Burton T. (et. al.) Wind energy Handbook

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accidental release and the occasional need to deactivate them for maintenance

purposes; the rotor locking device should act on the low-speed shaft, so that its

effectiveness is not dependent on the integrity of the gearbox – typically the

device consists of a pin mounted in a ﬁxed housing, which can be engaged in a

hole in a shaft-mounted disc;

• guards to shield any rotating parts within the nacelle;

• suitable ﬁxtures for the attachment of safety harnesses for personnel working

outside the nacelle.

The designer needs to assess the requirement for all-weather access to the nacelle

at an early stage. Lattice towers afford no protection from the weather when

climbing, so the number of days on which access for maintenance is possible will be

restricted. Similar restrictions will arise if the nacelle cover has to be opened to the

elements in order to provide space for personnel to enter.

Consideration also needs to be given to the means of raising and lowering tools

and spares. If the interior of the tower is interrupted by intermediate platforms,

these operations have to be performed outside, with consequent weather limit-

ations.

Standard rules for electrical safety apply to all electrical equipment. However,

particular care must be taken with the routing of electrical cables between tower

and nacelle, in order to avoid potential damage due to chaﬁng when they twist. If

the power tran sformer is located in the tower bas e or nacelle instead of in a separate

enclosure at ground level, it should be partitioned off to minimize the ﬁre risk to

personnel.

References

Armstrong, J. R. C. and Hancock, M., (1991). ‘Feasibility study of teetered, stall-regulated

rotors’. ETSU Report No. WN 6022.

Bossanyi, E. A. and Gamble, C. R., (1991). ‘Investigation of torque control using a variable

slip induction generator’ ETSU WN-6018, Energy Technology Support Unit, Harwell, UK.

Burton, A. L., Mill, P. W. and Simpson, P. B., (1990). ‘LS1 post-synchronization commission-

ing’. Proceedings of the 12th BWEA Conference, pp 183–193. Mechanical Engineering Pub-

lications, Bury St Edmunds, UK.

Coiante, D. et al., (1989). ‘Gamma 60 1.5 MW wind turbine generator’. Proceedings of the

European Wind Energy Conference, pp 1027–1032.

Corbet, D. C. and Morgan, C. A., (1991). ‘Passive control of horizontal axis wind turbines’.

Proceedings of the 13th BWEA Conference, pp 131–136. Mechanical Engineering Publications,

Bury St Edmunds, UK.

Derrick, A., (1992). ‘Aerodynamic characteristics of novel tip brakes and control devices for

HAWTs’. Proceedings of the 14th BWEA Conference, pp 73–78. Mechanical Engineering

Publications, Bury St Edmunds, UK.

Engstrom, S. et al., (1997). ‘Evaluation of the Nordic 1000 Prototype’. Proceedings of the

European Wind Energy Conference, pp 213–216.

REFERENCES 375

Falchetta, M. et al., (1996). ‘Structural behaviour of the Gamma 60 prototype’. Proceedings of

the European Union Wind Energy Conference, pp 268–271.

Fuglsang, P. and Thomsen, K., (1998). ‘Cost optimization of wind turbines for large-scale

offshore wind farms’. Riso National Laboratory Report No. R-1000. Riso National Laboratory,

Roskilde, Denmark.

Hindmarsh, J., (1984). Electrical machines and their applications. Butterworth Heinemann, UK.

Henderson, G. M., et al., (1990). ‘Synchronous wind power generation by means of a torque-

limiting gearbox’. Proceedings of the 12th BWEA Conference, pp 41–46. Mechanical Engi-

neering Publications, Bury St Edmunds, UK.

Jamieson, P. and Agius, P., (1990). ‘A comparison of aerodynamic devices for control and

overspeed protection of HAWTs’. Proceedings of the 12th BWEA Conference, pp 205–213.

Mechanical Engineering Publications, Bury St Edmunds, UK.

Jamieson, P. and Brown, C. J., (1992). ‘The optimization of stall-regulated rotor design’.

Proceedings of the 14th BWEA Conference, pp 79–84, Mechanical Engineering Publications,

Bury St Edmunds, UK.

Joose, P. A. and Kraan, I., (1997). ‘Development of a tentortube for blade tip mechanisms’.

Proceedings of the European Wind Energy Conference, pp 638–641.

Laithwaite, E. R. and Freris, L. L., (1980). Electric energy: its generation, transmission and use.

McGraw-Hill, Maidenhead, UK.

Law, H., Doubt, H. A. and Cooper, B. J., (1984). ‘Power control systems for the Orkney wind-

turbine generators’. GEC Engineering No 2.

Leithead, W. E. and Rogers, M. C., (1995). ‘Improving damping by a simple modiﬁcation

to the drive train’. Proceedings of the 17th BWEA Conference, pp 273–278. Mechanical

Engineering Publications, Bury St Edmunds, UK.

McPherson, G., (1990). An introduction to electrical machines and transformers. Second Edition.

John Wiley and Sons, New York, US.

Pedersen, T. K., (1995). ‘Semi-variable speed – a compromise?’. Proceedings of the Wind Energy

Conversion, 17th BWEA Conference, pp 249–260. Mechanical Engineering Publications, Bury

St Edmunds, UK.

Morgan, C., (1994). ‘The prospects for single-bladed horizontal axis wind turbines’. ETSU

Report No. W/45/00232/REP. Energy Technology Support Unint, Harwell, UK.

Petersen, T. P., et al., (1998). ‘Prediction of dynamic loads and induced vibrations in stall’.

Riso National Laboratory Report No. R-1045. Riso National Laboratory, Roskilde, Denmark.

Powles, S. J. R., (1983). ‘The effects of tower shadow on the dynamics of horizontal axis wind

turbines’. Wind Engng, 7, 1, 26–42.

Rawlinson-Smith, R. I., (1994). ‘Investigation of the teeter stability of stalled rotors’. ETSU

Report No. W.43/00256/REP. Energy Technology Support Unit, Harwell, UK.

Saad-Saoud, Z. and Jenkins, N., (1999). ‘Models for predicting ﬂicker induced by large wind

turbines’. IEEE Transactions on Energy Conversion, 14, 3, 743–748.

376

CONCEPTUAL DESIGN OF HORIZONTAL-AXIS TURBINES

Component Design

7.1 Blades

7.1.1 Introduction

A successful blade design must satisfy a wide range of objectives , some of which

are in conﬂict. These objectives can be summarized as follows:

(1) maximize annual energy yield for the speciﬁed wind speed distribution;

(2) limit maximum power output (in the case of stall regulated machines);

(3) resist extreme and fatigue loads;

(4) restrict tip deﬂections to avoid blade/tower collisions (in the case of upwind

machines);

(5) avoid resonan ces;

(6) minimize weight and cost.

The design process can be divided into two stages: the aerodynamic design, in

which objective s (1) and (2) are satisﬁed, and the structural design. The aerody-

namic design addresses the selection of the optimum geometry of the blade external

surface – normally simply referred to as the blade geometry – which is deﬁned by

the aerofoil family and the chord, twist and thickness distributions. The structural

design consists of blade material selection and the determination of a structural

cross section or spar within the external envelope that meets objectives (4) to (6).

Inevitably there is interaction between the two stages, as the blade thickness needs

to be large enough to accommodate a spar which is structurally efﬁcient.

The focus of Section 7.1 is on blade structural design. After a brief consideration

of the aerodynamic design in Section 7.1.2, practical constraints on the optimum

design are noted in Section 7.1.3 and forms of blade structure surveyed in Section

7.1.4. An overview of the properties of some potential blade materials is given in

Section 7.1.5 and the properties of glass-ﬁbre reinforced plastic (GFRP) and

laminated wood are considered in more detail in Sections 7.1.6 and 7.1.7. Governing

load cases are conside red in Sections 7.1.8 with reference to both stall- and pitch-

regulated machines. Subsequent sections touch upon blade resonance, panel buck-

ling design and blade root ﬁxings.

7.1.2 Aerodynamic design

The aerodynamic design encompasses the selection of aerofoil family and optimiza-

tion of the chord and twist distributions. The variation of thickness to chord ratio

along the blade also has to be considered, but this ratio is usually set at the

minimum value permitted by structural design considerations, as this minimizes

drag losses.

The process for optimizing the blade design of machines operating at a ﬁxed tip

speed ratio is described in Section 3.7.2, where analytical expressions for the blade

geometry parameter,



ºC

Nc()

2R

ºC

and the local inﬂow angle, , are derived as a function of the local tip speed ratio,

º ¼ ºr=R. (Equations (3.67a) and (3.68a)). If º  1, the expressions can be

approximated by



ºC

Nc()

2R

ºC

9º

and  ¼

3º

(7:1)

If it is decided to maintain the angle of attack, Æ, and hence the lift coefﬁcient, C

constant along the blade, then these relations translate to

c() ¼

16R

Nº



and  ¼

3º

 Æ (7:2)

so that both the chord and twist are inversely proportional to radius.

In the case of machines operating at constant rotational speed, and hence at

varying tip speed ratio, no parallel analytical solution for the optimum blade

geometry exists. Instead resort must be made to numerical methods based on blade

element – momentum theory, for example using Equations (3.51b) and (3.52a) in

Section 3.8.6.

For pitch-regulated machines, the annual energy capture attributed to the annular

ring swept out by each blade element is determined for the chosen wind speed

distribution, and the variation of energy capture with blade chord and twist at each

‘blade station’ computed. In this way the values of blade chord and twist at

each ‘blade station’ yielding maximum energy capture are identiﬁed.

For stall-regulated machines, the method is similar, but the total ann ual energy

capture has to be maximized within the constraint of limiting the maximum total

power output to the machi ne rating. The results of such an investigation are

reported by Fuglsang and Madsen (1995).

378 COMPONENT DESIGN

7.1.3 Practical modiﬁcations to optimum design

The result of the optimization desc ribed in the previous section is typically a blade

geometry in which both blade chord and blade twist vary appr oximately inversely

with radius, as illustrated in Figure 3.19. However, because the inboard section of

the blade makes only a small contribution to total power output (Figure 3.30), the

aerofoil section is generally not continued inboard of about 15 percent radius in

practice, and the chor d at this radius is substantially reduced, to perhaps half the

theoretical optimum. It is then often found expedient to taper the chord uniform ly

over the active length of the blade, with the tip chord and chord taper set so that the

chord distribution approximates closely to the optimum over the outboard half of

the blade (Figure 3.20).

The blade root area is normally circul ar in cross section in order to match up with

the pitch bearing in the case of pitchable blades, or to allow pitch angle adjustment at

the bolted ﬂange (to compensate for non-standard air density) in the case of stall-

regulated blades. The transition from the root section to the aerofoil section outboard

of 15 per cent radius should be a smooth one for structural reasons, with the result that

the latter section will have a high thickness to chord ratio of up to about 50 percent.

7.1.4 Form of blade structure

A hollow shell corresponding to the deﬁned blade envelope clearly provides a

simple, efﬁcient structure to resist ﬂexural and torsional loads and some blade

manufacturers adopt this form of construction (see Figure 7.1). However, in the case

of small and medium size machines, where the out-of-plane loads dominate, there

is greater beneﬁt in concentrating skin material in the forward half of the blade,

where the blade thickness is a maximum, so that it acts more efﬁciently in resisting

out-of-plane bending moments (see Figures 7.2 and 7.3). The weakened areas of the

Glass/Epoxy

25mm Wood veneers

Epoxy glue

Aluminium screen for lightening protection

(unusual)

Glass/Epoxy

Polyurethane point

Figure 7.1 Wood/Epoxy Blade Construction Utilizing Full Blade Shell (Reproduced from

Corbet (1991) by permission of the DT1 Renewable Energy R&D Programme)

BLADES 379

shell towards the trailing edge are then typically stiffened by means of sandwich

construction utilizing a PVC foam ﬁlling.

The hollow shell structure deﬁned by the aerofoil section is not very efﬁcient at re-

sisting out-of-plane shear loads, so these are catered for by the inclusion of one or more

shear webs oriented perpendicular to the blade chord. If the load-bea ring structure is

limited to a compact closed hollow section spar, consisting of two shear webs and the

skin sections between them (see Figure 7.4), then a GFRP blade lends itself to semi-

automatic lay-up on a rotating mandrel which can be withdrawn after curing.

7.1.5 Blade materials and properties

The ideal material for blade construction will combine the necessary structural

properties – namely high strength to weight ratio, fatigue life and stiffness – with

low cost and the ability to be formed into the desired aerofoil shape.

Glass/Epoxy web

Glass/Epoxy

4mm Wood veneers

Epoxy glue

Glass/Epoxy

Gel coat

Filler

Glass/Epoxy

Styrofoam

Glass/Epoxy

Gel coat

Figure 7.2 Wood/Epoxy Blade Construction Utilizing Forward Half of Blade Shell (Repro-

duced from Corbet (1991) by permission of the DT1 Renewable Energy R&D Programme)

Moulded GRP shear webs

UD Glass Fibre/polyester

Filler

CSM skins

CSM

PVC foam

CSM

Gel Coat

CSM ⫽ Continuous Strand Mat

Figure 7.3 Glass-ﬁbre Blade Construction Using Blade Skins in Forward Portion of Blade

Cross Section and Linking Shear Webs. (Reproduced from Corbet (1991), by permission of

the DT1 Renewable Energy, R&D Programme)

380

COMPONENT DESIGN

Table 7.1 lists the structural properties of the materials in general use for blade

manufacture and those of some other candidat e materials. For comparative pur-

poses, values are also presented of:

• compressive strength-to-weight ratio,

• fatigue strength as a percentage of compressive strength,

• stiffness-to-weight ratio,

• a panel stability parameter, E/(UCS)

It is evident that glass- and carbon-ﬁbre composites (GFRP and CFRP) have a

substantially higher compressive strengt h-to-weight ratio compared with the other

materials. However, this apparent advantage is not as decisive as it appears, for

two reasons. First of all, the ﬁbres of some of the plies making up the laminated

blade shell have to be align ed off-axis (typically at 458) to resist shear loads, giving

reduced strengths in the axial direction. Secondly, the rela tively low Young’s

modulus of these composites means that resistance to buckling of the thin skins

governs the design rather than simple compression yielding. The likelihood that

buckling will govern is inversely related to the panel stability parameter, E=(UCS)

given in the last column of the table, so that materials with high values, such as

wood composites will be least sensitive to buckling. As a result wood composite

blades are generally lighter than equivalent glass-ﬁbre composite blades. Design

against buckling is considered in Section 7.1.10.

It should be noted that the low strength of wood laminate compared with other

materials renders it unsuitable for blades with slender chords operating at high tip

speed, where the ﬂapwise bending moments during operation are inevitably high

in relation to blade thickness. For example, Jamieson and Brown (1992) have shown

that, in the case of a family of stall-regulated machines, the blade stress is highly

sensitive to rotational speed, increasing as the fourth power, if the skin thickness to

chord ratio is maintained constant. Although stresses can be reduced by increasing

the skin thickness, this represents a less and less efﬁcient use of the additional

Adhesive

TFT wound spar

CSM/polyester

PVC foam

Gel coat

CSM/polyester

CSM ⫽ Continuous Strand Mat

Figure 7.4 Glass-ﬁbre Blade Construction Using Compact Spar Wound with Transverse

Filament Tape (TFT) on Mandrel. (Reproduced from Corbet (1991), by permission of the DT1

Renewable Energy R&D Programme)

BLADES 381

Table 7.1 Structural Properties of Materials for Wind-turbine Blades

Material

(NB: UD denotes

unidirectional ﬁbres –

i.e., all ﬁbres running

longitudinally)

Ultimate

tensile

strength

(UTS) (MPa)

Ultimate

compressive

strength (UCS)

(MPa)

Speciﬁc

gravity

(s.g.)

Compressive

strength to

weight ratio

UCS/s.g.

Mean fatigue

strength at 10

cycles

(amplitude)

(MPa)

Mean fatique

strength as

percentage

of UCS

Young’s

Modulus,

E (GPa)

Stiffness to

weight ratio

E/s.g.

(GPa)

Panel stability

parameter

E/(UCS)

(MPa)

1

(Mean for composites,

minimum for metals)

1 Glass/polyester ply with

50% ﬁbre volume

fraction and UD lay-up

860–900

[1] [2]

720 [1] 1.85 390 140 [3] 19% 38 [2] 20.5 0.07

2 Glass/epoxy ply with

50% ﬁbre volume

fraction and UD

lay-up

Properties are generally very close to those for GRP given above

3 Glass/polyester laminate

with 50% ﬁbre volume

fraction and 80% of

ﬁbres running

longitudinally

690–720 580 1.85 310 120 21% 33.5 18 0.1

382 COMPONENT DESIGN

4 Carbon ﬁbre/epoxy ply

with 60% ﬁbre volume

fraction and UD lay-up

1830 [4] 1100 [4] 1.58 700 350 [5] 32% 142 [4] 90 0.12

5 Khaya ivorensis/epoxy

laminate

82 [6] 50 [6] 0.55 90 15 [7] 30% 10 [8] 18 4

6 Birch/epoxy laminate 117 [9] 81 [10] 0.67 121 16.5 [7] 20% 15 [10] 22.5 2.3

7 High Yield Steel (Grade

Fe 510)

510 510 7.85 65 50 [11] 10% 210 27 0.81

8 Weldable aluminium

alloy AA6082 (formerly

H30)

295 [12] 295 [12] 2.71 109 17 [13] 6% 69 [12] 25.5 0.79

Sources:

[1] Mayer (1996) Figure 2.4.

[2] Barbero (1998) Table 1.1.

[3] Mayer (1996) Fig. 14.4.

[4] Carbon ﬁbres exhibit a wide range of properties; ﬁgures given here are for one example only, taken from [2].

[5] Based on S– N curve index of m ¼ 14, taken from GL rules.

[6] Bonﬁeld and Ansell (1991) Moisture content ¼ 10%.

[7] Based on S– N curve index of m ¼ 13:4 for scarf-jointed wood laminates, taken from Hancock and Bond (1995).

[8] Bonﬁeld et al. (1992).

[9] Mayer (1996) Table 7.3.

[10] Hancock (Personal Communication). Moisture content ¼ 10%.

[11] Mean value for butt-welded joints with weld proﬁle ground smooth (Class C), taken from BS 5400, Part 10 (1980).

[12] CP 118: 1969 ‘Code of practice for structural aluminium’.

[13] Mean value estimated from mean minus two standard deviations value for ground butt welded joint with shallow thickness transition, Detail Cat 221, in

IIW ‘Fatigue design of welded joints and components’.

BLADES 383

material beyond a skin thickness to chord ratio of 3–4 percent, especially in the

outboard part of the blade, where the blade thickness to chord ratio is low.

Fatigue performance is conveniently measured by mean fatigue strength at 10

cycles, as a percentage of ultimate compressive strength. Clearly, carbon-ﬁbre and

khaya/epoxy perform best with a value of about 30 percent. The low value for

welded steel (10 percent), combined with steel’s low strength-to-weight ratio,

renders it uncompetitive for large diameter machines where gravity fatigue loading

becomes important, although it was chosen for some of the early prototype mega-

watt scale machines when the fatigue properties of composite materials were less

well understood.

The stiffness-to-weight ratio determines blade natural frequency. Apart from

CFRP, the values in the table are all in a relatively small range (18–27 GP a),

indicating that material choice will generally only have a marginal effect on

dynamic behaviour.

From the above brief survey, it is apparent that the material with the best all-

round structural properties is carbon-ﬁbre composite. However, it has not found

common use because it is an order of magnitude more costly than other materials.

Instead, the most popular material is glass/polyester, follow ed by glass/epoxy and

wood/epoxy.

Steel is the cheapest material in the raw state, and can be formed into tapering,

curved panels following the aerofoil proﬁle, except in the sharply curved region

near the leading edge. However, it is much harder to introduce a twist into such

panels, and this consideration, together with the poor fatigue properties, means that

steel is rarely used. By contrast, glass- and carbon-ﬁbre composites lend themselves

to wet lay-up in half-moulds proﬁled to give the correct aerofoil shape, planform

and twist. Laminated wood composite blades are built up in a similar way, but the

veneer thickness has to be restricted to enable the veneers to ﬂex to the required

curvature during lay-up.

In the following paragraphs, the properties of the materials in most common use

for blade manufacture are considered in more detail.

7.1.6 Properties of glass/polyester and glass/epoxy composites

As noted in Table 7.1, the properties of glass/polyester and glass/epoxy plies with

the same ﬁbre volume fraction and lay-up are generally very similar, i.e., the

inﬂuence of the matrix is slight. They will therefore be treated as the same material

in the discussion that follows, except in relation to fatigue, where some differences

have been noted. The glass used in blade construction is E-glass, which has good

structural properties in relation to its cost.

The plate elements forming the spar of a GFRP blade are normally laminates

consisting of several plies, with ﬁbres in different orientations to resist the design

loads. Within a ply (typically 0.5–1.0 mm in thickness), the ﬁbres may all be

arranged in the same direction, i.e. UD or unidirectional or they may run in two

directions at right angles in a wide variety of woven or non-woven fabrics.

Although the strength and stiffness properties of the ﬁbre s and matrix are well-

deﬁned, only some of the properties of a ply can be derived from them using simple

384 COMPONENT DESIGN