ai-aerospace/ams_data_full_2000-2020 huggingface dataset:
A two-dimensional section representative of a Bensen B8-M autogiro rotorblade was constructed and  installed in the tunnel, mounted on top of a 6-component strain gauge balance. Data from comparison of  the baseline section with a full-span tab system shows an increase in a few “counts” of drag on the section  (Figure 5). Later testing that included mock-ups of the supporting SMA wires and associated electronics  showed these added effects to be minimal, and thus the original data as shown here was used in  generating pre-test predictions for full-size rotor tests to follow.  0.07 +  -  0  +  w/tab  wlotab + 0.06  0.05  0.04  0.03 -  0  - +  0  0 + + -  0 +O  +O +O - +O  +O+a* +o +O  Typical tracking tab locations are near the 75% radius location, so for a typical helicopter tip speed  between 200 m/s and 213 m/s, the dynamic pressure at the tab location is approximately 40 N/m2.  Duplicating this dynamic pressure in a wind tunnel or other aerodynamic facility is difficult and expensive,  so an alternative testing method was employed. Since the density of water is approximately three orders  of magnitude greater than air, equivalent dynamic pressure may be achieved at speeds of approximately  7.7 m/s. Thus, dynamic pressure tests on a tab-equipped UH-60 helicopter blade airfoil section were  performed using a special “oar“ with a representative airfoil and two-position tab attached to the end, with  the oar placed into the water aside a small power boat. The airfoil was mounted on a long pole to permit  its being placed sufficiently far away from the side of the boat hull to eliminate any surface effects from the  boat displacement when traveling under power. Instrumentation consisted of a digital inclinometer for  measuring approximate angle of attack of the foil, and a handheld GPS unit for measuring boat hull  speed. Runs were taken in both upstream and downstream directions on the Delaware River, and despite  extremes of both dynamic pressure and high angles of incidence (up to 0.3 radian), the two-position tab  did not “pop” through to its alternate position from differential pressure on the tab.  Another operational concern for proper tab actuation was the anticipated high centrifugal environment  anticipated for tabs mounted at the 75% radial location. For UH-60 size rotors, this loading can approach  550 g‘s, and thus the tab may “pop through” from inertial loads, or the SMA wire may deform on the tab  under its own centrifugal load, effectively “using up” the available strain energy and rendering the tab  inoperative. The latter was surmised to be the case for early versions of SMA-based flap actuators,  developed by CDI for application to the V-22 tiltrotor blades on a model rotor set; ultimately, the wire-  based actuation mechanism had to provide mechanical guides for the chordwise-oriented actuator wires  to eliminate the centrifugal distortion. Although the prototype tab actuator developed here has wires  377
oriented chordwise, supporting calculations showed that they would not be subject to large strains due to  their minimal size and weight.  Testing was nonetheless performed on the tabs using a four-foot diameter model rotor, with the operating  rpm boosted to 550 in order to expose the tabs to representative centrifugal loads. In these tests, the tabs  were mounted at the blade tip to provide maximal centrifugal stresses for the wires. Testing revealed no  problems with the tabs reversing their positions at these g-levels, thus providing confidence for ultimately  testing the units on a full-size rotor system.  Full Size Rotor Testinq  The full-size rotor system used for evaluating the prototype actuation system was a Bensen B-8M autogiro  ("gyrocopter"), available for use at CDI (Figure 6). The advantage of using the autogiro test platform is the  ease of access to a man-rated rotor system, with all of the system integration issues associated with a full-  size flight vehicle. Since that aircraft is an autogiro, and not a helicopter, an alternate means of spinning  the rotor was required so to avoid having to fly the aircraft with the tab devices installed. This drive system  consisted of a pre-rotator assembly that used a flexible shaft that could be driven by a pulley mounted  inboard of the propeller hub on the autogiro's engine.  I  Figure 6. Autogir, .... .. ".. ... .....e r,u.e... uue.. ... .u.m-uize rotor tests.  Test objectives were twofold for the full-size rotor experiments: first, it was desired to demonstrate tab  effectiveness in controlling tracking angle of the rotor blades from positioning of individual tabs on the  rotor; and second, it was desired to assess the rpm range over which reliable electrical actuation of tab  displacement could be achieved.  Tabs were installed onto the Bensen blade set from the 0.71 to the 0.80 radius location, with each tab  assembly fabricated using an aluminum base plate holding four 2.5-cm by 5-cm tabs. Three tab plates  were mounted on each blade, with a Lexan separator plate bonded between the blade surface and the  mounting plate in order to facilitate removal of the tab system from the blade at the conclusion of the test  program (the removal agent for the tab system is known to attack certain plastics). To simplify the test  program, only one set of four tabs was fabricated using SMA wires, since the purpose of the test was to  demonstrate performance and not provide a complete system installation. As a result, the control  electronics were mounted at the rotor hub, and separate flat ribbon wires were attached to the bottom of  the blade to provide electrical power to each SMA tab. While a true prototype installation would include  local surface-mounted MOSFETs for control of individual SMA wire excitation at each tab plate, this  simplification helped expedite the completion of the full-size rotor test program. Details of the tab mounting  arrangement may be seen in Figure 7, and the hub electronics in Figure 8.  378
Figure 7. SMA-driven bi-modal tabs on underside trailing edge of Bensen B8-M blade section.  LU iftj Rotor tachometer  Figure 8. Details of hub on Bensen autogiro test platform.  In order to check blade track angle, a separate video camera installation was mounted in the rotor plane,  with a calibration card constructed and the blade ends colored (white/black) to differentiate them on the  videotape. The videotape was then viewed on a frame-by-frame basis to measure the blade tip position as  a function of tab deflection angle. For the extreme case of all tabs on the “black” blade deflected down,  and all tabs on the “white” blade deflected up, a tip deflection difference of 3.9 cm was measured from  scaling displacements from the video monitor. This compares favorably with a predicted value of 4.5 cm,  since the predicted displacement does not account for any lift rolloff at the most inboard and outboard  locations of the deflected tab assemblies. Screen capture of the rotor plane video may be seen in Figure  9.  379
Figure 9: Still video of white and black blade position from rotor plane camera.  Post-test analysis and further development testing revealed that the Lexan plastic material used to form  the base bender element eventually took a “set” due to the continued exposure to bending stresses, and  thus developed a preferential orientation that would make the tabs less robust. As a result, further  development produced a tab with equivalent stiff ness but no tendency to plastically creep, due to the use  of spring steel for its bender element. Shown in Figure 10 is a demonstration article that incorporates a  steel base with Kapton tape overcoat and flush attachments of the SMA wire assemblies, along with the  flatpack 6 volt battery used for its actuation.  Figure 10. Metal-based tab demonstration device with flatpack battery.  Distributed UAV Flight Control Application  High-altitude, long-endurance (HALE) remotely-operated aircraft (ROA), or unmanned aerial vehicles  (UAVs), are designed to operate at extreme altitudes for many days, in order to serve as platforms for  science missions, earth observation, or communications and data relay stations. In order to meet the long  duration requirements for these missions, these aircraft are designed to minimize drag and weight, and  thus are characterized by having large wing spans and significant aeroelastic response under both steady  and dynamic loading conditions. This inherent flexibility in these aircraft may sometimes lead to flight  dynamics issues and instabilities, as witnessed with the breakup of the HELIOS aircraft over Hawaii during  flight investigations while operating with alternate energy sources.  These aeroelastic issues may be mitigated to a considerable extent for these vehicles if the loading on the  large aspect ratio wings may be redistributed in flight, using distributed multiple actuators. Such an  approach provides enhanced robustness to the aircraft platform, in that multiple actuators would easily  accommodate the failure of a few to respond to commands, but if these actuators were conventional,  servo-type systems, this feature would be overshadowed by an unacceptable weight and power penalty.  380
join together  (b) assembled, % view (c) popped-through, 34 view join E together  (a) stamping  Figure 14a-c. Bi-stable actuator formed from a flat plate, incorporating three parallel strips.  Bender  - _.  d) demonstration device implementation and actuation response  Figure 14d. Bi-stable actuator formed from a flat plate, incorporating three parallel strips.  Partial views of  additional, co-planar  3-position actuators  Figure 15. Co-planar three-position actuators for compound geometry control.  Force applied from  SMA wires (not shown) snapped "up"  snapped "down"  Figure 16a. Three-position actuator showing base support bending from snap-through action.  385
Figure 16b. Three-position actuator in series arrangement for large displacement application.  Conclusions  An actuation system incorporating SMA wires to reorient a buckled member has been described, with a  focus on its use as a trailing edge control device for aerodynamic modifications. Applications to helicopter  rotor blade in-flight tracking, and UAV spanwise aerodynamic load control have shown that the small form  factor, minimal power requirements, and simplicity of design make it an attractive alternative to  conventional aerodynamic control effectors. Future use in other aerospace applications will benefit from  the attractive weight and size offered in such a device.  References  Asadurian, A., “Miniature Linear Actuator,” Proc. 34th Aerospace Mechanisms Symposium, NASA CP-  Bilanin, A. J., and McKillip, R.M., Jr., “Actuation Device with At Least Three Stable Positions,” U S. Patent  McKillip, R.M., Jr., “Remotely Controllable Actuating Device,” US. Patent No. 5,752,672, Max 19, 1998.  McKillip, R.M., Jr., “”Digital” Tracking Tabs for One-per-Rev Vibration Reduction,” Proc. 59 AHS Annual 2000-209895, May 2000.  No. 6,345,792, February 12, 2002.  Forum, Phoenix, AZ. Mav 2003. ..  McKillip, R:M., Jr., “Actuating Device with Multiple Stable Positions,” U.S. Patent No. 6,220,550, April 24,  2001.  386
YAW GONFIQURATW PITCH ooNpx3uwTIOw  Figure 1. Existing Forced Oscillation System (FOS)  The existing system is operationally inefficient thus keeping wind tunnel productivity low. Beside the need  to halting the wind-tunnel and the FOS system for amplitude change, the inherent one frequency at a time  data gathering is slow and time consuming. Many data points are needed to construct a frequency  response curve for any given aerodynamic derivative coefficient, and there are many coefficients for any  given vehicle configuration.  In addition to operational inefficiency, the motion generated by the existing system is limited to sinusoidal  profiles only. Other motion profiles including ramp and arbitrary waveforms are useful in aerodynamic  testing. Low-speed constant ramp may be useful for quasi-static testing; and arbitrary waveforms as used  in Schroeder sweeps can improve dynamic testing efficiency by measuring the full frequency response  spectrum of the aerodynamics behavior instead of the traditional method of one frequency at time  testing’.  The vision for a new FOS is to create a system that greatly enhances forced oscillation testing  capabilities. The new system should be able to oscillate not only in sinusoidal motion but also other  profiles that may be beneficial for aerodynamic testing. The system should enhance productivity such as  reducing model setup time and reducing run time needed to collect a data set.  System Requirements  The overall design goals of the new FOS included the ability and flexibility to oscillate a model in  sinusoidal as well as non-sinusoidal motion profiles. The system should be able to operate safety and  efficiently - it should be able to operate with different amplitudes and frequencies as well as different  motion profiles without the need to stop the test and mechanically change the system.  Technical requirements for use in the design of the system were derived from free-flight models. Often, it  is desirable to use the same model for lower development cost; Le., the same model design may be used  in wind tunnel, free flight, and drop model tests. Using the same model type of same scale has the added  benefit of avoiding potential inconsistencies which can occur from dynamic scaling.  The two model classes used to establish requirements were based on typical fighter configurations with  characteristics listed in Table 1.  388
Table 1. Model Characteristics  Traditionally, models are sinusoidally oscillated to determine the dynamics derivatives of a vehicle  configuration and then scaled for use in the full-size vehicle control and stability system studies. For the  model sizes in Table 1, Figure 2 shows the theoretical desired frequency versus amplitude of oscillation  for a non-dimensional frequency constant k of 0.18. For the selected constant k and model classes, the  dynamic free stream velocity is usually limited to low speed dynamic pressures'.  0 10 20 30 40 50  Oscillation Amplitude (Degrees)  Figure 2. Theoretical frequency vs. amplitude of oscillation  The 12-Ft Low-Speed Tunnel was selected as the primary facility for the new FOS. The tunnel has a  dynamic pressure of up to a Q of 48 kPa (7 psf) (V =23.5 m/s (77 ft/sec) at standard sea level conditions)  which meets the frequency scaling factor and satisfies the low speed requirement. In addition, this tunnel  has an arc-sector model support system that provides pitch and yaw static positioning of a model and can  be modified for use with the new forced oscillation system. However, the existing 12-Ft arc-sector support  model system was designed for static testing only with loading limitations as listed in Table 2. The model  support system needs to be analyzed for the additional dynamical forces to prevent overloading and to  prevent undesirable structural vibrations. From the static load limitations, the existing model support  system can sustain the 200-lb model class but with reduced normal force capacity (712 N or 160 Ibf).  Along with model support system limits, the balance used to measure dynamics and aerodynamic forces  during testing also imposes additional design constraints to the system. For both the 90-lb and 200-lb  model classes, the FF-10 balance was used. Table 2 also lists balance hading limits.  Additional derived requirements for performance and operational efficiency included:  Range of displacements * 170 deg  System accuracy * 0.05 deg  Finally, the new FOS must be integrated with the existing model support system, data acquisition system  and tunnel safety operations.  389
Table 2. Loading Limitations  Design Analysis  Concept studies were performed to arrive at a system design that would satisfy the design goals for  conducting dynamic derivative studies. Instead of using mechanical linkages to achieve motion profiles,  the new design is based on a tracking principle where a desired motion profile is achieved via a fast  closed-loop positional controller. The motion profile for the tracking system is numerically generated and  thus not limited to sinusoidal motion. It permits non-traditional profiles such as constant velocity and  Schroeder sweeps. This new design approach simplifies the mechanism design but requires more  emphasis on drive system and motion control design as well as system integration.  From the model classes and motion profile requirements, inverse analyses were performed to determine  the required actuating torque and speed to achieve the desired frequency and amplitude. Figure 3 is a  typical inverse dynamic analysis. Using the roll inertia for the 200-lb model, the peak torque of 246 N-m is  required to generate an oscillation of * 10 degrees at 2.5 Hz as shown. Analyses were performed for  frequency and amplitude combinations as shown in Figure 2. Table 3 summarizes the inverse dynamic  analysis results for sinusoidal motion profiles meeting the theoretical frequency versus amplitude of  oscillation curve.  Anaular Position  Q E -0.2 o.~~  0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 08 09 1  Angular velocity  92  ;o (r -2  -4 0 0.1 0.2 0.3 0.4 0.5 06 0.7 08 0.9 1  Angular Acceleration  N 50  50  -50 o 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 I  Torque  200  -200 z fo  0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1  Sec.  Figure 3. Inverse analysis result for a motion profile of 10 degrees at 2.5 Hz  390
i  Mode 3 Q 6.23 Hz  Twisting about Mode 1 @ 4.55 Hz  Side to Side Bending  u,  *  Mode 4 Q 6.41 Hz  Sting Bending Mode 2 @ 4.90 Hz  Fore & Aft Bending  Mode 5 Q 39.94 Hz  Side to Side Bending  w,  Figure 4. Modal analysis results for the arc-sector model support system  392
I 1 i veiouty Loop  ! PXI: Real Time Controller I I: I ; (W Feedback) I  ;( ---------------- - I-  Tandler -1 sumnorno Servo- I  Gearbox Match Gearbox I ~~g mcle  (FS 00) ineaa Right Angle -+ (~~25)  Figure 7. FOS block diagram  The mechanical subsystem of the new FOS is defined by the hardware located in the test section of the  wind tunnel. This subsystem consists of all the drive system components and mounting hardware. Figure  7 shows a FOS block diagram which provides the flow description for all the major components, and  Figure 8 shows a cross-section of the mechanical subsystem. The existing arc-sector model support  system provides static orientation in pitch and yaw; and the new FOS provides roll attitude and dynamic  motion control. The FOS is driven by a high performance Kollmorgen servo motor. This servo motor  orporates high energy rare earth neodymium-iron-boron magnets to provide a high torque-to-rotor  inertia ratio as well as exceptional continuous torque and peak torque performance. This motor has a  peak speed of 3600 RPM with a peak torque of 13.8 N-m and a continuous stall torque of 4.7 N-m.  Right angle spiral. bevel gearb  Angular contact bearing set  Existing model support system  Figure 8. Cross-section of the FOS assembly  Minimizing intrusion to the tail end of the model and potential aerodynamic intetference, the overall length  of the FOS was shortened by utilizing a right angle drive at the motor interface. Even though this  increased the frontal aerodynamic area of the FOS, the motor is still located within the frontal area of the  existing model support arc-sector and therefore overall tunnel flow is minimally affected. The right angle  drive is a Tandler spiral bevel, low backlash gearbox. The main function of this gearbox is to optimize  394
packaging. It provides a 1:l gear ratio. This gearbox was procured with a flanged input interface and a  hollow shaft output interface. The flanged input interface provides a bolt-on interface for the motor with  misalignments handled by an integrated flex coupling.  The output of the Tandler gearbox drives a Sumitorno Servo-Match Gearbox which is a low-backlash  precision cycloid torque multiplier designed for heavy duty industrial robotic applications. The Sumitomo  gearbox has a reduction drive ratio of 89:l. The Sumitomo input interface is a hollow shaft of the same  diameter of the Tandler gearbox. A precision manufactured shaft with machined keyway provides the  drive interface and a precision machined adapter plate provides the housing interface and alignment  between the two gearboxes.  A machined drive shaft is coupled to the output half of the Sumitomo drive. The drive shaft is supportea  by a pair of matched angular-contact ball bearings. The bearing set is sized to react to the cyclic bending,  side and axial loads resulting from aerodynamic and oscillatory dynamic forces. The drive shaft with the  bearing set effectively decouples the Sumitomo drive from all but rotational loads and provides an  interface for sting attachment. A 16-bit encoder measures the rotation of the drive shaft which provides  the sting’s angular position. It is used in the closed-loop motion control.  All of the mechanical components, except the servo motor are mounted inside a cylindrical, titanium  housing. Titanium was chosen for its strength and lightweight properties. The additional weight of the  FOS, compared to the weight of the existing model support sting and mounting hardware, decreases the  existing model support system model weight capability.  Attaching the FOS to the existing model support system required a rigid interface that could tolerate high  side and torsional loading. A tapered dovetail with bolts was chosen for this interface. The existing sting  adapter was also modified as shown in Figure 9 to match this same dovetail interface so both the FOS  and existing sting can be quickly interchanged for various testing needs. A removal bolt is also designed  into the interface to force the tapered dovetail joint apart for FOS and sting removal.  t  Figure 9. Comparison of existing to modified sting interface  The expected system performance for this design for a 200-lb model is shown in Figure 10. By generating  and plotting the various performance curves along with the desired theoretical curve from Figure 2, it was  determined that the first limiting loading factor for roll oscillation testing is the model support system; and  the next factor is due to the load limits of the balance used for measuring aerodynamic loads. The FOS  395
can conceivably exceed the load capability of the test infrastructure. It should also be observed that the  drive system is not limited by its speed but by the maximum motor torque.  I  h 5  E.-  t?:  E: N-  c 4- a. u-  LL  c-  a-  0- .e 3 - c - -. .-  2:  1- Similarly, the expected system performance for a 90-lb model is shown in Figure 11. However, for the  much smaller model class, the entire theoretical frequency versus amplitude curve (Figure 2) is  achievable. The limiting factor for the system is due to motor speed except at frequency greater than 3.4  Hz, where the limiting factor is the motor torque.  \ -Desired Theoretical Curve  *Limit Due to Motor Speed  --)(-Limit Due to Motor Torque  *Balance Limit in Roll Axis - 200 Ib Model  - FF-IO Balance (141 N-m in Roll)  - FA25, FS 00 Gearboxes  - 82048 Kollmorgen Motor  0  5 8 10 12 14 16 18 20 22 25 30 35 40 45 50  Oscillation Angle (degrees)  Figure 10. Predicted performance curve for a 200-lb model  System Integration and Testing  System integration testing was done outside the tunnel in a controlled environment by installing inertia  bars on the FOS to verify performance under simulated loading conditions. Various runs were made with  varying amplitude and frequencies gradually approaching the limitations of the hardware.  396