Blade sections for wind turbine and tidal current turbine applications

REVIEW PAPER

Blade sections for wind turbine and tidal current turbine
applications

—current status and future challenges

M. Rafiuddin Ahmed*

,†

Division of Mechanical Engineering, The University of the South Paci

ﬁc, Laucala Campus, Suva, Fiji

SUMMARY

The designers of horizontal axis wind turbines and tidal current turbines are increasingly focusing their attention on the design
of blade sections appropriate for speci

ﬁc applications. In modern large wind turbines, the blade tip is designed using a thin

airfoil for high lift : drag ratio, and the root region is designed using a thick version of the same airfoil for structural support.
A high lift to drag ratio is a generally accepted requirement; however, although a reduction in the drag coef

ﬁcient directly

contributes to a higher aerodynamic ef

ﬁciency, an increase in the lift coefﬁcient does not have a signiﬁcant contribution to

the torque, as it is only a small component of lift that increases the tangential force while the larger component increases the
thrust, necessitating an optimization. An airfoil with a curvature close to the leading edge that contributes more to the rotation
will be a good choice; however, it is still a challenge to design such an airfoil. The design of special purpose airfoils started with
LS and SERI airfoils, which are followed by many series of airfoils, including the new CAS airfoils. After nearly two decades of
extensive research, a number of airfoils are available; however, majority of them are thick airfoils as the strength is still a major
concern. Many of these still show deterioration in performance with leading edge contamination. Similarly, a change in the
freestream turbulence level affects the performance of the blade. A number of active and passive

ﬂow control devices have been

proposed and tested to improve the performance of blades/turbines. The structural requirements for tidal current turbines tend to
lead to thicker sections, particularly near the root, which will cause a higher drag coef

ﬁcient. A bigger challenge in the design of

blades for these turbines is to avoid cavitation (which also leads to thicker sections) and still obtain an acceptably high lift
coef

ﬁcient. Another challenge for the designers is to design blades that give consistent output at varying ﬂow conditions with

a simple control system. The performance of a rotating blade may be signi

ﬁcantly different from a non-rotating blade, which

requires that the design process should continue till the blade is tested under different operating conditions. Copyright ©
2012 John Wiley & Sons, Ltd.

KEY WORDS

wind turbine; tidal current turbine; blade section; lift; drag

Correspondence

*M. Ra

ﬁuddin Ahmed, Division of Mechanical Engineering, The University of the South Paciﬁc, Laucala Campus, Suva, Fiji.

†

E-mail: ahmed_r@usp.ac.fj

Received 8 August 2011; Revised 20 December 2011; Accepted 7 February 2012

1. INTRODUCTION

Research efforts directed at maximizing the power output of
horizontal axis wind turbines (WTs) and tidal current turbines
(TCTs) have increased signi

ﬁcantly during the recent years

providing impetus to extensive research on blades and blade
sections appropriate for speci

ﬁc applications. Advances in

the development of WTs and TCTs will have immense bene-
ﬁts in providing solutions to the global energy requirements.
The rotor blade is one of the most important components of
the WTs and TCTs, which is the primary energy conversion
device. For the turbine blade design, the selection of airfoils
for different sections and the distribution of chords and twists
are pivotal [1]. Most of the NACA airfoils are not appropriate
for WTs and TCTs because of the poor stall characteristics, low
structural ef

ﬁciency near the root, inconsistent performance

at varying Reynolds numbers (Re) and poor performance due
to the roughness effect resulting from leading edge (LE) con-
tamination for these pro

ﬁles. NACA airfoils are suitable mainly

for high Re and relatively small angles of attack (

a) [2,3]. The

magnitude, direction and turbulence levels of the atmospheric
wind are known to vary signi

ﬁcantly with time, which

adversely affects the performance of WTs if NACA airfoils
are employed for the blades. Figure 1 shows the variation of
the section lift coef

ﬁcient (C

) with

a at different Re for

NACA23012 airfoil [2]. It can be seen that the C

drops sharply

and signi

ﬁcantly as a increases beyond the stall angle for differ-

ent Re. Such an abrupt drop in C

will signi

ﬁcantly reduce the

output of a WT or a TCT. The trend for special-purpose blade
sections started in the early 1980s with the design of LS and
Solar Energy Research Institute (SERI) airfoils after the experi-
ence gained from employing aviation class NACA airfoils

INTERNATIONAL JOURNAL OF ENERGY RESEARCH

Int. J. Energy Res. 2012; 36:829–844

Published online 22 March 2012 in Wiley Online Library (wileyonlinelibrary.com). DOI: 10.1002/er.2912

829

highlighted the shortcomings of these airfoils for horizontal
axis wind turbines (HAWTs) [4]. Stall-controlled HAWTs
produced excessive power in stronger winds, which caused
generator damage. The need to gain a better understanding
of the airfoil performance near stall was felt, as some stall-
controlled turbines were operating with some part of the
blade in deep stall a lot of time; the predicted loads were less
than the measured loads, and the LE roughness was affecting
the turbine performance.

Figure 2 shows the

a versus C

behavior of traditional

airfoils and some of the presently used airfoils that have a
gradual upstream movement of the point of separation with
increasing angle of attack. A number of special-purpose air-
foils for WT applications have a gradual upstream movement
of the location of separation from the trailing edge so that C

does not drop sharply and the coef

ﬁcient of drag (C

) does

not rise sharply with an increase in the angle of attack. The
third type of behavior, in which the C

value essentially

remains constant over a wide range of angles of attack, is also
shown. With a growing demand for a reduction in energy
costs, the designers are now forced to think of simple passive
control techniques. An airfoil of this type of behavior will
de

ﬁnitely contribute to a reduction in the cost, as its perfor-

mance will not deteriorate with a change in the

ﬂow direction,

and it will help achieve the challenge of

‘lean design’ [5],

which is especially important for offshore WTs and TCTs
for which maintenance and repairs are expensive and time

consuming. Researchers are, at the moment, testing a number
of active and passive

ﬂow control devices to improve the

performance of the turbines and to control the load on the
rotor. The present paper discusses the different blade sections
that were used in wind turbines starting from the earliest
special-purpose airfoils that were designed by National
Renewable Energy Laboratory (NREL). The main perfor-
mance characteristics of the popular blade pro

ﬁles used in

WTs and TCTs are discussed; many of these are also tested
for their performance under different operating conditions.
Promising

ﬂow control techniques and devices for impro-

vement of turbine performance are discussed in brief. The
performance characteristics of rotating turbine blades are com-
pared with those of non-rotating turbine blades, and

ﬁnally, the

future challenges for blade designers are brie

ﬂy discussed.

2. LIFT AND DRAG

The main focus in the design of blade sections has been to
maximize lift : drag ratio (L/D) mainly by increasing C

. It

is still a generally accepted requirement. However, there are
different preferences among blade aerodynamicists regarding
the maximum C

depending on the type of control

—the stall-

controlled turbines restrict C

l,max

to serve two purposes: (i) to

reduce the peak power generated; and (ii) to keep the thrust
on the system small. One of the earlier airfoils that were
designed (in 1986) with a restricted C

l,max

was S809 [6]. This

airfoil was designed with another related objective of
reduced effect of LE roughness so that the value of C

l,max

does not reduce due to LE contamination. On the other
hand, the pitch-controlled turbines require a high C

l,max

The pitch-controlled system adjusts

a such that maximum

L/D is obtained for all the wind speeds up to the maximum
power. A high value of maximum lift coef

ﬁcient gives higher

aerodynamic ef

ﬁciency as long as the blade structure is satis-

factory. One of the recent designs [7] focused on increasing
the tangential force coef

ﬁcient (C

) rather than C

, as an

Figure 1. The section lift coef

ﬁcient of NACA23012 airfoil at differ-

ent angles of attack.

○ – Re = 3 10

;

□ – Re = 6 10

;

◊ – 8.8

;

– Re = 6 10

(standard roughness) [2].

Figure 2. Different

a versus C

behaviours.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

830

DOI: 10.1002/er

increase in C

contributes more to the thrust than to the rota-

tion. Having C

as the design objective can allow for more

attention on insensitivity to LE roughness rather than on high
lift : drag ratio caused by low C

. In most cases, it is desirable

that the design-

a region is close to C

l,max

because this enables

low rotor solidity and/or low rotational speed [7].

3. LOCATION OF TRANSITION
POINT

It is known that the location of transition of the boundary
layer (BL), especially on the upper surface, is a very impor-
tant factor that signi

ﬁcantly inﬂuences the performance of

the airfoil. The distance from the LE to the point where
transition occurs, x

, reduces with (i) increasing Re; and

(ii) increasing turbulence intensity (Tu). Figure 3 shows
the location of transition on the upper surface of SG6043
airfoil at a lower Re and 5% Tu. When Re is increased, a re-
duction in x

can clearly be seen from Figure 4. A reduction

in x

also can be seen from Figure 5 when Tu is increased to

10%. As x

reduces, the region of laminar boundary layer

reduces, and that of turbulent BL increases; this results in
an increase in the skin friction drag. For higher

a, the shift

in the location of transition is normally less as the transition
anyway occurs close to the LE. However, for higher

a, a

reduction in x

may shift the location of separation

towards the TE, resulting in a reduction in the wake thick-
ness and a lower momentum loss.

The freestream turbulence levels of the atmospheric wind

at heights at which wind turbines are normally installed are
generally higher compared with the levels achieved in stan-
dard wind tunnels. Studies performed in the past have
explored the effect of Tu on the airfoil characteristics
[8

–10] below one million Re. Hoffmann [8] studied the

effect of varying the freestream turbulence intensity from
0.25% to 9% for NACA0015 airfoil and reported an increase
in the peak lift coef

ﬁcient because of delayed ﬂow separation

at higher

a. Devinant et al. [9] varied the turbulence level

from 0.5% to 16% and studied its effect on NACA
65

-421; they found that the

ﬂow was separating at higher

angles of attack when the turbulence level was increased.
Most of the works on such studies are performed on NACA
airfoils. Recently, Maeda et al. [10] studied the effects of
turbulence intensity on the static and dynamic characteristics
of DU93-W-210 airfoil at Re = 350 000 and two turbulence
levels of 0.15% and 11%. They found that the

ﬂow separa-

tion is delayed at higher turbulence levels, and the stall angle
increased.

Also, when the blade LE gathers dust, dirt, and so on, it

causes early transition of the BL. This roughness effect is
known to reduce the C

l,max

[2]. Figure 6 shows the effect of

roughness on C

l,max

at different Re. It can clearly be seen that

increasing roughness reduces C

l,max

. Airfoils with a relative

thickness of 25% or more may suffer a signi

ﬁcant deteriora-

tion in performance because of roughness effects. LE rough-
ness adds thickness to the BL and shifts the location of
transition very close to the LE. The resulting thicker BL leads
to increased drag, a reduction in the effective camber and an
earlier stall because of the weakening of the BL. The associ-
ated reduction in C

l,max

depends on the airfoil geometry and

the degree of contamination of the airfoil LE. A study on the
effects of airfoil thickness and C

l,max

on roughness sensitivity

concluded that the roughness sensitivity is directly propor-
tional to the airfoil thickness [11]. Some turbine blade
designers choose a pro

ﬁle and a design a, which has the tran-

sition point close to the LE so that for any reason, if the BL
transitions earlier, the performance of the blade is not much
different from the case with free transition. The SERI series
of airfoils (discussed in the next section) were designed to

Figure 3. Location of transition on the upper surface of SG6043

airfoil at low Re and 5% Tu.

Figure 4. Location of transition on the upper surface of SG6043

airfoil at high Re and 5% Tu.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

831

Int. J. Energy Res. 2012; 36:829–844 © 2012 John Wiley & Sons, Ltd.
DOI: 10.1002/er

have the transition point close to the LE just prior to reaching
C

l,max

. It is interesting to note that below a Re of about 10

the roughened or turbulated airfoils perform better because
of the transfer of energy from outside the BL to the low
energy region inside the BL. However, above this Re, the
BL tends to become weaker for roughened or turbulated air-
foils and L/D drops [12].

4. SERI AIRFOILS

The development of special-purpose airfoils for HAWTs
began in 1984 as a joint effort between the NREL, formerly
the SERI, and Airfoils Incorporated. In all, nine airfoil

families were designed for different rotor sizes [13] using
the Eppler code [14,15]. By minimizing the energy losses
because of roughness effects, by optimizing airfoils

’ perfor-

mance characteristics for appropriate Re and thickness and
by limiting C

l,max

, an annual energy improvement of up to

35% was estimated with SERI series of airfoils. For small
to medium blade lengths, 11

–15% thick airfoils were

designed, whereas for large lengths, 16

–21% thick airfoils

were designed. Greater thicknesses were intended to provide
greater blade

ﬂap stiffness for tower clearance, lower blade

weight important for large machines, and to accommodate
aerodynamic overspeed control devices for stall-regulated
machines. Two of the airfoil families were thin and were suit-
able for downwind rotors of up to 10 m blade lengths and 20
to 100 kW power. Figure 7 shows the thin airfoil family
designed for pitch-controlled blades that did not have a
restrained C

l,max

. The tip-region airfoil (S803) is 11.5%

thick, whereas the root-region airfoil (S804) is 18% thick.
The maximum C

was found to be 1.5 at the respective Re.

For the stall-controlled blades of this size range, four thin air-
foils, S806A, S805A, S807 and S808, were designed. The
C

l,max

value for the tip-region airfoil, S806A, was only 1.1.

Another thick airfoils family, S820, S819 and S821, was
designed, which had similar characteristics; the tip-region
blade maximum thickness was 16%, which was to accom-
modate speed-control mechanism for stall-regulated turbines
and has a greater stiffness at the cost of slightly higher drag.

For blades of 10

–15 m length and with ratings of

100

–400 kW, a thick-airfoil family consisting of S826,

S825, S814 and S815 was designed. The geometries and
the design speci

ﬁcations are shown in Figure 8. The tip-

region airfoil was 14% thick and was found to have a
C

l,max

of 1.6. The root region (40% radius) airfoil, S814, is

24% thick and was designed with two primary objectives:
(i) to achieve a C

l,max

of at least 1.30 for Re = 1.5

which should not reduce with transition

ﬁxed near the LE

on both the surface; and (ii) to obtain low pro

ﬁle drag

Figure 5. Location of transition on the upper surface of SG6043

airfoil at low Re and 10% Tu.

Figure 6. Effect of surface roughness on the maximum lift coef

ﬁcient for NACA 63(420)-422 at different Re [2].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

832

DOI: 10.1002/er

coef

ﬁcients over the range of C

from 0.6 to 1.2 for the same

Re. With these objectives, the drag polar was plotted; the C

for the C

range of 0.5 to 1.2 increased only slightly because

of the elimination of signi

ﬁcant laminar separation bubbles

on the upper surface that are found on many laminar-

ﬂow

airfoils. The drag increases very rapidly outside the low-drag
range because the BL transition point moves quickly toward

the LE with increasing (or decreasing) C

. This feature results

in a LE that produces a suction peak at higher C

values,

which ensures that transition on the upper surface occurs
very close to the LE, and hence, the value of C

l,max

is insen-

sitive to roughness at the LE [16]. Based on these aspects, the
pressure distributions along the polar were deduced. The
desired pressure distribution for the higher C

angle is shown

Figure 7. Thin NREL airfoil family for medium blades (high tip C

l,max

) [13].

Figure 8. Thick NREL airfoil family for large blades (high tip C

l,max

) [13].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

833

Int. J. Energy Res. 2012; 36:829–844 © 2012 John Wiley & Sons, Ltd.
DOI: 10.1002/er

in Figure 9. As can be seen from the

ﬁgure, the suction peak

occurs just aft of the LE, which indicates that the transition
point moves very quickly towards the LE with increasing
C

leading to roughness insensitivity of C

l,max

5. DU AIRFOILS

The design objectives of the DU series of airfoils were to
keep sensitivity due to LE contamination and contour
imperfections of the nose as low as possible. The maxi-
mum lift capacity was held at moderate levels, to keep
the loss of lift due to surface contamination as small
as possible. It was found that the thicker versions of the
NACA airfoils that were employed for the root region were
suffering from a severe degradation of the performance
because of premature transition. It was felt that thick
airfoils in use that time had a restrained upper surface
thickness to avoid early separation. To compensate for
the resulting loss of lift, certain amount of lower surface
aft-loading was incorporated, giving the typical S-shape
to the DU series of airfoils [17]. Two of the airfoils
DU-91-W2-250 (25% thick) and DU-93-W-210 (21% thick)
are shown in Figure 10. The measured C

l,max

was found to be

1.37 for the DU-91-W2-250 airfoil, which was slightly less
than the design value. The maximum L/D was 128. The
sensitivity of the airfoil to distortion of the boundary layer
at the nose was investigated by applying zigzag tape of
0.35 mm thickness at the 5% chord station. Because of the

trip, C

l,max

dropped from 1.37 to 1.16, as shown in Figure 11.

However, after the stall, the C

value rose more sharply com-

pared with the case without the zigzag tape. To serve as an
intermediate airfoil between DU 91-W2-250 and outboard air-
foils with rather high camber, the 21% thick DU 93-W-210
was designed. The airfoil had a maximum L/D of 143 at
Re = 3

and a C

l,max

of 1.35. The airfoil model was exten-

sively used to experimentally verify the effect of vortex genera-
tors (VGs), Gurney

ﬂaps and trip wires. The VGs are solid tabs

mounted on the airfoil surface to promote mixing and avoid/
delay BL separation; the Gurney

ﬂaps are small tabs attached

to the lower surface to improve the lift, and trip wires are used
to trip the BL to turbulent.

6. RIS

AIRFOILS

The Ris

A1 series of airfoils were designed by Ris

National Laboratory for wind turbines with stall, active stall
or pitch regulation, for rotor sizes of about 600 kW [18].
Detailed wind tunnel testing was performed on three of these
airfoils: Ris

-A1-18, Ris-A1-21 and Ris-A1-24 [19].

The following were the characteristics of the airfoils:
(i) maximum lift coef

ﬁcient around 1.5 in natural conditions

for all airfoils; (ii) high lift : drag ratio also at high angles of
attack just below maximum lift; (iii) insensitivity to leading
edge roughness for the maximum lift coef

ﬁcient; and

(iv) trailing edge stall, smooth post-stall behaviour. Tests
conducted on all the airfoil sections with zigzag tape showed
that the airfoils were reasonably insensitive to LE roughness;
the value of C

l,max

reduced from about 1.4 to about 1.2

for Re = 1.6

. The combination of vortex generators

and Gurney

ﬂaps increased the C

l,max

to about 2.0. The

Ris

-A1-24 airfoil is also a popular choice for TCTs.

The Ris

B1 family of airfoils was designed to maxi-

mize the tangential component of the force (rather than
the lift) so that the contribution in the direction of rotation
is high. As described earlier, an increase in lift mainly
increases the rotor thrust, whereas the tangential compo-
nent increases only a little. The Ris

-B1 airfoils were

designed with the objective of maximizing C

rather than

. However, it was found that for most of the airfoils, C

and the C

peak at the same

a. Another design objective

was good geometric compatibility between the different
airfoil sections and good geometric properties for inboard
airfoils. The design point region was

(

a - a

) of the order

of 9

–14

. The design

a-region was chosen close to C

l,max

because it enables low rotor solidity and/or low rotational
speed. For the root region, a high C

l,max

allows a reduction

in solidity. However, the contribution to the overall torque

Figure 9. Desired pressure distribution for the higher C

angle for

the S814 airfoil [16].

Figure 10. Pro

ﬁles of DU-91-W2-250 and DU-93-W-210 airfoils [17].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

834

DOI: 10.1002/er

from the root region is limited, and aerodynamic shap-
ing for high C

should not happen at the expense of

the cross-section structure. They found that the applica-
tion of vortex generators in combination with Gurney
ﬂaps can increase C

l,max

. The Ris

B1 family of air-

foils with thickness greater than 24% were designed to
have a high C

l,max

, by ensuring the desirable cross-

section structure. Figure 12 shows the measured C

and C

for the 24% thick airfoil with vortex generators

at 20% chord alone and vortex generators in combina-
tion with triangular Gurney

ﬂaps of 1% height. The

combination of vortex generators and Gurney

ﬂaps led

to an increase of 34% in C

l,max

to a value of 2.17. This

could provide an attractive choice for the root part of a
wind turbine blade where reduction of solidity is a key
issue to reduce blade costs [7]. The airfoils discussed in
the above sections are employed in WTs of a number of

manufacturers such as Vestas, DeWind, GE-Wind,
Suzlon and RE-Power.

7. CAS-W1 AIRFOILS

The CAS-W1 family of airfoils was designed by the Chinese
Academy of Sciences. The airfoils have thicknesses of
15

–25%. The CAS-W1-250 airfoil has a maximum thickness

of 25% and a TE thickness of 0.6%. At Re = 3

, the air-

foil has very good aerodynamic characteristics with a C

l,max

of 1.7 at 15

in clean conditions and of 1.66 at the same

with LE roughness. The maximum L/D is 157.6 at

a = 6

[20]. Thus, it has better aerodynamics characteristics and
reduced sensitivity to LE roughness compared with the
Ris

-A1/B1-24 and DU-91-W2-250 airfoils. The airfoil is

also found to have stable stall characteristics.

Figure 11. Measured airfoil performance of DU-91-W2-250 [17].

Figure 12. The measured C

and C

values with vortex generators at 20% chord and a combination of vortex generators and Gurney

ﬂaps [7].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

835

Int. J. Energy Res. 2012; 36:829–844 © 2012 John Wiley & Sons, Ltd.
DOI: 10.1002/er

8. OTHER THICK AIRFOILS

Airfoils of 30% thickness are located at the inner 40% of the
blade, where both aerodynamics performance and structural
considerations are important factors. The location of the
maximum thickness and the thickness of the upper surface
depend on the structural design requirements. The airfoil
shapes of three dedicated 30% airfoils, FFA-W3-301,
DU-97-W-300 and AH-94-W-301, are shown in Figure 13.
The performance characteristics of DU-97-W-300 and
AH-94-W-301 airfoils were found to be similar except for
C

l,max

, which is slightly higher for the DU airfoil [21]. The

FFA-W3-301 airfoil had a lower C

l,max

recorded.

Two airfoils, one 18% thick and the other 25% thick, were

–30% compared with NACA64

-618 and DU-91-W2-

250 airfoils. This was achieved by modifying these pro

ﬁles

to increase the C

and to decrease the C

values.

9. ROUGHNESS SENSITIVITY OF
THICK AIRFOILS

Van Rooij and Timmer [21] reviewed the performance of a
number of airfoils: DU, FFA, S8xx, AH, Ris

and NACA

series for inboard region (thick airfoils). They used the
reference of zigzag tape with height of 0.35 mm at 0.05c to
simulate LE contamination. They also found that in the
inboard region, difference between two-dimensional (2D)

wind tunnel test results and actual turbine blade results are
signi

ﬁcant because of the effects of rotation. Table I presents

a comparison of the effect of roughness on the performance
of different 25% thick airfoils. Although there was a small
difference in the roughness simulation for different airfoils,
it can be seen from the table that the Ris

-A1-24, S814

and DU-91-W2-250 airfoils performed quite well with their
C

l,max

for rough conditions dropping only a little compared

with the clean conditions. The CAS-W1-250 airfoil,
discussed in section 7, performed very well with simulated
roughness, with its C

l,max

dropping from 1.71 to 1.66 at

a = 15

[20]. For the Ris

-B1 airfoils, insensitivity to LE

roughness was ensured by two additional design objectives:
(i) having suction side transition from laminar to turbulent
ﬂow in the LE region for angles of attack at C

l,max

; and (ii)

obtaining a high C

l,max

with simulated LE roughness [7].

For the 18% thick airfoil, the drop in C

l,max

was only 3.7%

at Re = 1.6

, whereas for the 24% thick airfoil, it was

7.4% with the standard zigzag tape. However, more severe
roughness caused reductions in C

l,max

of 12

–27%.

10. THICK FLATBACK AIRFOILS

For the inboard region of large WT blades, thick blunt TE or
ﬂatback airfoils were designed and tested [23]. These airfoils
were designed to provide structural and aerodynamic perfor-
mance advantages

—structurally, the sectional area for a given

maximum airfoil thickness increases, and aerodynamically,
C

l,max

and lift curve slope increase; also, the sensitivity to

roughness reduces [24]. The main advantage of TE thickness
is that it allows for a portion of the pressure recovery to take
place in the wake of the airfoil, reducing the severity of the
adverse pressure gradient on the suction side and, hence,
reduces the chance of

ﬂow separation. This results in improved

performance in both clean and soiled conditions and improves
lift characteristics. Figure 14 shows the pro

ﬁles and pressure

distributions of a sharp and a blunt TE airfoil. It can be seen
that the pressure gradient is mild for the thick TE case, and
the pressure recovery is continuing. This will alleviate the
tendency of early

ﬂow separation.

Figure 13. Pro

ﬁles of three 30% thick airfoils [21].

Table I. Effect of roughness on the performance of (about) 25% thick airfoils [21].

Con

ﬁguration

Clean

‘Rough’

Airfoil

(L/D)

max

l,max

(L/D)

max

l,max

Re = 3 10

DU-91-W2-250

127.6

1.37

61.8

1.16

NACA-63

421

-425

120

1.277

0.803

AH-93-W-257

120.7

1.41

1.04

S814

114.1

1.408

61.4

1.357*

Re = 1.6 10

FFA-W3-241

1.37

48.5

1.16**

Ris

-A1-24

1.36

1.17**

*Grit roughness at upper surface

– x/c = 0.02 and lower surface – x/c = 0.1.

**ZZ-tape at x/c = 0.05 on upper surface and at x/c = 0.1 on lower surface.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

836

DOI: 10.1002/er

Three thick TE airfoils were designed by adding different

thicknesses to both the sides of the TE of FB3500 airfoil
(35% thick), and their performance characteristics were
studied under different conditions. The improved perfor-
mance for the

ﬁxed transition case for the thick TE airfoils,

compared with the sharp TE airfoil demonstrated the reduc-
tion of LE soiling sensitivity [23]. However, a higher C

the obvious cost of a thick TE.

11. LOW RE AIRFOILS

A number of sites in many countries have regions of low
wind (4

–5 m/s). WTs for such winds require low Re airfoils.

Advances in the development of small WTs will provide
solutions to the energy requirements of many countries.
Researchers are making performance data sets of low Re
airfoils available, which will help in validating codes used
for design purposes. Aeroacoustic characteristics are also
studied by researchers [25] because many small WTs need
to be installed in populated areas where noise can be a major
issue. Smaller WTs normally rotate at high rpm, which
results in high tip speeds and hence higher noise. Figure 15
shows the pro

ﬁles of six of the airfoils that were studied in

detail for both aerodynamic and aeroacoustic performance,
and the reports are made available by NREL for designers.
All these are relatively thin airfoils with maximum thickness
varying from 8.5% to 16.5%. The suf

ﬁxes along with the

names of the airfoils refer to multiple versions of those
airfoils. The airfoils were tested over a range of low Re
(generally 100 000

–500 000). Some of these airfoils are

employed by WT manufacturers such as Aeromag, World-
Power Technologies and Southwest Wind Power. Apart
from these, a large number of low and high Re airfoils coor-
dinates, and data

ﬁles are available at UIUC’s Airfoil Data

Site [26]. Design of low Re airfoils is still ongoing [27].
One of the low Re airfoils, SG6043, was tested experimen-
tally in the low-speed wind tunnel at the University of the
South Paci

ﬁc and numerically with ANSYS-CFX for appli-

cations to WTs in the Paci

ﬁc Island countries. The Re was

varied from 38 000 to 300 000, and Tu was varied from 1%
to 10%. The

a versus C

and

a versus C

graphs are shown

in Figures 16 and 17, respectively [28]. It can be seen that
C

increases continuously up to 14

for the lowest Re of

38 000, after which it starts to decrease because of the sepa-
ration of the

ﬂow from the upper surface. For the Re of

100 000 and above, C

increases up to 16

and then starts to

decrease beyond this angle. It is interesting to note that the
drop in the C

values is gradual, as the point of separation

moves upstream from the TE. The highest values of C

were

recorded for the Re of 200 000. For this Re, the value of C

dropped only a little from 16

to 18

, indicating that the stall

for this airfoil is not sudden. The drag coef

ﬁcient increases

slowly with

a for lower angles; from the angle of 14

for

Re = 38 000 and of 16

for Re = 150 000, there is a signi

ﬁcant

increase in C

, indicating the onset of stalling process.

12. ACTIVE AND PASSIVE FLOW
CONTROL

Devices for

ﬂow control are commonly employed to improve

the aerodynamics performance of the blade section(s). The

Figure 14. Pressure distributions on the TR-35 and TR-35-10 airfoils

at Re = 4.5

and

a = 8

[24].

Figure 15. Pro

ﬁles of six low Re airfoils [25].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

837

Int. J. Energy Res. 2012; 36:829–844 © 2012 John Wiley & Sons, Ltd.
DOI: 10.1002/er

objective also may be to reduce extreme loads and to mitigate
fatigue loads. In general, the intent of these devices is to
delay or advance transition, to reduce or increase turbulence
or to prevent or promote separation. The ensuing effects
include lift enhancement, tangential force enhancement, drag
reduction, mixing improvement, heat transfer enhancement
and

ﬂow-induced noise reduction [29]. The overall beneﬁt

to the turbine performance with minimum tradeoff is nor-
mally the goal [30]. A review of the aerodynamic models
used to estimate the aerodynamic loads on WTs is presented
by Hansen and Madsen [31]. The commonly used passive
control devices are as follows:

ﬁxed VGs, which are small solid plates mounted on
the blade surface that promote mixing and mitigate
boundary layer separation. VGs that are appropriately
sized and correctly oriented produce coherent helical
vortex structures that cause mixing between the air in
the freestream and BL [32]. They are commonly used
to reduce

ﬂow separation and increase C

l,max

b) Gurney

ﬂaps, which are small tabs attached to the

lower surface of the airfoil in the vicinity of the TE
with a height that can vary from 1% to 5% of the

chord. Liebeck

’s [33] results showed a signiﬁcant

increment in lift compared with the baseline airfoil.
The increased lift was also measured and reported by
reference [17,34]. Liebeck suggested an optimal
Gurney

ﬂap height to be of the order of 1–2% of the

chord. The

ﬂap increases the pressure on the pressure

side, reduces pressure on the suction side and helps
the BL to remain attached till the TE. Gurney

ﬂaps

are preferable for the inner part of the blade to increase
C

, allowing for a smaller chord for the same lift. In

another work, Tongchitpakdee et al. [35] tested the
effect of Gurney

ﬂap at 7 and 15 m/s and yaw angles

of 0

, 10

and 30

for S809 airfoil. They found that

both C

and C

increase at the lower wind speed. The

increase in the tangential force because of the deploy-
ment of the

ﬂap led to an increase in torque and power.

Some of the active control devices are as follows:

a) circulation control devices, which involve tangen-

tially blowing a small high-velocity jet over a curved
surface such as a rounded TE causing the BL and the
jet sheet to remain attached to the surface because of
the Coanda effect and turn without separation. This
causes the rear separation point to move to the lower
surface resulting in enhanced circulation around the
airfoil and hence enhanced lift [35]. The enhanced
circulation resulting from the blowing resulted in
an increased loading over the entire blade section.
The pressure distributions showed enhanced suction
both near LE and TE. This resulted in an increase in
C

at the wind speed of 7 m/s and yaw angles of up to

[35].

b) traditional and nontraditional TE

ﬂaps, which change

the sectional camber by de

ﬂecting the TE portion of

the airfoil; the traditional

ﬂaps tend to be heavy and

slow and take up a large portion of the chord; nontra-
ditional

ﬂaps, on the other hand, have a quick activa-

tion, are lightweight and occupy less chord. With the
use of smart materials, the nontraditional

ﬂaps are

found to be aerodynamically superior giving improve-
ment in L/D [29].

c) active translational microtabs deployed normal to the

surface close to the T.E. a maximum deployment
height on the order of the BL thickness (1

–2% chord).

Deployment of the tabs effectively changes the
sectional camber of the rotor blade and the TE

ﬂow

conditions, thereby changing the aerodynamic charac-
teristics of the blade. The advantages of the microtab
are its low actuation power requirements, short actua-
tion time, simplicity and the fact that it requires
minimal changes in the way blades are manufactured.
Inclusion of gaps or serrations in the tabs was found to
yield higher gains in terms of L/D for a given C

compared with the solid tabs [36]. The overall drag
can be reduced with this concept. Similar concepts
are employed in miniature TE effectors (MiTEs) and
micro

ﬂaps [30].

Figure 16. Variation of C

with

a for SG6043 airfoil at different

Re [28].

Figure 17. Variation of C

with

a for SG6043 airfoil at different

Re [28].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

838

DOI: 10.1002/er

In addition to the above, various other passive and active

ﬂow control methods and devices are proposed and tested
[30,31]. MacPhee and Beyene [37] simulated the aeroelastic
response of a morphing airfoil subjected to variable loading
and found that superior lift to drag ratios are obtained over
a speci

ﬁed range of a. They claim that their passive pitch

control method would achieve the same end result as an
active pitch-control device without expensive control and
sensor equipment. Another method being investigated is
the modi

ﬁcation of the blade tip to enhance the performance

[38,39]. Thus, it is clear that numerous methods and devices
are developed and tested with the main objective of extract-
ing maximum power from the

ﬂowing wind. A rotor dynam-

ics model that predicts the rotor speed for various turbine
con

ﬁgurations, operating over a wide range of wind condi-

tions, is recently proposed by Pope et al. [40].

Of late, the concept of smart rotor control is catching up.

With increasing blade diameters, the need for more sophisti-
cated load control techniques has generated interest in locally
distributed aerodynamic control systems with build-in intelli-
gence on the blades. Such concepts are often named in
popular terms

‘smart structures’ or ‘smart rotor control’. A

comprehensive review of the smart rotor control research is
presented by Barlas and van Kuik [41].

13. BLADES FOR TCTS

One of the major advantages of TCTs is that, although the
source

ﬂuctuates, it is highly predictable, unlike wind which

is highly random and unpredictable both in magnitude and
direction. As the density of water is 800 times that of air,
the forces acting on the TCT rotor are greater compared with
WTs. Thus, the structural requirements for TCTs tend to lead
to thicker sections, particularly near the root, which will
increase the C

. A bigger challenge in the design of blades

for TCTs is to avoid cavitation and still obtain an acceptably
higher C

[42]. Cavitation occurs when the absolute value of

the pressure coef

ﬁcient is greater than the cavitation number,

. Unlike WTs, which are getting bigger and bigger in size,

TCTs are limited in size, mainly by the size of channels in
which they are placed. Fixed speed, passively stall-controlled
TCTs are likely to have a lower annual energy production
compared with variable speed, pitch-controlled ones.
However, the

ﬂow conditions for TCTs are much more

steady and predictable; hence, it would not be justi

ﬁed to

use complex control systems. Unlike the WT blades that
have undergone extensive research and development and
many airfoil families have been developed, the blades for
TCTs are either designed by the manufacturers with speci

ﬁc

objectives, or in many cases, relatively thick WT blades that
do not have strong suction peaks are chosen and modi

ﬁed or

used directly.

In a work to optimize the hydrofoil for a marine turbine,

an optimization method was employed by NREL with a
number of NACA and Ris

sections. The results again

converged only on thick hydrofoils. Thinner hydrofoils are
more susceptible to cavitation as they have large suction

peaks on the upper surface [43]. Figure 18 shows the pres-
sure distribution on the Ris

-A1-24 airfoil at a = 8

and Re =

1.6

. It can be seen from the

ﬁgure that the suction peak

for this airfoil is not large at this angle; thus, it will be free
from cavitation. The suction on the upper surface is higher
for the smooth airfoil compared with the case when LE
roughness is simulated. As a result, the value of C

l,max

drops

from about 1.4 to about 1.2 [18,19].

In a work by Coiro et al. [44], a hydrofoil for a marine

turbine was designed by modifying S805 section and tested
at Re = 500 000, which was the estimated Re at the tip. The
modi

ﬁed section was named GT1. Cavitation number was

estimated to be

= 4.1 at a velocity of 2.5 m/s at 10

Figure 19 shows the pressure distribution on the surface of
the hydrofoil at

a = 6

and Re = 500 000. The modi

ﬁcation

signi

ﬁcantly improved the pressure distribution and reduced

the suction peak. The maximum L/D of the GT1 airfoil was
122 compared with about 88 for the S805 airfoil at this Re.
Another popular airfoil that is used in TCTs is S814. Up to

Figure 18. Pressure distribution on the Ris

-A1-24 airfoil at a = 8

and Re = 1.6

[19].

Figure 19. Pressure distribution on the GT1 hydrofoil [44].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

839

Int. J. Energy Res. 2012; 36:829–844 © 2012 John Wiley & Sons, Ltd.
DOI: 10.1002/er

a = 8

, the suction peak is not very high and C

is about 1.3 at

Re = 1

, making it appropriate for these turbines. Goun-

dar et al. [45] designed and tested a hydrofoil that gives a
maximum C

of greater than 2 without causing cavitation.

Cavitation causes structural damage to turbine blades, the
pressures associated with bubble collapse are high enough
to cause failure of metals [46]. Cavitation deteriorates the
performance of the blades; it causes lift to decrease and drag
to increase. Batten et al. [42] developed models to show
when the cavitation inception occurs, how it will affect
the performance of marine current turbines and how cavita-
tion can be avoided by using a section with a higher camber
or by changing the pitch angle of the blades. They also
suggested the use of suitable blade materials such as

ﬁbre-

reinforced plastics. Mueller and Wallace [47] proposed the
development of new blade coatings to offer increased cavita-
tion resistance. However, it is better to avoid cavitation with
appropriate design [45

–47].

14. EFFECTS OF ROTATION

The effects of rotation on the performance of a turbine blade
have been a matter of great interest and investigation by
researchers for many years [48

–50]. However, a detailed

understanding of the effects of rotating is still lacking, as
the effects depend on a number of parameters such as blade
pro

ﬁle, orientation, length, rotational speed, wind character-

istics; the effects also vary along the length of a rotating
blade. Ronsten [48] from his wind tunnel measurements
found that at lower

a, there is a good agreement between

the pressure distribution and C

for rotating and non-rotating

blades up to moderate angles of attack at all radial locations.
The hub region experienced higher loadings at low tip speed
ratios than what was predicted using 2D data. Tangler [49]
evaluated the measured NASA Ames Unsteady Aerodynamic
Experiment post-stall blade element data and provided
guidelines for developing an empirical approach that predicts
post-stall airfoil characteristics. The three-dimensional effects
were found to cause delayed stall with higher C

values

compared with the 2D data, especially near the root region.
Sicot et al. [50] studied the effects of rotation and turbulence
on the turbine blade performance in a wind tunnel. The Re on
the blade was 1.5

to 4

. They found that at lower

angles of attack, the pressure distributions on the rotating and
non-rotating blades are similar, as can be seen from Figure 20
for

a = 4.8

. At stall and post-stall angles, they recorded a

stronger suction on the suction surface for a rotating blade
compared with the non-rotating case. Figure 21 shows the
pressure distributions for

a = 20

and Re = 4

. It can

be seen that the pressure on the lower surface is slightly higher
for the rotating case; on the other hand, the suction on the
upper surface is much stronger for the rotating blade com-
pared with the parked one, indicating an augmentation of
forces. Similar rotational augmentation of forces including
stall delay and lift enhancement was reported after extensive
wind tunnel experimentation by Schreck and Robinson [51].
They employed the S809 airfoil between 0.25R and the tip.

Local in

ﬂow angles (LFA) were measured using ﬁve-hole

probes. LFA was de

ﬁned as the angle between the local

ﬂow vector and the local blade chord line, measured at the

probe tip. The measured LFAs at three locations for values
of U

ranging from 5 to 25 m/s for zero yaw operation are

shown in Figure 22. At 5 m/s, LFA began at approximately
5

for all three radial locations. At higher U

values, the

LFA plots started diverging, with the maximum divergence
for the 0.34R location. Thus, higher values of U

prompted

higher LFA. The variations of C

with LFA for rotating and

non-rotating blades are shown in Figure 23. At lower LFAs,

Figure 20. Comparison of pressure distributions on a rotating

and non-rotating blade for

a = 5

and Re = 4

[50].

Figure 21. Comparison of pressure distributions on a rotating

and non-rotating blade for

a = 20

and Re = 4

[50].

Figure 22. Measured LFA as a function of freestream velocity

at three locations [51].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

840

DOI: 10.1002/er

the C

values rise linearly and the slopes for rotating and non-

rotating blades are similar. However, at higher LFAs, the C

values for the rotating blade rise steeply and attain a maxi-
mum at 45

. They observed considerable stall delay and stall

ampli

ﬁcation at other blade locations of 0.47R, 0.63R and

0.8R [52]. However, stall delay and C

ampli

ﬁcation were

consistently greater for radial locations farther inboard on
the blade.

When the in

ﬂow is yawed (the rotor plane yawed with

respect to U

), the vectors

Ωr and U

are no longer perpen-

dicular. Thus, local in

ﬂow magnitude and LFA vary during

the rotation cycle, depending upon whether the blade section
is advancing upstream into the wind or retreating down-
stream from it. As a consequence, LFA no longer remains
constant as the blade rotates in azimuth. For a U

13 m/s and yaw angle of 40

, the pressure distributions on

the blade at 0.47R are shown in Figure 24. As can be seen
from the

ﬁgure, the dynamic stall vortex (indicated by the

suction peak) was

ﬁrst detected just aft of the LE at 0.04c

and corresponding to

c = 72.8

. Thirty-

ﬁve msec later,

when

c = 57.8

, the suction peak had moved to 0.36c,

where it had broadened and magnitude decreased to 2.42.
After another 66 msec, when

c = 29.4

, the suction peak

had moved further downstream to 0.68c and was no longer
distinct (because of the growth of the vortex), and the

–C

magnitude was 1.42. Olorunsola [53] presented results of
yaw force at different wind speeds and found that at non-zero
yaw angles, the yaw force increases signi

ﬁcantly at high

wind speeds.

It is known that changes in wind magnitude and direction

are frequent; these can signi

ﬁcantly affect the performance of

the turbine, as discussed above and may cause rotational
augmentation or dynamic stall. Similarly, the turbulence
level of the wind changes, which is likely to compound the
adverse effects discussed above. Although the modern tur-
bines are equipped with control systems, the changes in wind
speed and direction may outpace the control actuation rates,
leading to the problems discussed. This particular detailed
study was performed on a particular blade under speci

ﬁc con-

ditions; the extent of force augmentations may be different
for different blade pro

ﬁles, sizes and operating conditions.

14.1. Tip speed ratio and wake rotation losses

A higher tip speed ratio (higher rotational speed) means the
desired power can be generated by a lower torque (reduced
weight of rotor shaft and gearbox). A three-bladed WT rotor
performs optimally at a tip speed ratio of 7 to 8. However, a
higher tip speed ratio means a higher noise level. At tip
velocities greater than 70 m/s for WTs, noise becomes a
major concern. For offshore WTs, it is possible to exceed
tip velocities beyond 70 m/s, as noise is not a major concern.
For TCTs, optimum performance is achieved at a tip speed
ratio of 3 to 4 [44]. However, for both WTs and TCTs, the
strength of the structure is a concern at high tip speed ratios
as the force on the hub increases. A challenge for the
designers is thus to design blades that are thin, light-weight
and still capable of taking the load.

Ideally, a turbine should have no wake rotation losses.

Slow-rotating, high torque turbines experience more wake
rotation losses compared with high-speed turbines with
low torque [54], which again requires that blades should
rotate at high speeds.

15. FUTURE WORK

• As discussed above, different blade proﬁles/blades and

ﬂow conditions will encounter different levels of force
augmentation and dynamic stall while in operation. It
would be good to incorporate some, if not all, of these
aspects into initial design. Design of blade should
continue till it is tested in the

ﬁeld for an extended

period and then modi

ﬁed by addressing the problems

encountered. A challenge is to reduce uncertainty levels
in

ﬁeld measurements.

• A problem with ﬁeld measurements is the high level of

unsteadiness in the wind that the researchers have no
control over. This is compounded by variations in
temperature and pressure. This not only introduces
uncertainties of a higher order but also calls for a more
accurate assessment of the available energy [55].

Figure 23. Measured C

values at different LFAs at 0.3R for rotating

and non-rotating blade [51].

Figure 24. Measured pressure distribution at three

c values for

= 13 m/s, yaw angle of 40

and at 0.47R [51].

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

841

• Looking at the proﬁles of most of the thick airfoils

(e.g., Figures 10 and 13), it can be seen that at slightly
higher angles, the pressures on both the surfaces do
not contribute signi

ﬁcantly to C

. The added lower

surface thickness does not help contribute much to C

although some contribution may come from the part
aft of the maximum thickness location. If the objective
is to get a strong component in the direction of rotation,
then the section design needs to be modi

ﬁed. A good

contribution to lift needs to come from the region close
to the LE such that there is a component in the negative
drag direction (this will also reduce drag). Some active
and passive devices, for example, in reference [35] have
shown positive results in terms of increasing C

, but the

increased as well.

• Although there are some attempts to get more contribu-

tion to the rotation from the root region (e.g., reference
[7]), a lot of work needs to be performed on the root
region airfoils and supplementary devices that enhance
the contribution from this region.

• With the help of a feedback control system, it is possible

to reduce unsteady loads with dynamically deployed
Gurney

ﬂaps [56]. However, more work is needed on

this type of

ﬂaps.

• With a growing concern about the adverse effects of

large WTs and WT farms on weather, future research
works also should pay equal attention to the turbulence
and mixing caused in the wake. At the same time, the
generation of rotational kinetic energy in the wake
results in less energy extraction by the rotor than would
be expected without wake rotation.

• With a growing interest in wind farms, detailed investi-

gation of wakes assumes more signi

ﬁcance and impor-

tance. This is another area, which needs more attention.

• For offshore WTs and TCTs, maintenance and repairs

are normally expensive and time consuming. Taking
this into consideration will lead to lean design.

• High-design lift coefﬁcients allow design of slender

blades while maintaining high aerodynamic ef

ﬁciency.

Also, the rated power will be reached at relatively low
wind speed.

• Another challenge is to design blades that give consis-

tent output at varying

ﬂow conditions with a simple

control system.

• It is disappointing to see some large wind turbines

employing very thin blades (almost like

ﬂatplates)near

the tip for structural reasons. The engineering commu-
nity needs to strive to overcome this problem.

16. CONCLUSIONS

• After nearly two decades of extensive research, a num-

ber of airfoils are available; however, majority of them
are thick airfoils as the strength is still a major concern.

• Many of the airfoil designs started with the intention of

reducing the effects of LE contamination. However, still
most of the airfoils show a deterioration in performance

with contamination. The effect of LE roughness
increases with thickness. For sections less than 18%
thick, effect of roughness on the lift-curve slope is small.

• There is still a concern about increasing C

(as it

increases the thrust more). Some of the designs started
with the intention of maximizing the tangential force
coef

ﬁcient; however, it is still a challenge.

• To maximize the tangential force component, a good

contribution to lift needs to come from the region
close to the LE such that there is a component in the
negative drag direction (this also will reduce drag).
It is achieved partially with active and passive

ﬂow

control devices, but the dif

ﬁculty of increasing C

without increasing C

much still remains.

NOMENCLATURE

= chord length, m

= coef

ﬁcient of drag

= coef

ﬁcient of lift

l,max

= maximum coef

ﬁcient of lift

= normal force coef

ﬁcient

= coef

ﬁcient of pressure

= tangential force coef

ﬁcient

= drag force, N

= lift force, N

= radial distance from hub, m

= blade length, m

= Reynolds number

= turbulence intensity, %

= freestream mean velocity, m/s

= distance along chord from the leading

edge, m

= distance along chord where transition

occurs, m

= angle of attack, degrees

= blade azimuth angle, degrees

= cavitation number

= blade rotation rate, rad/s

Acronyms

HAWT

= horizontal axis wind turbine

LFA

= local in

ﬂow angle

= leading edge

TCT

= tidal current turbine

= trailing edge

= vortex generator

= wind turbine

REFERENCES

1. Xuan H, Weimin Z, Xiao L, Jieping L. Aerodynamic

and aeroacoustic optimization of wind turbine blade
by a genetic algorithm. Paper no. AIAA-2008-1331.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

842

DOI: 10.1002/er

Proceedings of 46th AIAA Aerospace Sciences Meet-
ing and Exhibit Reno, Nevada; 2008.

2. Abbott IH, Van Doenhaff AE. Theory of wing sections

(1st edn). Dover Publications Inc.: New York, 1959.

3. Fupeng H, Yuhong L, Zuoyi C. Suggestions for

improving wind turbines blade characteristics. Wind
Engineering 2001;

25:105–113.

4. Hansen AC, Butter

ﬁeld CP. Aerodynamics of horizon-

tal-axis wind turbines. Annual Review of Fluid
Mechanics 1993;

25:115–149.

5. van Bussel GJW. Technology challenges for offshore

wind power. Paper no. O-Wd-9-1, Proceedings of
RE2010 International Conference and Exhibit, Yokohama,
Japan, 2010.

6. Somers DM. Design and experimental results for the

S809 airfoil. NREL Report NREL/SR-440-6918; 1997.

7. Fuglsang P, Bak C, Gaunaa M, Antoniou I. Design and

veri

ﬁcation of the Ris-B1 airfoil family for wind

turbines. Journal of Solar Energy Engineering 2004;
126:1002–1010.

8. Hoffmann JA. Effect of freestream turbulence on the

performance characteristics of an airfoil. AIAA Journal
1991;

29:1353–1356.

9. Devinant P, Laverne T, Hureau J. Experimental study

of wind-turbine airfoil aerodynamics in high turbu-
lence. Journal of Wind Engineering and Industrial
Aerodynamics 2002;

90:689–707.

10. Maeda T, Kamada Y, Murata J, Toki T, Tobuchi A.

Effect of turbulence intensity on dynamic characteristics
of wind turbine airfoil. Paper No. P-Wd-36, Proceedings
of RE2010 International Conference and Exhibit,
Yokohama, Japan. 2010.

11. Tangler JL. Effects of airfoil thickness and maximum

lift on roughness sensitivity, Proceedings of the 3rd
ASME/JSME Fluids Engineering Conference, CA,
USA. July 1999.

12. Lissaman PBS. Low-Reynolds-number airfoils. Annual

Review of Fluid Mechanics 1983;

15:223–239.

13. Tangler JL, Somers DM. NREL airfoil families

for HAWTs. online at http://wind.nrel.gov/airfoils/
Documents/AirfoilDocuments.html, 1995.

14. Eppler R. Airfoil design and data. Springer-Verlag:

Berlin, 1990.

15. Eppler R. Airfoil program system. User

’s guide. R.

Eppler, c. 1993.

16. Somers DM. Design and experimental results for the

S814 airfoil. NREL Report NREL/SR-440-6919; 1997.

17. Timmer WA, van Rooij PRJOM. Summary of the

Delft University wind turbine dedicated airfoils. AIAA
paper AIAA-2003-0352, 2003.

18. Fuglsang P, Dahl KS. Design of the new Ris

-A1 airfoil

family for wind turbines. Proceedings of EWEC

’99,

Nice, France. 1999.

19. Fuglsang P, Dahl KS, Antoniou, I. Wind tunnel tests of

the Ris

-A1-18, Ris-A1-21 and Ris-A1-24 airfoils.

Ris

report Ris-R-1112, 1999.

20. Xu Y. Introduction to fundamental research at IET.

Proceedings of APSOWET 2010, Mokpo, Korea,
2010; 175

–187.

21. van Rooij PRJOM, Timmer WA. Roughness sensitiv-

ity considerations for thick rotor blade airfoils. AIAA
paper AIAA-2003-0350, 2003.

22. Lee YG, Ahn SM, Yeom CH, Lee DS. Airfoil lift and

drag effects on ef

ﬁciencies of a wind turbine blade. Paper

no. S46_2, Proceedings of WWEC2009 Conference,
Jeju, Korea; 2009.

23. Baker JP, Mayda EA, van Dam CP. Experimental anal-

ysis of thick blunt trailing edge wind turbine airfoils.
ASME Journal of Solar Energy Engineering 2006;
128:422–431.

24. Standish KJ, van Dam CP. Aerodynamic analysis of

blunt trailing edge airfoils. ASME Journal of Solar
Energy Engineering 2003;

125:479–487.

25. Selig M, McGranahan BD. Wind tunnel aerodynamics

tests of six airfoils for use on small wind turbines.
ASME Journal of Solar Energy Engineering 2004;
126:986–1001.

26. Selig M. UIUC airfoil data site: http://www.ae.illinois.

edu/m-selig/ads.html

27. Singh RK, Ahmed MR, Zullah MA, Lee YH. Design

of a low Reynolds number airfoil for small horizon-
tal axis wind turbines. Renewable Energy 2012;
42:66–76.

28. Ahmed MR, Narayan S, Zullah MA, Lee YH.

Experimental and numerical studies on a low
Reynolds number airfoil for wind turbine applica-
tions. Journal of Fluid Science and Technology
2011;

6:357–371.

29. Gad-el-Hak M. Overview of turbulence control research

in U.S.A. Proceedings of the Symposium on Smart Con-
trol of Turbulence, Kasagi N (ed.). University of Tokyo:
Tokyo, Japan, 2

–3 Dec. 1999; 1–20.

30. Berg D, Johnson SJ, Van Dam CP. Active load control

techniques for wind turbines. Sandia National Labs
Report SAND2008-4809, 2008.

31. Hansen MOL, Mades HA. Review paper on wind

turbine aerodynamics. ASME Journal of Fluids Engi-
neering 2011;

133: 114001-1-12.

32. Lin JC. Review of research on low-pro

ﬁle vortex

generators to control boundary layer separation. Prog-
ress in Aerospace Sciences 2002;

38:389–420.

33. Liebeck RH. Design of subsonic airfoils for high lift.

Journal of Aircraft 1978;

15:547–561.

34. Li Y, Wang J, Zhang P. Effect of Gurney

ﬂaps on a

NACA0012 airfoil. Flow, Turbulence and Combustion
2002;

68:27–39.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

843

35. Tongchitpakdee C, Banjanirat S, Sankar LN. Numerical

studies of the effects of active and passive circulation
enhancement concepts on wind turbine performance.
ASME Journal of Solar Energy Engineering 2006;
128:432–444.

36. Mayda EA, Van Dam CP, Nakafuji DY. Computa-

tional investigation of

ﬁnite width microtabs for aero-

dynamic load control. AIAA Paper AIAA-2005-1185.

37. MacPhee D, Beyene A. Fluid

–structure interaction of a

morphing symmetrical wind turbine blade subjected to
variable load. International Journal of Energy Research
2011; doi: 10.1002/er.1925 (available online).

38. Shimizu Y, Ismaili E, Kamada Y, Maeda T. Power

augmentation of a HAWT by Mie-type tip vanes,
considering wind túnnel

ﬂow visualization, blade aspect

ratio and Reynolds number. Wind Engineering 2003;
27:183–194.

39. Gertz D, Johnson DA. An evaluation testbed for wind

turbine blade tip designs

– baseline case. International

Journal of Energy Research 2011;

35:1360–1370.

40. Pope K, Milman R, Naterer GF. Rotor dynamics correla-

tion for maximum power and transient control of wind
turbines. International Journal of Energy Research
2010;

34:736–742.

41. Barlas TK, van Kuik GAM. Review of state of the art

in smart rotor control research for wind turbines. Prog-
ress in Aerospace Sciences 2010;

46:1–27.

42. Batten WMJ, Bahaj AS, Molland AF, Chaplin JR.

Hydrodynamics of marine current turbines. Renewable
Energy 2006;

31:249–256.

43. Sale D, Jonkman J, Musial W. Hydrodynamic optimi-

zation method and design code for stall-regulated hy-
drokinetic turbine rotors. NREL Report NREL/
CP500-45021; 2009.

44. Coiro DP, Maisto U, Scherillo F, Melone S, Grasso F.

Horizontal axis tidal current turbine: numerical and
experimental investigations. Proceedings of Owemes,
Civitavecchia, Italy; 2006.

45. Goundar JN, Ahmed MR, Lee YH. Numerical and

experimental studies on hydrofoils for marine current
turbines. Renewable Energy 2012;

42:173–179.

46. Eisenberg P. Mechanics of cavitation, Hydronautics

incorporated, 1950.

47. Mueller M, Wallace R. Enabling science and technol-

ogy for marine renewable energy. Energy Policy
2008;

36:4376–4382.

48. Ronsten G. Static pressure measurements on a rotat-

ing and non-rotating 2.375 m wind turbine blade.
Comparison with 2D calculations. Journal of Wind
Engineering and Industrial Aerodynamics 1992;
39:105–118.

49. Tangler JL. Insight into wind turbine stall and post

stall aerodynamics. Wind Energy 2004;

7:247–260.

50. Sicot C, Devinant P, Loyer S, Hureau J. Rotational

and turbulence effects on a wind turbine blade, in-
vestigation of the stall mechanisms. Journal of Wind
Engineering and Industrial Aerodynamics 2008;
96:1320–1331.

51. Schreck S, Robinson M. Wind turbine blade

ﬂow

ﬁelds and prospects for active aerodynamic control.
NREL report, NREL/CP-500-41606; 2007.

52. Schreck S, Robinson M. Rotational augmentation of hor-

izontal axis wind turbine blade aerodynamic response.
Wind Energy 2002;

5:133–150.

53. Olorunsola O. On the free yaw behaviour of horizontal

axis wind turbines. International Journal of Energy
Research 1986;

10:343–355.

54. Hau E. Wind turbines (2nd edn). Springer-Verlag:

Berlin Heidelberg, 2006.

55. Sahin AD, Dincer I, Rosen MA. Thermodynamic anal-

ysis of wind energy. International Journal of Energy
Research 2006;

30:553–566.

56. Frederick M, Kerrigan EC, Graham JMR. Gust allevia-

tion using rapidly deployed trailing-edge

ﬂaps. Journal

of Wind Engineering and Industrial Aerodynamics
2010;

98:712–723.

Blade sections for wind and tidal current turbines—status and future

M. R. Ahmed

844

DOI: 10.1002/er

Wyszukiwarka

Podobne podstrony:
Compliant Blades For Wind Turbines
Synchronous Generator And Frequency Converter In Wind Turbine Applications System Design And Efficie
A Low Speed, High Torque, Direct Drive Permanent Magnet Generator For Wind Turbines
Wind Turbine 5 Metre Diameter Carbon Fibre Blades For Wind Turbine10Kwblades
(Wind) A Low Speed, High Torque, Direct Drive Permanent Magnet Generator For Wind Turbines
Design Of Direct Driven Permanent Magnet Generators For Wind Turbines
0 Power Control for Wind Turbines in Weak Grids H Bindner 1999
Adjustable Speed Generators For Wind Turbines Based On Doubly
Design Fatigue Test and NDE of a Sectional Wind Turbine Rotor Blade
[US 2006] D517986 Wind turbine and rotor blade of a wind turbine
[US 2006] D517986 Wind turbine and rotor blade of a wind turbine
Design Requirements For Medium Sized Wind Turbines For Remote And Hybrid Power Systems
Darrieus Wind Turbine Design, Construction And Testing
Plans For Wind Generator Pt250 Blade Plan10A
Innovative Solutions In Power Electronics For Variable Speed Wind Turbines
(WinD Power) Dynamic Modeling of Ge 1 5 And 3 6 Wind Turbine Generator {}[2003}
Modeling Of The Wind Turbine With A Doubly Fed Induction Generator For Grid Integration Studies

więcej podobnych podstron

Blade sections for wind turbine and tidal current turbine applications—current status and future challenges