20th AIAA Aerodynamic Decelerator Systems Technology Conference and Seminar<BR>
4 - 7 May 2009, Seattle, Washington
AIAA 2009-2999
Design of Subscale Parachute Models for MSL Supersonic
Wind Tunnel Testing
James Reuter 1, Walter Machalick 2 Allen Witkowski3 and Mike Kandis.4
Pioneer Aerospace, South Windsor, CT, USA
Anita Sengupta 5
Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA, USA
Richard Kelsch 6
NASA Glenn Research Center, Cleveland, OH, USA
Subscale models of the Mars Science Laboratory (MSL) parachute were tested under
conditions representative of the MSL deployment envelope at the NASA Glenn Research
Center’s 10ft x 10ft wind tunnel. Four percent (4%) scale (0.813 m) parachute models were
used to examine parachute performance in the wake of a similarly scaled Viking-type entry
vehicle at supersonic speeds spanning Mach 2 to 2.5. Design of subscale parachute models is
typically a trade off between scaling accuracy and survivability in the expected test
environment. High fidelity scaling of structural elements renders the models susceptible to
damage during operation at supersonic speeds, especially in the presence of high frequency
area oscillations. Deploying and inflating these models at relatively high dynamic pressures
and speeds also posed challenges, as evidenced by the catastrophic failure of several
parachutes during and shortly after deployment in the first scheduled entry. As a result, the
equipment and procedures used to protect and deploy the model parachute system were
revised over several generations until a reliable system was obtained and utilized in the
second available entry. Additional trades and design changes were incorporated for the
third available entry in the test program. Measures were also taken in regard to the test
environment to enhance the life-span of parachute models against the high frequency forces
associated with canopy area oscillations. This paper provides a detailed description of the
parachute models as well as the evolution of the ancillary equipment and test procedures
that ultimately resulted in a reliable test technique.
Nomenclature
CD
D0
DGB
JPL
Mach
MSL
GRC
CFD
FSI
SOW
=
=
=
=
=
=
=
=
=
=
Drag coefficient
Parachute nominal diameter
Disk Gap Band
Jet Propulsion Laboratory
Mach Number
Mars Science Laboratory
Glenn Research Center
Computational Fluid Dynamics
Fluid Structure Interaction
Statement of Work
1
Senior Engineer, AIAA Member.
Program Engineer, AIAA Member.
3
Director, Engineering Operations, AIAA Member.
4
Engineering Lead Analyst, AIAA Member.
5
MSL Subscale Parachute Contract Technical Manager, AIAA Senior Member.
6
Program Manager, AIAA Member.
2
1
American Institute of Aeronautics and Astronautics
Copyright © 2009 by Pioneer Aerospace Corporation. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.
I. Introduction
T
he Mars Science Laboratory (MSL) Parachute Decelerator System (PDS) qualification program included the use
of Computational Fluid Dynamic (CFD) and Fluid Structure Interaction (FSI) numerical simulations to explore
supersonic performance. These numerical simulations were validated by subscale supersonic wind tunnel testing
programs. This paper contains the technical basis for the design of a 4% scale Viking Disk-Gap-Band parachute to
be used in supersonic tests at the Glenn Research Center. It also describes several issues and design changes that
arose during the test program, and provides the results of static structural tests and measurements. The References
section of this paper also provides a list of several of the reports associated with the test and instrumentation
methods, analytical methods, and results based on the testing utilizing the models described herein (Note that the list
does not necessarily encompass all reported sources for this exceptionally detailed test program).
II. Requirement
Design requirements were defined by JPL MSL PDS Contract SOW, Exhibit V, Revision B, reprinted below along
with Table 1. Note that some of these requirements are not the same as the actual tested values, which can be found
in the referenced documents, but are reprinted here as the originating design requirements. Conversion of these
requirements into appropriate geometric relationships and a gore model are presented in Table 2.
The purpose of this task is to design and fabricate scaled MSL parachutes for the GRC 10x10 supersonic wind
tunnel test program. Wind tunnel tests are to be conducted to validate Fluid Structure Interaction (FSI) codes
under development for the MSL parachute program. The test program will provide experimental data to explore
the Mach, Re, trailing distance, and dynamic pressure dependence of Viking DGB parachute dynamics in
supersonic flow. The scaled parachutes are to be geometrically identical to the full scale flight parachutes in
terms of Viking scaling parameters. The scaled parachutes shall be materially representative to the extent
possible for the wind tunnel environment detailed below.
1.
Design 2.5% (sic) scale MSL parachutes to meet the operating conditions as detailed in Table 1. Design
shall be a geometric scaling from the Viking-type flight canopy but there is no requirement to match the
number of gores of the flight canopy. Design shall minimize material stiffness yet meet load requirements
derived from Table 2. Design data shall include:
a. Gore layout
b. Fabrication approach
c. Stress analysis
d. Materials selection
2. Design interface of scaled parachutes to the GRC test fixtures:
a. Provide means of attachment of the riser to GRC load transducer for tension measurement
b. Provide means of canopy restraint to constrain canopy motion in the lateral direction but not
deform the canopy due to axial translation...
Provide means for quickly and reliably removing canopy restraint without interfering with inflation
Test Description
Do
(m)
x/d
Mach
Q
(Pa)
Constrained 21.5 match Reflight
0.813
10
2.0
2.5
15440
17666
Constrained 21.5 match Reflight
0.813
10
2.0
2.5
15440
17666
Unconstrained 21.5 match Reflight
0.813
10
2.0
2.5
15440
17666
Unconstrained 21.5 match Qflight
0.696
10
2.0
2.5
650
1015
Table 1 - Design Test Conditions
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American Institute of Aeronautics and Astronautics
Symbol
Nominal diameter,
Canopy
reference
area
Geometric porosity
Disk area
Gap height
Band height
Vent diameter
Suspension
line
length
Number of gores
Gore included halfangle
Disk area per gore,
incl. vent
Disk gore height
Disk base
Disk radial length
Total gap area
Total band area
Trailing Distance
*Specification
requirement
Do
Value
32.01 in
Source
0.813 m
2
2
4% Viking
0.519 m
¼π Do2
(0.125)So
12.5% So*
426.46 in
1.34 in.
3.87 in.
2.24 in.
54.42 in.
0.275 m
0.034 m
0.098 m
0.057 m
1.382 m
0.53So*
0.042Do*
0.121Do*
0.07Do*
1.7Do*
N
Ǿ
24
7.5 deg.
24
7.5 deg
selected
360/2N
Adg
17.77 in2
0.011 m2
SD /N
hdg
bd
11.67 in.
3.05 in.
0.296 m
0.077 m
(Adg /sinǾ)2
2hdg tanǾ
rd
Agap
Aband
T
11.77 in.
98.01 in2
283.28 in2
66.9 in.
0.299 m
0.063 m2
0.183 m2
1.70 m
hdg/cosǾ
24bdHG
24bdHB
10d (d=0.209 Do*)
So
804.64 in
lG
(0.125)So
SD
HG
HB
DV
LS
2
Table 2 - Specified Viking Geometric Relationships and Gore
Layout for a 24-Gore, 4% Scale Model Before Adjustments for
Stitching Take-up and Vent Blockage
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American Institute of Aeronautics and Astronautics
III. Design Details
A. Design Loads and Material Selection
Design of the 32-inch DGB model involves an interplay between number of gores, a reasonable match
between strength requirements and available materials, flexibility and scaling of finished seams and hems,
symmetrical matching of suspension lines and confluence bridle legs, and manufacturing limits on miniaturization.
A worst-case steady-state axial load is preliminarily estimated by the expression
F = CDAq, where
CD = drag coefficient based on canopy area
A = canopy area, ft2
q = worst case dynamic press, lb/ ft2
F = (0.6) (5.59)(369) = 1237 lbf (5502N)
To the steady state-load is applied an estimated factor of 2.25 to account for
opening
shock+pulsing+fatigue, a joint and stitching degradation factor of 1.25, and a structural safety factor of 1.5, bringing
the required total axial/radial material strength requirement to 5219 lbf. (23,214N)
number
of gores
select
12
12
16
16
24
36
minimum
suspension
line & radial
member
material
strength req.
-lbf
435
435
326
326
217
145
number of
confluence
bridle legs
4
6
4
8
6
6
number of
lines per
bridle leg
3
2
4
2
4
6
minimum
bridle leg
material
strength
req.
width of
band and
disk base inches
width at
vent
6.64
6.64
4.85
4.85
3.05
2.03
0.50
0.50
0.37
0.37
0.29
0.20
-lbf
1304
870
1304
652
870
870
-inches
Table 3 - RELATIONSHIP BETWEEN NUMBER OF GORES AND MINIMUM
STRENGTH REQUIREMENT OF AXIAL LOAD-CARRYING ELEMENTS
Kevlar is selected as the material for axial/radial load-carrying members because of its high strength-todensity ratio and flexibility. (note: a Kevlar parachute structure with nylon fabric is also a specification
requirement.) Kevlar cord is currently available in rated breaking strengths of 100, 200, and 300 lbf. A 200-lbf cord,
when flattened by the sewing machine presser foot, provides the narrowest width that can be stitched with thread of
appropriate strength. This suspension line cord has an approximate diameter of 0.0195 inches. For the bridle
material, a small width dimension is desirable in order to minimize bunching at the loop-to-loop line and riser
connections. The smallest widths of Kevlar tape available are 0.5 inches in 250, 550, and 800 lbf rated strengths;
and 0.5625 inches in 500 and 700 lbf rated strengths. From evaluation of the above table, a 24-gore design is
selected. This choice is based on best material selections for line/radial and bridle members, the number of lines
appropriate for connection to each bridle leg, and avoiding main seam crowding at the vent.
A canopy material of low-permeability was a specification requirement in order to simulate the effective
permeability of parachute fabric operating in the low density Martian atmosphere. A standard 1.17 oz/yd2 parachute
fabric having a porosity range of 0.5 to 5.0 ft3/ft2/min under 0.5 in H2O pressure differential, PIA-C-44378C, was
selected.
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American Institute of Aeronautics and Astronautics
Confluence Bridle and Riser Extension Construction
A structurally efficient method of joining the suspension lines to the legs of the confluence bridle is by
forming fold-back loops at the end of the bridle legs and interlocking the loops. The six-legged bridle is formed
from three layers of 800-lbf Kevlar web and the riser from single 4000-lbf Kevlar cord. The fold-back configuration
effectively doubles the strength at the point of contact.
To provide the specified distance from capsule major diameter to canopy forward edge, 71.66 inches, it is
necessary to add an extension from the bridle confluence point. This is usually accomplished by extending the
bundled bridle legs to the desired length as a single member. However, an alternative approach was necessary
because of load-share issues. As an alternative, the bridle legs were not extended but were instead terminated in a
loop just below the confluence wrap. A length of Kevlar cord rated at 4,000 lbf breaking strength was added as a
riser extension to the swivel and attached to the bridle with a loop-to-loop connection and to the swivel by loop-topin, see figure 1.
Figure 1 - Web-type Bridle with Cord Riser Extension
Total blocked
area, sq. in.
Total open area,
sq. in.
Ratio of Open
Vent Area to
Canopy Ref. Area
804.2
804.2
Area blocked by
remainder of line
32.00
32.00
Area of crossover or vent
fitting blockage
804.2
Radius of crossover or vent
fitting blockage
32.00
Gore half-angle,
degrees
Proto. 32" DGB model
PN 11582 (trial unit)
New 32" DGB models
PN11673
32" DGB with 2" dia vent
plug PN 11708
Width of vent
line, inches
562728
Number of gores,
N
846.46
21.5 m MSL DGB
Vent Area,
inches2
Nomin
al dia.,
Do,
inches
59.25
*(.07Do)
2757
80
0.22
2.25
2.802
24.666
472.084
496.750
2260.446
0.00402
1.33**
2.56*
2.64***
3.30*
3.42***
1.39
5.15
5.47
8.55
9.19
24
0.07
7.5
0.268
0.226
24
0.07
7.5
0.268
0.226
24
0.12
7.5
1.000
3.142
0.667
1.700
1.767
1.728
2.386
0.893
1.926
1.993
4.870
5.528
0.497
3.221
3.481
3.173
3.660
0.00062
0.00401
0.00432
0.00401
0.00455
Vent Diameter,
inches
Description of Parachute
Nominal area, So,
inches2
Adjustments to Vent Area to Account for Vent Line Thickness
To aerodynamically scale the vent of the model, it was considered necessary to scale the unblocked, or
open, vent area rather than the conventionally-defined area based on the length of the vent lines. Calculation of
open vent area is straightforward geometry which considers the total vent area based on vent line length and
subtracts the area of total blockage where lines cross over (or the area of the vent fitting for the constrained
configuration) and the blockage of the exposed length of each vent line. A target ratio of open area to canopy
reference area equal to 0.00402 was calculated using a measurement of the vent line width of the 21.5 m MSL DGB
vent lines. Pattern vent line markings were determined by iterating vent diameter to match the target ratio for both
the constrained and unconstrained 32-inch DGB models. Table 4 contains the results and also shows an actual vent
open area ratio based on actual average finished vent line lengths. It should be noted that Viking designers did not
account for blockage in their reported vent details. The Viking canopy vent is typically listed as 0.07 x DO with vent
area = 0.005 x S0. In reality, the open area of the vent was approximately one half that when accounting for the
blockage due to vent tapes and the attached deployment bag.
*Pattern marking dimension **finished dimension ***Average finished vent line measurements, see Table 5
Table 4 - Calculation of Model Vent Diameter to Achieve Unblocked Vent Area Ratio Equivalent to
21.5m MSL DGB
5
American Institute of Aeronautics and Astronautics
B. Modifications and Changes
Bridle Design
Several failures of one or more bridle legs occurred during the first tunnel entry. In virtually all
instances, the fold-back stitching of the suspension line capture loop was the weak point. The cause was determined
to be an inability of the Kevlar web to equally spread the narrow, concentrated force of the suspension lines across
the width and into all stitching of the fold-back. Instead, the force was largely directed into the "point" of the two
outer rows of stitching and caused progressive failure of all stitching. The solution was to change the bridle material
from a wide web to braided Kevlar cord and employ the "finger trap" construction shown in Figure 2.
Riser Extension
For purposes of relocating the swivel from the capsule-attachment end to the mid-point of the riser, the
one-piece riser design was replaced by two shorter lengths. Forming independent attachment loops at each end of
the riser was no longer possible with the short design. Instead, a continuous loop of Kevlar cord was formed. The
center portion of the loop was confined by shrink tubing with the exposed ends functioning as attachment loops.
Swivel location
Heat-shrink covering
Figure 2 Cord-type Bridle with Two-Piece Riser Extension
Vent Line-to-Vent Ring Attachment Joint
During tests of the constrained configuration with apex fitting, several or all of the vent lines failed at the
inner-diameter contact point between cord and ring. Since the apex region of the parachute is subjected to severe
and rapid dynamic motions during test, a probable cause of vent line failure was wear due to relative sliding motions
of the cord and the inner diameter of the ring. To resist this type of motion, a silicone rubber sealant compound
(General Electric RTV 110) was applied to both sides of the ring between each vent line. A Teflon fixture was
machined to obtain and hold correct positioning of ring and lines during the curing process.
Figure 3 - Vent Line Alignment Fixture
Figure 4 - Vent Ring and Vent Lines in
Fixture During RTV Curing
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American Institute of Aeronautics and Astronautics
Thin Suspension Lines
Analysis of shock wave patterns observed during entry 1 test runs generated a requirement to modify, and
fabricate new, a certain number of models with suspension lines of thinner diameter. Modification of existing
canopies consisted of cutting off existing lines at a point just below the canopy skirt and folding back the remaining
"tail" to form a small loop coincident with the level of the skirt. Fabrication of new thin-line chutes employed the
same processes as original new parachutes, except that the radial reinforcement cords were cut and marked to extend
just past the skirt (instead of continuing to a confluence point). An attachment loop was then formed as described
above for the pre-existing models. Loops at each end of the replacement lines interlocked with these loops and with
existing bridle loops.
Two types of cord were selected for use as suspension lines: PIA-T-87128 size 5 Kevlar thread, untreated,
155 lbf rated minimum breaking strength and TKMB-20 braided Kevlar cord, untreated, 100 lbf rated minimum
breaking strength. These became identified as "Thin" and "Super Thin" on construction drawings and in test reports.
While under light tension, both of these cords had a measured diameter of about 0.014 inches. The TKMB braided
type cord constricted somewhat under tension and it did not widen by fraying or "fuzzing" during tests as did the
untreated Kevlar thread.
Because the replacement lines were too narrow to be stitched or "finger trapped" as in the base design, they
had to be attached to both canopy and bridle using some form of knot. In selecting a knot type, three factors were
investigated:
1. Strength loss at the knot
2. The amount of knot slippage under predicted tension
3. Avoiding a need to disassemble or replace the bridle legs
The knot that proved to be most repeatable and most accurately formed, and having minimum slippage,
was using a short fold-back and an overhand knot to form a small loop at each end of the replacement line. Each
end loop was formed on a fixture to control and hold its size while the knot was being seated. A short length of
tubing was then heat-shrunk over the knot as a means of resisting slippage and aerodynamically constraining the
"tail" of the cord. To join bridle legs and lines on existing parachutes, independently forming the suspension line
loops required either a new bridle assembly or "opening" the six end loops of existing bridles by removing stitches.
The latter was determined to be the more cost effective method.
C. Deployment Sleeve
Requirement
The function of the deployment sleeve is to constrain the folded canopy and to hold the suspension line
group plus sleeve under tension during the tunnel start-up period. The original requirement was to provide an
arrangement whereby a mechanically-generated axial force of 40 lbf applied to the downstream end of the sleeve
would tension both the sleeve and the suspension lines during run-up. When test conditions were reached and the
mechanical force increased to 60 lbf or more, the sleeve would autonomously open, separate from the test item, and
blow down the tunnel.
Initial "Daisy Chain" Design
The initial sleeve design, shown as the top image of Figure 5, consisted of a flat fabric assembly that, when
wrapped about the folded canopy, was held closed by a series of daisy chain interlocking loops. At the downstream
end, a cord from the mechanical retractor was looped about the first link in the chain. At the upstream end, the last
link was looped about a break tie holding the mouth of the sleeve closed. During the first tunnel entry, this approach
was found to be largely unsatisfactory because at the 90 degree bend, where the long active loop passed through the
short passive loop, the cord would "kink", or take a set that resisted the free unlacing process observed in prior
bench tests. There were apparently two reasons for the difference: (1) the dynamic environment to which the daisy
chain was subjected while under tension, and the extended period of time while under pre-release tension, were not
replicated in bench tests, and (2) the planned pre-release tension of 40 lbf did not adequately prevent excursions and
vibration of the suspension lines and was increased to a level well above that applied in bench testing.
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American Institute of Aeronautics and Astronautics
Rip Cord Design
As replacement for the daisy chain loops, a redesign, shown in the lower image of Figure 5, incorporated a
series of Teflon pins along a ripcord to hold the sleeve closed. During run-up, mechanical tension was applied to the
sleeve and suspension lines by means of a break cord attached directly to the sleeve. The end of the rip cord was
also attached to the tensioning cord so that, when the break tie failed, tension was abruptly transferred to it.
Residual elasticity and continued retraction of the tensioning cord caused the pins to be extracted. The length of the
pin locking the mouth closure at the upstream end of the sleeve was sized to ensure that it would be the last pin to be
removed.
Although the redesign was successful in completely unlocking the sleeve and the mouth closure, it was
observed during initial runs that the sleeve did not move downstream quickly enough by air blast alone to prevent
overtake by the inflating canopy, which occurred between 9 ms to 15 ms. To overcome this, an elastic cord was
attached between the sleeve and the retractor cord. In addition to being pre-stretched, this cord became additionally
stretched as the tensioning cord retracted a distance needed to fail the break cord. Thus, when the mouth tie pin was
extracted, contraction of the elastic cord "snatched" the sleeve off of the still-folded canopy quickly and positively.
This approach was successful, although it did require that both sleeve and rip cord remain attached to the retraction
mechanism instead of moving freely down stream as originally desired.
Figure 5 - Sleeve Designs
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American Institute of Aeronautics and Astronautics
D. Model Finished Measurements
Nine unused models and a surviving model (from run 20) were measured in the axial/radial direction as a
means of evaluating dimensional accuracy, consistency and repeatability. Five of these models were also measured
circumferentially for the additional purpose of calculating an average actual finished canopy area and geometric
porosity. This process required 144 axial/radial and 96 circumferential measurements per canopy, each made while
the segment involved was subjected to 10 lbf tension. Average segment measurements of all canopies are listed in
Table 5. Non-dimensional specification requirements are located in the column headings. It should be noted that
the disk diameters of the constrained units are smaller than the disk diameters of the unconstrained design. This
difference results from the necessity to enlarge the vent to account for the center fitting. Based on average
measurements, the calculated areas of each gore segment and the resulting geometric porosity are shown in Table 6.
Note that the surviving model is significantly larger than the unused ones. This is expected as past wind tunnel and
aerial drop test canopy measurements also show a “stretching” of the canopy to a slightly larger size post-test. This
effect was of concern during the Viking program, because the “stretching” resulted in a significant loss of elasticity.
This is of much less concern today, since the materials used do not have great elasticity to start with.
SUMMARY OF RESULTS OF MEASUREMENTS OF UNUSED* 32-INCH-Do MODELS
*except SN 626074 from run 20
Constrained?
Component (req'ment)
No
No
P/N
Configuration
11673501
New with
continuous 200lb lines, cordtype bridle
Average
of 24
lines+
bridle
inches
Band Length
(0.121Do)
Average
of 24
Ls/Do
gores
inches
HB/Do
Gap Length
(0.042Do)
Average
of 24
gores
inches
LG/Do
Disk Diameter
(0.726Do)
Average
of 24
gores
inches
DD/Do
Vent Dia. (.07Do)
Average of
24 vent
lines +
fitting if
applicable
inches
Dv/Do
Vent Area (0.005 So)
Total vent
area, in2
note 1
Area
blocked by
vent lines
and center
fitting in2
SV/SO unblocked
626006
56.00 1.75
3.89 0.122
1.34 0.042
24.06 0.752
2.67
0.083
5.600
2.018
0.0045
626014
55.95 1.75
3.95 0.123
1.34 0.042
24.10 0.753
2.69
0.084
5.683
2.035
0.0045
626021
55.87 1.75
3.94 0.123
1.32 0.041
24.06 0.752
2.60
0.081
5.309
1.959
0.0042
626017
54.75 1.71
3.75 0.117
1.35 0.042
24.06 0.752
2.68
0.084
5.641
2.027
0.0045
626018
54.68 1.71
3.85 0.120
1.32 0.041
24.06 0.752
2.47
0.077
4.792
1.850
0.0037
56.44 1.76
3.83 0.120
1.43 0.045
24.04 0.751
2.74
0.086
5.940
2.077
0.0047
626023
55.83 1.74
3.95 0.123
1.30 0.041
22.70 0.709
3.44
0.108
9.294
5.561
0.0046
type bridle
626070
56.06 1.75
3.99 0.125
1.29 0.040
22.76 0.711
3.44
0.108
9.294
5.561
0.0046
New with 100-lb
626078
56.14 1.75
3.93 0.123
1.29 0.040
22.74 0.711
3.40
0.106
9.079
5.494
0.0045
626079
56.24 1.76
3.95 0.123
1.28 0.040
22.76 0.711
3.40
0.106
9.070
5.494
0.0045
11673- Retrofitted with
155 lb "thin" lines
503
("5-cd")
New with 100-lb
No 11780-1 "super" lines
New with
Yes
S/N
suspension line
length (1.70 Do)
11708- continuous 200lb lines, cord501
626074
(run 20)
Yes 11770-1 "thin" lines, cordtype bridle
Note 1: Design vent area of vent was increased to account for non-scale blockage by vent lines and center fitting
Table 5 - Measurement Summary
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American Institute of Aeronautics and Astronautics
SUMMARY OF CALCULATED GORE-SEGMENT AREAS AND GEOMETRIC POROSITY
Constrained?
Component
(Spec.)
No
No
P/N
11673501
11673503
No
11780-1
Yes
11708501
Yes 11770-1
Band Area
(0.35So)
Configuration
S/N
New with
continuous
200-lb lines,
cord-type
bridle
Retrofitted with
155 lb "thin"
lines ("5-cd")
New with 100lb "very thin"
lines
New with
continuous
200-lb lines,
cord-type
bridle
New with 100lb "thin" lines,
cord-type
bridle
626006
626014
Gap Area
(0.12 SO)
Average
of 24
AB/So
gores
in2
Disk Area, incl.
vent (0.53SO)
Average
of 24
gores
in2
AG/So
Average
of 24
gores
in2
AD/So
Disk Circumference
(2.285Do)
Total
Area
(804.25)
Geometric
Porosity
(12.5%)
Average
base of
circ.D/Do
24 gores
inches
AT
in2
(AG+AV)/AT
11.83
0.353
3.970
0.118
17.58
0.525
3.00
2.250
801.2
12.55
11.55
0.345
3.907
0.117
17.49
0.522
3.01
2.258
790.8
12.25
11.62
0.347
4.290
0.128
17.88
0.534
3.04
2.280
807.1
13.54
626070
12.11
0.361
3.883
0.116
16.87
0.503
2.99
2.243
788.6
12.38
626078
11.65
0.348
3.851
0.115
16.67
0.497
2.96
2.220
777.2
12.25
626021
626017
626018
626074
(run
20)
626023
626079
Table 6 - Calculated “As Built” Geometric Properties
IV. Conclusion
A 24-gore model design, based on a conservatively-predicted steady-state drag force of 1237 lbf (5502N)
and a minimum structural safety factor of 1.5 was described. The tested efficiency of all joints exceeded those
assumed for design purposes, while averaged drag forces experienced during high-q runs were less than predicted,
ranging from 900 to 950 lbf (4003 to 4226 N). Sections removed from post-test damaged models showed that the
strength of suspension lines and radial joints had been degraded from 18% to 39% by high-dynamic-pressure tests
plus subsequent exposure to airflow during tunnel shut-down. Vent line connections to a vent ring were degraded
70% to 100%. When measured under 10-lbf tension, dimensional variations within and between nine new
parachutes were found to be small. Based on detailed radial and circumferential measurements of 5 units, the asmanufactured equivalent nominal diameter of the models ranged from 31.46 to 32.06 in. (0.80 to 0.81 m).
Geometric porosity (the open area of gap plus vent divided by canopy nominal area) ranged from 12.25 to 12.55%
for new parachutes, and was 13.54% for one used parachute.
Acknowledgments
The authors would like to acknowledge the efforts of all of the technicians and support personnel at Pioneer
Aerospace who helped make these high fidelity models possible, as well as Anthony Levay in support of this test
program. Additionally, everyone at JPL, NASA GRC, and NASA Ames who worked on the various phases of the
test entries associated with these models. This paper describes Pioneer Aerospace support to the Jet Propulsion
Laboratory, California Institute of Technology, under a contract with the National Aeronautics and Space
Administration.
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