Hydraulic performance, calculations

Overall coefficients, 332 Vertical plate coil, 331 Hindered settling velocities, 231, 236 Horsepower, centrifugal pump driver, 201 Hydraulic performance, calculations, 180-188... 

Brake horsepower, centrifugal pumps, 200 Driver horsepower, 201 Burst pressure, 405, 456 Cartridge filters, 274-278 Capture mechanism, 279 Edge filler, 278 Filter media, table, 278 Micron ratings, 277 Reusable elements, 281 Sintered metal, 280 Types, 276, 277, 279 Wound vs. pleated, 276, 277 Centrifugal pumps, operating characteristics, 177-180 Calculations, see hydraulic performance Capacity, 180... 

Tlie core-internal thermal-hydraulic performance of fuel temperature, core-internal structure, and core-internal coolant distribution were confirmed to be appropriate to their design during the full power operation. Tlie maximum temperature of the core support-plate measured at the upper surface of the centre core support-plate was 450°C that was sufficiently below its limited value of 530°C. Also, it was confirmed that other core-internal structure temperatures were well below their design criteria. From the result tliat no core-internal structure temperature measurement showed an abnormal value, it was confinned tliat there was no abnormal leak flow of coolant such as cross and bypass flows between fuel blocks, replaceable reflector blocks, permanent reflector blocks, etc. The maximum fuel temperature was evaluated to be 1 463°C prior to the high-temperature test operation. It was re-evaluated using the measured temperature data i.e. core-inlet and -outlet coolant temperatures and the calculated value of 1478°C does not exceed the normal operation hmit of 1 495°C. 

What is the hydraulic performance of the tower under new conditions The operating pressure drops can be calculated based on hydraulic calculations discussed in Chapter 12. 

Construction of a full size non-nuclear hydraulic module to verily hydraulic performance of the core and primary circuit and to determine basic thermal-hydraulic correlations Performance of neutron-physical, thermal hydraulic and structural calculations Adjustment of codes and performance of safety analysis ... 

In addition to these theoretical questions experimental investigations will be needed. The performance of mixed-oxide fuel with plutonium contents up to 45% under irradiation to high bumup will have to be determined. Reactor physics calculation methods suitable for burner cores will have to be validated by comparison with data from critical assemblies. There may also be a need for experimental work on the thermal hydraulic performance of heterogeneous cores containing both fuel and diluent material. 

It also became necessary to consider interactions between the nuclear and the thermal-hydraulic performance. Iterative calculations, notably between the AIMAZ (nuclear) and CUSH (core and primary circuit hydraulic models) were needed to reach final estimates of overall reactor power distributions and void coefficients. 

This flow parameter is the square root of the ratio of liquid kinetic energy to gas kinetic energy. The ordinate of this correlation includes the gas flow rate, the gas and liquid densities, the a/e ratio (which is characteristic of the particular tower packing shape and size), and a liquid viscosity term. Lobo et al proposed the use of a packing factor to characterize a particular packing shape and size [17]. They determined that the a/e ratio did not adequately predict packing hydraulic performance. Eckert further modified this correlation and calculated the packing factors from experimentally determined pressure drops [18]. 

Since the core-wide nuclear performance parameters and local power distribution within a fuel assembly depend highly on the moderator condition in LWRs, thermal-hydraulic coupled calculations have been regarded as essential in the core design procedure for thermal neutron spectrum cores. As the thermal-hydraulic... 

The pressure drops, described in the sections above, were used to calculate the percentage AP/P lor the piping system, defined in Equation 9-24. This percentage AP/P was calculated for various piping arrangements in order to compare hydraulic performance. 

The gas compression in practically all commercial machines is polytropic. That is, it is not adiabatic or isothermal, but some form peculiar to the gas properties and the hydraulic design of the compressor. Actual machines may be rated on adiabatic performance and then related to polytropic conditions by the polytropic efficiency. Other performance rating procedures handle the calculations as polytropic. For reference, both methods are presented. 

One of the possible ways to account for the effect of roughness on the pressure drop in a micro-tube is to apply a modified-viscosity model to calculate the velocity distribution. Qu et al. (2000) performed an experimental study of the pressure drop in trapezoidal silicon micro-channels with the relative roughness and hydraulic diameter ranging from 3.5 to 5.7% and 51 to 169 pm, respectively. These experiments showed significant difference between experimental and theoretical pressure gradient. 

Warrier et al. (2002) conducted experiments of forced convection in small rectangular channels using FC-84 as the test fluid. The test section consisted of five parallel channels with hydraulic diameter = 0.75 mm and length-to-diameter ratio Lh/r/h = 433.5 (Fig. 4.5d and Table 4.4). The experiments were performed with uniform heat fluxes applied to the top and bottom surfaces. The wall heat flux was calculated using the total surface area of the flow channels. Variation of single-phase Nusselt number with dimensionless axial distance is shown in Fig. 4.6b. The numerical results presented by Kays and Crawford (1993) are also shown in Fig. 4.6b. The measured values agree quite well with the numerical results. 

U.S. EPA bases its 1 gallon/acre/day leak detection sensitivity on the results of calculations that show that, theoretically, an LDS overlying a composite bottom liner with an intact FML component can detect, collect, and remove liquids from a top liner leak rate <1 gallon/acre/day. This performance standard, therefore, can be met with designs that include a composite bottom liner. Based on numerical studies, one cannot meet the leak detection sensitivity with a compacted soil bottom liner, even one with a hydraulic conductivity of 10-7 cm/s. Therefore, the emphasis of this standard is on selecting an appropriate bottom liner system. 

The sizing of piping is based upon a hydraulic analysis for the water distribution network for the WCCE. The main delivery pipe should be sized to provide 150% of the design flow rate. A residual pressure and flow requirement at the most remote hydrocarbon process or storage location from the supply source dictates the sizing for the remaining system. Normal reliability requirements usually suggest that minimum of two sources of supply be available that are in themselves remote from each other. Therefore two remote flow calculations must be performed to determine the minimum pipe distribution size. NFPA 24 requires that the minimum residual pressure available in a fire main not be less than 6.9 bars (100 psi.). Velocity calculations should be performed which verify flows are not more than the limits of the material that is employed. 

PEM resistance in operational PEFC as a function of the fuel cell current density, comparing experimental data (dots) and calculated results from a performance model based on the hydraulic permeation model for various applied gas pressure differences between anode and cathode compartments. (Reprinted from S. Renganathan et al. Journal of Power Sources 160 (2006) 386-397. Copyright 2006, with permission from Elsevier.)... 

The calculations were performed for a Reynolds number based on the channel hydraulic diameter (Ri ) of 1,800. This high Reynolds number was chosen since a more accurate determination of ( can be obtained at these flow conditions (either experimental or computational) keeping in mind that inertial effects are more pronounced at high Reh. The permeabilities of the filter wall... 

Compaction properties of each material were determined with a standardized test performed on a custom-built hydraulic compaction simulator using 8 mm (0.3150 in.) round flat-faced punches. A linear saw-tooth upper punch position profile was selected with a punch velocity of 300 mm/sec for both punch extension and retraction. The lower punch position was at a fixed position within the die during the compaction event. The powder weight loaded into the die for each compression was calculated from the equation below so as to form a cylindrical tablet having a thickness-to-diameter ratio of 0.30 at a theoretical SF of 1.0. These dimensions are typical of commercially elegant tablets. 

Example 10.2 compares data of Table 10.4 with calculations based on Figures 10.6 and 10.7 for all-liquid mixing. Power and rpm requirements at a given superficial liquid velocity are seen to be very sensitive to impeller diameter. When alternate combinations of HP/rpm are shown in the table for a particular performance, the design of the agitator shaft may be a discriminant between them. The shaft must allow for the torque and bending moment caused by the hydraulic forces acting on the impeller and shaft. Also, the... 

Hydraulic calculations of water supply systems are critical to confirming their adequacy. Often such calculations are not current and, hence, flows and pressures are not readily available to be used as part of evaluations of system performance in cases where different events necessitate their use. 

Related Calculations. Use the method given here for any type of pump whose variables are included in the Hydraulic Institute curves (Figs. 6.10 and 6.11) and in similar curves available from the same source. Operating specific speed, computed as above, is sometimes plotted on the performance curve of a centrifugal pump so that the characteristics of the unit can be better understood. Type specific speed is the operating specific speed giving maximum efficiency for a given pump and is a number used... 

Hydraulic analysis of the Aspen Plus simulator produces thermodynamic ideal minimum flow and actual flow curves for rigorous distillation column simulations. These types of calculations are performed for RADFRAC columns. Using the input summary given in problem 4.48 construct the stage-flow curves. Assess the thermodynamic performance of the column. 

Membrane performance characteristics in the hydraulic and diffusion limits are compared to each other in Fig. 9. Figure 9(a) illustrates that in the diffusion model considerable deviations from the purely ohmic performance of the saturated membrane arise already at small jv/Jj, well below the critical current density. This is in line with the comparison of the water-content profiles calculated in the diffusion model, Fig. 9(b), with those from the hydraulic permeation model, in Fig. 7. Indeed, membrane dehydration is much stronger in the diffusion model, affecting larger membrane domains at given values of jp/./j. Moreover, the profiles exhibit different curvature from those in Fig. 7. 

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