ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
Steel) using a thermally conductive epoxy. The second concept attempted to expose the RTD directly to  the lunar regolith and thermally isolate the sensor from the rest of the drill string (Figure 5, bottom).   Figure 5. RTD placement within the TRIDENT drill string ( top – 2016 design; bottom  – 2017 design).  There was a noticeable improvement in response time for the second concept, but the assembly of this  option is more complex. The increased assembly complexity for the second concept is because the  assembly requires bonding a thermal isolator to the base metal using a thermally insulating epoxy and then  bonding the RTD to the insulation using a thermally conductive epoxy. The goal here is ultimately to  minimize the mass or thermal inertia of the temperature sensor (including the epoxy used to bond and  protect the sensor). Though improvements to the design / assembly can still be made, the initial results  show great promise. When comparing the two concepts over a similar operational scenario, it is apparent  that the rate of change measured with the RTD is much faster in the 2017 tests (~8°C / min) compared to  the 2016 tests (~2°C / min) (Figure 6).   In tests performed in 2016 and 2017, the average regolith temperature was < -100°C with the ambient  chamber temperature being < -50°C and atmospheric pressure < 10-5 Torr. In both cases, the NU-LHT 3M  lunar regolith had 5% weight water added to simulate expected values from LCROSS data. Similar drilling  profiles were performed for these tests though the Rate of Penetration (ROP) during the 2017 tests were a  little faster than in the 2016 tests. This means that the 2016 tests were putting more energy into the regolith  per cm of regolith displaced. During both tests shown here, the power required to displace the regolith was  comparable and on average only a few Watts. It is unlikely that the additional energy used during the 2016  test would have account for the large differences in measured temperatures between the two tests. The  2017 test actually started with the drill bit approximately 20°C warmer than in the 2016 test, but was more  than 15°C colder by the time a 40-cm depth was reached. Another interesting observation is that the 2016  test never measured a temperature colder than the initial starting temperature of approximately -65°C  whereas the sensor in the 2017 test ended up measuring temperatures less than -105°C shortly after  reaching a 75-cm depth.  NASA/CP—2018-219887 18
Figure 6. Drill Bit Temperature vs. Depth (2016 & 2017- Original & Modified Design).  There are ongoing analyses being performed to help predict steady state temperatures from the response  of the temperature sensors over a relatively short time duration. Vetted models to predict temperatures are  required as extrapolation is almost certainly necessary to estimate the actual temperature of the lunar  regolith at various depths. Currently the time required for the sensor to reach a steady state temperature  measurement is on the order of several hours. Improvements to the sensitivity of the temperature sensor  in the form of improved thermal isolation from the drill bit, reduced mass, and reduced thermal resistance  between the sensor and the regolith will result in better estimations of the regolith temperature.  Slip-Ring Design  To enable the use of an embedded temperature sensor in the drill bit, it’s necessary to use an electrical  slip-ring to convey the electrical signal through the rotating auger joint. The drill design preceding TRIDENT  was originally designed and built under a different NASA program targeting more terrestrial field testing [4].  In the early stages of the RP program, the drill, then named RP15, underwent random vibration testing to  qualification levels. The drill mechanisms all survived and continued to function nominally, except for the  slip ring. During the Z axis vibration test (in-axis with the drill bit), the off-the-shelf slip-ring experienced  separation between the rotor and the stator at a bonded interface, as shown in Figure 7. Certainly, this  commercial-off-the-shelf (COTS) slip-ring was not designed for this environment and definitely not the large  response observed at ~110 Hz (Figure 8). Though the COTS slip-ring did survive several hours of testing  on the rotary percussive drill system up to this point, the random vibration environment proved to be more  than the bonded joint could handle. Given this failure, it was apparent that for the TRIDENT drill design,  slip-ring design and/or selection was going to require a more concerted effort. -150-100-50050100150200250300350400450500550600650700750800 -110-105-100-95-90-85-80-75-70-65-60-55-50-45-40Time Elapsed 2017 (HH:MM:SS) Drill Location (abs mm)Bit T emperature (Celcius) Time Elapsed 2016 (HH:MM:SS)Bit Temperature & Drill Location v.s. Time Elapsed   2016 GRC VF-13 Lunar Chamber Soil Bin #2, Hole #3  2017 GRC VF-13 Lunar Chamber Soil Bin #3, Hole #2Bit Temperature 2016 (Celcius) Bit Temperature 2017(Celcius) Drill Location 2016 (abs mm) Drill Location 2017(abs mm) 2016 Temp.  Response 2017 Temp.  Response NASA/CP—2018-219887 19
Figure 7. Slip-ring damage from system random vibration test ( left – undamaged Slip-ring before test;  right – damaged Slip-ring after test).  Figure 8. Response of Drill Head Accelerometer during Z-axis Random Vibe.  After researching availability of slip-rings that would meet the form, fit, and function for the TRIDENT drill,  it became evident that a custom solution was going to be necessary. To help save on time and costs, the  solution for TRIDENT was to design and build the slip-ring using in-house expertise. Fortunately, there  existed a heritage design that closely fit the size and signal requirements for TRIDENT. This was a  significant benefit to reduce the time required for materials selection, coatings, sizing, etc. However, the  heritage process used for space-flight applications to build up the rotor, secure the internal rings in place,  and isolate the rings (or channels) from each other is highly involved and quite expensive. The more  traditional process was not a financially viable option given the budget available to produce and test the  TRIDENT Engineering Test Unit (ETU). Fortunately, this problem had a relatively simple solution which was  to use Vespel spacers to isolate and constrain the conductive rings (Figure 9).  Separated  Rotor NASA/CP—2018-219887 20
Figure 9. Slip-ring assembly overview ( left – assembly cross section showing Vespel spacers;   right –  final assembly).  Once this solution was realized, the design and fabrication of the hardware was completed. The slip-ring  assembly was then assembled and integrated into the TRIDENT ETU. To date, the TRIDENT slip-ring  continues to perform as designed (the 2017 temperature response from Figure 6 is evidence of its  functionality). This mechanism, along with the rest of the TRIDENT drill, went through the same vibrational  test procedure as the RP 15 drill. However, improved designs for launch lock concepts greatly reduced the  response at the drill head (Figure 10). It’s conceivable that the COTS slip-ring may have survived the  significantly reduced loading experienced at the drill head during this test. However, other environmental  factors (temperature, atmospheric pressure, etc.) also dictate the necessity of a custom slip-ring for this  application.  Figure 10. Improved fixtures / launch lock concepts for the TRIDENT drill in random vibration testing  (left – drill on vibration table; right  – significantly reduced response at the drill head).  NASA/CP—2018-219887 21
Passive Brush for Sample Transfer  One of the key systems in ensuring enough sample is delivered to the OVEN subsystems is a passive  brushing system designed to help remove regolith from the drill auger and direct it to a chute leading to the  OVEN crucible. The RP environmental requirements have greatly reduced the trade space for materials  that can be used in the passive brush. To protect against electrostatic charge build up, the brush material  had to be conductive while also remaining flexible at cryogenic temperatures. The effectiveness of the brush  at removing regolith from the auger is dependent on a number of factors which include: cohesiveness of  regolith, size of brush bristles, length of brush bristles, stiffness of  material used for the brush bristles, wear rate of the bristle material,  and position of the brush with respect to the auger.   The trade space studied included brass brushes of various bristle  and brush diameters, Tampico, Thunderon, and conductive Nylon  66 (Table 1). Regolith used for the testing included JSC 1A and NULHT 2M Lunar regolith. Criteria used to evaluate the various brush  materials included electrical conductivity; brush life expectancy,  and effectiveness of cleaning debris from the drill string.  Observations from the testing were:  Bronze offered excellent electrical conductivity & was the most effective cleaner, but experienced significant wear after only a single hole which has a significant effect on the ability to clean the drill string for subsequent drilled holes. Thunderon fibers were too small to be an effective cleaner. Tampico cleaned sufficiently well, had sufficient life expectancy, but because it is a natural fiber, it will have more variance in conductivity. Conductive Nylon 66 cleaned sufficiently well, had sufficient life expectancy, and excellent conductivity. Table 1. Brush materials and physical dimensions.  Material Bristle Diameter  mm (in) Outside Diameter  mm (in) Brush Thickness  mm (in)  Bronze 0.152 (0.006) 63.5 (2.5) 9.53 (0.375)  Bronze 0.356 (0.014) 63.5 (2.5) 9.53 (0.375)  Bronze 0.356 (0.014) 76.2 (3.0) 15.88 (0.625)  Tampico 0.203 (0.008) 63.5 (2.5) 9.53 (0.375) Tampico 0.203 (0.008) 76.2 (3.0) 9.53 (0.375) Thunderon 0.051 (0.002) 63.5 (2.5) 9.53 (0.375) Conductive Nylon 66 0.203 (0.008) 76.2 (3.0) 9.53 (0.375) Though test results thus far have been instructive in selecting the appropriate brush parameters, there is  still much that can be improved. Other items to investigate are: alternative brush materials (other than what  has already been tested - i.e. titanium) and alternative brush mounting configurations. The end goal driving  the selection is ultimately based on the volume of material that the brush helps deliver to the instruments  and the consistency of the volume delivered. Figure 11. Keyed scraper for  IceBreaker (a TRIDENT  predecessor drill) [5].  NASA/CP—2018-219887 22
Traditional brushes were not the only thing considered to remove material from the auger flutes. A springloaded and rotating scraper design was also considered (Figure 11). This design was used on a preceding  drill design and was largely more effective at removing more cohesive cuttings than traditional brushes. A  prototype design was built and tested for the RP drill (Figure 12). One challenge in this case was designing  the rotating scraper to interact with both fine pitch and coarse pitch flutes on the auger. For the TRIDENT  drill, this variable pitch combination is necessary to help retain cuttings / regolith near the bit (fine pitch) for  delivery to other instruments and more efficiently transfer cuttings / regolith through the remaining auger  length (coarse pitch). Though the concept would work well while operating in either section of the auger  (fine or coarse pitch), the transition point between the fine and coarse pitch flutes was a failure point for this  mechanism. The primary reason causing the failure is that the coarse pitch flute requires the scraper to  rotate twice as fast as the fine pitch flute because the pitch of the coarse flute is twice that of the fine pitch  flute. This discrepancy could not be easily reconciled for this sample transfer approach.  Figure 12. Single Rotating Scraper design ( left – scraper interacting with fine pitch flutes; right – scraper  interacting with coarse pitch flutes).  Although the brush has an obvious direct impact in the performance of the sample cleaning system, the  mechanism supporting the brush is equally important. For both the 2016 and 2017 tests, the same  conductive nylon brush design was used. However, there were significant performance differences. In 2016,  a single brush lasted the entire test phase with minimal wear. In 2017, it was apparent that the brush  stopped rotating as significant wear on only one part of the brush was observed after only a few tests (after  which the brush had to be replaced). Two factors likely led to this change: one was an offset on the axis of  the brush in the 2017 test to bias the brush towards one side of the auger; the other factor was a design  modification in the 2017 design to help better protect the brush bearings by adding a seal between the  brush housing cavity and the bearings for each bearing (Figure 13).   Figure 13. TRIDENT brush mechanism cross-sections (left - 2017 TRIDENT brush mechanism;  right – 2016 RP15 brush mechanism).  NASA/CP—2018-219887 23
Adding the axial offset in the brush was originally intended to help improve removal of cuttings / regolith  from the drill string. However, this offset led to periodic pinching of bristles between the auger and auger  tube which would ultimately pull out or shear bristles in the brush. If this happens enough in one section of  the brush, there is no longer enough bristle engagement with the auger to cause the brush to rotate. Of  course, this also reduces the ability of the brush to clean the auger.  Using seals to better protect the bearings also had a negative  impact on the brush. For the 2016 configuration, the design  was more simplistic and only relied on bearing shields to  protect the bearings. Though this wasn’t ideal, given the  proximity of the bearings to regolith dust, these bearings had  never failed – low speeds, cycles, and loads minimized the  chances of failure in the bearings. The seals added in the  2017 design were entirely precautionary, but the increased  drag from the seals made brush rotation more difficult and  contributed to a reduction in the life of the brush Design improvements around the brush and brush support are  both necessary to help improve sample transfer efficiency.  One design that can be leveraged to support the brush may  be the design built to support the Atacama Rover Astrobiology  Drilling Studies (ARADS) project [8]. This particular design  (Figure 14) utilizes a rotating scraper (the ARADS drill has a  constant auger pitch) that is supported on a cantilevered arm.  The scraper could be replaced with a brush and the  mechanism can be simplified due to the more compliant  nature of a brush compared to the more rigid scraper (i.e.  spring loaded arm is not required). Figure 14. Alternative Support to use  for brush (scraper shown here).  NASA/CP—2018-219887 24
protect the drill and other subsystems from Electromagnetic Interference (EMI) and reduce signal crosstalk. The mass of the cable harnesses was also reduced by 85% (from 7 lb to 0.5 lb (3 kg to 0.2 kg)) in  comparison to the RP15 Drill (Figure 16 and Figure 17).   The FPC assembly has proven to be extremely robust since the base dielectric and coverlay material of  are all made from Polyimide (e.g. Kapton). Not only does this perform well in cryogenic application, it also  provides excellent electrical insulation and low outgassing profile in the vacuum environment.  Figure 16. Side by side comparison of RP15 cable harness with TRIDENT ETU flex cable harness.Figure 17. Reduced bulkiness and mass  from RP15 (left) to TRIDENT (right).  References  1. D. Andrews, et al. “Introducing the Resource Prospector (RP) Mission”, AIAA SPACE 2014 Conference and Exposition, AIAA SPACE Forum, (AIAA 2014-4378) 2. A. Colaprete, et al. "Detection of Water in the LCROSS Ejecta Plume", Science. 330 (6003): 463–468. 3. D. Hurley, et al “Lunar Polar Volatiles: Assessment of Existing Observations for Exploration.” https://ssed.gsfc.nasa.gov/dream/docs/PolarVolatiles_HEOMD_Hurley.pdf. Accessed October 4, 2017. 4. K. Zacny, et al. “Life in the Atacama: The Drill and Sample Delivery System.” 44th Lunar and Planetary Science Conference (2013). 5. Paulsen et al., “Testing of a 1 meter Mars IceBreaker Drill in a 3.5 meter Vacuum Chamber and in an Antarctic Mars Analog Site.” AIAA Space 2011. 6. Teijin Aramid, “Technora Catalog,” Technora datasheet. https://www.teijin.com/products/ advanced_fibers/aramid/contents/aramid/technora/pdf_1/Technora_catalog%28E%29.pdf. Accessed February 2, 2018. 7. S. Nagihara, et al. “Improved Data Reduction Algorithm for the Needle Probe Method Applied to InSitu Thermal Conductivity Measurements of Lunar and Planetary Regoliths.” Planetary and Space Science, Volume 92, p. 49-56. 8. D. Waller. “Mars rover tests driving, drilling and detecting life in Chile's high desert.” https://phys.org/news/2017-03-mars-rover-drilling-life-chile.html. Accessed January 31, 2018. NASA/CP—2018-219887 26
KaRIn Alignment Mechanism Design, Development and Testing  John Luke Wolff * Abstract  The Ka-Band Radar Interferometric System (KaRIn) will characterize the ocean mesoscale/sub-mesoscale  circulation and provide a global inventory of all lakes, reservoirs, wetlands, and major rivers. There are two  identical Alignment Mechanisms (AM) mirrored on the instrument to correct KaRIn Azimuth radar pointing  errors. Each AM is a one-degree-of-freedom precision mechanism. The mechanism must meet challenging  thermoelastic distortion requirements over a wide temperature range. The flight design was qualified in less  than two years, and the flight hardware is currently in acceptance testing. Discussed are key design details  along with supporting development and qualification test results.   Introduction  The KaRIn instrument uses radar interferometry to perform science measurements, which dictates tight  alignments between the respective radar elements. The instrument requires a series of deployments, as  shown in Figure 1. The AM enables the KaRIn instrument to correct residual post-deployment antenna  pointing errors. The choice to maintain pointing passively imposed challenging thermoelastic distortion  requirements on the AM. To simplify the mechanism overall complexity, launch locks are not used. Each  AM consists of a Mechanism Strut Assembly (MSA) and Rotation Flexure Assembly, shown in Figure 2.  The Rotation Flexure Assembly contains a hexfoil rotary flexure, which acts as the mechanism pivot axis  for the one-degree-of-freedom commanded motion. The mechanism pivot axis is configured such that  rotations about the pivot axis are linearly proportional to KaRIn Azimuth angle. The MSA consists of singleblade and coaxial double-blade flexures to create a kinematic five-bar linkage configuration, with the  diagonal strut element containing a Linear Actuator Assembly (LAA). The LAA can lengthen or shorten via  a geared stepper motor with an ACME leadscrew output. This linear length change in the diagonal strut  causes a parallelogram effect in the MSA. The parallelogram effect results in primarily a rotation about the  mechanism pivot axis. The small extraneous angular motion in the orthogonal axis is acceptable. The key  design requirements are summarized in Table 1.  Table 1. Summary of Key Design Requirements  Mechanical Environments 0.0 kHz to 0.1 kHz: Sine Input, Peak 38G, All 3-Axes  0.1 kHz to 2.0 kHz: Random Input, 6.8 G rms, All 3-Axes  Note: Structurally support additional mass (NTE 4.6 kg)  Thermal Environments Non-Operation: -95°C to +100°C  Operation: -35°C to +100°C  Note: Operation limit includes motor self-heating  Thermoelastic Distortions Relative Azimuth: 2 mdeg (1-sigma)  Note: Based on on-orbit temperature predictions  Commanded Motion Accuracy: <6% of commanded motion  Resolution: 0.6 mdeg  Range of Motion: 150 mdeg to 250 mdeg  Note: Step rate chosen to be 10 Hz based on electronics  *Jet Propulsion Laboratory, California Institute of Technology, Pasadena, CA Proceedings of the 44th Aerospace Mechanisms Symposium, NASA Glenn Research Center, May 16-18, 2018 NASA/CP—2018-219887 27
Figure 1. KaRIn Antenna Deployment  Figure 2. Alignment Mechanism Configuration Design Overview for Thermoelastic Distortions  KaRIn Azimuth thermoelastic distortions, caused by the MSA, are the primary concern due to the  instrument’s strong sensitivity to this parameter. Therefore, this paper focuses solely on those thermoelastic  distortions. The MSA 1D-thermoelastic distortion model, see Eq. 1, has been validated by FEM (Finite  Element Model) and by testing, which is discussed later in this paper. Eq. 2 is the uncertainty in the model.  For each key term, the partials are shown in Figure 3. Summarized in Table 2 are the model uncertainties  and associated mitigation approaches. The dominant error sources are the coefficient of thermal expansion  (CTE) uncertainty1 and 3D thermoelastic distortions due to bolt slip.   1Includes hysteresis between thermal cycles, measurement error, and unit-unit variability.  NASA/CP—2018-219887 28
Coefficient of Thermal Expansion (CTE)  To use reference sources ([1], [2]) for instantaneous CTE over temperature, the thermoelastic distortion  model would need to incorporate unacceptably high CTE uncertainties. To mitigate this risk, all flight piece  parts, except the LAA, were fabricated from controlled slabs of material whose CTE was characterized per  ASTM E2892. To build confidence in the CTE over temperature for the LAA, a flight-like LAA was built, with  the CTE measured per ASTM E289. Due to difficulties controlling all of the materials in the LAA, the team  continued to assume a large CTE uncertainty for the LAA until the flight LAAs were each measured. In  Figure 5, the measured instantaneous CTE over temperature is compared to several reference sources.  All CTE uncertainties used in the model are summarized in Table 3.  Thermoelastic Distortion due to Bolt Slip  As the temperature is changed for CTE mismatched materials bolted together, radial strain increases until  joint friction due to bolt preload is overcome, whereby radial strain projects into axial distortion errors. Where  possible, this error source was eliminated by adding radial flexures to one side of the bolted joint to eliminate  radial coupling at the bolted joint. At all other bolted joint locations, a low friction surface coating was applied  to improve our confidence in the design and the repeatability. At these bolted joints, thermoelastic distortion  uncertainty was empirically estimated by varying in the FEM the radial stiffness at each bolted joint location  for a 10°C bulk soak case. Then, the FEM results were linearly scaled based on closed form solutions  estimating the maximum temperature change required for each bolted joint to slip.   =1(⁄ )−ℎℎ− (1)  ()=1[(⁄ )({})]+ℎ({ℎ}) +[{}](2)  Figure 3. Mechanism Strut Assembly Thermoelastic Distortions Key Terms  Table 2. Thermoelastic Distortion Uncertainty Error Sources  Uncertainty  Sources Mitigation – Design Intent Development Test  (Prior to Mechanism Level Test)  Temperature  (Gradient) Top-to-bottom and left-to-right, materials  chosen are nearly symmetric. Minimize  CTE by utilizing High Purity (HP) Invar363 where possible. See Figure 4. None Temperature   (Bulk)  Coefficient of  Thermal  Expansion Metallic hardware and bolted joints to  improve confidence in incorporating test  results to model predicts. Characterize CTE for flight material  and prototype LAA. Final verification,  characterize CTE for each flight LAA.  3D Effect   (due to bolt slip) Radial flexures where configuration allows  OR low friction surface coating (SF-2). None  Linear Lengths Control fabrication/assembly tolerances. Inspect part/assembly tolerances.  2Five coupons machined from each slab of material, with three thermal cycles measured per coupon.  3HP Invar36 exhibits very good dimensional stability properties. While not discussed further in this paper,  HP Invar36 is critical to the mechanism temporal stability performance [4].  NASA/CP—2018-219887 29