CFD (Computational Fluid Dynamics) analysis of the temporal response of the microtab show that it has two disadvantages compared with a trailing edge flap, when considering it for active load control of a wind turbine. First of all the microtab has a reverse control phenomenon in the initial phase of deployment, and second of all the response is delayed, compared with the flap, due to buildup of flow structures. This motivated us to find an engineering model for the temporal response of the microtab to investigate the behavior of the microtab in an aeroelastic model of a wind turbine. Before we attempted to model the microtab in detail, we inserted temporal response of the microtab found with CFD directly into the aeroelastic code FLEX5. We found that the reverse control was not important for the response of the blade. We propose to use a simple queue delay to model the flow physics of the microtab. Finally, we calculated the load reduction potential of the microtab, assuming that we could model it as a delay, and compared it to that of a trailing edge flap. In our calculation the trailing edge flap had a load reduction potential twice as big as the microtab solution on the UPWIND reference turbine.
Nomenclature Cl = lift force coefficient [1] ΔClx = range of change in Cl, of x at given angle of attack [1] Cd = drag coefficient [1] Mx = Bending moment of the x component [Nm] U = wind speed [m/s] aoa = Angle of attack [deg] c = chord [m] t = time [s] τ = non-dimensional time, τ =t·U/c [1]
I. Introduction CTIVE load control for wind turbines is becoming increasingly important for the wind turbines, due to their increasing size and complexity. Especially for very large turbines, distributed active load control on the blades
may be useful for reducing the fatigue loads of the blades and the turbine. We use the term active aerodynamic device, AAD, for a device situated on a region of the blade, which can be controlled to modify the local aerodynamic forces. In this article two possible devices are considered
The trailing edge flap, which is equivalent to the aileron of a plane. The trailing edge flap has been extensively studied at Risø-DTU [1][2][3]. Recent studies have shown high load reduction potential for this technology [4]. In this article only the rigid, hinged trailing edge flap is considered.
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The microtab (MT) is a small tab or plate emerging from the surface of the airfoil near the trailing edge. It is similar to a Gurney flap, except that it is retractable. A MT deployed on the pressure side of the airfoil will increase lift, while a MT deployed on the suction side will decrease lift. The proper sizing and placement of the Microtab has been investigated in several studies [5][6][7].The microtab has the advantage of being very small and hence potentially require very little actuation power. Due to the small size, a fast actuation is expected. However, as we will see in the next section, the microtab interacts with the flow in some undesirable manners during deployment. First of all when deploying the microtab, the initial build up of lift, is in the opposite direction as the steady state value of the deployed configuration [5] [6] [7]. This may confuse a simple controller. Furthermore, the time to build up the lift is slower than that of the flap, and this may reduce the performance of the controller.
Before attempting to create a detailed engineering model for the temporal response of the microtab two questions
were raised. 1. Is the delay of any significance for the control? 2. Is the reverse lift phenomenon a problem when controlling the device on the turbine?
This paper will try to shed light on the questions, and furthermore give an indication of the load control potential
of the two devices. The reduction potential will of course also depend on other parameters, such as the controller. In this study a very simple controller is used.
II. CFD Study of the Microtab
A. The temporal response of a microtab deployment modeled with CFD Although the actuation speed is high, several CFD studies [5][6][7] has indicated that the microtab has an initial
delay due to a build-up of the flow structures near the microtab. If the actuation time is in the same order of magnitude as the time it takes the flow to travel one chord length, τ=1 c/U, then the delay in the lift force increase is also in the order of τ=1 c/U. An example could be to consider a blade section at 50m from root on the UPWIND 5MW Reference Turbine [8] at rated wind speed: (2.76m) /(62.5m/s) =44 ms. Then the delay between a trailing edge flap with the same deployment speed and the microtab is approximately 44 ms. In Figure 1, the temporal response of the microtab is compared to that of the flap, both cases calculated using CFD [7].
Furthermore in the very initial phase of the microtab deployment, the force is opposite of the desired for a short
while. The two devices have almost the same steady state lift value in the deployed state, which makes it apparent that the microtab lift increment is delayed a non-dimensional time Δτ≈1.
a) b)
Figure 1: Left plot (a) Cl as function of non-dimensional time for the microtab (blue) vs. a hinged flap (red). Right (b) is the Cd as function of time. From [7].
Δτ=1
Reverse lift response
Non-dimensional time τ=t·U/c Non-dimensional time τ=t·U/c
Cd
Cl
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B. Steady State Data The steady state polars for the two devices that we are considering, have been calculated using Ellipsys2D
[9][10][11]. We used the DU-96-W-180 profile, because it is placed on the outer region of the UPWIND 5MW Reference Turbine, which will be addressed later in the paper.
The grids were generated using Hypgrid2D [12], and were O-type grids, extending 20 chords from the airfoil. For the microtab case, the grid size was 640x128 and for the trailing edge flap, the grid size was 512x128. For modeling the boundary layer we used the γ-Reθ correlation based transition model by Menter [13], which was implemented by Sørensen [14] with a k-ω SST turbulence model. The grids are shown in Figure 2.
The microtab results that have been calculated using CFD, were validated against wind tunnel measurements, from the LM Wind Power Low Speed Wind Tunnel (LSWT). The LSWT was specifically designed for validation of airfoil sections under flow conditions very close those that are present on large wind turbine blades [15]. Unfortunately, no wind tunnel measurements were available for the rigid trailing edge flap. The CFD results and the experimental results are shown in Figure 3. The bandwidth between the upper and the lower lift curve for a given device is the ΔCl. On the turbine, the design lift angle is approximately 5 degrees, and hence the ΔCl near this angle is an important parameter for the performance of the device.
The CFD results fit very well with experiments for the case where the microtab was deployed on the pressure side, and for the undeployed (clean/normal) airfoil. For the case were the microtab was deployed on the suction side, the decrease in lift is slightly under-estimated in the CFD results. The lift decrease generated by the microtab on the suction side, is very sensitive to the state of the boundary layer. Hence, a small difference between the boundary layer that was modeled in the CFD, and the one that was in the experiment, will give this difference in the final result.
a. b.
c. d.
Figure 2: Grids for the DU96-W-180 profile. Left (a & c) are the grids for the microtab when deployed on the pressure side. The microtab was 1% of the chord in height, and was placed 5% of the chord from the trailing edge. The grid was 640x128 and extended 20 chord from the airfoil. Right colomn (b & d) show the grid for the flap, deflected 10 deg., with a hingepoint at 10% of the chord from the trailing edge. Grid size was 512x128, and extended 20 chords from the airfoil.
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Figure 3: The steady state polars for the microtab and flap on the DU96-W-180 profile. The CFD results for the microtab, can be compared with wind tunnel measurements. Reynolds number was 3 million for all cases.
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III. Implication on wind turbine dynamics
A. FLEX5 with Active Aerodynamic Devices FLEX5 is a widely used aeroelastic wind turbine simulation tool that works in the time domain. It was developed
by Stig Øye at the Technical University of Denmark. For more information, see [16]. In the present work, FLEX5 has been modified by adding the following features:
One or more AAD elements can be added to each blade. An element extends a certain span of the blade, e.g. a flap centered at 50m, with a width of 7 m, will extend from 46.5 m to 53.5 m from the root.
The aerodynamics of the element is determined by an input file, which contains the polars for different steady state settings of the AAD element. Linear interpolation is performed between states.
Each of the AAD elements has an independent P(I) controller. The input signal for the controller is a flapwise bending moment at a particular radial position on the blade. This could for example come from one or more strain gauge sensors. The reference signal for the controller is the same flapwise bending moment as the input signal, but low pass filtered so that the reference signal changes only slowly over time when the operational state of the turbine changes. The controller then acts on the difference between the input signal and the reference signal multiplied with a proportional gain, P. The gain and the time constant for the reference signal is tuned using an optimization routine, that finds the optimal gains to give the lowest flapwise root bending moment. No gain scheduling was applied, so the gain was the same for all operational states of the turbine. The control strategy is very similar to that applied in [4]. The control output, Q, is a number between -1 and 1, where
o Q=-1 corresponds to the polar in the set with the lowest lift, o Q=0, corresponds to the original, undeployed, polar. o Q=1, corresponds to the polar with the highest lift. o Otherwise, the polar is found by linear interpolation of the polars, based on Q value.
The effect of shed vorticity is modeled by superposition of a series of linear first order step responses, as described by Karman & Sears [17]. For the step responses, the flat plate approximation of Jones [18] was used. The original dynamic stall model of FLEX5 was kept as is, which means that the stall effects are calculated on the undeformed polar, and the effect of the AAD element is added afterwards, only considering the quasi-steady separation which is in the input file of the AAD element.
B. A simulated step response To investigate if the reverse control phenomena of the microtab had any importance for the dynamics of the
turbine a very simple case was set up, where the UPWIND 5MW reference turbine was modeled, except all turbulence and periodic effects such as tower shadow, wind shear and gravity were removed. Tilt and cone angles for the rotor were set to zero, to remove further dynamic effects. The controller of the AAD element, which was placed at 50m from the root, and had a width of 7m, was exchanged with a simple table lookup for the force on the AAD element. Then two cases were simulated using table lookup from:
1. The microtab response for the lift, from Figure 1. 2. A simple step response, using Jones approximation for a flat plate. The step was delayed by 40ms.
The results are shown Figure 4. For an entirely stiff turbine, the effect of the reverse lift phenomena of the
microtab, is clearly seen in the flapwise moment, because the aerodynamic forces are dominating the flapwise moment. For the stiff turbine, the response of the structure has no delay to the aerodynamic forces. For the flexible blade, however, the inertial forces of the blade dominate the response, and hence the reverse lift phenomenon is hardly noticed by the structure.
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IV. Fatigue Load Reduction Potential of Flaps and Microtabs
A. The UPWIND 5MW Reference Turbine In the following we have modeled the Upwind 5MW Reference Turbine using FLEX5. The foundation was stiff
at ground level, i.e. no substructure was modelled. The turbine has been modeled using the IEC Ed. 3, Wind Class 2A. Only the design load case 1.1 (normal operation, normal turbulence) have been considered. For every 2 m/s from a wind speed of 5 m/s to 25 m/s, 4 x 10 min time series were generated and analysed for fatigue. The fatigue loads were calculated by the standard procedure of IEC Ed.3, i.e. by rainflow counting the loads for each sensor for each wind speed. The damage from each wind speed was then weighted with Weibull distribution corresponding to the wind class. The fatigue load given is the damage equivalent load of the given sensor for a 20-year lifetime. For the blade loads a Wöhler-coefficient of m=10 was used, but for the tower bottom, yaw and tilt moment, a value of m=3 was used.
The following settings were used: The AAD element was centered at 50 m from the root of the blade, and had a 7 m extension. The rotation speed was fixed for each wind speed to the speed that was given by the generator
characteristics, and the torsional degree of freedom on the shaft was locked. Otherwise, all degrees of freedom were free.
The original PI controller for the collective pitch system was used. The proportional gain of the AAD controller was set to a high value of 4e-6, and the flapwise moment
sensor signal was sampled at a location 25 m from the blade root. The time constant for the reference signal was 6 s, which is close to 1P. A high gain means that the controller respons very rapidly to changes in the input signal, which generally gives a high fatigue load reduction. The disadvantage of using a high gain is the increased activity of the devices, which might cause excessive wear on a real turbine.
B. Investigation of the influence of AAD element ΔCl on the obtainable fatigue load reduction The ΔCl range of the AAD elements was varied from 0 to 0.8 (±0.4 from the undeployed state). ΔCd of the
elements was not changed, i.e. there was no additional drag, due to element actuation. The load reductions have been plotted in Figure 5. As the objective of the controller was to reduce the flapwise fluctuations of the blade, the fatigue load of the flapwise moment reduced when the ΔCl was increased. The only load which was seen to increase was the edgewise fatigue load. This was believed to be due to the coupling between the lift force and the edgewise
a) The response when the entire turbine is stiff. b). The response of a flexible blade on a stiff turbine.
Figure 4: The response of the microtab (red curve) and a step response (green). The step response is delayed by 40ms, compared to when the microtab movement begins.
Time, s
Mfla
p, k
Nm
Time, s
Mfla
p, k
Nm
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direction. When the lift force oscillates more, then the edgewise mode is excited as well. A ΔCl of 0.8 is also a very high value, because it requires a quite large flap of approximately 20% of the chord to attain.
C. Fatigue loads of a microtab versus a flap It was found that the ΔCl of a given device matters a lot for the load reduction potential. The polar sets for the
two devices, which are shown in Figure 3, were used on the UPWIND Reference Turbine.
Four cases were run using: 1. Flap polar set, found with CFD, and a normal controller. 2. Microtab polar set, from LSWT. The LSWT data was chosen, because the ΔCl was higher than for the
CFD data. The controller was run in a normal manner as in case 1. 3. Microtab polar set, from LSWT. The controller was then run in an on-off manner, where the controller
output signal was binned, in the following way a. if the normal controller value was between -0.5 and -1, the microtab was deployed fully on the
suction side. (low lift) b. If the normal controller value was between -0.5 and 0.5, the microtab was in the undeployed
state. c. If the normal controller value was between 0.5 and 1, then the microtab was fully deployed on
the pressure side (high lift.) 4. Microtab polar set from LSWT. Binned as in case 3, but also delayed in a queue for 40 ms, to simulate
the delay in the flow near the microtab. This is the most realistic model of the microtab we have used. The load reduction found for each case is showed in Figure 6. The flap polar set was found to give a much larger
load reduction than the microtab set, although they were run in the same way. This is believed to be mainly due to the lower ΔCl at the high angles of attack, where the AAD element needs to lower the lift. An investigation was made on the angle of attack that the section of the AAD element was experiencing. The histogram is showed in Figure 7, and although it was only a small fraction of the time, that the section was at a high angle of attack, these high loads are very significant for the flapwise fatigue loads and loads in general.
Looking at the microtab results alone, it was clear that binning the controller output (on-off mode) did not matter very much for the results. When we put the final feature of the microtab, namely the delay, onto the binned microtab controller, the load reduction potential fell slightly from -4.73% to -4.58% on the flapwise bending moment fatigue load. This was a clear indication that modeling the transient behavior of microtab in detail is not important for doing aeroservoelastic computations.
Figure 5: Fatigue load change, with varying freedom of the AAD elements. Calculated with FLEX5.
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Figure 7: A histogram of the angle of attack, for the station at 50 m from the blade root. The wind speed was 11 m/s, which was rated wind speed where the mean angle of attack was expected to be highest.
Figure 6: Load reduction potential for cases 1 to 4. 1. The flap with a normal controller. 2. The microtab, MT, with the normal controller. 3. The microtab which has been binned in 3 states – full suction side, full pressure side, and undeployed. 4. Same as case 3, but a delay of 40 ms was added to the binned controller for the microtab. Case 4 was the most realistic model for the Microtab.
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V. Conclusion Aeroservoelastic simulations on the UPWIND Reference Wind Turbine with active aerodynamic devices, was
conducted with unsteady aerodynamic models for the microtab and the hinged trailing edge flap. The model was based on a set of polars that define the ΔCl of a given device. The two devices had slightly different polar sets. The clear distinction between the two devices was that the flap was controlled in a continuous manner with instant aerodynamic response, and the microtab was modeled with the same controller as the flap, except it was binned, so the microtabs were in on-off states and a queue delay of the control-signal was used to model the aerodynamic lag of the microtab.
The key findings were The reverse control of the microtab, seen in CFD computations, does not matter for the structural
response. This is probably due to the very short time in which the reverse lift occurs, compared with the response time of the structure.
On the modeled UPWIND Reference Turbine, the flap had a load reduction potential more than twice as big as the microtab. This was believed to because the microtabs had a low lift reduction at high angles of attack, were the suction side microtab is in the separated zone of the profile.
Hardly any difference was seen in the load reduction potential of the microtab, when comparing a continuous controller with a binned controller. This could partly be explained by the high gain of the controller.
When comparing the binned controller of the microtab with the binned, delayed microtab controller, a small difference in the load reduction potential was seen. However, the difference was not significant compared to the difference that was seen between the polar sets of the microtab and flap respectively. Therefore, there is no reason to believe that it is important to model the transient behavior of the microtab deployment in more detail.
References
[1] Basualdo, S. Load Alleviation on Wind Turbines using Variable Airfoil Geometry, M.Sc. Thesis Project, Technical University of Denmark, 2004.
[2] Troldborg, N. Computational study of the Risø-B1-18 airfoil with a hinged flap providing variable trailing edge geometry. Wind Engineering, vol. 29, no. 2, 2005
[3] Buhl, T.; Gaunaa, M.; Bak, C.; Potential Load Reduction Using Airfoils with Variable Trailing Edge Geometry. Journal of Solar Energy Engineering. November 2005, Vol. 127, p. 503-516
[4] Andersen P. B et al. Deformable trailing edge flaps for modern megawatt wind turbine controllers using strain gauge sensors. Wind Energy 2010, 13. DOI: 10.1002/we.371
[5] Baker, J. P., Standish, K. J. and van Dam C. P., Two-Dimensional Wind Tunnel and Computational Investigation of a Microtab Modified S809 Airfoil, AIAA 2005-1186, 2005.
[6] Chow, R. and van Dam, C. P., On the temporal response of active load control devices, Wind Energy. 2010, 13. DOI: 10.1002/we.370
[7] Baek, P, Gaunaa, M., Sørensen, N.N, Fuglsang, P. Comparative Study of Distributed Active Load Control Concepts for Wind Turbine Blades, The Science of Making Torque from Wind Conference, Crete 2010.
[8] Jonkman J. NREL 5 MW baseline wind turbine. Technical Report, NREL/NWTC, 1617 Cole Boulevard; Golden, CO 80401-3393, 2005.
[9] Michelsen, J.A. Block structured Multigrid solution of 2D and 3D elliptic PDEs, Technical Report AFM 94-06, Technical University of Denmark, 1994.
[10] Michelsen, J. A. Basis3D - a Platform for Development of Multiblock PDE Solvers. Technical Report AFM 92-05, Technical University of Denmark, 1992.
[11] Sørensen, N. N. General Purpose Flow Solver Applied to Flow over Hills. Risø-R- 827-(EN), Risø National Laboratory, Roskilde, Denmark, June 1995.
[12] Sørensen N. N. HypGrid2D a 2-D Mesh Generator. Risø-R-1035 (EN), Risø National Laboratory, Roskilde, Denmark, Feb 1998.
[13] Menter F. R. et. al. A Correlation-Based Transition Model Using Local Variables, Part I - Model Formulation. In Proceedings of ASME Turbo Expo 2004, Power for Land, Sea, and Air , Vienna, Austria, June 14-17 2004. ASME. GT2004-53452.
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[14] Sørensen, N. N., CFD modelling of laminar-turbulent transition for airfoils and rotors using the gamma-(Re)over-tilde (theta) model, in journal: Wind Energy (ISSN: 1095-4244) (DOI: 10.1002/we.325) , vol: 12, issue: 8, pages: 715-733, 2009
[15] Fuglsang, P and Bove, S. Wind tunnel testing of airfoils involves more than just wall corrections. In EWEC 2008, Brussel, 2008
[16] Øye, S. FLEX4 - Simulation of Wind Turbine Dynamics. Proceedings of the 28th IEA Meeting of Experts “State of the Art of Aeroelastic Codes for Wind Turbine Calculations”, pp. 71-76, Technical University of Denmark, Lyngby, Denmark, April 11-12, 1996
[17] von Karman, T. H., Sears, W. R. Airfoil theory for non-uniform motion. Journal of the Aerodynamical Science 1938; 5
[18] Jones, R.T. The unsteady lift of a wing of finite aspect ratio. Tech. Rep. 681, NACA Report, 1940.
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