Deconstructing Hub Drag: Final Report on Experiments

Transcription

1 DISTRIBUTION STATEMENT A: Distribution approved for public release; distribution is unlimited. Deconstructing Hub Drag: Final Report on Experiments Alexander H. Forbes, Vrishank S. Raghav, Michael Mayo, Narayanan M. Komerath, Daniel Guggenheim School of Aerospace Engineering Georgia Institute of Technology Atlanta, Georgia phone: (404) fax: (404) Award Number: N

2 SUMMARY A progression of experiments deconstructed the hub drag of a helicopter. Starting with a cylindrical shaft, drag contributions from different elements were built up on a generic model of the hub of a 4-bladed rotor. The experiments were conducted in the 2.13m x 2.74m (7 ft x 9 ft) wind tunnel at Georgia Tech. The model included a generic hub, pitch links, blade shanks and scissor links. Load cells separating the 6 components of airloads, laser Particle Image Velocimetry (PIV), and hot film constant temperature anemometry were used to capture the values and physics of drag contributions at the level needed to refine predictions. The generic model had dimensions approximating a 25-percent scale model of a modern rotorcraft hub. Identified issues including the role of hub rotation in airloads and flow deviation were examined in greater detail. Mounting the motor driving the rotating hub directly to a 6-DOF load cell enabled azimuth-resolved measurement of all 6 airload components. This in turn enabled identification of interaction effects. Various hypotheses were explored to investigate reasons for cited discrepancies in the published literature. The side force loosely ascribed to the Magnus effect was studied for its contribution to the vehicle power demand, indistinguishable to the engine from power demand due to drag. The side force was found to be highly unsteady depending on flow conditions, and its azimuth-averaged value could change sign with advance ratio. The torque required to drive the hub increased with rotation speed and freestream speed, but was found to be of the same order as that expected from the drag difference between shanks on the two sides. The azimuth-resolved force and torque data suggest that transient torque values peak at certain portions of the azimuth that slightly lead those of highest drag, for the configurations studied. The addition of a scissors linkage to the generic hub resulted in a 10 percent increase in drag above the prior generic configuration, but reduced the amplitude of periodic drag variation. Wake surveys using a hot-wire anemometer probe, and using Particle Image Velocimetry, captured the flow deviation, wake deficits and frequency content of the wake. The large-amplitude fluctuations posed challenges to hot-film probe anemometry in the sharpest gradients of the wake shear regions. The frequency content was surveyed using both Fast Fourier Transform and Wavelet Transform techniques. The total power demand attributed to the hub is still mostly due to drag. Scaling to vehicle flight speed and size should improve this approximation, unless interactions with pylons, nacelles or other components cause asymmetries sufficient to change the result. In general, the results showed that the drag of individual components behaves as expected, and is explained to good accuracy by relatively simple estimation techniques. Interaction effects between flowfields of different components pose substantial issues to accurate prediction. An example was the finding that addition of the shanks did not increase drag, because the wake of the shanks reduced the dynamic pressure experienced by downstream components, and the flow deviation reduced the drag due to base separation. Diagnostic advances included high-accuracy, reliable 6-DOF load measurements in wind tunnels using load cells, and the ability to capture finely-resolved azimuthal variations of all 6 load components with a rotating setup. These set the stage for deeper studies of interaction effects. A computational effort was conducted under the same project using Navier-Stokes solvers and a PhD thesis is in progress from that effort at this writing. Those results are expected to be reported separately. 2

3 CONTENTS NOMENCLATURE 4 INTRODUCTION 5 LONG-TERM GOALS AND OBJECTIVES 6 PRIOR WORK 7 MODEL CONFIGURATIONS 10 AERODYNAMIC LOADS 16 VELOCITY FIELD MEASUREMENTS 33 WAKE VELOCITY FLUCTUATION MEASUREMENTS 39 SUMMARY OF DATA AND RESULTS 49 CONCLUSIONS 56 REFERENCES 58 WORK COMPLETED 62 IMPACT/APPLICATIONS 62 TRANSITIONS 62 RELATED PROJECTS 63 PUBLICATIONS 63 HONORS/AWARDS/PRIZES 63 DEGREES EARNED 64 ACKNOWLEDGMENTS 64 3

4 NOMENCLATURE b Cylinder radius, m C DC Cylinder drag coefficient D Drag, N. Also used to denote hub diameters as in 1D, 2D etc. D h ) Hub diameter Y Side Force, N τ z Torque around axis of rotation, N-m D a Drag of advancing side, N D r Drag of retreating side, N F M Side force due to Magnus effect, N F s Data Sampling Rate, Hz f f Fourier Frequency, Hz R Hub radius, m Re Reynolds number s Wavelet Scale U Free stream velocity, m/s α Spin Ratio µ Advance Ratio: ratio of freestream speed to shank tip speed, (1/α) Ω Hub rotation, RPM ω Hub rotation, RPS ρ Free stream density, kg/m 3 Ψ Azimuthal angle, degrees P Power, W 4

5 INTRODUCTION Parasite drag is a limiter of helicopter forward flight speed and efficiency. The rotor hub is a challenge to designers and aerodynamicists alike: it contributes up to 50 percent of the parasite drag of the vehicle Charles N. Keys (1975). It poses a large and complex obstacle in achieving, predicting and reducing drag. The drag contribution due to individual elements can be calculated to a reasonable approximation for conceptual design from the methods given by Hoerner (1965). However, experience in the literature appeared at the outset to suggest that the effect of a rotating hub assembly on the parasite power of a helicopter in flight, was substantially different from those obtained using the above approximations. This project sought to investigate these discrepancies using generic configurations and basic research. The experimental work was conducted in 3 stages. Initial experiments measured drag of a non-rotating hub assembly as it was built up piece by piece starting with a cylindrical shaft. A drag deconstruction process was used to gauge the difference between the drag on the assembly versus a linear superposition of the drag of individual components. Next, flowfield measurements were conducted. A Pitot probe rake was initially used to survey the dynamic pressure deficits in the wake. These results are not presented in this report, since they were superseded by Particle Image velocimetry maps. Interaction effects appeared for instance where the wakes of blade shanks reduced the onset flow dynamic pressure for separation at the aft portion of the hub, thereby reducing the drag below that expected from superposition. The flow features in the wake of a rotating hub were captured using Particle Image Velocimetry. As reported in Raghav et al. (2012), once the loads agreed with computations, the detailed wake structure behind a rotating hub agreed with surprising fidelity with computations, showing at least that any unmodeled support interference was negligible. The next area of major uncertainty was the effect of rotation, especially as relevant to high-speed flight of compound helicopters. Wind tunnel and flight test lore suggested much stronger effects of rotation on the overall power increment ascribed to drag, than cursory analysis would explain. Our experiments examined the question of whether the side force, or the power required to generate the torque to overcome asymmetric loads, may be substantial contributors to the vehicle power requirement (indistinguishable from power required due to drag). The magnitude and relevance vary widely with vehicle size and advance ratio, which translate to large changes in relative velocities and Reynolds numbers. Regimes with steady wakes as well as those with periodic shedding were encountered. Prior work had also shown the importance of the horseshoe corner vortices generated where the hub mast meets the vehicle hub fairing. These vortices can interact with the flowfields of the blade shanks, 5

6 generating periodic effects and pulsing effects on the drag and side force. Such vortices are not expected in the present experiments as there is no hub fairing, nacelle or fuselage present. Major challenges in measuring and predicting hub drag include the following: 1. Complex Geometry (a) The complex geometry leading to flow interactions over the hub create interference drag which is hard to interpret. (b) Interactions augmented by presence of fine structures such as tubes, wires, linkages and fasteners. 2. Effect of Rotation (a) Consideration of all hub structures in a typical assembly requires the analysis of a large range of Reynolds number effects. (b) Measurement and prediction of drag of rotating components is a challenge in itself. (c) For high speed applications, the hub may experience compressibility effects, adding a new dimension of complexity. LONG-TERM GOALS AND OBJECTIVES The long-term goal was to develop methods to understand the airloads on complex-shaped objects and vehicles. The objectives of this project were: 1. Isolate and quantify the different sources of hub drag. 2. Tighten tolerances of empirical upper-bound predictions suitable for conceptual design, by reference to the basic experiments. 3. Enable first-principles prediction of drag through computational aerodynamics, suitable for the preliminary design stage and beyond. 4. Advance the state of knowledge on flows around complex configurations involving flow separation and a wide range of Reynolds numbers. 6

7 PRIOR WORK Early work on hub drag is described in Churchill and Harrington (1959); Linville (1972); Sheehy (1977). Churchill and Harrington (1959) did full scale tests on a direct tilting hub with no blade shanks at 0, 100, 150, and 200RPM at a velocity of 50 m/s. Rotation was found to have no effect on parasite drag. Tests were also done on a fully articulated hub with a swash plate and control system but no shanks, at 50RPM and at velocities of 24, 29, 35 and 42 m/s. A very slight increase in drag was found with increase in advance ratio. Linville (1972) tested a Sikorsky S compound helicopter hub with no fairing, a floating fairing that incorporated a pylon and a boundary layer control system, and a rigid fairing with a sealed cover for the rotor head. The gross model drag increased slightly as RPM increased from 0-100%, for some configurations, attributed possibly to the Magnus effect. The data did not reveal any consistent significant effect of RPM on the model or rotor head drag. Sheehy and Clark (1976) and Sheehy (1977) reviewed available hub drag data, and pursued reduction of the interference drag between the hub and the pylon. Charles N. Keys (1975) laid out guidelines for reducing the parasite drag due to rotor hubs. Kerr (1975) studied the relationship between helicopter drag, loads and component life. Hoffman (1975) studied the relationship between helicopter drag, stability and control. Williams and Montana (1975) laid out a comprehensive NASA plan to reduce helicopter drag. The bluff body wakes due to rotor hub components have a large detrimental impact on the performance of many types of rotor powered vehicles. Work on the topic has used analogies with simple model geometries such as cylinders Roshko (1993); Chng and Tsai (2006). More complex hub geometries were studied recently by Bridgeman and Lancaster (2010a,b) using an unstructured Navier-Stokes flow solver for the bluff body problem, and sought to identify a physics based analysis methodology capable of accurately predicting drag on realistic geometries. Wind tunnel data showed that the typical computational resolution used at Bell Helicopter for fuselages was able to predict drag to within 5% of measured values. Much finer meshes were needed to achieve the same results for hub configurations. The addition of hub fairings is meant to reduce separation and interference drag from the rotor hub and fuselage, see Young et al. (1987). Drag reduction on a 1/5 scale Bell Textron Model 222 Helicopter of 20% was shown by Martin et al. (1993) by using a small hub fairing combined with a non-tapered pylon fairing. Without the small hub fairing the model s drag was reduced by 14.5%. Martin et al. (1993) additionally showed that for an unfaired hub there was negligible influence on the model forces due to the hub s rotation rate. This agrees with the Linville (1972) result that Ω had no noticeable or consistent effect on total model drag. Felker (1985) tested an XH-59A advancing blade concept helicopter rotor hub at four different 7

8 configurations. The hub had very few protuberances compared with conventional, articulated helicopter rotor hubs, with no flap hinges, lead-lag hinges, or lead-lag dampers. Pitch horns and pitch links for the upper rotor were installed in the rotor hub. The faired hub drag was not significantly affected by rotation. The unfaired hub drag was significantly influenced by rotation. The aerodynamic cleanliness of the rotor hub may have been a factor in how the rotation rate reduced the drag. The drag of a circular cylinder is initially reduced by rotation about its axis, and a similar effect could occur on the inter-rotor shaft of the XH-59A. Felker found that for the unfaired hub s of the XH-59A the drag area was actually reduced by the rotation of the hub, possibly as a result of the cleanliness of the XH-59A relative to other hub designs. Increasing the distance between the rotor hub and pylon has shown no significant reduction in drag due to an offsetting increase in rotor shaft and control drag, per Charles N. Keys (1975). This offsetting increase in rotor shaft and control drag could be countered by placing the pitch links in the shaft and adding a fairing to the shaft. Keys and Rosenstein showed that at low advance ratio, an increase in average shank dynamic pressure explains the drag increase. However, wind tunnel tests with and without rotation do not conform to expectations. Wind tunnel tests performed on compound rotorcraft show large uncertainty in source and magnitude of what is called interference drag, which must be well-understood if reductions in hub drag are to be consistently applied. While gaps in the understanding of the nonlinear physics remain, advances have been made in hub drag design. Fairing designs have been explored by industry, see Wake et al. (2009a), to reduce flow separation and interference drag between the hub and fuselage. To date, frontal swept area of the hub design has been the leading parameter tied to hub drag per Sheehy and Clark (1976), therefore the fairing of an existing hub design does not address the issue directly. This is especially true for articulated hubs, where empty space is required for the control hinges. While empirically corrected analytic estimates have been developed to predict hub drag based on frontal area, there is no consistent trend when accounting for interference effect and frontal swept area, per Sheehy (1977). NASA and industry researchers showed that a cambered, flat-bottomed hub fairing was effective due to elimination of separated flow between the hub fairing and pylon, and elimination of shedding from upper corners of the pylon. Gaps between the pylon and cambered fairing negate benefits from having a cambered flat lower surface. A survey by Carter Copter shows complex flow interactions. Hub/pylon gap interference can cause a superflow region where Mach numbers can reach high, even supersonic, levels in the case of compound rotorcraft. Components encountering this flow experience increased drag. More recently, in-flight data on the RSRA vehicle were compared to wind tunnel test data with some success. Bluff body drag has been studied on automobiles, however the complexities due to relative motion of 8

9 components and the rotor-airframe-wake interaction of hub drag are unique to rotorcraft. Configuration-based factors affecting hub drag include the center section, the shanks, hoses and cables, the pylon/ hub gap, the fuselage attitude, the drive shaft, swash plate, pitch link rods, the pylon opening beneath the hub, the number of blades, and fairing leakage. Several of these components are in the transitional regime of Reynolds number under typical full-scale rotorcraft operating conditions, and have large drag coefficients. The confusion regarding the effect of rotation becomes evident when dealing with compound rotorcraft that operate at high advance ratios. Under these conditions, the predicted contribution from rotation should be smaller than the data apparently indicated. Flow separation due to interactions between the hub, blade shanks and pylon is a large contributing factor to rotor hub drag, per De Gregorio (2012b). Recently, the FLUENT code was applied to dual rotor hubs for validation with wind tunnel data, per Wake et al. (2009b). Acceptable correlation was achieved for bluff-body geometries, but a hexahedral mesh was needed. Methods for drag reduction were presented including fairing reshaping to reduce flow separation and interference drag, splitters and vanes to reduce flow separation, flow control using steady and unsteady blowing concepts in the non-rotating frame of reference, streamwise vorticity generation to help the flow remain attached on the fairings, and reshaping of the pylon/fuselage to redirect flow away from the bottom hub fairing. The vane and splitter methods showed marginal benefits, but the fairing reshaping resulted in more than 70% relative drag reduction. Larger shaft fairings reduced drag on the top and bottom hub fairings. Reshaping of the shaft fairing airfoil shape reduced drag on all components of the hub, an hour glass shape of the shaft fairing in the vertical direction reduced top and bottom hub drag, and improved shaping to reduced interference drag. Several experimental efforts on hub drag have been directed towards improving drag characteristics of current hub designs by the addition of fairings. Considerations for hub displacement from the fuselage have been made, weighing the effects of increased frontal area to decreased interference Sheehy and Clark (1976). A typical helicopter hub is comprised of a myriad of bluff body components, mainly due to compromises in manufacturing procedures. The distinction between streamlined and bluff bodies is associated with the poor performance characteristics (such as lift to drag ratio) of the latter configurations. The primary drag component in these bluff bodies is due to flow separation rather than viscous effects. The pressure drag of bluff bodies results from vortices in the wake that are shed from the structure. Bluff body wakes associated with rotor hub components affect the performance of both commercial and military air vehicles Gregory et al. (2008). While many studies of two-dimensional (2D) bluff bodies 9

10 have been performed in the past, three-dimensional (3D) studies have been typically restricted to spheres and cylinders of varying aspect ratios and/or Reynolds numbers. Reynolds number sweeps provide a sequence of distinct bifurcations which produce significant and measurable differences in the flow field per Roshko (1993). Seidel et al. (2006) experimentally found two counter rotating helical modes in the wake of bullet shaped bodies. The current objective is the quantification of different sources of hub drag aimed at tightening the tolerances of empirical upper-bound predictions suitable for conceptual design by reference to basic experiments. Recent experimental investigation on a complete helicopter by De Gregorio (2012a) and coupled experimental and CFD investigations by Antoniadis et al. (2012) suggest that the main hub wake formation as a reason for the unsteady forces observed on the horizontal stabilizers. Hence, the flow field details could provide useful insights into the well known problem of helicopter tail shake Ishak et al. (2008); de Waard and Trouve (1999). Prior aerodynamic experiments in wind tunnels on rotating models (see Horanoff (1969); Mehta (1985); Smits and Smith (1994); Sayers and Hill (1999); Asai et al. (2007)) were unable to resolve the loads with high azimuthal precision. We were able to acquired azimuthally resolved, 6-DOF airloads on rotating hubs. MODEL CONFIGURATIONS Experiments were conducted in the 2.14m x 2.54m (7 x 9 ) test section of the John J. Harper low speed wind tunnel located at the Daniel Guggenheim School of Aerospace Engineering of Georgia Institute of Technology. A generic four bladed hub model was assembled to approximately one-quarter scale to that of a 10-ton helicopter (Figure 1). The model consists of the hub plates, shanks for the rotor blades, a swashplate with pitch links, drive shaft, and supports. The effect of capping the shanks was studied initially, and then blade shanks were capped for the rest of the experiments performed. In the final stage of experiments, two models of the scissors linkages used on modern hubs were also attached to the swash plate, shown in Fig. 5. A complete description of the model can be found in Raghav et al. (2012). The model includes structures to represent hub plates, blade shanks, a swashplate, pitch links, drive shaft and the required hardware for assembly. Though greatly simplified from the complexities of full-scale hubs with hydraulic lines and intricate mechanisms, this model provides the interference and flow separation attributed to the full-scale system while maintaining the principal geometric characteristics. The model was adapted into three configurations, as shown in Figure 2. Configuration 2(a) has unplugged blade shanks, configuration 2(b) has plugged shanks, while configuration 2(c) has plugged shanks and a capped region between the two hub plates 10

11 Figure 1: Hub model setup inside the John J. Harper wind tunnel test section (a) Unplugged (b) Plugged (c) Hubcapped Figure 2: Three model configurations for static hub experiments and computational simulations 11

12 Full Hub No Shanks No Hub No Pitchlinks No Swashplate Full hub + scissors Figure 3: Flowchart depicting progression of deconstruction. Figure 3 shows the progression of deconstruction tests, followed by final experiments with the full hub with and without a pair of scissors components added. In the final stage of experiments for rotating hub airload measurements and hot-film anemometer surveys, the hub was mounted on a stepper motor as seen in Fig. 4. The motor was mounted on a load cell for force and torque measurements. The load cell was supported by a tripod attached to the wind tunnel support structure. Unless otherwise stated on any of the following figures, a circular fairing was placed around the motor.the hub was accelerated to the desired Ω using the stepper motor. After the hub reached the desired Ω, the wind tunnel was turned on and taken to a specified speed. When finding azimuth resolved loads the hub was first set to 0 yaw and the tunnel was set to the desired speed. Once at the test speed the hub was then accelerated to Ω. Knowing the time taken to reach Ω permitted the instantaneous counter reading to be related to the instantaneous azimuth of the hub based on the yaw reference. Magnus force, the force perpendicular to U around the rotating hub is also of interest. The Magnus force on a cylinder can be negative over a small range of parameters per Fletcher (1972). The possibility of this occurring on the rotor hub was investigated. In addition to the Magnus force, the effect of varying Ω and tunnel speeds on drag and torque were investigated. The effects of the advance ratio, µ described in Eqn. 1 on drag, side force, and torque around the axis of rotation were also investigated. Some test conditions were selected to coincide with the work of Fletcher (1972), who described a critical Reynolds number regime in which negative Magnus effect was shown to occur over rotating cylinders. His results specified a Reynolds number range from 1 x10 5 to 5 x10 5. The critical Reynolds number regime corresponds to wind tunnel speeds ranging from 6 to 13m/s. In addition to focusing on the critical Reynolds number regime, critical spin ratio (α) values were also investigated. The critical α value is approximately 0.2, corresponding to an advance ratio of 5. The side force in these two sets of critical regions was specifically investigated. 12

13 Test Configuration Azimuth/RPM Increments Velocity Static 2(a),2(b),2(c) mph Rotating 2(b) 4, 80, 240 RPM mph Table 1: Summary of Experimental Test Conditions The progression of data acquisition was initiated with static hub tests performed on all the configurations to understand the difference in drag behavior. The model was swept through azimuthal orientations throughout a series of runs performed at zero angle of attack (pitch angle). Six-axis force transducer data were obtained for a 90 azimuthal sweep in 15 increments at a range of tunnel speeds. Subsequently, tests were performed on a rotating hub with the configuration 2(b). A range of rotation rates from 4 to 240 rpm was explored. The test conditions are summarized in Table 1. In addition, the plugged shank configuration was broken down in stages to measure the individual contribution of each structural component. Since most theoretical drag prediction methods have been developed for the interference effects between two streamlined bodies, or one streamlined body and one less so, for example a wing joint with a fuselage, measuring the interference effects of the hub components was left to experimental trials. Therefore, the hub drag was measured in progression as depicted in the flowchart (Figure 3). The effect of rotation on the deconstructed model was also measured and is presented in this report. Intermediate results were presented in several publications. Ortega et al. (2011) presented measurements and computations on the deconstruction of the drag of a generic hub in a wind tunnel freestream. The scaling of loads between model scale and full scale helicopter hubs was investigated Shenoy et al. (2011, 2012). Raghav et al Raghav et al. (2012) reported loads, wake velocity fields, turbulence spectra and rotation effects on the same generic hub model. µ = 60U 2πΩR (1) {bf Support and probe interference Support interference is always present in wind tunnel experiments. In the experiments conducted here, standard precautions were taken, such as ensuring that maximum local blockage of test section frontal area was well below the accepted 5 percent. Fairings were used around the stepper motor to minimize azimuthal variations in interference. Tare readings without the model mounted, and with the model mounted and no flow on, were used for linear subtractions, which are of course not perfect ways if there is substantial interference present. Comparisons with computations presented in the papers published 13

14 Figure 4: Rotor hub model and motor Figure 5: Swash plate and scissors link assembled with drive shaft 14

15 Figure 6: Sketch of the support structure for wake PIV measurements from this effort, for instance Shenoy et al. (2011, 2012) and Raghav et al. (2012) indicate that the support interference effects were indeed minimal, or were adequately accounted with the given support geometry. In the case of the wake measurements, downstream and in-plane interference are both serious concerns. Two different setups were used, one for the Particle Image Velocimetry data acquisition, and the other for the hot-film anemometer surveys. The PIV set up is considered first. It is sketched in Figure 6. The items placed inside the test section included a traverse system to move the laser sheet generating optics vertically and horizontally, and the PIV camera mounted on its standard camera rails and attachments, attached to the tunnel floor directly below the horizontal measurement plane. Concern about interference would be justified in this case. However, as shown in Raghav et al. (2012), near-perfect agreement was shown by the computational group with the shear-layer profiles provided to them, without considering the cameras and light sheet generating setup in the computational grid. In the case of the hot-film anemometer probe surveys, probe interference is again a concern, and unavoidable. This was minimized by pointing the probe upstream so that the flow encountered the sensor with minimal disturbance. However, the support attachment for the probe, and the traverse system, were in this case far less intrusive than the setup for the PIV wake surveys, as shown in Figure 7. Based on the above reasoning, we submit that there is no need to model any support interference in correlating with any of the data in this report. 15

16 Figure 7: Sketch of the support structure for wake hot-film probe anemometry AERODYNAMIC LOADS Force measurements used an ATI Gamma load cell, whose specifications are shown in Table 2. The hub setup was mounted on the load cell. Force and torque data from the load cell were recorded at the sampling frequency described in Equation 2, which is far above the expected 100Hz frequency response range of the load cell. The goal of the force measurements was to find the side force, drag, and torque around the axis of rotation for several and tunnel speeds varying from 0 to 23m s. F s = 3600Hz (2) Measurement Range Sensitivity Limit X and Y Force 130 N N Z Force 400 N 0.05 N X and Y Torque 10 N-m N-m Z Torque 10 N-m N-m Table 2: LOAD CELL SPECIFICATIONS For rotating hub experiments, load cell data acquired at a high digitization rate were phase-averaged with 1 degree azimuthal resolution, indexed to a selected position of the hub. The formula for number of data points in each set is shown in Equation 3. The F S for each was chosen in Equation 2, such that there would be 10 data points per rotation per degree for phase averaging. The technique was verified 16

17 through a static test case comparison to the quasi-steady rotation of an aspect ratio = 1 circular cylinder. Data points per degree = F S ω (3) Selected results are presented in a progression from integrated loads to flow field details. The results are broken down into three main categories: 1. Static and dynamic force measurements 2. Wake velocity characterization. 3. Wake frequency spectra First, the measurement of integrated drag and side force is presented from experimental results. In addition, the effects of rotation are inferred. Next, the wake velocity field is examined. Finally, spectra of speed fluctuations in the wake are presented. Figure 8 illustrates the static, rotating and deconstructed hub drag for a range of wind tunnel speeds. Several configurations are shown together to highlight the many qualities of the drag contributions. The unplugged configuration had the shank tubes open for through-flow: this was rejected as non-representative of helicopter applications. The capped configuration closes the remaining through-flow gaps in the hub. This is seen to have only a minor effect on drag. The capped hub configuration creates complete flow blockage through the center of the hub plates and results in the least drag at the 45 static orientation. Therefore, it appears that the drag due to flow separation on the capped configuration is less than the interference drag caused by the channel-like flow through the center of the hub on the non capped configuration. The maximum drag is obtained for the static hub when it is oriented at 0 azimuth (blade shanks normal and perpendicular to the free stream). This orientation corresponds to the maximum frontal area of the hub. The 45 static orientation and the hub in rotation at 240 rpm result in nearly the same drag. This result is not unexpected, as the 45 orientation is equivalent to the average frontal area of the hub in rotation. The result is also in agreement with previous findings by Sheehy (1977) for unfaired hubs. The drag forces for each test condition are shown in Figure 11. For the values of Ω tested, rotation did not have any noticeable effect on the drag of the rotor hub. Figure 12 displays the roughly 10 percent increase in drag caused by the scissors at each U. Drag normalized by the tunnel dynamic pressure, as shown in Figure 9, shows little variation due to change in Reynolds number in the speed range covered. The complete model at 0 orientation shows 17

18 (a) (b) Figure 8: Variation of the hub drag with azimuth and free stream speed. 18

19 Figure 9: Drag scaled by tunnel dynamic pressure. slight Reynolds number effects, however the scaling of the vertical axis may be understating the variation. Collecting load data for a rotating model presents the challenge of decoupling the drag component from the anti-torque of the rotating shaft and motor system. The force transducer is fixed just below a compact high torque stepper motor which drives the hub shaft. Data indicate negligible coupling of the measured torque about the shaft axis and the tunnel axis. This was verified by detailed load cell calibration with the mounted hub model and by comparison of torque measurements with static wind tunnel tests. In order to clarify trends, torque measurements are shown as positive values in Figure 10(a). However, due to the orientation of the force transducer with respect to the model, measured torques along the drive shaft axis were aligned along the negative z-axis of the load cell. The torque measured along the drive shaft axis is observed to increase in magnitude with increased 19

20 (a) Measured torque along drive shaft axis (b) Side force variation with hub rotation Figure 10: Variation in forces with changing rotational and free stream speeds. rotation rate as delivered by the motor (Figure 10(a). The static model also measures an increasing torque for greater tunnel speeds. The model is nearly axially symmetric, however the connection joints of the pitch links to the blade shanks provide some asymmetry, thus generating the measured torque. Further investigation with an axially symmetric model is planned to confirm this assumption. Initial measurements showed little variation in side force due to rotation, and even a reduction of measured side force at greater tunnel speeds (Figure 10(b). The side forces observed for each test condition are shown in Fig. 13. The side force on the hub appears to have some dependence on Ω, however there is no discernible trend due to Ω. Raghav et al. (2012) found there was a non-zero side force for the static rotor hub, which was attributed to model asymmetry (the pitch links are offset from the diameter). The side force followed a seemingly linear trend, with a shift in side force at 13 m/s followed by a return to the linear slope. This shift shows a sign reversal in side force for some range of Reynolds numbers between and While the side force changes sign for certain test conditions there is no evidence that this is due to Magnus effect. The torque was demonstrated to increase with not only Ω but also with U. This effect is seen in Fig. 14. Both Ω and U have a large effect on the torque required of the motor to maintain rotation. The reason for the torque dependance on Ω was needed to move forward. To simplify the problem, the blade shanks were considered independently to see if a reason for the 20

21 Figure 11: Drag Force Vs Freestream Speed for Various Ω rising torque could be found. The drag on the blade shanks were assumed to be major torque contributors due to their distance from the hub axis. Figure 15 shows the direction of drag at the shank tip for advance ratios less than 1. For µ <1 the drag on the shanks on the retreating side is aiding the rotation of the hub, however, after µ becomes >1 the drag on the retreating side opposes the direction of rotation. Torque due to the shanks is described by Equation 4 τ zshanks = (D a D r )R (4) Drag on the cylinder shanks was calculated using the average velocity that would be experienced on each shank. Since both shanks are cylinders, C DC needs to be adjusted depending on the Reynolds number, which is near the transition between laminar and turbulent flow for many of the test cases. Velocity used in the shank drag calculations is dependent on µ. The difference in the average velocity between the advancing and retreating shanks is fixed based on the hub Ω. However, since drag on each shank scales with U 2, the difference in drag between the two shanks, and by extension the torque, increases with increasing U or Ω. The drag difference experienced by the advancing and retreating sides was seen in previous experiments on rotating hubs by Raghav et al. (2012). 21

22 Figure 12: 120 RPM Drag Comparison with/without Scissors with no Motor Fairing and no Tare Removed 22

23 Figure 13: Side Force Vs Freestream Speed for Various Ω 23

24 Figure 14: Average Torque Around Axis of Rotation Vs Freestream Speed for Varying Figure 15: Direction of Drag Forces of Shanks ( < 1) 24

25 Figure 16: Aerodynamic Loads Through A Hub Cycle (80 RPM 8.94 M/S) 25

26 Figure 17: Drag During Two Hub Cycles With No Motor Fairing Phase Averaged Through 10 Seconds(120 RPM 11.2 M/S ) 26

27 The removal of the blade shanks to examine the deconstructed model indicates that the blade shank contribution to the total drag is negligible. Furthermore, removal of the hub plates reduced the total drag by approximately one-third. Consequently, two-thirds of the hub drag is due to the contributions of the drive shaft, swashplate and pitch link drag. Rotating each of these deconstructed models confirms that there is little variation in drag from the static configuration. The pitch link drag contribution may be computed from the shift in measurements from the prior deconstructed model. Experimental measurements indicate that the time averaged force transducer data do not vary with the onset of hub rotation, as noted previously. However, azimuthally varied static hub measurements do capture a variation in the drag with orientation of the model. Note that the measurements shown in Figure 8(a) do not include the drag due to the motor used in the drive shaft. Using the work of Raghav et al. (2012) on the same generic hub model the drag breakdown of each component was found. The addition of scissors was found to add 10 percent, leading to the drag breakdown shown in Fig. 18. Interference drag could impact the amount of drag from each component. To find if there was any impact of interference on the drag breakdown a simple estimate of drag of each component was made. The hub shaft, pitch links, and shanks were easily simplified as plain cylinders. The swash plate was considered to be a bluff body with a small thickness. Using these simplifications the drag could be estimated to a reasonable degree. The comparison between calculated and actual drag is shown in Fig. 19. The swash plate + shaft is largely under predicted, the reason for which is unknown. The differences in drag for the shanks may be due to a reduction in drag of the shanks due to interference with the rotor hub. Figure 16 shows the aerodynamic loads through 1 rotation of the hub. The resolution of the drag and side force is 1 degree, where as the resolution of the yaw moment is 5 degrees. The lower resolution is due to poor signal to noise ratio of the data resulting from the high stepping frequency of the motor. From the figure there are 4 orientations per rotation where the drag, side force and torque exhibit a maximum value. These 4 peaks correspond to the azimuthal orientations of the hub where a set of shanks is perpendicular to the flow. Additionally, at the 0 orientation the hub scissors are parallel to the flow. The drag variation with time through 2 revolutions of the hub, with and without scissors is shown in Fig. 17. The presence of the scissors linkages appears to cause higher drag during the azimuth intervals where drag was low in the absence of the scissors. The drag was previously shown to increase by 10 percent with the addition of the scissors. This is demonstrated in Fig. 17 by comparing the average drag across the 2 cycles. The average torque is plotted against advance ratio in Fig. 20. This plot shows a very interesting 27

28 Figure 18: Drag Contributions of Components Figure 19: Component Drag Expected Vs Measured 28

29 Figure 20: Average Torque Vs µ for varying Ω 29

30 behavior. At the low rotation speed of 80 rpm, the torque variation with advance ratio has a slope exponent less than one, i.e., the slope decreases and the curve flattens out as advance ratio increases. At 120 rpm, the slope is constant, so that the variation is linear with advance ratio. Above 120 rpm the torque rises with increasing slope as advance ratio increases. At the highest rpm tested, 240 rpm, the slope is very steep. Power required from the motor was investigated in this study. The power due to each aerodynamic load (drag (D), side force (Y) and torque (τ z )) at each condition was computed using the following expressions: P D = D U (5) P Y = S U (6) P τz = τ z 2πω (7) The comparison between power usage due to each aerodynamic load is shown in Fig. 21. The Power requirement to overcome drag on the rotor hub is significantly larger than either of the other 2 components, due to drag power scaling with U 3. The drag power requirement is not changed by the hub s µ and side force power, as with side force seen in Fig. 13, does not change consistently with µ. Torque, however, does have consistent changes in power requirement with µ, shown in Fig

31 Figure 21: Power Comparison at 240 RPM 31

32 Figure 22: Torque Power Comparison at Varying RPM 32

33 VELOCITY FIELD MEASUREMENTS A detailed velocity map of the hub wake was obtained via Particle Image Velocimetry. The light sheet illumination was provided with a Litron double-pulsed Nd:YAG laser(532 nm, 200mj/pulse) and the scattering from seed particles entrained in the flow was captured by a LaVision Imager Intense CCD camera (Imager Pro X 2M pixels, 14bit). The flow was seeded with 2µm-3µm droplets produced with a aerosol seeder. The recording rate yielded 14 velocity fields per second. The time separation between the two light flashes used to capture particle displacements was set based on the velocity range expected and adjusted as needed based on the velocity range measured. One hundred PIV image pairs were acquired for each field captured. The data presented were obtained by averaging the 100 velocity vector fields. In cases where the hub rotated, the averaging was done with phase locking to obtain azimuth-resolved velocity fields. Single realizations of the velocity field are usually inadequate, because the seeding may not be present adequately in all parts of the image in every instance, resulting in several areas where velocity vectors could not be obtained. This would hinder construction of derivative fields such as vorticity. Velocity deficits in the wake at a location one-half of the hub diameter (Figure 23) downstream of the hub axis were profiled for several model orientations. Figure 24 illustrates the typical variation observed in the wake profile for the hub in two static configurations and one case in rotation. The data show a contraction of the momentum deficit in the hub wake when the model is oriented at 45 azimuth, corresponding to a reduction of hub drag based on frontal area (Figure 24). The largest deficit at the static hub centerline is clearly defined by the strong velocity deficit just behind the hub main shaft. The two secondary wake deficits appear behind the pitch links to the left and right of the hub shaft. When the hub rotates in a counter clockwise direction, the primary velocity deficit is translated upward and to the right, appearing behind the right (aft looking forward) blade shank. The velocity deficit due to the main drive shaft appears to have coalesced with that of the right pitch link, leveling only a secondary velocity deficit from the left pitch link, which has also translated in the positive y-direction (to the right) of the flow induced by the counter clockwise rotation. The rotated wake is seen to shift with the hub rotation direction. Also, the free stream region for the rotating case shows an increase in velocity with respect to the tunnel speed setting of 20mph. The static cases of 0 and 45 show substantial asymmetry, which is interesting as the only asymmetry in the model arises from the pitch link joint with the blade shanks. The wake shows interesting features. One is the apparent convection of the low-speed zone (presumably due to separation and recirculation) that occurs downstream of the shank: this convects in a periodic manner and appears in phase-resolved, ensemble averaged contours of the flow speed in the horizontal 33

34 (a) Top view (b) Isometric View Figure 23: Map of PIV data collection plane comprised of overlapping stitches 34

35 Figure 24: Wake tunnel axis velocity profiles captured by PIV and scaled by free stream velocity of 20mph 35

36 Figure 25: PIV result ihowing a color map of the speed, and velocity vectors, in the horizontal (X-Y) plane downstream of the hub, at a hub phase of 30 degrees plane. This is illustrated in Figures 25, 26 and 27, taken at 30, 45 and 60 degrees phase with respect to the zero reference of azimuth. The position coordinates are referred to the PIV measurement locations shown in Figure 23. The primary axis of vorticity in the flowfield downstream of a shank is expected to be parallel to the shank axis. The plane being viewed here is probably parallel to the primary axis of the vorticity vector field, and hence vorticity contours in this plane do not produce useful visualization. However, it is interesting to note the large discrete region of low speed. As this convects, it may be expected to cause severe fluctuations in the speed being sensed by hot-film anemometer probes, addressed in the next section. The vorticity in the wake flowfield was studied using the velocity field shown in Figure 28. The Normalized Angular Momentum (NAM) algorithm was used to capture centers of vortical structures. A test case of the NAM is shown in Figure 29. A numerical test case velocity field was generated, with a few discrete vortices. This is shown in the left side of Figure 29. The vorticity field computed by the NAM is shown on the right of Figure 29. The technique was applied to the portion of the hub wake shown in Figure 23. The result is shown in Figure 30. The plot on the left side shows the vorticity field computed directly, while the right side shows the vorticity field computed using the NAM. 36

37 Figure 26: PIV result ihowing a color map of the speed, and velocity vectors, in the horizontal (X-Y) plane downstream of the hub, at a hub phase of 45 degrees Figure 27: PIV result showing a color map of the speed, and velocity vectors, in the horizontal (X-Y) plane downstream of the hub, at a hub phase of 60 degrees 37

38 Figure 28: PIV field showing Figure 29: Validation of the Normalized Angular Momentum algorithm to capture vortical structures in a velocity field. 38

39 Figure 30: Vortical structures in the indicated portion of the hub wake, computed directly as vorticity, and using the Normalized Angular Momentum algorithm. WAKE VELOCITY FLUCTUATION MEASUREMENTS A hot-film constant temperature anemometer probe was used to measure flow fluctuations in the wake at locations listed in Table 3. A hot-film sensor uses the principle of a hot-wire anemometer sensor. Feedback-controlled electrical resistance heating is used to set the sensor external temperature at a constant level in the range of 520K, well above the ambient flow temperature. The voltage needed to maintain this setting through a Wheatstone Bridge circuit, is empirically related to the flow speed over the sensor. In the case of the hot film, the active resistance is that of a thin film of metal coated over a strong non-conducting substrate rod, instead of using an extremely thin metal wire. Thus the sensor has greater mechanical ruggedness, and is appropriate where the micron-sized sensor wires are not needed. The anemometer voltages were digitized and sampled at 5000Hz, for 30 seconds at each station. A filtered signal was simultaneously recorded, with a high pass filter set at 3Hz and a low pass filter set at 2000Hz. An amplifier gain, typically 20dB, was used to optimize the amplitude of the fluctuating signal for digitization. Using these voltages the instantaneous flow speed could be calculated using the unfiltered voltage, while the small velocity fluctuations could be amplified for good resolution, and calculated from the filtered voltages. The full-signal speed calculation uses the nonlinear voltage-speed 39

40 relationship obtained from calibration of the anemometer probe against a Pitot-static probe. The conversion of the filtered signal uses the instantaneous slope of the voltage-speed calibration. This method still poses substantial issues when the fluctuation amplitude approaches the mean value of the speed. In regions where the flow may encounter transient stagnation or reversal, the use of such probes is extremely difficult, and prone to large error. Axis Locations Resolution X 1D h to 3D h 1D h Y D h to D h D h Z D h, 0, D h and D h Table 3: Summary of wake velocity measurement locations The full database of hot-film anemometer measurements is given on the Internet at the project website given on the front cover of this report. NOTE: The gain applied to the signal has been accounted (removed) for in the data presented. The data are presented in two columns, the first column is the mean velocity and the second column is the fluctuation velocity. Each column consists of 5000 samples/s 30s = 150,000 samples. The data were first collected as voltages and then converted to velocity using the calibration coefficients, and the data set is presented in m/s. Please refer to Table 3 for the specific file names corresponding to each measurement location. For example: X = 1D, Y = 1, Z = 1 corresponds to X = 1 D h, Y = D h, Z = D h. Time-averaged wake profiles obtained using the hot-film probe are given below. These are not expected to be the same as the velocity profiles obtained using PIV because the hot-film anemometer is sensitive to all components of velocity, although with dominant sensitivity to the components that are perpendicular to the axis of the cylindrical sensor element, whether wire or film substrate. Figures 31 and 32 compare the effect of having scissors present, on the time-averaged velocity profile 1 diameter downstream, at 30 mph freestream speed. Figure 33 presents the wake profile at the same location (1 diameter downstream) at 20 mph, and Figures 34 and 35 show what happens at 2 and 3 diameters downstream, respectively. The expected spreading of the wake is seen. Curiously, the case with no scissors has more features in the wake profile. This agrees with the azimuth-resolved load measurements, which show that the presence of the scissors causes the dips in the drag variation to be be filled in, reducing the frequency content of the azimuth-resolved loads. The frequency content at the above locations is next examined using spectra obtained from the hot-film data. The specific wake location examined below is X=1D h and Z=0 at Ω=120 and U =8.9 m/s from 40

41 Figure 31: Time -averaged speed profile with no scissors, at 1 diameter downstream. Figure 32: Time -averaged speed profile with scissors, at 1 diameter downstream. 41

42 Figure 33: Time -averaged speed profile with scissors, at 1 diameter downstream. Figure 34: Time -averaged speed profile with scissors, at 2 diameters downstream. 42

43 Figure 35: Time -averaged speed profile with scissors, at 2 diameters downstream. Y= 0.5D h to 0.5D h. Figure 42 illustrates the FFT power spectral density of the velocity fluctuations across the span of the hub wake. The power spectral density data presented are for the energy content in the wake at 8Hz, 4/rev. The 4/rev content shows a clear asymmetry between the advancing and retreating sides of the rotor hub. This was also observed in the investigation by Roesch and Dequin (1985). The effect of the scissors components added to the hub for the final round of experiments, can be seen from the following figures. Figure 36 shows the spectrum at X=1D (1 diameter downstream of hub centerline, Z=2 denoting the measurement location where Z is hub diameters as shown in Table 3, with 240 rpm rotation speed and freestream speed of 30 mph (13.41m/s). The frequency axis is expressed in units of per revolution. Thus the major spectral content is at 8 per rev, with large peaks downstream of the advancing side. When the scissors are added, the spectra change as seen in Figure 37. The fluctuation intensity at 8 per rev rises dramatically further inboard on the advancing side, but drops at outboard locations. The next set of sample figures trace what happens at a given Z location (Z0 as given in Table 3 as we move downstream, with the scissors present. The first, Figure 38, shows the spectra at Z0 and X=1D, which is directly above the location shown in the prior two figures, but the data were acquired at a lower 43

44 Figure 36: Spectral features in the hot-film anemometer data obtained across the wake behind the hub with no scissors, at 1 diameter downstream. Figure 37: Spectral features in the hot-film anemometer data obtained across the wake behind the hub with scissors added, at 1 diameter downstream. 44

45 Figure 38: Spectral features in the hot-film anemometer data obtained across the wake behind the hub with scissors added, at 1 diameter downstream, and Z=0, at 20mph. speed of 20mph (8.94m/s). At this point the 8 per rev peak occurs inboard, on the retreating blade side. Going downstream to 2 diameters (X=2D), we see from Figure 39 that the 8 per rev fluctuation intensity is spread across the entire retreating side. Figure 40 shows that a X=3D there is much more broadband content as the wake gets more chaotic. In an attempt to better comprehend the temporal nature of the hub wake, the wavelet transform was applied to the same wake turbulence data set. Table 4 shows some of the differences between the Wavelet and FFT approaches. The purpose of the wavelet transform was to find discrete fluctuations occurring in the wake Farge (1992). The wavelet used in the transformation was the complex Morlet, visualized in Fig. 41, which has harmonic parts, useful when comparing results to FFT, and is often used to analyze turbulent flow using constant temperature anemometry Hoa (2009). The wavelet transform was carried out using the wavelet toolbox built into Matlab. The signal is easily loaded in the interface into the 1-dimensional continuous wavelet feature. From there the wavelet type, complex Morlet in this instance, and the number of scales, 600 were used for the analysis here, and then the wavelet transformation was performed and generated the wavelet coefficients at each integer scale. The Wavelet transform coefficients are compared to the FFT of the signal, at the Y= D h, in Figure 43 for 2 seconds of a 30 second signal. The color legend indicates the correlation of the wavelet 45

46 Figure 39: Spectral features in the hot-film anemometer data obtained across the wake behind the hub with scissors added, at 2 diameters downstream, and Z=0, at 20mph. Figure 40: Spectral features in the hot-film anemometer data obtained across the wake behind the hub with scissors added, at 3 diameters downstream, and Z=0, at 20mph. 46

47 Figure 41: Real and Imaginary Components of Complex Morlet Wavelet Feature FFT Wavelet Transform Harmonic Signals High Quality Low Quality Time Resolved Frequency content Ensemble Averaged Frequency Time Resolved Content Table 4: Fourier and Wavelet Transform Comparison coefficients to the signal. The wavelet transformation shows the highest correlation with the complex Morlet at, or near, the low frequency peaks in the FFT. As the frequency increases the power spectral density computed from the Fourier Transforms decreases, at the same time the correlation with the complex Morlet also decreases significantly. From the wavelet transform it can be observed that the high correlation frequencies line up well with the peak frequencies seen in the FFT. The largest peaks in frequencies occurred at 4/rev and 8/rev, corresponding to the passage of the 4 shanks of the hub. There are also peaks offset from these, at 6/rev and 12/rev, due to parts of the hub passing by twice per rotation. It is conjectured that the two bolts connecting the hub plates could cause the 6/rev and 12/rev energy content in the hub wake. Once the wavelet transform is performed the scales of the wavelet must be related to a domain that is easier to understand. Because the comparison to the Fourier transform is presented here the scale is converted to the Fourier frequency( f f ). Equation 8 relates the scale of the wavelet transformation to f f. f f 1 s (8) 47

48 UNCERTAINTY ESTIMATES The main errors in constant temperature anemometry measurements arise due to the following reasons (adapted from Yavuzkurt (1984)): 1. Calibration error, ε C - caused due to inaccuracies during calibration measurements, curve fitting and calibration drift error. 2. Approximation error, ε A - Error caused by approximation techniques, mainly caused by standard deviation deviation. 3. Temperature variation, ε T - Error due to temperature variations in the flow. It can change air density slightly and affect the heat transfer between the flow and the wire. 4. Pressure uncertainty, ε P - Uncertainty due to not knowing ambient pressure exactly. Pressure uncertainty causes an uncertainty on assumed air density. The total error (ε) is then calculated by ε = k ε 2 C + ε 2 A + ε 2 T + ε 2 P, where k 2 for a confidence interval of 95%. Error estimates for wake velocity measurement are summarized in Table 5, along with uncertainties in all the parameters. Parameter Worst case estimate Dynamic Pressure 0.02% Dynamic Viscosity 0.01% Density 0.03% Velocity 0.03% Load Cell Calibration 1.37% Calibration 0.77% Temperature variation 0.45% Approximation 0.43% Ambient pressure 0.03% Total ±1.98% Table 5: Summary of experimental uncertainties 48

49 Figure 42: 4/Rev Energy content in the hub wake at X = 1D and Z = 0 SUMMARY OF DATA AND RESULTS Air loads data are given on the Internet at the project url given on the front cover of this report. The data organization is summarized in Table 6. The airloads data are published with a README text file. All force measuremements are reported in lbs, and all moment measurements are reported in lb-in. The loads data are in 4 categories of test cases: 1. General Tests. These give loads at specified rotation rates (RPM) and freestream speed (Meters per second). 2. No Scissor Tests. These are tests done with no scisssors and no motor fairing, taken at a digitization rate of 10,000 Hz. 3. Azimuth-resolved. 4. Fairing + Tare. Load measurements on the motor with a fairing on the motor. The report ADLR pdf summarizes the hot-wire anemometer data. Wake velocity measurements were conducted at locations summarized in Table 7 and the axis are illustrated in Fig. 45. The locations are normalized with hub diameter D h. The measurements were made across the span of the hub wake (Y axis) at a resolution of mm ( D h ). Specific Z planes were chosen in 49

50 Figure 43: Wavelet transform coefficients compared to FFT at Y= D h 50

51 Figure 44: Downstream hot-wire measurement locations order to understand the overall nature of the unsteady wake and in specific to investigate the effect of shanks and the scissors on the unsteady rotor hub wake. The measurement planes where D h Z D h were chosen in order to investigate the flow over the main hub assembly, whereas the Z = 0 204D h measurement plane is to investigate the effect of scissors and pitch links. The X locations (1D h X 3D h ) were chosen to investigate the evolution of the wake on moving downstream. The downstream locations are also referred to as near wake (1D h ), mid wake (2D h ) and far wake (3D h ). Some significant results are reiterated in the figures below. Figure 46 shows azimuth-resolved airloads data, with drag, side force and torque shown. Drag is the dominant airload, but there are measurable azimuthal variations in side force and torque. This result is not expected to change in scaling to full-scale helicopter cases; in fact drag is expected to dominate in a much stronger way. This aspect is studied in Figure 47. Figure 48 shows the azimuth-resolved drag, as well as the variation of drag with freestream speed. There is about a 10 percent drag increase due to these two elements. The azimuth-resolved drag shows that the effect of the scissors elements (which may be unique to how we placed them on the hub) is to fill in azimuthal intervals where the drag was otherwise lower. Note that the widely differing magnitudes of the loads is accommodated by using different vertical scales. Power demand due to side force and torque are referred to the right side axis. 51

52 Folder Name RPM Velocities(mph) Azimuth Resolved , 30, 50 Fairing + Tare 0 10, 15, 20, 25, 30, 35, 40, 45, 50 General Tests 80 10, 15, 20, 25, 30, 35, 40, 45, , 15, 20, 25, 30, 35, 40, 45, , 15, 20, 25, 30, 35, 40, 45, , 15, 20, 25, 30, 35, 40, 45, 50 No Scissors Tests , 25, 30, 35, 40, 45, 50, 55 With Scissors Comparison , 30, 35, 40 Table 6: Rotor Hub Aerodynamic Load Test Cases Figure 45: Illustration of axis referred to in the measurement locations Figure 49 shows an attempt to gauge the significance of interaction effects. The theoretical calculations are in fact empirical, using methods such as those given in Hoerner (1965). The actual ones are obtained from the drag deconstruction tests. The shaft in this case includes the swashplate. The reduced drag on the shanks when installed, is real: it is attributed to the shank wakes reducing the onset dynamic pressure and perhaps redirecting the flow that would otherwise cause large flow separation at the base. The next result is that there is substantial 4 per rev fluctuation content in the wake on the retreating blade side of the hub. This is indicated in Figure 50 where the power spectral density at 4 per rev (units are arbitrary) is shown, as the probe location was moved across the flow from the retreating blade side 52

53 Figure 46: Drag is the dominant airload on the hub Figure 47: Implications for the total power demand on a helicopter 53

54 Figure 48: Effect of the scissors elements on azimuth-resolved drag Figure 49: Significance of interactions between component flowfields, as gauged from simple predictions of component drag 54