DESIGN, MANUFACTURE, AND TESTING OF AN INTEGRAL TWIST-ACTUATED ROTOR BLADE
John P. Rodgers, Research Assistant, Active Materials and Structures Laboratory, MIT 37-331, 77 Massachusetts Avenue, Cambridge, MA 02139, (617) 258-5920
Nesbitt W. Hagood, Associate Professor, Active Materials and Structures Laboratory, MIT 37-315, 77 Massachusetts Avenue, Cambridge, MA 02139, (617) 253-2738
Presented at the 8th International Conference on Adaptive Structures and Technology, Wakayama, Japan, 1997
Keywords: Helicopter, Servo-flap, Rotor blade, Active fiber composite
An active blade designed for the control of rotor vibrations and noise has been developed. Active fiber composites have been integrated with the composite spar to induce shear stresses within the spar laminate and thus a distributed twisting moment along the blade. The design of an active blade model based on a 1/6th Mach-scale Chinook CH-47D is reviewed. The design goals included ±2° of blade tip twist with a maximum of 20% added blade mass. The fabrication of the actuators and a model blade section are described. Next, results from benchtop testing of the blade section including stiffness, twist actuation, and survivability under tensile load are presented. Results are compared with model predictions and design improvements are suggested for the complete blade.
Helicopter rotor blades experience significant vibration and noise levels as a result of variations in rotor blade aerodynamic loads with blade azimuth angle. Actively controlled rotor blades are currently being investigated as a means of reducing the detrimental vibrations and noise. A reduction in the vibration and noise levels will improve pilot effectiveness and passenger comfort, and reduce maintenance and operating costs [Derham and Hagood, 1996]. Performance benefits may also be achieved through increased payload and cruise speed.
Active twist control of the rotor blades may be used to reject the aerodynamic disturbances affecting the blades. These disturbances are the most severe during transitional and forward flight, manifesting in the form of N/rev vertical hub loads. Individual blade control (IBC) of blade twist or angle of attack in real time enables the implementation of active vibration control strategies.
This study focuses on directly twisting the blade using integrated Active Fiber Composite (AFC) actuators. Anisotropic active plies may be embedded within the composite spar of the blade, as shown in figure 1, to induce shear stresses which create the twist. In addition, the dense electroceramic active material is positioned around the quarter-chord of the blade section, thus minimizing the weight penalty. The benefits of this concept are the significant actuation authority and high bandwidth. This concept is an alternative to trailing edge flap concepts which induce an aerodynamic twisting moment to deflect the blade. The integral actuation concept eliminates the need for a complex, highly efficient actuation amplification mechanism.
Currently, several other means of reducing rotor vibrations are being pursued. Integral twist-actuated rotor blades using monolithic directionally applied piezoelectric wafers are also being developed [Chen and Chopra, 1997]. Another twist-actuated rotor uses a piezo-actuated bending-torsion coupled beam to twist the blade [Bernhard and Chopra, 1997]. Most other research in IBC involves flap actuation [Prechtl and Hall, 1997, Straub and Hassan, 1996, Giurgiutiu et al, 1995].
Several previous studies have formed the background for the current research. In order to evaluate the feasibility of the integral twist actuation concept, a Rehfield single-cell composite beam model was developed [du Plessis and Hagood, 1995]. Interdigitated electrode piezoelectric fiber composite actuators were selected and used in a 1/16 scale benchtop twist demonstration. A series of characterization tests have demonstrated the integrity of the actuators in a simulated rotor stress environment [Rodgers et al, 1996]. A more advanced rotor dynamic analysis of the integral actuation scheme was later performed [Derham and Hagood, 1996]. This included a systems-level cost-benefit analysis and demonstrated the potential impact of the integral actuation concept. This paper will review the actuator development and design and will describe updated manufacturing techniques. The benchtop testing of an active spar section will also be presented.
The Active Fiber Composite (AFC) is an anisotropic, conformable, composite actuator which was developed for embedding within composite laminates. The actuator consists of continuous, aligned, electroceramic fibers in an epoxy-based matrix which is sandwiched between two layers of polyimid film which have a conductive inner surface for applying the driving electric field. More recently, the performance of the AFC system was greatly improved with the change to an Interdigitated Electrode (IDE) pattern, which orients the applied electric field along the active fibers, enabling the use of the primary piezoelectric effect [Bent and Hagood, 1995]. A diagram of the actuator is shown in figure 2.
Several variations of the material system have been investigated for this and other related studies. First, several compositions of the active fibers have been tested, all within the PZT family. All prior publications have centered on the PZT-5H composition. Recently, other compositions have been evaluated which show promise for higher induced stress capability and compressive depolarization stress tolerances. Extruded round fibers and cut square cross-section fibers have been compared. Other tests have evaluated matrix additives and electrode types [Bent, 1997].
A series of characterization tests was performed on the active fiber composites in order to evaluate performance under realistic operating conditions [Rodgers et al, 1996]. In addition to providing mechanical, electrical, and actuation properties of the actuators, the tests qualify the actuators for mechanical strain levels within the active blade. The performance of the baseline PZT-5H fiber system with copper/kapton electrodes was shown not to significantly degrade under the following conditions:
• 3000 microstrain static load
• 2.3e8 electrical fatigue cycles
• 1e7 mechanical fatigue cycles at 1250±900 microstrain
with gradual degradation at more severe levels as shown in the reference. The test articles were 0.5 inch by 3.5 inch composites, which were laminated between single plies of E-glass fabric for laminate testing.
More recently, data were collected in order to compare fiber compositions and geometry. The existing data on PZT5H round fibers was augmented with data on PZT5A round fibers, PZT5A square fibers, and PZT4 square fibers. The tests performed were free actuation, laminate actuation (induced stress), and actuation under static tensile loading (damage tolerance). Based on the characterization tests and compressive stress depolarization limits, the PZT5A composition was selected. While the square fibers performed the best, especially in damage tolerance, performance of the blade-size packs has shown that the circular fibers are currently the best choice for the material system. Manufacturing difficulties have precluded the use of square fibers.
For the matrix, samples having an undoped epoxy matrix had better performance than doped samples. For the electrodes, etched copper/kapton material was selected over screen printed electrodes for greater integrity. Additional details of the material selection testing are available in a previous publication [Rodgers et al, 1997].
The effective properties of the final active ply configuration are presented in Table I.
density (kg/m3 )
d31 * (pm/V)
d32 * (pm/V)
Table I. Active Ply Properties
These stiffnesses were determined from a Uniform Fields microelectromechanical model [Bent and Hagood, 1995] using experimental stiffness data from the composite to adjust the active fiber compliance (1.22 times bulk PZT5A compliance). The longitudinal or '1' axis is aligned with the fiber direction. The density was determined from measurements. The piezoelectric d-constants represent high field effective properties and are based on experimental measurements.
The approach used to design the active rotor blade was to select an existing rotor blade design as a baseline configuration and then modify it to incorporate active plies. The baseline configuration selected was a 1/6th Mach scale CH-47D blade developed for wind-tunnel testing at Boeing. This configuration was selected because it was an appropriate size for the anticipated testing and because of the significant experimental data and manufacturing experience available at Boeing for this model system. Additional details of the design and the models used can be found in previous work [Rodgers et al, 1997].
The model CH-47D blade, shown in figure 3, has a span of 60.619 (measured from the center of rotation) and a chord of 5.388 inches. It is designed to be used on a fully articulated hub with a single pin located at 0.15R (15% radius). The blade has built-in 12° linear twist and tapers from a VR7 airfoil at 0.85R to a VR8 airfoil at the tip. The primary structural member of the model blade is a co-cured "D" spar , while the aft fairing is added in a secondary cure. For the active blade design, several of the materials used were updated to reflect current best practices. E-glass fabric, S-glass unidirectional, and IM7 unidirectional tapes are used with a Rohacell foam core. A film adhesive is also used around the core and for secondary bonds.
The baseline model blade is Mach scaled from the Boeing CH-47D with a geometric scaling of 1:5.939 (approximately 1/6th scale). Mach scaling was selected to provide actuator performance data which would be the most applicable to the development of actuators for the full scale blade. The mass distribution and torsional stiffness properties were allowed to vary in order to achieve the design goals for twist, as described in the design requirements section below. However, the Lock number of the active model blade ends up as 99.8% of the full scale CH-47D blade. This may be attributed to the modernization of the blade materials, elimination of a original tip mass balance fixture, and the concentration of the added active material in the forward section of the airfoil which results in the elimination of some nose balance mass. The first torsional mode of the active blade remains greater than 4/rev (baseline is 5/rev).
The basic design concept of the active blade is to replace some portion of the passive composite materials in the blade spar with active plies. The shaded regions in figure 3 illustrate the regions of the spar laminate considered. The active plies may be embedded in the spar laminates and possibly in the web.
For this application, the fibers are aligned at a 45° angle to the longitudinal axis of the actuator pack, while the longitudinal axis of the actuator pack is aligned with the longitudinal axis of the blade. In the blade spar, the active plies alternate from +45° to -45°. By actuating the +45° and -45° plies opposite to each other, i.e., extending the fibers in the +45° plies while contracting them in the -45° plies, the actuators will produce a shear deformation of the spar laminate. By then coordinating all the actuators in the spar, the result is a twisting of the entire blade.
The design process was directed at achieving a number of goals for the model blade. While the model blade will only be used for hover testing at the MIT Rotor Test Stand Facility, it was designed to meet strength requirements representative of a full-scale service environment in order to demonstrate the viability of this approach. The following is a summary of the design requirements:
• ±2° tip twist
• <20% mass increase over baseline model blade with updated materials
• 50% of nominal torsional stiffness of baseline model blade
• Equivalent axial, bending, and shear stiffness of nominal baseline blade
• Section cg at quarter-chord
• Ply strain levels do not exceed allowables
• Provide passive load transfer path in plies surrounding active plies
• Counter crack propagation and delamination failure modes
• Electrically insulate active plies
In updating the materials and lay-ups, the target stiffness and inertial properties of the blade were maintained with the exception of the torsional stiffness. The target torsional stiffness was set at 50% of the nominal baseline value to enhance the twist. Results from a rotor dynamic analysis suggest that the changes would have no detrimental effects [Derham and Hagood, 1996].
The variables in the design include the amount and placement of the active material and the passive ply lay-up of the spar. Other details of the design that are considered include manufacturability, actuator pack design, power distribution, and interconnections with the hub.
Candidate configurations for the active blade were developed by adding various numbers of active composite plies to the baseline spar laminate. For each configuration, the passive lay-up for the blade was then modified to match the target stiffness values at a representative section. Leading edge mass was added to properly locate the center of gravity of the section at the quarter-chord.
The static loads used in this analysis were estimated from the design loads for full-scale blades, including the assumption of 20% rotational overspeed in addition to a 1.5 ultimate load factor. The fatigue loads were based on analytical predictions from Boeing’s TECH-01 analysis [Shultz et al, 1994] for high-speed forward flight. The centrifugal loading includes the contribution of the added blade mass from the distributed actuators.
A single-cell composite beam model with active layers was used to evaluate prospective spar designs [du Plessis and Hagood, 1995]. The model was developed with a Rehfield-type framework including one warping degree of freedom. The results of preliminary analyses showed that two to four active plies could be integrated into the spar and web laminates. Active plies were excluded from the nose of the spar to avoid the high curvature and high strain levels resulting from chordwise bending.
In order to ensure the structural integrity of the blade section, maintain actuation capability after loading, and insulate the actuator electrodes from each other, passive plies are required to surround the active plies. Characterization studies have shown that passive composite plies aligned with and adjacent to active plies increase the effective strength of the active plies by providing a load transfer path around cracks in the ceramic fibers, and may also inhibit the propagation of cracks in the active plies.
Four designs were developed with differing levels of actuator authority in the blade. Each design attempted to meet each of the design constraints, while maximizing the induced twist. The four types are: 2 active plies in upper and lower spar walls, 2 active plies in upper and lower spar walls and the web, 3 active plies in the upper and lower spar walls, and finally 4 active plies in the upper and lower spar walls. Figure 4 depicts the spar lay-up for each case.
The model-predicted induced moment and twist are compared for each case in figures 5 and 6 using a nominal 1100 microstrain free actuation for the active plies. Figure 6 also includes dashed lines representing the initial design constraints. Note that the mass is stated relative to the baseline model blade with updated materials.
The induced twist moves toward an asymptote with increased active material. This occurs as the ratio of active mass to passive structural mass increases. The actuation of the blade spar approaches the induced strain limit of the active material.
In general, the 2-ply design did not have sufficient authority to meet the desired twist. The addition of active material to the web lay-up was found to be effective but was limited by manufacturability, increased torsional stiffness, and center-of-gravity related weight penalties. The 4-ply design was most effective, but involved excessive spar wall thickness. The 3-ply design nearly meets all of the requirements. One drawback of this design is the unbalanced spar laminate which results in a slight extension/twist-coupling blade which should not affect blade performance.
The active plies which are distributed in the upper and lower spar laminates are divided into actuator packs in order to simplify manufacturing and increase reliability. The fact that the packs will be independently wired creates an additional design constraint. The total length of each active ply is 1.047 m of which 6 are required. Segmenting each layer into 7 pieces results in a repeated length of about 15 cm and a total of 42 packs. This design maintains a pack size which is manufacturable while increasing the number of leads and internal connections to a reasonable level.
Another component of the active fiber composite is the kapton electrodes. The electrode layers sandwich the active fibers and matrix material in order to deliver the electric field to the fibers. The interdigitated electrode pattern was designed using the same rules as in previous actuator designs [Bent and Hagood, 1995]. Figure 7 illustrates the electrode pattern and pack size.
The electrical connections from the packs to the leads supplying the power are placed along the web of the blade. This arrangement allows direct access to the aft edges of all of the packs after the spar is cured. Individual connections to each electrode flap require 84 leads along the web of the blade. A lightweight, low volume solution for distributing the power to the packs is a flexible circuit. The flex circuit consists of multiple parallel lines of copper arranged in layers with kapton insulating layers in between. Each line of the flex circuit terminates with a solder pad to be connected to a particular electrode flap as shown in figure 8.
At the inboard end of the web, the flex circuit exits the blade and terminates at the hub. A printed circuit board serves as a matrix connector, connecting individual lines from the flex circuit to a total of 5 source wires. This enables signals of opposite phase to be used to drive the +45° plies and -45°. In addition, actuation of the upper and lower spar laminates is independently controlled.
The final active blade design features three active plies in the upper and lower spar laminates between 0.27R and 0.95R. The active plies are divided into 42 independently wired actuators. The predicted stiffness and mass of the blade are within 10% of the target values. The ultimate design strains predicted for the active plies are 5100 microstrain tension and 3600 microstrain compression, both of which occur at blade station 0.337R. The peak actuator strains predicted for steady hover testing are 819 microstrain tension and 173 microstrain compression in the extremes. If a load factor of 2 is applied to all loads except CF to account for unsteady effects, the peak actuator strains are 1277 microstrain tension and 685 microstrain compression. Under these conditions, minimal damage accumulation is expected for the active composites. The greatest risk for the hover testing of the model blade is localized compressive stress depolarization of the active fibers (PZT5A depolarization becomes significant above 450 microstrain). Note that the strain predictions do not consider the effects of actuation, thermal prestress, and applied electric field which would provide a more accurate gauge of the state of polarization in the piezoceramic.
Both the active plies and the integral blade are manufactured in the Active Materials and Structures Laboratory at MIT. The following subsections provide an overview of the manufacture of the active fiber composites and the integral blade.
The general concept for the manufacture of AFC’s is to form a composite with a single layer of aligned active fibers in epoxy matrix and interdigitated electrodes. Registration of the top and bottom electrode patterns and compression of the lamina are required to maximize performance. In addition, void content must be minimized to reduce dielectric breakdown risks.
The packs are produced using a hot press and vacuum combination. The upper and lower electrodes are aligned and taped to the top and bottom plates, respectively. A 5 mm wide strip of ±45° E-glass fabric is placed along the inner edge of the electrode rails of the bottom electrode. The fabric fills the dead area under the rails in order to provide a more structural interface with adjoining passive plies as well as reduced weight. The fibers are then placed within a kapton tape mold by hand, and are adjusted to achieve roughly uniform spacing and proper alignment. This is illustrated in figure 9. The matrix consisting of Shell Epon 9405/9470 epoxy and BYK A530 air release agent (0.5%) is added to the fibers in the mold. The fibers are adjusted to form a uniformly distributed, single layer with no crossed-over fibers.
Next, the top plate is suspended 5 mm above the fiber and epoxy on the cure plate using alignment pins to ensure registration. A vacuum is then pulled (2 mm Hg) on the sample to degas the matrix and prevent any bubbles from being trapped by the top electrode. A photo of this vacuum/hot press mechanism is shown in figure 10. The support beams are then lowered while the vacuum is applied until the top electrode comes in contact with the fibers and matrix below. Once this has occurred, the support beams are used to apply downward pressure (100 kPa) to the top cover as the vacuum is released, allowing any voids within the composite to collapse. The matrix is cured for 3 hours at 120°C. The cure plate allows for 4 packs to be manufactured simultaneously.
Each pack is subjected to a qualification test prior to fabrication of the model blade. The actuator is then poled for 20 minutes at 80°C and 4000 V in air. Next it is cycled to a representative work cycle of -1200 V to 2800 V for an analysis of induced strain capability which must surpass a peak-to-peak longitudinal strain level of 1100 microstrain at 10 Hz.
The model blade spar is manufactured using procedures developed by Boeing Helicopters. In general, the spar consists of an instrumented foam core wrapped with composite laminae to achieve the designed lay-up for the nose, upper and lower spar walls, and the web. The active plies are incorporated into this lay-up procedure such that the electrode flaps fold onto the outer surface of the web. The entire spar assembly is cured at 120°C for 90 minutes in a two-part mold which has a filler block in the fairing portion. The spar section and flex circuit are shown in figure 11.
BLADE SECTION TEST RESULTS
An active blade section was tested in order to evaluate the design and manufacture, and to allow for improvements in the full model blade. The blade section has 12 actuator packs incorporated in the typical model blade section at midspan. The total length of the section was 0.60 m from the root pin. The active length extended from 0.27R to 0.46R, or 0.30 m. Actuation tests also allow for comparisons with the model, but more importantly demonstrate the effectiveness of the design and manufacturing process. An aluminum tip fixture was cured into the outboard end of the spar section to interface with the grips on the tensile testing machine.
The twist capability of the active blade section was measured in a benchtop test. One end of the blade was clamped, while the other was free. A pair of laser displacement sensors was used to measure the twist angle at the blade tip as shown in figure 12.
The active plies were driven with a 20 Hz sinusoid having a 600 V DC offset and an amplitude of 2550 V peak-to-peak. The resulting twist performance is plotted in figure 13. The estimated constant twist rate over the active length of the spar is plotted in a hysteresis loop. The total tip twist was 0.38° peak-to-peak for the applied voltage cycle with an average twist rate of 1.26°/m.
The intended voltage cycle was 800 VDC + 4000 Vpp for the actuation testing. However, arcing occurred at about 2600 Vpp between 2 packs in the lower spar laminate. The cause was most likely inadequate insulation between the conductive electrode rails at the ends of the packs, which will be corrected in the next blade. The glass/epoxy plies which separate the active plies may not have provided sufficient insulation between the edges of +45° packs and the -45° packs. The large current capability (about 1 Amp) and lack of an automatic shut down on the amplifier contributed to the extent of the damage which resulted from the breakdown.
For this experiment, the modified Rehfield model predicts roughly 1.69°/m assuming a free strain of 460 microstrain for the active plies in the spar. The estimated pack free strain reflects the effects of reduced voltage and increased frequency. Scaling the data for the full active length and a pack free strain of 1100 microstrain yields a prediction of 2.99° of peak-to-peak twist capability.
Passive stiffness measurements were used to verify that the manufactured spar properties were within acceptable design limits. The measured axial stiffness was 6.5 MN while the torsional stiffness was 115 Nm2. Model predictions for the spar are 5.7 MN for axial and between 91.4 and 97.7 Nm2 for the torsional stiffness. Thus the spar section is stiffer than predicted by 14% in tension and 18% in torsion.
In another test, the performance of one AFC pack was monitored as a function of applied tensile load. Negligible degradation in performance was found in three repeated cycles to a static tensile strain level of 2400 microstrain or approximately 3000 lbs. (13.3 kN). The repeated loading to 3000 lbs., the full centripetal load at the blade root expected in the hover tests, also increases confidence in the blade manufacturing process.
A 1/6th Mach scale CH-47D model blade has been designed with active twist capability. The design and manufacture of the active model blade has been demonstrated in a benchtop spar section test. A twist rate of 1.26 deg/m was measured in a preliminary test. This was roughly 75% of the predicted value for that level of actuation. Dielectric breakdown damage in the spar resulting from insufficient insulation between packs limited further actuation testing. The next prototype will incorporate improved insulation and higher performance requirements for the packs in order to achieve the design objectives for the integral blade. The manufacture of the full Mach scale integral blade and a fatigue test section are currently in progress.
This work was supported by DARPA under the Smart Structures for Rotor Control contract with Dr. Spencer Wu of AFOSR and Dr. Robert Crowe of DARPA as the technical contract monitors. Additional support was received from the ARO with Gary Anderson as the technical contract monitor. The authors acknowledge Douglas B. Weems of Boeing Helicopters for contributions to the blade design. The project was also supported by Robert Derham and Richard Bussom at Boeing Helicopters. Special thanks to Aaron Bent, Alex Pizzochero, Seward Pulitzer, Paul Bauer, Eric Prechtl, Sang Joon Shin, Jaymee Johnson, and Jaco du Plessis from MIT.
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