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supported by two plates that flex only in the drag direction and hang from two more vertical plates attached to the fixed base of the force balance. Both the lift and drag platforms were connected to separate flexures on which foil strain-gauge bridges were mounted. The strain-gauge bridges provided voltage outputs proportional to the respective lift and drag forces. The voltages were amplified using custom-designed operational amplifier circuits that minimized offset drift and sensitivity to external electronic noise. Calibration of the force balance was done by applying known weights to a cable pulley system attached to the support sting. The average uncertainty in the force measurements was 0.63% in lift and 0.9% in drag. The experiment was controlled by a digital computer with a programmable analog-to-digital converter (ADC) and digital input– output (DIO) interface. The minimum voltage resolution of the ADC was 0.6 mV. The voltages proportional to the lift and drag forces were acquired along with a voltage proportional to the velocity at the entrance to the test section. The acquisition software was programmed to acquire 10,000 voltage samples over 10 s. This was found to provide repeatable time-averaged statistics that varied by less than 0.1%. The angular position of the airfoil was controlled by voltage pulses from the DIO into the stepper motor controller; with this, the angular position was repeatable to within 0:005 deg. A mechanical readout that was geared to the stepper motor shaft provided positive feedback on the angular position. Before making lift–drag measurements, values of the lift and drag voltages were first acquired at different angles of attack without flow. Any difference from the zero-force voltage that was due to eccentric loading was recorded and subtracted from the results at the same angles of attack with flow. This process was repeated any time the model was removed from the force balance. The freestream speed at the entrance to the measurement section was measured with a pitot- static probe connected to a pressure transducer. The output of the transducer was monitored on a dc volt meter and simultaneously acquired by the data acquisition computer when the voltages proportional to the lift and drag forces were acquired. Based on the pressure transducer calibration, the accuracy of the freestream speed measurement was 0:01 m=s. The combination of the uncertainties in the force measurements and voltages resulted in an average uncertainty in the lift and drag coefficients of approximately 1%. The high voltage leads to the plasma actuator were well shielded to minimize the noise effect, if any. The output from the lift and drag channels from the force balance were connected to an oscilloscope to check for RF noise interference when the plasma actuator is turned on. No noise interference was noticed on the scope. This confirmed that the plasma actuator did not corrupt the force data. These experiments began with lift and drag measurements of the model with traditional control surfaces. Then fluorescent-oil flow visualization experiments were conducted that were used to characterize the vortex breakdown and separation lines at different angles of attack. Following this, laser-smoke flow visualization experiments were conducted to capture off-surface flowfield information on the lee side of the wing model. Finally, lift and drag measurements were performed on a half-span model to identify the optimal location for plasma actuators for lift control. These involved plasma actuators on the lee side and wind side of the wing leading edge. In the present work, the plasma actuator consisted of two 0.05 mm (0.002 in.) thick copper electrodes separated by two layers of 0.05 mm (0.002 in.) thick Kapton film. The Kapton film has a breakdown voltage of approximately 7 kV per 103 -in. thickness and a dielectric constant of 3.3, which provide good electrical properties. The electrodes were arranged in an asymmetric arrangement, as shown in Fig. 1. They were overlapped by a small amount (on the order of 0.5 mm) to ensure a uniform plasma in the full spanwise direction. The plasma actuator was bonded directly to the surface of the wing. When the ac voltage amplitude was large enough, the air ionized in the region over the covered electrode. A 0.1-mm recess was molded into the wing model to secure the actuator flush to the surface. The two copper-foil electrodes were aligned parallel with the leading edge. The spanwise length of the actuators was 90% of the wing span. With this arrangement of electrodes, the body force produced by the actuator would induce a velocity component in the direction from the exposed electrode toward the covered electrode. With the actuator oriented on the leading edge, it induces a flow around the leading edge of the wing. Many different actuator arrangements were examined on both the lee-side and wind-side portions of the wing. Figure 3a shows an example of a plasma wing configuration examined with a continuous spanwise plasma actuator at the leading-edge and two trailing-edge split actuators. The effect of trailing-edge actuators are reported in a different study [32]. Figure 4 shows an example of a plasma wing configuration with multiple plasma actuators at the wing leading edge. PATEL ET AL. 1267 Fig. 3 Photographs of a plasma wing model in a wind-tunnel test section (left), with plasma on (right). Photograph of another plasma wing configuration with four Fig. 4 actuators at the wing leading edge.PDF Image | Plasma Actuators for Hingeless Aerodynamic Control of an Unmanned Air Vehicle
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