Typical entry vehicle aeroshells are limited in size by the launch vehicle shroud. Inflatable aerodynamic decelerators allow larger aeroshell diameters for entry vehicles because they are not constrained to the launch vehicle shroud diameter. During launch, the hypersonic inflatable aerodynamic decelerator (HIAD) is packed in a stowed configuration. Prior to atmospheric entry, the HIAD is deployed to produce a drag device many times larger than the launch shroud diameter. The large surface area of the inflatable aeroshell provides deceleration of high-mass entry vehicles at relatively low ballistic coefficients. Even for these low ballistic coefficients there is still appreciable heating, requiring the HIAD to employ a thermal protection system (TPS). This TPS must be capable of surviving the heat pulse, and the rigors of fabrication handling, high density packing, deployment, and aerodynamic loading. This paper provides a comprehensive overview of flexible TPS tests and results, conducted over the last three years. This paper also includes an overview of each test facility, the general approach for testing flexible TPS, the thermal analysis methodology and results, and a comparison with 8-foot High Temperature Tunnel, Laser-Hardened Materials Evaluation Laboratory, and Panel Test Facility test data. Results are presented for a baseline TPS layup that can withstand a 20 W/cm 2 heat flux, silicon carbide (SiC) based TPS layup, and polyimide insulator TPS layup. Recent work has focused on developing material layups expected to survive heat flux loads up to 50 W/cm 2 (which is adequate for many potential applications), future work will consider concepts capable of withstanding more than 100 W/cm 2 incident radiant heat flux. This paper provides an overview of the experimental setup, material layup configurations, facility conditions, and planned future flexible TPS activities.

The successful flight of the Inflatable Reentry Vehicle Experiment (IRVE)-3 has further demonstrated the potential value of Hypersonic Inflatable Aerodynamic Decelerator (HIAD) technology. This technology development effort is funded by NASA's Space Technology Mission Directorate (STMD) Game Changing Development Program (GCDP). This paper provides an overview of a multi-year HIAD technology development effort, detailing the projects completed to date and the additional testing planned for the future.The effort was divided into three areas: Flexible Systems Development (FSD), Mission Advanced Entry Concepts (AEC), and Flight Validation. FSD consists of a Flexible Thermal Protection Systems (FTPS) element, which is investigating high temperature materials, coatings, and additives for use in the bladder, insulator, and heat shield layers; and an Inflatable Structures (IS) element which includes manufacture and testing (laboratory and wind tunnel) of inflatable structures and their associated structural elements. AEC consists of the Mission Applications element developing concepts (including payload interfaces) for missions at multiple destinations for the purpose of demonstrating the benefits and need for the HIAD technology as well as the Next Generation Subsystems element.Ground test development has been pursued in parallel with the Flight Validation IRVE-3 flight test. A larger scale (6m diameter) HIAD inflatable structure was constructed and aerodynamically tested in the National Full-scale Aerodynamics Complex (NFAC) 40ft by 80ft test section along with a duplicate of the IRVE-3 3m article. Both the 6m and 3m articles were tested with instrumented aerodynamic covers which incorporated an array of pressure taps to capture surface pressure distribution to validate Computational Fluid Dynamics (CFD) model predictions of surface pressure distribution. The 3m article also had a duplicate IRVE-3 Thermal Protection System (TPS) to test in addition to testing with the Aerocover configuration. Both the Aerocovers and the TPS were populated with high contrast targets so that photogrammetric solutions of the loaded surface could be created. These solutions both refined the aerodynamic shape for CFD modeling and provided a deformed shape to validate structural Finite Element Analysis (FEA) models.Extensive aerothermal testing has been performed on the TPS candidates. This testing has been conducted in several facilities across the country. The majority of the testing has been conducted in the Boeing Large Core Arc Tunnel (LCAT). HIAD is continuing to mature testing methodology in this facility and is developing new test sample fixtures and control methodologies to improve understanding and quality of the environments to which the samples are subjected. Additional testing has been and continues to be performed in the With the successful completion to the IRVE-3 flight demonstration, mission planning efforts are ramping up on the development of the HIAD Earth Atmospheric Reenty Test (HEART) which will demonstrate a relevant...

Stress-rupture tests were conducted in air, under vacuum, and in steam-containing environments to identify the failure modes and degradation mechanisms of a carbon-fiber-reinforced silicon carbide (C/SiC) composite at two temperatures, 600°a nd 1200°C. Stress-rupture lives in air and steam-containing environments (50 -80% steam with argon) are similar for a stress of 69 MPa at 1200°C. Lives of specimens tested in a 20% steam/argon environment were about twice as long. For tests conducted at 600°C, composite life in 20% steam/argon was 30 times longer than life in air. Thermogravimetric analysis of the carbon fibers was conducted under conditions similar to the stress-rupture tests. The oxidation rate of the fibers in the various environments correlated with the composite stressrupture lives. Examination of the failed specimens indicated that oxidation of the carbon fibers was the primary damage mode for specimens tested in air and steam environments at both temperatures.

The failure sequence following crack formation in a chevron-notched four-point bend specimen is examined in a parametric study using the Bluhm slice synthesis model. Premature failure resulting from crack formation forces which exceed those required to propagate a crack beyond O,' min is examined together with the critical crack length and critical crack front length. An energy based approach is used to establish factors which forecast the tendency of such premature failure due to crack formation for any selected chevron-notched geometry. A comparative study reveals that, for constant values of a 1 and ~o, the dimensionless beam compliance and stress intensity factor are essentially independent of specimen width and thickness. The chevron tip position a0 has its primary effect on the force required to initiate a sharp crack. Small values for ao maximize the stable region length, however, the premature failure tendency is also high for smaller a0 values. Improvements in premature failure resistance can be realized for larger values of c~0 with only a minor reduction in the stable region length. The stable region length is also maximized for larger chevron base positions, a l, but the chance for premature failure is also raised. Smaller base positions improve the premature failure resistance with only minor decreases in the stable region length. Chevron geometries having a good balance of premature failure resistance, stable region length, and crack front length are 0.20 ~< a0 ~< 0.03 and 0.70 ~< at ~< 0.80. Nomenclaturea A aO = ao/W o~ 1 = al/W O~min B ~min = amin--~O cel--o~ 0 Cs( ) Cv( ) = crack length = Chevron crack area = dimension crack length = dimensionless tip position of the chevron ligament = dimensionless base position of the chevron ligament = dimensionless crack length at dG(a)/da = 0 = thickness of the bend beam = dimensionless crack width at dG(a)/da = 0 = dimensionless straight-through crack specimen compliance, (SE~B)/P = dimensionless chevron-notched speciment compliance, (6EIB)/P Ao~mi n = O~mi n -o~ 0 = dimensionless subcritical region length increment 6 = load-line displacement Aa = a -ao = dimensionless crack growth increment E' = plane stress, E = E I, plane strain, E ~ = E/(1 -u2), elastic modulus n 12 Af~ P P S~ $2 Sp SD O'cr 0 A U W Ys( ) Yv( ) z=B---7~ A.M. Calomino and L.J. Ghosn = elastic energy release rate = critical elastic energy release rate = shear correction term for Bluhm slice model = mode I stress intensity factor = critical mode I stress intensity factor = number of slices through cross section for Bluhm model = Poisson's ratio = fracture energy exhausted after crack formation and arrest = force applied to beam at the load-line = notch radius at chevron tip = load span distance = support span distance = fixed-force overload resistance factor = fixed-displacement overload resistance factor = critical stress for crack formation at chevron tip position = base angle of chevron cross-section = change in strain energy after crack formation and arrest = width of bend specimen = di...

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