ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
145 produces a radial force on the sphere.  When the Spool reaches its Latched status, the spheres remain  locked between the inside walls of the Panel module and the external surface of the Spool.  

Preload (reference Figure 3)  The Kick -Off springs force (V), applied to the Panel module through the Active sleeve,  is re-introduced  through the spheres vertically, into the Satellite module (V), and radially into the Spool (H). Such action  removes any residual play between the mated parts. Furthermore, during the unlatching sequence, the  Kick-Off force, which is applied all along the demating stroke, provides an additional deployment aid,  intended to insure absence of disengagement uncertainties . 

The preload introduced by the Kick -Off spring (0.5 kN) is considere d adequate for stiff, light  panels. Should  that be insufficient to ensure continuous contact  between Panel and Satellite modules at any point of the  operational envelope, the provision of an optional Ring nut  (6) to be screwed onto the Active sleeve (9) in  order to create further separation force has been provided.    The downward rotation of the Ring nut  (6) translates in  an upward axial motion of the Active sleeve against  the Panel module. With this feature,  an additional preload of up to 5 kN can be achieved.  The control of the  preload is obtained by putting a calibrated preload wave spring (8)  between the Ring nut and the Satellite  module.  

Figure 3.  SHREK section (Left= Latched – Right= Released)   Manual release (reference Figure 3)  On the ground, t he Spool can be manually rotated to the Unlatched position  through a dedicated operating  point accessible from the side of the Mechanism. The Spool rotation ends at a predetermined mechanical  stop,  where the Spool pockets are aligned with the spheres. In that position,  the radial component (H) of  the overall force (V), exchanged between Panel and Satellite, drives  the spheres into  the Satellite module.  With the spheres out of the way , the modul es are dis engaged. The Active sleeve, under the action of the  Kick Off spring, slides upward along the Satellite module, forcing the Panel module away and blocking the  spheres  into their pockets . 

Remote release (reference Figure 3)  A NiTiNOL bar , enveloped by two redundant heaters, is axially integrated into the Satellite module. The bar  is installed in a torsional ly deformed martensitic phase with the mechanism in the latched condition.
146        Figure 4.  SMA bar  with heater   The bar has one extremity firmly mated to the Satellite module, while the opposite end  is connected to the  Spool through the SMA link (3). When heated at a temperature above the austenitic phase transformation,  the bar rotates , recovering its original un- deformed shape. T he bar rotation, transferred to the Spool by the  SMA link, puts  the mechanism into the U nlatched status, thus releasing the Panel module.  

SMA Monolateral  spool control   Figure 5 shows the peculiarity of the monolateral constraint between the SMA bar and the Spool , allowing  the Spool to rotate without being affected or affecting the SMA bar  in the martensitic phase . Conversely,  the rotation of the SMA bar in the austenitic phase induces the rotation of the Spool.  

Figure 5 .  Spool control   The central picture of Figure 5 shows the SHREK in closed status. The picture on the left shows the Spool  manually rotated in Released status with the SMA bar unaffected by the Spool rotation. The picture on the  right shows how the rotation of the SMA bar is transferred to the Spool (remote actuation).  

Reset after remote actuation  Once the SMA bar is cooled below the martensitic phase temperature, and the mechanism is in the  Released condition, the bar can be re -deformed through the operating point located on the top of the
147 Satellite module (pin wrench interface visible in Figure. 5 ). Up to 50 reset operations without performance  degradation are possible.  

Assessment of Results  

Operability  features  evaluation  Self-latching , manual Unlatching, Remote actuation and Reset functions  have matched the expectations.   The additional preloading concept needs to be further improved. The thread pitch is  to be enlarged to make  it less likely to bind. Smooth m anual release requires the Ring nut to  be fully screwed down . The use of a  relatively large wrench is not seen as appropriate in close vicinity of solar panels. The wave spring  effectively allows seeing the applied preload at a glance. Conversely,  that spring was revealed to be the  single most  significant source of shock.    The NiTiNOL  bar proved it had more than adequate behavior. In particular , it was ascertained that the  actuating time is independent of  the preload  applied, a good indication of the adequacy of the motorization  margin. During development,  different alloy compositions and alloy training procedures were tried. It was  realized that the functional  properties, in terms of torque generation for a given torsion strain,  were  inversely  proportional to the transition temperature. This means that rising the austenitic transition temperature reduces the motorization margin. The selected solution has a transition at about 80°C in 55 s econds .  Vacuum tests will probably induce further, hopefully little, adjustments.  

Shock  Response Spectrum  

Figure 6.  Preliminary Shock test results   Within the SHREK mechanism,  there are four  sources of shock:   • The Spool hitting its mechanical stop under the drive of the torsion spring.   • The spheres hitting the bottom of the Spool pockets  driven by the action of the Kick Off spring  • The Active sleeve reaching its mechanical stop driven by  the action of the Kick Off spring.   • The release of the energy stored in the preload spring.   Care has been taken to physically separate as much as possible the events associated with shock  generation.  This has worked well with the first three sources.    From the SRS curves shown in Figure 6, reflecting the Release Shocks in different configurations, it can be seen that the wave spring increases the Shock level, rising it marginally beyond the Ultra- Low Shock  acceleration threshold. Replacing such a spring  with a standard washer brings back the shock level below
148 100 g.  With the removal of the wave spring, a different method to measure the applied preload needs to be  conceived. Integrated strain gauges are under evaluation.  

Motorization   SHREK concept is bas ed on sliding elements, thus friction plays an important role in performance and  motorization factors.  Several element tests were performed to support the selection of the materials, the  treatments, the roughness and geometry of the contacting surfaces as well as the number and size of the  locking bodies.  

The torque required to open the mechanism, with a friction coefficient of 0.15 and a preload of 5 kN,  is  expected to peak at 8 Nm during the initial 5- 8 degrees of Spool rotation. After that,  the mechanis m is  expected to become self -propelling with a null driving torque at about 10 degrees and a positive one after  that. Actual tests performed in air have shown a max torque of 5 Nm peaking at the same expected rotation  point.   The SMA bar is  dimensioned with a torque capability of not less than 12 Nm at the beginning of the shape  recover and not less than 8 Nm after 20 degrees rotation.  Actual motorization factor will be measured in  vacuum conditions.  

Tests   So far, three SHREK prototypes have been manufactured and successfully subjected to Vibration (Sine  and Random), Shock, High Temperature and functional endurance tests in air. T hermal  Vacuum  testing is  planned in the first quarter of 2016.  Opportunities for an in- orbit demonstration are being actively pursued . 

Affordability   At a mass of less th an 0.4 kg, S HREK is composed of  15 parts , obtain ed with standard machinin g  processes , a SMA bar with associated heater, plus standard parts . Its assembly takes roughly 20 minutes  and the us e of two simple  special tools . Its manual operation needs one simple,  to make and use, special  tool. A 30% cost reduction over the average compar ative existing mechanisms  is expected.  

Conclusions    The SHREK mechanism is maturing as a small, simple and reliable mechanism for a specific field of  application (solar panels of small satellites). Its unique features, simplicity and economic affordability should grant its adoption on a large scale.    The outstanding future tasks are:   • performance measurement in vacuum environment  (actuation time, motorization and life )  • alternative method of preload measurement (strain gauges integration)   • obtain flight experience with an in- orbit demonstration .  The first two tasks  are planned to be achieved during f irst half of 2016.  

References  

• Patent Pending “Shrek -  Domanda di brevetto italiano n. 102015000010459”
150 The Design Challenge  

The ‘MTV1’ D-strut® is designed to withstand 623 N ( 140 lbf) axial load, rotate 7o at the main spring and  flexure and stroke 0.99 cm (0.389 inch)  in compression and 0.68 cm (0.267 inch) in tension,  0.057- 0.083  stroke per unit length. More typical designs rotate 3o and stroke 0.025 stroke per unit length. To put this  into p erspective, the MTV D -strut® is  11.9 cm (4.7 inch) in length with a 5.08 cm (2 inch) diameter and  weigh s 860 g (1.9 lbm). More typical D -struts® would be 25.4 cm (10 inches ) in length with a diameter  greater than 7.62 cm ( 3 inches ) and weigh significantly more than 91 0 g (2 lbm). It was found that such a  soft spring in such a small package resulted in a surge mode, surge mode harmonics and lateral modes  at frequencies lower and gains higher than desired. This was detrimental to performance out to 1000Hz.  Figure 2 shows the actual performance of the hardware prior to applying a low outgassing viscoelastic  material ((self -fusing tape (SFT) ) and after the application  of SFT . With SFT, the strut performance met  customer specifications . 

Figure 1.  MTV Isolator Cross Section   Application and Test of a Visco elastic Coating  

Solution development   A machine spring coated with SFT  was used to successfully attenuate undesired modes to acceptable  levels of gain and meet performance requirements , as shown in Figure 2.  Alternative solutions include the  use of a tuned mass damper but this was not viable due to volume constraints and the presence of more  than one mode with different excitation directions and frequencies. A complete redesign was out of scope  and a coating of some sort was the only way to meet cost and schedule constraints. Several materials were tried and tested:  SFT, heat shrink, Sorborthane, RTV  Silicone  embedded between the strut blades,  constrained layer damping, neoprene sleeve,  and a silicone coating. It was paramount that any solution  applied was able to meet performance and stringent outgassing (max vol atile condensable material   content of  0.1% and a total mass loss of less than 1.0% by weight) criteria but should require no design  modification, not create any particulate,  and meet the life requirements of the design. Initial testing  showed SFT to be the most promising. ASTM E 595 testing for outgassing was carried out  at another  Honeywell location ( Clearwater ) and allowed us to test multiple coatings and iterate on bakeout  temperatures and durations in a relatively short time.  With bake out , the s elf-fusing tape met all the  requirements and a method to apply the SFT onto the struts consistently and repeatedly  was devised.  

Application of SFT   The material came in 2.5- cm (1 -inch) wide tape with backing. A lathe was used to apply the tape onto the  main spring with consistent tension,  as the tension affected stiffness, and even overlap. Through trial and  error the optimum lathe seed and feed rates, drag torque for the tape dispenser , and application angle                                                    1 Honeywell Program Name for this effort
151 were established. The setup is shown in Figure 3. After  the application of the SFT, the units were tested  for impedance, to confirm the required stiffness was met. The units were then baked for 40 hours, at   1013.26  Pa (1e-5 Torr ) at 204oC (400oF) and impedance testing was repeated. Impedance testing  confirms the static stiffness Ka, measures dynamic stiffness Ka+Kb2, and allows the damping magnitude  and peak damping frequency to be measured from which the damping coefficient Ca is determined.  Instrumentation measures non- contact displacement at the shaker head vibration input end of the strut  and load cell force at the fixed output end of the strut . Based on a sample of 4 struts wrapped in SFT , it  was found that the standard devi ation of the stiffness (Ka), stiffness ( Kb) and damping coefficient ( Ca)  was 0.62%, 4% and 1.1% of the measured mean respectively and after bakeout these values changed  2.1%, 4.7% and 5% respectively. These values are within the deviation accounted for in Honeywell’s systems analysis for Ka, Kb, Ca.  

Figure 2.  Left: Actual Transmissibility of Hardware (prior to coating in magenta, with a silicone  mold coating in blue and with SFT coating in green). A: Resonant Frequency, B: First lateral  mode, C: First Surge Mode, D: Second Lateral mode, E: Third Lateral Mode, F: Harmonic Surge  Mode . Right: Bipod Transmissibility Performance of Strut with SFT coating and Model Correlation.  

Figure 3.  Left: Tap Test Set up . Right: SFT wrap set up.                                                        2 Where Kb is the in- series spring with the damper
152 Testing over temperature   SFT has an operational temperature of - 51oC to 180oC. This was well within the temperature  requirements, however, thermal cycle testing was carried out to ensure temperature did not affect the  material robustness, in terms of life and performance. The MTV isolators were subjected to two survival  temperature cycles  of -40°C (cold) to 71°C (hot) and eight non- operational temperature cycles  of -24°C  (cold) to 50°C (hot) at cycle rate of 3°C/minute;  both test s carried out at ambient pressure. The strut was  subsequently tested in the bipod transmissibility configuration and it was ascertained that the strut did not  show any signs of damage and that performance did not deteriorate. Thermal impedance testing was  carried out to evaluate the performance of the SFT during the operational temperature range. Figure 5   shows the results of the test. The temp erature shift causes  the Ca to decrease and increase as the  temperature rises and falls, shifting the Phase Shift peak, left and right. The attenuation of the isolator modes exhibits very litt le change. The frequency shifts are as expected with visco elasti c behavior : getting  stiffer with lower temperature and getting softer with higher temperature. The frequency shift is less than 5%, which is within the variation kept during analysis of the machined parts. This was considered acceptable.  

Figure 4.  Top: X-X excitation; X axis results.  Bottom: Y -Y excitation; Y axis results  

Testing for linearity with forcing and for robustness   High-level and low -level bipod transmissibility testing was also carried out, forcing at random non- op  levels and random –op levels. It was found that performance was consistent, such that it was not  necessary to consider non -linearities in the range of interest.  Life testing included 12 full quasi -static  cycles and the full (100- hour) requirem ent for random non- op and thermal cycle testing. Bipod  transmissibility before and the life test was carried to check for any difference in performance. No changes were found and therefore we were confident that SFT was robust to withstand specification environments . The units were subjected to q ualification testing. This included single strut & bipod proof  load, random survival levels, shock levels and thermal cycle testing. Before and after data from the life  test and qualification testing shows that the material had not degraded.
153   Figure 5.  Thermal Impedance Results.   D-Strut Model Development and Model Correlation  

Typical applications have operational  requirements out to a couple hundred hertz (300 Hz typical) and  historically our ability to predict surge modes allows us to design them to be outside the requirement  frequency range.  In this case , the presence of surge mode harmonics and high gain lateral modes meant  the standard approach was not sufficient and more i nvolved testing and modeling correlation was  required. Tap testing was carried out  on the strut  and it was found that the presence of a nested flexure  and a male threaded interface meant the lateral modes were sensitive to flexure orientation.  The tap test   consists of shaking the strut laterally (X & Y direction in Figure 3) with both ends of the struts held and  several accelerometers placed on the isolator. Two rows of accelerometers were used, placed at 90 degrees from one another. This allows for lateral modes in the strut to be identified, including differences  in modes in X and Y. Figure 4 shows the presence  of two lateral modes, with the first lateral mode,  associated with the main spring having  higher gain closest to the spring (accelerometers 2&4), and the  second lateral mode,  associated with the flexure having higher gain closest to the flexure (accelerometers  1&3). The frequency 3 of the first lateral mode does not change with excitation direction and is 256.2 Hz  for accelerometers locations 2 and 4 (cl osest to the main spring) whereas the frequency of the second  lateral mode is sensitive to excitation direction and change from 434.6 to 418.8 Hz for X -X and Y -Y  excitation respectively for accelerometer locations 1 and 3 (closest to the flexure). These results showed  that flexure orientation had an effect of lateral mode placement. Subsequent bipod testing showed that the gain of the second lateral mode was dependent on flexure orientation; when flexures were orientated in the same direction the modes constructively interfered producing maximum gain. Based on tap testing and bipod transmissibility data for different flexure orientations , a detailed Simulink model was created  and tuned to test data; a good agreement between test data and model data is shown in Figure 2. As  mentioned above, it was important to characterize the material used, however, no attempt was made to create a non- linear model of the viscoelastic material as the change in characteristics was within the                                                    3 N.B. the frequencies reported are lower than those found in transmissibility testing due to the mass of  accelerometers. This testing was conducted for mode identification purposes.
154 predicted performance envelope tolerance due to machine part tolerance and bellows stiffness tolerance.  The model, shown in Figure 6,  shows the inclusion of a detailed surge model, which is represented by  12  masses simulating the sprung masses  of the main spring and damping to control the gains of these  modes.  

Figure 6.  Nastran and MATLAB /Simulink Bipod model  

Conclusions  

This paper presents an isolation system comprising of D -Struts ®, designed to meet limited volume  constraints with a 13-Hz 1st mode requirement with no permissible non- linearities in the operating range  of 0-1000Hz , and experiencing large vibration, shock and static loads. It was found that resulting design  was susceptible to surge and lateral modes as well as flexure orientation. This required a more detailed  approach to strut mode identification and strut model development. A  detailed Simulink model was  created and tuned to test data providing good agreement between the test data and model data is shown.  The model shows the inclusion of the surge model, which is simulated by  12 masses representing the  sprung masses  of the main spring and damping to control the gains of these modes . A machine spring  coated with self-fusing tape, a viscoelastic material,  was used to successfully attenuate undesired modes  to acceptable levels of gain and meet performance and life requirements. A big concern with viscoelastic  materials is their ability to meet life requirements and work repeatedly and predictably.  In this paper we  detail the tests carried out to ensure robustness and strut performance repeatability was met. The  presence of such a solution allows lighter and smaller struts to attenuate to high frequencies, such that challenging program requirements such as the one herein described (13- Hz 1st mode without stops, and  high associated loads) can be met.
155 Development of a  High -Performance, Low -Profile Translation Table with Wire  Feedthrough for a Deep Space CubeSat  

Alex Few* 

    Abstract    NEA Scout, a 6U cubesat  and secondary payload on NASA’s Space Launch System Exploration Mission  (EM-1), will use  an 85- m 2 solar sail to travel about 1 astronomical unit (about 1.5 x 108 km) to a near -Earth  asteroid  for observation and reconnaissance1. A combination of reaction wheels, reaction control system,  and a slow rotisserie roll about the solar sail’s normal axis were expected to handle attitude control and  adjust for imperfections in the deployed sail  during the 2.5- year mission . As the design for NEA  Scout  matured, one of the critical design parameters, the offset in the center of mass  (CM)  and center of pressure  (CP) proved to be sub- optimal . After significant mission and control analysis, the CP/CM offset was  accommodated by the addition of  a new subsystem to NEA  Scout . This subsystem, called the Active Mass  Translator (AMT), would reside near the geometric center of NEA  Scout and adjust the CM by moving one  portion of the flight system relative to the other . The AMT was g iven limited desig n space - 17 mm o f the  vehicle’s assembly height -  and was required to generate ± 8 cm by ± 2 cm translation to sub- millimeter  accuracy . Furthermore, the design was required to accommodate a large wire bundle and coax cables fed  through the center of the mechanism . The bend radius, bend resistance, and the exposure to deep space  environment complicated  the AMT design and operation and necessitated a unique design to mitigate ris ks  of wire bundle damage, binding, and cold welding during operation. This paper will outline the design  constraints for the AMT, discuss the methods and reasoning for the design, and identify the lessons learned  through the design, breadboard and test  phases  for the low-profile translation stages with wire feedthrough  capability . 

Figure 1 .  NEA  Scout concept design1 

                                                     * NASA Mars hall Space Flight Center,  Huntsville, AL  

Proceedings of the 43rd Aerospace Mechanisms Symposium, NASA Ames Research Center, May 4-6, 2016