ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
anhydrous ammonia would likely degrade conventional lubricants, cage materials, and steels. Ball bearings  that were made with ceramic zirconia races, zirconia balls, and a polytetrafluoroethylene (PTFE) cage were  used in this application. A key EET running gear design feature was the elimination of inner race radial  interference fits. A purely axial preload was used to maintain shaft bearing interface alignment and eliminate  potential slip through the use of an axial spring pack. This configuration enabled the ceramic to be placed  in compression during all phases of assembly and operation and eliminated the ceramic bearing tensile  stresses at the inner race failure mode.  Figure 3. EET Design  NASA/CP—2018-219887 502
Below the rotating disc was the test chamber, which consisted of a four-pin load block. Each pin ran on its  own wear track, which enabled four data points at unique speeds to be obtained with each test operation.  Contact load between the carbon pins and the rotating disc was maintained by a spring pack and a load  screw arrangement whose features were machined into the load block.  The entire pin-on-disc arrangement was enclosed in a pressure vessel to enable safe operation in a  thermodynamically controlled environment. A motor was mounted on the pressure vessel top, and torque  was transferred to the shaft through a magnetic coupling. Three types of materials were used for the  magnetic coupling dome to manage operational risk with eddy current heating of the pressure vessel.  Design calculations and operational experience provided the following constraints:  •Stainless steel dome: use to 300 rpm. •Titanium dome: use to 900 rpm. •Glass-filled polyetheretherketone (PEEK) dome: use to 3500 rpm. To maintain a constant temperature during test operations and to enable tribometer precooling or  preheating to simplify liquid filling operations, a thermal bath was constructed in which the sealed tribometer  could be submerged. This thermal chamber was designed to meet a –75°C to +125°C temperature range.  Operation  The procedure for operation of the EET was as follows:  Test Specimen Preparation  •Perform pretest photomicroscopy and profilometry of the chrome-coated disc and four carbon pins to document pretest configuration. •Install runner onto the shaft and verify axial runout is less than 12.6 µm. •Verify motor operation, magnetic coupling alignment, and set speed point in motor controller. •Install carbon pins in PTFE sabot and set spring preload. •Measure and record pretest spring preload for each pin. •Install load block into pressure. •Install bottom seal and flange. •Environmental Control •Prepare the constant temperature bath with appropriate heaters, chillers, and heat transfer fluids. •Immerse the tribometer in the temperature-controlled bath. •Monitor internal temperature of the EET through purge gas exit temperature. •At the desired temperature, fill the EET with a predetermined volume of test fluid. •Monitor the tribometer and fill vessel mass, and the EET pressure and temperature. •After pressure, temperature, and fluid levels have stabilized, initiate test. Test Operation  •During startup, monitor motor speed to ensure proper operation. •Continuously record motor speed and power draw at a minimum of 10 Hz. •Verify motor speed with external strobe every 10 minutes. •Monitor pressure, temperature, and speed for stability and trends. NASA/CP—2018-219887 504
•Terminate the test at the first onset of unexpected or unexplained variability in test conditions. •Terminate test at the desired sliding distance (time). •Collect post-test data. •Drain test liquid from test chamber and collect in specimen jar and any debris. •Purge test chamber with nitrogen and bring to room temperature while monitoring pressure temperature and gas exit conditions. •When safe conditions exist within the tribometer, remove the lower flange and seal. •Remove the load block. •Measure and record the post-test normal force on each carbon pin. •Perform photomicroscopy, profilometry, debris analysis, and additional diagnostics as the test conditions or results dictate. Wear Coefficient Data  Low-Speed Wear Data  In order to build an early database of wear coefficients while the EET was under fabrication and to gain  experience with the material couple, a series of low-speed wear tests were performed on commercial pin  on disc tribometers in both dry and flooded contact conditions. A photograph of typical test hardware is  shown in Figure 5 and results are captured in Table 1. Test results indicated a wear coefficient on the order  of 10–7 (mm3/N m) with strong sensitivity to contact area (i.e., pressure). The wear coefficient is defined as  cubic millimeters of worn material normalized by normal force (N) and sliding distance (m). The friction  coefficient of the wear couple was measured in these tests to be in the range of 0.14 to 0.22 in room  temperature air. Testing in flooded liquid conditions revealed a reduction of friction coefficient to the range  of 0.07 to 0.11 while the wear coefficient remained within the same order of magnitude.  Figure 5. Test Specimens for low speed wear testing  NASA/CP—2018-219887 505
18. Midgley, J. W. and Teer, D. G., “An Investigation of the Mechanism of the Friction and Wear of Carbon,” Transactions of the American Society of Mechanical Engineers (ASME), December 1963. 19. Sliney, H. E. and Jacobson, T. P., “Some Effects of Composition on the Friction and Wear of Graphite Fiber Reinforced Polyimide Liners in Plain Spherical Bearings,” NASA Technical paper 1229, 1997. 20. Demas, N. G. and Polycarpu, A. A., “Ultra High Pressure Tribometer for Testing CO2 Refrigerant at Chamber Pressures up to 2,000 psi to Simulate Compressor Conditions,” Tribology Transactions , Vol. 49, pp. 291–296, 2006. Table 1. Wear Coefficient Data from Low-Speed Tribometer  Fluid Load (N) Linear Speed Wear Coefficient (mm3/N-m) Sample Temperature Contact Pressure (MPa) 4%NH 3 Hydrate 10 100 mm/s 4.97865E-07 25.5 to 27°C 2.601 4%NH 3 Hydrate 10 200 mm/s 2.41698E-07 25.5 to 27°C 3.581 4%NH 3 Hydrate 10 200 mm/s 2.95354E-07 25.5 to 27°C 1.798 dry, air 10 300 mm/s 6.38000E-07 25.5 to 27°C 12.109 dry, air 10 300 mm/s 1.29384E-07 25.5 to 27°C 8.416 dry, air 10 900 mm/s 1.83617E-07 25.5 to 27°C 6.466 dry, air 10 900 mm/s 1.92718E-07 25.5 to 27°C 7.101 proplyene glycol 5 300 mm/s 0.23949E-07 25.5 to 27°C 8.043 dry, air 2 900 mm/s 1.57290E-07 25.5 to 27°C 1.378 dry, air 1 900 mm/s 1.33452E-07 25.5 to 27°C 0.842 4%NH 3 Hydrate 10 100 mm/s 3.23704E-07 25.5 to 27°C 2.011 4%NH3 Hydrate 10 600 mm/s 0.329972E-07 25.5 to 27°C 7.082 4%NH 3 Hydrate 10 400 mm/s 0.803432E-07 25.5 to 27°C 4.744 dry, air 10 600 mm/s 3.54035E-07 25.5 to 27°C 2.244 dry, air 10 400 mm/s 1.67000E-07 25.5 to 27°C 3.393 NASA/CP—2018-219887 509
Table 3. Wear Coefficient Data for 28 Percent Ammonia, 100 Percent Ammonia, and Dry Nitrogen  Environments  FluidLinear Speed  (m/s)Test Pressure  (psig)Test  Temperature  (°C)Contact  Pressure (MPa)Wear  Coefficient  (mm3/N-m) 28% NH 3 1.07 6.0-7.5 45 32 9.00E-09 28% NH 3 1.33 6.0-7.5 45 29 8.00E-09 28% NH 3 1.60 6.0-7.5 45 9.8 8.50E-09 28% NH 3 1.87 6.0-7.5 45 20 6.50E-09 100% NH 3 3.30 15 -43 31 6.20E-09 100% NH 3 4.16 15 -43 29 8.30E-09 100% NH 3 4.98 15 -43 20 5.40E-09 100% NH 3 5.80 15 -43 21 8.30E-09 100% NH 3 4.85 20 -43 22 6.90E-09 100% NH 3 4.85 20 -43 23 3.10E-09 100% NH 3 6.06 20 -43 34 2.33E-09 100% NH 3 6.06 20 -43 28 1.69E-09 100% NH 3 7.27 20 -43 34 1.99E-09 100% NH 3 7.27 20 -43 22 2.47E-09 100% NH 3 8.48 20 -43 35 1.48E-09 100% NH 3 8.48 20 -43 26 1.38E-09 N2, bone dry 4.85 20 -45 21 7.15E-05 N2, bone dry 6.06 20 -45 21 4.99E-05 N2, bone dry 7.28 20 -45 21 3.85E-05 N2, bone dry 8.50 20 -45 21 2.19E-05 NASA/CP—2018-219887 511
Accelerated Testing of Tribological Components - Uncertainties and Solutions  Simon Lewis*, Mike Buttery*, Oliver Poyntz-Wright*, Anthony Kent* and Achilleas Vortsellas*  Abstract  A fundamental requirement of any spacecraft mechanism development is to demonstrate the integrity of  the selected lubricant system by means of an appropriate life test. As mechanism lifetimes are often long,  and development times compressed some form of accelerated test procedure may be required for  programmatic reasons though the strict tribological validity of such tests both for fluid-lubricants and for  mechanisms employing solid lubricants or self-lubricating bearing is often a point of concern.   This paper discusses the current state of knowledge and limitations regarding accelerated testing, including  the influences on lifetime, torque and material wear of the accelerated conditions. The uncertainties of  presently available acceleration techniques and limitations of available data and methods are highlighted  together with some potential future solutions.  Introduction  The challenges of successful accelerated life testing of spacecraft mechanisms and lubricants have been  a recurring consideration for many years (for example [1], [2]). A simple review of the main classes of  spacecraft mechanisms, their typical operating speeds and typical un-factored lifetimes is shown in  Figure 1 and highlights the desirability of accelerated testing for missions where operational lifetimes  (measured in terms of cycles or revs) are “long” but development times may be rather short. However  accelerated tests may also be considered even for mechanisms with more modest life requirements for  example where confidence pre-tests are needed (prior to qualification), for mechanisms which have a dwell  (stationary period) or duty cycle with long stationary periods (e.g. refocus or calibration mechanisms), where  a fixed launch date exists (e.g. “missions of opportunity” for comets/planetary exploration), where  programmatic concerns dominate or where a compressed development cycle is being taken (e.g. low-cost  missions).   Figure 1. Typical unfactored lifetime requirements in revolutions for different mechanisms  *ESTL (European Space Tribology Laboratory), ESR Technology Ltd., Warrington, U.K. Proceedings of the 44th Aerospace Mechanisms Symposium, NASA Glenn Research Center, May 16-18, 2018 NASA/CP—2018-219887 513
upon assembly would be a high proportion (rapidly reaching 100%) suggesting that this phenomenon is  perhaps an overlooked lubricant loss route which could cause the onset of a bearing with very minimal  remaining free oil at the ball/raceway contacts (essentially starved/almost dry).   Figure 3. Oil mass uptake by cage and potential impact on proportion of free oil  (Basis: 28 µl free oil available, storage period up to 9 months)  Environment  Lubricant flow (including wake effects) and lubricating film thickness behaviors are clearly highly dependent  on environmental and local temperatures. Tribometer level tests in both sliding (e.g. Pin-on-Disc tribometer  (PoD)) and rolling/pivoting contacts (e.g. Spiral Orbit Tribometer (SOT)) demonstrate that the tribo-life of  common space lubricants is environmentally sensitive. SOT tests demonstrate that the assumption of parity  of environmental effects between air or nitrogen and vacuum would result in a factor 5-10 over-estimate of  the in vacuum lifetime achievable. Whilst the factor may be different in a specific real application, this  demonstrates clearly that the potential error from the erroneous assumption of parity can be large.   For example, the lifetimes of the PFPE grease, Braycote 601EF, as measured in the SOT (in terms of  normalized lifetime in orbits per rev) and in pure sliding pin-on-disc (PoD) testing (absolute life in revs) in  various environments [13] is shown in Figure 4. This data, is reinforced by more recent work for PFPE and  MAC oils [14] clearly shows that environment has a considerable impact on both measures of lifetime,  implying that life test should be executed in vacuum to avoid under-test.   Figure 4. Lifetime of Braycote 601EF in various environments as assessed on  SOT (left) and PoD (right) [13]  NASA/CP—2018-219887 516
become increasingly important later in the bearing lifetime. Since they have no lubricant pre-applied to balls  or races, self-lubricating bearings are entirely dependent on this kind of in-situ lubricant double-transfer  from the beginning of life   The most common material for MoS 2-lubricated or self-lubricating bearings is PGM-HT (a PTFE glass-fiber,  MoS 2 composite) though other transferring polymers and metals may be used also for specific applications  or with alternative primary lubricants.  Primary Solid Lubricant – MoS 2 Where the primary solid lubricant is sputtered MoS 2 there is a considerable body of work which documents  in general terms the expected performance at tribometer level (for example summarized in [17]) and while  life data at component level in absolute terms may be scarce and variable due to different definitions of  end-of-life or end-of-test, normalized lifetime plots show the trend to be expected.  In pure sliding, whereas there seems no relationship between the friction coefficient of any individual MoS 2 sample and the lifetime obtained, there is a strong relationship between sliding life and mean contact stress,  Pm, in which life is stated [17] to be proportional to P m-3.8, a value also shown in the best-fit line for coatings  on a range of substrate materials of different hardness [18] as shown in Figure 5. A similar relationship  exists for data from rolling contact experiments using a SOT in which the life of MoS 2 applied to a 52100  steel ball running against a 52100 disc is found proportional to P m-3.2.  A limited number of tests have also been carried out on spur gears [19], and whilst the initial test campaign  used low precision, relatively low hardness and non-hunting ratio gears, some tentative relationship of gear  lifetime versus mean contract stress from MoS 2-lubricated gears has begun to emerge.  Figure 5. Normalized in vacuum lifetimes of sputtered  MoS 2 in PoD test on various substrates [18] Figure 6. Normalized in vacuum lifetime of sputtered  MoS 2 lubricated spur gears [19]  Whilst the spread of results in Figure 5 is due to different material combinations, if any regression curve  approach is to be used to estimate life (for example Figure 6 for gear life), or as a basis of accelerated test  justification, then the spread of results needs to be minimized. In general. this spread could come from two  potential sources, the natural variability of the test (whether tribometer or gear test), or some supposed  variability in the sputtered MoS 2 material itself. However, recent work suggests the spread of lifetime  obtained from sputtered MoS 2 is broadly similar to that of PFPE or MAC lubricants (when compared in likefor-like test such as the SOT). The typical standard deviation on life of a batch of MoS 2 tested in the SOT  is ~15%, a figure which compares well with the variability found in the same test for PFPE and MAC  lubricants [20] (measured at 1.5 GPa mean stress and room temperature).  NASA/CP—2018-219887 518
Figure 7. Comparison standard deviation of invacuum lifetime of 3 lubricants tested in SOT Figure 8. Relative lifetime of sputtered lead as  compared to sputtered MoS 2 in vacuum (~1 µm)  as function of contact stress  Primary Solid Lubricant – Lead  Since its high ductility permits transfer and re-transfer to different surfaces, like that of other low shear  strength metals, the lifetime of lead in rolling contact tribometer tests (and in components) is very long. As  it seems to persist as a very thin layer (deposited at around 1-µm thickness but functional even when the  remaining film is only ~100 Å) [21]. In the SOT for example (with Pb applied to 52100 steel balls only running  against 440C plates) Pb achieves in vacuum lifetimes at least an order of magnitude greater than MoS 2 and appears relatively insensitive to contact stress, however its corresponding friction coefficient is  somewhat higher than that of MoS 2 in the same tests.   Primary Solid Lubricant - Summary  In general, primary solid lubricants, whether metallic or lamellar such as MoS 2 can be relatively well  characterized by tribometer tests in vacuum, nitrogen and air. For example, in tribometers at least, the  relationship between contact stress and life for MoS 2 seems well established, and the dependencies of life  on temperature and substrate hardness have also been studied [22, 17]. Nevertheless, the relationship  between life and sliding speed, though often stated to be low, is not well documented (although at very low  speeds similar to slip velocities in low speed bearings a known friction/speed dependency exists).   Secondary Lubricant – PGM-HT  Regarding the secondary lubricant there are typically two principle accelerated test concerns, firstly that the  material itself may display unexpected wear behaviors as a function of speed, stress or local temperature,  and secondly that the cage may become unstable at speed, again causing different wear behavior in the  accelerated test to the nominal speed application and thus invalidating an accelerated test.  ESTL carried out a program of PoD and SOT test work aimed at improving understanding of the behavior  of PGM-HT material. As originally supplied, the material was found to be dimensionally unstable at elevated  temperature, a problem subsequently overcome by introduction of a thermal conditioning pre-treatment  [23]. In characterizing the tribological behavior of the thermally conditioned material ESTL carried out a  number of PoD tests in vacuum [24] which showed that whilst in general higher friction tests correlated to  higher wear rate, a wide range of frictional behaviors and wear rates could be expected (Figure 9).  Interestingly also the legacy material RT Duroid 5813 (“Duroid”) seemed to have approximately a 4 times  higher wear rate than found for the PGM-HT.  NASA/CP—2018-219887 519
Figure 9. In vacuum PoD friction vs wear rate of PGM-HT and Duroid v Steel [24]   (Mean Hertzian stress 10 -16.2 MPa ,52100 steel disc, Ra < ~0.1 µm, sliding speed 0.01-0.1 m/s)  As a first experiment the SOT was used to characterize the initial formation of a transfer film by the double  transfer method described above [25]. In this work a single ball loaded between plates was initially  unlubricated but developed a transfer film of the PTFE/glass-fiber/MoS 2 due to its once per revolution  impact with the guide plate and in subsequent rolling this material was also transferred from the ball to the  flat plates (this was considered as analogous to a cage/land or cage/ball collisions within a ball bearing).  Tests were run for only 50,000 ball orbits as a trial, but nevertheless the results added somewhat to  understanding.   Figure 10. Schematic SOT with polymeric guide plate at start of test (left) and after material transfer from  guide via ball surfaces to upper and lower plates (right).  Figure 11. SOT Friction of PGM-HT (left) and Duroid (right)  (In vacuum, mean contact stress 600 MPa) [25]  NASA/CP—2018-219887 520
Tests were carried out in vacuum, moist air and dry nitrogen on three materials: RT Duroid 5813 (legacy  material), PGM-HT (current widely used cage material), and C29 [26] a potential future cage material  (containing PTFE, carbon nano-fibers and MoS 2). Standard test conditions were 600 MPa mean contact  stress (increased to 1000 MPa for high load tests), 50-rpm rotational speed (increased to 100 rpm for high  speed tests), 21°C (increased to 50°C for high temperature tests).   Figure 12. SOT guide plate wear (left) and XRF-measured residual Mo thickness (right) for PGM-HT,  Duroid and C29 materials  Results, presented in [27], showed similar characteristics for all materials with variable friction, apparent  rapid transitions from a low friction (µ~0.05-0.15) value to a higher value (µ~0.4 or higher). Subsequent  measurements of Mo thickness (made by X-ray fluorescence spectrometry (XRF)) showed a two-valued  population perhaps also suggesting rapid transitions from one state to another. Subsequent profilometry  showed the transfer film to be very thin and discontinuous (with thickness of order 1-3 µm in the “best”  cases and some small amount of metallic wear (~1 µm) in the “worst case”.   Given that transfer film formation was shown to be relatively rapid, but also that the film was observed to  be quasi-stable with evidence of a high rate of variability of thickness and frictional behavior with time, the  equilibrium wear rate of the material in sliding was also investigated. The main concerns were that  accelerated test might lead to quite different wear regime, and perhaps difference in transferred material  characteristics, especially when thermal effects are also considered. PoD tests were carried out in vacuum  under load and speed conditions considered appropriate for the simulation of cage/ball and cage/land  contacts for low-speed, nominal (~160 rpm) and accelerated bearing test conditions.   Figure 13. PoD Friction and Specific Wear Rate for PGM-HT v steel in vacuum as function of rotational  speed (4 N, ~17 MPa mean)  NASA/CP—2018-219887 521
As can be seen, for pin-on-disc tests there is quite a wide spread of friction and wear behavior and although  mean friction seems approximately constant with speed, there may be some small increase in equilibrium  wear rate with speed (but also a relatively high variability between tests).  When considering the impact of temperature (at temperatures typical of an accelerated test 21°C and 50°C),  again the changes in tribological behavior were modest.  Figure 14. Effect of temperature on in vacuum friction coefficient and specific wear rate of  PGM-HT v steel  When considering the impact of load, though there is clearly also some spread, results show that while the  mean friction is not greatly impacted by increased load, the wear rate does increase with increased load, a  result which substantiates the practice of accelerating test by increasing both load and speed when using  this material.  Figure 15. Effect of load on in vacuum friction and specific wear rate of  PGM-HT v steel  Whilst it is acknowledged that small changes in friction can impact cage stability, the above data suggests  that provided the bearing cage remains stable throughout test then acceleration of a test by increasing  speed, load or both could be appropriate. This is because in general the friction and wear behaviors for  PGM-HT have a quite wide spread and the effects of speed on wear, even over a quite wide speed range  are approximately “within the noise” for this material.   The detection of cage instability then becomes a key enabler for the accelerated test envisaged. A recent  test at ESTL [28] highlighted that a cage can be stable, quasi-stable or unstable at various times depending  on complex factors (friction coefficients, instantaneous geometry, mass distribution, speed etc). Whilst  validation of cage numerical models requires a detailed understanding of these phenomena, it may be  argued that the success of an accelerated test needs only to demonstrate the absence of instability in the  accelerated test if none is found at nominal speed operation.  NASA/CP—2018-219887 522