1 Copyright © 2011 by ASME

Proceedings of the ASME 2011 Conference on Smart Materials, Adaptive Structures and Intelligent Systems SMASIS2011

September 18-21, 2011, Scottsdale, Arizona, USA

SMASIS2011-4940 TECHNOLOGICAL SOLUTIONS FOR REALISING AN ACTIVE GURNEY FLAP FOR GREEN ROTORCRAFT APPLICATIONS, BASED ON PIEZOELECTRICS AND ELECTROMAGNETS.

Gianluca Diodati Monica Ciminello Antonio Concilio

Italian Aerospace Research Centre (CIRA) Via Maiorise, Capua (CE), Italy

email: g.diodati@cira.it; monica.ciminello@gmail.com; a.concilio@cira.it

ABSTRACT

Gurney flaps are small devices whose dimensions are usually percentage units of the airfoil chord. Liebeck, first hypothesized that the Gurney flap effectively could change the flow field in the region of the trailing edge by introducing two contra-rotating vortices behind the flap, therein altering Kutta’s condition and circulation. In turn, this helps avoiding flow separation and thus allowing the achievement of higher angles of attack before the stall condition occurs. This paper presents preliminary investigations performed on different, original and feasible technological solutions, addressed at realizing Gurney flaps for rotorcraft smart blades.

INTRODUCTION Gurney flap concept was first introduced by Liebeck [1] in 1978, providing wind tunnel results for an airfoil with and without a Gurney flap, 1.25% chord height, he showed increased lift and decreased drag, compared to the baseline airfoil. Liebeck also noted that increased flap height over 2% c, while leading to lift benefit, could lead to significant rises in drag; he then concluded that the optimal Gurney flap height was to be fixed on the order of 1-2%c. In [2], a microtab-based load control system was investigated. This system involved small tabs near the trailing edge of the rotor blade, where these tabs deployed approximately normal to the skin. This movement effectively changed the sectional camber of the rotor blade and the trailing-edge flow conditions, thereby changing aerodynamic characteristics of the blade. Microtab concept advantages are its low actuation power requirements, short actuation time, simplicity, and the fact that

it requires minimal changes in the way blades are manufactured. In [3], Yen et al. tested the load control capabilities of the microtab system, remotely controlled, for a fixed wing in a wind tunnel. Tests demonstrated both microtab robustness and lift enhancement effects. In [4], a variable droop leading edge airfoil was successfully used to control compressible dynamic stall and its associated adverse pitching moment variations. The price of this success was however a 10% loss of lift. A Gurney flap was then attached to the airfoil trailing edge lower skin to recover this “lost-lift”., A devoted study addressed the optimum flap height all along the retreating blade region of interest since this kind of devices is seldom used in compressible flow environments. Analysis of airfoil relative performance data with three separate Gurney flaps (1 to 3% c height) confirmed that the best performance was attained with a flap height set to 1%-chord. In [5], various technologies using commercial piezoelectric stack and patch devices were considered, and the respective performance were presented. In the following table, retrieved from [5], weight and power required per meter of wingspan and for the whole blade (6.55 m) span are presented:

The linear actuator solution was found to be much lighter than the stack actuator alternative. The latter devices however, has

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been shown [6] to have a better bandwidth, allowing both a full Gurney flaps deployment along the blade rotation and the implementation of active vibration reduction in a range up to 4 x rev. The problem of designing a Gurney flap system must take at least into account the aircraft type, reference flight conditions and airfoil. Several solutions for realizing a Gurney flap are presented in literature. Among the available actuators, however, an important selection may be done by considering bandwidth and weight characteristics Inside JTI GRC1 “Innovative Rotor Blade” project, Active Gurney Flap (AGF) was selected as a good trade-off between relatively low technological complexity and efficacy in improving blade performance in the retreating region. In this paper, different technological solutions for realizing an Active Gurney Flap are reported. Firstly, main requirements are collected and discussed, including needed mechanical energy, bandwidth, allowable weight and space, ability to sustain the centrifugal forces, etc. Then, smart material-based devices are selected, referring to electromagnets and piezoelectric stack. Merit figures considered: • required mechanical energy to actuate the flap in the worst aerodynamic case. • actuator bandwidth so large to allow a full actuation cycle within one blade revolution. • minimum weight and space requirements; • capability of sustaining expected centrifugal forces.

NOMENCLATURE AGF Active Gurney Flap NACA National Advisory Committee for

Aeronautics FEMM Finite Element Method Magnetics CS-JU CleanSky – Joint Undertaking GRC-ITD Green RotorCraft – Integrated Technology

Demonstrator RPM Revolutions Per Minute JTI Joint Technology Initiative bn Number of blades R Blade radius Vtip Tip speed c Blade chord δ Active gurney flap deployment δmax Maximum active gurney flap deployment L Beam length b Beam width ts Beam thickness tp Piezo thickness a Distance of piezo stack from the clamp y_stack Piezo stack stroke P Piezo stack actuation force h_stack Piezo stack height

Es Beam Young modulus Is Beam inertia momentum As Beam cross section area ρs Beam density m Beam mass per unit length mb Beam overall mass Y(x,t) Beam displacement q(t) Modal coordinate y(x) Beam modal displacement Eel electric field d31 piezoelectric reverse effect constant νs structure poisson ratio νp piezo poisson ratio ε strain

WORKING HYPOTHESYS

Rotor Definition Main rotor major characteristics, considered in the

herein reported investigation, are following summarized: • Number of blades: bn = 4 • Radius: R, about 8 m • Tip speed: Vtip, about 200.0 m/s

A centrifugal force of about 70g per blade span meter then resulted, giving rise to an acceleration of about 560g at the blade tip.

Geometrical requirements In a rotor blade, thicker profiles are generally used at

the root and thinner at the tip. Because stall control phenomena arises in the root region, only root profiles are referred to. In the current application, NACA23012 was selected (see Figure 1).

Figure 1: Profile thickness

As reported in [7], the best AGF position is at the

trailing edge; performance deteriorates as it is moved upstream. Good results can be however achieved up to 90% c, while no significant performance loss is shown up to 95% c.

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References [8] to [10] demonstrated that among a range between 1 and 2.5% c height devices, the shorter is the one showing the best performance.

In this work, following some preliminary aerodynamic investigations addresses the study towards a 2% c height AGF.

Dynamic requirements The law regulating the height of the actuator is

fundamental in order to identify the dynamic characteristics of the actuation system. Such a deployment law can have a shape as the one reported in Figure 2.

According to that, the AGF shall pass from a fully retracted to a fully extended state position in about 10° to 15° azimuth angle, corresponding to a period of about 10 ms to the referred angular speed.

It means a mean deployment velocity of about 1.5 m/s, and a corresponding acceleration of about 150 m/s2. The actuation frequency may be instead fixed at 4 Hz (once per round)

Figure 2: Tip deployment law

Deployment law A preliminary trapezoidal deployment law was assumed,

Figure 2. The adopted procedure consisted of modifying the reference signal into its main harmonics. In order to reduce bandwidth requirements, the first ten harmonics were used, with their own amplitudes and phases.

Further considerations did further lead to filter low energy level components. As reported in the normalized energy spectrum, Figure 3, in this specific case, the seventh and the eighth addenda could be removed from the Fourier series. This did not affect the highest component (about 35 Hz).

The transformed law of deployment is shown in Figure 4. It represents the “real” command to be given to the actuator.

Figure 3: Normalized Energy Spectrum

Figure 4: Deployment law reconstruction using the highest

first 10 components (excluding the 7th and the 8th)

TECHNOLOGICAL SOLUTIONS

This section illustrates a series of technological systems investigated from the authors to develop an Active Gurney Flap system. The basic structure adopted for the implementation of the following technological solutions is a cantilever beam with a flap at the tip. This choice relies in the possibility of attaining the best performances in terms of flap deployment with a very simple structure. Thanks to the elasticity of the beam, in fact, such a dynamic can be obtained without any kind of amplification system and avoiding hinge moments and sliding mechanisms giving problems in a centrifugal field encountered on a helicopter rotor blade.

Piezo patches A first approach is based on piezo patches actuators. Piezo

patches were considered to be bonded on both sides of a cantilever beam as reported in Figure 5. As an extreme option, the actuators could cover the whole 1D structural element. In more practical cases, a limited surface should however be considered. The actuation system should take into account some structural dynamics issues, as for instance:

• Resonance working conditions, that are very sensitive in terms of phase stability and motion amplitude;

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• Because magnitude and phase can be thought constant out of resonances, static-like system may be addressed;

• Centrifugal force effects can be considered constant in amplitude, direction and orientation.

Figure 5: Piezo patches architecture

According to these statements, the first beam

eigenfrequencies should be higher than the highest harmonic used for the deployment law reconstruction (35 Hz). Fixing this limit to 70 Hz (so to obtain a quasi-static behavior of the beam in the dynamic range of interest), the maximum length of the beam can be set, [11]:

ss

ss

AIE

Lρπ 702

875.1 22

⋅≤ (1)

Lecce-Concilio 2D strain actuation model, [12], can be

combined with Eq.1 and the other classical equations governing the static displacement field of a cantilever beam, to evaluate the achievable AGF displacement, following the action of a piezo actuator system

( ) ε

ρπν⋅

⎥⎥⎦

⎤

⎢⎢⎣

⎡

⋅−≤

s

s

sAGF

Ey

37041875,1

23 2

(2)

where

( ) ( )elEdttEt

tt3123

2

81266

++++

=ε, S

P

P

S

EE

Eνν

−−

=11

, P

S

tt

t = (3)

Eq.2 sets the max displacement as a function of the

transmitted strain, once the material characteristics are fixed. It is then possible to identify an optimal ratio between structure and piezo thickness. The following figures of merit are then obtained for different materials and reported in Figure 6. Resin-like materials resulted then to give the higher deployment value (herein intended as the sole merit figure). Nevertheless, the max predicted displacement (3.5 mm after an applied electric field of 1000 V/mm) is just partially matching the specifics (2%c AGF deployment, equivalent to about 10 mm).

A

B

C

D Figure 6: Comparison of strain actuation levels on Epoxy

(A), Wooden (B), Aluminium (C) and Steel (D) beam

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Piezo stack A second technological solution involves piezoelectric

stack actuators, able to generate very large forces but small deformations. Placing the actuator next to the clamp, large tip displacement may be assured (Figure 7).

Figure 7: Piezo stack architecture

Through the maximization of the work (dW/dy = 0 with

W=P*y), and imposing the condition of impedance matching, it is possible to determine the optimal operating point which corresponds to the maximum displacement (ref. [8]). This point appears to have the coordinates (ystack / 2, P / 2).

To proper express the tip displacement (in this architecture, coincident with the AGF envisaged motion), the equations are referred to, relative to a cantilever beam under a concentrated load. The already mentioned Eq. 1 is used to determine the upper limit of beam length. AGF deployment can be then expressed as a product of a term depending only on the beam physical and geometrical characteristics by a term depending only on the actuator performances:

( ) ( )3/23/13/2

2

3702

875.12

121

stackssss

ss

ss

yPIEAIE

IE ⎥⎥⎦

⎤

⎢⎢⎣

⎡

⋅≤

ρπδ

(4) Technical data of commercial piezo stacks (produced by

Physik Instrumente, series P-0XX PICA Power Actuator, [14]) were considered in order to select the best feasible solution. For each stack type, max stroke and actuation force are reported in Table 1:

Stack Type

Stroke (μm)

Force (N)

3/23/1stackyP

Height (mm)

P-010.40 60 2200 0.0199 56 P-010.80 120 2400 0.0326 107 P-016.40 60 6000 0.0278 54 P-016.80 120 6500 0.0454 101 P-016.90 180 6500 0.0595 150

Table 1: PI piezo stack performances

The attainable displacement, Eq.4, could be then evaluated for each stack, applied to beams made of different materials (resin, wood, aluminium, steel). Results are reported in the following figures

A

B

C

D Figure 8: Max tip displacement by a piezo stack action on Wooden (A), Epoxy (B), Aluminium (C), Steel (D) beam

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Considering the following beam dimensions: • Beam width b: 5 mm • Beam thickness ts: 0.5 mm • Beam length L: 72 mm

and using a piezo stack type P-016.90, 24 mm, AGF displacement was predicted. FE simulations were performed to further check the analytical estimations. In Table 2 some weights for the considered piezo stacks have been reported:

Stack Type Weight (gr)

P-010.80 64 P-016.80 150 P-016.90 200

Table 2: Piezo stack weight

Electromagnets A further alternative may be represented by the use of

electromagnets. Preliminary numerical simulations were performed to evaluate the characteristics of the requested electromagnetic driving force.

According to the classical beam theory, its general displacement field may be expressed as:

),,(),(),(2

2

4

4

txYFt

txYmx

txYEI =∂

∂+

∂∂

(5)

the approximated first mode shape is given by the

following expression, [11]:

2)1(3)1()( 3 +−−−=

Lx

Lxxy

(6)

Using Eq.5 and 6, the beam equation for a concentrated force can be written as (Y(x,t)=q(t)*y(x) ):

)(),,()()()()(2

2

4

4

Fss xxtxYFt

tqxmytqx

xyIE −=∂

∂+

∂∂ δ

(7)

By means of a modal projection, Eq.(7) becomes:

∫

∫

∫

∫ −=

∂∂

+∂

∂L

L

F

L

L

ss

dxxym

xxYFxytq

dxxym

dxx

xyxyIE

ttq

0

2

0

0

2

04

4

2

2

)(

)()()()(

)(

)()()(

δ (8)

from which it is obtained:

))(()()( 21 tqFKtqtq ⋅=+ω&& (9)

Where the parameters K and 1ω are:

MxyK F )(

= (10)

MLIE

mmLIE ss

bAGF

ss331

3)24.0(

3=

+=ω

(11)

Applied EM force was calculated by means of FEMM 4.2 open source code, [15]. Matlab ODE45 routine was used to simulate the beam time response, implementing Runge–Kutta algorithm. Results are reported in Figure 9, showing a 21 ms response.

Figure 9: Numerical estimation of AGF deployment time

In order to validate the numerical model, a complete and

simplified test rig, hosting a preliminary active beam prototype was realized. Experimental results are shown in Figure 10.

Deployment Time

-1,4

-1,2

-1

-0,8

-0,6

-0,4

-0,2

0

0,2

0 0,5 1 1,5 2 2,5 3 3,5 4 4,5 5

Time [s]

Nor

mal

ized

Str

ain

Figure 10: Experimental measurement of the electromagnet

AGF deployment time

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Comparison between numerical and experimental results is reported in Table 3. Numerical simulations are implemented with ideal condition in terms of perfect flat surface of the magnet.

In the experimentation, of course, facing with real condition of surface roughness and mechanical losses, measurements were just intended at verifying the capability of attaining the required displacements.

Numerical prediction Experimental measure 21 ms 28 ms

Table 3: Deployment time: num. Vs. exp. results

This last actuation mechanism seems to be then the most promising among those evaluated, in terms of whether weight, power consumption or price. Protrusion time can be further improved by optimizing geometrical and mechanical parameters of the system.

CONCLUSION AND FUTURE STEP

In this document, performance of three different AGF mechanisms was evaluated, referring to an architecture based on a simple cantilever beam.

Piezo patch actuators resulted to be inadequate to produce the bending tip required displacement.

Stack actuators proved to achieve the desired behavior, but power consumption and added weight seemed excessive penalties for actual rotorcraft applications.

The most promising actuation system resulted the one based on the use of electromagnets. At the first early stage of the study, they were measured to produce a 28 ms structural dynamic response, associated with low electrical power consumption (about 5W) and low weight penalty (about 50gr.).

In the next future, a full closed loop control system will be designed and implemented via DSP, implementing the deployment law signal herein described while controlling the dynamic response for the structural element.

Further experimental activities on a refined prototype will be aimed at optimizing the dynamic behavior of the device in terms of protrusion time while considering the effect of centrifugal forces.

ACKNOWLEDGMENTS

This research activity is partially funded by the Clean Sky Joint Undertaking under projects GAM-GRC, related to activities performed within the ITD Green Rotorcraft.

The authors would also like to thank Dr. Salvatore Ameduri for his guide in the very first approach to FEMM4.2 software and his support in the experimental setup realisation.

REFERENCES

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[2] Yen, D. T., van Dam, C. P., Br¨aeuchle, F., Smith, R. L.,

and Collins, S. D., “Active Load Control and Lift Enhancement Using MEM Translational Tabs,” Fluids 2000, No. 2000-2422, AIAA, Denver, CO, June 2000.

[3] Yen-Nakafuji, D. T., van Dam, C. P., Smith, R. L., and

Collins, S. D., “Active Load Control for Airfoils Using Microtabs,” Journal of Solar Energy Engineering, Vol. 123, 2001, pp. 282–289.

[4] Chandrasekhara MS. Compressible dynamic stall control

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[5] A. Paternoster , P. de Jong, R. Loedersloot, A. de Boer,

R. Akkerman, Initial Technology Review, IGOR - Actuator and Control System for Green Rotorcraft I.T.D., University of Twente, Engineering Technology, 2009. http://doc.utwente.nl/69643/1/Piezo_actuators_Gurney_flaps_2-2.pdf

[6] Yung H. Yu, Bernd Gmelin, Wolf Splettstoesser, Jean J.

Philippe, Jean Prieur, Thomas F. Brooks, “Reduction of helicopter blade-vortex interaction noise by active rotor control technology”, Progress in Aerospace Sciences, Volume 33, Issues 9-10, 1997, Pages 647-687.

[7] Van Dam et al., “Microfabricated Translational Stages for

Control of Aerodynamic Loading”, Patent No. US 7,028,954 B2, Apr. 18, 2006.

[8] Kentfield, J. A. C., “The Potential of Gurney Flaps for

improving the Aerodynamic Performance of Helicopter Rotors,” Paper AIAA 1993-4883, AIAA International Powered Lift Conference Proceedings, Santa Clara, CA, December 1–3, 1993.

[9] Maughmer, M., Lesieutre, G., and Kinzel, M., “Miniature Trailing-Edge Effectors for Rotorcraft Performance Enhancement,” American Helicopter Society 61st Annual Forum Proceedings, Grapevine, TX, June 1–3, 2005.

[10] Myose, R., Papadakis, M., and Heron, I., “Gurney Flap

Experiments on Airfoils, Wings, and Reflection Plane Model,” Journal of Aircraft, Vol. 35, (2), March–April 1998, pp. 206–211.

[11] R.D. Blevins, “Formulas for Natural Frequency and

Mode Shape”, Van Nostrand Reinhold Co, New York (1979).

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[12] A. Concilio, “Controllo Attivo del Rumore in Cabina Mediante Attuatori Piezoelettrici Distribuiti sulla Struttura”, PhD Thesis University of Naples Federico II, 1995.

[13] Eric F Prechtl, Steven R Hall, “Design of a high

efficiency, large stroke, electromechanical actuator”. Smart Mater. Struct. 8 (1999) 13-30. PII: S0964-1726(99)98849-3

[14] www.physikinstrumente.com/ [15] http://www.femm.info/wiki/HomePage