STATOR DIFFUSION ENHANCEMENT USING A RE-CIRCULATING CO-FLOWING STEADY JET

Transcription

1 Proeedings of ASME: ASME Turbo Expo 004 Power for Land, Sea and Air June 4-7, 004- Vienna, Austria GT STATOR DIFFUSION ENHANCEMENT USING A RE-CIRCULATING CO-FLOWING STEADY JET David Car AFRL/PRTF 950 Fifth St. Bldg. 8D Wright-Patterson AFB, OH Niholas J. Kuprowiz AFRL/PRTA 950 Fifth St. Bldg. 8D Wright-Patterson AFB, OH Jordi Estevadeordal Innovative Sientifi Solutions, In. 766 Indian Ripple Rd. Beaverreek, OH Geheng Zha University of Miami Dept. of Mehanial Engineering Coral Gables, FL 334 William Copenhaver AFRL/PRT 950 Fifth St. Bldg. 8D Wright-Patterson AFB, OH ABSTRACT This paper will outline a steady flow ontrol tehnique that augments the diffusion proess within a stator passage via a ontinuous o-flowing seondary flow stream along the sution surfae. The tehnique is similar to that used for flow vetoring in nozzles where a seondary flow stream is used to enhane the diffusion and vetoring of high speed ets. Diffusion fators in exess of 0.95 are simulated and the penalty for the seondary system is addressed with an availability and simple power analysis. Losses within the seondary flow stream were inluded in the availability analysis, but it did not aount for losses within a delivery system of this seondary flow. This was aomplished through the D power analysis whih assessed this tehnique s impat on the effiieny of an axial ompression stage and the sensitivity of this effiieny to the seondary flow system s effiieny. Also, a system level analysis is presented to assess the merits that may be realized in a notional engine with this type of flow ontrol. Partiularly, impats on speifi fuel onsumption and thrust-to-weight ratio were addressed. A asade experiment was performed to demonstrate the onept and was onduted in a blow-down asade tunnel. Signifiant improvements in diffusion were qualitatively seen from the DPIV measurements despite limitations in ahieving the desired seondary flow onditions. INTRODUCTION The trend to higher stage loading in axial turbomahines ontinues to put an inreasing demand on the stator to provide higher amounts of diffusion. Many researhers are addressing this issue through the use of varying flow ontrol tehniques [- 4]. Many of these tehniques involve the use of blowing ets or sution devies or a ombination of the two to ontrol the amount of separation and thereby inrease the diffusion and overall performane of the blade setion. There are others who are showing promise using unsteady tehniques with syntheti ets to ontrol boundary layer parameters [5,6]. These tehniques are very appealing in that they impart a zero net mass hange to the ore flow. Still others are using plasmas for boundary layer ontrol [7]. The flow ontrol tehnique outlined in this paper imparts no net mass flow hange to the ore flow through the use of a re-irulating o-flowing et that enhanes the diffusion and turning levels through its interation with the ore flow. This tehnique is similar to those used in et nozzle flow vetoring whih reate a seondary flow using a sution ollar and pump [8-]. The tehnique outlined here spawned from the work performed in fluidi vetoring in nozzles and is an attempt to reate a similar ondition in the stator passage with a reirulating o-flowing system. It is envisioned that this seondary flow is sustained with a re-irulating seondary flow system as shown in Figure. The penalty of this type of arrangement on the overall ompression system may not be too great if the required total pressure ratio of the seondary flow system and its effiieny requirement is reasonable. A power and availability analysis was onduted in an attempt to address this issue. Also, a system level analysis using a notional engine onfiguration is arried out to assess the impat on engine performane. The onern over heating effets aused by the re-irulating flow and the possible need for interooling the seondary flow was raised in private ommuniations []. This tehnique relies on momentum and heat exhange between the ore flow and the seondary flow as it traverses the blade so it may not be

2 neessary for an interooler, partiularly if the seondary flow system does not require a substantial power input. A asade experiment was onduted to evaluate the onept. Shortfalls within the seondary flow system limited the amount of seondary flow, falling short of the desired onditions of the CFD simulations, but a qualitative improvement in diffusion was apparent from DPIV measurements.. is made that the region ontaining the seondary flow et is a irular ar and that the inlet and exit veloity of the seondary flow et is equal. The resulting non-dimensional axial and tangential blade loading equations that result are as follows: Vx Fx ( sin( θ ) sin( θ )) C p V () F y Γ x + ( os( θ ) os( θ )) where F Γ ρ V ρ h m& V V xs m& Vx ( V y V y ) S Γ F x x ρv x C p V S x S F P P ρ V VxS Fy ρ V S y x x Figure Flow Control Conept Figure 3 Control Volume Figure Flow Control Implementation CONTROL VOLUME ANALYSIS A D ontrol volume analysis was performed to determine the relevant non-dimensional parameters and the influene this flow ontrol tehnique may have on the ore flow. The ontrol volume used is shown in Figure 3. The simplifying assumption Albei the reader will reognize these parameters as the stati pressure rise oeffiien ; the irulation, Γ ; the momentum oeffiient; F y C p ; and the non-dimensional tangential F x and axial blade loading, and respetively. The amount of turning a stator ross setion ahieves is proportional to the tangential blade loading. The momentum oeffiien, offers a positive ontribution to the tangential blade loading, F y, if and only if it effets Γ in a positive

3 Γ manner, i.e. > 0. Interestingly, if the seondary flow et diretion were reversed in the ontrol volume analysis, i.e. a ounter-flow diretion, then the equations remain the same. Also, F y is inreased if the angular extent over whih the seondary flow region ats is inreased. The momentum oeffiien, offers a saleable nondimensional parameter for the seondary flow system. It s seen in muh literature onerning flow ontrol devies; partiularly ones involving irulation ontrol airfoils [3,4]. The seondary mass flow must sale with the ore mass flow along with the et veloity to axial veloity ratio in order to hold the momentum oeffiient onstant. The seondary mass flow ould also sale with the engine shaft speed sine the ore mass flow in an axial ompression system is tied to the engine shaft speed and it s envisioned that the seondary flow pump would be tied to the same shaft as shown in Figure. Therefore it Figure 4 Hybrid CFD Grid seems reasonable that the type of onfiguration shown in Figure ould hold relative onstant over the engine operating The inlet and exit of the ore flow domain and the inset range. A omputational fluid dynamis investigation was avity were simulated in the same manner. Both speified the performed as a first step in understanding the influene this total temperature, pressure and flow angle at the inlet and the flow ontrol tehnique might have on the ore flow. stati pressure at the exit boundary. The exit stati pressures to the ore flow and the avity flow were iterated until the desired inlet Mah number to the ore flow was ahieved and the exit CFD SIMULATIONS mass flow to the inset avity was equal to the inlet mass flow to The geometry shown in Figure 3 was simulated using the avity. The inlet and exit to the inset avity are deoupled FLUENT, a ommerial CFD pakage. The geometry was in this way whih is different from the envisioned appliation reated from a series of three irular ars ensuring ontinuity requiring a re-irulating pumping system. Time onstraints of slope between them. The first irular ar extends from 0- fored us to use the existing boundary onditions provided in 0% axial hord; the seond from 0-90%; and the third from the CFD pakage rather than write ustom routines to represent 90-00%. The overall hord is 0.7 m with a solidity of.67. the pumping system and provide feedbak to the inlet of the The radius of urvature for eah setion is as follows: inset avity. Therefore, the total temperature for the avity m, m and 0.84 m. The thikness distribution for the inlet was set to the ore flow total temperature thereby not ontrolled blade was set to allow enough ross setional area to addressing the possible re-heating effets mentioned earlier. pass the required seondary flow for a future asade experiment whih would allow full span ontrol. This experiment will be disussed later in this paper. The simulation was a D simulation with a hybrid grid onsisting of triangular elements in the ore flow with a quadrilateral element O-grid around the blade. The grid system is shown in Fig. 4. The solver setup used k-epsilon turbulene modeling and wall funtions to model the boundary layer. For the flow ontrolled blade, the inset avity grid was x0 and there were a total of 7050 ells in the entire grid domain. The inlet onditions for this asade setion were a Mah number of 0.7 with an inlet angle of 68 degrees. An axial exiting flow with an approximate Mah number of 0. was the desired exit ondition. This ase would result in an extreme DF of approximately The inlet Mah number of 0.7 was hosen beause of a future asade test in whih the asade tunnel was limited to a maximum running Mah number of 0.8. Firs a baseline was run to ompare against and establish the effetiveness of the flow ontrol tehnique. The baseline blade was similar in shape to the flow ontrol blade in terms of blade urvature. The sution surfae of the baseline blade did not ontain the inset avity and the pressure surfae was altered on the flow ontrol blade to provide the neessary ross setional area for internal avities for the sution and blowing volume in a future asade experiment. The omparison between the flow ontrolled blade and the baseline blade shape is still valid even though the blade shapes are not exatly the same beause: ) If I were to simulate a true baseline, I would simulate the exat flow ontrol blade with the inset avity. No one would design a blade shape with a bakward faing step on the sution surfae. Therefore, I would have to simulate a blade shape without the inset avity whih hanges my flow ontrolled blade; ) This flow ontrol tehnique inherently modifies the blade shape by needing an inset avity, unlike other tehniques that seek to influene the flow by applying sution and blowing to an already existing profile. The tehnique would inextriably alter any blade shape it was applied to. With reasons ) and ) stated above, the goal should be to ompare the flow ontrol tehnique to a blade shape that tries to ahieve the same end. Sine the baseline shape was used to reate the flow ontrolled shape, partiularly 3

4 keeping the urvature reasonably the same, the omparison is reasonable. The baseline ase was simulated and the Mah ontours are shown in F igure 5. As an be expeted, the flow is badly separated on the sution surfae starting at approximately 50% hord. The flow ontrolled blade is shown in Figure 6 for a momentum oeffiient of The flow remains attahed for the entire blade region and appears ompletely diffused at the exit plane. Figure 5 Baseline Blade Mah Contours The zero momentum oeffiient ase is for the baseline solution. As the table shows, the irulation and hene the blade loading is inreased with inreasing values of inetion. The flow ontrolled blade was able to ahieve an axial exit ondition for an inetion between 0.79 and 0.97 for the h/s and between 0.34 and 0.46 for h/s The irulation and DF values in Table are plotted in Figure 7. DF will be disussed as a figure of merit throughout the body of text sine as mentioned in the introdution, the desire to inrease axial ompression stage loading plaes an inreased demand on the diffusion requirements of the stator. Table Flow Control Parameters Γ M α DF fx00 π h/sx C,3 C, J, Figure 6 Flow Control Blade Mah Contours Table summarizes the baseline and flow ontrolled blade for various momentum oeffiients (whih were varied by hanging the seondary flow inetion to ore flow total pressure ratios, ) and frational inetion slot heights (h/s). π Referening the loations highlighted in Figure 6, the definition of DF used in the table is: DF V J,,3 y + () V, V σv, where V y is the hange in tangential veloity between loations (,) and (,3). Mass averaged quantities were used for the alulation. Cirulation Γ DF Γ DF h/sx h/sx Momentum Coeffiient ( ) Diffusion Fator Figure 7 Cirulation and DF versus Momentum Coeffiient The slope of the irulation line as a funtion of is positive, i.e. Γ > 0. Therefore, as expeted the influene on the loading is benefiial and the overall diffusion is greatly inreased. The relationship is also seen to be quite linear in the range simulated, although more points would be neessary of ourse to establish a trend. The DF and irulation in Figure 7 do not ollapse on the inetion momentum oeffiient for the two different inetion slot heights simulated. As an be seen, an offset exists between 4

5 the two ases. This an be understood by realizing that the flow ontrol proess does not rely solely on inetion, but also on the applied sution and that the momentum oeffiients for these two areas are quite different. The preeding ontrol volume analysis assumed that the inlet and exit heights (i.e. areas) were the same for the blowing and sution ports in the inset avity. In reality this is not the ase. The sution area needed to be larger than the inlet area to prevent hoking and aommodate the required mass flow ineted and mixed with the ore flow. Therefore, the above plot is partiular to the geometry at hand sine the x-axis is for the inetion only and does not take into aount the separate for the sution. Another interesting result is that the 0.46 with h/s ahieves a similar DF as the 0.97 with h/s ase. Similar results were disovered in the past for flow ontrol airfoils [3,4] whih ahieved similar lift oeffiients for varying inetion slot heights and therefore various momentum oeffiients. It was a matter of whether you are pressure limited or flow limited in the seondary flow system. A low pressure with a higher flow rate ould ahieve the same end as a high pressure, low flow rate system. Of ourse there is most likely a limit in the range over whih this applies and ertainly there are seondary effets that would make one option more appealing over another, but this appears to be ase for the narrow range of geometries and momentum oeffiients simulated here. A parameter that helps to ollapse the urves is the isentropi power input of the seondary flow stream. To tie in this analysis with a subsequent stage analysis, Figure and Figure 6 will be used for the subsript nomenlature. The proetion of the loations speified in Figure to the CFD simulation is shown in Figure 6. The ore inlet and exit onditions are loations (,) and (,3) respetively. Likewise, (,) and (,) represent the inetion and sution of the seondary flow respetively. The fluidi power assoiated with the total pressure drop between the sution and inetion ports an be defined as: γ γ P, isen m& p T t,, π (3) The total pressure ratio, π, in Eq. (3) is the ratio of the inetion to sution slot total pressure. Non-dimensionalizing this equation using the inlet enthalpy flux, m & T, yields: where, isen fτ π γ γ p, P (4) τ is the temperature ratio between seondary flow sution slot and the inlet ore onditions and is the flow fration, i.e. ratio of seondary to ore mass flow. Plotting the DF against Eq. (4) using mass averaged values is shown in Figure 8. As is seen, the DF ollapses rather well for the two simulations with varying inetion slot height. It should be noted that the sution slot height was.6 times larger than the inetion slot for the h/s ase and.6 times larger for the h/s ase. Interestingly, this f suggests that the diffusion enhanement using this tehnique orrelates with the irreversibility (power loss) of the seondary flow stream. Reently, seondary flow stream irreversibility has been presented as a mehanism for the flow vetoring using a ounter-flowing seondary flow stream for nozzle flow ontrol [5]. The idea is presented that losses within the seondary stream due to the interation with the ore flow inrease the loading, thereby vetoring the flow. Looking bak at Table again also shows that a flow fration of 4.87% is required for a of A rule of thumb that has emerged over the years has been that a flow ontrol devie that uses 0.5 % of the ore flow is exessive. This stems from the understanding that 0.5% of the flow would need to be removed or added elsewhere in the yle and this would be detrimental on the overall engine performane. Bu unlike other tehniques, the flow is re-irulated rather than added or removed. In order to obtain a better understanding of whether this amount of flow fration would ause a large penalty on the ompression system, both an availability analysis and a simplified power analysis were performed. Diffusion Fator h/sx h/sx Isentropi Power Input Figure 8 Isentropi Power Input versus DF AVAILABILITY A treatise on availability analysis is given in [6]. The availability of a system is defined as: Ψ & & ) (5) H m( H T s where t is the hange in stagnation enthalpy, is initial starting total temperature and s is the hange in entropy. The availability is simply the available power extratable at the exit to a system if the system were brought isentropially bak to rest at its initial onditions. For a asade onfiguration, generally sine the asade is a only a loss produing mehanism with no power inpu i.e. H t 0 and s > 0. The total availability of a system is the sum of the availability at eah boundary to the system. Ψ & t < 0 t T t 5

6 As before, this analysis will refer to Figure and Figure 6 for the subsript notation. The availability for the CFD simulation at eah loation is as follows: 0,,3,, m& 0 m& (( H H ) T ( s s )),3 ( H H ) T ( s s ),,,,,,3,, Both Ψ &, and Ψ &, are zero sine the referene values for these two streams are the onditions at (,) and (,) respetively. Summing the above relationships to obtain the total availability, simplifying and non-dimensionalizing by the ore inlet enthalpy flux yields: m& T p,,, s p, Tt,, s τ + f τ p, p, Tt,, p, where p is the speifi heat at onstant pressure and T,3 τ Tt,, T T, τ., (6) (7) For the CFD simulation, the seondary flow inetion total temperature was equal to the ore flow total temperature, i.e.. Also, for simplifiation, let and the T, Tt,, p, p, Eq. (7) redues to: ( τ ) + f ( τ ) ( s + f s ) & (8) p, Ψ The first two terms in Eq. (8) represent the work input/loss of the ore and seondary stream respetively and the last term the ombined irreversibility. The work ours through the shear work between the ore and the seondary flow and although small, is still on the order of the irreversibility term [7]. Therefore, it must be inluded in the analysis. The ratio of availability to the baseline availability (i.e. the ase with no flow ontrol) is shown in Figure 9 for two different inetion slot heights. Mass averaged values were used for the alulations and the exit onditions were integrated at the onstant axial loation labeled (,3) in Figure 6. Note that a mass averaged value of the baseline ase in this region would be a onservative estimate of the loss sine the weighting would be higher for the higher momentum, i.e. nonseparated, regions of the flow. In other words, losses would be larger is this flow ondition were allowed to mix out. Ψ /Ψ base x h/sx h/sx Momentum Coeffiient ( ) Figure 9 Availability versus Momentum Coeffiient As you an see from Figure 9, for the highest momentum oeffiient the availability of the flow ontrol ase is approximately 50% less than that of the unontrolled ase. This simply means there is a 50% redution in the overall loss prodution with the entrainment and mixing proess involved for this flow ontrol tehnique as ompared to the separated baseline ase. Thermodynamially this flow ontrol proess is more effiient than the separated baseline ase, but losses within a delivery system ould make the two thermodynamially equivalent. The next setion will evaluate the performane of a ompression stage that may inorporate this flow ontrol sheme and assess the sensitivity of the stage performane to the seondary flow delivery system effiieny. Keep in mind though, even it were determined that the flow ontrol sheme was found to be thermodynamially equivalent with the separated baseline ase due to the inlusion of losses within the delivery system, the flow ontrolled ase still has the advantage of ahieving the desired diffusion levels and exit onditions over the unontrolled baseline. Another onsideration from Figure 9 is that the power loss for the lower is less than the higher values. Sine when 0 then there must exist a value of base where the power loss is a minimum, but does not oinide with the value of that produes the most diffusion and turning. This may be an artifat of the averaging proedure, but area averaging produed the same trend. POWER ANALYSIS An assessment of this tehnique used in the stage of an axial ompression system was performed by analyzing the power input of a stage using this flow ontrol tehnique in the stator setion as shown in Figure and omparing that with the isentropi power input for the same pressure ratio. To obtain the stage design harateristis, a meanline analysis was performed with the goal of a 68 degree rotor exit ondition, 6

7 .67 stator solidity and 0.98 stator DF in order to be onsistent with the previous results of the CFD analysis. The following meanline input parameters were used: Table Meanline Input Parameters m& 95.3 kg/s/m A Rotor Tip Speed 46.7 m/s Rotor Work Input.680e5 J/kg AR. rotor rotor 0.90 AVR 0.85 rotor rotor σ 3.0 σ.67 stator AVR 0.8 stator The resulting pressure ratio and reation for the stage was approximately 4. and 60/40 respetively. The high rotor solidity was neessary to limit the DF requirement of the rotor and even with this high a solidity, the rotor DF was approximately 0.7. As shown earlier, Figure 8 gives the non-dimensional power loss of the seondary stream for the flow ontrol proess. An adiabati effiieny may be assigned to this value to aount for losses within a hypothetial seondary flow delivery system as shown in Figure and give an atual power input for the entire seondary flow system, yielding: P, isen P, (9) atual Combining this with the power input of the rotor and using the definition of adiabati effiieny gives the total power requirement of the stage shown in Figure as: γ π γ P, isen & + & (0) P m T p, m p T, The first term in Eq. (0) is the power input of the ore flow from inlet (,) to exit (,3). The ore pressure ratio, π, inludes the apparent loss through the flow ontrolled stator. Sine this flow ontrol tehnique imparts momentum to the ore flow, the apparent total pressure loss through the stator is negligible. The CFD simulations of the stator setion showed a redution of only 0.7% in the ore mass averaged total pressure for the momentum oeffiient of Therefore, the ore total pressure ratio, π, ould be onsidered equivalent to the rotor pressure ratio. The ore effiieny,, is another matter. This effiieny ould be affeted by heat transfer from the seondary flow stream to the ore flow making the proess non-adiabati through the stator. This violates the assumption made in deriving the overall power input equation whih used the definition of adiabati effiieny for the ore flow power input. Also, the less effiient the seondary flow system, i.e. the smaller, the larger the seondary flow inetion temperature for a given seondary flow total pressure ratio. This would make the ore flow effiieny in this analysis a funtion of the seondary flow effiieny due to heat transfer effets. An assessment of this effet at this time is not aounted for in this analysis sine the CFD simulations presented earlier were performed with the inetion total temperature equal to the ore flow total temperature. For this analysis, this effiieny will be held at the rotor effiieny value of the meanline analysis, i.e. rotor. Continuing with the analysis, the adiabati, isentropi power input of the stage is the first term of Eq. (0) with the adiabati effiieny set to one: γ γ P isen m& p T t,, π () The ratio of Eq. (), the isentropi power, to Eq. (0), the atual power inpu defines an effiieny and is, upon substituting and re-arranging: P isen () P + P, isen + γ γ π Figure 0 shows Eq. () plotted against the seondary flow system effiieny,. The value for was hosen from Figure 8 for a DF of Two values of overall effiieny are highlighted in Figure 0. There is no partiular signifiane for these two values, they were hosen for omparison purposes only to assess the sensitivity of the overall effiieny to the seondary flow effiieny raised in private disussions [7]. As an be seen, the overall effiieny is for a seondary flow system effiieny of 0.7. There is.7 points in overall effiieny taken up in powering the seondary flow system. A 8.5% redution in the seondary flow effiieny, i.e. 0.50, results in an overall effiieny of P,isen , a redution of 0.75% over the ase. The sensitivity of the overall effiieny to the seondary flow system effiieny is not very large in the range for the seondary flow power input under onsideration and for this stage onfiguration. From this simple analysis, the penalty to the stage using this flow ontrol tehnique does not appear to have a severe impat on performane, even with low seondary flow system effiieny. Although, as mentioned earlier, this does not take into aount the non-adiabati effets due to heat transfer between the seondary flow and ore flow. 7

8 f hpt blade, ooling, 5.50% of ore flow π lpt.03 lpt 0.89 f lpt vane, ooling,.60% of ore flow f lp blade, ooling.40% of ore flow Nozzle CFG 0.95 Fn SLS-mil N Fn SLS-max 4784 N kg/hr/dan SFC SLS-mil Figure Notional Engine Cross Setion Figure 0 Overall Effiieny vs Seondary Flow System Effiieny SYSTEM LEVEL ANALYSIS The overall engine performane impliation of utilizing the flow ontrol tehnique for front fan appliations was investigated as part of this effort. A two-spool, mixed-flow, augmented turbofan with a year-000 prodution-apable level of tehnology is used for analysis purposes. Table 3 provides the basi thermodynami yle harateristis of this engine, and Figure provides a graphial representation of the geometry/flowpath. (It is important to note that this onfiguration is notional, and does not represent a fielded engine or an engine in development.) Table 3 Notional Engine Cyle Charateristis OPR 0 BPR 0.7 m& 4.7 kg/se Throttle Ratio.5 Inlet Reovery 0.97 π 4.44 fan fan 0.87 hp hp Maximum T K Maximum T K π.9 hpt 0.88 hpt f hpt vane, ooling, 0.00% of ore flow Table 4 desribes the pertinent aerodynami data related to the 3-stage front fan design in the notional engine. Of partiular relevane to the urrent study is the average amount of stator turning at the hub loations. Table 4 Fan Aerodynami Parameters of Notional Engine Correted Flow / Annulus Area 70.9 kg/se/m^ Number of Stages 3 Stage Loading (gj(delh)/u^) 0.83 Inlet Correted Tip Speed m/s Inlet Physial Tip Speed m/s Exit Hub Speed 34.9 m/s Average Hub Stator Turning 56.6 deg In general, without the use of ative flow ontrol, separation losses limit the amount of stator turning at the hub loation of any stage to ~65 degrees. Through the use of ative flow ontrol, this onstraint beomes less severe and fan aerodynami designs with redued stage ount beome feasible. The weight savings assoiated with a stage redution are ountered by thermodynami losses assoiated with powering the flow ontrol system. The magnitudes of these savings/losses are quantified as follows. The ompressor map of the fan itself, i.e. all stages, is assumed to be unhanged, and horsepower is extrated from the low-pressure turbine to power the flow ontrol system. Figure shows the effet of low-spool horsepower extration on thrust and speifi fuel onsumption (eah at sea-level stati onditions without the use of an afterburner). 8

9 Fn [lbf] Horsepower Extration Figure Thrust and SFC versus Seondary Flow Power Extration A reasonable assumption for the power required to drive the flow ontrol system is 5% of the fan power in the notional engine. This 5% requirement was arrived at by examining the ratio of the seondary flow power input (seond term of Eq. (0)) to ore fan power requirement (first term of Eq. (0)) from the D power analysis outlined in the previous setion: P P +, isen (3) γ P π γ Using the fan effiieny and pressure ratio from Table 3 along with setting the seondary flow effiieny to 0.5 and P,isen to yields the approximate value of 0.05 for P. Multiplying this by two for additional losses (e.g. P bearing, shaf et.) gives a onservative estimate of 0.05 or 5%. Will all this in mind, this yields a requirement of approximately 0 kw, and results in a (dry) net thrust loss of 463 N and speifi fuel onsumption (SFC) inrease of 0.03 kg/hr/dan. (Note that at sea level stati in full afterburner with 0 kw extration, the net thrust loss is 730 N when ompared to the notional engine in full afterburner.) A redution in the number of fan stages from 3 to will result in a fan weight savings in the -45 N regime. Figure 3 shows the influene of average stage loading oeffiient on st stage stator turning angle at the hub loation. For loadings above 0.95, Figure 3 shows that a two stage design will require an amount of hub stator turning greater than 65 degrees. The relative Mah number at the first stage rotor tip is shown for the and 3 stage designs as a funtion of loading oeffiient in Figure 4. The average stage loading oeffiient must inrease from 0.8 to.4 in order for the stage design to have the same st stage rotor tip relative Mah number as the 3 stage design. An inrease in average stage loading oeffiient implies a higher stator diffusion requiremen an environment in whih flow ontrol an help. The system level analysis suggests that ative fan flow ontrol an be used to inrease overall engine thrust to weight. The notional engine had a thrust-to-weight ratio of 5.6. A SFC [lbm/hr/lbf] Net Thrust SFC replaement of the 3 stage front fan with a stage design that inorporates ative flow ontrol will, despite the thermodynami losses, yield an overall engine thrust-to-weight improvement to the 5.80 regime. Further analysis is required to quantify the benefit for other engine yles, partiularly ones that involve fixing the engine volumetri size and allowing for higher OPR. It is in these yles that one may see a SFC benefit while maintaining or slightly inreasing the thrust to weight ratio. st Stage Hub Stator Turning [deg] Loading stage design 3 stage design Figure 3 First Stage Stator Hub Turning versus Loading Rel Mah Number Loading Figure 4 Relative Mah Number versus Loading EXPERIMENTAL RESULTS stage design 3 stage design Casade A asade was reated by simply extruding the ross setion outlined earlier with internal avities suffiiently sized to provide the neessary mass flow to feed the inetion and sution slots on the blade surfae. The blade setion with the larger inetion slot height to pith ratio (h/s) of.0974 was hosen. The ross setion of the asade is shown in Figure 5 and the solid model in Figure 6. Seondary flow to the asade is provided from only one side of the blade to allow viewing from the other side of the asade test setion for DPIV measurements. Also, this would be the intended onfiguration for a stage as shown in Figure. Although, the blade ross setion provided suffiient area to aommodate the neessary mass flow, it was equal to the respetive slot areas to 9

10 whih they supplied air flow. Ideally one would want the ross setional areas supplying the seondary flow to be larger, but that was not possible due to geometri onstraints and strutural rigidity onsiderations. Although manufaturing the asade using EDM tehnology is a fairly easy task, a minimum wall thikness is required to prevent warping and maintain toleranes. The blowing and sution supply avities need to be loated within the entral portion of the blade ross setion sine this portion of the blade provides suffiient metal thikness. With the blowing and sution slots loated to the ends of the blade, this requires the internal flow to negotiate an undesirable 80 degree turn as an be seen in Figure 5. Unfortunately, without very areful design onsiderations in this region, this will result in signifiantly more loss generation. Insuffiient time did not allow for any detailed analysis of this area. With all the geometri, manufaturing and time onstraints, the ross setion in Figure 5 is onsidered the best that ould be done. The asade test setion is six inhes tall and ontains six blades aross the test setion with a blade pith of meters. The blade hord measured as a straight line from leading to trailing edge is 0.7 meters giving a solidity of.67. The asade inlet Mah number was 0.7 with an inlet angle of 68 degrees. Figure 5 Flow Control Casade Cross Setion Figure 6 Flow Control Casade Seondary Flow System General Layout The seondary flow system is reated using a entrifugal automotive superharger powered by an eletri motor and ontrolled by an AC inverter. Figure 7 gives a shemati of this system along with measurements made at various stations along the flowpath of the system. The by-pass valve shown will be explained later. Figure 7 Experimental Seondary Flow Ciruit The superharger was hosen beause it ould provide the desired pressure ratio and power input for the momentum oeffiients of interest. It was diffiult to loate a superharger with suffiient surge margin for the flow rate of interest. The asade orreted flow rate is approximately 4.80 kg/se. Sine only 4 blades out of the 6 (i.e. 5 passages) are ontrolled, the flow rate that needs to be ontrolled is 3.84 kg/se. The desired fration of seondary flow to ore flow (i.e. flow fration, f) from Table is , whih means that the required orreted flow rate for the seondary flow system is 0.74 kg/se. Searhing all the available entrifugal superharger maps puts this flow rate very lose to the surge line for the even the smallest superharger that ould provide the neessary pressure ratio. A detailed map of the superharger hosen for this experiment was not available, but a map of a very lose ompetitor with similar performane was obtained and verified that it would be operating very lose to surge at this flow rate. The desired pressure ratio of the seondary flow system from Table is approximately.48. This pressure ratio is the ratio of the blowing et flow total pressure to the ore flow total pressure. In atuality, the seondary flow system pressure ratio is equal to the ratio of the sution slot total pressure to the inlet blowing slot total pressure. This pressure ratio is approximately.4. But this pressure ratio does not take into aount any flow losses within the seondary flow system. Due to short period of time afforded for the design of the seondary flow system, a simple dump manifold was reated to supply the air to the asade setion. Figure 8 shows a piture of the dump manifold with the over removed. Of ourse this type of manifold is not desirable from a loss perspetive, so an analysis was onduted to determine the required pressure ratio of the entrifugal ompressor needed to deliver a.4 pressure ratio on the asade blade. The loss analysis showed that the entrifugal ompressor needed to supply a pressure ratio on the order.9 to ahieve the desired blowing total pressure. This plaes the ompressor in surge for a orreted flow rate of 0

11 0.74 kg/se. Therefore, a by-pass valve was added to the system as shown in Figure 7 to re-irulate flow and keep the entrifugal ompressor from surging while providing suffiient pressure ratio to the seondary flow. This assumes that the inlet blowing slots for the flow ontrol are hoked and that reheating effets of the re-irulating seondary flow for the short blow down test time of approximately 0 seonds would not have a signifiant effet. Drive System The approximate amount of power required for the seondary flow system an be alulated from a simple D analysis as arried out in the previous setions and re-written here: γ π γ P m & p T t (4) The blow-down tunnel is a pressurized system. The total pressure is anywhere from x0 5 Pa from past experiene with the speified inlet ondition to this asade. The tunnel total temperature an be approximated as a standard day ondition. With this tunnel total pressure and temperature and a orreted mass flow of 0.74 kg/se, the atual mass flow in the seondary system may vary between 0.38 kg/se to 0.95 kg/se. Approximating the ompression effiieny to be 70% from a omparable ompressor map would lead to a power requirement of 8.64 to 3.86 kw. A 3400 RPM, 37.8 kw, 460 Vol 3 phase eletri motor was hosen to be suffiient for the experiment ontrolled by an AC inverter apable of ontinuous kw ontrol.. The power transmission to the superharger was aomplished by using a high preision timing pulley and belt arrangement. The superharger speed needed to be approximately 4,000 rpm in order to ahieve the required pressure ratio. There is an internal step up within the superharger of 4.: thereby requiring an external pulley arrangement of 3: using a maximum motor speed of 3400 RPM. At these speeds, stok steel pulleys are lose to their maximum rim speed, but a system was found that ould transmit the required power and was used suessfully in the experiment. DPIV Results The desired inlet onditions for the asade test were an inlet angle of 68 degrees and Mah number of 0.7. Unfortunately, the seondary flow system hoked at less than half the desired flow rate of approximately 0.38 kg/se. A simple test was performed to determine whih part of the seondary flow system was limiting the flow rate. Two of the blade surfae inetion slots were taped off and the flow rate remained unhanged. The tape was removed and two of the blade surfae sution slots were taped off whih in turn halved the flow rate. Therefore, from this simple analysis it was determined that the somewhere within the sution side of the seondary flow system hoking ourred. Currently a CFD analysis of the blade passages is being performed to determine where within the sution portion of the seondary flow system this ourred. Figure 8 Dump Manifold Piture With the limited amount of flow rate in the seondary system, the inlet Mah number to the asade was lowered to 0.3 in order to try and math the flow fration and momentum oeffiient of and 0.97 respetively of the CFD simulations outlined earlier. This was the lowest Mah number the asade tunnel ould run stably. Also, to ahieve the desired pressure ratio of.87 within the seondary flow system, the by-pass valve was slightly opened to re-irulate a fration of the flow prior to the manifold to keep the superharger from surging. An attempt was made to alulate a momentum oeffiient by measuring the stati pressure at the endwalls on eah side of the blade inetion slots. One ould then alulate a momentum oeffiient assuming the total pressure and temperature measurement within the blowing plenum portion of the dump manifold shown in Figure 7 (loation 3 in the figure) to be the total pressure and temperature at the inetion slot. Unfortunately, the thermoouple measurements used to determine the total temperature at the inlet to and within the blowing plenum of the dump manifold failed during the test. Data taken for the seondary flow system prior to the test showed that the superharger ould operate in the by-pass onfiguration at an approximate adiabati effiieny of 0.65 for a total pressure ratio of.8. This was the operating ondition hosen for the test. The measured inlet temperature to the superharger at this operating ondition during the test was 309 K. Assuming an adiabati effiieny of 0.65 with a total pressure ratio of.8 puts the exit temperature of the superharger at 396 K. With this temperature, along with the measured total and stati pressure prior to the blowing plenum (loation in Figure 7) to alulate the mass flow of the seondary system, yielded a seondary mass flow rate of kg/se. The unertainty in this alulation was determined to be ± kg/se. The ore flow rate for the blow down tunnel operating at a Mah number 0.3 was alulated to be.0 kg/se. This yields a flow fration of seondary to primary flow of This is approximately % of the ore flow less than the desired value of It was also noted that measurements taken prior to the failure of the thermoouples showed a signifiant heat loss in the dump manifold. That ombined with the failure of the temperature measurements during the test made any attempt to alulate the momentum oeffiient unreliable. Figure 9 shows the interrogation regions for the DPIV measurements. The interested reader is pointed to

12 referenes [,8] for a more detailed look at the DPIV setup. Only regions A, B and C were obtained due to time onstraints. There are many fators assoiated with the DPIV unertainty-alulation proess (laser, CCD, seeding, imaging, algorithms, osillosope, et). The highest unertainty was found to be assoiated with the veloity alulation whih involves x (the displaement in pixels of eah interrogation region), t (the time interval between the two exposures), and the magnifiation of the digital image relative to the obet (pix/m). The displaement in pixels obtained by peak-loator algorithms an provide sub-pixel auray (< 0. pixels) after orretion for various biases. The t was adusted to yield typial displaements of the main stream > 0 pixels, and the unertainty is thus < %. Values in the wake region, however, may have higher unertainties due to the lower x. The maximum unertainty in the t was alulated from the time interval between the two laser pulses with the aid of an osillosope (unertainty %). It was found that this unertainty inreases with lower laser power and with lower t. A onservative number for the present experiments, whih employed a t of about 4 s and powers around 0 mj, was found to be <%. The magnifiation was measured using images of grids loated in the laser-sheet plane to better than %. Combining these onservative measurements of unertainty yields a maximum error of < % for the free-stream regions and ~0% in the low speed areas suh as the wake region. A omposite image of the interrogated regions for the baseline ase (i.e. no flow ontrol) is shown in Figure 9. inetion would imply a redution in momentum oeffiient aligned with the ore flow. A subsequent redesign whih will seek to provide inetion and sution uniformity will follow the analysis. Figure 0 DPIV Experimental Baseline Composite DPIV Image Figure Experimental Flow Control Composite DPIV Image Figure 9 DPIV Interrogation Regions As you an see, a signifiant amount of separation ours along the sution surfae of the blade. Similarly, a omposite image of the ontrolled blade is shown in Figure 0. Again, separation is seen, but at a signifiantly lower level over the baseline ase, showing that the ontrol is having a desirable effet. Signifiant differene were seen in the inetion slot stati pressure measurements taken at eah endwall whih leads to the onlusion of very non-uniform inetion and a high probability for non-uniform sution. A detailed 3D CFD analysis is underway to model this asade setion and preliminary results are onfirming this hypothesis. Any radial CONCLUSIONS A ontinuous steady flow ontrol onept using an inset avity and pumping system to ontinually re-irulate a seondary flow whih in turn enhanes the ore flow diffusion was presented. A ontrol volume analysis was performed to find relevant non-dimensional parameters and a CFD simulation exeuted varying those parameters showed that an apparent DF of 0.98 ould be ahieved. The word apparent is used beause the seondary flow ontrol stream adds momentum to the ore flow with the seondary flow system requiring work input through a pumping system. The penalty for the seondary system is addressed with a simple power and availability analysis. The availability analysis showed that the proess involved with this tehnique resulted in approximately

13 a 50% lower power loss through the stator over the separated baseline ase. This analysis inluded the losses within the seondary flow stream, but did not aount for losses within a delivery system of this seondary flow. This means that thermodynamially this flow ontrol proess is more effiient than the separated baseline ase, but losses within a delivery system ould make the two thermodynamially equivalent. Even with thermodynami equivalene, the flow ontrolled ase still has the advantage of ahieving the desired diffusion levels and exit onditions over the unontrolled baseline. The D power analysis was used to assess this tehnique s impat on the effiieny of an axial ompression stage and the sensitivity of this effiieny to the seondary flow system s effiieny. A stage with a pressure ratio of 4. and a ompressor adiabati effiieny of 0.90, yielded an overall effiieny of for a seondary flow system effiieny of 0.7. Lowering the seondary flow system effiieny to 0.5 gave a overall effiieny. The sensitivity of the overall effiieny to the seondary flow system effiieny was not exessively large. Therefore, it may be feasible to ahieve a 4.0 pressure ratio with a 0.88 adiabati effiieny. Also, a system level analysis is presented to assess the merits that may be realized in a notional engine with this type of flow ontrol. The notional engine ontained a 3 stage fan whih was replaed by a stage fan using flow ontrol. The power to drive the flow ontrol system was taken off the low pressure turbine. For this preliminary study, an engine fuel onsumption penalty and a thrust-to-weight benefit were realized with a speifi fuel onsumption and thrust-to-weight ratio inrease of.5% and 3.% respetively. A asade experiment was performed to demonstrate the onept and was onduted in a blow-down asade tunnel. The seondary flow system was designed using an automotive superharger driven by an eletri motor to provide ontinuous re-irulating flow. The experiment was not suessful in providing the simulated levels of diffusion due to flow limitations, i.e. hoking, within the sution portion of the flow ontrol sheme. Even with this limitation, signifiant improvements in diffusion were qualitatively seen from the DPIV measurements. It is believed that non-uniform inetion and sution led to the less than expeted behavior of the flow ontrol tehnique. A 3D CFD analysis is ongoing and preliminary findings are showing large inetion and sution non-uniformities. A redesigned asade setion with improved inetion and sution harateristis will follow. ACKNOWLEDGMENTS The authors would like to thank Tehsburg in Blaksburg for their dediation during experimental testing and what turned out to be a very long week. NOMENCLATURE AR aspet ratio AVR axial veloity ratio BPR bypass ratio inetion momentum oeffiient C stati pressure rise oeffiient p DF diffusion fator f h/s F Fn H t m& m& A V OPR P S s SFC α Γ π σ ρ τ Ψ & M flow fration ratio of inetion slot height to pith integrated loading thrust stagnation enthalpy mass flow speifi mass flow veloity magnitude overall pressure ratio power pith entropy speifi fuel onsumption mass averaged flow angle irulation adiabati effiieny total pressure ratio solidity density total temperature ratio availability Subsripts ore flow seondary flow x axial diretion y tangential diretion t stagnation ondition fan Fan hp high pressure ompressor hpt high pressure turbine lpt low pressure turbine sls sea level stati onditions REFERENCES. Bons, J.P., Sondergaard, R., Rivir, R.B., 000, Turbine Separation Control using Pulsed Vortex Generator Jets, ASME Paper No. 000-GT-06.. Estevadeordal, J., Copenhaver, W., Car, D., Koh, P., Ng, W., Guillo S., Carter, C., 00, Maro- and milli-dpiv Studies of a Boundary-layer-based Flow-ontrol System for a Transoni Casade, th Intl. Symp. On Apll. Of Laser Tehniques to Fluid Mehanis, Lisbon. 3. Farokhi, S., 998, Propulsion system design with smart vortex generators, Airraft Design,, pp MCormik, D. C., 000, Boundary Layer Separation Control with Direted Syntheti Jets, AIAA Paper No MMihael, J.M., 996, Progress and Prospets for Ative Flow ontrol Using Mirofabriated Eletro- Mehanial Systems (MEMS), AIAA Paper No

14 6. Smith, B. and Glezer, A., 00, Jet Vetoring Using Syntheti Jet Atuators, J. Fluid Meh., 458, pp Roth, J. R., D. M. Sherman, and S. P. Wilkinson, 998, Boundary Layer Flow Control with a One Atmosphere Uniform Glow Disharge, AIAA Paper No Strykowski, P. J., Krothapalli, A., Forliti, D. J., 996, "Counterflow thrust vetoring of supersoni ets," AIAA J., 34, Strykowski, P. J., Krothpalli, A., 993, The Counterurrent Mixing Layer- Strategies for Shear-Layer Control, AIAA Paper No Strykowski, P. J., Krothapalli, A., Jendoubi, S., 996, The Effet of Counterflow on the Development of Compressible Shear Layers, J. of Fluid Meh., 308, pp Van der Veer, M. R., Strykowski, P. J., 997, Counterflow Thrust Vetor Control of Subsoni Jets: Continuous and Bistable Regimes, J. of Propulsion and Power, Vol. 3(3), pp Personal ommuniations, GE Airraft Engines, Cininnati, OH, MLahlan, B. G., 989, Study of a Cirlulation Control Airfoil with Leading/Trailing-Edge Blowing, J. Airraf 6(9), pp Shrewsbury, G. D., 989, Numerial Study of a Researh Cirulation Control Airfoil Using Navier-Stokes Methods, J. Airraf 6(), pp Forliti, D. J., Strykowski, P.J., Gillgris R. D., 00, The Role of Irrerversibility in Vetoring Thrust Using Counterflow Control, AIAA Paper No Van Wylen, G. J., Sonntag, R. E., 973, Fundamental of Classial Thermodynamis, nd Ed., John Wiley and Sons, New York, NY, pp Personal ommuniations, MIT Gas Turbine Lab, Boston, MA, Estevadeordal, J., Gogineni, S., Goss, L., Copenhaver, W. and Gorrell, S., 00, Study of Wake-Blade Interations in a Transoni Compressor Using Flow Visualization and DPIV, ASME J. of Fluids Engineering, 4, pp