Simulations and Analysis of Fuel Flow in an Injector Including Transient Needle Effects

Transcription

1 ILASS-Americas 24th Annual Conference on Liquid Atomization and Spray Systems, San Antonio, TX, May 2012 Simulations and Analysis of Fuel Flow in an Injector Including Transient Needle Effects K.D. Neroorkar 1, C.E. Mitcham II 2, A.H. Plazas 3, R.O. Grover, Jr 3 and D.P. Schmidt 1 1 University of Massachusetts Amherst 2 General Motors Powertrain 3 General Motors Research and Development Abstract A fully parallelized, 3 dimensional, transient flow solver based on the Homogeneous Relaxation Model was developed for modeling and analyzing flash boiling flows in fuel injectors. This model has been validated in past work on both cavitating and flash boiling flows in real injector geometries. The ability of this model to be coupled with different fluids has also been demonstrated. The focus of the current work is to analyze the flow of diesel fuel in a 5-hole microsac injector with tapered holes. Transient needle effects are captured by providing an initial mesh, and moving the needle through mesh topological changes involving addition and deletion of mesh cell layers. The experimental needle lift profile is used to impose the needle motion, and model validation is conducted against measurements of the injected mass flow profile over the injection duration. Additionally, it is observed that the transient motion of the needle greatly influences the characteristics of the internal flow. Corresponding Author: kneroork@engin.umass.edu

2 Introduction The performance of an internal combustion engine is greatly affected by the characteristics of the fuel injection system. The highly transient nature of the flow through these systems and their small size makes these systems extremely difficult to analyze experimentally. This difficulty is further compounded by the occurrence of multiphase phenomena like cavitation in diesel and flash boiling in gasoline injectors. Hence computational fluid dynamics (CFD) is often used to model the flow through fuel injectors and to provide insight into the physical processes taking place inside these systems [1, 2, 3, 4, 5, 6]. In the past, most of the CFD analyses of fuel injectors has been limited to steady state flows assuming the needle to be fixed at a given location due to the complexity of including the moving needle. Very limited studies have been presented to analyze these transient effects. The work of Margot et al. [7] and Payri et al. [8] used a method with pure mesh motion without changing the number of cells in the domain. They used the STAR-CD solver for their flow simulations. Lee and Reitz [9] used KIVA, which uses a snapper algorithm for needle motion similar to the method presented in the current paper. In KIVA, a structured hexahedral mesh is required in the entire flow domain which significantly increases the difficulty of mesh generation when considering arbitrary shaped injectors. The current work involves using the Homogeneous Relaxation Model (HRM), which was incorporated in the OpenFOAM CFD library for analyzing the flow of diesel fuel through a 5-hole microsac injector. The HRM was originally developed for flash boiling flows in gasoline injectors. However, its applicability to diesel flow has been validated in recent publications [10, 11]. The transient needle motion is captured by addition and removal of cell layers by using OpenFOAM s dynamic mesh library. An important feature of the OpenFOAM library is its ability to be used with polyhedral meshes. Model Formulation and Numerical method The flow solver used in the current work is called HRMFoam, and its detailed formulation and numerical method is presented in past work [12, 13, 14]. For the transient needle cases, the first order upwind method was used for the divergence terms in order to increase stability, and slip walls were assumed in the nozzles to allow for a coarser mesh to reduce computation time. A steady needle simulation at 1200 bar injection and 60 bar back pressure showed a mass flow rate decrease of 5% when boundary conditions were changed from slip to no-slip. The transient cases will therefore show a slightly higher mass flow rate due to the slip-wall assumption. Additionally, laminar flow was considered as a first step to investigate the effect of the transient needle on the flow. Results and Discussions Experimental data: The 5 hole microsac geometry simulated in the current work was taken from Payri et al. [15] and Benajes et al [16] and is shown in Fig. 1. The data presented in the work of Payri et al. included two configurations as shown in the figure: without needle (data from steady state flow test rig) and with needle (data from the Injection Discharge Rate Curve Indicator (IDRCI) using the Bosch method [17]). The injector holes are tapered with a k factor of 1.5. For the case without the needle, the injection pressure was 10 and 20 MPa, and the back pressure was varied to see the effect on mass flow rate. For the case with the needle, for different injection and back pressures, the energizing time was changed from 0.25 ms to 3ms. The lower energizing times (less than 1ms) demonstrate the injection rate curve in an injection event representative of an engine. The longer energizing times produce the steady results at full needle lift. In addition to the injection rate curves, the needle lift profiles were also recorded by Payri et al. [15] by using an injector holder fitted with a needle lift sensor. These profiles were used in the current study to govern the needle motion in the injector. Steady flow results: The diesel fuel was simulated using a surrogate obtained from the work of Weber et al. [18]. The comparison of the properties of this surrogate with industrial diesels was presented in Neroorkar et al. [11], and it was found that the density and vapor pressure were within the range of the industrial diesels. However, the variability in properties of diesel from region to region implies that the properties of the surrogate may not match exactly with the diesel used in any experiments. For example, it was found that the density of the diesel used in the experiments by Payri et al. [15] was different from the surrogate s density by around 1.5%. Also, it must be noted that the steady flow results still incorporate the transient terms from the transport equations however the boundaries are stationary. For the steady flow cases, a polyhedral mesh was generated in STARCCM+ with around 225,000 cells. The Fig. 2 shows the mass flow rate comparisons for the steady flow cases with and without the needle. The experimental results for the cases with the needle represent the stable mass flow rate obtained from 2

3 (a) with needle (a) with needle (b) without needle Figure 1. Geometry of microsac injector simulated in the current study the injection rate curves for the high energizing time case (3ms). From Bernoulli s equation, for non cavitating cases, the mass flow rate should be directly proportional to the square root of pressure drop. This behavior is well represented in both experimental and simulation results for both sets of cases. However, the absolute values are different and this can be due to the difference in the fuel properties, or some minor differences in the actual hydrogrinding radius of the nozzle versus the CAD model generated. The steady state results are important because they indicate that the HRMFoam model satisfactorily reproduces the data at full needle lifts for the microsac injector which will be used for transient calculations. It must be noted that the discrepancy in the mass flow rate seen in Fig. 2 will also affect the predictions for the transient cases when at full lift. Transient needle effects: The mesh for the transient cases was generated in two parts. The first part included the needle and seat region and was composed of hexahedral cells. The second part included the sac region along with the nozzles and was a tetrahedral mesh. Both these parts were gener- (b) without needle Figure 2. Mass flow rates for steady flow cases ated in the software ICEM CFD and are shown in Fig. 3. This configuration is used for the following reasons: since the needle motion is achieved by addition/deletion of cell layers in the mesh, a layered hexahedral mesh is required around the needle, however, the extension of this hexahedral mesh into the sac and nozzle regions for a multi-hole injector is non-trivial. It is believed that the current method can be applied to arbitrary injector designs due to the lack of constraint on the nozzle/sac mesh. Once the two parts of the mesh were generated, they were connected using the generalized grid interface (GGI) feature in OpenFOAM [19] which allows the coupling of two non-conformal regions through an interface similar to the one shown in Fig. 3. The initial mesh was generated with the needle being close to the seat to represent the fully closed position. In past work by Margot et al [7] and Lee and Reitz [9], the needle had to maintain a minimum lift from the seat in order to prevent degener- 3

4 Case a b Pinj (bar) Pback (bar) Pulse width (ms) Table 1. Test cases considered for transient simulations ate cells from forming. In the current work, a minimum lift of 25 microns was assumed. The current formulation allows the needle to be moved by providing a file of time versus the required displacement from the initial position. As the needle moves, the hexahedral cell layer closest to the needle boundary gets stretched or compressed, and when the cell layer size goes outside a user-specified bound, a new cell layer is introduced or deleted adjacent to the needle. It must be noted that the mesh motion is limited to the cells at the boundary, and no motion equation is solved for the interior points in the mesh. This improves the speed of the computations by avoiding the solution of an additional equation. Secondly, the mesh motion can distort the interior mesh and increase skewness for large displacements. In the current method, a good mesh resolution can be maintained at minimum lift as well as at full lift conditions since the cell sizes will always be within the user-specified bound. However, the addition and deletion of layers leads to small localized disturbances in the flow solution which get smoothed out during consecutive time steps, and do not affect the overall solution. The transient simulations were run in parallel on 4 processors. Transient needle simulation results: Two cases were considered as shown in Table 1. The Fig. 4 shows the comparison of the mass flow rate predicted by the model with the injection rate data presented in the work by Payri et al [15]. The point to be noted from the figures is that the simulated needle lift curve flattens at around 10 % of the full lift since this is equal to the minimum lift of 25 µm assumed in this study. It is seen that for the higher energizing time case a, the mass flow rate matches the behavior of the injection rate curve for the part of the profile corresponding to the needle opening. However, it is seen that as the needle closes, the mass flow rate predicted by the model reduces faster than the experimental data which seems to have a small delay. For the smaller energizing time case b, the predictions are very different as compared to the experimental mass flow rates. This can be attributed to the fact that the minimum lift of 25 µm is almost 50 % of the peak lift of the needle for this case which is around 70 µm. Also, past analysis of the (a) Two part mesh for transient simulations (b) Close-up of sac region after assembly Figure 3. Mesh setup for transient needle cases diesel injection system by Seykens et al [20] showed that the needle lift sensor measures the movement of the needle control plunger which may not correspond to the actual movement of the needle tip due to high elastic deformation prior to injection. They reported that the initial deformation of the control plunger and needle can be as high as 36% of the total measured displacement for injection pressures of 1400bar. Additionally, it was observed that the rail pressure also deforms the control plunger and needle from the bottom during injection further affecting the measurement. In the current study, we assume that the data from the needle lift sensor directly correlates to the needle tip movement. For future studies, the method presented by Powell et al [21] which uses X-ray phase-enhanced imaging to resolve the geometry of the needle tip as it moves during an injection event would be more valid for such transient simulations. In general, it is observed that the flow is highly 4

5 (a) needle just opening, time = 0.1ms (a) Case a (b) needle at full lift, time = 0.9ms (c) needle just closed, time = 1.9ms (b) Case b Figure 5. Streamlines colored by velocity at three different times during simulation for case a Figure 4. Comparison of mass flow rate predictions with injection rate data. Dashed lines - Payri et al [15], Solid lines - Simulations M assdispersion = s transient when the needle is at low lifts. This is mainly due to the presence of a swirling flow under these conditions as can be observed from the streamlines for case a shown in Fig. 5. The swirling flow only exists at low lifts when the flow in the nozzles is affected by the fuel in the sac volume. This behavior has also been observed in past work on transient simulations [8, 9]. In case b, due to the low energizing time and low maximum needle lift, the swirling motion exists almost over the entirety of the simulation. As was mentioned earlier, one of the main advantages of the current work is the ability to simulate a full 3D injector. As a result of this, the effect of the swirling flow on the mass flow rate through each nozzle hole can be analyzed. For this, the percent dispersion of mass flow rate was calculated as follows σ= σ 100 M Σ(Mi M )2 N (1) (2) In the above equations, Mi, M and N represent the mass flow rate per hole, the average mass flow rate through all holes at a given time, and the number of nozzles respectively. The Fig. 6 shows the mass dispersion over the entire simulation for the cases a and b. It is clearly observed that the dispersion is very high when the needle is at low lifts, whereas the dispersion reduces to negligible values as the needle approaches the full lift condition for case a. However, case b shows that since the maximum needle lift achieved in this case is small, high dispersion exists over almost the entire simulation. This indicates that in case of pilot injections in engines, large 5

6 (a) Case a Figure 7. Nozzles showing asymmetric vapor formation in case a when needle is just closed, time = 1.9ms (b) Case b Figure 6. Mass dispersion between holes for the transient cases amount of flow asymmetry can be expected consequently leading to spray asymmetry. It is also observed that the cavitation behavior is very transient and asymmetric under low lift conditions as seen from Fig. 7 which shows the void fraction contour at an instant when the needle is just closed in case a. Since the nozzles are tapered in this case, the amount of cavitation is very limited. On the other hand, it is expected that cylindrical nozzles have a higher amount of cavitation and will potentially show higher transient and asymmetric hole-to-hole behavior. Finally, it is important to note that some amount of asymmetry will exist in the nozzle CAD model and mesh which will potentially drive some asymmetry in the nozzle flow. Additionally, an iterative method is used for obtaining the pressure solution which can cause overshoots and undershoots. Hence, experimental verification of the reported phenomena are required. Conclusions The homogeneous relaxation model was used for simulating the flow through a 5 hole microsac injector, and the results were compared against the experimental data of Payri et al [15]. Steady state simulations were performed for cases without the needle and with the needle fixed at full lift to verify that the model reproduces the experimental data under these conditions. In both cases, the model was able to predict the trends of the experimental results accurately. Two transient cases were simulated and the mass flow rate predictions of the model were compared against injection rate data. In general, it was found that the model captured the trend of the data for an energizing time of 1ms, however, the smaller energizing time of 0.5ms showed large discrepancies. The reason for this can be attributed to the fact that the peak needle lift achieved in this case is just about twice the initial minimum lift assumed in the current study. Future studies will try to include the effect of the minimum needle lift on the predictions. The hole to hole dispersion in mass flow rates was plotted and it was found that the highest values of dispersion existed at low needle lift and could be as high as 10 % even though the injector was symmetric. However, it must be noted that the nozzle CAD model and grid will have some inherent asymmetry which can also drive some hole-hole variations and experimental verification of this phenomena is required. Acknowledgements The authors acknowledge the financial support of General Motors for this work. References [1] F.J. Salvador, S. Hoyas, R. Novella, and J. Martínez-López. Proceedings of the Institu- 6

7 tion of Mechanical Engineers, Part D: Journal of Automobile Engineering, 225: , [2] E. Giannadakis, D. Papoulias, M. Gavaises, C. Arcoumanis, C. Soteriou, and W. Tang. SAE Paper , [3] S Som, S.K. Aggarwal, E.M. El-Hannouny, and D.E. Longman. Journal of Engineering for Gas Turbines and Power, 132(4), [4] F. Peng Kärrholm, Henry Weller, and Niklas Nordin. ASME Conference Proceedings, 2007(42894): , [5] W. Yuan and G.H. Schnerr. Proc. CAV2001: Fourth Intl Symp. Cavitation, Pasadena, CA, [18] J. Weber, N. Peters, R. Diwakar, R.M. Siewert, and A. Lippert. SAE Paper , [19] M. Beaudoin and H. Jasak. Open Source CFD International Conference, Berlin, Germany, [20] X.L.J. Seykens, Somers L.M.T., and R.S.G. Baert. MECCA, III:30 39, [21] C.F. Powell, A.L. Kastengren, Z. Liu, and K. Fezzaa. Journal of Engineering for Gas Turbines and Power, 133, [6] A. Alajbegovic, G. Meister, D. Greif, and B. Basara. Experimental Thermal and Fluid Science, 26(6-7): , [7] X. Margot, S. Hoyas, P. Fajardo, and S. Patouna. Mathematical and Computer Modelling, 52: , [8] F. Payri, X. Margot, S. Patouna, F. Ravet, and M. Funk. SAE Paper , [9] W.G. Lee and R.D. Reitz. Journal of Engineering for Gas Turbines and Power, 132, [10] B. Shields, K. Neroorkar, and D.P. Schmidt. 23rd Annual Conference on Liquid Atomization and Spray Systems, Ventura, CA, May 2011, [11] K. Neroorkar, B. Shields, R.O. Grover jr, A.H. Plazas, and D. P. Schmidt. SAE Paper , [12] S. Gopalakrishnan and D.P. Schmidt. SAE Paper , [13] D.P. Schmidt, S. Gopalakrishnan, and H. Jasak. Intl. J. of Multiphase Flow, 36: , [14] K. Neroorkar, S. Gopalakrishnan, D. Schmidt, and R. O. Grover Jr. Atomization and Sprays, 21(2): , [15] R. Payri, A. Gil, A.H. Plazas, and B Giménez. SAE Paper , [16] J. Benajes, J.V. Pastor, R. Payri, and A.H. Plazas. Journal of Fluids Engineering, 126, [17] W. Bosch. SAE Tech. Paper ,