ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
337 In order to determine the mean operating pressure in the PMCs , it was necessary to make two sets of  frequency scans with a change of s ieve temperature between the two. The measurements of PMC motor  current versus  frequency were  used to determine the resonant frequency of the PMC and using  calibration runs obtained before and during spacecraft thermal -vacuum tests, a mean operating pressure   could be determined.  From the analysis done in March 2015, the PMC  2 pressure was 3.505 +/ - 0.015  kPa. The rate of PMC2 pressure decrease is constant and is ~ 0.03 kPa/y ear. The resonance shift of  PMC  2 is estimated to be ~ 1 Hz, having  minimal effect on operation or efficiency . The ability to perform  such in - orbit diagnostic s is essential as the knowledge is needed for computations of optical filtering provided by  the cell . This slow change in pressure has been accommodated through normal calibrations, allowing for  updates to the optical filter model s of this  channel  with little impact to data retri evals. 

Figure 8.  Pressure Modulator C ells Mean Pressure Drop Over Lifetime for PMC 2  

Stirling Cycle Cryocooler Performance  On 07 May 2001 the cooler on side B of the instrument failed suddenly, without warning, as previously  described [3]. One of the coolers had an out -of-limit condition for approximately 10 minutes before the  instrument was put into Safe Mode, turning the coolers off . Observations [11] included a heating effect  (2 K) at the displacer cold -tip B just before being turned off . Current telemetry indicated little change  during the anomaly , which conflicted with possible explanations for the failure. It was learned from the  vendor that the c urrent telemetry for the displacer was not functional, which  clarified the situation. The  instrument was  then power cycled, followed by a series of diagnostic tests  which  were run between   18 May and 26 Jun 2001 , including the mapping of  compressor to displacer stroke shown in Figure 9 .  

Phenomena consistent with pressure wave driving were observed in tests over varying amplitudes,  explaining a different coupling between compressor B and its displacer. All tests showed that displacer B  was not being driven electrically, but was coupled  with the pressure wave from the compressor (phase  control lost). Since that time, scientific data have been delivered by partially redundant channels 5- 8 on  side A, with the coolers working in an unbalanced mode to avoid end- stop collisions in the displac er. The  cooler operating frequency was adjusted down slightly from 43.96 Hz to 43.0 Hz to reduce the cold- tip  heating effect. This mode had  cooler B operating at 82.5% of the amplitude of cooler A and produced   some residual vibration, with the cooler worki ng in the ‘heating regime’, warming the detector due to the  lack of phase control ( displacer B running free). The c ompressor B amplitude was  fixed at   5.0 mm p-p, while the amplitude of displacer B was held in the range of  3.65 ± 0.07 mm p-p.  

PMC 2 Pressure [kPa]
338  

Figure 9.  In-orbit Displacer Stroke vs Compressor Stroke for C ooler (B) After Anomaly  

It has not been determined whether the displacer drive problem was related to the electronics, the  harness or the displacer mechanism, just that an open circuit was experienced. From the behavior  reported previously [6] and subsequent compressor operation, there was clearly no failure of gas seals,  with no indication of pressure drop. The pistons were free to move so there was no obvious failure of a  displacer coil bond to the coil former (a problem experienced due to silicone contamination affecting the bond of the motor coil, which was corrected in the subsequent batch of compressors built for the  INTEGRAL mission  [6]). It is notable that a lead- in spring failure was not completely ruled out and that the  MOPITT cooler application is a rare case of a launch in t he passive ‘unshorted’ condition (not actively  controlled) . Shorting had been reviewed as an option, but was not readily supported by the Lockheed  Martin electronics . Displacers have inherently low damping compared to that of a compressor, so there  was an acceptance of residual risk, with mitigation provided by testi ng to acceptance levels in this  configuration. To preclude any residual risk of this nature in the future, it is recommended that coolers  always be clamped by adequate shorting or active control during launch.  

The long- term cooler displacer / compressor  displacement data is shown in Figure 10. This shows the  slow increase in ‘ stroke’  needed to maintain the temperature control  compensation for contamination  build- up. The rate of increase reduces  gradually with time and further decontamination cycles.   

         Figure 10 .  Cryocooler Displace r and Compressor Displacement Over Lifetime   a) b) Compressor 2(A) Stroke mm p-p  Displacer 2(A) Stroke mm p-p Displacer B Stroke Telemetry [mm peak -to-peak]   Compressor B Stroke Telemetry [mm peak -to-peak]
339 The detector temperatures as maintained by the cooler s are shown in Figure 11 . Normally, the detector  temperature on the non- cooling side B is stable around a warm temperature of 326 K and not varying  more than 0.2 K (occasionally ~0.5 K). The detector temperature on side  A (coolers working with nominal  amplitudes) is stable at ~84.0 K. It continues to demonstrate a lessoning of the rates of change of  temperature after each decontamination cycle, as the cold- enclosure continues to be freed of moisture.  However, the presence of contamination build -up (condensation)  on the cold- tips is still noticeable, even  after 16 years in orbit and many decontamination cycles.              Figure 11.  Detector Temperatures Over Lifetime  

The ‘PDECs’ (position control) mode had been preferred prior to the failure as this mode was less   sensitive to drifting than the ‘ADECS’ (accelerometer feedback) mode. Axial vibration levels were reduced  to below 0.12 N rms in PDECs prior to t he failure [9]. After the anomaly, the compressors were intentionally  run at imbalanced strokes to allow the active detectors to cool, while keeping the failed displacer from being driven into its end stops . As a result, the unit has operated in this manual mode with temperature  control on the active side for most of the mission. Variations of the axial induced vibration (Z) are driven  primarily by seasonal parameters, but overall  they are very stable. Increasing displacement (Figure 10)  has little impact on residual axial vibration, as seen in Figure 12. Considering the overall situation,  the  vibration cancellation of the system is considered to be exceptional for the improvised balancing situation   which needed to be adopt ed.  

     Figure 12.  Cooler Accelerometer Readings [N rms]  a) Axial – Z axis, b) Lateral  – X axis  (Lateral Y-axis is very similar to X axis)    2000 2002 2004 2006 2008 2010 2012 20140.080.10.120.140.160.180.20.220.240.26      Time        2000 2002 2004 2006 2008 2010 2012 20140.150.20.250.30.350.4 Time      Decontamination  features   a) b) c) d) Displacer 2(A) Cold -tip Temperature [K]   Displacer 1(B) Cold -tip Temperature [K]
342 The l aunch configuration of the Stirling displacer  without shorting (or active control),  is thought to have  incurred some residual risk despite having been tested  at acceptance vibration levels . Where feasible, i t  is recommended to review such configurations closely and to ensure clamping  by adequate shorting or  active control  to limit stroke variations  during launch.   

Decontamination heaters and the venting design have been vital to the ongoing reliable performance of  the instrument, even after 16 years in orbit . It is emphasized that even with tight contamination controls,  the system design must have adequate heater capacity to remove water vapor  and other contaminants  and adequate venting to ensure long life functionality . The venting design, material selection,  and ground  purging with dry gas were key to ensuring that the required frequency of decontamination cycles was not too high.    A capillary bonding method for sealing of  germanium infra- red windows into the titanium housings  showed the capability for  use in a long duration mission, while  the depressurization witnessed for two of  the units  may indicate that the related process controls could be improved.    Accelerometer anomalies  continue to be caused by interaction of trapped particles  from the sun with the  neutral atmosphere. Other cooler applications with similar accelerometer arrangements  should take note  of the accelerometer phenomena cited h erein for both the cooler control integrity and scientific value.  

Acknowledgements  

The authors are grateful for the past contributions of the lead electrical designer, the late Dennis Henry,   and the lead software engineer Ron Irvine, who helped ensure success of the MOPITT  mechanisms via  their dedication and expertise. Furthermore, many  experts were consulted during the development  and  test phase of the mechanisms , including Casey de Kramer and Ed Devine (representing NASA), Richard  Coco and Richard Walker  (Draper Labs), Steve Gill  and Rob Rowntree (ESTL), Guy Peskett (Oxford),  Brian Jones (BAe), Ray von Savoye ( Lockheed Martin) , the late Derek Binge (Tecstar) , George Bailak,  Gurpreet Mand and Boyd Tolton (University of Toronto) . The MOPITT  instrument  team is thankful for their  guidance and sharing of expertise.
343 References  

1. Gille, J., Drummond, J.R., Edwards,  D.P., Deeter, M.N., Masters, N.S., Emmons, L.K., Pfister,  G.,  Yudin,  V.A. “What 10 Years of MOPITT Data Have Taught Us About Tropospheric Chemistry“,  American Geophysical Union, 2009 Fall Meeting, abstract #U32A -03  2. Drummond, J. R., Zou, J., Nichitiu, F., Kar, J., Deschambaul t, R., Hackett , J. “A review of 9 -year  performance and operation of the MOPITT instrument ”, J. Adv. Space Res.  (2010) ,  doi:10.1016/j.asr.2009.11.019   3. Caldwell,  D., Hackett,  J., Gibson,  A.S.,  Drummond, J.R., Nichitiu,  F. “The design and flight  performance of the MOPITT instrument mechanisms”, Proceedings of 11th European Space Mechanism & Tribology Symposium, ( 2005) ,  Proceedings 591: 99- 106  4. Gibson,  A.S., Hackett, J., Bailak ,G. "Design of a Length Modulated Cell for Remote Sounding of   Greenhouse Gases", NASA 33rd AMS , CP-1999- 209259, (May 1999), pp309- 324  5. Taylor,  F.W.,  “Forty years of satellite meteorology at Oxford”, Quarterly Journal of the Royal  Meteorological Soc ., Vol137 Pt B, Issue 655, pp 277– 285 (January 2011)   6. Venters,  P. "The Pressure Modulation System in the Improved Stratospheric and Mesospheric   Sounder", D.Phil.Thes is, University of Oxford (1991)   7. Gibson, A. S., Reed, J., Bradshaw, T. W.; Linder, M. “ Heritage Overview: 20 Years of   Commercial Production of Cryocoolers for Space” , Transactions of the Cryogenic Engineering  Conference - CEC, Vol. 52. AIP Conference Proceedings, (2008)  Volume 985, pp. 493- 505  8. Nichitiu,  F., Drummond,  J.R., Zou.J., Deschambault,  R. “Solar particle events seen by the MOPITT  instrument”,  Journal of Atmospheric and Solar -Terrestrial Physics 66 (2004) pp 1797 –1803   9. Mand,  G.S.,  Drummond,  J.R., Henry , D., Hackett,  J. “MOPITT On -Orbit Stirling Cycle Cooler  Performance”, Cryocoolers 11, Klu wer A cademic/Plenum, New York (2001)   10. NASA- Terra  Status  as per webpage report , http://terra.nasa.gov/wp - content/uploads/2014/03/ESMO_TerraWebStatus_June2014.pdf   11. MOPITT On -Orbit Cooler Anomaly Int erim Report, COM DEV RPT/MOP /50008/001- P0, 23 -Jul-2001
346 Noise characterization is an important test for any slip ring, as it can be the best indicator of the health of  the contact system. For circuits carrying power, those that are heavily averaged, and other robust signals  are generally insensitive to noise on the slip ring. However , some signals can be very sensitive. Stick -slip  behavior is evident in torque traces which results in noise peaks associated with these events (as shown  in Figure 2 for a typical slip ring assembly). The noise spikes in this figure correspond to the leading brush  sliding forward on the ring while the trailing brush remains in contact; the leading brush is momentarily open  for this brief period of time. Low current an alog signal circuits, for example, pose a particular challenge.  Other researchers have proposed slip ring noise has contributed to anomalous encoder readings and resulted in control problems on orbit [1] (Koss and Woolaway). Though the Windsat SRA discussed by Koss and Woolaway was of a unique configuration in terms of the materials used, rotational speed, and  total number of revolutions (life), their findings warrant further investigation into the integrity of low -current  signals on slip rings.  

Figure 1.  Cross -section of brush ring interface.  

Figure 2.  Typical signal (2 -brush) gold -on-gold slip ring resistance and drag torque as a function  of time. The behavior seen here is common at low rotation rates where stick -slip behavior is  prevalent. At higher speeds, those commonly used for noise testing, stick -slip is virtually non - existent. Brush design, lubricant selection and drive configuration  can drastically affect this  behavior. One of the products of this study will be noise level comparisons over a range of  rotation rates.      Dielectric  Potting Material   Ring:  Brass/Ni/Au  Brush
347 Experiment Design  

To characterize performance of a gold- on-gold slip ring we designed and fabricated a two- channel s lip ring  assembly using flight -like materials and processes. This SRA makes use of brass rings, cut with a 90degree v -groove (Figure 1) then  nickel and hard- gold plated. The rings are integrated onto an insulating  shaft and mounted to an aluminum housing on a pair of preloaded radial contact bearings. The brushes are  Neyoro 28A, a Au-Ag-Ni alloy commonly used in space flight slip ring applications [5] (Pitney). The brushes  are monofilament (homogeneous). They are mounted on a PWB brush block which is then assembled to the housing. The brushes are adjusted for appropriate cont act force with the rings and aligned with the  center of the v -groove.    The 2- channel SRA was lubricated with NYE 2001A and run- in in both directions under laboratory  conditions. After cleaning and inspection the SRA was re- lubricated with NYE 2001A and tes ted in air and  under vacuum conditions with the signals and speeds outlined in Table 3. Cleaning was performed using a  hexane rinse with the brush block and brush in- place so as not to disturb the contact system alignment.  After cleaning the unit was re- lubricated with NYE 2001 in the same quantities as before and the tests rerun. In a similar manner the rings are cleaned and re- lubed for NYE 2001- 3PbN p and tested.    The results presented herein will summarize the effect of rotational rate and lubricant selection on slip ring resistance and noise as a function of rotation rate.  

Experiment Configuration  

Testing was conducted on a two- channel slip ring. The slip ring itself was procured per a source control  drawing from Electro- Miniatures Corporation of Moonachi e, NJ . The slip ring was machined from brass tube  with 90- degree v -grooves which were nickel then hard gold plated per the vendor’s standard processes.  Brushes were procured from the same source and were produced by Deringer Ney of Vernon Hills, IL.  

Table 1.  Ring and brush details.  

 Testing was done in a thermal vacuum chamber  under ambient temperature conditions (20- 24°C) at 10 -4  to 10-5 torr pressure. The Unit Under Test (UUT) was driven from outside the thermal vacuum chamber  through a ferrofluidic feedthrough by a 1.8- degree commercial hybrid stepper motor using 256:1  microstepping and a 100:1 planetary gearhead. This yields a step size at the slip ring of 1.23 µrad. The test  slip ring in the chamber is shown in Figure 3.   Material/Dimension Specification/Units Ring Substrate Red Brass ASTM B36 Ring Diffusion Barrier Electrolytic Nickle Technic WS Ring Wear Surface Hard Gold MIL-G-45204 Type I, II Ring Pitch Diameter 70.61±0.051 (2.780±0.002) mm (inches) Ring V-Groove Angle 90 Degrees Brush Material Neyoro 28A ASTM B477 Brush Composition Au/Ag/Ni 75%/22%/3% by weight Brush Format Circular Cross-Section, Monofilament Brush Diameter 0.406 (0.016) mm (inches) Brush Force 6.0 gram-force nominal
348   Figure 3.  UUT in chamber before door closure. Slip ring assembly is shielded using aluminum foil  and aluminum tape. Drive shaft exits to the right o f the photo to chamber feed through . 

Figure 4.  Slip ring brushes in v -grooves. Heavier than typical lubricant application is present in  this photo taken during run -in. 

Procedure  

The slip ring was assembled and the brush blocks aligned with corresponding rings. Figure 4 shows the  brushes in the slip rings used in this test.  

The rings were lubricated with neat Pennzane® (NYE 2001A) and run- in under standard laboratory  conditions (atmospheric pressure and ro om temperature) for 2000 revolutions minimum in each direction.  After run- in the system was tested under lab conditions to verify performance and ensure test protocols  were appropriate. Lubricant quantity was judged visually by evaluating the meniscus at t he brush/ring  interface. Material and configuration information for the slip ring is given in Table 1.
349   Power was applied to the slip rings using a bat tery and voltage splitter arrangement. Figure 5 shows the  electrical schematic for this setup. A series of resistors was used to set the current applied to the slip ring pair. This battery arrangement was used because c ommercial power supp lies proved to be too noisy for  these tests as the noise levels resulted in unacceptable signal -to-noise ratio.    Using the battery pack approach we were able to effectively eliminate power supply noise and deal primarily with environmental and A/D noise in the DAQ . Our background noise level was on the order of   80-100 µV pk-pk. This background noise level dictat es the peak -to-peak noise that can be resolved in the test  and therefor e our effective signal -to-noise ratio.    Typical noise levels in this type of slip ring are on the order of tens of milliohms per circuit, meaning that for  a circuit pair in series that value is multiplied by  √2. Therefore, we expected approximately 3 0 milliohms of  noise. Table 2 shows the voltage level of the expected noise as a function of applied current via resistor  value in the power supply  circuit ( Figure 5). 

Table 2.  Power supply and slip ring test electrical parameters, including environmental  (background) noise levels and calculated signal to noise (S/N) ratios and resolvable noise levels.  

Figure 5.  Power supply circuit used to control current to the slip rings under test.  

Current was measured and recorded continuously by measuring the voltage drop across a high- power  resistor for each measurement . In this way , the current across the slip rings was directly measured,  variations of which was accounted for in our resistance measurements. Two sense resistors were used depending on the current applied and tailored to the test case. A 10- Ω resistor was used for high current  cases (120 mA and 12 mA) and a 100- Ω resistor was used for all other cases. In this paper , we will refer to  these current levels by their nominal values (e.g. , 100 mA instead of 110 mA) for convenience and Resistor  ValueBattery  VoltageSense  ResistorNominal  UUT  ResistanceActual  CurrentNominal  CurrentBackground  Noise LevelSlip Ring  Pair  ResistanceNoise  Spec  LimitSR  Voltage  DropS/N  RatioUnfiltered  Resolvable  Noise 100Ω 12.3V 10Ω 0.35Ω 110mA 100mA 100µV 131mΩ 30mΩ 14.6mV 146 900µΩ 1kΩ 12.3V 10Ω 0.35Ω 12mA 10mA 100µV 131mΩ 30mΩ 1.6mV 15.9 8mΩ 10kΩ 12.3V 100Ω 0.35Ω 1.2mA 1mA 100µV 131mΩ 30mΩ 160µV 1.6 82mΩ 100kΩ 12.3V 100Ω 0.35Ω 120µA100µA 100µV 131mΩ 30mΩ 16µV 0.16 0.88Ω 1MΩ 12.3V 100Ω 0.35Ω 12µA 10µA 100µV 131mΩ 30mΩ 1.6µV 0.016 8Ω
350 simplicity. Our DAQ was set up with two A/D channels, one for the voltage drop across the slip ring and  one for the sense resistor. Resistance was calculated in real time for each test as opposed to a nominal  current used and voltage being recorded. A sample resistance data set is gi ven in Figure 11. Data were  acquired at 100 kHz on both channels, a 100- Hz second- order low -pass filter  was used on the current  channel while the nois e channel was recorded at 100 kHz broken into three levels, each defined by a  second order low pass digital filter. One was set at 10 kHz, another at 1 kHz and the last at 100 Hz. Filtering  the raw noise data was also beneficial in reducing the resolvable noise and ef fectively increasing the signal - to-noise ratio, allowing us to decrease the resolvable noise levels from approximately 8  Ω for the unfiltered  signal to approximately 1.25 Ω using the 1- kHz filter.  

Figure 6.  Schematic representation of test setup.  Switch at resistor is normally closed, it is only  open for transient detection tests described herein.
351   Figure 7.  Annotated sample of DAQ test interface showing ongoing test.   Because the background noise levels in our production environment and test setup could not be reduced  below approximately 100 µV, we were unable to measure true slip ring contact noise directly at current  levels below around 1 mA. Although our threshold noise levels effectively  raise our measurement noise  floor and reduce our signal to noise ratio,  we can still resolve noise events of a certain magnitude. Filtering  the raw noise data helps to improve our signal -to-noise ratio and allows us to easi ly see a 1- ohm change  in the sli p ring resistance as witnessed by our test for sensitivity on the UUT in- situ at ambient pressure. To  do this, we connected a 1- ohm resistor as shown in Figure 8 and ran our data acquisition program to collect  resistance and noise data ( Figure 9) while our lowest (~10  µA) current level was applied.  The resistor was  attached as shown in Figure 8; connections were sprung against one another so that short duration  transitions in resistance could be induced. The oscilloscope trace in Figure 8 captures one of these transitions at 32 µs. The screen shot in Figure 9 shows the impact of this test on the captured data. Despite  the short duration of the resistance transition induced by the test , all filtered groups recorded the event;   also of note is that the resistance trace also recorded the corresponding increase in measured circuit resistance. This gives confidence that transient open circuit events are captured in the test data for very low current level s.  

Results  

Lubricant  and Rotation Rate Trials   Our results indicate that in our ambient temperature, high vacuum test environment the NYE 2001A and 2001 lubricant formulations behaved similarly at all tested rotation rates, with the formulated version (2 001)  providing slightly better performance across most rates and currents (see Figure 10 for a subset  of these  results). Neither of these lubricants exhibited measureable performance differences between operation in air at standard pressure and hard vacuum.