CNWRA 2003-04

Revision 1

ANALYSIS OF RAIL CAR COMPONENTS EXPOSED TO

A TUNNEL FIRE ENVIRONMENT

Prepared for

U.S. Nuclear Regulatory Commission

Contract NRC–02–02–012

Center for Nuclear Waste Regulatory Analyses

San Antonio, Texas
CNWRA 2003-04

Revision 1

ANALYSIS OF RAIL CAR COMPONENTS EXPOSED TO

A TUNNEL FIRE ENVIRONMENT

Prepared for

U.S. Nuclear Regulatory Commission

Contract NRC–02–02–012

Prepared by

Andre S. Garabedian

Darrell S. Dunn

Asadul H. Chowdhury

Center for Nuclear Waste Regulatory Analyses

San Antonio, Texas

November 2002
iii

ABSTRACT

Rail car components recovered from the train involved in the July 18, 2001, Howard Street

Tunnel, Baltimore, Maryland train derailment and fire were used to estimate the fire duration and

temperatures achieved by the components. Steel samples including sections of the box car

panels and a bolt from an air brake assembly were analyzed using standard metallurgical

methods to determine oxide layer thickness and the amount of metal lost as a result of the

elevated temperature exposure. Aluminum alloy air brake valve assemblies, which melted as a

consequence of the fire, were analyzed using a heat transfer model. 

Analyses of the recovered components suggest the surface temperature of the steel reached

700 to 850 °C [1,292 to 1,562 °F] assuming an exposure time of 4 hours at the elevated

temperatures. Independent assessment of fire duration could not be obtained from the steel

components because the oxide scale thickness and metal loss are dependent on both time and

temperature. Several limitations to the assessment of temperature were noted including the

effects of oxide-scale spalling and post-fire atmospheric exposure for a period of more

than 1 year. 

Thermal analysis of the aluminum air brake valve body located approximately 10 m [33 ft] from

the fire (at the brake end of Car 52) indicated melting occurred early in the fire event, and the

temperature achieved by this component was at least 600 °C [1,112 °F]. A similar aluminum

cover located at approximately 20 m [66 ft] from the fire (at the mid-point of Car 51) was only

partially melted indicating its temperature may have reached 600 °C [1,112 °F] for a limited time,

where an aluminum cover located approximately 30 m [98 ft] away from the spill site (at the

brake end of Car 53) was not affected by the fire exposure at all. This temperature profile

indicating a decrease in exposure temperature with distance away from the spill site was further

substantiated by the lack of damage to other components, such as railcar exterior paint.

The analyses conducted suggest the temperatures achieved by materials present in a confined

space fire are strongly dependent on the proximity of the component of interest to the fire

source. Gas temperatures near the fire source were likely in excess of 800 °C [1,472 °F] for

more than 30 minutes and the reaction of components in this region were likely influenced by

the direct radiation from the fire. At a distance of approximately 20 m [66 ft] from the fuel

source, where the dominant mode of heat transfer was convection, the exposure was capable of

generating surface temperatures as high as 600 °C [1,112 °F], however, only for a much shorter

duration. 
iv

CONTENTS

Section Page

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

FIGURES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vii

ACKNOWLEDGMENTS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . ix

1 BACKGROUND . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 1-1

2 OBJECTIVES AND SCOPE OF WORK . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 2-1

3 SAMPLES COLLECTED TO DATE . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1

3.1 Valve Assembly . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-1

3.2 Valve Materials . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-4

3.3 Steel Samples . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-4

3.4 Sand Samples . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-4

4 METALLURGICAL ANALYSES . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1

4.1 Analysis Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-1

4.2 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-3

4.3 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-11

5 ALUMINUM SAMPLE THERMAL MODEL . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-1

5.1 Assumptions . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-1

5.2 Basic Equations . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-2

5.3 Analysis Method . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-3

5.4 Results . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-3

5.4.1 Ramp-Plateau Function . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-3

5.4.2 Ramp-Plateau-Decay Function . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-5

5.5 Discussion . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-6

6 CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 6-1

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

FIGURES

Figure Page

3-1 The ABDX-L Control Valve . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-2

3-2 ABDX-L Service Portion Outline and Photograph . . . . . . . . . . . . . . . . . . . . . . . . . . . 3-3

4-1 Oxidation Rate Constants for Fe-21⁄4Cr-1Mo Steel as a Function of Temperature . . . 4-2

4-2 Reduction in Metal Thickness as a Function of Time for Isothermal Exposures in

the Range of 827 to 1,177 °C [1,520 to 2,150 °F] . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-2

4-3 Metal Oxide Thickness as a Function of Time for Isothermal Exposures at

Temperatures Ranging from 527 to 827 °C [980 to 1,520 °F] . . . . . . . . . . . . . . . . . . 4-3

4-4 Raman Spectra for Scales Recovered from Car 51 . . . . . . . . . . . . . . . . . . . . . . . . . . 4-4

4-5 X-Ray Diffraction of Orange Colored Scales from Car 51 . . . . . . . . . . . . . . . . . . . . . 4-5

4-6 X-Ray Diffraction of Black Colored Oxide Scales from Car 51 . . . . . . . . . . . . . . . . . . 4-5

4-7 Cross Section of Car 50 Roof Approximately 40 cm [16 in] from the Opposite End

of the Roof Panel Section . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-6

4-8 Cross Section of Car 50 Roof Approximately 3 cm [1.2 in] from the Edge Directly

Exposed to the Fire . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-7

4-9 Cross Section of Car 50 Roof at Edge Directly Exposed to the Fire . . . . . . . . . . . . . 4-7

4-10 (a) Secondary Electron Image from Car 50 Roof at Edge Directly Exposed to the

Fire, (b) Iron Image Map, (c) Oxygen Image Map . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-8

4-11 Scanning Electron Microscope Secondary Electron Image of Car 52 Air Brake

Valve Bolt Showing Metal and Oxide Layer . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-9

4-12 Scanning Electron Microscope Secondary Electron Image of Car 52 Air Brake

Valve Bolt Head Showing Iron and Aluminum and (b) Close-Up of Iron with

Aluminum . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 4-10

5-1 ASTM E 1529 (1993) Ramp Function and Corresponding Aluminum Thermal

Profile . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-4

5-2 Published Flame Temperature Ramp Function and Corresponding Aluminum

Thermal Profile . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-4

5-3 Ramp-Plateau Function Demonstrates Melting at Published Flame Temperature

for Exposure Times Greater than 360 Seconds . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 5-5

5-4 Ramp-Plateau-Decay Function and Corresponding Aluminum Thermal Profile . . . . . 5-6

6-1 Air Brake Valve Found 20 + m [66 + ft] from Fuel Spill Site . . . . . . . . . . . . . . . . . . . . 6-1
vi

ACKNOWLEDGMENTS

This report was prepared to document work performed by the Center for Nuclear Waste

Regulatory Analyses (CNWRA) for the U.S. Nuclear Regulatory Commission (NRC) under

Contract No. NRC–02–02–012. The activities reported here were performed on behalf of the

NRC Office of Nuclear Material Safety and Safeguards, Spent Fuel Project Office. The report is

an independent product of the CNWRA and does not necessarily reflect the views or regulatory

position of the NRC. 

The authors thank Yi-Ming Pan and Jason Huczek at Southwest Research Institute® for their

technical assistance provided at various stages of the project. The authors thank Yi-Ming Pan,

Jason Huczek, and B. Sagar for their technical and programmatic reviews of this report and

C. Cudd and A. Woods for editorial reviews. The authors are thankful to C. Patton for assisting

with the word processing and preparation of the report.

QUALITY OF DATA, ANALYSES, AND CODE DEVELOPMENT

DATA: No CNWRA-generated original data are contained in this report. Sources for other data

should be consulted for determining the level of quality for those data.

ANALYSES AND CODES: Spreadsheet calculations were accomplished using Microsoft®

Excel 2000. Analyses are documented in Scientific Notebook 550. 
1-1

1 BACKGROUND

On July 18, 2001, a 60-car train, carrying flammable liquids, corrosive acids, paper products,

and other commodities derailed while passing through the Howard Street Tunnel in Baltimore,

Maryland. This incident derailed 11 cars, Cars 46 through 56, while on an upward slope of

approximately 0.8 percent.

The derailment caused structural damage to Car 52, carrying approximately 106,000 L

[28,000 gal] tripropylene, puncturing the base of the tank and initiating a fuel spill. A fire

ensued, which burned freely for 3 hours. After approximately 3 hours into the event, a water

main ruptured above the tunnel, introducing cooling water to the tunnel. 

On August 27, 2002, the U.S. Nuclear Regulatory Commission (NRC) directed the Center for

Nuclear Waste Regulatory Analyses to support the NRC and the National Transportation Safety

Board in Washington, DC, to review and analyze the information obtained from the site visit of

the stored, fire damaged cars from the Baltimore tunnel accident. Information obtained from the

site visit and specimens from the rail cars were analyzed to estimate the duration of exposure

and temperatures achieved by the fire-damaged cars. The objectives and scope of work

discussed in this report are based on this directive.
2-1

2 OBJECTIVES AND SCOPE OF WORK

The goal of the project was to compile forensic evidence about the materials damaged during

the tunnel fire and to use engineering methods to gain insight into the thermal environment the

materials experienced during the fire. An estimation of the thermal environment at

approximately 20 m [66 ft] on each side of the fire source will also be provided.

The thermal environment is defined by a temperature-time history. Two analyses needed to be

conducted to establish this history:

Establish a Rough Temperature Achieved by Samples—Determine the rough-temperature

profile by analyzing the phase change in materials collected from the site. These data will

indicate the maximum temperatures achieved and may give a rough estimation of the duration.

The source of such data is the metallurgical analysis of the exposed steel and aluminum

components in the vicinity of the fire.

Predict the Duration—Correlate the rough-temperature history gained by metallurgical analysis

to physical observations (e.g., melting). Using heat transfer equations and the metallurgical

estimation of temperature, the predicted thermal environment can be imposed on the sample to

determine whether damage, consistent with the damage observed, could have been caused by

the exposure. In this case, the predictor will be the total phase change (melting) of a die-cast

aluminum cover plate.

The analysis will focus on the vicinity of the tripropylene car before the water introduction

because the thermal environment conditions are assumed to have been the worst. Analyses of

the fire environment were directed, specifically, to the fire environment surrounding three air

brake valve assemblies in the vicinity of the spilled tripropylene.

The prediction of the thermal environment in the vicinity of the air brake valve served to provide

a better estimate for the thermal environment in the zones of interest {15–20 m [49.21–65.62 ft]

to either side of a presumed spill site}. Data in the vicinity of the air brake valves will be

extrapolated to other zones of interest, because there was no readily observable evidence of

the thermal environment in these areas (e.g., no distinct melting, charring, or obvious

deformation).
3-1

3 SAMPLES COLLECTED TO DATE

Several samples were collected from the site. Samples were chosen because of their proximity

to the fire source and their level of damage.

These materials included

Steel Scale, Taken from the Brake End of Car 51—This location was adjacent to the brake

end of Car 52 and was approximately 10–12 m [32.81–39.37 ft] uphill from the fire source.

Section of Roof Plate from Car 50—The steel section was taken from the roof of Car 50

because this location was the source of a substantial, secondary fire as a result of the

prolonged paper fire.

Remainder of ABDX-L Air Brake Valve from Car 52—This air brake valve exhibited evidence

of prolonged fire exposure. The valve includes several die-cast aluminum covers, which had

completely melted as a result of the fire. The valve was found approximately 10 m [33 ft] uphill

from the spill site, and approximately 1 m [3.3 ft] from the tunnel floor. A similar valve assembly

was found at the midpoint of Car 51, approximately 20 m [66 ft] uphill from the fire source. 

Although the exact model type was not determined, the valve cover was clearly from the same

family of products and was composed of the same aluminum alloy. A third air brake valve was

found approximately 30 m [100 ft] from the spill site, and analyzed.

Exposed Bolts Remaining on Air Brake Valve Assembly—The bolts used to hold the

die-cast aluminum covers were retrieved for subsequent analysis.

Sand Sample Adhered to the Base of Rail #44—Rail #44 (as identified by National

Transportation Safety Board) was in the vicinity of the derailment and the fire source. The sand

sample was taken from the south face of the track. The sand was of interest because it had

exhibited some form of phase change, which transformed it from a powder to a consolidated

mass. 

New ABDX-L Air Brake Valve Cover and Attaching Bolts—A new valve cover and mounting

hardware were provided by the manufacturer, WABTEC. These components were collected

with the intent of defining a baseline for subsequent metallurgical analyses. Neither the

aluminum cover nor the bolts were fully analyzed as part of this project because these samples

were not received in time to be included in this report.

3.1 Valve Assembly 

The valve assembly found on Car 52 was identified as an ABDX-L valve. The valve assembly

located on Car 51 was also assumed to be an ABD valve assembly. A complete ABDX-L valve

assembly is shown in Figure 3-1.
Figure 3-1. The ABDX-L Control Valve (Clockwise from Top: Valve Body Schematic, Photograph of D

Car 52, Corresponding Outline of the Valve Body)

3-2
Figure 3-2. ABDX-L Service Portion Outline and Photograph

3-3
1

Dimensions scaled from WABTEC Drawing. 0592776,000,00

3-4

The standard ABDX-L-style valve consists of a normal service valve and an emergency valve. 

For this analysis, the damage observed on the service portion of the valve was analyzed (see

Figure 3-2).

The primary area of interest in the analysis was the aluminum alloy cover over the release valve

in the service portion of the valve assembly shown in Figure 3-2. The location of the cover is

marked by the four bolts remaining on one side of the valve, and the associated

aluminum residue.

The following assumptions were made regarding the aluminum component based on

observations made during the site visit and photographic documentation of the rail cars after

the accident:

Material: Die-Cast Aluminum Alloy (ASTM B 85-99 Grade 360, 

aluminum-silicon alloy)

Dimensions1

: 127 mm [5 in] diameter, 25.4 mm [1 in] thick

Heat Transfer Area: Circumference and one side 228 cm2 [35 in2]

Volume: Assuming full depth of 322 cm3 [20 in3]

Orientation: Vertical

3.2 Valve Materials

As stated by the manufacturer, the aluminum cover plate is composed of ASTM B 85-99

Alloy 360. It is assumed that the alloy is an aluminum-silicon alloy, the properties for which are

common for a range of alloy percentages. The alloy has the following properties

(Holman, 1990):

Thermal Conductivity (k): 177 W/m K

Density (ρ): 2,707 kg/m3

Specific Heat: 0.892 kJ/kg K

Melting Temperature: 600 /C [873 K]

3.3 Steel Samples

Samples were taken from Car 51 (scale), the air brake valve of Car 52 (bolt shown in

Figure 3-2) and the roof of Car 50. Samples from the roof panel of Car 50 and the bolt from the

airbrake valve of Car 52 were cut from their respective pieces using a low-speed saw. Care

was taken not to disturb or heat the oxide scales on these samples. 

3.4 Sand Samples

Samples of sand were taken from the Howard Street Tunnel away from the proximity of the fire

as well as sand that was adhered to the base of Rail #44.
4-1

4 METALLURGICAL ANALYSES

Metallurgical analyses were performed on each of the samples described in Section 3. The

analytical methods were selected to identify the phases present, determine the composition of

the phases, and where appropriate, measure important dimensions of the components. 

Observed properties were compared to published properties for each of the materials, and an

estimate of the thermal conditions required to produce these phase changes was derived.

4.1 Analysis Method

Sand samples collected from the Howard Street Tunnel and oxides on the steel specimens

were analyzed using x-ray diffraction and Raman spectroscopy. In addition, specimens were

mounted in epoxy and cross sectioned for metallurgical analyses. The cross-sectioned

specimens were examined using both an optical microscope and a scanning electron

microscope with an energy dispersive spectrometer. 

Estimations of component temperature and fire duration were made by assessing the extent of

iron oxidation and metal loss. Oxidation of iron and carbon steels is dependent on temperature. 

At temperatures as high as 1,600 °C [2,912 °F], both magnetite (Fe3O4) and hematite ("–Fe2O3)

are stable (Rapp, 1980). Wustite (Fe1!yO) is stable in the temperature range of 567 to 1,400 °C

[1,052 to 2,552 °F]. Wustite formed at elevated temperatures is not stable at lower

temperatures and is transformed to Fe + Fe3O4. The rate of oxidation is also dependent on

temperature. At a temperature of 250 °C [482 °F] or greater, the oxide on iron and steels grows

at a parabolic rate (Szlarska-Smialowska and Jurek, 1976; Runk and Kim, 1970). The thickness

of the oxide scale, X, is a function of time according to Eq. (4-1)

X kt (4-1) 2 = p

where kp is the rate constant and t is time. A similar expression can be used to determine the

metal recession rate as a consequence of oxidation (Simms and Little, 1987). 

Figure 4-1 shows the rate constants for metal reduction (kr

), and scale thickness (kp) obtained

for an Fe-21

/4Cr-1Mo steel and iron. Rate constants for Fe-21/4Cr-1Mo were reported by Simms

and Little (1987) and Larose and Rapp (1997) over a temperature range from 550 to 700 °C

[1,022 to 1,292 °F]. Kubaschewski and Hopkins (1962) reported rate constants for iron from

500 to 1,100 °C [932 to 2,012 °F]. The composition and microstructure of the steels are known

to affect oxidation kinetics. Additions of carbon increases the oxidation kinetics of steel. 

Oxidation kinetics are faster for fine pearlite compared to coarse pearlite or a spheroidized

microstructure (Runk and Kim, 1970). Based on data published by Kubaschewski and Hopkins,

(1962) , the reduction in metal thickness as a consequence of oxidation under isothermal

conditions is shown in Figure 4-2. The change in the rate of metal loss as a function of time is a

result of oxide spalling. Similarly, the metal oxide thickness as a function of time and

temperature can be calculated and is shown in Figure 4-3. It should be noted that the

calculated values for reduction in metal thickness and oxide scale thickness shown in Figures

4-2 and 4-3 do not consider the effects of composition because the calculations are based on

results observed for iron. The effect of microstructure was also not considered (Runk and Kim,

1970). Oxide cracking and spalling are assumed when the oxide scale is greater than 85 μm in

thickness (Simms and Little, 1987;(Kubaschewski and Hopkins, 1962). Error in the assessment

of temperature as a consequence of neglecting compositional and microstructural effects
4-2

0.0007 0.0008 0.0009 0.001 0.0011 0.0012 0.0013 0.0014 0.0015

1/T, 1/K

-14

-12

-10

-8

-6

-4

Log kr or kp, cm2 s-1

kr, Metal Reduction, Fe-21/4Cr-1Mo

kp, Scale Thickness, Fe-21/4Cr-1Mo

kp, Scale Thickness, Fe

kr, Metal Reduction, Fe

Figure 4-1. Oxidation Rate Constants for Fe-21⁄4Cr-1Mo Steel and

Iron as a Function of Temperature (Simms and Little, 1987;

Kubaschewski and Hopkins, 1962)

should be considered. Increasing the carbon content of steel from 0.2 to 0.8 wt% increases the

rate constant for oxide film thickness by 42 percent (Runk and Kim, 1970). Microstructure has

similar effect. Runk and Kim (1970) reported the rate constant for 0.81 wt% carbon steel with a

fine pearlite structure to be 52 percent greater than that for a spheroidized microstructure. 

Oxide film spalling is known to occur at temperatures as low as 700 °C [1,292 °F]

(Kubaschewski and Hopkins, 1962). Spalling results in exposure of the underlying metal and,

as a result, can significantly increase the rate of metal loss. Metal loss and oxide scale

thickness shown in Figures 4-2 and 4-3 are calculated assuming that the oxidation kinetics after

scale spalling are identical to a clean metal surface. It should be noted that the calculated oxide

scale thickness shown in Figure 4-3 is the total oxide scale thickness which considers the

change in oxidation rate as a result of spalling. The observed scale is likely to be much thinner

as a consequence of sloughing after cracking and spalling.

Additional metallurgical analyses of the undamaged steel components may have been useful to

assess the temperature and time of exposure. To perform these analyses, corresponding

components that were not exposed to elevated temperatures are needed to assess the effects

of the fire temperature and duration on microstructural changes. Because the corresponding

components were not available for comparison, however, these analyses were not performed. 
4-3

0 2 4 6 8 10 12 14 16 18

Time, Hours

0

100

200

300

400

500

600

R ed u ction in me ta l th ickne ss, microme te rs

527 oC

627 oC

677 oC

727 o

C

777 oC

827 o 877 C oC

Figure 4-2. Reduction in Metal Thickness as a Function of Time

for Isothermal Exposures of Iron in the Range of 527 to 877 °C

[980 to 1,610 °F]

0 2 4 6 8 10 12 14 16 18

Time, Hours

0

200

400

600

800

1000

Me tal o xid e th ickness, micro meters

527 oC

627 oC

677 oC

727 oC

777 oC

827 o

877 oC C

Figure 4-3. Metal Oxide Thickness as a Function of Time for

Isothermal Exposures of Iron at Temperatures Ranging from 527 to

877 °C [980 to 1,610 °F]
4-4

0 500 1000 1500 2000 2500

Raman Shift, cm-1

In te n sity, Arb itrary un its

226

294

414

507

613

1321

249

379

527

643

1059

1299

Back side of flake 

alpha-Fe2O3

Front side of flake 

gamma-FeOOH 

Figure 4-4. Raman Spectra for Scales Recovered from Car 51 

4.2 Results

Both Raman spectroscopy (Figure 4-4) and x-ray diffraction (Figures 4-5 and 4-6) of the scale

obtained from Car 51 indicate the presence of leproidocrocite ((–FeOOH), hematite ("–Fe2O3)

and possibly magnetite (Fe3O4). It should be noted that magnetite can easily be oxidized in air

to hematite. The presence of (–FeOOH, which covered large sections of the scales, indicates

that the specimens were exposed to either a humid environment or water. Formation of

(–FeOOH may have occurred as a result of water contact after rupture of the water main, fire

fighting efforts, or subsequent storage of the train cars involved in the derailment.

As indicated in Section 4.1, the thickness of the iron oxide scale is a function of temperature

and time. Although thickness of the scales obtained from Car 51 varied, a thickness of 400 to

430 μm [0.016 to 0.017 in] was obtained by examination of the specimen cross section. Initial

calculations using Eq. (4-1) and rate constants from Kubaschewski and Hopkins (1962) show

that an oxide of 420 μm [0.017 in] can be formed at 777 °C [1,430 °F] in 4 hours. At 827 °C

[1,520 °F], a 420-μm [0.017 in] thick scale can be produced in 2 hours. 
4-5

10 20 30 40 50 60

2 Theta, Degrees

0

100

200

300

Intensity, Counts

FeOOH

FeOOH

FeOOH

FeOOH

FeOOH

FeOOH

FeOOH

Fe

O3 4 or Fe

O2 3

Fe

O3 4 or Fe

O2 3

X-ray diffraction 

Car 51 Scale

Orange color

Figure 4-5. X-Ray Diffraction of Orange Colored Oxide Scales from Car 51

10 20 30 40 50 60

2 Theta, Degrees

0

400

800

1200

1600

Intensity, 

Counts

F e

O2 3

X-ray diffraction 

Car 51 Scale

black color

Fe

O2 3

Fe

O3 4

Fe

O3 4

Fe

O3 4 Fe

O2 3

Figure 4-6. X-Ray Diffraction of Black Colored Oxide Scales from Car 51
4-6

Figure 4-7. Cross section of Car 50 Roof Approximately 40 cm

[16 in] from the Opposite End of the Roof Panel Section that May

Not Have Been Exposed to the Fire. Thickness of the Metal was

Approximately 1,761 μm [0.069 in].

Specimens from the roof section of Car 50 were analyzed for oxide scale thickness as well as

the thickness of the remaining steel section. The first specimen was cut from the opposite end

of the roof panel section that may not have been directly exposed to the fire (Figure 4-7). The

second specimen was cut from the edge of the roof panel section that was directly exposed to

flames from the paper inside the box car (Figures 4-8 and 4-9). From the cross section of the

roof shown in Figure 4-7, the remaining metal thickness was determined to be 1,761 μm

[0.069 in]. Near the edge of the roof, the metal thickness varies from 829 μm [0.033 in] at the

edge (Figure 4-9), to 1,464 μm [0.058 in] at a distance approximately 3 cm [1.2 in] from the

exposed edge (Figure 4-8). Although it is apparent that the metal thickness of the roof section

shown in Figure 4-7 has been reduced by oxidation, the metal thickness of this section was

used as a basis to determine metal loss. Assuming that metal loss occurred at equal rates on

both sides of the specimen shown in Figures 4-8 and 4-9, the metal loss from each side varied

from 148 to 466 μm [0.006 to 0.018 in]. Based on the calculated reduction in metal thickness as

a function of time and temperature shown in Figure 4-2, temperatures in the range of 750 to

850 °C [1,382 to 1,562 °F] for approximately 4 hours would be required to achieve the reduced

metal thickness. 
4-7

Figure 4-9. Cross Section of Car 50 Roof at Edge Directly

Exposed to the Fire. Thickness of the Metal Ranged from 829 to

930 μm [0.033 to 0.037 in]. 

Figure 4-8. Cross Section of Car 50 Roof Approximately

3 cm [1.2 in] from the Edge Directly Exposed to the Fire. 

Thickness of the Metal was Approximately 1,464 μm [0.057 in].
4-8

Figure 4-10. (a) Secondary Electron Image from Car 50 Roof at Edge Directly Exposed to

the Fire, (b) Iron Image Map, (c) Oxygen Image Map 

Figure 4-10(a) shows a closeup of the metal oxide scale of the Car 50 roof section near the

edge exposed to the fire. Cracks and spalling of the oxide are apparent from the cross section.

Figure 4-10(b) shows the scanning electron microscope-energy dispersive spectrometer

element map for iron, and Figure 4-10(c) shows the element map for oxygen which confirms the

scale is an iron oxide. Based on the analyses of the cross-sectioned specimen, the oxide scale

on the roof section exposed to the fire ranged from 50 to 81 μm [0.0019 to 0.0032 in] in

thickness. According to the calculations shown in Figure 4-3 scales of this thickness can be

produced at temperatures below 700 °C [1,292 °F] for a period of 4 hours. However, the oxide

scale shown in Figure 4-10(a) has clear signs of spalling which may have reduced the scale

thickness that remained attached to the damaged roof section. 

The cross section of the air brake bolt is shown in Figure 4-11. The oxide scale on the bolt

shank was determined to be 53 μm [0.0021 in]. Little spalling or cracking of the oxide layer was

observed, suggesting the temperature was less than that achieved by the roof panel of Car 50. 

Based on the predicted metal oxide thickness shown in Figure 4-3, a 53 μm [0.0021 in] oxide

scale could be expected in 4 hours at 627 °C [1,160 °F]. At higher temperatures, spalling of the

oxide would be expected. 
4-9

Figure 4-11. Scanning Electron Microscope Secondary

Electron Image of Car 52 Air Brake Valve Bolt Showing

Metal and Oxide Layer 

Figure 4-12(a) shows a cross section of one of the air brake assembly bolts. During the

analysis of the specimen, several regions were found where an aluminum–iron phase had

formed. A higher magnification of such a region is shown in Figure 4-12(b). Using an energy

dispersive spectrometer, the composition of the bolt was determined to be 98.4 atomic percent

iron with no measurable concentration of aluminum. In the mixed aluminum–iron region, the

composition was found to be 66 atomic percent aluminum, 4 atomic percent silicon, and

30 atomic percent iron. The liquidus temperature for a 66 atomic percent Aluminum and

34 atomic percent iron binary is 1,180 °C [2,156 °F] and the solidus temperature is 1,165 °C

[2,129 °F] (ASM International, 1992). The aluminum–iron binary phase diagram does not

contain low melting point eutectic compositions. The addition of 4 atomic percent silicon is not

expected to significantly alter the solidus or liquidus temperatures based on examination of the

iron–silicon and aluminum–silicon binary systems. 

The aluminum–iron phase identified on the air brake valve bolt may have been formed as a

result of diffusion of aluminum into the iron at elevated temperatures. Eq. (4-2) describes

simple diffusion

x Dt = (4-2)

where x is the diffusion distance (cm) [in], D is the diffusion coefficient (cm2 s!1) [in2 s!1] and t is

time (seconds). Assuming a diffusion distance of 100 μm [0.0039 in] and a time of 10 hours, the

calculated value of D is 2.8 × 10!9

 cm2 s!1 [4.3 × 10!10 in2 s!1]. The value of D is dependent on

temperature. For face center cubic iron ((-Fe), a self-diffusion coefficient of 2.8 × 10!9

 cm2 s!1

[4.3 × 10!10 in2 s !1] corresponds to a temperature of 1,464 °C [2,667 °F] (Geiger and

Poirier, 1980). 
4-10

Figure 4-12. (a) Scanning Electron Microscope Secondary Electron Image of Car 52

Air Brake Valve Bolt Head Showing Iron and Aluminum and (b) Closeup of Iron

with Aluminum 

The sand samples were identified as quartz (SiO2) using Raman spectroscopy. Results of x-ray

diffraction analyses also indicated quartz and small amounts of ankerite [Ca(Fe,Mg)(CO3)2]. 

The sample obtained from Rail #44 also had Raman peaks consistent with the presence of

(–FeOOH. As previously indicated, the presence of (–FeOOH is consistent with corrosion of

iron and steel rather than oxidation at elevated temperatures. 
4-11

4.3 Discussion

Analyses of the scale obtained from Car 51 suggest the oxide was initially formed as a result of

high-temperature oxidation. Subsequent exposure to water or humid air resulted in the

formation of iron hydroxides that are typical of corrosion products for steel structures exposed to

a humid or aqueous environment. Calculations of the time and temperature required to form

oxide scales of similar in thickness to the specimens obtained from Car 51 or from the roof of

Car 50 do not consider changes in thickness that may result from subsequent exposure to

humidity or water. 

The thickness of the oxide scales recovered from Car 51 and the reduced metal thickness

observed in the roof panel of Car 50 correspond to temperatures in the range of 750 to 850°C

[1,382 to 1,562 °F] ] for a period of approximately 4 hours. The estimation of the temperature

does not consider the effects of steel composition and microstructure (Runk and Kim, 1970;

Hauffe, 1965) or the partial pressure of oxygen (Kubaschewski and Hopkins, 1962) on the

oxidation kinetics. Cross sections of the roof panel from Car 50 clearly indicate significant

cracking of the oxide, suggesting the roof panel reached temperatures where spalling of the

oxide may be expected. Cracks in the oxide scales tend to increase the oxidation rate of the

underlying metal. Assessment of the oxide layer thickness is compromised by scale thickness

variation and scale spalling. As a result, assessments of the temperature of the component and

fire duration using the thickness of the oxide scales may have significant errors. 

Estimations of component temperature and fire duration from the analysis of the oxide scale

thickness and the reduction in the metal thickness are dependent on the selection of kinetic

parameters. As indicated in Section 4.1, the rate constants are dependent on several factors,

including microstructure and composition. Additional calculations were performed to determine

the effects of uncertainty in the values of the rate constants. Multiplying the rate constants

shown in Figure 4-1 by a factor of 2 increased the oxide scale or reduced metal thickness by

approximately 50 percent. 

Spalling of the oxide may also be problematic for the assessment of temperature because of

uncertainty in the conditions necessary for spalling to occur and the uncertainty in the oxidation

rates after spalling of the oxide scale. On the other hand, the adhesion of iron oxides to iron is

a function of temperature and scale thickness (Kubaschewski and Hopkins, 1962) and may

provide a marker for the temperature achieved by the steel components. For scales of 200 μm

[0.0079 in] thickness, the adhesion increases as temperature increases from 650 to 775 °C

[1,202 to 1,427 °F] and then decreases to zero at 950 °C [1,742 °F]. When the adhesion

strength decreases, the oxide can be easily removed from the metal, and spalling can occur. 

Spalling of oxide scales on iron and steel is generally observed at temperatures of 850 °C

[1,562 °F] but has been observed as low as 700 °C [1,292 °F] (Kubaschewski and Hopkins,

1962). The scale recovered from the brake end of Car 51 adjacent to Car 52 was completely

separated from the underlying steel, suggesting the steel reached temperatures above 700 °C

[1,292 °F]. 

The oxide scale on the bolt recovered from the air brake valve of Car 50 appeared to be more

intact than the scale on the roof panel of Car 50, suggesting that the oxide on the bolt was

formed at a lower temperature. The observation of the aluminum–iron–silicon phase on the

underside of the bolt head is inconsistent with the lower component temperature suggested by

the intact, 53-μm [0.0021-in] thick oxide scale on the shank of the bolt. It is unlikely that the
4-12

formation of aluminum–iron–silicon occurred as a result of diffusion of aluminum into the steel

bolt. No low melting point eutectic exists in the aluminum–iron, aluminum–silicon, or the

iron–silicon systems that would result in the formation of this phase. 

Initial analyses of the sand collected from the tunnel did not provide a means to estimate

temperature. Quartz has a melting temperature of 1,610 °C [2,930 °F], which is greater than the

solidus temperature of carbon steel. Although not completely conclusive, the x-ray diffraction

analyses indicated the presence of ankerite, which contains carbonate. Carbonate minerals

tend to decompose and evolve carbon dioxide at elevated temperatures. Specific information

on ankerite was not available but ankerite is similar to dolomite [CaCO3AMgCO3], which

decomposes at temperatures above 730 °C [1,346 °F]. Because ankerite appears to be present

in both sand samples collected at locations close to and away from the fire, it appears that the

sand collected from Rail #44 did not achieve temperatures in excess of 730 °C [1,346 °F]. 
5-1

5 ALUMINUM SAMPLE THERMAL MODEL

A fundamental heat transfer analysis of the aluminum air brake cover found on Car 51 was

conducted. The procedure followed is commonly used to predict the length of exposure required

to bring a structural element to a critical temperature and the failure times in steel sections,

based on strength reductions at elevated temperatures.

5.1 Assumptions

To simplify our initial analysis, the following assumptions were made:

Lumped Mass—The aluminum components were assumed to be lumped masses of solid

aluminum. It is also assumed that the temperature of the relatively small volume of aluminum is

uniform. This assumption is conservative because it will over predict the amount of aluminum

present for phase change.

Full Volume of Aluminum—In the absence of an accurate values of aluminum volume

contained in the valve cover, a calculation of the volume of the cover was made based on

dimensions provided in WABTEC drawings. This assumption is conservative because it will

predict higher temperatures to cause melting of the larger volume.

Complete Melting—The aluminum section reached its liquidus temperature. This is a good

assumption considering the condition of the covers in Figure 3-2.

Zero Conductive Losses to Adjacent Valve—This analysis considers the aluminum portion

separate from any adjacent material. This condition will predict slightly faster melting times than

the actual case, which is, however, a fair assumption when the cover deforms and drips away

from the main valve assembly. 

Convective Heat Transfer Coefficient—The heat transfer coefficient was assumed to be

50 kW/m2K. This value was taken from the Eurocode, as referenced in Buchanan (2001). This

heat-transfer coefficient is suggested when calculating the exposure of steel elements in a

hydrocarbon fire. According to Buchanan, heat transfer is not strongly heat-transfer coefficient

dependent because the primary heat-transfer mode is radiation.

Flame Temperature—A common flame temperature for a typical hydrocarbon fuel is between

810 and 925 °C [1,490 and 1,697 °F] (DeHaan, 1991). The analysis utilized two types of ramp

functions; ramp to a maximum temperature and plateau; and ramp-plateau-decay temperature

profiles. The ramp-plateau decay fire profile most accurately represents the conditions in the

tunnel because of the ventilation constraints of the tunnel geometry. Results for ramp-plateau

and ramp-plateau-decay functions are presented in Section 5.4.

Radiative Exposure—The added exposure because of radiation from the luminous flame was

not considered. The radiative fraction is typically taken to be approximately 30 percent of the

total heat release of the burning fuel. This fraction changes with changes in oxygen availability

and burning efficiency of the fuel. Appropriate values for radiant exposures range between

20 and 50 kW/m2. These values were chosen based on published radiative fractions of known

hydrocarbons (Tewarson, 1995; Drysdale, 1985). The exclusion of radiative inputs will lead to

longer predicted exposure times.
5-2

5.2 Basic Equations

The analysis is based on the principle that heat entering the aluminum section over a period of

time (∆t) raises the temperature by some amount (∆T). This concept is represented as

q Ft c V T &" (5-2) Al Al Al • • •• ∆ ∆ = ρ

where

q hT T T T &" ( ) ( ) (5-3) c f Al f Al = −+ − σ ε 4 4

Substituting Eq. (5-2) into Eq.(5-1) yields:

∆ ∆ T { } ( )( ) (5-4)

F

V c hT T T T t s

s s

c f Al f Al = −+ − • • 1 4 4

ρ

σ ε

where

∆TAl — Change in aluminum temperature (K)

F — Surface area (m2)

V — Volume (m3)

ρ — Density of aluminum (2,707 kg/m3)

cAl — Specific heat of aluminum (0.892 kJ/kgK)

hc — Heat transfer coefficient (50 W/m2K)

Tf — Flame temperature (K)

Tal — Aluminum temperature (K)

σ  Stefan-Boltzman constant (56.7× 10!12 kW/m2K4

)

Emissivity (ε) is given by

ε (5-5)

ε ε

=

+





 



 − 



 





1

1 1 1

S R

where

 εS — Emissivity of the source

 εR — Emissivity of the receiver

Emissivities were assumed to be 0.67 for both the emitter and the receiver (Buchanan, 2001).
5-3

5.3 Analysis Method

A timestep (∆t) was chosen, and the equations were solved to provide the ∆T at each timestep.

A Microsoft® VisualBasic program was used for the repetitive calculations. A brief sensitivity

analysis was performed on hc and ε, because these two variables are easily changed (as

opposed to the material property variables).

Two temperature functions were tested, as described in Section 5.1. The analysis focused on

ramp-plateau and ramp-plateau-decay temperature profiles. Standard values for nonene

(or similar hydrocarbon fuels) indicate a flame temperature between 810 and 925 /C [1,697 /F]. 

These values were used as starting points for the thermal analysis. Where possible, data

obtained from metallurgical analyses were used to further adjust the temperature profile.

5.4 Results

Preliminary model results for several possible fire scenarios are provided. All the results

presented assume the same sample geometry, material properties, and heat transfer

coefficients. The constant inputs are provided in Sections 3.1 and 5.1. Fire temperature

profiles were varied and predicted aluminum profile performance was noted.

5.4.1 Ramp-Plateau Function

The ramp-plateau function assumed the maximum temperature was realized in 3 minutes, and

was maintained for the duration of the exposure. This ramp function is typical of hydrocarbon

fires modeled in test furnaces. The American Society for Testing and Materials (1993) suggests

a temperature of 1,093 /C for 3 minutes and 1,180 /C [2,156] for 5 minutes, and all times

thereafter. This is a very conservative assumption considering the limited availability of fuel

(based on spill rate) and the limited supply of oxygen for combustion (based on

tunnel geometry).

When modeled using this fire exposure, the sample aluminum element exhibits melting at

300 seconds into the exposure (see Figure 5-1).

A more realistic ramp function would use the upper and lower bound for flame temperature, as

found in the literature. These temperatures are 810 and 925 /C [1,490 and 1,697 /F],

respectively. These exposure conditions predict melting at approximately 500 seconds.

Both of these models predict the aluminum melted before a reasonable fire duration could

be obtained.

Figure 5-2 illustrates that, when taking the published flame temperature as the maximum

temperature experienced by the valve, any exposure time greater than 360 seconds will

produce melting conditions.
5-4

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

0

200

400

600

800

1000

1200

Aluminum Temperature

Flame Temperature

Bulk Aluminum Temperature, °C

Model Run Time, seconds

Ex

posure Tem

perature, °C

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

0

200

400

600

800

1000

1200

Aluminum Temperature

Flame Temperature

Bulk Aluminum Temperature, °C

Model Run Time, seconds

Ex

posure Tem

perature, °C

Figure 5-1. ASTM E 1529 (1993) Ramp Function and Corresponding Aluminum Thermal

Profile (Melting Predicted at 300 seconds)

Figure 5-2. Published Flame Temperature Ramp Function and Corresponding Aluminum

Thermal Profile (Melting Predicted at ~500 seconds) 
5-5

0

200

400

600

800

1000

1200

0 100 200 300 400 500 600 700

0

200

400

600

800

1000

1200

Aluminum Temperature

Flame Temperature

Bulk Aluminum Temperature, °C

Model Run Time, seconds

Ex

posure Tem

perature, °C

Thermal Soak Time

Figure 5-3. Ramp-Plateau Function Demonstrates Melting at Published Flame

Temperature for Exposure Times Greater than 360 Seconds

5.4.2 Ramp-Plateau-Decay Function

A more realistic approach is to take into account a slower growth rate and the decay in flame

temperature as oxygen becomes scarce in the tunnel. Flammable liquids burning in pool fires,

in open air, can be expected to provide temperature profiles consistent with Figures 5-2

and 5-3. Fires burning under limited oxygen conditions will decay following the initial rise and

stabilize at some equilibrium temperature until ventilation conditions change or fuel is

exhausted. It is also safe to assume the temperatures at the valve {roughly 10 m [33 ft] from

the source} would have taken longer to reach the flame temperature. According to ASTM E

1529 (1993), a 3-minute ramp will completely immerse an element in an open-burning fire. A

7.5-minute initial ramp was chosen here. A decay to half the fire temperature was assumed to

have occurred 18 minutes into the fire exposure (assuming oxygen starvation).
5-6

0

200

400

600

800

1000

1200

0 200 400 600 800 1000 1200

0

200

400

600

800

1000

1200

Aluminum Temperature

Flame Temperature

Bulk Aluminum Temperature, °C

Model Run Time, seconds

Ex

posure Tem

perature, °C

Figure 5-4. Ramp-Plateau-Decay Function and Corresponding Aluminum Thermal Profile

(Melting Predicted at ~800 seconds) 

The conditions depicted in Figure 5-4 may be more difficult than originally assumed because

810 /C [1,490 /F] is the actual, published flame temperature of the fuel. The valve would have

had to be completely engulfed in flames for it to realize these temperatures. Considering the

location of the value, 10 m [33 ft] uphill, and its position in the tunnel (low to the floor), total

flame engulfment may have only occurred intermittently and, possibly, for only a short duration.

5.5 Discussion

The model was run with several variations in hc and ε. In all cases, the model predicted the

melting occurred between 793 seconds (assuming hc = 25 W/m2K and ε = 0.26) and

463 seconds (assuming hc = 50 W/m2K and ε = 0.5). The predictions of melting at far less than

1,800 seconds could not be used to ascertain the duration of the exposure.

Even when not considering the radiative exposure of the sample to the flame luminosity, the

melting time of the valve cover is predicted to have occurred early in the exposure. Considering

the radiative inputs would have considerably decreased the melting time. In relation to

laboratory testing, in order to simulate a radiative heat flux of 50 kW/m2, samples are exposed to

a furnace environment of approximately 800 °C [1,472 /F].

Had the model predicted melting at over 1,800 seconds, one could deduce the fire duration was

at least 1,800 seconds. The data indicated the aluminum section melted very early in the fire

duration and the aluminum was severely damaged as a result of the fire exposure before the

aluminum performance could be used to establish fire duration. 
6-1

6 CONCLUSIONS

Metallurgical analyses were conducted to assess the fire temperature and fire duration. The

analyses were partially confounded by several factors including the atmospheric exposure of the

rail car components for a period of more than 1 year following the fire. Another limitation of the

analyses was the lack of suitable controls, or components that were not exposed to the fire, that

could be used to assess extent of fire damage to the recovered components. Metallurgical

analyses suggest that the fire temperature at the end of Car 51, adjacent to Car 52, was at least

700 /C [1,292 /F] and likely closer to 850 /C [1,562 /F], based on the observation of spalling of

the oxide scale. Based on the reduction in the thickness of the roof of Car 50, temperatures for

this component were estimated to be in the range of 750 to 850 /C [1,382 to1,562 /F] for a

period of 4 hours. The oxide formed on the steel bolt from the air brake valve assembly of Car

52 suggests the temperature in this region was estimated to be 627 /C [1,160 /F] for an

exposure period of 4 hours. 

The preliminary data reviewed, including a first-look metallurgical analysis coupled with a

simplified thermal model, indicate conditions causing melting in the aluminum could only occur if

temperatures in excess of 600 /C [1,112 /F] were achieved in the sample, for an appreciable

duration.

A second aluminum witness was found approximately 20 m [66 ft] uphill of the fire source. This

piece is shown in Figure 6-1. It is interesting to note that this piece did not melt completely,

which would indicate the very specific conditions required to melt the valve assembly on Car 52

were likely not encountered at approximately 20m [66 ft] away. 

Figure 6-1. Air Brake Valve Found 20+ m [66 ft] from Fuel Spill Site (Note Only Partial

Melting at This Location)

Closer analysis of the damage caused to the valve cover found at Car 52, indicates that a

separation distance of approximately 20 m [66 ft] was sufficient to limit the exposure conditions

in this area, as evidenced by the level of damage to the valve cover. Considering only partial
6-2

melting in this area, temperatures at the surface of the aluminum were only slightly above 600/C

[1,112/F] (enough to initiate melting) but were not maintained for any appreciable duration. 

A third valve, of similar construction to the valves found on Cars 51 and 52, was also found on

Car 53 (approximately 40 m [downhill] from the tripropylene spill point). This valve exhibited no

signs of damage at all, and was found completely intact. Thus, at separation distances of 40 m

or greater, the exposure conditions were demonstrated to be less than 600/C.

The exposure profile derived from the post-fire condition of a number of aluminum valve covers

is further supported when one considers other damage noted on the railcars. Railcars (e.g. Cars

50 and 51) were painted with Dupont paint No. 909M-22416, an air-dried, high solid, alkydenamel paint. Given the observed damage to the steel and paint on Cars 51, 52, and 53, the

blistering temperature of 700 /C was only realized within 15 m of the spill site. These distances

were measured from the tops of the railcars (worst case), where both heat and corrosive

products of combustion were concentrated. There was no damage to the paint or steel on Car

53, adjacent to the fuel spill car (52), and located 12 m from the spill site.

The exposure conditions of standard compliance tests are based on a diesel, or similar

hydrocarbon, pool fire burning in an open pool configuration. For testing purposes, this fire

scenario provides exposure temperatures of approximately 800 /C [1,472 /F] for 30 minutes. It

is also assumed in this scenario the item being tested is in close proximity to the fuel spill

source and is completely immersed in the fire environment. Administrative controls can ensure

conditions that would limit the proximity of fuel sources to important cargo. The data obtained

during this review of the Baltimore Tunnel Fire indicate the separation of cargo from potential

fuel sources will be an acceptable method of limiting their exposure.
7-1

7 REFERENCES

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ASM International. 1992. 

ASTM. “Standard Test Methods for Determining Effects of Large Hydrocarbon Pool Fires on

Structural Member and Assemblies.” ASTM E 1529–93. 1993.

Buchanan, A.H. Structural Design for Fire Safety. West Sussex, England: John Wiley and

Sons, Limited. 2001.

DeHaan, J.D. Kirk’s Fire Investigation—3rd Edition. Englewood Cliffs, New Jersey: 

Brady—A Prentice Hall Division. 1991.

Drysdale, D. An Introduction to Fire Dynamics. New York City, New York: John Wiley and

Sons, Limited. 1985

Geiger, G.H. and D.R. Poirier. Transport Phenomena in Metallurgy. Reading, Massachusetts

Addison-Wesley Publishing Company. 1980. 

Hauffe, K. Oxidation of Metals. New York City, New York: Plenum Press. 1965

Holman, J.P. Heat Transfer—7th Edition. New York City, New York: McGraw-Hill, Inc. 1990.

Kubaschewski, O. and B.E. Hopkins. Oxidation of Metals and Alloys. New York City, New York:

Academic Press Inc. 1962. 

Larose, S. and R.A. Rapp. “Review of Low Temperature Oxidation of Carbon Steels and Low

Alloy Steels for Use as High-Level Radioactive Waste Package Materials.” CNWRA 97-003. 

San Antonio, Texas: Center for Nuclear Waste Regulatory Analyses. 1997. 

Rapp, R.A. High Temperature Corrosion (Reference Manual for Audio Course). 

Washington, DC: The American Chemical Society. 1980.

Runk R.B. and H.J. Kim. “The Oxidation of Iron-Carbon Alloys at Low Temperatures.” 

Oxidation of Metals. Vol. 2, No. 3. p. 285. 1970.

Simms, N.J. and J.A. Little. “High-Temperature Oxidation of Fe-21

/4Cr-1Mo in Oxygen.” 

Oxidation of Metals. Vol. 27, Nos. 5/6. p. 283. 1987.

Szlarska-Smialowska Z. and J. Jurek. “Ellipsometric Studies on Iron Oxide Film Growth at

100–350 °C.” Corrosion. Vol. 32, No. 7. p. 294. 1976.

Tewarson, A. Generation of Heat and Chemical Compounds in Fires—Society of Fire

Protection Engineers (SFPE) Handbook. Section 3, Chapter 4. pp. 3-99–100. Quincy,

Massachusetts: National Fire Protection Association. 1995.