ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
71 [] []11 1 ()sin 2 , ( ) 0,22 2() 1 1 () 1sin 2 , ( ) ,1 ,2 2 2(1 )tAA t qqut tAA tqqαπα απα +∈  =− +∈  −   and the other three phases are shifted 90° , 180°  and 270° . The actuator phase 𝛼𝛼(𝑡𝑡) denotes a specific  moment in the waveform cycle, and 0<𝑞𝑞<1 is the asymmetry factor  where 𝑞𝑞=0.5 is a pure sine wave.  After some open- loop testing,  𝑞𝑞=0.6 has been chosen, since this value returned  the smallest velocity  variations during a complete waveform  cycle.  

Figure 4.  Measured  open -loop mirror rotation and its estimated velocity using a constant   0.02- Hz waveform frequency.  

Two specific open- loop results using these waveforms  are shown in Figure 4, depicting the mirror rotation  measured by the interferometer and the derived velocity estimate.  Both experiments use a constant  frequency of  0.02 Hz, i.e. the actuator phase 𝛼𝛼(𝑡𝑡)(t) increa ses linearly,  so that a complete cycle takes  50 s. Indeed, the mirror  rotation is not perfectly linear, but the bottom plot shows  that there are no stick - slip effects or reversed directions,  which would be show -stoppers for accurate long- stroke closed- loop  motion. The velocity variations are at most a factor 3, which are relative easy to account for in a controller   design.  

Frequency response function measurements   Further characterization of the IFPM behavior has been done in the frequency domain, via local   frequency response  function (FRF) measurements around various actuator  operating points [8]. To this  end the waveforms  have been set at different nominal 𝛼𝛼nom ,𝑘𝑘, around which a small amount of additional  noise 𝛼𝛼�(𝑡𝑡) has been added,  which results in a mirror response 𝑦𝑦. The local dynamics  around operating  point 𝑘𝑘 can then be calculated as the ratio between the cross -power density 𝑆𝑆𝑦𝑦𝛼𝛼� and auto- power  density  𝑆𝑆𝛼𝛼�𝛼𝛼� of the tw o signals
72 ()() .()y kSjHjSjα ααωωω=   The resulting FRF s for 21 different nominal phases 𝛼𝛼 nom ,𝑘𝑘 along a full actuator cycle are shown in   Figure 5, using either the interferometer or the encoder as sensor  for 𝑦𝑦 (notice the gain difference  between the two, since one is expressed in nrad and the other in nm). The latter  obviously has a w orse  coherence and is noisier , since it has a much coarser resolution than the interferometer.  Apart from that,  both results are very much alike.  All FRFs show the same resonances and more or less the same anti - resonances, which implies that the dynamics i s nearly constant over a full actuator cycle. The FRFs only   differ in their gain, which is particularly clear for low frequencies.  These local DC- gain variations of at  most 9  dB can directly be linked to the observed velocity variations in Figure 4, since  this is ro ughly a  factor 3. Moreover, all FRFs have the same low -frequent phase, which confirms that there are no  direction reversals.  

Figure 5.  Measured FRFs as a function of the actuator phase 𝜶𝜶.   Note that 360 ° is equivalent to 𝜶𝜶 𝐧𝐧𝐧𝐧𝐧𝐧 =𝟏𝟏. 

Closed -Loop Performance  

The friction- based Piezo LEGS actuators are not meant  to provide accurate open- loop motion, which is  indeed confirmed by the above measurements. Hence, to meet  the extreme accuracy requirements, the  IFPM has to be operated in closed loop.  

Controller design  Since the DC -gain variations along the actuator cycle are relatively small, a fixed linear controll er can be  sufficient  to demonstrate the IFPM performance. The foreseen feedback loop is illustrated in Figure 6,  where the proposed controller structur e encompasses   • a double integrator to allow perfect tracking of linearly  increasing setpoints;   • a zero to create phase- lead around the cross -over frequency  (i.e. for closed- loop stability);   • a second- order low -pass filter to suppress higher -order  dynamics.
73  

Figure 6.  Feedback loop for controlling the mirror angle  𝒚𝒚(𝒕𝒕) via the actuator phase 𝜶𝜶(𝒕𝒕). 

The controller parameters have been tuned specifically on the measured FRFs in Figure 5(a), i.e., based  on the high- resolution  interferometer measurements. The parameters  have been chosen such that all  local FRFs yield a robustly  stable closed loop with a high as possible bandwidth.  In this case the  robustness has been defined by both a maximum  phase margin of 55°  and a modulus margin (maximum   value of the closed- loop sensitivity function) of at  most 3 .5 dB. The resulting open loops 𝐻𝐻𝑘𝑘(𝑗𝑗𝑗𝑗)𝐶𝐶(𝑗𝑗𝑗𝑗) are  depicted in Figure 7, both as a Bode and a Nyquist diagram. The latter shows that the closed loop is  indeed stable for all local FRFs, since all curves are on the right side of the point ( -1,0) [9]. Moreover, all  loops indeed satisfy the set modulus margin (black circle) and phase margin (black arc). Figure 7(a)  shows that t he achieved bandwidth lies between 26 and 62 Hz.     

Figure 7.  Frequency domain representations of the open  loop 𝑯𝑯𝒌𝒌(𝒋𝒋𝒋𝒋)𝑪𝑪(𝒋𝒋𝒋𝒋). 

Measured tracking performance   The designed feedback controller has been implemented in the Simulink / dSpace environment, and used  to validate the IFPM performance. As was originally the intention,  first the external interferometer been  selected as  feedback sensor .  Figure 8 shows the results in terms of the amplitude spectral  density (ASD) of the closed- loop error during  various  different experiments. The dashed line indicates the 5 nrad/√Hz pointing jitter requ irement. During  these experiments  both forward and backward movements has  been carried out  with three different  velocities ranging  from 12 .5 nrad/s (specified science mode velocity)  to 2250 nrad/s. Strictly speaking the  performance does  not have to be demonstrated for velocities larger than 12.5 nrad/s, but these are  included anyway to demonstrate the IFPM behavior over a larger stroke where the piezosteppers go  through a number of cycles. Clearly the requirement is met over all frequencies in all closed- loop  experiments.
74  

Figure 8.  Measured  amplitude spectral density (ASD) of the closed -loop error using the  interferometer as feedback sensor.  

However, interferometer measurements are known to be very sensitive to items such as temperature  variations and turbulence, which usually results in low -frequency drifts.  Hence, when the interferometer is  used as feedback sensor,  the mirror rotation is actually tracking these drifts.  As an alternative,  it has  therefore been decided to try to close the loop on the Micro- E encoder and use the interferometer  as a  reference sensor . Since Figure 5 showed that the dynamics are the same for both sensors, the same  controller as discussed above can be used for this purpose (only an additional gain of 1/ 0.03106 had to  be included to convert the encoder nm -readout to nrad).  

Figure 9.  Two closed -loop responses with 12:5 nrad/s mirror rotation (science mode) using the  internal encoder as feedback sensor. The setpoints are drawn on top of the measured rotations.    Two specific closed- loop time responses using this approach are dep icted in Figure 9. In these  experiments the  setpoint is changing with the science mode velocity, both forward and backward, which  the encoder is tracking successfully.  The encoder resolution steps are clearly visible in the response.   The effect of these steps on the ASD of the error (as  measured by the encoder) is however very limited,  as is  shown in Figure 10(a). Due  to the coarser resolution, the level of the ASD is much higher than in  Figure 8, but still the requirement is met quite easily, both for the science- mode and t he large- stroke  scenario. This implies that IFPM can be used with high accuracy in a stand- alone configuration,  i.e.
75 independent of any external metrology, which is  very beneficial in the current breadboarding phase of the  IFP project.  

Figure 10.  Amplitude  spectral densities of the closed -loop error using the internal encoder   as feedback sensor.  

Notice that the shape of the ASD in Figure 10(a) (especially  for the 12:5 nrad/s case) resembles the  designed closed- loop sensitivity  function    ()1 1( )( )SjHj Cjωωω=+  which is typically small for low frequencies and constant  for high frequencies. In the transition between  these two regions the bandwidth can be estimated, which is indeed in the order of 30 Hz.   Figure 10(b) shows the ASD of the interferometer readout  for the same experiments. This plot shows a  violation of  the requirement below 2 Hz. As pointe d out before, this  could partly be attributed to drift in the  reference sensor  itself, which is difficult to rule out in a normal unconditioned laboratory environment. As  such, we can conclude that the IFPM performs compliant, at least up to what  can be measured in a
77 Suomi -NPP Mission On-orbit Experience with Toroid Ball Bearing  Retainers under Unidirectional and Reversing Motion 

Otto Bruegman*, Kamal Thakore**, Stu Loewenthal+ and John Cymerman++ 

Abstract  

The Advanced Technology Microwave Sounder (ATMS) instrument scan system on the Suomi National  Polar-orbiting Partnership (S NPP) spacecraft has experienced several randomly occurring increased  torque ‘events’ since its on- orbit activation in November 2011. Based on a review of on- orbit telemet ry data  and data gathered from scan m echanism bearing life testing on the ground, the conclusion was drawn that  some degradation of Teflon toroid ball retainers was occurring in the instrument Scan Drive M echanism. A  life extension program was developed and executed on -orbit with very good results to date. The life  extension program consisted of reversing the mechanism for a limited number of consecutive scans every  day.  

Introduction  

The ATMS scan mechanism has been described in references  1 and 2, so only a cursory  description is  included . This paper is devoted to root identification of the ATMS scan bearing torque events , evaluation  of sensor impacts, and mechanism life evaluation by resolving the differences between the ATMS bearing  experience vs other more positi ve on -orbit experience. In addition, part of the paper discusses the  implementation of an  on-orbit life extension mitigation on the SNPP spacecraft.  

Figure 1 depicts the scan mechanism in cross section, which identifies the major components of the  assembl y. Note that there are two sets of bearings in this assembly . The problem is only observable with  the Main Motor bearings for ATMS . The Compensator Motor bearings are identical to the Main Motor  bearings, but run at a rate of 1.8X the Main Motor rotation rate and have not shown any misbehavior . At  this time, we don’t know why the Compensator Motor bearings are acting differently than the Main Motor bearings.    Figure 2 depicts the bearings in question, showing the arrangement of the balls and the retainers . The  bearings are an angular contact pair manufactured by Timken. The bearing pairs have an OD of 57.15 mm  (2.0 inches ), with each beari ng having a complement of 38 3. 175-mm (0.125- inch) balls and 19 toroids .  The nominal contact angle is 20 deg. Bearing pr eload is 44.5 newtons (10 pounds).    The bearings are lubricated with Nye Rheolube 2000B . The design life of the ATMS instrument is 7 years . 

                                                     * NOAA, Silver Spring, MD   ** NASA Goddard Space Flight Center, Greenbelt, MD   + Lockheed Martin Space Systems Company, Sunnyvale, CA   ++ Aerospace Corporation, El Segundo, CA  

Proceedings of the 43rd Aerospace Mechanisms Symposium, NASA Ames Research Center, May 4-6, 2016
78  

Figure 1 .  ATMS Scanner Cross -Section  

Figure 2 .  Nominal ATMS Bearing Configuration  

On-orbit Indications of a Problem  

The SNPP spacecraft was launched on October 28, 2011 . The ATMS instrument started operations on   Nov 8, 2011 and has since been in continual operations . The instrument experienced its first indications of  scanner related problems about a year later . The indications of problems were based on evaluation of main  motor current spikes and related motor control telemetry.  

Compensator Motor Bearings   Main Motor Bearings
80   

Figure 4 .  Test Unit #2 bearing condition from  2009 disassembly 

The other units (1 and 3) continued to run beyond their 7- year design lives , but test unit #3 exceeded its  torque limits on 6/29/2012 and was taken off line. At this point in time, it was decided to switch out the drive  motor of the unit to more closely match the torque capability of the flight motor . The original test set motors  had a torque capability of 69 in- oz (0.49 N -m) continuous duty, with a 120 in- oz (0.85 N -m) peak capability .  While the flight motor had a torque capability of 120 in- oz (0.85 N -m) continuous duty with a 300 in- oz   (2.1 N -m) peak capability . The replacement motor chosen for test station 3 closely matched the capability  of flight motor. This switch  of the motors was completed in March, 2014 and the unit was restarted.  

Recent Developments    In late 2014, the SNPP ATMS instrument started to experience some more elevated motor current spikes .  This lead to t he re- institution of the Sc anner Drive Working Group. The working group was tasked with the  evaluation of the on- orbit situation and to provide recommendations for mechanism life extension  mitigations.    This lead to a variety of work, starting with a re-evaluation of the work on bearing life prediction, applicability  of life test data on other programs, work  done in 2009 on the test unit 2 etc. 

A re-evaluation of related life test data by Stu Loewenthal of Lockheed Martin suggested that the differences  in life test results were the result of mechanism direction of motion. In life test data from other programs  that used toroid type bearings, it was noted that the particular mechanism motion profile included periodic  and numerous reversals .    A comparison of the Test Station #2 life test bearing with rotation in one direction and a similar life test bearing that rotated equally in both directions is shown in Figure 5. The bi -directional bearing is clearly in  much better condition than the ATMS bearing des pite achieving six times the number of revolutions.
81  

Figure 5 .  Comparison of Test Station#2 bearing with a similar bearing in relatively good condition  with equal direction reversal s after 6 X times the number of revolutions.   Left Figure: ATMS uni -directional test bearing at 63 Mrevs   Right Figure: Similar bi- directional bearing at 362 Mrevs  

Reversing bearing rotational direction minimizes toroid wear because it reduces the chance for the toroids  to get stuck/pinned in one orientation. The toroid  that is pinned by the driving ball will tend to rub in one  spot causing the lubricant to dry out and also inhibits the supply of fresh lubricant. The photos of the worn toroids from post -test inspection (see the next section) clearly show that the toroids were locked in one  orientation forming ball pocket wear divots. From the steel ring- on-block tests reported in Reference 3 , the  wear rate of unlubricated (dry), PTFE test samples was >1000 times the wear rate of those lubricated. This obviously has a major eff ect on toroid wear life.    The major benefit from reversing direction is that it allows the toroids to reorient due to clearances between  the balls and the release of load during reversal . The unloaded toroids have the opportunity to reorient  themselves not only presenting a fresh rewetted surface but also distributing wear more equally around the toroid perimeter .    The ATMS bearing application, by comparison, did not involve any reversals in its operation. The ATMS  motion profile is uni -directional for th e entire life of the instrument . There are accelerations and  decelerations in the ATMS profile, but no opportunities for reversal of motion. A description of the ATMS  profile is provided in Figure 6. Note that  the primary data collect occurs during the E arth scan at const ant  speed as does the cold and hot calibrations . In order to minimize revisit time to the required 8/3 of a second  per revolution, the scanner must accelerate and br ake at extremely high levels ( 4500 to 5000 deg/s2) three  times each revol ution . At first it was thought that these high acceleration levels were responsible for ball  skidding leading to premature life. However subsequent analysis showed that the dynamic loads were very small compared to the bearing preload making skidding unlik ely.
82  

Figure 6 .  ATMS Scan Profile Description  

Test Station Bearing  Post -Mortem  

To gain insight into the mechanism of the toroid wear seen in the original test station 2 bearings, the bearings were removed from storage for a more detailed inspection. Detailed measurements were taken of  the surviving toroids and some 3D modeling was performed to gain insight into the mechanics of the wear phenomena. Figure 7 and 8 present an example of one of the surviving toroids and modeling results  showing the conf iguration of the toroids during the wear process.  

Figure 7 .  Sample Bearing Toroid –  Test Station #2
83   Figure 8 .  Modeling of Toroid Wear  of Sample Pictured in Figure 7  – Test Station #2  

The wear patterns noticeable in the test station #2 toroids  had a variety of configurations, but the one shown  above is most telling. It shows that the bearing races had a significant contribution to the wearing of the  toroids . However, the major mode of wear that lead to the destruction of the toroids was the one where the  balls wore an ever deepening ‘dimple’ into the toroid that lead to the disintegration of the circular geometry .  In other words, the ‘dimple’ went so deep that the toroid split in two. Subsequently, the toroid halves were  ground up by the continu ously rotating bearing element s – thereby causing the current spikes that were  observed in telemetry . The grinding up of the toroid material lead to the creation of a slurry of Teflon particles  and lubricant outside of the bearings . There were no intact toroids in either of the tw o bearings from test  station #1 indicating the test station bearings were run well beyond their useful lives.  

Only a slight indication of residual lubrication on rolling elements  was observed . There were also i ndications  that lub ricant  was creeping out of the bearing assembly  forming the slurry observed external to the bearing.    Pre-load of the bearing pair was checked and found to be unchanged if not a little higher than from assembly  records . This may have been due to measurem ent uncertainty or to a very thin film of slurry on the metallic  surfaces that minutely changed bearing dimensions.    It should be pointed out th at it was anticipated that many  of the toroids in test stations 1 and 3 were  compromised at this point . However, the units were run to gain experience with running damaged bearings  and also to see what benefit might be gained from our mitigation strategy of periodically reversing the motion of the scanner.    For the purposes of showing the variety of the wear  patterns of the toroids from test station #2, Figure 9 is  offered. The first picture shows a concave toroid wear pattern suggesting that this toroid was rotating and  distributing wear around the toroid circumference. The remaining pictures show ball wear pockets indicative of toroids that got stuck in one position accelerating wear
84  

Figure 9 .  Toroid Wear Patterns (test station # 2) 

Putting the Puzzle Together  

So we had some information suggesting that the on- orbit motor current spikes were similar  to the bearing  life test station results . Also, we had some 3D modeling results that was able to explain the wear patterns  on a sample of the intact toroids . On-orbit motor current data was obtained at a high sampling rate    (148 samples/second) that could be used to perform some analyses –  including FFT, to get a better  understanding of the on- orbit situation. In addition, wear rate analysis was performed that suggested that  Teflon wear rate was highly dependent on lubrication of the interfacing surfaces –  ball to toroid in this case.  Figure 10 shows a sample of an FFT that was obtained from on- orbit data that suggested a response at  the Ball Spin Frequency . This was another clue in suspecting an interaction between the bearing ball(s)  and its cage  (toroid) .  

Figure 10 .  FFT Results showing Ball Spin Frequency  response  

The end result of all th e investigation and analysis lead to the conclusion that the ball toroid interface was  running dry . The operation of the mechanism in one direction was not allowing for the re- orientation and  hence re- lubrication of those surfaces of the toroids . Thus began a rapidly advancing wear of the softer  material (Teflon toroids) . As the ‘pockets’ in the toroids got deeper, the harder it was for any re- orientation  of the toroids to occur.  

Ball Spin Frequency
87 Life Extension Mitigation Implementation  

The on- orbit reversal process was initiated on July  14, 2015 . It consisted of a table load and its verification  and the n followed by the first reversal sequence . The first reversals were done by ground command, which  put restrictions on how fast people and command systems can operate. This resulted in approximately 26  reverse direction rotations of the scanner before the r eturn of normal forward scan operation. This reversal  resulted in a motor current spike that died away after about 3 days, which is faster than previous ‘naturally  occurring’ current spikes . A result like that was anticipated based on bearing test set reversals . The reversal  command process was implement ed via the spacecraft  Daily Activity Schedule command load on   August 24, 2015. This simplified the command implementation and had the effect of implementing the  reversals as part of normal spacecraft operation – basically making it seamless . It also had the result of  fixing the number of reversals at the desired number of six  contiguous reversals per day.  

Many thanks go out to the people in the satellite control center for their work in developing and  testing of  the scripts related to the implementation of the reversals.    The daily reversals have been perfor med without interruption since August 24, 2015 . Only one noticeable  motor current ‘event’ has occurred since then and that happened early in the pr ocess on August 26, 2015.  Since then the motor current has been rather stable with only small perturbations.    Figure 1 3 shows the transition from reverse and then back to normal . The figure shows the scan angle vs  scan telemetry sample number of the datase t. There are 148 data samples per complete scan rotation.  

Figure 13 .  Scanner Recovery from Reverse to Normal Scan  

Figure 14 is provided to capture the scanner main motor current for the past year . It should be noted that  the motor current level is now  at start of mission values . This would indicate that the reversals are effective  in re -lubricating the bearing/toroid surfaces . Also, the reversals may allow the toroids to re- orient  themselves exposing freshly lubricated Teflon surfaces to contact the adjacent balls.