaseous products. With most of the oxygen and 

much of the nitrogen consumed, much of the ambient air drawn into the assembly was removed 

and only part exited the top. The large drop in the flow of gas exiting the assembly may explain 

why ambient air infiltrated the gas sampled from the top of the assembly during this period.

0.0

0.5

1.0

1.5

2.0

2.5

0 20 40 60 80 100 120

Normalized Argon or Helium Ratio

Time (hr)

Gas removed Gas released

N2/Ar

O2/Ar

CO2/Ar

N2/He

stage 1 stage 2

Figure A.6 Normalized argon ratios for nitrogen, oxygen and carbon dioxide.

A.3.1 Stage 2 Oxidation

The Stage 2 reaction primarily involved the oxidation of ZrN. The Stage 2 oxidation period 

begins at about 17 hours when the N2/He ratio rises above unity signifying the transition from 

nitrogen removal to nitrogen release by reaction:

ZrN + O2 → ZrO2 + 1⁄2 N2 A.4

Note that the transition from Stage 1 to Stage 2 occurs five hours earlier than indicated by the 

nitrogen to argon ratio due to the excess argon released by the rod ballooning.

Time (h)
A-8

Carbon dioxide is consumed by the oxidation of ZrN to form zirconium oxide and zirconium 

carbide and release nitrogen gas:

ZrN + 1⁄2CO2 → 1⁄2ZrO2 + 1⁄2ZrC + 1⁄2N2 A.5

At about 57 hours the CO2/Ar ratio rises above unity signifying the release of carbon dioxide gas 

by the following reaction:

ZrC + 2O2 → ZrO2 + CO2 A.6

The Stage 2 oxidation period was almost complete just after 80 hours, but was not completely 

over until about 105 hours.

During the secondary oxidation period gaseous products from the reaction were formed. 

Enough nitrogen gas was generated to offset half the oxygen removal. Since 90 percent of the 

gas drawn into the bottom of the assembly exited the top, there is little impetus for air infiltration 

into the gas sample as was seen during the primary oxidation period.

A.3.3 Implications

There are important implications of the staged oxidation progression that should be considered

when modeling the post-ignition behavior of the burning assembly. The addition of significant 

energy from the partial oxidation of most of the zirconium by nitrogen coupled with the 

substantial reduction of convective energy exiting the assembly likely leads to a greater 

temperature increase during the primary oxidation period than can be predicted by oxidation 

with oxygen alone. The elevated temperatures significantly increase the propensity for the burn 

to propagate to adjacent assemblies. The oxidation of zirconium by nitrogen and subsequent 

oxidation of zirconium nitride need to be included in any models that assess burn propagation to 

adjacent assemblies.

A.4 Summary

A residual gas mass spectrometer (RGA) was used to monitor the concentrations of the gas 

constituents exiting the center assembly over a 130 hour period. The gas sample stream was 

conditioned by filtering particulates and the sample pressure was held constant by an electronic 

back pressure controller. Ambient air was automatically sampled at two hour intervals to correct 

instrument drift, which drifted by a factor of four. Calibration gases containing argon and 

nitrogen were used to correct the argon and nitrogen concentrations for the periods when 

oxygen was absent from the gas sample. The calibration correction reduced the mass balance 

error from about 3 percent to less than 1 percent.

The ratio of nitrogen to argon and helium was used to monitor the reaction of nitrogen with 

zirconium. The analysis clearly shows two distinct stages of oxidation over the course of the 

test as summarized in Table A.2. The primary oxidation stage was very energetic and 

continued for 11 hours after ignition. During this stage all the oxygen entering the assembly 

was consumed by oxidation of zirconium and a substantial portion of the nitrogen was 

consumed by production of the nitride from much (if not most) of the remaining unoxidized 

zirconium. The nitride reaction releases about a third of the energy of complete oxidation. All 

reactions consumed gaseous reactants so the flow exiting the assembly was substantially 

reduced. The added energy release from wholesale production of the nitride coupled with the 

significant reduction in convective energy removal are important phenomena to consider when 

assessing the propensity of an ignited assembly to propagate ignition to adjacent assemblies. 
A-9

During the second oxidation stage, all of the oxygen entering the assembly was consumed by 

the oxidation of zirconium nitride as indicated by the release of nitrogen gas. This stage 

continued for 88 hours. The lower energy release of zirconium nitride oxidation (two thirds of 

complete zirconium oxidation) coupled with increased convective energy removal resulted in a 

less energetic burn during the second stage.

Table A.2 Summary of reaction regimes.

Start Time 

(hr)

End Time 

(hr)

Duration 

(hr) Description

0 6.3 6.3 Pre-ignition ambient conditions.

6.3 17 10.7 Primary oxidation stage. Initial reaction of 

zirconium metal, oxidation, nitride production, 

carbide production. Complete oxidation by 

oxygen and partial oxidation by nitrogen and 

carbon dioxide. No gaseous products of 

reaction formed resulting in substantial 

reduction of outlet gas flow. The low outflow 

allows air to mix with sample gas. Down flow 

of air to the reaction may be possible. The 

addition of significant energy from the partial 

oxidation of most of the zirconium inventory 

coupled with the substantial reduction of 

energy convection out of the assembly leads 

to much greater temperature increase than 

can be predicted by oxidation by oxygen 

alone. 

17 105 88 Secondary oxidation stage. Subsequent 

complete oxidation of partially oxidized 

zirconium, primarily ZrN, resulting in the 

release of nitrogen and carbon dioxide gases. 

105 130 25 Return to ambient conditions.





NUREG/CR-7216 Spent Fuel Pool Project Phase II: Pre-Ignition and Ignition Testing of a 1x4 Commercial 17x17 Pressurized

Water Reactor Spent Fuel Assemblies under Complete Loss of Coolant Accident Conditions