Monitoring of velocity and pressure fields within an axial turbine Pierre Duquesne1, Monica Iliescu1, Richard Fraser1, Claire Deschênes1, Gabriel Dan Ciocan1 1
Hydraulic Machinery Laboratory, Faculté des sciences et de génie, Université Laval 1065 Avenue de la Médecine, Québec, G1V 0A6, CANADA,
[email protected] Abstract In the framework of the consortium of R&D on hydraulic machines launched by the LAMH (Hydraulic Machines Laboratory of Laval University) in November 2007, a five holes unsteady pressure probe using embedded pressure sensors has been developed. Such probe allows us to obtain the unsteady static pressure, total pressure and the three flow velocity components. This paper aims to present this new five holes unsteady pressure probe, used and results for a propeller turbine. Keywords: Pressure probe, unsteady phenomena, hydraulic propeller turbine
1. Introduction Axial-T, the first project of the Consortium on Hydraulic Machines aims to document the flow in an axial hydraulic turbine. These measures are used for validation of CFD simulations. Different measurement instruments were used such as PIV, LDV and an unsteady pressure probe with 5 embedded sensors. This pressure probe developed internally gives access to three-dimensional velocity measurements and total and static pressure measurements. This pressure probe has been used at the rotor-stator interface and runner outlet for different points of operation. This paper focuses on both runner outlet measurement and the energetic balance of the runner. The main characteristics and use of the probe are first described. The experimental setup and positioning system will be presented. Results found for three operating points will finally be presented.
2. Experimental means 2.1 Unsteady pressure probe design and technical specifications A Pitot probe gives simultaneous information on the static and dynamic components of a steady pressure field. However, there are two major limitations in using standard Pitot tubes in hydraulic turbomachinery. The first one is the necessity to be perfectly aligned with the flow stream to provide correct readings on both pressure and velocity. The second one is the impossibility to get information on unsteady phenomena because the pressure is transmitted from the pressure taps to sensors through a fluid column acting as a damper absorbing fluctuations. The average value is correct but the amplitude of fluctuations includes a bias error. The first limitation is overcome by adding 4 pressure taps to the classical Pitot tube to determine the alignment in the flow [1]. The lateral pressure taps are used to position the probe relatively to the flow and the central one to measure the total pressure. In this configuration, referred as the nulling technique, the probe is oriented in the flow until equal values are obtained on each peripheral pressure port, i.e. the probe is aligned with the fluid. The total pressure is then directly the reading of the central sensor, while the static pressure is derived from the lateral sensors. Geometric asymmetries are considered in the calibration. Achieving accurate data requires, however, complex positioning systems not readily adapted to turbomachinery. E.g. the size of the insert would modify locally the boundary layer topology, and possibly impact on the mean flow field, forcing a dissymmetry. The external size of the displacement system has also to be considered, mainly in areas with restrained access such as spiral casing discharge ring. To overcome the second limitation and extend the investigation range to unsteady flows, a new probe has been designed (Fig.1) incorporating embedded sub-miniature piezoresistive transducers with a dynamic response of over 50 kHz. Those sensors are typically 10 times more sensitive than the traditional ones. Their diameter of 2 mm allows a compact design into a pyramidal enclosure of 8 mm equivalent diameter, close to the size of standard five holes probe. At the design stage, the four lateral sensors have been placed at 45˚ from the supporting arm axis, in order to minimize the effect of the probe arm on the top sensor. More flexibility on the probe positioning in the flow is necessary to obtain unsteady values. The method uses a parametric dependency of the differential pressure obtained with lateral sensors on the flow direction. Those functions are modelled through normalised calibration coefficients, measured in a well controlled flow. In steady flows, the relative orientation of the flow © Pierre Duquesne, 2010. Published in Engineering Conferences Online (ECO): http://eco.pepublishing.com DOI: to be inserted by the publisher
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Fig1. Geometry of the 5-sensors pressure probe and reference frame
Fig2. Sensor unbalance concept
2.2 Probe calibration and dynamic range At calibration stage, four normalised coefficients, characterizing the unbalance of pressure on the 5 sensors with respect to the flow direction, are mapped for a range of positions (α, τ) of the probe in the flow: 5
P −P P −P P − Pm P −P F (α ,τ ) = 2 3 ; G (α ,τ ) = 4 5 ; L (α ,τ ) = t ; H (α ,τ ) = 1 t ; Pm = P1 − Pm P1 − Pm Pt − P0 Pt − P0
∑P i
i=2
[2]
(1)
4
According to Fig.1, α denotes rotation around the axis parallel to the probe supporting arm, and τ denotes rotation around an axis perpendicular to the first one. The origin of the spherical reference frame is located in the centre of mass of the probe. The coefficients F and G are defined as differences between pressures of opposite lateral sensors (Pi with i=2,3,4,5), normalized by the difference between the pressure of the central sensor (P1) and the mean pressure in the plane passing through the 4 lateral sensors (Pm). The coefficients L and H characterize the deviation of the central sensor reading (P1), and of the average pressure of lateral sensors (Pm), respectively, from the local stagnation pressure. Both L and H are normalised with the local dynamic pressure (Pt – P0). The probe can be calibrated either in a fully developed flow such as a hydrodynamic tunnel, either in a water or air jet with potential velocity profile. For the present study, both air and water potential jets have been used [3] [4]. Reference value of total pressure had to be adjustable and was continuously monitored throughout the calibration. An angular domain from –25 ° to +25 ° was swept in both directions α and τ , with a resolution of 2.5°, see [3]. An example of the 4 parametric surfaces obtained is shown in Fig 3. The angular limits have been set according to the shape of the calibration charts for an extended domain. Peaks in the F extended map (Fig. 3-right), indicate that flow separation occurs for angles higher than 25°.
Fig3. Normalized calibration coefficients charts and validation of the angular range on F (Morachioli [3]) A significant requirement to use the probe in unsteady flows is the independence of the calibration coefficients on the Reynolds number. This has been verified by performing calibration for the range of velocities expected in the axial turbine: 2 to 10 m/s. Based on the calibration, the accuracy of the probe in terms of velocity is ±2% for angles smaller than 10°, and ±3.8% for an angular range of [–25°, +25°]. In terms of pressure, the accuracy is ±3.4% and ±5.8%, respectively, for the same angular ranges [4].
3. Test case and accuracy assessment 3.1 Experimental setup Experiments were performed on a low-head axial turbine with 24 stay vanes, 24 wicket gates and 6 blades propeller – see [5]. Both the intake and the elbow draft tube are asymmetrical and are split in two channels by off-axis piers. The global characteristics of head, flow and efficiency have been measured according to the IEC 60193 [6] standard. Across the operation range, three conditions were investigated close to the peak efficiency, at constant head and variable flow rate. The specific speed nQE ranges from 0.418 at 94% rated opening, to 0.424 for the design point, and up to 0.443 for overload – see fig. 5. The global parameters were controlled systematically within 0.2 % during the measurement campaign. Static and dynamic pressure surveys were carried-out in two sections, at the runner inlet and outlet: 25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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Upstream the runner, the probe moved on a radius from the shroud toward the hub with a spatial resolution of 5 mm. Considering the reference frame of the machine (right-handed frame with the runner axis Z pointing upwards, X downstream, and the horizontal plane XY crossing the distributor at mid-height), the measurement axis is tilted at 30° from the vertical axis, and is located at 123.5 ° from the X axis in the circumferential direction – see Fig. 4. - At the runner outlet, the probe was installed at 65% of the runner diameter under the blade’s axis – see Fig. 4. The circumferential position was 33.5° from the X axis, in the runner rotation direction. The probe traversed from the cone wall toward the axis by 5 mm steps. To move the probe, a positioning system with two degrees of freedom has been developed. Maximal error is 0.64° in rotation and 0.08 mm in translation. This system is automated with stepper motors and controlled with Labview. As previously discussed, the probe is constrained to an angular range of ±25° for both α and τ, to prevent flow separations. However, in the actual measurement setup, the orientation in τ direction is limited by the traverse system to a fixed position. It is thus a constraint of the access to the measurement zone. The position of the traverse system has been optimised such as to keep the probe within the calibrated τ angular range with respect to the flow direction, along the full travel range of the linear drive and for all operating conditions. Operating point
2 3 4
N11 Q11 Guide vane N11op 3 Q11op 3 opening [°] [%] [%] 31 100 96 33 100 100 38 100 107 Fig5. Operating points
Relative efficiency [%] 97 100 91
Fig4. Experimental setup: positioning of the probe traverse system at the runner outlet (A) and inlet (B). Bottom left inserts show the probe position in the flow for the two locations: below the runner (C) and in the rotor-stator gap (D) 3.2 Data acquisition and reduction Data were sampled at a rate of 1 kHz at the runner outlet and 2 kHz at the rotor-stator interface, by blocks of 2 sec duration. The relative runner position was recorded simultaneously, through a 0-5 Volts ramped signal with an accuracy of 0.1% of the full scale reset by the top encoder signal delivering 1 pulse/rotation. Thus the unsteady fluctuations of the velocity and pressure fields, synchronous with the runner position, are captured. An angular resolution of 4°on runner position was suitable to insure a good convergence of the phase averaged values. For each spatial position, the nulling method is used to determine the optimum orientation of the probe in the average flow direction with respect to α, while the τ position is fixed. However, if the unsteady fluctuations of the flow direction exceed the probe range, the average value may be biased by these out of range values: - All instantaneous data fall in the probe’s angular range – accurate estimation of the mean value - Less than 10% of the unsteady values are out of range – biased average, but the value can be corrected at post-processing stage by rejecting the outliers - Flow direction exceeds the calibration range for more than 10% of the unsteady values - bias error on the average value and statistics are no longer reliable, measure at this location should be discarded. To overcome this limitation, data are recorded for three different α positions for a same radial position of the probe. First, the probe is oriented in the mean flow direction, α=0°, then rotated to α=40° and α=-40°. The span of the probe is thus enhanced from 50° to 130° in the α direction, but remains 50° in the τ direction. This procedure is called the scan procedure. Thresholds are set for F and G coefficients according to the calibration charts, and all values out of range are rejected. The percentage of unsteady values in the probe range is calculated for each spatial position and the scan procedure is applied when necessary. 3.3 Uncertainty analysis Fig.6 presents circumferentially-averaged velocity profiles measured with the 5-sensors probe and 2D-LDV at the runner outlet, on the same axis and in the same operating conditions. Only the axial the axial and tangential components were captured by LaserDoppler measurements, therefore the comparison is made for those two components only. The radial positions are normalized by the local radius of the measurement cross-section, while velocities are referred to the free stream velocity in the reference section of the machine – discharge ring. The LDV-measured velocity profile is statistically accurate on 98% of the section, with 2% uncertainty. In the median region, from 30% to 80% of the local radius of the measurement section, the velocity profiles match the LDV mean value within 5%, while differences are noticeable near the wall and near the rotation axis, corresponding to higher gradients. Close to the wall, from 80 to 100% of the local radius, the velocity profile measured by LDV shows much steeper gradients. 25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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DOI: to be inserted by the publisher This behaviour is mainly justified by the physics of the decelerating swirling flow in a straight diffuser, in the vicinity of the boundary layer. Given the axial location of the measurement section with respect to the runner, adverse pressure gradients are likely to occur at that level in the cone, and the flow is decelerating near the wall. 1,2
Normalized velocity : C/Cref
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Fig6. Comparison of 2D-LDV and probe velocity field Fig7 Number of valid data used to get the averaged downstream the runner phase-resolved values For the pressure probing, the hydraulic profile is respected, but due to its intrusive nature, the probe induces a hydraulic resistance proportional to its penetration depth into the flow. The shapes and sizes of both the probe head and the supporting arm have been optimised for generating minimal losses, while insuring the required mechanical strength with respect to the pressure levels encountered in the measurement sections. The drawback is a lower performance in high gradient regions: near the wall, the velocity gradients are virtually undetectable, mainly due to the probe head size. Below the hub, for radii less than 30%, the viscous effects are predominant and force the turbulence level to rise. A deficit of 79.4 % of the freestream velocity is present in the time-averaged axial velocity profiles with respect to the axial distance from the hub. On the pressure probe-measured profiles, the radial velocity shows a change in direction towards the axis at 30 % of the local radius. In this zone, the axial velocity deficit is underestimated by 50%, while the tangential velocity profile shows an accelerated swirling motion of twice the one measured by LDV, followed by a deceleration. The lower gradient of axial velocity, combined with the overestimated swirl intensity, triggers a deviation of 45% of the pressure probe-measured profiles from the LDV profiles. Furthermore, the stagnation region in the wake of the hub, extending up to 7 % of the local radius, is not detected by the probe. Globally, the difference between the probe measurements and the laser-measured profiles is attributable to the unsteady variations in the flow direction in zones of gradients in the mean flow field, which may exceed the calibration range. This behaviour is due to the dissymmetry of the swirling flow downstream the runner. Using the scan procedure in the centre of the section to increase the angular range of the probe (Fig.6) proved to be effective in recovering the axial velocity deficit, but discrepancies in the angular component profiles still persist. This behaviour is supported by the rate of valid data used in the calculation of the phase-averaged values, displayed in Fig.7 and Fig.11 for the optimum point. In regions close to the axis, 20% Rref, data are statistically less accurate due to the smaller amount of valid data for averaging.
4. Flow analysis 4.1 Flow field at the runner outlet A pressure probe survey of the unsteady velocity and pressure fields has been performed in a cross-section downstream the runner for three operating conditions. Both sides of the efficiency curve for a constant head were investigated for three guide vanes opening, 31°, 33° near BEP and 38°. The axial component of the velocity field and the total pressure field synchronous with the runner rotation frequency are presented in fig.8. Data are normalised by the free stream velocity in the reference section at the runner outlet, and the pressure values are referred to the specific kinetic energy in this same section. For the point nearest to the peak efficiency, 33° opening, the axial velocity component and total pressure are evenly distributed in the radial direction. Velocity deficits of 50 % have been observed in the axial velocity near the wall and at the center. Two elementary Batchelor vortices superimpose on the base swirling motion: one located at 10% Rref, and a second one centered at 20%Rref [8]. The swirl motion is amplified near the axis, indicating a vortex rope with low axial velocity, but no stagnation region. At partial load, 31° opening, the flow is concentrated in an annular section between 60 % and 80 % of the local radius with 35% of total flow. Proper alignment of the probe with the flow was difficult at an extent of 25% of the local radius, consistent with a low pressure (11.7%Pref) below the hub. It indicates possible radial fluctuations in the flow direction, which may exceed the calibration range for τ. This issue cannot be corrected in the current measurement configuration, since the traverse system supporting the probe does not allow tilting. This area is less extended for the other two operation conditions. For overload conditions, guide vanes opening of 38°, a counter-rotating vortex develops near the axis, consistent with a drop in axial velocity below the hub and a low static pressure (10.1 Pref). The steady axial velocity profile is smoother than for the previous conditions. The circumferential velocity vanishes for radii between 40-50% Rref, indicating a purely axial flow in that region. A co-rotating swirl is present from 50% Rref to the wall. 25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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Fig.8. Axial velocity and total pressure fluctuation fields downstream the runner Fig.9. Average axial and tangential field. Fig.10 displays the radial distribution of turbulent kinetic energy (ktur) in the survey section, for the 3 guide vane openings. Data are calculated with the random unsteady fluctuations of the velocity components and referred to the average specific kinetic energy at each radial location. Differentiation is made between radial positions for which a single position of the probe in the flow was necessary (square markers), and locations for which the probe was rotated at +/-40° (plain dot markers). 16
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Fig. 11. Rate of valid phase-averaged values for the Fig.10. Turbulent kinetic energy evolution downstream the runner reference operating point for the 3 operating conditions under investigation The contribution of velocity components to the mean turbulent kinetic energy is unevenly distributed along the radius, as well as in the azimuthal direction depending on the runner position. Unsteady fluctuations of the flow direction are often accompanied by significant increase of the random fluctuations, which is the case from 50% to 85% of the local radius. The rise in turbulent kinetic energy in this region is consistent with the turbulent mixing process necessary to dissipate the shear flow generated at the runner outlet. The non-uniformity in the velocity profiles incoming to the diffuser are due to combined effects such as the circumferential dissymmetry of the flow at the leading edge, the unbalanced blade loading from hub to tip and the shear layer formed at the blade trailing edge. The decay of circumferential defect of the relative velocity depends on the inflow conditions and thus varies with the flow rate. During the project, PIV measurements conducted in the cone [8] showed that the runner exit sheared flow is strongly dissipated in the current measurement position. While for 31° and 33° openings, the runner wake has not vanished completely, higher mixing is present at overload 38° and higher spatial diffusion is observed.. A similar behaviour is detected in the centre of the section. The peak of the turbulent kinetic energy, with respect to the local kinetic energy, indicates high fluctuations of the flow direction in the hub wake. Spectral analysis under the hub (below 20% of Rref) indicates a vortex frequency of 20% the runner rotation frequency. The radial diffusion of the wake, the vortex rotation frequency and the axial velocity defect vary with the operating point. Nevertheless, unsteady fluctuations in the flow direction may bias the phase-averaged data. The percentage of valid phase-average data considered for the estimation of the time-averaged quantities for the optimum operating condition is presented in Fig.11. The threshold value is calculated according to the Student 25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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Fluctuations corresponding to the runner blades passage are noticeable at the bottom of the measurement section on both static and total pressure profiles, up to 15 % of the local cross-section
-
The profiles become uniform from 85 % of the cross-section depth up to the hub wall, the distance along the meridian streamlines increasing gradually from both the guide vanes trailing edge and the runner blades leading edge.
Static effects of the non-uniform incoming flow and dynamic effects induced by the runner are difficult to discriminate in the current experiment, since the measurements are done in the inertial frame and a single circumferential location. The circumferential fluctuations synchronous with the runner position are, however, predominant at the bottom of the section, where the probe head is in the proximity of the blade tip. Few alignment issues have been encountered, and globally more than 98% of the acquired data were in the calibration range.
Fig.12. Turbulent kinetic energy profile at the rotor-stator interface and frequency analysis for the 3 velocity components at 20%Rref near the hub, and 94%Rref near the blade tip
Fig.13. Evolution of the fluctuation amplitude at the blade passing frequency: tangential velocity upstream the runner
The radial distribution of the turbulent kinetic energy upstream the runner is represented in Fig.12. The measurement locations cover 80% of the cross-section, starting at 15% from the hub wall to 95% near the bottom ring. The mean turbulent kinetic energy in the inlet cross-section represents 1.52% of the kinetic energy of the free flow in the reference section, with a standard deviation of 0.25%. The turbulence repartition is balanced along the local radius, but phenomena that contribute to this turbulence level are distinct, as mentioned above. The spectral analysis of each velocity components, Fig.13, shows the domination of the blade passage frequency close to the periphery, at 94% of the local radius. The amplitudes vary between 2.5%, 3% and 4% of the freestream velocity, for the binormal, tangential and normal components, respectively. However, in absolute values, the tangential velocity is the most sensitive component to blade passage in the rotating frame. According to Fig.14, the circumferential fluctuations in tangential velocity vanish rapidly towards the hub, with increasing distance from the runner blades. Amplitudes decrease from 3.5% near the bottom ring, when the blade is near the probe, to 0.5% in the upper 60% of the section. The amplitude spectra near the hub show no 25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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Fig.14. Normalised total (A) and static (B) pressure fields upstream the runner for the near BEP point. Measurement zone and total pressure result (C).Contribution of potential energy, kinetic and static pressure terms to the energy balance of the runner (D) 4.3 Energy balance of the runner The probe measurements provide all the necessary information for estimating the specific energy available to the runner. Integration of the energy conservation law over a control volume limited by the two measurement sections upstream and downstream the runner is thus possible. Steady axisymmetric flow at both inlet and outlet sections is considered for the energy balance. The unsteady effects and circumferential non-uniformities are thus not taken into account. The specific energy transferred to the runner is, in this case, the difference of averaged total pressure profiles integrated over the inlet/outlet cross-sections, weighted by the mass flow through these sections (4):
Aoutlet P ρ ⋅ Cz ⋅ n ⋅ dA t ∫ ∆E = A = ∫ Cz ⋅ n ⋅ dA A
Ainlet
∫ (P A
0
Aoutlet / ρ + Cz 2 / 2 + gZ ) ⋅ Cz ⋅ n ⋅ dA ∫ Cz ⋅ n ⋅ dA A
(2)
Ainlet
Based on a ratio of 90% of the total surface covered by the experiments in the inlet and outlet sections, the specific energy extracted by the runner is 87.2%. Extrapolating the results to 100% of the sections, it means an efficiency of 96.8% for the runner. These estimations are possibly biased because several effects are not considered: - The extra 10% portion of the sections not covered by the experiments are localised near walls and in the axis of the hub. Strong gradients of axial velocity and the hub vortex are thus not accounted for in the present calculation - Measures were made for a single azimuthal position. Previous studies on the same turbine give evidence of a circumferential dissymmetry of the steady flow field, both at the interface [8] and downstream the runner [10]. Head losses in the runner with respect to the design assumptions are typically caused by: - The blade loading is non-uniformly distributed from hub to tip at the inlet and outlet sections - The viscous effects of the sheared flow at the runner inlet and outlet, of vortices propagating in the interblade channels [9], and of tip vortices are greatly damped, with respect to the distance of the measurement sections from the runner leading and trailing edges According to the IEC code 60193, three terms contribute to the energy balance: the specific pressure energy, the specific kinetic energy and the specific potential energy, as indicated by the expression: 1 1 (3) E = ( Pabs1 − Pabs 2 ) + (C12 − C2 2 ) + g ( Z1 − Z 2 ) 2 ρ Since the probe gives the static pressure field; the specific pressure energy can thus be easily computed. With the spatial position for each measurement, the potential energy is accessible. To estimate the kinetic energy, only the velocity normal to the section is considered. We estimate the kinetic energy term as the difference between the total energy obtained by integration of the total pressure profiles and the sum of the other static pressure and the spatial position. For operation point 3, the kinetic energy is estimated at 37% of the total energy, the pressure energy at 57% and finally the potential energy at 6% for the measurement section. See Fig.14.
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5. Conclusion In this paper, unsteady pressure probe developed and calibrated at Laval University with an accuracy of 3.8% and 5.8% on velocity and total pressure was described. The results for a propeller turbine measurement campaign at both runner inlet and outlet were presented, data and discussed. A comparison with LDV probing shows a superposition within 5% except for high velocity gradient and important variation of flow direction zones. Special procedure has been developed to overcome this last limitation by extending the probe range from ±25° to ±65° for α angle. Flow analysis for each position and operating condition provides information on both blade's wake and passage, turbulent kinetic energy on each component and both velocity deficit and low static pressure zone under the hub. Finally, through the energetic balance based on the measurements performed with this dynamic five holes pressure probe, the runner efficiency has been found to 96.8%. The repartition of the energy is 6% of potential, 36% of kinetic and 58% of pressure for the near peak efficiency condition.
Acknowledgments The authors would like to thank the participants of the Consortium on Hydraulic Machines for their support and contribution to this research project: Alstom Hydro Canada, Andritz Hydro LTD, Edelca, Hydro-Quebec, Laval University, NRCan, Voith Hydro Inc. Our gratitude goes as well to the Canadian Natural Sciences and Engineering Research Council who provided funding for this research.
Nomenclature
g
Alpha angle like in Fig1 Average water density Tau angle Global area of measurement zone Best Efficiency Point Radial velocity component Reference velocity, define for the runner diameter Standard deviation for radial velocity Tangential Velocity components Axial and debiting Velocity component Specific hydraulic energy of machine Gravity constant
ktur
0.5 Cu' 2 + Cr' 2 + C'z2
LDV
Laser Doppler velocimetry
α ρ
τ A
BEP Cr Cref C r'
Cu Cz E
(
)
nQE N11 N11OP3 Pabsi
PIV Pi Pt P0 Kturb Q11 R Rref Zi 1 2
Specific speed Turbine speed at nominal gate opening N11 for operation point 3 Absolute static pressure at surface i (i=1,2) Particle Image Velocimetry Pressure at sensor i (i=1,2,3,4,5) see fig1 Total pressure Static pressure Turbulent kinetic energy Turbine flow rate at nominal gate opening Radius Total radius for measurement zone (X translation) Vertical spatial position Surface of high pressure (interface) Surface of low pressure (runner outlet)
References [1] Bryer, D.W. and Pankhurst, R.C., 1971, Pressure-probe methods for determining wind speed and flow direction, National Physical Laboratory, London. [2] Ciocan, G.D., 1992, "Contribution à l’analyse des écoulements 3D complexes en turbomachines”, Ph. D. Thesis, Laboratoire des Ecoulements Géophysiques et Industriels de Grenoble - Equipe Turbomachines, Grenoble. [3] Morachioli, L., 2006, "Développement d’une sonde de pression 5 trous instationnaire avec capteurs embarqués", Msc. Thesis, Hydraulic Machines Laboratory of Laval University, Québec. [4] Duquesne, P., Deschênes C., Iliescu, M., Ciocan, G.D., 2009 "Calibration in a water potential jet of a five-holes pressure probe with embedded sensors for unsteady flows measurement", Int. Conf. of Experimental Mech., ICEM 2009, Singapore. [5] Deschênes, C., Ciocan, G.D., De Henau, V., Flemming, F., Huang, J., Koller, M., Arzola Naime, F., Page, M., Qian, Ruixia., Vu, Thi., "General overview of the AxialT Project: a partnership for low head turbine developments",2010, 25th IAHR Symposium on Hydraulic Machinery and Systems, Timisoara, Romania. [6] International Standart IEC 60193, 1999. [7] Susan-Resiga, R., Ciocan, G. D., Anton, I., Avellan, F., 1984, "Analysis of the Swirling Flow Downstream a Francis Turbine Runner", Journal of Fluids Engineering, Vol. 128, January 2006, pp. 177-189 [8] Gagnon, J.M., Iliescu, M., Ciocan, G.D., Deschênes, C., 2008 "Experimental investigation of runner outlet flow in axial turbine with LDV and stereoscopic PIV" 24th IAHR, FOZ DO IGUASSU [9] Beaulieu, S., Deschênes, C., Iliescu, M., Ciocan, G.D., 2009 "Study of the Flow Field Through the Runner of a Propeller Turbine Using Stereoscopic",3rd IAHR International Meeting of the Workgroup on Cavitation and Dynamic Problems in Hydraulic Machinery and Systems, Brno, Czech Republic [10]Gagnon, J.-M., Flemming, F., Qian, R., Deschênes, C., Coulson, S., 2010,"Experimental and Numerical Investigations of Inlet Boundary Conditions for a Propeller Turbine Draft Tube", ASME- FEDSM 2010, Montreal, Qc
25th IAHR Symposium on Hydraulic Machinery and Systems, September 20-24, 2010, Timisoara, Romania
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