ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
23 Magnetic Chuck Assembly (MCA) Design & Testing 

MCA Overview  Figures 10 and 11 provide an overview of the Magnetic Chuck Assembly (MCA). The main design driver for the MCA was associated with the desire to allow for a passive breakaway of the Core Bit Assembly under a rover slip condition during drilling. The implementation of a passive release also required that a low-profile separation interface be maintained to prevent an off-nominal loading condition in the event of a hardware hangup or snag occuring during interface separation As such, the MCA design utilizes two permanent magnets (one fixed and one with a rotational degree of freedom) to create a magnetic interface that can be turned on (engaged) or turned off (disengaged) while maintaining a very low-profile  separation interface.   

Figure 10.  General overview of MCA within SAT  Upon testing, the MCA was shown to function as intended and the bit could be passively released under  a given separation load of 100 N at the bit tip. In order to utilize the magnetic chuck approach, several compromises were required in terms of material selection. In order to prevent unwanted interactions between the MCA and adjacent mechanisms, several torque coupling interfaces and drive shafts were machined from non-magnetic A286 which could not be hardened as much as desired. In anticipation of the accelerated wear that would occur on these compromised interfaces, additional material was provided to allow for increased torque coupling wear during operation while preserving the tool’s functional lifetime. As expected, accelerated wear was noticed during testing at the specified torque coupling interfaces (see  Figure 12).  MCA Lessons Learned  In order to employ a new and novel approach for bit retention it became apparent that several common  mechanism design issues would need to be considered:   More often than not some type of compromise (e.g. accelerated wear of low hardness materials  at torque interfaces due to material selection contraints) often results that may not present itself  until the hardware is well into its verification and validation testing. As such, recognizing this
24 compromise early in the design processes may allow for a better anticipation of the actual  hardware performance and possible mitigation paths.   This also clearly highlights the need to address hardware durability early in the design lifecycle as  well as providing a means for graceful degradation. 

Figure 11.  General schematic of MCA  Figure 12.  Accelerated wear at torque couplings 

Core Bit Assembly (CBA) Design & Testing 

CBA Overview  Figures 13 and 14 provide an overview of the Core Bit Assembly (CBA). The CBA is used to provide the mechanical constraints for the Core Break-Off Mechanism and to accept the impact energy and the rotational input from the Spindle Percussion Assembly. The main component of the CBA is the custom drill bit which is supported by a duplex bearing pair and is free to translate axially within a sleeved bushing to allow maximum transmission of the applied impact energy to the rock. The drill bit is based on a COTS coring bit tooth configuration with two external helical flutes for cuttings removal.    Due to the choice of implementing a magnetic chuck, a large overall chuck interface was required for the CBA. This in turn resulted in a large bit housing which inadvertently reduced the tool’s capability to accommodate large surface irregularities in the rock as well drilling full-depth holes at off-normal angles (see Figure 15). Furthermore, the large diameter has downstream impacts to the overall IMSAH architecture volume and mass as the SHEC houses the spare bits.
26 During the tool development it was presumed that the amount of rock that needs to be removed as well  as bit tooth wear play a significant factor in the tool’s rate of penetration performance. Since the ID of the bit is based on a desired core size of 9.9 mm (0.39 inch), the amount of material that is removed during drilling (and the corresponding bit face surface area) is driven by the coring bits outer diameter. Furthermore, because the coring bit houses the Core Break-Off mechanism understanding the tool’s sensitivity to bit diameter would be useful for follow on development efforts. Although resources were not  available to fully investigate the sensitivity of the tool to bit diameter, two bounding cases were developed:  1. Min Bit OD – 22.6 mm (0.89 inch)  2. Max Bit OD – 24.4 mm (0.96 inch)  It is clear from the test data shown in Table 1 that there appears to in fact be a significant improvement in  rate of penetration even with a minor reduction in the bit face surface area of only 16%. For the purpose of this investigation relatively unworn bits were used during the rate of penetration investigation in order to  limit the amount of test variables.  Table 1.  Rate of penetration of coring tool utilizing 121g striker at 500 rpm      Avg Drilling Rate [mm/min]  22.6 mm  (0.96") Bit 24.4 mm  (0.89") Bit  Limestone 26.2 59.6  Kaolinite 49.3 162  Siltstone 6.3 27.9  Saddleback  Basalt 2.4 7.9  Volcanic Breccia 1.3 2.7 

CBA Lessons Learned  During the CBA design effort, the modularity of the tool’s sub-elements allowed for a relatively independent development from its mating hardware; however, it also resulted in a failure to check the implications of an increase of the mating chuck interface at the tool’s assembly level resulting in a  significant reduction in capability to drill holes that are off normal from the rock surface.     As such, remember the big picture – resultant interface growth that is satisfied at the mechanism  level, i.e. black box interfaces to black box as intended, might still allow proper mechanism function, but upset higher level functionality  In addition, it was recognized that since this was an early development effort with a low maturity level, a  significant portion of the verification and validation effort should focus on understanding the tool’s  performance and capability sensitivity as it relates to bit diameter.   Therefore, during early development efforts with hardware that has a low maturity level,  significant emphasis should be placed on the desire to understand the hardware’s performance  sensitivity to varying requirements, especially if the results may be used as inputs on subsequent  higher fidelity tools.
27 Core Break-Off (CBO) Mechanism Design & Testing 

Figure 16.  General schematic of CBO  CBO Overview  Figure 17 provides an overview of the Core Break-Off (CBO) Mechanism. The core break-off device is a  pinching/cleaving mechanism that uses two opposing 45o wedges that are symmetrically split about the  cleaving plane. The mechanism is actuated through the CBO drivetrain located within the SPA and is  designed to fracture cores that range in diameter from 9 mm to 10 mm.  As alluded to above, the need to fracture the generated core close to the parent rock required the CBO to  be nested within the ID of the drill bit. As such, during nominal drilling operations the CBO is disengaged from the CBO drivetrain in order to allow for unobstructed operation of the Core Bit Assembly. Once the desired core is generated, the CBO drivetrain interfaces to the CBO mechanism thru a novel axially compliant lead screw that translates axially to interface with the CBO torque interface as well as provide the rotational input to fracture the core. During testing, the CBO was able to fracture all generated cores  as intended.  However, a significant problem did present itself during testing. In addition to providing a means for core  fracture and capture, the CBO also provides a method for retaining the sample tube within the pinch tube by way of flexure retention fingers. During testing, these retention features began to fail at high stress concentration regions located at the base of the flexures. This failure method was initially missed during the design phase as the flexures were only analyzed under a low static load which showed significant margin. However, this load case did not appropriately account for the repeated cyclical loading generated  by the SPA and the resultant failure of the flexure by fatigue.  CBO Lessons Learned  As previously mentioned, due to the vibratory environment generated by the percussive mechanism  several retention features ended up failing to retain the sample tube. Therefore:   In a vibratory environment, stress concentration regions should be analyzed for fatigue failure  even if the resultant cyclical load is small and static results show high margin.
28 Conclusion   In support of the development of the Integrated Mars Sample Acquisition and Handling (IMSAH)  architecture, a low-mass Sample Acquisition Tool (SAT) has been developed. As with most R&D efforts, the project’s schedule and resources were limited so a modular approach to the tool’s development was implemented such that individual subassemblies/mechanisms could be independently developed and tested from the others prior to integration. In addition, these same schedule constraints resulted in early design choices that limited the tool’s flexibility to investigate some anomalous behavior due to several linked mechanisms as well as a premature attempt at performance optimization. As such, during early development efforts with low maturity levels, it became readily apparent that greater emphasis should be placed on the desire to understand the hardware’s performance sensitivity to varying requirements. In doing so, more useful feedback can be provided to allow for better informed decisions on subsequent higher fidelity developments efforts. With that said, the developed SAT has demonstrated the necessary capability to autonomously generate, fracture, and capture rocks cores as necessitated by the proposed  sampling architecture.  Acknowledgements   The research described in this publication was carried out at the Jet Propulsion Laboratory of California Institute of Technology under contract from the National Aeronautics and Space Administration (NASA).  References  [1] R. Mattingly, S. Matousek, and F. Jordan, “Continuing Evolution of Mars Sample Return,”  IEEE  Aerospace Conference, paper #1238, March 2004  [2] R. Mattingly and L. May, “Mars Sample Return as a Campaign”, IEEE Aerospace Conference, paper  #1805, March 2011  [3] P. Backes, J. Aldrich, D. Zarzhitsky, K. Klein, and P. Younse “Demonstration of Autonomous Coring  and Caching for a Mars Sample Return Campaign Concept”, IEEE Aerospace Conference, March 2012  [4] P. Backes, R. Lindemann, C. Collins, and P. Younse, “An Integrated Coring and Caching Concept,”  IEEE Aerospace Conference, paper #1675, March 2010  [5] P. Younse, C. Collins, and P. Backes, “A Sample Handling, Encapsulation, and Containerization  Subsystem Concept for Mars Sample Caching Mission,” International Planetary Probe Workshop (IPPW-7), June 2010  [6] N. Hudson, P. Backes, M. DiCicco, and M. Bajracharya, “Estimation and Control for Autonomous  Coring from a Rover Manipulator,” IEEE Aerospace Conference, March 2010  [7] P. Backes, P. Younse, M. DiCicco, N. Hudson, C. Collins, A. Allwood, R. Paolini, C. Male, J. Ma, A.  Steele, P. Conrad, “Experimental Results of Rover-Based Coring and Caching,” IEEE Aerospace Conference, March 2011  [8] H. Price, K. Cramer, S. Doudrick, W. Lee, J. Matijevic, S. Weinstein, T. Lam-Trong, O. Marsal, and R.  Mitcheltree, “Mars Sample Return Spacecraft Systems Architecture,” IEEE Aerospace Conference, March 2000  [9] A. Okon, “Mars Science Laboratory Drill, “40th Aerospace Mechanisms Symposium, Cocoa Beach  Florida, May 2010
30 For long duration exposure, Dyneema fiber can be used from cryogenic conditions up to a temperature of  70°C. A not very high melting temperature (~150°C) of the fiber makes it possible to cut the Dyneema string using a heating element operated by a small amount of electric power. Preliminary investigations have shown that good results can be achieved using a small resistor as a heating element and a 60° string angle of contact. Tests have shown that Dyneema string shrinks during the warming up process providing better contact to the resistor. The actuation time depends mainly on the local temperature. High reliability and repeatability ratio was observed in the presented actuation method.   Table 1.  Comparison of the Dyneema fiber and steel selected properties   Material Density  kg/m 3 Tensile  strength  GPa Young’s  modulus  GPa Elongation  at break  % Melting  temperature  °C  DYNEEMA®SK65 970 3 95 3.6 149  Steel 316L 8000 0.6 200 45 1370   The general rule in our designs is using Dyneema string only as an actuator element not as a main holder element. For a string of 0.5-mm diameter, the tensile load capability is about 350 N but we reduce the cord’s intensity of stress using leverage to eliminate the necessity of using very tough heating elements. A simple way of putting into use the listed principles is shown on the Figure 1 in MUPUS lock and release mechanism pictorial drawing.    

Figure 1.  MUPUS penetrator head lock & release mechanism principle of operation  

The MUPUS penetrator head, being a part of the counter-mass, has three cylindrical appendages that are placed in three bearings of the structure. The bearings are crossed to provide a zero degrees of freedom grip. When stowed, the penetrator head is held by the clamping of the latch component into its bracket. The proper tension between the cylindrical appendages and bearings is provided by an elastic flat spring and adjusting the gap between beam and structure. Melting of the Dyneema string releases the opening of the latch component and the unlocked bracket allows the motion of the penetrator head in the deployment direction, as it is indicated in Figure 1 by the red arrow.
31 CHOMIK lock & release mechanism description 

CHOMIK L&R mechanism is modified and improved compared to MUPUS’. First of all, it has been designed for different dynamic loads during launch, able to withstand 10g of vibration acceleration and 40g shocks. Such loads result in a relatively high penetrator holding force, which in this case was about 400N. General requirements also influenced the instrument structure that has to contain electronics boards, penetrator guide, and the lock and release unit as a compatible part. The movement direction of the penetrator head after releasing also underwent changes. In MUPUS, the penetrator head has to retreat from the locked elements. In CHOMIK, the penetrator head has to operate inside its holding elements. The instruments with the penetrator head’s movement directions pointed out and their lock & release mechanisms are shown in Figure 2 and Figure 3.   The selection of the melting elements determined the parameters of Dyneema string: 0.5-mm diameter with 10-15 N tension force safe loading. Such a system required 1:40 leverage to withstand the penetrator’s head holding force. Figure 4 illustrates a kinematic diagram of the developed symmetric multi-arms system. The penetrator holding force is transmitted through the crank arms to a slider by the binary links. An angular shaped slider has a sliding joint with a possible displacement of 18 mm. It is blocked by the holders in the stowed position. During release, initially connected by Dyneema string, holders rotate, setting the slider free. The slider’s linear movement cause the crank arms’ rotation. Three pairs of adjustable holding pins separate from the penetrator’s sockets creating 10-mm gaps between the lock and the penetrator.             

Figure 2.  MUPUS penetrator with deployment system (left) and  its lock and release mechanism in the stowed position (right)
32          

Figure 3.  CHOMIK penetrator with electronic box (left) and  its lock and release mechanism in the stowed position (right) – penetrator is removed 

     Figure 4.  A planar kinematic diagram of the lock and release mechanism 

There are three types of springs responsible for the mechanism release: (I) torsion springs generate  torque T to the crank arms rotating them by 18°, (II) flat C-shaped springs guide and produce a pushing  force F1 to a slider, (III) flat springs, which are integrated with holders and execute two functions: kick-off  and Dyneema string tensioning with F2 force.
33  

1 - PENETRATOR        2 – SOCKETS        3 - HOLDING PINS        4 - CRANK ARMS        5 - BINARY LINKS  6 - DYNEEMA STRING        7 - SLIDER        8 - HEATERS (RESISTORS)         9 - HOLDERS        10 - STRUCTURE 

Figure 5.  Practical realization of the mechanism 

The tight string protects the holders after their displacement caused by the Dyneema cord elongation of about 3.5%. The design of the mechanism is shown in Figure 5. The complete lock and release mechanism weighs 150 grams and occupies a volume of 81x82x36 mm. Two redundant melting elements have 110 Ω of resistance and the actuation power voltage is 28V. Two connectors provide confirmation of  the holding arms release.  The main challenge during the mechanism development was the elimination of the self-locking effect. To avoid such a situation, the design process was preceded by detailed kinematic studies, and then a structural-thermal model was assembled and tested. The investigation showed some problems with slider elements that potentially could produce a non-symmetric release. This issue has been solved by blocking one of the kick-off spring’s motions and implemented in the qualification model.   Tests also showed that it is very important to place all moving arms in their suitable positions during the arming process. Figure 6 presents the mechanism’s arming equipment which is helpful to provide symmetric arms’ settings and to generate the preliminary tension making it possible to install the Dyneema string.  The mechanism successfully passed the vibration resistance test, shock test and linear overload test. A functional test has been carried out in the vacuum-thermal chamber. The temperature on the instrument structure was about -150°C and vacuum of about 10 -6 Pa. The Dyneema string application needs 6  seconds for release in those simulating Phobos environmental conditions, compared to less than 1 second at room temperature (laboratory air). Shock generated during release is very low compared to those created by a pyrotechnic actuator.
34    

Figure 6.  Mechanism arming equipment (left) and  L&R mechanism with Dyneema string installed (right) 

High-speed camera record has shown that the maximum speed of the holding pins is about 4 m/s during  releasing (at room temperature). The time period, from Dyneema string melt to the crank arm full rotation, is about 5 milliseconds; then the crank arms rebound several times dissipating energy, and stop after about 52 milliseconds. As it is shown in Figure 7, the first crank arms movement is very equal, which is the result of accurate positioning of the mechanism element during Dyneema string installation using the arming equipment. Figure 8 shows selected moments of the mechanism action. Frame A shows the locked mechanism. In Frame B, the Dyneema string is melted but holders still occupy the slider lateral surface. In Frame C, the slider is shifting resulting in the binary links pull and that makes the crank arms rotation possible. Frame D shows a mechanism in the fully released position. The moments captured on the presented frames has been marked in Figure 7.  

Figure 7.  Holding pins displacement during mechanism release     3436384042444648505254 0 0,005 0,01 0,015 0,02 0,025 0,03 0,035 0,04 0,045 0,05 0,055 0,06 0,065 0,07Time [s]Displacement [mm] QM left crank arm position QM right crank arm positionABD C
35  

Figure 8.  Lock & release mechanism action - penetrator is removed  

Tribological tests 

All rotating and sliding joints in the mechanism (Figure 9) have to work in very low temperatures and  vacuum, with a small movement range and with only one, short time actuation. This influenced the slide bearings selection. Titanium alloy with titanium nitride layer and polyimide-based plastic Vespel SP1 were chosen as mating parts materials. Tribo-components made of Ti6Al4V alloy (shafts, pins, latches) with Vespel SP1 (bushings, sliders, guides) have been used in many SRC PAS mechanisms, including the CHOMIK device. This selection was based on our experience, previous material properties tests, and the ability to manufacture good quality layers on parts with a complicated shape by our partners.
36    Figure 9.  L&R mechanism with its exemplary  plastic-metal slide componentsFigure 10.  Pin-on-Disk tribometer mounted  in the SRC thermal-vacuum chamber

To develop a unit without a self-locking effect, a coefficient of friction for selected mating plastic-metal  parts had to be known. Another unknown was a repeatability of the mechanism behavior during room and low temperature operation. Adequate tests were conducted in a newly built vacuum pin-on-disc tribometer (Figure 10). Ti6Al4V alloy with titanium nitride layer as a pin while Vespel SP1 as a disk were used.  

Figure 11.  Coefficient of friction versus sliding distance for Ti6Al4V pin with TiN layer and Vespel  SP1 disk for vacuum environment: room temper ature (left), very low temperature (right)   The tests were conducted in the following conditions: (1) pressure: 0.01-0.04 Pa, (2) load: 50N, (3) sliding speed: 0.1  m/s, (4) contact area: 10 mm2. Figure 11 (left) shows coefficient of friction versus sliding  distance at room temperature, whereas Figure 11 (right) is a diagram of coefficient of friction for -80°C. Tribological tests in the Phobos-simulated environment showed that the coefficient of friction for Vespel SP1 and Titanium alloy with titanium nitride layer is low, about 0.22. Moreover, friction properties are stable and on the same level, both at room or low temperature.      00,10,20,30,40,5 0 5 10 15 20 25 30 35 Distance [m]Coefficient of Friction 00,10,20,30,40,5 0 5 10 15 20 25 30 35 Distance [m]Coefficient of Friction
37 Conclusions 

Development of the CHOMIK L&R mechanism provided successive arguments that thin Dyneema string melting method can be applied for releasing medium-high loaded space devices. The alike systems are developed at SRC PAS taking into account the superior reliability of the melting process. Medium-higher loads need a high leverage mechanism, which is the main subject of the presented paper. The conducted tests proved it. One of the lessons learned is that in case of the high leverage mechanism, additional springs counteracting the self-locking of the mechanism  must be applied. Usage of a system that  counteracts the cord elongation is also recommended. The investigations that have been made confirm that the Dyneema string melting actuation method is very reliable and Vespel SP1 and titanium alloy with titanium nitride layer are suitable mating materials for low temperature applications. 

References   1. Grygorczuk J., Dobrowolski M., Wisniewski L., Banaszkiewicz M., Ciesielska M., Kedziora B., Seweryn  K., Wawrzaszek R., Wierzchon T., Trzaska M., Ossowski M. “Advance mechanisms and tribological  tests of the hammering sampling device CHOMIK. ” The Proceedings of ESMATS”  (September  2011).  2. Grygorczuk, J., Banaszkiewicz, M., Seweryn, K., Spohn, T. (2007). MUPUS Insertion device for the  Rosetta mission. Journal of Telecommunications and Information Technology (1/2007), pp50-53.  3. Grygorczuk, J., Banaszkiewicz, M., Kargl, G., Kömle, N., Ball, A.J., Seweryn, K. (2009). Use of  hammering to determine cometary nucleus mechanical properties. Penetrometry in the Solar System  II (Eds: Günter Kargl, Norbert I. Kömle, Andrew J. Ball, Ralph Lorenz), Proceedings of the 2nd International Workshop on Penetrometry in the Solar System, Graz, 25th-28th September, 2006, Austrian Academy of Sciences Press, pp93-107.     The paper is supported by the Polish national grant project 791/N-ROSJA/2010/0  and ESA PECS project  No. 98 105