ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
214 each bearing (each of which also contains a phenolic retainer). Braycote 601EF was manually added onto  all circular and Flexspline teeth, with excess removed after assembly.  

Endurance Failure  

The First Life   The test began with completion of start -up torque measurements of the Drive Output subassembly, and an  average value of  11.1 mNm (1.57 in- oz) was observed, landing nicely within the acceptance criteria of  <14.1 mNm ( 2.0 in -oz), per the original size 10 HD source control drawing. The required sustained static  output load of 11.3 Nm ( 100 in- lb) was maintained for 60 seconds with no ratcheting behavior observed, in  both directions. The subsequent running portion of the test kicked off, and for the first 27 hours, the input  torque required to drive th e output torque load of  2.3 Nm (20 in- lb), alternating directions every 6.5 seconds  at a rate of 8.0  deg/s , was an average of 38.1 mNm ( 5.4 in- oz). The input torque dropped to 35.7 mNm  (5.05 in- oz) (-6.5%), shown in the left panel of Figure 3, corresponding with an increase in speed of 0.58  deg/s (+6.7%) across a 52 minute period. The speed returned to its approximate initial condition (there was  no speed controller in the configuration) starting around ~170 hours, a nd the input torque increased, to  around 41.7 mNm ( 5.9 in- oz). Then, in the final 16.5 second dither time  test event , at around 210 hours,  the output speed dropped to about 87% of its starting value, and the input torque increased to a maximum  of 45.2 mNm (6.4 in- oz), about 18% higher than the starting value.  

However, after about 30 hours, the speed and input torque seemed to level back out, ending with an  average input torque of ~ 38.1 mNm ( 5.4 in- oz) and an output rate of about 8 .0 deg/s. The first life w as  complete, reaching over 250 hours of as -run test time. Despite the  variation along the way, the consistent  starting and ending input torque and speed measurements seemed like indicators of a successful first life.  The startup torque was re -measured, and  an average value of 4.8 mNm ( 0.68 in- oz), still defined as  passing, was observed . The sustained 11.3 Nm ( 100 in- lb) no-ratchet test was successful again, and life  two was ahead.  

Figure 3. Comparison of input torque observed over the initial 45 hours of the first and second lives.  

The Death March   At the onset of life two, the average input torque was back down to 35.7 mNm (5.06 in- oz), which  corresponded nicely to the run- in value seen after the first life settled.  As shown in Figure 3  and, although  the speed was around 12% higher than it was at the close of the previous test, the lower measured torque  reduced concern associated with the higher speed, and the test proceeded. Both measured values climbed  slowly to 38.1 mNm ( 5.4 in- oz) and 9.0  deg/s over the course of the following 75 hours. Then, another  change appeared with about a 0.5 deg/s speed drop over a half hour period, accompanied by an input  torque increase to 41 mNm ( 5.8 in- oz).
215 This marked the beginning of the roller coaster ride to c ome. Speed increased, then decreased, then  increased again over the following 15 hours, reaching a local high of 9.77 deg/s. After another 10 hours,  the speed had dropped back down to 8.4 deg/s. The input torque followed the pattern of increasing with the  speed decreases, and vice versa, over this 56.4 hour run with a 16 second direction change timing. At the  profile end, input torque was an average of  43.4 mNm (6.15 in- oz). After this, the death march began.  

At the start of each event type to follow, as well as in between, the motor voltage was readjusted to obtain  the desired speed, and the input torque would swell in response and subside again as the speed would  drop back down, like a snake swallowing prey (or a system without speed control). The effect  of one such  adjustment can be seen in the middle of the 1.5 second dither time test event in the shaded center panel  of Figure 4. After 34 hours, the sy stem stop ped cycling , and the measured input torque seemed to change  exclusively in the uphill direction, until it reached an average of almost 72 mNm ( 10.2 in- oz), or double the  observed value at the start of life two. See Figure 4 and Figure  5. 

Figure 4. Output Speed Variation          Figure 5. Input Torque Variation   Observed During the Second Life         Observed During the Second Life  

In the final stretch of life two, a marathon run with a 6.5- second dither time delay was embarked upon. The  speed had dropped from the initially set value, but seemed to settle out at just over 7.0  deg/s, which lasted  almost 15 hours. Then, things declined. Over a three hour period, the speed dropped from 7.0  to 4.5 deg/s,  which  was the lowest yet observed speed, shown in  Figure 6,  and the input torque increased to the highest  yet seen value of 91.1 mNm ( 12.9 in- oz), shown in Figure 7. The passing of 5 more hours brought hope,  then another 5 hours brought despair. When it seemed like the system had settled itself at 4.5 deg/s and a  drag torque of just over 84.7 mNm ( 12 in- oz), the motor voltage was driven back up to obtain the desired  output speed.  

   Figure 6. Output Speed Degradation i n                  Figure 7.  Input Torque Increase in the  the Final 6.5 Second Dither Time Test Event               F inal 6.5 S econd Dther Time Test Event 

This time, when the speed was increased to the desired value, the system faught harder than before. Input  torque climbed quickly , exceed ing previous test maximums, as seen in the far right of Figures 6 & 7. In
216 addition to the behavior was a simultaneous temperature spike of 10 °C on the test article housing in under  a half hour, with no change to the ambient environment. When the data for both lives was concatenated,  as shown in Figure 8,  it was clear that this final 6.5 second dither time test event showed unacceptable  performance. The test was halted for further assessment.  

Figure 8. From Top to Bottom, Measured Motor Cur rent, Speed, Housing Temperature, Input Torque and  Output Torque from the Endurance Test, Throught the Final 6.5 Second Dither Time Test Event  

The Final Surrender   For the endurance test that had been completed, data had been logged at 1 Hz. It was decided to perform  a final test run with a higher sample frequency, holding the output speed to the 8.0 deg/s target value, and  running with a 16- second dither time delay under the full 2.3 Nm (20 in- lb) load. This one last batch of data  was desired before the unit was to be disassembled, never to be run in the same configuration again. This  final run followed the pattern of the input torque and speed rising and falling in terrifying opposition, growing  increasingly worse as time processed. The test reached a climax when the input torque values climbed to  a whopping 141 mNm (20 in- oz), where it sat for about 26 hours, with the speed continuing to descend  throughout.  

Then, the system turned a strange corner. The input torque dropped, about 40 minutes before which t he  titanium housing temperature climbed 5 degrees Celsius in about 20 minutes. The output speed  experienced a major noise reduction at the onset of this temperature increase, and proceeded to exceed  10 deg/s, the highest value observed yet. A few blips fol lowed, before the input torque necked down to  77.7 mNm (11 in- oz). See Figure 9. After seven hours, the rollercoaster ride ended, as the motor began to  stall. See Figure 10 for the full test dataset. The time had definitely come for disassembly. Beforehand, the  start-up torque and no- ratchet torque tests were conducted. The no- ratchet torque was successful again,  and the start -up torque test pushed the acceptable limits, with a peak measured value of 24.7 mNm   (3.5 in-oz).
217   Figure 9. From Top to Bottom, Speed, Input Torque, and Temperature up to Termination of the Final 16  Second Dither Time Test Event  

Figure 10.  From top to bottom, speed, input torque, and housing temperature for entire test, including  finally executed 16s  dither time added test event. Vertical lines indicate test event separation points.  Sawteeth in temperature data induced by local thermostat.  

Removal of a single piece part revealed severely degraded lubricant in the Drive Output subassembly. The  grease within the Wave Generator bearings looked like copper colored glitter, and felt like the glitter was  mixed with dried glue. The material surrounding the tooth interface of the Flexspline and the Circular Splines  was black in color and crumbly, with a texture similar to clay mixed with brown sugar. Refer to Figure 11,  panels 1 and 2 show the blackened sample within the gear mesh, and panels 3 and 4 show the glittery  Wave Generator bearing samples. Each sample was sent to the analytical chemistry lab for analysis, and  the HD assembly was  ultrasonically cleaned using Vertrel XF, then sent for Scanning Electron Microscope  (SEM) imaging and Energy Dispersive X -Ray (EDX) analysis .
218   Figure 11. Panel 1: Drive Output Subassembly After Removal of Dynamic Spline, Showing Output Tetth  of Flexspline, and Static Spline. Panel 2: Dark, Crumbly Material Gathered in the Static Spline Feature  During Disassembly. Panel 3: Wave Generator Subassembly with Flexspline Installed. Panel 4: Opened  Drive Output Subassembly  Before Distrubance of HD Components  

The HD showed wear on all teeth, biased with the worst damage on the dynamic spline tooth interface, and  on most Wave Generator bearing surfaces. Damage on the teeth and bearing races was aligned with the  line of actio n of the bearings. SEM revealed wear to the Flexspline teeth, shown in Panel 4 of Figure 12,  which corresponded with Dynamic Circular Spline tooth wear, shown in Panel 1 of Figure 12. The Flexspline  had been slipping on the wave generator bearing outer rac es, with polymerized grease product in between,  and as a result, damage was seen on the inner diameter of the Flexspline (shown in Panel 2 of Figure 12)  and the outer diameters of the bearing outer races (shown in Panel 5 of Figure 12). Flexspline material  was  removed at that interface, combined with the bearing grease product, and redeposited onto the Wave  Generator bearing balls (see Panel 3 of Figure 12) and ball tracks.  

Figure 12.  Panel 1: Dynamic Spline tooth wear. Panel 2: Damaged Flexspline, input side inner diameter -  damage pattern reflects ball track load path. Panel 3: Flexspline material adhesion onto bearing ball  (increased chromium & nickel seen as dark patches via EDX). Inset shows representative “normal” ball  surface at similar scale. Panel 4: Flexspline output side tooth wear. Panel 5: Damaged bearing outer  race-  pattern reflects ball track load path.
221 of the two bearings. A total of 0.11 g of the 50/50 mass ratio slurry was added to each Wave Generator  bearing.  

The HD and duplex bearings were then reinstalled into the Drive Output Subassembly, which was installed  into the Endurance Test fixture. The syst em was run with no applied load at speeds of 8.0 and 4.5 deg/s  and temperatures of -20°C, -8°C, and 30°C. Startup torque tests were also conducted over temperature.  The results of this testing are shown in the panels indicated as “Cold Test 1”, of Figure 13.  

Another test was run with updated ratios, to understand the effect of reducing the oil percentage while  increasing overall mass of applied slurry. The Nye Synthetic Oil 2001/Rheolube 2000 ratio this time was  39/61 by mass for the Wave Generator Bearin gs, and 36/64 for the duplex bearings. The resultant drag  and startup torque values obtained over temperature are shown in the panels indicated as “Cold Test 2”,  of Figure 13. See Table 2 for lubricant parameters related to both Cold Tests.  

Figure 13.  Test Data from the C old Tests Conducted to D efine Grease & Oil Fill Quantities.  

The second tested lubricant configuration was chosen due to the improved performance, combined with  the additional slurry mass of the assembly. The desire for a higher slurry mass was to minimize any future  risk of lubricant starvation. The drag torque values were acceptable, thanks to the high- torque output of the  re-wound Duplex motors, and the relatively high current available from the instrument power bus.  

Table 2. Lubricant Application Parameters for Cold Tests 1 & 2.   Cold  Test Item Nye Synthetic Oil 2001/Rheolube  2000 Ratio (% mass)  Slurry Mass per  Bearing (g)   1 Wave Generator Bearings  50/50  0.11  1 Duplex Output Bearings  70/30  0.077   2 Wave Generator Bearings  39/61  0.137   2 Duplex Output Bearings  36/64  0.115
227 Multi -Mission Deployable Boom: Spring Mechanism Design, Failure Investigation,  and Resolution  

Christine A. Gebara* and Paul D. Lytal * 

Abstract  

As more ambitious missions are being pursued to  better  understand Earth and our Solar System, aerospace  mechan isms and deployable structures are being approached in new fas hions. NASA’s Jet Propulsion  Laboratory is developing two low Earth orbit satellites  that will use radar instruments to better understand  temporal changes in the Earth’s surface . Both the Surface Water Ocean Topography (SWOT) and the  NASA- ISRO Synthetic Aperture Radar (NISAR) spacecraft utilize large  carbon fiber  deployable antennae  to conduct such science . These large antennae have been des igned with similar mechanisms. During  testing  of the spring mechanisms  that deploy the antennae , a hardware failure was found. The source of  the hardware failure was traced back to the custom torsion springs used within the mechanism. Because  of the mechan ism volume constraints , the springs were designed with high aspect ratio rectangular cross  sections to maximize the spring constant for  the me chanism. Ultimately, a failure investigation and testing  campaign  led to  spring mechanisms  that have  been successf ully integrated into both spacecraft . 

Introduction  

NASA’s Jet Propulsion Laboratory (JPL) is currently developing two earth orbiting satellites. The Surface  Water Ocean Topography (SWOT) mission will conduct the first global survey of Earth’s surface water. The  NASA- ISRO Synthetic Aperture Radar (NISAR) mission will study temporal changes to Earth’s land and  ice-sheets using advanced radar techniques. Both missions serve to better understand how the Earth is  changing over time using radar -based instruments. The SWOT mission is a collaboration with the French,  Canadian, and UK  Government  Space Agencies and will launch from Vandenberg Air Force Base on a  SpaceX Falcon 9 Rocket in 2021. NISAR is a  partnership between NASA and the Indian Space Research  Organization (ISRO) and will launch in 2022 from Satish Dhawan Space Center on an Indian  Geosynchronous Satellite Launch Vehicle.  

Figure 1. A) SWOT with hinge locations shown on one mast B) NISAR with hinge locations shown.  

Both SWOT and NISAR use deployable radar reflector mast designs  developed at JPL. These deployable  masts, while different in geometry, have similar components and sub- assemblies. Both masts are  constructed from bonded Invar and carbon fiber composite structures  and employ analogous flight 

* NASA Jet Propulsion  Laboratory, Pasadena, CA   Christine.A.Gebara@JPL.NASA.gov ; Paul.D.Lytal@JPL.NASA.gov  

Proceedings of the 45th Aerospace Mechanisms Symposium, NASA Johnson Space Center, 2020
228 deployable hinge mechanisms. The SWOT mission has two identical reflector masts, each with two  deployable hinges. NISAR has a single mast with four deployable hinges. These masts can be seen in  Figure 1. Operationally, the deployable masts are launched in a stowed state wi th a launch restraint system  composed of separation nut devices. When commanded, the launch restraints release a pre- tensioned  spring and damper mechanism which deploys each hinge. Hinge deployment progress is monitored on the  ground using a potentiometer as well as a limit switch  on each hinge. Upon completion of the deployment,  an actuator -driven latching mechanism preloads precision alignment features on either side of the hinge  together. Figure 2 displays a n overview of  the mechanisms.  

Figure 2. Hinge Deploy & Latching Mechanisms (NISAR configuration s hown)  

Mechanism Design and  Fabrication  

Mechanism Design   Each deployable hinge for  the SWOT and NISAR masts is outfitted with a spring, damper and potentiometer  mounted co-axially with each hinge line. The NISAR mast is  composed of 7-inch (18-cm) square composite  tubing. The SWOT mast is composed of 10-inch (25-cm) squar e composite tubing. Figure 3 displays the  spring mechanisms for each mission. The smaller, 7-inch (18-cm) mast cross- section of NISAR became  the driving factor in the design of the spring mechanism to maximize mechanical  commonality between  projects . Common mounting interfaces were design ed for both p rojects . Ultimately, this led to a cylindrical  volume allowance of 7 inches  (18 cm)  in length and 1.75 inches  (44 mm)  in diameter for the NISAR spring  mechanism.  Because of the differences between the SWOT and NISAR stowed hinge angles , as well as  differences in hinge angles at different locations on each mast , four different torsion spring configurations  were developed, each with the spring arms located at different angles relative to each other in the relaxed  position.  This can also be seen in Figure 3 when comparing both images.
229 A)     B)  

Figure 3. View of A) NISAR and B) SWOT spring mechanisms  

The spring  mechanism is  required to meet JPL design requirements  for mission critical  spring design. As  such, springs are required to have a minimum no test yield factor of safety ( FS) of 1.50 and an ultimate FS  of 1.65.  Further more, JPL design principles impose a minimum mechanism torque margin of 100% in worst  case environments at end of life. These dri ving requirements meant the torsion  springs needed to  produce  a minimum deployment -direction torque of 28 inch- pounds  (3.2 N -m) at hinge closure. A standard round  wire 17-7 precipitation -hardened stainless -steel  torsion spring would not produce adequate torque in the  volume available without  violating mission critical factors of safety.  Alternative materials such as Elgiloy and  MP35N were considered, but all of the vendors considered for fabrication of these springs had a significantly  higher volume of experience working with 17- 7 stainless  steel and developmental risk was deemed higher  with these alternative materials.  Therefore, a geometric solution was developed:  a rectangular cross section  spring to maximize the moment of inertia within the available volume . 

After developmental fabrication test runs, the spring wire height -to-width ratio selected for the spring cross  section was 3.88 :1. This  value was determined to be the highest ratio achievable with available CNC spring  winding manufacturi ng capabilities . Spring manufacturing still included many challenges given the  propensity of the spring wire to rotate about the axis of the wire during winding and inconsistencies in spring   back resulting in non- uniform torsion spring inner and outer diameters. The CNC spring winding  configuration is shown in Figure 4 . Guide support features were added to the flight spring mandrel design  to prevent twist about the axis of the spring wire at either end of each spring.  

Figure 4. Torsion Spring on CNC Coiling Machine 

The rectangular cross section caused early manufacturing issues  for the flight units . The springs did not  initially meet axial length requirements. Further the wire was prone to unexpected twisting during winding.  The initial inclination of the team was to attempt to relax  the overall spring length requirement, but that  would have had significant ripple impacts into the mature design of the hinge and mast structures.  To  address length requirement non- compliance, t he initially  baselined spring with 29 coils was modified to a  baseline design of 27 coils. With this change, the spring design violated  JPL design requirements  minimum  factor of safety requirements . Reducing the number of windings increased the stress in the spring.  In  consultation with JPL materials expert s, material coupon testing for the flight lot of material was conducted
230 to establish higher strength allowables for the flight lot of material to address the slight negative strength  margins. Ultimately, the final flight springs w ere successfully manufactured with a variation of less than  0.007 inch (178 µm) in diameter and 0.012 inch (305 µm) in length across twenty -eight units . The torsion  spring design that was developed met all requirements , as verified via tensile test witness  coupons of the  material, destructive winding testing, dye penetrant inspection,  and other verification techniques.  

Figure 5. Cross section of NISAR Spring Mechanism  

Once the spri ng mechanism design solution was reached, prototype units were built . A prototype test  program was successfully completed prior to flight hardware fabrication to reduce risk of issues in the flight  hinge and latching mechanism development. The prototype test program included both ambient and  thermal functional testing and thermal characterization testing on a flight -like hinge fixture. The prototype  program did not include vibrational testing or life testing due to programmatic constraints . The lack of these  prototype tests  prevented design issues described in the next section from being uncovered prior  to the  qualification t est program.  

Mechanism Integrati on and Hardware Failure  

Hardware Failure   Upon successful completion of the prototype test program, fabrication of qualification, flight spare, and flight  piece parts  ensued.  Seven  SWOT spring mechanisms were assembled with a qualification unit slated for  thermal testing to characterize torque output at the worst case cold, ambient, and hot qualification   temperatures. The qualification unit was of a SWOT design but  was deemed similar to the NISAR design.  Therefore, a single qualification  unit was used for both missions.  Thermal life testing was conducted after  the qualification unit had undergone vibrational testing.  Thermal test temperatures and vibrational test  levels  were set to encompass  the environments for  both missions.   

During  thermal testing, the spring was wound and unwound manually through its operational range of  motion using a rotary turn table. T orque output and rotary angle were tracked with a transducer and  encoder, respectively.  At the qualification  hot temperature, hardware failure was observed. From repeated   torsion springs c ycling (winding and unwinding), fragmented Teflon  Foreign Object Debris (FOD) was  generated.  This can be seen in Figure 6.  The source of the FOD was determined to be from two glass -filled  Teflon  sleeve bearings in contact with the inner diameter of  each spring inside the mechanism. The spring  mechanism continued to function and torque performance was not measurably altered by the  fragmentation. Upon further investigation it was determined that the sleeve bearing had begun to fail prior  to hot thermal testing.