【文章內(nèi)容簡介】 aphy setup was employed. The Xray transform is then no longer dependent on the projection angle of the setup. Thus, from a single measurement at source position r S,0 and from a set of virtual source positions r S, virt on a virtual circular or bit including r S,0 the rotationally symmetric distribution can be reconstructed. Due to rotational symmetry, the FDK algorithm .was implemented in such a way that only the radial and axial gas holdup distributions in the central vertical slice were puted with considerable reduction in putational load. Finally, it is important to note that the reconstruction from a single projection at virtual source positions is equal to rotational averaging of the reconstruction result. This, however, results in a space variant noise distribution in the reconstructed data. The signal to noise ratio of voxels residing at the outer boundary of the supporting circular reconstruction grid will ultimately rise due to the greater circumference length and thus of the greater number of voxels during averaging.3. Numerical investigationsThe putational fluid dynamics analyses of the stirred tank reactor were performed with CFX numerical the nonbaffled vessel exhibits to some extent an axisymmetric behaviour on a macromixing scale, the process was regarded as threedimensional in order to demonstrate the local variations in the gas holdup associatedwith the blades positions. In such a way, cavity presence behind the impeller blades was theoretically predicted. An unstructured tetrahedral mesh, which explicitly includes the impeller geometry, with about 1,500,000 elements was used for the solution of the flow field in the stirred vessel. The mesh was globally refined since a detailed view in the whole tank was required. Additional simulations were performed on differently sized grids to establish grid independence of the results. There was no additional refinement of the mesh at wall since the wall turbulence damping function are known to work well at high Reynolds numbers as in the particular stirred tank was broken down into four domains: threerotating and one stationary. The choice of having four domains instead of two, one stationary and one rotating was made because of some difficulties when meshing the geometry in 3D. The three rotating domains prise the impeller, the part of the stirrer shaft in the impeller region and the stirrer shaft above the impeller. A multiple frames of references interaction scheme was used in the simulations to bond the stationary and the rotating domains. Wall boundary conditions with free slip for the gaseous and no slip for the liquid phases were the only ones required as the vessel was closed at the top. Four steady state simulations at stirrer speed from 200 to 800 rpm were conducted to obtain an initial guess of the flow field and the phase distribution for the simulation at 1000 rpm. Zero velocity fields for both phases have been taken as initial conditions for the simulation at stirrer speed of 200 rpm. The numerical predictions above 1000 rpm used the previous simulation results as an initial guess. Starting from 1000 rpm, five simulations have been performed at stirrer speed intervals of 50 rpm. The gas phase was modelled as dispersed fluid with a mean bubble diameter of 1 mm and the liquid phase as continuous fluid. According to Laakkonen et al. (2005), the mean bubble diameter for an air–water system in a stirred reactor agitated by a Rushton turbine at stirrer speed around 250 rpm was close to 3 mm. However, in the current experimental setup the stirrer was rotated with a speed above 1000 rpm and the holes on the impeller shaft had a diameter of 2 mm. The choice of 1 mm in size bubbles was also based on additionally performed highspeed camera measurements. For an isothermal twophase flow with no mass transfer, the mass (4)and momentum(5)governing equations can be applied to both phases. The phases in this case have separate velocities and other related to the phases fields, but share the mon pressure field. Additionally, the volume conservation equation, expressing the constraint that the volume fractions sum to unity , is relevant. In the above equations, r denotes the density, m the dynamic viscosity, p the pressure and U the velocity. A quantity subscribed with or refers to the value of the quantity of the particular phase or . S Ma denotes the momentum source due to external body forces. M a describes the interfacial forces acting on phase a due to the presence of phase , for which the drag force, was taken into account. The drag between the phases was modelled with a constant drag coefficient model, C D = as the bubble Reynolds number was large enough to ensure domination of the inertial over the viscous effects. The interfacial transfer of momentum is determined by the contact surface area between the two is characterised by the interfacial area per unit volume between phase and phase , known as the interfacial area density . The interfacial area density between the continuous phase, the isopropanol, and the dispersed phase, the air,was modelled assuming that the air (phase ) is present as spherical particles of mean diameter = 1 mm according to (6)The turbulence was considered using phasedependent turbulent models. For the gas phase the dispersed phase zero equation (AEA Technology and CFX International, 2004) was adopted. Different turbulence models, available in CFX TM software package, and their suitability have been considered for the liquid phase but the turbulence model was finally implemented because of numerical stability. Although the model is generally acknowledged as not suitable for rotating flow, the applied numerical setup scheme was able successfully to predict the free surface deformation as well as the bubble dispersion in the stirred vessel. The Sa