Community Research and Development Information Service - CORDIS

FP7

ACcTIOM Report Summary

Project ID: 298187
Funded under: FP7-JTI
Country: France

Final Report Summary - ACCTIOM (Advanced Pylon Noise Reduction Design and Characterization through flight worthy PIV)

Executive Summary:
Clean Sky, the most ambitious European aerospace research programme ever, is focused on developing future concepts for aircraft with substantially reduced environmental impact. The CROR engine with uncased blades on two stages that rotate in opposite directions could cut fuel consumption and associated carbon dioxide emissions by 30 %. Despite its promise, previous development efforts have been hampered by higher noise levels.
EU-funded scientists from the ISAE-SUPAERO and Aéroconseil consortium, in association with Airbus, have been working on the project 'Advanced pylon noise reduction design and characterization through flight worthy PIV' (ACcTIOM). They have addressed this challenge by inventing new active flow control strategies aiming at minimising CROR-induced noise through a combination of aerodynamic optimisation of the propeller pylon shape and the development of an innovative active flow control system to erase the pylon wake before it interacts with the CROR blades.
During the first phase of the ACcTIOM project, the active flow control system, a combination of scooping/blowing strategies, was designed and optimised via a coupled approach combining exhaustive Computational Fluid Dynamics (Reynolds-averaged Navier-Stokes) simulations and test bench experiments. The wind tunnel (WT) test model of the CROR pylon, including equipment, instrumentation and equipped with the so-developed embedded prototype of the active flow control system, was finalised as well.
Exhaustive series of WT tests were completed for the operating validation of the WT model of the CROR pylon equipped with its embedded flow control system, on-board instrumentation, and data acquisition and control systems. The WT tests were comprised of wall pressure coefficient taps, transverse profiles of total pressure at various streamwise and spanwise positions in the wake of the pylon and on stereoscopic Particle Image Velocimetry (3C-PIV) transverse plane measurements at various streamwise positions in the wake of the pylon. These WT tests confirmed the strong efficiency of the developed embedded active flow control system in erasing the pylon wake, when operated in the region of optimal flow control parameters numerically and experimentally identified by the consortium scientists in charge of ACcTIOM. They have also highlighted the robustness of the flow control system in mitigating the wake of the pylon despite moderate variations of the flow conditions. This embedded active flow control system acts as an aerodynamic stealth system.
In the second phase of the ACcTIOM project, the objectives of the consortium scientists have been to develop advanced optical methodologies, based on vibration-controlled 3C-PIV, able to be in-flight operated and dedicated to the validation of the efficiency of the above mentioned CROR pylon design and associated active flow control system in erasing the pylon wake on the CROR-propelled Flying Test Bench (FTB). To this avail they have first developed numerical models of the expected vibratory environment inside the cabin of the FTB. These models have permitted to design experimental test benches, hereafter denoted Vibrational Environment Simulator or VES, able to reproduce, in laboratory, the vibrational environment of the FTB and its influence on the optical misalignment of the different 3C-PIV subsystems and resulting 3C-PIV measurement issues. Further, the team has defined the hybrid, passive/active, vibration control strategy and required equipment for the design of a Vibration Correction Methodology (VCM) dedicated to the implementation and confident operation of in-flight 3C-PIV. The VCM will attenuate vibrations experienced by the 3C-PIV subsystems under VES influence during testing.
ACcTIOM technologies for active reduction of noise associated with the innovative CROR engine design will speed up certification and commercialisation of more energy-efficient aircraft. In-flight use of advanced vibration-controlled 3C-PIV will enhance understanding of mechanisms in other airframe elements as well. The outcomes should significantly reduce the environmental impact of air travel.

Project Context and Objectives:
Given the increasingly stringent regulations governing the air transport sector, driven by environmental and economic concerns, efforts to reduce air traffic fuel consumption and to limit pollutant and noise emissions encourage both aircraft and powerplant manufacturers to develop innovative solutions for future aircrafts. Clean Sky, the most ambitious European aerospace research programme ever, is focused on developing future concepts for aircraft with substantially reduced environmental impact. In this context, Counter Rotating Open Rotor (CROR) propeller technology is appearing as a promising solution, as it could cut fuel consumption and associated carbon dioxide emissions by 30 %. However the promotion of this not yet fully mature technology, although it was initiated in the early 80’s, raises numerous issues and imposes to address significant scientific and technological challenges.

Amongst these challenges, the interaction of the wake of the CROR engine pylon with the counter rotating blades, positioned downstream of the pylon (pusher configuration), promote strong total pressure fluctuations on the powerplant structure. The latter are responsible for both airframe noise and vibrations that penalize aircraft certification. ‘Erasing’ the pylon wake such as to recover a strictly uniform flow upstream of the rotating blades would suppress this major source of airframe noise and vibrations. It therefore appears as a promising flow control strategy for the certification of CROR propeller-equipped aircrafts.

The 'Advanced pylon noise reduction design and characterization through flight worthy PIV' project, also referred to as ACcTIOM project, is put forward in that context. This collaborative research project, funded by the European Commission through the Cleansky/SFWA (Smart Fixed Wing Aircraft) Demonstrator program, is led in partnership with ISAE, Aéroconseil and Airbus. It first aims to design an active flow control system embedded in the optimally-shaped CROR pylon, and dedicated to the reduction of noise emission through the active erasing of the pylon wake. Second, it aims to develop an advanced experimental methodology, based on vibration-controlled stereoscopic Particle Image Velocimetry (3C-PIV), able to be flight-operated and that will serve the validation of the above mentioned active flow control system when operated on the Flying Test Bench.

The main objectives of the ACcTIOM project are listed below.

Objectives 1 to 5 refer to the first major objective of the ACcTIOM, related to the design, development, conception and validation of the optimized CROR pylon and of its embedded active flow control system dedicated to the annihilation of the pylon wake.
1/ On the basis of the providing by Airbus of the reference geometry of the CROR-propeller pylon, to define an optimum design, via Computational Fluid Dynamics (CFD) simulation-based optimization process, of the aft part of the CROR engine pylon and to develop, via exhaustive CFD simulations, an innovative active flow control system embedded in the pylon and able to erase the pylon wake (WP2). This optimized and controlled pylon should comply with Airbus acoustic requirements (WP2).
2/ To develop and manufacture, on the basis of WP2 specifications, a rigid WT model representative of the optimized pylon shape and equipped with the active flow control system designed in WP2. This WT model will be operated in the ISAE-SUPAERO S4 WT. As such it will take into account all the constraints of this WT (limited size and flow velocity, optical accesses for PIV measurements, scooping mass flow rate limitations). These tasks are part of WP3.
3/ To complete an exhaustive series of WT tests for the operating validation of the flow control system-equipped CROR pylon WT model, on-board instrumentation, and data acquisition and control systems. The WT tests will be comprised of wall pressure coefficient taps, total pressure multi-probe and advanced stereoscopic Particle Image Velocimetry (3C-PIV) measurements in the wake of the CROR pylon. They should highlight the efficiency and robustness of the developed embedded active flow control system in erasing the CROR pylon wake upstream of the actual location of the CROR blades (WP3).
4/ To post process and to fulfill the physical analysis of the WT test results.
5/ To conclude on the efficiency and the robustness of the embedded active flow control system in annihilating the CROR pylon wake before it interacts with the CROR blades, and to provide guidelines for the transposition of the WT results to flight conditions. This closes WP3.

The second major objective of the ACcTIOM project is to develop an advanced experimental methodology, based on vibration-controlled stereoscopic Particle Image Velocimetry (3C-PIV), able to be flight-operated inside the Flying Test Bench in order to validate the efficiency of the active flow control system. Objectives 6 to 10 refer to the latter.
6/ In the context of WP4, to develop and validate a Vibrating Environment Simulator (VES) able to reproduce the vibration spectrum representative of vibration levels experienced in the implementation zone of the stereoscopic Particle Image Velocimetry (3C-PIV) system to be flight-operated in the cabin of the CROR powerplant equipped FTB.
7/ Once the VES designed, to manufacture it and to conduct a series of tests dedicated to the validation of its ability in imposing the above-mentioned vibration spectrum on the different PIV subsystems (cameras, laser sheet), and in determining the envelop of acceptable vs. prohibitive misalignments of the PIV subsystems along each of the 6 misalignment axes.
8/ To quantify the influence of displacements and misalignments imposed by the vibratory environment in the FTB cabin, simulated via the VES, on 3C-PIV measurement errors, through the determination of the 6-axes envelop of acceptable vs. prohibitive misalignments.
9/ To complete the definition and the conception of a vibration correction methodology (VCM) able to suppress the above-mentioned vibrations at the location of the PIV optical subsystems. The final goal is to achieve accurate and usable in-flight PIV measurements dedicated to the in-flight validation of the active flow control system developed in the context of WP2 and WP3.
10/ In the context of WP5, to propose and to detail various strategies, relying on the VCM, that will permit to implement and operate 3C-PIV measurements in the FTB in a confident way.

Project Results:
First objective of the ACcTIOM project: to design, develop, manufacture and validate an active flow control system embedded in an optimally-shaped CROR pylon, dedicated to the reduction of CROR noise emission through the active erasing of the pylon wake (WP2 and WP3).

WP2 and WP3 synthetic description of the actions and results obtained

Pylon and active flow control system design optimization (WP2)

• 2D aeroline optimization of the pylon and scooping/blowing slot shapes and relative positions
• study of prospective concepts of flow control strategies (trailing edge blowing, blowing temperature effects)
On the basis of the optimized 2D shape of the pylon equipped with blowing and scooping slots, designed during the first period of the ACcTIOM project and fully detailed in deliverable D2.1 and periodic report 1, WP2 has been complemented with the following tasks during period 2:
• exhaustive characterization, through 600 2D-RANS computations per flow condition, of the pylon wake response to flow control parameters, through the evolution of the acoustic optimization criterion, the C2 acoustic criterion, as a function of the scooping and blowing mass flow rates (MFR), hereafter denoted as Qscoop and Qblow respectively, for both {M∞ = 0.23 (ISA conditions), Re/m = 4.3 × 10^6, alpha = 0deg}, corresponding to the in-flight operational point, and {M∞ = 0.12 (ISA conditions), Re/m = 2.7 × 10^6, alpha = 0deg}, corresponding to the wind tunnel operational point, flow conditions
• on this basis, determination of the [Qscoop,Qblow] zone to focus on for the WT tests campaigns. In particular, extraction of the acoustic efficiency surface of the active flow control system in mitigating the wake of the pylon as a function of these flow control parameters

3D design, conception and manufacturing of the active WT model of the pylon (WP3)

• design of a modular concept, composed of the association of several identical active modules, ended by non equipped 3D parts (i.e. pylon reference profiles without flow control system) and end-plates
• ability to study the influence of the mixing of uncontrolled and controlled zones of the pylon wake, downstream of the pylon trailing edge
• taking into account of Airbus request to reconsider the flow control system in order to, ultimately, be able to be operated in the presence of a slipped flow (not considered for the call for proposal). This has led to the re design of the 3D active modules such as to ensure to separately operate both sides of both blowing and scooping sub-modules (see definition of active module, sub-module and flow control system in periodic report 1)
• definition of the parameters to be considered for the control of both scooping and blowing mass flow rates (inner and outer cavity pressure taps, mass flowmeters, pressure regulators, number and diameter of the holes constituting the shocked plates for the blowing mass flow rate control, pressure taps for the validation of both scooping and blowing homogeneity along the span of the active control modules, etc.)
• definition of WT model equipment, in terms of pressure taps (Kp lines, cavity flow control, flow homogeneity along the span of the scooping and blowing slots) and temperature sensors
• manufacturing of the WT model subcontracted to Rossi Aerospace (reasons for the delay in the manufacturing detailed in periodic report 1)
• receipt of the WT model achieved in 2 consecutive phases. 1/ receipt of a single active module, for validation of the design and manufacturing, in Feb.2013. 2/ receipt of the full WT model in May. 2013
• equipment of the WT model subcontracted to Ramonville Productique (June-July 2013) and supported with Airbus funding
• assembly of the WT model, surface finishing, checking of pressure taps operation, electronic connection and data acquisition system validation, etc.: Dec. 2013 (WT availability required)

Pre validation tests of the active module (WP3)

• validation of the blowing mass flow rate homogeneity along the span of the blowing slots validated in Feb. 2013, on the basis of the active module taken delivery in Feb. 2013
• 2C-PIV conducted at the exit and downstream of the blowing slots
• planar laser tomoscopy (same location as PIV)
• Results: blowing homogeneity along the span validated. Design of the active modules validated

WT tests at M∞ = 0.12 (WP3)

• four tests campaigns conducted up to the end of Feb. 2014
• pre tests for the operating validation of all the pressure taps and temperature sensors, data acquisition systems, pressure regulators, scooping and blowing mass flow rate temporal stability and accuracy
• Kp profiles test campaigns for various uncontrolled and controlled configurations
• Pt rake profiles at 2 transverse X-positions downstream of the pylon trailing edge (X/C = 7.5% and X/C = 15%) and along the span of the active module, for various uncontrolled and controlled configurations
• 3C-PIV test campaigns: transverse flow fields at 3 positions X/C downstream of the pylon trailing edge (1%, 3.75% and 7.5%), for the most pertinent uncontrolled and controlled configurations (determined from the Pt rake profiles tests campaigns). 7 configurations investigated. Extraction of instantaneous velocity fields, time-averaged velocities, Reynolds tensor, cross-correlations, skewness, flatness, etc.

WT test results post processing and analysis – WP3 completion

• WT test results post processing and physical analysis
• exhaustive determination of the operating surface of the active flow control system
• validation of the efficiency of the active flow control system in annihilating the wake of the pylon, thus satisfying the acoustic objective of the CROR pylon
• highlighting of the robustness of the active flow control system in erasing the wake of the pylon under moderate variations of the flow control parameters.

WP2 and WP3 detailed description of the actions and results obtained

I. Definition of the Active Flow Control System

Considering integration and manufacturing constraints on the future CROR-propelled aircraft, as well as robustness in a given flight domain with a low drag impact, the flow control strategy to be developed for the mitigation of the CROR pylon wake is first identified. It will consist in an innovative scooping/blowing coupled active flow control system. The scooping device should first ingest the already developed boundary layer, while the blowing device should then re energize the close-to-wall flow in order to mitigate the boundary layer that develops downstream of the scooping device up to the trailing edge of the pylon. Indeed, previous studies involving the control of the pylon wake through the sole use of a blowing device revealed a moderate impact on the wake mitigation, not sufficient to comply with the required acoustic performance of the CROR powerplant.
Due to pylon mechanical constraints, the implementation area of the flow control system has to be restricted to the aft part of the pylon, beyond a downstream-to-stagnation point position equal to 75% of the full-scale pylon chord C (see Fig. 1). Furthermore, the active flow control system has to be optimized for standard takeoff and landing flight conditions, i.e. for a Mach number M∞ = 0.23 at ISA conditions, and a Reynolds number per meter Re/m = 4.3×10^6. Here the upstream airflow is considered with no incidence relative to the pylon (alpha=0deg). The flight or airflow operating conditions will hereafter be provided in the form {M∞, Re/m, alpha}.

A. Criterion for the Design and Optimization Process

The design and optimization process of the flow control system relies, amongst others, on the minimization of an integral criterion, hereafter referred as C2 criterion. The optimization procedure also considered the minimization of both the pylon drag and the scooping and blowing mass flow rates, while taking into account WT model and full-scale, aircraft-dedicated pylon integration constraints, but with a lower relative weight in the optimization process. C2 quantifies the distortion of normalized total pressure profiles in the pylon wake Pt / Pt0 (where Pt0 and Pt stand for the free stream total pressure and total pressure in the pylon wake respectively), at a given transverse station, equal to X/C=7.5%, downstream of the pylon trailing edge. Note that this transverse position is located at mid distance from the trailing edge of the pylon to the first stage of the pylon counter rotating blades. It is given by C2 = S.A.Ptmax.(Ptmax-Ptmin)/Pt0^2, where S stands for the pylon wake span, A its area relative to unity, Ptmax and Ptmin the maximum and minimum total pressure in the wake respectively.
C2 should range below a target value, hereafter referred as C2t=0.25×10^(-10), when computed at the location of the first blade stage of the future CROR powerplant. Here it is computed at midway between the pylon trailing edge and the first blade stage of the CROR. This imposes a more critical and binding optimization process.

B. Definition of the WT Model 2D Shape Baseline

Due to the limited size of the WT test section (3m x 2m elliptical test section, 2.4m long, see next section), the global scale of the WT model has to be limited to a 1:2 scale of the actual full-scale pylon, based on its chord C. Indeed, the first half of the Airbus concept-based full-scale pylon geometry was truncated and a new leading edge has been designed between 50% and 65% of chord C (Fig. 1). The aft part of the geometry, located beyond 65% of chord C was kept identical to the original geometry. In particular, this new C/2 chord-length baseline geometry depicts the same boat tail angle and trailing edge as the original pylon geometry. Within this scope the geometrical coherence along the last 35% of full-scale chord ensures integration similarity between the WT model and the full-scale aircraft-dedicated aft pylon.

C. Flow Control Definition and Optimization Process

The definition and optimization process of the flow control system was conducted on the 2D WT geometry illustrated in Fig. 1, with the following airflow conditions: {M∞ = 0.23, Re/m = 4.3 × 10^6, alpha=0deg}. It relied on 2D RANS computations conducted with the CFD++ Metacomp Tech finite volume software. Taking advantage of the planar symmetry of the geometry, only half of the geometry was modeled. The 2D computational domain was composed of 250.000 hexahedral cells. The mesh was highly refined close to the WT model walls, ensuring y+ values below 1. A characteristic boundary condition, fixing free-stream Mach number M∞ and static conditions was imposed to the outer boundaries of the domain. The surface of the WT model was modeled as adiabatic no-slip surface. The spatial discretization was achieved using second-order upwind scheme method. An implicit second-order algorithm was employed for the pressure-velocity coupling and a Spalart-Allmaras turbulence model provided closure for the Reynolds stress tensor in the compressible RANS equations.
Several input parameters were considered for the optimization process: streamwise and normal-to-wall positions of the scooping and blowing zones, height of the scooping and blowing slots (note that the shape of the scooping slot leading edge was designed with respect to state of the art), boat tail angle and trailing edge thickness, scooping and blowing mass flow rates. Part of the scooping and blowing inner ducts were modeled and ended on a mass flow inlet-type boundary condition, where the normalized scooping and blowing mass flow rates per span meter, hereafter referred as Qscoop and Qblow, were imposed respectively (Qscoop and Qblow correspond to the scooping and blowing mass flow rate (MFR) per meter along the span of the WT model respectively, normalized by rho∞.U∞.F. Here rho∞ and U∞ are the free stream density and velocity respectively. F stands for the frontal area of a 1m span-extruded WT model section). The optimum configuration was then determined by minimizing the following criteria, thanks to an iterative approach: C2 criterion, scooping and blowing mass flow rates, drag impact, flow control system integration constraints for both the WT model and the full-scale pylon. Fig. 2 depicts the resulting optimized configuration of the flow control system-equipped model geometry.
At last the inner air ducts and cavities of both scooping and blowing devices (see deliverable D2.1 and periodic report 1) were partially optimized via 3D RANS computations in order to minimize the pressure losses, from the scooping and blowing supply valves (vacuum valves and compressed air valves respectively) up to the scooping and blowing slots. This inner design optimization was completed thanks to pre series of experiments in the WT. The latter allowed guaranteeing the spatial and temporal uniformity of the flow at both the inlet and outlet of the scooping and blowing devices respectively.

The aerodynamic shapes of the WT geometry and of the scooping and blowing devices obtained from the optimization process are illustrated in Fig. 2. The location of these devices has been determined to 80% of chord C for the scooping inlet slot and 95% of chord C for the blowing outlet slot.

Considering the {M= 0.23, Re/m = 4.3 ×10^6, alpha=0deg} flow conditions, the optimum scooping MFR (required to ingest the boundary layer) is 3.4kg/s/m (global mass flow rate for both sides of the WTT geometry). The optimum blowing MFR value associated with the above-mentioned optimum scooping MFR is 0.4kg/s/m. It is interesting to notice the high ratio between blowing and scooping optimum MFR. This tends to indicate that a flow control system that would solely use the energy of the blowing MFR in order to generate the scooping MFR, through a Venturi effect apparatus for instance would probably not be sufficiently efficient to ensure the target scooping MFR. One should also notice the high sensitivity of the C2 criterion to blowing MFR variations. As a consequence, the target maximum value C2t may not be guaranteed beyond blowing MFR variation the order of ±4g/s/m around the optimum value. The operational MFR has thus to be precisely managed.

Considering the optimized geometry and the impact of optimum scooping and blowing MFR on the resulting WTT geometry wake topology, in terms of normalized total pressure transverse profiles at X/C = 7.5% and X/C = 15%, the actuation of the flow control system with the optimized parameters (size, positioning, design of the scooping and blowing devices, MFR, etc.) promotes a dramatic mitigation of the distortion of the wake. It results to very low C2 values, much lower than the target value C2t (4 orders of magnitude lower).

This promising results tend to indicate that the proposed aft part design of the CROR-propeller pylon associated with the innovative scooping/blowing active flow control system developed during this WP will be efficient in complying with the Airbus acoustic requirements.

Once this optimum configuration was determined for the actual flight conditions {M∞ = 0.23, Re/m = 4.3 × 10^6, alpha=0deg}, the resulting optimized geometry was fixed and optimum scooping and blowing mass flow rates were then estimated for the WT airflow conditions {M∞ = 0.12, Re/m = 2.7 × 10^6, alpha=0deg}. To this avail, complementary computations were achieved by varying the normalized scooping mass flow rate Qscoop from 0 to 0.083/m, discretized into 30 uniformly distributed values, and the blowing mass flow rate Qblow from 0 to 0.014/m, discretized into 20 uniformly distributed values. This amounted to a total of 600 computations, solved with the StarCCM+ CD-Adapco finite volume solver following a fully automated parameterization / computation / post processing procedure. For each computation, the transverse profile of total pressure Pt was extracted at X/C = 7.5% downstream of the WT model trailing edge and the C2 criterion was computed. These results provided guidelines for targeted WT tests. In particular, it allowed identifying an experimental zone to focus on for the investigation of the influence of the flow control parameters [Qscoop, Qblow] on the mitigation of Pt profiles in the close wake of the pylon. Finally, the experimental database also served the validation of the computational procedure for {M∞ = 0.12, Re/m = 2.7 × 10^6, alpha=0deg} and its efficient transposition to {M∞ = 0.23, Re/m = 4.3 × 10^6, alpha=0deg}.

II. Experimental Approach

A. Overview of the WT Model

The design of the WT model is illustrated in Fig. 3. It consists in a C/2 wingspan, C/2 chord length profile complemented with two tip-wall sets at the two ends of the model. In order to ensure uniform flow rates of the scooping and blowing devices along the span of the model, the scooping and blowing device slots are split into separated flow control modules of shorter span, distributed along the wingspan of the model. The WT model is thus primarily composed of three flow control system-equipped modules, hereafter denoted active modules, each 0.11C span, positioned in the central part of the WT model (labeled (1) in Fig. 3). The advantage of this strategy relies in its possible transposition to the future aircraft-dedicated pylon. Moreover, since the aircraft-dedicated pylon will probably not incorporate scooping and blowing slots on its whole span, due to integration issues close to the fuselage and engine junctions, it is interesting to reproduce and to study the potential impact of the local mixing of controlled/non controlled wakes via experiments on the WT model. To this avail, two complementary passive end-modules (not equipped with the scooping and blowing devices), each 0.085C span and labeled (2) and (3) in Fig. 3, ensure the junction of the active part of the pylon with the two tip-wall sets (labeled (4) and (5)). These tip-wall sets aim at isolating potential 3D boundary effects out of the region of interest for the study of the pylon wake.

The initial objectives of the ACcTIOM project, WP2 and WP3, as defined in the Description of Work (DoW) associated with Grant Agreement n°298187, were to design an optimized pylon shape and a dedicated active flow control system able to mitigate the wake of the CROR-propeller pylon in order to comply with Airbus acoustic requirements for {M= 0.23, Re/m = 4.3 ×10^6, alpha=0deg} flow conditions. However, considering actual low velocity flight conditions, such as those encountered during takeoff and landing phases, it is legitimate to consider that the pylon can temporarily be submitted to locally slipped airflow. In order to guarantee the profitability of the present work for the future CROR-propelled aircraft, the consortium ISAE/Aeroconseil has accepted Airbus request to partially re define the specifications of the WT model into a more complex operability domain imposed by {M= 0.23, Re/m = 4.3 ×10^6, alpha=”few deg”} slipped flows, out of the scope of the project at its starting date. This re orientation of the demonstrator specifications has increased its technical complexity, in particular in terms of scooping/blowing modules internal design, parts of the aeraulics circuits, pneumatics elements and MFR control.

At last and in order to comply, in the longer term, with potentially slipped flow configurations past the aircraft, resulting in alpha ≠ 0deg flow past the CROR pylon, the separate control of the pylon wake on both sides of the pylon results in the splitting of each active module into two independent sub-modules, distributed on each side of the pylon.

The WT model is equipped with 83 wall pressure taps for the determination of pressure coefficient (Kp) profiles along the chord of the WT model. The pressure taps are uniformly distributed along three lines, positioned at mid span of each of the three active modules, along the chord and on both sides of the WT model. Each scooping and blowing device is equipped with total temperature sensors (model NS LM35, accuracy ±0.5°C in the range 0 to 100°C). Each blowing device is also equipped with a total pressure sensor (model Keller series 33, accuracy 0.05% of full scale [0-10bar]). These measurements, in association with the measure of the free stream total temperature (accuracy ±0.1degC) and of the free stream total and static pressures (accuracy ±3Pa) are used to fix and monitor the mass flow rate imposed at the outlet of each blowing device when operated during WT test campaigns. In addition, the differential pressure measured at several uniformly span-distributed positions along the span and on both sides of each blowing and scooping device complement the monitoring of the mass flow rate at the inlet and outlet of the scooping and blowing slots respectively. Both scooping and blowing devices are also connected to mass flow meters (normalized accuracy 7.10^(-4), relative to rho∞.U∞.F) ensuring redundancy for mass flow determination. At last, each scooping and blowing device is equipped with several pressure taps, uniformly span-distributed on both sides and along the scooping and blowing slots respectively. These pressure taps allow monitoring the flow homogeneity along the span of each scooping and blowing slots. The previously described pressure taps are all connected to three pressure scanners, model ESP-64HD (accuracy 3Pa). The full set of data produced by the sensors equipping the WT model and the WT test section are acquired via the WT-dedicated data acquisition and control system (24bits A/D conversion).

B. Experimental Setup

The WT test campaign was conducted in the ISAE-S4 subsonic wind tunnel facility. It is an Eiffel-type closed loop wind tunnel, with an opened 3m × 2m elliptical test section. The wind velocity U∞ can be varied from 10m/s to 40m/s at sea-level ISA conditions. The fluctuating intensity in the working test section is below 0.4%. The WT model was positioned in the centre of the WT test section, on a remote-controlled incidence-adjusting device. For a free stream velocity U∞ equal to 40m/s, at sea-level ISA, the 1:2nd scale of the pylon, based on the WT model chord length C/2, results in relatively small chord-based Reynolds number ReC/2, about 4 times smaller than the actual Reynolds number on a full-scale CROR pylon operated at flight conditions {M∞ = 0.23, Re/m = 4.3 × 10^6}. Consequently, to force the boundary layer turbulent transition (for free stream velocities ranging from 20 to 40m/s), two triggering strips -0.5mm thick, 4mm width- positioned along the whole span of the WT model, at curvilinear abscissa s = 0.075C and s = 0.1C from stagnation point respectively were used on the front part of the WT model.
Three configurations, partly illustrated in Fig. 4, were considered during the WT test campaign. The two first configurations, referred to as conf0 and conf1 respectively (Fig. 4(a)), aim at characterizing the reference, uncontrolled flow past the WT model, and the influence of opened (conf0) or closed (conf1) scooping slots on the state of the WT model close wake. They also serve the study of the influence of a triggered vs. natural laminar-to-turbulent transition of the boundary layer on the flow past the model, through the adding, or not, of the triggering strips.

The third configuration, referred to as conf2 (Fig. 4(b)), corresponds to the active control configuration. Here, the central module scooping slots are maintained opened such as to operate the scooping devices while upper and lower module scooping slots are closed. The two triggering strips are fixed as described above. Note that this restricted configuration was imposed by the failure of two of the four available vacuum pumps during pre validation tests that limited the suction capacity of the vacuum supply system. As such conf2 allowed reaching the predicted scooping mass flow rate per meter. Moreover, it is important to mention that the sole use of the central module does not undermine the validation procedure of the active flow control system efficiency since it is based on the replication of several identical modules. This configuration will be deeply analysed by varying the main flow control parameters -scooping and blowing mass flow rates [Qscoop, Qblow]- of the flow control system.

C. Experimental Methods

Two complementary series of experiments have been conducted during the WT tests campaign. The first series was comprised of pressure coefficient (Kp) distribution profiles along the chord of the model and transverse profiles of total pressure (Pt) at various streamwise and spanwise positions in the wake of the pylon. Conf0, conf1 and conf2-{control off} were investigated for three values of free stream velocity U∞, namely 20m/s, 30m/s and 40m/s, and for a unique incidence alpha = 0deg. Investigation of conf2-{control on} was limited to U∞ = 40m/s at sea level ISA conditions (corresponding to a Mach number M∞ = 0.12), and alpha = 0deg. Several flow control parameters [Qscoop, Qblow] were tested for conf2 such as to partly replicate the numerically explored functioning surface of the active flow control system. In the present study the flow past the WT model is expected to be symmetrical relative to the chord of the WT model (alpha = 0deg). Consequently, the scooping (resp., blowing) mass flow rates are fixed to the same value on each side of the model.
The second series of experiments was comprised of 3C-PIV transverse plane measurements for a free stream velocity U∞ = 40m/s and alpha = 0deg. The investigation has concentrated on conf1 (with triggering strips) and conf2 with a limited number of [Qscoop, Qblow] flow control parameters. The latter were selected on the basis of the results obtained from the first series of experiments and considered as the most representative of the influence of the flow control system on the topology of the wake.
Fig. 5 illustrates the 3C-PIV setup in the WT test section. The PIV system was a stereoscopic, High Definition Dantec Dynamics system. It was composed of a frequency- doubled, double-cavity Nd:YAG laser illumination source (wavelength 532nm, 200mJ/7ns per pulse), two highly sensitive digital imaging devices (sensor resolution 4000×2672 pixels2) equipped with two Nikon 180mm f/2.8D lenses mounted on Scheimpflug devices, and dedicated hardware and software for laser/cameras synchronization, data acquisition and analysis.
For each of the conf1 and conf2 reviewed configurations, three transverse planes (parallel to (Y, Z) plane in the reference frame. The origin of the orthonormal reference frame is located at the trailing edge of the WT model, at mid span, corresponding to the centre of the mid active module. X-axis is aligned with the free stream velocity. Y-axis is aligned with the span of the WT model), located at X/C = 1%, X/C = 3.75% and X/C = 7.5% and centered at (Y/C = 0, Z/C = 0) were investigated. The laser was positioned on one side of the test section, out of the wind, on a mono-axial computer-driven displacement system. The two PIV cameras were positioned out of the wind, upstream the model, on both sides of the WT diffuser (Fig. 5). They were fixed on two mono-axial computer-driven displacement tables, aligned with the centerline of the test section.
The measurement plane dimensions were approximately 0.11C by 0.15C, with a measurement grid of normalized spatial resolution (dX, dY, dZ) equal to 5.10^(-4)×3.6.10^(-4)×3.6.10^(-4). The time lap dt between two laser pulses used for the determination of a velocity vector map was fixed to 20μs. The resulting uncertainty on instantaneous three component velocity measurements is about 0.3m/s. Statistics were obtained on the basis of 500 pairs of images per camera. This ensured the convergence of both first and second order statistics with a variance of less than 2% at each point of the PIV measurement grid (variance was below 15% for third and fourth order statistics). Statistical calculations included mean velocities UX, UY, UZ and r.m.s. fluctuating velocities u'X, u'Y and u'Z along X, Y and Z directions, Reynolds stress tensor, covariances and correlation coefficients, skewness and Kurtosis coefficients.

III. Results and Discussion

A. Kp Profile Results

First series of tests was achieved on conf0 (scooping slots opened), conf1 (scooping slots closed) and conf2 with control off, for three values of free stream velocity U∞ (namely 20m/s, 30m/s and 40m/s) and alpha = 0deg. They aimed at i) determining the influence of the triggered vs. natural, laminar-to-turbulent transition of the boundary layer (BL), ii) validating the uniformity of the flow along the span of the WT model, through Kp distribution.
These test results, not illustrated here for the sake of conciseness but fully detailed in deliverable D4.2 and periodic report 3, draw analogous conclusions for the three geometrical configurations. They reveal first that triggered vs. natural BL transition does not significantly affect Kp distribution. This tends to indicate that the triggering strips used to trigger the BL laminar-to-turbulent transition are positioned close to the natural transition zone. Anyway, from now on and in order to strictly control the position of the laminar-to-turbulent BL transition, the triggering strips will be definitively positioned on the WT model, as described in Sec. 3.B. Second, slight Kp profile differences are observed between positive- and negative−Z sides of the WT model, in the lowest pressure zone of the profile, around X/C = Xlp (visible on Fig. 6). This may be potentially due to slight and very local geometrical differences of the WT model, or to pressure tap leaks. Third, comparison of the Kp profiles distribution measured along the three Kp lines, uniformly spanwise-distributed at mid-span of each of the three active modules, are similar along the span and on both sides of the model, for both conf0 and conf1. This confirms the uniformity and the symmetry of the flow along the span of the WT model. At last very good accordance is obtained between numerical and experimental results, except very close downstream of the scooping slot for conf0, around X = Xsc (also visible on Fig. 6, conf2-{control off} for instance, but to a larger extent since this configuration is not geometrically uniform along the span) where a 2% difference is observed between measured and predicted Kp, probably due to a slightly underestimated separation zone in this region. This constitutes a first step in the validation of the computational approach.

Second series of tests were dedicated to the influence of the flow control parameters [Qscoop, Qblow] on the Kp distribution along the WT model profile for conf2. Fig. 6 provides a comparison of Kp distribution between conf2-{control-off}, and conf2-{[Qscoop, Qblow] = [0.058/m, 0.008/m]}, which was identified as the optimal control point in terms of the reduction of the C2 criterion (see below). For these above mentioned flow control parameters, the strong influence of the BL scooping on Kp distribution is clearly visible here, in particular in the second third of the Kp profile, in the vicinity of the scooping slot. In this area, upstream of the scooping slot and due to the reset of the BL by the scooping device, the scooping reinforces the low pressure zone on the profile and annihilates the recompression bubble normally imposed upstream of the scooping slot leading edge when scooping is off. In addition, due to the consecutive strong increase of the close to wall velocity, a strong recompression occurs at the stagnation point on the scooping slot leading edge, that extends while rapidly decreasing downstream of the scooping slot leading edge.

At last, one can still notice a very good accordance between CFD-predicted and experimental Kp distributions, for both control-off and control-on configurations. This reinforces the validation process of the computational approach developed in the design and optimization phase.

B. Pt Profile Results

Results of the total pressure (Pt) transverse profiles test campaign, conducted in parallel to the Kp distribution test campaign, are now discussed. They complement the Kp distribution analysis.
A first series of Pt transverse profile measurements (not illustrated here) was achieved on conf0 and conf1, for the three previously mentioned free stream velocities and for various streamwise and spanwise positions in the wake of the pylon. The latter confirmed the uniformity and the symmetry of the flow along the span of the WT model.
From now on and for the sake of conciseness, only results obtained for flow conditions {M∞ = 0.12, Re/m = 2.7 × 10^6 (V∞ = 40m/s), alpha = 0deg} and at streamwise location X/C = 7.5% will be discussed.

Figure 7 represents (Pt - Pt(out-wake)) profiles (Pt(out−wake) stands for the total pressure out of the wake, and is experimentally determined as the total pressure for which variations by less than 1% are measured for five consecutive Pt probes of the Pt rake.) at mid-span of the central module (Y/C = 0) for conf1, conf2-{control- off} and conf2-{control-on} with three specific [Qscoop,Qblow] flow control parameters. First one can observe the positive influence of the scooping slot shutters on the state of the WT model wake. The wake is much narrower, over 30%, when the active module scooping slots are closed in comparison with opened. Second, controlling the wake solely via a blowing strategy, without scooping the boundary layers upstream of the blowing slots, turns out to be ineffective in terms of C2t target acoustic criterion, whatever the applied blowing mass flow rate (illustrated here for [Qscoop,Qblow] = [0/m, 0.011/m]). Third, flow control parameters [Qscoop,Qblow] = [0.058/m, 0.008/m] demonstrate the strong efficiency of the active flow control system in fully flattening Pt profile in the wake of the pylon, leading to C2 value two orders of magnitude lower than the C2t target value. At last the exhaustive experimental scanning of various [Qscoop,Qblow] flow control parameters has also permitted to identify other very efficient scooping and blowing mass flow rate couples, satisfying the target acoustic criterion C2t. Indeed, operating point [0.026/m, 0.01/m] deeply affects Pt profile which appears to be strongly flattened. This operating point could be a good compromise between wake mitigation and a lower energetic impact than [0.058/m, 0.008/m] in terms of air bleeding in the aircraft systems.

C. 3C-PIV Results

As mentioned in Sec. 3.B, 3C-PIV transverse plane measurements were completed at three positions downstream of the pylon trailing edge (X/C = 1%, X/C = 3.75% and X/C = 7.5%), for a free stream velocity U∞ = 40m/s and alpha = 0deg. The investigation has concentrated on conf1, conf2-{control-off} (hereafter denoted conf2/off) and conf2-{control-on} with a limited number of [Qscoop,Qblow] flow control parameters. The latter were selected on the basis of the results obtained from Pt transverse profile measurements and considered as the most representative of the influence of the flow control system on the topology of the wake. They are listed in Tab. 1.
Figure 8 and Fig. 9 depict UX and UZ mean velocity and u'X, u'Y and u'Z r.m.s. fluctuating velocity cross-planes respectively, measured at streamwise location X/C = 7.5% and centered around Y/C = 0, for conf1 and conf2/off. The comparison between conf1 and conf2/off highlights the positive impact of the scooping slot shutters, when closed, on the mean wake topology and on the fluctuating velocity levels experienced downstream of the pylon trailing edge. Indeed, conf1 exhibits a narrower and more uniform wake along the span of the central module than conf2/off for which both mean UX velocity field and u'X, u'Y and u'Z fluctuating velocity fields are noticeably distorted, in particular in the upper-to-central module junction zone, for Y/C ≈ 0.05, and to a lesser extent in the lower-to-central module junction zone, for Y/C ≈ −0.05 (see Fig. 8 and Fig. 9). This distortion may potentially be induced by junction defaults between the closed scooping slots of the upper and lower modules and the opened scooping slots of the central module.
These defaults may promote unsteady flow separation in these zones. This results in an increase of the r.m.s. fluctuating velocity levels in the wake of conf2/off. However it is interesting to notice i) reasonably similar UX mean velocity profiles in the central part of the central module (Y/C ≈ 0, see also Fig. 12) with a noticeable velocity deficit in the centerline of the wake (Z/C = 0), ii) isotropic fluctuating velocities u'Y and u'Z along Y− and Z− directions (see also Fig. 13), iii) u'X fluctuating velocity along X− direction up to 50% higher than fluctuating velocity in transverse-to-flow directions Y and Z, for both conf1 and conf2/off configurations (Fig. 13).
At last, one can note that UZ mean velocity cross-fields remain quite similar between conf1 and conf2/off, even if slightly distorted in the upper- and lower-to-central module junction zones for conf2/off. This is clearly visible on UZ mean velocity profiles that depict quite similar trends, for both conf1 and conf2/off (Fig. 12). The same trends will also be observed later, for the other tested configurations (Fig. 12). As such the latter will not be further discussed. One can also notice that UX mean velocity profiles out of the wake, i.e. for |Z/C| > 0.016, do not strictly recover the free stream velocity (U∞ = 40m/s). Indeed, at this distance from the centerline, the potential field, and so the velocity field, are still moderately affected by the proximity of the pylon, independently from viscous effects. However, the associated transverse velocity gradients are too low to be considered as an issue for airflow/rotating blades interaction-induced acoustic concerns.

Figure 10 and Fig. 11 compare UX and u'X statistical cross-planes for conf2/off vs. conf2/3 and conf2/5, and for conf2/off vs. conf2/4 and conf2/6 respectively. The strong efficiency of the flow control system in mitigating the wake of the pylon is clearly highlighted, in particular for conf2/4 and even more for conf2/6 which both lie in the flow control parametric region of operating points where the acoustic criterion was found to be optimal, i.e. below the target value C2t, based on Pt transverse profile measurements. Interestingly enough, it can be noticed that the initial wake, corresponding to conf2/off and which, it is worth remembering, is disturbed and non uniform, is brought back to a much more uniform state when the active flow control system is activated, for the four controlled configurations investigated here (see Fig. 12 and Fig. 13). The efficiency of the flow control system is also well demonstrated when considering the u'X r.m.s. fluctuating velocity levels, which are strongly attenuated when the active flow control system is turned on, in particular for conf2/4 and above all for conf2/6, in comparison with control-off (conf2/off) or with conf1 (Fig. 11). These observations are confirmed by the analysis of mean and fluctuating velocity profiles extracted from the PIV cross-planes (Fig. 12 and Fig. 13). Therefore, i) UX mean velocity profiles are strongly flattened when flow control system is turned on to optimal flow control parameters, in particular for conf2/6, along the span of the central module, except at its periphery, for Y/C ~ ±0.055, where the controlled wake mixes with the uncontrolled wakes generated by upper and lower modules (the latter are well identified in Fig. 10 and Fig. 11, on UX mean velocity cross-planes, where they appear as blue pockets, localized around Y/C ≈ ±0.06); ii) the wake is strongly mitigated and narrowed for conf2/4 and conf2/6 in comparison with conf1 (see Fig. 9) and conf2/off (Fig. 11, see also Fig. 12 and Fig.13); iii) r.m.s. fluctuating velocities u'X and u'Y recover isotropy while, contrary to conf2/off, u'Z fluctuations increase, by up to 50% for conf2/4, and by up to 75% for conf2/6.

At last Fig. 12 and Fig. 13 synthesize the comparative review of conf1, conf2/off, conf2/4 and conf2/6 in the form of UX/U∞, UZ/U∞ normalized mean velocity, u'X/U∞, u'Y/U∞ and u'Z/U∞ normalized r.m.s. fluctuating velocity as a function of Z/C, at X/C = 7.5% and Y/C = 0. They confirm the strong efficiency of the embedded flow control system in erasing the pylon wake, when operated in the region of optimal flow control parameters numerically and experimentally identified.

IV. Conclusion

This study aimed at developing and validating an innovative concept of active flow control system, dedicated to the attenuation of airframe noise and vibrations on a CROR propeller through the pylon wake mitigation. It relies on a boundary layer scooping/blowing strategy. First a multi parametric iterative design and optimization process was conducted thanks to 2D and 3D Reynolds Averaged Navier Stokes computations. It allowed designing the wind tunnel model of the flow control system-equipped pylon and pre defining the flow control parameters - scooping and blowing mass flow rates - to apply such as to mitigate the pylon wake. By experimentally varying these parameters, and taking advantage of the WT experiment results, based on pressure coefficient distribution along the WT model profile, transverse profiles of total pressure at various streamwise and spanwise positions in the wake of the pylon and on 3C-PIV transverse-plane measurements at various streamwise positions, an exhaustive cartography of the flow control system efficiency was determined. It allowed identifying nominal operating points, where the target acoustic criterion is satisfied. It has also highlighted the robustness of the embedded flow control system in fully mitigating the wake of the pylon, even under moderate scooping and small blowing mass flow rate variations. Finally, the computational approach has been validated and its transposition to the prediction of the wake mitigation for flight conditions can be assessed.

Second objective of the ACcTIOM project: to develop an advanced experimental methodology, based on vibration-controlled stereoscopic Particle Image Velocimetry (3C-PIV), able to be flight-operated inside the Flying Test Bench in order to validate the efficiency of the active flow control system in annihiltaing the pylon wake.

WP4 and WP5 synthetic description of the actions and results obtained

VES development, manufacturing, validation and operation (WP4)

• determination of the estimated Flying Test Bench (FTB) vibration spectrum to be experienced in the cabin of the CROR-propelled FTB aircraft, through exhaustive literature review, Airbus standards for vibratory test procedures and series of vibratory tests (not initially planned in the Grant Agreement) on the ISAE heavy load shaker (tests campaign completed in Nov. 2013)
• Definition, manufacturing, validation and operation of the static VES dedicated to the determination of PIV measurement errors under the influence of optical misalignments induced by imposed vibrations.
• Definition, manufacturing, validation and operation of the dynamic VES, consisting in a mechanical system able to generate multi axis vibrations satisfying the previously determined vibration spectrum.
• Development of a simulation model of this mechanical system, ultimately serving the conception of the VCM active control laws.

Vibration induced PIV errors quantification (WP4)

• Quantification, via the VES, of the envelop of acceptable vs. prohibitive displacement/misalignments of PIV subsystem relative to the other ones, along the 6 misalignment axes, beyond which the VCM should be activated.

VCM Development (WP4)

• Definition of a hybrid vibration control strategy: passive isolation coupled to an active control system
• Global specification of the passive isolation system and manufacturing
• Validation of the passive isolation system through dynamic VES tests
• Detailed specification and purchase of the active control system
• Receipt of the active control system and characterization of its dynamics
• Mechanical interfacing of the active control system with the passive isolation system
• Definition, development and validation of control laws for the active control system

Validation of VCM for in-flight 3C-PIV (WP5)

• definition of several strategies to ensure the implementation and operation of in-flight 3C-PIV measurements in the cabin of the CROR-powerplant equipped FTB.

WP4 and WP5 detailed description of the actions and results obtained

Approach and Modelling of the problem
This section presents the methodology developed to propose strategies enabling the implementation and operation of in-flight 3C-PIV measurements in the FTB. It relies on the reductionism principle and involves various modelling levels, both for the experimental and predictive approaches.

First of all, let's recall the guiding principle that served the definition of the methodology developed in WP4.
The vibratory environment inside the cabin of the FTB is expected to be potentially prohibitive for the in-flight implementation and operation of stereoscopic PIV (3C-PIV). The application of in-flight 3C-PIV is thus highly challenging, while its domain of applications and the inherent advantages to use such efficient flow characterization method during flight tests is evident.
This vibratory environment and more particularly the vibration spectrum experienced inside the cabin of the FTB are however still nowadays unknown. Indeed neither the integration of the CROR powerplant on the FTB nor the numerical prediction of the structural interaction between the operating CROR engine and the future FTB are yet available. Consequently a first assumption had to be made at the early stage of WP4 by the consortium in charge of the ACcTIOM project, in order to propose a probable and consistent vibration spectrum able to confidently represent the FTB vibratory environment when in-flight operated. The tasks that led to the definition of this spectrum were previously detailed in deliverable D4.2 and in periodic report 2. This assumption will hereafter be denoted Assumption#1.

If submitted to the abovementioned vibration spectrum, the different PIV subsystems, namely the PIV cameras, lenses, double-pulse YAG laser and laser sheet generating lenses, may be subjected to time-varying relative displacements, leading to optical misalignments. The latter may affect more or less, depending on the amplitude of the misalignments and on the involved axis/axes of misalignment, the ability of the stereoscopic PIV algorithm to faithfully rebuild 3C velocity fields on the basis of the 2C displacement fields computed for each of the two cameras implemented in the 3C-PIV measurement system. It is indeed easy to appreciate that the 3C-PIV method will probably still be operable and efficient for moderate misalignments of a given subsystem, but not necessarily with the same level of misalignment for each of the considered misalignment axis, while it will not be operable anymore for larger misalignments.

At this stage of the discussion, it is thus necessary to design an experimental test bench dedicated to the determination of the maximal displacements, along each of the 6 axes (3 translations, 3 rotations), below which the 3C-PIV measurements can still be considered as acceptable, and above which a vibration correction methodology is required. This will hereafter be referred to as the envelop of acceptable misalignments. Furthermore this test bench should guarantee a high degree of confidence, in terms of accuracy, repeatability and reproducibility, as it will involve a statistical approach to quantify the measurement errors induced by 3C-PIV subsystem misalignments. In particular, the test bench will not be based on the measurement of a real flow, seeded with submicronic particles, as such flow would naturally produce uncontrollable fluctuations due to turbulence. This would not permit to satisfy the required repeatability for the statistical approach without a prohibitive amount of data. To this avail, an original synthetic flow field, made of a granular coating covering a multi panel plate has been designed. The latter, hereafter denoted synthetic flow device or SFD, perfectly imitates a PIV image when observed at the distance to be considered between the PIV cameras and the laser sheet in the future FTB during flight tests. Moreover this synthetic flow field enables to reproduce the displacement, in a depth equivalent to the one of the pulsed laser sheet to be used during flight tests for the illumination of the measurement plane, of particles that would normally be induced by the ow past the FTB, between the two instants separating the acquisition of two images used for the determination of an instantaneous displacement field through image cross-correlation algorithms. Furthermore the SFD provides a still more complex flow than the one expected in the wake of the CROR pylon on the FTB, by superimposing an abrupt shear displacement on a restricted part of this plate, and a rotating displacement, faking a vortical flow (known as highly challenging for PIV misalignment issues !), on another part of this plate. As such the robustness of the approach developed for the definition of the VCM is still enhanced. The simplification of the real flow field illuminated by the laser sheet to a synthetic one, the SFD, constitutes the second assumption, referred to as Assumption#2.

Moreover, as this approach requires a large number of PIV images of the synthetic flow field to both ensure the convergence of the statistics for a given measure, but also for the determination of the envelop of acceptable misalignments, the time needed to conduct the analysis would be more than prohibitive. Indeed an estimation of more than 35 thousands steps of displacements (cumulated value for the determination of the acceptable envelop of displacements along the 6 axes), and so of more than 35 thousands 3C velocity fields to compute are required per configuration of cameras/SFD positioning, leading to series of cumulated test campaigns of more than 100 hours. Conducting such huge test campaigns on the basis of a real flow field would be inconsistent with the requirement to perfectly control the experimental parameters, from the beginning to the end of the test. This still reinforces the pertinence of Assumption#2, based on the SFD approach. This is detailed in Sec. 1.

The determination of the maximal displacements of a given 3C-PIV subsystem relative to the other ones along each of the 6 axes (3 translations, 3 rotations), above which a vibration correction methodology is mandatory, requires to perfectly control the step by step displacements of this subsystem along each of the 6 axes, with a high accuracy. To this avail a high precision Stewart platform, also referred to as hexapod, will be used. Its characterization was provided in Deliverable D4.2 and periodic report 2. In practice, both cameras should be misaligned in a controlled way on the calibration and misalignment test bench, as they will be similarly implemented inside the cabin of the FTB. However, due to restricted funding constraints, only one hexapod could be purchased. Consequently the influence of misalignment on the cameras will only be investigated for 1 of the 2 cameras. This simplification corresponds to Assumption#3. This assumption does not undermine the analysis since the global measurement error induced by the 2 cameras misalignment is a cumulative effect of both camera misalignment. As such the strategy to adopt for the VCM definition will be based on the worst case, where opposite misalignment occur for the 2 cameras, decreasing the maximal acceptable displacement per camera by a factor of 2. The influence of the laser misalignment will not be analyzed here. The reason is twofold. First, and as further described in Sec. 1, the laser and its laser sheet generating lenses (rigidly fixed on the laser main structure) will be positioned on a highly rigid structure, positioned in a down position, on the main structure of the cabin floor. As such and contrary to the cameras, that should be fixed in a high position, close to the roof of the cabin in order to have an easy optical access to the outside of the cabin, on the laser sheet illuminating the flow in the wake of the CROR pylon, the risk to amplify any vibrations due to the flexibility of the supporting structure is much lower for the laser. Second, a displacement of the laser sheet relative to its reference position would only translate into uncertainty in the positioning of the measurement plane in the FTB reference frame. In this context the exhaustive characterization of the flow in the wake of the CROR pylon, provided during WP2 and WP3, has shown that the variation of the flow field in the out-of-laser plane direction is order of magnitude lower than in the in-plane direction, strongly limiting the impact of this potential out-of-plane displacement on the measurement. While the in-plane variation of the laser sheet position would not affect at all the measurement, as far as the laser sheet remains in the field of view of the cameras (which is easy to ensure by generating a laser sheet quite larger than the camera field of view), the out-of-laser plane potential variations of the laser sheet position would only be sensitive if the laser sheet would move out of the depth of field of the lenses. As explained later, this is unlikely to occur, especially since the reference distance between the cameras and the laser sheet will be quite large, of the order of 4m.

A fourth simplification of the problem is proposed regarding the time scales involved. For in-flight PIV measurements in the FTB, the time lapse separating the acquisition of two consecutive images constituting a pair of images and providing an instantaneous displacement field of the flow when cross-correlated, is of the order of a few microseconds. During that time lapse, the relative displacement of a 3C-PIV subsystem relative to another one can faithfully be neglected, considering the previously evoked vibration spectrum, and more generally the time scales representative of structure vibrations. As such, a static approach can be proposed for the calibration and misalignment test bench, consisting in acquiring at 2 perfectly uncorrelated instants 2 images of the SFD, the latter being moved in a controlled way to 2 different positions for the acquisition of the 2 images, such as to simulate the particle displacements in the real flow during the time lapse separating the acquisition of real PIV images. During this time lapse, the different PIV subsystems should thus not be moved. Considering now the time lapse separating the acquisition of 2 consecutive pairs of images, the latter is of the order of a few tenths of a second. During that time lapse, the different PIV subsystems may have moved, one relative to another one, under the influence of vibrations promoted by the CROR engine and modelled via the previously mentioned vibration spectrum. As such, the controlled displacement of the hexapod along the 6 axes will simulate these potential displacements. The modelling of the vibration-induced displacements through this static approach for the calibration and misalignment test bench refers to as Assumption#4.

Once determined the acceptable misalignment limits along each of the 6 axes for a given 3C-PIV subsystem relative to the other ones, these displacement values should be translated in terms of corresponding vibration frequencies in the vibration spectrum. The so-identified maximal displacements, and associated minimal frequencies should then be used as input for the definition of the Vibration Control Methodology (VCM), able to annihilate detrimental vibrations, reproduced via the Vibration Environment Simulator (VES), designed in WP4 and detailed in Sec. 2, and responsible for ineffective 3C-PIV measurements. This is described in Sec.2.
The following diagram summarizes this model approach. At the center of the Figure A stands the ultimate goal, the implementation of in-flight 3C-PIV measurements in the CROR powerplant-equipped FTB. The modelling axes are:
• Vibratory modelling: How to simulate an unknown vibratory spectrum from in-flight to laboratory. It corresponds to Assumption#1,
• Environment modelling rate: cabin environment or laboratory environment. How to define the PIV subsystems relative position, and to simulate a real flow, such as to confidently represent what the in-flight 3C-PIV tests will be? It corresponds to Assumption#2 and Assumption#3,
• Temporal modelling rate: dynamic vs static approach (block static PIV and block Vibration Control Methodology). It corresponds to Assumption#4.

1 Definition of the displacement limits – Stereoscopic 3C-PIV measurements

1.1 In-Flight to Indoor PIV Measurements

In order to know the required performance that must be satisfied by the vibration control system (passive isolators + hexapod), the maximum authorized displacement values (the envelop of acceptable misalignments) that still enable to reconstruct the correct 3C velocity fields from the 3C-PIV measurements have to be specified. More specifically, the limit values shall be used to specify the required position accuracy of the active control. Performing all required PIV measurements manually would take a long time due to the total number of tests that must be done. In order to avoid such a lost of time, both numerical prediction (based on a ray tracing approach) and experiments relying on a dedicated test bench have been conducted so as to define the limits of displacements in an automatic way. We first describe the ray tracing approach, then the experiment setup one. Figure B shows a schematic view of the CROR engine in situation and the proposed test bench configuration used for the numerical and experimental approach.

1.1.1 Effect of optical component misalignment on the quality of PIV images

In the framework of the PIV technique implementation in vibrating environment, it can be relevant to study the effect of optical component misalignment on PIV images quality, in order to define the less sensitive imaging system. This has been achieved with the ZEMAX OpticStudio software, by considering a geometrical configuration between the PIV camera sensor and the PIV measurement plane, corresponding to the laser sheet, in accordance with what the relative implementation of the various 3C-PIV subsystems and the position of the laser sheet in the wake of the CROR pylon wake should be in the FTB, during 3C-PIV in-flight measurements. The advantage of a large F-number optic has been highlighted insofar as it can be considered as less sensitive to vibrating environment. This assumption has been confirmed by a parametric study showing the effect of image plane misalignment on the specifications of the imaging system. The assessment of the optical specifications (RMS spot radius, Point Spread Function) has been achieved with ZEMAX OpticStudio software, applied on a generic stereoscopic PIV configuration obeying the Scheimpflug law. This preliminary study has to be pushed on to estimate the effect of individual optic misalignment inside the objective, on the optical characteristics of the imaging device. Furthermore, the influence of a transparent window inserted on the optical path (optical access in the aircraft), should be taken into account before studying the effect of the light scattering provided by the seeded flow.

1.1.2 Experimental setup and Method

The previous section has permitted to apprehend the influence of optical parameters such as the optical arrangement between the PIV camera sensor and the PIV laser sheet, the tilt angle of the camera sensor relative to the lens, the F-number parameter of the lens, the image plane location and tilt angle on the PIV system sensitivity to vibration-induced misalignment. On this basis, it is now interesting to deepen the study by exploiting the efficiency of the dedicated 3C-PIV calibration and misalignment test bench in providing the envelop of acceptable vs. prohibitive misalignment along the 6 axes (translations and rotations). The fundamental principles of the approach are:
• Quantify measurement errors (mean and standard deviation) of the optical system + synthetic flow field generation system, also referred to as SFD, when the 3C-PIV subsystems are at their reference positions i.e. when they are not yet submitted to similar to vibration-induced misalignments. This corresponds to the baseline,
• Simulate a 3CP-PIV subsystem misalignment due to vibrations and analyze the impact on 3C-PIV measurements errors (Squint-eyed view) by comparing mean and standard deviation with the baseline,
• Repeat the previous tasks by varying the height of the 3C-PIV cameras relative to the PIV synthetic flow field to quantify the influence of the cabin implementation of the 3C-PIV subsystems on misalignment-induced measurement errors.
Figure C describes the context of the proposed method. The top view of the Figure C highlights positions of the laser-sheet and the hypothetical distances of cameras from the latter inside the cabin and behind glass windows. The bottom left view focus on heights of cameras to match the related angles described in the top view as well. Finally, the bottom right view puts lights on the stereoscopic view to perform the 3C-PIV measurements.

In this study, as visible on the bottom left view of the Figure C, several heights of cameras related to the PIV measurement plane, hereafter denoted target, will be used in order to investigate on the influence of the camera position. Indeed, at this moment, the integration of cameras in the cabin is not yet defined, this leading to what will be the best implementation of camera considering the angle between the laser-sheet and the CCD sensor of the camera. To do so, in this section, the Flying Test-Bench is modelled following a static approach point of view. To this avail the images of particles that should normally be obtained by illuminating the seeded flow with the laser pulses during in-flight tests in the FTB are modelled using a solid target made of sand with a specific size fixed with a black resin in lieu of the PIV pair of particle images [Figure D]. Two moving parts are used to simulate rotating and shear flow displacements. In fact, this approach comes from the time scale between a 2-picture acquisition (per camera). Indeed, in PIV, a pair of pictures being acquired within less than, typically 10 microseconds, the displacement and deforming of the structure on which the cameras are fixed to may confidently be neglected during that time. Consequently, time independent measurements may be used on the static PIV test bench, without imposing any misalignment between the first and the second acquired images, constitutive of a pair of images used for the computation of an instantaneous flow field measure. However, the time separating the acquisition of a pair of images to the consecutive one is much larger, typically a few tenths of a second, depending on the PIV camera technology. As a result, during that time the deformation of the structure (hence its displacement) supporting the cameras is no longer negligible (spectrum based on the RTCA/DO-160E standards, curve C, provided by Airbus), and this may have an impact on the PIV measurement errors through stereoscopic rebuild issues. As the determination of the time varying displacements imposed by the vibrations of the structure supporting the various 3C-PIV subsystems is challenging, if not still unknown, the use of an incremental displacement of the hexapod modelling the structure deformation along the 6 axes will serve the characterization of acceptable vs. prohibitive misalignment of a given 3C-PIV subsystem vs. the other ones along these both independent or combined 6 axes (translations and rotations).

Knowing those limits imply to carry out successive images of a flow of particle with a relative displacement of one camera or the laser sheet between the two sets of images, and to do such records for every kind (translation or rotation) and values of relative displacements. Moreover, for each type of relative displacement, the procedure must be reiterated for several amplitude of displacements so as to find the limit for the computation of the correct velocity fields. In order to simulate the relative displacements of one camera compared to the other two elements of the PIV system, one camera is fixed on top of the hexapod which is used in this case to generate static displacements, while the other camera remains still as shown in Figure E. Relative displacements of the laser sheet are simulated by translations of the SFD, also referred to as the target, as shown in Figure D and without modifying the camera's focus. The algorithm that is used for the automatic generation of the limit displacements shape is based on the use of softwares such as Dynamic Studio and Matlab. Dynamic Studio is used to control the cameras and to make the acquisition of the particle images and Matlab is used to control the camera and target's displacements. Thus, this algorithm requires also the ability of interfacing both softwares.
Consequently, those remarks lead to the quantification of limit displacements of the camera, itself being divided in 3 milestones:
• The calibration of the PIV bench deals with both the Scheimpflug conditions and the Soloff method i.e. multi-planes calibration in the camera's depth of field,
• The error quantification of 3 different kinds of target's moves recreating a typical flow. This deals with moving only the target (stationary cameras) in order to get errors due to:
o Optical acquisition system,
o Dantec Algorithm (DynamicStudio software),
o Move accuracy of rotation and linear (translation and shear) stages through repetitions for
different values,
o SFD moving in the depth of field of cameras, for displacements representative of the actual
displacements of seeding particles in the typical flow generated past the FTB during flight tests.
• Moving the camera embedded on the hexapod within the 6 degrees of freedom to set the maximum acceptable displacements allowing the reconstruction of the velocity/displacement field. This method is similar to the eyesight. Indeed, each eye catches 2D pictures, the brain makes them 3D and if an eye is moving from its reference position (squinting eye) then the vision is blurred. Here, we want to determine at which position this phenomena is happening with the camera.

The main milestone is the final one and those milestones will be repeated over different heights of cameras still targeting the SFD. Pictures are not including the Scheimpflug rule and the bi-Scheimpflug rule. Moreover, 4 configurations have been tested (z0, z1, z2 and z3).

1.2 Quantification of PIV measurement errors through picture processing

Method: baseline characterization and camera misalignment
The baseline characterization (for each height Z0, Z1, Z2, Z3) is the measurement errors due to the SFD displacements (synthetic flow expected in-flight conditions), optical system and post-processing algorithm:
• Translation : half of a typical laser-sheet thickness (a few millimeters is a common value for the laser sheet thickness),
• Rotation : modelling a vortical flow,
• Shear: modelling an abrupt shear layer.
Once errors being quantified, a reference displacement (target's moving parts) representative of the flow is defined to reproduce in-flight conditions at 0.23 Mach (285 km/h). It is used in Sect. 1.3.2 while the camera on the hexapod moves following values in Figure F in order to determine limits of maximum acceptable displacements for the camera.

Figure G illustrates the procedure of the camera moving along 1 axis (for instance Z) and capturing the reference displacement representative of in-flight conditions, no matter which configuration (either 'z = 0 mm' or 'z 0 mm') is considered.
For a given camera position defined by the Scheimpflug rule, the target moves from a position 1 to a position 2 (Figure G) while the camera acquires images for each target position, and this gives one pair of images leading to a displacement field. Then the camera moves in the range defined by Figure F, the target moves from a position 1 to a position 2 (Figure G) while the camera acquires images for each target position, and this gives a second pair of images leading to another displacement field...and so on following the camera displacement with a fixed step defined. This methodology is adequate because of the time scale between picture acquisitions. Indeed, between pictures corresponding to target position 1 and 2, only 5 μs separate them in order to giving enough time to particles crossing the laser-sheet at Mach 0.23. Meanwhile the structure where the camera is based on has no time to get deformation, so the camera does not move between the pictures corresponding to the SFD (target) in position 1 and 2.

However, from a pair of images to another one (Figure G), time flies (the order of a few tenths of a second) and from a pair of images to another the structure may move and the camera as well. In PIV, several acquisitions are needed to do averages, typically between 50 and 150 pairs of images. In order to compare all these pair of images, the camera need to stay the same position (along the 6 degrees of freedom) to make results relevant. In facts, the camera cannot stay in this very particular position accurately due to vibrations of the in-flight conditions. Here, we want to define the maximum acceptable camera displacement to reconstruct the displacement field, hence the displacement of the camera, for instance on thetaZ (Figure G) with a fixed step, and for sure the target displacement from position 1 to 2.

1.3 Results

1.3.1 Definition of a reference flow / displacement - Baseline characterization

This section sets a reference displacement of the moving parts which will be used in the Sec.1.3.2. Indeed, this section refers to the context of the aircraft; the mean flow at Mach 0.23 around CROR's pylon. As a result, a reference displacement is set while cameras are not moving since the structure of the airplane (where the cameras are based on) is considered as rigid. This reference displacement is the one expected for the wake around the Counter Rotative Open Rotor. Finally, these reference displacements are still the values to reach out when the camera moves in the next Sec.1.3.2 and as soon as the reference displacement is not retrieved with the algorithm, then the limit of the acceptable camera displacement is defined for each one of its 6 degrees of freedom. Thus, for this reference displacement, set values of moving parts are defined in the Figure H.

1.3.2 Technical specifications for Vibration Control Methodology

Sec.1.2 lead to the definition of maximum acceptable displacements allowing a reconstruction of velocity/displacement fields. Consequently, the whole methodology developed in the Sec.1.2 is performed several times by taking into account various elevations of cameras relatively to the center of the target, hence the cabin configuration with the transposable issue from heights to different distances between laser and cameras. So, it has been performed other baseline characterizations and camera misalignment (squint-eyed view) due to vibrations for various heights. We recall that the baseline characterization (for each height Z0, Z1, Z2, Z3) is the measurement errors due to the SFD displacement (synthetic flow expected in-flight conditions), optical system and post-processing algorithm, while maintaining the cameras at their reference positions:
• Translation : half of the laser-sheet thickness,
• Rotation : vortical flow representative of, e.g., a vortex generator,
• Shear : an abrupt shear layer.
The Figure I sets relative displacements (translation and rotation) of 1 camera compared to the other one along its 6 degrees of freedom.

Explanation In low configuration (Z0), the main drawback is the displacement sensitivity along axes of freedom thetaZ thetaY. Indeed, the Scheimpflug rule is quickly not completed since the 3 planes do not cross any more. For instance, a 0.170deg camera rotation on thetaZ corresponds to a 11.8 mm shift on target's z-direction. Thanks to the distance between cameras and the target (i.e. 4m), this high displacement tolerance is possible contrary to classical short distance PIV measurements. Moreover, a displacement on the X-axis does not affect the reconstruction of displacement fields for any elevation if the diaphragm of the lens is quite closed (F/8D). Indeed, a wide F-number allows a significant camera's depth of field and the second Scheimpflug angle due to the elevation (to set the focus on the whole target) of the camera is no longer required. For in-flight PIV conditions, the solution would be a powerful laser and a diaphragm of the lens wide (at least F/8D) opened in order to increase the depth of field and to be less pessimistic on the rotation angles due to F/1.8D to correctly reconstruct the velocity fields.
The expected motion in-flight conditions lead the camera to move on many degrees of freedom. Moreover, rotation angles are considered the most pessimistic drawback to not complete the Scheimpflug rule and to not correctly reconstruct displacement fields. Consequently, the Figure J presents the maximum rotation angles for different heights.

2 Vibratory Environment Simulator

2.1 Definition of a vibratory environment

The design of the vibratory environment simulator relies on the estimation of the vibrations that will affect the PIV system during flight tests. These vibrations are commonly characterized in terms of an acceleration Power Spectral Density (PSD). The characteristics of the latter, referred to as the input data for WP4, is a major challenge since no vibratory data are yet available in the context of a CROR powerplant-equipped aircraft. Indeed, to the author's knowledge, no quantitative information about the vibratory sources of the CROR powerplant have been produced yet and the FTB is not yet operational. Consequently, and so as not to penalize WP4 progress, series of experiments, detailed in deliverable D4.2 and periodic report 2, have been conducted in order to define the acceleration PSD that will be applied to the PIV system. Since this PSD should remain consistent with the expected, but paradoxically unknown, vibratory environment of the FTB, a pragmatic approach was adopted, based on both a comprehensive literature search and Airbus standards for vibratory test procedures. This frequency spectrum, illustrated in Figure K, is based on the RTCA/DO-160E standards, curve C, provided by Airbus. It was partly reshaped in order to better comply with vibratory levels encountered in normal flight conditions, while still ensuring a high safety factor, over 40 for the root mean square (rms) value of the vertical acceleration, in comparison with effective vibration levels measured in various commercial aircrafts of different generations under smooth cruise conditions.

One should however notice that the existing bibliography regarding such topics is poor. The limited data available in the literature indicate maximum measured rms and peak values of the vertical acceleration of the order of, e.g., 0.0085g and 0.03g respectively, where g stands for the gravitational acceleration, in the cabin of a Boeing 717 or Douglas DC-9 for smooth cruise conditions. These values reach up to 0.12g and 0.67g respectively for touchdown conditions. In the cabin of a Dash8-300, the measured rms value of the vertical acceleration equals 0.03g for smooth cruise conditions. Furthermore, it is taken for granted that the maximum relevant frequencies for vibration perception lie below 200Hz. Here, the rms and peak values of the vertical acceleration associated with the proposed PSD equal 1.2g and 4g respectively and the frequency range of the acceleration lies between 10 and 1000Hz. This acceleration PSD will henceforth be the input data for the design of the vibratory environment simulator (VES). It is important to note that this PSD corresponds to the vibratory content along the vertical direction, hereafter denoted z. Publications detailed in deliverable D4.2 mention that, for smooth cruise conditions, the lateral accelerations experienced in the cabin of a B717 or a DC-9 do not exceed 15% of the vertical accelerations, and that these lateral accelerations are nearly imperceptible below a few tens of Hz. Getting no further vibratory specifications, neither in the available literature nor in Airbus standards, and in order to still ensure a high safety factor for the VES in comparison with in vivo vibration conditions in the FTB, the vibrations imposed by the VES along the x and y directions, and more particularly the low frequency part of their respective PSD (the most critical for PIV measurements, in comparison with higher frequencies, since these low frequencies impose the largest parasite displacements) are assumed to be weaker or, at worst, similar to the z-direction PSD. A benchmark study on commercial systems able to generate tri axial vibrations in agreement with the previously defined PSD, as well as possible rotations, revealed the lack of systems whose costs are in line with the ACcTIOM budget. This is mainly due to the large frequency bandwidth that is required. This benchmark study has brought to light the necessity to engineer a VES system that meets the required technical specifications at reasonable cost. The chosen solution uses a shaker that generates vibrations in the z direction, and a spring-based mechanical system to transform this mono axial displacement into both translations and rotations along the three x, y and z axes. Such a mechanical system, with three different springs, has been developed. (Figure L) shows the designed system. In order to give more degrees of freedom in the final spring settings, the mechanical assembly of each individual spring is designed to give the ability of adapting the stiffness. Moreover, ball-joints are used in order to provide the required rotation and horizontal translation degrees of freedom. Finally, a central mechanical rod is used so as to ensure the weight support of the upper plate and its payload and is fixed to both plates by means of ball-joints as well.

2.2 Experiment

The system presented on Figure L has been tested thanks to a shake table. Unfortunately, the shaker table's coil turns out to be very sensitive to the torque caused by the mass rotation around the ball-joint. Indeed, the mass leads to a torque certainly desired but too much significant for the ISAE's shaker table to handle it.
As a result, it has been not possible to generate lateral acceleration from vertical ones. However, the ACcTIOM's partner (Airbus) has the possibility to reproduce the vibratory environment on 5 axes and to ensure the certification of the system through a vibratory environment. Moreover, a second contact (Alliantech) has agreed to help us on this issue with a similar 5-axis shaker table. The Alliantech's shaker table is located at Munich (Germany). Consequently, an experiment was set up to reproduce and measure the vibrations that will be encountered into the aircraft. A shaker was used and controlled so as to get accelerations according to the same frequency spectrum of Figure K on vertical (z) direction and horizontal (x) direction as well in spite of the fact lateral (x-direction or y-direction) accelerations are expected to be 15% of the vertical one (z-direction).
Real accelerations and velocities carried out by the shaker were measured and are shown in Figure M. Then, the displacement frequency spectrum had been computed from velocity temporal records. From the Figure M, it can be seen that the displacement frequencies are mostly in the frequency band 5 Hz to 20 Hz. This means that the frequency components greater than 20 Hz in the acceleration spectrum do not give rise to significant displacements. However, small displacements still exist for greater frequencies than 20 Hz and must be cut-off by mean of a passive isolation. This leads to protect camera's electronics. Regarding the lower frequencies than 20 Hz harmful for the PIV measurements, an active control has to be implemented.

2.3 Conclusion

Conclusion on the analysis of the vibratory environment An investigation and the definition of a vibratory environment representative of a CROR-propelled aircraft has been done. This task is not obvious since CROR-propelled aircraft do not exist yet and the vibratory environment spectrum can only be estimated from literature. This study, based on experiments, gives an idea of the displacements, velocities and accelerations of the aircraft's structure which the PIV system will be based on. It gives three important indications for the specification of the vibration control methodology (made of a passive system and an active one) that should be added up in order to control the cameras:
• a maximum acceleration 4 g and a rms acceleration 1.2 g,
• a maximum displacement 0.9 mm mostly due to the frequency range [ 0 - 20 Hz ],
• a maximum velocity 67 mm/s.
Conclusion for the Vibratory Environment Simulator A Vibratory Environment Simulator has been designed to obtain from 1 degree of freedom in the vertical excitation, many degrees of freedom. Due to reasons detailed in this Sec.2.2, this system is not efficient. However, in order to overcome this drawback and to not underestimate in-flight vibratory results, expectations for degrees of freedom x and y (lateral direction) would be equal to results got in the z (vertical direction). Given the literature and Sec.2, this leads to pessimistic assumptions since in the aircraft's cabin, the lateral excitations (x-direction and y-direction) do not exceed 15% of the vertical excitation (z-direction). Tests models and simulations developed in the Sec.3 take this assumption into account.

3 Vibration Control Methodology – Strategy for the mitigation of the vibration-induced PIV measurements errors

3.1 Introduction

The vibration-induced PIV measurement errors will depend upon the isolation system that will be used to limit vibrations of the PIV cameras and laser. From the frequency spectrum given in Figure M, a combined passive-active isolation (hybrid) must be used. Indeed, a passive isolation is necessary to get rid of the high frequency vibrations but cannot filter the low frequency components. In first approach, the remaining low frequency vibrations should thereby be managed by an active vibration control. Thus, the hexapod, used in Sec.1.3 for the quantification of PIV measurement errors, may be used to perform a such active control. In order to specify the technical solution for passive and active isolations, a study has been led to analyze the displacements and velocities caused by vibrations whose frequency spectrum is given by Figure K. We address in this section the design of a hybrid system for vibration control of PIV subsystems. The Hexapod PI, used in Sec.1 for PIV measurement errors, is evaluated to perform such a hybrid system. The solution of vibration control will be assessed by comparing displacements and accelerations without any vibration control system (Sec.2) and with vibration control. First, we focus on the passive isolation system (elastomer), than on active isolation system (hexapod). Finally, we will conclude on the Vibration Control Methodology based on a hybrid system.

3.2 Passive isolation system

3.2.1 Introduction of the passive isolation platform

Passive isolation systems enable to avoid high frequency vibrations. Among the various available technologies, elastomer or mechanical isolators offer the advantages to be cheap, small and to give both vertical and horizontal isolations. From the performance and mechanical specifications implied by the control of the 3C-PIV measurement system, it was chosen to develop a passive platform based on the use of elastomer isolators.

3.2.2 Definition of an intermediate structure between isolators and the hexapod

We recall that the aim of the Sec.3 is to provide a hybrid Vibration Control Methodology made of passive isolators and an active system (hexapod). Both systems cannot be directly linked due to the shape of the base of the hexapod (subjected to plate modes in high frequencies) and the required payload on isolators to get in its work range. From this assumption, a definition of an intermediate structure between isolators and the hexapod has been investigated. This leads to the structure and the passive isolation platform presented on Figure N.

3.2.3 Results - Impact of the passive isolation platform on dynamic behavior

The efficiency analysis of Paulstradyn-7 isolators has been based on measurements and Simulink models in order to get the expected behavior of the passive isolation platform in terms of accelerations, velocities and displacements in-flight conditions. The Figure O shows time responses and Power Spectral Densities of accelerations, velocities and displacements in-flight conditions for a vertical excitation (z-direction), i.e. input spectrum in the frequency range [10Hz - 1000Hz]. The vertical excitation is considered as the most pessimistic criteria to overcome.
Values associated to these time responses will be considered as input for the active isolation system (Sec.3.3).

3.2.4 Conclusion on the passive isolation system

We recall that the passive isolation platform aims at protecting electronics of the hexapod or the camera by filtering high frequency accelerations, which are not a drawback for the PIV measurements. The bandwidth of the three quantities is restricted to about 15 Hz with high amplitudes below this limit due to the passive platform resonance. More specifically, the acceleration spectrum is now much more narrowed since almost all its frequency content is cut-off. Finally, the impact of the filter on the acceleration becomes very low. Moreover, the maximal velocity is reduced from 67 mm/s to 34 mm/s and the remaining maximum displacements are lower than 0.65 mm. However we can see the negative impact on the displacement due to the filter resonance at 6.84 Hz. Indeed the maximum displacement went from 0.58mm at the base of the system (input spectrum [10Hz - 1000Hz]) to 0.63mm on the passive platform. This resonance affects displacements in the bandwidth [0-15Hz]. This effect should be erased by using a complementary active control system. In order to be pessimistic, results on vertical excitation (z-direction) have been considered for the two lateral excitations (x-direction and y-direction), in spite of the fact that the literature gives lateral excitations amplitudes equal to 15% of vertical excitations. Regarding the rotations (pitch, roll and yaw) results on the pitch have been considered for the roll and the yaw as well, always keeping in mind an pessimistic point of view with an input spectrum [10Hz - 1000Hz] quite significant. This leads to maximal rotations of 0.42deg and maximal angular velocities of 33.1deg/s.

3.3 Active isolation system

The objective of the active vibration control is to reject any remaining vibrations in order to get fixed PIV cameras and laser. Since the vibrations can affect the PIV elements by creating both translations along and rotations around these three directions, the dedicated system has to move in the 6 degrees of freedom. Since the Hexapod PI-840 has been used for Sec.1, the following study will evaluate this hexapod as an active isolation system. According to the simulation results presented in the conclusion on the passive isolation platform, the hexapod must comply with the dynamic specifications that are recalled in Figure P.

The following study aims at testing the hexapod PI-840 through its controller (no modification on the fly of the controller settings) and using its buffer trajectory mode. As a result, the identification of the hexapod is performed in closed loop through a step at 0.1mm which could be a typical displacement for the active control of the camera. From this step response, a model of the hexapod is performed. Conclusion on the Hexapod PI-840: For a reference input, the top plate of the hexapod reaches the target position within at least 25 ms. In order to perform a real-time active control in the bandwidth [0 - 15 Hz], the sampling time should be equal to 10 ms maximum. In addition to that sending target position from the Matlab environment adds few milliseconds. Consequently, without being able to modify on the y controller settings and without having a real-time system, the active control cannot be possible with this hexapod PI-840 with this specific controller. The manufacturer is currently developing a real-time system with a better dynamic.

3.4 Control with hybrid system (passive isolator and active control)

The last part of this study concerns the design of an active control law. This one must control the 6 DOF displacements. However, in order to validate the chosen methodology, the active control is considered only on vertical vibrations since they provide the highest vibration amplitudes. We now focus on a hybrid solution composed of dampers to attenuate high frequencies and an active device for low frequencies. The hexapod used for PIV measurements has a bandwidth of 20Hz and is evaluated in combination with the dampers studied for the passive control of the vibrations.
On Figure Q, the vibratory spectrum inserted in the loop is the one set in Sec.2 (Figure K) with vibratory amplitudes. The block corresponding at the passive isolation platform takes into account results presented in Sec.3.2.3.

Figure R shows the simulation results along the Z-axis. Vibration amplitude expectations appear (bottom Figure R) to be quite low (max 81 μm). This solution fits with the requirements defined in the Sec.1.3.2 in order to correctly reconstruct the velocity/displacement fields measured by the PIV sub-systems. Indeed, the remaining low frequencies are greatly attenuated compared to the solution with dampers only (Figure O). This solution allows increasing the performance of the vibration attenuation system but is costly in terms of weight and complexity. Moreover, the same procedure may be applied for the other degrees of freedom. Indeed, as mentioned before, all axis may be considered as uncoupled.

3.4.1 Conclusion on the hybrid isolation system

The methodology developed in this Sec.3.3 appears fitting with the displacement requirements (Sec.1.3.2) if hexapod's controller settings were commutable. However, due to the complexity to add an extra control loop, immutable controller settings and high settling times, this hexapod cannot be used as an active control system for now. The manufacturer is currently developing a real-time system with a better dynamic.
In order to find a solution available, this drawback has led us to find a suitable commercial solution more dynamic. For the past few years, Safran Electronics and Defense (formerly Sagem Defense Securite) has been developing the Euroflir gyrostabilized electro-optic but the cost of it does not fit in the ACcTIOM's budget.

4 Conclusion

To sum up, one recalls that the objectives of the ACcTIOM project were twofolds.
First, it aimed at developing and validating an innovative concept of active flow control system, dedicated to the attenuation of airframe noise and vibrations on a CROR propeller through the pylon wake mitigation. It relies on a boundary layer scooping/blowing strategy. First a multi parametric iterative design and optimization process was conducted thanks to 2D and 3D Reynolds Averaged Navier Stokes computations. It allowed designing the wind tunnel model of the flow control system-equipped pylon and pre defining the flow control parameters - scooping and blowing mass flow rates - to apply such as to mitigate the pylon wake. By experimentally varying these parameters, and taking advantage of the WT experiment results, based on pressure coefficient distribution along the WT model profile, transverse profiles of total pressure at various streamwise and spanwise positions in the wake of the pylon and on 3C-PIV transverse-plane measurements at various streamwise positions, an exhaustive cartography of the flow control system efficiency was determined. It allowed identifying nominal operating points, where the target acoustic criterion is satisfied. It has also highlighted the robustness of the embedded flow control system in fully mitigating the wake of the pylon, even under moderate scooping and small blowing mass flow rate variations. Finally, the computational approach has been validated and its transposition to the prediction of the wake mitigation for flight conditions can be assessed.
Second, it aimed at developing and advanced experimental methodology, based on vibration-controlled stereoscopic Particle Image Velocimetry (3C-PIV), able to be flight-operated inside the Flying Test Bench in order to validate the efficiency of the active flow control system in annihiltaing the pylon wake.
• 3C-PIV Calibration Test Bench developed and validated - Versatility transposable to various experimental test benches,
• Characterization of the error envelop of the 3C-PIV under influence of vibrations in FTB provided,
• Specifications for Vibration Control Methodology defined,
• VCM available: hardware stage made of hybrid system (passive isolators and active control) and software stage leading to clean database from still spurious images.
More specifically, the study has been performed as follow. Based of vibratory spectrum discussed with Airbus and on an exhaustive bibliographic investigation, we have defined a pessimistic vibratory spectrum for the cabin environment of the CROR engine-equipped FTB. On this basis, we wondered what would be the influence of displacements induced by this vibratory spectrum on PIV measurements. Consequently, we have developed a static PIV test-bench able to reproduce dynamic misalignment on a PIV subsystem (camera or laser) compared to another one for a given measurement. We recall that the most pessimistic criteria in PIV measurements is not the displacement of the laser-sheet but one camera compared to the other one, this is purely due to optical and aero-dynamical reasons. Indeed, if we admit that the laser moves while the two cameras are motionless, then acquired pictures are simply unfocused. Consequently, the software part of the Vibration Correction Methodology may easily identify and delete these ones. So, the static PIV test-bench only takes care of the camera displacement and may be performed thanks to different time scales between measurements (Sec.1.2). With this static PIV test-bench, a typical flow expected in-flight conditions has been synthesized and reproduced through the target displacements. Moreover, it allows moving one PIV subsystem (camera) from the other one along the 6 degrees of freedom and for a given step. From it, we have been able to create an envelop with a (standard deviation) uncertainty where the velocity/displacement fields are correctly reconstructed by the post-processing software. Beyond this limit of motion (translations or rotations), motion control is required. These displacements are related to the vibratory spectrum and lead to a crossover frequency setting two actions in order to mitigate displacements. Beyond this limit, we have used a passive filter and before, an active control system. However, we recall that a passive filter stands for protecting camera/hexapod electronics. Indeed, displacements in high frequencies are quite low and do not affect PIV measurements. In addition to that, the passive filter may increase low frequency amplitudes already high due to the pessimistic vibratory spectrum. Consequently, the amplification of the motion in low frequencies leads to penalize the active control system by giving more displacement constraints. If the chosen cameras used in the aircraft cabin can handle high frequency vibrations and the real vibratory spectrum amplitudes are lower than our pessimistic one, then the problem may simplified by not taking care of the high frequencies and the passive filter may be unused, leading to use only the active control system. However, if the vibratory spectrum in-flight conditions presents similar vibration amplitudes, then the passive filter and an active control system must be embedded (hybrid system).
Finally, depending on the in-flight vibratory results, embedded PIV subsystems might be well fixed in aircraft's cabin without any vibration control but with a powerful laser, optical lens widely opened (at least F8). The pessimistic version where the vibratory spectrum looks like the one defined in this ACcTIOM's project, consequently passive filter (if camera's electronics sensitive) and active control (hybrid solution) must be implemented. The proposed solution leads to a model and a control loop. To this avail, a hexapod with a better dynamic than the one used in this study (PI-840) is required or an Euroflir gyrostabilized electro-optic systems.

Potential Impact:
Potential impact

The expected direct impacts of the ACcTIOM project results are dual. First the development of an optimized CROR-propeller pylon and of its associated active flow control system will permit to mitigate the wake of the pylon before it impacts the first blade stage of the CROR propeller. As such it will alleviate the fluid-structure interaction-induced emission of noise of the future CROR-propelled aircraft. This work will directly serve the potential certification of this new generation of more energy-efficient aircrafts. This would thus reinforce the leadership and competitiveness of the main european aeronautical stakeholders, equipment manufacturers and subcontractors.
Second, the development of an advanced experimental methodology, based on vibration-controlled stereoscopic Particle Image Velocimetry (3C-PIV) able to be flight-operated, will enhance knowledge on the mechanisms driving the wake mitigation by the active flow control system. The latter could support the definition of future innovative engineering systems based on this technology. It will also participate to the certification process of this innovative flow control technology. Moreover mastery of this advanced measurement technique in harsh flight conditions will have widespread applications both for the characterization of complex airflows past specific airframe elements (wing/pylon/nacelle interaction for instance) or for other certification issues.

During the ACcTIOM project, several French or foreign companies (most of them were European, a few of them in the US) have been contacted as potential suppliers for innovative technological solutions dedicated to both the two major objectives of the ACcTIOM project. Amongst them, one can cite Physik Instrumente GMBH, Newport, Bilz Vibration Technology AG, SBG Systems, KVH Industries Inc., B&K, Dantec Dynamics. Other companies (LEICA, PI, Symétrie, Safran Electronics and Defense) were also contacted and several meetings were organized such as to define with them precise specifications for the adaptation and the modification of still existing products, but also early stage products, in particular gyrostabilized electro-optic. These actions may lead to future availability of commercial products able to satisfy the requirements of in-flight 3C-PIV measurements, as these companies have well understood the high potential of such new activities to expand their market.

Dissemination activities

The whole set of deliverables (D2.1, D4.2, D4.3, D5.4) and complementary intermediate reports provided during periods 1, 2, 3 and 4, the numerous meetings organized during these periods and the four periodic reports provided to both Airbus and the JU have strongly participated to the dissemination activities. Moreover several papers have been presented to international conferences and the submission of, at least, 2 articles to international peer-reviewed journals should be completed within the period Q4 2016 - Q1 2017. The latter are listed below.

still effective

A poster has been presented at the French identification working group(Groupe de Travail en Identification) which depends on the Modelization, Analysis and Control of dynamic Systems research group (GDR MACS).
2 papers have been presented at the 50th 3AF International Conference on Applied Aerodynamics, held in Toulouse the 30th, 31st of March and 1st of April 2015.
An article has been published in the Proceedings of the European Control Conference, held on July 16th 2015 in Linz, Austria.
2 papers have been presented at the Aviation 2016 Conference of the American Institute of Aeronautics and Astronautics on June 13th-17th 2016 in Washington, United States of America.
1 paper has been presented at the Applied Aerodynamics Conference of the Royal Aeronautical Society on July 19-21 2016, in Bristol, UK.
1 paper has been presented at the Greener Aviation of 3AF on October 11th-13th 2016 in Brussels, Belgium.

list of communications

- Bury, Yannick and Bordron, Alban and Belloc, Hervé and Prat, Damien. Development of an Innovative Active Flow Control System for CROR Powerplant Noise Reduction through Pylon Wake Mitigation. In 50th 3AF International Conference on Applied Aerodynamics, 29-30 March – 01 April 2015, Toulouse - France.

- Gourdain, Nicolas and Bury, Yannick and Dupont, Louis and Bodart, Julien. Large Eddy Simulation of a flow control device for noise reduction due to a CROR/pylon interaction. In 50th 3AF International Conference on Applied Aerodynamics, 29-30 March – 01 April 2015, Toulouse - France.

- Vayssettes, Jérémy and Mercère, Guillaume and Bury, Yannick and Budinger, Valérie. Structured model identification algorithm based on constrained optimisation. In Proceedings of the European Control Conference, Linz, Austria, 2015.

- Budinger, Valérie and Bury, Yannick and Michon, Guilhem and Napias, Gaël. In-flight PIV for CROR flight-test demonstration. In AIAA Aviation 2016, AIAA Aviation and Aeronautics Forum and Exposition, 13-17 June 2016, Washington DC, USA.

- Bury, Yannick and Bordron, Alban and Belloc, Hervé and Prat, Damien. CROR-powerplant pylon wake mitigation for noise reduction through innovative blowing/suction-based active flow control system. In AIAA Aviation 2016, AIAA Aviation and Aeronautics Forum and Exposition, 13-17 June 2016, Washington DC, USA.

- Gourdain, Nicolas and Bury, Yannick and Bodart, Julien. Large-Eddy Simulation and analysis of the controlled turbulent wake generated by a thick profile. In 2016 Applied Aerodynamic Conference of the Royal Aeronautical Society, 19-21 July 2016, Bristol, UK.

- Napias, Gaël and Bury, Yannick and Budinger, Valérie. In-flight PIV for CROR flight-test demonstration. In Greener Aviation 2016, 3AF Conference, 11-13 October 2016, Bruxelles, Belgium.

planned

1 article related to WP3 results, related to the active flow control system efficiency in erasing the CROR pylon wake, will be submitted at the AIAA Journal within the period Q4 2016-Q1 2017.
1 article related to WP4-5 results, related to the SFD approach, will be submitted at the Experiments in Fluids Journal within the period Q4 2016-Q1 2017.

Exploitation of results

1/ An operable and highly efficient active flow control system dedicated to the mitigation of the noise generated by CROR engines through CROR pylon wake alleviation

The optimized Counter Rotating Open Rotor (CROR) powerplant pylon and its associated innovative active flow control system developed in the context of the ACcTIOM project permit to fully erase the wake of the CROR powerplant pylon before it impacts the first blade stage of the CROR propeller. As such it alleviates the fluid-structure interaction-induced emission of noise of the future CROR-propelled aircraft. The prototype developed in the context of the ACcTIOM project directly serves the potential certification of this new generation of more energy-efficient aircraft, and consequently its commercialization. This concept has been patented by Airbus. The exhaustive description of the prototype has been transferred to Airbus. The technological adaptation from a wind tunnel operable prototype to an in-flight operable system is currently ongoing by Airbus. It should be exploited in the context of potential commercialization of the SFWA-based future aircraft.

2/ knowledge advancement on wall jet/outer flow interaction and underlying physics

From a fundamental point of view, results obtained during the ACcTIOM project, WP2 and WP3, foster knowledge on the physical mechanisms driving the mixing of wall jets with outer flows. The latter are involved in several engineering applications, e.g. flow control for drag reduction of automotive or flying vehicles, fuel mixing, anti-icing systems, demisting/defrosting systems, etc. There is no direct exploitation of products expected here, but the enhancement of the physical understanding of underlying phenomena directly supports the definition of future innovative engineering systems dedicated to the aforementioned applications.

3/ knowledge advancement on advanced optical methodology for the experimental characterization of fluid flows and their interaction with structures and systems.

Results of the ACcTIOM project, WP4 and WP5, are directly transposable to the application of in-flight, or more generally on-board, stereoscopic Particle Image Velocimetry (3C-PIV). There is no direct exploitation of products expected here, but an immediate impact on R&D activities involving the experimental characterization of airflows (or more globally fluid flows) past systems, through wind tunnel or onboard, potentially in-flight, test campaigns.

List of Websites:
relevant contacts :
Dr Yannick Bury - ACcTIOM project coordinator
Institut Supérieur de l'Aéronautique et de l'Espace (ISAE)
Département Aérodynamique Energétique et Propulsion
10, avenue Edouard Belin BP 54032 - 31055 Toulouse CEDEX 4
Tel : (+33)5 61 33 84 98 - Fax : (+33)5 61 33 84 63
yannick.bury@isae.fr

Related information

Reported by

INSTITUT SUPERIEUR DE L'AERONAUTIQUE ET DE L'ESPACE
France
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