ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
Conclusions The premature failure of the ISS external thermal control system provided an unprecedented opportunity to retrieve and examine spaceflight hardware that has operated for years in a microgravity environment.The ground-based inspections of the returned pump and ensuing failure investigation revealed five potential bearing forces that were un-accounted for in the design phase of the pump. While these forces  are relatively small compared to the deadweight load of the rotating group in standard gravity operation, when on orbit, these forces could combine in a number of random orientations to overload the pump bearings leading to solid-surface contact, wear, and premature failure. The recovery plan identified the  lubricant jet force on the rear bearing as being the damage-initiating event for the failure. Since this force  is directly related to the square or cube of the operating speed of the pump, there exists a crossover  speed at which bearing margin becomes negative and carbon wear begins to occur. Design life for the  remaining pumps was recovered through a change in flight rules for the operation of the pump modulethat limits rotational speed and minimizes speed changes and start stop cycles. Through the course of the  failure investigation, recovery, and follow-on assessment of pump wear life, design guidance has been  developed to improve the life of future mechanically pumped thermal control systems for both human and robotic exploration missions. The following six recommendations are proposed for future development of  high speed pumps designed for long-term microgravity operation: /g120Maintain an axially symmetric journal bearing design /g120Ensure that cylindrical rigid body modes dominate over conical modes /g120Perform vertical testing to unmask dead weight loading effects /g120Perform qualification tests that quantitatively validate bearing performance /g120Perform qualification tests that quantitatively validate rotordynamic performance /g120Maintain ground assets to experimentally probe and verify in-space anomalies 461
References 1. Active Thermal Control System (ATCS) Overview, The Boeing Company,  http://www.boeing.com/assets/pdf/defense-space/space/spacestation/systems/docs/ISS% 20Active%20Thermal%20Control%20System.pdf 2. Birur, G.C., Bhandari, P., Gram, M.B., Durkee, J., “Integrated Pump Assembly – An Active Cooling  System for Mars Pathfinder Thermal Control”, Society of Automotive Engineers, 26th International Conference on Environmental Sciences, Monterey, CA, 1996 3. Bloch, P.E., “Understanding Canned Motor Pumps”, Maintenance Technology, 01 September 2008. 4. Duchesne, S.M., Sweterlitsch, J. J., Son, C. H., Perry, J. L., “Impacts of an Ammonia Leak on the  Cabin Atmosphere  of the International Space Station”, American Institute of Aeronautics and Astronautics, International Conference on Environmental Systems (ICES); San Diego, CA; 15-19 Jul. 2012; United States, 2012. 5. Lockwood, M.K., Ercol, C.J., Cho, W.L., Hartman, D., Adamson, G., “An Active Cooling System for the  Solar Probe Power System”, American Institute of Aeronautics and Astronautics, International  Conference on Environmental Systems, AIAA-2010-6294, 2010.  6. Ticker, R., “Engineering Test Beds on the International Space Station”, IEEE A&E Systems Magazine,  August 2009. 7. Swanson, T.D., Birur, G. C., “NASA Thermal Control Technologies for Robotic Spacecraft”, National  Aeronautics and Space Administration, N20030031332, 2003. 8. Westheimer, D. T., Tuan, G. C., “Active Thermal Control System Considerations for the Next  Generation of Human Rated Space Vehicles”, American Institute of Aeronautics and Astronautics AeroSpace Sciences Meeting, AIAA-2005-342, 2005. 9. Paris, A.D., Bhandari, P., Birur, G.C., “High Temperature Mechanically Pumped Fluid Loop for Space  Applications – Working Fluid Selection”, SAE International, 04ICES-282, 2004. 10. Jimenez, A.P., Hoberecht, M.A., “ Space Station Freedom Photovoltaic Power Module  Design  Status”, National Aeronautics and Space Administration, 899271, IEEE , 1989. 11. Shen, F., Drolen, B., Prabhu, J., Harper, L., Eichinger, E., Hgyuen, C., “Long Life Mechanical Fluid  Pump for Space Applications”, American Institute of Aeronautics and Astronautics AeroSpace Sciences Meeting, AIAA-2005-273, 2005. 12. Son, C.H., “ Numerical Study of Ammonia Leak and Dispersion in the International Space Station”,  American Institute of Aeronautics and Astronautics, International Conference on Environmental Systems (ICES); San Diego, CA; 15-19 Jul. 2012; United States, 2012 13. Baldauf R. W., Kawecki, T., Purdy, W., Hoang, T.T., “Development of a Bearingsless Ammonia Pump  for Spacecraft Thermal Control”, NRL Review, 2008 14. Midgley, J.W., Teer, D. G., “An Investigation of the Mechanism of the Friction and Wear of Carbon”,  Transactions of the ASME, American Society of Mechanical Engineers, December, 1963. 15. Fritz, R.J., “The Effects of an Annular Fluid on the Vibrations of a Long Rotor, Part 1 – Theory”,  Journal of basic Engineering, December 1970. 16. Fritz, R.J., “The Effects of an Annular Fluid on the Vibrations of a Long Rotor, Part 2 – Test”, Journal  of basic Engineering, December 1970. 462
the Martian atmosphere at pressure 10 Torr or greater. Similar demands for high precision manufacture  and assembly of the molecular drag pump exist as for the turbo-molecular pump.    In collaboration with GSFC, Creare has been developing and demonstrating the technologies required to design and build miniature high vacuum pumps. In particular, we have designed and built two of the smallest available high vacuum pumps that have the following pumping characteristics:  a compression ratio for air that is greater than 10e8; a pumping speed of about 5 L/sec; 10-W power consumption for an exhaust pressure of 10 Torr; and operation over a very wide range of temperatures (from -40°C to 65°C).The smallest of these pumps (Figure 1a) has a mass of 130 g, a diameter of 3.3 cm (1.3 in), and an  overall length of 5.8 cm (2.3 in) (i.e., the size of a C-cell battery). The slightly larger pump (Figure 1b) has  a mass of 500 g, a diameter of 5.1 cm (2.0 in), and an overall length of 11.7 cm (4.6 in) (i.e., the size of a soda can). The larger version was space qualified for use on the Mars Science Laboratory (MSL) SampleAnalysis at Mars (SAM) instrument that was launched in 2011 and is currently being used operationally  on Mars. The smaller version was tested to TRL 6 in the laboratory for potential use on ExoMars, a  reformulated Mars mission, or other astrobiological missions. (a) (b) Figure 1. (a) Smallest Turbomolecular/Mol ecular Drag Pump. ( b) Space-Qualified  Turbomolecular/Molecular Drag Pump That is  Part of the SAM Instrument on MSL. Vacuum Pump Fabrication Challenges As described above, in order to greatly reduce the size and mass of a TMP/MDP, very high rotor speeds  are required since the pump tip speed must remain constant as the pump diameter is decreased. In order  to achieve the desired size and mass for space missions, the miniaturized pumps that we have developed  require motor operation at 100,000 rpm (for the 5.1-cm-diameter pump) and 200,000 rpm (for the 3.3-cm-diameter pump). Table 1 summarizes the features enabled by miniaturization and those required for space qualification and the corresponding consequence and fabrication challenge. While the need for tight dimensional control, the selection of proper materials, and accommodating the consequences of thermal and structural interactions require careful engineering design, the most difficult challenge that we faced was the impact of very high speed operation of the motors on lifetime of the pumps. In particular, since the miniaturized pumps need to operate at such high speeds, the combined motor and rotor structural resonances are below the operating speed of the pump, which results in super-critical  operation. This, combined with the need for relatively long-life under harsh operating conditions (i.e.,  many starts/stops) resulted in the key limitation that needed to be overcome during space qualification. In  fact, TMP/MDPs designed for use on Earth are turned on once and rarely, if ever, turned off and on again, in order to avoid the wear-and-tear associated with starting and stopping the pumps. In addition,terrestrial TMP/MDPs are operated at sub-critical speeds. 464
Miniaturized TMP/MDP Design The design of the miniaturized, space-qualified TMP/MDP is shown in Figure 2. Figure 2a provides an  exploded view and Figure 2b provides a cross sectional view of the pump. This figure shows that the pump rotor is cantilevered over the pump motor and that the turbomolecular blades are on the outside of the pump rotor near the gas inlet and that the smooth molecular drag pump rotor begins on the outside of the pump rotor and continues on the inside of the pump rotor, making use of the bottom end of the pump stator and the outside of the pump motor for the molecular drag pump “grooves”. Table 1.  Summary of Miniaturized Turbomolecu lar/Molecular Drag Pump Fabrication Challenges. Feature Consequence Fabrication challenge Lower pumping speed Good internal leakage control Tight dimensional control Smaller dimensions Small clearances in flow channels Tight dimensional control High speed Super critical operation Advanced balancing procedures High speed, lightweight Material structural limits Use high strength alloys Wide temperature range Compatible material selection Incorporate heater and heat rejection Wide temperature swing CTE mismatch Careful thermal/structural design (a) (b) Figure 2. (a) Exploded, Three-Dimensional View of Space-Qualified 100,000 rpm TMP/MDP (b) Cross Section of Space-Qualified 100,000 rpm TMP/MDP Resonant Frequency Analysis Optimization Analysis of the motor rotor/pump rotor combination (see Figure 3) show that the lowest natural frequency of the combined system is approximately 985 Hz (e.g. approximately 59,000 rpm) and that the mode  shape is such that the pump rotor acts like a bell on the “stiff”, tungsten-carbide shaft with the bearings  providing the support points. In order to make the natural frequency of this mode as high as possible, we  have designed the pump so that the bearings are as close as possible to each other given the other constraints of the motor design (namely the electro-magnetic requirements of the low-power, high-speed  motor). In order to enable the pump motor to pass through this resonance, we have mounted the bearings in an elastomer, which provides both greater damping and lower stiffness as compared to hard-mounted bearings. While these properties enable super-critical operation, they also presented operational  465
temperature range challenges and necessitated the introduction of a heater on the pump to ensure that  the elastomer was soft enough during startup to provide these important properties. Lifetime Optimization As described above, high-speed operation, super-critical operation, and the need for many start and stop  cycles contributed to the need to optimize the lifetime of the pump to meet the mission requirements. In addition, since science considerations limited the choice of lubricant to non-hydrocarbon greases, the bearing-lubricant combination selection became the focus of an effort to optimize the pump lifetime. We used a combination of screening tests at room temperature with hardware that consisted of the pump motor and a pump rotor simulator (i.e., “Life Test Simulators”) as well as complete pumps that were tested over the complete operational temperature range of the mission. Figure 3. Fundamental Structural Resonance M ode Shape From Finite Element Analysis A summary of the test configurations and the results of the life testing using the Life Test Simulators are shown in Table 2. Based on the results of similar previous tests and input from bearing and lubricant experts, we focused our testing on two retainer materials and various pre-load values. We focused our  testing on these variations because results of our previous testing indicated that cracking or wear of the retainers was the primary, life-limiting, failure mechanism. In addition, based on results of our testing, the bearing manufacturer revised the design of their standard Meldin retainer to provide additional a larger  cross section of material. Finally, we considered the effect of pre-load since the overall design required  that the bearings be pre-loaded and past experience showing a significant effect of pre-load on bearing  and lubricant life. The results show that the Vespel retainers, which are strong, but which do not absorb much lubricant, can be made to last twice the desired lifetime specification. However, the performance of this material and the lubricant is strongly related to the pre-load. The Meldin retainer material, being  porous, stores lubricant and can therefore achieve a longer life time – up to and beyond three times the  desired lifetime specification. 466
Table 2.  Summary of Life Test Simulator Configurations and Results Test Unit # Retainer Material Pre-Load (N )/(lb) Lifetime (h r) Fraction of Spec 3 Vespel 2.7 / 0.6 801 1.59 4 Vespel 1.8 / 0.4 1027 2.04 5 Vespel 4.0 / 0.9 575 1.09 6 Meldin (original) 2.7 / 0.6 1160 2.21 7 Meldin (revised) 1.8 / 0.4 1736 3.31 8 Meldin (revised) 2.7 / 0.6 1690 3.22 9 Meldin (revised) 2.7 / 0.6 1343 2.56 Vacuum Pump Measured Performance The Martian atmosphere, though much lower pressure than Earth (e.g., approximately 10 torr of mostly  CO 2), is still much too dense for mass spectrometers to work. The SAM instrument requires two vacuum  pumps that are able to exhaust directly to the Martian atmosphere and much achieve high vacuum levels. In order to meet the high compression ratio, we developed a miniaturized hybrid turbomolecular and  molecular drag pump with unique features. The pump is shown in Figure 4 after installation in the SAM  instrument. The pump is less than half the size of the smallest turbo drag vacuum pump available commercially and consumes only one-tenth of the power. As a consequence of the smaller size, this  turbo drag spins at 100,000 rpm. Figure 4. Wide Range Pump Installed in SAM Instrument During the course of the pump development, we measured the ultimate pressure and flow rate of the  pump design. The ultimate pressure test measures high vacuum pressure on the upstream side of the WRP while maintaining a given exhaust pressure on the downstream side of the WRP and the flow is  only that due to outgassing in the test chamber (assumed to be very small). The tests are conducted first  with the pump purged with CO 2and CO 2on the exhaust side and subsequently with a thorough Helium  purge and Helium on the exhaust side. The pressure measurements are conducted in an environmental chamber with the chamber temperature set at 22 °C, +50 °C and -40 °C respectively. The pump was started at room temperature, brought up to full speed, 100,000 rpm, and maintained at this  speed during temperature cycling. The pump was initially brought up to +50 °C to allow as much of the entrained gas in the high vacuum chamber to outgas as possible. We did observe a steady decline of the high vacuum pressure throughout the test as the chamber was cleaned of outgassing material. The  chamber had been rough pumped for several days prior to testing, but even so improved high vacuum  performance will probably be possible with a more thorough bake out and cleaning with a turbo pump. So, the results reported here can be seen as conservative and provide a lower limit on vacuum pump  467
compression ratio. At each of the three temperatures, we recorded the steady-state high vacuum  pressure, the pump foreline pressure, the DC voltage and current into the motor controller, and the pump housing temperature. The high vacuum pressure was collected with a Bayard-Albert (BA) tube. A correction factor was applied to the indicated pressure to account for the difference in gas performance of the BA tube. The foreline pressure is measured with a Convectron gauge and properly corrected for the difference in gas convection properties. The measured ultimate pressure and resulting compression ratio for CO 2as a function of foreline pressure are given in Figure 5. We see that the pump maintains a virtually constant high vacuum pressure, but that this pressure is highly dependent on the temperature. The pumping ability of the turbo molecular pump is expected to be reduced with increased mean molecular speed. However, this only accounts for a very minor amount (<10%) of change in no-flow pressure. Barring any increase in leakage across the pump, we expect the main reason for increased ultimate pressure at high temperature to be due to increased outgassing in the vacuum chamber, which is also supported by the fact that the pressure at similar operating conditions kept dropping during the test as the chamber was baked out. We see that the compression ratio at Mars operating conditions exceeds 5e6 at high temperature and is close to 10e8 at low temperature. Compression Ratio CO2 1.00E+051.00E+061.00E+071.00E+08 02468 1 0 1 2 Foreline Pressure [Torr]Compression Ratio50°C22°C- 40 °CVacuum Pressure CO2 1.00E-071.00E-061.00E-05 02468 1 0 1 2 Foreline pressure [Torr]Vacuum Pressure [Torr]22°C50°C - 40 °C Figure 5. Ultimate Pressure a nd Compression Ratio Pumping CO 2 The flow test measures the pumping speed of the WRP while maintaining an exhaust pressure on the  downstream side of the WRP similar to Martian conditions of 12 Torr. The tests were conducted in three configurations: (1) the pump purged with CO 2and CO 2on the exhaust side; (2) a thorough Helium purge  and CO 2on the exhaust side; and (3) N 2on both high vacuum and exhaust side of the pump. During the  tests CO 2, Helium, or N 2is bled into the high vacuum chamber through an orifice and the pressure drop  across the orifice is used to determine the flow rate and thereby the pumping speed. The measurements are conducted in an environmental chamber with the chamber temperature set at +20°C, +50°C and  -40°C respectively. The pump was started at room temperature, brought up to full speed, 100,000 rpm, and maintained at this  speed during temperature cycling. The chamber was initially brought up to +50 °C to allow as much of the entrained gas in the high vacuum chamber to outgas as possible. We did observe a steady decline of the  high vacuum pressure throughout the test as the chamber was cleaned of outgassing material. The chamber had been rough pumped for several days prior to testing, but even so improved high vacuum performance will probably be possible with a more thorough bake out and cleaning with a turbo pump. 468
The pumping speed measurements will be affected by the outgassing rates, in particular at low  temperatures where it will underestimate the flow rate. The pumping speed is derived by bleeding the gas through a short circular duct of known dimensions into the high vacuum chamber. Assuming that the gas  flow through the duct is much larger than outgassing rates, the duct flow rates will equal pumping speed.The duct flow rate can be determined from measurements of gas pressure on either side of the duct. At each of the three temperatures, we recorded the steady-state high vacuum pressure, the gas leak  pressure upstream from the metering duct, the pump foreline pressure, the DC voltage and current into  the motor controller, and the pump housing temperature. We also measured the temperature of the  chamber wall upstream of the orifice and used this temperature as a measure of the temperature of the gas flowing through the orifice. Since this chamber is large both in length and diameter relative to the gas mean free path, and since the gas flow through this chamber is very slow, the gas has ample time to reach thermal equilibrium with the chamber walls. If the real gas temperature is actually different, the duct conductance will be wrong but we estimate that this error will be less than a few percent. The high vacuum pressure was measured with a BA tube. A correction factor was applied to the indicated pressure to account for the difference in gas performance of the BA tube. The foreline pressure is measured with a Convectron gauge and properly corrected for the difference in gas convection properties. The bleed pressure was measured with a calibrated Pirani gauge, and corrected for the difference in gas conductivity of Helium, CO 2, and N 2. The main uncertainty in the flow measurement is due to the error in  the BA tube measurement. This will directly affect the pumping speed measurement. We estimate the total relative error on the pumping speed to be 20-25%. We measured the pumping speed by adjusting the setting of a leak valve on the upstream side of the  metering duct. We covered flow rates through the duct from zero to a value where the pressure upstream  of the metering duct (the leak pressure) gave a mean free path of the gas molecules ten times longer than the duct diameter. This ensures that the duct conductance can be treated as purely molecular flow far from transitioning to continuum flow. The measured pumping speed is shown in Figure 6 for the threegasses and three gas temperatures spanning the temperature extremes for the WRP operation. We see that the pumping speed at high vacuum pressure (i.e., small differential pressure across the orifice) reaches a constant level of approximately 6 L/sec. This level is consistent with model values for the turbo-molecular stage at zero differential pressure across the turbo molecular stages. When the high vacuum  pressure (i.e. inlet to the pump) is lowered the pumping speed decreases, since the probability of amolecule hitting the pump inlet goes down to the point where as many molecules enter the pump inlet as  exit it. 469
Figure 6.  WRP Pumping Speed When Pumping CO 2 With Exhaust of CO 2at 12 torr The pumping speed is clearly affected by temperature, the pump being less efficient at higher  temperature. The main difference between high and low temperatures is believed to be the level of  outgassing from the vacuum chamber wall. At higher temperatures, the outgassing will increase  substantially, thereby adding leak flow to the orifice flow that the WRP must pump. The pumping speed at high temperatures will therefore be artificially low. By proper bake out procedures we believe that the outgassing leak rate may be lowered substantially, which will result in increased measured pumping speed of the WRP. Conclusions In this paper, we described the main design challenges of designing and developing a space-qualified,  miniature turbomolecular/molecular drag pump and the ways in which they were solved. This included a discussion of the custom design of a miniaturized high speed motor to drive the turbo drag pump rotor,  analysis of rotor dynamics for super critical operation, and bearing/lubricant design and selection. Measured performance and lifetime data proved that the pump design achieved the mission specifications and current operation on Mars proves that the pump is meeting mission requirements. References 1. Mahaffy, P.R., et al. “The Sample Analysis at Mars Investigation and Instrument Suite.”  Space Science  Reviews, September 2012, Volume 170, Issue 1-4, pp 401-478.Pumping speed vs vacuum pressure CO 2 0.01.02.03.04.05.06.07.0 1.00E-07 1.00E-06 1.00E-05 1.00E-04 Vacuum pressure [Torr]Pumping speed [L/sec] 22°C 50°C - 40°C 470