Wind tunnel tests

After this first phase of numerical simulations to define slot location and blowing angle for cover manufacturing, wind tunnel tests on a model with removable leading and trailing edge parts have been performed [9]. Fig. V.30 shows the CAD of the model and the model in the ON-ERA L1 wind tunnel in Lille.

Fig. V.30: CAD of model (left) and picture of the model in the ONERA L1 wind tunnel (right).

To illustrate, among other results, the validation of the numerical simulations, Fig. V.31 shows a comparison of the wall pressure distributions on the unslotted flap configuration between numerical simulation and wind tunnel test data without (in blue) and with control by a continuous blowing slot located at the flap knee (in red). A very good agreement is found.

Fig. V.31: Comparison of wall pressure distributions on the unslotted flap configuration between numerical simulation and wind tunnel test data without (in blue) and with control by a continuous

blowing slot (in red).

PIV measurements have also been performed by the PIV team in Lille to compare with the numerical simulations. PIV velocity fields without control (left) and with control by a continuous blowing slot are plotted in Fig. V.32. The velocity peak due to the momentum injection is clearly visible close to the wall in Fig. V.32 (right).

Fig. V.32: PIV velocity field without control (left) and with control by a continuous blowing slot at 150g.s^-1 (α = 20°, U∞ = 63 m.s^-1).

To summarize all the wind tunnel tests results obtained by continuous or pulsed blowing slots with different widths (0.25, 0.37 and 0.5mm) and different freestream Mach numbers (0.115, 0.146 and 0.175), the lift gain for the slatless configuration is plotted as function of the momen-tum coefficient. It is remarkable to see a linear increase of CNmax with Cµ. Moreover, the Cµ defini-tion proposed in eq. (3) allows collapsing all results even for different DC values of the pulsed blowing. As explained in section 1.2, since the Cµ definition allows collapsing all results, it means that pulsed blowing allows reducing the mass flow rate requirement by 30% to have the same lift gain.

Peak of velocity

near the wall

Fig. V.33: Lift gain for the slatless configuration as function of the momentum coefficient for continuous and pulsed blowing at different freestream velocities and slot widths.

Unfortunately, even with this 30% reduction in the mass flow rate, the objective to reach a 30% lift gain still requires too high values of the mass flow rate for aircraft manufacturers (two times too high in fact). Moreover, as observed in the numerical simulations with the appearance of a shock at the leading edge, this value of 30% lift gain is only achievable by reducing the freestream Mach number to values which are not realistic (M = 0.1175 instead of 0.2).

To solve the issue of mass flow requirements, another project has been launched to replace continuous/pulsed blowing slots by synthetic jets which are zero-net-mass-flux. Since the variable stroke synthetic jet presented in Chapter I section 1.1 did not allow reaching the objective of hav-ing sonic peak blowhav-ing velocities, I have proposed another technical solution based on commer-cial valves. To mimic the effect of a synthetic jet with alternating suction and blowing phases, Matrix valves are used to switch between a compressed air feeding and a vacuum source. Since the velocities are now given by the air feeding pressure, it is now possible to reach peak sonic velocities. Of course, this is not a true synthetic jet but the objective is to see the effect of actua-tors which mimic the effect of a synthetic jet on the flow. Another advantage of this technical solution is that it allows studying continuous/pulsed blowing solution but also continuous suction and synthetic jet with the same leading edge. Fig. V.34 (top left) shows the CAD of the leading edge where HP means high pressure duct (connected to a compressor) and LP means low pressure duct (connected to a vacuum pump). Following previous numerical studies showing the interest the potential of blowing on the lower side, it is possible to blow through tangential slots on the upper side but also on the lower side. Since for the suction phase, the air is coming from the ex-ternal flow close to the leading edge where the local static pressure is low (≈ 0.7 bar), to have a zero-net-mass-flux, i.e. the same mass flow rate for the blowing phase as for the suction one, this is necessary to have six suction valves (in blue, in Fig. V.34 (top right)) for one blowing valve which has a feeding pressure of 6×0.7 = 4.2 bars. Fig. V.34 (bottom) shows two pictures of the open leading edge with 162 valves to cover the 2.4 m span of the model.

Fig. V.34: CAD of the actuated leading edge (top) and pictures of the model (bottom).

This actuator has been characterised by hot-wire anemometry. Fig. V.35 shows the com-mand signal (in black) and the velocity at the slot exit (in purple: measured by hot-wire, in cyan:

estimated from a Kulite in the cavity). As usual, the suction velocity is underestimated by hot-wire anemometry since the air is coming from all directions and not just one like during the blow-ing phase. A good agreement is found between the measured velocity and the estimated one from the pressure sensor during the blowing phase which validates the use of this sensor to estimate the blowing/suction velocities in the wind tunnel where it is not possible to put a hot-wire. Fig. V.36 shows the DLR F15 model in the ONERA L1 wind tunnel.

Fig. V.35: Command signal (in black) and velocity at the slot exit (in purple: measured by hot-wire, in cyan: estimated from unsteady pressure sensor).

-250 -200 -150 -100 -50 0 50 100 150 200

0 0.01 0.02 0.03 0.04 0.05 0.06 0.07 0.08 0.09 0.1

Velocity (m/s)

-1 0 1 2 3 4 5 6

Measured velocity Computed velocity Excitation

Fig. V.36: DLR F15 model in the ONERA L1 wind tunnel.

Fig. V.37 illustrates some of the wind tunnel results. It shows the lift curves with continu-ous blowing or suction (left) and the comparison with synthetic jet on the upper or the lower side (right). Fig. V.37 (left) shows that, as expected, the larger the mass flow rate, the larger the stall delay. It also shows that a continuous suction with qm = 40 g.s^-1 (in cyan) gives the same maxi-mum lift as a continuous blowing with qm = 90 g.s^-1 (in green) which means a reduction of 55% of the mass flow rate requirement. It is well-known in the literature that a continuous suction is more effective than a continuous blowing (see ref. [44]). So, why suction is not used instead of blowing for AFC? Because, since the air density is lower in a vacuum pipe, it requires larger pipes for the same mass flow rate as blowing. So, in terms of volume requirements, it is more demanding which is a problem in the wings of an aircraft. Secondly, there is a source of high pressure on the engines but not of vacuum.

Fig. V.37 (right) shows that the same maximum lift can be obtained with a synthetic jet on the lower side (orange curve) which is zero-net-mass-flux as with a continuous blowing with a mass flow rate of 50 g.s^-1. A synthetic jet is as effective as a continuous/pulsed blowing with low-er Cµ values (0.56% instead of 0.96%) thanks to the suction phase which is the most effective one.

Moreover, a synthetic jet is more effective on the lower side (orange curve) than on the upper side (dark blue curve).

Fig. V.37: Lift curves with continuous blowing/suction (left) and synthetic jet (right).

Fig. V.38 summarizes all the results obtained with continuous/pulsed blowing and synthet-ic jet. On the left plot, the lift gain is plotted as function of Cµ. As it can be observed, the use of the Cµ coefficient does not allow collapsing all results especially for high Cµ values. The

agree-ment is improved by using the moagree-mentum ratio instead of the moagree-mentum flux ratio which means, as observed previously on a different model, that the slot width seems to have no effect on the lift gain.

Fig. V.38: Lift gain as function of the momentum coefficient (left) and of the non-dimensionalised momentum (right).

After this first wind tunnel test on a 2D airfoil, two other campaigns have been performed in the DNW-NWB wind tunnel to study the effect of sweep and in the cryogenic DNW-KKK wind tunnel to study the effect of the Reynolds number. Fig. V.39 (a) shows the DLR F15 model in swept conditions. In summary, Fig. V.39 (b) shows that the lift gains are smaller (8% maxi-mum) with sweep than previously (23% maximaxi-mum). Fig. V.39 (c) shows the DLR F15 model in the cryogenic DNW-KKK wind tunnel. Fig. V.39 (d) shows that the lift gain decreases when the Reynolds number increases. This is due to the fact that lift without control increases with the Reynolds number so the difference with the maximum achievable lift with control (Cp.M² < -1 criterion) decreases and the lift gains are smaller.

a) DLR F15 model in swept posi-tion in the DNW-NWB wind

tunnel.

b) Lift gain as function of the momentum coefficient.

c) DLR F15 model in the DNW-KKK wind tunnel.

d) Lift gain as function of the Reynolds number and the momentum coefficient.

Fig. V.39: Pictures of the F15 model in wind tunnels (left) and associated lift gains (right).