THE EFFECT OF PYLON TRAILING EDGE BLOWING ON THEPERFORMANCE AND NOISE PRODUCTION OF A

PROPELLER

L.L.M. Veldhuis∗ , T. Sinnige∗∗Delft University of Technology

Keywords: propeller flow, noise, pylon interaction, trailing edge blowing, flow control

Abstract

This paper discusses a study of the effects of py-lon trailing edge blowing on pusher propeller per-formance and noise emissions. Experimental in-vestigations were performed in a low speed openjet wind tunnel, using a powered propeller modeland a generic pylon model. A novel pylon blow-ing system was integrated in the aft part of thepylon to provide a uniform outflow. The numer-ical analysis was executed using a combinationof a propeller model that is based on an extendedBlade Element Model (BEM) and a set of ana-lytic methods from the literature, to assess the ef-fects of pylon installation on the propeller per-formance and noise emissions. Measurementsof the velocity distributions in the blown py-lon wake showed that application of the blow-ing system reduced the integrated velocity deficitby up to 60% compared to the unblown config-uration. Evaluations of the propeller forces andmoments showed that the effects of installationon the time-averaged propeller performance aresmall, with differences of at most 2% for ad-vance ratios below J=1.4. Furthermore, excellentagreement was obtained between the computedand measured performance for the isolated pro-peller. With respect to the propeller noise emis-sions it was observed that installation of the py-lon upstream of the propeller strongly increasesthe noise levels. Depending on the propelleroperating point, noise penalties of up to 15dBwere measured. The application of pylon blow-ing clearly reduced the propeller noise emissions

over the entire advance ratio range, with reduc-tions of up to 7dB compared to the unblown case.The largest noise reductions were found at thehighest blowing rate, indicating that the most ef-fective blowing rate might not have been reachedyet.

1 Introduction

The environmental impact of aircraft operationsand increasing fuel prices have led to the demandfor more fuel-efficient aircraft. One of the tech-nologies with the potential to offer a significantreduction in fuel burn is the open rotor engine,which allows the bypass ratio to be increased tovalues unattainable by turbofans. This results ina step change in propulsive efficiency, with esti-mations for the corresponding reduction in fuelburn at around 25−30% [11, 17, 26]. The majordisadvantage of the open rotor engine is its asso-ciated high level of noise emissions. In case of acontra rotating open rotor (CROR) configuration,tonal noise is generated by each rotor individu-ally as well as by the aerodynamic interactionsbetween the two rotors. In addition to isolatednoise sources, installation of the open rotor onthe aircraft introduces another noise generatingmechanism. For interior noise and ground clear-ance reasons, a pylon-mounted placement at theaft end of the fuselage, with a pusher configura-tion is considered in most of the studies presentedin the literature. In this case the wake shed fromthe upstream pylon impinges on the front rotor,resulting in unsteady blade loading and associ-ated noise emissions [2, 3, 10, 12, 17, 19, 23].

1

VELDHUIS, SINNIGE

The pylon - open rotor interaction mainly affectsthe noise levels associated with the front rotortones, while the rear rotor tones show only asmall effect and the interaction tones remain un-affected [18, 19, 23, 25]. Furthermore, apart fromthe impact on the propeller noise emissions, thefluctuating blade loads may also affect the pro-peller performance.

2 Pylon Trailing Edge Blowing

The pylon installation effects originate from theimpingement of the pylon wake on the propeller.The resulting non-uniform propeller inflow leadsto unsteady blade loading with associated per-formance and noise emission penalties. Basedon the interaction mechanism it can be expectedthat the installation effects can be reduced oreven completely removed by eliminating the py-lon wake. Ricouard et al. [19] show that espe-cially the active technique of pylon blowing canbe very effective in reducing the pylon - propellerinteraction effects.

The concept of pylon wake eliminationthrough pylon blowing is illustrated in Fig.1,which presents an idealized example in which thepylon wake is filled up completely.

The potential of pylon trailing edge blowingto reduce the adverse installation effects experi-enced on rear-fuselage mounted open rotor en-gines forms the main topic of this paper. Bothexperimental and numerical investigations wereperformed. It should be noted that the availableexperimental equipment limited the current studyto single-rotating propeller applications, in con-

+ =

Pylon wake + Pylon blowing = Filled pylon wake

Fig. 1 Schematic illustrating the concept of py-lon wake elimination through pylon blowing.

trast to the contra-rotating technology typicallyprojected for future applications of open rotor en-gines on next-generation passenger aircraft.

3 Experimental Setup

The experiments were performed in a low speedopen jet wind tunnel, using a powered propellermodel and a typical pylon model. A novel ”Uni-form Blowing Rod” (UBR) was integrated in theaft part of the pylon. A number of measurementtechniques were applied to measure the pylonwake profiles, propeller performance, and pro-peller noise emissions.

3.1 Wind Tunnel Facility

All experiments were performed in Delft Uni-versity of Technology’s Open Jet Facility (OJF)[6]. This closed circuit wind tunnel with opentest section has an octagonal outlet with widthand height equal to 2.85m and provides a maxi-mum wind velocity of 32m/s. To remove spatialvelocity deviations and to reduce the flow’s tur-bulence level, the settling chamber is equippedwith a honeycomb flow rectifier and five screens.This results in velocity deviations smaller than0.5% in the vertical plane at 2m from the out-let, with a longitudinal turbulence intensity levelaround 0.24%. To reduce noise levels, the insideof the entire tunnel is covered with perforatedplates installed on mineral wood and sound ab-sorbing foam. The fan however does not featureany special noise reduction measures.

3.2 Wind Tunnel Models

The wind tunnel tests involved the use of twoscale models: a powered propeller and a pylon.To simulate a pusher configuration the tractorpropeller model was placed behind the pylon.

3.2.1 Propeller Model

The powered single-rotating propeller scalemodel was made by Fokker Aircraft Companyduring the development of the conceptual FokkerF29. The model has a diameter of Dprop =0.3043 m, a hub diameter of Dhub = 0.084 m,and is equipped with up to eight blades. In the

2

EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

450

2822

Left view

114

Isometric view

Fig. 2 Overview of the installed propeller setup(dimensions in millimeters).

current study the eight-bladed configuration wasselected. The blade angle at 75% of the bladespan can be adjusted as desired, and was set toβ0.75R = 41◦. The propeller is driven by a TechDevelopment (TDI) 1999A pneumatic motor. Afour-component Rotating Shaft Balance (RSB),capable of separately measuring thrust, torqueand in-plane components, is integrated in the pro-peller model. The entire propeller test rig is uti-lized and maintained by Delft University of Tech-nology and it is obtained from the German-DutchWind Tunnels (DNW) on a lease basis.

The propeller model was positioned at 30%of the pylon chord behind the pylon during themeasurements performed in the installed andblown configurations. Note that this spacing is atthe high end of the spectrum of pylon - propellerspacings considered during comparable pylon -propeller interaction studies available in the liter-ature [2, 3, 4, 13, 23, 25, 28]. However, takinginto account the geometry of the powered pro-peller model it was not possible to position thepropeller any closer to the pylon. A technicaldrawing of the complete setup of the propeller inthe installed configuration is presented in Fig. 2.Photographs of the propeller test setup in the OJFwind tunnel are shown in Fig. 3.

3.2.2 Pylon Model

The pylon was designed based on typical pyloncharacteristics used in comparable pylon - pro-peller interaction studies [2, 3, 4, 13, 23, 25, 28]and taking into account the minimum dimensionsrequired for a successful integration of the blow-ing system in the trailing edge region of the py-lon. As a result, the pylon had a straight, unta-

Pylon

Propeller

U∞

Motor nacelle

Side view

Eight-bladed propeller in close-up

Fig. 3 Propeller-Pylon setup in the Delft Univer-sity OJF windtunnel.

pered planform with the cross-section formed bya NACA 0012 profile modified to have a trail-ing edge thickness of 1.1% of the chord. Thepylon chord length was equal to 1.33 times thepropeller diameter, while the span of the pylonwas selected based on the available space in thetest setup, resulting in a span of bpyl = 0.450 m.Transition was fixed at 25% of the chord at bothsides of the pylon.

An overview of the pylon characteristics isgiven in Table 1; a technical drawing of the pylonmodel is depicted in Fig. 4.

3.3 Uniform Blowing Rod

The core of the pylon blowing system is formedby a Uniform Blowing Rod (UBR). Developed

3

VELDHUIS, SINNIGE

Table 1 Pylon model characteristics.Parameter Symbol Value

Chord cpyl 1.33/DpropSpan bpyl 1.48/Dprop

TE thickness tTE 0.011/cSweep angle Λpyl 0◦

Taper ratio λpyl 1.0Airfoil - NACA 0012 (mod)

in conjunction with FlowMotion1, the UBR isdesigned to provide a uniform outflow from itsoutlet. The UBR basically consists of two com-ponents: an interior air channel with a vari-able cross-sectional shape along the span, and anoutlet channel with a constant cross-section andvanes placed at constant spacing to align the flowwhich exits the UBR.

The UBR design is characterized by the ge-ometry of the initial cross-section (at the inletplane, i.e. the UBR’s root), after which the shapeof the cross-sections of the interior air channel inspanwise direction towards the tip are computedsuch that in theory a uniform outflow profile isobtained. This is done by taking into account theboundary layer development along the UBR’s in-terior air channel in determining the optimal localcross-sections. A simple rectangle was chosen ascross-section at the root of the UBR. Vanes areplaced in the outlet segment to align the flow be-fore it exits the UBR. The final shape of the vaneswas determined based on analyses of the UBR’soutflow profiles performed with ANSYS Fluentusing steady-state RANS simulations of the flowinside the UBR. A technical drawing of the finalUBR design, as integrated in the aft part of thepylon, is presented in Fig. 5.

3.4 Measurement Techniques

3.4.1 Pressure measurements

The wake profiles behind the unblown and blownpylons were measured by traversing a total and astatic pressure tube over a range of lateral posi-tions behind the pylon. During all measurements

1FlowMotion - Consultancy for Heat Transfer and FluidDynamics: www.flowmotion.nl

173

277

450

2 67 139

406

Right view

4.5

Top view

Fig. 4 Geometry of the pylon model (dimensionsin millimeters).

277

266

1 845

11420

Right view

6

1Rear view

4.5

6

Bottom view

Fig. 5 Uniform Blowing Rod (UBR) geometry(dimensions in millimeters).

4

EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

the pylon was set to zero degrees angle of at-tack. The tubes were installed on a computer con-trolled traversing system which could translate inall three directions.

3.4.2 Noise measurements

Two LinearX M51 high performance low volt-age electret condenser microphones were usedto measure the acoustic pressures induced dur-ing the experiments. Both microphones were po-sitioned outside the open jet stream, at a side-line distance of 2.25 m from the centerline of thepropeller. The frequency response of the micro-phones is practically flat in the 50-2000 Hz range;for the response at lower and higher frequenciescalibrated correction factors provided by LinearXwere used to correct the microphone data. Fur-thermore, a correction was applied to account forrefraction effects due to the presence of the shearlayer following the method outlined in Ref. [16].The microphone measurements were performedat the same sampling frequency as used for theRSB measurements: 50 kHz. To convert the rawmicrophone signals into sound pressure levels themicrophones were calibrated every measurementday using a pistonphone.

3.4.3 Particle image velocimetry

A stereoscopic PIV system was used to measurethe velocity fields around the propeller blades atseveral radial stations. The laser light was pro-vided by a Qantel EverGreen CFR200 Nd:Yaglaser with a pulse energy of 200 mJ/pulse. Tracerparticles were introduced downstream of thewind tunnel test section with a SAFEX Twin Foggenerator using a SAFEX inside nebelfluide mix-ture of diethylene glycol and water. Two LaVi-sion Imager Pro LX 16 Mpix (4,872 x 3,248 pix-els) cameras were used with Nikon 105mm f/2.8AF Micro lenses (set to f/5.6-8) to record theimage pairs. Data sets consisting of 200 imagepairs were recorded phase-locked at a measure-ment frequency of about 0.5 Hz.

A photograph depicting an overview of thePIV setup is presented in Fig. 6.

Support table

PIV window

Laser headU∞

Propeller

Fig. 6 Overview of PIV setup.

4 Numerical Setup

A numerical scheme was developed to computethe effects of pylon installation on the perfor-mance and noise emissions of pusher propellers,following the schematic flowchart depicted inFig. 7.

4.1 Pylon Wake Profiles

To compute the effects of installation of the py-lon on the propeller performance and noise emis-sions the pylon wake profiles need to be deter-mined first. For this purpose the Schlichtingwake model was used, which is characterizedby the following two governing equations (Ref.[20]):

∆uU∞

(Xw,Yw) =

√10

18β

√cdcXw

[1−∣∣∣∣Yw

b

∣∣∣∣ 32]2

(1)

bw (Xw) = β√

10cdcXw (2)

with bw the wake semi-width, c the chordlength, cd the pylon 2D drag coefficient, ∆u thelocal velocity deficit in the pylon wake, U∞ thefreestream velocity, Xw and Yw longitudinal andlateral coordinates measured from the center ofthe pylon’s trailing edge, and β an empirical con-stant equal to β = 0.18 [20]. The pylon drag co-efficient was computed using XFOIL [7].

Note that the blown pylon wake profiles can-not be computed using the developed numericaltool. Instead, for the computations related to the

5

VELDHUIS, SINNIGE

Propellergeometry

Installedperformance(Sears)

Isolatedperformance(XROTOR)

Isolated Noise(Hanson)

Installed Noise(Hanson)

Pylongeometry

Pylon wake profileSchlichting

2D analysispylon airfoil(XFOIL/RFOIL)

Fig. 7 Flowchart of the numerical scheme de-veloped to analyze pusher propeller performanceand noise emissions.

blown configuration wake profiles were used asobtained from the experimental evaluations.

4.2 Propeller Performance

The numerical assessment of the propeller per-formance is built around the propeller analysisand design program XROTOR [8]. The instal-lation effects are accounted for by correcting theisolated blade response for the change in the dy-namic pressure and the angle of attack in the py-lon wake.

4.2.1 Isolated Configuration

The isolated (steady-state) propeller performancewas computed using the propeller lifting lineprogram XROTOR. The propeller blade sectioncharacteristics which are required as inputs toXROTOR were determined using RFOIL [27].To correct for the effects of rotation on the bladesection characteristics the empirical model devel-oped by Snel et al. was used (Ref. [24]):

crotl = cl + tanh

{A(c

r

)B}(

cl− cllin

)(3)

with A and B tuning parameters, set to their de-fault values of A= 3 and B= 2, respectively [24].The drag coefficient was not corrected for rota-tional effects.

Computation of non-uniform propeller inflow

Computation of ∆cl and∆cd due to ∆α

Computation of ∆cl and∆cd due to ∆q

Computation of total un-steady cl and cd

Computation of installedcl and cd

Isolated cl, cd

Computation of installedpropeller performance

Fig. 8 Flowchart of the installed propeller per-formance computation routine.

4.2.2 Installed Configuration

Following the isolated performance computa-tions the installation effects for a single-rotatingpusher propeller are considered. In the installedconfiguration the inflow at the propeller disk ischaracterized by a non-uniform velocity field dueto the velocity deficit in the pylon wake. Theassumption is made that the final installed bladeloading can be computed by following the prin-ciple of superposition, hence the effects of thechanges in the dynamic pressure and the angle ofattack are considered separately. A flowchart ofthe installed propeller performance computationroutine is presented in Fig. 8.

The change in blade loads due to the varyingdynamic pressure is evaluated at the local anglesof attack computed for the isolated case. The as-sumption is made that the isolated lift and dragcoefficients per radial station are constant overthe full rotation, hence Reynolds number effectsare neglected. Furthermore, it is assumed that theinduced velocities corresponding to the steady-state solution apply at each polar angle φ withoutmodification. The changes in the lift and dragcoefficients due to the variation in dynamic pres-sure at constant angle of attack ∆c∆q

linstand ∆c∆q

dinstare computed using:

6

EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

∆c∆qlinst

(η ,φ) = cSSl (η)

{W 2

inst (η ,φ)

W 2iso (η)

−1}(4)

∆c∆qdinst

(η ,φ) = cSSd (η)

{W 2

inst (η ,φ)

W 2iso (η)

−1}(5)

with cSSl and cSS

d the steady-state (isolated)lift and drag coefficients, Wiso the undisturbed ef-fective velocity, Winst the local effective velocityin installed conditions, η the non-dimensional ra-dial coordinate, and φ the polar angle.

The change in blade loads due to the variationof the angle of attack in the pylon wake regionis computed using Sears’ method, which is de-scribed in Refs. [21, 22]. Only the lift coefficientis considered; the effects on the drag coefficientare neglected. The computation of the unsteadyblade response is performed in the frequency do-main. Therefore, the velocity deficit in the pylonwake is rewritten as a periodic gust in a directionnormal to the blade sections’ upper surfaces andthen expressed as a complex Fourier series.

The harmonics of the unsteady blade lift co-efficient due to the change in angle of attack inthe pylon wake region ∆c∆α

linstkare computed using

the Sears function [5, 9]:

∆c∆αlinstk

(η) = 2πvgnk

(η)

Wiso (η)S (6)

with vgnkthe Fourier coefficients of the normal

gust velocity and S the Sears function as definedin Refs. [22, 21]. A (low-frequency) compress-ibility correction is applied to the result obtainedfrom to the incompressible Sears function [1].

Having computed the harmonics of the un-steady lift coefficient, the local unsteady lift co-efficients are obtained as a function of the blade’spolar angle φ by taking the inverse Fourier trans-form of the harmonics.

The changes in the lift and drag coefficientsdue to the effects of the reduced dynamic pres-sure and the increased angle of attack in the py-lon wake region are superimposed to obtain thefinal unsteady propeller blade loads cUS

l and cUSd .

With the unsteady blade loads known, the result-ing unsteady thrust and torque are obtained by in-tegrating the contributions of all blade elements:

T US1B (φ) =

Nr

∑i=1

12

ρW 2i (c

USli (φ)cosϕi−

cUSdi

(φ)sinϕi)ci∆ηiR (7)

QUS1B (φ) =

Nr

∑i=1

12

ρW 2i (c

USli (φ)sinϕi+

cUSdi

(φ)cosϕi)ciηi∆ηiR2 (8)

with ϕi the advance angle (including induced ef-fects) of blade segment i. Note that the assump-tion is made that the additional lift and drag re-sulting from the installation effects act perpen-dicular and parallel to the local effective velocityincluding induced effects.

Having computed the unsteady thrust andtorque for a single-bladed propeller using Eqs. 7and 8, the results are generalized to a B-bladedpropeller by taking into account the proper phaseshifts between the various blades. Finally, the un-steady thrust and torque are added to the steady-state (isolated) results to obtain the propeller per-formance in the installed configuration.

4.3 Propeller Noise Emissions

The propeller noise emissions are computedusing the methods developed by Hanson [14,15]. The isolated propeller is analyzed with thescheme discussed in detail in Ref. [14]. Forthe installed configuration the method for contra-rotating propellers discussed in Ref. [15] isadopted. The current case with a fixed inlet dis-tortion due to the presence of the pylon followsfrom the contra-rotating problem by assuming animaginary front rotor with zero angular velocityand unity blade number. In all cases the assump-tion of a uniform lift distribution was made, whilethe blade loads obtained from the isolated pro-peller performance analysis were corrected forthe shift due to the induced angle. Broadbandnoise emissions were neglected in all analyses.

5 Results

The experimental and numerical methods wereused to analyze the UBR’s outflow velocity pro-

7

VELDHUIS, SINNIGE

Q = 680 L/min

Q = 600 L/min

Outflow

Velocity

Ue[m

/s]

Spanwise Coordinate Z [mm] (root to tip)0 20 40 60 80 100 120 140 160 180 200 220 240 260

0

2

4

6

8

10

12

14

16

18

Fig. 9 Uniform Blowing Rod outflow profiles fortwo blowing rates; Xw = 0.1cpyl.

Q=680 L/min

Q=600 L/min

Q=500 L/min

Q=400 L/min

Q=0 L/min

Non-D

imen

sionalVelocity

U/U

∞[-]

Lateral Position Yw [mm]−40 −30 −20 −10 0 10 20 30 40 50

0.60

0.70

0.80

0.90

1.00

1.10

1.20

Fig. 10 Pylon wake velocity profiles for variousblowing rates; U∞ = 19m/s, Xw = 0.3cpyl.

files, the pylon wake profiles, the propeller per-formance, and the propeller noise emissions.

From Fig. 9 it is observed that the UBR’s out-flow is not as uniform in spanwise direction asdesired. This is likely the result of non-uniforminflow due to flow separation in the divergent in-let of the UBR.

5.1 Pylon Wake Profiles

The velocity profiles measured in the pylon wakeat a freestream velocity of U∞ = 19 m/s are pre-sented in Fig. 10. The wake profiles were mea-sured at the position of the propeller plane usedin the evaluations of the installed and blown pro-peller performance and noise emissions (corre-sponding to a longitudinal wake-based coordi-nate of Xw = 0.3cpyl). The pylon blowing sys-tem was operated for a number of blowing rates Qranging from zero (unblown) up to the maximumblowing rate of 680 L/min. Note that the wakeprofiles were obtained for the isolated pylon, i.e.the presence of the propeller was neglected.

The non-dimensional wake velocity profilesdisplayed in Fig. 10 show a continuously in-creasing reduction in wake depth with increas-ing blowing rate. At blowing rates of 500 and

U∞=26 m/s, Q=680 L/min

U∞=19 m/s, Q=680 L/min

U∞=26 m/s, Q=0 L/min

U∞=19 m/s, Q=0 L/min

Non-D

imen

sionalVelocity

U/U

∞[-]

Lateral Position Yw [mm]−40 −30 −20 −10 0 10 20 30 40 50

0.60

0.70

0.80

0.90

1.00

1.10

1.20

Fig. 11 Unblown and blown pylon wake velocityprofiles for different freestream velocities; Xw =0.3cpyl.

600 L/min a reduction in the integrated velocitydeficit of around 60% is obtained when comparedto the unblown configuration. The application ofblowing from the extended pylon’s trailing edgedoes not have an appreciable effect on the totalwake width. This indicates that the jet blown intothe flow from the UBR’s outlet does not fully mixwith the external flow before reaching the axialposition of Xw = 0.3cpyl. As a result, the veloc-ity profiles measured at this position are not uni-form but instead display a profile with one localmaximum on the wake’s centerline and two localminima left and right of the centerline.

The influence of the freestream velocity onthe amount of wake fill-up achieved using the py-lon blowing system is shown in Fig. 11, whichcontains data measured at 19 m/s and 26 m/s.

Figure 11 shows that the amount of wakefill-up obtained by applying the blowing sys-tem decreases with increasing freestream veloc-ity. At a given blowing rate the increase inthe non-dimensional velocity on the wake cen-terline is smaller when the freestream velocity ishigher. This is directly related to the ratio of thefreestream flow velocity and the velocity of theflow blown into the pylon wake. The latter needsto be larger than the freestream velocity for theblowing system to be effective.

5.2 Propeller Performance

The propeller performance was analyzed to as-sess the effects of installation and pylon blow-ing on pusher propeller performance. The perfor-mance was expressed in terms of the thrust coef-ficient CT , the torque coefficient CQ, and the pro-

8

EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

NUM rot. cor.NUMEXP

CQ

CT

η

Pro

pel

ler

Per

form

ance

CT,C

Q,η

[-]

Advance Ratio J [-]

0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.00.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

Fig. 12 Experimental and numerical propellerperformance diagrams; isolated configuration,U∞ = 26m/s, showing the effect of rotation cor-rection according to eq. (3).

peller efficiency η , defined as:

CT =T

ρn2D4 (9)

CQ =Q

ρn2D5 (10)

η =J

2π

CT

CQ(11)

with D the propeller diameter, J the advance ratio(J = U∞

nD ), n the rotational rate of the propeller inrevolutions per second, Q the torque, and T thethrust.

The RSB used to acquire the experimentaldata suffered from a low signal-to-noise ratio dueto interference effects with electromagnetic radi-ation from the wind tunnel motor. As a result,time-accurate experimental evaluations were notpossible. The time-averaged measurement dataon the other hand proved to be reproducible. Toremove high frequency noise from the experi-mental data a zero-phase digital filter with cut-offfrequency of 2500 Hz was used.

5.2.1 Isolated Configuration

The computed and measured propeller perfor-mance diagrams for the isolated configuration aredepicted in Fig. 12.

As expected a quasi-linear behavior is foundfor high advance ratios, while at the lower ad-vance ratios the non-linear blade section response

Velocity

Defi

cit∆U/U

∞[-]

VerticalCoordinate

Z[m

]

Lateral Coordinate Y [m]−R 0 R

0.00

0.05

0.10

0.15

0.20

R

0

−R

Fig. 13 Velocity deficit at the position of the pro-peller plane; U∞ = 26m/s, Xw = 0.3cpyl.

results in a non-linear behavior. Excellent agree-ment is obtained between the experimental andnumerical results for the thrust coefficient, withdifferences of at most 1% for advance ratiosabove J = 0.7. The larger differences at lower ad-vance ratios are as expected considering the highblade angles of attack experienced in this regime,reducing the accuracy of the numerical analysisof the blade section characteristics. The corre-spondence between the computed and measuredtorque coefficients is not as good as for the thrustcoefficient. This is likely the result of inaccura-cies in the drag coefficient data used in the XRO-TOR computations.

5.2.2 Installed Configuration

Based on the physics of the pylon - pusher pro-peller interaction it can be expected that the time-averaged thrust and torque increase in the in-stalled configuration. When rotating through thepylon wake the propeller blades experience an in-crease in angle of attack resulting from the lo-cally reduced inflow velocity. This leads to anincrease in lift produced by the blade in the wakeregion, resulting in distinct peaks in the thrustand torque signals. Note that this increase is par-tially offset by the reduction in dynamic pres-sure in the pylon wake. To illustrate the extentof the pylon wake region on the propeller disk,Fig. 13 presents an example of a computed veloc-ity deficit profile at the position of the propellerplane.

9

VELDHUIS, SINNIGE

The propeller inflow depicted in Fig. 13 wasused to compute the installed propeller perfor-mance according to the methods outlined in Sec-tion 4.2. The experimental data are not treatedhere since the measured data sets showed an in-consistent behavior between results obtained atdifferent freestream velocities. This is likely theresult of the fact that the small changes in the pro-peller performance due to installation fell withinthe expected variability of the measurements per-formed using the RSB.

A comparison between the numerical isolatedand installed propeller performance is presentedin Fig. 14.

INST

ISO

CQ

CT

η

PropellerPerform

ance

CT,C

Q,η[-]

Advance Ratio J [-]0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

0.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

Fig. 14 Computed effects of pylon installationon the propeller performance. Isolated (ISO) andinstalled (INST) configurations, U∞ = 26m/s.

From Fig. 14 it is clear that the effects ofinstallation on the time-averaged propeller per-formance are small. For advance ratios belowJ = 1.4 the increase in the thrust coefficient dueto installation is less than 2% of the isolatedvalue. For the same advance ratio range thetime-accurate thrust and torque signals displayedpeak-to-peak variations of less than 4%. Thesmall effects of installation on the propeller per-formance are as expected considering the lim-ited extent of the polar region in which the pylonwake is present.

5.2.3 Blown Configuration

The propeller performance in the blown config-uration was computed with the same methodsas used for the installed configuration, but nowwith a measured blown pylon wake profile (corre-sponding to the profile at Q = 600L/min shown

Vel

oci

ty|V

|[m

/s]

Y-c

oor

din

ate

(fre

estr

eam

dir

.)[1

/c]

X-coordinate [1/c]

−1.0 −0.5 0.0 0.5 1.0 1.5 2.035

40

45

50

55

−1.0

−0.5

0.0

0.5

1.0

Fig. 15 Velocity field obtained from PIV measur-ments in a fixed frame of reference at r/R = 0.87and J = 1.3.

in Fig. 10) as input. It was found that the ap-plication of pylon blowing further reduces the ef-fects of the presence of the pylon on the propellerperformance. For the same advance ratio rangeas considered before (J < 1.4), the computedtime-averaged thrust and torque coefficients inthe blown configuration were equal to their iso-lated counterparts. The corresponding peak-to-peak variations in the time-accurate blown pro-peller performance results displayed fluctuationsof at most 2%.

5.3 PIV analysis

PIV was performed to obtain velocity fieldsaround the propeller blade which could be usedfor additional analyses. In this case the velocitydata were used to determine the lift coefficient atdifferent radial positions by computing the localcirculation via a contour integration around thepropeller blade.

Fig. 15 shows a typical example of the veloc-ity field as determined by the PIV measurements.The dark blue area around the airfoil correspondsto a masked area that was removed to prevent un-wanted effects from erroneous datapoints (due toreflections) close to the blade surface. To obtainvelocity data in the shadow region, in front of theairfoil, interpolation was performed. From Fig.15 both the periodicity in azimuthal direction andthe wake area can clearly be recognized.

The radial distribution of the blade lift coef-

10

EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

J1.7 EXP

J1.7 NUM

J1.3 EXP

J1.3 NUM

J0.9 EXP

J0.9 NUM

Bla

de

Lift

Coeffi

cien

tc l

[-]

Radial Station r/R [-]

0.5 0.6 0.7 0.8 0.9 1.00.0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1.0

1.1

Fig. 16 Radial distribution of the blade lift co-efficient. Comparison of experimental values ob-tained from PIV data and computed results.

ficient computed from the PIV data is presentedin Fig. 16. The zero values close to the tip of thepropeller are very likely indicative for the con-traction of the slipstream that is not accounted forin the numerical propeller model. The cause forthe relatively large discrepancy between numeri-cal and experimental data for an advance ratio ofJ = 1.7 is yet unclear to the authors.

5.4 Propeller Noise Emissions

As mentioned previously, the unsteady bladeloads resulting from the installation of the pylonupstream of the propeller lead to additional noiseemissions. Application of the pylon blowing sys-tem should fill up the pylon wake, thereby reduc-ing the noise penalty due to installation. To as-sess the magnitude of the noise penalty due to in-stallation and to verify whether the pylon blowingsystem could indeed result in reduced noise lev-els, the propeller noise emissions were computedand measured in the isolated, installed, and blownconfigurations. The experimental total noise lev-els were obtained by extracting the SPL of thefirst ten propeller tones and subsequently sum-ming these.

5.4.1 Unblown Configuration

The propeller noise emissions in the installedconfiguration are shown as a function of the ad-vance ratio in Fig. 17. Both the experimental andnumerical data are considered.

NUMEXP

SoundPressure

Level

SPL

[dB]

(re20µPa,R

=3.3D

pro

p)

Advance Ratio J [-]0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

55

60

65

70

75

80

85

90

95

Fig. 17 Total sound pressure levels versus ad-vance ratio; installed configuration, U∞ = 19m/s,θ = 110◦, φ = 90◦.

As can be seen, a good agreement is ob-tained between the numerical and experimentaldata over the entire advance ratio range. How-ever, the experimental data show a distinct dropin noise levels for the advance ratio range 1.0 ≤J ≤ 1.3. This is likely the result of cancellationof the sound fields associated with the steady andunsteady blade loads.

A comparison between the isolated and in-stalled noise levels at an axial directivity angle ofθ = 110◦ (see Fig. 20, presented in the discussionof the blown results) shows that the installation ofthe pylon upstream of the propeller significantlyincreases the propeller noise emissions. Depend-ing on the advance ratio, noise increases of up to15 dB were measured. Note that in the advanceratio range 1.0 ≤ J ≤ 1.3, the isolated noise lev-els are higher than the installed values. This cor-responds to the observation, made before, that inthis advance ratio range the installed noise lev-els show a distinct drop due to cancellation of thesound fields associated with the steady and un-steady blade loads.

The effects of installation significantlychange the directivity pattern of the propellernoise emissions. Computations performed at afreestream velocity of 50 m/s led to the axial andcircumferential directivity patterns of the soundpressure levels of the propeller noise as depictedin Fig. 18 and Fig. 19, respectively. Again theSPL was defined at an observer distance of 3.3times the propeller diameter.

Figure 18 displays that the effects of instal-lation of the upstream pylon are especially pro-nounced in the up- and downstream directions.Whereas for the isolated propeller the noise emis-

11

VELDHUIS, SINNIGE

110 dB

110 dB

100

100

90

90

80

80

70

70

60

60

50

50

40

40

30

30

20

20

10

10

INST

ISO

AxialDirectivityAngle

θ[deg]

150◦

120◦

90◦60◦

30◦

0◦

330◦

300◦270◦

240◦

210◦

180◦

Fig. 18 Computed axial directivity of isolatedand installed noise emissions; U∞ = 50m/s, J =0.9, φ = 90◦.

sions vanish towards the propeller axis, in the in-stalled configuration this is no longer the case.Furthermore, the installation of the pylon intro-duces distinct lobes in the circumferential direc-tion as shown in Fig. 19. The installed noiselevels are highest in the circumferential range ap-proximately perpendicular to the pylon plane.

5.4.2 Blown Configuration

The goal of the pylon blowing system is to fillup the pylon wake, thereby reducing the instal-lation effects hence reducing the propeller noiseemissions compared to the unblown case. Toverify whether application of the pylon blowingsystem indeed led to noise reductions Fig. 20presents the measured noise levels in the isolated,installed, and blown (Q = 660 L/min) configura-tions.

Figure 20 shows that the application of thepylon blowing system indeed reduced the pro-peller noise emissions. Depending on the ad-vance ratio noise reductions due to blowing ofup to 7 dB are observed when compared to theunblown, installed configuration. In general thepropeller noise emissions in blown conditions arestill higher than for the isolated configuration, in-dicating that the installation effects are not com-pletely eliminated by the application of blowing.This is as expected considering the blown pylon

110 dB

110 dB

100

100

90

90

80

80

70

70

60

60

50

50

40

40

30

30

20

20

10

10

INST

ISO

Circu

mferentialDirectivityAngle

φ[deg]

330◦

300◦

270◦240◦

210◦

180◦

150◦

120◦90◦

60◦

30◦

0◦

Fig. 19 Computed circumferential directivityof isolated and installed noise emissions; U∞ =50m/s, J = 0.9, θ = 90◦.

wake profiles discussed before, which did not be-come completely uniform due to insufficient mix-ing of the external and blown flows.

All noise results discussed so far consideredthe combination of all propeller tones. Further-more, only a single blowing rate was consid-ered. To further increase insight in the effectsof blowing on the propeller noise emissions, Fig.21 presents the tonal noise reductions measuredat all blowing rates considered, for the first sixpropeller tones separately. Note that these val-ues are only valid for a single operating point(U∞ = 19m/s and J = 0.9). The orange dashedlines indicate the expected variability in the noisemeasurements.

From Fig. 21 it is concluded that the applica-

BLOW (Q = 660 L/min)

INSTISO

SoundPressure

Level

SPL

[dB]

(re20µPa,R

=3.3D

pro

p)

Advance Ratio J [-]0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0

55

60

65

70

75

80

85

90

95

Fig. 20 Total sound pressure levels versus ad-vance ratio (experimental results); isolated, in-stalled, and blown configurations, U∞ = 19m/s,θ = 110◦, φ = 90◦.

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EFFECT OF PYLON TRAILING EDGE BLOWING ON A PROPELLER

Q=660 L/min

Q=600 L/min

Q=500 L/min

Q=400 L/min

Changein

SoundPressure

Level

SPL

BLOW

−SPL

INST[dB]

BPF Tone Index [-]1BPF 2BPF 3BPF 4BPF 5BPF 6BPF

−18

−16

−14

−12

−10

−8

−6

−4

−2

0

2

Fig. 21 Tonal noise reductions due to blowing:effects of the blowing rate; blown configuration,U∞ = 19m/s, J = 0.9, θ = 110◦, φ = 90◦.

tion of blowing reduces the sound pressure levelsof the tones for all but one of the blowing ratesconsidered. The SPL of the propeller tones de-creases with increasing blowing rate, with the re-ductions becoming significant for all six tones atthe blowing rates of 600 and 660 L/min. Whenblowing at a rate of 660 L/min the 1BPF tone isreduced by almost 4 dB, while the reductions forthe higher BPF multiples are even larger.

Considering the wake profiles presented inFig. 10 it might be surprising that thelargest noise reductions were achieved at Q =660 L/min. The wake measurements showed thatat a comparable blowing rate (Q = 680 L/min)the application of blowing resulted in the intro-duction of a jet with higher than freestream ve-locity in the center of the wake region. However,it should be noted again that the wake measure-ments were performed using the isolated pylonmodel only. With the thrusting propeller present,the external velocity at the position of the blow-ing outlet will be increased. Considering that theeffectiveness of the blowing system reduces withincreasing external flow velocity (see Fig. 11),it is concluded that it might be the case that thelocally increased velocities ahead of the thrust-ing propeller demand a higher blowing rate to fillup the pylon wake than required for the isolatedcase. As such, it might be possible that in pow-ered conditions the wake profile at the propellerplane shows a minimum integral velocity deficitfor a blowing rate of Q = 660 L/min, thereby

explaining the best performance at this blowingrate. To test this hypothesis wake surveys shouldbe performed with the rotating propeller presentbehind the pylon.

It is expected that there exists a certain blow-ing rate for which the noise reductions due toblowing display a maximum, after which thenoise levels increase again for higher blowingrates. Considering that the noise levels continuedto decrease with increasing blowing rate for allblowing rates considered, it is concluded that themost effective blowing rate might not have beenreached. Additional tests should be performed athigher blowing rates to verify this.

6 Conclusions

Experimental and numerical analyses of the ef-fects of pylon blowing on pusher propeller per-formance and noise emissions have been per-formed successfully. From the results obtainedso far the following conclusions can be drawn:

• Application of the pylon blowing systemresulted in reductions in the integrated ve-locity deficit of up to 60% compared to theunblown configuration. However, the mix-ing of the external and blown flows was notoptimal, as a result of which the velocityprofiles in blown conditions were not com-pletely uniform.

• The agreement between the experimen-tal and numerical propeller performancefor the isolated configuration was good,with differences between the computed andmeasured thrust coefficients smaller than1% for advance ratios above 0.7. It wasconcluded that the effects of installation onthe time-averaged propeller performanceare small, with increases in the thrust andtorque coefficients due to installation ofless than 2% for advance ratios below 1.4.Accordingly, it was found that the effectsof blowing on the time-averaged propellerperformance are small.

• PIV measurements around the propellerblade basically confirmed the loading dis-

13

VELDHUIS, SINNIGE

tributions that were found from the numer-ical model. However, no explanation fordiscrepancies at the large advance ratio ofJ = 1.7 was found, so far.

• From the experimental and numerical eval-uations of the propeller noise emissions itis concluded that the installation of an up-stream pylon strongly increases the noiselevels, with measured noise penalties of upto 15 dB compared to the isolated case.The pylon blowing system was successfulin reducing the propeller noise emissionscompared to the unblown case. Depend-ing on the advance ratio noise reductionsof up to 7 dB were observed. Consider-ing that the noise reductions increased withincreasing blowing rate, it was concludedthat the most effective blowing rate mightnot have been reached during the experi-ments.

The results presented in this paper confirm thepotential of pylon trailing edge blowing to re-duce the adverse installation effects experiencedby pusher propellers. Considering the signifi-cant fuel savings promised by future engine con-cepts employing propellers in a pusher configu-ration, this is an important result which can beused to develop potential solutions for the rela-tively high noise emissions associated with suchpropulsion systems. Follow-up research usingadditional computational and measurement tech-niques is required to increase the understandingof the mixing characteristics of the blowing sys-tem and its effects on the propeller performanceand noise emissions.

Acknowledgments

The authors would like to thank Christian Potmafrom FlowMotion for the efforts he put into thedesign of the “Uniform Blowing Rod”, a devicewhich was essential for the successful comple-tion of the experimental test campaigns. Further-more, DNW is kindly acknowledged for provid-ing the propeller test rig under a lease contractwith Delft University of Technology.

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