Document [original]

Wind Energ. Sci., 5, 1645–1662, 2020

https://doi.org/10.5194/wes-5-1645-2020

the Creative Commons Attribution 4.0 License.

Aerodynamic effects of Gurney flaps on the rotor blades

of a research wind turbine

Jörg Alber1, Rodrigo Soto-Valle1, Marinos Manolesos2, Sirko Bartholomay1, Christian Navid Nayeri1,

Marvin Schönlau1, Christian Menzel1, Christian Oliver Paschereit1, Joachim Twele3, and

Jens Fortmann3

1Technische Universität Berlin, Hermann-Föttinger Institut, Müller-Breslau-Str. 8, 10623 Berlin, Germany

2College of Engineering, Swansea University, Bay Campus, Fabian Way, Swansea, SA1 8EN, United Kingdom

3Hochschule für Technik und Wirtschaft Berlin, Wilhelminenhofstraße 75A, 12459 Berlin, Germany

Correspondence: Jörg Alber ([email protected])

Received: 10 February 2020 – Discussion started: 21 February 2020

Revised: 13 September 2020 – Accepted: 21 September 2020 – Published: 26 November 2020

Abstract. This paper investigates the aerodynamic impact of Gurney flaps on a research wind turbine of the

Hermann-Föttinger Institute at the Technische Universität Berlin. The rotor radius is 1.5 m, and the blade con-

figurations consist of the clean and the tripped baseline cases, emulating the effects of forced leading-edge

transition. The wind tunnel experiments include three operation points based on tip speed ratios of 3.0, 4.3, and

5.6, reaching Reynolds numbers of approximately 2.5×105. The measurements are taken by means of three

different methods: ultrasonic anemometry in the wake, surface pressure taps in the midspan blade region, and

strain gauges at the blade root. The retrofit applications consist of two Gurney flap heights of 0.5 % and 1.0 % in

relation to the chord length, which are implemented perpendicular to the pressure side at the trailing edge. As a

result, the Gurney flap configurations lead to performance improvements in terms of the axial wake velocities,

the angles of attack and the lift coefficients. The enhancement of the root bending moments implies an increase

in both the rotor torque and the thrust. Furthermore, the aerodynamic impact appears to be more pronounced

in the tripped case compared to the clean case. Gurney flaps are considered a passive flow-control device worth

investigating for the use on horizontal-axis wind turbines.

1 Introduction

The energy yield of modern horizontal-axis wind turbines

(HAWTs) is supposed to be optimal while keeping the main-

tenance costs as low as possible over a lifetime of around

20 years. However, the performance of rotor blades faces

serious challenges, two of which are early separation and

roughness effects. Early separation is a problem especially

in the inner blade region towards the root, where the angles

of attack (AoAs) are elevated due to structural constraints,

such as limited chord length and twist angles (see Fig. 1a).

Over time, the resulting dynamic loads contribute to the ma-

terial fatigue of the blade (Mueller-Vahl et al., 2012). For

this reason, passive flow-control (PFC) devices, such as vor-

tex generators (VGs), are implemented in the inner blade re-

gion of different-size HAWTs aiming at stall delay (Pechli-

vanoglou et al., 2013). At the same time, the long-standing

surface erosion causes roughness effects, especially close to

the leading edge (LE; see Fig. 1b). LE roughness is relevant

throughout the entire blade span and especially in the outer

region towards the blade tip. Apart from the broad range of

weather conditions, surface roughening is aggravated by rain

and insects as well as sand or salt particles (Pechlivanoglou

et al., 2010). Consequently, the energy yield of HAWTs is of-

ten found to be lower than predicted or regressing over time

(Wilcox et al., 2017).

This paper investigates the retrofit application of Gurney

flaps (GFs) in order to improve the aerodynamic performance

of rotor blades. This PFC device consists of a wedge or right-

angle profile that is attached perpendicularly to the pres-

sure side at the trailing edge (TE). The Gurney flap height,

Published by Copernicus Publications on behalf of the European Academy of Wind Energy e.V.

1646 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Figure 1. Rotor blades of utility-scale wind turbines. (a) Flow indicators to detect early separation in the inner blade region, reproduced

from Pechlivanoglou et al. (2013). (b) Leading-edge erosion, with permission from Seilpartner Windkraft GmbH.

Figure 2. (a) Position of the Gurney flap at the trailing edge of a Clark-Y airfoil section. (b) CFD simulation of the HQ17 airfoil at

Re =1.0×106, reproduced and modified from Schatz et al. (2004a).

in relation to the chord length, c, is the main design pa-

rameter, illustrated in Fig. 2a. It is usually in the range of

0.5 %c< GF < 2.0 %cwithout taking the TE thickness into

account.

The research on TE flaps of airplane wings dates back to

the early 20th century (Gruschwitz and Schrenk, 1933). The

GF itself is named after the race car driver Dan Gurney, who

discovered the significant gain in downforce when applying

the device on the rear spoilers. Following from that, GFs have

been implemented on high-lift-dependent transport airliners

(Bechert et al., 2000) and helicopter stabilizers (Houghton

et al., 2013). More recently, Vestas®has started offering

GFs in combination with VGs as so-called aerodynamic up-

grades of HAWTs, predicting annual yield improvements of

up to 2.0 % (Vestas, 2020). The design of the DTU 10 MW

reference wind turbine includes smooth wedge-shaped GFs

in the first half of the blade length, 0.05R<r< 0.4R, using

GF heights in the range of 1.3 %c< GF < 3.5 %c(Bak et al.,

2013).

Figure 2b illustrates the changes in the flow field of

the laminar airfoil HQ17 when implementing different GF

heights, as reported by Liebeck (1978) by means of the New-

man airfoil. Key to the aerodynamic understanding is the

development of one vortex upstream and two counterrotat-

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1647

Figure 3. (a) Closed-loop wind tunnel in top view, reproduced and modified from Klein at al. (2018). (b) BeRT setup in front view looking

downstream.

ing vortices downstream of the GF, as such entailing a low-

pressure region in the TE wake. As a result, the downwash

angle of the flow becomes steeper, the requirements for pres-

sure recovery on the suction side milder, the local boundary

layer thinner, and the suction peak higher. Additionally, the

flow on the pressure side decelerates, leading to a positive

pressure built up in the TE region. The resulting shift of the

Kutta condition leads to increased circulation and thus to el-

evated lift forces, which is the main Gurney flap characteris-

tic. At the same time, the low-pressure region aft of the TE

induces additional drag, especially if vortex shedding is ini-

tiated in the form of a Kármán vortex street. Hence, the lift

increase is accompanied by a certain drag penalty that affects

the lift-to-drag (L /D) ratio accordingly.

This is why various experimental and numerical research

projects aim to limit the adverse drag increase while main-

taining the beneficial lift enhancement. Giguère et al. (1995)

and Kentfield (1996) conclude that the GF height is sup-

posed to be submerged into the local boundary layer (BL)

in order to keep the drag at an acceptable level. Bechert et

al. (2000) demonstrate that additional holes, slits, and espe-

cially the pattern of dragonfly wings lead to reduced drag

on the HQ17 airfoil (thmax =15.2 %c,Re =1.0×106). In

addition, promising results are presented for very small GF

heights in the range of 0.2 %c< GF < 0.5 %c, i.e., substan-

tially smaller than the BL thickness at the TE. Following

from that, wake simulations based on computational fluid dy-

namics (CFD) of Schatz et al. (2004b) reveal that the amount

of induced drag depends on the GF height, in fact, in a dispro-

portionate manner, as illustrated in Fig. 2b: for GF =1.5 %c

a vortex street is triggered, while for GF =0.5 %cthe wake

is shed in a relatively smooth way. In a similar manner, Alber

et al. (2017) suggest the use of very small GF heights of ap-

proximately half the local BL thickness in order to maintain,

or even improve, the airfoil L /D ratio of different DU and

NACA airfoils.

The aforementioned design principle, GF < δ, is applied on

the rotor blades of the Berlin Research Turbine (BeRT) using

GF heights of 0.5 %cand 1.0 %c. In addition, forced LE tran-

sition is triggered in order to emulate the effects of leading-

edge roughness.

The aerodynamic impact of GFs is investigated by means

of the following measurement methods:

–3D ultrasonic anemometry in the turbine wake to deter-

mine the local AoA;

–chord-wise pressure taps to calculate the local pressure

distribution and the lift performance;

–strain gauges at the blade root to measure the flapwise

and the edgewise root bending moments.

In summary, the objective of the experiments is to assess

the suitability of retrofit GFs in order to alleviate the follow-

ing adverse effects:

–early separation due to the high-AoA regime, relevant

in the inner blade region (see Fig. 1a);

–decreasing lift forces due to leading-edge erosion, rele-

vant in the outer blade region (see Fig. 1b).

In the remainder of this paper, the experimental setup is

described in detail, followed by the presentation and the dis-

cussion of the results. The main conclusions are summarized

in the final section of this report.

2 Experimental setup

2.1 Berlin Research Turbine

The BeRT is a test bench of the closed-loop wind tunnel of

the Hermann-Föttinger Institut at the Technische Universität

Berlin. It is a unique wind turbine demonstrator to explore

specific fluid-dynamic phenomena based on a fully equipped

rotating system (Vey et al., 2015).

Figure 3a depicts the wind tunnel facility consisting of the

high-speed (2.0m×1.4 m) and the low-speed (4.2m×4.2 m)

https://doi.org/10.5194/wes-5-1645-2020 Wind Energ. Sci., 5, 1645–1662, 2020

1648 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Figure 4. (a) Zigzag tape at the leading edge of the suction side. (b) Gurney flap and ZZ tape at the pressure side of the trailing edge.

test section. The BeRT is situated in the low-speed test sec-

tion downstream of the flow-conditioning screens and up-

stream of the wind tunnel contraction. The maximum in-

flow velocity is 10 m s−1. The third screen upstream of the

rotor plane is equipped with an additional turbulence filter

mat (Vildedon P15/150s) in order to reduce the turbulence

intensity to 1.0 % < Ti < 1.5 %, as reported by Bartholomay et

al. (2017). Figure 3b displays the BeRT setup and the mea-

surement methods applied. The rotor radius is R=1.5 m,

producing a relatively large blockage ratio of approximately

40 % in relation to the test section area. The blockage effects

on both the flow and the rotor performance are discussed

in Sect. 3.1. Relative distances are expressed in relation to

the rotor radius, R, and the zero position at the center of

the rotor plane at X=Y=Z=0. The blades consist of the

low Reynolds profile, Clark-Y, with a maximum thickness

of thmax =11.9 %cand a modified TE thickness of 0.75 %c.

The blade geometry is optimized aerodynamically, including

a linear decrease in both the chord lengths and the twist an-

gles from root to tip alongside most of the blade span. The

root section is contiguous to the round rotor hub, and the tip

section is pointy (see Fig. 4). The tip speed ratio (TSR) at

rated conditions is 4.3, developing a span-wise Re number

range from root to tip of 1.7×105<Re < 3.0×105. The ax-

ial inflow velocity is captured by two parallel Prandtl tubes

that are permanently installed at approximately one rotor ra-

dius upstream, close to each wind tunnel wall and slightly

above hub height. At rated conditions, the inflow velocity

is 6.5 m s−1at a rotational frequency of frot =3.0 Hz. The

data acquisition system of the rotating sensors, such as pres-

sure taps and strain gauges, is installed within the rotational

spinner (see Fig. 6a). The electrical power is transferred to

the rotating system through a slip ring. Communication with

the host PC is established via Wi-Fi connection in order to

set and modify the rotational speed. The signals are captured

on all channels simultaneously at a rate of 10 kHz, generat-

Table 1. Blade configurations.

Tripped case Clean case

Baseline

GF =0.5 %cOperation points

GF =1.0 %c

ing around 6.0×105data points per measurement, which are

streamed to a host PC via network connection.

2.2 Blade configurations and operation points

The test matrix consists of six blade configurations (Table 1)

and three operation points (Table 2), which are specified

throughout this section.

2.2.1 Forced transition

Following Klein et al. (2018), the principal baseline config-

uration of the BeRT includes zigzag (ZZ) turbulator tape: in

short, the tripped case. ZZ tape is applied in order to initi-

ate the laminar-to-turbulent transition of the boundary layer

(BL) at a fixed location. In practical terms, it is used to emu-

late LE roughness effects on both airfoil sections (van Rooij

and Timmer, 2003) and rotor blades (Zhang et al., 2017). Its

height is slightly smaller than the local BL thickness, δ, in

order to trigger the BL transition while avoiding a dispropor-

tionate drag increase or even turbulent separation. The ZZ

tape is implemented on all BeRT blades at a chord-wise LE

position of both the suction side (SuS) at xSuS =5.0 %cand

the pressure side (PrS) at xPrS =10.0 %c. The BL thickness

of the clean baseline is calculated with the software XFOIL

(Drela, 1989) based on the Re number, the AoA, and the N

criterion (Ncrit) modeling the transition location. The design

conditions of the Clark-Y airfoil are defined by αopt =5.0◦,

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1649

Table 2. Summary of operation points.

Stall Rated Feather

TSR 3.0 4.3 5.6

Inflow velocity in m s−16.5 6.5 5.0

Rot. frequency frot in hertz 2.1 3.0 3.0

Re number (Sect. 3.2) 2.2×1052.8×1052.7×105

AoA in ◦(tripped baseline; Sect. 3.1) 16.3 8.8 4.8

AoA in ◦(clean baseline; Appendix A) 16.5 8.6 4.6

Re =2.5×105, and Ncrit =6 accounting for the elevated Ti

inside the test section (Sect. 2.1). As such, the attached flow

at prestall conditions is assumed two-dimensional in order to

estimate δby means of the XFOIL code. The absolute height

of the ZZ tape is adjusted in various steps in relation to the

chord length, as depicted in Fig. 4a. In addition, all experi-

ments are also performed under the consideration of the free

BL transition, i.e., without including ZZ tape: in short, the

clean case.

2.2.2 Gurney flaps

The GF height is supposed to be submerged into the BL at

the TE in order to keep the drag penalty at an acceptable

level, as discussed in Sect. 1. Hence, it is important to es-

timate δbefore dimensioning the GF height since the aero-

dynamic impact depends on the GF / δ ratio. Apart from the

AoA and the transition location, δis related to Re. The Re

number range of the BeRT is significantly lower compared

to the blades of multi-MW HAWTs. At design conditions

(Re =2.5×105), the XFOIL code predicts the BL thickness

at the TE to be δTE =1.0 %c. Additionally, another GF height

of half the local δis chosen so that the GF configurations

consist of GF =1.0 %cand GF =0.5 %c. For comparison,

the FFA-W3-241 airfoil (thmax =24.1 %c,Re =12.0×106),

which is used in the outer blade region of the DTU 10 MW

reference wind turbine (Bak et al., 2013), generates a BL

of δTE ≈0.30 %c. As such, the application of GF > 0.30 %c

would be likely to cause the L /D ratio to decline, as illus-

trated in Fig. 2b.

Apart from the very tip section, the GFs are implemented

in the form of thin angle profiles made of brass. One side

of the angle profiles is cut in a linear way in order to match

the chord decrease, as shown in Fig. 4b. The other side of

the profile is attached with thin double-sided adhesive tape

adjacent to the TE.

2.2.3 Operation points

The operation points (OPs) include the so-called stall, rated,

and feather conditions, which are characterized by low,

medium, and high TSRs or AoAs, respectively (see Table 2).

Each measurement has a total duration of 60 s. No blockage

correction is applied so that the results refer to the conditions

Figure 5. (a) Ultrasonic anemometer, with permission from Thies

CLIMA. (b) Definition of the azimuthal blade positions looking

downstream.

inside the closed test section. All sensors are calibrated, and

a zero-offset measurement is performed before each test run

in order to reduce experimental errors. The uncertainty of the

results is evaluated in Appendix B.

2.3 Measurement methods

The experimental approaches are summarized in Table 3 and

explained in detail throughout this section.

2.3.1 Ultrasonic anemometry

Three-dimensional ultrasonic anemometers (UAs) are widely

spread in the wind energy industry. The technology is rec-

ognized by different wind industry standards, such as the

IEC 61400 to determine the power curve of wind turbines or

the Association of German Engineers (VDI) for turbulence

measurements. There are numerous references for the use of

UAs in the context of wind tunnel campaigns, such as Weber

et al. (1995), Hand et al. (2001), and Cuerva et al. (2003).

The UA is a commercial product of Thies CLIMA (version

4.383). According to the manufacturer, they are precalibrated

and free from maintenance.

Figure 5a displays the three separate acoustic transmitter–

receiver pairs that are installed orthogonally to each other.

The velocity vectors, u,v, and w, are determined by six in-

dividual measurements based on the bidirectional time-of-

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1650 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Table 3. Measurement methods and quantities.

Sensor Measured quantity Derived quantity Blade position

Ultrasonic anemometer 3D wake velocities AoA 0.56R

Pressure taps Pressure distribution Lift coefficients 0.45R

Strain gauges Flapwise and edgewise bending moments Blade root

Figure 6. (a) BeRT blade and pressure taps, with permission from SMART BLADE GmbH. (b) Chord-wise position of pressure taps at

r=0.45R.

flight principle, i.e., the duration of each signal to be sent

and received.

u=L

21

t1−1

t2,(1)

where Lis the exact running length between each sensor

pair so that the measurement volume amounts to 200mm ×

200mm ×100mm. The velocity vectors vand ware deter-

mined accordingly. Equation (1) shows that the 3D velocity

calculation depends solely on the average propagation time

of the ultrasound, t1and t2, depending on the specific air-

flow passing through the measurement volume. As such, the

output values already imply the density and temperature of

the air. Subsequently, the velocity vectors are transformed

into a natural coordinate system so that the output time se-

ries consist of the axial, lateral, and vertical velocity compo-

nents (u,v, and w). The device-internal data acquisition sys-

tem is a half-duplex interface that is completely independent

of both the wind tunnel and the BeRT system. According to

the manufacturer, the measurement accuracy is 0.1 m s−1per

integrated value and 0.01 m s−1with respect to each of the

three velocity components. The data are recorded at a sam-

pling rate of 60 Hz, thus providing around 3600 data points

per measurement. Considering the relatively large measure-

ment volume and the low sampling rate compared to, e.g.,

hot-wire or laser-based devices, the UA is not adequate for

the investigation of complex or high-speed flow structures.

However, the BeRT wake flow is expected to consist of an

axial and a tangential velocity component due to the forma-

tion of a rotating wake tube. The impact of complex tip and

root vortices is considered negligible in the midspan blade

region, as shown by Herráez et al. (2018).

The UA is installed at one static position downstream,

X=1.3R; in the midspan region, Y=0.56R; and at hub

height, Z=0R(see Fig. 5b). It is positioned vertically with a

spirit level and turned around its own axis towards the undis-

turbed axial inflow so that the lateral and the vertical compo-

nents, vand w, tend to 0. The setup is fixed at its final posi-

tion for all measurements, which are presented in Sect. 3.

2.3.2 Pressure taps

The pressure distribution is extracted by means of 18 pres-

sure taps (PTs) on the SuS and 12 on the PrS, located along

the chord length at r=0.45R(see Fig. 6b). Each orifice is

connected via silicone tubing to its corresponding differen-

tial pressure sensor (HCL0025E), i.e., the pressure box in-

side the spinner. The sensor accuracy is given with 0.05 % of

the full-scale range of ±2500 Pa under nominal conditions.

The experimental procedure and the data postprocessing are

based on Soto-Valle et al. (2020).

The differential pressure values are transformed into the

pressure coefficient,

cpi =1psti +prot

pdyn,ref =psti −pst,∞+0.5ρ·(ωr)2

pdyn,ref

,(2)

where

–1psti is the static pressure difference between each PT

and the inflow Prandtl tube, pst,∞;

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1651

–prot refers to the pressure due to the rotation of the blade

element (it is added to 1psti in the form of a constant

correction term in accordance with Hand et al., 2001);

–pdyn,ref describes the referential dynamic pressure, i.e.,

the effective flow velocity experienced by the blade el-

ement (following Hand et al., 2001, it is determined by

the maximum pressure that is recorded on the pressure

side, the frontal stagnation point, where cpi =1.0; ac-

cording to Eq. 2 the referential dynamic pressure is then

calculated with pdyn,ref =1pst,ref +prot).

The cpvalues are phase-averaged over an azimuthal an-

gle of ϕ=10◦(see Fig. 5b). Each PT provides a total of 36

pressure values at the following blade positions: ϕ=[0, 10,

20 . . . 350◦] so that ϕ=270◦contains the average of all data

points between 265◦and 275◦. The pressure difference, 1cp,

is calculated by subtracting the integrated cpdistribution be-

tween the PrS and the SuS in order to determine both the

normal coefficient, cn, and the tangential coefficient, ct. Per

definition, cnis orthogonal to the chord line pointing towards

the SuS, while ctis parallel to the chord line pointing towards

the LE.

According to Hand et al. (2001), the axial and tangential

coefficients are calculated with

cn=1

2·X30

i=1cpi +cpi+1·(xi+1−xi)(3)

and

ct=1

2·X30

i=1cpi +cpi+1·(yi+1−yi),(4)

where xand yare the normalized chord positions of each PT.

The numbering starts at the TE (x=0.9) with the PTs on the

SuS, moving counterclockwise until the LE (x=0) and back

to the TE on the PrS.

Subsequently, the lift coefficient, cl, and the pressure drag

coefficient, cdp,are determined by (Fuglsang et al., 1998)

cl=cn·cos(α)+ct·sin(α)(5)

and

cdp =cn·sin(α)−ct·cos(α).(6)

The required AoAs, α, are adopted by the uncorrected in-

flow and wake velocity measurements (Sect. 3.1). At pre-

stall conditions, i.e., considering small AoAs, ctcnso that

cn≈cl(Barlow et al., 1999). It is noted that Eq. (6) de-

scribes the pressure drag, which does not account for the

skin-friction drag component. Hence, it is not possible to ex-

tract the total drag, cd, of the blade element via the local cp

distribution (Houghton et al., 2013).

2.3.3 Strain gauges

The strain gauges (SGs) are mounted at the clamping of the

blade detecting the root bending moments (RBMs) in the out-

of-plane or flapwise and in-plane or edgewise direction (see

Fig. 6a). They are connected in a full-bridge configuration

aiming at the mitigation of temperature and cross-talk effects

(FAET-A6194N-35). The experimental procedure to deter-

mine the RBMs is based on Bartholomay et al. (2018). For

the purpose of the presented baseline measurements, a sim-

plified postprocessing protocol is applied without including

the data-based cross-talk correction.

Before testing each blade configuration, the offset signal

is recorded in slow motion at the lowest rotating frequency

available, frot =0.1 Hz. In this way, the gravitational RBMs

are subtracted from the results, which are otherwise regis-

tered as a sinusoidal signal in the edgewise direction. At op-

erational frequencies, the axial forces due to the blade rota-

tion are causing a material deformation directed towards the

blade tip. They are quantified as a combination of centrifugal

and gravitational forces by

Faxial =Fcent −Fgrav

=(mblade ·rcg ·ω2)−(mblade ·g·cos(ϕ)),(7)

where mblade =5.67 kg, the center of gravity is located at

rcg =0.31R,gis the gravitational constant, and ϕrefers to

each phase-locked blade position. The rotational frequency,

ω, is kept constant during each test run so that the centrifu-

gal force Fcent becomes a constant correction term at each

OP. The effective flapwise and edgewise RBMs, which are

related exclusively to the aerodynamic loads acting on the

blade, are determined by

Mflap (ϕ)=Uf,raw(ϕ)−Uf,off(ϕ)·Kf1 −(Faxial ·Kf2)(8)

and

Medge (ϕ)=Ue,raw(ϕ)−Ue,off(ϕ)·Ke1 −(Faxial ·Ke2),(9)

where

–Mflap and Medge are the aerodynamic flapwise or edge-

wise RBMs in Nm;

–Uf,raw and Ue,raw stand for the raw data signal in V;

–Uf,off and Ue,off describe the slow-motion offset signal

in V;

–Kf1 and Ke1 refer to constant calibration factors to

transform V into Nm;

–Kf2 and Ke2 refer to constant calibration factors to

transform the axial forces from N into Nm.

Applying Eqs. (8) and (9), both the out-of-plane and the in-

plane RBMs are computed for each of the 36 blade positions

(see Sect. 3).

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1652 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Figure 7. Tripped case at r=0.56Rand ϕ=270◦.(a) Mean axial and tangential (tan) wake velocity normalized by the inflow velocity.

(b) Standard deviation of the wake velocity normalized by the average wake velocity.

Figure 8. Angles of attack in the tripped case at r=0.56Rand ϕ=270◦.(a) Stall and rated conditions. (b) Rated and feather condi-

tions. (c) AoA difference between Gurney flap configurations and the baseline.

3 Results

The measurement results of both the tripped and the clean

cases are presented and discussed. For space economy, the

clean case is only presented in terms of the concluding re-

sults, such as the lift performance in Sect. 3.2 and the root

bending moments in Sect. 3.3, but otherwise accessible in

Appendix A for completeness.

3.1 Wake velocities and angles of attack

Following Snel et al. (2009), Fig. 7a shows the average axial

and tangential wake velocity normalized by the axial inflow

velocity at each OP, uu−1

∞and wu−1

∞.

Starting from the baseline, Fig. 7a shows that the axial

wake velocities are found to be significantly higher com-

pared to typical free-flow conditions without wind tunnel

walls. According to the steady-state blade element momen-

tum (BEM) method, the optimum axial wake velocity is

supposed to be around one-third of the inflow (Burton et

al., 2011). In this case, it amounts to more than two-thirds

at all OPs. This phenomenon is caused by the wind tun-

nel blockage effects, previously shown by CFD simula-

tions using the fluid-dynamic code FLOWer. At rated con-

ditions of the BeRT, Klein et al. (2018) conclude that the

flow decelerates to an axial wake velocity in the range of

0.62u∞<uCFD < 0.77u∞, which is in agreement with the ex-

perimental results, uexp =0.69u∞.The corresponding tan-

gential velocity, on the other hand, is similar to the steady-

state BEM simulation of QBlade (Marten et al., 2013), with

wBEM =0.18u∞compared to wexp =0.17u∞. According to

Eq. (11), wdepends primarily on the rotational speed of the

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1653

Table 4. Comparison of approximate AoA results at rated conditions and ϕ=270◦.

Method Blade position Case AoA Reference

Pressure taps 0.45RClean 8.0◦Soto-Valle et al. (2020)

Ultrasonic anemometry 0.56RTripped 8.8◦Present study

Three-hole probe 0.65RTripped 8.5◦Klein et al. (2018)

CFD simulation 0.65RTripped 8.2◦Klein et al. (2018)

Figure 9. Pressure coefficients in the tripped case with respect to different scales at r=0.45Rand ϕ=270◦.(a) TSR =3.0. (b) TSR =4.3.

blade. The tangential wake velocity is therefore less affected

by the wind tunnel blockage effect.

Regarding the impact of the GFs, Fig. 7a illustrates the

consistent decrease in the axial and the consistent increase

in the tangential wake velocity in relation to the GF height.

The lateral velocity component is neglected as it amounts

to v0.1ms−1. Figure 7b shows the standard deviation

normalized by the corresponding average velocity compo-

nent describing the 1D turbulence intensity, expressed in per-

cent (Burton et al., 2011). As expected, the flow separation,

TSR =3.0, is captured by the UA in the form of a more tur-

bulent wake field, especially regarding the tangential compo-

nent. The GF configurations do not influence the wake turbu-

lence considerably, except for the tangential velocity compo-

nent at stall, where the GFs appear to mitigate the turbulence

level.

According to the BEM method (Hansen, 2015), the wake

velocity is converted into the axial and tangential rotor in-

duction factors,

a=1

21−u

u∞(10)

and

a0=w

2ωr .(11)

The induction factors, aand a0, describe the decrease in

the axial and the increase in the tangential velocity compo-

nent from a reference point sufficiently far away from the

rotor plane rather than the rotor plane itself (Burton et al.,

2011). The wake measurements are recorded at a distance of

X=1.3Rdownstream in order to avoid the influence of the

wind tunnel contraction (see Fig. 3a).

Subsequently, the AoAs are derived by means of Eqs. (10)

and (11) with

α=arctan(1−a)u∞

(1+a0)ωr −β=arctanu∞+u

2ωr +w−β, (12)

where the twist angle at the radial location of the UA is

β(0.56R)=9.8◦.

At rated conditions, the AoA of the baseline case is

αZZ =8.8◦(see Fig. 8a and b). This outcome is in agreement

with different experimental and numerical investigations of

the BeRT, gathered in Table 4.

The relatively small deviations between the results are due

to the different measurement methods as well as blade con-

figurations (Table 4). The AoA is therefore considered con-

stant in the midspan region within the range of 0.45R≤r≤

0.65R. In all cases, the AoAs are significantly higher com-

pared to the original blade design of the BeRT, αopt =5.0◦.

Figure 8c displays the consistent AoA decrease caused by

the GF configurations. The AoA differences between GF and

https://doi.org/10.5194/wes-5-1645-2020 Wind Energ. Sci., 5, 1645–1662, 2020

1654 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Figure 10. Lift coefficients over angles of attack at r=0.45Rand ϕ=270◦.(a) Tripped case. (b) Clean case. (c) Relative lift increase in

Gurney flap configurations in relation to the corresponding baseline.

baseline configurations amount to 1αGF=0.5 %c=0.5◦and

1αGF=1.0 %c=0.9◦, i.e., to a level that is closer to the op-

timum blade operation. The results quantify an important ef-

fect of retrofit GFs on the blade performance: decreasing ax-

ial wake velocities and thus reduced AoA.

In Sect. 3.2, the AoAs are correlated with the normal-force

coefficients in order to obtain the lift coefficients.

3.2 Pressure distribution and lift performance

Figure 9 shows the distribution of the pressure coefficients,

cp, in relation to the different OPs.

The cpcurves shown in Fig. 9b and c represent the prestall

cases at αTSR=4.3=8.8◦and αTSR=5.6=4.8◦, respectively.

At stall (see Fig. 9a), the separation at the SuS is not yet com-

plete despite the elevated AoA, αTSR=3.0=16.3◦. The curves

indicate the effect of stall delay due to the blade rotation, as

discussed hereafter.

The GF configurations cause an expansion of the pres-

sure differences between the PrS and the SuS, 1cp, along

the complete chord length and regarding all OPs. This ef-

fect is particularly visible in terms of the aft loading to-

wards the TE at 0.5 < x< 0.9. The increased circulation due

to the GF applications is reflected by 1cp, as reported by

Storms and Jang (1994) based on the clean NACA 4412 air-

foil (thmax =12.0 %c,Re =2.0×106).

In order to quantify the results, the cpdistribution is trans-

formed into the local lift curve based on Eq. (5). The required

AoAs are adopted from Sect. 3.1 so that the lift coefficients

combine the results of both the wake velocity and the pres-

sure measurements.

Figure 10a and b depict the lift coefficients of both the

tripped and the clean cases. Starting from the baseline, the

tripped case shows smaller clat 4◦<α< 5◦because of the

forced BL transition at the LE. At 8◦<α< 9◦, this is not the

case anymore, while in the stall region, 15◦<α< 17◦, the ZZ

tape appears to develop a beneficial effect on the lift perfor-

Figure 11. Lift coefficients of the Clark-Y airfoil including Gurney

flap, reproduced and modified from Kheir-Aldeen (2014).

mance. This phenomenon is probably caused by the tripped

and more turbulent BL that remains attached until it is closer

to the TE. In the clean case, however, the less energetic BL

separates earlier, thus leading to smaller clat elevated AoA.

This observation is confirmed by comparable airfoil exper-

iments on the FX 63-137 airfoil section (thmax =13.7 %c,

Re =2.0×105) using ZZ tape with a thickness of 0.75 mm

(Holst et al., 2016). Despite the decrease in the prestall, the

lift coefficients are found at a similar level in the poststall

region.

Looking at the GF configurations, the clperformance in

the tripped case is at a similar or even higher level consider-

ing the complete AoA range, 4◦<α< 17◦. Hence, forced LE

transition does not neutralize or mitigate the GF effect. In

fact, the GF configurations appear to alleviate the adverse ef-

fects of forced LE transition by improving the local clperfor-

mance. Figure 10c highlights the relative lift increase, 1cl,

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1655

Figure 12. Lift over pressure drag coefficients at r=0.45Rand ϕ=270◦.(a) Tripped case. (b) Clean case. (c) Pressure drag coefficients in

relation to the corresponding baseline.

Figure 13. Flapwise and edgewise root bending moments in the tripped case. (a) TSR =3.0. (b) TSR =4.3. (c) TSR =5.6.

between the GF and the corresponding baseline configura-

tions. At rated conditions, TSR =4.3, 1cl,GF=0.5 %c=0.11

or 9.3 %, and 1cl,GF=1.0 %c=0.19 or 16.9 %, illustrating the

main characteristic of retrofit GFs: the considerable lift in-

crease.

The level of both cl,baseline and 1cl,GF=1.0 %cis in agree-

ment with comparable wind tunnel experiments based on a

similar Clark-Y airfoil section, as depicted in Fig. 11.

Figure 11 compares the lift coefficients of the clean

Clark-Y airfoil section (thmax =14.0 %c,Re =2.1×105,

GF =1.2 %c) and the clean Clark-Y blade element of the

BeRT (thmax =11.9 %c,Re =2.5×105, GF =1.0 %c). The

results demonstrate similarities for both the baseline and the

GF configurations. The elevated clin the case of the BeRT

are due to the thinner Clark-Y blade element. At cl,max, the

blade performance is furthermore characterized by the radial

flow due to the blade rotation causing stall delay. This be-

havior is in agreement with experiments on the field rotor

at the Delft University of Technology. Van Rooij and Tim-

mer (2003) report a significant shift of cl,max compared to

2D airfoil simulations.

For completeness, the lift over the pressure drag coeffi-

cients (Eq. 6) is displayed as an indicator of the drag perfor-

mance. It is reiterated that cdp <cd, as previously discussed

in Sect. 2.3.2.

Figure 12a and b illustrate the dependency of cdp on

the OP, reaching values of 0.024 < cdp,prestall < 0.04 and

cdp,stall ≈0.25. In general, the baseline results are compa-

rable to the clean S809 airfoil (thmax =21.0 %c,Re =3.0×

105) that is used for the NREL Phase VI test turbine (Hand

et al., 2001). Figure 12c visualizes the increase in cdp in the

tripped case due to the implementation of the ZZ tape. The

GF configurations, on the other hand, influence the cdp values

in a less noticeable way.

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1656 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Figure 14. Flapwise (flap) and edgewise (edge) root bending moments. (a) Tripped case. (b) Relative increase to tripped baseline. (c) Clean

case. (d) Relative increase to clean baseline.

After evaluating one area of the midspan blade region, the

impact of GFs over the complete blade span is presented in

Sect. 3.3.

3.3 Root bending moments

The integration of the aerodynamic loads, i.e., the lift and the

drag forces acting along the blade span, yields the RBMs.

The in-plane or edgewise RBMs are proportional to the rotor

torque and thus the mechanical power output. They are di-

rectly related to the out-of-plane or flapwise RBMs, which

are proportional to the rotor thrust and thus the structural

loads (Hansen, 2015).

Figure 13 displays the aerodynamic RBMs that are

recorded over one blade revolution in the form of 36 phase-

locked blade positions. The impact of the GF configurations

is registered as an overall increase in both the flapwise and

the edgewise RBMs. In order to quantify and to discuss the

results, the RBMs are presented as average values for both

the tripped and the clean cases.

The results of Fig. 14a confirm the increment of the av-

erage RBMs in relation to the GF height in accordance with

Fig. 13. In the clean case, the overall trend is similar to the

tripped case considering all OPs (see Fig. 14c). This means

that the impact of the Gurney flaps, previously quantified in

terms of the local lift coefficients, is now registered in the

form of increased RBMs in both the flapwise and the edge-

wise direction.

In Fig. 14b, the performance of the GF configura-

tions is quantified in relation to the tripped baseline.

At rated conditions, the average increase in the flap-

wise RBMs amounts to 1Mflap,GF=0.5 %c=3.8 Nm or

6.7 % and to 1Mflap,GF=1.0 %c=7.0 Nm or 12.4 %.

At the same time, the edgewise RBMs are en-

hanced by 1Medge,GF=0.5 %c=1.0 Nm or 11.2 % and

1Medge,GF=1.0 %c=1.8 Nm or 19.7 %. In the clean case

(see Fig. 14d), the overall trend is similar though less

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1657

pronounced. In both cases, the GF configurations generate

performance improvements regarding the rotor torque, albeit

at the expense of the inherent increase in the rotor thrust.

Overall, the results reinforce the observation that GFs are

more effective in relation to the tripped compared to the clean

baseline. Looking at the relative increase shown in Fig. 14b

and d, the GF configurations appear to alleviate the effects of

forced LE transition, especially on the edgewise RBMs, as

previously discussed in Sect. 3.2 with respect to the local lift

performance.

4 Conclusions

The aerodynamic impact of Gurney flaps is investigated on

the rotor blades of the Berlin Research Turbine. The test ma-

trix consists of the clean and the tripped baseline cases as

well as two GF configurations of 0.5 %cand 1.0 %c. Three

measurement methods are applied, including 3D ultrasonic

anemometry, surface pressure taps, and strain gauges.

The baseline measurements confirm the influence of

the prevailing wind tunnel blockage. At rated conditions,

TSR =4.3, and in the midspan blade region, the axial wake

velocity is approximately double in comparison to ideal free-

flow conditions without wind tunnel walls. The correspond-

ing angle of attack is elevated in comparison to the opti-

mum blade design and amounts to αexp =8.8◦rather than

αopt =5.0◦.

The impact of the Gurney flaps is registered regarding all

blade configurations and operation points. In the tripped case

and at rated conditions, the axial wake velocities are reduced,

and the angles of attack are decreased by 1αGF=0.5 %c=0.5◦

and 1αGF=1.0 %c=0.9◦. At the same time, the local lift co-

efficients are enhanced by 1cl,GF=0.5 %c=0.11 or 9.3 % and

1cl,GF=1.0 %c=0.19 or 16.9 %, which is the main character-

istic of Gurney flaps. The effect of the aerodynamic loads

over the complete blade span is analyzed by means of the

root bending moments. The average increase in the out-

of-plane direction amounts to 1Mflap,GF=0.5 %c=3.8 Nm

or 6.7 % and to 1Mflap,GF=1.0 %c=7.0 Nm or 12.4 %.

Simultaneously, the in-plane bending moments are el-

evated by 1Medge,GF=0.5 %c=1.0 Nm or 11.2 % and

1Medge,GF=1.0 %c=1.8 Nm or 19.7 %. Hence, decreasing

angles of attack and increasing lift coefficients appear to be

correlated with the enhancement of both the rotor torque and

the thrust. Overall, the aerodynamic effect is found to be

more pronounced in the tripped case compared to the clean

case.

The experimental results demonstrate the potential of

retrofit Gurney flaps to improve the rotor blade performance

in the following ways:

–decreasing angles of attack to a level that is closer to the

optimum blade operation;

–elevated lift forces compensating for the adverse effects

of forced leading-edge transition.

In summary, Gurney flaps are considered a passive flow-

control device worth investigating for the use on horizontal-

axis wind turbines of different sizes. However, the design of

the Gurney flap height in relation to the local boundary layer

thickness is crucial in order to achieve performance improve-

ments while avoiding detrimental effects such as additional

drag forces. Future research is required to quantify the im-

pact of Gurney flaps on dynamic loads, surface roughness,

and the power output of rotor blades that operate in open-

field conditions and at high Reynolds numbers.

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1658 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Appendix A: Results of the clean case

Figure A1. Clean case at r=0.56Rand ϕ=270◦.(a) Axial and tangential (tan) wake velocity normalized by the inflow velocity. (b) Stan-

dard deviation of the wake velocity normalized by the average wake velocity.

Figure A2. Angles of attack in the clean case at r=0.56Rand ϕ=270◦.(a) Stall and rated conditions. (b) Rated and feather condi-

tions. (c) AoA difference between Gurney flap configuration and the baseline.

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1659

Figure A3. Pressure coefficients in the clean case with respect to different scales at r=0.45Rand ϕ=270◦.(a) TSR =3.0. (b) TSR =4.3.

Figure A4. Flapwise and edgewise root bending moments in the clean case. (a) TSR =3.0. (b) TSR =4.3. (c) TSR =5.6.

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1660 J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades

Appendix B: Uncertainty estimation

The experimental uncertainty of the raw results is expressed

by means of the standard deviation,

σ=v

n−1

i=1|µi−µ|2,(B1)

where nis the number of samples, and µrefers to the average

result. The values of σare rounded up conservatively and are

considered representative for both tripped and clean baseline

cases as well as the GF configurations.

As expected, the scatter of both the velocity and the

pressure data depends on the OP; i.e., it is higher at stall

(TSR =3.0; see Table B1). Looking at the RBMs, however,

the experimental uncertainty of σ(Mflap) and σ(Medge)

is influenced by the structural impact of the rotational fre-

quency that the SGs register simultaneously to the aerody-

namic forces. Overall, the standard deviation is not signifi-

cantly influenced by either of the GF configurations.

Table B1. Standard deviation and reference values in brackets.

Section Quantity TSR =3.0 TSR =4.3 TSR =5.6

3.1 σ(u∞; m s−1) 0.02 (6.57) 0.02 (6.57) 0.01 (5.02)

σ(u; m s−1) 0.20 (4.87) 0.06 (4.55) 0.04 (3.49)

σ(w; m s−1) 0.20 (1.06) 0.06 (1.12) 0.03 (0.71)

3.2∗σmin (1p; Pa) 2.8 (21.8) 2.6 (102.5) 1.7 (6.1)

σmax (1p; Pa) 30.0 (−193.6) 5.8 (−269.1) 3.2 (−41.6)

3.3 σ(Mflap; Nm) 1.9 (36.6) 2.9 (56.5) 2.2 (42.9)

σ(Medge; Nm) 1.0 (8.5) 1.1 (9.1) 0.6 (4.4)

∗Minimum and maximum standard deviation of pressure taps.

Table B2. The 95 % confidence interval and reference values in brackets.

Section Quantity TSR =3.0 TSR =4.3 TSR =5.6

3.1 ε(u∞; m s−1) 5.0×10−5(6.57) 5.0×10−5(6.57) 2.8×10−5(5.02)

ε(u; m s−1) 6.1×10−3(4.87) 2.1×10−3(4.55) 1.2×10−3(3.49)

ε(w; m s−1) 7.1×10−3(1.06) 1.8×10−3(1.12) 1.1×10−3(0.71)

3.2∗εmin (1p; Pa) 4.3×10−2(21.8) 4.0×10−2(102.5) 2.7×10−2(6.1)

εmax (1p; Pa) 5.1×10−1(−193.6) 8.8×10−2(−269.1) 4.8×10−2(−41.6)

3.3 ε(Mflap; Nm) 2.9×10−2(36.6) 4.5×10−2(56.5) 3.4×10−2(42.9)

ε(Medge; Nm) 1.5×10−2(8.5) 1.6×10−2(9.1) 9.6×10−3(4.4)

∗Minimum and maximum confidence interval of pressure taps.

Subsequently, the 95 % confidence interval or so-called

random error is computed with

ε=t·σ

√n≈1.96 ·σ

√n,(B2)

where tis the Student’s tdistribution (Barlow et al., 1999).

The values of the 95 % confidence interval (see Table B2),

are significantly smaller compared to those of the standard

deviation (Table B1). The reason is the relatively large num-

ber of samples: n≈3.6×103in terms of the wake velocities,

uand w, and n≈1.7×104per azimuthal angle in the remain-

ing cases. Hence, the presented average results are contained

by a reasonably small confidence interval.

Wind Energ. Sci., 5, 1645–1662, 2020 https://doi.org/10.5194/wes-5-1645-2020

J. Alber et al.: Aerodynamic effects of Gurney flaps on the rotor blades 1661

Data availability. Measurement data and results can be provided

by contacting the corresponding author.

Author contributions. JA performed the wind tunnel experi-

ments together with RSV counting on the support of all coauthors.

JA processed the data and prepared the manuscript with the support

of MM and RSV, both of whom contributed with important com-

ments and suggestions to all sections of the paper.

Competing interests. The authors declare that they have no con-

flict of interest.

Special issue statement. This article is part of the special issue

“Wind Energy Science Conference 2019”. It is a result of the Wind

Energy Science Conference 2019, Cork, Ireland, 17–20 June 2019.

Acknowledgements. The authors would like to acknowledge the

constant support of the BeRT project by the researchers and the

technicians of the Hermann-Föttinger Institut at the Technische

Universität Berlin. The authors also appreciate the technical sup-

port of SMART BLADE GmbH. Rodrigo Soto-Valle would like to

thank ANID PFCHA/Becas Chile-DAAD/2016-91645539 for the

support. Marinos Manolesos would like to acknowledge the contri-

bution of the EPSRC Supergen ORE Hub Early Career Researcher

Research Fund.

Financial support. This open-access publication was funded

by Technische Universität Berlin.

Review statement. This paper was edited by Mingming Zhang

and reviewed by Athanasios Barlas, Galih Bangga, and two anony-

mous referees.

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