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DESIGN, STRUCTURAL TESTING, AND COST EFFECTIVENESS OF SECTIONAL WIND TURBINE BLADES

(PUBLISHABLE FINAL REPORT)

A.G. Dutton Energy Research Unit

CLRC Rutherford Appleton Laboratory UK

C. Kildegaard

LM Glasfiber A/S Denmark

T. Dobbe, R. Bensoussan

LM Glasfiber Holland Netherlands

C. Kensche, F. Hahn

Institute of Structures and Design German Aerospace Centre, DLR, Stuttgart

Germany

D.R.V. van Delft, G.D. de Winkel Delft University of Technology

Netherlands

Contract No. JOR3-CT97-0167

PUBLISHABLE FINAL REPORT 1 August 1997 - 30 November 2000

Research funded in part by THE EUROPEAN COMMISSION

in the framework of the Non-Nuclear Energy Programme

JOULE III

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DESIGN, STRUCTURAL TESTING, AND COST EFFECTIVENESS OF SECTIONAL WIND TURBINE BLADES

Abstract

As wind turbines are designed and constructed to ever increasing sizes, so the component sizes must increase in proportion. In the case of the blades, for MW scale machines, handling and transportation start to become major problems since, for example, the road clearance beneath bridges is no longer sufficient and the length of trailers becomes unwieldy. This could be a particular problem in developing unconventional sites in complex terrain.

One solution would be to design blades in sections, which could then be assembled on site. As well as easing transportation problems, such a design has the potential to improve ease of construction and maintenance, and may allow the tailoring of wind turbines for specific sites. However, it is of prime importance that structural integrity and blade performance are not diminished by the connections and that any extra costs incurred by the design are less than the subsequent savings.

This project therefore carried out a comprehensive review of potential blade connection concepts, tested the most promising principles at coupon scale, and then designed and built a prototype sectional blade. The prototype sectional blade was tested for both static and fatigue strength and the results compared with the equivalent conventional blade design. The implications for manufacture and transportation were assessed and the overall cost effectiveness of sectional blade design was estimated.

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Project partnership

The roles of the partners and their areas of expertise are shown in Table 1.

Organization Abbrev. name

Function

Energy Research Unit, CLRC Rutherford Appleton Laboratory

ERU/RAL Project co-ordination Project management of final sectional blade concept evaluation Condition monitoring and evaluation

LM Glasfiber A/S LMG Project management of prototype sectional blade design Coupon, model, and prototype blade manufacture Assessment of cost effectiveness of sectional blade concepts

LM Glasfiber Holland BV LMGH Prototype sectional blade design and construction Deutsche Forschungs-anstalt fur Luft und Raumfahrt

DLR Project management of blade connection concept development Assessment of blade connection mechanical integrity

Delft University of Technology DUT Project management of prototype sectional blade test Prototype sectional blade test

Table 1 : Responsibilities of the partners

Objectives

The objectives of the project were to:

(i) review possible connection assembly concepts,

(ii) test the most promising connector designs,

(iii) design and build a prototype sectional blade,

(iv) test the prototype blade for both static and fatigue strength and compare the results with the equivalent conventional blade design.

List of publications

[1] Geiger, T., Olesen, M., Connection concepts for sectional wind turbine blades, Deliverable report, EC project no. JOR3-CT97-0167,16 March 1998

[2] Dutton, A.G., Geiger, T., Hahn, F., Olesen, M., Kensche, C., Korsgaard, J., van Delft, D.R.V., de Winkel, G.D., Design concepts for sectional wind turbine blades, European Wind Energy Conference, 1-5 March 1999, Nice, France

[3] Hahn, F., Kildegaard, C., Development of a bending spar connection concept for a "sectional rotor blade", DEWEK 2000

[4] Hahn, F., Kensche, C., Paynter, R., Dutton, A.G., Kildegaard, C., Korsgaard, J., Design, fatigue test, and NDE of a sectional wind turbine rotor blade, 15th Technical Conf. of the American Society for Composites, September 2000

[5] Dutton, A.G., Kildegaard, C., van Delft, D.R.V., de Winkel, G.D., Kensche, C., Dobbe, T., Bensoussan, R., The potential of sectional wind turbine blades, to be presented at European Wind Energy Conference, 2-6 July 2001, Copenhagen, Denmark

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SCIENTIFIC AND TECHNICAL DESCRIPTION

1. INTRODUCTION

The sectional blade concept was first used in the context of wind turbine technology on the Debra turbine in 1982 [1]. The blades were divided in two equal lengths which fitted inside a standard 6 metre container in order to facilitate transport and make cost savings. The T-bolt load connection (also known as the IKEA connection following its adoption by the Swedish furniture manufacturer) was used and verified. The blades have been in operation for 12 years without any major structural problems, although some additional maintenance has been required to verify the pre-stress of the bolts and the quality of protection against water ingress. (Note, however, that loss of pre-stress is not generally a problem in fatigue tests of this type of joint or in root connections.)

The Debra blade connection technology was complicated, took a relatively long time to manufacture, and required special devices. Understanding of joints in glass fibre composite materials has progressed considerably since this early design and connections based on other principles may now be more advantageous. Moreover, manufacturing techniques have changed markedly, in particular the recent shift to resin infusion moulding. The incorporation of this kind of technology into the mid-span of a modern wind turbine blade represents a major innovation, involving detailed assessment of the transfer of loads within the blade and how these may be carried in the most cost effective manner.

The sectional blade concept could be exploited for a range of different purposes, each of which would dictate the desired location for dividing the blade. The primary objective envisaged here is to facilitate transportation of very large blades, and hence the break in the blade is likely to be placed as far from the root as a standard container will allow. However, building blades with equal mass sections might be required in future in order to limit the lifting loads to those for a crane of given capacity (either for initial construction or maintenance purposes). Alternatively, it may be desired to exploit a single root moulding for a range of blade tip sections of differing length, thereby increasing product range and permitting tailoring of individual wind turbines to local wind conditions.

A range of varied concepts for constructing a sectional blade have been evaluated against a set of criteria balancing the increased benefits of handling with the increase in complexity and cost. Finite element modelling and coupon and scale model testing have been performed in the cases of the three most promising concepts.

Two designs have been incorporated in actual blade constructions - one at "scale model" size and the other in a prototype "full scale" blade. The two sectional blades were tested both statically and in fatigue and an assessment made of the mechanical integrity of the designs and their overall cost effectiveness.

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2. INITIAL CONCEPT REVIEW

Initial design concepts for a blade division were proposed based on experience from the wind turbine blade industry and the light aircraft industry. The initial concept review developed a total of 18 different designs, which fell into three categories:

(i) designs based on current rotor blade technology (9 concepts all incorporating some kind of bolted connection),

(ii) designs based on light aircraft wing design (5 concepts all involving some kind of interconnecting structure),

(iii) laminated and bonded connections.

All designs were evaluated against the manufacturer's specification, which comprised a set of invariable demands and an array of weighted variable assessment parameters [2,3].

(a) (b)

Figure 2.1 : Embedded bushing with stud bolt blade connection principle and T-bolt blade connection principle.

Title: sectional5.epsCreator: fig2dev Version 3.2 Patchlevel CreationDate: Wed Jan 28 17:00:46 19

Figure 2.2 : Blade sections with connecting tubes

Three designs chosen for further evaluation, namely the embedded bushing with stud-bolt and the T-bolt connection (from current blade technology - see Figure 2.1), and blade sections with connecting tubes (from light aircraft design - see Figure 2.2).

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The laminated and bonded connections were all rejected on the grounds of much increased on-site assembly work, with considerable effort required to maintain surface finish tolerance (for aerodynamic reasons), and the lack of any experience with such a connection within a highly loaded shell.

From the manufacturer's specification for a sectional blade, it was required that each design should fulfil a set of basic invariable demands::

• transfer the loads into the blade shell,

• be suitable for incorporation in existing blade designs,

• be connectable from outside (by means of a hatch, if necessary),

• conform to certification requirements,

• withstand specified extreme and fatigue loads,

• in principle be suitable for incorporating anywhere in the blade, but in practice between 0.33 R and 0.67 R.

Parameter Weighting T-bolt Embedded bushing with stud bolt

Sections with connecting tubes

Weight 1 6 7 5

Disconnect-ability 1 8 8 10

Aerodynamics 6 9 9 10

Costs 10 5

(more GRP)

5

(more steel parts)

3

(more parts)

On-site assembly 5 6 6 9

Quality control 5 8 5

(can't see parts)

7

Maintenance 5 8

(initial creep)

9 9

Tolerance requirements

5 7 5

(difficult with resin infusion)

6

Mould 2 7 8 4

Production complexity

8 7 5 4

(more pieces)

Reliability 10 7 6 8

Total performance 403 360 380

Table 2 : Sectional blade concept evaluation - revised measures of performance against the variable specifications.

In addition, the designs were assessed against the variable specifications, weighted 1-10, shown in Table 2. The performance of each of the selected designs against the variable specifications is shown in the table. These "scores" were adjusted at the end of the project to reflect the experience gained in the practicalities of manufacturing sectional wind turbine blades. In particular, it was recognised that the importance of minimising production complexity had not been sufficiently emphasised in the original study. The T-bolt connection, which scored only moderately well initially, was given more credit for the greater ease of production compared with the embedded bushing design. Also, after the

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experience of making a blade with this configuration, the manufacturer was more confident about quality control issues and meeting tolerance requirements.

The embedded bushing with stud bolt design scored highly on cost grounds, but was later marked down more severely for the tight tolerance requirements and the difficulty of achieving suitable quality control, particularly when manufactured using resin infusion processes rather than hand lay-up as was originally assumed.

Two designs derived from sail-planes were considered to have merit. The connecting tubes design was given priority on the grounds that the technology was already familiar in a less critical application (i.e. tip-braking mechanisms) and this approach was therefore considered to be more reliable.

Thus, the T-bolt, embedded bushing with stud bolt, and the blade sections with connecting tubes designs were selected for further development using finite element analysis and coupon and scale model testing.

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3. SECTIONAL BLADE CONNECTION CONCEPT DEVELOPMENT

The partners established basic design principles for the three candidate connections. Finite element analyses were carried out by DLR and RAL to establish basic dimensions and balance loads appropriately between GRP and steel parts, and then coupons and a model blade were manufactured by LMG. Testing of bolted connections was carried out by DUT and of the scale model blade by DLR. Thermoelastic stress measurements were made by RAL during several fatigue tests.

3.1 Embedded bushing with stud bolt connection

Initial design work was carried out to estimate the appropriate configuration of embedded bushing with stud bolt connections in a sectional blade (Figure 3.1). A simple test coupon was designed and two static and four fatigue tests carried out to determine the pull-out strength of the bushing connection under static and fatigue loading.

The basic geometry of the embedded bushing with stud-bolt test coupon is shown in Figure 3.2. Only one of the symmetric halves of the coupon is shown. The coupon consists of a GRP block with a steel bushing embedded at one end and an assembly hole to provide access to the bolt head. The two halves of the coupon are then secured by a stud-bolt which is pre-tensioned to ensure the desired transfer of loads through the joint. The design aim was to reduce the overall width of the specimen in order to maximise the number of connections that could be accommodated within any given blade section and hence to reduce the load carried by each connection.

Figure 3.1 : Initial estimated location plan for embedded bushing with stud-bolt connections in a sectional wind turbine blade

Figure 3.2 : Basic geometry of embedded bushing with stud-bolt test coupon

Finite element analysis of the test coupon was carried out to:

(i) verify the load balance between the bolt, the bushing, and the GRP,

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(ii) compare the loads in the connection between the bushing and the GRP with the loads in the GRP around the assembly opening and hence to optimise the width of the connection so steel and GRP parts were equally utilised.

The results of the static tests are summarised in Table 3. Two different failure modes were observed, the first a failure of the bonding between the bushing and the GRP, the second a shear failure in the threaded rod.

Test No. Failure load [kN] Failure mode 1 450 Shear failure of the bonding between the steel tube and the

GRP 2 493 Shear failure of the thread of the rod

Table 3 : Embedded bushing with stud-bolt coupon static test summary

Table 4 shows results from the four fatigue tests. Three different failure modes were obtained, two of which are shown in Figure 3.3. During test 3, at 138,500 cycles, the test specimen was overloaded in tension due to a power supply failure, which caused loss of most of the strain gauges and misalignment of the two halves of the test specimen.

The bolt load factor gives an estimate of the external load range that is carried by the bolt. A low load factor is desirable in order to reduce the load range seen by the steel bolt, which is often the critical component. Based on the measured bolt forces, the average load factor was 0.134, compared to the value of 0.249 estimated from the FE model. The reasons for this discrepancy are not fully understood, although the trend might be expected due to the stiffer behaviour of the real laminate compared to the model prediction.

Test No. Applied load [kN]

Pre-load [kN]

No. cycles to failure

Failure mode

3 100 265 1.3 x 107 Shear failure of the bonding between the bushing and the GRP

4 200 / 1 Hz 265 1.15 x 105 Shear failure of the bonding between the bushing and the GRP

5 100 / .75 Hz 265 5.6 x 104 Failure of the GRP beside the slot in the lower part of the specimen

6 130 265 2.0 x 106 Failure of the rod by fatigue

Table 4 : Embedded bushing with stud-bolt fatigue test summary

(a) (b)

Figure 3.3 : Embedded bushing with stud-bolt fatigue test failure modes - shear failure of bonding between bushing and GRP in coupon 3 (a) and GRP failure in coupon 5 (b)

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Thermoelastic stress measurements were carried out on the embedded bushing with stud-bolt test coupons using the Deltatherm system, manufactured by Stress Photonics Inc. in the USA. This system can take images of the change in total stress in a body during fatigue loading in as little as a few seconds (as opposed to several hours for the previous generation SPATE equipment). During the research programme, it was found that the apparent measured stress distribution was affected by relative changes in surface temperature to a much greater extent than expected. However, the measurements did confirm the stress distribution predicted by the FE model and showed early indications of developing damage. Work is continuing to develop a suitable method to correct the thermoelastic stress map for both spatial and temporal temperature fluctuations.

3.2 T-bolt connection

The T-bolt connection can be considered as a straightforward alternative to the stud bolt connection and so the design configuration considered in the previous section remains valid (see Figure 3.1).

A finite element model of the T-bolt test coupon was made in order to:

(i) assist in the interpretation of the thermoelastic stress measurements,

(ii) understand the transfer of load through the connection.

The T-bolt connection test coupon before and after assembly is shown in Figure 3.4. The two parts are connected with a threaded steel rod of diameter 30 mm. The T-bolt connection test programme consisted of two static tests and three fatigue tests. The two static tests were performed under increasing tension until failure.

(a) (b)

Figure 3.4 : T-bolt connection test coupon before assembly (a) and after assembly (b)

The results of the static tests are summarised in Table 5. The same failure mode was obtained in both cases, namely tensile failure of the threaded rod.

Test No. Failure load [kN] Failure mode 1 493 Tensile failure of the threaded rod 2 491 Tensile failure of the threaded rod

Table 5 : T-bolt connection coupon static test summary

The results of the fatigue tests are summarised in Table 6. Specimen 3 failed in the GRP part, at the edge of the circular slot in the top half of the coupon. Specimen 4a failed in the bolt by the development of a fatigue crack in the thread of the bolt, just beneath the nut. The bolt was replaced

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and the test was continued, as test coupon 4b. The test ended in a similar manner to test 4a after a very similar lifetime.

Test No. Applied load [kN]

Pre-load [kN]

No. cycles to failure

Failure mode

3 200 250 1.38 x 106 Failure in GRP, beside the circular slot 4a 250 300 9.88 x 104 Fatigue failure of bolt 4b 250 300 8.38 x 104 Fatigue failure of bolt

Table 6 : T-bolt connection coupon fatigue test summary

A comparison between the predicted thermoelastic (total) stress distribution and the measured distribution is shown in Figure 3.5. Zero stress is represented by black, with red-yellow for stresses in phase with the load and blue-green for stresses 180o out of phase with the load. The view of the test coupon was limited by the surfaces of the load rig, which extended below the top of the circular pin. Strain gauges and their lead wires can be seen on the coupon centreline and on the wings of the elliptical cut-out. There is good agreement between the FE model and the thermoelastic stress measurement with respect to the stress concentration on the 90o line from the circular roller and beside the elliptical cut-out. The magnitude of the zero loaded or slightly compression loaded areas above and below the bolts on the coupon centre-line are well reproduced, although these zones spread out laterally in the measurement. This may be an indication that the FE mesh needs further refinement to properly represent the contact stresses.

(a) (b)

Figure 3.5 : T-bolt connection test coupon thermoelastic stress distribution from finite element prediction (a) and calibration load (b)

3.3 Blade sections with connecting tubes

The principle of the connecting tubes connection means that blades may be manufactured in two sections, which are later connected by a tube element acting as a load bearing spar [4]. The structural part of the connecting tube is built as two tapered glass polyester tubes with metal bearing points at

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each end. This is built into the outer section of the blade during manufacture and assembled with the inner section by slotting into mating elements and fixing with a clamping bolt. The gap between the two shells (about 10mm) will not take any stress and would have to be filled with a flexible sealant for weather proofing.

Since it was impossible to design a simple test coupon for 3- or 4-point bending to evaluate the design principles, the connection was tested using a reduced scale model (Figure 3.6). This had the bonus that the performance could be more directly compared with that of the prototype sectional blade, tested in the final phase of the project.

Figure 3.6 : Blade sections with connecting tubes - scale model blade being prepared for testing in the laboratory at DLR

(a) (b)

Figure 3.7 : FE model of blade sections with connecting tubes showing nodes (a) and the elements in a cutaway section of the blade (b) [blue = shell on pressure side; violet = main shear webs; red = new connecting parts]

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An FE model (Figure 3.7) of the blade sections with connecting tubes was developed to reveal the possible weak points of the tube connection when inserted into a LM 13,4 m blade and to determine the thickness of the lay-up for the new connecting structure. The shaft is bedded into two steel rings (located between 3,0 – 3,2 m and 4,3 – 4,5 m) by contact elements. A 50 cm long steel tube is bonded to the inside of the GRP tube and the inner nodes coupled to the nodes of the steel ring to represent the locking device against the fly forces. At the connection plane (4,5 m) one node of each of the adjoining bulkheads is coupled in the vertical (flapwise) direction, preventing the two parts from rotating in relation to each other.

Three load cases were considered [5]: flapwise, edgewise, and combined flapwise and edgewise loading. Flapwise bending was predicted to be the most critical load condition. The central vertical parts of the new "half-moon" webs around the shaft and tube sections, where the two halves of the webs are bonded together, are the weakest point of the design; if the connection is not sufficiently stiff, the shear load exceeds permissible levels. After several iterations, the optimum component thickness to carry the loads with a reasonable safety factor was determined (Figure 3.8a). The resulting shear strain distribution along the critical joining line is presented in Figure 3.8b, where strain values are presented as the percentage of maximum allowable strain. Thus, the maximum shear strain in the new webs is predicted to be 60% of the allowable value. The shaft within the blade connection is also highly loaded in shear at a wingspan of 3,5 m, where the steel retaining ring inside the GRP tube ends (Figure 3.9). The strain in the rest of the blade also remains well within the required limits.

Figure 3.8 : Cross-section of final "half-moon" web design with associated shear strain distribution under flapwise loading

(a) (b)

Figure 3.9 : Blade sections with connecting tubes finite element results showing shear stress in shaft (a) and longitudinal strain in tube (b) for flapwise loading case

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15

8

8

8

8

Foam core

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The blade manufactured for the scale model test is the inner 9 metres of an otherwise standard LMG 13,4 m blade. The join is 4,5 m from the root and the connection tube extends 1,5 m either side of this. The loads for the static tests were derived from the design loads, inclusive of all recommended safety factors, as supplied by LMG for both flapwise and edgewise directions.The entire area of the new joint, from 3 m to 6 m, was loaded under more severe bending conditions than required by the design.

The test procedure was:

• 3 static load tests, one in flapwise (towards suction side) and two in edgewise (towards leading and trailing edges) directions,

• 5 million cycle fatigue test in flapwise direction, as above,

• Repeat initial static loading.

For the static test in flapwise direction only small load-displacement hysteresis effects were observed and there was no departure from linearity such as might be indicative of likely early failure. The maximum compressive strain (0,32%) was found in the blade skin at 6,2 m from the root, on the suction side of the profile at the 35 % chord line; the maximum tensile strain (0,29%) is found on the opposite side of the blade. The strains in the tube are slightly lower: the maximum compressive strain is 0,29% (TO_0°); the maximum tensile strain is 0,22% (TU_0°).

All strains were observed to return to the origin at the end of the test, suggesting that the structural integrity of the blade was not effected by the test, at least in the areas where the strains were monitored. However, the strain gauges at 3m from the root showed some abnormal behaviour. The strain gauge on the suction side (loaded compressive) had several kinks during upward loading, where the strain dropped rapidly before continuing again in a linear fashion. The tensile loaded strain gauge at this section registered a loss of strain during the hold at maximum load. This behaviour can be explained by considering the performance of the locking device for fly loads. The screw delivered by LMG was of quite poor quality and could only be tightened up to 150 Nm torque. From later dynamic testing, it was observed that all the transverse load from the end of the tube was carried into the rest of the blade through this locking device, resulting in the steel parts starting to move against each other above a certain load. This had the effect of causing discrete drops in the surface strain, as seen here.

Figure 3.10 : Blade sections with connecting tubes - flapwise loading and deflection

For the static tests in the two edgewise directions the blade was mounted with a pitch angle of ±90° at the tip of the blade and the load applied with a crane in a similar way to the flapwise test. The two edgewise tests gave essentially similar results.

Figure 3.11 shows that the blade divide became unloaded at a load of 2.65 kN, which is the self-weight of the blade. Trend lines were applied to the graph above and below this value to give the "measured" play, p*, at 8,5 m from the root of 1,6 mm. Assuming that the tube is fixed at 3.0 m from the root (by the locking device for the fly forces) and all the play therefore comes from the ring at 4.5 m, the play, p, at the mid-section was determined to be 0,44 mm. The magnitude of this play was found to correspond to a lack of fit between the steel ring and the outer diameter of the GRP tube at the blade connection, due to the steel ring being slightly ovalized.

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0

2

4

6

8

10

12

14

16

18

20

22

24

26

28

30

32

34

36

38

40

42

44

0 500 1000 1500 2000 2500 3000 3500 4000 4500 5000 5500 6000 6500

Load [N]

Dis

plac

emen

t [m

m]

1,6 mm

Figure 3.11 : Blade sections with connecting tubes - measurement of play in blade connection

To produce results comparable with the undivided LM 13,4 m blade, the tube-connected sectional blade was tested under constant amplitude load for 5 million cycles. Since the principal aim of the test was to evaluate the performance of the tube connection under bending and the rest of the blade was only required to achieve a realistic load introduction to the tube, it was decided to perform the test only in the flapwise direction. Bending moments in the tube itself are the same whether the blade experiences either edgewise or flapwise loading, but the flapwise loads are more severe, both at the connection and for the other major structural parts. The blade was mounted on the test rig with the suction surface towards the floor and was excited with a rotating mass. The blade survived 5,000,000 load cycles without any sign of failure (except for a broken screw in the locking device, which was attributed to poor material specification). The resonance frequency remained constant throughout the test, which, since resonance frequency is very sensitive to blade stiffness, can be taken as evidence that no stiffness degradation occurred.

On dividing the blade after the test, the steel rings showed fretting corrosion at the areas in contact with the GRP tube (Figure 3.12). The rust on the tube was used to indicate where the parts were not properly in contact and hence how the loads were transmitted through the connection. The evidence of fretting corrosion indicates that, at the end of the tube, the load was transferred through the locking device for the fly forces instead of via the fitting of the tube in the inner steel ring. It appears that there was only a small area on the under side of the tube that was in permanent contact with the steel ring. It is suggested that if this contact had been better maintained (through improved manufacturing tolerances) then the screw would have had to take less load and so would not have failed. However, this minor failure was adequately addressed by using higher quality steel for the replacement screw.

(a) (b)

Figure 3.12 : Blade sections with connecting tubes - fretting corrosion on GRP tube from steel ring at 4,5 m (a) and fretting corrosion on steel connector at 3,0 m top surface (b)

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After completion of the fatigue test, the original static tests were repeated to compare the stiffness characteristics and strain distributions. No significant change was observed and the stiffness of the blade was found to be unchanged.

Again thermoelastic stress measurements were made on the blade. Figure 3.13 shows a detail of the measured thermoelastic stress distribution at the extremities of the connection. More detail is evident than in the initial FE model. In particular, at the inboard joint there is a reduction in total stress immediately inboard of the joint with higher levels to either side, most notably on the tip side. The points of highest thermoelastic stress occur at the conjunctions between the reinforced shear webs and the blade shell. Additional strain gauges, placed in one of these locations, exhibited a double-period strain response, implying that some local buckling effects might be occurring. In general, it is concluded that there are no unreasonably high stress concentrations on the outside of the shell, although more detailed FE calculations around the conjunction between the blade skin and the reinforced shear web may be merited. Data from the tipmost connection shows a similar distribution but "mirrored" and smaller in lateral dimensions because the distance between the shear webs is narrower at this section. The stress concentration effects appear reduced in this case.

Figure 3.13 : Blade sections with connecting tubes - thermoelastic stress distribution at inboard and outboard connection

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4. EXPERIENCE WITH A PROTOTYPE SECTIONAL WIND TURBINE BLADE

LMGH designed and manufactured a prototype sectional rotor blade incorporating the T-bolt connection. This blade was then tested statically and in fatigue by DUT.

4.1 Blade design

The basis of the prototype blade, designated SB23.3, was the existing LM23.3 blade, without its tip-brake mechanism. The blade separation was incorporated at a distance of 7.3 m from the root flange, rather than at the midpoint of the blade (11.5 m), in order that the connection should be as relevant as possible to larger blades. The chord length at this point (about 1.7 m) is considered representative of the chord length of larger blades at mid span.

The basic design was modified to incorporate the connections by:

• thickening of the basic UD laminate at the joint by a factor of about 3, incorporating UD and ±45° fabrics, to achieve more isotropic properties about the holes,

• drilling of circular and elliptical bores for mounting the cross-nuts and pre-tensioning the bolts,

• drilling of bolt holes parallel to the blade skin for the bolts themselves,

• separation of the shear webs and reinforcement of sections using bulkheads.

Figure 4.1 shows the main details of the connection designed for the SB23.3 blade. The longitudinal thickness distribution shown in Figure 4.2, with a "classical" slope of 1:40 for the increase and decrease of the skin thickness, and a central region of constant thickness, was introduced to support the connection. To increase the thickness and get satisfactory mechanical properties in all directions (anisotropy increases the stress concentrations around a hole), roughly equal quantities of ±45° and 90° fabrics were added to the original UD material layers. The construction was made as symmetric as possible to get a smooth distribution of the forces through the connection. Thus, the original fabrics were integrated layer by layer with the additional fabrics needed for the connection, resulting in a modified area with total length of 4.3 m, between 5.1 m and 9.4 m from the root flange. The rest of the blade was built as a standard LM23.3 blade.

15

5

SEPA

RATI

ONAT

L730

0

200

Ø70

200

Ø32

M30

70

150TIP PART ROOT PART

Figure 4.1 : Prototype sectional blade - major T-bolt dimensions

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Although it would have been simpler to drill the longitudinal holes (32 mm diameter) perpendicular to the cross section at L7300, it was decided that they should be made to follow the angle of the blade (about 1.8o maximum) due to its "conical" form (Figure 4.3). This arrangement avoids asymmetric loading conditions at each joint, while increasing manufacturing complexity since the angles of each hole had to be programmed separately.

8001800

20

60

1800

L7300

Figure 4.2 : Prototype sectional blade - longitudinal thickness distribution

60-6

5

Figure 4.3 : Prototype sectional blade - final T-bolt design incorporated into blade section

The distribution of holes around the cross section was determined by the requirement to include the largest number possible in the blade. Consideration had to be given to space restrictions around the leading edge, where a glue flange is present between the two shells, and in the trailing edge, where the laminates join together. A harmonic positioning of the T-bolts has been reached by using 18 pieces (9 for each side) with a variable distance between the bores.

The flapwise loads had been determined to be critical, so the separation of 155 mm was imposed in the mid-blade span to reproduce exactly the geometry of the test coupons. Towards the edges, the distance between consecutive bores increases progressively to 170 mm, 185 mm, and finally 200 mm, resulting in the distribution of holes shown in Figure 4.4.

At the location where the blade was divided, the construction around the shear webs was modified to transmit the shear forces correctly. Some transverse GRP bulkheads were designed and placed at the ends of the shear webs. The connection between these components was made with an additional laminate. The bulkheads were designed using the extreme shear load present in the blade and were glued in the shells, immediately adjoining the intended separation point, with a separation of 20 mm, which was estimated to be the minimum space required for sawing the blade without damage to the bulkheads.

It was calculated that three steel cross nuts would be sufficient to carry the weight of the tip part in the situation where no pre-tension is available in the connection, as, for instance, during the assembly of the two parts.

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Hole pattern for SB23.3

Figure 4.4 : Prototype sectional blade - distribution of bolted connections around blade section

4.2 Blade manufacture

The SB23.3 sectional blade was produced using the resin infusion process, as for the standard LM23.3 blade, as follows:

(i) Dry fabrics were placed in the mould according to the book of instructions containing drawings of the different layers (material name, dimensions, and position in the mould) which have to be stacked on top of each other to produce the two shells. The laying up of the dry glass fabrics in the two shell moulds was accomplished in about one day, which is longer than is usual for the standard LM23.3 blade.

(ii) The fabrics were infused with resin, following a process as similar as possible to the strategy used for the normal blade, with some additional injection and suction points added in the area of the thickened laminate and connected to the existing infusion tubes. The critical points for the infusion process were to inject the laminate as completely and thoroughly as possible, particularly in the extremities of the thickened part. The injection of the resin was completed in about two hours, with only very minor problems and the gel time of the delayed resin was long enough to allow the infusion of the whole blade in a single process.

(iii) The two shear webs were produced as single pieces using standard moulds by putting some foam pieces with different thickness against the mould flanges to produce the required thickness reduction. Each web was then cut in two parts after positioning of the bulkheads, in order to ensure the correct fit.

(iv) The inner components were then fitted in the infused shells. Two GRP bulkheads had been produced for the separation point, using a special mould. After fixing the bulkheads, the shear webs had to be cut in two pieces and glued in the same shell. To connect the transverse bulkheads with the shear webs, some ±45° laminate was inserted by hand. The gluing flange, normally laminated on the leading edge of the LM23.3 blade, would no longer fit due to the increased thickness of the blade and was therefore laminated only as far as the modified area. Thus, where the laminate was thicker than for the normal blade, no internal flange was built in the mould and this area had to be laminated afterwards on the outside surface.

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(a) (b)

(c) (d)

(e) (f)

Figure 4.5 : Prototype sectional blade production

(v) The downwind and upwind shells were glued together. Special attention had to be paid to the bulkheads during the closing of the mould to avoid intrusion of glue between them.

(vi) After removing the blade from the mould, some areas on the outer surfaces (both sides) were found to be dry and had to be repaired before the blade could be divided.

21

(a) (b)

(c) (d)

(e)

Figure 4.6 : Prototype sectional blade production

22

Figure 4.7 : Prototype sectional blade - assembly of the blade at DUT

(vii) The blade was cut using a special tool, designed and produced for the purpose with the objective of reducing the stresses in the laminate due to the blade's own weight. Two wooden clamps were connected together with a stiff steel frame, so that the two parts would still be supported when completely cut. A special disk saw with diamond edges was purchased to cut the blade. This had the capacity to cut to a maximum depth of 80 mm, which was more than sufficient for the section of the SB23.3 blade. Cooling was not required.

(viii) After cutting, the mating surfaces had some irregularities, which had to be removed in order to avoid having gaps between the two sections. This was achieved by attaching a thin laminate, comprising some woven fabrics and chopped strand mat, by hand lay-up to each surface and curing while under compression with a flat steel plate.

(ix) The holes (longitudinal and transverse) were drilled using a numerical-controlled machine with diamond drill and water cooling. The drilling machine was positioned manually at each hole location, using three location studs in the bulkheads as reference. The co-ordinates of the holes (centreline) were then given, together with the different angles. The only difference

23

between the two parts was the presence of slots in the root part. After each drilling action, the tolerances were checked with standard or special measurement tools.

(x) The blade parts were transported separately and final assembly of the blade was made on the test rig at DUT. The root section was fixed on the test rig first and then the tip section was lifted up with the crane to bring it in front of the mounted part. The bolts were first tensioned by hand and then with a wrench. By measuring the strain in some of the bolts, it was possible to be determine the average torque needed to obtain the required pre-tension. An identical torque rating was finally used for all the 18 bolts of the connection. After assembly, a gap of about 2 mm was observed towards the trailing edge of the blade, extending for a length of about 300 mm. This was due to some irregularities in the flatness of the cut surfaces. Even under full pre-tension, the gap could not be closed.

4.3 Static blade test

The static loads for the testing of the prototype sectional blade were selected to correspond as closely as possible to previous tests performed on the standard LM23.3 blade. The blade was mounted with the leading edge upwards and loaded via a wooden block at 14.6 m. Strain gauges were placed on the blade surface, at five radial stations and at the connection between the two halves (L = 7.3 m). Six bolts were instrumented with a single strain gauge to measure the bolt force and two bolts with three strain gauges to be able to determine bending in the bolt (Figure 4.8)

The load cases for the prototype blade were (load at L = 14.6 m):

• Flapwise 137 kN (compression DWS)

• Edgewise 49 kN

• Edgewise -75 kN

The maximum load was applied to the blade in 30-40 seconds, the load was kept constant for 10 seconds, and then the blade was unloaded in about 30-40 seconds. Figure 4.9 shows the prototype sectional blade with the edgewise static loads of 43 kN and -137 kN. No damage was observed during any of the static tests.

The distribution of bolt loads in the flapwise load case (Figure 4.10), and, to a much lesser extent, the edgewise load case, shows some marked hysteresis effects and non-linearities. However, since in all cases the curves return to the origin, there are no permanent effects and this behaviour is probably attributable to geometric effects.

Since the bolt loads and, in particular, the load factors with respect to the externally applied load are critical to the integrity of the connection, the maximum bolt forces during each of the three static tests are presented in Table 7. The bolt forces are converted to stress ranges based on a cross section of 561 mm2 and then normalised to give the stress range for each load increment of 10 kN at L = 14.6 m. The location of the bolts is given in Figure 4.8.

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Figure 4.8 : Prototype sectional blade test - strain gauge locations

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Figure 4.9 : Prototype sectional blade static edgewise load tests (43 kN and -137 kN)

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Measured force range Estimated stress range Normalised stress range Load case

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B01 62.2 10.8 -19.0 111.0 19.0 -34.0 8.1 4.5 -4.5 B05 94.4 -3.2 5.1 168.0 -6.0 9.0 12.3 -1.3 1.2 B09 17.9 -21.7 52.8 32.0 -39.0 94.0 2.3 -9.0 12.5 B10 -21.7 -17.3 45.0 -39.0 -31.0 80.0 -2.8 -7.2 10.7 B14 -84.7 6.2 -9.9 -151.0 11.0 -18.0 -11.0 2.6 -2.4 B18 -21.8 14.8 -25.0 -39.0 26.0 -45.0 -2.8 6.1 -5.9 BB01 121.8 -1.0 0.3 217.0 -2.0 1.0 15.8 -0.4 0.1 BB02 126.1 -1.2 0.9 225.0 -2.0 2.0 16.4 -0.5 0.2 BB03 109.0 -0.1 -0.4 194.0 0.0 -1.0 14.2 0.0 -0.1 BB04 -96.3 9.5 -15.1 -172.0 17.0 -27.0 -12.5 3.9 -3.6 BB05 -74.2 8.8 -13.8 -132.0 16.0 -25.0 -9.7 3.6 -3.3 BB06 -88.1 9.2 -14.9 -157.0 16.0 -27.0 -11.5 3.8 -3.5

Table 7 : Prototype sectional blade static test bolt forces

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The prototype sectional blade was tested with the same static loads as the standard blade without the T-bolt connection, in both flapwise and edgewise directions, without any damage occurring in the region of the connection or elsewhere on the blade. It can therefore be considered to have passed the static load test.

4.4 Fatigue blade test

The fatigue loads for the testing of the prototype sectional blade were selected to correspond with those used for certification testing of the standard LM23.3 blade.

The following load cases were tested:

Edgewise loading : 5.0 x 106 cycles, min. = -61.5 kN, max. = +15 kN at L = 14.6 m

Flapwise loading : 2.5 x 106 cycles, min. = -13.2 kN, max. = +56.8 kN at L = 14.6 m

The test frequency for the edgewise test was 1.3 Hz, for the flapwise test 0.9 Hz.

For the edgewise test, the blade was mounted in the test rig with the leading edge pointing upwards. The negative minimum load therefore represents compression in the leading edge. The load range was 76.5 kN. At the beginning of the test there were significant vibrations and out-of-plane movements in the blade which limited the test frequency which could be used. To resolve this situation, the blade was cut at L = 19 m and the load introduction was rotated by approximately 6o. This action was carried out after the first 1.3x106 cycles and resulted in a lower actuator force to achieve the required displacement and less noise in the strain gauge readings.

The blade survived the required number of cycles (5.0 x 106), during which time no damage or stiffness degradation (see Figure 4.11a) was observed. The load ranges in the bolts of the connection did not change during the test (see Figure 4.11b).

For the flapwise test the blade was mounted in the test rig with the down wind side facing downwards. The minimum load was -13.2 kN (DWS under tension) and the maximum load was +56.8 kN (DWS under compression), resulting in a load range of 70 kN. To prevent out-of-plane movements the load introduction was slightly rotated.

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Figure 4.11 : Prototype sectional blade edgewise fatigue test - strain gauge ranges at L = 7.025 m (a) and bolt force ranges (b)

The blade survived the required number of cycles (2.5 x 106), during which time no damage and almost no stiffness degradation was observed. The load ranges in the bolts of the connection did not change during the test. A typical strain range distribution for the section at L = 7.025 m is shown in Figure 4.12a and the bolt force ranges are shown in Figure 4.12b.

The prototype sectional blade was tested with the same fatigue loads as the standard undivided blade, in both flapwise and edgewise directions without any damage occurring in or around the connections, nor in the rest of the blade. Also no stiffness degradation of the blade was measured, and the load ranges of the forces in the bolts in the connection did not change during the test.

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Figure 4.12 : Prototype sectional blade flapwise fatigue test - strain gauge ranges at L = 7.025 m (a) and bolt force ranges (b)

The thermoelastic stress measurements were strongly influenced by the large displacements of the structure. However, they did appear to confirm the higher stress concentrations around the circular holes compared to the slots found with the test coupon. This finding is also in line with the one failure in the test coupons which occurred in the GRP part due to a crack running outwards from the circular hole. Thus, the use of two slots rather than a slot and a hole should be investigated for future applications.

4.5 Evaluation of structural integrity

Since the prototype sectional blade SB23.3 survived both its static and fatigue tests, it can be considered to have fulfilled the structural integrity criteria. However, the specification for the loading of the T-bolts is insufficient and this remains a weak point of the structural design.

The design constraints for the T-bolt connection represent a compromise between providing sufficient strength in the bolts to survive the extreme static load and making them with enough compliance that the fatigue loads are mostly carried by the surrounding laminate. The static test results clearly indicated that the joint would be capable of surviving the extreme load case. The first fatigue test (coupon 3) resulted in failure of the laminate part, giving a very optimistic impression of the joint's fatigue strength. However, during the test a large static overload was inadvertently applied to the coupon, which may have had the effect of increasing the fatigue lifetime of the steel part. The importance of the overload was not appreciated until the test on T-bolt connection test coupon 4 failed in the steel part after a very low number of cycles. The test was repeated with a replacement bolt, but the test again ended with failure of the steel part after a similarly short lifetime.

To improve the fatigue lifetime of the bolts, it would be necessary to modify the stiffness of the connection so that a higher proportion of the load is carried by the laminate. Such a change would transfer more load through the glass/polyester laminate part of the joint (and, in particular, the holes permitting access to the T-bolts). This change can be expressed in terms of the load factor of the bolts, which is the proportion of the externally applied load which is carried through the bolts.

The bolt load factors for the prototype blade have been determined and analysed. It is concluded that several bolts in the blade exceeded their allowable GL design lives by several times. The highest loaded bolts (B09 and B10 in the trailing edge) were loaded more highly than expected because the two blade sections did not fit together properly in this region. The actual load factor in these bolts was 0.33 and 0.36 under edgewise loading; the load factors for all other bolts were less than 0.21. These values should be compared with the theoretical value (neglecting geometric and contact stiffness effects) of 0.314 and the value of 0.24 measured in the coupon tests.

It is concluded that the alignment of the two blade sections must be improved in order to reduce the bolt load factors in the trailing edge zone. Notwithstanding this improvement, calculations indicate that the bolt load factors need to be reduced to 0.14 in order to satisfy GL certification requirements. Alternatively, it would be necessary to use bolts with better fatigue properties.

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5. OVERVIEW OF SECTIONAL BLADE DESIGN

5.1 Structural integrity

Both the scale model blade with connecting tubes and the prototype sectional blade with T-bolt connections survived their static and fatigue tests with no obvious signs of damage.

The connecting tubes design requires high manufacturing tolerances to achieve proper transfer of loads between the tube and the steel rings supporting it. This may incur higher costs during manufacture, but the circularity and quality of fit can be monitored as part of the quality assurance procedure. The thermoelastic stress measurements indicated stress concentrations at the join between the bulkheads and the main blade skin, particularly at the inboard connection. This is a zone of potential weakness that should be considered in future design work and monitored during certification tests. Since the two blade sections in the connecting tubes design do not close, it would be necessary to use some flexible sealant in the gap, which may be a source of potential water ingress and would need regular inspection.

The load factors in the bolts of the prototype sectional blade were difficult to predict and were ultimately too high (compared with GL design recommendations). It is interesting, therefore, to compare the performance of the T-bolt connection and the embedded bushing with stud-bolt connection.

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Mean regression line: T-bolt Bushing

Figure 5.1 : Comparison of embedded bushing with stud bolt and T-bolt connections - external load range vs. number of cycles

Figure 5.1 compares the fatigue test results for the embedded bushing with stud bolt and T-bolt connections. The figure shows the externally applied load range versus the number of cycles to failure. At first sight, the T-bolt connections appear to perform better than the bushing connections. However, the width of the T-bolt connection coupons was significantly larger than the width of the bushing specimens (155 mm vs. 109 mm), which means that fewer T-bolt connections could be placed in the circumference of a rotor blade cross section, and hence each connection must carry a proportionately higher load.

Figure 5.2 shows the same data as Figure 5.1, but now the external load has been divided by the specimen width. This figure gives an indication of the maximum load ranges that can be transferred by the connections over a given rotor blade circumference length. From the figure, it can be concluded that the two types of connection would achieve approximately the same connection strength (in terms of fatigue life) in an actual rotor blade. However, the embedded bushing connection would involve more discrete connections than the T-bolt connection.

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Mean regression line: T-bolt Bushing

Figure 5.2 : Comparison of embedded bushing with stud bolt and T-bolt connections - external load range per unit section circumference v. cycles

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S-N lines: GL, Steel, K=52.3 Based on DUT bolt tests (mean)

Figure 5.3 : Comparison of embedded bushing with stud bolt and T-bolt connections - bolt stress range vs. number of cycles and design S-n lines

Figure 5.3 shows the comparison of the measured stress ranges in the bolts with the GL S-N line (K=52.3) and an S-N line based upon TU-Delft test results. For the embedded bushing connections, the GRP part is the most critical part for all but one of the tests, whilst the steel bolt is most critical for the T-bolt connections where there has been no overload. The only bushing specimen with failure of the bolt, shows a relatively low fatigue life of the bolt, but no lower than the GL S-N design line.

5.2 Economic evaluation

The main incentive for producing a sectional wind turbine blade is to make transportation and handling of the blade easier and therefore to reduce overall costs. This is particularly relevant where the blade must be transported by road, possibly to inaccessible locations (sharp corners, small roads, weight limitations for bridges etc.). To evaluate the economic viability of a blade made in two parts compared

29

to the standard integral item, the almost inevitable increase in production costs has to be offset against the reduction in transport costs.

The selection of the cutting location (7.3 m from root) for the prototype sectional blade was chosen to ensure that the chord length at the division would correspond to the chord length found at the mid span of a much larger blade, for example, with a length of 50 m. In this way, it can be considered that a large blade will be divided into two parts of equal length, which was not the case for the LM23.3 blade. This corresponds to the most effective reduction of the transported length.

To estimate the increase of costs in the prototype sectional blade, SB23.3, a comparison has been carried out with the original LM23.3 blade. For the standard blade, the total cost can be separated into material costs and production costs. The material costs are largely based on the mass of glass fabric, resin, glue, core reinforcement etc. put into the blade. Some parts have to be considered by item: bushings for the connection at the root, lightning protection system, tip-brake mechanism, if needed, etc. The production costs are related to the production hours needed for building the blade, cutting and grinding, finishing, quality control, etc. The costs for the buildings, overheads, investments, etc., are also included in the production costs. For the LM23.3 blade, the material costs are almost equal to the production costs (49% for material costs and 51% for production costs).

During the production of the prototype sectional blade, the material costs were registered. A total augmentation in material costs of 43% has been observed compared to the LM23.3 blade. The increase in material costs is shown in diagramatic form in Figure 5.4a.

The augmentation in production costs for the Sectional Blade has been estimated at about 94% (Figure 5.4b), compared to the standard LM23.3 blade.

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(a) (b)

Figure 5.4 : Prototype sectional blade - increase of material costs (a) and production costs (b)

For the total cost price of the two blades, the costs for materials and production have been combined together and show that the SB23.3 blade with T-bolt connection was 68% more expensive to produce than the classical LM23.3 blade (Figure 5.5).

This same comparison, between the total cost price for a standard blade and the equivalent blade with mid-span T-bolt connection, has been repeated for larger blades. For this estimate, the standard formulae used within LM have been applied, giving the total cost price of the blade as function of its length [6]. Note that the design of the T-bolt connection is made with a chosen width and can thus be related linearly to the chord length, which means that a larger chord can simply be equipped with more T-bolts.

For a blade of 60 m length, the chord length at the connection and hence the number of T-bolt components can easily be determined and thus the material cost price for the connection can be calculated. The augmentation of the total cost price of a 60 m long blade has been estimated to be about 19%. The sharp reduction in the cost is mainly due to the fact that a separation in the blade using the T-bolt connection affects a relatively short part of the blade, immediately around the division, whereas the rest of the blade remains standard. Relative to the blade length, the modified area is thus smaller for a longer blade.

30

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Figure 5.5 : Prototype sectional blade - overall increase in materials and production costs

For large rotor blades, length is a potential problem for road transportation for two major reasons:

• the truck has to be long and strong enough to support the blade,

• the truck should be able to turn all corners leading to the destination.

To quantify the advantage in transport costs from a sectional blade, a study has been carried out, considering a blade with a length of 60 meters. A comparison of costs has been done for the transport of this blade, with and without division, from LM Denmark to Hannover, Germany. However, it should be noted that transport regulations vary between European countries, which could result in significant price fluctuations between countries. Furthermore, special transportation is strongly case dependent and has to be determined accurately for each individual offer.

Notwithstanding these objections, nominal cost prices have been supplied by a Danish truck company for transportation of a 30 m long blade from LM Denmark to Hannover. The cost per 30 m blade is about 2600 euros, but the costs increase by a factor of 2 for a 40 m long blade and by a factor of 5 for a 60 m long blade. The respective costs per 40 m or 60 m blade for the chosen destination are therefore: 5200 and 13000 euros.

If a 60 m. long blade has to be transported from LM Denmark to Hannover, three possibilities are considered with corresponding cost prices:

entire blade - no division 13000 euros

blade in two parts - one truck 2600 euros

blade in two parts - one truck per part 5200 euros (2*2600)

The increase of total cost (material and production costs) for a 23.3 m blade with the T-bolt connection has been estimated to be about 68%. For a larger blade (60 m.) this augmentation will be less, say about 19%.

The transport cost for the large blade has been investigated for a trip between LM Denmark and Hannover. The saving in transport costs (assuming that the two blade halves can be accommodated on a single truck) is 10500 euros, which represents about 5% of the cost price of the blade.

It therefore appears that for this case the gain in transport costs for a divided blade cannot compensate for the additional costs in production of the sectional blade. However, a divided blade might give the customer a major advantage regarding the accessibility of isolated sites where larger blades could not otherwise be transported.

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5.3 Other implications

The extension of the T-bolt connection concept to larger blades is considered in the light of likely trends in blade development.

T-bolt connection for larger blades

The design of the prototype sectional blade was based on the loads calculated for the LM23.3. For larger rotor blades, however, the loading situation will probably be different than for smaller blades. With increasing blade length and weight, the relative importance of fatigue loads compared to extreme loads increases. As the loads are strongly related to a lot of parameters, such as wind climate, power regulation system of the turbine, rotor diameter, number of blades, aerodynamic design, etc. no clear trend line can be derived. For this reason, if a much larger blade had to be produced in two parts joined with a T-bolt connection, as discussed herein, a redesign of the surrounding laminate and bolts would have to be performed, which could result in some variation from the cost estimates given above.

Furthermore, the trend followed today by almost all blade manufacturers is to reduce the chord length as much as possible in order to limit the loads, essentially on the tower. This means that fewer T-bolts could be placed in a longer blade and that the loads they would have to carry will probably increase. Due to the difficulties in defining some general trends for the loads, it is difficult to draw any further realistic conclusions relating to the design of the T-bolt connection for long blades.

Another consideration is the additional weight of the connection, particularly the reinforced laminate and steel parts, which will tend to increase the cyclic loading on the blade root generated by the rotation of the blade. If the connection is placed far from the root flange, a situation could be reached where the root diameter in combination with the existing blade root stud connection is no longer sufficient and the blade root would also have to be redesigned for the sectional blade.

Markets for sectional wind turbine blades

Today, the largest blades in production approach 40 m in length and can still be transported on the roads, albeit with special measures and trucks. For larger blades to be produced in the future, perhaps exceeding 50 m length, the principle of making the blade in two parts could be attractive for locations on land. Even if the transport of the two parts has to be realised with two trucks this may be the only option if the blade could not otherwise be transported in one piece. The option to use sectional blades will give some advantages to the customer for the installation of new land wind farms.

For road transport, it should also be noted that the height is becoming more and more of a limitation, as the blade root diameter must also increase (currently between 2 m and 3 m for the largest blades). A typical height for European bridges on major land routes is 4.2 m.

Sectional blades are probably not attractive for offshore wind energy installations. In this case, the limitations due to transport by truck disappear if the blades are produced beside a suitable harbour. When transportation by barge is a possibility, the limitations on length, width, height, or weight are no longer critical. For offshore wind energy, therefore, a division of the blades is unlikely to be interesting.

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6. CONCLUSIONS

Three connection concepts have been considered in detail during the project:

(i) embedded bushing with stud-bolt connection,

(ii) T-bolt connection

(iii) blade sections with connecting tubes.

The blade sections with connecting tubes concept has been demonstrated on a scale model blade and the structural integrity of the design was proven in static and fatigue loading. The only problem was the failure of the screw for the locking device, which can be solved by using steel of higher quality in association with general design improvements. Thus, it is concluded that a sectional blade based on the connecting tubes principle is a sound design approach.

The embedded bushing with stud-bolt connection was shown to perform comparably with the T-bolt connection in coupon tests, but after much consideration it was concluded that it would be very difficult to manufacture to the required tolerances in a large, resin-infused blade.

The T-bolt connection performed satisfactorily in coupon tests and was selected to be incorporated in a prototype sectional blade of 23 m length. This blade survived a static extreme load test and then 5 million cycles edgewise fatigue loading and 2.5 million cycles flapwise fatigue loading. However, the fatigue loads carried by the bolts is close to the limit of their strength and more accurate models of the performance of the joint are required before serious production could be considered.

A prototype sectional blade was designed and manufactured containing a chordwise array of T-bolt connections. This was selected in preference to the embedded bushing design, because of the perceived greater ease of manufacture.

The prototype sectional blade was tested with the same static loads as the standard undivided blade of the same size, in both flapwise and edgewise directions, without any damage occurring in the region of the connection or elsewhere on the blade. The prototype sectional blade was then tested with appropriate fatigue loads without any damage occurring in the region of the connection or elsewhere in the blade. No stiffness degradation of the blade was measured, and the load ranges of the forces in the bolts in the connection did not change during the tests.

A review of the costs of sectional blades found that, for a 60 m blade, manufacturing costs would increase by 15% - 20%. This can be offset against savings on transport costs, which may be up to 5% of blade cost for an average journey. Since manufacturers see no real problems in manufacturing single, integral blades up to this size, sectional blades are only likely to be viable where there are particular problems of access to a site.

Thermoelastic stress measurements were made during the test phases of the project using new, state of the art instrumentation. Difficulties of calibration have arisen due to changes in local temperature distribution, to which the TSA system is unexpectedly sensitive. Despite these shortcomings, areas of high stress concentration have been identified in the T-bolt test coupon and prototype sectional blade, which suggest that a circular hole may be more damaging than a slot, and in the connecting tubes blade at the points of attachment of the bulkheads. TSA is therefore a promising technique for design validation during full scale blade testing in the laboratory.

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7. FURTHER WORK

The work reported here has necessarily involved the insertion of a connection within the geometry of an existing blade. If the sectional blade concept is to be utilised on a commercial scale then it will need to be included in the blade design from the outset in order to:

• optimise the mass distribution in the blade and ensure that the blade root design is strong enough to accommodate the increased edgewise fatigue loading,

• optimise the lay-up of the fabrics to reduce the extra production time incurred in manufacture of the prototype sectional blade,

• define the trailing edge geometry, the leading edge flange, and the bulkhead geometry to avoid additional effort in unnecessary modification,

• consider the use of separate moulds for the two blade halves (with possibly easier introduction of resin in the thickened part of the laminate) and new moulds or filling pieces to make the shear webs,

• optimise cutting and drilling procedures to allow operations to be carried out in-house in a few hours.

In addition it will be necessary to:

• develop appropriate procedures for sealing around the connection (inside and outside) to protect the steel components from rust and to avoid humidity infiltration into the laminates,

• design appropriate ways of incorporating a new lightning protection system and tip-brake mechanism, if present,

• develop new balancing procedures and tools.

This work will only be worth carrying out once an identifiable market has been found. Certainly, the sectional blades project has not identified any insuperable technical problems against the commercialisation of sectional wind turbine blades.

The project has experienced some difficulty in accurately determining load factors for the bolts in the connection. These difficulties can be attributed to the specific geometry of the connection and contact stiffness effects. Improved safety margins, more accurate design tools, and, possibly, monitoring of bolts in the field would be required to give greater confidence in this critical point of the design.

While the tube connection is possibly the more efficient design, blade manufacturers have more experience with bolted connections. Some further work is required to evaluate the viability of the tube connection in a full size (e.g. 60 metres) blade.

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REFERENCES

1. Hald, H., Kensche, Ch.W., Development and tests of a light-weight GRP rotor blade for the

DFVLR 100 kW WEC, Proc. Wind Power '85, 27-30 August 1985, San Francisco, CA

2 Dutton, A.G., Geiger, T., Hahn, F., Olesen, M., Kensche, C., Korsgaard, J., van Delft, D.R.V., de Winkel, G.D., Design concepts for sectional wind turbine blades, European Wind Energy Conference, 1-5 March 1999, Nice, France

3 Geiger, T., Olesen, M., Connection concepts for sectional wind turbine blades, Deliverable report, EC project no. JOR3-CT97-0167,16 March 1998

4 Hahn, F., Kildegaard, C., Development of a bending spar connection concept for a "sectional rotor blade", DEWEK 2000

5 Hahn, F., Kensche, C., Paynter, R., Dutton, A.G., Kildegaard, C., Korsgaard, J., Design, fatigue test, and NDE of a sectional wind turbine rotor blade, 15th Technical Conf. of the American Society for Composites, September 2000

6 Bensoussan, R., Parametrical cost functions - Dowec task 9, Document no. TN99-364, 09/12/99