NASA Technical Memorandum 85682AVRADCOM TECHNICAL REPORT 83-B-4

PARAMETRIC TIP EFFECTS FORCONFORMABLE ROTOR APPLICATIONS

HAYNE R, MANTAYWILLIAM T, YEAGER~ JR,

AUGUST 1983

NASA-TM-85682

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NJ\SI\National Aeronautics andSpace Administration

Langley Research CenterHampton. '/irginia 23665

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PARAMETRIC TIP EFFECTS FOR CONFORMABLE ROTOR APPLICATIONS

ABSTRACT

This research study was initiated to systematicallydetermine the impact of selected blade tip geometric parameterson aeroelastically conformable rotor performance and loadscharacteristics. The model articulated rotors includedbaseline and torsionally soft blades with interchangeabletips. Seven blade tip designs were evaluated on the baselinerotor and three tip designs were tested on the torsionally softblades. The designs incorporated a systematic variation inthree geometric parameters: sweep, taper, and anhedral. Therotors were evaluated in the NASA Langley Transonic DynamicsTunnel at several advance ratios, lift and propulsive forcevalues, and tip Mach numbers. Based on the test results, tipparameter variations generated significant rotor performanceand loads differences for both baseline and torsionally softblades. Azimuthal variation of elastic twist generated by thetip parameters strongly correlated with rotor performance andloads, but the magnitude of advancing blade elastic twist didnot.

1.0 Introduction

Reducing helicopter vibratory loads while improvingperformance through passive control has been the goal of theAeroelastically Conformable Rotor (ACR) concept. Initial ACRstudies (ref. 1) examined the potential of a conformable rotorto alter the unfavorable blade spanwise and azimuthal loaddistributions which lead to increased vibratory bending loadsand power requirements. In reference 1, test results on amodel hingeless rotor indicated that elastic ,twist measurablychanged blade loads on a torsionally soft bla'de. Promisingmethods of achieving this passive control concept, incorpora­ting time varying elastic twist, have been identified analyti­cally (ref. 2), and blade design features producing thatdesired elastic control were sought for an articulated rotor.

The effect of blade tip shape on rotor performance andloads has received much attention for application to multi­bladed helicopters (refs. 3-5). Experimental data have alsobeen obtained (ref. 6) which initiated identification of theblade tip shape as a promising passive control concept. Thereference 6 test utilized a model rotor blade with conventionaltorsional stiffness. Although the resulting loads and perform­ance of the configurations were tip-shape-dependent, theidentification of which parameter caused each load or perform­ance change was elusive. This was due, in part, to multipleparameter variations occurring with each tip change. Neverthe­less, the concept of passive control to achieve better rotor

1

performance while reducing loads was encouraged by theseresults and several conformable designs were pursued. Theresulting studies (refs. 7-8) considered variations in bladetorsional stiffness, airfoil section, mass distribution, andtrailing edge tab deflection, as well as tip geometry, in thedesign. Again, the wind-tunnel tests of these ACR conceptsproduced encouraging loads and performance data, but the aero­elastic mechanism for design success or failure was notobvious.

Expanded testing and analysis of the configurations ofreference 6 resulted in identification of several key issuesfor future ACR application and development (ref. 9). For thetorsionally stiff rotor used in that test, the parametricvariations of tip sweep, taper and anhedral did measurablychange the elastic twist and integrated performance, but theredid not appear to be a strong connection between elastic twistand performance. Additional tests on the blades of ref. 8which incorporated large tip spans and trailing edge tabdeflections (refs. 10-11) showed performance and'loadsvariations which, though significant, were not easilyexplainable by individual parameter effects.

The parameters most effective in improving conformablerotor performance and loads characteristics have thus not beensystematically determined. Although it has been shown thatchanges in adjustable trailing edge tabs have significanteffects on conformable rotor behavior (ref. ll), the rotorblade tip operates in a very influential portion of the rotordisk and thus provides significant research impetus. This isespecially true if ACR success is dependent on elastic twistcontrol. Consequently, the research study described herein wasinitiated to systematically determine the impact of selectedblade tip geometric parameters on ACR performance and loadscharacteristics.

2.0 Notation

a speed of sound, ft/sec

b blade number

CD rotor drag coefficient, DpnR2(nR}2

CL rotor lift coefficient, LpnR2(nR}2

CL nominal rotor lift coefficient

CQ rotor torque coefficient, QpnR3(nR}2

2

c

c.g.

a.c.

D

H

L

Q

r

R

v

p

a

w

blade chord, in.

section measured center of gravity location, in.

section computed aerodynamic center location, in.

rotor drag, lb.

longitudinal rotor force perpendicular to control axis,1b.

rotor lift, lb.

rotor blade tip Mach number, n:

rotor torque, ft-lb.

blade radial station, ft

rotor radius, ft

free-stream velocity, ft/sec

angle of attack of rotor shaft, deg.

elastic twist angle, positive nose-up, deg.

d t o Vrotor a va nce ra 10, nR

mass density of test medium, slug/ft3

nominal rotor solidity ratio, bc/nR = .082

. . . total blade arearotor area SOlldlty ratlo, rotor disk area

azimuth angle of rotor blade, deg

rotor rotational speed, rad/sec

natural frequency of rotating blade, rad/sec

2.1 Abbreviations

R rectangular

S swept

T tapered

A anhedra13

3.0 Apparatus

3.1 Wind Tunnel

The experimental program was conducted in the LangleyTransonic Dynamics Tunnel (TOT) shown in figure 1. The TOT isa continuous flow tunnel with a slotted test section. Thetunnel test section is 16 ft square with cropped corners andhas a cross-sectional area of 248 ft2• Freon-121 was usedas the test medium in the TOT for this program. Because of itshigh density and low speed of sound,the use of Freon-12 aidedthe attempt to match full-scale Reynolds number while achievingfull-scale Mach number values. Also, some restrictions onmodel structural design are eased, while dynamic similarity isstill maintained. The heavier test medium permits a heavierstructural design to obtain the required stiffness characteris­tics and thus eases the design and fabrication requirements ofthe model (refs. 12, 13). For this investigation, Freon-12 ata nominal density of .006 slug/ft3 was used as the testmedium.

3.2 Model Description

All the experimental blades described herein were testedon the Aeroelastic Rotor Experimental System (ARES) shown infigures 2 and 3. The ARES has a generalized helicopter fuse­lage shape enclosing the rotor controls and drive system. Itis powered by a variable frequency synchronous motor connectedto the rotor shaft through a belt-driven two-stage speedreduction system. The ARES rotor control system and pitchattitude (as) are remotely controlled from within the wind­tunnel control ·room. The ARES pitch attitude is varied by anelectrically controlled hydraulic actuator. Blade collectivepitch and lateral and longitudinal cyclic pitch are input tothe rotor through a conventional swashplate. The swashplate ismoved by three hydraulic actuators.

3.3 Description of Rotor Blades

The rotor models used in this investigation were 0.175­scale, four-bladed articulated rotors with coincident lead-lagand flapping hinges. The blade geometry was the same for bothbaseline and soft torsional rotors tested and is shown infigure 4. The blades were designed so that the tip configu­ration could be changed at the 89 percent radius. The

1Freon: Registered trademark of E.I. du Pont de Nemours &Co., Inc.

4

bl '._ IL :;c.:umentedwith four-arm strain-~age bricJ_-' Lo measure loads and deflec­tions at four blade radial stations. Flapwise (out-of-plane)moments and chordwise (in-plane) moments were measured at 26,39, 53, and 81 percent radius, while torsional moments weremeasured at 29, 37, 52, and 78 percent radius. The rotatingblade data are transferred through a 3D-channel slip-ringassembly. Rotor forces and moments are measured by a six­component strain-gage balance mounted below the pylon and drivesystem. The balance is fixed with respect to the rotor shaftand pitches with the fuselage. Fuselage forces and moments arenot sensed by the balance.

3.5 Description of Parametric Tips

Seven blade tip designs were evaluated on the baselinerotor and three of these tip designs were tested on thetorsionally soft blades. The tip designs incorporated asystemmatic variation in three geometric parameters: sweep,taper, and anhedral. These parameters were varied while alltips had the same target values of inertial properties, airfoilcontour, and twist. The magnitude of the sweep angle and taperratio chosen for ACR application were representative of

5

current design values for modern helicopter rotors. Themagnitude chosen for tip anhedral is typical of previousinvestigations. Figure 5 presents the geometry of the tipdesigns, while Table II lists the measured tip characteristicsand compares them to the design targets. Airfoil sectiontolerances were held to ±.002 in. The value of a usedthroughout this report for normalizing performance coefficientsis 0.082, based on a blade nominal chord of 3.6 inches and aradius of 56.2 inches. Table II also gives the values ofweighted solidities.

4.0 Test Methodology

The test procedure was designed to facilitate comparisonbetween tip configurations. Each rotor configuration was firsttracked and balanced in hover in the Freon test medium toremove first harmonic fixed system loads. At each forwardflight test point, the rotor rotational speed and tunnelconditions were adjusted to give the desired tip Mach numberand advance ratio, at the desired shaft angle of attack. Bladecollective pitch was changed to obtain the target rotor liftand propulsive force; and at each collective pitch setting, thecyclic pitch was used to remove rotor first-harmonic flappingwith respect to the rotor shaft. Data were then recorded ateach value of rotor task. The maximum rotor task attained ateach shaft angle of attack was determined in most cases byeither blade load limits or ARES drive system limits.Aerodynamic rotor hub tares were determined with the bladesremoved throughout the ranges of shaft angle of attack andadvance ratio investigated. Both deadweight and hub aero­dynamic tares have been removed from the data presented herein.

4.2 Accuracies

Based on replicated data points, the repeatability ofthe data for constant shaft angle of attack, control angles andadvance ratio has been estimated to be within the followinglimits:

CL-- ± 0.0025a

C~ ± 0.0004

a

Co-- ± 0.00015

a

The accuracy for angle measurements is estimated to be within±0.25°.

6

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4.3 Test Conditions

All the tip configurations shown in Figure 5 were testedon the baseline stiffness blade for the target conditions shownin Table III. The magnitudes of lift and propulsive forceparameters, tip r~ch number, and advance ratio were chosen asrepresentative of a modern utility helicopter. The ACR lowtorsional stiffness blades were used to test three tips for thetarget conditions of Table III. These tips were chosen becauseof their interesting performance and loads characteristics whentested on the baseline blades, as will be described later inthis paper. The ACR tip planforms are those of Figure 5c, f,and g: swept, swept anhedral, and swept tapered anhedral.Within the scope of this paper, the performance and loads datapresented for analysis emphasizes the target lift and shaftangle combinations of Table III, at one rotational tip Machnumber (0.65), and two advance ratios (0.35, 0.40).

5.0 Results

5.1 Rotor Performance

Fixed system forces and torque were obtained using theprocedures and limits described earlier for all tip configura­tions and conditions of Table III. Parametric performanceresults for selected conditions are presented in Figure 6. Theadvance ratios and lift coefficients were selected for presen­tation because they showed the most significant difference inrotor performance between configurations.

The effect of tip shape on baseline rotor performancefor the complete set of tips is shown in Figure 7. Eachvertical vector represents the percent reduction in torquecoefficient for a given rotor task for each tip shape. Eachhorizontal vector serves to label the added parameter for eachtip. This method of presentation of rotor performance allowsthe separation of parametric geometry effects to be easilyvisualized. For the baseline blades tested and the conditionsshown, the rotor's performance was enhanced by the addition ofanhedral to a rectangular planform and the addition of sweep tothe tapered planform. Tip taper alone enhanced rectangularrotor performance at ~ = .35 conditions but not at higherspeeds (~ = .40). The independent effect of torsional stiff­ness is not included in Figure 7.

The tip configurations which were tested on the torsion­ally soft blade were chosen because of the good performancethese tips gave the baseline blade. As shown in figure 6 how­ever, these configurations exhibited higher torque requirementsthan their baseline counterparts for the same rotor tasks.

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5.2 Rotor Loads

Blade oscillatory loads are important not only fromvibratory and fatigue considerations but also because theyprovide insight into blade loading environment and elasticdeformation trends. Torsional loads and flapwise oscillatoryloads are associated with local blade loading and twist (ref.8). Figure 8 presents 1/2 peak-to-peak flapwise loads at fourspanwise stations for all configurations tested. These oscil­latory loads are data points taken at the ~; MTIP, CLioand as values listed for each tip configuration. Theconfigurations are also ranked in Figure 8 according to theirperformance at the Cola values shown. The rank of 1 is thelowest torque required for the rotor task. Examination offigure 8 shows a significant relationship between performanceand oscillatory flapwise loads. Specifically, the configura­tions which exhibited the lowest flapwise loads were also thebest performers with regard to CQla while the poor perform­ance configurations generated the highest flapwise loads.

5.3 Elastic Twist

Spanwise distributions of blade torsional moment timehistories were converted to elastic twist distributions throughknown torsional stiffness properties of the blade, controlsystem, and UH-60 model hub. This is shown in Figure 9 for allconfigurations tested at the ~,MrIP, CLio and asvalues listed. Some interpolation of the inboard torsionalloads was occasionally necessary. Figure 9 indicates theelastic twist is configuration-dependent for each rotor taskand, as might be expected, varies with rotor environment. Notethe compressed vertical scale for the ACR configurations. Theelastic twist waveforms are comprised of several harmonics, butare dominated by the one per rev torsional moments.

The amount of azimuthal activity in the elastic twistplots is significant when compared with the integrated rotorperformance for each configuration. Elastic twist activity isnoted by the number and magnitude of waveform peaks around theazimuth. Figure 9 waveforms have been ordered according toeach configuration's torque coefficient for the rotor tasksshown with the lowest torque configuration appearing first, andthe highest torque configuration last in each case. A correla­tion between rotor performance and elastic twist is evident inthe data shown. Specifically, the configurations whichexhibited small azimuthal activity in elastic twist were thebest performers.

8

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6.0 Analysis of Results

The performance and loads data for the baseline and ACRconfigurations were examined to provide insight into themechanism by which the tip planform and torsional stiffnessparameters affected the aeroelastic environment of the rotorblades. The controlled differences between configurations wereevaluated for the fundamental changes they caused in therotor's performance and response in light of past and currentconformable design concepts.

6.1 Rotor Performance Analysis

Several of the tip shape configurations, namely thoseincorporating taper, possessed slightly less area solidity thanthe untapered tips. This alone might be expected to produceperformance differences which may be significant, depending onthe rotor task involved. Additionally, several configurationsutilized anhedral in the tip area, which can project adifferent area to the airstream than the comparable straightblade. In order to isolate and quantify these inelasticeffects, a rigid blade performance analysis was utilized.

Rotor performance characteristics and azimuthaldistributions of rotor-blade-section angle of attack werecalculated with a computer program using a strip-theoryimplementation of the equations presented in reference 14. Inthe analysis, the blade was assumed to be rigid with pitch andflap degrees of freedom but no lag degree of freedom. Therotor airfoil section characteristics used were obtained fromreference 15. Changes in section aerodynamic coefficients withangle of attack and Mach number were included in the analysis.All calculations were made by using both a uniform inflow modeland the nonuniform inflow model from reference 16. Reference16 considers a rotor load distribution which closely resemblesthat of a typical rotor and yields a non-iterative solution forthe induced velocity at any point on the rotor. The validityof this model was largely confirmed by reference 17.

For selected rotor task conditions, Figure 10 shows pre­dicted performance changes due to solidity and anhedral, forboth uniform and non-uniform inflow models. The effect oftaper alone on rotor performance trends is predicted by bothinflow models, although the magnitude of each configuration'storque coefficient is fairly well predicted only by the non­uniform inflow model. The effect of tip anhedral on rotor per­formance trends and torque magnitude is not well predicted bythis rigid blade analysis for either inflow model. Becausethis analysis does not incorporate a vortex wake even for itsnon-uniform inflow model, and because anhedral is expected tomodify the location of tip vortices in the wake (ref. 6), it is

9

not surprlslng that tip anhedral effects cannot be adequatelymodeled. The inability of a rigid blade analysis to totallyaccount for performance changes due to tip solidity variationshas been noted previously, for example, in reference 9. Itshould be noted that past efforts in correlating model-scaleand full-scale rotor performance have been less than successfulwhen blade dynamics was not adequately modeled (ref. 18).Therefore, inclusion of blade dynamics effects in predictingperformance is thought to be important for most rotorplanforms.

6.2 Blade Elastic Twist Magnitude

Past aeroelastically conformable rotor design conceptshave considered the magnitude of advancing blade elastic twistas a solution to a potentially unfavorable angle of attackenvironment (ref. 2, for example). Depending on the tip air­foil section and advancing blade Mach number, a nose-up elastictwist was thought to be desirable on the advancing side toachieve lower rotor torque and blade loads. Figure 11 presentselastic twist magnitudes on the advancing side (~ = 90 0

) foreach configuration and rotor task shown. Figure 11 also con­tains the total geometric pitch angle (which is comprised ofelastic twist, built-in twist, collective and cyclic pitchangles) at ~ = 90 0

• Both types of blade angle data are alsoranked according to their configuration's performance, withnumber 1 requiring the least torque for a given rotor. task.

As is evident from Figure 11, there is no strongcorrelation between the magnitude of each configuration'sadvancing blade elastic or total pitch angle and the perform­ance of the rotor. It is recognized that configurationperformance and loads depend on local angle of attack which isaffected by inflow distribution as well as pitch angle and thatnon-uniform inflow velocity can be very sensitive toplanformconfiguration. Nevertheless, aeroelastically conformablerotors have been designed to achieve specific azimuthalplacement of elastic twist magnitudes and the present studydoes not support the effectiveness of this design goal.

6.3 Conformable Rotor Control

Conformable rotors which effect significant bladetorsional response may generate rotor control characteristicswhich should be evaluated for their contributions to rotorstability and control (ref. 8). Throughout the test programdescribed herein, all configurations were easily controlledthrough the model actuator-swashplate system for all testconditions. The amount of control needed to achieve each rotortask was configuration-dependent. Figure 12 shows, for arepresentative rotor task, the longitudinal cyclic pitch

10

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required to remove first harmonic flapping with respect to therotor shaft for several configurations which differ in bladetorsional stiffness. The differences in elastic twist measuredfor these configurations, also included in Figure 12, is offsetby control input differences of nearly the same magnitude inorder to remove the first harmonic flapping with ,respect to therotor shaft. The differences in longitudinal cyclic pitch forthese configurations is significant not so much for controltravel considerations, but for what these angles reveal aboutthe rotor behavior for these tips and torsional stiffnesses.

An interesting phenomenon was observed in both the pitchcontrol required to trim the rotor and the rotor task achieved,in particular, the rotor propulsive force. For a given advanceratio, tip Mach number, force normal to the trimmed tip pathplane, and shaft angle of attack, the torsionally soft rotorconfigurations consistently exhibit less propulsive force.This can be seen in the performance data of Figure 6. Resolu­tion of the rotor balance forces reveals that this decrease inrotor propulsive force occurs for two primary reasons. First,the control axis for the torsionally soft rotor has tilted aftdue to the changes in longitudinal pitch mentioned above.Secondly, the rotor longitudinal force perpendicular to thecontrol axis (H-force) is greater for the torsionally softblade while the thrust was maintained constant. The controlaxis aft tilt is due to the rotor control used in the test andthe nose-down elastic twist magnitude observed. The H-forceincrease for the ACR configurations is probably due to anincrease in integrated drag loading on the advancing side ofthe azimuth. This drag increase would also manifest itself indecreased rotor efficiency, a fact which was shown earlier inthis paper for these configurations.

6.4 Blade Loading

It is well known that the radial and azimuthal distribu­tion of rotor blade loading can affect both performance andloads. The potential of the conformable rotor concept totailor these airloads has, in fact, been viewed as a key to theoptimization of rotor performance (ref. 2). Specifically, aredistribution of airloads which avoids sharp radial and azi­muthal gradients in loading has been investigated for rotorperformance improvement (ref. 19).

As previously shown, the rotor configurations describedin this paper which exhibited good performance and low vibra­tory loads generated the least activity in elastic twist aroundthe azimuth. Since several configurations provided significantaerodynamic center-elastic axis offsets, the elastic twistvariations observed may be primarily due to oscillatory tiplift. Although section pitching moment variations may add toelastic twist perturbations around the azimuth, these wouldalso be lift dependent.

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It is therefore possible that the success of thoseconfigurations which exhibited low vibratory loads andincreased performance is based on a redistribution of lifteither radially or azimuthally, or both. The clue to theapparent airload redistribution may be found in the parametercombinations which aeroelastically complement each other. Forexample, as has been shown previously in figure 7, anhedralseems to aeroelastically help a rectangular tip planform morethan it does a swept-tapered planform. Furthermore, theaddition of sweep seems to enhance the aerodynamic environmentof a tapered planform more than it does a rectangular tip forthe configurations tested.

These examples illustrate the complex relationshipbetween tip geometry and blade aeroelastic environment whichmay be largely responsible for determining rotor performance.The beneficial parameter couplings may, for example, alleviatecompressible effects and adverse loadings on a rotor blade via3-D aerodynami cs or angl e of attack adj ustments. In any case,the reduction in tip loading gradients is manifested in theelastic twist activity and results in good rotor performance.The use of an aeroelastic analysis would be necessary to qual­ify this observation, but the test results included herein sup­port the hypothesis that both performance improvements and loadreductions are generated by the same aeroelastic mechanism.

7.0 Conclusions

Based on the data obtained, and for the test conditionsand model configurations investigated, the followingconclusions have been reached:

1. Significant performance and loads differences weregenerated among the seven tip parameter variations.

2. Torsionally soft rotor (ACR) applications for the tipshapes tested resulted in poorer performance and usuallyhigher loads than for the baseline configuration.

3. The configurations which exhibited the lowest oscillatoryflapwise loads also had the best performance while theconfigurations with poor performance generated the highestflapwi se loads.

4. There does not exist a strong correlation of advancingblade elastic twist magnitude with rotor performance andflapwise loads.

5. There exists a strong correlation between azimuthalvariation of elastic twist and rotor performance andloads. The configurations which exhibited small azimuthalactivity in elastic twist were the best performers.

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9.0 References

1. Doman, Glidden S.; Tarzanin, Frank J.; and Shaw, John,Jr.: Investigation of Aeroelastically Adaptive RotorSystems. Proceedings of a Symposium on Rotor Technology,American Helicopter Society, August 1976 •

2. Blackwell, R. H.; and Merkley, D. J.: The AeroelasticallyConformable Rotor Concept. Preprint No. 78-59, 34thAnnual Forum Proceedings, American Helicopter Society, May1978.

3. Berry, John D.; and Mineck, Raymond E.: Wind-Tunnel Testof an Articulated Helicopter Rotor Model With Several TipShapes. NASA TM 80080, 1980.

4. Stroub, Robert H.; Rabbott, John P., Jr.; and Niebanck,Charles F.: Rotor Blade Tip Shape Effects on Performanceand Loads From Full-Scale Wind Tunnel Testing. Journal ofthe American Helicopter Society, Volume 24, No.5, October1979, pp. 28-35.

5. Philippe, J. J.; and Vuillet, A.: Aerodynamic Design ofAdvanced Rotors with New Tip Shapes. 39th Annual ForumProceedings, American Helicopter Society, May 1983, pp.58-71.

6. Weller, William H.: Experimental Investigation of Effectsof Blade Tip Geometry on Loads and Performance for anArticulated Rotor System. NASA TP 1303, 1979.

7. Sutton, Lawrence R.; White, Richard P., Jr.; and Marker,Robert L.: Wind-Tunnel Evaluation of an AeroelasticallyConformable Rotor. USAAVRADCOM TR-81-D-43, 1982.

8. Blackwell, R. H.; Murrill, R. J.; Yeager, W. T., Jr.; andMirick, P. H.: Wind Tunnel Evaluation of AeroelasticallyConformable Rotors. Preprint No. 80-23, 36th Annual ForumProceedings, American Helicopter Society, May 1980.

9. Yeager, William T., Jr.; and Mantay, Wayne R.:Wind-Tunnel Investigation of the Effects of Blade TipGeometry on the Interaction of Torsional Loads andPerformance for an Articulated Helicopter Rotor. NASA TP1926, 1981.

10. Yeager, William T., Jr.; and Mantay, Wayne R.: Loads andPerformance Data From a Wind-Tunnel Test of ModelArticulated Helicopter Rotors With Two Different BladeTorsional Stiffnesses. NASA TM 84573, 1983.

13

11. Blackwell, R. H.; and Fredrickson, K. C.: Wind-TunnelEvaluation of Aeroelastically Conformable Rotors.USAAVRADCOM TR-80-D-32, 1981.

12. Lee, Charles: Weight Considerations in DynamicallySimilar Model Rotor Design. SAWE Paper No. 659, May 1968.

13. Hunt, G. K.: Similarity Requirements for AeroelasticModels of Helicopter Rotors. C.P. 1245, Royal AircraftEstablishment, 1973.

14. Gessow, Alfred; and Crim, Almer D.: A Method for Studyingthe Transient Blade-Flapping Behavior of Lifting Rotors atExtreme Operating Conditions. NACA TN 3366, 1955.

15. Noonan, Kevin W.; and Bingham, Gene J.: AerodynamicCharacteristics of Three Helicopter Rotor Airfoil Sectionsat Reynolds Numbers From Model Scale to Full Scale at MachNumbers From 0.35 to 0.90. NASA TP-1701, AVRADCOM TR80-B-5, 1980.

16. Mangler, K. W.; and Squire, H~ B.: The Induced VelocityField of a Rotor. R. &M. No. 2642, British A.R.C., May1950.

17. Heyson, Harry H.; and Katzoff, S.: Induced VelocitiesNear a Lifting Rotor With Nonuniform Disk Loading. NACARep. 1319, 1957. (Supersedes NACA TN 3690 by Heyson andKatzoff and TN 3691 by Heyson.)

18. Yeager, William T., Jr.; and fv1antay, Wayne R.:Correlation of Full-Scale Helicopter Rotor Performance inAir With Model-Scale Freon Data. NASA TN D-8323, 1976.

19. Moffitt, Robert C.; and Bissell, John R.: Theory andApplication of Optimum Airloads to Rotors in Hover andForward Flight. 39th Annual Forum Proceedings, AmericanHelicopter Society, May 1982.

14

,

Table I. Model Rotor Blade Rotating NaturalFrequencies at n = 68.07 radlsec

Baseline Mode Identity ACRwin win

2.68 Flapwise 2.65Torsion 4.48

4.98 Flapwise 4.935.08 Chordwise 4.987.06 Torsion8.17 Flapwise 8.17

15

Table II. Model Rotor Blade Tip Characteristics

Tip e.g. location (in.) c.g.-a.c. (pos. e.g.Design Target Tip wei ght Tip twi st c forward)

Chordwise Spanwise (grms) (deg)1.236 2.774 71 1.35 .96R .98R LOR

Tip Configuration

Rectangular 1.30 2.75 73.1 1.27 .028 -.05 .02

Tapered 1.24 2.82 73.4 1.27 -.014 -.056 .007

Swept 1.50 2.85 73.6 1.27 .096 -.04 .019

Swept Tapered 1.31 2.94 71.4 1.27 .096 -.017 .008

Rectangular Anhedral 1.31 2.75 71.1 1.14 .028 -.05 .02

Swept Anhedral 1.48 2.96 70.4 .93 .096 -.04 .019

Swept Tapered Anhedral 1.25 3.00 71.8 1.27 .096 -.017 .008

Rotor Solidity

......en

Area sol i di ty

Thrust-weighted solidity

Torque weighted solidity

Tapered Configurations

.08127

.07905

.07793

Non-tapered Configurations

.08252

.08263

.08259

..

Table III. Target Test Conditions

MTIPCL CL CL

lJ as - as - as -a a a

.30 .65 _6.0°,_7.8° .06 _4.5°,_5.9° .08 _3.6°,_4.7° .10.68.70

.35 .65 -8.2°,-10.5° .06 _6.1°,_7.9° ~08 _4.9°,_6.3° .10.67

.40 .63 -10.6°,-13.6° .06 _8.0°,_10.3° .08 _6.4°,_8.3° .10.65

17

Top view

SCALf .ft

Figure 1.- Langley Transonic Dynamics Tunnel~

Figure 2.- Aeroelastic rotor experimental system (ARES) modelin Langley Tran~onic Dynamics Tunnel.

BALANCECENTROID

MODEL PIVOT.----L..--t----:7"--__POI NT

1

_·------8.37-----~

}+----- 4. 99--~

r7.78

~w~Figure 3.- Schematic diagram of aeroelastic rotor experimental

system. All dimensions are in feet.

NOTE, LINEAR BUILT-IN TWI ST, iii = -120

SC ID95 AIRFOIL~FLAPPINGAXIS rSC 1095 AIRFOILl 1- 0.5 TIP JOINTI J--SC 1095-R8 AIRFOIL

idd-f!;~~=====-:::':::====i=i

I~-------CENTER OFROTATION

Figure 4.- Rotor blade geometry. Blade dimensions are in inches.18

·1/4C

1.8125 in.

STA 50STA 50

l/4C

1.8125 in.

STA 50

I O.94R O.96R O. 98R'--- ---I..~ I

l/4C

STA 52.924 STA 56.224I I

O. 94R O. 96R O. 98R RI I

STA 50I

O.89RI

(a) Rectangular (c) Swept (d) Swept Tapered

RO.94R

STA 50

I

SCI095 l/4C

AIRFOIL 1.8125 in.

T.E··~~~~~~~~.95~R~I~()02:~Ar

1/4C

R

SCI095AIRFOIL 3.625 In.

O.94RI

T.E.=====

1/4C

RI

SCI095AIRFOil 3.625 in.

O.95R A

T.E======~

(e) Rectangular Anhedral (f) Swept Anhedral (g) Swept Tapered Anhedral

Figure 5.- Geometry of tips tested •. Dimensions are in inchesunless otherwise indicated.

CQ/a

.007 .~5 .0075 .0080 009 .0075 .0080 •ross .0090 .0095 .0100 .0105I I • r I I I I ( I

o SWEPT

.008 o RECTANGULAR .010 .Oll<> TAPEREDtJ. SWEPT TAPERED

.009 ~ RECTANGULAR ANHEDRAL .Oll .012Cl SWEPT ANf(DRAL

,D SWEPT TAPERED ANHEDRAL.012.010

~-CO -CO .013

-CO • ACR SWEPT

aIII ACR SWEPT AN~DRAL a a

.Oll • ACR SWEPT TAPERED ANHEDRAL .013 .014

.012 .014 .015

.013 .015 .016

.014 .016 .017

.07885

(c)lJ = .40M,-IP = .65t"L

.098--a

(b)J.l = .35~IP = .65t"L

.07872-=a

(allJ = .35

'R'TIP = .65CL

Figure 6.- Experimental rotor performance.

No

-.

SWEEP

r- RECl.....-....--,RECT

SWIEP

TAPER

RECT

TAPER

TAPER

SWEEP

TAPER +SWEEP

TAPER + TAPER +SWEEP +

~S\·._[E_? ANHEDRAL

ANHEDRAL +r SWEEP

RECTANGULARANHEDRAL

TAPER + .SIIIEE? +

6------,. ANHEDRAL

RECTANGULARANHEDRAL

ANI£DRAL +SWEEP

TAPER +S\\'EE? +ANHEDRAL

TAPER +SWEEP

RECTANGULAR ANHEDRAL +ANI-£DRAL , SWEEP

(a)lJ = .35Mr = .65

CL .07872--aCo

-.0110--a

(b) (c)lJ = .35 lJ .40r~T = .65 MT

= .65CL .098

CL .07885-- --a aCD

-.01325CD

-.0145-- --a a

Figure 7.- Parametric tip shape effect on experimental baselinerotor performance.

N.......

.26R

.39R

~UulwL:lliluJll53R.

20o

~liuluLL:lWuui26R

20

o12345618

PERFORMANCE RANK AT Cola = -0.0135

STA ST SA RA STA S R TACR

1/2P-PFLAPWISE

LOADAT

STATION,in-Ib

~LilluLL=lliwlLL53R

20o

~~=lilluLLL012345618PERFORMANCE RANK AT CD = -0.0125

STA ST SA RA T S R STAACR

1/2 P-PFLAPWISE

LOADAT

STATION,in-Ib

60 .81R4D

20oI--.l....-..L-..L-....L-....L-""",,--....I...--"--"--"-_

60

40

20oI--..L-..L-....L-....L..--'--....I...-....L.-....L......L......L.._

60

4D

20o1--"'---"'---"'---"'---"'---"'---"'---"'---"-'---

60

4D

20

ol.--.L.1

---1-2

---1-3

---1-4

---1-5-6.L..-.L.1

-8.L-9~I":-O­

PERFORMANCE RANK AT COla = -0.0110

ST STA SA RA T R S S STA SAACR ACR ACR

1/2P·PFLAPWISE

LOADAT

STATION,in-Ib

(a)~ = .35M,- = .65

'C"L.07872--a

a = -6.1 0

s

(b)= .35= .65

.098aa = -4.90

s

(c)~ = .40Mr = .65

"C"L.07885--a

a = -8.0 0

s

Figure 8.- Flapwise oscillatory blade loads versus rotorperformance.

NN

•

.. -.J

Z··r 1 r 2 I 3 I 4 I 5La

I.Z

ELASTIC .6TWIST,tIJ.

do9- pes.nost-L9 •. 6

-LZ

-1.1

-Z.•0 JO 60 90 120 150 ISO 210 240 210 m 3J) ) JO 60 90 120 150 180 210 240 210 JOO no 0 30 60 90 12'0 150 180 210 240 210 XXl nJ 0 30 60 90 120 150 180 Z10 Z40 Z10 JOO 3JO 0 30 60 90 120 ISO HlO 210 240 210 )00 330

" deg " deg " deg " deg " deg-ST STA SA RA T

Z••V\ 416 z.. 7 •. a 8 . r 9La 16

l.a 1.6

~"':: A A J----"' J ::TWIST.tIJ.

deg.pes.

-~.~ v l VV .1.Z-1.Z -1.Z I.Z.•

-1.1 ·11 .1.6 .1.6

·Z.. I I , , I , , I , I I -Z:. I I I I I I I I r I I ..·at I I I I I I I I!! ! ..:ao lO6090m~mm~mJOOm OJ0609Om~~mmmJOOm '0lO609Om~mm~mJOOm 0 lO609Om~mm~mJOOm

" deg " deg " deg " deg

R S STA SAACR ACR

(a) J.I = .35 MT = .65 CL = .07872a as = -6.1 0

• Performance rank at CD/a = -.0110'I'\)w Figure 9.- Radial distribution of elastic twist versus azimuth at .78 R.

2'·1 1 I 2 I 3 I 4 I 51.8

1.2

ELASTIC .6TWIST.M.

dog. po<.nose·uP -,6

-L/

-1.1

-2'·0 lO 60 QO 120 110 180 210 240 210 JOO llO 0 lO 60 QO 120 ISO ISO 210 240 210 3((l )30 0 lO 60 QO 120 110 180 210 240 210 JOO llO 0 lO 60 QO 120 110 180 210 240 210 JOO llO 0 lO 60 QO 120 110 ISO 210 240 210 JOO llO,. dog ,. de<] ,. de<] ,. dog ,. dog

-

ST RA STA SA T

2··

t2.V\ ..86 . 7 •. 8 8 I 9

1.8 1.8 l.616

1.2 12 I..

""",'/\ • •• A /\ ::~ "TWIST.

~~~ ~ t v~ r--- A.nose-up -.6 -.6 -1.21 "" ~/ "-

.1.2 -1.2 -2••

-1.8 -1.8 -)6

-2.4 I , I I r , , ! r I I -2.4 I r I I I I I I I ! I .... 8' ! ! I ! I ! ! I I ! '..c.SOlO60QO~~~m~mJOOm OlO60QO~~~m~mJOOm OlO60QOrn~~m~mJOOm 0 lO60QO~~~m~mJOOm

,. dog ,. de<] ,. de<] ,. de<]

S R STA SAACR ACR

(b) IJ = .35 H.r = .65 CL = .07872a as = -7.9°. Performance rank at Col a = -.01250

N.,. Figure 9.- Continued .

--, •

• •

2.41 2 3 4

1.8

1.2

ELASTIC.6

TWIST.M.

del}. pas.nose-up

-L2

-1.8

-2.40 210 240 270 300 330 0 30 W 90 120 150 180 150 180 210 240 270 JOO 330

I. del} I. dl!9

STA ST SA RA

4.82.4 85 61.8

3.6

1.2 2.4

ELASTIC .6 1.2

TWIST.M.

deg. pas.nose-up. -.6

-1.2

-1.8

-2.40 30 60 90 120 150 180

I. deg

T S R STAACR

(C) JI = .35 M,- = .65 CL = .098 as = -4.9°. Performance rank at Cola = -.0125a

N Figure 9.- Continued.U1

2.4

1.81 2 3 4

1.2

ElASTIC .6TWIST,tP.

deq. pos.nose-up -.6

-1.2

-1.8

.24' I ! I I I I I I I I , , I I , , I I I I I I I , I I I I I I I I I I I

. 0 30 60 90 12lJ I~O ISO 210 240 270 300 330 0 30 60 90 12lJ I~ 180 210 240 210 300 330 0 30 60 90 120 I~ 180 210 240 210 300 330 0t. deq t. deq t, d")

30 60 90 12lJ I~ 180 210 240 210 300 330t, deq

STA ST SA RA

752.4

1.2

1.8

-1.2

-1.8

ELASTIC .6TWIST,tP, 0

deq. pos.nose-up -.6

-2,4. _ •• _.... .~ 180 210 240 270 300 330 0 30- '60 90 120 I~ 180 210 240 2m 300 330 0 30 60 90 1~T80 210 240 210 300 330t. d") t. deq t, deq

S R T

(d) p = .40 MT = .65CL _--a .07885

I

as = _8.0°. Performance rank at CD/a = -.01350

N0\

Figure 9.- Concluded. !.!' )

• .. -"'\ •

•

...

Co Co

4 6 8x 10-4 4 6 8 x 10-4j j i , I ,

4

5

"6 """7 "'ll "-CO 8 ... "<>O'll ~l:l. "9 'll "'ll"10 ° ... "<> ......

~ "11....

12 x 10-4- THEORY} RECTANGULAR - TI£ORY} RECTANGULAR° DATA ~ DATA

- - TfEORY} TAPERED -- MORYI<> DATA l:l. DATA TAPERED

NO ANfEDRAL WjTH ANfEDRAL

(a) Uniform inflow model

Co Co

4 5 7x10-4 3 4 5 7x10-4r--Tj----r,-----i------',j r-,--.,.j---i,'---r-------;j

4

5 " ........ ....

6 .... ....

" ....7 " ....

" ....-CD " ....

8 ~"I ~....

....9 ... l:l. ....

....10 ° "~ "11 l:l.

12 x 10-4 - THEORY I RECTANGULAR - TfEORYI RECTANGULAR° DATA ~ DATA

- - THEORY I TAPERED -- TI£ORYITAPERED<> DATA l:l. DATA

NO ANfEDRAL WITH ANHEDRAL

(b) Non-uniform inflow model

Figure 10.- Comparison of rotor performance with rigid bladeanalytical model p = .35, ~L/a = .08, Mr = .65 •

27

-.01250

I I I I I ! I

2345678PERFORMANCE RANK

,...l-

I-

l-

I

5.6

5.4

5.2

TOTAL 5.0PITCH 4.8

AT 46; = 90", •deg. pas. 4.4nose-up 4.2

4.03.816

Performance rank at

I , I I I I I I I

12l45678QPERFORMANCI RANK

ST RA STA SA T S R STA SAACRACR

1.4

1.2

1.0

.8

.6

.4

.2oI-'-........"'T-.L.-~L.-J"'-T--r-

-.2-.4

-.6

-.8-1.0

-1.2-1.4

ELASTICTWIST

AT• =90",deg. pas.nose-up

-.0110

I I I ! I I I

2345678PERFORMANCE RANK

I

a

4.8

4.6

4.4TOTALPITCH 4.2

AT 4.0I =90", l.8deg. pas. l 6nose-up .

14l.23.0

erformance rank at

123456789PERFORMANCE RANK

ST STA SA RA T R S STA SAACRACR

P

II

l-

f-

l-

I-

I-

L., I

1.41.2

1.0

.8

.6

ELASTIC .4TWIST .2

AT 0• = 90",deg. pas. -.2nose-up -.4

-.6'--.8

-1.0

-1.2-1.4

-.01350a

a = -7.9°s

, I ! , I , I

1234567PERFORMANCE RANK

"C"L = .07872a

5.8

TOTAL 5.6PITCH 5.4

AT, = 90", 5.2deg. pas. 5.0nose-up 4.8

4. 6'L.L-....I..-...........L.-..L-.L-...lL...

Performance rank at

.35(b) II

I , I I I I I

1214567PERFORMANCE RANK

STA ST SA RA S R T

1.4

1.21.0

ELASTIC .8TWIST .6

AT• = 90", .4deg. pas••2nose-up 0 t-T--J-....L..."""--..L-~L.

-.2-.4

-.6

-.01250--a

as = _6.1 0

I I r I I r I I

12145678PERFORMANCE RANK

.07872

Performance rank at

CL= .35' - ­

a(a) II

ELASTICTWIST

AT "II = 90".deg. pas.nose-up

1.8 ,..

1.6-1.4­

1.2-

1.0 - 5.8.8 5.6

.6 I- TOTAL 5.4

.41- I PITCH 5.2

.2 I- AT 5.0oI-r-........-'--~L...-JI---I.--r- ,; = 90". 4.8

-.2 t- deg. pas. 4 6nose-up .

-.4 I- 4.4

-.61- 4.2-.8 4.0 L-L--L-....L..-"'-L...-JI---I.---L..

-1.0 L..! I I I I I I I

12145678PERFORMANCE RANK

STA ST SA RA T S R STAACR

~ ~( C) II = .35 - = .098 a = -4.90 (d) lJ = .40 - - .07885 a = -8 0°~ a S as· .

Figure 11.- Advancing side tip elastic twist and total p~tch versus configuration performance (r/R'=~7a ""Mr = .65).--NCO

..

•

•

NOSE-DOWNLONGITUDINAL

CYClICPITCH

AT 1/1 ='K1',deg

D BASELINE

~ ACR8.6

f8.2~

7.8

~----......7.4 IIr--

ElASTIC ~:::t...----.TWISTAT -1.0

1/1 = 'K1',deg. pas. -0. 6nose-up

~2 D D

-6 -8 -6 -8

(a) Swept-Tapered Anhedral Tip

-8

__---0

Cl BASELINE

• ACR-I. 8

8.8 -1.4 •6-

~

R.4 ELASTIC -1.0NOSE-DOWN TWISTLONGITUDINAL

8.0 AT -.6CYCLIC I/J =90",PITCH

7.6 de<]. pas. -.2AT 1/1 = 90", nose-upde<]

7.2 .2

6.8~ .6Q

I I !

-6 -8 -6

as, deg

(b) Swept Anhedral Tip

Figure 12.- Advancing blade control angle and elastic twistversus as (~ = .35 MT = .65 CLla = .08).

29

1-------------1.,•

1L262209AH76

3. Recipient's Catalog No.

5. Report DateAugust 1983

14. Army Project No.

11. Contract or Grant No.

I 2. Government Accession No.

7. Author(s)Wayne R. Mantay and William T. Yeager, Jr.

4. Title and SubtitleParametric Tip Effects for Conformable RotorApplications

1. Report No. NASA TM-85682AVRADCOM TR 83-8-4

12. Sponsoring Agency Name and AddressNational Aeronautics and Space AdministrationWashington, DC 20546

andU.S. Army Aviation Research &Development CommandSt. Louis, MO 63166

9. Performing Organization Name and AddressStructures LaboratoryAVRADCOM Research and Technology LaboratoriesNASA Langley Research Center

~-H_am~p~t~0~n~,~V-A--2-3-66-5-----------------~11~~~R~~~d~i~Co~~Technical Memorandum

6. Performing Organization Code505-42-23-05

I----------------.:...------+-=..=..:.-.:..=:.-::.:~~---__l.8. Performing Organization Report No.

1--- -1 10. Work Unit No.

15. Supplementary NotesWayne R. Mantay and William T. Yeager, Jr., Structures Laboratory, AVRADCOMResearch and Technology Laboratories

16. Abstract A research study was initiated to systematically determine the impact ofselected blade tip geometric parameters on aeroelastically conformable rotorperformance and loads characteristics. The model articulated rotors included baselineand torsionally soft blades with interchangeable tips. Seven blade tip designs wereevaluated on the baseline rotor and three tip designs were tested on the torsionallysoft blades. The designs incorporated a systemmatic variation in three geometricparameters: sweep, taper, and anhedral. The rotors were evaluated in the NASA LangleyTransonic Dynamics Tunnel at several advance ratios, lift and propulsive force values,and tip Mach numbers. Based on the test results, tip parameter variations generatedsignificant rotor performance and loads differences for both baseline and torsionallysoft blades. Azimuthal variation of elastic twist generated by the tip parametersstrongly correlated with rotor performance and loads, but the magnitude of advancingblade elastic twist did not cQrre1~te~

17. Key Words (Suggested by Authorls))HelicoptersTip ShapesConformable RotorsRotor Aeroelasti'city

18. Distribution StatementUnclassified~Unlimited

Subject Category 39

19. Security Classif. (of this report)Unclassifi.ed

20. Security Classif. lof this page)

Unclassified21. No. of Pages

3022. Price"

A03

"For sale by the National Technical Information Service, Springfield, Virginia 221 Sl

•

1

•~.