ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
The mechanism drive system consists of a three-phase stepper motor with an integral RA and gear-head  with a 96:1 gear ratio (N). The gear-motor is coupled to a drive shaft with a flexible coupling. Both the  motor and the drive shaft are secured to the flexible coupling with match-drilled pins. A preloaded,  separated bearing pair supports the drive shaft; and the drive shaft in turn supports the balanced load  inertia. Because the load is balanced, the detent torque of the stepper motor is sufficient to hold the load  in place during launch and no launch lock is necessary. Figure 1 shows a schematic of the drive system.  Figure 1.  The Drive  The total mass of the suspended load is 1.36 kg and the mass moment of inertia of this rotating load (J L) is 0.027 kg•m2. When reflected through the gear-head, the load inertia (J LM) [3] is:  JLM = (J L) / N2 = (0.027 kg•m2) / 962 = 2.9 X 10-6 kg•m2….…….........(1)  The mass moment of inertia of the 3.8-cm (1.5-inch) diameter stepper motor (J M) is 2.8 X 10-5 kg•m2. The  inertia factor (J F) [3] is defined as:  JF = (J M + J LM) / J M……………………………………….………………...(2)  For this system the inertia factor is 2.64. A good rule of thumb is that the inertia factor of a stepper motor  system should be less than three. Additionally, torque margin was calculated to be 126% using a factor of  3 for all friction values, meeting the MIL-A-83577 requirement of 125% for mechanism at their Critical  Design Review. Analysis indicated a functional design.  Position Knowledge Several options were considered to provide rotational feedback of the mechanism, including contacting  and non-contacting switches, encoders, potentiometers and RAs. Contacting switches were ruled out because they are inherently limited life items, while encoders and potentiometers were eliminated  because of their relatively large size. The RA was selected over the non-contacting switches, such as Hall  Effect sensors, because the RA was easier to mechanically integrate to the system since it is integral to  the motor, requires fewer lead wires, has higher reliability numbers, higher positional knowledge and is  easier for software to implement.  The RA provides the added feature of monitoring the heath of the drive system. If the peak acceleration  decreases, the drive is degrading; and this knowledge can enable the operations team to modify their use  of an on orbit mechanism before it fails. A technical disadvantage of using a RA for feedback is that it derives position of the motor shaft, not the output of the gear-head shaft. As a result the accuracy of the  position feedback provided by the RA is limited by the backlash of the gear-head, which for our  application was sufficient for positional knowledge. There is also a procurement disadvantage, as the  technology is patented by CDA InterCorp, and thus is only available from one supplier. There are other  30
suppliers with acceleration feedback technology, but they require DC excitation, which results in  additional power loss and performance variation with supply voltage fluctuations.  EDU Performance  The EDU was necessary for calibrating and determining the gains in the electronics to determine proper  step verification. More on this process will be discussed later in the paper. The Responsible Engineer for  the control electronics observed the proper direction of rotation and displacement for stepping the  actuator, however, the observed output of the RA showed significant overshoot and ringing of the stepper  motor. Figure 2 shows the RA output from a single step of this development unit.  RA Output for Single Step -0.2-0.15-0.1-0.0500.050.10.150.2 0 0.02 0.04 0.06 0.08 0.1 0.12 Time (sec)Volts DC Figure 2.  Oscilloscope Output of the Underdamped EDU Mechanism  This plot of data captured with an oscilloscope shows the significant overshoot and ringing of the  mechanism. Figure 3 shows the RA output of 6 steps being driven at the operational pulse rate of 26  Pulses Per Second (PPS).  31
Figure 3.  Six Steps at the Operational Pulse Rate of 26 PPS  This oscilloscope trace shows that the motion of the mechanism was erratic and significantly  underdamped. The ringing was both audible and visible in the lab, raising the concern that loads might be  high enough to cause a premature mechanical failure of the drive. The specific concerns of the team  centered on if a resonance was excited a bearing retainer could fail, or the gears could micro-pit due to  the vibration and reversing. Also problematic in this performance was that the overshoot acceleration  voltage was nearly equal to the primary step voltage. Our “Original Transition Voltage Counting” method  yielded false positional information as the threshold voltage is realized at the primary step as well as the  overshoot. The processed logic would falsely determine the motor was simply stepping back-and-forth  between two pulses. (Original Transition Voltage Counting is described in the later section Processing the  Rotary Accelerometer Signal .)  Investigation of the Problem  Engineering Review Board Process Due to schedule constraints, EDU testing and flight assembly occurred concurrently. By the time, the  EDU testing showed that the designed method for step verification did not work as intended, the flight  stepper motor had already been received, almost all of the mechanisms piece parts had been completed  and the flight electronics were in fabrication. Re-work of flight hardware was going to be required to solve  this problem, thus an Engineering Review Board (ERB) was convened to determine the best system  solution for the program. The first step in the ERB process was to brainstorm every possible root cause.  Figure 4 shows the fishbone diagram developed by the ERB team as a result of this brainstorming effort.  32
Figure 4.  ERB Fishbone Diagram  An audit of the EDU electronics and Special Test Equipment’s documentation and certification logs  determined the electronics were built to print and certified. Review of the test procedure and set-up  eliminated the test procedure and operator error. Review of test data supplied by the motor supplier  eliminated backlash, motor windings and the accelerometer. A pin-to-pin test of the wire harness verified  its manufacture. Torque margin of the EDU was measured using the RA to be greater than 400%,  however, when a reduced voltage was applied to the EDU the system still exhibited significantly  underdamped actuation. Therefore, excessive torque margin was eliminated. While the inertial factor of  2.64 was below the desired value of three, because the rotating mass was so large, at 1.36 kg, the initial  determination of the ERB was that the system was under-damped because the inertia factor was too high, and the trade table of solutions for this problem was created.  A Note On Torque Margin Torque margin is measured using an RA by increasing the load friction until the drive begins to miss steps; the ratio between the nominal RA output and the RA output when a step is missed is the torque  margin [1]. For the measurements made using the EDU, the load inertia was increased by applying hand  pressure to the rotating load, and an operational torque margin of greater than 400% was measured. The  calculated torque margin at CDR was estimated to be 126%. This large discrepancy between the  calculated and actual margins is a result of performance projection without hardware to empirically test  our CDR analysis. Standards require the assumption of worst case gear-head efficiency, motor detent  torque, and most significantly a factor of safety on all friction values. This is sufficient justification for using  EDUs early in a system’s development. Actual torque margin measurements can be made, potentially enabling the use of a smaller stepper motor or increasing efficiency by reducing power consumption.  Troubleshooting and Root Cause Determination Table 1 below delineates the trades considered for solving the excessive inertia factor problem. We  sought the most technically robust solution that minimized cost and schedule impacts to the program. The  trades were complicated by the fact that both flight and EDU hardware were in house, and any change to  the mechanical configuration required an in-process change to existing hardware. Mechanism Electronics Test Configuration Operational Error Backlash  Motor Windings  Wiring Harness  Accelerometer  Excessive Torque Margin  Excessive Inertia Factor Design Error  Fabrication Error  Part Failure Drive Current  Pulse Modulation  Step Rate Test Procedure Error  Test Script Error  PWA STE Design/Fab Error Test Set-Up Incorrect  Incorrect Commanding  PWA STE S/W Error  33
Figure 5.  Six Steps at 64 PPS  Another interesting result was that when the system was driven faster than half the resonance of the  ringing the amplitude of the acceleration of the second step decreased (see Figure 6). This behavior was  consistent for all rates greater t han 64 PPS. Finally, if the system wa s driven at any rate between the  operational pulse rate of 26 PPS and the sweet spot of 64 PPS, the only change was that there were  fewer bounces the closer the rate got to 64 PPS. 

Figure 6.  Six Steps at 92 PPS  The next best, solution from a cost and schedule standpoint, was minor modification of the electronics,  therefore our testing moved onto shaping the pulse to reduce the acceleration. One of the modifications  tested was to add large capacitors to the drive circuitry to “soften” the edges of the voltage pulses. This  technique is known as the “Miller Effect”. Several one /g1F capacitors were added, and while there was  some benefit to the mechanical resonance, the underdamped situation still persisted. Also power losses  on the electronics board increased, and there was no “real estate” on the board to add the sizable  capacitors. Therefore, this option was eliminated. This, and other methods, did reduce the audible noise  35
becomes more complicated at high velocities, but for applications where the rotor velocity is below 200  rad/sec, this relationship is valid.  Bm /g2 Km2……………………………………………………….(3)  Armed with this information, we calculated what winding modifications were required to achieve reduced  holding torque and increased damping, by adding a separate set of shorted windings. The reduction of  one gage of magnet wire reduces the volume of the turns and increases DC Resistance by 26%. Since  we wanted to reduce torque constant as well, we needed to reduce the number of turns into the motor  phases. A straight reduction in turns will reduce the torque constant by the same percentage. Table 2  describes the original EDU performance (EDU#1) and our desired “New” motor performance.  Table 2.  Motor Design Characteristics Comparison  Parameter Units EDU#1  (Baseline) “New” Motor  (EDU#2 & Flight) Comment  Torque Constant Nm/Amp 0.205 0.164  Turns Reference Reference 100% 80% 20% reduction in number  of turns = 20% reduction  in torque constant  Winding Resistance Ohms 40 45Use magnet wire size 1.5  AWG finer than baseline  Percentage Pack for  power windings Reference 100% 57% 20% less turns, and 1.5  gage finer wire  (100•0.8•1.26 -1.5) Motor Constant Nm/ /g1watt 0.0324 0.0244 Kt / /g1R Holding Torque at  18 VDC Nm 0.092 0.066 ~30% reduction in torque  desirable.  Stall Power at  35 VDC Watts 30.6 27.2 ~10% reduction in power  Reducing the power winding pack (slot fill percentage) to 57% of EDU#1, allows us to use up to 43% of  the slot area as shorted coil damping windings. Table 3 compares the shorted coil requirements and  performance to the power winding for EDU#2 and Flight versions.  Table 3.  EDU#2 & Flight Motor Performance Relative to EDU#1  Parameter Units Power  Winding Shorted Coil Comment  Turns Reference  (Percentage of  EDU#1 turns) Reference 80% 80% Use same percentage of turns on  shorted coils as primary power.   Note:  This is not a requirement.  Resistance Ohms 45 63 1.5 wire sizes finer than the  power winding  (45•1.261.5) Motor Constant Nm/ /g1watt 0.0244  0.02066 (Ref) Reference only, used to calculate  damping rate  Shorted Coil  Damping Rate  Nm-sec/rad N/A 4.27 X 10-4 Two sets of damping coils  working together  Percentage Pack  of EDU#1 Slot Fill Reference 57% 40% Providing an overall slot fill of  97% with respect to EDU#1  All of our desired characteristics are met with this scenario. The torque margin reduces from 400% to  214% with the new proposed motor, and the damping rate with the shorted coils is 40% of a fully shorted  redundant winding. If there is a failure of one set of coils, the damping rate will be 20% of a fully shorted winding set, still within our desired requirements, so our new motor is one fault tolerant. Since we have  sufficient of torque margin, we wanted to add as much damping as possible.  37
In order to simulate the “New” motor performance, we conducted a test with our EDU test set. We set the  supply voltage down to 12.6V to simulate the new motor at 18V, our minimum voltage, and we added  shorting resistors across the redundant winding to simulate the magnitude of damping of the “New”  motor’s shorted windings. Figure 7 shows a scope trace of our test set simulation. This plot shows  significant difference between the primary pulse RA signal and the overshoot. Additionally, the  mechanical performance of the system was adequately damped and torque margins were exactly as  predicted from the previous analysis.  Figure 7.  Performance Simulation of the Stepper Motor with Integral Eddy Current Damper  Step Performance Simulation Stepper motors perform like classical underdamped second order systems. It is a straightforward  procedure to analyze the RA output to determine the step kinematics of a mechanism by varying the  Damping Ratio ( /g4) and time factor ( /g1/g3) of classical second order response equations. This allows us to  correlate actual performance to theoretical step performance.  Motor position ( /g2m) of a second order response system may be simulated from the following equation [4]:  /g11/g12/g10 /g8/g9/g7+ • /g1/g2=•/g1 ) sin( 1 /g4/g3/g3/g5/g6 temt t………………………………….(4)  Where:  • /g2mt= Motor Position at time “t” in radians  • E = Step Angle in radians  • /g4 = Damping Ratio (dimensionless 0< /g4<1)  21/g3 /g2 /g1 = ………………………………………………………..(5)  )( cos1/g3 /g2/g1= ……………………………………………………….(6)  38
/g8/g8 /g9/g7 /g5/g5 /g6/g4 /g1=/g1 21 1tan /g3/g3/g2 ……………………………………………….(7)  Equation 4 is a time-normalized equation. You may establish a new time base by varying a “time factor”  (/g1/g3). Simply vary the time base of the simulation to achieve an equivalent crossover time as observed  from the empirical data. You may then calculate the velocity ( /g4) by taking the derivative of position with  respect to time (d /g2/dt), and the acceleration by taking the derivative of velocity with respect to time  (d/g4/dt). Then vary /g5  and/g1/g3in a spreadsheet until your model matches the observed acceleration data  from the RA. From this process, you have detailed kinematical position, velocity and acceleration  information of the mechanism simply by matching the model to the observed acceleration signals.  Referring back to Figure 2, the oscilloscope output of EDU#1 for a single step of the mechanism with load  inertia, in order to simulate the performance, we used the equations above to match the overshoot and  settling characteristics. Figure 8, is our simulation with the damping ratio and time factor noted.  Acceleration vs. Time   Damping Ratio  /g3 =  0.11 - Time Factor /g1/g2 = 0.0030  -70000-50000-30000-1000010000300005000070000 0 0.02 0.04 0.06 0.08 0.1 0.12 time (sec)Acceleration (Rad/sec^2) -0.25-0.2-0.15-0.1-0.0500.050.10.150.20.25 RA Output (VDC) Figure 8.  Simulation of EDU#1 Acceleration Profile  EDU / Flight Modifications The elegance of this design change was that it was transparent to the rest of the system, both mechanical  and electrical. The new motor had the same electrical interface and would be controlled as originally  planned. The new motor had the same mechanical interface, the same envelope and would meet our  torque margin requirements. Additionally, in order to minimize both the cost of and the time to fabricate  the new flight motors, the existing flight motors’ gear-heads were reused. The performance of the motor  supplier, CDA InterCorp, during this re-build effort was extraordinary: delivering two flight motors only  eight weeks after the contract was signed. The total time from discovery of the problem until the new  motors were received was just under five months. Because other parts of the assembly were able to  proceed in parallel, the actual schedule delay to the mechanism was only two months. The program was  able to plan accordingly and there was no impact to the overall system schedule.  Modified Actuator Performance The modified actuators with integral eddy current damping performed better than expected. The  magnitude of the acceleration, measured torque and inherently damped characteristics were as  predicted, or better than predictions. The scope shot in Figure 9 shows the motor phase voltages along  39
with the new mechanism acceleration profile, with the internal damping. Notice how the RA signal shows  a distinct pulse per step with minimal overshoot providing desired monitoring characteristics. Robust  acceleration characteristics, minimal overshoot and consistent performance highlights the improved  mechanism. This scope trace was taken at nominal voltage (+24VDC) and pulse rate. The acceleration  signal is raw and unfiltered. The difference between the overshoot pulse and the first transient pulse is a  robust 400%, making the step counting reliable and consistent. Most importantly, the stepping  characteristic was smooth, quiet and dependable. The erratic motion and audible rattle was no longer  present.  Phase A Voltage Phase B VoltagePhase C Voltage RA Output 50 mV

Figure 9.  Flight Actuator Performance with Internal Damping   Table 4 delineates the difference in performance between the original, EDU#1, and the new, EDU#2 and  Flight, units. From our kinematical analysis, we detail the differences in performance from our initial EDU#1 and “final” configuration. Most notable in the differences in performance, are the reduction in Peak Positional Overshoot, which reduced by 47%, and the Peak Velocity Overshoot, which reduced from -113 rad/sec (-1079 RPM), down to -22 rad/sec (-210 RPM). Since the kinetic energy varies with the square of the velocity, the Overshoot Velocity Kinetic Energy was reduced by an astonishing 96%.  Table 4.  Motor Performance Comparison  Parameter Units EDU#1 EDU#2 and  Flight Comment  Power (Max) Watts 29 25.6 At Maximum Voltage  (34VDC), Lower = Better  Torque Margin % >400% >200% At Minimum Voltage (18VDC) Higher not necessarily better  Peak Positional  Overshoot % 72% 25% At Maximum Voltage  (34VDC), Lower = Better  Peak Velocity at Motor  (during step transient) rad/sec 161 108At Maximum Voltage  (34VDC), Lower = Better  Peak Velocity Overshoot  at Motor rad/Sec -113 -22Kinetic Energy at Velocity  Overshoot Reduced by  96%  40
Reaction Torque at Load  During Peak Acceleration Nm 13.4 9.6 At Maximum Voltage  (34VDC), Lower = Better  Drive Methods and Control Electronics  Method of Three-Phase Control We decided to use three-phase stepper actuators, due to several tangible benefits such as fine step  resolution, simple electronics with fewer power switching Field-Effect Transistors (FETs) and increased  torque capacity under high friction conditions. There are many ways to control stepper motors, but the  most common are wave and bipolar operation. Wave operation (also known as “Line-to-line” or 2/3  phase) and bipolar (also known as two-leads-tied or 3/3 phase) are represented in Figure 10.  B AA+ AC+ CB- B++2 8VDC +2 8VDC +2 8VDCA+ A- B+ B- C+ C1 xx 2 xx   3 xx   4 xx 5 xx 6 xx 7 xx A+ A- B+ B- C+ C1 xx x 2 x x x 3 xx x 4 xx x 5 xx x 6 xx xStepActive FETsWave Ste pping Sequence Bipolar Stepping SequenceStepActive FETs

Figure 10.  Three Phase Control Methods 

For wave operation, two of the three phases are active at one time. In the Stepping Sequence Chart, we  see that for the first step, Phase A is connected to the +28V through the A+ FET, and Phase B is  connected to ground through the B- FET. Under this condition, current is flowing from A through B to  ground. Phase C is open and is not connected through either FET. For the transition from step one to  step two, Phase A remains connected to +28V, but now Phase C is connected to ground through C- FET  and Phase B is open. The terminology of wave comes from this open leg shifting, and is analogous to  two-phase wave operation. Referring back to Figure 7, this is a scope trace of the three phase stepping  sequence of a wave drive with a regulated power supply, notice how the voltage in the open leg oscillates  during the overshoot transient. This is the back-emf circulating and causing the underdamped  characteristics when no damping coils are present.  For bipolar operation, all legs are active during each stepping sequence. For bipolar step one, Phase B is  “high” through the B+ FET, and both A and C are connected to ground through the A- and C- FETS  respectively. At the second pulse, Phase B and C do not change state, but Phase A transitions from  ground to +28V. This controlled and continuous  reversal of current flow is why we call this method bipolar,  and is analogous to two-phase bipolar operation.  Control Method Trades For either control method, during the transition of each pulse, you should never  have the situation where  the high and low side FETs of a common node (e.g. A+ and A-) are simultaneously active. If they were to  41
RA +- - +-+- VbVb Comparator #1 CW Pulse  OutputCCW Pulse Output

Figure 11.  Simple Transition Voltage Counting Circuitry   Figure 12.  Oscilloscope Output of an Adequately Damped Single Step  Conclusion  The first and most obvious conclusion is that EDUs proved invaluable to the process. EDU testing  enabled the discovery of the underdamped condition at a stage in the process when corrective action did  not negatively affect the program. Thanks to cooperation and teamwork between the Mechanisms,  Electrical Engineering, System Engineering and Reliability Teams, as well as a responsive supplier, there  was no impact to the program master schedule, even with a major in-process modification to the actuator.  The most significant lesson learned is that the drive control methodology and power supply regulation  have a much more significant impact on mechanism performance than anticipated. Wave or unipolar  drive methods result in a much more underdamped system, when compared to bipolar drive methods.  Additionally, power supply voltages in test sets must duplicate on-orbit conditions, whether the supply is  regulated or unregulated, because the supply voltage type affects the fundamental step kinematics and  damping.  We discussed detrimental implications of an underdamped stepping mechanism, and the reasons to  require an adequately damped system. We covered many common industry practices and potential  options that could yield desired damping and mechanism characteristics. We took advantage of an  excess of torque margin capacity within the motor, and used shorting coils to provide desired damping  and mechanism performance. Our solution was efficient and robust, making use of existing technologies  and capabilities while maintaining critical schedule requirements.  The utilization of a Rotary Accelerometer (RA) proved extremely useful in characterizing performance and  troubleshooting potential solutions. Determining operational torque margin and conducting the simple  kinematical analysis of the mechanism performance were also integral to the development and solution process. Utilization of an RA in the design increased mechanism reliability, improved position knowledge,  verified step performance, and provided state of health information for this mission critical application.  43