Computational fluid dynamics (CFD) analysis for the reduction of impeller discharge flow distortion

https://ntrs.nasa.gov/search.jsp?R=19970018371 2019-05-20T21:40:21+00:00Z NASA-CR-204260 Computational Fluid Dynamics Impeller Analysis (CFD) Discharge /'Y/_/_ _/_f_ for the Reduction of Distortion Flow R. Garcia and P. K. McConnaughey Marshall Space Flight Center A. Easfland Rocketdyne Division, Rockwell International Introduction The use of CFD in the has increased in recent Marshall The Space team's methodologies and and form that baseline designed high the consortium accuracy applying the appropriate NASA/MSFC CFD risk. of this stages The design Team efficiency because used STME have lead team membership, PSTT is one of three CFD designed in the for the The STME a two-stage design was fuel pump design to a reduction predicted was as predicted with minimal risk. characteristic pattern at the the high head of the design and The that pump deemed was to pose 2). The weight, and consortium data showed cost impeller that the that a two-stage design for the STME test data also verified another CFD was not desirable. impeller discharge was strengthened coefficient necessary for the required classical The by two aspects pressure rise fuel "jet-wake" of the design: and by the by relatively few impeller exit blades, 12, necessary to reduce manufacturing cost (figure 2). This "jet-wake" pattern produces an unsteady loading on the diffuser vanes and has, in past rocket engine programs, lead to diffuser structural failure. In industrial applications, impeller reduction performance significant typically this problem the diffuser and of diffuser losses increases makes this vane is typically to ,,'low unsteady avoided the by increasing dissipation loading. the space of this pattern and, This and, more importantly in rocket in the pump's size and weight. approach unacceptable in high approach leads between hence, the the to small engine applications, to This latter consideration performance a whose (table in parts, of the fuel an impeller had the STME reached production. One sample of the baseline (figure 1) was manufactured and tested in a water rig. The test impeller performance pump was possible in Propulsion design Space turbopump. to design 1). (table to understand for Applications applied CFD ('PSTT) of several methodology(s) was fuel PSTT would first impeller (STME) for a two-stage allowed availability This Engine's impeller three developmental performance Technology and Consortium The PSTT impeller. Main with Stage the (McConnaughey3). Transportation Pump (MSFC) flow inside a pump. The PSTT's objectives, are discussed in Garcial and Garcia2. The improve the past activities teams then of high performance rocket engine pumps has been aided by the activities of the assessing include and Technology Center Flight goals and analysis This increase design years. rocket engines. T r After all, to reduce one of the motivations pump weight. The baseline impeller, diffuser radius while totally vanes if the typical were to be maintained. and performance loss, performance losses is especially This not in the unacceptable, impeller The because STME's have pump directly To achieve goal, this decided that a parametric study members using six different parameter codes. CFD same inlet and exit baseline consortium analyzed had the same head as the on of comparison flow the During diameters impeller. could distortion for subsequent individual team calculations. All experimental Laser Space Shuttle impeller the study in all but area vane leading it was design, actual used Engine (SSME) case. one vaneless edge. The be conducted. the accessed. that sizes Pressure benchmark domain In the actual space between of this ignoring hardware Fuel impeller vaneless space it and instead had the case vaneless was space run of the expansion baseline downstream consortium of the effect on the impeller. two the (HPFTR) that well. boundary conditions and studied passage impeller Typical expansion edge been by impeller from data very is a rapid modeling on demonstrated trailing base determined against using the of the of the as the code Turbopump and there the were their recently width with "slip" boundaries was adequate to ensure that these earlier conclusions were one used activities analyzed A listing benchmarked most on designs an analysis were used with previously impeller the in the as having in the diffuser PSTT a for analyzing the impeller. applicable to the current impeller impeller. which The included inclusion vaneless space, while not changing the impeller flow significantly, did effect on the velocity profiles at the radius corresponding to the diffuser leading edge. However, none of the calculations included the diffuser potential the for the PSTT became to and a minimized blade , flow distortion were predicted for the REACT3D code. the affect grid experience datasets, High These The been (L2F) impeller exit by Prueger5 found have Velocimeter schematically in the continuous However, times designed to produce the done so that the effect of and were This was be adequately on past based (Brozowski4). 3 shows flow and codes the and reported used calculations. member Main tests performance results are Figure three should cases analyzed appears on table 3. Each member performed baseline consortium impeller and the results of that analysis each losses. performance consisted of evaluating the effect of six different geometric parameters performance. The study included the participation of seven PSTT This study the impeller each it was thick cycle, approximately horsepower of the LOX pump. Therefore, the new challenge design a high head coefficient impeller with low blade count to blade (b-t-b) velocity distortion. was of the impeller entail a to engine absorbs which impeller for the necessitated engine gas-generator leads turbopumps fuel would diffuser gap of 3-5 percent thicker diffuser vanes would of the of the true the baseline consortium predicted and measured for designing "jet-wake" pattern the of the have a large vane vanes or its ! Results The effect of a given parameter on the impeller both distributions of key performance variables variables. Figure coefficient and 4 shows the flow Ideally, split Figure should Any mass flow imbalance the diffuser vanes. Not partial blade, but as possible. h-t-s Non because Small h-t-s From figure b-t-b distortion studied studied task. space not did International chordwise is often Further, significantly Research vaneless affect exit impeller to the other The cases conditions. simplification. ARC that the the form Loading (Rkdn) was blade loading in distortion design but large of the diffuser vane. distortions an inverse relationship by comparing efficiency, figures which ARC was the b-t-b was 4 and very to perform is not the Rocketdyne on high was the be sections, discussed. downstream it as a constant an analysis INS3D-UP is affected distribution can most which the subsequent will 5 it the task of evaluating downstream of the by modeling code preserving immediately in the work done that occurs when Distributiorf: assigned distribution geometry distribution h-t-s of the was assigned immediately simplified used while distortion Center (ARC) space geometry performance. boundary Blade there distortion. 6b). There is a slight increase blade trailing edge unloading analysis. Chordwise that variation design on be as uniform only a few significantly changed the flow split, impeller at 49 percent/51 percent. Therefore, in all the but blades. and efficiency are not necessarily proportional. One and an accompanying increase in efficiency. Overall, major 6a shows Figure vaneless (figure in the in the slip angle for in the In the impeller (or short) the grouping on of this flow affects impeller. impeller with be seen h-t-s direction variation Ames actual validity h-t-s parameter Space: of the head parameter. non-uniformity consortium Vaneless the effect area of the baseline six global should geometric of the was direction of the properties one impeller b-t-b the each the partial changed desirable the of the is a measure a spanwise and impeller, and in the baseline parameters parameter in the 5 it can can be seen that distortion have increased distortion concepts on the distortion side in the can be accounted between the baseline nearly ideal on either parameter on global direction. Similarly, the hub-to-shroud (h-t-s) of the distortion of the flow vector in the it represents distortions effect the distribution uniformity cannot. of the be equal distortion in the b-t-b is a measure direction. diffuser b-t-b of the various 5 shows was evaluated using "global" performance across the partial blades will generate a dynamic load only should the flow be balanced on either side of the velocity the The the velocity parameter impact the efficiency. performance and integrated span to assess (Kiris6) by the of inclusion significantly the to perform this of the affected by the impeller due to a delay the vaneless space is included Division, Rockwell task of evaluating the effect on the impeller performance of changing the and distortion. Rkdn axial the CFD code REACT-3D (ChanT) to perform this task and the study to be discussed in the following section. The variations used length where: (1) shifting shifting the was one there either blade's no clear of these distortion does lead blade's Axial loading loading towards change in the two changes. was not significantly to a more favorable Impeller impact the long long and leading the trailing edge. also either. (figure varied the distortion. The low The 5). impeller's baseline axial decrease or eliminate impeUer's discharge Because degrees this axial low momentum region. B2, was decreased width, edge has by 20 percent of this latter change, the impeller blade exit angle was from the baseline's 38 degrees to maintain a consistent again, performance the was the b-t-b was decreased figure 5, however, that distortion. Also, not significantly increasing impeller the rotordynarnics of a pump. Therefore, used as a method for improving the to b-t-b distortion affected by these weigh will axial length of the on determining the its study, 8) but (figure from h-t-s is a detriment definitive the diffusion. (figure 9). Note has the worst if increasing the axial length of the baseline design, then considerations heavily case to 41.5 Once increaseed head rise. generally to momentum to reduce changes by the increased impeller axial length the case with the lowest b-t-b distortion due b-t-b proposed that an of curvature and portion this that to determine a low It was radius (2) loading length design For and 8 shows efficiency that the leading near the long blade leading edge along the shroud. in the impeller's axial length would increase the region increase edge, Figure impeller's head coefficient and From figure 9 it is also evident affected flow split Rkdn Length: on performance towards the impeller modeled axial to the impeller is in addition length. Tandem Blading: SECA was assigned the task of evaluating the effect of cutting the long "impeller blades near the leading edge and clocking one portion of the blade relative to the other. Earlier studies in the PSTT had indicated that tandem blades were to be successful, SECA ran two degrees, the cut cases where respectively, performed in the use high side The of the long performance energy 11). (figure of the opposite from to the tandem the because that should were be near rotated of the impeller's blade's blade. SECA used the decreased significantly (Notice blade blades to the remainder relative direction flow long tandem the the long blade. side This to energize of the grid mapping, the angular general small conclusion concepts dominated design is that do not. are in other applications by rotational forces. were caused relative In retrospect, by either the clockings this because The low rapid rotation being was to suction the reference This will be corrected for the final paper.) penalty in that the h-t-s distortion increased clockings 22.5 code FDNS3D (ChenS) to perform this task. (figure 10) and the b-t-b distortion increased same for these three cases. changes suffered a double large edge. and The logic rotation. pressure leading 7.5 degrees concept may may improve not of the streamline the And these two as well. The flowfield, be as effective the flow in a radial momentum regions bending the is not but as similar impeller is primarily seen in the baseline from the axial to ¢ the radial direction pressure gradients. or by the secondary flows set up by the pressure-to-suction Partial Blade Chord Length, Location: Location of the partial blade leading edge was also studied. Scientific Research Associates (SRA) studied two cases where the partial blade's chord length was increased. SRA used the code MINT (Briley 9) to perform this task. Increasing the length of the partial increased the impeller exit b-tb distortion (figure 12) as well as the h-t-s distortion (figure 5). Based on the trend in the flow split, a small increase in the partial blade length may provide the even mass split desired across the partial blade. The second portion of this part of the study involved maintained the partial blade's chord length constant, but varied the location of the partial's leading edge. In the baseline, the leading edge of the partial blade bisects the angle between adjacent long blade. For this task Lewis Research Center used the code HAH3d (Hahl0). Initially, two variations about the baseline were run involving a 5 degree shift in the leading edge of the partial towards the suction side of the long blade, and a 5 degree shift towards the pressure side of the long blade. Based on these results, a third case involving a 2.5 degree shift towards the pressure side of the long blade was run. Figure 13 shows that the last two cases reduced the distortion in the b-t-b direction. The 2.5 degree shift has an ideal predicted flow split of 50/50. (The results figures 4 and 5 of the final paper). of these calculations will be included in Blade Trailing Edge Lean: The trailing edge of the blades are nearly axial in the baseline design. Virginia Polytechnic Institute (VPI) studied the effect of varying the blade trailing edge circumferential location from h-t-s. Backward lean is defined as the shroud trailing edge leading the hub trailing edge as the impeUer rotates. Forward lean is the reverse; the hub trailing edge leads the shroud trailing edge. The initial results indicated that backward lean did significantly reduce b-t-b distortion. However, the initial lean cases modeled also changed the blade trailing edge angle. Therefore, to understand whether it was the backward lean or the blade exit angle distribution that lead to the improvement, two additional cases were rtm which studied these two variations independently. Figure 14 shows that the backward blade lean is the dominant cause for the decrease of b-t-b distortion and from figure 5 it appears that a combination of blade lean and blade exit angle variations may be more effective than lean alone (VPI #2 vs. VPI #5). This concept looks promising but presents two potentially negative aspects: (1) the h-t-s distortion rises proportionally to the decrease in b-t-b distortion, and (2) the blade lean c_,ncepts may be more difficult to manufacture. Conclusions The PSTT impellers. has, with the use of CFD, improved the performance of rocket engine An example of the baseline consortium impeller has been tested and its performance with goal the at the exit the significantly affect the clockings the viable concepts detail and length, for reducing a new impeller and b-t-b the while still backward one length lean tend results or several of these blade. to increase leading edge, edge trailing will did design partial blade of the blade These chord blade of the partial the do not in a tandem chord distortion that blade the length of the distortion. incorporating maintaining blade flow impeller along chord location designed, manufactured, and tested. results, the team has already decided possible partial conducted indica4e of the ciockings in the circumferential axial impeller in the increases the relative been the b-t-b Results distribution increases have decrease downstream work Small as small or large Changing increasing space the distortion. as well which performance. vaneless Changing studies Further concepts sacrificing in the results. be beneficial distortion. at design without made compromise Large to that predicted. of arriving impeller simplifications may close is very be studied concepts will are in be Based on the preliminary assessment of the that further increases in head coefficient are efficiency acceptable and levels of distortion. References 1. Garcia, of the NASA/MSFC International Garcia, 28th Joint 5-8, 1992, Stage Pump 4. 5. Brozowski, Prueger, HPFTP L. and the Rojas, G. and DC A., AIAA Dynamics 6-8, "Activities 92-3222, 1992, Schutzenhofer, Fourth of Rotating Marshall Space TN. "Overview of the NASA/MSFC Technology," ALAA 92-3219, Conference, Propulsion of the AIAA/SAE/ASME/ASEE Nashville, L. A., in Propulsion final L., Paper at the ASME (complete Eastland, impeller, Conference, Propulsion for Easfland, July 28th Joint to be presented Washington, SSME of the Proceedings and Activities of the Hawaii. P., and for Applications "A Summary July 6-8, 1992, TN. testing, 1992, Team," Phenomena Technology," AIAA/SAE/ASME/ASEE Nashville, Technology Honolulu, P. K. and Consortium L. A., Schutzerthofer, Transport Conference, Propulsion McConnaughey, CFD on R., McConnaughey, Center Flight 3. April and Stage Pump Symposium Machinery, 2. E. D., R., Jackson, A., Paper 1993, results Engineering Fluid of SSME to be available reference to be presented July discussing 1993 comparing at the Monterey, CFD HPFTP for the results CA J (complete June final reference 21-24, manuscript). to test data AIAA/SAE/ASME/ASEE manuscript). impeller Conference, for the 29th Joint to be available 6. Kiris, the Flow C., Rogers, Artificial Bioengineering, 7. Chan, ASME D. C. and Compressor 8. Chen, Pressure January 9. Tip S., Kwak, Through 9-12, 1989, W. R., Clearance," 10. Hah, C., Bryans, Computations Reno, ASME Orlando, July Analysis Flow AIAA 25th Joint for Paper FL,. June Flow 91-GT-56, Annual in Meeting. of a Double Circular Propulsion Conference. Arc Flow Computations with a 27th Aerospace Sciences Meeting, Past a Turbine International Gas Blade with and without Turbine and Aeroengine 1991. A. C., Moussa, Aerodynamic Conditions," 1988. for Probing Advances Nevada. et al, "Computational Congress, 311, 89-2569, 1991, Winter Y. S., "Compressible and Incompressible Based Method," AIAA 899-0286, AIAA Briley, Operating AIAA December Division, K., "Turbulent Approach I., "Computational Devices," Bioengineering Tran, Cascade," D., Chang, Heart ]'ournal Z., Tomsho, Performance M. E., "Application of Backswept of Turbomachinery Trans. Impeller ASME,Vol. of Viscous Flow at Viscous 110, pp. 303- Definition of Performance c u = Cu / Utip where Cu = absolute tangential c_.= C ./U+=. where C III III tllJ = meridional velocity velocity III U. = wheel tip velocity tip " 13= relative flow angle, in degrees, referenced ¢x= absolute flow angle, in degrees, TI -efficiency Parameters - head rise/Euler _V= head coefficient from tangential referenced dir. from tangential dir. head rise - AH tg/U t2p RSHL = rotor stagnation head loss coefficient 2 = (Euler head - head rise) g/U tip Relative Radius = (R i " R hub) / (R shroud - R hub ) Relative X = (X i - X shroud Relative Angle = (Angle i ) / (X hub - Angle -X ) shroud suction ) / (Angle pressure - Angle suction ) Distortion hub-to-shroud where (Xk= mass blade-to-blade where = [max((z averaged = C _max C = average (x. = mass J averaged k) - min flow ((x.)j Parameter Definitions (ek)l hub-to-shroud angle, - min total discharge flow angle, averaged in the blade-to-blade direction ((x j)] blade-to-blade velocity averaged in the hub-to-shroud direction TABLE 1. PUMP TEAM MEMBERS • NASA Marshall Space Flight Center (MSFC) • NASA Ames Research Center (ARC) • NASA Lewis Research Center (LeRC) • David Taylor Research Center • Rocketdyne (RDYN) * Pratt & Whitney (P&W) • Aerojet • Ingersoll-Rand • CFD Research Corporation • SECA • Scientific Research • The University • Pennsylvania • University • Virginia • California Associates of Alabama (SRA) in Huntsville State University (UAH) (PSU) of Cincinnati Polytechnic Institute Institute of Technology TABLE 2 - IMPELLER RPM Impeller 30108 InletTip D Impeller Inlet Hub D Impeller Inlet B Impeller Outlet D Impeller Outlet 13 Impeller SPECIFICATION B 2 Width Impeller Tip Speed Impeller Specific Speed Impeller W2/W 1 9.38 6.097 17.9 14.14 38.0 1.12 1857 1141 0.690 Impeller C_/C 1 0.377 Impeller Cu2/U 2 0.726 Impeller Blade Number D: Diameter 6+6 in Inch, 13: RMS Blade Angle from Tangential. Table 3. List of Cases Analyzed Cases Postprocessed Organization Size Description ARC #1 328K Baseline ARC #2 540K Baseline with exit cavity at imp. exit Rkdn #1 20K Baseline Rkdn #2 20K Baseline envelope, heavy Rkdn #3 20K Baseline envelope, light Rkdn #4 20K Axial length +37%, B2 -20%, Beta2 = 41.5 Rkdn #5 20K Axial length +20%, B2 -20%, Beta2 = 41.5 Rkdn #6 20K Axial length +00%, B2 -20%, Beta2 = 41.5 I.e. loading 1.e. loading _;ECA #1 71K Baseline SECA #2 75K Tandem blade, 7.5 degrees SECA #3 75K Tandem blade, 22.5 degrees SRA #1 160K Baseline SRA #2 160K Increases SRA #3 160K Longest #1 33K Baseline VPI #2 20K Backward VPI #3 20K Forward VPI #4 33K Base. envelope, no blade VPI #5 33K Base. envelope, backward VPI partial partial blade blade blade blade lean, lean, clocking against clocking against rotation rotation length length Moore Moore distribution distribution lean, blade Beta2 lean, 46-32 Beta2 degrees = 38 degrees JelleduJI uJn!posuoo eu!lese8 "r eJn6!-I 09°0 - 01,'0- 0_'0- I 00"0 dSH 11 INCH SSME LASER HPFTP VELOCIMETER IMPELLER TEST DATA R=5,570" R=5.701" R=5.833" ".° Figure 2. SSME HPFTP impeller exit radial velocity measurements FLOW EXITWALL RECIRCULATION FLOW WALL. INLETSHROUD RECIRCULATION FLOW _E,VY,,NES-_P,C,L,MPE,_LER F,OW P,','MSI FIG. 3. SCHEMATIC OF IMPELLER GEOMETRY AND DOMAIN ANALYZED ,o o o o o ,,I 1 I I I I I ARC #I ARC #I ARC #2 ARC#2 RKDN #I RKDN #I n. RKDN #2 f_ U •RKDN #3 t_ RKDN #2 "11 RKDN #3 g RKDN #4 =Z n. RKDN #4 _=q RKDN#5 RKDN #5 RKDN#6 RKDN #6 o SECA#I B SECA #I SECA#2 {D m m o SECA ..q O SECA#3 3 SRA #1 SECA#3 S.RA#I 00 O O SRA #2 g SRA #3 O ¢ n _o 2-. O ..I {h VPI #1 VII VPI #3 VPI #4 VPI #5 SRA#2 SPA#3 VPI#I i.d. fl VFI#2 U fl VPI#3 VPI#4 ooo0 0ooo ,o,0 o,,, ARC#I _I I I 0 ARC #I I 0 0 0 0 0 ARC II ' ' ' I \xx\ \\\xx\xxxxx ARC #2 ARC#2 RKDN #I RKDN#]' RKDN #2 RKDN#2' e- RKDN#3' .fl RKDN#4' -I ID $=l- t_ rl 3 "u o N ID -I m x __xxxxm RKDN #3 RKDN 13' I RKDN #I RDN#4' ! xx\xxxx\xx\xxxxx RKDN#5' RKDN#6' _\\%%\\%%x_ SECA #I' M RKDN#5 _ RKDN#6 " SECA #1 M SECA#3' SECA #3 SPA#I SRA #I m, 0 SPA#2 "U 0_' _, o t'D el __ ?i .. s_ II "I II I VPI#2 _\_%\\%\\\\\\\%\_ SRAI_" SRA#3 VII#1 Wi #I VP!#2 _I Q :I_. SRA_2 SPA#3 3 r SECA#2 SECA #2' ID 0 _xx_ RKDNI6 _xxxm RKDN IS' o ' V_III" 0 _II_" I ff M, eP VPi#3 0 I Vfl#3 VPI#4 VPI#4 VPI#5 VII#5 ffll_" _t ffll_ I I t, Advanced Impeller Averaged 40.0 ' •I H _ ,- 30.0 ¢.. 20.0 I Parametrics: Exit Cavity Effect Rel. Flow Angle: R/Rtip = 1.0275 ' I ' I ' I ' arc, no exit cavity cavity arc, w/exit v O " 10.0 rr 0.0 0.0 0.2 0.4 0.6 0.8 1.0 Relative X Averaged 0.20 ' I HH Cm vs. Relative X: R/Rtip = 1.0275 I I I I arc, exit cavity cavity arc, no w/exit 0.15 E 0.10 O 0.05 0.00 0.0 0.2 Figure 6a. Circumferentially 0.4 0.6 Relative X averaged hub-to-shroud 0.8 1.0 impeller exit radial velocity distribution Advanced Impeller Parametrics: Rel. FI0w Angle: PJRtip = 1.0275, 40.0 1 _ Exit Cavity Effect Rel X = .5 ' I ' if) 30.0 "O v 20.0 < O LL 10.0 H H n-- 0.0 0.0 arc, no exit cavity arc, w/exit cavity 0.2 0.4 0.6 0.8 1.0 Relative Angle Cm vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 0.20 I I H arc, no exit H arc, w/exit cavity cavity 0.15 (3. E 0.10 O 0.05 0.00 0.0 0.2 0.4 0.6 0.8 1.0 Relative Angle Figure 6b. Blade-to-blade impeller exit radial velocity distribution, at 50% of the blade span Advanced Impeller Parametrics: Performance 0.80 ' I Predictions: ' I Blade Loading Head Coefficient _ I ' I T I t I ' 0.70 tQn :t= 0.60 - H rkdn baseline H O rkdn rkdn rkdn rkdn rkdn _J "O 0.50 - _ q) "1- heavy I.e. loading low I.e. loading XL +37% XL +20% XL +00% 0.40 0.30 I I 0.0 f 0.2 I 0.4 0.6 1.0 0.8 Relative X Performance 1.10 I Predictions: I Efficiency I ' I 1.00 0.90 >., rkdn baseline C -- 0.80 H - _ 0.70 - _ o_ =: LU rkdn rkdn rkdn rkdn rkdn 1 heavy I.e. loading low I.e. loading XL +37% XL +20% XL +00% 0.60 0.50 I 0.0 J 0.2 I I , 0.4 I , 0.6 , 0.8 .0 Relative X Figure 8. Performance prediction for blade loading and impeller axial length study iD , Advanced Impeller Parametrics: Blade Loading Rel. Flow Angle: R/Rtip = 1.0275, Rel X = .5 30.0 ' I ' I ' I ' I ' ° ",L v 20.0 E_ E: <C O eline E_. 10.0 (1) > 7"/4//_ " // _1_/ .@..i (l) _ / l-----m _ G----G , _ _ v- rr 0.0 , 0.0 0.2 _ 0.4 , ) 0.6 0.8 rkdn rkdn rkdn rkdn rkdn heavy I.e. Ioadi low I.e. loading XL +37% XL +20% XL +00% 1.0 Relative Angle Cm vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 0.20 L rkdn baseline H rkdn heavy I.e. loading rkdn low I.e. loading rkdn XL +37% 0.15 E I ' I I , I rkdn XL +20% rkdn XL +00% O.. 22) ' H 0.10 O 0.05 0.00 0.0 0.2 0.4 0.6 0.8 1.0 Relative Angle Figure 9. Blade-to-blade impeller exit radial velocity distribution at 50% of the blade span ql Advanced Impeller Parametrics: Tandem Blades Performance 0.80 ¢- ._ ' I Predictions: ' I H seca baseline H seca 7.5 degrees seca 22.5 degrees Head ' Coefficient I ' I ' 0.70 e_ :=:: O O "t3 • 0.60 "-rv 0.50 , 0.0 I 0.2 v , v v ' I ' H seca baseline H H seca 7.5 degrees seca 22.5 degrees v i , i 0.4 0.6 Relative X Performance 1.10 v Predictions: v v , v v i 0.8 , 1.0 Efficiency I ' I I l I I 1.00 UJ 0.90 0.80 I 0.0 i 0.2 0.4 0.6 i I 0.8 Relative X Figure 10. Performance prediction for the tandem blade study 1.0 Advanced Impeller Parametrics: Tandem Blades Rel. Flow Angle vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 30.0 ' ' ' ' ' ' ' ' / J (/) <D (P "(3 ' 20.0 | ._ [ l--'c-t seca baseline /H seca7.5 degrees II v i E_ ¢- < 10.0 O LL I (P > 0.0 rr 110.0 / 0.0 0.2 0.4 0.6 / 1.0 0.8 Relative Angle Cm vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 ' I i I t ' I _ I ' I i I I I , 0.10 Q. °i E O 0.00 10.10 0.0 l 0.2 0.4 0.6 0.8 1.0 Relative Angle Figure 11. Blade-to-blade impeller exit radial velocity distribution at 50% of the blade span Advanced Impeller Parametrics: Partial Blade Length Rel. Flow Angle vs Rel Angle: R/Rtip = 1.0275, Rel X = .5 40.0, I , , , , , , a , , , I 30.0 "O 20.0 v t- 10.0 < O LL rr 0.0 110.0 120.0 0.0 0.2 0.4 0.6 0.8 1.0 Relative Angle Cm vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 ' I I ' t 0.10 E (3 H 0.00 H H sra, baseline sra, long partial sra Ion 10.10 0.0 0.2 0.4 0.6 0.8 1.0 Relative Angle Figure 12. Blade-to-blade impeller exit radial velocity distribution at 50% of the blade span Advanced Impeller Parametrics: Splitter L.E. Location Rel. Flow Angle . , 50.0 vs. Rel Angle: . , . ¢/) 40.0 "10 v R/Rtip = 1.0275, Rel X = .5 , . , H Irc baseline H Irc 5.0 degrees suct. Irc 5.0 degrees press. o_£) Irc 2.5 degrees press. 30.0 t- < O LL 20.0 10.0 i n- 0.0 0.0 0.2 0.4 0.6 Relative Cm vs. Rel Angle: 0.20 I ' 0.8 Angle R/Rtip = 1.0275, I 1.0 ' Rel X = .5 I Irebaseline H Irc 5.0 degrees suct. Irc 5.0 degrees press. o--0 Irc 2.5 degrees press. 0.15 (3. E3 E 0.10 O 0.05 I 0.00 0.0 0.2 13. Blade-to-blade I 0.4 Relative Figure , impeller exit radial , 0.6 I 0.8 1.0 Angle velocity distribution at 50% of the blade span Advanced Impeller Parametrics: Blade Lean Rel. Flow Angle vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 35.0 | , , , , H ___ L 30.0 vpi baseline H _ _ vpi backward lean vpi forward lean vpi n.l., beta2=46 to 32 25.0 ,_ 20.0 o 15.0 " " = 10.0 5,0 / , 0.0 I , 0.2 I , 0.4 I , 0.6 t , 0.8 1.0 Relative Angle Cm vs. Rel Angle: R/Rtip = 1.0275, Rel X = .5 0.20 , ' 0.15 H vpi baseline H vpi vpi vpi v i backward lean forward lean n.l., beta2=46 to 32 back. lean, beta2=38 _ i i - olo O 0.05 0.00 t 0.0 ' ' 0.2 ' ' 0.4 ' ' 0.6 ' ' 0.8 ' 1.0 Relative Angle Figure 14. Blade-to-blade impeller exit radial velocity distribution at 50% of the blade span