By R. Varatharajoo, E. J. Abdullah, D. L. Majid, F. I. Romli, A. S. Mohd Rafie
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Additional resources for Aerotech IV: Recent Advances in Aerospace Technologies: Selected, Peer Reviewed Papers from the AEROTECH IV, November 21-22, 2012, Kuala Lumpur, Malaysia
14) θ = θa +θw . (15) where θ a is the angle of the flapping axis with respect to the free- or mean-stream velocity, U, and θ w is the mean angle of the chord with respect to the flapping axis. As depicted in Fig. 6, in Eq. 2, α’ is given by AR * C '(k ) w α' = α− 0 U (2 + AR) (where w0 is the downwash term ). U (16) The relative angle of attack α at the ¾-chord location due to the wing’s motion is given by 3 h cos (θ − θ a ) + cθ + U (θ − θ ) 4 . α= U (17) where C’(k), F’(k) and G’(k) relate to the Theodorsen function  which are functions of reduced frequency k.
8]. The graph also shows that the air core diameter increases with increasing injection pressure for each atomizer tested. However, at higher range of injection pressure (4 bar and above), the air core diameter has become less dependent on the injection pressure, which is in agreement with the previous research . This is because increasing injection pressure will increase the tangential velocity of injection to the atomizer and increases the strength of swirling motion inside the swirl chamber.
5 Kg Rear Front Rotation Fig. 1 Baseline brake disk Ventilated blade passage modifications. The curved blade is design with close reference to the results from baseline design. The incidence angle of about 11° is estimated from the baseline air flow miss-alignment. The misalignment is addressed when constructing the helical curved super impose onto the baseline brake disk drawing. The geometries of the new designs are shown in Fig. 2 and Fig. 3 for curved 1 and curved 2, respectively. Table 2 illustrates the mass and wetted area variation for the modified geometries.