Design, Fabrication, Test, and Evaluation of Small-Scale Tiltrotor Whirl FlutterWind Tunnel Models

Guillermo J. Costa, Sandilya Kambampati, and Samuel C. JohnsonGraduate Research Assistants

Penn State VLRCOEUniversity Park, PA, USA

Edward C. SmithProfessor of Aerospace Engineering

Penn State VLRCOEUniversity Park, PA, USA

ABSTRACTThree generations of small, semispan, subscale tiltrotor wind tunnel models were designed, fabricated, and windtunnel tested. The goal of this project was to develop a series of flutter tests at small scale and low cost that wouldexperimentally validate a series of analytical models developed in-house; this required a wind tunnel model that hadas low a flutter speed as was feasible, in order to permit it to be tested to the point of instability within the operatinglimits of the current facility. The first-generation model consisted of a hollow plastic wing rapid-prototyped fromABS plastic, with a three-bladed rotor consisting of constant-chord wooden blades. Five different configurations ofthis first-generation model were tested, but only one configuration exhibited whirl flutter within the test facility, at aspeed of 115 ft/s; however, this configuration was only able to exhibit whirl flutter through the use of a one-poundsteel mass mounted aft of the wing trailing edge. For the unstable configuration, the center of gravity (c.g.) of theaft-mass was located 5.5 in. aft of the wing elastic axis. The second-generation model used the same wing, but featuredcomposite rotor blades; this second-generation model exhibited whirl flutter at a tunnel speed of 113 ft/s, but was alsoincapable of experiencing an instability without the use of the aft-mass. The third-generation model consisted of acomposite wing and composite rotor blades, with an integrated wing spar that acted as a flexure; two configurationsof this third-generation model exhibited whirl flutter within the test facility, at tunnel speeds of 95 and 105 ft/s, andshowed excellent correlation with the analytical model.

NOTATION

a = airfoil section lift-curve slope, /radc = blade chord, ftEIbeam = beamwise stiffness, lb f − f t2

EIchord = chordwise stiffness, lb f − f t2

GJ = torsional stiffness, lb f − f t2

Ib = single-blade bending inertia, slug− f t2

[M] = wing/rotor system mass matrix[C] = wing/rotor system damping matrix[K] = wing/rotor stiffness systemR = rotor radius, ft[X ] = wing/rotor degrees of freedom vectorV∞ = freestream velocity, ft/sβ = flapwise degree of freedomγ = Lock number, dimensionless

Presented at the AHS 71st Annual Forum, Virginia Beach,Virginia, May 5–7, 2015. Copyright c© 2015 by the AmericanHelicopter Society International, Inc. All rights reserved.

δ3 = kinematic pitch-flap coupling angle, deg.λ = eigenvalueµ = advance ratio, V∞/ΩR, dimensionlessνβ = rotor flapwise frequencyρ = freestream density, slug/ f t3

Ω = rotor shaft speed, rad/sω = circular natural frequency, rad/sζ = lagwise degree of freedom; damping ratio

INTRODUCTION

Tiltrotor aircraft require rotor flapping, achieved via flaphinges or a gimbal, for good flight control characteristics andlow loads; however, the coupled wing-pylon-rotor system canbecome unstable in airplane mode at high forward speeds.This instability is commonly known as whirl flutter. Whirlflutter involves interactions of aerodynamic forces on the ro-tor, and cyclic flap and lag motions. The rotor dynamics canimpart aerodynamic and inertial forces and moments to thewing, which then cause beamwise, chordwise, and torsionalperturbations of the wing. These wing perturbations furtherdisturb the rotor, which can then lead to positive feedback andinitiation of the whirl mode. Traditionally, the flutter mar-

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gin of tiltrotors has been increased through the use of thickwings to maximize the wing stiffness; this comes at the costof increased structural weight and decreased aerodynamic ef-ficiency.

The first rotor/pylon instability of a tiltrotor was discov-ered during flight testing of the Bell XV-3, which crashed dueto an instability very much like whirl flutter (Ref. 1). Test-ing at the NASA Ames 40x80 wind tunnel in 1962 (Ref. 2)were conducted as part of an effort to find a solution. Fur-ther stability tests, such as the NASA Langley WRATS model(Refs. 3,4), validated the concept of using a cantilevered semi-span wing with attached rotor and pylon. The results of theWRATS tests were used for design improvement, paramet-ric studies, and the validation of analytical methods. How-ever, these WRATS studies were done at a much larger scale(1/5-scale), and with much more expensive equipment, thanthe present work. Further studies (Ref. 5) investigated theeffects of wing, control system, blade stiffness, and kine-matic pitch-flap coupling on whirl flutter stability, and wereused to improve the whirl flutter margin of the V-22 Osprey.Whirl flutter testing of tiltrotors having wings with compos-ite bending/torsion couplings (Ref. 6), and soft-inplane ro-tors (Ref. 7) has also been conducted using variants of theWRTAS test system. More recent investigations, in Europeand China (Refs. 8,9), have focused on the validation of whirlflutter prediction analyses and stability augmentation schemathrough the use of smaller wind tunnel models.

In 2013-2014, several new approaches to expanding whirlflutter envelopes for high-efficiency and high-speed tiltro-tors were explored at Penn State via numerical simulation.Among these include aeroelastic tailoring of tiltrotor wingswith wing extensions and winglets (Refs. 10–12). Emergingconcepts, such as unique “stepover” control systems that en-abled the design and testing of four-bladed tiltrotor hubs, werealso fabricated and tested within the industry (Ref. 13). Thepresent work describes the development of a new computa-tional model of tiltrotor whirl flutter, which was used to guidethe creation of several small-scale, semispan, tiltrotor whirlflutter models designed specifically for use in the Penn Statelow-speed wind tunnel. These wind tunnel models were de-signed, developed, and tested in the 2013 - 2015 timeframe. Abrief overview of the numerical tools developed to predict thestability of the wind tunnel models is given, and the evolutionof the wind tunnel models are traced from the initial efforts totheir current form.

The semispan wind tunnel models that were tested are cat-egorized into three generations (termed Gen-1, -2, and -3),each of which incorporated upgrades to improve the testabil-ity of the model within the operating capabilities of the windtunnel. Although these models were not dynamically scaledfrom a specific aircraft, they were intended to incorporate sev-eral features of full-scale tiltrotors, including a forward-sweptwing, gimbaled hub, collective pitch control, and the incor-poration of pitch-flap coupling (δ3). The specific features ofeach model are detailed in later sections.

All of the models tested to date were unpowered, and col-

lective pitch was used to maintain a constant RPM across theentire range of tunnel speeds. The models were excited inthe wing beamwise mode using compressed air. The modelswere all excited near their fundamental wing beamwise fre-quency, and root forces and moments were recorded using aload cell mounted to the wind tunnel wall. The sampling fre-quency was held constant at 1,000 Hz across all tests; modaldamping of the wing was calculated using a moving-blockmethod (Ref. 14) and compared with the numerical predic-tions.

AEROELASTIC MODEL

An analytical model consisting of a three-bladed rotor and py-lon mounted on a cantilevered wing is developed and vali-dated by the authors in Ref. 10. The detailed derivation of therotor aeromechanical equations can be found in Ref. 15, and16 describes the derivation of the equations for the wing aswell as the coupling of the wing and rotor systems.

Rotor and pylon model

A schematic of the rotor and pylon system is shown in Fig. 1.The rotor has three flap degrees of freedom (one coning, β0,and two cyclic, β1C and β1S), and three lag degrees of freedom(ζ0,ζ1C,ζ1S). Because the rotor examined in the present workdoes not contain precone and the control system is assumed tobe rigid, the blade lagwise and torsional degrees of freedomis also assumed to be rigid. Because the hub has no precone,the stiffness of the control system is not of primary impor-tance. The pylon has six degrees of freedom: three transla-tional (xP,yP,zP), and three rotational (αx,αy,αz). The result-ing hub forces (T,H,Y ) and moments (Mx,My,Q) due to therotor and pylon degrees of freedom are transferred on to thewing. A linear strip theory is used to model the aerodynamics,and the rotor is trimmed to be windmilling (Q = 0) at a prede-fined RPM. All the equations are derived in the non-rotatingframe of the rotor.

Fig. 1. Aeroelastic model of rotor and wing.

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Wing finite element model

A schematic of the wing finite element model is shown in Fig.2. Each element has two nodes, and each node has five de-grees of freedom: two beamwise bending (w,w′), two chord-wise bending (v,v′), and one torsional (φ ) degree of freedom.The moment-curvature relationship is given by

MyMzMφ

=

EIb 0 00 EIc 00 0 GJ

w′′

v′′

φ ′

(1)

Fig. 2. Wing finite element model, from Ref. 16.

Coupling of rotor and wing systems

The wing motion affects the rotor motion, and the rotor hubforces affect the wing motion. To simulate this phenomenon,the finite element model couples the pylon rotational andtranslational degrees of freedom (αx,αy,αz,xp,zp) with thewing bending and torsion degrees of freedom (w,v,φ ,w′,v′) atthe wing tip. The resulting combined equations of motion forthe rotor and wing can be written as a set of standard second-order ordinary differential equations:

[M ¨[X ]+ [C] ˙[X ]+ [K][X ] = 0 (2)

The frequency and damping of the system can be calcu-lated from the eigenvalues (λi) of Eq. 2. The circular fre-quency (ωi) and damping (ζi) of the ith mode are given, re-spectively, by the imaginary and real parts of the ith eigen-value, λi:

ωi = Im(λi) (3)

ζi =Re(λi)

|λi|(4)

A particular mode is unstable if the damping of that modeis negative. The lowest wind tunnel speed at which the damp-ing of any of the modes is zero is defined as the whirl flutterspeed.

WIND TUNNEL MODEL DESCRIPTIONSAND PHYSICAL PROPERTIES

Three generations of wind tunnel models were designed andfabricated in-house, with each successive model generationattempting to address at least one limitation of the one thatpreceded it. A Selig S4233 airfoil was used for the wing of allwind tunnel models. The rotor used for all wind tunnel modelswas a three-bladed, stiff-inplane rotor with a gimbaled hub.Initial efforts focused on the use of commercially-availablepropellers offered by conventional hobby vendors; however,these plastic propellers were too heavy to result in realisticLock numbers. Instead, the tail rotor blades from a 600-size radio-controlled helicopter were used for the Gen-1 rotor;these blades were untwisted and consisted of a balsa/sprucecore with a plastic overwrap, had a 1.28 in. constant-chordplanform. The rotor blades of the Gen-1 model were symmet-rical and of a constant NACA 0011 section, based on the mea-sured chord to thickness ratio of the blades; this blade airfoilremained constant for all wind tunnel model configurations.

The rotor hub for all models features kinematic pitch-flapcoupling, with a δ3 angle of ±47; no blade flexure or pre-cone was included in the rotor hub. The maximum rangeof the gimbal is 10 in any direction. For the present work,the rotor’s pitch-flap coupling is defined as negative (flap-up,pitch-down) when the sense of the δ3 angle is positive; con-versely, the pitch-flap coupling is defined as positive (flap-up,pitch-up) when the sense of the δ3 angle is negative. Thisdefinition of δ3 is in accordance with standard practice, as de-fined in Ref. 13. The geometry of the wind tunnel model’shub (Fig. 3) defines the geometry of the δ3 angle, which forthe present work was fixed at ±47.

Fig. 3. Rotor kinematic pitch-flap coupling (δ3) geometry.Negative sense of δ3 (pitch-up, flap-up) shown.

The components housed within the nacelle (Fig. 4) includea digital servo, which was used to control the collective pitch

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of the rotor; this permitted each of the models to operate ata fixed RPM regardless of tunnel speed. A shaft collar at-tached to the rotor shaft was used to mount a small disc mag-net, which was used in conjunction with a Hall effect sensor tomeasure rotor RPM during testing. Root forces and momentsas measured by the load cell were recorded using a customLabVIEW code developed in-house on a test computer; thepitch control servo was controlled by a pulse-width modula-tion control board located outside of the wind tunnel test sec-tion.

Fig. 4. Nacelle components.

Gen-1 model

The Gen-1 model was developed and tested during the Jan-uary 2012 - May 2013 timeframe. Gen-1 also set the standardfor many of the features and test methods that would be in-corporated into subsequent models. The semispan length ofGen-1 was constrained to 9.38 in. due to the limitations ofthe rapid-prototyping machines available at the time of themodel’s construction. Other parameters, such as rotor radiusand wing chord, were selected based on the geometric rela-tionships of similar components of the XV-15 (e.g. rotor ra-dius normalized by wing semispan, wing chord normalized bywing semispan, etc.); for instance, the wing chord was chosento be 3 in., or 32% of the semispan, based on the XV-15’schord-to-semispan ratio of 32.6% (Ref. 1).

The Gen-1 model featured a hollow monocoque wing thatwas rapid-prototyped in-house from ABS plastic. The wing,root mounting plate (for attachment to the load cell), and anacelle (which provided mounting points for the rotor mecha-nisms) were all integrated into a single structure, which madethe Gen-1 wing very stiff in chord and torsion. A contouredchannel was incorporated into the Gen-1 wing to permit thepass-through of the wires and tubing necessary for rotor con-trol and model excitation. The wing was swept forward 5

in order to accommodate flapping of the gimbaled rotor. Anassembly of the Gen-1 model is shown in Fig. 5.

Fig. 5. Gen-1 wind tunnel model (Config. 1 shown).

A one-pound steel mass was used to alter the stability ofthe Gen-1 model. Two of the Gen-1 configurations (Con-figs. 2 &3) had this mass mounted directly to the nacelle(Fig. 6), and two configurations (Configs. 4 &5) had this massmounted aft of the wing trailing edge (Fig. 7) such that thec.g. of the steel mass was located 5.5 in. aft of the wing’s elas-tic axis. Specifications of the Gen-1 model are summarized inTable 1.

Fig. 6. Gen-1 Configs. 2 &3, with end-mass.

Gen-2 model

The primary function of the Gen-2 model was to examinethe effect of higher-Lock number blades on the stability ofthe ABS monocoque wing. Thus, the only difference be-

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Fig. 7. Gen-1 Config. 4, with aft-mass.

Fig. 8. Effect of added mass on Gen-1 modal frequenciesvs. tunnel speed.

tween Gen-2 and Gen-1 was the use of custom compositerotor blades that were fabricated in-house; these compositeblades were slightly longer (0.5 in.) and approximately 40%lighter than the wooden Gen-1 blades, which increased theLock number from 2.36 to 3.23. The Gen-2 blades were lin-early tapered (1.28 in. root chord, 0.896 in. tip chord) in orderto maintain a geometric solidity of 0.1 between Gen-1 andGen-2. A summary of the Gen-2 model is given in Table 2.The two Gen-2 configurations tested (Gen-2a and -2b) corre-spond, respectively, to the most- and least-stable configura-tions of the Gen-1 model (Gen-1 Config. 1 and Gen-1 Config.4).

Gen-3 model

Unlike Gen-2, which was in many ways a modification ofGen-1, the Gen-3 model was developed specifically to addressthe need to experience an instability without the need for a

Table 1. Gen-1 model summary.

Item Configuration1 2 3 4 5

Wing type Rapid-prototype ABS monocoqueSemispan 9.38 in.Rotor radius 7.55 7.55 7.55 7.55 0Lock number 2.36 2.36 2.36 2.36 0δ3 +47 +47 −47 −47 −47

Added mass None End End Aft Aft

Table 2. Gen-2 model summary.Item Gen-2a Gen-2bWing type Rapid-prototype ABS monocoqueSemispan 9.38 in.Rotor radius 8.05 in.Lock number 3.23δ3 +47 −47

Added mass None Aft

one-pound steel tip mass to be mounted aft of the wing trail-ing edge. This required the wing modal stiffnesses to be tunedto specific values, which needed to be low enough to permitflutter within the operating limits of the facility; this necessi-tated the Gen-3 wing to incorporate a rectangular aluminumspar bonded to an internal channel within the wing (Fig. 10; aswith the ABS monocoque wing, a pass-through channel wasincluded in the Gen-3 wing to permit the necessary controlhardware to be routed through the wing.

The composite wing was fabricated in-house, and con-sists of a rectangular aluminum spar bonded to a foam wingcore with epoxy; the nacelle was secured to the wing coreusing machine screws and epoxy. A single layer of TorayT300 3x1 twill fabric (5.7 oz/yd2) was used as a wing over-wrap, securing the nacelle to the wing core while providinga smooth aerodynamic surface. Spar stiffnesses were calcu-lated based on material properties and cross-sectional inertia,using the methods given in standard mechanics of materialstexts (Ref. 17), whereas the stiffnesses of the carbon/foam-core section of the wing were calculated using classical sand-wich panel theory (Ref. 18) and composite box-beam the-ory (Ref. 19). For the purposes of calculating wing stiffnessvalues, the composite section of the wing was treated as asingle-cell composite box beam. A comparison of the sparand wing stiffnesses is presented in Table 3.

Table 3. Gen-3 model wing properties.Item Spar WingEIbeam, lb f − f t2 9.841 63.16EIchord , lb f − f t2 17.50 1,626GJ, lb f − f t2 42.12 440.4Spar cross-section 0.250 x 0.190 in.Spar material 6060-T6

The use of a composite wing allowed the Gen-3 wing toextend farther into the test section of the wind tunnel (Fig. 11),

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Fig. 9. Gen-1 & -2 fan plot.

Fig. 10. Gen-3 wing components.

which includes a flexure region inboard of the wing (3.080 in.,as given in Fig. 11). This flexure region allows the stiffnessof the composite wing to be neglected in analysis, and thewing modal deflections can thus be treated as solely due tothe stiffness of the spar. The flexure region also allowed theradius of the composite blades to be increased by an additional0.5 in. relative to the Gen-2 model (to a maximum rotor radiusof 8.55 in. for Gen-3b), and resulted in a Lock number of 3.70for the Gen-3b configuration.

One configuration of the Gen-3 model (Gen-3b) featuredtwisted blades, which were designed specifically to minimizethe amount of the blades that were in potential stall; Giventhe range of operating Reynolds numbers for the wind tun-nel model (100,000 – 120,000 at 0.75R), a static stall angleof 10 was assumed (Ref. 20) for the blade sections. Forthe untwisted Gen-1 and Gen-2 blades (Fig. 12), it is read-ily apparent that the inner 45% of the blade span is above thestatic stall angle of attack at all tunnel speeds of interest. Thestalled, driving, and driven regions of the blade span shown inFigs. 12 and 13 are defined as follows:

• Stalled: the region of the blade that operates above thestatic stall angle of attack and causes drag, slowing therotation of the blade.

• Driving: the autorotative region wherein the total aero-dynamic force is inclined slightly forward of the axis of

Fig. 11. Gen-3 model mounted in wind tunnel.

rotation, which supplies thrust and accelerates the rota-tion of the blade.

• Driven: the region of the blade wherein the total aerody-namic force is inclined slightly aft of the axis of rotation,resulting in a drag force which slows the rotation of theblade.

Fig. 12. Local angle of attack by blade station, θtw = 0;Ω =1,600 RPM (Gen-1 & Gen-2 models).

A blade element momentum theory analysis was used toselect a twist rate for the blades that would minimize the

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amount of the blade that was above the static stall angle ofattack throughout the tunnel range of interest (approximately70 - 120 ft/s), and a twist rate of −25/R was selected; bladetwist rate was kept constant across the blade span for ease offabrication.

Fig. 13. Local angle of attack by blade station, θtw =−25/R; Ω = 2,000 RPM (Gen-3 models)

The Gen-3 model consisted of three configurations: Gen-3a, Gen-3b, and Gen-3c, the details of which are summarizedin Table 4. The Lock number presented in Table 4 is definedin the usual fashion:

γ =ρacR4

Ib(5)

The Gen-3c variant was a blades-off configuration simi-lar in purpose to Gen-1 Config. 5, and was tested to ensurethat any instabilities that were experienced during testing werecaused solely by the influence of the rotor, and thus indicativeof whirl flutter.

Table 4. Gen-3 model summary.Item Gen-3a Gen-3b Gen-3cWing type Composite with integrated sparSemispan 14.35 in.Rotor radius 8.05 in. 8.55 in. 0Twist rate 0 −25/R 0Lock number 3.23 3.70 0δ3 −47 −47 0Added mass None None None

For ease of reference, Table 5 summarizes how key fea-tures vary across all three generations of the wind tunnel mod-els.

TEST METHODOLOGYThe setup for wind tunnel testing is shown in Fig. 14. A pneu-matic solenoid valve is connected to an external compressed

Table 5. Key wind tunnel model features by generation.Item Gen-1 Gen-2 Gen-3Semispan 9.38 in. 14.35 in.Rotor radius 7.55 in. 8.05 in. 8.55 in.Twist rate 0 0 −25/RBlade taper 0 0.7 0.7Single-blade mass, grams 9.35 5.65 5.91Lock number 2.36 3.23 3.70

air supply and is wired to the DAQ used for data acquisition;the solenoid is plumbed to two perturbation jets (Fig. 15) thatare bonded to the nacelle and are aligned with the wing’s elas-tic axis.

Fig. 14. Test setup; blades shown were not flutter tested.

The wind tunnel is then brought to the desired speed, andthe servo controller is used to adjust the collective pitch of themodel in order to hold the rotor fixed at a desired RPM. Theair supply valve is then opened, and the solenoid is activated;the data acquisition code then drives the solenoid switching ata user-specified frequency, which for all tests was taken to bethe fundamental beamwise frequency of the wing. This exci-tation is allowed to continue for several seconds, after whichthe solenoid is deactivated and the air supply valve is closed,allowing the model’s motion to dampen; for stable cases, themodel will return to its original state. The data acquisitioncode is used to record the decay in root forces and momentsvia the load cell.

Table 6. Wing root moment by load cell channel.Wing mode Load cell channel

Beam RollChord Yaw

Torsion Pitch

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Fig. 15. Pneumatic perturbation jets housed in nacelle cap.

A moving-block analysis was used to determine the damp-ing ratio of each test point, which used the time history of theappropriate root moment to determine the wing modal damp-ing ratio (Table 6). A sample time history of a stable caseis shown in Fig. 16. A fast Fourier transform (FFT) is takenof a ’block’ of data (usually known as a data window) afterthe model excitation has ended, and the natural log of funda-mental modal frequency magnitude is computed; the block isincremented forward by one time step and the process is re-peated, creating a plot of modal frequency magnitude versustime (Fig. 17). The slope of the linear region of this magni-tude plot is the damping ratio for the model at the tunnel speedunder consideration.

Fig. 16. Time history of stable case; block of data outlined.

The process can then be repeated at progressively highertunnel speeds, until the model experiences an instability or themaximum operating speed of the tunnel is reached. At eachtunnel speed, the collective pitch of the model was readjusted

Fig. 17. Damping ratio.

in order to maintain a constant RPM.

EXPERIMENTAL RESULTS

Generation 1 test results

The predicted pitch settings showed good correlation with theexperimental values. Although the inner 45% of the untwistedblade is above the static stall angle (Fig. 12), the outer portionis not and provides most of the torque to cause rotation. Thisimplies that the values of the lift and drag coefficients wereclose enough to those actually encountered to come within afew degrees trim.

Fig. 18. Gen-1 predicted and measured trim pitch settings.

The Gen-1 modal frequencies were determined via stick-rap prior to wind-on testing. The moving-block method wasused to calculate the modal damping of the Gen-1 model priorto stability testing; because the chord and torsion modes of the

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Gen-1 wing are considerably stiffer than the beam mode, onlythe wind-off modal damping of the wing beam mode is ofinterest. These properties are presented in Table 7. Becausethe Gen-2 model consisted of the same wing and placementsof the end-mass and aft-mass as the Gen-1 model, the modalfrequencies and damping of the Gen-2 model are assumed tobe the same as the Gen-1 model. The only difference betweenthe Gen-1 and Gen-2 models was the use of composite rotorblades, which changed the total rotor mass by 0.02 lbm; thisdifference in mass is between 1−4% of the all-up mass of theGen-1 and Gen-2 models (depending on configuration), andwas considered insignificant.

Table 7. Gen-1 and Gen-2 wing modal frequencies andwind-off damping.

Config. Mode fn, Hz fn/Ω ζbeam

1Beam 10.38 0.389

0.025Chord 44.56 1.670Torsion 30.40 1.140

2 & 3Beam 7.405 0.278

0.012Chord 32.96 1.236Torsion 30.40 1.140

4 & 5Beam 5.909 0.222

0.012Chord 32.96 1.236Torsion 10.40 0.390

Each of the Gen-1 configurations (Table 1) was in the windtunnel at least twice to ensure consistency and repeatability.The experimentally-measured damping ratios for each of theGen-1 configurations are plotted in Figs. 19-23, which com-pares each configuration of the Gen-1 model to the previousone; the maximum tunnel speed of 140 ft/s is shown in a heavydashed line. At a tunnel speed of 120 ft/s, it is immediatelyapparent that the addition of the end-mass was found to re-duce the modal damping of Gen-1 Config. 2 to approximatelyhalf that of Gen-1 Config. 1. When the sense of δ3 was re-versed (Gen-1 Config. 3), the modal damping of the wing wasnot found to change in any appreciable fashion (Fig. 20) upto a tunnel speed of approximately 110 ft/s; however, it maybe seen in Fig. 20 that at higher tunnel speeds, the dampingof Gen-1 Config. 3 was experiencing a downward trend. Itmust therefore be concluded that the sense of δ3 did not playa significant role in the stability of the Gen-1 model, but it ispossible that a negative sense of δ3 would radically alter thestability of the model at higher tunnel speeds that were outsidethe testing capabilities of the facility.

The addition of a one-pound end-mass to the nacelle (Con-figs. 2 & 3) reduced the modal damping of the model, but notenough to cause an instability. An instability was only en-countered when the end-mass attachment was relocated aft ofthe wing trailing edge (Config. 4) at a distance of 5.5 in. aftof the wing elastic axis. At lower tunnel speeds, the modaldamping of Config. 4 was very near that of Config. 3; how-ever, as the tunnel speed was increased, the modal damping ofConfig. 4 remained relatively constant at about 1% of criticalup to a tunnel speed of 110 ft/s. At a tunnel speed of 115 ft/s,

Fig. 19. Gen-1 damping for Configurations 1 and 2.

Fig. 20. Gen-1 damping for Configurations 2 and 3.

the wing oscillations continued to increase even after the exci-tation had ceased (Fig. 22); the wing oscillations increased inmagnitude, requiring the tunnel speed to be reduced in orderto prevent damage to the model.

To ensure that the observed instability was truly due todestabilizing forces of the rotor (and thus actually whirl flut-ter), the Config. 4 velocity sweep was repeated with the bladesremoved: except for the absence of the blades, Configuration5 was identical to Configuration 4. As seen in Fig. 23, themodal damping for Config. 5 was higher at all wind speedswith the rotor blades removed, and no instability was encoun-tered at any tunnel speed. This suggests that the instabilityencountered with Config. 4 was due solely to the influence ofthe rotor, rather than an unstable wing configuration.

Correlation between the predicted and experimentally-measured damping values for the Gen-1 model varied be-tween configurations, but generally showed poor correlation,as evidenced in Fig. 24

The Gen-1 Config. 4 tests were repeated, and a high-speed

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Fig. 21. Gen-1 damping for Configurations 3 and 4.

Fig. 22. Time history of Gen-1 unstable case, V∞ = 115 ft/s.

camera was used to capture the motion of the rotor during per-turbation. Examination of this video revealed that Gen-1 Con-fig. 4 experienced a strong beam/torsion coupling of the wing,which resulted in the gimbal bucket striking the rotor shaftduring excitation (Fig. 26). This impacting of the gimbal onthe rotor shaft effectively increases the flapwise frequency ofthe rotor. Artificially increasing νβ to 1.44 improves the cor-relation between the predicted and measured damping ratiosof Gen-1 Config. 4 (Fig. 27), but this is a crude way to capturea nonlinear phenomenon. Additionally, this is not representa-tive of real tiltrotors, and thus the Gen-1 variants cannot beconsidered the final iteration of the wind tunnel model: theresults were not predictable, the gimbal/shaft impacts causedby the large wing motions are not representative of real air-craft, and the use of the aft-mass to induce whirl flutter is nota realistic method of causing an instability.

Fig. 23. Gen-1 damping for Configurations 4 and 5.

Fig. 24. Gen-1 Config. 1 predicted and measured dampingversus tunnel speed.

Generation 2 test results

The predicted and measured trim pitch settings for Gen-2showed good correlation, and most of the measured trim pitchvalues were within 2 of the predicted values (Fig. 28).

One of the necessary improvements to the Gen-1 modelcentered on the fabrication of composite rotor blades thatwould increase the Lock number of the rotor. The use of com-posite rotor blades was the primary difference between theGen-1 and Gen-2 models. Two configurations of the Gen-2model (Gen-2a and Gen-2b, see Table 2) were tested. TheseGen-2 configurations may be thought of as ’surrogates’ oftheir Gen-1 counterparts (Table 8), with the only differencebetween them being the use of composite blades to increasethe Lock number of the Gen-2 models (Table 5).

The modal damping of the Gen-2 models was comparedto that of their Gen-1 surrogates at several identical tunnelspeeds. As was the case with Gen-1, the Gen-2 tests onlyexamined the beamwise mode of the wing. The modal damp-ing of the Gen-2a model (Fig. 29) closely followed that of

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Fig. 25. Gen-1 Config. 4 predicted and measured dampingversus tunnel speed.

Fig. 26. Gen-1 model experiencing wing beam/torsion cou-pling.

Gen-1 Config. 1 at the lower tunnel speeds (Fig. 30), andneither model experienced an instability within the operat-ing limits of the facility. Both Gen-1 Config. 1 and Gen-2aexperienced modal damping that increased in a relatively lin-ear fashion with increasing tunnel speed up to approximately100 ft/s; above this speed, the damping of Gen-2a was foundto be consistently less than that of Gen-1 Config. 1. Sincethe only difference between these two models was the higherLock number of Gen-2a, it is possible that Gen-2a would haveexperienced an instability at a lower tunnel speed than Gen-1Config. 1; however, this was not testable within the currentfacility’s operating limits.

The addition of the aft-mass (Gen-1 Config. 4 and Gen-2b) lowered the effective beamwise damping of both the Gen-

Table 8. Gen-1 and Gen-2 model comparisons.Item SurrogateGen-1 Config. 1 Gen-2aGen-1 Config. 4 Gen-2b

Fig. 27. Gen-1 Config. 4 with νβ = 1.44.

Fig. 28. Gen-2 predicted and measured trim pitch settings.

1 Config. 4 and Gen-2b models; as was the case with Gen-1Config. 4, the sense of δ3 was reversed for Gen-2b (from +47

to −47). The modal damping of Gen-2b closely followedthat of Gen-1 Config. 4 at lower tunnel speeds (Fig. 31); above80 ft/s, however, the damping of the Gen-2b model was foundto be consistently lower than that of Gen-1 Config. 4 for anygiven tunnel speed.

Once the Gen-2b model experienced an instability, itsmodal damping remained consistently negative. This wasfound to be in direct contrast to the modal damping of Gen-1Config. 4, which was found to experience a brief uptick near115 ft/s before remaining negative. The Gen-2b modal damp-ing remained consistently more negative than that of Gen-1Config. 4 at comparable tunnel speeds above 110 ft/s, how-ever the flutter speed of each of these models was virtuallyidentical (Table 9).

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Fig. 29. Gen-2a predicted and measured damping versustunnel speed.

Fig. 30. Gen-2a damping versus Gen-1 Config. 1 average.

Generation 3 test results

Three configurations of the Gen-3 model were tested. TheGen-3 model was an entirely new design, which was intendedspecifically to experience flutter without the need for an aft-mass. All of the Gen-3 model variants featured a compositewing with an integrated spar; the stiffness of this spar was se-lected so that the model would flutter at as low a tunnel speed

Table 9. Gen-1 and Gen-2 flutter speeds.Model δ3 Added mass Vf lutter, ft/sGen-1 Config. 1 +47 None > 140Gen-1 Config. 2 +47 End > 140Gen-1 Config. 3 −47 End > 140Gen-1 Config. 4 −47 Aft 115Gen-1 Config. 5 −47 Aft > 140Gen-2a +47 None > 140Gen-2b −47 Aft 113

Fig. 31. Gen-2b damping versus Gen-1 Config. 4 average.

Table 10. Gen-3 variants.Designation Rotor radius, in. Twist δ3 γ

Gen-3a 8.05 0 −47 3.23Gen-3b 8.55 −25/R −47 3.70Gen-3c 0 N/A N/A 0

as possible. As described previously, and the only differencebetween the Gen-3 variants is the rotor installed for testing.As with Gen-1 testing, the final configuration tested had therotor blades removed in order to ensure that the instabilitiesmeasured during testing were due solely to the influence ofthe rotor rather than any inherent characteristic of the wing.The Gen-3 model properties of the three Gen-3 variants aregiven in Table 10.

The wind-off modal frequencies and damping ratios for theGen-3 model were determined via stick-raps and the moving-block method prior to wind-on testing, in a fashion identicalto what was used for the Gen-1 model. These properties aregiven in Table 11, along with the masses of the wing and spar.Because the difference in the rotor mass between the Gen-3a and Gen-3b models is 0.002 lbm, or 0.24% of the Gen-3bmodel’s total mass, the differences in modal frequencies andwind-off damping between the Gen-3a and Gen-3b modelswas considered insignificant.

The predicted and measured trim pitch settings for theGen-3 model showed an acceptable correlation, although thetrim pitch settings for Gen-3b were slightly higher than thecalculated values (Fig. 33). It was noticed that the minimumtunnel speed necessary for windmilling of the Gen-3a modelwas slightly higher than that of the Gen-1 model (V∞ = 60 ft/svs 50 ft/s, respectively), and the minimum tunnel speed neces-sary for windmilling of the Gen-3b model was slightly higherstill (V∞ = 70 ft/s). Below these tunnel speeds, the Gen-2 andGen-3 models were found to either be unable to maintain astable RPM, or would operate with the gimbal in contact withthe rotor shaft.

The Gen-3 model was intentionally designed to be less stiff

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Table 11. Gen-3 wing modal frequencies and wind-offdamping.

Config. Mode fn, Hz fn/Ω ζbeam

Gen-3aBeam 5.316 0.160 0.0129Chord 7.526 0.226 0.0139

Torsion 26.49 0.795 0.0082

Gen-3bBeam 5.316 0.160 0.0129Chord 7.526 0.226 0.0139

Torsion 26.49 0.795 0.0082

Gen-3cBeam 5.619 0.169 0.0220Chord 7.940 0.238 0.0185

Torsion 27.95 0.839 0.0130

Fig. 32. Gen-3a predicted and measured trim pitch set-tings at blade root.

than that of the Gen-1 model, and as such the operating RPMof the rotor had to be reexamined in order to avoid a 1/revimbalance. The stick-rap tests performed on the Gen-1 modelwere repeated, and a fan plot was created to select the oper-ating speed of the Gen-3 rotor. From this fan plot (Fig. 34),it can be seen that operating the Gen-3 rotor at 1,600 RPMwould risk a 1/rev excitation driving the wing torsion mode.As such, the operating speed of the Gen-3 rotor was increasedto 2,000 RPM.

The Gen-3 damping as a function of airspeed was com-pared to the predictions of the numerical analysis. For each ofthe Gen-3a and Gen-3b tests, the model was perturbed at itsbeamwise natural frequency through the compressed air ex-citation jets; however, the small cross-sectional geometry ofthe wing spar allowed all three wing modes (beam, chord, andtorsion) to influence the stability of the wind tunnel model.These wing modes were measured using the roll (beam), yaw(chord), and pitch (torsion) channels of the load cell. Theanalysis predicted that the chord mode of Gen-3a would be thefirst wing mode to attain zero damping, at a tunnel speed ofapproximately 105 ft/s. In reality, the model experienced onetest point of negative damping at a tunnel speed of 99 ft/s; themodel did, however, experience repeated negative damping inthe chord mode beginning at 101 ft/s, and this was considered

Fig. 33. Gen-3b predicted and measured trim pitch set-tings at blade root.

Fig. 34. Gen-3 fan plot.

to be the flutter speed of the model.

The modal damping of the Gen-3b model was measuredin a fashion identical to that described previously for Gen-1, Gen-2, and Gen-3a. Unlike Gen-3a, the torsion mode ofGen-3b was predicted to go unstable first at approximately 97ft/s. The Gen-3b wind tunnel model experienced instabilityslightly before the predicted value, at a tunnel speed of 94-95ft/s (Fig. 36).

Near the flutter speed of Gen-3b, the model exhibited verylow damping, as expected in some cases requiring on the or-der of ten seconds for the oscillations to dampen (Fig. 37);however, once the flutter speed was reached, the oscillationsof the Gen-3b wing quickly increased, especially in torsion.

The Gen-3b model had a higher Lock number than Gen-3a(approximately 15% greater) and a much higher Lock numberthan Gen-1 (on the order of 50% greater), so the differencesin flutter behavior between the models is not surprising. Thefact that the untwisted blades of Gen-1, Gen-2, and Gen-3a

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Fig. 35. Gen-3a predicted and measured damping versustunnel speed.

Fig. 36. Gen-3b predicted and measured damping versustunnel speed.

had large regions of the blade span potentially stalled meansthat the forces imparted to the wing from the rotor were likelyreduced; it stands to reason that the forces imparted to thewing by the twisted blades of the Gen-3b rotor were thereforesignificantly higher.

The Gen-3c model was a blades-off configuration(Fig. 39), and served to demonstrate that the instabilities ob-served were indeed caused solely by the presence of the rotorand thus indicative of whirl flutter. The removal of the ro-tor blades resulted in the amplitudes of the chord and torsionbeing barely distinguishable from the noise floor.

No instability was predicted for Gen-3c, and no instabil-ity was encountered during testing. The damping of Gen-3c,in fact, was found to increase with increasing tunnel speedin a relatively linear fashion (Fig. 40). This would not re-main the case if the tunnel speed continued to increase indef-initely; however, within the operating range of the facility, it

Fig. 37. Time history of stable Gen-3b test point; 90 ft/stunnel speed.

Fig. 38. Time history of unstable Gen-3b test point; 99 ft/stunnel speed.

was not possible to make the blades-off configuration expe-rience an instability. Therefore, the instabilities encounteredby the Gen-3a and Gen-3b models were due solely to the in-fluence of the rotor, and are thus indicative of whirl flutter.There appeared to be good correlation between the calculatedand experimental damping ratios, with most (approximately70%) of the test points showing a mean damping ratio thatwas within 5% of the predicted values. The flutter speed ofall of the configurations tested is summarized in Table 12.

CONCLUSIONS

Three generations (Gen-1, -2, and -3) of semispan, small-scale tiltrotor whirl flutter models were designed, fabricated,and tested in a low-speed wind tunnel. Five configurationsof the Gen-1 model were tested, four of which incorporated asteel mass attached to the end of the nacelle or mounted aft ofthe wing’s trailing edge (Configs. 2–5). Acceptable correla-

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Fig. 39. Gen-3c model mounted to load cell.

Fig. 40. Gen-3c predicted and measured damping versustunnel speed.

Table 12. Flutter speeds by model generation and configu-ration.

Model δ3 Added mass Vf lutter, ft/sGen-1 Config. 1 +47 None > 140Gen-1 Config. 2 +47 End > 140Gen-1 Config. 3 −47 End > 140Gen-1 Config. 4 −47 Aft 115Gen-1 Config. 5 N/A Aft > 140Gen-2a +47 None > 140Gen-2b −47 Aft 113Gen-3a −47 None 101Gen-3b −47 None 95Gen-3c N/A None > 140

tion was observed between the predicted and measured damp-ing values of the first three Gen-1 configurations; however,the damping of Gen-1 Config. 4 showed very poor correlationwith the predicted values. The Gen-1 model exhibited whirlflutter only through the use of the aft-mass, but the couplingof the wing beam and torsion modes resulted in the gimbalstriking the rotor shaft during excitation. To ensure that the in-stability of Gen-1 was caused solely by the rotor (and thus anindication of whirl flutter), a fifth configuration (Gen-1 Con-fig. 5) was tested with the rotor blades removed. Gen-1 Con-fig. 5 did not exhibit any instability within the operating limitsof the facility.

The Gen-2 model was developed to measure the effect ofincreased Lock number on the stability of the Gen-1 wing.Two configurations of Gen-2 were tested and compared totheir Gen-1 counterparts: Gen-2a (comparable to Gen-1 Con-fig. 1) and Gen-2b (comparable to Gen-1 Config. 4). Gen-2a showed similar behavior to Gen-1 Config. 1, and neithermodel exhibited whirl flutter within the operating limits of thewind tunnel. Gen-2a exhibited damping very similar to Gen-1Config. 1 up to a tunnel speed of 120 ft/s; above this speed,Gen-2a’s damping was found to be consistently less than Gen-1 Config. 1, and at 140 ft/s (the tunnel’s maximum operatingspeed) the damping of Gen-2a was approximately 26% lowerthan that of Gen-1 Config. 1.

Gen-2b was the counterpart to Gen-1 Config. 4, which fea-tured a negative sense of δ3 and the use of the aft-mass. Thestability of the Gen-2b model closely followed the Gen-1 re-sults at lower tunnel speeds; above a tunnel speed of 80 ft/s,the damping of Gen-2b was found to be consistently lowerthan that of Gen-1 Config. 4. The Gen-2b model exhibitedwhirl flutter at a tunnel speed of 113 ft/s; this is comparableto Gen-1 Config. 4, which exhibited whirl flutter at a tunnelspeed of 115 ft/s.

Gen-3 was an entirely new model designed to experiencean instability without the use of the aft-mass. The Gen-3models had Lock numbers of 3.23 and 3.70, and was sub-divided into three configurations: Gen-3a, Gen-3b, and Gen-3c. The Gen-3 models also featured wing frequency ratiosthat were more representative of full-scale aircraft. The Gen-3a model used the untwisted composite blades used from theGen-2 model, and the Gen-3b model incorporated twisted ro-tor blades. The Gen-3c model was a blades-off configuration(similar to Gen-1 Config. 5), and was tested to ensure thatany instability measured during testing was due solely to theinfluence of the rotor, and thus indicative of whirl flutter.

The Gen-3a and -3b models showed good correlation withthe predicted values, and were able to exhibit whirl flutterwithout the use of the aft-mass: the Gen-3a model at approx-imately 101 ft/s, and the Gen-3b model at approximately 97ft/s. The Gen-3c model was not predicted to experience aninstability at any tunnel speed attainable with the current fa-cility, which was indeed validated with experimental results.

The work performed to date has further validated the fea-sibility of whirl flutter testing at a small scale. The Gen-3model was able to experience flutter at a low enough tunnel

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speed to permit preliminary testing of devices designed to en-hance the whirl flutter speed of tiltrotors within the operatingcapabilities of the current facility.

Author contact: Edward C. Smith, ecs5@psu.edu

ACKNOWLEDGMENTS

The authors wish to thank Dr. Jose Palacios of the Penn StateVLRCOE for his assistance with the data acquisition modulesused during wind tunnel testing. The authors would also liketo thank Mr. Rick Auhl for his support in test setup. G. J.Costa thanks the Penn State STEM Fellowship and NDSEGFellowship for their continued support.

This research is partially funded by the Government underAgreement No. W911W6-11-2-0011. The views and conclu-sions contained in this presentation are those of the authorsand should not be interpreted as representing the official poli-cies, either expressed or implied, of the U.S. Government.

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