ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
Likewise, small changes in the roll angle, /g84R, of the shoulder hinge provided variable deployment paths.  This was evident during the -Y Solar Array's post-vibe deployments. The deployment dolly was inadvertently misaligned and the post-vibe deployment took a much different path than the pre-vibe  deployment, resulting in a stalled deployment on a highly warped g-negation table that was specifically  placed in a spot where we did not expect the deployment to reach based on pre-vibe results. In Figure 24, at the identical time in deployment, the different deployment trajectories can be seen. With /g84 Rbeing out of  spec by approximately ±0.10 /g113, we can see that the wing is not deployed as much as in the case with less or virtually no roll angle. This deployment with excessive roll angle continues on a path uncharacteristic of the expected deployment. To adjust for pitch and roll, the dolly jacks at each corner of the spacecraft dolly were manipulated.                   Figure 25.  Roll Angle Measurement at the Shoulder Hinge To achieve an optimal setup for a post-vibe deployment, the pitch and roll angles were measured and  adjusted closest to the deployment root hinge, the shoulder hinge, while the arrays were stowed. Figure  25 shows the location that we could measure directly on the shoulder hinge to assess roll. Due to the tight  packaging of the stowed array, the closest area we could measure pitch was across the two solar array booms between the Shoulder Hinge and the Elbow Hinge, as seen in Figure 26. A Digi-pas brand 2-axis  precision digital level with an accuracy of 0.01 /g113was used for these measurements.                   Figure 26.  Pitch Angle Measurement on Deployment Boom The air hose configuration of the g-negation system, as seen in Figure 27, is nebulous in design and was  photographed repeatedly from different vantage points to achieve a repeatable system. Slight changes in the air routing had the same effects as slight changes in the pitch and roll angle with respect to changing the deployment path and deployment time by generating friction at points in the system different from the desired baseline. By keeping accurate records of the pitch and roll angles as well as the g-negation hose routing, the solar  array deployment times and deployment paths were in family. Telemetry graphs from the +Y Solar Array  Wing observatory level pre-vibe and pre-ship deployments, illustrated in Figure 28 and Figure 29, show only a 3 second difference in deployment time, a 2.91% difference, and similar angular velocity characteristics between these two different deployment setups performed 3 months apart 346
discussed. Different Motorization Ratio computations are compared. Finally, a new motorization ratio  computation is proposed.   Background The three main institutional bodies that overlook space-based mechanisms are: NASA, ESA and AIAA.  They each recommend slightly different methods of computing the margin on electrically driven drives.  The three methods are presented below:  Presentation of Different Mo torization Rati o Requirements (1) NASA-STD-5017        Available Driving Torque   Operating Torque Margin =                                          - 1      Total Resisting Torque /g3 When test verified, an operating torque/force margin of 1.0 or greater is required at all applicable points of  travel. Only verification by analysis shall require prior review and approval. (2) ESA - ECSS-E-ST-33-01C Table 1.  Recommended Theoretical  and Measured Uncertainty Factors Component of  resistanceSymbolTheoretical  FactorMeasured  Factor Inertia /g44/g3 1.1 1.1 Spring /g54/g3 1.2 1.2 Motor magnetic losses /g43/g48/g3 1.5 1.2 Friction /g41/g53/g3 3 1.5 Hysteresis /g43/g92/g3 3 1.5 Others (Harness) /g43/g36/g3 31 . 5 Adhesion /g43/g39/g3 33 The minimum actuation torque (Tmin) shall be derived by the equation: Tmin= 2 x (1.1 I+1.2S+ 1.5 HM+ 3FR+ 3HY+ 3HA+ 3HD) +1.25 TD+TL  Where: /g16/g3 /g44 is the inertial torque applied to a mechanism subjected to acceleration in an inertial frame of  reference (e.g. spinning spacecraft, payload or other). /g16/g3 /g55/g39is the inertial resistance torque caused by the worst-case acceleration function specified by  the customer at the mechanism level. /g16/g3 /g55 Lis the deliverable output torque, when specified by the customer. 350
(3) AIAA–S-114-2005 Table 2. Minimum Recommended Static To rque or Force Margin for New Designs Design Phase Force or Torque Margin Conceptual Design Review 175% Preliminary Design Review 150% Critical Design Review 125% Acceptance/Qualification Test 100% The static torque margin shall be expressed in percentage and is calculated as follows: Drive torque % =       - 1   x 100                Resisting torque The three methods of calculating Motorization Ratios (MR) as described above are the most widely  accepted in the space industry. However, each method results in different ratios. It is quite obvious that defining a standard way of calculating MR to be applicable to every motor technology, design maturity  level, and duty cycle is very challenging and not practical, resulting in unnecessary high level of  conservatism, especially for well proven heritage designs. Effects of Power Off Time on Output Torque and MR Extensive testing has shown that output torque drops from the 100% duty cycle value to the unpowered  holding torque and below as the time off duration is increased. As expected, when power is removed from the motor, only the magnetic torque prevents the motor rotor from backdriving under the drive train windup torque. It all becomes a question of timing. If the power is removed long enough for the rotor to move back by more than one step, then when the power comes back on to command the rotor to move to the next step, it is too late and the motor loses synchronization. Effects of Maturity and Heritage and Available Power on Required MR The rotary actuators discussed are of MOOG Type 3 and Type 5 family and have extensive in flight  heritage and have gone through several life tests. Each of the life tests have proven that the rotary  actuator internal losses do not increase with time in vacuum or mechanical use. The most recent life test  on a Moog Type3 actuator accumulated 5.3 million output degrees of motion without any sign of degradation. The consistency of measured threshold voltage/current throughout the life program at hot, cold and  ambient conditions, especially the consistency of the initial functional and final functional is indicative of the fact that the internal losses did not increase after exposure to thermal vacuum and mechanical environments of the life test.  Rotary actuators of this design are also very simple, only one reduction stage (160:1) with the minimum  number of bearings and gear meshes.  351
Table 3. Summary of Life Test Threshold Voltage Results (Type 3 actuator)  Test ConditionPrimary Redundant CW CCW CW CCW Initial Functional  @ 22.3°C 11.01 11.00 11.11 11.10 One Times Life @ 107.1°C 13.00 14.00 12.40 13.20 One Times Life @ -54.3°C 10.00 9.50 9.50 9.50 One and a Half Times Life @ 25.1°C 11.61 11.92 11.60 11.62 Final Functional @ 24.2°C 11.30 11.30 11.30 11.20 Partial Disassembly and  Examination Initial Functional @ 24.6°C 11.40 11.70 11.20 11.60 Initial Functional in Vacuum @ 19.7°C 11.30 11.30 11.50 11.70 One Times Life @ 106.2°C 12.70 13.50 12.60 13.60 One Times Life @ -54.1°C 9.50 9.40 9.50 9.40 One and a Half Times Life @ 105.4°C 12.50 13.20 12.80 13.50 One and a Half Times Life @ -53.7°C 9.80 9.70 9.80 9.80 One and a Half Times Life @ 22.6°C 11.70 11.60 11.50 11.70 Final Functional @ 21.7°C 11.00 11.20 11.00 11.10 Table 4. Summary of Life Test Threshold voltage Variations ConditionMax. Threshold  VoltageMin. Threshold  VoltageAve. Threshold  VoltagePercent  Variation Ambient 11.7 11.00 11.4 3.0 HOT 14.0 12.50 13.1 5.2 COLD 10.0 9.40 9.6 4.2 352
Table 5. Summary of Life Test Threshold Current Results Test StepMotor   HousingThreshold  VoltageThreshold  CurrentDelta/mean value Temperature CW CCW CW CCW CW CCW /g113/g113C (V) (V) (mA) (mA) (mA) (mA) Initial functional testing 22.3 11.0 11.0 175 175 -1.1 -2.6 One Times Life testing @ HOT 107.1 13.0 14.0 152 164 4.3 4.7 One Times Life testing @ COLD -54.3 10.0 9.5 212 203 4.3 -0.3 One and a Half Times Life @ AMB 25.1 11.6 11.9 179 184 2.9 6.4 Final functional testing 24.2 11.3 11.3 178 177 1.9 -0.6 Initial functional testing 24.6 11.4 11.7 173 178 -3.1 0.4 Initial functional testing in vacuum  @ AMBIENT19.7 11.3 11.3 174 174 -2.1 -3.6 One Times Life testing @ HOT 106.2 12.7 13.5 144 158 -3.7 -1.3 One Times Life testing @ COLD -54.1 9.5 9.4 202 199 -5.7 -4.3 One and a Half Times Life @ HOT 105-4 12.5 13.2 147 156 -0.7 -3.3 One and a Half Times Life @  COLD-53.7 9.8 9.7 209 206 1.3 4.7 One and a Half Times Life @ AMB 22.6 11.7 11.6 180 178 3.9 0.4 Final functional testing 21.7 11.0 11.2 174 177 -2.1 -.06 Figure 1.  Life Test Threshold Current ResultsThreshold Current (mA) 050100150200250 Initital functional testing 1 x life @ HOT1 x life @ COLD 1.5 x times life @ AMB Final functional testing Initital functional testing Functional testing in vacuum @ AMB 1x life @ HOT1x life @ COLD 1.5 x life @ HOT1.5 x life @ COLD 1.5 x times life @ AMB Final functional testing Test CaseCurrent (mA)CW CCW 353
Space Robotics Example Other important parameters to take into account when determining what motorization ratio is sufficient is  the available power compared to the size of the mechanism and the simplicity of the electric motor used.  When relatively large amounts of power are available and the motor is a brushless DC from which the  torque versus speed curve is well known, a reduced motorization ratio might be contemplated. As an example of this, the drive trains of the Latching End Effector (LEE) of the Space Station Remote  Manipulator System (SSRMS) are presented with their characteristics and margins. The SSRMS has  been in orbit since April 2001 and has been performing nominally even if its motorization ratios during the design phase were relatively low compared to current requirements. Figure 2. SSRMS LEE in orbit /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g48/g39/g36/g3/g38/g82/g85/g83/g12 The LEEs have three separate mechanisms each containing complex drive trains. The first mechanism is  the snare, it rotates a ring on which cables are attached and permits to snare the grapple probe of the Grapple Fixture (GF) affixed to a payload. The second mechanism is the rigidized mechanism, it pulls  back a 45-lbf (200-N) carriage inside the LEE which contains the snare mechanism and applies up to 1,800-lbf (8,000-N) axial preload between the LEE end ring and the GF. This is sufficient to handle small payloads. The third mechanism is the latch; it adds 15,000 lbf (67,000 N) of axial preload between the  LEE end ring and the GF to permit handling large payloads and also mates and umbilical. The complete system is designed to fully capture a 100,000-kg space shuttle in 30 seconds. The available electrical  power to perform this feat is 300 W. Pictures of each drive train gearbox are presented in Figures 3 and 4.  Figure 3. SSRMS LEE Snare Gearbox, Scale: OD = 4.86 inches (0.123 m)  /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g48/g39/g36/g3/g38/g82/g85/g83/g12 354
Figure 4. SSRMS LEE Rigidized Gearbox With Ball screw, Scale: ballscrew = 18-in long (46 cm)  /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g48/g39/g36/g3/g38/g82/g85/g83/g12/g3 /g3 Figure 5. SSRMS LEE Latch Gearbox, Scale: OD = 5.4 in (14 cm) /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g48/g39/g36/g3/g38/g82/g85/g83/g12 The characteristics of each drive train are presented in Table 6. As can be seen in Table 6, during the  design phase, the required worst-case torque margins were only 10%. The estimated worst-case torque  margins were only 16% for two of the three mechanisms. This compared to the institutional minimum requirement of between 100% and 200%. This is evidence that it is possible to design successful  mechanisms with much lower motorization ratios than currently required.  The authors do not believe the same Motorization Ratio should apply to the above designs as that of a newly designed complex multi-stage low-power drive train that has never flown. Effects of Different External Load Characteristics External loads may have different characteristics when applied to the output of the actuators. Pure  resistive torques (friction), only resist to the actuator applied output torque. As soon as the actuator stops, the resistive torque drops to zero, as the pure frictional torque only opposes motion.  Spring, dynamical oscillatory response, and gravity torques are very different than the pure friction  torques; they remain even when the motion command has stopped. That is why when the applied torque is removed by removing power to the windings, the gravity or the wind-up spring torque may overcome  the unpowered holding torque and back-drive the stepper motor. This phenomenon has been investigated  by conducting multiple tests at reduced duty cycles when power is removed for a proportion of the pulse  period. The effect of the temperature on the performance of actuator for reduced duty cycle has been studied to select an acceptable and appropriate correction factor. 355
Table 6. Drive Train Characteristics Gear   Ratio No of  StagesNo of  gearsNo of  bearingsRequired  Torque  Margin  on  nominal  caseEstimated  Torque  Margin on  nominal  caseRequired  Torque  Margin  on Worst  CaseEstimated  Torque  Margin on  Worst  Case Snare 230:1 3 planetary,  2 spur21 19 100% NA 10% 252% Rigidize 11.6:1 1 spur, 2  planetary,1  ballscrew*15 19 100% 414% 10% 16.7% Latch 10.2:1 2 planetary,  2 spur, 4  ballscrews*19 25 100% 134% 10% 16% *ballscrew pitch: 0.2 in/rev (5 mm/rev) Measuring Torque at Reduced Duty Cycle on Moog Type 3 Actuators Background Moog has measured torque outputs of Type 3 Actuators using two different heritage testing methods.    The first method was to measure the reduced duty cycle torque using a standard torque transducer and  magnetic particle brake. In this method, the load applied is increased on the magnetic particle brake until the actuator stalls. This method tends to measure a torque a little less than unpowered holding torque  and is a result of the unpowered holding torque and wind-up in the actuator and the test stand. The second heritage method to measure torque of the actuator at reduced duty cycle was to add weight  to a moment arm and gradually add more weight until the actuator backdrives. The results of this test  correlated with the unpowered holding torque. When in the power OFF phase, the moment works against  the unpowered holding torque. Beyond this level, backdriving occurs. This testing was done at numerous reduced duty cycles at 1 PPS and at 10 PPS. It was shown that the motor and actuator unpowered  detents correlate. The values obtained were in the 22.6 N-m (200 lb-in) range (motor level testing showed 14 N-cm, which with a 160:1 gear reduction would result in 22.6 N-m at the output). Discussion Both of the methods described in the background section above are thwarted because the dynamic, static  and loss loads are combined into a torque that has a 3 times factor on it. This requires the actuator to  overcome the torque at both the powered portion of the pulse and the unpowered portion of the pulse. Since movement of the actuator occurs only during the powered portion of the pulse, the demonstrated torque should be 3 X (IL + RT + IT), where IL is internal loss, RT is resistive torque, and IT is a calculatedtorque from the inertia load. During the unpowered portion of the pulse, the unpowered holding torque  needs to hold only the maximum static load, which is the resistive torque and the addition of the inertial  torque. To further pinpoint the problem, internal loss (actuator output turning) is only present during the powered  pulse and is not present when the actuator is at the unpowered portion of the pulse (non-dynamic). Inertia is present during actuator powered potion (motion) and the unpowered portion (dynamical response) of  the pulse. Accurate measurement of torque during the powered portion of the pulse is needed to be able to show compliance.   356
CF *(UHT) SMR = ------------------- >3      ( I T  +  R T ) Where: UHT= Actuator Unpowered Holding Torque Nm (in-lb)  IT = Inertial Torque at output  RT = Resistive Torque  CF =           Correction Factor. /g38/g41/g3/g68/g70/g70/g82/g88/g81/g87/g86/g3/g73/g82/g85/g3/g79/g82/g86/g86/g3/g82/g73/g3/g75/g82/g79/g71/g76/g81/g74/g3/g87/g82/g85/g84/g88/g72/g3/g90/g75/g72/g81/g3/g83/g82/g90/g72/g85/g3/g76/g86/g3/g85/g72/g80/g82/g89/g72/g71/g3/g90/g75/g76/g79/g72/g3 /g87/g75/g72/g85/g72/g3/g76/g86/g3/g86/g82/g80/g72/g3/g85/g72/g86/g76/g71/g88/g68/g79/g3/g78/g76/g81/g72/g87/g76/g70/g3/g72/g81/g72/g85/g74/g92 One can notice that the minimum SMR will correspond to the motor step/location featuring the  minimum detent torque. To that extent, the SMR shall be determined step by step over at least one  motor rev (2 revs preferably), with a sufficient OFF time between each step. The DMR and SMR are therefore relying on complementary properties of the motor, in particular the  mapping of the detent torque over one motor rev. The ideal motor would feature a ratio close to 1 between maximum and minimum values of the detent torque over one rev – as the maximum detent is  detrimental to DMR and the minimum detent is detrimental to the SMR. This reflects how accurately the motor is controlled by design and by manufacturing. Testing and Results: Unpowered Detent Torque As stated above, moment arm testing at 1 PPS and 10 PPS at numerous reduced duty cycle values  yielded the minimum unpowered holding torque values as the actuator’s torque output. These results showed that the actuator could only demonstrate torques less than the unpowered holding torque levels  during the un-powered portion of a reduced duty cycle. The test results have fidelity to 0.5 N-m and do  not take into account the friction, any wind-up of the torque transducer and couplings between the moment arm and the actuator.   It was demonstrated that the decrease of torque below the unpowered holding torque was caused by a dynamic component with regards to powered cardinal position versus unpowered holding position, test equipment (couplings and torque transducer), harmonic drive wind-up, possible inertia dynamics, and  temperature of operation.   Table 7.  Moment Arm (Pendulum) Test Results SPEED    (PPS)DUTY  CYCLE  (%)DUTY  CYCLE  (msec)VOLTAGE  (Vdc)TORQUE  PRIMARY  (CW) Nm (lb-in)TORQUE  PRIMARY  (CCW) Nm (lb-in)TORQUE RED.  (CW) Nm (lb-in)TORQUE  RED.  (CCW) Nm (lb-in) 1 4 40 23.13 21.5(190) 21.5(190) 21.5(190) 21.5(190) 1 5.1 51 23.13 21.5(190) 22.6(200) 21.5(190) 22.0(195) 1 8 80 23.13 22.6(200) 22.0(195) 21.5(190) 22.0(195) 1 8 80 27.39 22.6(200) 22.6(200) 22.0(195) 22.0(195) 10 40 40 23.13 21.5(190) 20.9(185) 21.5(190) 20.3(180) 10 51 51 23.13 20.3(180) 19.8(175) 20.9(185) 19.8(175) 10 80 80 23.13 21.5(190) 20.9(185) 21.5(190) 20.9(185) 10 80 80 27.39 21.5(190) 20.9(185) 22.0(195) 21.5(190) 358
The range of results at 1 PPS through duty cycles of 4%, 5.1% and 8% is between 21.5 to 22.6 N-m of  unpowered detent torque. The range of results at 10 PPS, keeping the duration of the powered portion of the pulse equivalent to the 1 PPS data, shows a decrease in the measured torques to 19.8 to 22.0 N-m. The need for a correction factor for the Static Motorization Ratio calculation can be seen in the difference  between the moment arm measured unpowered detent torque and the motor screening values. This particular actuator’s motor had a 14 N-cm (22.6 N-m at actuator output) minimum unpowered detent torque during motor screening testing.   For this reason Moog recommends using a .75 correcti on factor (CF in SMR formula) for the Static  Motorization Rati o calculation. Powered Torque Using a KistlerTMTorque Sensor, a motor was tested to determine the actual torque output of the motor at  various speeds and duty cycles. This sensor measures reactive torque of the stator housing when the rotor takes a step. The motor housing was attached to the Kistler TMtorque table and the rotor was pulsed.  The motor housing resultant (equal but opposite) torque was recorded. This torque was compared to the 100% duty cycle torque measured during motor screening testing. Correlation between the motor screening test at 100% duty cycle and the tests results taken with the Kistler TMTorque Sensor are  presented in Table 8. Table 8. KistlerTMTorque Sensor vs. Herit age Torque Measurement SPEEDDUTY  CYCLE  (%)ON  time  (msec)KISTLER  MEASURED  TORQUE       N-cm (in-oz)HERITAGE  MEASURED  TORQUE       N-cm (in-oz) 1 4 40 38.3(54.2) 1 5.1 51 38.0(53.9) 1 8 80 38.3(54.2) 1 50 500 38.4(54.4) 1 96 960 38.2(54.1) 1 100 1000 37.5(53.1) 10 80 80 38.9(55.1) 32.5(46.0) 10 96 96 38.3(55.0) 10 100 100 37.1(52.5) 35.9(50.8) 16 80 50 38.2(54.1) 16 96 60 38.3(54.3) 16 100 62.5 37.2(52.7) The consistency of measured torque values show that the torque remains a function of current and  windings throughout the testing regimes as measured with the KistlerTMTorque Sensor. See a graph of  the results in Figure 6.  359
Figure 6. KistlerTMTorque Sensor vs. Heritage Torque Measurement Interestingly, the comparison of wave forms between the slowest speed and shortest powered pulse  (1 PPS and 4% duty cycle) and all three of the 100% duty cycle tests (1, 10 and 16 PPS) showed correlation. See the comparison of the wave forms in Figure 7. Figure 7. Wave Forms of Motor Reaction Torques 1stPeak=Powered torque minus UHT 2ndPeak=Reduced level powered torque plus UHT contributionWave Forms of Motor Reaction Torques Aligned to within 0.1 msec 360
Actuator torque capability is reached after unpowered detent torque is overcome. The first peak shows  this resulting torque which was used to show the torque capability of the motor. The second peak is the overshoot of the rotor to the cardinal powered position in combination with the unpowered detent torque helping to pull the rotor into the cardinal detent position. It should be noted that the major oscillations dampen out within approximately 45 msec of command at all three speeds and at the one reduced duty cycle compared. The results above show that there is little change in torque between the powered portions of a pulse  when comparing torques at 100% duty cycle and reduced duty cycle pulses. U npowered Detent Tor que is  constant throughout the temperature extremes. Heritage testing at 100% duty cycle should be adequate to use in the powered portion  Similar tests were conducted independently on two (2) Moog Rotary Actuators at Astrium. Following is a  summary of the tests and the conclusions.  The rotary actuator was gravity loaded using a moment arm. Fine and coarse potentiometers were  monitored by the electrical test equipment during the test. To evaluate the dynamic performance, the  Back Emf voltages of the redundant phases were plotted using an oscilloscope. Temperature was monitored throughout testing using an ohmmeter measuring the redundant winding resistance. The  torque test was performed at 30 to 60 percent duty cycle. Figure 8. Mechanical Setup at Astrium /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g36/g86/g87/g85/g76/g88/g80/g12 The test method was to increase the gravity loading as long as no miss-stepping is observed (over 240  steps). B emfvoltage was checked on the redundant winding. Figure 9 shows a sample of gathered data. 361
Figure 9. A Sample of Gathered Data /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g36/g86/g87/g85/g76/g88/g80/g12 Figure 10 highlights the behavior of the motor during transition from ON to OFF state. The testing  concluded that the static torque is at its minimum for OFF time greater than 10 msec and it is independent of the applied voltage Vcc and of duty cycle,   Figure 10.  B emfScope Plot of T_On & T_OFF /g11/g38/g82/g88/g85/g87/g72/g86/g92/g3/g82/g73/g3/g36/g86/g87/g85/g76/g88/g80/g12/g3

362