ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
The first failure for the bearings from the second manufacturer was also an inner race spall of one  bearing. The total test time before failure was 27 hours. Again, the root cause is uncertain, but due to the short fatigue life of the bearings and the similarity in spall characteristics, the failure most likely initiated in the same fashion for the two manufacturers’ bearings.   The test that produced the failure was restarted. The Si 3N4balls were reused and a new pair of inner and  outer races was replaced in the failed bearing. The goal again was to observe further progression of the  ball scars. The additional test time before the next failure was 15 hours. Combined with the first test, the  total test time was 42 hours for all the Si 3N4balls and the original bearings. Failure for the restarted test  was an inner race spall of one of the original bearings. The root cause of the failure of the restarted and  initial tests is likely the same because both inner race spalls looked very similar. All RCF tests resulted in inner race spalls with a similar characteristic “surface initiation site” pattern  (Figure 9). The spalls were most likely caused by the stress riser from rolling over the scar features on the  Si3N4balls.  The reduction in the number of scars by 75% (64 scars to 16 scars) delayed, but did not eliminate, the  onset of damage leading to inner race spall. Figure 10 shows the SEM images of the resulting damage  that occurred after 120 hours and 668 hours of operation. The failure after 120 hours is perhaps the earliest initiation of a spall that has ever been captured before any significant material had been expelled. Our experience in setting the vibration limit of the RCF testers enabled us to capture the initiation stage. The feature that caused the second failure after 668 hours resembles the spalls previously discussed.  Scanning electron microscopy coupled with our experience in setting the vibration limit of the RCF testers enabled us to capture the progression of damage leading to inner race spalls in a RCF test. The root cause of the main spall is still unclear, but a few possibilities based on post-mortem  characterization are: 1) debris shedding from Si 3N4ball, 2) subsurface void in inner race steel because  the spall is uncharacteristically deep for typical debris-induced damage, and 3) scar-induced localized  stress riser on surface of steel races. The first scenario is unlikely because the RCF testing of “small”  scars (Test #5 in Table 3), which is still in progress with over 4500 hours, would also have shed debris.  Furthermore, the lubricant oil in the testers is constantly filtered. The latter scenario is the most likely explanation because of the observation of scar witness marks in the raceway and insights from finite-element modeling showing high subsurface stress of the raceway near the edge of the scars. Figure 9.  Inner race spalls of hybrid bear ings with “large” scars on every ball. 137
Figure 10.  Inner race spalls of hybrid bearings  with “large” scars on only 2 balls in each bearing. First failure occurred after 120 hours (left), a nd second failure occurred  after 668 hours (right). Progression of Scars on Si 3N4Balls At the conclusion of each test, SEM was used to document the progression of the scars on the Si 3N4 balls, and to assess the resulting damage to the bearings and identify the potential cause of the failure. Figure 11 shows the typical progression of a scar after testing, which provides two significant findings. First, all the scars on the Si 3N4balls (64 scars in 4 bearings) had been run over, as evidenced by the  apparent halo around and widening of each scar. This key finding suggests that, contrary to initial assumption, balls with scars do not track or maintain their orientation inside a bearing. This finding is  further supported by the lack of a wear band on each ball, regardless of the number of hours in operation, which is typically observed in unscarred balls after RCF testing. The reason for the lack of ball tracking could be due to the ball being pivoted each time the scar enters the contact patch. Another reason could simply be because at such a high rate of rotation, any minute change in an otherwise pristine surface of the ball creates a misalignment between the ball’s center of mass and its axis of rotation, leading to rotational imbalance. Figure 11.  Representative artificially-generated scar: before (left) and after (right) RCF testing.  This particular test lasted 28 hours at 5000 rpm with a mean contact stre ss of 330 ksi (2.3 GPa).  The second finding is the burnishing of the scars during testing. From Figure 11, it is apparent that the  peaks within the scars have either been flattened or completely removed, leaving behind only the valleys and remnants after material removal. 138
Pre-Test Scar Profile: (left) along horizontal centerline, (right) along vertical centerline Post-Test Scar Profile: (left) along horizontal centerline, (right) along vertical centerline Figure 12.  Optical surface profilometry trace of a scar before (top) and after (bottom) RCF testing Further investigation of the apparent burnishing of the scars was conducted using optical (white-light)  surface profilometry before and after testing (Figure 12), which provides a quantitative comparison of the  roughness and morphology. Traces were taken near the major and minor axes of the scar ellipse, and the corresponding profiles are shown in the plots in the left column and right column, respectively. As the roughness decreased, the scar effectively became more benign, which could suggest that most of theraceway damage occurred very early in testing. Data Analyses and Statistical Methods Data Processing and Distribution Models The purpose of data analysis is to estimate the distribution of strength, the distribution of fatigue life, and  to, if possible, combine the two distributions to create a reliability forecast chart that accounts for the effects of load and service life. The cumulative plotting position used here is the median rank estimate [8] given by: /g41/g3/g32/g3/g11/g77/g3/g16/g3/g19/g17/g22/g12/g18/g11/g49/g3/g14/g3/g19/g17/g23/g12/g3/g32/g3/g20/g3/g177/g3/g54 ,          ( 1 ) where/g41is the cumulative failure probability and is the complement of /g54= 1 - /g41,/g77is the rank number, and /g49is the sample size. The model for rolling contact fatigue life that accounts for both load level and fatigue life is the bivariate  model [9] given by: /g54/g3/g32/g3/g11/g54 /g85/g12/g13/g13/g11/g53/g18/g53 /g85/g12/g80/g11/g87/g18/g87/g85/g12/g69,           ( 2 ) where /g54is, again, the cumulative survival probability, /g53is the applied ball contact stress scaled by /g53/g85the  reference strength /g15/g3/g87is the observed fatigue life scaled by /g87/g85the reference life (such as the median life), and/g80/g3and /g69are the Weibull modulus for strength and life distributions, respectively. Subscript /g85denotes  reference scaling values for cumulative survival, strength and fatigue life, typically taken from the  observed data and representing median reference values wherever possible. Solving for /g53gives: /g47/g82/g74/g3/g11/g53/g12/g3/g32/g3/g47/g82/g74/g3/g11/g53 /g85/g12/g3/g16/g3/g69/g18/g80/g3/g47/g82/g74/g3/g11/g87/g18/g87 /g85/g12/g3/g14/g3/g11/g20/g18/g80/g12/g62/g47/g82/g74/g11/g62/g20/g18/g54/g12/g18/g11/g20/g18/g54 /g85/g12/g64/g17      ( 3 ) 139
Equation (3) produces a family of straight lines as shown in Figure 15A, constant reliability contours, of  slope /g16/g69/g18/g80/g3 on/g47/g82/g74/g11/g53/g12 versus/g47/g82/g74/g11/g87/g12/g3 coordinates, provided that the distributions of strength and fatigue life  are each consistent with Weibull distributions having time-invariant shape factor. However, this is not the  case for the RCF data examined here, and some modification of the model is explored. Following the line in [9], this equation is modified to account for augmented degradation to the bearings: /g53/g3/g32/g3/g20/g3/g16/g3/g73/g62/g20/g3/g16/g3/g72/g91/g83/g11/g87/g18/g87 /g70/g12/g64/g3/g135/g3/g53/g85/g11/g87/g85/g18/g87/g12/g69/g18/g80/g62/g47/g81/g11/g20/g18/g54/g18/g47/g81/g11/g21/g12/g64/g11/g20/g18/g80/g12,        (4)/g3/g3 where the reference survival probability /g54/g85/g3= 0.5 and parameters /g73and/g87/g70represent the fraction of strength  being reduced by a first order corrosion or degradation process with time constant /g87/g70. This equation  underlies the probability contours of Figure 15B, with parameters shown in the text box. This graph  suggests that, within perhaps the decade of time beyond the observed data for the parameters chosen, that the life is predicted to drop.. Substantial amount of data, at low load levels and relatively long lives, are required to get an approximate value for the turn-down model parameters. We would also caution that though such degradation has been observed in glass composite materials [9], we are not aware of any reports of similar degradation in silicon nitride bearings in ambient or vacuum environments. That is, it is not clear if the turn-down is real or of practical significance. Ball Strength Distribution The strength distribution shape factor estimate, /g80, was obtained from experimental data using  Aerospace-developed strength tester [10] of Si 3N4balls. Although based on a partial sample of the first  17 out of 131 specimens, this /g80/g3/g32/g3/g27 is a very good data fit, especially since it conforms to the weakest  data (Figure 13). The estimated median strength is 725 ksi (5.0 GPa).  Figure 13. Strength distribution of Si 3N4balls Rolling Contact Fatigue Tests of As-Manufactured Hybrid Bearings The individual data (first of four) for sizes 101 and 304 are collected and shown in Figure 14. The  similarity of the distribution shapes allows these data to be pooled, forming a mixed population distribution. Unlike the expected inflection near the central values of life, these data show a monotonic  increase in failure probability to the longest observed life. Fortunately the service life design is defined by  the short observed lives, and, while the distribution at long life is of academic interest, it has little impact  on the practical design operating life. The strength shape factor, /g80/g3= 8/g15/g3and life shape factor /g3/g69= 0.55 were combined to create a Reliability Forecast Chart shown in Figure 15, that gives a projected view for  the combination of load, life, and reliability at very long service lives. 140
Figure 14. Individual probability distributions (le ft) and combined RCF lifetim e distribution (right)  of as-manufactured hybrid bearings for 101H and 304H sizes Rolling Contact Fatigue Tests of Hybrid Bearings with Scarred Si 3N4 Balls Figure 16 shows the pronounced reduction of fatigue life and lowered life scatter. However, the  monotonic increase of failure probability seen with the pristine bearings is not present with the damaged bearings. The median life of the scarred bearings was about 5% of the pristine bearing life, and there was  no similarity observed in the distribution plots A                                          B  Figure 15. Turn-down model for as-manufactured 101H and 304H bearings. A) No turn down,  B) Notional turndown for one particular set of  parameters shown here . More data would be  needed to confirm or refute if a turn down effect is present in these materials. Turn-Down Model for 101/304 Hybrid Bearing RCF 10 Yrs 15 yrs 101&304 at 400 k s i 304 at 330 k s i Median 304 at 400 k s i 50% contour 1:100 1:10000 1:million 1 10 100 1000 10000 100000 1 0.111 0.1 0.01 RCF  Lifetime   -  hoursFraction of Median Normalized Strength ( 725 ksi) 141
damage effects in hybrid bearings is planned. To date, no infant mortality of the hybrid bearings has been  observed in tests with “short” (40-μm wide) scars on the Si 3N4balls.   Some test time savings may result from testing only slightly more than half the test specimens, thus  avoiding the longest lived members of the populations. The Reliability Forecast Chart presented in this  paper could be a useful tool in estimating load conditions of high projected reliability at long lives. Acknowledgements This work was funded by the Independent Research and Development program of The Aerospace  Corporation supported by USAF Space and Missile Systems Center (contract no. FA8802-09-C-0001). References [1] Park, W., et al., “Rolling Contact Fatigue and Load Capacity Tests of M62 Bearing Steel”, 32nd Aerospace Mechanisms Conference, NASA/CP-1998-207191, pp. 237-251, NASA Kennedy  Space Center, FL, 13-15 May 1998. [2] Popp, M. and Sternagel, R., “Hybrid Ceramic and All Ceramic Anti-Friction Bearings”,  Proceedings of the 8thEuropean Space Mechanisms and Tribology Symposium , ESA SP-438,  Sept., 1999. [3] Hutchings, I. M., Tribology: Friction and Wear of Engineering Materials , Arnold, London, 1992. [4] Gibson, H., Moore, C., and Thom, R., “Marshal Space Flight Center High Speed Turbopump  Bearing Test Rig”, Proceedings from the 34thAerospace Mechanisms Symposium , 10 May, 2000. [5]  Kang, J. and Hadfield, M., “Comparison of 4-ball and 5-ball Rolling Contact Fatigue Tests on Lubricated Si 3N4/Steel Contacts”, /g48/g68/g87/g72/g85/g76/g68/g79/g86/g3/g68/g81/g71/g3/g39/g72/g86/g76/g74/g81 , Vol. 24, No. 8, pp. 595-604, 2003. [6]  Warrier, S. Jarmon, D.C., and Chin, H.A., “Finite Element Analysis of the Critical Flaw Size in  Hybrid Silicon Nitride Ball Bearing”, and references therein, Proc. ASME Turboexpo 2000 , Paper  2000-GT-65, Munich Germany, 8-11 May 2000.  [7] Hamilton, G. M., “Explicit equations for the stresses beneath sliding spherical contact”, Proc.  Instn. Mech. Engrs , 197C, 53-59, 1983. [8] Johnson, L. G., /g55/g75/g72/g3/g48/g72/g71/g76/g68/g81/g3/g53/g68/g81/g78/g86/g3/g82/g73/g3/g54/g68/g80/g83/g79/g72/g3/g57/g68/g79/g88/g72/g86/g3/g76/g81/g3/g40/g81/g74/g76/g81/g72/g72/g85/g76/g81/g74/g3/g40/g91/g83/g72/g85/g76/g80/g72/g81/g87/g86 , McGraw-Hill,  New York, 1973.    [9] Robinson, E. Y., “A Universal Bivariate Weibull Model for Static and Dynamic Fatigue Reliability Forecasting,” 2011 ASTM International Symposium on Rolling Element Bearings, STP 1542, pp. 26-46, Anaheim, CA, April 13–15, 2011.   [10] O’Brien, M. J., de la Cruz, A. R., and Nguyen, E. A., “A Novel Proof Test for Silicon Nitride Balls”, /g45/g82/g88/g85/g81/g68/g79/g3/g82/g73/g3/g87/g75/g72/g3/g36/g80/g72/g85/g76/g70/g68/g81/g3/g38/g72/g85/g68/g80/g76/g70/g3/g54/g82/g70/g76/g72/g87/g92 , Vol. 94, Issue 2, pp. 597-604, February 2011. 143
Design Overview Requirements on the REXIS radiation cover flow from instrument measurement objectives as well as interface and environmental requirements. In order to meet REXIS science goals, the cover must provide a sufficient level of radiation protection for the CCDs. However, as a student instrument, the mass and  power available to the REXIS instrument is limited. The “not to exceed” values for REXIS are 6.5 kg   (14 lbm) mass and 15 W average power. Therefore a low mass and low power actuation device is  desired. During the cruise to the asteroid, REXIS will experience very cold temperatures. All components  of the mechanism must be kept warm enough to ensure successful cover opening. Lastly, as an asteroidsample return mission, OSIRIS-REx has strict contamination requirements. These constraints flow to  REXIS, and it is imperative that the opening of the door does not release particulates or molecular contaminates that could endanger the integrity of the asteroid sample. The REXIS radiation cover is located on top of the instrument and covers the coded-aperture mask so  there is no direct line of sight for radiation to impinge on the CCDs. The cover design is shown in Figure  1a in the closed configuration. The 4-mm-thick (0.2-in) aluminum radiation cover rotates about a custom  spring hinge with a 3/16-in (4.76-mm) diameter stainless steel shaft. The mechanism rotates about  Vespel SP-3 bushings. Additional bushings made of Rulon J are also placed in between the torsion  springs and the shaft to reduce friction and prevent cold welding of the springs to the shaft. The door is held in the stowed position by an FD04 Frangibolt actuator and custom notched titanium (Ti) bolt. The  FD04 was largely chosen due to its large stroke performance over equivalent products in the Standard Frangibolt family (FCx), along with its low power requirements, low mass, and low cost.  Figure 1.  (a) The REXIS radiation cover in cl osed configuration (b) detail of the Frangibolt  assembly The FD04 is part of the recently introduced Mini Frangibolt product family and has flight heritage on the  DICE CubeSat [4]. It consists of a Copper-Aluminum-Nickel shape memory alloy cylinder and provides ~550 lb (2450 N) of force to fracture the fastener in tension during actuation. This model is the smallest in  size and requires the least power of the TiNi Frangibolts. It has a compressed length of 0.500 in (12.7 mm), and a stroke of 0.040 in (1.0 mm). Although shorter in length than a typical Nitinol alloy actuator (such as the FC2, which at a length of 1.000 in (25.40 mm) is the next size up), the FD04 has a larger  stroke and is therefore less sensitive to preload [3]. The Frangibolt fastens two tabs together, one on the cover and one on the mask frame to hold the door closed. When power is applied to the Frangibolt, an  internal heater (with redundant capacity) heats up the memory shape alloy, placing the bolt under tension  and causing it to fracture. Once the Ti bolt is fractured, the door is released and opened by stainless steel  B A  146
Figure 3.  (a) Large T-Vac chamber used fo r ETU-2 (b) ETU-2 mounted  to cold baseplate Results The FD04 successfully fractured the fastener and opened the radiation cover in 19 of 22 tests attempted. The tests in which actuation did not occur were all early -20°C attempts before the thermal isolation issues were completely understood. In order not to risk damaging the FD04 Frangibolt, the power to the Frangibolt was shut off at 70 seconds. Once the thermal isolation issues were understood and power was applied for greater duration, the same FD04 Frangibolts actuated successfully. The actuation times for  the successful tests can be seen in Figure 4. This figure shows the test temperature and the time required to achieve actuation. In general, we see a trend of longer actuation times with colder temperatures. However, the results also show that at a given temperature, actuation time decreases as the power applied to the FD04 increases. Actuation at -40°C took more than 120 seconds with Ti washers and 9 V. The same test configuration at a similar temperature actuated in only 80 seconds when 10 V was applied.During the testing period, TiNi Aerospace provided the REXIS team with an alternate washer design made from G10 in order to reduce the actuation time at low temperatures. The results show that the use  of G10 washers consistently reduces the Frangibolt actuation time when compared to the use of Ti washers at similar temperatures and the same input voltage. Near -40°C and with an input power of 10 Vthe actuation time was reduced from 80 seconds to 40 seconds by changing from Ti to G10 washers.These results verify expectations based on the thermal conductivities between G10 and Ti and correlate with similar tests performed by TiNi. Finally, one test was performed at a very low preload (torqued to ~0.5 in-lb (0.06 N-m) as opposed to the specified 2.7 in-lb (0.3 N-m)) to confirm that the stroke of the  FD04 could fracture the fastener under these conditions. This test was performed with G10 washers at room temperature and with 9 V applied. The fastener broke in less than 20 seconds. As shown in Figure  4, this actuation time is in family with all other room temperature tests conducted with both G10 and Ti  washers at the specified installation preload. 148
Figure 4.  Actuation times for ETU-1 and ETU-2 as a function of temperature   Despite successful bolt fracture upon Frangibolt  actuation, there were a few post-actuation  observations worth noting. During the test protocol,  four distinct FD04 actuators were used. In the course  of testing, the first two actuators developed a tear along the seam of the insulation as shown in Figure 5.In addition, after the change from ETU-1 to ETU-2 a  chip was observed in the actuator jacket (also shown  in Figure 5). Upon actuation, the FD04 does move  significantly, and it is believed that this damage to the Frangibolt jacket was due to impact with the aluminum housing around the Frangibolt. The tearing and  damage to the jacket is cosmetic in nature and does not affect the performance of the FD04. Further tests with both actuators resulted in successful cover openings. However, given the strict contamination  requirements on OSIRIS-REx it is important to avoid this type of damage to the actuator on REXIS. To mitigate the issue, the Al housing was modified to remove sharp edges, allow more movement clearance  for the actuator, and include cushioning. Once these changes were made, no further damage to the  jacket occurred.   After two particular tests, the power to the Frangibolt actuator was not cut immediately after actuation and white smoke was observed emitting from the actuator in one of the tests. In these cases the actuators  were irreparably damaged. Once the fastener has broken, the FD04 is essentially floating free in the  housing and the thermal path to the rest of the structure is broken. In this state, the actuator heats up very  quickly and can easily exceed its high temperature limit. The FD04 as designed does not include a temperature sensor. Therefore, a limit switch of some type must be used to confirm actuation and ensure that power is cut. Discussion and Lessons Learned The testing of the REXIS radiation cover prototypes validated the functionality of the deployment  mechanism. The cover actuated successfully in 19 tests at a range of temperatures from -48°C to 25°C and both in air and in vacuum. It was shown that the springs open the cover once the fastener has fractured and that the cover remains open. It was noted that a large kick-off force is provided by the FD04  actuation, and systems that are sensitive to shock or jitter may require damping. The REXIS radiation  cover design was modified to include bumper tabs with Poron 4701-40 to reduce the impact of the door -50 -40 -30 -20 -10 0 10 20 30  0 20 40 60 80 100 120 140 160 180 Temperature (°C)  Time (s) Frangibolt Actuation Time vs Temperature  Ti Washer, 9V  Ti Washer, 10V  G10 Washer, 9V  G10 Washer, 10V Reduced  preload  Figure 5. Cosmetic damage to the FD04  jacket post-actuation. The numbers are the  actuator serial number. 149
References 1. Allen, B. et al. “The REgolith X-Ray Imaging Spectrometer (REXIS) for OSIRIS-REx: identifying  regional elemental enrichment on asteroids.” /g54/g51/g44/g40/g3/g38/g82/g81/g73/g72/g85/g72/g81/g70/g72/g3/g54/g72/g85/g76/g72/g86/g3/g21/g19/g20/g22 , 8840.  2.Strüder, L., et al. "The European photon imaging camera on XMM-Newton: the pn-CCD camera." /g36/g86/g87/g85/g82/g81/g82/g80/g92/g3/g68/g81/g71/g3/g36/g86/g87/g85/g82/g83/g75/g92/g86/g76/g70/g86 365 (2001): L18-L26. 3. Martynov, V., TiNi Aerospace, Inc. “Cu-Al-Ni Single Cystalline Shape Memory Alloy (SCSMA),  Properties and Prospective for Application” /g21/g23/g87/g75/g3/g36/g81/g81/g88/g68/g79/g36/g72/g85/g82/g48/g68/g87/g3/g38/g82/g81/g73/g72/g85/g72/g81/g70/g72/g3/g68/g81/g71/g3/g40/g91/g83/g82/g86/g76/g87/g76/g82/g81/g17 4. Fish, et al. “DICE Mission Design, Development, and Implementation: Success and Challenges”.  /g21/g25/g87/g75/g3/g36/g81/g81/g88/g68/g79/g3/g36/g44/g36/g36/g18/g56/g54/g56/g3/g38/g82/g81/g73/g72/g85/g72/g81/g70/g72/g3/g82/g81/g3/g54/g80/g68/g79/g79/g3/g54/g68/g87/g72/g79/g79/g76/g87/g72/g86 , SSC-12-XI-1.  152