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The experiments were carried out in order to compare ‘real life’ results with those determined through the numerical finite element method of analysing a bend twist coupled composite beam. This was to enable the FE analysis to be used as part of a design tool for bend-twist coupled tidal turbine blades. In order to consider whether the numerical model is accurate the results of both analyses have been compared.

Comparison of the theoretical and experimental bend achieved along the length of the beam for a representative range of masses is illustrated in Figure 4-29. The suffixes T and E denote the theoretical and experimental results respectively. At the lower tip loads it is apparent that there is good correlation between the numerical and experimental results. With tip loads in excess of 149.3 N there is considerable divergence is evident which increases with increasing tip load. It is generally accepted that the perfect nature of numerical modelling can lead to an overestimation of stiffness, however the level of divergence shown in Figure 4-29 is too significant to be explained through numerical perfection. The numerical analysis used material properties supplied by the pre-preg manufacturer. A sensitivity analysis was conducted to investigate the influence of material properties on the beam response. It was found that a 10% change in the values of EY and EZ (the transverse and through thickness properties of the UD laminate respectively) result in an increase of beam bending of 3% for an applied tip load of 208.82N. The theoretical analysis also assumed that the geometry of the beam and orientation of each layer were perfect; i.e. the mid-layers at 60o _{are assumed to}

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the mid-layer ply angle, from 60o to 54o, resulted in a 14% increase in bending deflection for a tip load of 208.82N. These analyses indicated that the level of bend the beam experiences for a given load is heavily influenced by accuracy in the manufacturing process, and is the most likely reason for the divergence of the experimental and numerical results. At the highest tip load this discrepancy is greater; however, observations in the experiments indicated the occurrence of internal ply failure which is not modelled in the numerical analysis.

Figure 4-29: Comparison of theoretical and experimental results for bend along the length of the beam for a range of tip loads.

The comparison of the relationship between induced twist and beam length over a representative range of tip loads for the numerical and experimental results is shown in Figure 4-30. Again the divergence of the numerical and experimental results is apparent with increasing tip load. The results seem to correspond more closely at the shorter beam lengths and diverge towards the tip of the beam for each tip load. The sensitivity analysis on ply angle again showed that the response is highly sensitive to the accuracy of the manufacture.

0 10 20 30 40 50 60 400 600 800 1000 1200 Ben d ( mm)

Distance along beam (mm)

3.038 kg T 3.038 kg E 9.154 kg T 9.154 kg E 15.221 kg T 15.221 kg E 21.287 kg T 21.287 kg E 28.895 kg T 28.895 kg E

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Figure 4-30: Comparison of theoretical and experimental results for induced twist against beam length over a representative range of masses

Figure 4-31 illustrates the relationship between theoretical and experimental bend and induced twist at station 2 on the beam for each tip load. While the results are relatively consistent for the lower tip loads (up to 124 N) there is marked divergence after this point. The number of load steps in the non-linear analysis was increased to 20; however the theoretical results still maintained a linear profile and did not capture the non-linear trends apparent in the experiments. The same sensitivity analyses, regarding material properties and accuracy of manufacture, showed that induced twist was similarly affected as with the beam bend and a considerable amount of this divergence is attributed to this. However, the discrepancy is also considered to have been caused by fibre breakage occurring in the beam when the heavier masses were hung from the free end. While the sounds of fibres breaking were only heard to start under a tip load of 208.8 N, after divergence of the theoretical and experimental results occurs, it is possible that failure may have started at lower loads but been less audible with the cumulative effects of ply failure and imperfections in manufacturing causing the difference in experimental and theoretical results.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 400 600 800 1000 1200 In duced Twist (d eg rees)

Distance along beam (mm)

3.038 kg T 3.038 kg E 9.154 kg T 9.154 kg E 15.221 kg T 15.221 kg E 21.287 kg T 21.287 kg E 28.895 kg T 28.895 kg E

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Figure 4-31: Comparison of theoretical and experimental bend and induced twist for each tip load at station 2

Judging by the fact that the numerical and experimental results, presented in Figure 4-31 diverge so markedly, and not all of this can be attributed to inaccurate manufacturing, the theoretical analysis may not account for fibre breakage and first layer (first ply) failure. Figure 4-32 illustrates the behaviour of a laminate upon first layer failure.

Figure 4-32: Behaviour upon first layer failure [116]

0.0 0.3 0.6 0.9 1.2 1.5 1.8 0 10 20 30 40 50 60 0 5 10 15 20 25 30 35 Twist ( d eg rees) Ben d ( mm) Mass (kg) Theoretical Bend Experimental Bend Theoretical Twist Experimental Twist

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First layer failure can be defined as the layer or ply group that fails in a multidirectional laminate; the load corresponding to this failure is the design limit load. In Figure 4-32 the postfailure loading slopes are identical, but the curves are offset from one another. In reality, failures between fibres do not occur at precisely the same load and therefore the actual load-deformation curve is a series of small events as illustrated in Figure 4-33. This shows the progressive failure typical of a laminated composite.

Figure 4-33: Typical laminate load-deformation curve [116]

The load-deformation (mass-bend data) presented in Figure 4-31 is plotted in a similar format to the classic load-deformation curve, Figure 4-34. It can be observed that it takes a similar form to that of the load-deformation curve in Figure 4-33, thus leading to the assumption that the reason for the divergence of results between the theoretical and numerical results in Figures 4-30 to 4-32 is partly due to laminate failure. It is apparent that that a progressive damage modelling approach may be required to model the material degradation due to excessive load application.

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Figure 4-34: Load-deflection curve for the composite bend-twist coupled beam

Analysis of the Von Mises stresses experienced by the theoretical model of the beam indicated that the maximum stresses are at the root, where the beam is clamped. The maximum Von Mises stress in the beam when a load of 150N (15.22kg) was applied was 1.38GPa, whereas the maximum tensile stress of the SE84LV/HEC is 2.844GPa – over twice the value seen in the theoretical analysis. It is apparent, however, that failure of layers in the physical specimen occurred at a tip load just less than 150N. This could be due to the presence of voids and wrinkles in the laminate. Judd and Wright [117] state that the interlaminar shear strength of a composite decreases by about 7 percent for each 1 percent of voids up to a total void content of about 4 percent. Other mechanical properties are also affected [118]. Strong [119] suggests that a good quality laminate should have a maximum void content of 0.5%. Many large voids, visible to the naked eye, can be observed in the cut face of a section of the manufactured beam suggesting that the void content is considerably higher than 0.5% despite the beam having been cured at elevated temperatures and pressures in an autoclave. The presence of such defects may have contributed to the discrepancies between the theoretical and experimental results.

Initial studies (Sections 4.3 and 4.4) have indicated that the combination of a fixed pitch and passively twisting blade is likely to increase the profitability of a free stream tidal turbine in comparison with a rigid fixed pitch or variable pitch turbine blade [120]. Increases in annual energy capture of around 2.5% and a reduction in thrust coefficient of around 10%

0 50 100 150 200 250 300 350 0 10 20 30 40 50 60 Tip Lo ad (N ) Bend/Deflection (mm)

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were observed for a blade which was able to produce a tip twist in the region of a few degrees. These studies were undertaken assuming a distribution of twist of the form 0.5Svf½, where S is the blade span and V is the tidal flow velocity. This twist distribution is very similar to those produced as a result of both the FE analysis and experiments suggesting that performance improvements comparable to those mentioned previously may be expected.

Considering the carbon bend-twist coupled spar under experimentation, it may be assumed that a turbine of diameter 2.6m may have a similar length blade, with a similar thickness/chord ratio at the stock. Knowing that the thrust coefficient for a HATT, CT, is:

2 1 2 T T C Av   (4.12)

Assuming a reasonable value of CT of 0.8, the flow velocity, v, can be calculated to be 2.3m/s.

This is an acceptable value for the mean spring peak velocity for a tidal stream energy extraction site.

The experimental method used is considered to be a viable method for determining the induced twist of a bend-twist coupled beam. It will be used in future tests; however, methods to reduce the void content, optimise the repeatability of the fibre angle during layup, attachment of laser pens of higher quality will be included.

It is evident that there are limitations to the numerical model, and more care and consideration should be given to both theoretical modelling (ply wise failure) and manufacturing perfection in order to bring the results closer together. This work, however, clearly identifies a route to modelling passive adaptation for the design of HATT blades.