NASA/TM-2000-209022

Experimental and Numerical 

Characterization of a Steady-State 

Cylindrical Blackbody Cavity at 

1100 Degrees Celsius

Thomas J. Horn

NASA Dryden Flight Research Center

Edwards, California

Amanie N. Abdelmessih

Saint Martin’s College

Lacey, Washington

August 2000
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NASA/TM-2000-209022

Experimental and Numerical 

Characterization of a Steady-State 

Cylindrical Blackbody Cavity at 

1100 Degrees Celsius

Thomas J. Horn

NASA Dryden Flight Research Center

Edwards, California

Amanie N. Abdelmessih

Saint Martin’s College

Lacey, Washington

August 2000

National Aeronautics and

Space Administration

Dryden Flight Research Center

Edwards, California 93523-0273
NOTICE

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Space Administration.

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1

Copyright  2000 by The American Society of Mechanical Engineers. No 

copyright is asserted in the United States under Title 17, U.S. Code. The U.S. 

Government has a royalty-free license to exercise all rights under the copyright 

claimed herein for Governmental purposes. All other rights are reserved by the 

copyright owner.

KEYWORDS: Blackbody calibration furnace, Heat flux 

calibration, Heat transfer, Numerical thermal analysis, 

Temperature measurement.

ABSTRACT

A blackbody calibration furnace at the NASA Dryden Flight 

Research Center is used to calibrate heat flux gages. These gages 

are for measuring the aerodynamic heat flux on hypersonic flight 

vehicle surfaces. The blackbody is a graphite tube with a 

midplane partition which divides the tube into two 

compartments (dual cavities). Electrical resistance heating is 

used to heat the graphite tube. This heating and the boundary 

conditions imposed on the graphite tube result in temperature 

gradients along the walls of the blackbody cavity. This paper 

describes measurements made during steady-state operation and 

development of finite-difference thermal models of the 

blackbody furnace at 1100 °C. Two configurations w

one with the blackbody outer surface insulated and 

without insulation. The dominant modes of heat tra

identified for each configuration and the effect of v

material properties and electric current that was pa

the blackbody were quantified.

NOMENCLATURE

A area, mm2

ATJ an isostatically molded grade of graphite w

grain and high strength

E emissive power, W/m2

F emissive power scaling factor

FLL Flight Loads Laboratory at NASA Dryden 

Research Center, Edwards, California

I electric current, amperes

kW kilowatt

L length, mm

MPa megapascal

MW megawatt

NASA National Aeronautics and Space Administ

NIST National Institute of Standards and Techn

EXPERIMENTAL AND NUMERICAL CHARACTERIZATION OF A STEADY-STAT

CYLINDRICAL BLACKBODY CAVITY AT 1100 DEGREES CELSIUS

Thomas J. Horn

Aerostructures Branch

NASA Dryden Flight Research Center

P.O. Box 273, MS 48202A

Edwards, California 93523

Tel (661) 276-2232, Fax (661) 276-2852, tom.horn@dfrc.nasa.gov

Amanie N. Abdelmessih, Ph. D.

Associate Professor

Mechanical Engineering Department

Saint Martin’s College

5300 Pacific Avenue SE

Lacey, Washington 98503-1297

Tel (360) 438-4532, Fax (360) 438-4548, abdelmessih@stmartin.edu
2

OFT optical fiber thermometer

P power, watts

PID proportional integral derivative

R electrical resistance, ohm

T absolute temperature, K

V volume, mm3

W watt

x distance from center partition, mm

temperature change

ρ electrical resistivity, ohm⋅mm

σ Stefan-Boltzman constant, 

Subscripts:

corr corrected

ind optical fiber thermometer indicated reading

ref reference pyrometer

tot total

INTRODUCTION

During hypersonic flight, high heat fluxes are generated on 

the surfaces of vehicles. It is important to understand what these 

heat fluxes are in order to better understand the aerodynamics of 

hypersonic flight. Heat flux gages can be used to measure these 

heat fluxes but they need to be properly calibrated before they 

are put into service. In addition, the furnace used to calibrate 

these gages must be characterized in order to determine its 

effect on the accuracy of the calibration.

The National Institute of Standards and Technology (NIST), 

Gaithersburg, Maryland, has made significant strides in recent 

years towards developing well-understood calibration systems 

which operate in convection, conduction, and radiation heat 

transfer modes (Holmberg, et al., 1997; Grosshandler and 

Blackburn, 1997; Murthy, et al., 1997; Holmberg, et al., 1999). 

The most powerful of the NIST facilities is a radiation facility 

capable of producing heat fluxes of 200 kW/m2

. This is still well 

below the heat fluxes that can be generated in hypersonic flight, 

which can be in excess of 1 MW/m2.

The Flight Loads Laboratory (FLL) at NASA’s Dryden Flight 

Research Center, Edwards, California is equipped with a 

calibration furnace capable of calibrating heat flux gages up to 

1.1 MW/m2 and temperature sensors of up to 2200 °C. This 

calibration system uses a cylindrical dual-cavity blackbody for 

calibrating a reference heat flux gage. It can also be fitted with a 

flat plate heating element to calibrate heat flux gages (having 

geometries that can not be inserted into the blackbody cavity) 

against the reference heat flux gage. The calibration furnace 

manufacturer (Thermogage, 1991) recommends that the 

blackbody temperature at the center of the middle partition be 

measured with a NIST traceable optical pyrometer, then the heat 

flux gage can be inserted into the blackbody cavity for a few 

seconds. The calibration of the heat flux gage is determined by 

dividing the emissive power of the blackbody by the peak output 

∆T

5.670 10–8× W⋅m–2⋅K–4

generated by the heat flux gage during the insertion

process may be repeated at other temperatures (and 

fluxes) as required.

There are two key questions which must be answ

to produce a quality heat flux gage calibration with 

uncertainty intervals. First, what effect do the vario

conditions (such as conduction, radiation, convecti

effect of axial temperature gradients) have on the b

cavity (Abdelmessih, 1998). Secondly, what effect 

presence of the reference heat flux gage in the blac

have on the calibration. The measurement uncertai

associated with these questions have not been spec

addressed in the literature and no detailed thermal 

been performed. Therefore, it is the purpose of this 

establish the basis of addressing these questions.

Research at the FLL aimed at quantifying the un

associated with this calibration system and process 

by performing experimental and numerical charact

both the blackbody cavity (which is discussed in th

the flat plate heater (Jiang, et al., 1998). Proper cha

of the system will require transient analyses of the 

heat flux gage insertion process.

This paper discusses the initial efforts to charact

blackbody cavity under steady-state operations at 1

These efforts include detailed experimental measur

two-dimensional, axisymmetric, numerical, therma

The temperature of 1100 °C was chosen because it 

temperature at which the heat flux calibration syste

manufacturer recommends switching from the unin

blackbody configuration used at lower temperature

insulated configuration used at higher temperatures

steady-state models will help quantify the relative im

the various boundary conditions present in the blac

system. They will also serve as an important step to

transient analysis defining both the relevant physic

phenomena and the initial conditions required at th

transient thermal analyses.

Note that use of trade names or names of manufa

this document does not constitute an official endor

such products or manufacturers, either expressed o

the National Aeronautics and Space Administration

EXPERIMENTAL OVERVIEW

Various experimental measurements were made 

blackbody calibration system characterization proc

section describes the calibration system hardware, 

instrumentation used to acquire the various measur

the procedures used during testing.

Test Setup

Figure 1 shows the calibration system. This syste

a power supply, a dual-cavity cylindrical blackbody
3

optical pyrometer for temperature control which views one side 

of the partition in the graphite tube. Other components not 

shown in the figure include a computer used for data acquisition 

and temperature control, a cooling system, and a supply of 

compressed argon gas.

The blackbody cavity (Fig. 2a) is a hollow cylinder measuring 

25.7 mm inside diameter. It is manufactured from ATJ graphite 

and is 28.9 cm long, with a 5.4-mm-thick partition at the 

middle. ATJ is an isostatically-molded grade of graphite of fine 

grain and high strength. At the top center of the graphite tube is 

an oval slot (Fig. 2b) that is 9.4 mm long by 3.1 mm wide and 

serves as a bleed port for argon purge gas during pretest purging 

of the blackbody cavities. The graphite tube is assembled with 

other components, shown in Fig. 3. This graphite tube is held in 

place at each end by a graphite bushing assembly installed in a 

copper ring (Fig. 4). A quartz tube, which serves as a 

containment barrier for argon purge gas, is held between the 

copper rings. The space between the outer surface of the 

blackbody and the quartz tube is insulated with graphite felt and 

foil when temperatures at 1100 °C or above are desired. This 

space is left uninsulated when the temperature does not exceed 

1100 °C. Unheated graphite extension tubes, 15.2 cm long, are 

installed at both ends. Laminar flow of argon gas exiting the 

extension tubes prevents air from diffusing into the heated 

graphite tube.

Figure 1. Blackbody Calibration Furnace.

Figure 2(a). Overview.

Figure 2(b). Partition Detail.

Figure 2. Graphite Blackbody Tube (dual c

Figure 3. Blackbody Assembly Componen

Blackbody axis

Center

partition

2.00 m

3.10 mm

2.00 mm

5.40 mm 

9.40 mm

CL
4

Figure 4. Copper Ring and Bushing Detail.

Figure 5 shows this entire assembly placed atop two 

electrodes that are connected to the power supply. Stable 

temperatures between 800 °C and 2200 °C can be attained at the 

center of the graphite tube by passing regulated electric current 

from the power supply through the blackbody assembly. The 

copper electrodes and rings which hold the blackbody assembly 

are water cooled. An aluminum water-cooled reflector (Fig. 5) 

surrounds the quartz cylinder. This reflector serves as the 

radiation boundary for the external surface of the graphite tube 

during tests run without graphite insulation in the quartz 

cylinder, thereby reducing heat lost to the ambient surroundings 

and ensuring more uniform temperatures in the blackbody 

cavity. The reflector is left in place during insulated runs as a 

safety device to prevent contact with potentially hot surfaces.

To minimize oxidation of the graphite, argon is used to purge 

the interior and exterior of the graphite tube. The argon is 

introduced through ports in the copper rings at each end of the 

tube. The argon is forced into the blackbody cavity before 

heating, as well as into the space between the graphite tube and 

Unheated

extension

000008

Copper ring

Bushing

Quartz tube

Graphite

insulation

Blackbody tube

Figure 5. Assembled Blackbody on Electrodes.

quartz tube, including the insulation when installed

argon flows out the unheated extensions during the 

prevent oxygen from entering the hot graphite tube

examination indicates that oxidation should not be 

long as sufficient purging is performed before and 

experimentation.

The furnace temperature is computer controlled, 

standard PID algorithm in a commercially availabl

acquisition and control software package (Laborato

Technologies Corporation, 1994). An optical pyrom

the center of the mid-partition from one end of the 

and provides temperature feedback to the control s

thermal control system is capable of maintaining th

temperature within ±0.5 °C of the desired steady-st

from 800 °C to 2200 °C.

Instrumentation

Various measurements were required to characte

blackbody cavity. These included the temperature a

of the partition in the middle of the blackbody (hen

referred to as blackbody temperature), partition em

electrical current passing through the blackbody as

blackbody surface temperature profiles, and locatio

temperature sensor.

Optical Pyrometers. Three optical pyrometer

were utilized during the experimental characterizat

blackbody cavity. The control pyrometer and refere

OFT

detector

unit

Crossbar

Power sup

OFT light

Referenc

Blackbody asse

Control pyromete

Slide to alig

 instruments

Slide to

insert OFT

Emissivity

measuring

pyrometer

OFT fiber

 optic cable

Deflection

 measurement

 potentiometer

Deflection

 potentiometer

 string

Figure 6. Top View of Pyrometer and OFT 
5

pyrometer measured blackbody temperature. The third 

pyrometer, which included an emissivity measurement system, 

was used to measure the partition emissivity at both room 

temperature and at the test temperature. 

The control and reference pyrometers were identical radiation 

pyrometers capable of measuring temperatures between 800 °C 

and 3100 °C. Their measurement area covers approximately 

one-quarter of the diameter of the blackbody partition. The 

response time of these pyrometers is 0.5 seconds for a step 

change from room temperature to test temperature. The control 

pyrometer was used by the control system to maintain stable 

temperatures and its calibration was accurate to within ±1 °C at 

1100 °C. The reference pyrometer was calibrated at NIST prior 

to testing. The NIST calibration included a 95-percent 

confidence uncertainty of 0.7 °C at 1100 °C. The temperature 

measurements of the center partition measured by the reference 

pyrometer will be referred to as the blackbody temperature

throughout this paper.

The third pyrometer included a system for measuring surface 

emissivity (Cameron, 1989) at both room temperature and test 

temperatures. The pyrometer measures emissivities from 0.1 to 

1.00 by determining how much energy, from a low-power laser, 

was reflected off the center partition. This pyrometer could 

complete an emissivity measurement approximately once per 

second. These emissivity measurements are accurate to within 

±0.01. The pyrometer operation is based on the assumption that 

the emissivity measurement is being taken on a diffuse 

reflecting surface at a temperature of less than 1500 °C. 

Optical Fiber Thermometer (OFT). Axial surface 

temperature profiles of the blackbody cavity were obtained 

using an OFT, (Luxtron, 1990) with a 90° bend at the tip 

(Fig. 7). The OFT system includes a single crystal sapphire 

lightpipe, a fiber optic cable, and a receiving unit consisting of 

an optical detector and signal conditioning. The lightpipe is the 

only piece of the system which enters the blackbody cavity. The 

light is transmitted from the lightpipe to the receiving unit 

through the fiber optic cable. The receiving unit senses the 

3-4 mm

≥ 4.1 mm

Lightpipe

To fiber and

detector

000011

Blackbody tube

52° view angle

Purge gas bleed port

Figure 7. OFT Insertion Geometry.

incoming infrared energy and converts that energy 

voltage proportional to temperature, which is route

acquisition system. 

The OFT system is capable of measuring temper

400 °C to 1900 °C with a resolution of 0.4 °C. Resp

the OFT system is 0.04 seconds. The OFT was mo

slide track and crossbar assembly (see Fig. 6). The 

(Fig. 8) was locally designed and built to include th

of rotating the sensor to measure the axial tempera

distribution at various locations around the interior 

circumference of the blackbody cavity. This mount 

of the sensor between 3 mm and 4 mm from the bla

(Fig. 7). The axial temperature profiles were obtain

the OFT at the desired angular location and then sl

inserting the OFT into the blackbody. The OFT wa

several axial locations as it was moved into the bla

order to obtain steady-state readings. The axial loc

sensing tip was measured using a deflection potent

connected to the crossbar.

During testing, we noticed that a significant amou

was entering the OFT through the bend in the senso

was obvious when the indicated temperature rose a

minimum reading of 400 °C by simply placing the 

the entrance and in line with the hot blackbody wit

tip aimed at ambient surroundings, such as the wal

The effect disappeared when the OFT sensor was m

Figure 8. OFT and Mount.
6

from the axis of the hot blackbody while still viewing the 

ambient surface.

A Monte Carlo ray trace analysis was performed on the OFT 

sensor to obtain a rough estimate of the energy traveling down 

the axis of the sensor, which entered through the bend instead of 

the sensing tip. The analysis included a model of the blackbody 

interior, which served as the emitting surface. This analysis 

demonstrated that the light entering the sensor at the bend was 

on the order of 10 percent of the total energy traveling down the 

axis of the OFT. Also, the analysis showed that this percentage 

remained essentially constant at all axial locations in the portion 

of the blackbody between the copper electrodes.

Some of the energy entering the bend was likely to be present, 

and thereby accounted for, during the manufacturers calibration 

process. Therefore, it was inappropriate to apply a 10-percent 

reduction to the total emissive power measured by the sensor. 

The total emissive power for a blackbody is represented by

(1)

A correction scheme was developed which utilized the known 

temperature (from the reference optical pyrometer) at the center 

partition to correct the OFT measurements. The OFT should 

indicate nearly the same temperature as the pyrometer when the 

OFT is located near the center partition. Assuming this to be the 

case, a correction factor can be determined from

(2)

The corrected temperature for any axial location in the hot zone 

of the blackbody can then be obtained from

(3)

The measurement uncertainty for the corrected temperature 

obtained when using equations (2) and (3) is taken to be ±5 °C. 

This uncertainty includes consideration of the uncertainties of 

the measurements used in the equations and the presence of 

temperature gradients within the blackbody, as demonstrated in 

the thermal analysis.

Etot σT4 =

F

Eref

Eind

---------- Tref ( )4

Tind [ ] ( ) x 4mm = 4 = = ----------------------------------------------

Tcorr( ) x F Tind [ ] ( ) x 4 { }

1 4⁄ =

OFT Position Measurement. The position of 

sensing tip was measured using a deflection measu

potentiometer. The range of the device was 0 to 10

an accuracy of ±1.6 mm. Measurement response tim

60 microseconds. The deflection measurement pote

was mounted on the top of the calibration system p

cabinet (Fig. 6). The spring-loaded string from the 

measurement potentiometer was connected to the c

assembly. The measurement system was calibrated 

0 mm was indicated when the OFT was in contact 

center partition.

Electric Current Measurement. The electrica

passing through the blackbody assembly was meas

current transformer with an input-to-output reducti

1200:5. The output from the current transformer w

to a current transducer, which converted the reduce

DC voltage. This DC voltage could then be acquire

acquisition system and converted into the value of 

blackbody assembly current. The measurement unc

the current transformer and transducer combination 

±13 amperes and the response time was 0.1 second

Measurement Uncertainty. Uncertainties for 

measurements are summarized in Table 1. The unc

have been divided into two groups. The first group 

measurement uncertainty inherent in the sensor, sh

top row of data. The second group is the uncertaint

acquisition systems measurement of the sensor outp

uncertainty is obtained by the square-root-of-the-su

squares method. All of these uncertainties are expr

terms of engineering units appropriate for the parti

measurement.

Test Procedure

The test process began by assembling and install

blackbody on the power supply electrode cradles. T

process included connecting hoses for cooling wate

purge gas to the ports on the copper rings. The only 

the assembly for the insulated and uninsulated tests 

installation of graphite felt and foil insulation betw

Table 1: Measurement Uncertainty Summary at 1100 °C.

Source Reference 

pyrometer, °C

Current, 

amperes

OFT, °C Axial location, mm Emissivity

Sensor 0.7 13.0 5.0 1.6 0.01

Data acquisition 0.1 0.7 0.2 0.3 N/A

Total 0.7 13.0 5.0 1.6 0.01
7

blackbody and quartz tube for insulated tests. The pyrometers 

and OFT were then installed and aligned with the blackbody.

Tests of the insulated and uninsulated blackbody followed the 

same process once assembly was complete. Each test began by 

establishing cooling water and purge gas flow to the blackbody. 

Establishing cooling water flow was a simple matter of 

configuring valves and turning on the closed-loop water system. 

Argon purge gas was supplied from a standard 17.2 MPa tank. 

The initial argon flow rate was greater than 2 m3/hr in order to 

blow air out of the system. Argon was forced into the central 

portion of the blackbody and between the blackbody and quartz 

tube by closing off the ends of the blackbody extensions. The 

extensions remained blocked for at least one minute. The total 

argon flow was then reduced to 0.28 m3/hr, which was sufficient 

to maintain a slow flow of argon out of various openings in the 

blackbody assembly and prevent air from re-entering the 

assembly.

Once the cooling water and purge gas flows were established, 

the data acquisition and control software was started and the 

blackbody furnace was placed in manual power control. The 

manual power setting was manipulated to maintain a 

temperature rise rate of 3 °C to 5 °C per second until the 

measured temperature approached the desired test temperature. 

The control software was switched into automatic PID control 

once the measured temperature was within 25 °C of the desired 

test temperature. The control software was then allowed to 

stabilize control at the desired test temperature. The set point 

temperature was stabilized to within ±0.5 °C of the desired test 

temperature in less than 3 minutes.

Once the blackbody temperature was stabilized according to 

the control pyrometer, the blackbody temperature was checked 

with the NIST-calibrated reference pyrometer. These 

temperatures were measured on opposite sides of the center 

partition and were within ±0.4 °C, well within the uncertainty of 

the NIST pyrometer calibration. The NIST-calibrated pyrometer 

was then moved away from the blackbody axis in order to 

provide a clear view for the emissivity measuring pyrometer, 

and the emissivity measurement was taken.

Once the emissivity measurement was recorded, the ninetydegree-bend OFT was positioned in front of the blackbody. The 

initial angular position of the OFT sensing tip was pointing up 

(0° rotation), as shown in Fig. 8. The OFT was then slowly 

pushed into the blackbody while its indicated temperature and 

tip location were recorded by the data acquisition system. 

During the uninsulated test, which occurred early in the test 

program, the OFT was inserted into the blackbody with steady 

motion. Approximately 30 seconds elapsed between the time of 

first motion and when the sensing tip reached maximum depth. 

The sensor was left at maximum depth for approximately 

6 seconds and then pulled out of the blackbody. 

The insertion process for the insulated case was slightly 

different, in that the sensor was stopped for 5 to 10 seconds at 

several axial locations. The locations were nomina

136 mm, 129 mm, 117 mm, 91 mm, 66 mm, 40.6 m

and 4.1 mm from the center partition. The slide tra

the OFT was mounted was pushed against a stop fo

insertion at 4.1 mm and this axial position was repe

the measurement uncertainty previously stated. Th

was manually positioned at marks representing the 

locations and these locations were repeated within 

Insertions were performed in a clockwise directi

around the circumference in both the uninsulated a

configurations. The sensor was rotated manually us

pointer and radial marks shown in Fig. 8 as guides

THERMAL ANALYSIS

Numerical thermal analyses of the insulated and 

blackbody configurations were developed to aid in t

understanding of the relevant heat transfer mechan

in the calibration system. This section discusses the 

used, the model geometry and grid, the boundary co

assumptions applied to the model, how heat genera

incorporated, and the material properties used in th

Software

Numerical models of the blackbody assembly w

developed, for both the insulated and uninsulated 

configurations, using commercially available therm

software (MacNeal-Schwendler Corporation, 1999

software includes a user interface for model genera

radiation view factor analysis program in addition 

solver. The thermal solver converts the model mesh 

thermal resistor-capacitor network and uses a weig

explicit-implicit finite difference scheme to genera

solution. Details of the numerical solution techniqu

found in Volume 1, Chapter 7 of the MacNeal-Schw

Corporation manual.

Geometry and Grid

The numerical model geometry was input as a tw

dimensional cross section, bounded by the centerlin

bottom) and including both the left and right caviti

blackbody assembly. Figure 9 shows the left side o

The graphite tube, end cap bushings, and extension

modeled. The aluminum reflector and copper rings 

modeled as constant temperature surfaces in the un

model. Options selected in the thermal analysis sof

instruct the solver to interpret the geometry as a thi

axisymmetric object and compute the radiation and 

heat transfer accordingly.

The mesh shown in Fig. 9 represents the initial m

both the insulated and uninsulated cases. The node 

the blackbody wall was nominally 2 mm axial by 1 

The spacing was reduced to 1 mm axial at the conn
8

bushing. The radial spacing was varied as required in the 

tapered region of the blackbody to maintain 4 nodes in the radial 

direction. Node spacing was 2 mm x 2 mm in the center 

partition. The bushings, where higher thermal gradients were 

expected, were meshed with a nominal 1 mm by 1 mm node 

spacing. A variable node spacing was used in the mesh of the 

end extensions. This mesh was approximately 1 mm by 1 mm 

near the bushings and expanded to approximately 9 mm x 9 mm 

as distance from the bushings increased.

Insulated Model. Graphite felt surrounds the graphite tube 

in the insulated configuration but was not included in the model 

geometry. The effect of the insulation was modeled as an 

adiabatic boundary condition on the exterior tube wall and 

reduced electric current passing through the graphite tube, as 

discussed below.

Uninsulated Model. The water-cooled copper ring at each 

end cap and the water-cooled aluminum reflector were 

represented as surfaces at fixed temperatures for the uninsulated 

model geometry and mesh. The copper caps and aluminum 

reflector were included as boundaries in the model to properly 

define the radiant heat transfer from the external surface of the 

uninsulated graphite tube.

Mesh Refinement. Grid refinement was performed in the 

blackbody node spacing in order to verify grid-independent 

results. The blackbody mesh was reduced to a nominal 1 mm 

axial by 1 mm radial node spacing throughout the blackbody. 

Reduction of the radial node spacing in the tapered region of the 

tube was maintained. Mesh refinement was not performed on 

the bushings and end extensions. Variations resulting from mesh 

refinement in the axial temperature profile were less than 2.5 °C 

for the uninsulated case and less than 1 °C for the insulated 

case. These variations are negligible (<0.25 percent) compared 

to the nominal 1100 °C blackbody temperature and are within 

experimental measurement uncertainty. Therefore, the nominal 

2 mm axial mesh was used to generate data for this paper in 

order to reduce computation time.

Blackbody tube

Extension

Bushing

Copper ring (radiating

 surface only)

000013

Aluminum reflector

(radiating surface only)

Axis of symmetry Center partition

Figure 9. Thermal Analysis Model, Left Side of 

Blackbody.

Modeling Assumptions

Certain assumptions had to be made in the therm

Simplifying assumptions were made concerning the

resistance, effect of the quartz tube, temperature of 

and copper rings, and internal convection of the bla

cavity.

Thermal Contact Resistance. Thermal conta

is a significant factor affecting the conduction heat 

the material interfaces in both end caps. These inte

include blackbody-to-bushing, bushing-to-copper r

between bushing components. However, it is extrem

to make high-quality estimates of contact resistanc

analysis, contact resistance has been neglected and 

bushing contact temperature was varied as required 

experimental temperature measurements at the end

graphite tube. It was assumed that this technique w

sufficiently account for the effect of contact resista

heat transfer through the bushing.

Effect of Quartz Tube. The quartz tube, being 

transparent to infrared radiation, was assumed to h

effect on the radiant heat transfer between the exter

of the graphite tube and the reflector and copper en

uninsulated configuration. Likewise, the quartz wa

to have no effect on the heat transfer from the grap

the insulated configuration. Therefore, the quartz tu

included in either the insulated or uninsulated mod

Reflector and Copper Ring Temperature. Th

reflector and copper rings were water cooled and p

external radiation boundary for the graphite tube in 

uninsulated configuration. The temperature distribu

components was modeled as isothermal and at the 

cooling-water temperature (40 °C).

Blackbody Cavity Internal Convection. Con

inside the blackbody cavity was assumed to be neg

on experimental observations. The argon flow-rate 

end cap was 0.14 m3/hr for all runs. Because of the 

resistance to flow, most of the argon exits through t

tubes. Therefore, forced convection is not expected 

significant in the blackbody cavity. In addition, exa

circumferential temperature variations showed insi

changes, verifying that natural convection was also 

insignificant.

Heat Generation

Heat was input into the insulated and uninsulated 

models by using volumetric heat generation. The th

computed the heat generation per unit volume base

given total current passing through the blackbody a

local geometry of the assembly, and the electrical r

The total electric current passing through the assem
9

measured during each test and the measured value was used in 

the model. The models included temperature-dependent 

electrical resistivity for the ATJ graphite tube and constant 

electrical resistivity for the pyrolytic graphite and carboncarbon components. The equation used in the model for power 

dissipated per unit volume was

(4)

However,

(5)

and

(6)

So,

(7)

Here ‘A’ represents the total area through which current flows 

in the actual hardware. The area and electrical resistivity were 

computed at each node where heat generation was defined.

There were two areas in each end cap where the electric 

current flowed from the copper ring into the bushing. The total 

current was divided between these two areas in proportion to the 

local contact area between the bushing and copper ring.

Insulated Model Adjustments. Some of the electric 

current passing through the blackbody assembly passed through 

the graphite felt and foil insulation instead of the graphite tube 

in the insulated configuration. Measurement of the actual 

currents passing through the graphite tube and insulation was 

not possible because the configuration of the system and the 

extreme temperatures involved did not allow it. The current 

passing through the graphite tube was reduced as required in the 

insulated model in order to obtain agreement with experimental 

results. Reasonable agreement, within 6.5 °C, was obtained 

between the numerical and experimental results with 

49.3 percent of the total measured current passing through the 

insulated portion of the blackbody.

Boundary Conditions

The thermal boundary conditions in the blackbody assembly 

were symmetric about the center partition for the steady-state 

analysis presented in this paper. However, the entire assembly 

was modeled in order to support future work, which would 

involve transient, asymmetric conditions. Modes of heat transfer 

common to both the insulated and uninsulated configurations 

include: (1) fixed isothermal temperatures where the bushings 

and extensions contact the copper rings (Fig. 10), (2) radiation 

P

V

--- I

2

R

V = --------

R ρL

A = ------

V LA =

P

V

--- ρI

2

A2 = -------

on the interior of the blackbody, bushings, and exte

(3) natural convection on the exterior of the extens

Unique boundary conditions existed on the extern

the graphite tube for uninsulated and insulated con

The model of the uninsulated configuration include

exchange between the graphite tube and the water-

copper rings and aluminum reflector. The adiabatic 

boundary condition on the external surface of the g

was used for the insulated model.

Bushing-Copper Ring Boundary Temperatu

isothermal temperature boundary condition was en

points where the bushings contacted the copper rin

end cap. These temperatures were adjusted as requ

the insulated and uninsulated models in order to att

agreement between the experimentally and numeri

determined tube surface temperatures. The lack of 

contact resistance in the thermal models resulted in 

boundary condition temperatures being set artificia

External Surface Radiation from the Graphi

(Uninsulated Blackbody). Radiation exchange wa

the uninsulated model between the exterior of the b

and the copper rings and aluminum reflector. The e

values for the copper rings and aluminum reflector 

adjusted, within literature-reported values, to obtai

results in good agreement with experimental result

copper and aluminum surfaces had significant surfa

and their emissivities were assumed to be the same

reflector emissivity of 0.26 produced numerical res

5.2 °C of the measured axial temperature. The emi

0.26 is within the range presented in Siegel and Ho

for oxidized aluminum.

Extension

(ATJ graphite)

C

c

Carboncarbon

Pyrolytic

graphite

Fixed temperature

Black

(ATJ 

Figure 10. Bushing Geometry Detail.
10

Adiabatic Wall (Insulated Blackbody). The adiabatic wall 

boundary condition was enforced on the exterior of the graphite 

tube in the insulated configuration model. An insignificant 

amount of heat was expected to flow radially through the 

insulation due to the thermal resistance of the insulation and 

heating of the insulation by a portion of the electric current 

passing through the blackbody assembly.

Material Properties

Properties, including thermal conductivity and electrical 

resistivity, for the various materials involved (Fig. 10) were 

obtained from manufacturers data whenever possible. Thermal 

conductivity and electrical resistivity for ATJ graphite were 

available from the manufacturer as a function of temperature 

throughout the range of interest (Fig. 11 and 12). The 

manufacturer of the carbon-carbon composite provided thermal 

conductivity data from room temperature to 120 °C and 

electrical resistivity data at room temperature (Table 2). The 

carbon-carbon thermal conductivity value at 120 °C was used in 

the analysis. Values for the thermal conductivity and electrical 

resistivity of pyrolytic graphite at room temperature were 

obtained from Clauser et al. (1963), and are summarized in 

Table 2. Temperature dependence of pyrolytic graphite thermal 

conductivity (Fig. 11) followed the same relative curve as ATJ 

graphite.

Figure 11. Thermal Conductivity for ATJ Graphite and 

Pyrolytic Graphite.

.40

.35

.30

.25

.20

.15

.10

.05

Thermal conductivity, W/mm • °C

1500 2000 2500 3000

Temperature, °C

0 500 1000

000015

Pyrolytic graphite

ATJ graphite

Figure 12. Electrical Resistivity of ATJ Gra

Table 2: Material Thermal Conductivit

and Electrical Resistivity.

The emissivity of the ATJ blackbody center parti

measured as part of the test and was used as the em

all graphite radiating surfaces. Published values (S

Howell, 1992) were used for initial estimates for th

reflector and copper ring emissivity (Table 3). The 

Material Thermal

conductivity

(W/mm ⋅°C)

Electr

resistiv

(Ω ⋅m

ATJ Graphite See Fig. 11 See Fig

Carbon-Carbon

(parallel to 

laminate at 120 °C)

0.0865 0.00

Carbon-Carbon

(perpendicular to 

laminate at 120 °C)

0.04152 0.02

Pyrolytic Graphite

(parallel to grain) See Fig. 11 0.00

.0110

.0105

.0100

.0095

.0090

.0085

.0080

.0075

.0070

.0065

.0060

Electrical resistivity, 

Ω • mm

1500 2000 

Temperature, °C

0 500 1000

Table 3: Material Emissivity.

ATJ Graphite (measured, 

used for all graphite 

components)

Emissivity: 0.88 – 0.91

Oxidized Aluminum 

(Siegel and Howell, 1992)

Emissivity: 0.20 – 0.31

Copper

(depending on level of

polish and oxidation)

(Siegel and Howell, 1992)

Emissivity: 0.15 – 0.78
11

aluminum emissivities are heavily dependent on the degree of 

oxidation, so they are used only as a guide. Measurement of the 

highly specular copper and aluminum emissivities was not 

possible because the available emissivity measuring equipment 

required diffuse reflection for proper operation.

RESULTS

This section discusses both the experimental and numerical 

analysis results. These results are discussed and compared for 

both the uninsulated and insulated blackbody configurations.

Experimental Results

Uninsulated Configuration. The corrected measured axial 

temperature profiles taken with the OFT at 0°, 90°, 180°, and 

270° rotation during the uninsulated test are shown in Fig. 13. 

The temperature profiles are nearly isothermal up to 20 mm 

from the partition and then drop off slightly to around 70 mm 

from the partition, which is the approximate location where the 

graphite tube begins to taper. There is another essentially 

isothermal region between 70 mm and 100 mm after which a 

sharp temperature decrease begins. The blackbody walls remain 

within 20 °C of the nominal blackbody temperature, 1100 °C, 

up to approximately 100 mm from the center partition. 

A slight temperature decrease of 3.5 °C was observed in the 

OFT measurements taken at 0° position at the 4.1 mm location. 

This decrease resulted from the presence of the purge gas bleed 

slot at the top center of the graphite tube. This observation was 

consistent with findings in Abdelmessih (1998). The end of this 

slot is 2 mm from the partition and just comes into the view 

angle of the OFT centered at approximately 4 mm from the 

partition (Fig. 7).

1200

1100

1000

900

800

700

600

Temperature, °C

80 100 120 140

Distance from partition face, mm

0 20 40 60

000017

Start of graphite

 tube taper

Blackbody end

 at 142 mm

Rotation

0°

90°

180°

270°

Figure 13. Uninsulated Blackbody Axial 

Temperature Distribution.

The blackbody was held at a constant temperatur

±0.1 °C as indicated by the control pyrometer. The 

indicated by the NIST-calibrated reference pyrome

viewed the end of the blackbody measured by the O

within ±0.3 °C of the temperature indicated by the 

pyrometer. The emissivity of the center partition m

during uninsulated tests varied from 0.88 to 0.90. A 

0.89 was used as the graphite emissivity in the therm

The measured electric current passing through the 

was 602 amperes.

Insulated Configuration. The corrected meas

temperature profiles taken with the OFT at 0°, 90°, 

270° rotation during the insulated test are shown in 

These profiles show the typical nonlinear temperat

distribution expected from heat transfer that is dom

heat generation and conduction. 

A decrease in the 0° position temperature at the 4

location was also observed in the insulated tests. T

was approximately 2.5 °C, 1 °C less than that obse

uninsulated test. This is appropriate since the OFT 

hot insulation wrapped around the blackbody throu

slot, but the OFT would view the cool aluminum re

through the slot in the uninsulated case. 

The insulated blackbody temperature was held c

within ±0.1 °C as measured by the control pyrome

NIST-calibrated reference pyrometer indicated a te

within 0.4 °C of the control temperature value. The 

emissivity of the partition during the 1100 °C insul

0.91. This value was used in the thermal analysis o

insulated blackbody. The measured total electric cu

1200

1100

1000

900

800

700

600

Temperature, °C

80 100 1

Distance from partition face, mm

0 20 40 60

Start of graphite

 tube taper

Blackbod

 at 142 m

Rotation

0°

90°

180°

270°

Figure 14. Insulated Blackbody Axial Tem

Distribution.
12

through the blackbody assembly was 334 amperes. This value 

includes current that passed through the insulation and the 

graphite tube.

Thermal Analysis Results

Uninsulated Model. The numerical results from the 

uninsulated blackbody model were brought into agreement with 

experimental results by adjusting the reflector emissivity and the 

bushing-copper ring boundary temperature. The computed 

temperatures are within ±3 °C at locations less than 90 mm from 

the center partition (Fig. 15). The maximum difference between 

computed and measured temperatures between 90 mm and 

130 mm from the partition is less than 7 °C. Temperature 

differences are less than 15 °C from 130 mm to the end of the 

graphite tube at 141.8 mm.

Figure 15. Uninsulated Blackbody Axial Temperature 

Distribution from Experimental Measurement and 

Numerical Model.

The computed temperature profile includes a hump between 

75 mm and 110 mm from the partition that is not found in the 

experimental data. Axial mesh density has been ruled out as a 

cause of this anomaly since doubling the axial mesh density had 

minimal effect on the magnitude of the increase. The external 

blackbody radiation boundary condition has been found to have 

a significant impact on the magnitude of the increase. A 

decrease of 0.02 in the reflector emissivity virtually eliminated 

the hump, reducing it to only 0.5 °C. However, decreasing the 

reflector emissivity increased the partition temperature 15 °C 

above the measured result. Small changes within manufacturing 

tolerances in the blackbody taper, particularly in the heat 

generation calculations, have also had significant effect on the 

increase. These small changes in the heating calculation have 

been shown to change the magnitude of the increase by at least 

5 °C. The factors affecting the anomalous temperature increase 

 Measurement

 Numerical model

1200

1100

1000

900

800

700

600

Temperature, °C

80 100 120 140

Distance from partition face, mm

0 20 40 60

000019

have been adjusted to minimize the deviation from t

temperature profile.

The thermal analysis provided insight regarding 

uniformity of the center partition temperature. The 

temperature varied less than 1 °C in the uninsulated 

(Fig. 16). 

Figure 16. Center Partition Temperature D

The thermal analysis data may also be used to es

effective emissivity of the center partition. For this 

effective emissivity is defined as the total energy em

and reflected by the center partition divided by the 

would be emitted if the partition were a perfect em

effective emissivity depends on the length of the tu

cylindrical wall included in the calculation. Figure 

variation in computed effective emissivity as a func

length of tube wall included in the calculation. The 

emissivity computed for the uninsulated model is w

1102

1101

1100

1099

1098

Temperature, °C

0 2 4 6 8 10 

Radial distance from center of partitio

Maximum

 internal rad

Insulated blackbody

Uninsulated blackbody

1.00

.99

.98

.97

.96

.95

.94

.93

.92

.91

.90

Effective partition emissivity

0 20 40 60 80 100 

Length of tube wall included in

calculation, mm (0 mm = partition fa

Uniform temperat

Uninsulated mode

Insulated model

Figure 17. Effective Emissivity Computati
13

0.1 percent of its final value of 0.983 when including 85 mm of 

the tube wall extending from the partition. This indicates that 

temperature variations beyond 85 mm will have little effect on 

the partition effective emissivity. Also note that the effective 

emissivity computed for the model is slightly higher than that 

calculated for a perfectly isothermal blackbody cavity at 

1100 °C. This results from the fact that the temperatures near 

the center partition, as computed in the model, were slightly 

higher than the ideal condition.

Insulated Model. The bushing-copper ring contact 

temperature and the portion of electrical current passing through 

the blackbody cylinder were adjusted to bring the insulated 

blackbody model into agreement with the experimental results. 

This adjustment process resulted in 49.3 percent of the total 

measured current actually passing through the blackbody and a 

bushing-copper ring contact boundary temperature of 645 °C. 

The remaining 50.7 percent of the current is assumed to travel 

through the graphite felt insulation which surrounds the 

blackbody. The model reproduced experimental measurements 

within ±5 °C (Fig. 18), except at the blackbody-bushing 

interface. The model produced temperatures 20 °C above the 

measured results in this region as a result of the complex heat 

transfer and electrical current flow at the contact boundary 

between the graphite tube and bushing.

The variations in center partition temperature were greater in 

the insulated model than they were in the uninsulated model. 

Figure 16 shows the insulated model partition temperature 

varied nearly 3° from the center to the outer edge. This is 

because of the lower temperatures in, and therefore lower 

energy arriving from, the blackbody cavity walls.

1200

1150

1100

1050

1000

950

900

850

800

750

700

650

600

Temperature, °C

140

000022

0 20 40 60 80 100 120

Distance from partition face, mm

 Measurement

 Numerical model

Figure 18. Insulated Blackbody Axial Temperature 

Distribution from Experimental Measurement and 

Thermal Model.

The cooler temperatures in the insulated case led 

somewhat lower effective partition emissivity as w

effective partition emissivity was 0.979 when using 

graphite tube in the calculation. The lower value is 

the cooler temperatures in the insulated blackbody 

the graphite tube is approached. The effective parti

emissivity was within 0.1 percent of its final value 

of the tube had been included in the calculation.

Sensitivity Analysis. An analysis has been pe

assess the sensitivity of the temperature profiles com

insulated and uninsulated models to changes in var

parameters used in the thermal model. Table 4 show

parameters studied, the variation in each parameter 

applied to each model, and the resulting maximum 

the axial temperature distribution for both the insul

uninsulated cases. The amount of variation for blac

emissivity and total electrical current was based on 

measurement uncertainty. Engineering judgement w

choose reasonable variations in other parameters.

The axial temperature distributions are most sen

total current passing through the blackbody. Both e

resistivity and reflector emissivity produced signifi

in the computed axial temperatures of the uninsula

blackbody. The maximums shown in Table 4 for th

parameters occurred at the center partition. Howev

temperature computed near the mouth of the blackb

only ±3.5 °C and less than ±1 °C for the variation o

resistivity and reflector emissivity, respectively.

The bushing-copper ring boundary temperature v

±25 °C caused changes in the computed temperatu

mouth of the blackbody of ±21 °C and ±25 °C for 

uninsulated and insulated cases, respectively. The p

temperature changed less than ±1 °C when this bou

condition was changed by ±25 °C.

The change in blackbody emissivity resulted in a 

small change in uninsulated blackbody temperature

maximum temperature change for the variation of t

parameter occurred near the center partition. 

The change in uninsulated blackbody thermal co

resulted in a nearly insignificant temperature chang

its maximum near the mouth of the blackbody. The 

caused by variation of this parameter was only abo

the center partition.

The model of the insulated configuration demon

significant sensitivity to other parameters, such as t

conductivity, electrical resistivity, and bushing-cop

boundary temperature, which influence the heat ge

conduction within the model. The changes induced 

parameters were of the same order of magnitude th

blackbody cavity.
14

The insulated model was nearly insensitive to small changes 

in blackbody emissivity. A blackbody emissivity change of 0.02 

(about 2 percent of the nominal value), caused <0.2 °C change 

in the computed axial temperatures.

CONCLUDING REMARKS

A cylindrical, dual-cavity blackbody has been characterized 

through experimental measurements and numerical models for 

steady-state operation at 1100 °C. Two hardware configurations 

were explored. The first configuration, recommended for 

operation at temperatures below 1100 °C, allowed the exterior 

of the graphite tube to radiate to a water-cooled aluminum 

reflector. The second configuration, employed for calibration at 

temperatures above 1100 ° C, accounted for graphite felt 

insulation wrapped around the graphite tube.

Measurement of the axial temperature profile in the 

uninsulated configuration revealed a nearly isothermal section 

extending 20 mm from the center partition. Tube wall 

temperatures remained within 20 °C of the center partition 

temperature up to 100 mm from the partition. The axial 

temperature distribution in the insulated configuration was 

much less isothermal. 

Measurements of the axial temperature profile, at preset axial 

locations, were taken every 45° around the inner circumference 

of the blackbody cavity. These temperature measurements 

confirm that internal natural and forced convection are 

essentially insignificant in both the insulated and uninsulated 

models.

The numerical models of the steady-state operation of the 

blackbody, in both the insulated and uninsulated configurations, 

have been tuned to experimental results. The models for both 

configurations capture the relevant heat transfer mechanisms 

and produce computed axial temperature profiles within 8 °C of 

the measured results.

The heat transfer mechanisms with significant im

uninsulated model are (1) radiation exchange affec

external surface of the graphite tube, and (2) heat g

the graphite tube. Variations of the properties and b

conditions affecting conduction and radiation in th

the blackbody had little effect on the blackbody tem

Variations in the properties and boundary conditi

to heat generation and conduction had the largest im

calculation of blackbody temperature for the insula

The effect of radiation in the insulated model was r

small.

The computed axial temperature distributions, as 

computation of temperature gradients in the center 

enabled the calculation of the effective partition em

effective partition emissivity in both the insulated a

uninsulated configurations was approximately 0.98

is nearly identical to the value computed for a perfe

isothermal blackbody cavity.

The steady-state experiments and thermal model

in this paper provide a sound foundation for additio

of the blackbody calibration system. It is recomme

work presented in this paper be extended to the ful

operational temperatures from 800 °C to 2200 °C. A

of the transient insertion of the reference heat flux 

blackbody cavity, if explored, could lead to quantifi

uncertainties involved in that process.

ACKNOWLEDGMENT

A portion of this work was performed as part of 

ASEE Summer Faculty Fellowship Program. The e

were performed in the Flight Loads Laboratory, at 

Dryden Flight Research Center, Edwards, Californ

Table 4: Sensitivity of Blackbody Temperature to Changes in Various Parameters.

Property or Boundary Condition Variation

Maximum Change

Uninsulated Insulated

Location , °C Location 

Blackbody thermal conductivity ±5.0 percent mouth ±1.0 partition ±

Electrical resistivity ±5.0 percent partition ±19.0 partition ±

Blackbody emissivity ±2.2 percent partition ±4.5 partition 

Reflector emissivity ±8.5 percent partition ±15.0

Total current ±26 amperes partition ±32.0 partition ±

Bushing-copper ring boundary temperature ±25.0 °C mouth ±21.0 mouth ±

∆T ∆
15

REFERENCES

Abdelmessih, A. N., 1998, “Experimental Measurements of 

Temperature and Heat Flux in a High Temperature Black Body 

Cavity,” NASA Technical Reports, NGT 2-52212, Document ID 

19990021026, PP A:1-3.

Cameron, E. S., 1989, “Laser Assisted Remote Temperature 

Measurement,” Optomechanical Design of Laser Transmitters 

and Receivers, SPIE Vol. 1044, pp. 155 – 162.

Clauser, H. R., Fabian, R., Peckner, D., and Riley, M. W., 1963, 

The Encyclopedia of Engineering Materials and Processes, 

Reinhold Publishing Corp., New York, New York.

Grosshandler, W. L,. and Blackburn, D., 1997, “Development of 

a High Flux Conduction Calibration Apparatus,” HTD-Vol. 353, 

Proceedings of the ASME Heat Transfer Division, V. 3, 

pp. 153–158.

Holmberg, D., Steckler, K,. Womeldorf, C,. and Grosshandler, 

W., 1997, “Facility for Calibrating Heat Flux Sensors in a 

Convective Environment,” HTD-Vol. 353, Proceedings of the 

ASME Heat Transfer Division, V. 3, pp. 165 – 171.

Holmberg, D. G., Womeldorf, C. A., and Grosshandler, W. L., 

1999, “Design and Uncertainty Analysis of a SecondGeneration Convective Heat Flux Calibration Facility,” HTDVol. 364, Proceedings of the ASME Heat Transfer Division, 

V. 4, pp. 65 – 70.

Jiang, S., Horn, T. J., and Dhir, V. K., 1998, “Nume

Analysis of a Radiant Heat Flux Calibration System

361, Proceedings of the ASME Heat Transfer Divis

pp. 609 – 616.

Laboratory Technologies Corporation, 1994, LABT

Acquisition and Process Control Software, Window

Guide, Wilmington, Massachusetts.

Luxtron Inc., 1990, Accufiber Models 10 &100 Op

Thermometers Manual, Santa Clara, California.

MacNeal-Schwendler Corporation, 1999, MSC/PA

THERMAL Users Guide, Volumes 1 and 2, Softwar

Online Documentation.

Murthy, A. V., Tsai, B. K., and Saunders, R. D., 19

“Radiative Calibration of Heat Flux Sensors at NIS

Overview,” HTD-Vol. 353, Proceedings of the ASM

Transfer Division, V. 3, pp. 159 - 164.

Siegel, R., and. Howell, J. R., 1992, Thermal Radia

Transfer, 3rd edition, Hemisphere Publishing Corp

Washington, D.C.

Thermogage Inc., 1991, Thermogage Operation M

Model 48 kW Calibration Furnace, Frostburg, Mar
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NSN 7540-01-280-5500 Standard Form 298 (Rev. 2-8

Prescribed by ANSI Std. Z39-18

298-102

Experimental and Numerical Characterization of a Steady-State

Cylindrical Blackbody Cavity at 1100 Degrees Celsius

WU 522-32-24-E8-RS-0

Thomas J. Horn and Amanie N. Abdelmessih

NASA Dryden Flight Research Center

P.O. Box 273

Edwards, California 93523-0273

H-2403

National Aeronautics and Space Administration

Washington, DC 20546-0001 NASA/TM-2000-2090

A blackbody calibration furnace at the NASA Dryden Flight Research Center is used to calibrate hea

gages. These gages are for measuring the aerodynamic heat flux on hypersonic flight vehicle surfaces

blackbody is a graphite tube with a midplane partition which divides the tube into two compartments 

cavities). Electrical resistance heating is used to heat the graphite tube. This heating and the bou

conditions imposed on the graphite tube result in temperature gradients along the walls of the black

cavity. This paper describes measurements made during steady-state operation and development of 

difference thermal models of the blockbody furnace at 1100 °C. Two configurations were studied, one wi

blackbody outer surface insulated and the other without insulation. The dominant modes of heat transfer 

identified for each configuration and the effect of variations in material properties and electric current tha

passed through the blackbody were quantified.

Blackbody calibration furnace, Heat flux calibration, Heat transfer, Numerical

thermal analysis, Temperature measurement A03

21

Unclassified Unclassified Unclassified Unlimited

August 2000 Technical Memorandum

Presented at 34th National Heat Transfer Conference, Pittsburgh, Pennsylvania, August 20–22, 2000, paper NHTC

12140. Dr. Abdelmessih is an Associate Professor, Mechanical Engineering Department at St Martin’s College, L

Washington and a participant in the NASA/ASEE Summer Faculty Fellowship Program.

Unclassified—Unlimited

Subject Category 34

This report is available at http://www.dfrc.nasa.gov/DTRS/