ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
by driving the percussor cam through the second stage of the spur gear with a torque watch and then  dividing the maximum reading by the first stage spur gear ratio.  Table 5. Venus drill motor torque and speed under various test conditions  * estimates based on measurement with a torque watch Whereas it has been shown that surface sliding friction at the bit to rock interface increases with temperature  under both Earth and Mars atmospheric pressures [10], the results of the VMTF drilling trials with the Venus  drill (see Table 6) suggest that temperature would not significantly impact specific drilling energy at Venus  conditions, at least for rotary-percussive machines. On the other hand, the rate of penetration at high  temperature is expected to be slower because the stiffness of the percussive spring (and, therefore, the  percussive energy per blow) is reduced. Furthermore, increased electrical (i2R) losses are expected at high  temperature because the resistance of the motor coils is increased. However, it should be noted that  specific energy, as defined herein, reflects mechanical performance only and does not include electrical  losses. Accordingly, SE is calculated using Total Power Output in Table 6, which is the combined  mechanical power output from both motors, rather than Estimated Power Consumption, which applies  efficiencies read from dynamometer test results in Figure 7.  Table 6. Specific energy for drilling measured at room temperature and high temperature  Motor Torque Margin and Thermal Considerations  Per spaceflight mechanism design requirements in AIAA S-114-2005 Moving Mechanical Assemblies, the  flight drill would need actuators capable of providing at least twice the nominal torque (both peak and  continuous) needed to operate the drill in worst-case environmental conditions. This margin allows for  modelling uncertainty, such as the detrimental effects of operating in the Venus atmosphere, and  unexpected failures that could cause additional drag on the mechanism, such as a seized bearing. Drilling Test RT = Room Temp VTP = Venus Temp & PressMotorMotor Peak  Torque* (Nm)Motor Current,  Ave. (A)Kt (Nm/A)Motor Cont.  Torque, Ave. (Nm)Motor Velocity,  Ave. (RPM)Power Output,  Ave. (W) Auger 1.561 0.115 0.180 2945 55 Percussor 0.57 2.816 0.115 0.324 3217 109 Auger 1.362 0.115 0.157 2900 48 Percussor 0.68 3.469 0.115 0.399 3521 147 Auger 1.217 0.084 0.102 2931 31 Percussor 0.57 3.874 0.084 0.325 2867 98RT3 (Sloped Basalt, HT Spring) VTP3 (Flat Basalt, HT Spring)Bench (Flat Basalt, RT Spring) Drilling Test RT = Room Temp VTP = Venus Temp & PressTotal Power  Output, Ave. (W)Est. Power  Consumption (W)Drilling  Depth (cm)Drilling Time (sec)ROP (mm/min)Specific  Energy (Whr/cc) 2.36 373 3.79 1.754.19 429 5.86 1.714.81 568 5.08 1.67 164 195 129RT3 (Sloped Basalt, HT Spring) VTP3 (Flat Basalt, HT Spring)Bench (Flat Basalt, RT Spring)261 358 202 NASA/CP—2018-219887 50
Like conventional DC motors, the continuous output torque of the HT motors is limited by their ability to  dissipate waste heat generated in the coils. The motors could be operated intermittently at higher torques  if they were periodically turned off and allowed to cool down but such operational constraints would not be  compatible with the unusually short life (only a few hours!) of a Venus surface mission. The motor peak  torque, on the other hand, may be limited by either the demagnetization potential of the PMs, which can be  determined from the characteristic B-H curve of the magnetic material, or the magnetic permeability of the  magnet steel used in the motor stator and rotor. The continuous and peak current capacity of the motor  controller and power supply should not be overlooked, either, though they are remotely located and not  affected by the test chamber environment.  Since the Sm 2Co17 PMs mounted in the rotor should be kept below a temperature of 525°C to minimize the  risk of demagnetization, the allowable temperature rise over the 10-minute drilling duration specified in the  proposed VISAGE mission timeline is 63°C (starting at an ambient temperature of 462°C). The difference  between the continuous average torque needed to drill to the target depth of 5 cm in 10 minutes and the  controlled constant torque load at which the motors reach their maximum allowable operating temperature  in the same period of time is the motor continuous torque margin. This margin can be accurately determined  using a dynamometer.  Figure 9 presents drill depth and motor thermocouple data from a successful drilling trial performed at VTP.  In this test, the drill reached a depth of 5 cm in 13.5 minutes and the drill motors reached steady-state  temperature with a 25°C temperature margin left on the percussor motor. If thermal constraints were  removed, however, a quick inspection of the high temperature torque vs. speed curve in Fig. 5 shows an  available torque of 0.4 Nm at 3000 rpm, or a margin of 23% over the average torque required to drive the  percussor from Table 5.  Figure 9. Drill depth and motor thermocouple data from a drilling test performed at full  Venus conditions  These results highlight the need for better thermal modelling of the motor, since the thermal coupling  between the stator, where the heat is generated in the coils, and the rotor, where the temperature-sensitive  PMs are mounted, has not been characterized. If it can be demonstrated that the rotor stays significantly  cooler than the stator while the motor is operating, the maximum allowable operating temperature of the  motor could be increased. In addition, passive heat sinks (e.g., fins, etc.) could be added to the motor  housing to improve heat dissipation to the ambient environment. After these thermal effects have been  accounted for and the motor torque margins have been accurately measured, the flight drill motors can be  correctly sized to achieve the desired drilling performance. 0123456 0 200 400 600 800Drill Depth (cm) Time (s)Drill Depth vs. Time (VTP Trial #5) NASA/CP—2018-219887 51
Summary  The HT motor model has been validated with test data, confirming that conventional methods can be used  to size these motors and account for losses at high temperature. The VMTF drilling trials have confirmed  the feasibility of drilling into rock at Venus temperature and pressure to generate fines suitable for  pneumatic transport per the proposed VISAGE mission concept of operations. They have also produced  drilling performance and average power consumption benchmarks useful for mission resource planning.  The motor torque margins are limited by the maximum operating temperature of the PMs. However, these  margins can be improved by accurate thermal modelling of the motor and the addition of passive heat sinks.  Acknowledgements  This work was funded by NASA’s HOT (High Operating Temperature) Technology Program. We owe our  sincere thanks and appreciation to Program Officer Dr. Quang-Viet Nguyen and Technology Coordinator  Dr. Gary Hunter. Part of this research was carried out at the Jet Propulsion Laboratory, California Institute  of Technology, under a contract with the National Aeronautics and Space Administration.  Appendix  Measured and simulated back-EMF waveform plots at 3000 rpm and two different operating temperatures  are included here for reference.  NASA/CP—2018-219887 52
References  1. Rehnmark, F. et al. (2017). VISAGE Rock Sampling Drill, Venus Modeling Workshop, May 2017, Cleveland, Ohio. 2. Kumar, N. (2014). High-Temperature Motors. In High Temperature Materials and Mechanisms (Ed. Y. Bar-Cohen), CRC Press, Boca Raton, Florida, 281-295. 3. Zacny, K. et al. (2017). Development of Venus Drill, Proc. of IEEE Aerospace Conf., Big Sky, Montana. 4. Rehnmark, F. et al. (2017). Performance Characterization of HT Actuator for Venus. VEXAG 2017, Laurel, Maryland. 5. Montone, D. Temperature Effects on DC Motor Performance, downloaded from http://www.pittmanmotors.com/Resources.aspx 6. Hanselman, D. (2006) Brushless Permanent Magnet Design, Magna Physics Publishing, 206-209. 7. Paulsen, G. et al. (2010). Rotary-Percussive Deep Drill for Planetary Applications, Proc. of ASCE Earth and Space 2010, Honolulu, Hawaii. 8. Rehnmark, F. et al. (2017). High Temperature Electromagnetic Actuator and Sampling Drill for Venus Exploration, Proc. of ESMATS 2017, Hatfield, UK. 9. Zacny, K. et al. (2013). Reaching 1 m Deep on Mars: The Icebreaker Drill. Astrobiology 13(12), 11661198. 10. Zacny, K. A., and G. A. Cooper (2007). Friction of drill bits under Martian pressure, J. Geophys. Res., 112, E03003, doi:10.1029/2005JE002538.-15-10-5051015 0 5 10 15 20 25Phase Voltage (V) Time (ms)Motor Back-EMF Simulation (3000 RPM, 500°C) Phase A Phase B Phase C NASA/CP—2018-219887 54
Electrical Discharge Mechanism  The problem starts with increasing surface roughness of the balls and races. The mechanism for  increasing surface roughness of the ball and races in a bearing is due to electrical discharge across the  ball to race interface. Friction increase has been observed in bearings at potentials as low as 6 V as  discussed in Reference 2. Increased coefficient of friction is the key ingredient necessary for the  discussion of anomalous friction signatures and plausible causes. The most significant variables with  electrical discharge in bearings are the Elasto-Hydro-Dynamic (EHD) film thickness (electrode gap), the  lubricant electrical properties (resistance and breakdown voltage), and the characteristics of the  differential voltage across the bearing, as depicted in Figure 2.   Figure 2. Electrical Discharge variables in an oil lubricated ball bearing  These variables are constantly changing. The EHD film thickness is highly dependent on the bearing  rotation rate, and goes to zero at zero speed. Therefore, at rest, the resistance across the bearings is  shorted and no electrical discharge can occur. The lubricant electrical properties are the most important  variable since a high electrical resistance is necessary to avoid bleeding off any charge developed across  the lubricant film. Synthetic lubricants, or highly refined lubricants, are especially susceptible to electrical  discharge due to their high electrical resistivity.   The differential voltage across the bearing has been measured to be +/-1 to 2 volts during normal  operation in some reaction wheels, which are induced by the quiescent motor current. It is assumed that  higher voltage differentials are present during space weather events, when friction events have been  observed. The cause of higher voltage differentials is not yet understood, but could be induced by single  event electrical effects and transients in the motor driver during periods of spacecraft charging.  NASA/CP—2018-219887 58
It is important to note that the effects of electrical discharge are microscopic due to the scale of the EHD  film thickness (~0.1 micron), which is a key variable. It is also important to note that the electrical  discharge damage mechanism is not a single event, but likely a series of discharge events happening at  high frequency. The induced voltage waveform measured in a reaction wheel revealed voltages induced  at many kHz, so discharges can occur thousands of times per second for significant periods of time.  Electrical discharge across the bearing lubricant film at sufficiently high voltages results in crater  formation. An example of the results of a discharge across a lubricant film is shown in Figure 3, which  also presents the inevitable hardness increase of the rim of the formed crater since the molten race  material will be instantly oil quenched by the lubricant in the bearing, resulting in formation of untempered  Martensite has a hardness as high as Rockwell C65, compared to the tempered hardness of C58 and  C60 for the race and ball, respectively. The height of the rim of the crater increases the roughness and  the higher hardness increases the abrasive properties of the roughened bearing surfaces.   Figure 3. Electrical Discharge crater formation resulting in increased roughness   The electrical discharge mechanism is also capable of improving the surface finish of bearings with low  voltage discharges less than 3 volts. The cases of friction recovery from an elevated state are believed to  be the result of electrical discharge while operating at high speeds when the EHD film thickness  increases, and large asperities formed by higher voltage electrical discharge events are the path of  discharge for the lower voltage differential induced by the motor. The result is the erosion of the  asperities, similar to the routine manufacturing process of electro-polishing of metal surfaces as shown in  Figure 4.   NASA/CP—2018-219887 59
Figure 4. Low Voltage Electrical Discharge electro-polishing resulting in improved surface finish  If the composite roughness has degraded to the point that it exceeds the film thickness even at the high  speed, the asperities continuously short circuit the lubricant film and the bearing can never go into  electrical isolation. In this condition, the re-polishing mechanism is ineffective and the bearing may  continue to degrade until the retainer is weakened by wear and fractures.   The in-situ bearing polishing mechanism is documented in more detail in a United Technologies  Corporation patent application listed as Reference 3.  Smallest gap between ball and asperities  results in low voltage electrical discharge  directly to highest asperities  Highest asperities are reduced and low  voltage discharges are directed to smaller  asperities  Smaller asperities are reduced and low  voltage discharges are directed to smallest  asperities  Smallest asperity is reduced and surface  finish is polished to the point that low  voltage differential does not discharge  NASA/CP—2018-219887 60
Anomalous Friction Signatures  Most observed friction anomalies can be segregated into dry friction (Coulomb kinetic friction) or wet  friction (Newtonian viscous friction). The magnitude of a dry friction increase is independent of rotational  speed and therefore remains present while crossing zero speed. This is a characteristic of rubbing friction  between two surfaces under load. Viscous friction increase is proportional to rotational speed, with no  change exhibited at zero speed, which is typically a response to Newtonian shearing of a fluid between  two surfaces moving relative to each other. The characteristic signatures of both of these types of friction  anomalies are represented graphically in Figure 5. It is typical in a reaction wheel application for a dry  friction increase to occur gradually over a period of a few days and then stabilize at the elevated friction  level, while the viscous friction increase is instantaneous.   Figure 5. Dry and Wet Friction Increase Signatures  Coulomb (Dry) Friction Increase  Dry friction in reaction wheel applications is typically characterized by a 1-2 mN-m increase in friction,  which is independent of speed. A perfect example of this type of friction increase is Case 1 from  Reference 4, shown in Figure 6, which clearly shows this type of friction increase occurring starting after  Day 39 from launch.  Figure 6. Example of dry friction increase in time domain (a) and relative to wheel speed (b).   The solid blue arrow identifies the inception of the friction anomaly, showing a gradual increase over a  few days.  NASA/CP—2018-219887 62
The most likely mechanism for this signature is retainer windup combined with a roughened surface finish  on the balls. Retainer windup occurs when sufficient asymmetry is present in the bearing raceway to  result in ball speed variation displacement exceeding the ball pocket clearance. This results in binding of  the retainer and rubbing in the ball pockets and between the retainer pilot diameter and outer race guiding  land, as shown in Figure 7. It is believed that the asymmetry is caused by an asymmetric increase in  surface roughness of the race due to electrical discharge. In laboratory tests, the increased friction  signature could not be duplicated by severe retainer windup alone, which supports the claim that  increased roughness of the balls rubbing in the ball pockets is also necessary for the friction to increase. Figure 7. Retainer Windup Rubbing Points due to excessive ball speed variation.  Viscous (Wet) Friction Increase  Viscous friction increases are a less common friction event for reaction wheels and are characterized by  an instantaneous increase in viscous friction, which is speed dependent. A perfect example of this type of  friction increase is Case 2 from Reference 4, shown in Figure 8, which clearly shows this type of friction  increase occurring instantaneously while operating at a constant speed of 1400 rpm on Day 45 from  launch. Subsequent on-orbit testing varied the wheel speed to characterize the friction signature, which  proved that the change is clearly speed dependent.  Figure 8. Example of viscous friction increase in time domain (a) and relative to wheel speed (b). The  solid red arrow identifies the inception of the friction anomaly, showing an instantaneous increase which  is dependent on speed, decreasing to no change at zero RPM Days 31-45 Day 45On-orbit  Testing  Days 52 & 53  Days 58 & 59 (a) (b)  NASA/CP—2018-219887 63
This friction change signature is believed to be due to a change in ball control from inner race spin control  to outer race spin control. The balls in an angular contact ball bearing have a spin component, and in a  stable bearing, the controlling race will have pure rolling at the contact and the other race will have a  slight amount of spin at the contact. Which race is the controlling race is determined by the values of  resisting spin torque at each contact ellipse, with the higher resisting torque becoming the controlling  race. Since the load on each ellipse is the same, the resisting torque will be primarily a function of the  contact ellipse major axis dimension ‘a’ and the coefficient of friction. Most bearings have a larger contact  ‘a’ dimension on the inner race, which results in inner race control for a normal bearing. If an electrical  discharge event occurs which quickly increases the roughness of the outer race, thereby increasing the  coefficient of friction, the spin control could quickly change to the outer race causing the inner race to take  the spin component as shown in Figure 9.  Figure 9. Inner race control changing to outer race control due to the increase in friction coefficient of the  outer race resulting from an electrical discharge   The result of the change in ball control is a change in the viscous friction torque in the EHD film at the  non-controlled contact ellipse due to ball spin. The spin of the ball on the thin EHD film will result in  Newtonian shear, which can be imagined as the friction torque resulting from a disk spinning on a plate  separated by an extremely thin viscous fluid. The larger ‘a’ dimension from the ball spin on the inner race  instead of the smaller value on the outer race would result in an increase in the viscous friction value. An interesting anecdote supporting this model occurred during a  long-term reaction wheel test when a torque change with a signature  nearly identical to the on-orbit example in Figure 6 occurred. The  new friction level was stable and the wheel was operated for a  considerable period of time following the friction change. At the end  of the test, the bearings were removed and inspected and it was  observed that a ball had two circumferential ball wear bands with two  different widths, one the approximate width of the inner race ball  track and one the approximate width of the outer race ball track, as  sketched in Figure 10. This is what would be expected if the spin  control shifted from inner race to outer race and spent considerable  operating time in each state.  The viscous friction anomaly has not been known to result in a reduction in lifetime, which is also  consistent with the model. There is no progressive wear encountered, no binding or windup of the cage,  merely an increase in surface roughness on one race, which changes the balance of forces and torques  within the bearing.  Figure 10. Artist sketch of  ball inspected after viscous  friction change NASA/CP—2018-219887 64
Oil Jog  Viscous friction spikes in some applications have been attributed to ‘oil jogs’. Tests have shown that  these “oil jog” torque signatures can be a result of electrical discharge; a theory for the phenomena will be  presented. “Oil jogs” are characterized by an abrupt increase in friction, usually while operating at a  constant speed, followed by a decay lasting seconds, minutes or hours, as approximated in Figure 10.  They have been frequently attributed to a sudden migration of lubricant into an area of the bearing where  viscous churning results in a temporary increase in torque, and which decays as the lubricant is  redistributed. These instantaneous jogs were duplicated in tests by merely applying a small voltage  across the bearings, with both positive jogs and negative jogs, as depicted in Figure 11, which does not  strongly support the common hypothesis that they are the result of lubricant dynamics. While it is hard to  imagine a mechanism for the sudden burst of lubricant migrating into a bearing, it is even harder to  imagine a mechanism for a sudden migration of bulk lubricant out of a bearing.   Figure 11. Positive and Negative Oil Jog Signatures  An alternative scenario of the mechanism  generating these events is the result of electrical  discharge. The same mechanism described in the  previous section for the viscous friction increase  may apply to this transient event. The abrupt  increase can be due to an electrical discharge event  increasing the coefficient of friction on the outer  race, which transfers ball control to the outer race,  and the gradual recovery to inner race control  occurs as the low voltage electrical discharge from  the motor re-polishes the outer race, as depicted in  Figure 12. This scenario can apply to a negative oil  jog as well, where the bearing is in a stable outer  race control condition, and an electrical discharge  event increases the coefficient of friction on the  inner race, resulting in an abrupt decrease in  viscous friction and similar gradual recovery.  Figure 12. Electrical discharge scenario for  Oil Jog Signature NASA/CP—2018-219887 65