AIAA 2009-6438
AIAA SPACE 2009 Conference & Exposition
14 - 17 September 2009, Pasadena, California
An Airbag-Based Crew Impact Attenuation System Concept
for the Orion CEV – First Generation System Development
Sydney Do*, Olivier de Weck†, and Ricardo Robles Jr.‡
Department of Aeronautics and Astronautics, Massachusetts Institute of Technology, Cambridge, MA, 02139
and
Joseph Pellicciotti§ and Tim Brady¶
NASA Engineering Safety Center, Hampton, VA, 23681
Airbag-based methods for crew impact attenuation have been highlighted as a potential
means of easing the mass constraints currently affecting the Orion Crew Exploration
Vehicle. An analog airbag test article has been designed, fabricated, and subjected to a drop
test campaign to investigate the potential performance of such a system. Through a fusion of
accelerometer measurements and photogrammetric analysis of high-speed camera data; it
was found that the system performed adequately up to impact velocities of 4.7m/s, where a
low-injury risk was maintained throughout the impacting event. This paper presents the
background and development behind this first generation system, as well as the results and
lessons learnt from this preliminary drop test campaign.
I. Introduction
O
VER the past two years, increasing mass constraints have forced multiple downgrades to the originally
baselined performance of the Orion Crew Exploration Vehicle (CEV). In late 2007; in an attempt to reduce the
vehicle’s then overall mass to within its predefined mass allocation, a decision was made to change the vehicle’s
baseline landing profile of landing on land to the water-based “splashdown” concept employed during the Apollo
era1,2. In early 2008, it became apparent that thrust oscillations occurring in the Ares-I launch vehicle could
potentially induce dangerously high dynamic G-force loads in the Orion crew module during launch3. Currently, this
issue remains unresolved, with all available solutions requiring the addition of mass to the system4. More recently,
NASA has announced that the first generation of the Orion CEV will be baselined to carry four crew members to the
International Space Station, rather than the originally specified crew of six5,6. Like the decision to move to nominal
water landings, this performance decrement has been attributed to continuing difficulties in keeping the overall
system mass to within its mass allocation.
Hence, it appears evident that the only way to regain this lost system capability is by implementing lightweight,
synergetic concepts which satisfactorily meet or exceed system performance and safety requirements. One area
where this may be achieved is the Crew Impact Attenuation System (CIAS), along with its accompanying seats. As
the name suggests, the function of this system is to attenuate all impact loads subjected to the crew throughout the
mission to within tolerable levels of injury-risk. Currently, its design consists of crew seats mounted to a strutsupported pallet, cushioned underneath by crushable materials, as shown in Fig. 1b).
*
Graduate Research Assistant, Department of Aeronautics and Astronautics, Massachusetts Institute of Technology,
Cambridge, MA, 02139, and AIAA Student Member
†
Associate Professor, Department of Aeronautics and Astronautics & Department of Engineering Systems,
Massachusetts Institute of Technology, Cambridge, MA, 02139, and AIAA Associate Fellow
‡
Undergraduate Student, Department of Aeronautics and Astronautics, Massachusetts Institute of Technology,
Cambridge, MA, 02139, and AIAA Non-Member
§
NASA Technical Fellow, NASA Engineering and Safety Center Mechanical Systems, NASA Goddard Space
Flight Center, Greenbelt, MD, 20771, and AIAA Non-Member
¶
Engineer, NASA Engineering and Safety Center Systems Engineering Office, NASA Johnson Space Center,
Houston, TX, 77058, and AIAA Non-Member
1
American Institute of Aeronautics and Astronautics
Copyright © 2009 by the American Institute of Aeronautics and Astronautics, Inc. All rights reserved.
To further investigate alternative design
concepts for the CIAS, the NASA Engineering
and Safety Center (NESC) has been requested
to review the baseline design of the Orion
vehicle’s landing attenuation system and to
recommend potential design alternatives.
During a workshop on innovative engineering
conducted by the NESC in the summer of
2008, the idea of personal airbag systems was
first conceptualized. Inspired from the
structure of seeds in nature and their ability to
protect their embryos from mechanical loads;
this concept involves using an inflated airbag Figure 1. a). Orion CEV Baseline Configuration
“seat” to protect the occupant from the
b). Orion Baseline Crew Impact Attenuation System
anticipated impact loads induced during
landings of the Orion crew module. The ideation process that led to this final concept is summarized in Fig. 2.
1
Orion
Descent
2
Current Crew
3 Configuration
4
Personal “Bubbles”
Maintain Relative
Position of Crew
5
Cross Strapping
Personal Airbag
Seat Concept
Adjustable Size
Figure 2 – Personal Airbag System Ideation Process7
The main advantage of such a configuration is its significantly lower mass relative to the existing CIAS and seat
design, whilst potentially being able to attenuate impact loads to a low injury-risk level. An additional benefit
includes the ability of the system to be deflated and stowed away, hence providing additional in-cabin volume once
the spacecraft is in orbit. Initial estimates have found that these savings equate to a potential 36% reduction in CIAS
mass without the crew, and an increase of 26% of in-orbit habitable volume7.
Consequently, in an effort to further develop this concept and investigate its load attenuation potential; a first
generation airbag-based CIAS has been designed, fabricated, and tested. It is the purpose of this paper to present this
initial system, describe the testing campaign undertaken, and discuss the related results and their implications on the
overall landing performance of the Orion vehicle.
II. System Design
During the initial stages of system design, it became apparent that the mass-optimal system would not only
protect the crew from the various dynamic loads imposed upon them during a mission; but also perform the role of
supporting the crew during launch and reentry. To accomplish this latter function, the system would have to meet
NASA’s anthropometric requirements (ie. accommodate body types ranging from approximately a 5th percentile
Japanese female to a 95th percentile American male)8, whilst also accounting for spinal elongation of the crew after
long-term exposure to a micro-gravity environment9. Hence, the potential for system reconfigurability, as well as
landing attenuation capability, would be the key driving metrics used to define a baseline system configuration. The
final result of this effort yielded a two-tiered system definition, consisting of a top level operational concept, along
with a baseline airbag-seat layout.
As can be seen in Fig. 3a), this final operational concept consists of three distinct phases, being: an inflated
configuration, a stowed configuration, and a deflated configuration. Over a typical mission, the system would
transition through each of these phases in the following manner:
- During pre-launch activities, the system would be in an inflated configuration so that it could function as a seat to
support the occupant as the final stages of launch vehicle fuelling and checkout were being performed
- During launch, the system would remain inflated to continue supporting its occupant
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American Institute of Aeronautics and Astronautics
-
-
Once in space, where support of the seat occupant was no longer required, the system would be deflated and
stowed away to increase the available cabin working space. This configuration is referred to as the stowed
configuration
On completion of the mission and prior to Earth reentry, the system would be reinflated to again support the crew
during the descent and subsequent landing phases
On landing, the seat would deflate as the spacecraft impacted with the landing surface, attenuating the dynamic
loads imposed upon the occupant. More details regarding the impact mechanics of airbags will be discussed in
Section II-B.
2Stowed Configuration
Freed up cabin space
Floor panel cover
Deflated airbag system
1
Reentry
Launch
Orion Floor
Inflated Configuration
3Deflated Configuration
Orion Floor
Head
Airbag
Hard point for
system attachment
Orion Floor
Lower Back
Airbag
Upper Back
Airbag
a).
Lower Leg
Airbag
Thigh
Airbag
b).
Figure 3. a). Top Level Airbag-based CIAS Concept b). Baseline Airbag Seat Layout
With regard to the baseline layout of the system, a multi-airbag arrangement was chosen to improve system
stability upon impact, by mitigating the gas-shifting effects which would be experienced in a single airbag
configuration. This arrangement would also be much easier to manufacture, as well as provide an inherent degree of
redundancy within the system by maintaining some functionality if one airbag was to fail. Additionally, using
multiple airbags would have a comparable mass penalty to that of a single airbag configuration10, whilst also
allowing for the introduction of an additional degree of reconfigurability in the system via the ability to vary the
design of each of the airbags.
As an initial estimate of this multi-airbag arrangement, airbags were placed under key mass concentrations in the
human body when in the recumbent position – the safest position for an astronaut during an impact-type landing
event. This is depicted in Fig. 3b), where airbags are situated to support the head, upper and lower back, thigh, and
lower leg. Also shown in Fig. 3b) is the fact that these airbags were chosen to initially be cylindrical in shape. This
decision was based primarily on their ease of manufacture, as well as their common use in load attenuating
applications.
It should be noted that the methodology employed for this development effort consists of multiple design-buildtest (DBT) cycles, where lessons learnt from previous testing cycles will be brought forward into the next
generation’s design. As a consequence, several design decisions made for the first generation test article were based
on practical issues such as manufacturability, cost, and time. This will become more evident in subsequent sections
of this paper. Naturally, however, as in any development effort; these design decisions will be accompanied by
increasingly detailed analysis and increased willingness to invest in more costly development as the DBT cycles
progress and more is understood about the system.
A. System Performance Requirements
To ensure that the Orion CEV, along with other Constellation vehicles, are capable of accommodating all of the
safety and basic comfort needs of their crew during all phases of flight, NASA has released the “Constellation
Program Human-Systems Integration Requirements” (HSIR) document9. This document lists a set of requirements
which must be met by a spacecraft for it to be considered human-rated.
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In relation to crew impact attenuation systems, the primary metric specified
X
by the HSIR for gauging system performance is the Brinkley Direct Response
Index (DRI). This index measures the risk of injury to an occupant given a
measured acceleration profile by comparing the output of a dynamics model of
the human body, with limiting values representing varying levels of risk to
injury (Tables 2 and 3). When originally composed, the NASA HSIR dictated
that a system must maintain the level of injury-risk to its occupant in the “very
low” range throughout a transient acceleration event for it to be considered
human-rated. Recently however, this requirement has been relaxed to a “low” Y
limiting injury risk level, as will be later explained.
Z
Figure 4. Brinkley Direct
With regard to the dynamics model used to determine the Brinkley DRI, a
Response Index Model11
lumped parameter model is utilized. This model approximates the dynamic
response of a human as the response of a spring-mass-damper system to a given acceleration profile in each of the
three orthogonal axes, referenced to the center of the torso. A simplifying assumption made is that the effects of the
applied acceleration profile in each of the three axes are uncoupled. This dynamic system is modeled with the
following relationship:
X (t ) 2[Z n X (t ) Z n2 X (t )
(1)
A(t )
Where:
X is the relative acceleration of the dynamic system with respect to the reference point (or center of the torso) in
either one of the x-, y-, or z-directions
X is the relative velocity of the dynamic system with respect to the reference point in either one of the x-, y-, or zdirections
X is the relative displacement of the of the dynamic system with respect to the reference point in either one of the
x-, y-, or z-directions. Here, a positive value corresponds to a compression
A is the measured acceleration profile from the reference point in either one of the x-, y-, or z-directions
[ is the damping ratio of the dynamic system representing the response in the given x-, y-, or z-direction
Z n is the natural frequency of the dynamic system representing the response in the given x-, y-, or z-direction; and
t
is a time coordinate
The Brinkley DRI is obtained by solving the system given by Eq. (1) and inputting the result into the following
relationship:
DRI (t )
Z n2 X (t )
(2)
g
Where g is the acceleration due to the Earth’s gravity, used here as a normalizing factor.
Furthermore, the damping ratio and natural frequency values to be used in the aforementioned lumped parameter
model as specified by the NASA HSIR are as follows:
Table 1 – NASA HSIR specified natural frequencies and damping ratios to be used in the Brinkley Dynamic
Response Model9
x
y
z
Zn
62.8
58.0
52.9
[
0.2
0.09
0.224
Table 2 – NASA HSIR Specified Brinkley DRI Limits9
x
Brinkley DRI Limit Level
DRIx < 0
DRIx > 0
-22.4
31
Very Low (Previous Nominal)
-28
35
Low (Current Nominal)
y
DRIy < 0
-11.8
-14
z
DRIy > 0
11.8
14
DRIz < 0
-11
-13.4
DRIz > 0
13.1
15.2
As was mentioned earlier, NASA has recently changed its maximum acceptable injury-risk level requirement
from “very low” to “low”. This decision was based on recommendations made by an independent group of industry,
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American Institute of Aeronautics and Astronautics
academic, and NASA experts tasked to determine if the HSIR specified occupant protection criteria were too
conservative. It was realized that the “very low” Brinkley limit requirement was originally mandated so that
deconditioned crew members could be safely returned to the ground. This capability was deemed as not necessary
for Constellation Program vehicles during the original specification of the system requirements, hence allowing for
the initial Brinkley limit to be relaxed to a “low” injury-risk level1.
In addition to the DRI limit values specified by the HSIR, limiting values corresponding to higher levels of
injury-risk have also been derived for use in evaluating the safety of systems subjected to high transient
accelerations. These were obtained from the original Brinkley study, which focused on determining human
acceleration tolerance measures for use in aircraft ejection seat design; and are given below in Table 3.
Table 3 – Brinkley Limits for Moderate and High Risks of Injury12
x
y
Brinkley DRI Limit Level
DRIx < 0
DRIx > 0
DRIy < 0
DRIy > 0
-35
40
-20
17
Moderate
-46
46
-30
22
High
z
DRIz < 0
-12
-15
DRIz > 0
18
22.8
Since the DRI is a time-dependent function, a parameter called the ȕ-function is commonly used to determine if
a system has remained below a given Brinkley limit over the duration of an acceleration event. As shown in Eq. 3,
this function corresponds to the root sum square of the relative DRI values in each of the three orthogonal axes.
max^E `
­
°
max ®
°
¯
2
2
2 ½
§ DR x (t ) · § DR x (t ) · § DR x (t ) · °
¨
¸ ¨
¸ ¨
¸ ¾ 1
¨ DR lim ¸ ¨ DR lim ¸ ¨ DR lim ¸
x
x
x
©
¹ ©
¹ ©
¹ °
¿
(3)
Here, a maximum beta value of less than one corresponds to the system satisfying a given injury-risk level over
the duration of an impacting event.
B. Past Airbag Applications
Airbags attenuate impact loads by converting the kinetic energy of an impacting event into the potential energy
of a compressed gas, and then venting this gas to remove energy from the system. Depending on the amount of gas
vented, the system will either experience a bounce after the initial impact or come to an immediate rest. Because of
this seemingly straightforward and elegant process, airbags have been consistently considered as viable options for
landing attenuation systems in spacecraft since the early 1960’s, when the need for safely returning a vehicle from
space first became apparent.
This is evident in NASA’s first human spaceflight program,
Project Mercury, where a passive airbag installed in a skirt
between the heat shield and the capsule was used to attenuate
landing loads. This system relied on air being drawn through
holes in the skirt after heat shield separation, and filling the
airbag. On impact, the drawn air would be forced back out
through the holes in which they entered, removing some of the
kinetic energy from the system13.
Airbags were also prominent in the early days of Soviet
space exploration, with the Luna 9 lunar lander in 1966 using
airbags to achieve the first ever soft landing of a human-made
vehicle on an extraterrestrial planetary body. This system
involved completely enveloping the spacecraft in a protective
Passive
airbag layer moments before impact. Upon initial contact, the Figure 5. Project Mercury
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Airbag
System
gas inside the airbag compressed and expanded. This caused
the entire system to experience several bounces before finally coming to rest16. It should be noted that unlike the
Mercury system, no means of gas expulsion was built into this airbag system. This meant that the only method in
which to dissipate the kinetic energy of the initial impact was through heat and frictional losses occurring during
each subsequent impact, where kinetic energy would be converted into potential energy and back into kinetic energy
for the next bounce.
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This “bouncy-ball” type landing has since been used
on all spacecraft containing airbag-based landing
attenuation systems, including Luna 9’s follow-up
mission, the Luna 13; as well as NASA’s Mars
Pathfinder and its famed cousins, the twin Mars
Exploration Rovers. The only exception to this is the
European Space Agency proposed ExoMars lander,
currently planned for launch in 2016. Unlike the
previously mentioned spacecraft, ExoMars will use a
vented airbag system to protect it in a manner similar to
that of the Mercury spacecraft. Here, the airbag system
will be designed to dissipate the vehicle’s kinetic energy
via expulsion of compressed gas upon impact, resulting
Figure 6. Luna 9 and 13 Entry, Descent, and Landing
in a stuck, “dead” type landing.
Sequence15
In addition to spacecraft landing load attenuation,
airbags have also gained popularity in a variety of
terrestrial applications over the second half of the last
century. These have appeared most commonly in the
form of steering wheel and side skirt mounted airbags
within automobiles; and within protective garments, as
seen in Fig. 7. Because the likelihood of impact events
occurring in these applications is usually uncertain; a).
b).
these systems commonly come in active forms, Figure 7. a). Automobile Airbag System7
requiring a fast initiation triggered by a predetermined
b). Garment-based Airbag System7
set of dynamic criteria. For reentering spacecraft however, the onset of an impact is known well beforehand, and
hence a rapid inflation is not necessary. As a consequence, such active systems were deemed impractical for
incorporation into the proposed CIAS. This is because these terrestrial systems are designed to operate in a different
environment, and hence add little improvement to the overall performance of the proposed system.
C. Airbag Detailed Design
It is obvious that for human-rated impact attenuation systems, the aforementioned “dead” type landing profile is
preferred over the more commonly used “bouncy-ball” type due to reasons of both comfort and safety. Because of
this, airbags for the proposed system will require venting profiles and orifices designed to completely dissipate the
kinetic energy of the occupant on impact; such that no bouncing occurs and the resulting Brinkley DRI remains
beneath the “low” injury-risk level limit. Coupled to this is the choice of operating gas, the initial pressure of this
gas, and airbag geometry; parameters which all play a role in the energy conversion processes which occur during
impacting events.
To determine these parameters for the first generation airbag system, a one degree-of-freedom model based on
Refs. 17 and 18 was created to allow for the generation of quick, first-order accurate results; thus allowing for a
preliminary parameter optimization to be performed. Specifically, this involved determining parameters for each of
the five airbags by discretizing a 50th percentile male aviator human model into five key segments, corresponding to
the regions specified in Section II. With this, a range of parameters was input into the model for each airbag to
develop a design space over which to optimize over. These were obtained from spacing constraints about the human
body as well as initial simulation runs with the dynamics model, and are given in Table 4. Note that as a first
approximation, these values were obtained for the system with a zero pitch angle at the nominal Orion impact
velocity of 7.62m/s (25fps), and with a cabin pressurized with atmospheric air at sea level pressure. Also, the airbag
vents were constrained to stay open once opened in this analysis, thus replicating the same venting profile proposed
for the ExoMars rover. As a result, the opening condition (measured here by acceleration) as well as the size of the
vent; are parameters which were also considered as variables in the design space.
With a set of dynamic performance values for each design parameter combination obtained for each airbag, these
design spaces were optimized to minimize the velocity at the end of the airbag stroke, over the set of solutions where
the magnitude of the largest induced acceleration was less than 45G’s. The choice of this cost function was based on
both energy dissipation and injury-risk level considerations, with the minimal velocity criterion aimed at minimizing
the resulting bounce as much as possible; and the acceleration criterion based on minimizing the resulting Brinkley
DRI (whose maximum value was found to be directly proportional in magnitude to the maximum acceleration
experienced during the impact event). Here, the acceleration limit of 45G’s was chosen based on data given in Ref.
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19. The final results of this preliminary design effort and the corresponding performance as predicted by the 1-DOF
airbag model are shown in Table 5 and Fig. 8 respectively.
Table 4 –Parameter Values Forming the Optimization Space for Each Airbag
Airbag Dimension Range
External Vent/Orifice
Airbag
Mass (kg)
Area to cross
Opening g
Radius (m)
Length (m) Pressure (atm)
sectional area ratio Loading (g’s)
4.74 0.09:0.005:0.11
0.3:0.05:0.6
0.9:0.05:1.1
0.1:0.05:0:6
-5:-5:-50
Head
Upper Back
14.43
0.13:0.01:0.2
0.3:0.05:0.6
0.9:0.05:1.1
0.1:0.05:0:6
-5:-5:-50
Lower
Back/Pelvis
Thighs
26.05
0.13:0.01:0.2
0.3:0.05:0.6
0.9:0.05:1.1
0.1:0.05:0:6
-5:-5:-50
15.83
0.1:0.01:0.2
0.3:0.05:0.6
0.9:0.05:1.1
0.1:0.05:0:6
-5:-5:-50
Lower Leg
8.37
0.1:0.01:0.2
0.3:0.05:0.6
0.9:0.05:1.1
0.1:0.05:0:6
-5:-5:-50
Table 5 – Airbag Parameter Values for First Generation CIAS Design
Head
Upper Back Lower Back
Airbag
Airbag
Airbag
0.11
0.17
0.18
Radius (m)
0.35
0.6
0.6
Length (m)
106.4
101.3
106.4
Internal Pressure at Impact (kPa)
0.008
0.009
0.01
Total Orifice Area (m2)
120
130
150
Orifice Opening Pressure (kPa)
Thigh
Airbag
0.14
0.6
101.3
0.006
145
Lower Leg
Airbag
0.18
0.6
101.4
0.01
120
It is interesting to note that in Table 5, the internal pressure at impact is very close to that of the cabin pressure. This
implies that the solution to the trade-off between an airbag which is likely to bounce due to high pressure, and one
which is likely to cause the occupant to impact the ground due to low pressure; is such that the airbag maintains its
approximate equilibrium shape. However, this trend may change with varying cabin pressure and operating medium,
and hence requires further investigation.
Body Region Acceleration vs Time
Body Region Vertical Velocity vs Time
4
2
0
0.05
0.1
Time (s)
0.15
x-Axis Brinkley DRI vs Time
10
2
0
1
-10
-20
Upper Back
Lower Back
Thigh
Lower Leg
-30
-40
-50
b).
0
0.05
0.1
Time (s)
0.15
x-Axis Brinkley DRI
6
0
a).
Upper Back
Lower Back
Thigh
Lower Leg
Acceleration (Earth gs)
Vertical Velocity (m/s)
8
0
-1
Upper Back
Lower Back
Thigh
Lower Leg
-2
-3
-4
c).
0
0.05
0.1
Time (s)
0.15
Figure 8. Predicted Performance for the First Generation CIAS a). Vertical Velocity Time History
b). Vertical Acceleration Time History c). x-Axis (Vertical) Brinkley DRI Time History
Notice that in Fig. 7, the performance of the head airbag is not plotted as it was found that it would never impact
the ground in the current system configuration.
It can be seen in Fig. 8a) that although a large amount of velocity has been dissipated from the system
throughout the initial phases of the impacting event; some residual velocity (and hence kinetic energy) still remains,
causing it to commence bouncing after approximately 0.04 seconds. One method to further dissipate this remaining
energy whilst maintaining the same spacing constraints is to implement an active venting scheme, whereby the flow
rate through the orifice can be varied depending on the acceleration profile. This would provide an added degree of
adaptability into the system, hence increasing its robustness to various crew members as well as to off-nominal
impact events.
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Furthermore, another feature which could improve the performance of the baseline airbag concept is the antibottoming bag - an additional bag whose purpose is to prevent the occupant from ever coming into contact with the
landing surface. This capability becomes particularly useful in the case of a “bottoming-out” event, where the
system still has a downward velocity component after all of the stroke in the main airbag has been consumed.
With regard to the vertical acceleration and Brinkley DRI plots shown in Figs. 8b) and 8c) respectively, it can be
seen that the maximum acceleration induced on the system remains below values corresponding to a low injury-risk
level. This is particularly evident by the fact that the maximum DRI values in Fig. 8c) are well within the low injuryrisk level ranges given in Table 2. It should be noted here that because the model used only simulates dynamics in
one degree of freedom, it was only possible to calculate the Brinkley DRI in the x-direction.
Therefore, because of this expected level of system safety, it was decided to defer the implementation of the
abovementioned features until the second generation system. This would also allow for the determination of the
performance of a minimum capability system anticipated to meet top level requirements, which would in turn give
an indication of the added value of these improvements relative to their contribution to the overall system
complexity. It is this information which would be used to ensure that the system complexity; and hence its mass and
number of failure modes remains at a minimum as the design matures.
D. Test Hardware and Infrastructure Design
With a complete airbag configuration defined, a means of facilitating tests to evaluate their performance was
required. To achieve this, a seat structure was designed to support both the airbags and an occupant, and a test rig
was designed to simulate the range of impact angles and vertical impact velocities anticipated during Orion crew
module landings. Details of these are shown below in Fig. 9.
Figure 9 – Test Article and Infrastructure Design Features
Additionally, to emulate the dynamic response of a human occupant, a 50th percentile male Hybrid II
Anthropomorphic Test Device (ATD); more commonly referred to as a crash test dummy, was used. To measure its
dynamic response and risk to injury, tri-axial accelerometers were installed inside its chest. In regards to securing
the ATD into the test article, a five point harness similar to that proposed for the existing Orion CIAS was used.
E. Airbags to Beanbags
As the supporting infrastructure for the first generation CIAS was being designed and built, a separate effort to
fabricate airbags took place. It soon became apparent that this task had several additional design variables not
accounted for during the preliminary modeling effort, including material type, stitch patterning, and inflation and
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venting valve selection. These parameters were eventually defined, with the exception of the venting valve, which
became a prohibiting factor.
In the majority of previously implemented airbag-based landing attenuation systems, venting has been facilitated
by some form of burst valve, which would open either passively or actively when a predefined acceleration level or
internal pressure was reached. Active systems used a pyrotechnic cutter to rip a designed pattern within the bag;
whilst passive systems used disks with precision machined weak points designed to fail at a predetermined pressure
load. Clearly, pyrotechnic cutters would not be appropriate for a system located in close proximity to an astronaut in
a pressurized cabin; so the passively actuated burst disk was originally baselined for the first generation CIAS.
Upon further investigation into the procurement of commercially available burst disks, it was found that they
would be prohibitive in cost and lead time given the pressing schedule of the project and its available resources. As a
result, it was decided to perform the first test campaign with an analog airbag system, rather than with that originally
designed. This analog was chosen to be beanbags – fabric bags filled with polystyrene beans. By moving to this
option, the issue of burst disks could be avoided as there was no longer a requirement to have a hermetically sealed
vent prior to impact. Material porosity and sealing was also no longer an issue as the beans forming the analog
operating medium had a substantially larger volume than that of their predecessor, making them easier to contain.
Hence with this decision, the objectives of the first generation test campaign also changed. Whilst previously, the
primary objective was to evaluate the performance of an airbag-based impact attenuation system; it had now
transformed to evaluating the performance of the test equipment and procedures, whilst also obtaining operational
and results analysis experience. Although these sets of objectives are not directly related, some valuable lessons
were still able to be learnt from this initial test campaign, as will be discussed in Section IV.
III. System Testing
With the decision to move to an analog airbag system finalized, a test
plan structured to achieve its newly defined test objectives was required.
The final test plan involved performing drop tests with the fully integrated
system from increasing heights until a failure occurred. Specifically, this
consisted of performing drops at a zero impact angle at increasing
increments of one foot. Note that this zero impact angle used was based on
approximations made in the optimization exercise discussed in Section IIB. This was deemed appropriate based on the newly formed test
objectives. Hence by testing in this manner, trends in the system dynamics
and its load attenuation capability could be ascertained at varying impact
velocities, thus allowing for a performance envelope to be generated.
Additionally, this would also enable for the failure modes of the system to
be investigated.
With regard to data acquisition, a combination of accelerometers a).
embedded in the chest of the ATD, and high speed camera footage was
used. Accelerometer data allowed for the Brinkley DRI time history of
each drop to be calculated, whilst high speed camera footage allowed for
the dynamics of the system to be studied. Together, this data allowed for
the injury-risk level of the system to be related to its configuration via
comparison of the DRI time history with the system dynamics. Figure 10
shows the test setup used to perform these drops.
To obtain system dynamics information from the high speed camera
footage, a technique called photogrammetry was employed. Commonly
b).
used in the analysis of impact tests, this method involves placing a series Figure 10. a). Integrated Test Setup
of illuminated dots on the test setup. This illumination can be achieved by b). Lighting Conditions used for Testing
a color contrast between the dot and the background to which it is attached; or in the case here, via reflectivity.
During post-processing of the high speed camera footage, the position of these points is tracked frame by frame
using the dot illumination as an identifying property. By numerically differentiating the position of these tracking
points, important information such as impact velocity and maximum system acceleration can be gleaned. This
process is summarized below in Fig. 11.
In addition to linear dynamics information, photogrammetry can also be used to yield attitude data of the system,
by comparing the position of multiple strategically placed tracking points. For the case here however, the small
frame size of the high speed camera available meant that this was not possible. To overcome this, the Hough
9
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transform method20 was used to detect lines in the image frames corresponding to the seat structure, thus allowing
for its pitch angle to be directly obtained. With both attitude and position data obtained, it was then possible to
resimulate the impact dynamics of the system, thus further aiding in the analysis of each drop.
Resimulated Dynamics
Adjusting Lighting
Conditions
High Speed Camera
Display
Photogrammetric
Analysis
Dynamic Response
Detected Line
Tracking Point
Corresponding to
Seat Structure
Figure 11. Photogrammetric Analysis Technique Employed
IV. Test Results and Analysis
In late April 2009, testing of the first generation analog airbag-based CIAS commenced. As drops were being
performed, the limitations of the 1-DOF model became immediately apparent, with unforeseen two-dimensional
effects significantly contributing to the system dynamics. Specifically, this refers to a consistent pitching moment
which was experienced during all drops, where the system would bounce and pitch forward after impacting the
ground, and experience a forward sliding motion before coming to rest. As the drop height increased, this pitching
and sliding motion became more prominent. Figure 12 shows a breakdown of this motion.
Figure 12. Typical Dynamic Response Seen During Testing a). Impact b). Bouncing and Pitching c). Sliding
It was evident from these initial observations that this motion was due to a moment imparted on the system by an
offset distance between its center of gravity and the location of the impact. Later analysis revealed that as the system
bounced and pitched, its center of rotation changed from being at the lower back beanbag to the lower leg beanbag.
As the system was coming to rest after the maximum pitch angle was achieved, its geometry caused the lower
portion of the system to slide forward as the center of gravity was returning to the ground. Clearly, because the
model used for this first generation design was based on one degree of freedom in the vertical direction, this
imparted moment and subsequent pitching motion was not captured.
Additionally, further comparison of the 1-DOF model output and the observed system dynamics indicated that
the shape function used to relate the stroke of the bag to its geometry was not entirely accurate, thus influencing the
behavior of the subsequent results. This has since been resolved via refinement of this shape function.
Hence, as a result of these discrepancies, it was decided to inspect the system for damage after a preliminary set
of drops without the ATD. During this inspection, no damage was observed, resulting in the decision being made to
proceed with drop testing with the ATD until system failure. After a series of successful and uneventful drops (refer
to the Appendix), the system experienced a failure from a drop height of 5 feet, with the lower back beanbag
rupturing on impact. Upon detailed analysis of the high speed camera footage for this drop, it was concluded that
this failure was due to high pressure in the bag causing a rupture at a local stress concentration generated at the
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American Institute of Aeronautics and Astronautics
interface between the bag and its fill cap. At the
initiation of this rupture, a tear formed in a direction
45 degrees to that of the fiber – the direction in which
the beanbag material is the weakest. This conclusion is
supported by experience with local stress
concentration induced tears gained during preliminary
testing, as well as by studying the beanbag material’s
failure properties post-testing. Figure 13 shows the
system immediately after experiencing this failure.
Additionally, the reason as to why it was the lower
back beanbag that failed was because the system
configuration and the test impact angle meant that it
was the first to contact the ground, causing it to bear Figure 13. Test Failure from a 5 foot Drop Height
the majority of the impact load of the system.
Hence with the testing campaign completed as a result of this failure, a detailed analysis of the dynamics of each
drop commenced. This consisted of using both information obtained from photogrammetric analysis of the high
speed camera footage and accelerometer data, processed with a Brinkley model in the following manner:
- Using photogrammetry, time histories of vertical displacement, velocity, acceleration, and pitch angle were
obtained
- Based on the obtained time histories, time stamps were calculated for important stages in the system dynamics
throughout the drop. These stages included the time of impact, the moment at which the end of stroke of the
lower back beanbag occurred, and the time at which the maximum forward pitch angle was achieved
- Using the fact that the maximum vertical acceleration occurs at the end of the first stroke, the timestamp for this
event was used to time synchronize the accelerometer data (ie. the time at the end of the first stroke corresponds
to the time of the maximum vertical acceleration as measured by the ATD’s accelerometers)
- With time synchronized accelerometer and photogrammetrically obtained system dynamics data (and hence
resimulated dynamics), dynamically significant information can be extracted. Because the resulting Brinkley
DRI time histories are also time synchronized with the accelerometer data, this allows for comparison between
dynamic events and their injury-risk level
Note that a detailed comparison of dynamic histories at the photogrammetrically tracked point, and the position
corresponding to the Brinkley frame of reference, showed that a negligible time lag between dynamic events
occurring at these locations existed between them, thus justifying the validity of this analysis approach. Also, the
fact that the photogrammetric analysis and the accelerometer data are referenced to different regions within the test
article means that a direct comparison between their outputs cannot be made directly.
A. Considerations Regarding Measurement Uncertainty
As with any type of experiment, the uncertainty in the obtained measurements is an important consideration to
make when performing analysis of the subsequent results. For the case here, these primarily arose from the
photogrammetric techniques used to analyze the high speed camera footage, and the accuracy of the accelerometers
embedded in the ATD.
With regard to the accelerometer derived error, independently performed calibration tests indicated that their
output was accurate to within ±0.4G’s. Contrastingly, errors arising from the photogrammetric analysis were
estimated instead, based on observations made during post-processing of the high speed camera data. This was so
because changing lighting conditions and minute angular distortions affecting the image frame prevented a more
analytical method of obtaining error values from being used. Specifically, these errors appear in the form of
inaccurately tracked pixels shaping the tracking point, and misdetection of lines corresponding to the seat frame.
These correspond respectively to the point tracking and line detection algorithms implemented as part of the
photogrammetric analysis.
For the error arising from the point tracking algorithm, it is estimated that the calculated position of the
illuminated tracking point is within ±2 pixels of the true location as viewed by the high speed camera. This
estimation corresponds to a displacement error of approximately ±4mm and is based on close examination of the
final output of the algorithm used. Correspondingly, the attitude data obtained via line detection was estimated to
have an accuracy of ±4°, with this value being based on a later study performed to investigate the differences
between the method used here and other point tracking based attitude estimation techniques. Moreover, the
implementation of the human inspection of image frames, where a line corresponding to the seat structure could not
be directly detected by the Hough transform algorithm; also contributed to the aforementioned estimated error value.
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American Institute of Aeronautics and Astronautics
Scenarios where such an inspection was required occurred in frames where a loose cable, stray polystyrene beans, or
other such objects obstructed part of the camera’s view.
Related to these error sources is the period in time at which their values peak. As expected, the uncertainty in the
photogrammetric data is at a maximum at the turning points in the obtained dynamics profiles; where significant
dynamic events take place. Specifically, this refers to the fact that the largest point-tracking derived position error
occurs at the end of the stroke of the lower back beanbag, which is the first to contact the ground. Similarly, the
largest line detection derived angular error occurs at the peak of the first bounce of the system, where it obtains its
maximum pitch angle.
B. Typical Results
Since the results of each test indicate largely the same trends, the results of a single test case will be presented
here to show what was typically obtained. Here, this test case will be a drop from an initial height of 3 feet. A
summary of the results for all drop tests conducted can be found in the Appendix. Figure 14 shows the
photogrammetrically obtained data obtained from this test.
Test 6 - 3ft Drop - Vertical Displacement Time History
350
Vertical Displacement
300
Ground Plane
Test 6 - 3ft Drop - Pitch Angle Time History
50
X: 0.354
Y: 48.45
Vertical Displacement (mm)
45
Pitch Angle (deg)
40
35
30
25
20
15
10
250
200
150
100
50
5
0.5
a).
X: 0.354
Y: 244.3
1
Time (s)
1.5
2
X: 0.08
0
Y: 33.89
0.5
b).
1
Time (s)
1.5
2
Test 6 - 3ft Drop - Vertical Acceleration Time History
Test 6 - 3ft Drop - Vertical Velocity Time History
2
Vertical Velocity (m/s)
Vertical Acceleration (Earth gs)
25
1
X: 0.354
Y: 0.0322
0
-1
-2
-3
-4
-5
c).
X: 0.034
0.5
Y: -4.708
1
Time (s)
1.5
2
d).
20
X: 0.08
Y: 23.19
15
10
5
0
-5
-10
0.5
1
Time (s)
1.5
2
Figure 14. Photogrammetrically Obtained Test Results for a 3 foot Drop a). Pitch Angle Response
b). Vertical Displacement Response c). Vertical Velocity Response d). Vertical Acceleration Response
Using this information, the timestamps of significant dynamic events and their associated dynamic values can be
calculated. This in turn allows for a time synchronization with the data obtained from the ATD embedded
accelerometers to be performed, hence enabling for the effect of each event on the acceleration response in the
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American Institute of Aeronautics and Astronautics
Brinkley frame to be studied. Additional information can also be obtained from comparison of the acceleration
response with dynamics resimulated from the original photogrammetrically obtained data. This is summarized
below in Table 6 and Figs 15 and 16.
Table 6 – Time Breakdown of Significant Dynamic Events for a 3ft Drop Test
Event (Numbered Time Method of Identification
Reference
Dynamic Value
event in Fig. 14)
(s)
Figure
Time of Impact (1) 0.034 x Time of max. negative velocity
x Fig. 13c) Impact velocity = 4.7m/s
Time of End of
0.08
x Time of min. vertical displacement
x Fig. 13b) Max. acceleration =
First Stroke (2)
23.19G’s
x Time of max. spike in acceleration
x Fig. 13d)
profiles
Time at Height of
0.354 x Time where vertical velocity = 0
x Fig. 13c) Max. pitch angle = 48.5°
Bounce (6)
x Time at max. vertical displacement
x Fig. 13b) Height of bounce =
225.65mm
after the end of the first stroke
x Time of max. pitch angle
x Fig. 13a)
Acceleration Response - Test6 3ft Drop
4
4. First Contact with
Lower Leg Airbag
Drop Initiation
2
0
Acceleration (Earth Gs)
6. Height of Bounce
-2
1. First Contact with
Ground (Lower back
-4
airbag)
Sliding of Frame
-6
5. End of stroke of
Lower Leg Airbag
-8
2. End of stroke of
-10 combined upper and
lower back airbags
3. First contact
with thigh airbag
x
y
z
-12
-14
5.2
5.4
5.6
5.8
6
6.2
Time(s)
6.4
6.6
6.8
7
Figure 15. Dynamic Breakdown of Accelerometer Data obtained from a 3 foot Drop Test
As can be seen in Table 6, the system impacts the ground at approximately 4.7m/s, a velocity far less than the
expected Orion landing velocity of 7.62m/s (25fps). Between 0.04 and 0.05 seconds after this impact, the maximum
ATD experienced acceleration of -14.45G’s occurs in the vertical x-direction at the end of the stroke, as depicted in
Fig. 14. Following this, the system begins to pitch forward, as previously discussed. Approximately 0.02 seconds
after this, the instantaneous center of rotation of the system moves to the lower leg bag, as it comes into contact with
the ground. When the stroke of this bag is depleted and the system reaches its maximum pitch angle of 48.5°, it
starts to slide forward as it comes to its final resting position. For the current test case, this final resting position is
achieved approximately one second after the initiation of the drop. Also of interest is the fact that the x-direction
acceleration noticeably increases during the period over which the system is falling. As will be mentioned below,
this is the result of an additional force imparted on the system as the quick release mechanism supporting it was
triggered.
Hence with a basic understanding of the system dynamics obtained, a comparison can be made between it, and
the resulting injury-risk level. Figure 17 shows the Brinkley DRI for this test case, normalized with the low injuryrisk Brinkley limit (as found in Table 2); and its corresponding low, moderate, and high injury-risk beta numbers.
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American Institute of Aeronautics and Astronautics
1. First Contact with Ground
2. End of First Stroke
3. First Contact with Thigh Bag
4. First Contact with
Lower Leg Bag
5. End of Stroke of
Lower Leg Bag
6. Height of Bounce
Figure 16. Resimulated System Dynamics for a 3 foot Test Drop
E Number for Test6 3ft Drop
0.2
0
-0.2
0.8
Low
Moderate
High
X: 5.913
Y: 0.9198
0.7
0.6
0.5
0.4
-0.4
0.3
-0.6
0.2
0.1
-0.8
5.4
5.6
5.8
6
6.2
6.4
6.6
5.6
5.8
6
6.2
6.4
b).
Time (s)
Time (s)
Comparison between Acceleration and Brinkley DRI in the x-axis
15
Acceleration (Earth Gs) & Brinkley DRI
a).
0.9
x
y
z
0.4
E Number
Scaled Low Injury-Risk Brinkley DRI
Scaled Low Injury-Risk Brinkley DRI for Test6 3ft Drop
c).
6.6
6.8
7
10
5
0
-5
-10
-15
x-axis Acceleration
x-axis Brinkley DRI
-20
5.4
5.6
5.8
6
6.2
Time (s)
6.4
6.6
6.8
Figure 17. Injury-Risk Responses for a 3 foot Test Drop a). Scaled Brinkley DRI b). Low Injury-Risk ȕNumber c). Comparison between Acceleration and Brinkley DRI Response
It can be concluded from Fig. 17b) that the injury-risk level for this drop is low, as demonstrated by the fact that
the value of the low injury-risk ȕ-number remains less than one for the duration of the impact event. More
importantly however, is the fact that the largest contribution to the injury-risk is from the acceleration in the x14
American Institute of Aeronautics and Astronautics
direction at the end of the first stroke of the system. This can be seen in Fig. 17a), where the contribution from the xdirection is clearly the most significant; as well as in Fig. 17c), where it is observed that the largest magnitude spike
in the Brinkley DRI corresponds to the moment at which the initial stroking capability of the system has been
depleted. This indicates that the load attenuating capability of the bag which first contacts the ground has the
greatest influence on the resulting injury-risk of the impact event, which is to be expected. However, the fact that the
Brinkley DRI experiences another spike of comparable magnitude after this initial spike indicates that some
substantial residual kinetic energy remains in the system. Close inspection of the high speed camera footage
suggests that this is due to the lower back bag bottoming out and allowing the seat structure to impact the ground.
Only after subsequent bouncing, pitching, and sliding, is this energy removed entirely.
Although the magnitudes of the peaks in the DRI response remain below the low injury-risk threshold for this
test case, this bottoming-out event will become significant for drops which more closely emulate the impact velocity
of the Orion vehicle’s landing scenarios, as will be seen in the next section. It is anticipated that the implementation
of venting-type airbags will significantly reduce the magnitude of this secondary DRI peak, and hence the influence
of the resultant bottoming-out event. Incorporating the previously discussed anti-bottoming airbags would also
greatly assist in the load attenuation performance of the system. With these additions, it is expected that the
magnitude of the initial acceleration spike which occurs upon impact will decrease; whilst the time period over
which it occurs will increase. This would result in a lower initial DRI peak and a significant reduction in magnitude
of the DRI’s secondary peak, thus improving the overall occupant protection capability of the system.
C. Trends with Varying Drop Height
With the typical system dynamic response characteristics now identified; variations in each of these
characteristics with varying drop heights, and their effect on the injury-risk level to the occupant can be investigated.
Figure 18 summarizes these trends.
Brinkley DRI in the x-Direction for the Gen 1 System Testing
15
0
-5
1ft
1ft
2ft
2ft
3ft
3ft
4ft
4ft
-10
-15
-20
a).
c).
Drop (2.5m/s)
Drop (2.7m/s)
Drop (3.7m/s)
Drop (3.7m/s)
Drop (4.5m/s)
Drop (4.7m/s)
Drop (5.3m/s)
Drop (5.5m/s)
Brinkley DRI in the x-Direction
x-Direction Acceleration (Earth Gs)
Acceleration in the x-Direction for the Gen 1 System Testing
10
5
0
1ft
1ft
2ft
2ft
3ft
3ft
4ft
4ft
-5
-10
-15
-20
-25
-30
4.2
4.4
4.6
4.8
Time (s)
5
5.2
5.4
b).
4
4.5
5
Drop (2.5m/s)
Drop (2.7m/s)
Drop (3.7m/s)
Drop (3.7m/s)
Drop (4.5m/s)
Drop (4.7m/s)
Drop (5.3m/s)
Drop (5.5m/s)
5.5
Time (s)
d).
Figure 18. Trends in System Response with Increasing Drop Height a). x-Direction Acceleration
b). x-Direction Brinkley DRI c). Low Injury-Risk ȕ-Number d). Impact Velocity
Note that in Figs. 18a) and b), the same color plots correspond to drops from the same initial height and that in
Fig. 18d); the experimentally obtained impact velocity values are higher than those calculated based on the initial
15
American Institute of Aeronautics and Astronautics
drop height. This is due to the manner in which the drop was initiated. Specifically, this refers to the fact that the test
rig operator must pull down on a quick release mechanism supporting the test article to initiate a drop. This
downward pull imparts an additional acceleration on the system, thus increasing its impact velocity. Also, this
discussion will focus mainly on the system response in the x-direction as it is the most critical in terms of
minimizing the overall injury-risk level at the zero impact angle tested (as was found in the previous section).
Upon first inspection of Fig. 18a), it can be immediately observed that the drop tests performed are indeed
repeatable, as indicated by the consistent trends in the system response when the drop height is fixed. Additionally,
it is also seen that the maximum impact acceleration, and hence injury risk, increases with drop height, which is to
be expected. More significant however, is the fact that at drop heights of 4 feet (corresponding to impact velocities
of approximately 5.4m/s), the low injury-risk ȕ-number exceeds a value of 1, signifying that a moderate risk of
injury exists for this case (Fig. 18c)). Based on the increasing injury-risk trend with increasing impact velocity, this
result clearly indicates that the current system would not attenuate impact loads to an acceptable degree in the
nominal landing case. This was indeed expected prior to the commencement of this test campaign due to the fact
that analog airbags were used.
A more detailed examination of the acceleration responses shown in Fig. 18a) confirms the fact that all drops
display largely the same dynamic trends. These are reflected in the corresponding DRI response, where all drops
exhibit a secondary spike of comparable magnitude to an initial impact-induced spike, indicating the consistent
occurrence of a bottoming-out event (Fig. 18b)). As was the case for the 3 foot drop case previously discussed, this
consistent bottoming-out was verified by close inspection of high speed camera footage, providing further evidence
for the need for anti-bottoming bags to be introduced into the next generation system design. Furthermore, this
bottoming out was also found to be attributed to the location of the seat structure relative to the bags. In its current
configuration, the structure reduces the available bag stroke by one half, causing the bags to bottom out earlier.
Consequently, future systems will be designed such that the seat structure enables for the bags they support to
expend their entire stroke.
V. Conclusion
A first generation analog airbag-based CIAS test article was developed and subjected to a series of drop tests. It
was found that at impact velocities of higher than 5.3m/s (a velocity below that of the Orion nominal landing
velocity) the system was unable to maintain a low-injury risk level. Although this was the case, several lessons were
learnt which will contribute significantly to future, airbag-based versions of the system. These include the need for
the seat supporting structure to be positioned such that it is conducive to maximum bag stroke expenditure, as well
as the potential need for anti-bottoming bags to further reduce injury-risk by preventing direct impact between the
occupant and the ground.
To address the issue of implementing cost-effective and low lead time burst valves to enable further
development of airbags, an investigation is currently underway, aimed at developing and testing a range of venting
schemes at a component level. Once a suitable scheme is finalized, detailed design and fabrication of a second
generation test article will commence. This will involve the following tasks:
- A redesign of the seat frame assembly to maximize airbag stroke
- Further refinement of the system model to better capture airbag dynamics, as well as two-, and three
dimensional dynamic phenomena
- The design and implementation of an airbag system consisting of anti-bottoming airbags, the finalized venting
mechanism, and features to mitigate the formation of local stress concentrations; and
- An investigation into methods of self-reconfiguration of the system to accommodate various mission phase
requirements
With these tasks and the testing of the subsequent system complete, it is anticipated that further improvements
will be implemented as more is understood about the relationship between the system’s design parameters and its
performance. As the design matures, the system complexity will be actively managed to ensure that the final product
satisfies its original objective of maintaining a low-injury risk landing environment in a lightweight, reconfigurable,
and robust manner – objectives aimed at facilitating the return of lost capability to the Orion vehicle’s baseline
design.
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American Institute of Aeronautics and Astronautics
Appendix
Results summary of all successful drop tests conducted:
Drop
Height (ft)
1
Impact
Velocity (m/s)
2.5
Height of
Bounce (mm)
83.8
Max Pitch
Angle (deg)
48.5
Max x-Acceleration
(ATD) (Earth G's)
-6.932
Max Brinkley
DRI-X
-12.44
Risk to Injury
(HSIR Specs)
Very low
1
2.7
94.7
42.5
-7.629
-13.13
Very low
2
3.7
147.8
47.1
-12.15
-19.93
Very low
2
3.7
155.2
47.5
-12.3
-19.67
Very low
3
4.5
211.4
51.0
-16.34
-24.45
Low
3
4.7
225.7
48.5
-14.45
-23.59
Low
4
5.3
N/A
N/A
-20.91
-29.82
Moderate
4
5.5
242.95
50
-20.28
-29.54
Moderate
NB. A “N/A” indicates that it was not possible to obtain the indicated data due to surrounding debris obstructing the
high speed camera lens.
Acknowledgments
Financial support for this research was provided by the NASA Engineering Safety Center (NESC) and the
Constellation University Institutes Program (CUIP) under prime award number NCC3989 and subaward Z634013.
The authors would like to thank Dr. Charlie Camarda and Dr. Ed Fasanella for their support of this work; as well as
Lisa Jones and her team at NASA LaRC for their assistance with the loan of a Hybrid II ATD. Many thanks also go
to Jackson Siu and Peter Cheung from the Pennsylvania State University for their invaluable contribution to this
project.
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