New Airfoils for Small Horizontal Axis Wind Turbines

Transcription

1 P. Giguere Graduate Research Assistant. M.S. Selig Assistant Professor. Department of Aeronautical and Astronautical Engineering. University of Illinois at Urbana-Champaign. 36 Talbot Laboratory. 14 S. Wright Street. Urbana. IL New Airfoils for Small Horizontal Axis Wind Turbines In a continuing effort to enhance the performance of small wind energy systems, one root airfoil and three primary airfoils were specifically designed for small horizontal axis wind turbines. These airfoils are intended primarily for 1-5 kw variable-speed wind turbines for both conventional (tapered/twisted) or pultruded blades. The four airfoils were wind-tunnel tested at Reynolds numbers between 1, and 5,?~. Tests with simulated leading-edge roughness were also conducted. The results zndlcate that small variable-speed wind turbines should benefit from the use of the new airfoils which provide enhanced lift-to-drag ratio performance as compared with previously existing airfoils. 1 Introduction Over the last 12 years, a considerable numbe~ of ~oils ~ave been specifically developed for horizon~ axis. w~nd turbi?es (HAWfs). For example, 25 specially tailored airfmls.for wmd turbines originated from the joint work between the NatJ.onal Renewable Energy Laboratory (NREL) and Airfoils, Inc. (Tangier and Somers, 1995). For wind-turbine applications, the advanced NREL airfoils have been shown to provide aerodynamic and structural advantages as compared with airfoils originally designed f?r airplane applications. In addition, BjOrk ( 1988, 1989, 199), Hill and Garrard ( 1989), Madsen and Rasmussen ( 1992), and Timmer and Rooy (1992) have also designed airf~ils for ~ v.:s.. There are, however, only a limited number of wmd-turbme rurfo!ls that have been developed exclusively for small blades. Typically, the entire span of the blades?f small ~wrs operate at low Reynolds numbers where lanunar.sep~tton effects can severely degrade the performance of airfmls not designed for this flow regime. In additi~n, the ty~ically f~t r~tational speeds of small HA wrs provide centrifugal stj.ffemng that reduces the blade bending loads and thus, allows for the use of thinner airfoils as compared with those designed for larger wind turbines. Reductions in laminar separation effec~ and in airfoil thickness provide an increase m aerodynarmc performance, which yields better energy.capture. The~efore, ~wing to low Reynolds number aerodynanu.cs and centnfu?al stiffening effects, wind-turbine airfoils designed for medmm and large blades are not particularly suitable for smal~er?lades.. A number of existing low Reynolds number rurfods, which were designed for small unmanned aerial vehicles, have been shown to be applicable to small wind turbines (Giguere and Selig, 1996, 1997a). Specially tailored airfoils for small HA Wfs, however, should provide enhanced performance. From the literature, it appears that the S822 and S823 airfoils are the only two airfoils that have been specially designed for small blades. These two airfoils are principally intended for stallregulated HA wrs and form one of the nine advanced NREL airfoil families (Tangier and Somers, 1995), Consequently, there is a need for new wind-turbine airfoils for small HA Wfs, especially for those operating in variable-speed ~ode. In a continuing effort to support new blade designs for.sm~ll HA Wfs and improve their performance, a total of four rurfotls were designed and wind-tunnel tested. These airfoils are intended primarily for 1-5 kw rated power variable-speed wind turbines for conventional (tapered/twisted) or pultruded blades. Contributed by the Solar Energy Division of THE AMERICAN SOCIETY OF ME CHANICAL ENGINEERS for publication in the ASME JOURNAL OF SoLAR. ENERGY ENGINEERING. Manuscript received by the ASME Solar Energy DIVISIOn, July 1997; final revision. Dec Associate TechnicB.l Editor: P. S. Veers. Section 2 describes the airfoil design objectives and methodology. A brief description of the wind-tunnel facility, measurement techniques, and data reduction method is provided in Section 3. Section 4 provides a discussion of the wind-tunnel test results, including tests with fixed transition to simulate leadingedge roughness caused by blade erosion, accumulation of insect debris, etc. Finally, conclusions are given in Section 5. 2 Airfoil Design Objectives and Methodology In this design work, the focus has been on designing airfoils that can be used along the entire blade span of small variablespeed HA Wfs. Three airfoils, referred to as the primary airfoils, were designed to provide optimum performance over a broad range of operating conditions. Considering the low operating Reynolds numbers and beneficial centrifugal stiffening effects of small HA Wfs, the airfoil thickness for the primary airfoils was fixed at ten percent. A 16 percent root airfoil was also designed to accommodate possible large root bending moment and large blade-stiffness :equirements. With the root airfoil over the inboard 3 percent of the blade, the primary airfoils are recommended for use over the last 25 percent of the blade (from the 75 percent station to the tip), with blending between 3 percent and 75 percent of the blade span. The root airfoil (SG64) and the three primary airfoils (SG641, SG642, and SG643) are depicted in Fig. l. During normal operation, variable-speed HA Wfs operate below stall and over a relatively limited lift range as compared with constant-speed wind turbines. For these operating conditions, minimizing leading-edge roughness effects, such as no loss in the maximum lift coefficient, is not particularly critical (Tangier and Somers, 1995; Tangier, 1997). For optimum aerodynamic performance during variable-speed operation, the low-drag lift range (drag bucket) can be reduced in favor of greater lift -to-drag ratios. Accordingly, the primary airfoils were designed to achieve as high a lift-to-drag ratio as possible at the design lift coefficient, ranging from.6 to 1.2. The primary airfoil having the lowest design lift coefficient was expected to yield large blade chords and consequently, the root airfoil was designed to be mainly used with the two other primary airfoils. To account for possible variations in the tip-speed ratio (TSR) caused by atmospheric turbulence, the best lift-to-drag conditions were designed to occur over a range of lift coefficients centered about the design lift ~fficient. The operational Reynolds number range for HA WTs having a rated power between 1-5 kw is typically below 1,,. At these Reynolds numbers and particularly below 5,, minimizing laminar seperation effects is an important design driver. Accordingly, low design Reynolds numbers were favored to control laminar seperation effects. The design Reynolds numbers for the four airfoils were selected according to the 18 I Vol. 12, MAY 1998 Copyright 1998 by ASME Transactions of the ASME

2 -~- Thickness: 16% : Cambe~ 2.5% --C~- Thidcness: 1%: Cambe~ 2.% _ c-== SG642 ==-= Thidcness: 1% : Camber: 3.8% -L---sG643 ~- Thidcness: 1% : Camber: 5.5% Fig. 1 Contours of the root and primary airfoils Table 1 Airfoil design objectives Airfoil tic Design C 1 Design Re SG64 16% 1.1 2, SG641 1%.6 5, SG642 1%.9 333,333 SG643 1% , E 1 SG64 Avg. accuracy:.23 mm (69 in) 5 ~ ~ - / --\;;.:.,:.-:;.,;;.,---- i:s xlc Fig. 2 Contour accuracy plot for the SG64 airfoil tradeoff between operating lift coefficient C 1 and Reynolds number Re. It has been shown that for an optimum variable-speed rotor, the product of the lift coefficient and Reynolds number is nearly constant for a particular blade station (Giguere and Selig, 1997b), i.e., (l) Here, R is the reduced Reynolds number which is constant so long as the TSR and the local velocity, which includes the axial induction factors, of a given blade station remain constant. Therefore, the potential increase in the maximum lift-to-drag ratio from operating at a high lift coefficient is reduced by a lower operating Reynolds number according to Eq. ( 1) and vice versa The design Reynolds number for each primary airfoil was determined using Eq. (I ) and a reduced Reynolds number R of 3,. For the root airfoil, a design Reynolds number of 2, was used. The objectives for the four new airfoils are shown in Table 1. To achieve the design objectives, the new airfoils were systematically designed using PRO FOIL (Selig and Maughmer, 1992), which is an inverse airfoil design method. In addition, the Eppler code (Eppler, 199) and XFOIL (Drela, 1989) were used to obtain preliminary airfoil performance characteristics to guide the design. With PRO FOIL, the desired velocity distribution is prescribed, from which the corresponding airfoil shape is determined. In tailoring the velocity distribution for each airfoil, a gradual transition ramp was introduced on the upper surface to minimize laminar separation effects such as bubble drag. The need for a shallow pressure gradient combined with a high-lift requirement lead to aft-loading and, consequently, a relatively high pitching moment coefficient ( -8 < Cm.cl < -.14). Note that the reduction in the size of the drag bucket caused by the use of aft-loading is not a drawback for variablespeed HA WTs because, as mentioned previously, they operate over a relatively limited lift range under normal operation. The primary airfoils have similar velocity distributions that differ mainly by their respective amount of aft-loading. Therefore, the aerodynamic characteristics of the three primary airfoils are related and these airfoils can be considered as an airfoil series. The addition of the root airfoil to the primary airfoils provides a unique airfoil family for variable-speed HA WTs. 3 Testing Apparatus and Methods The experiments were performed in the University of Illinois at Urbana-Champaign low-turbulence subsonic wind tunnel having a.857 m (2.81 ft) high and m (4ft) wide test section. For the airfoil tests, two m ( 6 ft) long Plexiglas splitter plates were inserted into the test section and reduced the test section width to.854 m (2.82 ft). With an empty test section, the turbulence intensity of the tunnel is less than.1 percent for the test Reynolds numbers (Guglielmo, 1996; Selig et al., 1995). The.35 m (1 ft) chord airfoil models were made of foam with a fiberglass finish and were inserted horizontally between the splitter plates with nominal gaps of 1-2 mm (4-8 in.). Using a coordinate measuring machine, each airfoil model was digitized at midspan. As an example of the digitized results, Fig. 2 presents the accuracy plot for the model of the SG64 airfoil, which was the least accurate of the four models used in these tests. The average accuracy of the SG64 airfoil model is, nevertheless,.23 mm (89 in.) over the 35 mm ( 12 in.) chord. Similar plots for the primary airfoils can be found in Lyon eta!. ( 1997a). The airfoil coordinates are tabulated in the Appendix. The lift was directly measured using a strain-gage load cell, and the drag was determined from the average of four different span- SG64 v Re=1, o Re:3, o Re=15, "' Rt!=<IOO,OOO '. ~ ~ ! ~ ~ : : :. : -----~~-~:~. ' :. : Ct --: r----- r T r w......l-.4...l-j:...::o...u::.'<>'--iu...j...o...l~w...j 1 O.o c... a (deg) Fig. 3 Drag polars and corresponding lift curves for the SG64 root airfoil (free transition) Journal of Solar Energy Engineering MAY 1998, VoL 12 I 19

3 SG641 "' Re-1, o Re-3, o Re-15, "' Re-4, o Ae=2, t> Re-5, 2, f-.r-.-,--,c"t"'-.-."""t""t"t't-rt-,-,-.,-,-:-r,-...-,-.,...,..,...,~ c, r r c/ r -- -~------~ L,c_...._._..._._._._, o.oo om o.o2 o.o3 o.o4 o.os o.os -1 o 1 2 a(deg) SG642 c... "' Re-1, o Re-3, o Ae-15, A Re-4, T ~ r ~ Ce r-- T r r r w,...,_......w a(deg) SG643 c... v Re-1, o Ae-3, o Ae-15, A Re-4, 2. j-r-,..,...,...,...,""t"'!..,...,...,_,...,.;.r""t.,...,..",..rer- rsoo,..,..:.;,ooo..t-1 : ' : 1 : j : t t !!! r ---- ~~T=- ----r r---- r r ().5 L.;..._...L..L....r...L... '-L c... a(deg) Fig. 4 Drag polars and corresponding lift curves for the SG641, SG642, SG643 primary airfoils (free transition) wise wake surveys spaced 76.2 mm (3 in.) apart (Selig et a!.,!995; Lyon eta!., 1997a). The overall uncertainty in both the lift and drag measurements was estimated to be percent ( Guglielmo, 1996; Selig eta!., 1995; Lyon eta!., I997a). All measurements were corrected for wind.. tunnel interference effects according to a method that has been validated with data from the NASA Langley Low Turbulence Pressure Tunnel (Giguere and Selig, 1997c; Lyon eta!., 1997b). Finally, to simulate leadingedge roughness effects, a 8 mm (23 in.) zigzag trip was fixed to the upper and lower surface of the airfoils (Lyon et a!., I997a). In this paper, this trip size is referred to as the "standard" trip height. The trips were positioned in the vicinity of the suction peaks at two percent chord on the upper surface and at five percent chord on the lower surface. Note that transition does not occur immediately behind the trips but rather takes place over a finite distance (Lyon et a!., I997a). The use of trips instead of grit 11 I Vol. 12, MAY 1998 roughness elements has the advantage of being a more repeatable method, but the results obtained with the standard trips should bc considered as a worst case scenario. 4 Wind Tunnel Test Results This section provides an overview of the airfoil data obtained from the wind-tunnel tests conducted with the four new airfoils. The complete data set for a Reynolds number range of 1,- 5, can be found in Lyon et al. ( 1997a). Eiren though this Reynolds number range might not covers the entire operational range of smalll-5 kw HA WTs, data above a Reynolds number of 5, can generally be obtained quite accurately from logarithmic and linear extrapolation for the drag and lift, respectively. The drag variations are largest below a Reynolds number of 3, and thus, data in that range is more critical to document for small wind turbines. Transactions of the ASME

4 Re = 3, SG64x Airfoil Family 1 o Previously existing airfoils FX (Vc=13.6%)'-.g 9 ('! Ol ~ ''6 SG642 (Vc=1%)'-, /,// ' /' E387 _-'' (Vc=9%) \ / o-a18 (Vc=7.3%) ~ 8 / Q"Clari<-Y (Vc=11.7%) E S662,..... "' ::I E (Vc=B%)'\ / '-SD737 (Vc=9.2%) -~ &' :E 7 SG641/ (Vc=1%)/ \ ~SD66 (Vc=1.4%) o's822 (Vc=16%) 5712 (Vc=8.8%) RG15 (Vc=8.9%) 6 SD732 (Vc=1%) / /GO 417a (Vc=2.9%) /SG64 (Vc=16%) '-BW-3 (Vc=S%) '-NACA2414 (Vc=14%) '-se23 (Vc=21%) u 1.4 Lift coefficient for maximum lift-to-drag ratio Fig. 5 Maximum lift-to-drag ratio versus the corresponding lift coefficient of various airfoils for small HAWTs (Re = 3,) Tests with free transition (no trips applied to the airfoil) were conducted at six Reynolds numbers from 1, to 5,. Figures 3 and 4 present the drag polars and corresponding lift curves for the root and primary airfoils, respectively. These figures show that the design objectives have been met. For example, the maximum lift-to-drag ratio of the SG642 occurs near a lift coefficient of.9 at its design Reynolds number and also over most of the Reynolds number range considered. Furthermore, the objective of providing a range of lift coefficients for which best liftto-drag ratio performance occurs has also been satisfied. Therefore, performance penalties owing to off--design conditions should be relatively small except for a Reynolds number of 1, where laminar separation effects caused a large increase in drag. It should be emphasized, however, that such performance at a Reynolds number of 1, is typical of most low Reynolds airfoils (Giguere and Selig, 1996; Selig et al., 1995). Figure 5 indicates the maximum lift-to-drag ratio and corresponding lift coefficient under clean conditions of the SG64x airfoil family and other low Reynolds number airfoils applicable to small HA WTs. The results are shown for a Reynolds number of 3,, which is representative of the data for other Reynolds numbers, and the thickness of each airfoil is also shown. Note that all the data shown in Fig. 5 is based on wind tunnel experiments conducted with the testing apparatus and methods described in Section 3. As indicated in Fig. 5, the new airfoils provide lift-todrag ratios that are equivalent or exceeding those of previously existing low Reynolds airfoils over a wide range of design lift coefficients. Therefore, small variable-speed HA WTs are likely to benefit from enhanced energy capture from the use of the new airfoils. Note, however, that comparing data for constant Reynolds number can be misleading based on the previously mentioned tradeoff between operating lift coefficient and Reynolds number. A better figure of merit to use in comparing the airfoils would be the maximum lift-to-drag ratio for a given reduced Reynolds number (Giguere and Selig, 1997a). Nonetheless, the results shown in Fig. 5 are indicative of the potential of the new airfoils for small variable-speed wind turbines. For the tests with ''fixed'' transition, three Reynolds numbers were considered for each airfoil: 15,, 3,, and 5,. An overview of these results is shown in Fig. 6 with results for a Reynolds number of 3,. To facilitate comparison, the drag polar of each airfoil is shown for both free and Re=3, o SG64 v SG641 "- SG642 2 _,..._,,...,...,.,..,...,...,.--r-r~~t"'"" ,...,...,...,~~..,.;;.sgs-t-t-4..:.:-t3 Re=3.ooo o SG64 v SG641 Fixed transition A SG642 zigzag trip@ 2%us, 5%1s c, - l...l.j'-'--'-'-'......l-d..l..l..l..l..w...j-l.l...l '--'-'... -'-'-' c, c.. Fig. 6 Drag polars for the root and the primary airfoils at a Reynolds number of 3, (free and fixed transition) Journal of Solar Energy Engineering MAY 1998, Vol. 12 I 111

5 2 _ SG64 Ct Re=3 SG641 Re=3 2. c, c, 2. SG642 Re=3 2. r~'{ c, SG643 Re=3 L*-"'-'-"W...-'-'...LJ ,..() J..L... "-'..... '-L.) = o: (deg) o: (deg) o: (deg) o: (deg) Fig. 7 Lift curves for the root and the primary airfoils at their design Reynolds numbers (free and fixed transition): dark symbols, increasing a and open symbols, decreasing a fixed transition. In these tests, the standard trip as defined in Section 3 was used. As expected, the higher the design lift coefficient and relative thickness of the airfoil, the higher is the loss in performance caused by forcing transition. Furthermore, the lift coefficient corresponding to the maximum lift-to-drag ratio is also influenced by the simulated roughness at the leading edge. The change in lift coefficient for maximum lift-to-drag ratio can be explained by the results presented in Fig. 7. Even though the lift coefficients for maximum lift-to-drag ratio vary between the free and fixed transition cases (Fig. 6), it can be seen from Fig. 7 that the angle of attack for which maximum lift-to-drag ratio occurs remains, however, relatively constant. For example, Fig. 6 indicates that the lift coefficient for maximum lift-to-drag ratio of the SG643 airfoil with free and fixed transition is 1.16 and.93, respectively. The lift curves for the SG643 airfoil shown in Fig. 7 indicates that these lift coefficients correspond to an angle of attack of 4 deg. Therefore, a blade designed with the new airfoils should yield optimum performance even with leading-edge roughness. Based on the loss in maximum lift coefficient shown in Fig. 7, the use of the new airfoils should not be extended to small stall-regulated HA WTs. Airfoils such as the NREL S822 and S823 are better candidates in that case. The new airfoils, due largely to lower thickness, provide much better lift-to-drag ratio performance as compared with the S822 and S823 airfoils. Table 2 summarizes the maximum lift coefficient, maximum lift-to-drag ratio and the corresponding lift coefficient for both the free and fixed transition cases of the SG64x airfoil family. As expected, the results presented in Table 2 indicate that the sensitivity to roughness of the airfoil performance increases with the design lift coefficient (Tangier, 1997). Note, however, that for a given Reynolds number, the maximum lift-to-drag ratios with fixed transition of the airfoil having the highest design lift coefficient ( SG643) remain the largest of all the new airfoils. In addition, the lift-to-drag ratio characteristics of Table 2 SG64 root airfoil Performance summary for the SG64x airfoil family SG641 (primary airfoil FIH tnmaltlon Fixed lnmaiuon Free trlnsttlon Fixed transition Re c..., IC>:{Udl- {Ud)- c..., ;(Ud),. (UdJ- c..., :(UdJ- {Ud),.. C...tt C,;{Udl- (Ud),. 1, 1.21l r SG642.JJ>rlm!ry airfoil SG643(prtmary airfoil Free trancltlon Fixedtnnaltlon Fixed transition er~ ;{Ud)_ (1/d)_ c..., ;(Ud). (Ud\- c..., C,;(Vd\- {Ud)- C...tt ;(1/d\-((Ud), , , I Vol. 12, MAY 1998 the airfoils with fixed transition are for the most part independent of the Reynolds number. Consequently, blades using the SG643 airfoil are likely to yield the best energy capture in the presence of leading-edge roughness elements. Such blades, however, are also likely to be most affected by Reynolds number effects owing to the high design lift coefficient of the SG643 airfoil. The data shown so far with fixed transition has been for a fixed trip height, namely the standard trip. The effect of reducing the height of the trip was also investigated, and the results for the SG642 airfoil at a Reynolds number of 3, are shown in Fig. 8. A total of five different zigzag trip heights were used ranging from.13 mrn (5 in.) or 4 percent chord to 6 mm (23 in.-standard trip) or.19. percent chord. The results indicate that trips thicker than the smallest one caused a rapid degradation in airfoil performance. Further tests were performed with plain trips to investigate whether or not the new airfoils were optimized for low drag at their respective design points. The tradeoff to consider involves reducing bubble drag at the cost of increased device dragdrag caused by the trip itself. For the SG642 airfoil, a.15 mm (6 in.) thick plain (tape type) trip was positioned at different chord wise positions on the upper surface at the respective design lift coefficient. The results shown in Fig. 9 clearly indicate that the addition of the plain trip did not lead to any drag reduction beyond the experimental uncertainty of the measurements. In addition, three additional plain trip heights were also tested at 35 percent chord and the same Reynolds number of 3,. As indicated in Fig. I, trips thicker than the baseline (.15 mrn) were not beneficial in reducing the overall drag. Similar results were also obtained for the SG643 airfoil at Reynolds numbers of 2, and 3,. Therefore, the airfoils do have low bubble drag that likely cannot be reduced through the use of a trip or turbulator. Finally, the four airfoils considered as a group provide excellent lift-to-drag ratios over a broad range of lift conditions and are well suited for their intended application. 5 Conclusions The airfoils designed during this work form a unique airfoil family for small variable-speed HA WTs. Wind-tunnel tests of the root airfoil ( SG64) and three primary airfoils ( SG641, SG642, and SG643) provided an extensive airfoil data set that can be used in the design of small blade~; The two primary airfoils having the highest design lift coefficients (SG642 and SG643) yielded enhanced lift-to-drag performance over many other low Reynolds number airfoils. Consequently, small variable.. speed wind turbines are likely to benefit from the use of these two airfoils. The SG64 root airfoil and low-lift SG64l airfoil are also likely to enhance the energy capture of small variable-speed HA WTs owing to structural requirements and Transactions of the ASME

6 ~ SG642 v No trip o hlc=.12% Re=3, o hlc=4% t> hlc=.15% zigzag trip 2%us, 5%1s <> hlc=o.os% I> hlc=.19% r'""'~.,...,.~...-~r"t'"] Ct ' ' ' ' ' r r : , : r r ; r r t : ~ : 'f ----~ C.t a (deg) Fig. 8 Drag polars and corresponding lift curves for the SG642 airfoil with varying zigzag trip heights the tradeoff between operating lift coefficient and Reynolds numbers. Overall, the new airfoils provide excellent lift-to-drag ratios over a broad range of lift conditions and are well suited for small variable-speed wind turbines E- ~ ~ Baseline drag coefficient (13) 8 with experimental uncertainty of % (J SG642 6 Plain trip h=5%c (6") Re = 3, a =4.2 deg 4 ' Trip location on upper surface (% chord) Fig. 9 Drag coefficient as a function of the chordwise position of a hi c = 5 percent plain trip on the upper surface 12 ~ 1 t=~;;:;.::r::;p~~~i:i'iss~ts-ee'~-s-.-;:j:-:;]3. ~~ ~Baseline drag coefficient (13) 8 with experimental uncertainty of % SG642 6 Plain trip at 35%c upper surface Re =3, a =4.2 deg 4 2 Q5 Q1 Q15 Trip height(% chord).2 Fig. 1 Drag coefficient as a function of the normalized height of a plain trip positioned at 35 percent chord on the upper surface Journal of Solar Energy Engineering Acknowledgments The support from the National Renewable Energy Laboratory under subcontract XAF is gratefully acknowledged. Also, the authors would like to thank James L. Tangier of NREL for many useful discussions throughout this work. Finally, the authors are also thankful to Mark Allen of Allen Developments for building the four wind-tunnel models used in this study. References Bjork, A. 1988, "Airfoil Design for Horizontal Axis Wind Turbines," Proceedings of the Second lea Symposium on Joint Aclion on Aerodynamics of Wind Turbines, Lyngby, Nov., pp BjOrk, A., 1989, "Airfoil Design foc Variable RPM Horizontal Axis Wind Turbines," Proceedings EWEC 89, European Wind Energy Conference, Glasgow, Scotland. Bjork, A., 199, "Coordinates and Calculation for the FFA-Wl-xxx, FFA-W2- x.u, FFA-W3-xxx Series of Airfoils for Horizontal Axis Wind Turbines," The Aeronautical Research Institute of Sweden, Technical Report, FFA TN Drela, M., 1989, "XFOIL: An Analysis and Design System for Low Reynolds Number Airfoils,'' Lecture Notes in Engineering: Low Reynolds Number Aerodynamics, T. J. Mueller, ed., Vol. 54, Springer-Verlag, New York, June. Eppler, R., 199, Airfoil Design and Data, Springer-Verlag, New York. Giguere, P., and Selig, M. S., 1996, "Aerodynamic Performance of Small Wind Turbines Operating at Low Reynolds Numbers," American Wind Energy Association WIND POWER '96 Conference, June 23-27, Denver, CO. Giguere, P., and Selig, M. S., 1997a, "Low Reynolds Number Airfoils for Small Horizontal Axis Wmd Turbines," Wind Engineering, 1997, Vol. 21, No.6, Giguere, P., and Selig, M. S., 1997b, "Desirable Airfoil Characteristics for Large Variable-Speed Horizontal Axis Wind Turbines," ASME JoURNAL OF SO LAR ENERGY ENGINEERING, Vol. 119, August, pp Giguere P., and Selig, M.S., l997c, "Velocity Correction for Two Dimensional Tests with Splitter Plates," AJAA J., Vol. 35, No.7, pp Guglielmo, J., 1996, "Spanwise Drag Variations at Low Reynolds Numbers," Department of Aeronautical and Astronautical Engineering Department, University of Illinois at Urbana-Champaign, Urbana, IL, May. Hill, G. C., and Garrard, A. D., 1989, "Design of Aerofoils for Wind Turbine Use," Proceedings of the European Wind Energy Conference, Glasgow, Scotland. Lyon, C. A., Broeren, A. P., Giguere, P., Gopalarathnam, A., and Selig, M. S., 1997a, Summary of Low-Speed Airfoil Data-Volume 3, SoarTech Publications, 154 N, Horseshoe Circle, Virginia Beach, VA. Lyon, C. A., Selig, M. S., and Broeren, A. P. l997b, "Boundary-Layer Trips on Airfoils at Low Reynolds Numbers," AIAA Paper 97-51l. Madsen, H. A., and Rasmussen, F., 1992, "Experiences in Airfoil Analysis and Design," Proceedings of the Sixth lea Symposium on the Aerodynamics of Wind Turbines, Petten, Netherlands. Selig, M.S., and Maughmer, M.D., 1992, "Generalized Multipoint Inverse Airfoil Design," AIAA J., Vol. 3, No. II, pp Selig, M.S., Guglielmo, J. J., Broeren, A. P., and Giguere, P., 1995, Summary of Low Speed Airfoil Data- Volume 1, SoarTech Publications,. 154 N. Horseshoe Circle, Virginia Beach, VA. Tangier, J. L., and Somers, D. M., 1995, "NREL Airfoil Families for HA WTs," American Wind Energy Association WIND POWER '95 Conference, May 9-12, Washington, DC. Tangier, J. L., 1997, "Influence of Pitch, Twist and Taper on a Blade's Performance Loss Due to Roughness," ASME JOURNAL OF SoLAR ENERGY ENGI NEERtNG, Vol. 119, pp Timmer, W. A., and Rooy, R. P. J.. M. van, 1992, "Thick Airfoils for HA WTs," J. of Wind Engineering and Industrial Aerodynamics. Vol. 39, May. MAY 1998, Vol. 12 I 113

7 [SG64 APPENDIX: Airfoil Coordinates Jl:ili~U4 ISG642 x/c [SG643 x/c XJC ylc xlc yc Yfc y/c o.ouuuu I.OUUUU.. 1.llOi:lOO u.ooooo O.Q ~, I Vol. 12, MAY 1998 Transactions of the ASME