1

Multiple Actuators Flow Control over a Glauert-Goldschmied type Airfoil at Low Reynolds Numbers

Jonathan Yom-Tov1 and Avi Seifert2

Dep. of Fluid Mechanics and Heat transfer, School of Mechanical Engineering Faculty of Engineering, Tel-Aviv University, Ramat Aviv 69778, ISRAEL

Abstract

Thick airfoils are desirable for optimized structure, large volume, lower weight and more. Low Reynolds numbers flight of Miniature Aerial Vehicles (MAV) is complex, with thick boundary layers and laminar separations. However, the slow flight requires lower thrust, therefore the applicability of active flow control methods due to the significant control authority that zero-net-mass-flux actuators can exert on the flow become appealing. Active flow control can modify the flow as required to perform all functions to sustain and to control flight at low Reynolds numbers. An experiment on a modified Glauert-Goldschmied type airfoil, with a thickness ratio of 20% and Zero-net-mass-flux Piezo-electric fluidic actuators located at the leading and trailing edges of the airfoil was performed at low Reynolds numbers. At high angles of attack (>20°) the leading edge actuators, using very low momentum pulsed modulation input, were able to reattach the flow. Delay of stall and significant increase of Cl,max was achieved. Using the trailing edge actuators with higher momentum coefficients and high frequency pure sine wave excitation enabled both reduction of Cd at all angles of attack and increasing of Cl at low angles of attack (<14°). Significant control of pitching moment was also demonstrated. Combination of the two actuator placements allows efficient combination of the control authority that was found for each actuator operating alone.

Introduction

Thick airfoils are highly desirable from many applicable reasons: optimized structures, large internal volume, lower weight and cost, to name only a few. Traditional aerodynamic design methods, even when optimized, are limited by boundary layer separation and do not account for unsteady effects. The accurate prediction of boundary layer separation is dependent on reliable laminar-turbulent transition prediction, a capability that is still limited. Traditionally, these shortcomings in the design capability of thick airfoils have been overcome using steady suction [1-6] to delay or eliminate boundary layer separation, and even generate form-thrust [4]. Low Reynolds numbers flight of Miniature Arial Vehicles (MAV) is more complex, due to the absence of boundary layer transition, in most common cases, and the more significant effects of viscosity, thickening the boundary layer and promoting its separation. On the other hand, the slow flight speed requires lower thrust to sustain the flight. Therefore, enhances the applicability of active flow control (AFC) methods due to the significant control authority that zero-net-mass-flux (ZNMF) actuators can exert on the flow. This would allow replacement of the cumbersome suction systems by light-weight, efficient actuators. ZNMF Piezo fluidic actuators are suitable candidates to affect the flow significantly and therefore perform all functions required to maintain flight at low Reynolds numbers. Significant lift [7] and maintenance of attached flow, resulting in low drag could be achieved if the actuators will effectively control boundary layer separation [8]. All maneuvering affecting moments could be tailored as needed, provided that the airfoil and the 3D layout of the wing would be appropriately tailored. This could be achieved if the flow state could be altered between fully separated to fully attached, at all incidence angles [9]. Sufficient thrust could be generated using either the direct thrust generated by the actuators [e.g., Yehoshua and Seifert, Ref. 10] or form-thrust [i.e., Goldshmied’s work, Refs. 4-6] with fully attached flow distribution on aft facing, negatively sloped regions. All of the above 2D aspects are studied in the current experimental investigation.

� M.Sc. student � � �� � Corresponding author: HYPERLINK "mailto:Seifert@eng.tau.ac.il" Seifert@eng.tau.ac.il , Senior lecturer, Associate Fellow AIAA

35th AIAA Fluid Dynamics Conference and Exhibit6 - 9 June 2005, Toronto, Ontario Canada

AIAA 2005-5389

Copyright © 2005 by the author(s). Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.

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The Experiment The experiments were conducted on a modified Glauert airfoil (Fig. 1, its upper surface was the basis for the Hump model experiments, Ref. 8) at low chord Reynolds numbers, Re=3.6 to 18x104. The airfoil has a maximum thickness of 20%c and a chord (c) of 140mm. The airfoil spanned the width (b=609mm) of the Meadow-Knapp low-speed wind tunnel test section, at the Meadow aerodynamics laboratory of Tel-Aviv University. The tunnel test section height is 1500mm and the tunnel speed range is 4-65m/s. Turbulence level is below 0.2%.

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Fig. 1a The modified Glauert airfoil, 20% thick with streamwise row of pressure tap locations indicated by circles.

Fig. 1b The airfoil as installed on the bench-top for actuators calibration.

Fig. 1c A hot-wire positioned at the actuator’s exit slot during calibration. Slots are 0.5mm wide.

The airfoil angle of attack was monitored by an analog tilt sensor. The tilt sensor output was periodically

verified by an inclinometer positioned on a calibration surface mounted on the upper surface of the airfoil, parallel to the chord.

The airfoil has 33 pressure ports in the streamwise direction, located in the middle of the airfoil. Seventeen (17) additional ports measured pressure in the transverse direction. These ports were in two rows: one at 56.5%c, having 9 ports, and the other at 81.6%c and having 8 ports. The airfoil pressure ports were connected to a pressure scanning system with 128 sensors with full scale of 10” water column and a resolution of 0.001psi.

The wake total pressure distribution was measured by a wake rake (used for total drag calculations), with 28 pressure sensors, separated 30mm from each other in the vertical direction and positioned 120cm (8.6 chords) downstream of the airfoil trailing edge. When better than 30mm resolution in the wake was desired, the wake was automatically moved by a traverse system, and additional measurements were performed until convergence and desired resolution was obtained. The wake rake pressure sensors were connected to a pressure transducer (10mmHg, 0.06%) via a multiplexer

Free stream velocity was measured both upstream and downstream of the airfoil using 10” water column transducer with 0.01% resolution. The upstream measurement was conducted using a differential Pitot tube connected to a pressure transducer, while the downstream measurement was conducted using two Pitot tubes positioned alongside the wake rake’s position and connected to the wake scanning pressure system.

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Three rows of compact, cavity installed, piezoelectric actuators (with 14 actuators in each row) spanned the airfoil in three rows. The first row was located at the leading edge (LE, x/c=0), and the additional two rows at x/c=0. 63 and 0.67. The periodic excitation was ejected at a shallow downstream directed angle from 1mm high, (H≈0.7%c) by 42mm wide slots (≈30%c or ≈6.9% the span which is ≈4.35c). The actuators input excitation signals were provided by several, computer controlled two channel arbitrary function generators connected to the computer via a GPIB bus. The signals were amplified using 9 audio amplifiers and transformers. The 9 amplified signals were routed to the actuators via a switch matrix that allows the operation of a large number of combinations of actuation modes. Total voltage gain was between 15 to 25.

Two kinds of actuation signals were used: pure sine wave (PS), in which the actuators received a continuous AC signal at their Helmholtz resonance frequency, and burst mode (BM), in which the actuators received either two (for the trailing edge actuators) or three (for the leading edge actuators) bursts (cycles) of an AC signal at their Helmholtz frequency [10, 11]. These bursts were repeated at different low repetition rates, generating low frequencies. For example, the trailing edge actuators could have received two cycles of a sine wave at their Helmholtz frequency ten times a second.

The actuators were calibrated, after installation in the airfoil, on a dedicated bench-top calibration rig. The Helmholtz resonance frequency of the forward row of cavities (hereafter termed LE) was 1435±60Hz, while that of the upper trailing edge cavities (hereafter termed TEU) was 1255±70Hz. The actuators in the lower TE row were not used in this experiment. The frequency and amplitude response (peak slot exit velocity) of the LE and TEU actuators were measured by a hot-wire positioned at each actuator exit slot. Different voltages were supplied to the actuators so as to achieve uniform momentum coefficient across the span. This process was performed in an iterative manner, until satisfactory spanwise uniformity of the measured pressures was achieved. Another calibration was performed at the tunnel, after the test completion, using a miniature Pitot tube . The definition of the excitation momentum coefficient is based on the total excitation momentum due to the active

slots, normalized by a reference free-stream momentum in the form: 2

_

2∞∑= UAuAC wing

slotsactivepeakslotµ , where

slotA is a single excitation slot exit area and wingA is the plan-view wing area. The dimensionless excitation

frequency is defined as: ( ) ∞+ −= UXXfF slotte / , where f is the excitation frequency (1255Hz for all the TEU

actuators and 1435Hz for the LE actuators, using PS or the modulation frequency when BM was used), the distance between the excitation slot and the trailing edge is ( ) =− slotte XX 55mm mm and ∞U is the free-stream velocity (4 to 20 m/s in the current study, resulting in chord Reynolds number (Re) in the range 3.6x104 to 18x104. The side walls of the model are transparent, allowing future flow visualization and PIV interrogation of the flow above the airfoil. Pressures were measured by a PSI Inc. pressure scanner with a resolution of 0.02psi. Lift, pitching and rolling moments and form-drag were calculated from the pressure distributions while total drag was calculated from integrating the wake velocity distribution, measured about 1m downstream of the airfoil trailing edge. Experimental uncertainty for the data at Re=100,000 and below: ∆α=±0.5°, ∆U∞ ±2%, T=24±1°C, ∆Re=±5%, ∆Cp=±5%, ∆(Cl, Cm, Cdp) =±4%, ∆Cd=±10% or ± 0.005 (the bigger), ∆Cµ=±25% (these values depend on the extent of reverse flow residing on the airfoil and on the Reynolds number).

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Results Baseline Airfoil Performance Figure 2a describes the baseline lift vs incidence of the modified Glauert airfoil at low chord Reynolds numbers (Re). The maximum lift of the baseline configuration (Cl,max) is 1.1, with the exception of the Re=36k data, for which Cl,max is lower, approximately 0.8. The lift increases linearly with α, indicated by a slope of dCl/dα ≈0.071/deg, in the range -5°< α <5° and for Re≥90k. Between α=7° and α=15°, a lower lift slope is followed by a sharp rise in Cl, due to laminar-turbulent transition before separation takes place off the upper surface. As the Reynolds number increases, the Cl jump takes place at lower α due to earlier natural transition. For larger incidence angles, beyond the lift jump, Cl is almost constant. A slight additional increase in lift can be observed at α=22-26°, where Cl,max occurs. At negative angles of attack and low Reynolds number a separation bubble appears on the lower surface.

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Fig. 2a The effect of the chord Reynolds number (Re) on the modified Glauert airfoil baseline lift vs. α.

Fig 2b The effect of the chord Reynolds number on drag vs. incidence (α).

Figure 2b presents the baseline drag, measured by the wake integration method. It can be seen that Cd is fairly constant for –5°<α<3° and is about 0.03 with the exception of Re≤45k. Thereafter Cd increases slowly with α, then decreases abruptly to Cd≈0.08 due to transition, after which it rises sharply as the airfoil stalls. Transition takes place between rather large α≈15° for Re=36k to small α≈8° for Re=135k, causing abrupt drag reduction. This transition takes place at decreasing α as Re increases, but the magnitude of the drag reduction gets smaller as Re increases. Future tests will include tripped boundary layer experiments, but previous experience teaches that roughness has very little effect at these chord Re numbers. Leading Edge Excitation

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Fig. 3a The effect of the low pulsation frequency scan on post-stall lift and drag. Re=45k, Cµ=0.04, leading edge excitation, BM3.

Fig. 3b The effect of BM excitation from the LE at post stall condition over a wide range of Strouhal numbers at Re=70k,.LE BM3, Cµ=0.002

We first consider excitation from the leading edge in order to affect boundary layer transition and improve post-stall lift. An incidence angle of α =25° was selected for the performance of frequency and amplitude sensitivity tests, based on previous experiments and the current baseline airfoil lifting performance (Fig. 2a). A frequency scan using

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only 3 cycles of the actuator output with a low repetition rate (BM3), resulting in F+=O(1) is shown in Fig. 3a. Lift (RHS ordinate) is optimally increased with F+≈0.5 while drag (LHS ordinate) is optimally reduced using 2-3 times higher F+. It could be seen that while there is only one lift maximum, there are two drag minima, i.e., at F+=1.5 and 3, as can be seen for a slightly smaller Re (Fig. 3b) allowing a wider range of F+. Post-stall Cµ sensitivity at low and high F+ are shown in Fig. 4a, For Cl increase, burst mode (BM3) is much more effective than pure sine (PS), though the latter used significantly larger Cµ. Using burst mode, there is an initial sharp rise in Cl, which effectively is doubled using Cµ≈0.002, after which no further lift increase occurs. For pure sine wave excitation, there is no consistent response before Cµ>0.1, after which there is a gradual increase in Cl throughout the range of measured Cµ.

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Fig. 4a Lift alternation vs Cµ due to low (F+=0.35) and high frequency (F+=40) excitation emanating from the leading edge amplitude scanslot at post-stall low Re number (Re=36k), α=25°, ∆ϕ=0.

Fig. 4b Drag alternation due to low (F+=0.35) and high frequency (F+=40) excitation emanating from the leading edge slot at post-stall low Re number (Re=36k).

The data presented in Figure 4b shows that burst mode (BM) actuation from the leading edge is less effective than pure sine (PS) wave excitation for significant drag reduction. This is in contrast to the lift alternation where BM was more effective than PS. Pure sine operation results in an initial, and significant, decrease in drag, possibly due to transition to turbulent flow prior to separation, after which no further changes occur. It is evident that pure sine wave excitation with a small Cµ (≈1%) roughly halves the drag and –∆Cd/∆Cµ≈0.25/0.01≈25. However, the excitation effectiveness is maximal for burst mode with Cµ≈0.2%. For this condition, ∆Cd/∆Cµ≈40. Note that the Cµ scales are logarithmic in Figures 4.

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Fig. 5a Cl vs. α for baseline, high frequency high amplitude, low frequency low amplitude post stall leading edge excitation. Re=45k.

Fig. 5b Cd vs. α for baseline, high frequency (F+=40) high amplitude, low frequency (F+=0.56) low amplitude post stall leading edge excitation. Re=45k.

Figure 5a demonstrates the importance of the excitation frequency when the control is applied from the leading edge. When high frequency pure sine wave excitation is applied, with F+≈40 and Cµ=0.35, only a marginal

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improvement in lift is observed. When using excitation in burst mode with F+≈0.56 and Cµ=0.004, significant gains can be noticed, probably due to earlier transition between 8°<α<12° and at post-stall incidence (α>22°), despite the two orders of magnitude lower BM Cµ. In contrast to Cl, pure sine (PS) excitation has a stronger effect on Cd than burst mode (Fig. 5b). This is expected due to the two orders of magnitude larger Cµ of the PS excitation. The effect occurs almost throughout the range of α, but it is especially evident for 4°<α<14° and 16°<α<20°. However, the maximum momentum efficiency (at α=12° and 20 deg°) is low, only about one third.

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Fig. 6a LE excitation effect on Cp, α=25°, LE BM excitation (BM3, F+=0.56, Cµ=0.004). Re=45k

Fig. 6b Flow visualization image of controlled flow at α=25deg. Leading edge BM3 excitation, Re=23k.

Figure 6a shows pressure distributions at post-stall condition. LE burst mode excitation with 3 active cycles (BM3) at α=25° results in effective reattachment of the flow to the upper surface from the leading edge to the TEU slot region. Separation moved from x/c=0.2, for the baseline, to x/c=0.7 with LE burst mode excitation. However, it is most likely that the mean Cp shown for LE burst mode operation is due to a large and unsteady separation bubble and not due to stationary reattachment. The above hypothesis is validated by lower Re flow visualization image shown in Figure 6b. Figure 6b presents smoke flow visualization at α=25 and Re=23k. Leading edge excitation with BM3 and low Cµ is used to reattach the flow and generate large coherent structures that can clearly be seen in the image above the surface. Note that these structures were generated by pulsed modulated excitation and there is no evidence to the high frequency carrier wave. Excitation from the Trailing Edge Slot After studying the effect of the LE slot we now shall focus our attention to the trailing edge slot with the aim of partial or full control of the massive separation off the cusp region. Figure 7 shows the lift alteration due to pure sine wave excitation, with F+≈O(10), introduced from the upper trailing edge (TEU hereafter) slots with unison phase across the span. The AFC effect is evident over the entire Re range tested. It is most efficient at Re=90k (i.e. largest dCl/dCµ≈0.25/0.01≈25). The lift alternation effect reaches saturation at Cµ=0.07 for Re=45k and an intermediate saturation at Cµ=0.03-004 for Re=68k. The lift increment is almost linear with Cµ between 0.04>Cµ>0.01 with dCl/dCµ>10. Figure 8 shows the strong but complex effect the excitation magnitude, issuing from the TEU slot with PS signal, has on drag. For Cµ<0.03-0.04 there is a tendency to increase Cd. The drag coefficient decreases almost monotonically with Cµ at Re=45k. The drag reduction effects seem to saturate at Cµ(0.08 and Cd(0.015.

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Fig. 7 Lift response to trailing edge upper (TEU) slot excitation at low incidence and Re as indicated in the legend, F+=15, 10 and 7.5 for Re=45, 68 and 90k respectively, α=0.

Fig. 8 Drag response to trailing edge upper (TEU) slot excitation at low incidence. Re as indicated in legend, F+=15, 10 and 7.5 for Re=45, 68 and 90k respectively, α=0.

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Fig. 9a Lift and drag alternation vs Cµ for high and low frequency excitation emanating from the TEU slot, F+=15 (PS) and 2 (BM), Re=45k, α=6°.

Fig. 9b Wake velocity distributions for several Cµ’s that are shown in Fig. 9a.

The results of an amplitude scan performed at α=6° are presented in Fig. 9a. The excitation issues from the TEU slot at a flow condition where baseline massive separation takes place at 0.4<x/c<0.6 on the upper surface, certainly upstream of the slot. The lift is unaffected, or even slightly reduced, by pure sine (PS) wave excitation AFC until Cµ>0.04 is reached, after which there is a sharp increase in Cl. The lift data shows that burst mode (BM) operation is either slightly detrimental or ineffective to Cl, at the available range of Cµ, in good agreement with the PS data, at a similar Cµ. The drag response to Cµ is gradual in the low Cµ range, still the drag is reduced by more than 30% for Cµ<0.04. Operation with pure sine excitation (PS) signal results in a sharp decrease in Cd, at Cµ=0.06. The optimal drag reduction is obtained at the above Cµ with about 70% Cµ efficiency and a minimum Cd of approximately 0.01 (using span uniform phase). Burst mode (BM with F+≈2) has minimal effect on Cd, in the available range of Cµ and flow conditions (not shown). Figure 9b presents the wake velocity distributions for the conditions of the data presented in Fig. 9a. The wake moves down, in agreement with the increased lift and circulation on the airfoil, as the control reattaches the flow to the TEU region. The maximum wake velocity deficit is reduced and obtains a maximum of only 1%U for Cµ≈0.066.

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Fig. 10a Lift data due to low and high frequency excitation from the TEU slot, vs. α .PS: Cµ=0.084, F+=15, BM2: Cµ=0.013, F+=2

Fig. 10b Trailing edge upper (TEU) PS excitation effect on Cp at α=0 and Re=45k. F+=15, Cµ=0.065, Ph=0.

Figure 10a shows baseline Cl and the TEU slot effect on Cl-α distributions for two excitation modes. Pure sine operation (PS, high F+ and large Cµ) significantly increases Cl at α<6°, but loses efficiency between 6°<α<12°, presumably because laminar separation moved upstream of the active slot. Operation in the PS mode promotes transition to take place at 12° instead of 14° at the baseline, rendering the excitation beneficial again. Burst mode actuation slightly reduces Cl over most of the α range that was studied. At α>12° the TEU slot loses efficiency again due to the motion of separation upstream of the slot. In that condition, as for 6°<α<12°, the TEU excitation is ejected into still or “dead” air, rendering it inefficient. Figure 10b compares baseline and controlled pressure distributions acquired at low incidence while the TEU actuator attempts to control the massive separation off the cusped TE region. When the TEU actuators are operated, the suction on the upper airfoil surface is stronger than in the baseline (up to x/c=0.8). Note also that the entire lower surface Cp becomes significantly more positive. At the point where the flow reaches the discontinuity in the airfoil upper surface slope, a sharp pressure rise occurs. Due to the weakening excitation magnitude when progressing downstream from the slot, a secondary bubble appears at 0.7<x/c<0.85. Note that the flow reattaches to the TE, since mean Cp=0 there. The second TE slot located at downstream of the TEU (Fig. 1c) will be combined with the TEU slot to eliminate this secondary bubble.

Fig. 10c Baseline flow visualized by smoke showing TEU slot and cusp region. U~2.5m/s (Re=23k). α=0

Fig. 10d Controlled flow excitation from TEU slot (shown). Note the change in streaklines direction above and below the airfoil. (=0

Figures 10c and 10d present flow visualization images for the baseline and controlled flows respectively. Control was applied from the TEU slot. For the baseline it is evident that the flow separates from the upper surface upstream of the TEU slot. The streaklines leaving the lower surface are curved up, indicative of the negative lift at this condition. When TEU excitation is applied, the smoke pattern indicates a complete reattachment to the entire upper surface and significant alternation of the streaklines direction also under the airfoil. There is no evidence for the generation of coherent structures with the high F+ TEU excitation.

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Combined Excitation from the Leading and Trailing Edge Slots

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Fig. 11 Lift vs. α for baseline, LE BM excitation (BM3, F+=0.56, Cµ=0.004), TEU excitation (F+=15, PS, C(=0. 084, Ph=0) and combination thereof. Re=45k

Fig. 12a Drag vs alfa for baseline, LE BM excitation (BM3, F+=0.56, Cµ=0.004), TEU excitation (F+=15, PS, C(=0. 084, Ph=0) and combination thereof. Re=45k.

It is desirable to attempt utilizing the beneficial effects of each of the available excitation slots, operating modes and effective flow conditions of each mode and slot to optimally use the capabilities of the current multivariable-multi-component AFC system. The lift resulting from a combination of the LE and TEU actuators is shown in Fig 11 and discussed below. LE actuators, operating in burst mode at low Cµ, increase Cl in the range 8°<α<12° (due to promoted laminar-turbulent transition) and at α=23°-30° (due to post-stall leading edge separation control). The TEU actuators, operating in pure sine mode (high F+ and large Cµ) alone, increase lift at –8°<α<6° and at 10°<α<12°, as compared to the baseline. Operation of TEU actuators at pure sine with the LE actuators in burst mode, produces no additional lift benefit when compared to the operation of TEU actuators alone, except for the range 6°<α<12° where the LE excitation promotes transition and eliminates the performance degradation of the TEU PS data around α=10° . In fact, the operation of the TEU actuator with pure sine wave excitation nullifies the LE excitation beneficial effects at α=23°-30°. When both LE actuators (in burst mode) and TEU actuators (in pure sine) are operated together, Cl is affected uniformly throughout the range of measured angles of attack, unlike the case where only the LE actuators are operated, which results in an increase of Cl only for α> 6°, or when only TEU actuators are operated, in which case there is a fall followed by a rise in Cl in the range of 6°-12°. However, at post-stall conditions and with the given Cµ for the TEU actuators, only the LE burst mode excitation should be used for optimal lift improvement. This could be understood because when the flow reattaches to the region downstream of the TEU actuator, Cp there becomes more positive, i.e., reduces Cl locally. Very high Cµ is required to compensate for this positive Cp region by a negative Cp alternation upstream of the slot. Drag vs. incidence data is shown in Fig. 12a for the conditions of Figure 11. It is evident that LE burst mode lowers Cd only at α greater than 25°, where massive separation takes place at the leading edge. TEU operation with pure sine wave excitation is the most effective drag reduction approach, better than LE burst mode and LE burst mode combined with TEU pure sine. Reduction of the efficiency of TEU actuation can be seen above α=25°, although TEU pure sine actuation mode is still the most efficient. However, for α<20°, the TEU actuators alone provide most of the drag reduction with negligible effect due to the LE actuators operating in burst mode.

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Cm baseCm LE BMCm TEU PSCm LE BM & TEU PS

Fig. 12b Lift vs quarter chord pitching moment for baseline, LE BM excitation (BM3, F+=0.56, Cµ=0.004), TEU excitation (F+=14, PS, Cµ=0. 084, Phase=0) and combination thereof. Re=45k.

Fig. 13 Cm1/4 vs α for baseline, LE BM excitation (BM3, F+=0.56, Cµ=0.004), TEU excitation (F+=14, PS, Cµ=0.084, Ph=0) and combination thereof. Re=45k, Phase=0.

The effect of the different excitation possibilities on the pitching moment is shown in Figures 12b and 13. The leading edge excitation has mostly a forward loading increase capability, therefore rendering Cm more positive. The TEU and LE excitations combined, have a strong negative contribution to the pitching moment. It is evident that at low lift coefficients and low angles of attack, the baseline pitching moment can be altered both more negative and to a lesser extend more positive, just by activating the different available slots. Using the Cµ control, one can tailor Cm in any intermediate points between the extreme cases, resulting from varying levels of flow reattachment to different regions of the airfoil. As the lower free stream velocity becomes smaller, it deemed feasible that the PS excitation emanating from the TEU slot at low incidence will be able to propel the airfoil. This is demonstrated in Fig. 14. The wake velocity distributions and Cd that was calculated from it are significantly reduced at α=3° and U=4m/s. The wake almost disappears when operating the TEU actuators with pure sine wave input and high Cµ. The wake contains a jet region, but poor mixing between the remaining wake and jet regions probably indicates the low thrusting efficiency of this approach.

3.5

4

4.5

5

5.5

6

0.92 0.94 0.96 0.98 1 1.02U/Uinf

y/c

baseline Cd=0.074Cµ=0.1534 Cd=0.019Cµ=0.1531 Cd=0.010Cµ=0.1528 Cd=0.006

-0.2

0

0.2

0.4

0.6

0.8

1

1.2

1.4

0 0.03 0.06 0.09 0.12 0.15Cd

Cl

Cl baseline

Cl afc

Fig. 14 Wake velocity profiles at Re=36k and (=3deg demonstrating drag nullification via ZNMF TEU actuators. Cµ and Cd are indicated in the legend. F+=18, Phase=180deg.

Fig. 15 Lift vs. drag for baseline and TEU slot PS control. Cµ=0.283, F+=18, Phase=180 deg.

11

The Cl-Cd plot shown in figure 15 shows how AFC can turn a useless airfoil (in terms of lift-to drag ratio with useful lift) to one that provides performance comparable to high Re number (Re>106) airfoils. The drag is essentially constant and very close to zero with L/Dmax≈100 due to activation of the TEU actuators operating with a phase shift of 180 deg between each pair.

Summary, Conclusions and Recommendations The baseline flow over the modified Glauert airfoil contains a massively separated flow region located on the upper-aft surface. As the chord Reynolds number is decreased below 100,000, complex phenomena due to the interplay between transition, separation and reattachment dictates the airfoil performance. The current study made use of two rows of zero-net-mass-flux (ZNMF) actuators. One row that is located at the leading edge, had a favorable effect in promoting transition on the upper surface, therefore delaying premature, laminar boundary layer separation. At post-stall condition, low frequency, pulsed modulated excitation emanating from the leading edge actuator, with Strouhal number of about 0.5, increased Cl,max by about 25-30% using minute oscillatory momentum coefficient. Flow visualization validated that the mechanism for post-stall lift enhancement was due to the generation of large coherent structures corresponding to the low pulsating frequency and not the high resonance frequency of the actuator. Excitation emanating from the aft slot, located just upstream of the separated flow region, where the airfoil slope discontinuity takes place, dramatically altered the airfoil performance at all pre-stall angles of attack. Significant lift alternations, e.g., from –0.2 to 0.5, and drag nullification were measured. The mechanism is attributed to complete flow reattachment to the aft cusp region, generating form-thrust and access wake momentum (i.e., jet). Quarter chord moment was significantly altered as well. As in other recent studies, non-uniform spanwise phase distribution resulted in increased efficiency of the excitation system. Pulsed modulation did not prove as effective on the trailing edge region as it did for the leading edge region. The reason is unclear and requires further study. A preliminary effort to combine the leading edge excitation and the trailing edge region excitation proved beneficial and allowed to identify which actuator optimally affects the flow in a desirable manner for a given task and flow condition. This effort is a first step in a focused research aimed at enabling sustained low Reynolds number flight without moving parts. It should be supplemented by a flow field interrogation, a parallel CFD effort and a feedback control study. Higher Re applications are dependent upon new generations of actuators that will enable utilizing the same approach to efficient thick flying wings with distributed propulsion and optimized feedback flow control system.

Acknowledgment The authors would like to thank Ilan Fono, Eli Nevo, Shlomo Paster, Eli Ben-Hamou, Eli Kronish, Avram Blas and Mark Vasserman for the support.

References 1. Glauert, M. B., “The Design of Suction Aerofoils with a Very Large CL-Range”, Aeronautical Research

Council, R.&M. 2111, November 1945. 2. Saeed, F. and Selig, M. S., “Multipoint Inverse Airfoil Design Method of Slot-Suction Airfoils”, Journal of

Aircraft, Vol. 33, No. 4, 1996, pp. 708-715. 3. Glauert, M. B., Walker, W. S., Raymer, W. G., and Gregory, N., “Wind-Tunnel Tests on a Thick Suction

Aerofoil with a Single Slot”, Aeronautical Research Council R. & M. No. 2646, October 1948. 4. Goldschmied, F. R., “Airfoil Static-Pressure Thrust: Flight-Test Verification”, 1988 AIAA Aircraft Design and

Operating Meeting: Applied Aerodynamics Area, 1988. 5. Goldschmied F. R., “Integrated Hull Design, Boundary-layer Control, and Propulsion of Submerged Bodies”,

Journal of Hydronautics, Vol. 1, No. 1, 1967, pp. 2-11. 6. Goldschmied, F. R., ”Fuselage Self-Propulsion by Static-Pressure Thrust: Wind Tunnel Verification”, AIAA

Paper 87-2935, 1987. 7. Seifert, A., Darabi, A. and Wygnanski, I., “On the delay of airfoil stall by periodic excitation”, Journal of

Aircraft, Vol. 33, No. 4, 1996, pp. 691-699. 8. Seifert, A. and Pack, L.G., “Active Control of Separated Flow on a Wall-mounted “Hump” at High Reynolds

Numbers”, AIAA J., V. 40, No. 7, July, 2002, pp. 1363-1372. (Part of AIAA paper 99-3430). 9. I. Timor, E. Ben-Hamou, Y. Guy, and A. Seifert, “Maneuvering Aspects and 3D Effects of Active Airfoil Flow

Control”, AIAA paper 2004-2614, 2nd AIAA Flow Control Conference, Portland, Oregon, June 28-1, 2004. 10. Yehoshua, T. and Seifert, A. “Boundary Condition Effects on Oscillatory Momentum Generator”, AIAA Paper

2003-3710, June 2003 (MSc thesis Tel-Aviv Uni. 2004, J. papers submitted). 11. Margalit, S., Greenblatt, D., Seifert A. and Wygnanski , I., “Delta Wing Stall and Roll Control using Segmented

Piezoelectric Fluidic Actuators”, (AIAA paper 2002-3270), Accepted for publication, AIAA J. of Aircraft, March 2004.