EXPERIMENTAL INVESTIGATION OF ENDWALL AND SUCTION SIDE BLOWING IN A HIGHLY LOADED COMPRESSOR STATOR CASCADE

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1 Proceedings of ASME Turbo Expo 21: Power for Land, Sea and Air GT21 June 14-18, 21, Glasgow, UK GT EXPERIMENTAL INVESTIGATION OF ENDWALL AND SUCTION SIDE BLOWING IN A HIGHLY LOADED COMPREOR STATOR CASCADE Daniel Nerger Horst Saatoff Rolf Radespiel Institut für Strömungsmecanik Tecnisce Universität Braunscweig Bienroder Weg 3, 3816 Braunscweig, Germany d.nerger@tu-bs.de Volker Gümmer Carsten Clemen OE-1 Compressors and Fans Rolls-Royce Deutscland Ltd. & Co. KG Escenweg 11, Dalewit Blankenfelde-Malow, Germany volker.guemmer@rolls-royce.de ABSTRACT Te following paper describes an experimental investigation of a igly loaded stator cascade wit a pitc to cord ratio of t/l =.6. Experiments witout as well as wit active flow control by means of endwall and suction side blowing were conducted. Five ole probe measurements in pitcwise and spanwise direction as well as endwall oil flow visualiations were carried out in order to determine te performance of te cascade and to analye te flow penomena occuring. To quantify te effectivity of te active flow control metod, taking te additional energy input into account, corrected losses and an efficiency, wic relates te difference of flow power deficit wit and witout active flow control to te flow power of te blowing jet itself, were evaluated. Even toug an increase of static pressure rise could be acieved, a decrease of te total pressure losses was possible for a few operating points only. NOMENCLATURE c momentum coefficient d, blade tickness/eigt H k sape factor Address all correspondence to tis autor. meanwile Witt & Son AG meanwile Siemens AG, Energy Sector l cord lengt ṁ mass flow p, p t static/total pressure P flow power p/q 1 static pressure rise q dynamic pressure Q V volume flow t pitc W velocity β flow angle δ, δ boundary layer/displacement tickness ζ V 1 nondimensional total pressure loss ζ V 1,corr corrected nondimensional total pressure loss η B efficiency according to Bae [1] Θ momentum tickness λ stagger angle µ axial velocity ratio or outflow coefficient ρ density ϕ velocity coefficient ψ contraction number Subscripts 1,2 inlet/outlet conditions j jet quantities loc local sec secondary flow system 1 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 1 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

2 INTRODUCTION In te past few years several papers on active flow control in compressors via blowing ave been publised. Te classical metod applied is a tangential blowing over te profile suction surface to increase boundary layer momentum and terefore delaying or preventing separation, see e. g. [2, 3]. A more effective metod to avoid boundary layer separation is te use of vortex generator jets, wic not simply use te momentum input, but also increase mixing between te boundary layer and te outer flow. Tis metod until now mainly as been investigated in turbines and rater seldom in compressors, see e. g. [4]. Anoter metod used is trailing edge blowing. One option is to blow into te main flow direction filling up te wake. Tis approac reduces noise and provides a more uniform flow to te following blade row. Blowing under a certain angle to te main flow creates te so called jet flap. Working like a mecanical flap, circulation is increased, wile te losses can be reduced significantly. Fiscer et al. [5, 6] sow te effectiveness of tis metod. Tat active flow control via continuous blowing also is able to reduce secondary flow effects was sown by Döge [7] and Sceugenpflug [8]. In recent years tis idea was picked up and developed furter for example by Nerger [9], Mertens et al. [1] or Zander et al. [11]. Te present paper sows results for a tangential blowing over te profile suction surface and te endwalls applied to a igly loaded linear stator cascade, wic is caracteried by significant suction side boundary layer separation and secondary flow at baseline conditions witout blowing. Cascade Geometry and Boundary Conditions Te investigated stator cascade consists of tree igly loaded controlled diffusion airfoils (CDA), see Fig. 1. Tese profiles were designed for a turning angle of up to β = 6 wile te outlet flow angle is β 2 = (axial outflow). To reac tis ig flow turning originally a comparatively small pitc to cord ratio of t/l =.4 was cosen. Tis cascade sowed a rater good performance even toug strong secondary flow effects occured [12]. Performance improvements were acieved by reducing tese secondary flow effects introducing endwall blowing [9 11]. To test te blowing metod at increased blade loading te pitc to cord ratio was increased by fifty percent to t/l =.6 olding te oter parameters constant. Te geometric details of te cascade may be summaried as follows: blade section CDA, pitc to cord ratio t/l =.6, stagger angle λ = and aspect ratio /l =.8. Te blade section (cord lengt l = 25mm) as a maximum tickness to cord ratio of d/l =.8. Te wind tunnel was run at a constant speed of W 1 = 65m/s giving an inlet Mac number of Ma 1 =.2 and a Reynolds number, based on cord lengt and inlet velocity, of Re = Te inlet boundary layer is collateral wit a tickness δ = 1mm (δ/ =.5) at one tird of cord upstream of te cascade, and te corresponding integral parameters are δ = 1.25mm, Θ =.79mm and H k = EXPERIMENTAL FACILITY Cascade Test Rig Te cascade investigations were carried out in te new low speed cascade test rig [12] at te Institute of Fluid Mecanics of te Tecnisce Universität Braunscweig. Te oriontal test section widt (blade span) of te wind tunnel is = 2mm, its vertical eigt varies between 17 mm and 33 mm depending on inlet flow angle β 1, pitc to cord ratio t/l and blade count. Te inlet flow angle can be canged continuously by rotating te blade carriers, wile te stagger angle λ is usually eld constant at some prescribed value. Te cascade test rig can be installed into te.5 m wind tunnel of te institute. A periodic cascade flow is acieved by applying boundary layer suction upstream of te cascade at te ends of te oriontal test section walls, just above and below te end blades of te cascade. Additionally tailbords are installed above and below te cascade to guide te flow. Te suction mass flow and te position of te tailboards is controlled by te requirement of a constant static pressure distribution in a plane about one tird of cord upstream of and parallel to te cascade front. Note tat te boundary layers on te vertical test section walls (endwalls) are not removed by suction. Endwall Slots Blade Figure 1. STATOR CASCADE WITH GEOMETRICAL PARAMETERS AND MEASUREMENT PLANES 2 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 2 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

3 4mm Slot Cover 2mm 3mm 2mm 3mm 2mm 4mm Profile Suction Side 117.5mm 2mm Secondary Air Supply F T P Endwall Slots Passage 1 Figure 4. SLOT CONFIGURATION FOR SUCTION SIDE BLOWING Endwall Slots Passage 2 F T P Blade 1 Blade 2 Blade 3 Figure 2. TEST ARRANGEMENT WITH SECONDARY AIR SYSTEM; F VOLUME FLOW, T TEMPERATURE, P PREURE Profile Suction Side 22mm Blowing Direction 15 12mm Figure 5. CASCADE BLADE WITH SLOTS Figure mm Endwall SLOT CONFIGURATION FOR ENDWALL BLOWING Secondary Flow System and Blowing Configuration For te investigation of flow control by blowing te cascade test rig was equipped wit a secondary flow system, consisting of a blower and a main air tube. Tis tube divides into two smaller tubes, one for te suction surface blowing and one for te end- wall blowing, bot being trottled separately. Eac tube consists of a swirl flow meter (Trio-Wirl S, ABB), static pressure taps and a temperature probe. Finally bot tubes are furter divided up and connected to te four endwall slots (one at eac endwall and pitc) and to te six blade connections (one at eac end of te tree blades). In Fig. 2 te full test arrangement can be seen. Te endwall slots (Fig. 3) are positioned at 12% cord, wic is close to te suction peak, te starting point of te secondary flow. Te slot widt to local main flow direction is 12 mm and te widt to local main flow direction 22mm. Perpendicular to te blowing direction te slot eigt is 3.1mm. Te blowing direction is inclined by 15 to te endwall and parallel to te main flow direction, wic means tangential to te blade suction side at tis position. Te slots directly start at te profile suction side and are oriented towards te middle of te passage. Te position 3 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 3 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

4 and sie of te endwall slot configuration used ere was found to be most effective to reduce secondary flow for te cascade wit a pitc to cord ratio t/l =.4 [9]. Witin te profile suction surface of eac blade tree single slots are located at 47% cord lengt, a position resulting from a numerical sensitivity analysis (Figs. 4 and 5). Eac slot lengt is 2mm and te corresponding widt is 2 mm. Tese slots are distributed uniformly along blade eigt. It was found tat tis configuration, aving individual slots starting at a certain distance from te endwalls, gives te best performance for suction surface blowing, because very close to te endwall te effectivity of te blowing decreases rapidly. Te blowing direction again is inclined by 15 to te tangent of te suction side at te blowing position and parallel to te endwalls. EXPERIMENTAL PROCEDURES Five Hole Probe Traverses Traverse measurements were performed wit a calibrated five ole probe in pitcwise (y direction) as well as in spanwise direction ( direction) one tird of cord downstream of te cascade for different inlet flow angles β 1 (Fig. 1). Te step sie between measured points was y = 4mm and = 1mm, respectively. Additionally two sections close to te endwalls at = 4 mm and = 196 mm were analyed. First local values, depending on te y and location were calculated from te measured pressure differences. For te cascade caracteristics te following parameters were considered: i) outlet flow angle β 2, ii) nondimensional static pressure rise p/q 1 wit p = p 2 p 1, iii) nondimensional total pressure loss ζ V1 = p t /q 1 wit p t = p t1 p t2 and iv) axial velocity ratio µ= W 2x /W 1x. Te inlet flow pressures were measured wit a Pitot static probe (q 1 = p t1 p 1 ) about one tird of cord upstream of te cascade. Te inlet flow angle is fixed by te setting of te cascade. Te accuracy of te measured data analyed for te present paper is ±.1 for flow angles and ±.2% for pressures. In Tab. 1 te uncertainties of all investigated parameters are given. To take te additional total pressure input wit blowing into account te loss coefficient as to be modified as follows Table 1. Parameter Unit Statistical error β 2 1.3% ±.1 p/q % ζ V % 1 5.2% µ 1 2.5% η B 1 9.% ṁ 1 kg/s.3% ṁ j kg/s 1.1% W 1 m/s.3% W j m/s 7.% UNCERTAINTIES OF MEASURED PARAMETERS in pitcwise and spanwise direction. For te present cascade flow it is appropriate to calculate an area average for te axial velocity ratio and te static pressure rise, a mass average for te total pressure loss and a mixed area mass average for te outlet flow angle [13, 14]. Additionally, te blowing efficiency was evaluated wit a metod suggested by Bae [1], comparing te flow power deficit wit active flow control to flow power deficit witout blowing related to te flow power of te blowing jet itself: η B = P de ficit P de ficit, j P j 1% (3) P de ficit = ṁ1 ρ (p t1 p t2 ) (4) (y,) = p t1 p t2 (y,) q 1 (1) P de ficit, j = ṁ1 + ṁ j (p ρ t1 p t2 ) (5) wit p t1 = ṁ1 p t1 + ṁ j p t j ṁ 1 + ṁ j. (2) Te jet total pressure p t j is defined at te slot outlet. Based on te local values described above integral values were derived P j = 1 2 W 2 j ṁ j. (6) 4 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 4 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

5 Slot location Loss coefficient Outflow coefficient ζ Vs µ ϕ Endwall Suction side Velocity coefficient 32 p t s [Pa] 24 Experiment Correlation Experiment Correlation Table 2. CORRELATED OUTFLOW COEFFICIENTS AND LO CO- EFFICIENTS OF THE SLOTS 16 Suction Side Endwall Calculation of Jet Quantities For te calculation of te loss coefficient wit blowing ζ V 1,corr and te efficiency η B te jet total pressure and te jet velocity ave to be known. Tese parameters can not be measured directly during te experiment, instead te volume flow Q V, te temperature T sec and te static pressure p sec witin te secondary flow system ave been determined. From tese parameters te missing variables may be calculated. Tis can be done using a semi empirical metod based on equations for compressible outflow in combination wit experimentally determined slot coefficients. Te parameter describing te slot losses is te total pressure loss coefficient ζ V s. Tere are two main loss sources for te outflow troug a slot: i) a jet contraction wic reduces te mass flow being described by te contraction coefficient ψ and ii) te friction described via te velocity coefficient ϕ. Te product of bot gives te outflow coefficient µ. Te coefficients ζ Vs, µ and ϕ are known for a lot of simple geometric orifices like oles [15], but not for te special slots used ere. Tey ad to be determined experimentally performing a comparison between values measured witin te secondary flow system and values measured directly at te slot outlet using a tree ole probe for five different blowing mass flows [9]. Te flow downstream of all tree slots at te suction side were measured and te results averaged afterwards. Te same measurement was done for te two endwall slots. Losses between te measurement position and te slots witin te secondary flow system can be neglected in comparison to te slot losses. Slot loss coefficient and outflow coefficient are cosen in tat way, tat te experimental total pressure losses and te teoretical and experimental contraction number are equal. Te results are given in Tab. 2. In Fig. 6 te slot total pressure losses in dependency of te blowing mass flow are sown. Te black symbols mark te measured values and te wite symbols mark te results calculated wit te determined slot coefficients. Next to te good agreement a strong gradient of loss increase can be found wit increasing mass flow especially for te (small) suction side slots. Te relation between mass flow and outlet velocity is given in Fig Figure m j [kg/s] JET MA FLOW m j /m 1 3. [%] Figure 7. SLOT TOTAL PREURE LOES IN DEPENDENCY OF Endwall Suction Side W j [m/s] RELATION BETWEEN JET MA FLOW & JET VELOCITY Oil Flow Visualiations For te visualiation and analysis of te boundary layers at te endwalls and on airfoil suction as well as pressure surfaces oil flow pictures were produced for different inlet flow angles β 1. Oil flow pictures are potograps of specially prepared sur- 5 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 5 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

6 faces sowing te surface pattern of te sear stress lines. Te direction of tese lines is almost identical wit te flow direction of te fluid next to te surface. Following te criteria described by Tobak and Peake [16], te more important features of te flow may be identified, as for example: i) tree dimensional separation lines (convergence of te sear stress lines), ii) tree dimensional (re)attacment lines (divergence of te sear stress lines) and possibly iii) so called singular points (saddle and nodal points, foci). Wit tese features, a first idea of te flow field may be developed. RESULTS Basic Flow witout Active Flow Control Five Hole Probe Traverses Considering te passage averaged static pressure rise and total pressure losses in Fig. 8 in comparison to te small pitc to cord ratio [12] it is remarkable tat at β 1 < 56 te performance of te cascade is better, even te loading as increased. Losses are smaller up to an angle of β 1 = 58. A reason for tis is a distinctly reduced secondary flow, wic will be discussed later but also te smaller ratio of wetted to passage area. On te oter and te available working range moves to smaller inlet flow angles wit increasing pitc to cord ratio [6]. For ig inlet flow angles β 1 a massive suction side boundary layer separation results in decreasing static pressure rise and increasing total pressure losses. As mentioned before cascade performance is affected by different flow caracteristics at low and ig inlet flow angles. Fig. 9 sows te distribution of te cascade caracteristics along blade span for different inlet flow angles β 1 = 6 (reference), β 1 = 56 and β 1 = 58 (low incidence) as well as 62 (ig incidence). Te outlet flow angle distribution sows overturning close to te endwalls for all inlet flow angles affecting nearly 25% of blade span. At midspan β 2 = 8 is significantly larger tan te design intent of β 2 =. Te total pressure loss distribution sows, next to an increase wit increasing inlet flow angle, a redistribution witin te passage. Wile at β 1 = 56 te loss level close to te endwalls is larger tan at midspan, for increased angles tis distribution reverses. For β 1 > 58 te losses at midspan are larger tan close to te endwall. As will be sown later te reason for tis is a massive suction side separation wic becomes more and more dominant compared to te classical secondary flow. Te iger pitc to cord ratio decreases te secondary flow developement but promotes suction side boundary layer separation, especially at ig incidence angles. Tis effect also results in reduced static pressure rise and stronger static pressure variations along blade span. p/q 1 is smallest at midspan due to iger outlet flow angles and iger total pressure losses. Te differences become more serious wit increasing inlet flow angle. Te axial velocity ratio µ is caracteried by an increased trougflow close to te endwalls and.7 p q Figure 8. t/l =.4 t/l =.6 ζ V1 p.32 ζ V β 1 PAAGE AVERAGED CASCADE CHARACTERISTICS FOR DIFFERENT INLET FLOW ANGLES β 1 WITHOUT ACTIVE FLOW CON- TROL blockage at midspan. Wit increasing inlet flow angle β 1 tis difference is more and more distinct. In Figs. 1 and 11 contour plots for te outlet flow angle β 2 and te total pressure loss ζ V 1 are sown for β 1 = 56 and β 1 = 6, respectively. Te profile suction side is oriented downwards as in te experimental setup and te trailing edge position is indicated. Te measurements were conducted for one pitc only around te central blade of te cascade and duplicated for presentation. Discrepancies visible at te interface, especially for te β 1 = 6 results are due to te influence of te Pitot static probe mounted just upstream of te cascade inlet. Wen moving from te endwall towards midspan close to te blade trailing edge, first a region wit outlet flow angles of β 2 < is passed and afterwards a region wit very ig flow angles of β A classical secondary flow wit corresponding passage vortices is responsible for tis beaviour. Te vortex axis is perpendicular to te measurement plane and te vortex core is located at 25mm. Additionally a more significant over- and underturning, respectively can be observed parallel to te trailing edge. Tis is a separation vortex, due to a suction side boundary layer q1 6 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 6 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

7 8 β ζ V1 β 1 = 56 β 1 = 58 β.16 1 = 6 β 1 = 62 β ζ V1 [1] β 1 = 56 β 1 = 58 β 1 = 6 β 1 = p q 1 µ β 1 = 56 β 1 = 58 β 1 = 6 β 1 = β 1 = 56 β 1 = 58 β 1 = 6 β 1 = Figure 1. CONTOUR PLOTS OF OUTLET FLOW ANGLE β 2 (LEFT HAND SIDE) AND LO COEFFICIENT ζ V 1 (RIGHT HAND SIDE) WITHOUT ACTIVE FLOW CONTROL AT β 1 = 56 Figure 9. DISTRIBUTIONS OF CASCADE CHARACTERISTICS ALONG BLADE SPAN WITHOUT BLOWING separation, wit an axis parallel to te measurement plane [17]. Te vortex axis as to be upstream of te measurement plane because no backflow was measured. Especially te suction side boundary layer separation results in a region wit ig total pressure losses wic nearly affects 5% of te blade passage. Oil Flow Visualiations In order to gain furter insigt into te tree dimensional flow field of te present cascade, oil flow visualiations were conducted on te profile suction and pressure surface as well as on te endwalls for all inlet flow angles investigated. Representative results for β 1 = 54 and β 1 = 6 will be discussed ere. Potograps of te profile suction side and one endwall are depicted in Fig. 12 were selected flow caracteristics ave been igligted. Te oil flow visualiations sow a distribution of te wall sear stress lines, wic represent a good approximation of near wall streamlines. Te flow direction is from top to bottom in te suction side potos (top frames) and from left to rigt in te endwall potos (bottom frames). Te first dominant flow feature occuring at β 1 = 54 (left and side) on te suction side is a separation bubble between 1% cord (laminar separation) 15% cord (turbulent reattacement). At te suction peak position a separation line (SL1) starts at te endwall (saddle point SP1), due to a collision of te near suction side flow wit an endwall cross flow directed from pressure to suction side. Te endwall cross flow is driven by te β Figure 11. ζ V1 [1] CONTOUR PLOTS OF OUTLET FLOW ANGLE β 2 (LEFT HAND SIDE) AND LO COEFFICIENT ζ V 1 (RIGHT HAND SIDE) WITHOUT ACTIVE FLOW CONTROL AT β 1 = 6 pitcwise pressure gradient sifting low momentum fluid along te endwalls towards te suction side, forming a system of classical secondary vortices. Furter downstream tese vortices are situated above te blade suction surface away from te endwalls due to te ig aerodynamic loading and te sape of te CDA pressure distribution. Te separation line SL1 develops towards 7 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 7 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

8 Figure 12. OIL FLOW VISUALIZATION AT PROFILE SUCTION SIDE AND ENDWALL WITHOUT BLOWING FOR β 1 = 54 (LEFT HAND SIDE) and β 1 = 6 (RIGHT HAND SIDE) Mass flow Velocity Momentum Coolant Density ratio ratio coefficient Mac ratio Number ṁ j /ṁ 1 W j /W 1 c j Ma j ρ sec /ρ loc Endwall.8% % % % Suction Side.8% % % % Table 3. BLOWING PARAMETERS midspan ending in te focus F1 togeter wit a second separation line SL2 wic is caused by a collision of te main flow and an axially backward directed trailing edge flow (between 25% and 75% blade span). Near te endwall te flow direction is axially forward towards blade midspan. Approacing te separation line SL1 te direction canges to axially backward until te flow meets te focus F1. Tis region is caracteried by low momentum fluid and can be seen as a wite line in te oil flow picture. Due to te symmetry of te flow field te same flow features can be found close to te opposite endwall. Bot endwall flows meet at midspan indicated by one saddle point SP2 on te spanwise oriented separation line SL2. At te endwall (Fig. 12, bottom frame) it can be seen tat te pitcwise extension of te separated flow region, wic is limited by separation line SL3, is comparatively small due to te severe pressure gradient. Tis flow beaviour is similar to te corresponding flow found for a small pitc to cord ratio as described in [12] and [1]. For β 1 = 6 (rigt and side) an unsymmetric flow field can be observed. On te suction side only one focus is visible and te backflow region at midspan already starts at 45% of cord. Additionally flow directed axially backward occures at te rigt endwall, wile at te opposite side te flow is still directed axially forward. Te backward flow interacts wit a separation vortex at te endwall. A similar endwall separation in a compressor stator was described by Saatoff et al. [18]. Te extention of te separation region in pitcwise direction as grown but is again relatively small. Te transition via a separation bubble still persists. Flow wit Active Flow Control Five Hole Probe Traverses In Figs te cascade performance caracteristics are sown for suction side blowing only, endwall blowing only and blowing at four blowing mass flow rates, respectively. Te reference values at baseline condition are marked as oriontal lines. All blowing parameters are given in Tab. 3, were te momentum coefficient is defined as c j = (ṁ j W j )/(q 1 l). Te results for te static pressure rise coefficient (Fig. 13) sow tat using endwall blowing alone to reduce te secondary flow is not able to increase p/q 1. In contrast, suction side blowing results in a pressure rise increase by up to 8%, but tese improvements only can be observed for blowing mass flows ṁ j /ṁ 1 > 1.3%. Te described increase of te static pressure rise first sows a very steep gradient wic flattens for iger blowing mass flows aving a maximum at ṁ j /ṁ 1 = 2%. Wit te blowing an even furter increased static pressure rise can be acieved. However tis is only possible wit anoter additional 1% of blowing mass flow. Te reason for te current performance improvement wit endwall blowing in comparison to only endwall blowing is tat te suction side blowing reduces suction side boundary layer separation and due to tis te classical secondary 8 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 8 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

9 .75 p/q % EW.8% 2% EW 1% 1% EW 2% 2% EW 2% η B % EW.8% 2% EW 1% 1% EW 2% 2% EW 2% m j /m 1 [%] Figure 13. STATIC PREURE RISE p/q 1 AT β 1 = m j /m 1 [%] Figure 15. BLOWING EFFICIENCY η B AT β 1 = 6.42 ζ V1,.36 β % EW 2% % EW 1% % EW.8% % EW 2% p q µ m j /m 1 [%] Figure 14. TOTAL PREURE LOES ζ V 1,corr AT β 1 = 6 ; WHITE SYMBOLS CORRECTED, GREY SYMBOLS UNCORRECTED flow becomes more dominant again. Furter rising te blowing mass flow at te endwalls results in a decreasing static pressure rise. Using a smaller amount of air for te suction side blowing no increase of p/q 1 above te baseline level is possible. A net loss reduction (Fig. 14), taking into account te jet total pressure, was almost impossible. Only for te suction side Figure 16. DISTRIBUTIONS OF CASCADE CHARACTERISTICS ALONG BLADE SPAN WITH BLOWING AT β 1 = 6 ; ONLY SUCTION SIDE ṁ j /ṁ 1 = 1.6%, ONLY ENDWALL ṁ j /ṁ 1 = 1.6%, COM- BINED ṁ j /ṁ 1 =.8%+.8% 9 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 9 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

10 β 2 [1] β 2 [1] Figure 17. CONTOUR PLOTS OF OUTLET FLOW ANGLE β 2 (LEFT HAND SIDE) AND LO COEFFICIENT ζ V 1,corr (RIGHT HAND SIDE) WITH ACTIVE FLOW CONTROL AT β 1 = 6, ONLY ENDWALL ṁ j /ṁ 1 = 1.6% Figure 18. CONTOUR PLOTS OF OUTLET FLOW ANGLE β 2 (LEFT HAND SIDE) AND LO COEFFICIENT ζ V 1,corr (RIGHT HAND SIDE) WITH ACTIVE FLOW CONTROL AT β 1 = 6, ONLY SUCTION SIDE ṁ j /ṁ 1 = 1.6% blowing te ζ V 1,corr for ṁ j /ṁ 1 > 1.7% is marginally below te baseline level. If te jet total pressure is not taken into account te uncorrected total pressure losses ζ V 1 are reduced for all configurations wit increasing blowing mass flow. For pure endwall blowing te uncorrected losses first sligtly increase before decreasing, wile at pure suction side blowing continuously decreasing losses can be obtained. Te reason for tis are different velocity ratios between jet velocity and local velocity. Up to ṁ j /ṁ 1 = 1.2% te velocity ratio at te endwall is W j /W loc < 1.3 (W loc 9m/s) (see Fig. 7). Te additional low momentum fluid results in an additional weakening of te endwall boundary layer and terefore increases te total pressure losses. At te suction side te velocity ratio for all blowing mass flows investigated is ig enoug, due to iger jet velocities as well as due to a slot position furter downstream wit W loc 36m/s. For minimal losses te velocity ratio of jet velocity to local main flow velocity sould be sligtly above one. Except for suction side blowing wit ṁ j /ṁ 1 > 1.6% te efficiency of all blowing configurations is below ero (see Fig. 15). Tis means tat a worsening of te cascade flow wit blowing is present. In te following te distributions of te cascade caracteristics along blade span given in Fig. 16 for a constant total blowing mass flow of ṁ j /ṁ 1 = 1.6% will be discussed. Te static pressure rise p/q 1 not only sows a cange in sie compared to te baseline but also a more uniform distribution. Te increase towards te endwalls is limited to 5% span wile along te remaining blade span te level is nearly constant. A more uniform total pressure loss distribution and a reduced overturning close to te endwalls are reasons for tis enanced uniformity. An outlet flow angle β 2 = occures up to 1% span closer towards te endwall for all tree configurations. Te canges at midspan owever are very different. For pure suction side blowing te flow angle stays at te reference value of 4, wile it drops for te oter two configurations. In particular using pure endwall blowing very small outlet flow angles of 5 are reaced. Responsible for tis beaviour is an increased suction side separation vortex wic will be sown later in te oil flow visualiations. As mentioned before also te total pressure loss distribution is more uniform for all tree blowing configurations. Especially for endwall blowing te results sow a very smoot distribution altoug at a ig level. At te wake position of te blowing jets a sligtly decreased total pressure loss can be observed in form of a small dent at 1% blade span. For te blowing at te same position a local loss maximum can be seen, resulting from te lower momentum endwall jet in tis case. For te suction side blowing a local total pressure loss minimum can be found at midspan, witin te wake of te middle slot. Tis minimum is due to te fact, tat te outer jets bend towards midspan, driven by te secondary flow, ence amplifying te middle jet. Tis again will be sown later by oil flow visualiations. Te distribution of te axial velocity ratio µ sows a maximum trougflow at midspan for suction side blowing and at te endwalls for endwall blowing. For te blowing local trougflow maxima can be found at midspan as well as at te endwalls. Looking at te contour plots for te outlet flow angle β 2 for 1 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG 21 1 Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

11 β [1] comb., EW+,.8%+.8% comb., EW+, 1%+2%, 2% p q p q Figure 19. CONTOUR PLOTS OF OUTLET FLOW ANGLE β 2 (LEFT HAND SIDE) AND LO COEFFICIENT ζ V 1,corr (RIGHT HAND SIDE) WITH ACTIVE FLOW CONTROL AT β 1 = 6, COMBINED ṁ j /ṁ 1 =.8%+.8% te tree configurations in Figs , a region wit underturning and overturning is visible along te blade trailing edge. As mentioned above in te discussion of te baseline results tis is due to te separation vortex. Wit blowing te vortex is less uniform and distorted by te blowing jets. Close to te endwalls reduced overturning can be seen. Despite te outlet flow angle variation te total pressure loss distribution sows only little variations along blade span. Small regions wit reduced losses can be detected, wic mark te position of te blowing jets. Wile in te measurement plane te endwall jets stay very close to te endwalls te profile jet as separated from te suction surface forced by te corresponding boundary layer separation. Since no clear optimum for a blowing configuration could be determined, for tree different configurations a variation of te inlet flow angle β 1 was investigated. Te results are sown in Fig. 2. Te area averaged static pressure rise p/q 1 distribution sows tat until β 1 = 58 for all configurations a rise in p/q 1 can be acieved, wereas a pure suction side blowing is less effective. As mentioned above in te discussion of te baseline results at tis operating point te classical secondary flow is dominant and endwall flow control is necessary. A blowing mass flow of ṁ j /ṁ 1 = 1% already is enoug to acieve a good performance, wile furter increasing te mass flow only sows little effect. Wen increasing te inlet flow angle to iger values more blowing mass flow at te suction side is necessary to improve te static pressure rise, in doing so tese improvements are significant. If te suction side blowing mass flow is too small a reduced β 1 Figure 2. PAAGE AVERAGED CASCADE CHARACTERISTICS FOR DIFFERENT INLET FLOW ANGLES β 1 WITHOUT AND WITH BLOWING performance compared to baseline conditions can be observed. Area averaged corrected losses are far above te baseline. Only wit pure suction side blowing around β 1 = 6 a small reduction is possible. For very ig inlet flow angles te losses severely increase, wile te static pressure rise decreases significantly at te same time. Summing up blowing gives te best performance, wereas te ratio between suction side blowing and endwall blowing as to be increased towards part load conditions. Oil Flow Visualiations Finally te oil flow visualiations on te profile suction side will be presented for te tree configurations wit a total blowing mass flow of ṁ j /ṁ 1 = 1.6% (Fig. 21). Te flow direction in eac frame is from top to bottom. For suction side blowing only (top frame) a reduced extension of te separated flow region at midspan can be observed. It starts at 75% cord in comparison to 4% cord at baseline conditions. Te extension in spanwise direction is comparable. Downstream of te slots areas wit ig wall sear stresses sow te direction of te blowing jets. It can clearly be seen tat te outer jets are bend towards 11 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

12 midspan by te secondary flow coming from te endwalls. Tis results in an amplified mid region jet, wic could also be observed in te five ole probe measurements. Close to te endwalls at bot sides axially forward flow is found. Te results for blowing (bottom frame) are similar, wereas te extent of te suction side separation is less reduced due to suction side blowing. Because of tat te outer jets are bend earlier towards midspan. A totally different picture can be found for pure endwall blowing (middle frame). Te separation region already starts at 5% cord and extends nearly along te wole blade span. Reducing cross passage flow at te endwalls results in an increased loading of te profile suction side at mid span. SL 1 SL 2 CONCLUSION Experimental investigations of a igly loaded linear stator cascade wit a pitc to cord ratio t/l =.6 were conducted. At reference conditions two different flow patterns could be observed. For lower inlet flow angles up to β 1 = 58 a classical secondary flow wit a passage vortex was dominant. Even toug te total pressure losses due to tis secondary flow were ig a good static pressure rise was acieved. Wen increasing te inlet flow angle te flow beaviour canged to a massive suction side boundary layer separation. To improve te cascade performance an active flow control in terms of blowing was introduced. Since te flow beaviour canges wit increasing incidence a combination of blowing at te endwalls and at te profile suction side were used. Wile te endwall blowing effectively reduces te classical secondary flow, te suction side blowing reduces te profile boundary layer separation. A blowing was found to be most effective wen canging te blowing mass flow rate wit te cascade operating point. Altoug an increase of te static pressure rise above te reference level was possible in a wide range, te corrected losses, taking into account te blowing jet total pressure, could be reduced for a few operating points only. Wit te presented level of knowledge an application of a blowing configuration in a real compressor to reduce secondary flow effects seem to be far away. Te difficulty is, to determine te rigt parameters, wic are necessary for an effective flow control via blowing. Te most important ting is to know te dominant secondary flow features of te macine at baseline condition, wic are mostly affect te macine performance. ACKNOWLEDGMENT Te work was funded by te German Government and te Ministerium für Wirtscaft und Tecnologie as part of te Lufo III programme, Förderkenneicen 2T38. Te management of Rolls-Royce Deutscland Ltd. & Co. KG is gratefully acknowledged for supporting te work and permitting te presentation of results. Figure 21. SL 1 SL 1 SL 2 SL 2 OIL FLOW VISUALIZATION OF PROFILE SUCTION SIDE WITH ACTIVE FLOW CONTROL AT β 1 = 6 ; TOP ONLY SUCTION SIDE, ṁ j /ṁ 1 = 1.6%; MIDDLE ONLY ENDWALL, ṁ j /ṁ 1 = 1.6%; BOTTOM COMBINED, ṁ j /ṁ 1 =.8%+.8% REFERENCES [1] Bae, J., 21. Active control of tip clearance flow in axial compressors. P. D. tesis, Massacusetts Institute of Tecnology. [2] Fottner, L., Teoretical and experimental investigations on aerodynamically igly loaded compressor bladings wit boundary layer control. Proceedings of te 4t ISABE. [3] Kirtley, K. R., Graiosi, P., Wood, P., Beacer, B., and 12 c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG

13 Skin, H.-W., 24. Design and test of an ultra-low solidity flow-controlled compressor stator. Proceedings of ASME Turbo Expo, Paper No. GT [4] Culley, D. E., Brigt, M. M., Prast, P. S., and Straisar, A. J., 24. Active flow separation control of a stator vane using embedded injection in a multistage compressor experiment. ASME Journal of Turbomacinery, Vol. 126, January, pp [5] Fiscer, S., Saatoff, H., und Radespiel, R., 25. Verdicterprofile mit Stralklappen für den Einsat in aerodynamisc oc belasteten Statorgittern. Tagungsband ur 21. Strömungstecniscen Tagung, TU Dresden, S [6] Fiscer, S., Saatoff, H., and Radespiel, R., 28. Two dimensional RANS simulations of te flow troug a compressor cascade wit jet flaps. Aerospace Science and Tecnology, Vol. 12 (No. 8), pp [7] Döge, K. H., Hocdruckaxialventilatoren mit Grenscictbeeinflussung. Dissertation, TU Dresden. [8] Sceugenpflug, H., 199. Teoretisce und Experimentelle Untersucungen ur Reduierung der Randonenverluste ocbelasteter Axialverdicter durc Grenscictbeeinflußung. Dissertation, Universität der Bundeswer Müncen. [9] Nerger, D., 29. Aktive Strömungsbeeinflussung in ebenen Statorgittern oer aerodynamiscer Belastung durc Ausblasen. Zentrum für Luft- und Raumfart, TU Braunscweig, ZLR Forscungsberict 29-8, (Dissertation, TU Braunscweig). [1] Mertens, D., Swoboda, M., Huppert, A., and Tiele, F., 28. Transition modeling effects on te simulation of a stator cascade wit active flow control. Proceedings of ASME Turbo Expo, Paper No. GT [11] Zander, V., Hecklau, M., Nitsce, W., Huppert, A., and Swoboda, M., 29. Active control of corner vortices on a igly loaded compressor cascade. Proceedings of 8t European Turbomacinery Conference, Paper No. C157. [12] Nerger, D., Saatoff, H., and Radespiel, R., 27. Experimental and numerical analysis of a igly loaded low aspect ratio compressor stator cascade. Proceedings of 7t European Turbomacinery Conference, Paper No. A129. [13] Greiter, E. M., Tan, C. S., and Graf, M. B., 24. Internal flows Concepts and applications. Cambridge University Press. [14] Cumpsty, N. A., 26. Averaging nonuniform flow for a purpose. ASME Journal of Turbomacinery, Vol. 128, January, pp [15] Bol, W., 21. Tecnisce Strömungslere, 12. Auflage Vogel Bucverlag. [16] Tobak, M., and Peake, D. J., Topology of treedimensional separated flows. Annual Review of Fluid Mecanics, 14, pp [17] Ivey, P. C., and Swoboda, M., Leakage effects in te rotor tip-clearance region of a multistage axial compressor, part 1: Innovative experiments. Proceedings of ASME Turbo Expo, Paper No. 98-GT-591. [18] Saatoff, H., Deppe, A., Stark, U., Rodenburg, M., Rokamm, H., Wulff, D., and Kosyna, G., 23. Steady and unsteady casingwall flow penomena in a single-stage low-speed compressor at part-load conditions. International Journal of Rotating Macinery, Vol. 9, pp c Copyrigt Rolls Royce Deutscland Ltd. & Co. KG Copyrigt 21 by Rolls-Royce Deutscland Ltd. & Co KG