Dynamic Stall Control on a Vertical Axis Wind Turbine ...

50th AIAA Aerospace Sciences Meeting including the New Horizons Forum and Aerospace Exposition 09 - 12 January 2012, Nashville, Tennessee

AIAA 2012-0233

Dynamic Stall Control on a Vertical Axis Wind Turbine Using Plasma Actuators

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David Greenblatt1, Amos Ben-Harav2 and Magen Schulman3 Technion - Israel Institute of Technology, Technion City - Haifa 3200, Israel

Dynamic stall was controlled on the blades of a small high-solidity vertical axis wind turbine, by means of dielectric barrier discharge plasma actuators installed on the blade leading-edges. A parametric study was conducted with the objective of increasing the net power output resulting from control in an open-loop manner. For the majority of experiments, the actuators were configured to control separation on the upwind half of the turbine azimuth while limited experiments were conducted on the downwind half. Turbine power was measured using a specially-designed dynamometer that allowed full characterization of its performance. Actuator duty cycle dependence observed previously on static airfoils was also observed on the turbine. However, optimum reduced frequencies showed substantially different dependence and this was traced to the importance of the plasma pulsation frequency relative to the turbine rotational frequency. Overall turbine power performance improvements of 38% and 20% were measured for upwind and downwind dynamic stall control, respectively. Based on the data acquired, up-scaling the turbine by a factor of 5 and 10, the percentage of plasma power required to produce comparable improvements was conservatively estimated at 3.3% and 1.7% respectively. A new switching-direction actuator was developed, together with high-speed plasma triggering, in order to produce combined upwind and downwind actuation that is phase-locked to turbine rotational frequencies. Flowfield measurements using particle image velocimetry are presently being performed.

Nomenclature c CP

airfoil or turbine blade chord length turbine power coefficient, Pt / qU  HD

CT

static turbine torque coefficient T / q cHR plasma momentum coefficient Fp / q cH

C D DC fion fp Fp F Ft ,u  t ,d

F

H N Pt Pp R Re

turbine diameter, 2R duty cycle: ton/tp plasma ionization frequency plasma pulsation frequency, 1/tp plasma thrust reduced frequency f p c / U  turbine upwind reduced frequency, f p c / (  1)U  turbine downwind reduced frequency, f p c / (  1)U  turbine height number of blades turbine power, T  plasma actuator power input turbine radius Reynolds number

1

Senior Lecturer, Faculty of Mechanical Engineering, Senior AIAA Member. Graduate Student, Faculty of Mechanical Engineering. 3 Graduate Student, Faculty of Mechanical Engineering. 2

American Institute of Aeronautics and Astronautics 1 Copyright © 2012 by David Greenblatt. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission.

ton T U, U q V

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 stall    

time that plasma is active within a cycle turbine average torque wind tunnel speed wind tunnel dynamic pressure wind speed airfoil or blade angle of attack airfoil or blade stall angle of attack blade speed to wind speed ratio, R/U blade azimuth turbine solidity, Nc/R turbine rotational speed

I.

Introduction

In recent years there has been a resurgence in vertical axis wind turbine (VAWT) development for both urbanscale or off-grid applications as well as off-shore alternatives to horizontal axis wind turbines [1-7]. VAWTs have some advantages over horizontal axis machines in that they are insensitive to wind direction, the generator can be located closer to the ground and they produce lower sound emissions due to lower blade speeds [8]. A major problem associated with VAWTs is the phenomenon of dynamic stall [9-11]; namely the boundary layer separates dynamically from the turbine blades. Blade stall is a two-fold problem: on the one hand it dramatically reduces turbine power through loss of lift and increases in drag; on the other hand it produces large unsteady loads that compromise the structural integrity and also impact on the drive-train and generator selection [12,13]. In many cases the problem is compounded in the built environment or on buildings [14] where VAWTs are often required to operate at low rotational speeds and hence low blade speed to wind speed ratios ( ωR/V). This forces the turbine to operate in a partially stalled state that is exacerbated at high wind speed when the potential for power generation is maximized. Historically, the approach to avoiding dynamic stall was to pitch the blades using mechanical levers [15,16]. Attempts are also being made presently to improve turbine performance using circulation control [17]. Additionally, recent investigations have focused performance measurements, flow visualization and PIV measurements of the dynamic stall phenomenon [8]. However, no attempts have been made to directly control turbine blade dynamic stall using active flow control. Although dynamic stall has not been controlled on VAWTs, airfoil dynamic stall control under incompressible conditions has been studied using a variety of techniques, including leading-edge slats, deformable leading-edges, surface suction, slot blowing and zero mass-flux blowing (see references in [18]). The introduction of period perturbations, in the form of zero mass-flux blowing, is a particularly attractive because it requires far less slot momentum when compared to steady blowing. Significant increases in lift and reductions in drag were observed when the reduced perturbation frequency was in the range 0.6  F   1.1 [19] and this renders it potentially viable for application on VAWTs. However, the problem is complex because stall occurs alternately on the inboard or the outboard surfaces of the blades depending on whether they are upwind or downwind. A schematic showing an idealized VAWT plan view in fig. 1 illustrates the differences between inboard and outboard dynamic stall. Here Ua is the wind velocity immediately upwind of the turbine, W is the velocity relative to the blade and  is the local instantaneous angle of attack. The green arrow indicates dynamic stall occurring on the inboard (blue) surface of the blade when the blade is in the upwind half of the azimuth and occurring the outboard (red) surface of the blade when the blade is in the downwind half of the azimuth. Consequently, a flow control strategy with the greatest potential for increasing turbine power would require a method that can control stall alternately on both surfaces of the blade. Furthermore, any technique must be energy efficient in that annual energy yield should far exceed the energy required to drive the actuators. Based on a system study conducted by Sasson and Greenblatt [20,21], effective and energy efficient control could have three important benefits: firstly, the annual energy yield can be increased by up to a factor of three; secondly, the power increases will occur in the lower  regime and hence work in favor of noise reduction and lower structural loads; and thirdly, oscillatory aerodynamic loads imposed on the structure can be reduced. The objective of this study was to record performance improvements on a VAWT by controlling dynamic stall on the blades using periodic perturbations. This was achieved by constructing a small, high solidity, turbine equipped with surface mounted dielectric barrier discharge (DBD) plasmas actuators. The results of a parametric study and a turbine performance assessment are reported here. The application of DBD plasma actuators were demonstrated on stationary and dynamically pitching airfoils by Corke et al [22] and Post and Corke [23]. The American Institute of Aeronautics and Astronautics 2

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actuators are typically driven in the kilovolt range at several hundred hertz to several kilohertz and these produce plasma microdischarges, at time-scales O(10–9s) during the positive- and negative-going drive signals. The net result on the surrounding air is a wall-jet, which is considered analogous to the jet blown from a two-dimensional slot. Post and Corke [23] also considered aerodynamic effects when driving the actuators in a pulsed mode. Pulsing the actuator brings with it two significant advantages. Firstly, it has long been recognized that periodic perturbations can achieve aerodynamic performance benefits that are comparable or superior to steady blowing at one to two orders of magnitude less momentum input [24]. Secondly, Greenblatt et al [25] and Bachmann et al [26] observed that DBD plasma duty cycles (DC) of 1% and 50% produced comparable aerodynamic performance. Because power supplied to the actuator is proportional to duty cycle, it is possible that this plasma power input can be reduced sufficiently so as to render it viable for turbine applications. Finally, DBD plasma actuators have wide appeal because they have no moving or mechanical elements and are robust and reliable.

R

Upwind half

Ua

R

R

Ua

R

Ua

downwind half R

Fig. 1 Top view schematic of a hypothetical four-bladed VAWT showing the velocity components relative to the blade depending upon the position of the blade.

II.

Experimental Setup

The turbine used in this investigation consisted of a vertical shaft rotating on two self-aligning bearings, two horizontal adapters used for the attachment of connecting rods and two vertical blades attached to the connecting rods (see figs. 2a-2c). The blades were connected symmetrically on the shaft but shifted upwards to produce a long free end, intended for future flow particle image velocimetry (PIV) measurements. The turbine was installed in a low-speed blow-down wind tunnel with a 1mx0.61m test section area (2m long) and a maximum wind speed of 55m/s. The maximum turbulence level in the center of the tunnel was 0.3% of the free-stream velocity. A photograph of the turbine mounted in the wind tunnel is shown in fig. 2a. A custom dynamometer was coupled to the shaft and mounted below the turbine outside the tunnel. The dynamometer comprised of a magnetic brake, used to load the turbine, and a Kistler torque/rotational speed transducer (see fig. 2b). Flexible couplings were employed to compensate for misalignments. The turbine dimensions were diameter D=0.5m and height H=0.6m; these dimensions were dictated by the test section dimensions. The blades employed were symmetric NACA0015 profiles of chord length c=150mm, constructed from extruded aluminum. The resulting turbine solidity, namely =Nc/R=1.25, was relatively high.

American Institute of Aeronautics and Astronautics 3

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(a) (b) Fig. 2 (a) Photograph of the turbine mounted in the test section. Also visible are the dynamometer and lower slip-ring below the test section; (b) Schematic of the custom-designed dynamometer with shaft adapter for the slip-ring (not shown).

DBD plasma actuators were attached to the leading-edges of the blades. The actuators used were similar to those used in previous studies, namely the upper (exposed) and lower (encapsulated) electrodes (both 70m thick) separated by three layers of 50m thick Kapton® tape. These were wrapped around the leading-edge of the blade as shown schematically in fig. 3a. A layer of 50m thick Kapton® tape was applied to the blade to isolate the actuator from the wing. The actuators were driven at an ionization frequency of fion=8kHz at V=8kVpp (2.83kVrms). Pulsation frequencies were chosen to represent the F+ range between O(0.1) and O(5), corresponding to 10Hzfp500Hz and the duty cycle range tested was 1%DC50%. An identical actuator was calibrated independently by mounting it on sensitive load cell while recording its developed thrust (Fp ) as a function of power input (see schematic in fig. 3b). The results of this calibration are shown in fig. 3c which shows that the thrust developed at V=8kVpp was 1.5mN/m with a power consumption of 45watts/m at 100% duty cycle. encapsulated electrode (70µm)

insulation

dielectric (Kapton tape) 50µm x3 layers

aluminum blade

exposed electrode (70µm)

Fig. 3 (a) Schematic of the blade leading-edge region showing attachment of the DBD plasma actuator. American Institute of Aeronautics and Astronautics 4

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Fig. 3(b) Schematic of the actuators calibration setup and plasma power input calculation.

Fig. 3(c) Results of the actuator calibration (acknowledgement: Mr. S. Vey). In order to transfer the high voltage AC power required to operate the DBD actuators, slip-rings were installed on both turbine’s shaft ends. An end-type slip-ring was mounted at the top of the shaft and a through-bore-type was mounted below, above the dynamometer. In order to avoid large voltage differences within the each slip-ring, the high-voltage and earthed connectors were connected via the upper and lower slip-rings respectively (see fig. 4a-4c).

III.

Controlling Parameters and Experimental Methods

The reduced frequency definition F+ appropriate to the VAWT blades does not have a straight-forward definition as for example its definition on stationary airfoil studies, which is: f pc F  (1) U where U is the free-stream velocity (wind tunnel velocity) and f p is the plasma pulsation frequency. In the case of the turbine, we can extend this definition by using the velocity relative to the blade namely W, such that f pc (2) Ft   W In all subsequent discussion we shall assume that the chord length is constant, as it is in this study, for simplicity. If we attempt to set a constant Ft  , much like in airfoil studies, we encounter three main difficulties: 1. The instantaneous value of W varies around the azimuth . 2. The variation of W as a function of  depends on  The instantaneous value of W must be estimated theoretically or measured.

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(b)

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(c)

(a)

Fig. 4 (a) High voltage power conveyance schematic; (b) end-type slip-ring at the top of the shaft for the highvoltage signal; (c) through-bore-type slip-ring between the bearing and dynamometer for the signal ground. Thus in order to maintain a constant F+, the physical pulsation frequency f p must be variable. For the sake of simplicity, in this investigation we did not attempt to change the frequency of actuation as a function of  at any of the ’s tested. Instead, a constant frequency was selected for each set of experiments. Apart from this practical simplification, the approach can also be justified on physical grounds because static thick airfoil lift coefficient data show only a weak dependence on the reduced plasma pulsing frequency in the range 1  F   5 [26]. Thus a constant physical frequency can be selected such that Ft  varies within this range. To distinguish between dynamic stall commencing on the advancing and retreating blades, two nominal reduced frequencies were defined: one for experiments conducted on the upwind half of the turbine; and one for experiments conducted on the downwind half of the turbine, namely: f pc f pc Ft ,u   (3)  R  U  (  1)U  and f pc f pc Ft ,d   (4)  R  U  (  1)U  respectively (see fig. 4). Here it is assumed that the so-called induced velocity through the turbine and the tunnel velocity are equal (U=Va), which is a reasonable assumption at low . Although these two definitions do not take into account the variability of the reduced frequency, nor the interference effect of the turbine, they represent nominal values just prior to the blade pitching into the post stall regime. In a similar manner, we also define the nominal Reynolds numbers: ( R  U  )c (  1)U  c (5) Ret .u  



and

Ret .d 

( R  U  )c







(  1)U  c



(6)

to estimate of the range of Reynolds numbers encountered by the blades. Initial experiments were carried out with the plasma actuators configured to control dynamic stall on the upwind half of the turbine. Experiments were performed at tunnel speeds in the range 4.0m/sU7m/s that are American Institute of Aeronautics and Astronautics 6

representative of commonly encountered wind speeds. Limited experiments were also performed with the actuators configured to control dynamic stall on the downwind half of the turbine. For upwind control, two main types of experiments were performed: (a) parametric studies; and (b) turbine performance evaluations. Parametric studies were performed by setting a particular tunnel speed and dynamometer load and then varying the plasma control parameters systematically. The parameters were selected based on corresponding parameters known to control separation on corresponding static airfoils. Overall turbine performance was assessed by fixing a particular tunnel velocity and then varying the load gradually from the fully unloaded condition until the turbine stopped. The identical procedure was conducted for the baseline (uncontrolled) case and controlled case with a particular parameter combination. These data were augmented with static tests and surface flow visualization.

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IV.

Results

A.

Baseline Turbine and the Effect of Slip-Rings Prior to evaluating the effect of plasma actuation, baseline turbine performance data was acquired as shown in fig. 5a and 5b. The figures show the baseline power coefficient ( CP  Pt / qU  HD ) and torque without slip-rings as well as the effect of adding the upper end-type slip-ring and both the upper end-type and the lower through-bore slip-rings. The relatively small  range shown in fig. 5a is typical for a turbine with high solidity such as this (Nc/R=1.2). The power coefficients measured here are expected to be greater than the actual turbine values due to the blockage effect of the wind tunnel walls. The data was not corrected for this because the primary objective here was to assess the relative change between baseline and controlled scenarios (see sections 3.2 and 3.3). The figures illustrate the significant effect of torsional slip-rings friction: the upper slip-ring reduced the torque by approximately 0.04Nm while the combination produced almost no power at U=4.3m/s. An increase in tunnel velocity to U=4.9 was required to produce torque values comparable to those obtained with one slip-ring. Hence it was concluded that the total friction associated with both slip-rings together was approximately 0.14Nm. For these small-scale wind tunnel tests the slip-rings have a significant effect. However, up-scaling the turbine dramatically reduces this effect as discussed in section 3.4 below. 0.35

0.16

CP

T (Nm)

0.3

0.14

0.25

0.12 0.1

0.2 0.08

0.15 0.06

0.1 0.04

0.05

0

both slip-rings, U=4.9m/s upper slip-ring, U=4.3m/s no slip ring, U=4.4m/s

both slip-rings, U=4.9m/s 0.02

upper slip ring, U=4.3m/s no slip-rings, U=4.4m/s

0

20 30 40  (rad/s) 50 1.0 1.5 2.0 2.5 R/U 3.0 (a) (b) Fig. 5 Baseline performance of the turbine showing the effect of slip-rings on performance: (a) power coefficient versus ; (b) turbine mean torque versus rotational speed.

B.

Parametric Study Following the baseline evaluation of the turbine, a parametric study was performed for upwind stall control to assess turbine performance improvements subject to plasma parameters duty cycle and pulsation frequency as well as wind tunnel velocity. Figs. 6a and 6b show the transient turbine rotational speed when subjected to plasma American Institute of Aeronautics and Astronautics 7

pulsations for different duty cycles. These data are typical of all parametric studies conducted here. Two groups of data are presented: one for U=4.9m/s (fig. 6a) and one for 7.0m/s (fig. 6b). In all cases, data was acquired initially for baseline condition until t=10s and at t>10s, the plasma actuators are initiated. In all cases shown here, a frequency fe=200Hz was selected. This corresponded to the range of reduced frequencies 1.3  Ft ,u  2.1 observed to

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increase lift on thick non-symmetric airfoils at Re100,000 [25.26]. Following initiation of the actuation, the turbine rotational speed increases and eventually settles to a higher rotational speed. With the dynamometer load constant, the measured torque remains unchanged, and hence the gross increase in turbine power is directly proportional to the increase in rotational speed. For each data set, it can be seen that, although the variation with time is somewhat different, the duty cycle does not have a significant effect on the final rotational speed. This result is consistent with static airfoil data [25,26] and is significant because it means that the same increase in turbine power can be achieved with much less power to drive the actuators, thereby increasing the net turbine power. The effect of this on overall performance will be discussed below. 420

620

 (rpm)

U=4.9m/s

 (rpm)

400

600

380

580

1% 200 Hz 2% 200 Hz 5% 200 Hz 10% 200 HZ 20% 200 Hz 50% 200 Hz

360

1% 200 Hz 2% 200 Hz 5% 200 Hz 10% 200 HZ 20% 200 Hz

560

plasma initiated 340

U=7.0m/s

plasma initiated 540

0 40 80 0 40 80 time (s) 120 time (s) 120 (b) (a) Fig. 6 Transient turbine rotational speed resulting from initiation of plasma actuation for different duty cycles at two wind tunnel velocities. The reduced frequency range is: 1.3  Ft ,u  2.1 .

It can further be seen by comparing figs. 6a and 6b that the “rise time” is longer at 4.9m/s than at 7.0m/s: typically between 40s-80s and 20s-40s, respectively. This is due to the larger loads resulting from the combination of higher tunnel speed and rotational speed. For each run, sufficient time was allowed for stabilization of the turbine and, following that, long time averages (approximately 30s) of the data were acquired. The turbine response to plasma actuation exhibited substantial hysteresis. This can be seen by comparing the difference in the initial baseline speeds. Each of these “initial conditions” followed termination of actuation from the previous experiment, followed by slowing-down and stabilization of the rotational speed. To account for this hysteresis, the changes in turbine power were calculated from the changes in turbine rotational speed, namely Pt=T , based on the initial and final conditions for each experiment. Examples of turbine power coefficient changes as a function of duty cycle are shown in figs. 7 for 4.9, 5.9 and 7.0 m/s. As expected from the transient data shown in figs. 6a and 6b, large variations in duty cycle have virtually no effect on the turbine power coefficient. If we consider only the difference between the final and initial conditions of the turbine power, we can account partially for hysteresis. The differences in the changes to the power coefficient at the different speeds shown in fig. 7 result from the different  of the turbine. For example, the U=5.9 data set has an initial =1.7, which is lower than the others, and hence the blades are stalled for a greater fraction of the azimuth. Consequently, when the plasma is initiated under conditions where the boundary layer is separated for a greater fraction of the azimuth, the percentage increase in power is the largest.

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25

∆CP(%) 20 U=4.9m/s,l=1.9 U=5.9m/s,l=1.7

15

U=7.0m/s,l=2.1

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10

5

0 1

10

100 DC (%)

Fig. 7 Effect of duty cycle on changes to the power coefficient for different wind tunnel velocities and the corresponding range of reduced frequencies 1.3  Ft ,u  2.1 .

16.0

CP(%) 12.0

8.0 U=4.8m/s, DC=10% 4.0

U=5.0m/s, DC=5% U=7.1m/s, DC=5%, increasing fp U=7.1m/s, DC=5%

0.0 0

1

2

3

+

F

t,u

4

5

Fig. 8 Effect of reduced frequency on changes to the power coefficient for different wind tunnel velocities. Transient data was acquired as a function of frequency in the same manner as that acquired as a function of duty cycle described above. Percentage changes in power coefficient for data acquired at different wind speeds and  are shown in fig. 8 as a function of Ft ,u . Three data sets represent individual experiments, while the one U=7.1m/s set was acquired by increasing the pulsation frequency fp without allowing the turbine to return to its baseline state. Comparing the two data sets at U=7.1m/s shows another aspect of the turbines hysteresis: higher power coefficients American Institute of Aeronautics and Astronautics 9

can be obtained by successively increasing the pulsation frequencies rather than running the turbine each time from its baseline state. One puzzling aspect of the turbine’s frequency dependence is that its performance increases continuously with increasing frequency. Similar observations were also made at DC=50%. This observation is totally at odds with data acquired on static airfoils at post-stall angles-of-attack under steady free-stream conditions where peak lift coefficient changes are typically in the reduced frequency range 0.5 to 1. It should also be noted that our convenient definition of Ft ,u will under-predict the actual value on the blade for 0°>1. However, in that study the airfoil oscillation amplitudes were small; the turbine case is significantly different. Consider a typical case where Ft ,u  2 and =2. By substitution of dimensionless groups, we can write the ratio:

2 f p

 1 R  (9) F   c t ,u Under these conditions 2fp/=31 which would appear to fulfill the condition suggested by Greenblatt et al  2

[27]. However, on the turbine this translates to one perturbation every 11.6 degrees on the azimuth. The lack of an enforced phase relationship between the forcing and the turbine rotation means that the perturbations are introduced at arbitrary angles, possibly even after the flow has already begun separating from the surface. By increasing the perturbation frequency the differences in the azimuthal angles is smaller and this causes the perturbations to be introduced closer to onset of dynamic stall. Hence, although the higher frequencies are not the optimum for performance improvements, the fact that they are introduced more consistently closer to dynamic stall onset renders them more effective. Phase-locking the perturbations with the turbine rotation has the potential to overcome this limitation. Moreover, designing a turbine where the optimum reduced frequencies correspond to 2fp/ >O(102) would clearly also have potential to improve performance. Effects of scaling-up the turbine on the frequency ratio parameter described in section 3.4.

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C.

Turbine Performance Evaluation Fig. 11a shows the overall turbine performance for two wind speeds, U=4.9 m/s and 5.9m/s, for both the baseline turbine as well as upwind stall control in the range 4.4  Ft ,u  5.6 . For each tunnel speed, plasma produces an increase in performance over almost the entire operational range. The peak power coefficients are increased by 38% and 31% respectively. The baseline turbine performance is strongly dependent on the tunnel speed and there are two main reasons for this. The first is the turbine torque produced relative to the friction in the bearings and sliprings and the second is the different Reynolds numbers encountered by the blades.

0.25 4.4=