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On the mode transition of a double bypass variable cycle compression system
Aerospace Science and Technology 98 (2020) 105743
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On the mode transition of a double bypass variable cycle compression system Baojie Liu a,b,c , Ruoyu Wang a , Xianjun Yu a,b,c,∗ a b c
School of Energy and Power Engineering, Beihang University, Beijing, China National Key Laboratory of Science and Technology on Aero-Engine Aero-Thermodynamics, Beijing, China Collaborative Innovation Center of Advanced Aero-engine, Beijing, China
a r t i c l e
i n f o
Article history: Received 18 June 2019 Received in revised form 5 December 2019 Accepted 22 January 2020 Available online 24 January 2020 Communicated by Dionysios Angelidis Keywords: Double bypass engine Mode transition Matching Flow recirculation
a b s t r a c t By establishing a three dimensional model of a double bypass variable cycle compression system, the flow patterns and matching characteristics of each working component during mode transition are investigated using numerical simulation. Results show that transition from the single bypass mode to double bypass mode by opening the mode selector valve (MSV) alone would increase the fan operation point while decreasing that of the core driven fan stage (CDFS) and the high pressure compressor (HPC). It would also incur outer bypass flow recirculation and bring radial inflow distortions to the CDFS. Deviations of the fan aerodynamic performance lie mainly in its aft stage, while the HPC first stage undertakes most of the inlet distortion. Reducing the bypass backpressure during mode transition is an effective way to alleviate the outer bypass flow recirculation but would further choke the CDFS. Closing the forward variable area bypass injector (FVABI) could raise the CDFS matching point so as to improve the stator performance without influencing the outer bypass ratio. It is recommended to decrease the bypass backpressure and close FVABI simultaneously in the real transition process. © 2020 Elsevier Masson SAS. All rights reserved.
1. Introduction As one of the most promising technologies in aircraft industry, variable cycle engine (VCE) could adapt to various flight conditions while maintaining the optimum fuel consumption. Double bypass engine (DBE) is a typical VCE concept. It is also a welldeveloped configuration which has been evaluated comprehensively and adopted by the famous variable cycle engine F120 [1–5]. By introducing the CDFS and the inner bypass into the engine compression system, DBE is able to achieve the tailoring of internal airflow, thus combining both the features of turbofan and turbojet [3,6]. To be specific, there are two operation modes for the DBE, and the ‘turbofan’ mode is called the double bypass mode where a large portion of the inflow air is bypassed through the fan bypass, whereas the ‘turbojet’ mode is named as the single bypass mode where the fan bypass is shut off hence most air enters the core engine. Mode transition is undoubtedly a key point for DBE. Report of the NASA VCE test bed engine points out that the mode transi-
*
Corresponding author at: School of Energy and Power Engineering, Beihang University, Beijing, China. E-mail address: [email protected] (X. Yu). https://doi.org/10.1016/j.ast.2020.105743 1270-9638/© 2020 Elsevier Masson SAS. All rights reserved.
tion has to be conducted skillfully to avoid component or bypass mismatch [6,7]. The report also states that improper transition process could incur outer bypass flow recirculation within the engine compression system, thus worsening the overall aerodynamic performance. Such conclusions have been verified by later investigations [8]. Over the years, researchers have made progress in VCE performance prediction by proposing a number of simulation models [9–16]. However, the precise control methodology and the matching characteristic of the working components are rarely found in the open literature. As a matter of fact, in order to get the overall engine performance quickly, past investigations are mostly based upon the mathematical relations of component operation point and are unable to acquire the real flow conditions inside the engine. On the other hand, the internal flow is actually of great significance, as a well-organized flow field guarantees a wellperformed engine. To shed some light on the flow details of the variable cycle engine, in this study, three-dimensional numerical simulation is employed to calculate a DBE compression system. A series of conditions during mode transition is considered so as to investigate the influence of different control parameters. Based on the results, component operation points and inter-stage matching characteristics are summarized, and recommendations for the transition process are proposed.
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Nomenclature Variables B tot B in B out i l m Ma n p
α η∗ π∗ σ∗
Subscripts/Superscripts Total bypass ratio Inner bypass ratio Outer bypass ratio Incidence angle FVABI length Mass flow Mach number Rotational speed Pressure MSV opening angle Isentropic efficiency Total pressure ratio Total pressure recovery coefficient
is ref rel
∗
Isentropic Reference value Relative condition Total condition
Abbreviations CDFS DBE FVABI HPC MSV PS SS VCE
Core driven fan stage Double bypass engine Forward variable area bypass injector High pressure compressor Mode selector valve Pressure surface Suction surface Variable cycle engine
2. Model and numerical methods
3. Results and discussion
As shown in Fig. 1, the variable cycle compression system to be investigated consists of a two-stage fan, a one-stage CDFS, and a five-stage high pressure compressor. Both fan and CDFS have inlet guide vanes to adjust the inflow directions for the downstream rows. Main system control mechanisms including mode selector valve (MSV) and forward variable area bypass injector (FVABI) are also modeled in the present study. System transition between single and double bypass mode is accomplished through MSV, which stays closed in the single bypass mode hence all the fan air flows into the CDFS. On the other hand, MSV is opened during the double bypass mode, allowing a portion of the fan air to go through the outer bypass. Setting the position of MSV under single bypass mode as a reference, the angle between the reference position and the MSV real position is defined as α . Meanwhile, FVABI is installed at the outlet of the inner bypass. The axial movement of FVABI could modulate the outflow area of the inner bypass, thus ensuring efficient mixing of the inner and outer bypass streams. The axial length of FVABI is denoted by l. In the calculations, atmospheric total pressure and temperature were given at the inlet of the fan, while static pressure p 1 and p 2 were imposed at the outlet of HPC and the bypass, respectively. The physical rotational speed of the fan is set as n1 , while CDFS and HPC are operated under the rotational speed of n2 . Schematic of the geometrical model is presented in Fig. 2. The compression system was constructed by Unigraphics NX, NUMECA Autogrid5 and IGG were then utilized to generate the mesh for the blade region and the complicated bypass ducts, respectively. The O4H topology was used for each blade channel, whereas Otopology was adopted for the skin block around the blade. Based on the turbulent model to be selected, the value of the first grid off wall is set as y+ < 10. The grid number totals 15 million with approximately 0.8 million for each blade row, which has been proved to be adequate for simulating multistage compressor performance [17–21]. Commercial flow solver NUMECA FINE/TURBO was employed for the numerical simulations. As a prevalent multistage turbomachinery solver, validations of FINE/TURBO can be found in lots of open literature [22–26]. In the present study, three dimensional compressible Reynolds-averaged Navier-Stokes equations were solved to obtain the steady solutions. The SpalartAllmaras turbulent model was adopted, the temporal discretization scheme was Four-stage Runge-Kutta, while the spatial terms were discretized by the second-order central difference scheme. Cases were considered converged when the residuals became lower than 10−6 . Generally, solutions tended to converge after 2000 iterations.
3.1. The influence of the MSV position In order to get the matching characteristics of the compression system during mode transition, in this part, transition is conducted by simply opening the MSV. As the system changes from single bypass mode to double bypass mode, six MSV positions are considered, detailed control parameters are listed in Table 1. To get an overview of the system mass flow distribution, three bypass ratios are defined as follows:
B in =
mBYPASSin mHPC
,
B out =
mBYPASSout mCDFS
,
B tot =
mBYPASStot mHPC
where B in is the mass flow ratio between the inner bypass and the HPC, B out stands for the ratio of mass flow between the outer bypass and the CDFS, and B tot represents the mass flow of the total bypass divided by that of the HPC. As can be seen in Fig. 3, with the opening of MSV, the values of both B in and B tot tend to increase, while B out tends to decrease, indicating obvious mass flow redistribution. It should also be noted that for all the cases, B out has a negative value, which means that there is stronger and stronger flow recirculation in the outer bypass as MSV opens up. Furthermore, the outer bypass flow recirculation will reduce the mass flow in the total bypass, hence the variation range of B tot is smaller than that of B in . Fig. 4 gives the absolute Mach number in the vicinity of the MSV, Cases 1.1, 1.3 and 1.5 are illustrated. Streamlines are also depicted to show the flow field more clearly. By comparing these three cases, the evolution of flow recirculation can be recognized easily: The Mach number in the outer bypass increases continuously with the opening of MSV, while the outer bypass airflow gets into the CDFS inlet. Due to the outer bypass flow recirculation, in the downstream of the MSV, flow first separates from and then reattaches to the casing wall, forming a low-speed separation region at the upstream of the CDFS. The separation region will enlarge with the increase of α , as can be seen in Fig. 4(b) and (c). Meanwhile, the injection of the air accelerates the flow outside the separation region, hence the CDFS inlet Mach number near the casing will get higher if mode transition is conducted in the present way. The variations of component matching point are presented in Fig. 5, characteristics of each component alone (with uniform inlet flow conditions) are also given for the convenience of observation. It should be noted that the flow interaction within the compression system will actually influence the inflow and outflow con-
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Fig. 1. Schematic of the variable cycle compression system.
Fig. 2. Geometric model and mesh for the compression system.
Fig. 3. Bypass ratios under different MSV opening angles.
Table 1 Cases to investigate the influence of MSV position. Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case1.1 Case1.2 Case1.3 Case1.4 Case1.5 Case1.6
0 2 4 6 8 10
1.0
1.0
1.0
1.0
1.0
ditions of each component, hence the component real matching point will not coincide with their original characteristics. According to Fig. 5(a), the fan is matched at a near-choke point under the single bypass mode, and opening the MSV will push the fan operation point towards the surge line. On the contrary, Fig. 5(b) shows that the CDFS will get into deep choke during the mode transition, which is detrimental for its aerodynamic performance. Meanwhile, the operation point of HPC will move towards the surge line as MSV opens up, see Fig. 5(c).
To compare the influence of MSV position on different components quantitatively, Fig. 6 presents the component performance with the variation of α . It can be seen that the aerodynamic performance deviation of the fan and CDFS is much greater than that of the HPC. As α increases from 0◦ to 10◦ , the fan total pressure ratio is increased by nearly 10%, whereas π ∗ HPC only changes 3%. Additionally, the variation of η∗ HPC is lower than 1%, as shown in Fig. 6(b). As depicted in Fig. 7, the radial distributions of total pressure ratio under different MSV positions (Case1.1, Case1.3 and Case1.5) reveal the fan stage flow characteristics more clearly. It can be seen that when MSV is opened from 0◦ (Case1.1) to 4◦ (Case1.3), the increase in fan total pressure ratio comes mainly from its second stage, whereas the first stage pressure ratio barely changes. However, as the opening angle changes from 4◦ to 8◦ , total pressure ratios for both stages improve distinctly. Compared to the second stage, in which the pressure ratio increase occurs along the whole blade span, the total pressure ratio increase appears only at the upper blade region in the first stage. Fig. 8 further gives the fan relative Mach number at 90% blade height. For the convenience of comparison, shock wave positions for Case1.1 are depicted in all three cases with dashed lines. As can be seen in Fig. 8(a), being a typical transonic stage, the flow at the tip region in the first stage rotor is made up of an oblique shock wave at the inlet and a normal shock wave, i.e. the passage cutoff shock wave, in the blade channel. As for the second stage rotor, a relatively weak passage cutoff shock wave also exists, which means that the fan is matched at a deep choke state under the single bypass mode. According to Fig. 8(b), when MSV is opened by 4◦ , the normal shock wave in the second stage rotor channel undergoes an obvious upstream movement and merges largely with the inlet
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Fig. 4. Absolute Mach number in the vicinity of MSV. (For interpretation of the colors in the figure(s), the reader is referred to the web version of this article.)
Fig. 5. Component matching point at different MSV positions.
Fig. 6. Component total pressure ratio and efficiency under different MSV openging angles.
oblique shock wave. Hence, the second stage rotor nearly retreats the choke state. Nevertheless, the flow structures in the first stage remain almost unchanged. Generally speaking, during this procedure, the total pressure ratios should increase for the 2nd stage and stay nearly unchanged for the 1st stage. However, when the MSV is further opened (Fig. 8(c)), the shock waves in both the 1st and the 2nd stage rotors are pushed upstream obviously, and the total pressure ratios for both stages will increase, i.e. the operating points move to the left along the fan operating line as shown in Fig. 5(a). The occurrence of the outer bypass flow recirculation will inevitably bring about distortions to the downstream components. To investigate the influence of the distortion, flow characteristics at three positions are evaluated. As shown in Fig. 1, Plane1 is located at the near downstream of the MSV, Plane2 is at the inlet of the CDFS rotor, and Plane3 is at the outlet of the CDFS stator. The radial distribution of the normalized total pressure and total temperature at the three planes are given in Fig. 9. Cases to be considered are the same as those in Fig. 7 and Fig. 8. At Plane1, the total pressure in Cases 1.3 and 1.5 drops dramatically in the tip region and then increases quickly between 60% and 90% span. Meanwhile, the total pressure in the regions below 60% span is reduced, thus aggravating the distortion in the radial direction. The total temperature at the tip region of Plane1 is increased because of the mixing effect caused by the recirculation flow. As flow evolves downstream, at Plane2, the pressure and temperature dis-
tortions induced by the outer bypass flow recirculation are still remarkable, though weakened slightly by radial mixing effects. Due to the inflow distortions, flow field at the CDFS outlet distributes differently from the design point. In comparison with the Case1.1, the areas between 60% and 80% span witness a serious total pressure reduction in Case1.3, further opening the MSV will magnify the distortion. In Case1.5, the total pressure below 20% span is reduced significantly, implying flow separation in these areas. Last but not least, the total temperature distortion incurred by the outer bypass recirculation will also spread far downstream, the temperature variation still exists at plane3, although not as pronounced as that at the upstream. Fig. 10 gives the flow field comparison between Case1.1 and Case1.5, contours of the absolute Mach number nearby the inner bypass splitter ring are presented. As mentioned above, the inner bypass ratio tends to rise with the opening of the MSV, hence the inner bypass Mach number will get higher as mass flow increases, which coincides with the flow field. Besides, in Fig. 10(b), distortions caused by the outer bypass recirculation could spread so far away that the velocity distribution in the downstream of the CDFS is still influenced, and Mach number upstream of the inner bypass splitter ring is reduced remarkably. Additionally, it can be observed that in Case1.5, a separation region is formed in the hub region at the CDFS outlet, thus again explaining the total pressure reduction in Fig. 9.
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Fig. 7. Radial distribution of the fan stage total pressure ratio.
Fig. 8. Relative Mach number at fan 90% blade height.
Fig. 9. Radial distribution of normalized total pressure and total temperature at given positions.
Fig. 10. Absolute Mach number nearby the inner bypass splitter ring.
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Table 2 Cases to investigate the influence of the bypass backpressure.
Fig. 11. Aerodynamic performance of CDFS rotor.
Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case2.1 Case2.2 Case2.3 Case2.4 Case2.5
8
1.0
1.0
1.15 1.09 1.00 0.91 0.83
1.0
1.0
the CDFS matching point. Moreover, as presented in Fig. 14, with the strengthening of flow recirculation, the pressure surface of the CDFS stator will be occupied by large-scale separation flow, and the flow at the hub region on blade suction surface will also separate, which is harmful to the overall aerodynamic performance of the compression system. The outer bypass recirculation will also influence the aerodynamic performance of the high pressure compressor. Radial distributions of the HPC incidence angle and the total pressure ratio under different MSV opening conditions are illustrated in Fig. 15. According to Fig. 15(a), with the increase of the MSV opening angle, incidence angle at HPC inlet becomes lower at midspan while gets higher in the hub and tip regions, indicating mass flow redistribution along the spanwise direction. Correspondingly, the total pressure ratio of the HPC first stage presents a similar trend. Nonetheless, as shown in Fig. 15(b), after the mixing through the first stage, the incidence angle at the second stage inlet from different cases are in accordance with each other, not to mention the total pressure ratio. That is to say that the influence of the outer bypass flow recirculation on the HPC is largely confined to its first stage, and that the aft stages could work steadily regardless of the inlet condition. 3.2. The influence of the bypass backpressure
Fig. 12. Aerodynamic performance of CDFS stator.
Radial distributions of the CDFS rotor incidence angle and total pressure ratio are shown in Fig. 11. Cases presented are the same as those in Fig. 4. According to Fig. 4, the outer bypass flow recirculation will cause the flow at the casing region of the CDFS inlet to separate, thus weakening the local through flow capability. As a result, the rotor incidence angle above 90% span becomes higher as MSV opens up, and the total pressure ratio in the corresponding area improves. On the other hand, with the opening of the MSV, recirculation flow added into the CDFS inlet will increase the local mass flow rate, hence the incidence angle between 60% and 90% span is reduced, and the total pressure ratio below 90% span is decreased significantly, as shown in Fig. 11(a) and (b). Moreover, the reduction of the rotor total pressure ratio will cause the incidence angle at stator inlet to drop dramatically, as demonstrated in Fig. 12(a). Consequently, the stator is unloaded and the total pressure recovery coefficient shown in Fig. 12(b) exhibits a tendency to decrease due to the deterioration of the channel flow. Closer inspection of Fig. 10(b) and Fig. 12(b) shows that the radial distortion caused by the flow recirculation has a negative effect on the stator performance, which adds the total pressure loss between the 60∼90% span. The comparison of the relative Mach number at 80% span of the CDFS is presented in Fig. 13. To illustrate the variation of flow field more clearly, the position of the normal shock wave in rotor channel from the Case1.1 is depicted in the other two cases. As shown in Fig. 13(a), in order to satisfy the performance requirements of the double bypass mode, under the single bypass mode, the CDFS is matched at a low operation point, hence the normal shock wave in the rotor blade channel is close the blade trailing edge. When the MSV is opened, see Fig. 13(b) and (c), the normal shock wave in the rotor blade channel is pushed downstream, further lowering
According to former discussion, transitioning from the single bypass mode to the double bypass mode by opening the MSV alone will incur the outer bypass flow recirculation and influence the matching of the compression system. Therefore, in this part, methods to improve the system matching characteristic are investigated. Firstly, the influence of the bypass backpressure is studied. As shown in Table 2, the backpressure at the bypass outlet is varied from 1.15p 2ref to 0.83p 2ref , while the other control parameters are kept constant. System bypass ratios are presented in Fig. 16. It can be seen that all the cases to be considered have a negative B out , indicating the outer bypass flow recirculation. However, with the decrease of the bypass backpressure, the absolute value of B out tends to reduce, which suggests that the recirculation is suppressed. Meanwhile, the mass flow rate of the inner bypass is increased, leading to the increase of B tot and B in . Moreover, the variation of the bypass ratio will slow down with the decrease of p 2 , so it is not enough to eliminate the outer bypass recirculation by controlling the bypass backpressure alone. Variations of the component matching point are illustrated in Fig. 17. Results show that reducing the bypass backpressure will push the fan towards its choke line. Since opening the MSV will raise the fan operation point, decreasing the bypass backpressure will benefit the fan by increasing its stall margin. Nevertheless, the corrected mass flow of the CDFS tends to increase with the decrease of the bypass backpressure, which would push the CDFS to the deep chocking condition. Additionally, decreasing p 2 allows more air to go through the inner bypass, hence the mass flow rate of the HPC is reduced, causing the working point to transfer to the left on the operating line, as can be seen in Fig. 17(c). The aerodynamic performance of different compression components are further demonstrated in Fig. 18. According to Fig. 18(a),
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Fig. 13. Relative Mach number at CDFS 80% blade height.
Fig. 14. Surface streamlines and isentropic Mach number on CDFS stator under different MSV position.
Fig. 15. Radial distributions of the HPC aerodynamic performance.
backpressure could barely influence the total pressure ratio of the CDFS. This may be misleading because the matching characteristic of the CDFS is actually changed greatly, which has been discussed above. As shown in Fig. 18(b), the efficiency of the CDFS suffers a significant drop as the bypass backpressure decreases from 1.15p 2ref to 1.09p 2ref , and then keeps virtually identical at lower backpressures. On the other hand, the isentropic efficiency of the fan is first improved and then declined slightly with the reduction of p 2 , during which the HPC efficiency decreases slightly. The definition of the corrected rotational speed is as follows: Fig. 16. Bypass ratios under different bypass backpressure.
with the decrease of the bypass backpressure, the fan sure ratio is reduced by approximately 10%, while the sure ratio of the HPC is increased by nearly the same other noteworthy phenomenon is that the variation of
total prestotal presrange. Anthe bypass
ncor = n
288.15 ∗ T in
∗ denotes the total pressure at the inlet of each component. Here, T in It can therefore be seen in Fig. 18(c) that the corrected rotational
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Fig. 17. Component matching point at different bypass backpressure.
Fig. 18. System matching characteristics under different bypass backpressure.
Fig. 19. Relative Mach number at fan 90% span.
speed of both the CDFS and the HPC are raised with the reduction of p 2 owing to the pressure ratio decrement of the front fan. The real matching characteristic of the compression component depends on the combined effect of both the mass flow and the rotational speed. In order to further reveal the changing details of the flow field, Fig. 19 provides the Relative Mach number at 90% fan blade height. In Case2.1, the fan is operated near the stall line and the shock waves are detached from the leading edge of both rotor stages. The relatively high pre-shock Mach number adds to the shock loss. Yet in Case2.5, shock waves are swallowed deeply into the rotor blade channels, which will not only improve the component efficiency but also extend the stall margin. Given that opening the MSV will transfer the operation point of the fan close to the surge point, it is safe to say that lowering the bypass backpressure during mode transition will benefit the aerodynamic performance of the fan. Fig. 20 depicts the variation of the CDFS efficiency. Both the rotor efficiency and stage efficiency are illustrated. It can be seen that the CDFS stage efficiency is much lower than the rotor efficiency, which indicates that losses in the CDFS stator channel should be considerably high. Additionally, the difference between the stage efficiency and rotor efficiency turns out to increase with the decrease of the bypass backpressure, implying the promotion of the
Fig. 20. CDFS efficiency under different bypass backpressure.
stator loss. The stator total pressure recovery coefficient presented in Fig. 21 confirms the former deduction. In Case2.1 (case with the highest p 2 ), the total pressure recovery efficiency is already below 0.98, which is a relatively low value for a stator row. Yet reducing p 2 will further contribute to the drop of the total pressure recovery coefficient, thus deteriorating the stator flow. Flow profiles of the CDFS rotor along the radial direction are shown in Fig. 22. Cases presented are the same as those in Fig. 19. According to Fig. 22(a), under the high bypass backpressure condition (Case2.1), the incidence angle at the rotor tip is comparably
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high, and it drops quickly at lower blade heights. The radial distribution of the incidence angle is highly nonuniform due to the
Fig. 21. Total pressure recovery coefficient of the CDFS stator under different bypass backpressure.
Fig. 22. Aerodynamic performance of CDFS rotor.
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outer bypass flow recirculation. Meanwhile, for the case with the low backpressure condition (Case2.5), the incidence angle at the tip region is lowered with the weakening of the outer bypass recirculation and is improved for the region below 80% blade height. Correspondingly, in comparison with Case2.1, the rotor blade in Case2.5 gives lower total pressure ratio above 80% span, and higher total pressure ratio at lower blade heights, as shown in Fig. 22(b). Changes in the rotor performance will bring about changes in the stator channel. As shown in Fig. 23(a), the negative incidence angle at the CDFS stator inlet suggests that the stator row is matched at a near-choke point. Although decreasing the bypass backpressure could suppress the outer bypass flow recirculation, which is beneficial for the compression system, the positive effect is accompanied by a remarkable reduction of the incidence angle at the stator tip. As a result, the stator tip in Case2.5 is further choked, worsening the flow conditions on the blade pressure surface. Streamlines demonstrated in Fig. 24 tells that the separation region on the CDFS pressure surface is still distinctive even if the recirculation mass flow is reduced by decreasing the bypass backpressure, and that lowering the bypass backpressure will result in suction surface separation at the stator hub areas. Taken the loss characteristics into consideration (Fig. 23(b)), the near-hub total pressure recovery coefficient in Case2.5 is much lower than that in Case2.1 due to the suction surface flow separation. Also, since the separation region on blade pressure surface will move upward at the lower bypass backpressure condition (Case2.5), the total pressure recovery coefficient decrement caused by the pressure surface separation will happen at higher span areas. In summary, reducing the bypass backpressure has turned out to be an effective method to alleviate the outer bypass flow recirculation, but will further choke the CDFS. Controlling the bypass backpressure alone is far from enough to ensure smooth mode transition. 3.3. The influence of the FVABI position
Fig. 23. Aerodynamic performance of CDFS stator.
The substantial reason for the outer bypass circulation is the static pressure discrepancy at the junction of the inner and outer bypasses. It is therefore of great importance to modulate FVABI so as to optimize the mixing process. This part will focus on the effect of the FVABI position during the mode transition. As shown in Table 3, four FVABI positions are considered with the other parameters held constant. Variations of the bypass ratio for different FVABI positions are presented in Fig. 25. It can be seen that due to the weakening of the inner bypass through flow capability, the inner bypass ratio B in tends to decrease with the closure of the inner bypass (increase l). However, B out is barely changed in the given cases, hence the total bypass ratio B tot presents a similar trend to the B in . It should be noted that the slightness of B out variation does not manifest the uselessness of the FVABI. As a matter of fact, closing
Fig. 24. Surface streamlines and isentropic Mach number on CDFS stator under different bypass backpressure.
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the FVABI will improve the component matching characteristics. See Fig. 26, the closure of the FVABI could raise the operation point of the CDFS and the front fan, the total pressure ratio of the CDFS is increased significantly. Meanwhile, the performance of the HPC is influenced by two aspects: the increase of the inlet pressure (because of the total pressure ratio rise for the fan and the CDFS) and the increase of the mass flow rate (because of the reduction of B in ), both of which add to the decrement of its total pressure ratio. The comparisons of the component operation points are provided in Fig. 27. Results show that the increment in CDFS total pressure ratio during the closure of the FVABI adds up to 5%, which does great benefit to its efficiency. Compared to Case3.1, the stage efficiency of the CDFS in Case3.4 is improved by more than 5 points. Further inspection of CDFS efficiency shows that the improvement mainly comes from the stator row. As shown in
Table 3 Cases to investigate the influence of the FVABI position. Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case3.1 Case3.2 Case3.3 Case3.4
8
1.0 3.0 5.0 7.0
1.0
1.0
1.0
1.0
Fig. 25. Bypass ratios under different FVABI positions.
Fig. 27(c), the total pressure recovery coefficient of the CDFS stator is increased from 0.965 to 0.984 with the closure of the FVABI. The total pressure ratio of each fan rotor stage in Fig. 28 reveals that the closure of the FVABI will upload the tip region of the 1st stage rotor, which is a typical character for a transonic fan stage. Influenced by the front stage, the loading at the hub region of the 2nd stage rotor is increased. Contours of the relative Mach number in Fig. 29 confirm the prior trend. The shock wave positions for Case3.1 are depicted for comparison. It can be observed in Fig. 29(a) that at 10% span, the shock wave in the 2nd stage rotor would move upstream when the FVABI is closed. Similarly, the channel shock wave in the 1st stage rotor tip (90% span) is pushed upstream in Case3.4, as shown in Fig. 29(b). Comparisons of the CDFS radial aerodynamic performance between Case3.1 and Case3.4 are illustrated in Fig. 30. Fig. 30(a) shows that the rotor total pressure ratio is enhanced along the whole blade span with the closure of the FVABI. As a result, the incidence angle at the stator inlet is improved remarkably (Fig. 30(b)), saving the stator blade from the near-choke operating condition. In addition, the pressure rise ability of the stator row will recover with the increase of the incidence angle, hence the De Haller number in Fig. 30(c) decreases dramatically in Case3.4. Closer inspection of Fig. 30(c) tells that the De Haller number declination at 70% blade height and the hub corner has disappeared as the FVABI closes up, manifesting the elimination of the flow separation in corresponding regions. Consequently, flow loss in the stator row is decreased, as the total pressure recovery coefficient in Case3.4 rises significantly (Fig. 30(d)). The surface streamlines and the isentropic Mach number on the CDFS stator blade for Case3.1 and Case3.4 are presented in Fig. 31 so as to display the flow characteristics more clearly. In Case3.1 (Fig. 31(a)), flow separation not only takes place at the middle region of the blade pressure surface, but also at the hub corner of the suction surface, which is in accordance with the former discussion. On the other hand, the flow field in Case3.4 shows that the separation regions on both blade sides are removed, indicating the beneficial effect of the FVABI closure on the CDFS stator.
Fig. 26. Component matching point at different FVABI positions.
Fig. 27. System matching characteristics under different FVABI positions.
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Fig. 28. Radial distribution of the fan rotor total pressure ratio.
Fig. 29. Relative Mach number at 10% and 90% fan span.
Fig. 30. Radial distribution of CDFS aerodynamic performance.
Fig. 31. Surface streamlines and isentropic Mach number on CDFS stator under different FVABI positions.
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4. Conclusions By establishing a three dimensional model of a double bypass variable cycle compression system, the flow patterns and matching characteristics of the fan, the CDFS, and the HPC are studied using numerical simulation. The influence of the MSV opening angle on the system aerodynamic performance is investigated, and methods to ensure smooth mode transition are discussed. Transition of the compression system from the single bypass mode to the double bypass mode by opening the MSV alone would increase the fan operation point while decreasing that of the CDFS and the HPC. It would also incur the outer bypass flow recirculation and bring radial inflow distortions to the CDFS. Influences on the fan lie mainly in its aft stage, while the HPC first stage undertakes most of the distortion. Reducing the bypass backpressure during mode transition is an effective way to alleviate the outer bypass recirculation, but would further choke the CDFS. The CDFS stator would consequently be matched at a quite low point and suffer serious flow separation. Closing the FVABI could hardly suppress the outer bypass recirculation, but would raise the CDFS total pressure ratio and eliminate the flow separation in the stator row. Hence, it is recommended to decrease the bypass backpressure and close the FVABI simultaneously in the real transition process. Declaration of competing interest No competing interest. Acknowledgements The authors would like to acknowledge the supports of National Natural Science Foundation of China (Grant No. 51776010 and 51790511), National Science and Technology Major Project (Grant No. 2017-I I-0001-0013). References [1] R. Brown, Integration of a variable cycle engine concept in a supersonic cruise aircraft, in: 14th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1978. [2] E. Willis, A. Welliver, Variable-cycle engines for supersonic cruising aircraft, in: 12th Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1976. [3] R. Allan, General Electric Company variable cycle engine technology demonstrator programs, in: 15th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1979. [4] M.A.R. Do Nascimento, P. Pilidis, The selective bleed variable cycle engine, in: Vol. 2 Aircr. Engine, Mar. Microturbines Small Turbomach, ASME, 1991, V002T02A033. [5] T. Dabney, M. Hirschberg, Engine wars – competition for U.S. fighter engine production, in: 34th AIAA/ASME/SAE/ASEE Jt. Propuls. Conf. Exhib., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1998. [6] M. French, C. Allen, NASA VCE test bed engine aerodynamic performance characteristics and test results, in: 17th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1981.
[7] J.J.E. John W. Vdoviak, Paul R. Knott, Aerodynamic/Acoustic Performance of YJ101/Double Bypass VCE with Coannular Plug Nozzle, 1981, Cincinnati. [8] R.K. Denney, J.C. Tai, B.K. Kestner, D.N. Mavris, Variable cycle optimization for supersonic commercial applications, in: SAE Tech. Pap. Ser., 2005. [9] Y. Lyu, H. Tang, M. Chen, A study on combined variable geometries regulation of adaptive cycle engine during throttling, Appl. Sci. 6 (2016) 374, https://doi. org/10.3390/app6120374. [10] B. Liu, S. Jia, X. Yu, An integrated throughflow method for the performance analysis of variable cycle compression systems, Int. J. Turbo Jet-Engines (2018), https://doi.org/10.1515/tjj-2018-0010. [11] S. Adibhatla, K. Johnson, Evaluation of a nonlinear PSC algorithm on a variable cycle engine, in: 29th Jt. Propuls. Conf. Exhib., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1993. [12] X. Zhang, Z. Wang, J. Shi, Optimization of cycle parameters of variable cycle engine based on response surface model, in: 53rd AIAA/SAE/ASEE Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virginia, 2017, pp. 1–11. [13] E. Garvin, T. Sloss, R. Ress, S. Brooks, M. Murugan, Revolutionary concepts for multi-mode adaptive advanced cycle gas turbine engine, in: 2018 Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virginia, 2018. [14] H. Chen, H. Zhang, Z. Hu, Q. Zheng, The installation performance control of three ducts separate exhaust variable cycle engine, IEEE Access 7 (2019) 1764–1774, https://doi.org/10.1109/ACCESS.2018.2886369. [15] Y.S. Li, Y.C. Chen, Q. Zhao, Steady state calculation and performance analysis of variable cycle engine, in: Proc. 2018 9th Int. Conf. Mech. Aerosp. Eng. ICMAE 2018, 2018, pp. 352–356. [16] M. Chen, J. Zhang, H. Tang, Interval analysis of the standard of adaptive cycle engine component performance deviation, Aerosp. Sci. Technol. 81 (2018) 179–191, https://doi.org/10.1016/j.ast.2018.07.004. [17] X. Zheng, H. Yang, End-wall boundary layers and blockages of multistage axial compressors under different conditions, J. Propuls. Power 33 (2017) 908–916, https://doi.org/10.2514/1.B36251. [18] M. Banjac, M.V. Petrovic, A. Wiedermann, Multistage axial compressor flow field predictions using CFD and through-flow calculations, in: Vol. 2C Turbomach, ASME, 2016, V02CT39A047. [19] F. Wang, M. Carnevale, L. di Mare, S. Gallimore, Simulation of multistage compressor at off-design conditions, J. Turbomach. 140 (2017) 021011, https:// doi.org/10.1115/1.4038317. [20] H. Lu, Q. Li, T. Pan, Optimization of cantilevered stators in an industrial multistage compressor to improve efficiency, Energy 106 (2016) 590–601, https:// doi.org/10.1016/j.energy.2016.03.109. [21] S. Kim, K. Kim, C. Son, Three-dimensional unsteady simulation of a multistage axial compressor with labyrinth seals and its effects on overall performance and flow characteristics, Aerosp. Sci. Technol. 86 (2019) 683–693, https://doi. org/10.1016/j.ast.2019.01.055. [22] Y. Wu, G. An, B. Wang, Numerical investigation into the underlying mechanism connecting the vortex breakdown to the flow unsteadiness in a transonic compressor rotor, Aerosp. Sci. Technol. 86 (2019) 106–118, https://doi.org/10.1016/ j.ast.2018.12.040. [23] L. Zhang, S. Wang, W. Zhu, Application of endwall contouring in a highsubsonic tandem cascade with endwall boundary layer suction, Aerosp. Sci. Technol. 84 (2019) 245–256, https://doi.org/10.1016/j.ast.2018.08.041. [24] V. Matveev, E. Goriachkin, A. Volkov, Increase of gas-turbine plant efficiency by optimizing operation of compressors, IOP Conf. Ser., Mater. Sci. Eng. 302 (2018) 012030, https://doi.org/10.1088/1757-899X/302/1/012030. [25] S. Sun, S. Wang, S. Chen, C. Tao, L. Cai, J. Chen, The impact of various forward sweep angles on the performance of an ultra-high-load low-reaction transonic compressor rotor, Appl. Therm. Eng. 150 (2019) 953–966, https://doi.org/10. 1016/j.applthermaleng.2019.01.045. [26] C. Zhang, J. Hu, J. Li, Z. Wang, Three-dimensional compressor blading design improvements in low-speed model testing, Aerosp. Sci. Technol. 63 (2017) 179–190, https://doi.org/10.1016/j.ast.2016.12.033.
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Aerospace Science and Technology www.elsevier.com/locate/aescte
On the mode transition of a double bypass variable cycle compression system Baojie Liu a,b,c , Ruoyu Wang a , Xianjun Yu a,b,c,∗ a b c
School of Energy and Power Engineering, Beihang University, Beijing, China National Key Laboratory of Science and Technology on Aero-Engine Aero-Thermodynamics, Beijing, China Collaborative Innovation Center of Advanced Aero-engine, Beijing, China
a r t i c l e
i n f o
Article history: Received 18 June 2019 Received in revised form 5 December 2019 Accepted 22 January 2020 Available online 24 January 2020 Communicated by Dionysios Angelidis Keywords: Double bypass engine Mode transition Matching Flow recirculation
a b s t r a c t By establishing a three dimensional model of a double bypass variable cycle compression system, the flow patterns and matching characteristics of each working component during mode transition are investigated using numerical simulation. Results show that transition from the single bypass mode to double bypass mode by opening the mode selector valve (MSV) alone would increase the fan operation point while decreasing that of the core driven fan stage (CDFS) and the high pressure compressor (HPC). It would also incur outer bypass flow recirculation and bring radial inflow distortions to the CDFS. Deviations of the fan aerodynamic performance lie mainly in its aft stage, while the HPC first stage undertakes most of the inlet distortion. Reducing the bypass backpressure during mode transition is an effective way to alleviate the outer bypass flow recirculation but would further choke the CDFS. Closing the forward variable area bypass injector (FVABI) could raise the CDFS matching point so as to improve the stator performance without influencing the outer bypass ratio. It is recommended to decrease the bypass backpressure and close FVABI simultaneously in the real transition process. © 2020 Elsevier Masson SAS. All rights reserved.
1. Introduction As one of the most promising technologies in aircraft industry, variable cycle engine (VCE) could adapt to various flight conditions while maintaining the optimum fuel consumption. Double bypass engine (DBE) is a typical VCE concept. It is also a welldeveloped configuration which has been evaluated comprehensively and adopted by the famous variable cycle engine F120 [1–5]. By introducing the CDFS and the inner bypass into the engine compression system, DBE is able to achieve the tailoring of internal airflow, thus combining both the features of turbofan and turbojet [3,6]. To be specific, there are two operation modes for the DBE, and the ‘turbofan’ mode is called the double bypass mode where a large portion of the inflow air is bypassed through the fan bypass, whereas the ‘turbojet’ mode is named as the single bypass mode where the fan bypass is shut off hence most air enters the core engine. Mode transition is undoubtedly a key point for DBE. Report of the NASA VCE test bed engine points out that the mode transi-
*
Corresponding author at: School of Energy and Power Engineering, Beihang University, Beijing, China. E-mail address: [email protected] (X. Yu). https://doi.org/10.1016/j.ast.2020.105743 1270-9638/© 2020 Elsevier Masson SAS. All rights reserved.
tion has to be conducted skillfully to avoid component or bypass mismatch [6,7]. The report also states that improper transition process could incur outer bypass flow recirculation within the engine compression system, thus worsening the overall aerodynamic performance. Such conclusions have been verified by later investigations [8]. Over the years, researchers have made progress in VCE performance prediction by proposing a number of simulation models [9–16]. However, the precise control methodology and the matching characteristic of the working components are rarely found in the open literature. As a matter of fact, in order to get the overall engine performance quickly, past investigations are mostly based upon the mathematical relations of component operation point and are unable to acquire the real flow conditions inside the engine. On the other hand, the internal flow is actually of great significance, as a well-organized flow field guarantees a wellperformed engine. To shed some light on the flow details of the variable cycle engine, in this study, three-dimensional numerical simulation is employed to calculate a DBE compression system. A series of conditions during mode transition is considered so as to investigate the influence of different control parameters. Based on the results, component operation points and inter-stage matching characteristics are summarized, and recommendations for the transition process are proposed.
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Nomenclature Variables B tot B in B out i l m Ma n p
α η∗ π∗ σ∗
Subscripts/Superscripts Total bypass ratio Inner bypass ratio Outer bypass ratio Incidence angle FVABI length Mass flow Mach number Rotational speed Pressure MSV opening angle Isentropic efficiency Total pressure ratio Total pressure recovery coefficient
is ref rel
∗
Isentropic Reference value Relative condition Total condition
Abbreviations CDFS DBE FVABI HPC MSV PS SS VCE
Core driven fan stage Double bypass engine Forward variable area bypass injector High pressure compressor Mode selector valve Pressure surface Suction surface Variable cycle engine
2. Model and numerical methods
3. Results and discussion
As shown in Fig. 1, the variable cycle compression system to be investigated consists of a two-stage fan, a one-stage CDFS, and a five-stage high pressure compressor. Both fan and CDFS have inlet guide vanes to adjust the inflow directions for the downstream rows. Main system control mechanisms including mode selector valve (MSV) and forward variable area bypass injector (FVABI) are also modeled in the present study. System transition between single and double bypass mode is accomplished through MSV, which stays closed in the single bypass mode hence all the fan air flows into the CDFS. On the other hand, MSV is opened during the double bypass mode, allowing a portion of the fan air to go through the outer bypass. Setting the position of MSV under single bypass mode as a reference, the angle between the reference position and the MSV real position is defined as α . Meanwhile, FVABI is installed at the outlet of the inner bypass. The axial movement of FVABI could modulate the outflow area of the inner bypass, thus ensuring efficient mixing of the inner and outer bypass streams. The axial length of FVABI is denoted by l. In the calculations, atmospheric total pressure and temperature were given at the inlet of the fan, while static pressure p 1 and p 2 were imposed at the outlet of HPC and the bypass, respectively. The physical rotational speed of the fan is set as n1 , while CDFS and HPC are operated under the rotational speed of n2 . Schematic of the geometrical model is presented in Fig. 2. The compression system was constructed by Unigraphics NX, NUMECA Autogrid5 and IGG were then utilized to generate the mesh for the blade region and the complicated bypass ducts, respectively. The O4H topology was used for each blade channel, whereas Otopology was adopted for the skin block around the blade. Based on the turbulent model to be selected, the value of the first grid off wall is set as y+ < 10. The grid number totals 15 million with approximately 0.8 million for each blade row, which has been proved to be adequate for simulating multistage compressor performance [17–21]. Commercial flow solver NUMECA FINE/TURBO was employed for the numerical simulations. As a prevalent multistage turbomachinery solver, validations of FINE/TURBO can be found in lots of open literature [22–26]. In the present study, three dimensional compressible Reynolds-averaged Navier-Stokes equations were solved to obtain the steady solutions. The SpalartAllmaras turbulent model was adopted, the temporal discretization scheme was Four-stage Runge-Kutta, while the spatial terms were discretized by the second-order central difference scheme. Cases were considered converged when the residuals became lower than 10−6 . Generally, solutions tended to converge after 2000 iterations.
3.1. The influence of the MSV position In order to get the matching characteristics of the compression system during mode transition, in this part, transition is conducted by simply opening the MSV. As the system changes from single bypass mode to double bypass mode, six MSV positions are considered, detailed control parameters are listed in Table 1. To get an overview of the system mass flow distribution, three bypass ratios are defined as follows:
B in =
mBYPASSin mHPC
,
B out =
mBYPASSout mCDFS
,
B tot =
mBYPASStot mHPC
where B in is the mass flow ratio between the inner bypass and the HPC, B out stands for the ratio of mass flow between the outer bypass and the CDFS, and B tot represents the mass flow of the total bypass divided by that of the HPC. As can be seen in Fig. 3, with the opening of MSV, the values of both B in and B tot tend to increase, while B out tends to decrease, indicating obvious mass flow redistribution. It should also be noted that for all the cases, B out has a negative value, which means that there is stronger and stronger flow recirculation in the outer bypass as MSV opens up. Furthermore, the outer bypass flow recirculation will reduce the mass flow in the total bypass, hence the variation range of B tot is smaller than that of B in . Fig. 4 gives the absolute Mach number in the vicinity of the MSV, Cases 1.1, 1.3 and 1.5 are illustrated. Streamlines are also depicted to show the flow field more clearly. By comparing these three cases, the evolution of flow recirculation can be recognized easily: The Mach number in the outer bypass increases continuously with the opening of MSV, while the outer bypass airflow gets into the CDFS inlet. Due to the outer bypass flow recirculation, in the downstream of the MSV, flow first separates from and then reattaches to the casing wall, forming a low-speed separation region at the upstream of the CDFS. The separation region will enlarge with the increase of α , as can be seen in Fig. 4(b) and (c). Meanwhile, the injection of the air accelerates the flow outside the separation region, hence the CDFS inlet Mach number near the casing will get higher if mode transition is conducted in the present way. The variations of component matching point are presented in Fig. 5, characteristics of each component alone (with uniform inlet flow conditions) are also given for the convenience of observation. It should be noted that the flow interaction within the compression system will actually influence the inflow and outflow con-
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Fig. 1. Schematic of the variable cycle compression system.
Fig. 2. Geometric model and mesh for the compression system.
Fig. 3. Bypass ratios under different MSV opening angles.
Table 1 Cases to investigate the influence of MSV position. Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case1.1 Case1.2 Case1.3 Case1.4 Case1.5 Case1.6
0 2 4 6 8 10
1.0
1.0
1.0
1.0
1.0
ditions of each component, hence the component real matching point will not coincide with their original characteristics. According to Fig. 5(a), the fan is matched at a near-choke point under the single bypass mode, and opening the MSV will push the fan operation point towards the surge line. On the contrary, Fig. 5(b) shows that the CDFS will get into deep choke during the mode transition, which is detrimental for its aerodynamic performance. Meanwhile, the operation point of HPC will move towards the surge line as MSV opens up, see Fig. 5(c).
To compare the influence of MSV position on different components quantitatively, Fig. 6 presents the component performance with the variation of α . It can be seen that the aerodynamic performance deviation of the fan and CDFS is much greater than that of the HPC. As α increases from 0◦ to 10◦ , the fan total pressure ratio is increased by nearly 10%, whereas π ∗ HPC only changes 3%. Additionally, the variation of η∗ HPC is lower than 1%, as shown in Fig. 6(b). As depicted in Fig. 7, the radial distributions of total pressure ratio under different MSV positions (Case1.1, Case1.3 and Case1.5) reveal the fan stage flow characteristics more clearly. It can be seen that when MSV is opened from 0◦ (Case1.1) to 4◦ (Case1.3), the increase in fan total pressure ratio comes mainly from its second stage, whereas the first stage pressure ratio barely changes. However, as the opening angle changes from 4◦ to 8◦ , total pressure ratios for both stages improve distinctly. Compared to the second stage, in which the pressure ratio increase occurs along the whole blade span, the total pressure ratio increase appears only at the upper blade region in the first stage. Fig. 8 further gives the fan relative Mach number at 90% blade height. For the convenience of comparison, shock wave positions for Case1.1 are depicted in all three cases with dashed lines. As can be seen in Fig. 8(a), being a typical transonic stage, the flow at the tip region in the first stage rotor is made up of an oblique shock wave at the inlet and a normal shock wave, i.e. the passage cutoff shock wave, in the blade channel. As for the second stage rotor, a relatively weak passage cutoff shock wave also exists, which means that the fan is matched at a deep choke state under the single bypass mode. According to Fig. 8(b), when MSV is opened by 4◦ , the normal shock wave in the second stage rotor channel undergoes an obvious upstream movement and merges largely with the inlet
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Fig. 4. Absolute Mach number in the vicinity of MSV. (For interpretation of the colors in the figure(s), the reader is referred to the web version of this article.)
Fig. 5. Component matching point at different MSV positions.
Fig. 6. Component total pressure ratio and efficiency under different MSV openging angles.
oblique shock wave. Hence, the second stage rotor nearly retreats the choke state. Nevertheless, the flow structures in the first stage remain almost unchanged. Generally speaking, during this procedure, the total pressure ratios should increase for the 2nd stage and stay nearly unchanged for the 1st stage. However, when the MSV is further opened (Fig. 8(c)), the shock waves in both the 1st and the 2nd stage rotors are pushed upstream obviously, and the total pressure ratios for both stages will increase, i.e. the operating points move to the left along the fan operating line as shown in Fig. 5(a). The occurrence of the outer bypass flow recirculation will inevitably bring about distortions to the downstream components. To investigate the influence of the distortion, flow characteristics at three positions are evaluated. As shown in Fig. 1, Plane1 is located at the near downstream of the MSV, Plane2 is at the inlet of the CDFS rotor, and Plane3 is at the outlet of the CDFS stator. The radial distribution of the normalized total pressure and total temperature at the three planes are given in Fig. 9. Cases to be considered are the same as those in Fig. 7 and Fig. 8. At Plane1, the total pressure in Cases 1.3 and 1.5 drops dramatically in the tip region and then increases quickly between 60% and 90% span. Meanwhile, the total pressure in the regions below 60% span is reduced, thus aggravating the distortion in the radial direction. The total temperature at the tip region of Plane1 is increased because of the mixing effect caused by the recirculation flow. As flow evolves downstream, at Plane2, the pressure and temperature dis-
tortions induced by the outer bypass flow recirculation are still remarkable, though weakened slightly by radial mixing effects. Due to the inflow distortions, flow field at the CDFS outlet distributes differently from the design point. In comparison with the Case1.1, the areas between 60% and 80% span witness a serious total pressure reduction in Case1.3, further opening the MSV will magnify the distortion. In Case1.5, the total pressure below 20% span is reduced significantly, implying flow separation in these areas. Last but not least, the total temperature distortion incurred by the outer bypass recirculation will also spread far downstream, the temperature variation still exists at plane3, although not as pronounced as that at the upstream. Fig. 10 gives the flow field comparison between Case1.1 and Case1.5, contours of the absolute Mach number nearby the inner bypass splitter ring are presented. As mentioned above, the inner bypass ratio tends to rise with the opening of the MSV, hence the inner bypass Mach number will get higher as mass flow increases, which coincides with the flow field. Besides, in Fig. 10(b), distortions caused by the outer bypass recirculation could spread so far away that the velocity distribution in the downstream of the CDFS is still influenced, and Mach number upstream of the inner bypass splitter ring is reduced remarkably. Additionally, it can be observed that in Case1.5, a separation region is formed in the hub region at the CDFS outlet, thus again explaining the total pressure reduction in Fig. 9.
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Fig. 7. Radial distribution of the fan stage total pressure ratio.
Fig. 8. Relative Mach number at fan 90% blade height.
Fig. 9. Radial distribution of normalized total pressure and total temperature at given positions.
Fig. 10. Absolute Mach number nearby the inner bypass splitter ring.
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Table 2 Cases to investigate the influence of the bypass backpressure.
Fig. 11. Aerodynamic performance of CDFS rotor.
Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case2.1 Case2.2 Case2.3 Case2.4 Case2.5
8
1.0
1.0
1.15 1.09 1.00 0.91 0.83
1.0
1.0
the CDFS matching point. Moreover, as presented in Fig. 14, with the strengthening of flow recirculation, the pressure surface of the CDFS stator will be occupied by large-scale separation flow, and the flow at the hub region on blade suction surface will also separate, which is harmful to the overall aerodynamic performance of the compression system. The outer bypass recirculation will also influence the aerodynamic performance of the high pressure compressor. Radial distributions of the HPC incidence angle and the total pressure ratio under different MSV opening conditions are illustrated in Fig. 15. According to Fig. 15(a), with the increase of the MSV opening angle, incidence angle at HPC inlet becomes lower at midspan while gets higher in the hub and tip regions, indicating mass flow redistribution along the spanwise direction. Correspondingly, the total pressure ratio of the HPC first stage presents a similar trend. Nonetheless, as shown in Fig. 15(b), after the mixing through the first stage, the incidence angle at the second stage inlet from different cases are in accordance with each other, not to mention the total pressure ratio. That is to say that the influence of the outer bypass flow recirculation on the HPC is largely confined to its first stage, and that the aft stages could work steadily regardless of the inlet condition. 3.2. The influence of the bypass backpressure
Fig. 12. Aerodynamic performance of CDFS stator.
Radial distributions of the CDFS rotor incidence angle and total pressure ratio are shown in Fig. 11. Cases presented are the same as those in Fig. 4. According to Fig. 4, the outer bypass flow recirculation will cause the flow at the casing region of the CDFS inlet to separate, thus weakening the local through flow capability. As a result, the rotor incidence angle above 90% span becomes higher as MSV opens up, and the total pressure ratio in the corresponding area improves. On the other hand, with the opening of the MSV, recirculation flow added into the CDFS inlet will increase the local mass flow rate, hence the incidence angle between 60% and 90% span is reduced, and the total pressure ratio below 90% span is decreased significantly, as shown in Fig. 11(a) and (b). Moreover, the reduction of the rotor total pressure ratio will cause the incidence angle at stator inlet to drop dramatically, as demonstrated in Fig. 12(a). Consequently, the stator is unloaded and the total pressure recovery coefficient shown in Fig. 12(b) exhibits a tendency to decrease due to the deterioration of the channel flow. Closer inspection of Fig. 10(b) and Fig. 12(b) shows that the radial distortion caused by the flow recirculation has a negative effect on the stator performance, which adds the total pressure loss between the 60∼90% span. The comparison of the relative Mach number at 80% span of the CDFS is presented in Fig. 13. To illustrate the variation of flow field more clearly, the position of the normal shock wave in rotor channel from the Case1.1 is depicted in the other two cases. As shown in Fig. 13(a), in order to satisfy the performance requirements of the double bypass mode, under the single bypass mode, the CDFS is matched at a low operation point, hence the normal shock wave in the rotor blade channel is close the blade trailing edge. When the MSV is opened, see Fig. 13(b) and (c), the normal shock wave in the rotor blade channel is pushed downstream, further lowering
According to former discussion, transitioning from the single bypass mode to the double bypass mode by opening the MSV alone will incur the outer bypass flow recirculation and influence the matching of the compression system. Therefore, in this part, methods to improve the system matching characteristic are investigated. Firstly, the influence of the bypass backpressure is studied. As shown in Table 2, the backpressure at the bypass outlet is varied from 1.15p 2ref to 0.83p 2ref , while the other control parameters are kept constant. System bypass ratios are presented in Fig. 16. It can be seen that all the cases to be considered have a negative B out , indicating the outer bypass flow recirculation. However, with the decrease of the bypass backpressure, the absolute value of B out tends to reduce, which suggests that the recirculation is suppressed. Meanwhile, the mass flow rate of the inner bypass is increased, leading to the increase of B tot and B in . Moreover, the variation of the bypass ratio will slow down with the decrease of p 2 , so it is not enough to eliminate the outer bypass recirculation by controlling the bypass backpressure alone. Variations of the component matching point are illustrated in Fig. 17. Results show that reducing the bypass backpressure will push the fan towards its choke line. Since opening the MSV will raise the fan operation point, decreasing the bypass backpressure will benefit the fan by increasing its stall margin. Nevertheless, the corrected mass flow of the CDFS tends to increase with the decrease of the bypass backpressure, which would push the CDFS to the deep chocking condition. Additionally, decreasing p 2 allows more air to go through the inner bypass, hence the mass flow rate of the HPC is reduced, causing the working point to transfer to the left on the operating line, as can be seen in Fig. 17(c). The aerodynamic performance of different compression components are further demonstrated in Fig. 18. According to Fig. 18(a),
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Fig. 13. Relative Mach number at CDFS 80% blade height.
Fig. 14. Surface streamlines and isentropic Mach number on CDFS stator under different MSV position.
Fig. 15. Radial distributions of the HPC aerodynamic performance.
backpressure could barely influence the total pressure ratio of the CDFS. This may be misleading because the matching characteristic of the CDFS is actually changed greatly, which has been discussed above. As shown in Fig. 18(b), the efficiency of the CDFS suffers a significant drop as the bypass backpressure decreases from 1.15p 2ref to 1.09p 2ref , and then keeps virtually identical at lower backpressures. On the other hand, the isentropic efficiency of the fan is first improved and then declined slightly with the reduction of p 2 , during which the HPC efficiency decreases slightly. The definition of the corrected rotational speed is as follows: Fig. 16. Bypass ratios under different bypass backpressure.
with the decrease of the bypass backpressure, the fan sure ratio is reduced by approximately 10%, while the sure ratio of the HPC is increased by nearly the same other noteworthy phenomenon is that the variation of
total prestotal presrange. Anthe bypass
ncor = n
288.15 ∗ T in
∗ denotes the total pressure at the inlet of each component. Here, T in It can therefore be seen in Fig. 18(c) that the corrected rotational
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Fig. 17. Component matching point at different bypass backpressure.
Fig. 18. System matching characteristics under different bypass backpressure.
Fig. 19. Relative Mach number at fan 90% span.
speed of both the CDFS and the HPC are raised with the reduction of p 2 owing to the pressure ratio decrement of the front fan. The real matching characteristic of the compression component depends on the combined effect of both the mass flow and the rotational speed. In order to further reveal the changing details of the flow field, Fig. 19 provides the Relative Mach number at 90% fan blade height. In Case2.1, the fan is operated near the stall line and the shock waves are detached from the leading edge of both rotor stages. The relatively high pre-shock Mach number adds to the shock loss. Yet in Case2.5, shock waves are swallowed deeply into the rotor blade channels, which will not only improve the component efficiency but also extend the stall margin. Given that opening the MSV will transfer the operation point of the fan close to the surge point, it is safe to say that lowering the bypass backpressure during mode transition will benefit the aerodynamic performance of the fan. Fig. 20 depicts the variation of the CDFS efficiency. Both the rotor efficiency and stage efficiency are illustrated. It can be seen that the CDFS stage efficiency is much lower than the rotor efficiency, which indicates that losses in the CDFS stator channel should be considerably high. Additionally, the difference between the stage efficiency and rotor efficiency turns out to increase with the decrease of the bypass backpressure, implying the promotion of the
Fig. 20. CDFS efficiency under different bypass backpressure.
stator loss. The stator total pressure recovery coefficient presented in Fig. 21 confirms the former deduction. In Case2.1 (case with the highest p 2 ), the total pressure recovery efficiency is already below 0.98, which is a relatively low value for a stator row. Yet reducing p 2 will further contribute to the drop of the total pressure recovery coefficient, thus deteriorating the stator flow. Flow profiles of the CDFS rotor along the radial direction are shown in Fig. 22. Cases presented are the same as those in Fig. 19. According to Fig. 22(a), under the high bypass backpressure condition (Case2.1), the incidence angle at the rotor tip is comparably
B. Liu et al. / Aerospace Science and Technology 98 (2020) 105743
high, and it drops quickly at lower blade heights. The radial distribution of the incidence angle is highly nonuniform due to the
Fig. 21. Total pressure recovery coefficient of the CDFS stator under different bypass backpressure.
Fig. 22. Aerodynamic performance of CDFS rotor.
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outer bypass flow recirculation. Meanwhile, for the case with the low backpressure condition (Case2.5), the incidence angle at the tip region is lowered with the weakening of the outer bypass recirculation and is improved for the region below 80% blade height. Correspondingly, in comparison with Case2.1, the rotor blade in Case2.5 gives lower total pressure ratio above 80% span, and higher total pressure ratio at lower blade heights, as shown in Fig. 22(b). Changes in the rotor performance will bring about changes in the stator channel. As shown in Fig. 23(a), the negative incidence angle at the CDFS stator inlet suggests that the stator row is matched at a near-choke point. Although decreasing the bypass backpressure could suppress the outer bypass flow recirculation, which is beneficial for the compression system, the positive effect is accompanied by a remarkable reduction of the incidence angle at the stator tip. As a result, the stator tip in Case2.5 is further choked, worsening the flow conditions on the blade pressure surface. Streamlines demonstrated in Fig. 24 tells that the separation region on the CDFS pressure surface is still distinctive even if the recirculation mass flow is reduced by decreasing the bypass backpressure, and that lowering the bypass backpressure will result in suction surface separation at the stator hub areas. Taken the loss characteristics into consideration (Fig. 23(b)), the near-hub total pressure recovery coefficient in Case2.5 is much lower than that in Case2.1 due to the suction surface flow separation. Also, since the separation region on blade pressure surface will move upward at the lower bypass backpressure condition (Case2.5), the total pressure recovery coefficient decrement caused by the pressure surface separation will happen at higher span areas. In summary, reducing the bypass backpressure has turned out to be an effective method to alleviate the outer bypass flow recirculation, but will further choke the CDFS. Controlling the bypass backpressure alone is far from enough to ensure smooth mode transition. 3.3. The influence of the FVABI position
Fig. 23. Aerodynamic performance of CDFS stator.
The substantial reason for the outer bypass circulation is the static pressure discrepancy at the junction of the inner and outer bypasses. It is therefore of great importance to modulate FVABI so as to optimize the mixing process. This part will focus on the effect of the FVABI position during the mode transition. As shown in Table 3, four FVABI positions are considered with the other parameters held constant. Variations of the bypass ratio for different FVABI positions are presented in Fig. 25. It can be seen that due to the weakening of the inner bypass through flow capability, the inner bypass ratio B in tends to decrease with the closure of the inner bypass (increase l). However, B out is barely changed in the given cases, hence the total bypass ratio B tot presents a similar trend to the B in . It should be noted that the slightness of B out variation does not manifest the uselessness of the FVABI. As a matter of fact, closing
Fig. 24. Surface streamlines and isentropic Mach number on CDFS stator under different bypass backpressure.
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the FVABI will improve the component matching characteristics. See Fig. 26, the closure of the FVABI could raise the operation point of the CDFS and the front fan, the total pressure ratio of the CDFS is increased significantly. Meanwhile, the performance of the HPC is influenced by two aspects: the increase of the inlet pressure (because of the total pressure ratio rise for the fan and the CDFS) and the increase of the mass flow rate (because of the reduction of B in ), both of which add to the decrement of its total pressure ratio. The comparisons of the component operation points are provided in Fig. 27. Results show that the increment in CDFS total pressure ratio during the closure of the FVABI adds up to 5%, which does great benefit to its efficiency. Compared to Case3.1, the stage efficiency of the CDFS in Case3.4 is improved by more than 5 points. Further inspection of CDFS efficiency shows that the improvement mainly comes from the stator row. As shown in
Table 3 Cases to investigate the influence of the FVABI position. Cases
α (◦ )
l/lref
p 1 /p 1ref
p 2 /p 2ref
n1 /n1ref
n2 /n2ref
Case3.1 Case3.2 Case3.3 Case3.4
8
1.0 3.0 5.0 7.0
1.0
1.0
1.0
1.0
Fig. 25. Bypass ratios under different FVABI positions.
Fig. 27(c), the total pressure recovery coefficient of the CDFS stator is increased from 0.965 to 0.984 with the closure of the FVABI. The total pressure ratio of each fan rotor stage in Fig. 28 reveals that the closure of the FVABI will upload the tip region of the 1st stage rotor, which is a typical character for a transonic fan stage. Influenced by the front stage, the loading at the hub region of the 2nd stage rotor is increased. Contours of the relative Mach number in Fig. 29 confirm the prior trend. The shock wave positions for Case3.1 are depicted for comparison. It can be observed in Fig. 29(a) that at 10% span, the shock wave in the 2nd stage rotor would move upstream when the FVABI is closed. Similarly, the channel shock wave in the 1st stage rotor tip (90% span) is pushed upstream in Case3.4, as shown in Fig. 29(b). Comparisons of the CDFS radial aerodynamic performance between Case3.1 and Case3.4 are illustrated in Fig. 30. Fig. 30(a) shows that the rotor total pressure ratio is enhanced along the whole blade span with the closure of the FVABI. As a result, the incidence angle at the stator inlet is improved remarkably (Fig. 30(b)), saving the stator blade from the near-choke operating condition. In addition, the pressure rise ability of the stator row will recover with the increase of the incidence angle, hence the De Haller number in Fig. 30(c) decreases dramatically in Case3.4. Closer inspection of Fig. 30(c) tells that the De Haller number declination at 70% blade height and the hub corner has disappeared as the FVABI closes up, manifesting the elimination of the flow separation in corresponding regions. Consequently, flow loss in the stator row is decreased, as the total pressure recovery coefficient in Case3.4 rises significantly (Fig. 30(d)). The surface streamlines and the isentropic Mach number on the CDFS stator blade for Case3.1 and Case3.4 are presented in Fig. 31 so as to display the flow characteristics more clearly. In Case3.1 (Fig. 31(a)), flow separation not only takes place at the middle region of the blade pressure surface, but also at the hub corner of the suction surface, which is in accordance with the former discussion. On the other hand, the flow field in Case3.4 shows that the separation regions on both blade sides are removed, indicating the beneficial effect of the FVABI closure on the CDFS stator.
Fig. 26. Component matching point at different FVABI positions.
Fig. 27. System matching characteristics under different FVABI positions.
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Fig. 28. Radial distribution of the fan rotor total pressure ratio.
Fig. 29. Relative Mach number at 10% and 90% fan span.
Fig. 30. Radial distribution of CDFS aerodynamic performance.
Fig. 31. Surface streamlines and isentropic Mach number on CDFS stator under different FVABI positions.
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4. Conclusions By establishing a three dimensional model of a double bypass variable cycle compression system, the flow patterns and matching characteristics of the fan, the CDFS, and the HPC are studied using numerical simulation. The influence of the MSV opening angle on the system aerodynamic performance is investigated, and methods to ensure smooth mode transition are discussed. Transition of the compression system from the single bypass mode to the double bypass mode by opening the MSV alone would increase the fan operation point while decreasing that of the CDFS and the HPC. It would also incur the outer bypass flow recirculation and bring radial inflow distortions to the CDFS. Influences on the fan lie mainly in its aft stage, while the HPC first stage undertakes most of the distortion. Reducing the bypass backpressure during mode transition is an effective way to alleviate the outer bypass recirculation, but would further choke the CDFS. The CDFS stator would consequently be matched at a quite low point and suffer serious flow separation. Closing the FVABI could hardly suppress the outer bypass recirculation, but would raise the CDFS total pressure ratio and eliminate the flow separation in the stator row. Hence, it is recommended to decrease the bypass backpressure and close the FVABI simultaneously in the real transition process. Declaration of competing interest No competing interest. Acknowledgements The authors would like to acknowledge the supports of National Natural Science Foundation of China (Grant No. 51776010 and 51790511), National Science and Technology Major Project (Grant No. 2017-I I-0001-0013). References [1] R. Brown, Integration of a variable cycle engine concept in a supersonic cruise aircraft, in: 14th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1978. [2] E. Willis, A. Welliver, Variable-cycle engines for supersonic cruising aircraft, in: 12th Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1976. [3] R. Allan, General Electric Company variable cycle engine technology demonstrator programs, in: 15th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1979. [4] M.A.R. Do Nascimento, P. Pilidis, The selective bleed variable cycle engine, in: Vol. 2 Aircr. Engine, Mar. Microturbines Small Turbomach, ASME, 1991, V002T02A033. [5] T. Dabney, M. Hirschberg, Engine wars – competition for U.S. fighter engine production, in: 34th AIAA/ASME/SAE/ASEE Jt. Propuls. Conf. Exhib., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1998. [6] M. French, C. Allen, NASA VCE test bed engine aerodynamic performance characteristics and test results, in: 17th Jt. Propuls. Conf., American Institute of Aeronautics and Astronautics, Reston, Virigina, 1981.
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