Mexnext-II: The Latest Results on Experimental WindTurbine Aerodynamics

K. Boorsma and J.G. SchepersEnergy Research Centre of the Netherlands, ECN, P.O. Box 1, 1755 ZG Petten, The Netherlands

boorsma@ecn.nl

Sugoi Gomez-IradiNational Renewable Energy Center, CENER, Pamplona, Spain

Helge Aagaard Madsen, Niels Sørensen and Wen Zhong ShenTechnical University of Denmark, DTU, Lyngby, Denmark

Christoph SchulzInstitute of Aerodynamics and Gasdynamics, University of Stuttgart, Stuttgart, Germany

Scott SchreckNational Renewable Energy Laboratory, NREL, Colorado, USA

AbstractA selection of results from IEA Annex 29 Mexnexton analysis of wind tunnel measurements is pre-sented. A convincing example illustrates the im-portance of detailed aerodynamic measurements.The influence of MEXICO blade shape deviationsbetween design and manufactured geometry wasassessed by scanning the blade geometry andperforming comparative CFD simulations with thisgeometry. Generally speaking the differences be-tween the results for design and scanned geome-try do not justify the differences observed betweenexperiments and computations A comparison be-tween calculations and unexplored measurementsfrom the famous NREL UAE PHASE VI experi-ment at a relatively high rotational speed is per-formed. It was found that as long as prescribed air-foil data is used, a good agreement exists betweenlifting line code results. The tip effect remains diffi-cult to predict, although it is questioned in how farthe limited blade aspect ratio is representative forlarge commercial wind turbines. CFD RANS simu-lations generally perform better in this respect, al-though separated flow features remain a challengefor these models as well. Preparations for a sec-ond experiment on the existing MEXICO test rigare discussed. New configurations will be testedand new apparatus including an acoustic array willbe used, by which an even higher quality data setcan be assured than the first data set. A stand-still test of the MEXICO blades in the Delft Lowspeed tunnel allowed to determine an appropriate

roughness configuration and to make sure that theblades and their data acquisition are in good shapefor the New MEXICO campaign.

Keywords: Rotor aerodynamics, measurements,predictions

1 IntroductionThe subject of aerodynamics is extremely impor-tant in wind energy. It does not only determinethe energy production of a wind turbine, but italso determines the loads, stability, and noise ofa wind turbine and not to forget the wake be-hind the turbine and the consequent wind farmlosses. As such a thorough understanding of thedetailed wind turbine aerodynamics is of crucialimportance. The present paper describes the lat-est results from Mexnext [1], a large joint interna-tional project in which 19 parties from 9 countriescooperate in understanding the wind turbine aero-dynamic behaviour using advanced aerodynamicmeasurements. The experiments herein rangefrom a wide variety of sources, including measure-ments at the Japanese Mie university [2], the Chi-nese CARDC [3] and old FFA measurements [4].Firstly the importance of sufficient detail in mea-surements is illustrated in section 2. One of theoutcomes as discussed in section 3 deals with theinfluence of deviations between design and man-ufactured blade shape. Section 4 shows a com-parison between measurements and calculationsbased on the NREL UAE PHASE VI experiment

[5]. Finally the preparations for a new MEXICOexperiment [6] are discussed in section 5, followedby the conclusions.

2 Importance of detailed mea-surements

Two of the most representative wind tunnel experi-ments featuring relative large diameter and exten-sive instrumentation are the MEXICO and NRELUAE PHASE VI experiments. Both MEXICO andNREL UAE PHASE VI feature 5 instrumented ra-dial sections with pressure sensors. Both ex-periments feature a Horizontal Axis Wind Turbine(HAWT), where the first is a 4.5 m diameter 3-bladed upwind rotor and the second a 2-bladed 10m diameter upwind rotor (for most configurations).The need for this extensive instrumentation is illus-trated in Figure 1, which shows a comparison be-tween measured and predicted normal force distri-bution for the MEXICO turbine in axial flow condi-tions. Here the predictions are made using a bladeelement momentum (BEM) and a RANS Compu-tational Fluid Dynamics (CFD) code and the mea-surement results are obtained by integrating sec-tional pressure distributions over the chord. Thelegend indicates that there is an excellent agree-ment in rotor axial force, which for this exampleis obtained by integration the sectional forces overthe span. Hence, without any further informationon the local aerodynamic loads this would indicatea good quality of the prediction code. However theavailability of measured local aerodynamic loadsmade it possible to assess the agreement on a lo-cal level which showed a very poor performanceof the BEM code. As illustrated, the good agree-ment in rotor axial force is misleading and causedby compensating errors only. Since the operatingangle of attack for the mid- to outboard section is inthe separated flow region for this case, it is ques-tionable whether the BEM approach using 2D air-foil data is still applicable. Although the CFD codeperforms better on a local level, the agreement ofthe rotor integrated loads is less good in compari-son to the measured values.

3 Influence of blade shape de-viations

In a first phase of Mexnext several comparisonswere made between measurements and calcula-tions. In order to understand some of the differ-ences it was considered important to know theactual blade shape and to compare it with thespecified (design) shape. Thereto all three blade

Figure 1: Radial distribution of sectional normalforce, MEXICO turbine (U∞=24 m/s, 424 rpm, -2.3 ◦ pitch)

shapes have been scanned digitally. Comparingthe geometries generally yielded some small dif-ferences [7], the most important quantities dic-tating the aerodynamics being the profile shapeand twist angle. Nevertheless a CFD investiga-tion (RANS simulations) from the Mexnext partnersCENER and University of Stuttgart showed somenon-negligible differences [8], see also Figure 2(a)for a grasp of the results. Hereto it should be real-ized that the measured loading originates from thefive different instrumented sections that were dis-tributed over the three blades. This means that theCFD results of the scanned geometry (_SCANNED,dashed lines) were taken from blade 1 for the in-ner part (25%R and 35%R), from blade 2 for themidboard part (60%R) and from blade 3 for theoutboard part (82%R and 92%R). The results fromCENER and University of Stuttgart are comparedto the experimental data (black diamonds) and totheir CFD results of the design geometry (solidlines). Although these simulations indicate smalldifferences between the results for the scannedand design geometry (roughly ± 5% in normalforce), these differences are smaller and some-times in opposite direction compared to the overprediction of the design geometry CFD comparedto the experimental data (around 15% in normalforce). Therefore it is concluded that a blade ge-ometry deviation can not solely be held responsi-ble for the shown over prediction in comparison tothe measurements.

Figure 2(b) zooms in on the differences betweenthe three blades, which are more pronounced to-wards the outboard sections. It is expected thatgeometric differences at these sections stand outmore easily due to the higher dynamic pressure. Itappears that the blade 1 results often show an off-set in comparison to blade 2 and 3, which is in con-tradiction with the geometry comparison [7] that re-ported the shape of blade 3 to deviate the most incomparison to the other blades. It is also notedthat phase locked velocity measurements usingPIV from the same experiment revealed an excel-lent agreement between the induced flow fields di-rectly downstream from blade 2 and 3. This couldindicate that the deviations mainly impact the loadsdirectly but not so much the underlying flow field.

To judge the value of these differences it shouldbe realized that to perform a CFD simulation fromscanned geometry, the measured cloud of pointsmust be converted to a surface geometry con-sisting of curves (usually an IGES file). Thisconversion process can be done in many dif-ferent ways and the resulting curves are oftensmoothed, thereby disregarding small scale irreg-ularities. Therefore the accuracy of the resultinggeometry actually used for the CFD prediction canbe questioned. Nevertheless, the results give anindication of uncertainties due to geometry differ-ences between design and actual blade geometry.

4 Comparison between mea-surements and calculations

An extensive comparison between measurementsand calculations on the MEXICO campaign waspreviously reported [1], where in addition to loadsalso velocities and lifting line variables (e.g. angleof attack) for both axial and yawed flow conditionswere subject of investigation. It is however notonly the MEXICO measurements which are usedin the comparison but a main aim of Mexnext-II isalso to consider other experiments. Thereto a newcomparison round is defined on the NREL UAEPHASE VI experiment [5] at a rotational speed of90 rpm. Even though the NREL UAE PHASEVI experiment has been used in many validationcases before [9], the rotor speed was usually only72 rpm where the cases at 90 rpm remained unex-plored. A summary of the cases under investiga-tion is given in Table 1.

4.1 Description of codes and mea-surements

Most of the codes used in the current investiga-tion are described in the final report of Mexnext-I

Table 1: NREL UAE PHASE VI comparison cases(axial flow)Case U∞ Pitch Rot. α@80%R‡ a†

angle speed (estimate) (estimate)[m/s] [ ◦] [rpm] [ ◦] [-]

I5 5.08 0 71.7 4.5 0.21X5 5.02 3 90.2 1.5 0.20

X10 10.04 3 90.9 8.0 0.15X12 12.02 3 91.6 10.0 0.11

‡ Angle of attack† Rotor averaged axial induction factor

[1]. A brief description of the experiment itself andthe codes is given below. For the lifting line codesused, two calculation rounds were performed; onewith prescribed 2D airfoil data and one with 3D cor-rected airfoil data. The latter can be distinguishedby the extension of _3D in the legend name of eachdataset. The 3D corrections used therefore are de-scribed below as well.

4.1.1 NASA-AMES

The comparison features a 2 bladed turbine witha 10 m rotor in upwind configuration without coneand tilt, placed in the large test section of theNASA-AMES wind tunnel. More details can befound in [5]. To obtain normal and tangential force(defined normal and parallel to the chordline re-spectively), the measured pressures are integratedover the chord by applying linear interpolation. Thepressure value taken for each sensor is time aver-aged over the relevant datapoint. Axial force is ob-tained by linear interpolation of these forces overthe span using the 5 instrumented sections, as-suming zero load at the root and tip.

4.1.2 DTU_BEM

This code features BEM theory, where drag is in-cluded in the axial and tangential momentum bal-ances and the Prandtl tip factor is applied to theinduced velocities. To further correct the tip effectsbetween 2D and 3D airfoil data, the tip loss factoris applied on the force obtained by applying 2D air-foil data. More details can be found in [10]. The3D airfoil data caused by rotation is derived fromthe 2D airfoil data and the corrections are made in3 regions of angles of attack, roughly to be the at-tached flow region until maximum lift, post stall andpast leading edge separation. The coefficients thatare needed to perform the correction are predeter-mined, i.e. they are derived prior to the calculation.

(a) Design and scanned geometry (b) Individual blades

Figure 2: Radial distribution of sectional normal force, MEXICO turbine (U∞=15 m/s, 424 rpm, -2.3 ◦ pitch)

4.1.3 DTU_HAWC2

The code is based on the BEM theory but extendedfrom the classical approach to handle dynamic in-flow, dynamic stall, skewed inflow, and shear ef-fects on the induction [11]. There is no general3D correction model included in the software butthe airfoil input data are corrected, based on gen-eral experience with 3D effects and in particular thefield rotor experiments at the former RISØ and atNREL [12, 13]. These main effects are a stronglydelayed stall on the inner part of the blade, but alsowith a strong increases in drag. In addition to thatthere is some increase of the lift coefficient in poststall so that there is little or no negative slope onthe lift curve in this region.

4.1.4 ECNAero

The ECN AERO-MODULE [14] includes both BEMas well as a lifting line free vortex wake formula-tion, allowing the same external input (e.g. wind,tower, airfoil data) to be used for both models. TheBEM formulation is based on PHATAS [15], includ-ing state of the art engineering extensions whichhave matured over decades of research in windturbine rotor aerodynamics. The free vortex wakemethod is based on the AWSM code [16]. The 3Dcorrection is based on the model of Snel [17] asmodified in PHATAS, dependent on chord over ra-dius and tip speed ratio. As such it is embedded inthe overall code, applied during the calculation andrestricted to the inboard region below 50 ◦ angle ofattack.

4.1.5 CENER_CFD

The Wind Multi-Block (WMB) code is used as de-veloped by University of Liverpool and CENER[18]. The compressible Reynolds AveragedNavier-Stokes (RANS) flow equations are solvedon multiblock structures grids using a cell-centredfinite volume method for spatial discretization. Thek−ω SST turbulence model of Menter is used. Thegeometry has been self-defined based on airfoilshape and blade planform table, featuring a sharptrailing edge.

4.1.6 DTU_EllipSys3D

The EllipSys3D code is a multiblock finite volumediscretization of the incompressible RANS equa-tions in general curvilinear coordinates [19]. Thesimulation includes transition computations basedon the γ − ReΘ correlation based transition modelof Menter and turbulence modelling using the k−ωSST turbulence model of Menter. The geometry isbased on a purposely developed IGES file as dis-tributed within the Mexnext group, which features afinite trailing edge thickness.

4.2 Results

A selection of results is displayed in Figure 3.Judging by the axial force results in Figure 3(a),the results are generally in good agreement witheach other and the experiment for the first twocases featuring attached flow conditions on theblades. This was confirmed by inspection of theradial distribution of normal force along the blade.

(a) Axial force for the different cases (b) Normal force, X10, 2D airfoil data

(c) Normal force, X10, 3D airfoil data and CFD (d) Normal force, X5, 3D airfoil data and CFD

(e) Pressure distribution, X5, 80%R (f) Pressure distribution, X10, 80%R

Figure 3: Comparison between measurement and predictions, NREL UAE PHASE VI experiment

There is an overprediction for both lifting line codesas well as CFD for the X10 and X12 cases fea-turing separated flow conditions. Zooming in onthe lifting line results with prescribed airfoil datafor the X10 case in Figure 3(b), shows excellentagreement between predicted and measured val-ues below 80%R. It is noted that a previous com-parison round on this experiment [9] displayed awide variation between predictions, possibly dueto the differences in airfoil data used. In additionto that it can be observed that the tip correctionmodel of DTU_BEM yields different results than thecorrections used in DTU_HAWC2, ECNAero-BEM andECNAero-AWSM. Vortex theory (ECNAero-AWSM) de-livers the same result as the conventional Prandtltip factor for this case, probably because the tur-bine is lightly loaded at an axial induction factoraround 0.15. However, the differences in tip load-ing between ECNAero-BEM and ECNAero-AWSM re-main relatively small also for the other cases. Thisindicates that the dominating tip effect is not thedifference between local and annulus averaged in-duction but rather the lift fall-off due to the finiteblade radius, which is expected to be more presentfor low aspect ratio rotors in comparison to modernturbines.

Application of 3D corrections and also includingthe CFD results then yields Figure 3(c). Here thespreading between the results increases and it canbe observed that the DTU_HAWC2 3D correction isalso applied to the outboard region of the blade.Overviewing all the cases, the currently applied 3Dcorrections mostly affect the X10 and X12 casesince they exhibit separated flow along the span.

The CFD results are not in good agreement witheach other, also not for the X5 case (Figure3(d)), which could well be attributed to the dif-ferent geometry used as addressed in section4.1. The pressure distributions for these cases(Figure 3(e) and 3(f)) reveal differences in suc-tion level and peak (X10) and slight ’kinks’ in theDTU_EllipSys3D X5 upper and lower side pres-sure distribution aft of x=0.2 m due to boundarylayer transition. The measured pressure distri-bution from the inevitable finite number of sen-sors just miss the bend in the pressure distributionaround the maximum airfoil thickness. An estimateof the effect of the limited experimental chordwiseand spanwise resolution was obtained by applica-tion of this resolution to the CFD results in the datareduction procedure and compare this to the re-sults obtained with the original resolution. The in-fluence of the limited spanwise resolution is foundto be responsible for half the discrepancy in axialforce between measurements predictions for the

X10 and X12 case, while the agreement for the I5and X5 case remains good due to the low absolutedifferences. On a relative scale, the finite numberof pressure sensors in chordwise direction are re-sponsible for a 10% decrease of normal force at80%R for the X5 case, whereas this effect is foundto be negligible for large angles of attack (X10 andX12 case).

5 New MEXICO experimentThe first MEXICO measurements led to variousinsights on the added value of CFD, on modelingimprovements (e.g. tip effects, 3D-effects, yawedflow) and it led to an enhanced understanding ofthe 3D-flowfield around wind turbines [1]. Still theanalysis of the first MEXICO measurements led tosome open questions which need to be answeredin a new experiment: MEXICO was the first ex-periment in which the detailed pressure distribu-tion as well as the underlying flow field was mea-sured but the relation between the two opposedthe momentum relations (at least at design con-ditions). Consequently the relation between ve-locities and load needs further attention. More-over some results suffered from data uncertaintiesand faulty instrumentation and some parts of thedatabase (e.g. the measurements on dynamic in-flow and standstill) were considered incomplete. Afurther motivation for New MEXICO are the en-hanced PIV capabilities from DNW which becameavailable recent years, where in particular the reso-lution and/or the size of the PIV sheets is increasedleading to a much more complete mapping of theflow field. Further flow visualization is planned byapplication of oil to the blade surface and smokecandles from the blade tips. In addition to that, theacoustic noise sources will be measured with anacoustic array in order to establish the acoustic-aerodynamic link. Several new test cases will beadded as well, such as the application of Guerneyflaps and fast pitching actions.

5.1 Non rotating test

In preparation for this test the blades were placedin the TUDelft Low Speed Tunnel. In addition tochecking the status of the blades and its instru-mentation, this test also enabled the measurementof quasi-standstill sectional characteristics. Sincethis experiment finished very recently, much of thedata analysis still needs to be performed and thehere shown results are preliminary. The tunnel fea-tures an octagonal cross section of 1.25 m high by1.8 m wide, in which the blades are positioned ver-tically pointing downward. Since the blade lengthof 2.04 m exceeds the tunnel height two configura-

(a) Inboard set-up (b) Outboard set-up

Figure 4: Test set-up in the Delft tunnel

tions were employed, one focusing on the outboardpart (with a free tip) and one on the inboard bladepart (with the tunnel floor cutting off the 69%R sec-tion), see also Figure 4. With a maximum tunnelspeed of 100 m/s, the Reynolds numbers of therotating experiment could be matched.

The data acquisition and pressure sensors werebrought back to life successfully after an inactiveperiod of more than seven years. Having a closelook at the apparatus allowed for fixing some ofthe pressure signals in the inboard sections whichwere faulty during the first MEXICO campaign. Awake rake was positioned downstream of the bladeto measure the velocity deficit. In addition to ob-taining sectional drag, traversing the rake along theblade span for all three blades gave a possibilityto further investigate the agreement between theblades.

5.1.1 Sectional characteristics

Using the ECN AERO-MODULE free vortex wakecode AWSM, a first survey was performed to inves-tigate the angle of attack variation along the bladeusing prescribed airfoil data. By this approach anestimate of the degree of ’two-dimensionality’ ofthe experimental set-up can be obtained. Althoughfull details such as wall effects are not taken intoaccount here, this approach yields an approxima-tion for the order of magnitude of the induced ve-locities causing different inflow angles than the lo-cal geometric twist angle. The results shown inFigure 5 reveal that for this particular blade pitchangle induced angles of attack exceeding 2 ◦ canbe expected, not only confined to the tip area.Hence the trailing vorticity is not only concentratedin the tip vortex but also plays a role in the remain-der of the span due to the varying circulation alongthe blade span as a consequence of the radial twist

(a) Inboard set-up (b) Outboard set-up

Figure 5: Calculated induced angle of attack varia-tion for a geometric tip angle of 15 ◦

and chord distribution. Increasing or decreasingthe blade pitch angle from this value is found to re-duce the induced velocities. For the 15 ◦ pitch an-gle, the geometric angle of attack along the bladein combination with the chord distribution results ina circulation distribution where the trailing vorticityis most effective in inducing velocities perpendicu-lar to the chordline.

Although it is clear that the test set-up cannotbe used to determine sectional characteristics di-rectly, the set-up can be considered comparable toparked rotor conditions. Combining the measure-ments with a planned CFD simulation including thetunnel will possibly reveal information on the under-lying two dimensional sectional characteristics.

5.1.2 Roughness strips and visualization

A stethoscope was traversed over the blade sur-face to determine the location of laminar to turbu-lent transition. In a clean configuration, small sur-face irregularities were found to dictate the chord-wise position of transition along the span. This isregarded as unwanted taking into account com-parison to CFD calculations. The minimum width(5mm) and thickness (0.2 mm) zigzag strip to yieldtransition for the range of angles of attack andReynolds numbers was determined, to keep par-asitic drag due to the strip itself to a minimum.The chordwise position of the strips was kept at10%chord for both pressure and suction side of theblades. The zigzag strips positioning close to thepressure sensors were modified in such a way thatdistortion of the measured pressure distribution bylocal pressure changes due to the small vorticesemanating from the zigzag shape was prevented.

An oil flow visualization in the set-up for the out-board part of blade 3 confirmed that the roughnessstrips indeed provoke transition. Figure 6(a) shows

laminar to turbulent boundary layer transition (di-rectly aft of the zizag strip), illustrated by the lightand respectively darker colours due to the differ-ent friction coefficient between them. The differ-ence with the tip region which does not feature azigzag strip is clearly noticeable. The tape cover-ing the sensors at 60%R, 82%R and 92%R can beobserved from top to bottom respectively. Due tothe twist distribution, the geometric angle of attackis 4.8 ◦, 2.4 ◦ and 1.2 ◦ larger for these sectionsrespectively than the tip angle. Because of thisthe outboard sections already exhibit trailing edgeseparation, which can be observed by the brightcolours due to the oil not being transported overthe surface. Generally speaking the flow patterncan be considered two-dimensional. Figure 6(b)then shows a configuration at a higher tip pitch an-gle, where spanwise flow features can be observedin the separated flow region. In addition to that asmall laminar separation bubble can be observedby a bright coloured line before the roughness stripjust aft of the leading edge, approximately betweenthe tip and 70%R. Figure 6(c) shows the lower sidesurface for a negative angle of attack, where themidboard part of the blade shows separation in thecusp of the RISØ profile.

6 ConclusionsIn the Mexnext project detailed aerodynamic mea-surements are used. These measurements pro-vide local aerodynamic loads along the bladewhere in addition the underlying flow field whichdrives these loads is mapped in some experi-ments. It is proven that such very detailed infor-mation is needed to better understand the aerody-namic behaviour of a wind turbine. With this betterunderstanding it is possible to improve wind tur-bine design codes which eventually leads to morereliable and more efficient wind turbines. Further-more the detailed measurements form essentialand unique validation material for wind turbine de-sign code by which the the validity of these codescan be assessed and which enables the identifica-tion of areas where improvement is needed. Muchprogress has been made in the project wheremeasurements are used from a large variety ofsources, mainly wind tunnel measurements butalso field measurements. The paper puts empha-sis on a comparison between calculations and un-explored measurements from the famous NRELUAE PHASE VI experiment at a relatively high ro-tational speed. Moreover results from the MEX-ICO measurements are shown and the differencesare assessed between the measured and speci-fied blade geometry including the effect of thesedifferences on the loads. It is shown that much

(a) 10.4 ◦upper side

(b) 16.4 ◦upper side

(c) -3.6 ◦lower side

Figure 6: Blade oilflow visualization for a varietyof geometric angles of attack referenced to the tipchord (flow from right to left), U∞=60 m/s

more work is needed. Thereto the so called NewMEXICO experiment is performed, i.e. a secondexperiment on the existing MEXICO test rig in or-der to fill the missing gaps and to take into accountthe lessons learnt from the first MEXICO experi-ment. Hereby an even higher quality data set canbe obtained compared to the data set from the firstcampaign.

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AcknowledgementsThe authors would like to thank the participants ofthe Mexnext group for the pleasant cooperation.The EU Innwind project is acknowledged for finan-cial support in preparing the new MEXICO ex-periment. Finally the DUWIND department of theTUDelft, in particular G.J.W. van Bussel, W.A. Tim-mer, Y. Zhang and D. Baldacchino, are acknowl-egded for their support during the standstill testof the MEXICO blades in the Delft University lowspeed tunnel.

1 Introduction2 Importance of detailed measurements3 Influence of blade shape deviations4 Comparison between measurements and calculations4.1 Description of codes and measurements4.1.1 NASA-Ames4.1.2 DTU_BEM4.1.3 DTU_HAWC24.1.4 ECNAero4.1.5 CENER_CFD4.1.6 DTU_EllipSys3D

4.2 Results

5 New MEXICO experiment5.1 Non rotating test5.1.1 Sectional characteristics5.1.2 Roughness strips and visualization

6 Conclusions