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Defense Technical Information Center 
Compilation Part Notice 

ADPO 10740 

TITLE: Measurement Requirements for Improved 
Modeling of Arcjet Facility Flows 

DISTRIBUTION: Approved for public release, distribution unlimited 

This paper is part of the following report: 

TITLE: Measurement Techniques for High Enthalpy 
and Plasma Flows [Techniques de mesure pour les 
ecoulements de plasma et les ecoulements a haute 

To order the complete compilation report, use: ADA390586 

The component part is provided here to allow users access to individually authored sections 
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Measurement Requirements for Improved Modeling of Arcjet 

Facility Flows 

Presented by 
Douglas G. Fletcher 
Reacting Flow Environments Branch 
NASA Ames Research Center 
Moffett Field, CA, 94035-1000 

1. Introduction 3A-2 

1.1 Historical Development of NASA Ames Arcjet Facilities 3A-2 

1.2 Overview of Current NASA Arcjet Facilities 3A-4 

1.3 Ames Aerodynamic Heating Facility Arcjet 3A-4 

1.4 Arcjet Characterization Using Conventional Instrumentation 3A-5 

2. CFD for Arcjet Flows 3A-8 

2.1 Motivation for Arcjet Flow Modeling 3A-8 

2.2 CFD Requirements for Arcjet Flow Simulations 3A-8 

2.3 Strategies for Arcjet Flow Simulations 3A-9 

3. Measurement Requirements for Arcjet Flow Modeling 3 A- 10 

3.1 Enthalpy 3A-11 

3.2 Arc Heater 3 A- 11 

3.3 Arcjet Nozzle and Free-Stream Flow 3A-12 

3.4 Blunt-Body Shock-Layer Flow 3A-12 

3.5 Measurement Accuracy Requirements 3A-13 

4. Experimental and Computational Investigation of Shock-Layer Flows 3 A- 14 

4.1 Objectives of Investigation 3A-14 

4.2 Experimental and Computational Approach 3 A- 14 

4.3 High Pressure Case 3A-16 

4.4 Low Pressure Case 3A-20 

4.5 Spatially Resolved Measurements 3A-23 

4.6 Lessons Learned 3A-23 

5. Summary and Recommendations 3A-24 

6. Acknowledgements 3A-25 

7. References 3A-25 

Paper presented at the RTO AVT Course on “ Measurement Techniques for High Enthalpy and Plasma Flows”, 
held in Rhode-Saint-Genese, Belgium, 25-29 October 1999 , and published in RTO EN-8. 


1. Introduction 

Current efforts to develop new reusable launch vehi- 
cles and to pursue low-cost robotic planetary missions 
have led to a renewed interest in understanding arc- 
jet flows. Part of this renewed interest is concerned 
with improving the understanding of arcjet test results 
and the potential use of available computational-fluid- 
dynamic (CFD) codes to aid in this effort. These CFD 
codes have been extensively developed and tested for 
application to nonequilibrium, hypersonic flow model- 
ing. It is envisioned, perhaps naively, that the appli- 
cation of these CFD codes to the simulation of arcjet 
flows would serve two purposes: first, the codes would 
help to characterize the nonequilibrium nature of the 
arcjet flows; and second, arcjet experiments could po- 
tentially be used to validate the flow models. These 
two objectives are, to some extent, mutually exclusive. 
However, the purpose of the present discussion is to 
address what role CFD codes can play in the current 
arcjet flow characterization effort, and whether or not 
the simulation of arcjet facility tests can be used to 
evaluate some of the modeling that is used to formu- 
late these codes. 

This presentation is organized into several sections. 
In the introductory section, the development of large- 
scale, constricted-arc test facilties within NASA is re- 
viewed, and the current state of flow diagnostics using 
conventional instrumentation is summarized. The mo- 
tivation for using CFD to simulate arcjet flows is ad- 
dressed in the next section, and the basic requirements 
for CFD models that would be used for these simula- 
tions are briefly discussed. This section is followed by 
a more detailed description of experimental measure- 
ments that are needed to initiate credible simulations 
and to evaluate their fidelity in the different flow re- 
gions of an arcjet facility. Observations from a recent 
combined computational and experimental investiga- 
tion of shock-layer flows in a large-scale arcjet facility 
are then used to illustrate the current state of develop- 
ment of diagnostic instrumentation, CFD simulations, 
and general knowledge in the field of arcjet character- 
ization. Finally, the main points are summarized and 
recommendations for future efforts are given, 

1.1 Development of NASA Ames Arcjet Facili- 

Development efforts that led to what we now classify 
as arcjet test facilities began in the late 1950’s with the 
goal of producing a continuously operable hypersonic 
ground test facility. This need was driven by both 
US Department of Defense and NASA mission plan- 
ning requirements. From the NASA side, planetary 
missions and the manned space program were push- 
ing aerospace vehicles to higher aerothermodynamic 
heating rates. Several excellent texts have been writ- 
ten that include a much broader treatment of the his- 

torical development of arcjet facilities and plasma arc 
devices for propulsion. 1,2 However, for the purpose of 
introducing the current topic, a brief recapitulation of 
arcjet facility development activities at NASA Ames 
Research Center is given below. 

Fig. 1. NASA Ames concentric ring arcjet. 

The rather ambitious target capabilities for develop- 
ing the first Ames arcjet were: 1) 32 MJ/kg enthalpy; 
2) 100 atm pressure; 3) 1 MW input power; and 4) 
continuous and contaminant-free operation. The first 
successful arcjet that even partially met some of these 
goals was the Ames Concentric Ring Arcjet, 3 which 
is depicted in Fig. 1. While the device could operate 
at the intended high pressures, it had a very low effi- 
ciency in terms of coupling the electrical energy to the 
flow. As can be seen in Fig. 1, the arc region is quite 
small, and most of the incoming air stream bypasses 
the arc. This resulted in relatively low deposition of 
energy into the test gas stream. 

Fig. 2. Early Ames constricted- arc heater. 

To improve the coupling of electrical energy into the 
flow, the next round of heater configurations featured 
more widely spaced electrodes separated by an orifice 
plate that is intended to constrict the arc to a rela- 
tively small region. It was hoped that forcing the flow 
and arc through the same small region would improve 
the electrical energy deposition and raise the stream 
enthalpy. Figure 2 shows an example of this device, 
which did show an improvement in energy deposition. 
However, it proved to be nearly impossible to prevent 
the arc from attaching at the edge of the orifice plate, 


and excessive arc- induced failures produced further de- 
sign modifications. 

Subsequent efforts resulted in the development of su- 
personic arcjets, 4 which achieved high enthalpies and 
low heat loss by extending the arc through the throat 
region before attachment downstream in the low pres- 
sure, expanded flow region. A schematic of one of the 
earlier versions is shown in Fig. 3. Erosion of the down- 
stream attachment point was minimal for this type 
of arcjet because of the diffuse nature of the arc at 
the low pressures of the supersonic flow region. The 
constrictor diameter was only 6.4 mm, but the heater 
performance was pretty much as predicted, and there 
appeared to be a substantial gain in electrical energy 
deposition. Shortly thereafter a second supersonic arc- 
jet was developed with a 25.4 mm diameter constric- 
tor and this device delivered enthalpies on the order of 
900 MJ/kg on the flow centerline. 5 Unfortunately, the 
stream was highly nonuniform and the excessive radial 
gradients limited the application range of this heater. 

Fig. 3. NASA Ames supersonic, constricted- arc facil- 

In one of the more interesting developments that has 
particular relevance to the current discussion, the 
ARCFLO code was developed in 1967 to model pro- 
posed arc heater configurations. 6 For the instrumental 
technology available at the time, comparisons between 
arcjet performance measurements and ARCFLO pre- 
dictions were satisfactory. This led to the use of the 
code in the development of new heater configurations. 

Although impressively high enthalpy levels were gen- 
erated in some of these early devices, there was no 
great demand for routine operation at those condi- 
tions. Instead, the emergence of the shuttle as the 
primary launch and payload capability for NASA gen- 
erated a significant demand for test capability in the 
20 to 30 M J /kg range to develop and qualify shuttle- 
related thermal protection materials. With the ex- 
ception of meteor ablation studies and work involved 
with the development of heat shields for planetary- 
entry missions, 7 this test condition range has proven 
satisfactory for a majority of the aerospace commu- 
nity’s needs. Progress in providing robust test facil- 
ities in the required performance range was enabled 
by timely improvements in magnetically driven elec- 
trode technology. 8 All of the successful heater designs 
relied on magnetic fields to spin the arc attachment 

point around the electrode to reduce the local heat- 
ing. However, an optimal combination of geometry, 
current load, and magnetic field strength leading to ex- 
tended electrode lifetime could only be found through 
trial and error, since theoretical models of the com- 
bined fluid and plasma dynamics of the electrode were 
inadequate at that time. 

Using the new magnetically driven electrodes, the 
Ames 20 MW Constricted Arc Jet was built in 1972. 
A schematic rendering of the constrictor, downstream 
electrode package and nozzle configuration is shown in 
Fig. 4. This basic constricted-arc heater configuration 
has been used continuously, with relatively little vari- 
ation, in the Ames arcjet facilities since that time. An 
excellent description of the electrode and constrictor 
design and performance evaluation is given as part of 
the report on the Ames 60 MW arcjet, 9 which is still 
in use today. 

Fig. 4. Current version of Ames constricted- arc heater. 

In the late 60’s and early 70’s arcjets were in use at 
aerospace companies and research centers around the 
world. It appeared that arcjets would find extensive 
use as aerot her mo dynamic test facilities where funda- 
mental investigations of real gas phenomena could be 
conducted. Although they did not provide perfect sim- 
ulation of atmospheric flight environments, arcjets had 
a significant advantage over impulse facilities in that 
they could be operated at high enthalpy levels for long 
periods of time. Unfortunately, it proved to be very 
difficult to establish just what operating enthalpy level 
was actually reached. In fact, the inability to charac- 
terize the arcjet stream conditions ultimately limited 
arcjets mainly to applications where complete knowl- 
edge of stream conditions was not a requirement for 
evaluating test results. 

In a review of ground- test facility simulations of 
poorly understood real-gas phenomena, Park 10 iden- 
tified seven important problems: 1) determining aero- 
dynamic parameters; 2) viscous/shock interactions; 3) 
boundary layer transition; 4) understanding leeward 
or base region flows; 5) nonequilibrium radiation; 6) 
nonequilibrium ionization; and 7) surface catalysis. 
Park then examined the capabilities of three types of 
hypersonic ground-test facilities that could be used for 
fundamental investigations of these problems: 1) im- 
pulse facilities (including shock tunnels); 2) ballistic 
ranges; and 3) arcjets. Arcjets were only deemed suit- 


able for studies of nonequilibrium radiation and sur- 
face catalysis, and even then adequate specification of 
the stream conditions was mentioned as a requirement 
for improving the analysis of test results. 10 

Before discussing the issue of stream conditions fur- 
ther, it is useful to examine the current status of large- 
scale arcjet facilities and their role in thermal protec- 
tion material test and development. In addition, it is 
instructive to examine the use of conventional stream 
characterization instrumentation and how it is used in 
the interpretation of test measurements. 

1.2 Overview of Current NASA Arcjet Facili- 

Today, NASA’s large-scale arcjet facilities are used 
mainly to simulate aero thermal heating environments, 
although there is still some limited use in evaluating 
supersonic air-breathing propulsion concepts. Our dis- 
cussions will focus exclusively on facilities, modeling, 
and measurements that relate to the principal appli- 
cation: aerothermal heating simulation. Two NASA 
Centers, Johnson and Ames, are currently operating 
segmented- type constricted-arc heater facilities for this 
application. This facility is the workhorse for the 20 
to 30 MJ/kg enthalpy range of long-duration thermal 

Fig. 5. Current version of JSC TP-1 constricted-arc 
heater and nozzle. 

Arcjet facilities at Johnson Space Center support TPS 
testing requirements for manned missions. All of 
the thermal protection materials for shuttle, includ- 
ing tiles, coatings, and fillers, are qualified for use on 
the basis of tests in these arcjets. Johnson has two 10 
MW facilities JSC TP-1, which became operational in 
1973, and JSC TP-2, which was upgraded to 10 MW in 
1991. 11 Both facilities have segmented, constricted-arc 
heaters. The TP-1 facility is usually arranged with a 
conical nozzle configuration for st agnation- point test- 
ing, while TP- 2 is typically configured with a rectan- 
gular channel for flat-plate testing. A schematic of 
the Johnson TP-1 heater and nozzle is presented in 
Fig. 5, and it shows two noteworthy features. First, 
because of the tungsten cathode, Oo is injected sep- 
arately from N 2 further downstream in the heater to 
prolong the useful life of the electrode. Although they 
are injected separately, the two gases are thought to 
be mixed by the time the downstream electrode pack- 
age is reached. The second interesting feature is that 

the throat diameter is larger than that of the constric- 
tor, which causes some uncertainty regarding the sonic 
location. Axial velocities in the arc column could actu- 
ally be quite high, which may inhibit mixing of the Oo 
and No streams. The facility is equipped with energy 
balance instrumentation, which provides a measure of 
the bulk enthalpy for each test. 

The Arcjet Complex at Ames Research Center sup- 
ports Ames’ role as lead NASA Center for thermal 
protection material development. There are currently 
three operating segmented, cons trie ted- arc facilities: 
the Aerodynamic Heating Facility (AHF) and the 
Panel Test Facility (PTF) are both rated at 20 MW ; 
and the Interactive Heating Facility (IHF) is rated at 
60 MW. There are also two operable Hue Is- type heater 
facilities, 2x9 Turbulent Flow Facility (TFF) and the 
Direct Connect Arcjet Facility (DCAF). Two arcjet fa- 
cility buildings house the different arcjets, which share 
common steam-ejector vacuum and water-cooling sys- 
tems. With a shared vacuum system, only one facility 
can operate at a time. However, facilities can operate 
sequentially throughout the day with up to 8 runs dur- 
ing a single operating shift. Note that the operating 
frequency for an arcjet is greater than that of typical 
large-scale impulse facilities. 

A cross section of a typical Ames constricted- arc 
heater configuration was shown above in Fig. 4. The 
configuration is different from the JSC TP- 1 configura- 
tion that was shown in Fig. 5. For the Ames heater, the 
throat diameter is smaller than the constrictor diame- 
ter, so the sonic point will always be located between 
the converging and diverging sections of the nozzle. 
Also, both the upstream and downstream electrodes 
are copper, so oxygen does not need to be injected 
separately for air tests. Since the overwhelming ma- 
jority of arcjet tests at Ames Research Center are per- 
formed using segmented- type, constricted- arc heaters, 
Huels-type heaters will not be discussed further. 

Even though they are both classified as segmented, 
constricted- arc heaters, the different designs of the 
JSC and Ames heaters illustrate the variety of elec- 
trodes and nozzles that are in use today. There is 
no standard design. Consequently, performance will 
vary widely from facility to facility and characteriza- 
tion of the performance of one facility is by no means 
applicable to others unless the configuration is exaetty 

1.3 Ames Aerodynamic Heating Facility Arcjet 

The Aerodynamic Heating Facility (AHF) Arcjet at 
NASA Ames Research Center is an example of cur- 
rent large-scale, constricted-arc heater test facilities. 
A schematic of the facility is shown in Fig. 6. Facil- 
ity operation is initiated by evacuating the arcjet and 
then striking an arc in a low-pressure argon stream. 12 


The test gas flow, usually air or nitrogen, is then in- 
troduced through the segmented disks along the col- 
umn, and the arc current is adjusted to achieve the 
test conditions. Within the arc column, heating by 
the electrical discharge causes substantial dissociation 
and ionization of the test gas. The argon start-gas 
stream is maintained during operation, and additional 
argon is injected to protect the downstream electrode. 
Each electrode package is made up of a series of al- 
ternating copper rings and spacer disks. The rounded 
rings are the actual electrodes, and they protrude into 
the stream to move the arc attachment away from the 
wall (see Fig. 4). Magnetic windings inside the elec- 
trodes rotate the arc attachment point to reduce the 
heat load on the electrodes. Each electrode can carry 
up to 500 A of current. Typically, the anode is placed 
at the upstream end of the arc column to benefit from 
further cooling by the test gas. 

Test Chamber 

Fig. 6 Schematic of the NASA Ames AHF Arcjet. 

Upon leaving the heater, the flow is accelerated to hy- 
personic speed through a conical, converging-diverging 
nozzle. During the expansion-driven acceleration, the 
collision frequency decreases rapidly in the nozzle and 
the thermochemical state of the flow departs from 
equilibrium. At some point, the flow chemistry be- 
comes frozen, and this may be followed by freezing 
of the internal energy distribution of the molecular 
species. Various nozzle sections can be used to pro- 
vide expansion ratios ranging from 64 to 576. The 
flow exits the nozzle and continues expanding into a 
cabin where material tests are conducted. Material 
samples are typically inserted into the stream 36 cm 
downstream of the nozzle exit. Test durations of up to 
20 minutes are possible, depending on the particular 
conditions. During the tests, the stagnation pressure, 
cabin pressure, and arc heater conditions are continu- 
ously monitored. 

In a typical test cycle, a preliminary analysis of the ex- 
pected heat load in a flight application has been per- 
formed and a candidate thermal protection material 
has been selected for testing in an arcjet flow. The 
test conditions are chosen to attempt to match the ex- 
pected heat flux for a particular point on a predicted 
trajectory, such as the peak heating point. Conven- 
tional instruments, which will be discussed below, are 
used to verify the test conditions. The test results 

are then interpreted without the benefit of full knowl- 
edge of the stream conditions. Currently, relating test 
results from the arcjets to the intended flight applica- 
tion is more of an art than a science, because the arcjet 
stream conditions are not sufficiently characterized. 

1.4 Arcjet Characterization Using Conven- 
tional Instrumentation 

The words “Arcjet Characterization” are typically un- 
derstood to mean specifying the state of the arcjet test 
stream, and they are referred to throughout this dis- 
cussion in that context. Although it is important in 
flow modeling, the need for arcjet characterization is 
driven primarily by the needs of thermal protection 
material developers, who need better specification of 
the stream conditions to relate the results to flight en- 
vironments. In addition, an improved understanding 
of arcjet stream conditions in general may also make 
arcjets more suitable for fundamental studies of real 
gas phenomena. 

The state of arcjet stream characterization in the early 
90’s was summarized in an excellent and thoughtful 
review article by Scott. 13 Both established and novel 
instrumental techniques were critically reviewed in the 
article. The article focused mainly on how various di- 
agnostic techniques could be used to characterize the 
most important stream variables: enthalpy and the de- 
gree of nonequilibrium in the stream. Rather than re- 
peat this review, some of the more widely used conven- 
tional diagnostics are reviewed briefly below. The lim- 
itations of these measurement techniques are discussed 
to provide background for considering what measure- 
ments are required to improve arcjet flow modeling. 
Newer, less widely used spectroscopic techniques, such 
as multiphoton spectroscopy will be mentioned later, 
and are discussed more fully in the second article. 

Traditional instruments that are used to obtain flow 
property measurements include pitot probes and 
calorimeters. Additional instruments, such as thermo- 
couples and flow meters are used to measure coolant 
flow rates and temperature rise to perform an energy 
balance on the facility. Stream surveys are usually per- 
formed with a traversing, sting- mounted probe, since 
the facility can operate continuously and at a level 
where the instrument can give an equilibrated response 
to the quantity being measured. 

Pitot measurements yield the stagnation pressure be- 
hind a shock wave that is generated by the probe. 
For much of the operating range of today’s large-scale 
arcjet facilities, the pitot, or impact, pressure can be 
related to the dynamic pressure of the flow, pv 2 / 2, 
through the Rayleigh supersonic pitot relation 14 , 

M » L 



In the above expression, 7 is the ratio of specific heats 
for the gas, M is the Mach number, p P is the pitot 
pressure, p is the stream density, and v is the veloc- 
ity. Although p and v are both important stream vari- 
ables for arcjet flow characterization, a determination 
of each variable cannot be made without an additional 
measurement. For typical facility operating conditions 
the flow velocity is a considerably larger quantity than 
the stream density, so a strategy for determining both 
variables should involve a velocity measurement. 

Energy Balance - Most arcjet facilities are equipped 
with instruments that can be used to perform an en- 
ergy balance on the arcjet facility as a whole. Owing 
to its simplicity, the energy balance approach remains 
by far the most commonly used for characterizing the 
arcjet stream. The basic principle of the measurement 
is illustrated in Fig. 7, which shows the arc-jet opera- 
tion measurements that must be acquired to perform 
the energy balance. A simple first law relationship is 
invoked for the system, 

m h av g = VI - mc p (A T on - A T off ) , (2) 

where m is the mass flow rate, /i av g is the bulk en- 
thalpy, V is the arc voltage, I is the arc current, T 
and c p are the coolant temperature and specific heat, 
and the subscripts of AT refer to a measurement of 
the temperature rise with the arc on and with the 
arc off. This is required to account for the coolant 
temperature rise that results from pumping a viscous 
fluid through the cooling lines. An uncertainty anal- 
ysis for typical measurement errors can be performed, 
and this indicates that the average total gas enthalpy 
can be determined fairly accurately. 15 However, there 
are some important considerations. First, the larger 
the facility, the more difficult it is to accurately mea- 
sure the coolant temperature rise. Either a large num- 
ber of measurements must be made in the smaller 
coolant lines or the temperature distribution in a large 
manifold must be resolved to determine the coolant 
temperature rise. Second, the energy balance does 
not account for further heat losses beyond the nozzle 
that may reduce the bulk enthalpy value of the free 
stream. Finally, although knowledge of the enthalpy 
determined from an energy balance is important and 
useful from a facility perspective, it is still an aver- 
age, or bulk value. This average enthalpy value may 
not be representative of that part of the test stream 
actually impinging on the test article since gradients 
in flow enthalpy that may develop in the arc column 
persist owing to short residence times in the high pres- 
sure region of the nozzle. Perhaps more importantly, 
the energy balance approach provides no information 
about the degree of nonequilibrium or how the energy 
is apportioned in the free stream. 

Fig. 7. Energy balance on a large-scale, constricted- arc 
arcjet test facility. 

Energy balance measurements can also be used to de- 
termine heater efficiency values during facility opera- 
tion. The heater efficiency, which is generally a func- 
tion of arc pressure and current, is defined as 

*Ih(p,I) = rh h wg /{VI). (3) 

Once this is determined for the particular heater con- 
figuration, it can be used to quickly estimate the bulk 
enthalpy using the mass flow rate of the gas and the 
arc voltage and current by simply rearranging the 
equation. Because the efficiency is a function of the 
arc current and the stagnation pressure, this mea- 
surement must be carried out over the full range of 
facility operation to develop an empirical correlation 
that accounts for the dependence. 16 It is important 
to understand that changes in electrode configuration, 
or indeed, variation in electrodes themselves will di- 
rectly influence the heater efficiency. Moreover, the 
electrodes are typically the most frequently replaced 
component of the facility, so efficiency values, and this 
approach to estimating bulk enthalpy, should be used 
with caution. 

Electrode Package 

Fig. 8. Sonic flow method for determining enthalpy in 
an arcjet facility (after Winovich 17 ). 

Sonic Flow - Another method used to determine the 
total enthalpy is the sonic- flow method that was de- 
veloped by Winovich. 17 The basic physical principle of 
this method is that for any given equilibrium thermo- 
dynamic state there is a unique value of the sonic mass 
flow. Thus for a given enthalpy and pressure there is 


only one value of the choked mass flow. Conversely, for 
a known pressure a measurement of the mass flow de- 
termines the enthalpy. The graphical representation 
of this approach is shown in Fig. 8. Assuming that 
the flow is one-dimensional and in equilibrium, then 
for both real and ideal gases a simple expression re- 
lating mass flow and reservoir enthalpy can be derived 
from the equations governing the flow from a reservoir 
through a choked nozzle, 

\/2l^ |> A v\ 1/2 ' 

(R T 0 ) [p 0 t h 0 J 

In the above expression, A is the cross-sectional area, 
the subscript 0 refers to stagnation conditions, and the 
superscript * refers to conditions at the throat. 

Simplified versions of this equation can be derived for 
the case of thermally and calorically perfect gases, as 
well as for calorically imperfect gases. For real gases 
both 7 and R vary with temperature and pressure and 
there are no simple closed form expressions that repre- 
sent this variation. Consequently, the governing equa- 
tions for the choked nozzle flow were solved iteratively 
using an equation of state representing a dissociating 
gas for a range of pressures and enthalpies. For the 
range of pressures investigated ( 0.25 to 100 atm) all 
solutions for the mass flow fell within 4 % of a mean 
curve. A curve fit procedure was then used to develop 
the empirical correlation, 

777 _ C 

(A p T ) h £ 97 

( 5 ) 

where C is a constant factor whose value depends on 
the system of units. The effects of boundary layer, 
nonequilibrium (or frozen) chemistry, and variable 
heat loss to the nozzle walls were examined in the orig- 
inal work pertaining to this measurement approach. 17 
While boundary layer and heat loss effects appear 
to be small, the existence of nonequilibrium flow at 
the throat leads to a systematically low estimate of 
the flow enthalpy. As with the energy balance ap- 
proach, the total enthalpy determined with the sonic- 
flow method represents an average value, and there 
is no information about the degree of nonequilibrium 
where testing takes place beyond the nozzle exit. 

presented an experimental investigation of Goulard’s 
theoretical results for arcjet flows. During the same 
time period, empirical correlations for stagnation point 
heat transfer in any gas were published: 21,22 

( 6 ) 

where k is a gas species dependent constant, R e jj 
is the effective radius of the blunt-body article, and 
Ah is the difference between the stream and cold 
wall enthalpy. A clear advantage of this approach 
is that it gives a spatially resolved measure of the 
stream enthalpy at the test location. However, an im- 
portant assumption in the use of the above correla- 
tion is that the cat aly city of the surface of the heat 
flux gauge is essentially full, i.e. all atoms impinging 
on the surface recombine and deposit the excess en- 
ergy from the exothermic reaction on the surface as 
heat. It should be noted that full catalycity is rarely 
achieved for calorimeters, and heat flux measure- 
ments with gauges of different catalycity show wide 
variation. 23,24 Oxidized, uncleaned surfaces, which are 
typical on calorimeters that are in service, have sig- 
nificantly lower catalycity. This means that calorime- 
ters will generally under-measure the incident heat flux 
when significant dissociated species are present at the 
calorimeter surface. Since the inferred enthalpy is lin- 
early dependent on the measured heat flux, this ap- 
proach will lead to a lower estimate of the stream en- 
thalpy level. 

Fig. 9 . Stagnation point heat transfer measurements. 

Stagnation Point Heat Flux - With certain as- 
sumptions the total stream enthalpy can be inferred 
from a simultaneous measurement of heat transfer and 
impact pressure at the stagnation point of a blunt 
body, such as a sphere or cylinder as depicted in Fig. 9 . 
Boundary layer equations for stagnation point heat 
transfer were developed by Fay and Riddell, 18 and 
a subsequent modification of these results to include 
nonequilibrium boundary layer chemistry and surface 
catalytic effects was given by Goulard. 19 Later, Pope 20 

All of the conventional approaches to arcjet stream 
characterization that have been discussed in this sec- 
tion share common attributes in that they infer en- 
thalpy from other flow property measurements and 
they provide no information on the degree of non equi- 
librium. The ability of measurements made using 
these approaches to guide and inform flow modeling 
is therefore limited. Furthermore, it is not possible 
to use these measurements to relate the arcjet stream 
conditions to the intended flight application, because 


they do not address the nonequilibrium state of the 
free stream. 

New spectroscopic techniques that are currently under 
development may improve this situation, and some of 
these approaches will be discussed in the second lec- 
ture. Judicious application of CFD codes may also 
improve this situation by providing more insight into 
the thermochemical state of the flow. However, any 
CFD codes that are developed for this purpose must 
be guided by experimental results, and this topic is 
addressed in the following sections. 

2. CFD for Arcjet Flows 

Today, CFD is an important resource for aerospace 
vehicle design, testing, and development. Investiga- 
tions into new, or poorly understood, flow problems 
are often undertaken with a combined experimental 
and computational approach. Both the experiment 
and the modeling benefit from the collaboration, since 
the CFD simulations can evaluate a wide parameter 
space quickly and efficiently, while the experimental re- 
sults provide guidance for developing assumptions and 
improving model fidelity. The general state of CFD for 
a particular discipline in the wider field of aerospace 
applications is periodically reviewed. Recent reviews 
that are relevant to simulating arcjet facility flows can 
be found in Refs. 25 and 26, which examine CFD for 
high enthalpy test facilities and external flows, respec- 

The present discussion is concerned mainly with the 
impact of experimental measurements and instrumen- 
tation on modeling, so detailed examinations of numer- 
ical methods, particular models, and grid resolution, 
which are familiar topics in the literature regarding 
CFD, will not be covered. Rather, the intent is to 
discuss shortcomings in current instrumentation and 
available experimental data that make the task of pro- 
ducing credible arcjet flow simulations exceedingly dif- 
ficult, if not impossible. Although the conservation 
equations and general numerical method are discussed 
briefly below, they are invoked only to frame the dis- 
cussion about what must be measured and how well. 
The perspective is that of an experimental approach 
to flow modeling that examines assumptions, model 
inputs, and constraints in order to propose better ex- 
perimental tests to resolve ambiguities and uncertain- 

2.1 Motivation for Arcjet Flow Modeling 

A major driving force behind arcjet flow modeling is 
the desire to extract the most information from tests 
of thermal protection systems in large-scale arcjet fa- 
cilities. Testing costs are always a concern, and an 
investment in computational resources to avoid test 
article failures or to conduct a more efficient test cycle 

represents a prudent strategy. Computational investi- 
gations can often be undertaken at lower expense than 
experimental efforts. Unless a complete computational 
capability is being started from scratch, the costs of 
employing state-of-the-art instrumentation for experi- 
mental investigations is usually much higher, assuming 
that manpower for both efforts is equivalent. If more 
and better information could be obtained from arc- 
jet testing, then substantial development cost savings 
may be realized from a reduced dependence on flight- 
experiments (e.g. FIRE 27,28 and Apollo 29 ) that are 
often required to establish thermal protection system 

Minimizing thermal protection mass for current, low- 
budget planetary missions is also an important motiva- 
tion for generally improving the state of knowledge of 
arcjet flow stream conditions. For these missions there 
is neither time nor budget for flight testing a proto- 
type before launching. If results from arcjet tests can 
be extrapolated to flight conditions with quantifiable 
uncertainties, then it may be possible to reduce the de- 
sign safety margins that currently added to heat-shield 
thickness. 30 It may ultimately be possible to estab- 
lish flight performance of thermal protection materials 
through arcjet testing if a sufficient understanding of 
arcjet flows is developed. CFD modeling would play 
an indispensable and enabling role in this effort. 

Facility improvements and optimization for particular 
test configurations could also benefit from the develop- 
ment of CFD tools tailored to arcjet flow modeling. As 
noted above, ARCFLO played an important role in the 
early development of large-scale arcjet test facilities, 
and there is a need for modern computational tools to 
improve electrode designs, optimize heater configura- 
tions, and design new nozzles for flat- plate test con- 
figurations. These tools could also be used to design 
test configurations that would provide the necessary 
information at reduced cost and effort. 

Additional motivation derives from the desire to im- 
prove the general state of nonequilibrium flow mod- 
eling and the understanding of real gas effects. As 
mentioned above, the stable, relatively long-duration 
arcjet operation at high enthalpies creates opportu- 
nities for studying complex chemical and thermal in- 
teractions that cannot be easily analyzed in impulse 

2.2 CFD Requirements for Arcjet Flow Simu- 

Before discussing measurement requirements for im- 
proving computational simulations of arcjet facility 
flows, it is useful to examine the CFD requirements 
that have evolved from previous and ongoing efforts to 
model arc-heated flows in large-scale facilities. Arcjet 
flows are typically not in thermal and chemical equilib- 
rium, except possibly in the constrictor and electrode 


package regions. Consequently, any attempt to model 
the flow requires a CFD code that models nonequilib- 
rium processes. 

The conservation equations for hypersonic flows in 
thermal and chemical nonequilibrium that are solved 
by the LAURA CFD code 31 have been compiled 
in a single reference publication by Gnoffo and his 
co workers. 3 2 Eleven species that are typically encoun- 
tered in simulations of hypersonic air flows were in- 
cluded in the model: N 2 , O 2 , NO, N, O, Nj, O^", NO + , 
N + , 0 + , and e“ . Thus, eleven species continuity equa- 
tions and three momentum equations must be solved 
by the code. For this particular CFD code, three 
separate energy equations are modeled to account for 
nonequilibrium effects: vibrational energy, electronic 
energy, and total energy. Thermodynamic data for 
the eleven species and reaction rates for two different 
models, Park 33 and Dunn and Kang 34 , were also given 
in the report. In this CFD approach, which is repre- 
sentative of those currently in use for nonequilibrium, 
hypersonic flows, only the ground electronic states of 
each species are modeled. When radiative energy flux 
is important, it is typically treated separately or in a 
loosely-coupled fashion. It should be emphasized that 
there is no universally agreed upon model formulation, 
particularly when it comes to nonequilibrium processes 
and chemical reaction and energy transfer rates. In- 
terested readers are referred to Refs. 35-38 for other 
computational model formulations. 

For arcjet flows in large scale facilities, argon must 
also be considered since it is often added to the test 
gas flow to protect electrode surfaces. If only the neu- 
tral state is considered, this brings the total number 
of species for air/argon flows to twelve. In addition, 
thermodynamic and chemical reaction rate data must 
also be included for argon. 39 

2.3 Strategies for Arcjet Flow Simulations 

Just as there is no universally accepted model for 
nonequilibrium, hypersonic flows, there is no single 
CFD code that can simulate the complete arcjet fa- 
cility flow from the heater to the test article. Thus, 
some reasonable modeling strategy must be developed 
that matches available CFD models to flow regions in 
an advantageous manner. To illustrate this point, sev- 
eral modeling efforts that were concerned with either 
arcjet characterization or interpretation of arcjet test 
results are surveyed below. The presentation is or- 
ganized by flow region, starting from the heater and 
moving through the nozzle to the test article. 

Arc Heater - The flows within the arc heater and 
electrode packages are special cases, since the elec- 
trodynamic processes occurring within these typically 
subsonic flow regions are usually absent in hypersonic 
flows (with the exception of MED device flows). How- 
ever, a discussion of numerical studies of arc heaters 

is included here for two reasons: first, the flow may be 
in thermal and chemical equilibrium within the down- 
stream electrode package; and second, if the flow and 
discharge physics can be modeled correctly, then it 
may be possible to compute inlet conditions for use 
in nozzle calculations. 

Within an arc- heater, the flow is typically subsonic 
and is more properly described as a plasma owing to 
the presence of the electrical discharge. To model this 
portion of the flow accurately, a coupled solution of 
the fluid dynamics, radiation, and electrodynamics is 
required. The development of a CFD model for the 
arc heater that included the necessary coupling was 
undertaken at Ames Research Center several years 
ago, 40,41 but the effort was eventually abandoned. In- 
stead, the flow within constricted-arc heaters is still 
modeled with either the ARCFLO code, which was 
mentioned above, or a derivative. One of the deriva- 
tive codes, SWIRL ARC, 42,43 has been modified to in- 
clude the tangential component of gas injection that is 
typically used to help stabilize the discharge in high- 
pressure facilities. It should be noted that in any form, 
ARCFLO does not attempt to fully simulate the physi- 
cal processes within the heater. Rather, ARCFLO and 
its derivatives employ a semi-empirical approach to 
perform comparative studies and indicate trends that 
might be useful for design studies. 

Recently, there has been renewed interest in improv- 
ing computational models of constrictors. A Navier- 
Stokes formulation for a constrictor was developed and 
implemented by Kim et al, 44 and an improved, fully- 
coupled radiation model was applied to the study of 
an arc heater by Sakai et al. 45 The main motivation 
for this renewed activity is the need to increase the ef- 
ficiency and performance capabilities of existing arcjet 
facilities. Obviously, measurements will be required to 
validate these newer flow models. 

Nozzle and Free Stream - For studies relating to 
arcjet characterization, arcjet test interpretation, or 
general nonequilibrium flow modeling, the expanding 
flow in an arcjet facility nozzle presents a challenge to 
the computational community. The general strategy 
for modeling arcjet nozzle flows relies on some means 
for estimating the inlet conditions for the nozzle, par- 
ticularly the stagnation enthalpy, and then using what- 
ever experimental information is available from the 
free stream to assess the fidelity of the simulation. De- 
pending on the particular computational model, the 
inlet conditions can be specified either in the subsonic 
flow region upstream of the throat or in the super- 
sonic portion of the nozzle. As was mentioned in the 
description of a typical arcjet facility, the nozzle flows 
are not in thermal or chemical equilibrium. Therefore, 
the computational approach must model the thermo- 
dynamic and chemical kinetic processes that govern 
hypersonic, nonequilibrium flows. 

3 A- 10 

There have been several efforts aimed at simulating 
flows in arcjet nozzles. At Ames Research Center 
alone, three different numerical approaches have been 
used recently to simulate nozzle flows in conical 46-48 
and semi-elliptic 49 geometries. These particular stud- 
ies were undertaken specifically to address arcjet char- 
acterization issues. Additional investigations of con- 
ical nozzle flows have been carried out in support of 
arcjet surface catalysis experiments. 50 Details of the 
different numerical approaches are given in each, of the 
references. However, it is interesting to note the pro- 
gression of the numerical models used in these stud- 
ies. Babikian used a quasi- one- dimensional, multi- 
temperature flow model, NOZNT, 51 to compare with 
free stream temperature measurements in the Ames 
AHF Arcjet Facility. 46 Gokgen performed simulations 
of the nozzle flow with an axisymmetric, nonequilib- 
rium Navier-Stokes solver in support of shock layer 
experiments. 4 ' More recently, Loomis and his cowork- 
ers used GASP, which is a general three-dimensional, 
flow solver to simulate both conical and semi-elliptic 
nozzle flows in support of thermal protection material 
tests for the X-33. 49 

Concurrent experimental and computational studies of 
expanding, No/Ar plasma flows have also been car- 
ried out by Schonemann and co workers. 52 The note- 
worthy aspect of this particular investigation was the 
use of experimental measurements at one axial loca- 
tion to start the calculations and predict the rapidly 
expanding flow properties at a second, downstream lo- 
cation. This approach has the advantage of avoiding 
some of the uncertainties that result from estimating 
inflow conditions. 

Flow Over a Test Article - As the current use of 
large-scale arcjet facilities is aimed primarily at simu- 
lating aerothermal heating, it is extremely important 
to be able to model the flow over a test article ac- 
curately. Test article flows can be classified into two 
basic types: shock-layer flows over a test article in 
a conical nozzle flow and bound ary- layer flow over a 
flat plate for semi-elliptic, or rectangular, nozzle flow. 
Since the flat plate is usually an extension of the nozzle 
wall, the modeling requirements for simulating bound- 
ary layer flows are identical to those for nozzle flows, 
although the angle of attack is typically varied as part 
of an experimental investigation. Shock-layer flows are 
different, particularly for studies of stagnation point 
heating. For this configuration, the flow undergoes 
compression by a shock wave, whose strength depends 
on the particular test conditions and geometry, before 
impinging on the test- article surface. Thus, the free 
stream conditions, which largely determine the char- 
acteristics of the shock layer flow, must somehow be 
known to carry out the simulation. 

Typically, there are no stream measurements, other 
than pitot pressure and heat flux, that could be used 

to specify the stream conditions. For certain arcjet test 
conditions, it is possible to estimate the stream condi- 
tions using a combined equilibrium and frozen-flow an- 
alytical approach, and then carry the analysis through 
the shock layer based on measurements of the pitot 
pressure, heat flux, and model surface temperature. 50 
However, a more general approach involves simulating 
the nozzle flow (again, an estimate of the initial en- 
thalpy is required) with a numerical model and then 
using those conditions as input to a shock layer calcula- 
tion. An example of this latter approach can be found 
in the work of Gokgen, 4 '’ 48 which will be discussed in 
detail below. Inevitably, inaccuracies in modeling the 
expanding nozzle flow affect the simulations of shock- 
layer and boundary-layer flows in arcjet facilities. 

The response of the test article to the shock layer flow 
is also of considerable interest to the arcjet test and 
material development communities, where much can 
be gained by understanding the interaction between 
the shock layer flow and the material. A review arti- 
cle by Milos and Rasky 53 outlines the importance of 
properly defining the boundary conditions that gov- 
ern the interactions at the fluid/surface interface. The 
authors also point out that since boundary processes 
define the interaction of the fluid and solid computa- 
tional models, their boundary conditions must agree. 
This issue is especially important for understanding 
the performance of charring and ablating thermal pro- 
tection materials. Although it is very interesting, this 
topic is outside the scope of the present discussion. 

Measurement Requirements fur Arcjet Flow Modeling 

Arcjet Flow Measurements j 

Model Development Measurements 

Starting Conditions 

Simulation Validation 

Three-body recombination rales 
Third- body efficiencies 
Spontaneous emission rates 
Laser-excitation rates 
Col lisional-radi alive model rates 
Species thermodynamic data 
Energy Transport Rates 

Mass flows of 
test gases 
Inflow velocity 

Turbulent or 
Assessment of 


Species concentrations 
Translational T 
Rotational T 
Vibrational T 
Electronic state 
Post <no// le expansion 

Stream profiles 

Fig. 10. Classification of measurements for arcjet flow 

3. Measurement Requirements for Arcjet Flow 

Measurements that can be used to improve compu- 
tational models of arcjet flows can be separated into 
general categories, which are illustrated in Fig. 10. 
The first classification distinguishes between direct 
measurements of properties of arcjet flows and more 
generic measurements that can influence the develop- 
ment of models for nonequilibrium, hypersonic flows. 
Measurements of thermodynamic properties, species 
concentrations, velocity, and enthalpy in arcjet flows 

3 A- 11 

would all fall into the first category of direct measure- 
ments. More accurate determinations of important 
reaction or energy transfer rates, which need not be 
measured in arcjet flows, would fall into the second 
category. While this category is probably of equal im- 
portance in the improvement of arcjet flow modeling, 
the majority of the discussion below is concerned with 
direct measurements of primary arcjet flow quantities. 

Within the first category of direct arcjet flow prop- 
erty measurements, a further distinction can be made 
between measurements that would be used to define 
starting, or inflow-boundary, conditions and measure- 
ments that could be used to assess the fidelity of the 
simulation. Since the success of any flow modeling ef- 
fort is inextricably linked to the accuracy with which 
the starting conditions for the calculation are defined, 
measurements of the input parameters are considered 
to be of greater importance. Of the inflow parame- 
ters for arcjet flows, the total enthalpy is the most, 
important because it defines the total flow energy and 
the initial composition and temperature. Despite its 
importance, enthalpy has proven to be the most diffi- 
cult parameter to characterize accurately, as discussed 
above. Typically, the settling chamber pressure is 
measured to reasonable precision for most arcjet tests, 
so it is assumed herein that pressure is given. Other 
primary measurements that define the starting con- 
ditions are the mass flows of the test gases and the 
configuration aiid geometry of the facility. 

Velocity, species concentrations, temperature(s), static 
pressure, and density are examples of flow property 
measurements that can be made at various locations 
in the arcjet to assess the performance of a computa- 
tional model. Flow quantities that are derived from 
measurements of primary flow variables, such as the 
dynamic pressure, specific heat ratio, Mach number, 
and Reynolds number, are less important from the 
perspective of evaluating computational models. How- 
ever, these quantities are quite useful in specifying the 
performance of the arcjet facility and for relating the 
test conditions to the expected flight environment. 

3.1 Enthalpy 

It is readily apparent from even a casual reading of 
the previous section on modeling requirements that all 
simulations of the most important arcjet regions, the 
nozzle and shock- layer, or boundary-layer, flows, re- 
quire knowledge of the stagnation enthalpy. The state 
of enthalpy determination using conventional instru- 
mentation was examined in the introductory section, 
and it was found to be inadequate for several reasons. 
First, the conventional means for determining the flow 
enthalpy can only give an estimate of the total value, 
which does not specify the state of the essentially 
frozen free-stream flow. Second, for the energy bal- 
ance and sonic flow approaches, only the bulk enthalpy 

value can be determined. While this is useful for mon- 
itoring facility performance, the enthalpy value proba- 
bly does not represent the free-stream core flow, where 
stagnation- point tests are conducted, unless there are 
no span wise enthalpy gradients. The assumption of 
gradient-free flow appears to be questionable . 25 Third, 
even when great care is taken with the treatment of 
the calorimeter surface, enthalpy values derived from 
heat flux measurements are likely to be systematically 

It should be noted that the measurements required 
for determining flow enthalpy vary with flow region. 
Moreover, the influence of the enthalpy determination 
on the outcome of the flow simulations also depends 
on where the enthalpy measurement is made. For arc 
heater and nozzle flow simulations, a measurement of 
the stream enthalpy within the downstream electrode 
package, which also functions as a nominal settling 
chamber, is appropriate. However, for shock layer sim- 
ulations, inaccuracies are accumulated from simulating 
both the nozzle flow and the shock layer flow. A more 
appropriate enthalpy measurement location would be 
the free stream, provided that the measurement could 
quantify both the total enthalpy and the nonequilib- 
rium state of the gas. With this information about 
the free stream, the shock layer flow could be sim- 
ulated independently of the nozzle flow. Obviously, 
owing to the nonequilibrium nature of the flow, more 
flow property measurements are required to determine 
the thermodynamic and chemical state of the flow in 
the free stream. 

Finally, any enthalpy measurement must be spatially 
resolved, and enthalpy gradients must be quantified to 
remove potential ambiguity from the specification of 
the starting conditions. This issue will be discussed 
further below and the applicability of nonintrusive op- 
tical diagnostics to enthalpy measurements will be ad- 
dressed fully in the following lecture. 

3.2 Arc Heater 

The flow within the arc heater and electrode package is 
usually subsonic and the enthalpy is mostly static, be- 
ing comprised of thermal and chemical mode contribu- 
tions. Because pressure is reliably known, a measure of 
total density or temperature would permit a determi- 
nation of the total enthalpy. Of the two variables, tem- 
perature is more amenable to measurement through 
optical means. Assuming that the flow is in thermal 
equilibrium, then determination of a single tempera- 
ture is sufficient for determining enthalpy. If there is 
optical access to either the heater or electrode pack- 
age region, then a spectrally resolved emission mea- 
surement can be used to determine temperature. The 
specific procedures for determining temperatures from 
spectrally resolved emission are discussed in the fol- 
lowing lecture. 


Useful information could also be derived from addi- 
tional measurements of other flow variables in the elec- 
trode package. These other flow properties include: 
the axial flow velocity; the azimuthal velocity compo- 
nent, which would quantify the amount of swirl at the 
nozzle inlet; the total heat flux and radiative heat flux 
to the wall; the amount of copper, which is introduced 
into the stream by the process of electrode erosion; 54 
and the electron number density downstream of the 
arc. Although the axial extent of the electrode pack- 
age region is usually not that large, the flow is cooling 
as it moves toward the throat. Consequently, a de- 
termination of the axial variation in any flow quantity 
would provide some insight into the evolution of the 
flow as it begins to accelerate. 

Owing to limited accessibility, flow probes are not a vi- 
able option. Their survival at typical large-scale arcjet 
facility operating conditions is also an issue. Optical 
access to the downstream electrode package can of- 
ten be realized, 55 and measurements in this region are 
particularly useful because this region provides the in- 
flow to the nozzle. It may be possible to implement 
optical measurements at two different axial locations 
downstream of the arc termination to assess the rate of 
evolution of the stream properties. Furthermore, since 
large fluctuations in the magnitude of emission from 
atomic transitions have been observed in the electrode 
package, 55 it may be possible to develop a two point 
correlation approach for velocity measurement. 

3.3 Arcjet Nozzle and Free- St ream Flow 

Nozzle - For the purposes of this discussion, the start- 
ing point for nozzle flow is defined as the end of the 
electrode package. Unlike the segmented arc heaters 
and electrode packages of todays constricted-arc facili- 
ties, the nozzle assemblies are typically fabricated in a 
more monolithic manner. Because they are fabricated 
with integral water cooling, there is little hope for in- 
strumenting existing large-scale arcjet nozzles. This 
essentially precludes in situ monitoring of the onset of 
chemical and thermal freezing, which could then be 
used as a starting point for frozen flow analysis. 

Using smaller scale arcjet devices fabricated with seg- 
mented nozzles it may be possible to address the on- 
set of chemical, and possibly thermal, freezing for flow 
conditions of interest in aerothermal testing applica- 
tions. Note that the fluid dynamic expansion rate 
plays an important role in determining the location 
at which the flow freezes. Whatever studies are un- 
dertaken in smaller facilities must address this issue. 

Free Stream - Although the flow is usually chem- 
ically and thermally frozen by the time it exits the 
nozzle, the free stream region is often optically ac- 
cessible, and measurements of many flow properties 
are possible. Spatially resolved measurements of ve- 
locity, translational temperature, density, pressure, 

and species concentrations have ail been made using 
laser-induced fluorescence (LIF) techniques. 56 ” 59 Re- 
cently, measurements of enthalpy and its distribution 
among thermal, chemical, and kinetic modes, were 
demonstrated in N 2 /A 1’ 58 and air/Ar 59 flows using 
two-photon LIF of atomic nitrogen. Although more 
property measurements are required to determine en- 
thalpy for nonequilibrium flow, the approach of using 
LIF of the dissociated species to determine multiple 
flow parameters appears capable of providing this in- 
formation with the aid of certain assumptions. For- 
tunately, the validity of the assumptions that are cur- 
rently invoked can be evaluated experimentally. 59 Flow 
property measurements using LIF techniques will be 
discussed extensively in the second lecture. 

Although further development of this approach is re- 
quired to assess the assumptions and establish the 
range of applicability, LIF based stream property mea- 
surements may ultimately prove sufficient to establish 
the enthalpy and degree of nonequi librium of the free 
stream flow. This would provide a set of inflow con- 
ditions that could be used to calculate the shock-layer 
flow. A computational simulation of the flow over a 
test article that was started from known free stream 
conditions and compared to shock-layer property mea- 
surements would allow a better assessment of the com- 
putational modeling. Determinations of free stream 
rotational and vibrational temperatures and assess- 
ments of possible metastable atomic state populations 
are needed to establish the validity of the LIF-based 
approach. 59 

In addition to establishing inflow conditions for shock 
layer simulations, the two- photon LIF measurements 
provide stream property information that can be used 
to evaluate the fidelity of nozzle flow simulations. 
Since total enthalpy is specified by the LIF measure- 
ments, with a quantified uncertainty, that, value can be 
used along with the constrictor pressure to start the 
nozzle simulation. If the model used in the nozzle flow 
simulation is accurate, it should reproduce the mea- 
sured distribution of the total enthalpy into kinetic, 
thermal, and chemical contributions in the nonequi- 
librium free stream. Comparisons between nozzle sim- 
ulations and free stream measurements are underway 
for the chemically simpler N 2 /argon flow cases. 

3.4 Blunt- Body Shock- Layer Flow 

Even with the free stream conditions specified, much 
is required in order to improve the general understand- 
ing of shock- layer flows in the stagnation- point heating 
configuration for aerothermal test applications. Finite 
rate effects that vary in significance depending on the 
test conditions and model geometry still control the 
chemical and thermal state of the shock layer and im- 
pact issues such as the difference in catalytic heating 
between the arcjet test conditions and the flight envi- 
ronment. Moreover, depending on the test geometry 


and conditions, merged shock layer and rarified flow 
effects may be important, and may complicate the in- 
terpretation of heat transfer data.. 

Instrumentation and techniques for making spatially 
resolved flow property measurements are therefore re- 
quired to establish the thermochemical state of the 
gas in the shock layer as it moves toward the surface 
of a test article. Spatial resolution is important be- 
cause the flow is generally evolving from a nonequilib- 
rium state toward an equilibrium state as it approaches 
the surface. Understanding this evolution is important 
from a modeling perspective as well as for aiding in the 
interpretation of test results. Again, for nonequilib- 
rium situations, multiple flow properties, including ve- 
locity, species concentrations and temperatures must 
be measured to specify the flow state. In contrast to 
free steam conditions, shock layer temperatures (T r , 
T„, and T e ) can reach levels in excess of 5000 K and 
pressures can be orders of magnitude higher. Many 
internal energy levels of a number of species will be 
populated, and the distributions of populations over 
these energy levels may differ for different species (and 
possibly electronic states). With an ablating material, 
the situation is even more complex. 

However, the goal of understanding the shock layer 
thermochemistry is important because that is the envi- 
ronment that must ultimately be related to flight con- 
ditions. In addition, if instrumental approaches that 
determine the thermochemical state of the shock-layer 
can be developed, then it may be feasible to test all- 
body vehicle configurations in the long-duration, arcjet 
flow facilities. 

3.5 Measurement Accuracy Requirements 

The uncertainty in experimental measurements and in 
computational predictions is an important consider- 
ation in arcjet flow investigations. For the present 
discussion, only uncertainties in experimental mea- 
surements are considered. Experimental uncertainties 
are estimates of errors in measurements that typically 
arise from either systematic or random contributions, 
or more typically, both. The systematic and random 
contributions are manifestations of the more general 
measurement attributes: accuracy and precision. Def- 
initions of measurement accuracy and precision, which 
are frequently confused, can be found in a variety of 
reference publications, including a text on the sub- 
ject by Bevington. 60 In the introductory discussion, 
Bevington indicates that the accuracy relates to how 
close a measured value is to the “true” value, while 
precision provides information on how well something 
can be measured, regardless of what that measurement 
means. For arcjet flow property measurements that 
would be used to improve flow modeling capabilities, 
both accuracy and precision are important. However, 
because flow properties are generally unknown, instru- 
mentation must be developed to make measurements 

with a minimal reliance on assumptions that cannot 
be tested; otherwise it is impossible to assess accu- 
racy. How closely a measured flow property represents 
the real situation is important from the perspective 
of predicting absolute flow property magnitudes, as in 
the simulation of a single arcjet test. 

In contrast, it is often easier to establish the precision 
of a particular measurement from a number of different 
observations at similar conditions. This can be done 
without evaluating all of the assumptions that may go 
into a particular measurement, and it may then be pos- 
sible to use the measured quantity constructively with- 
out knowing the absolute accuracy. Once the precision 
is established, then the measurement could be made 
for a number of different flow conditions where a single 
control parameter, such as the arc current, is varied. 
Simulations of these different flow conditions and com- 
parisons with the experimental results would test the 
ability of the model to predict trends and might ulti- 
mately do more to establish confidence in the model- 
ing than single-condition comparisons. Obviously, the 
model would have to be optimized for an appropriately 
chosen test condition. This approach may prove to be 
more effective in advancing both modeling and instru- 
mentation development, especially when one considers 
the number of measurements required to document a 
single test case completely. As the instrumental tech- 
niques mature and as more is learned from parametric 
comparisons, then it may be feasible to pursue a single, 
well-documented test case. 

From an experimental perspective, given the scarcity 
of data and the fact that measurement results from one 
facility cannot be directly transferred to another unless 
the configuration is identical, any property measure- 
ments that also have quantified uncertainty estimates 
are useful in advancing the general state of knowledge. 
Enthalpy measurements, as well as other inflow con- 
ditions that are required to initiate simulations, are 
examples of this type of flow property. Measurements 
that are used to evaluate the validity of computational 
models must be held to a higher standard, since their 
determination may influence changes in the model for- 
mulation. It is difficult to formulate a general state- 
ment as to how high the standard should be, given the 
complexity of present day CFD models. For some pa- 
rameters an accuracy requirement can be postulated 
easily. As an example, consider LIF-measured atomic 
nitrogen concentrations, which currently have an esti- 
mated uncertainty of ^ 30 % . 59 Although this uncer- 
tainty appears to be large, the recombination rate for 
the reaction, 

A + A + M->A 2 + M , 

which largely determines the N atom concentration in 
the chemically frozen free stream, is currently uncer- 
tain by up to a factor of three. 62 Clearly even the 


relatively uncertain N atom concentration measure- 
ment can be used to evaluate flow model performance. 
To determine accuracy requirements for other flow 
properties, an effort should be made to evaluate un- 
certainties in quantities currently used in the models 
and parametric studies with the computational model 
should be performed to evaluate sensitivities. 

4. Experimental and Computational Investiga- 
tion of Shock-Layer Flows 

Recent attempts to simulate arcjet flows and compare 
the numerical predictions with experimental measure- 
ments illustrate the current status of both simulations 
and measurements. Knowledge gaps that affect the 
comparisons between simulations and measurements 
are readily apparent. The combined experimental and 
computational investigation of blunt- body, shock- layer 
flows in the Ames AHF Arcjet Facility was chosen for 
this purpose because the investigation was motivated 
by the need for improved characterization of arcjet 
flows, and understanding the shock- layer flow is di- 
rectly relevant to aerothermal testing of thermal pro- 
tection materials. Although experimental investiga- 
tions have also been undertaken in the electrode pack- 
age and free stream regions of the same arcjet facility, 
comparisons between measurements and simulations 
for those studies are ongoing. Consequently, more can 
be learned from examining the process and the results 
of the documented shock-layer flow property compar- 
isons, and from the results of those comparisons. 

4.1 Objectives of Investigation The objectives of 
this investigation were to: 1) determine whether a re- 
gion of thermal and chemical equilibrium exists in the 
shock layer formed over a flat- faced cylinder; 2) de- 
termine the conditions required to establish the equi- 
librium region; and 3) determine whether or not en- 
thalpy measurements could be derived from spectrally 
resolved emission emanating from the equilibrium re- 

Early investigations of arcjet facility flows included ef- 
forts to characterize the shock layer flow using emis- 
sion spectroscopy. 62,63 If the flow is in thermal and 
chemical equilibrium, then a measurement of the tem- 
perature from spectrally resolved emission and a con- 
current pressure measurement would uniquely specify 
the thermo chemical state of the flow and its enthalpy. 
A relatively recent analysis of shock- layer emission ap- 
peared to indicate the presence of an equilibrated re- 
gion within the shock layer at a lower pressure than 
had been expected. 46 Therefore, a major goal of the 
present investigation was to verify the existence of the 
equilibrated region, and begin the task of defining the 
conditions that produce the equilibrium flow. By un- 
dertaking this investigation, issues associated with the 
development of an “enthalpy meter” based on mea- 

surements of shock layer emission could also be as- 

At the outset, several areas of uncertainty were iden- 
tified that had to be addressed in the investigation. 
These areas included: 1) what criteria are used to 
identify a region of thermal and chemical equilibrium; 
2) how to interpret emission measurements with cer- 
tainty; and 3) the effects of spatial intensity gradients 
on measurements that are integrated along the line-of- 
sight. It quickly became apparent that computational- 
flow modeling could address some of these issues and 
help guide the experimental investigation. Conversely, 
it was realized that the experimental measurements 
might also provide some assessment of the computa- 
tional model validity, but this was not the primary 

4.2 Experimental and Computational Approach 

Experiment - The shock layer emission experiments 
were carried out in the Ames AHF Arcjet Facility, 
which was previously described in the introductory 
section. To generate the highest shock layer pressure 
values the facility was configured with the 30.5 cm di- 
ameter nozzle, which produces the least free stream 
expansion. A 15.2 cm diameter, flat- faced cylinder 
made of copper was placed in the stream to generate 
the shock layer. Two different test conditions were 
surveyed, and these are referred to as the high pres- 
sure and low pressure cases. Test conditions for the 
two cases are summarized in Table 1. 

Table 1. Arcjet test conditions for shock layer inves- 

Test Conditions 
















To obtain as much information from a single facility 
run as possible, line-of-sight emission spectra were ac- 
quired from multiple axial locations along the central 
stagnation streamline using a spectrograph and CCD 
camera. The model was placed at two different axial 
locations in the stream, 34.5 and 36.9 cm downstream 
of the nozzle exit (forward and back positions, respec- 
tively), to allow full coverage of the shock layer emis- 
sion with the finite viewing area of the CCD and spec- 
trograph system. Thus, two separate facility runs at 
the same nominal operating conditions were required 
to obtain the full shock layer emission profile for each 
test case. 

Emission spectra were acquired at several grating po- 
sitions covering the UV to near- 1 R wavelength range 
and the measured signals were converted to absolute 
intensities through calibration with standard spectral 
lamps. Each grating position was chosen to measure 


certain spectral features that could be used to ascer- 
tain temperature or species information using spectral 
analysis techniques. 64,65 Particular attention was given 
to developing methods for determining rotational and 
vibrational temperatures using spectral feature ratios 
that minimized the influence of uncertainties in the 
measurements. This was desirable because agreement 
between the measured temperature values was thought 
to be a good indicator of the presence of a thermally 
equilibrated flow region. Portions of the Nj (1,2), 
(0,1), and (0,0) vibrational bands were found to yield 
vibrational and rotational temperatures with mini- 
mum uncertainty based on spectral simulations. Note 
that the temperature values are derived from emis- 
sion that is integrated along the line-of-sight, so the 
inferred flow properties actually represent intensity- 
weighted, spatially averaged values. Further details of 
the experimental configuration and the spectral anal- 
ysis can be found in Refs. 65 and 66. 

Computational Modeling Approach - The ex- 
perimental measurements consisted of spectrally re- 
solved, absolute intensities from multiple axial loca- 
tions within the shock layer during a single facility 
run. Consequently, the computational simulation had 
to be able to address issues relating to emission, which 
required the use of a radiative transport code. In ad- 
dition, two different CFD models were required to 
predict the shock layer flow. The first model was 
used to simulate the nozzle flow to determine the free 
stream conditions ahead of the shock layer, which was 
then simulated with a second, separate computational 
model. Flow properties predicted by the shock layer 
model were then used to calculate the radiative trans- 

The two flow models that were used for the simulation 
were developed by Gokgen. 67,68 Both models use an 
a xisym metric formulation, which is appropriate for the 
conical nozzle flow and the shock layer flow over a flat- 
faced cylinder. Twelve chemical species: N 2 , O 2 , NO, 
N, 0, Nj , O*, NO + , N + , 0 + , e” , and Ar; are modeled 
for these flows, and three temperatures: translational, 
rotational, and vibrational; are used to represent the 
thermal state of the gas mixture. The reactions and 
rate coefficients that are used in the model are derived 
from the multi- temperature model of Park and Lee. 51 
Turbulence is not included in either flow model; the 
flow is assumed to be laminar throughout the facility. 
Further details of the computational models can be 
found in Refs. 67 and 68, and more information about 
the nozzle and shock layer computations can be found 
in Refs. 47 and 48. 

To calculate the emissive flux for the shock layer flow, 
the NEQAIR 69 radiative transport model was used. 
Inputs to the model, which include species concentra- 
tions and temperatures, were obtained from the flow 
solution by interpolating between calculated quantities 

at the known measurement locations. For all of the cal- 
culations, the emitting level populations were assumed 
to follow Boltzmann distributions, albeit with poten- 
tially differing values of rotational and vibrational tem- 
peratures. The electronic temperature was assumed to 
equal the vibrational temperature. 

Starting Conditions for the Calculations The 
most important inputs to the CFD model are the start- 
ing conditions, and, as discussed above in the intro- 
ductory section, the flow enthalpy and its spatial dis- 
tribution are not known in the free stream or at the 
entrance to the nozzle. Therefore, some means of esti- 
mating the starting conditions based on facility mea- 
surements was required for the nozzle and shock layer 
simulations. For this investigation, measurements of 
stagnation point heat flux and shock layer pitot pres- 
sure were used to estimate the total stream enthalpy 
based on an empirical correlation of the form presented 
in Eq. (6). At the time of this work, the argon mass 
flow rate had not been recently measured, so it was 
assumed to be equal to 5 % of the total mass flow for 
the high pressure case. Using the estimates of flow 
enthalpy and argon mass flow along with the facility 
measurement of the arc heater pressure, the nozzle 
flow could be calculated. Based on the low level (« 
3 ppm) measured in the stream of the 60 MW arcjet 
facility, 9 copper was not included in the simulations of 
the AHF arcjet nozzle and shock layer. 

Upon exiting the nozzle, the flow regions that are 
within the local Mach angle near the periphery ex- 
pand into the test box at a rate that is different from 
that in the nozzle, unless the nozzle exit static pressure 
matches the ambient pressure. Owing to this further 
expansion, the calculated flow property values at the 
exit of the nozzle could not be used directly as the in- 
flow conditions for the shock layer simulations. This 
effect is well documented for perfect gas flows, 70 but 
the rates of expansion for the nonequilibrium arcjet 
flows at various operating conditions are not known. 
Therefore, the calculations were continued in the axial 
direction at the nozzle expansion rate until the cal- 
culated dynamic pressure matched the value derived 
from the shock layer pitot pressure measurement us- 
ing Eq. (1). Although this procedure does not exactly 
reproduce the fluid dynamics of the full free stream, 
it produced acceptable inflow conditions for the shock 
layer simulations without requiring a complete simu- 
lation of the flow in the test box. 

It should be noted that the enthalpy value that was 
ultimately used in the simulations was greater than 
the value derived from the stagnation point heat flux 
and pressure measurements for both the high and low 
pressure test cases. After performing initial compar- 
isons between the calculated and measured emission 
spectra, the enthalpy was increased for both test cases 
to bring the calculated spectra into closer agreement 


with the measurements. The totai increase for the high 
pressure case was limited arbitrarily to 10 %, although 
it could have been increased further, as will be seen be- 
low. Adjusting the total flow enthalpy estimate to a 
higher value is justified if the calorimeter surface has a 
low catalycity and the recombining atom flux is signifi- 
cant, because correlations such as Eq. (6) apply to fully 
catalytic surfaces. While the catalycity of the partic- 
ular calorimeter used for these measurements is not 
precisely known, the surface of the constantan foil was 
known to have a ceramic oxide coating, so the gauge 
was considerably less than fully catalytic. Given the 
uncertainty in the total flow enthalpy, the poor agree- 
ment found in the initial comparisons between calcu- 
lated and measured spectra could not be attributed to 
the flow model. This clearly illustrates the importance 
of quantifying the stream enthalpy, since using the ex- 
perimental measurements to guide the adjustment of 
the input conditions compromises any assessment of 
the flow model. 

4.3 High Pressure Case 

Calculated Shock- Layer Properties - Using the 
inflow conditions and two-model computational ap- 
proach the flow properties were calculated for the 
shock layer at the high pressure case conditions. Axial 
profiles of the flow properties along the central stream- 
line where the measurements were made are shown in 
Fig. 11 and 12. The axial distributions of pressure and 
temperature are shown in Fig. 11 as a function of non- 
dimensional ized distance from the test article surface. 
Note that the nozzle solution predicts that the flow 
is vibration ally frozen ahead of the shock, but the ro- 
tational energy is predicted to be in equilibrium with 
translation, The rotational and translational temper- 
atures rise significantly near the shock and decrease 
as the blunt- body surface is approached. A lesser in- 
crease is exhibited by the vibrational temperature, and 
all three temperatures are seen to converge to a value 
that is very close to the equilibrium temperature for 
these conditions at about 0.1 R upstream of the sur- 
face. Thus, the calculations suggest that there is a 
region of thermal equilibrium within the shock layer 
at the high pressure conditions. 

Axial profiles of species mass fractions are shown in 
Fig. 12, also as a function of the normalized distance 
upstream of the test article surface. Outside of the 
surface boundary layer, at the point where the tem- 
perature values converge, the species mass fractions 
are quite close to their equilibrium values, which are 
indicated on the right-most vertical axis. According to 
the simulation, the flow is also very nearly in chemical 
equilibrium at these test conditions. 

Distance, x/R 

Fig. 11. Temperature and pressure profiles within the 
shock layer for the high pressure case. 

Fig. 12. Species mass fraction profiles within the shock 
layer for the high pressure case. 

In the description of the experimental effort above, the 
use of Nj spectral features to determine vibrational 
and rotational temperature values within the shock 
layer was noted. For an isolated rovibrational transi- 
tion, the measured intensity for thermal equilibrium 
conditions can be expressed as 

h = ^e- E «' kT , (7) 

where I\ is the spectral intensity, K\ represents the 
line shape function and the transition strength, n is 
the species density, Q is the partition function, L is 
the line-of-sight path length, and E u is the total upper 
state energy. This expression shows that the inten- 
sity has a linear dependence on the emitting species 
density and an exponential dependence on tempera- 
ture. From the computed axial property profiles of 
Figs. 11 and 12, it is apparent that all temperatures 
and the Nj concentration are higher near the shock 
front. Consequently, the measured N* emission could 


easily be dominated by contributions from emission at 
the shock front, where nonequilibrium effects are more 
likely to be present. Temperatures are derived from 
intensities with the implicit assumption that the ro- 
tational and/or vibrational level populations are each 
thermally equilibrated. If the measured intensity is 
dominated by emission from nonequilibrium regions, 
then this assumption is untenable. 

The prediction of significantly higher temperatures 
and N* concentrations near the shock front led to 
a further analysis of the flow property distributions 
along the emission measurement sight lines. Recall 
that one of the objectives was to use temperature mea- 
surement comparisons to determine whether or not the 
flow reached thermochemical equilibrium. This re- 
quires that the temperature values derived from the 
spectral analysis actually represent the central, core- 
flow region, rather than the shock front. To assess this 
potential problem in interpreting the spectral data, 
computed, flow property profiles along the spanwise 
flow direction were extracted from the shock layer so- 
lution at selected axial measurement locations. 

Distance along the line of sight, m 

Fig. 13. Temperature variation along the line-of-sight 
for the high pressure case. 

Profiles of rotational and vibrational temperatures 
along the line-of-sight direction, which is normal to 
the flow axis, are shown in Fig. 13 for two of the mea- 
surement positions. The axial location of the sightlines 
is given on the figure in terms of the distance from the 
surface of the test article. Again, this will make it dif- 
ficult to draw conclusions about the state of the flow 
from temperatures derived from the emission spectra. 

Distance along the line of sight, m 

Fig. 14. number density along the line-of-sight for 
the high pressure case. 

Computed Nj number densities are shown in Fig. 14 as 
a function of distance along the line-of-sight at three 
different axial measurement locations, two of which 
correspond to the locations of the temperature profiles 
in the previous figure. As with temperatures, there is 
a significant increase in N* number density near the 
shock front. For the two flow properties that govern 
emission from Nj, the shapes of the spanwise pro- 
files are predicted to be far from the idealized top hat- 
distribution that is assumed to exist in the spectral 

Comparisons With Measurements - Despite the 
indication that strong spatial gradients would com- 
promise the derivation of flow properties from the 
measured intensities, values of line-of-sight integrated 
rotational temperature, vibrational temperature, and 
number density values were extracted from the spec- 
tral data. To make a meaningful comparison, the cal- 
culated flow field emission was averaged in the same 
manner as the measured intensity. The approach used 
to derive these values from the flow property and emis- 
sion calculations is described at length in Ref. 47. 

Comparisons of the measured and calculated line- 
of-sight (LOS) averaged temperatures are shown in 
Fig. 15. Measured and calculated vibrational and ro- 
tational temperatures are shown as a function of the 
normalized distance from the surface of the test arti- 
cle. Despite the predicted strength of the shock- front 
region emission, the computed, LOS- averaged temper- 
atures still appear to nearly converge near the test ar- 
ticle surface. Compared to the measured temperature 
values, the computed LOS averaged values appear to 
approach convergence faster and to a greater degree. 
Although the overall trends appear to be consistent 
between the measured and calculated values, the two 
sets of LOS-averaged temperatures do not agree. In 
view of this disparity and because the measured tem- 
perature values do not converge, the existence of an 


equilibrated flow region could not be ascertained from 
the comparison. 

Fig. 15. Predicted and measured line-of-sight averaged 
temperatures for the high pressure case. 

By assuming that the vibrational temperature deter- 
mined from the spectral analysis of the measured in- 
tensities represented the electronic temperature, val- 
ues of LOS-integrated N2+ number density could be 
determined. As was done for temperature, an ap- 
proach for deriving a comparable quantity from the 
calculated flow properties was also developed. 71 Mea- 
sured and calculated values of the LOS-integrated N* 
number density are compared in Fig. 16. Some of the 
disagreement between the two sets of number densi- 
ties can be attributed to differences in spatial gradi- 
ents along the line-of-sight. Until the spatial gradient 
effects are investigated experimentally, it is not pos- 
sible to determine whether the difference seen in the 
comparison nearer the test- article surface is caused by 
spatial averaging from the optical system or by inac- 
curate modeling of the dissociative recombination 

Fig. 16. Predicted and measured LOS-integrated 
number density for the high pressure case. 

Based on the comparisons between the measured and 
calculated flow properties above, it was not possible to 
determine whether an equilibrated flow region exists 
within the shock layer for these conditions. Such a de- 
termination clearly requires an approach that resolves 
the spatial intensity gradients to extract information 
from the relevant flow region in the core of the shock 
layer. Once that is done, then the impact of other 
assumptions can be examined, and the processes that 
lead to equilibration can be investigated. Although 
the question of an equilibrated region was not conclu- 
sively resolved, it was still possible to address whether 
or not emission- based measurements could be used to 
determine the thermo chemical state of the flow. 

For these investigations, the experimental instrumen- 
tation did not actually measure flow properties. In- 
stead, flow property information was derived from an 
analysis of measured emissive intensities. To address 
the issue of using emission to evaluate the flow en- 
thalpy and to understand how the experimental ap- 
proach and the computational modeling might be im- 
proved, comparisons were made between calculated 
and measured spectral data at selected measurement 

Fig. 17. Measured and computed emission spectra for 
the 240 nm grating position, 20.7 mm upstream of test 

Comparisons between the measured and calculated 
emission spectra are presented below at several grating 
positions for a single axial location, 20.7 mm upstream 
of the test article. For the 240 nm grating position, the 
comparison is shown in Fig. 17. At this spectral loca- 
tion, the emission is mainly from NO 7 and S with 
probable contributions from the j3 and e systems. Ow- 
ing to the overlap of the emission from the different 
electronic states of NO, extracting temperature infor- 
mation from this spectral region is not feasible. Except 
for emission below 2100 A, the calculated intensity is 
less than the measured value. 

Fig. 18. Measured and computed emission spectra for 
the 310 nm grating position, 20.7 mm upstream of test 

A similar comparison of measured and calculated emis- 
sion spectra at the 310 nm grating position is shown 
in Fig. 18. The off-scale spectral features are Cu atom 


transitions. Copper is present in the stream because 
of electrode erosion, and it is not included in the com- 
putational model. Emission from molecular species in 
this spectral region is mainly from NO at shorter wave- 
lengths, and N 2 (2+) and N 2 (1-) systems at the longer 
wavelengths. The strongest emission peaks aside from 
those due to Cu emission are from N 2 (2+). Agree- 
ment between the calculations and the measurements 
is reasonable good in shape, but the overall signal level 
from the computational spectra appears to be low. Re- 
call that the calculated signal levels are exponentially 
dependent on the electronic temperature. If the calcu- 
lated electronic temperature, which is nearly 6000 K 
at this measurement location, was increased by 200 K, 
the N 2 (2+) emission would nearly double. 

Fig. 19. Measured and computed emission spectra for 
the 345 nm grating position, 20.7 mm upstream of test 

For the 345 nm grating position comparison, which is 
shown in Fig. 19, the emission is mainly from molec- 
ular species: N 2 (2+), N 2 (1-), and CN violet. The 
CN in the flow comes from dissociation of CO 2 that 
is present naturally in air and the subsequent recom- 
bination of C and N. Although it is truly a minor 
species, the transition strength is large and it is a sig- 
nificant emitter, as seen in the region near 3550 A 
where several of the measured peaks are not repro- 
duced by the calculated spectra. Cyanogen was not in- 
cluded as a species in the computational model. Some 
of the under-prediction of the intensity magnitude can 
therefore be attributed to the exclusion of CN from 
the calculation. For the N* (1-) emission, a 200 K 
increase in the electronic temperature would produce 
a roughly 20 % increase in the calculated intensity. 

Comparisons were also done for the 415 and 450 nm 
grating positions, which contained mostly molecular 
emission, and the agreement between the measured 
and calculated intensity magnitude is better, although 
the calculated levels are still low. The improved agree- 
ment for these grating positions is likely due to their 
use to guide the adjustment of the estimated stream 

The shock layer flow also contained significant O and 
N atom populations, and atomic transition intensi- 
ties were recorded at two near-IR grating positions. 
The measured and computed emission from O atomic 
transitions at 777 nm and 845 nm are compared in 
Figs. 20a and 20b. As with the grating positions at the 
shorter wavelengths, the calculated intensity is gener- 
ally lower than the measured intensity. At a calculated 

electronic temperature of 6000 K, an increase of 200 
K would nearly double the atomic emission. 









7700 7720 7740 7760 7780 7800 

Wavelength, A 


8400 8420 8440 8460 8480 



Wavelength, A 

Fig. 20. Measured and computed emission from the 
a) 777 nm O transitions and from the b) 845 nm O 
transition, 20.7 mm upstream of test article. 

Similar comparisons were done for N atom transitions 
and these are shown in Fig. 21a and 21b for the 744 
and 868 nm N transitions, respectively. The calculated 
emission for the N atom transitions is also low and 
because the emitting states are at energy levels that 
are similar to those of the O atom transitions above, an 
increase in the electronic temperature of 200 K would 
also result in a near doubling of the intensity for these 


Wavelength, A 


Fig. 21. Measured and computed emission from the 
a) 745 nm N transition and from the b) 868 nm N 
transition, 20.7 mm upstream of test article. 


Based on these spectral comparisons, it appears that 
the calculated emission was generally low for all grat- 
ing positions at this measurement location. Since the 
emission is exponentially dependent on temperature, 
it is likely that the calculated flow temperature was 
low. Again, the most likely culprit for this discrep- 
ancy is the enthalpy, which was probably not raised 
to the proper level. Clearly, the exponential depen- 
dence of the emission on temperature makes emission 
a very sensitive indicator of flow temperature. From 
an instrumentation development perspective, this im- 
plies that emission- based measurements have both the 
signal magnitude and sensitivity that are necessary 
to measure temperature, and ultimately flow enthalpy 
(with the assupmtion that velocity is negligible within 
the shock layer), reasonably well. However, the spa- 
tial gradients must be resolved for this approach to 

A different perspective on the LOS-integrated 
number density distribution that was presented in 
Fig. 16 can be obtained by comparing measured and 
calculated (1-) emission for a single grating posi- 
tion at each of the measurement locations on the cen- 
tral stagnation streamline. This comparison is shown 
for the 426 nm grating position in Fig. 22a for the 
back position of the test article, and in Fig. 22b for 
the forward position. There is an easily distinguish- 
able difference between the evolution of the measured 
signal and the calculated signal. By performing a di- 
rect comparison of measured and calculated emission 
spectra, uncertainties introduced in the analysis that 
was performed to derive flow properties from measured 
intensities are avoided. However, possible differences 
between measured and calculated flow property gradi- 
ents are still present and will influence the comparison. 
Of particular concern is the possibility of additional 
averaging of the measured intensities that may have 
been caused by the optical collection system, 65 This 
possibility has not been accounted for in these com- 
parisons, so only qualitative statements can be made 
regarding the differences between the calculated and 
measured spectra. 


b4=35.3 mm 

b3-33.2 mm 

b 1-27.3 mm 
Distance from 
model surface 


Fig. 22. Nj emission at 420 nm grating position for all 
of the axial measurement locations for the high pres- 
sure case: a.) back position; b.) forward position. 

In Figs. 22a and 22b, it is apparent that the measured 
rate of evolution and decline of the Nj emission as the 
flow goes from the shock to the test article is less rapid 
than predicted by the calculations. Although the lim- 
itations in the spatial resolution of the optical system 
preclude further statements about the axial distribu- 
tion of the Nj emission, the comparison does illustrate 
the possibility of using emission spectra to evaluate the 
population dynamics of important shock layer species. 

4.4 Low Pressure Case 

Starting Conditions - The lower pressure case test 
conditions were chosen to maximize the change in 
shock layer pressure, which was reduced by a factor of 
« 3. At the lower pressure, the collision frequency in 
the shock layer is reduced significantly. Consequently, 
the flow is less likely to be in thermal or chemical equi- 
librium. If the degree of departure from equilibrium 
could be determined at these test conditions, then 
progress could be made in defining the test conditions 
that lead to equilibration within the shock layer. 

As was found in the comparisons between the simu- 
lated and measured emission spectra for the high pres- 
sure case, the comparisons for the low pressure case 
indicated that the stream enthalpy value derived from 
the stagnation point heat flux and pressure measure- 
ments was probably low. Therefore, additional calcu- 
lations were performed at total enthalpy values that 
were 14 % and 32 % higher than the estimated values. 


In addition to the uncertainty in the stream enthalpy 
that was present for both the high and low pressure 
test cases, the uncertain argon mass flow rate became 
an issue for the simulations of the low pressure con- 
ditions. The total mass flow rate for the low pressure 
case was reduced by a factor of « 4, based on the re- 
duction in pressure, while the start and shield argon 
mass flow remained constant. For the high pressure 
case, the argon mass flow was assumed to be 5 % of the 
total mass flow. This implied that the relative argon 
mass flow could be 20 % of the total mass flow for the 
low pressure case. At this level, the uncertainty in the 
argon mass flow becomes more important because of 
its increased participation in the reaction kinetics. For 
example, in three-body recombination reactions, Ar is 
less efficient than N 2 as the third body. 61 To address 
this additional uncertainty and attempt to bound its 
influence, flow simulations were performed for three 
different argon mass fractions: 5 %, 10 %, and 30 %. 

Thus, owing to the uncertain starting conditions a to- 
tal of five different simulations of the shock layer flow 
for the low pressure case were computed. The start- 
ing conditions and computed free stream properties for 
each of these simulations are summarized in Table 2. 
At nearly constant enthalpy, increased argon mass flow 
is seen to increase the temperatures and free stream 
velocity slightly, while increasing the dissociation frac- 
tion for nitrogen substantially. Comparing cases that 
have the same argon mass fraction, increasing the total 
enthalpy produces results that are similar to increasing 
argon mass fraction at constant enthalpy. This uncer- 
tainty in the starting conditions clearly create difficul- 
ties for comparisons of simulations and experimental 

Table 2. Starting conditions and free stream proper- 
ties for the low pressure test case simulations 

Parameter Case 1 Case 2 Case 3 Case 4 Case 5 

po, atm 












w Ar 






«oo, km/s 






Poo, Pa 






Too, K 






T v 00 , K 






Wn 2 






w N 












Calculated Shock-Layer Properties - Using the 
Case 3 conditions, axial profiles of temperatures and 
pressure for the shock layer flow were computed, and 
these are shown in Fig. 23 as a function of the nor- 
malized distance from the test article surface. While 
the rotational and translational temperatures are still 
higher near the shock, the increase over the levels 
nearer the test article is not as great as was seen for 

the high pressure case (see Fig. 11). At the lower 
shock layer pressure, the vibrational and translational- 
rotational temperatures do not appear to converge 
outside of the boundary layer of the test article. This 
is in contrast to the results of the simulation for the 
high pressure case, where the temperatures clearly 
converged as the model surface was approached. Ac- 
cording to the simulation, the shock layer is not in 
thermal equilibrium, except within the boundary layer 
at these simulated conditions. 

Fig. 23. Axial profiles of temperature and pressure in 
the shock layer. 

Fig. 24. Axial profiles of species populations in shock 

For the same starting conditions, the axial profiles of 
the neutral species and Nj are plotted in Fig. 24, again 
as a function of the nondimension al distance from the 
surface. As was seen in the high pressure case, there 
appears to be a significant peak in the Nj concen- 
tration near the shock front. However, for the low 
pressure conditions of this simulation, the mass frac- 
tion near the shock front is only « 5 times higher than 
the mass fraction nearer the test article surface (as 
opposed to « 50 times for the high pressure case, see 


Fig. 12). The mass fractions of N and N 2 do not reach 
a limiting value as the surface is approached, so the 
simulation indicates that the shock layer is also not in 
chemical equilibrium. 


Distance along the line of sight, m 


Distance along the line of sight, m 

Fig. 25. Computed flow property profiles for Case 3 
conditions at selected axial locations for a) tempera- 
tures and b) N2+ number density. 

Although the starting enthalpy and argon mass flow 
values were less certain for the low pressure case, the 
gradients in the Nj and temperature profiles appeared 
to be less severe than found for the high pressure case. 
To assess the spatial gradients along the optical sight 
lines at the measurement locations, temperature and 
species profiles were extracted from the shock layer so- 
lutions, and these are shown for selected axial locations 
in Figs. 25a and 25b, respectively. For both temper- 
atures and N t number density, the computed profiles 
along the lines-of-sight are much closer to the ideal- 
ized top-hat distributions that are required to derive 
temperatures that are representative of the core flow 
region from the spectral analysis of the measured in- 
tensities. Unfortunately, it appears that the majority 
of the shock layer flow is likely to be in nonequilib- 
rium, which may violate the other major assumption 

of the spectral analysis. The degree of departure from 
equilibrium and its impact on the distributions of pop- 
ulations over the various energy levels is difficult to 

Comparisons With Measurements - As with the 
high pressure case, rotational and vibrational tem- 
perature values were derived from an analysis of 
spectral features. 66 Using an intensity- weighted aver- 
aging approach, 47 temperature values that could be 
compared with the experimental values were extracted 
from the computed flow properties at the axial mea- 
surement locations for some of the different simula- 
tion cases. Comparisons between the computed and 
measured LOS-averaged temperatures are shown in 
Figs. 26 and 27, for the simulation conditions of Case 
3 and Case 4, respectively. For these two cases, the 
enthalpy levels were in the middle of the range of sim- 
ulations and were nearly in agreement. However, for 
Case 3 the argon mass fraction was 0.05, while for Case 
4, the argon mass fraction was 0.3. In Fig. 26, for the 
Case 3 condition, there are significant differences in 
both trends and magnitudes between the computed 
and measured temperatures. First, for the experimen- 
tal values, the rotational and vibrational temperatures 
do not appear to overlap, except perhaps accidentally 
at one or two measurement locations. Owing to low 
signal levels near the shock front, the measured val- 
ues are highly uncertain. Consequently, the discussion 
of trends will be restricted to the positions nearer to 
the surface than x/R — -0.4. For those locations, the 
trends in the computed and measured vibrational tem- 
peratures appear to be reasonably similar, although 
the measured values are generally greater in magni- 
tude. In contrast, the measured rotational tempera- 
tures do not appear to decrease significantly from the 
values near the shock front, while the computed rota- 
tional temperatures clearly show evidence of relaxation 
going toward the test article. 

Fig. 26. Computed and measured LOS-averaged tem- 
peratures for the conditions of Case 3 (ho = 17.15 
MJ/kg, w Ar = 0.05). 


Fig. 27. Computed and measured LOS- aver aged tem- 
peratures for the conditions of Case 3 (ho = 17.26 
MJ/kg, 10 Ar = 0.30). 

At the high argon mass flow conditions of Case 4, the 
comparison between measured and computed L OS- 
averaged vibrational temperatures shows improved 
agreement on the magnitude, while maintaining rea- 
sonable agreement on the shape of the distribution 
at locations closer than x/R = -0.4. The agreement 
between magnitudes of the computed and measured 
LOS- averaged rotational temperatures also appears to 
be improved with the increased argon mass flow, but 
the difference between the distributions is unaffected. 
It should be noted that the influence of streamwise 
and spanwise spatial averaging by the optical collec- 
tion system on the measured values has not been fully 
accounted for in these comparisons. Considering the 
uncertainty in the starting conditions for the simula- 
tions and the uncertainty in the unresolved spatial in- 
tensity gradients for the experiment, the general lack 
of agreement is not surprising. 

4.5 Spatially Resolved Measurements 

In a recent set of experiments conducted at the low 
pressure test conditions, Park acquired emission spec- 
tra from multiple locations along the spanwise direc- 
tion as a single axial position within the shock layer. 71 
Several separate emission measurements were recorded 
simultaneously by the spectrograph and CCD system, 
and a series of adjusted collection mirror positions were 
used to cover the radial extent of the shock layer dur- 
ing a single facility run. An Abel-inversion was then 
used to obtain spatially resolved emission spectra from 
the LOS-integrated intensities. Finally, temperatures 
were derived from the Abel-inverted spectra using an 
analytical method that involved ratios of N* spectral 
features and ratios of 0 atom transitions. A fuller 
description of the experiment and the analytical ap- 
proach is given in Ref. 71. 

Radial distributions of the rotational, vibrational, and 
electronic temperatures that were derived from the 

spectrally resolved emission are shown in Fig. 28. The 
rotational and vibrational temperatures were deter- 
mined using the same analytical approach that was 
used to derive temperatures from the line-of-sight in- 
tensities, above. As explained in Ref. 71, two sets of 
atomic oxygen transitions were used to calculate elec- 
tronic temperature. Thus, the two electronic temper- 
ature distributions are labeled by the shorter wave- 
length transition used in each intensity ratio. Ow- 
ing to an unresolved background contribution at the 
8446.5 A transition, electronic temperatures derived 
from the intensity of that transition are systematically 
low. Electronic temperature values obtained using the 
7773.4 A transition are believed to be valid. 


. 6500 



£ 6000 




,CD 5500 



0 2 4 6 8 10 

Distance from the center, cm 

Fig. 28. Radial temperature profiles in the shock layer 
at 12.7 mm upstream from the test article. Labels for 
T e indicate the shorter wavelength 0 transition of the 

The most striking aspect of the radial temperature dis- 
tribution shown in Fig. 28 is that the rotational, vibra- 
tional, and electronic (for the 7773.4 A pair) tempera- 
tures appear to overlap within their respective uncer- 
tainties. This may indicate a region of thermal equilib- 
rium within the shock layer for the low pressure condi- 
tions, contrary to the prediction of the computational 
simulation. The shapes of the rotational and vibra- 
tional temperature distributions are reasonably similar 
to the shapes of the predicted radial temperature dis- 
tributions at 9 mm that were shown in Fig. 25a. (Note 
that the 9 mm axial location of the predictions was 
closest to the 12. 7 mm axial position of the measure- 
ments.) The temperatures and the emission spectra 
from these experiments are still being analyzed 

4.6 Lessons Learned 

First, and foremost, the futility of attempting to per- 
form detailed computational simulations of arcjet flows 
without adequate specification of the starting condi- 
tions, mainly the enthalpy, has to be recognized. Un- 



L L L LMff, 

- k Te_7773.4 

I I I — 


til this issue is resolved, the knowledge gained from 
performing combined experimental and computational 
investigations will be marginal. The limitations of us- 
ing stagnation point heat flux and impact pressure 
measurements to estimate enthalpy were clearly illus- 
trated. Although the unspecified argon mass flow rate 
did affect the low pressure test simulations, this inflow 
parameter is a more tractable problem. Most large- 
scale arcjet facilities routinely measure the mass flows 
of the test gases, and the Ames Research Center Arcjet 
Facilities have recently been equipped with improved 
mass flow sensors and control capability. 

The second important lesson to derive from this ex- 
ercise concerns the use of diagnostic techniques and 
approaches that do not resolve spatial gradients. At- 
tempts to compare simulation predictions with mea- 
surements of spatially integrated quantities can lead to 
misleading conclusions. For the low pressure case, the 
LOS-averaged temperature distributions suggest that 
the flow is not in thermal equilibrium, while the radial 
distribution of spatially resolved temperatures suggest 
the opposite. An investment of additional effort into 
acquiring emission spectra in the radial direction to 
obtain Abel- inverted intensities has a far better (and 
more certain) return than deriving the comparable in- 
tegrated flow properties from a number of computa- 
tional simulations. It was fortunate that the com- 
parisons between the predicted and measured L OS- 
averaged temperatures agreed so poorly for the high 
and low pressure cases; otherwise, the temptation to 
“correct” the measurements using the computational 
results may have proven overwhelming. 

Finally, the possibility that a portion of the shock layer 
is in thermal equilibrium at the low pressure test condi- 
tions contradicts the computational predictions, which 
showed extensive thermal nonequilibrium for all of the 
different low pressure simulations. Further knowledge 
of the chemical state of the shock layer, which is cur- 
rently under investigation, and verification of equilib- 
rium would provide much- needed insight into the na- 
ture of the shock layer flow. 

5. Summary and Recommendations 

The question of the state of CFD simulations of arc- 
jet flows is still dominated by the lack of knowledge 
about the flow enthalpy. Any other consideration is 
secondary. Conventional methods for estimating the 
flow enthalpy, including energy balance, sonic flow, 
and stagnation point heat transfer, all provide insuffi- 
cient specification of the flow enthalpy for simulation 
purposes. If adequate resources and dedicated effort 
are brought to bear on this problem, then eventually 
it will be resolved and arcjet facility simulations will 
become much more meaningful. New LIF-based ap- 
proaches to enthalpy measurement may improve this 
situation. 58 ’ 59 

Although this premise cannot be rigorously tested un- 
til enthalpy can be more accurately determined, it 
appears that current nonequilibrium, hypersonic-flow 
computational models are able to provide reasonable 
simulations of arcjet flows. This observation is based 
on a qualitative assessment of the comparisons be- 
tween the shock layer measurements and predictions, 
combined with the fact that no obvious shortcomings 
in the flow models could be identified. Despite hav- 
ing to estimate the enthalpy and use two CFD models 
and a radiative transport code to predict intensity, the 
comparisons with measured values were generally fa- 
vorable; particularly for the high pressure case. 

While the flow enthalpy must be accurately specified to 
enable detailed comparisons between simulations and 
measurements for a single arcjet facility test condition, 
measurements and simulations of relative trends in ar- 
cjet characteristics are not similarly constrained. Mea- 
surements and simulations of the response of sensitive 
flow properties to changes in arcjet control parameters 
during a single facility test can add substantial infor- 
mation to the knowledge base at the present time. The 
approach to this involves using diagnostic instrumen- 
tation to monitor stream parameters, when conditions 
have stabilized, as a single control variable, such as 
the arc current, is changed. (Examples of this type 
of experiment will be discussed extensively in the sec- 
ond lecture.) Making the best possible estimate of the 
total enthalpy for a single condition, a computational 
simulation is essentially calibrated at that condition. 
Subsequent conditions are then simulated using fur- 
ther estimates of the total enthalpy, without chang- 
ing the other parameters of the model. Comparisons 
are then made between the measured and predicted 
trends. This approach avoids the larger uncertainties 
that pertain to measurements of absolute quantities. 

Some of the more obvious parametric studies to per- 
form include: 1.) varying pressure to assess impact on 
chemistry; 2.) varying the arc current, which varies 
the initial ionization level; and 3.) varying the test 
gas composition to investigate relative third-body ef- 
ficiencies in N 2 recombination. 

To establish the validity of diagnostic approaches, 
comparisons between measurements made using multi- 
ple independent instruments would be extremely use- 
ful. This observation is particularly relevant to de- 
veloping new approaches for determining the flow en- 
thalpy, which is the most important parameter to mea- 
sure accurately, and which ultimately determines how 
useful and relevant arcjet testing will become. 

At present, it is too early to propose code validation 
experiments for large-scale arcjet facility flows. How- 
ever, it is appropriate to begin thinking about how to 
develop diagnostics and strategies that may eventually 
enable code validation experiments in these facilities. 
The most important advantage that arcjets have over 


impulse facilities is the test duration. Steady state 
flow conditions and material response can be achieved 
and documented. The long test time allows for mul- 
tiple property measurements that can be temporally 
averaged. Because there is no diaphragm, these mea- 
surements can be repeated in multiple runs to assess 
facility repeatability with a rapid turn-around time be- 
tween tests. 

Test guidelines for validation experiments have been 
suggested by Mehta in his descriptions of methods for 
producing credible computational simulations. 72 Per- 
forming validation experiments in arcjet facilities will 
require considerable maturation of currently available 
diagnostic techniques to ensure adequate specification 
of inflow conditions for either nozzle or shock-layer flow 
simulations. In addition to improving the accuracy of 
inflow condition measurements, sufficient spatial cov- 
erage must be attained to allow approximate integra- 
tion of the flow properties for comparison with other 
measurements of mass flow, energy balance, etc. With- 
out this type of internal accuracy check of the experi- 
mental results, confidence in the measurements would 
not be sufficient to motivate extensive computational 
simulations or efforts to improve physical models. 

6. Acknowledgements 

Many colleagues have contributed to the lecture ma- 
terial contained herein, but special thanks are due to 
Tahir Gokgen and Chung Park of Thermosciences In- 
stitute (Eloret), Mark Newfield of Ames Research Cen- 
ter, and James Donohue of UTRC for their excellent 
work on the combined computational and experimen- 
tal investigation of arcjet flows. John Balboni of the 
Thermophysics Facilities Branch at Ames Research 
Center provided many useful references on arcjet fa- 
cilities in general. The entire staff of the Aerodynamic 
Heating Facility Arcjet provided vital test support and 
Frank Hui’s efforts as test engineer deserve special 
recognition. Many of the thoughts expressed in this 
document resulted from conversations with colleagues 
about arcjet flows and testing. While any erroneous 
statements are attributable solely to the author, the 
contributions of Chul Park, Joan Pallix, Raj Venkat ap- 
athy, Ellis Whiting, and Jochen Marschall of Thermo- 
sciences Institute (Eloret) to what has been written are 
greatly appreciated. Similar conversations with Paul 
Kolodziej, Jeff Bull, Dave Stewart, Joe Hartman, Joe 
Olejniczak, Dave Olynick, Stephanie Langhoff, Sur- 
rendra Sharma, and George Raiche of NASA Ames 
Research Center have also proven to be very helpful in 
writing this document. 

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