ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
ix• The Mars Climate Sounder In-Flight Positioning Anomaly Bruno Jau & David Kass, Jet Propulsion Laboratory, Pasadena, CA 11:45 LUNCH BREAK  – Box lunch is provided   1:00  SESSION VI—MARS & WAY TO GET THERE   Neil Otte, Session ChairNASA Marshall Space Flight Center, Huntsville, AL • Icy Soil Acquisition Device for the 2007 Phoenix Mars Lander Philip Chu, Jack Wilson & Kiel Davis, Honeybee Robotics Spacecraft Mechanisms Corp., New York, NY; Lori Shiraishi & Kevin Burke, Jet Propulsion Laboratory, Pasadena, CA  • Sample Manipulation System for Sample Analysis at Mars Erik Mumm, Tom Kennedy, Lee Carlson & Dustyn Roberts, Honeybee Robotics Spacecraft Mechanisms Corp., New York, NY • A System for Suspending and Vibration-Isolating a Large Spacecraft for Testing in Vacuum David Keinholz, CSA Engineering Inc., Mountain View, CA • Gas Strut Separation Alternative for Ares I Brian Floyd & James Owens, NASA Marshall Space Flight Center, Huntsville, AL • Evaluation of Separation Mechanism Design for the Orion/Ares Launch Vehicle Kevin Konno, Daniel Catalano & Thomas Krivanek, NASA Glenn Research Center, Cleveland, OH 3:45  BREAK 4:00  SPECIAL PRESENTATION  – MSFC Launch Vehicles, History, Lessons Learned,   Projections to ARES Bob Ryan, NASA Marshall Space Flight Center (retired)  The presentation will cover the history of MSFC launch vehicles including Redstone,   Saturn, Space Shuttle, and NLS/NGLT. Issues discussed will focus on technical integration,  uncertainties and margins, risks and trades. A discussion of the Ares mission conﬁgurations  will lead to projections of possible issues on the ARES I and V conﬁgurations. 5:30– 10:30 SYMPOSIUM BANQUET AT U.S. SPACE & ROCKET CENTER   5:30  Bus leaves Embassy Suites to Rocket Center Spacedome IMAX Theater – “Magniﬁcent Desolation”Dinner under the Saturn VTour the museum 10:00  Bus leaves Rocket Center for hotel
xFRIDAY, 9 MAY 2008 7:30  Friday Presenters’ Breakfast – Embassy Suites Hotel Redstone Boardroom8:00  Light Refreshments – V on Braun Center South Hall Meeting Room8:30  SESSION VII—POSITIONERS  – V on Braun Center South Hall Meeting Room Scott Tibbitts, Session ChairSpaceDev Inc., Louisville, CO • Focus Mechanism for Kepler Mission Kraig Koski, Ball Aerospace & Technologies Corp., Boulder, CO • Precision Linear Actuator for Space Interferometry Mission (SIM) Siderostat Pointing Brant Cook, David Braun, Steve Hankins, John Koenig & Don Moore, Jet Propulsion Laboratory, Pasadena, CA • Development of a Low-Cost Fine Steering Mirror Steve Wassom & Morgan Davidson, Space Dynamics Laboratory/Utah State University Research Foundation, North Logan, UT 10:00  QUICK MID-SESSION BREAK • Cryogenic Focus Mechanism for the Spitzer Space Telescope William Schade, Ball Aerospace & Technologies Corp., Boulder, CO • Development of a Spacecraft Antenna Pointing Gimbal Chuck Monroe & Peter Rossoni, NASA Goddard Space Flight Center, Greenbelt, MD 11:15  SPECIAL PRESENTATION : An Overview of Orion CEV Mechanisms  Lance Lininger, Orion Service Module Mechanism Design Lead, will give a preview  of all the mechanisms on the Crew Exploration Vehicle. May see these again as the topic  of a paper in the future!  11:45  TECHNICAL SESSIONS CONCLUSION   • Herzl Award Presentation • Closing Remarks 12:00 LUNCH BREAK  – Box lunch is provided   1:30  BUSES DEPART HOTEL FOR TOUR 2:00– 4:00 FACILITY TOUR at ULA or MSFC
xiSYMPOSIUM ORGANIZING COMMITTEE   Donald H. McQueen, Jr., Host Chair, NASA MSFC Stuart H. Loewenthal, General Chairman, Lockheed Martin   Edward A. Boesiger, Operations Chairman, Lockheed Martin Carlton L. Foster, NASA MSFC Claef F. Hakun, NASA GSFC Christopher P. Hansen, NASA JSC Wayne Jermstad, NASA JSC Gérard Migliorero, ESA/ESTeC   Alan C. Littleﬁeld, NASA KSC Edward C. Litty, JPL Fred G. Martwick, NASA ARC Wilfredo Morales, NASA GRC Robert P. Mueller, NASA KSC Fred B. Oswald, NASA GRC Minh Phan, NASA GSFC Donald R. Sevilla, JPL Mark F. Turner, NASA ARC Robin Tutterow, NASA LaRC James E. Wells, NASA LaRC SYMPOSIUM ADVISORY COMMITTEE Obie H. Bradley, Jr., NASA LaRC (retired) Ronald E. Mancini, NASA ARC (retired) Stewart C. Meyers, NASA GSFC (retired) William C. Schneider, NASA JSC (retired)
Linear High-Force “Step And Repeat” Piezoelectric Motors  J.S.N. Paine*, M.E. Johns*, J.J. Sesler*, M.T. Stefanick* and J.A. Kennedy* Abstract  Dynamic Structures and Materials, LLC (DSM) has designed a patent-pending piezoelectric linear motor  for demonstration in two aerospace applications related to the operation of a cryogenic 2” isolation valve and an environmental controls air handling system for next-generation space vehicles. The scaleable  actuator technology provides a combination of force, stroke, and speed not previously demonstrated in  other piezoelectric motor developments. The IMPULSE PiezoMotor™ technology is presented as a viable  and superior replacement for heritage pneumatic, hydraulic, and electromagnetic actuation devices. A key  feature of the IMPULSE PiezoMotor™ technology resulting fr om its friction-based drive architecture is a  power-off-lock characteristic (fails in last position). A description of the mechanism and key performance  parameters (step size, drive force, and resolution/accuracy), influence of materials on friction/wear, and  power requirements are discussed.   Introduction to Piezoelectric Motors  A development effort to produce linear proportional electromech anical actuators for valve control has  resulted in a new class of linear actuation mechanisms based on piezoelectric prime movers. The piezoelectric driven actuators are based on a step and repeat motor topology that enables relatively large  strokes compared to typical piezoelectric driven actuators. This paper describes the mechanism and how  it operates as well as discusses de sign and performance issu es relative to using this device in precision  motion control applications.   Linear “Step and Repeat” piezoelectric motors or SRMs are electro-mechanical energy transducers that  convert electrical energy to mechanical motion through the expans ion of piezoelectric ceramic materials.  Piezoelectricity is the ability of certain crystals to generate a voltage in response to applied mechanical  stress. The piezoelectric effect is reversible in that piezoelectric crystals formed into thin layers and  stacked with electrodes can expand a small amount when subjected to an externally applied voltage.  Typical expansion of these “stacks” can be on the order of 0.1 to 0.2% of length (i.e. a 20 mm long sample will extend ~ 20 to 40 microns (0.001 to 0.0016”) at an applied voltage of 100 to 200 VDC. Figure  1 illustrates the expansion for a simple prismatic piezoelectric ceramic stack actuator. The expansion can  occur at very fast response times (20+ kHz) and relative high force (50 MPa). This high force can be harnessed to produce very high ba ndwidth actuators and motors.   Figure 1 – Exaggerated Expansion Of A Piezoelectric Stack In The Column Direction  * Dynamic Structures and Materials, LLC, Franklin, TN  Proceedings of the 39th Aerospace Mechanisms Symposium, NASA Marshall Space Flight Center, May 7-9, 2008  1
Direction of  motion Direction of  motion Figure 4 - Prototype Deliverable Of DSM’s 2000 N IMPULSE Piezomotor™ (IMPULSE 2000) At Top  And DSM’s IMPULSE 30 At Bottom  A third factor in motor drive force is the friction coefficient of interfaces between the ends of the clamps  and the clamper plate frame. DSM has found that the drive force is linearly proportional to the friction  coefficient of the two materials in contact. Therefore, increasing the coefficient of friction between the two  materials leads to an increase in drive force. DSM has performed a detailed experimental study of interface materials friction coefficients. A summary of this study is included in the next section.   Figure 5 shows the motor step size versus drive force for one of the IMPULSE 2000 prototypes and for  two different IMPULSE 1000 prototypes that have been produced for NASA applications. The results were  generated by driving a friction load with the motors while recording step size. Peak drive force scales  approximately with the size of the clamp stacks used in the motors and with friction coefficient differences in the two systems. In this graph, the IMPULSE 2000 has mo re than twice the projected blocked force of  the IMPULSE 1000 actuators. The IMPULSE 2000 has twi ce the piezoelectric clamp stack size of the  IMPULSE 1000 and a coefficient of friction that is approximately 50% greater. While it is difficult to  attribute specific quantitative differences to clamp size difference versus coefficient of friction, DSM has  found that the trends are consistent. When a motor has twice the clamp stack size of a previous version,  the available drive force can be double. When friction coefficient increases, the drive force follows.   Step Size and Drive Speed  The curve of the IMPULSE 2000 step size (Figure 5) shows a no load step of 27 microns and a near  linear drop to 8-micron step size at a friction load of 2000 N. For all three prototypes, the step size  consistently decreases with an increase in friction force. The difference in the step size of the two  4
IMPULSE 1000 motors reflects the variability in the friction between these early prototypes and the  importance of achieving consistent fit and finish in the motors at the micron tolerance level.   Generally, the motor step size is dictated by the length of the piezoelectric stack used in the extender and  the external load applied to the extender. Since the piezoelectric stacks are effectively spring elements,  when the extender stack expands against a load, the compressive reaction force effectively reduces the  stack’s step size. The nature of the load also affects step size. A weight directly applied to the end of the extender will cause the extender to compress and slip backward during clamper transitions when moving  against the weight. A friction load will stay in place and simply diminish the extender stroke as a function  of the friction magnitude. A spring-type load will compress the extender to reach force equilibrium during each step and will release the stored energy backward during each clamper transition. In this manner, the  step size becomes dependent not only on the magnitude of the external load, but also the direction and  type of load.  051015202530 0 200 400 600 800 1000 1200 1400 1600 1800 2000 2200 Friction Load (N)Step Size (micron)Impulse 1000 prototype 1 Impulse 2000 prototype 1 Impulse 1000 prototype 2 Figure 5 - Step Size Vs. Drive Force (Friction Load) For The IMPULSE 2000 And 1000 Prototypes  For example, the ideal step size and measured step size for DSM’s IMPULSE 30 is shown in Figure 6.  The unloaded step size of the extender using a 3.5 x 3.5 x 18 mm3 long piezoelectric stack is 23 microns.  The stiffness of this extender element is approximately 29 N/micron. Pushing a static load (weight) with  this extender reduces ideal step size by the slope of the curve (i.e. 29 N static load reduces the stroke by 1 micron). This ideal step size ignores the compliance of other frame components and the non-ideal  nature of the friction force that permits the motor to function. The compliance of the motor frame is a  linear value that can be placed in series with the extender compliance. The non-linear friction effects are more difficult to predict but follow a recognizable trend. At this point in the SRM development, DSM has  structural techniques that can be applied to reduce the motor frame compliance but has not completed a  predictive model. Instead, each motor is designed with a balance of mass and stiffness to achieve the application objective while recognizing that the force versus step size response may be far from ideal.   5
0510152025 0 5 10 15 20 25 30 35 Static Load (N)Step Size (micron)Measured/g3step/g3size Ideal/g3step/g3size Figure 6 - Step Size Vs. Drive Force (Static Load) For The IMPULSE 30 Prototype  Drive speed is the product of the motor step frequency (number of steps per second) and the extender  step size. The motor step frequency is the rate at which the electronics driver sequences the piezoelectric extender and clamper elements through the entire motor step cycle. Typical quasistatic clamping motors  require that all three piezoelectric elements (extender and two clamps) oscillate through a complete  square or triangle wave period without overlap in the rise or fall segments. DSM uses a special motor electronics driver that can produce six rise and fall segments fast enough to drive a 6-segment full step in  less than 125 microseconds or step rate of 8000 Hz.   In practice, DSM has been able to operate motors up to 80 mm/s. The limitation results from the natural  frequency of internal motor components (clamps and extender elements) being too low to respond to the  highest possible step frequency. DSM can run the IMPULSE 30 and IMPULSE 2000 at approximately 2600 and 2000 steps per second respectively. The limitation in the IMPULSE 30 is the natural frequency  of the extender element moving the lumped mass of the extender frame and active clamp elements. The  limitation of the IMPULSE 2000 stems from the limitation in maximum piezoelectric stack preload that currently can be applied to the extender piezoelectric elements.   The fact that the piezoelectric actuator is controlled by a sequence of driving waveforms means that the  motion profile can be optimized for different step frequencies within portions of the overall motion profile.  This characteristic is valuable for proportional control applications.   Characteristic Impulse Force Profile An interesting aspect of the IMPULSE PiezoMotor™ technology is the profile of the force that results  during the step action. The extensio n step of the IMPULSE PiezoMotor™ may produce significant impact  force on undamped loads. To quantify this effect, DSM tested three coupling conditions at a drive step rate of 800 motor steps per second (800 Hz). The resulting force profile measured by a high-stiffness load  cell in-line with the motor extender and ground is shown in Figure 7.   At the point in the extender waveform where the extender ad vances the load cell in to a solid aluminum  block a very sharp resonance peak of 2000 N peak to peak occurs. The RMS average force is  approximately 2200 N. This demonstrates the rapid hammering effect that the piezoelectric prime mover can produce against very stiff elastic loads. Placing layers of rubber betwe en the extender and the  aluminum block softens the hammering effect as shown by the blue and red curves. The impulse peaks  for the 1/8 th inch rubber (blue) are 1100 N peak to peak and approximately 1750 N RMS average. The  impulse peaks for the 1/4th inch rubber (red curve) are 600 N peak-to-peak and approximately 1400 N  RMS average.  6
10001500200025003000350040004500 0 0.001 0.002 0.003 0.004 0.005 Time (s)Force (N)Running against Aluminum 1/8" rubber layer in line 1/4" rubber layer in line Figure 7 – IMPULSE 2000 Peak Output Force Profile With Three Different Coupling Configurations  To Ground As Measured By An “In-Line” Load Cell  It is interesting to note that twice the thickness of the rubber (red curve) appears to reduce the impulse  peaks to ½ the value of the blue curve. The drop in  impulse peak-to-peak value s appears to follow the  thickness of the rubber layers. In this experiment, black neoprene rubber was used for the damping  barrier but other damping and spring materials have been tested with similar results.   Characterizing the Friction Interface Materials   Because the IMPULSE PiezoMotor™  architecture is a friction-based drive mechanism, DSM has  performed extensive experimental work with various material sets to understand the static and dynamic  friction response. DSM has also investigated how wear affects surface roughness properties for the  friction interface materials. For the IMPULSE PiezoMotor™ technology to operate properly, the friction  interface materials must be very flat and smooth or have a very low surface roughness. Since the active  clamp elements have stroke le vels on the order of 10 to 20 mi crons, surface roughness values that are  greater than 1 micron (40 microinch) can be significant. Additionally, the design goal is to achieve high  friction at the interface materials in order to produ ce a high level of motor drive force. DSM has also  determined that to maintain a rated load capacity, the surface roughness, flatness and parallelism of the  friction pair must be maintained to an acceptable level throughout the useful life of the actuator. Therefore, minimizing wear is also a design goal.  Working with University of South Florida (Mudhivarthi, 2006) and Brigham Young University, DSM has  performed a number of friction materials characterization studies for piezomotor applications. The testing  consists of a series of physical tests on numerous friction pairs that was designed to simulate the loads,  motions and conditions typical to the various motor applications. DSM has tested friction pairs including various combinations of steels, ceramics and surface treated steels. High-cycle wear tests were  performed and optimum friction pairs for this application were identified based on high friction, low wear  and low roughness. Much of this data is repeated from the work that DSM funded at University of South Florida (USF) that is recorded in the Mudhivarthi (2006) paper.  7
Figure 8 - Friction Versus Time Plot For A Solid Croblox /g147 Against A Solid 52100 Steel Specimen   Table 1A - High Cycle Tests Summary For Coated 52100 Steel Friction Interfaces   (repeated from Mudhivarthi, 2006)  No. Material Pair COF Post Test Condition  1 TiN / TiN 0.18-0.22 Coating removal/severe abrasion after 600  cycles; Rtm=41-364 /g80in 2 MoST / MoST 0.10-0.32 Severe coating removal after 100 cycles;  Rtm=45-177 /g80in 3 WC / ceramic gage block 0.09 Cancelled due to low friction  4 WC / 52100 steel gage block 0.18-0.21 Abrasion after 20,000 cycles; Rtm=33-65 /g80in 5 WC / Croblox /g147 gage block 0.11 Cancelled due to low friction  6 DLC / ceramic gage block 0.09-0.11 Cancelled due to low friction  7 DLC / 52100 steel gage block 0.13-0.15 Abrasion after 40,000 cycles; Rtm=14-96 /g80in 8 Tetrabond / ceramic gage block 0.17-0.21 Abrasion after 20,000 cycles; Rtm=79-94 /g80in 9 Tetrabond / 52100 steel gage  block0.32-0.36 Severe abrasion after 10,000 cycles;   Rtm=79-156 /g80in 10 TiN / A2 steel 0.16-0.24 Severe abrasion after 10,000 cycles;   Rtm=280-1,257 /g80in The solid material friction interface pairs (Table 1B) performed very well in terms of wear and surface  roughness. The steel feeler gages cycled against each other were the only pair that manifested severe  abrasion after a relatively low number of cycles. Co mpared to the coated spe cimens, the solid friction  interface materials had much lower abrasion and lasted to levels of 300,000 cycles in the case of ceramic on ceramic. When comparing the response of the materials, DSM found that the following pairs had  reasonable coefficient of friction levels in the range of 0.12 to 0.18 and indicated very mild abrasion:    9
/g120 Sapphire on 52100 steel gage block (test no. 14) at 0.13 to 0.15 COF,  /g120 Sapphire on ceramic gage block (test no. 15) at 0.15 to 0.17 COF,  /g120 52100 steel gage block on ceramic gage block (test nos. 12 and 17) at 0.12 to 0.17 COF, and  /g120 Ceramic gage block on ceramic gage block (test no. 18) at 0.13 to 0.15 COF.  For environments where resistance to corrosion is an issue, the two different ceramic pairs would perform  better than either of the two pairs with steel. Fine pa rticles generated during te sting of the 52100 steel  gage block on ceramic gage block we re found to be oxidized and  some of the fine abrasive scratches in  the steel also showed oxidation.  Table 1B - High Cycle Tests Summary For Solid Material Friction Interfaces  (repeated from Mudhivarthi,  2006 with updates*)  No. Material Pair COF Post Test Condition  11 Steel feeler gage / steel feeler  gage0.30-0.60 Severe abrasion after 600 cycles;   Rtm= 1.4 – 3.1 /g80m (57-123 /g80in) 12 52100 gage block / ceramic gage  block0.14-0.17 Mild abrasion after 200,000 cycles; Rtm= 0.1 –  1.2/g80m (4 - 47/g80in) 13 Croblox/g147 gage block / ceramic  gage block 0.13-0.2* Cancelled due to low friction  14 Sapphire / 52100 steel gage  block0.13-0.15 Mild abrasion after 40,000 cycles; Rtm= 0.1 –  0.2/g80m (5 - 10 /g80in) 15 Sapphire / ceramic gage block 0.15-0.17 Mild abrasion after 40,000 cycles; Rtm= 0.08 –  0.2/g80m (3 - 7 /g80in) 16 Croblox/g147 gage block / 52100  steel gage block 0.15-0.2* Abrasion after 10,000 cycles; Rtm= 0.08 – 1.1  /g80m (3 – 46 /g80in) 17 52100 steel gage block / ceramic  gage block 0.14-0.2 Mild abrasion after 200,000 cycles; Rtm= 0.08 –  1.4/g80m (3 – 55 /g80in) 18 Ceramic gage block / ceramic  gage block 0.13-0.15 Mild abrasion after 300,000 cycles; Rtm= 0.08 –  0.1/g80m (3 – 5 /g80in) Example of Wear at the Friction Interface  To characterize the material wear of friction surfaces within a prototype piezoelectric motor environment,  DSM assembled an extender element using two 52100 steel gage blocks separated by a piezoelectric  extender stack. The extender was actively driven within an early IMPULSE PiezoMotor™ prototype  frame. The mating friction surfaces contacting the extende r’s steel gage blocks we re fabricated from a  hard ceramic. Following a cumulative travel distance of 10 km, surface profile scans were made on the  extender’s blocks in order to quantify material loss. The mating ceramic surfaces were not profiled. The  surface scans were performed along the direction of the lines indicated in light blue in Figure 9.  Figure 9 - Sketch Of Extender Assembly Used In Long Endurance Friction Wear Testing  Figure 10 illustrates the surface height measured with a profilometer for each of ten scans. Note the  original flatness of the block on the outside edges. Over the course of ten kilometers, as much as ~8  microns of depth in the steel was worn away in certain locations, but the average material loss appears  closer to 4-5 microns of depth. One attr active feature of DSM’s IMPULSE PiezoMotor™ patent-pending  10
architecture is its ability to compensate for this material loss, thus avoiding a decrease in the motor’s  available driving force.  Figure 10 - Surface Heights For Each Of Ten Scans Across The 52100 Steel Extender Components  After 10 km Of Cumulative Linear Travel In A Prototype Piezoelectric Motor  Conclusions from the Study of Friction Interface Materials A number of coated and solid friction interface materials have been studied and tested. The solid friction  materials held up substantially better than the coated friction materials. Three types of materials held up notably well: a ceramic gage block; a hardened 52100 bearing steel; and, a sapphire substrate. All  indicated very little abrasive wear in a friction test at over 40 ,000 wear cycles. Of pa rticular interest is the  nature of the coefficient of friction in the hardened specimens. Regardless of the initial COF, all of the  materials migrated to a similar range of 0.11 to 0.18 COF after sufficient wear cycles had occurred.   Accuracy and Resolution of the SRM Mechanism   As described previously, the step size of DSM’s SRM mechanism is dependent on the size of the load  applied to the motor output. The accuracy of the step size is therefore dependent on the load magnitude  and the consistency of the load. Generally, DSM recommends  that these mechanisms be used with an  encoder or other type of position feedback device to provide a means of updating the position to the  controller and closing the position loop on a specific target.   In a recent example of using an IMPULSE 1000 to drive a NASA valve, DSM closed the loop on a  position target using an LVDT device for feedback. An LVDT produces a linear variable voltage as a  function of position over the travel length. In this application, the SRM was used to drive the valve to a set location (voltage). The controller is programmed to stop th e motor at the nearest step position within a  designated deadband around th e voltage position target. In this case, th e resolution of the position control  is approximately one motor step or less than 25 microns.   11