Experimental evaluation, liquid phase

Carey van P (1992) Liquid-vapor phase-change phenomena. An introduction to the thermophysics of vaporization and condensation processes in heat transfer equipment. Hemisphere, New York Celata GP, Cumo M, Mariani A (1997) Experimental evaluation of the onset of subcooled flow boiling at high liquid velocity and subcoohng. Int J Heat Mass Transfer 40 2979-2885 Celata GP, Cumo M, Mariani A (1993) Burnout in highly subcooled water flow boiling in small diameter tubes. Int J Heat Mass Transfer 36 1269-1285 Chen JC (1966) Correlation for boiling heat transfer to saturated fluids in convective flow. Ind Eng Chem Process Des Develop 5 322-329... 

Nitrogen - Water System. The interaction parameters for the nitrogen - water system have been evaluated using the data of Wiebe and Gaddy (10), Paratella and Sagramora (Vj ), Rigby and Prausnitz (12)and O Sullivan and Smith (13). As with the two previous systems, only one constant interaction parameter was necessary to correlate the vapor phase composition while the interaction parameter for the aqueous liquid phase increased monotonically with temperature. A comparison of the calculated and experimental vapor phase and aqueous liquid phase compositions is given in Table I. 

Plug-flow assumption Using die Michell and Furzer correlation (eq. (3.417)), the liquid-phase Peclet value is 0.66 and die minimum value of Z/d, evaluated by using the Mears criterion (eq. (3.421)) 4.81, which is much lower that the experimental one, which is 100. Then, condition (c) is satisfied. 

To determine Sb in marine sediments by ETAAS, a direct method was developed based on quantitating the analyte in the liquid phase of the slurries (prepared directly in autosampler cups). The variables influencing the extraction of Sb into the liquid phase and the experimental setup were set after a literature search and a subsequent multivariate optimisation procedure. After the optimisation, a study was carried out to assess robustness. Six variables were considered at three levels each (see Table 2.13). In addition, two noise factors were set after observing that two ions, which are currently present into marine sediments, might interfere in the quantitations. In order to evaluate robustness, a certified reference material was used throughout, BCR-CRM 277 Estuarine Sediment (guide value for Sb 3.5 0.4pgg ). Table 2.13 depicts the experimental setup. 

The fractional cesium release during any given time period may be calculated, provided that all of the constants in Equation 8 are known. Unfortunately, however, liquid-phase diffusivities can be estimated only to within an order of magnitude. An alternative approach, therefore, is to evaluate the constants D and a from two sets of experimental release data and to use these values to predict the fractional release for other conditions. 

With these parameters we can set-up a mass balance on the system, which is the basis for evaluating the experimental results. The mass balance for the absorption (of any gas, e. g. ozone) in a continuous-flow stirred tank reactor (CFSTR) under the assumption that the gas and liquid phases are ideally mixed (cL = cLe, cG = cGe), are as follows ... 

The data can be evaluated using any commonly available non-linear regression program or with a linear regression, in which k,a is the slope from the plot of the natural log of the concentration difference versus time. Linearity of the logarithmic values over one decade is required for the validity of the measurement. Of course the assumptions inherent in the model must apply to the experimental system, especially in respect to completely mixed gas as well as liquid phases and reactions are negligible. Two common problems are discussed below. Other common pitfalls and problems are summarized in Table 3-3. 

From a Solution Model. Calculation of the difference in reduced standard-state chemical potentials by methods I or III in the absence of experimental thermodynamic properties for the liquid phase necessitates the imposition of a solution model to represent the activity coefficients of the stoichiometric liquid. Method I is equivalent to the equation of Vieland (106) and has been used almost exclusively in the literature. The principal difference between methods I and III is in the evaluation of the activity coefficients... 

A general formulation of the problem of solid-liquid phase equilibrium in quaternary systems was presented and required the evaluation of two thermodynamic quantities, By and Ty. Four methods for calculating Gy from experimental data were suggested. With these methods, reliable values of Gy for most compound semiconductors could be determined. The term Ty involves the deviation of the liquid solution from ideal behavior relative to that in the solid. This term is less important than the individual activity coefficients because of a partial cancellation of the composition and temperature dependence of the individual activity coefficients. The thermodynamic data base available for liquid mixtures is far more extensive than that for solid solutions. Future work aimed at measurement of solid-mixture properties would be helpful in identifying miscibility limits and their relation to LPE as a problem of constrained equilibrium. 

The objective of this portion of the research was to experimentally evaluate surfactant effects on the liquid-liquid separation of hydrophobic oils from a surfactant system. For pump-and-treat subsurface remediation in the absence of surfactant, contaminated ground water would be pumped from the subsurface and through a liquid-liquid extraction column where the contaminant partitions from the aqueous phase into an extraction solvent phase. In the absence of surfactant, the driving force for partitioning is a function of the contaminant hydrophobicity. In the presence of surfactants, the contaminant is subject to competitive partitioning (i.e., into the micelles and into the extracting oil). 

Knowing an experimental value of k, it is possible to evaluate the diffusion coefficient of the atoms of a dissolving solid substance across the diffusion boundary layer at the solid-liquid interface into the bulk of the liquid phase using equations (5.6) and (5.7). Its calculation includes two steps. First, an approximate value of D is calculated from equation (5.6). Then, the Schmidt number, Sc, and the correction factor, /, is found (see Table 5.1). The final, precise value is evaluated from equation (5.7). In most cases, the results of these calculations do not differ by more than 10 %. Values of the diffusion coefficient of some transition metals in liquid aluminium are presented in Table 5.9.303... 

The model optimized with regard to numerical and physico-chemical parameters has been tested with experimental data from a pilot plant, and used to evaluate industrial operation data. Here, a good agreement between experimental and simulated values is established, both for the gas-phase concentration of CO2 (Fig. 9.19) and for the temperature in the liquid phase (Fig. 9.20). 

Vapour phase enthalpies were calculated using ideal gas heat capacity values and the liquid phase enthalpies were calculated by subtracting heat of vaporisation from the vapour enthalpies. The input data required to evaluate these thermodynamic properties were taken from Reid et al. (1977). Initialisation of the plate and condenser compositions (differential variables) was done using the fresh feed composition according to the policy described in section 4.1.1.(a). The simulation results are presented in Table 4.8. It shows that the product composition obtained by both ideal and nonideal phase equilibrium models are very close those obtained experimentally. However, the computation times for the two cases are considerably different. As can be seen from Table 4.8 about 67% time saving (compared to nonideal case) is possible when ideal equilibrium is used. 

We studied these phenomena experimentally in a wetted wall column and two stirred cell reactors and evaluated the results with both a penetration and a film model description of simultaneous mass transfer accompanied by complex liquid-phase reactions [5,6], The experimental results agree well with the calculations and the existence of the third regime with its desorption against overall driving force is demonstrated in practice (forced desorption or negative enhancement factor). 

The mass transfer model. In our previous work [6] the mass transfer model equations and their mathematical treatment have been described extensively. The relevant differential equations, describing the process of liquid-phase diffusion and simultaneous reactions of the species according to the penetration theory, are summarized in table 1. Recently we derived from this penetration theory description a film model version, which is incorporated in the evaluation of the experimental results. Details on the film model version are given elsewhere [5]. 

Three different approaches have been presented for estimating the partitioning of solutes between plastics and liquids. In the context of evaluating the output from these different approaches it is also useful to define the expected experimental ranges and limits for partition coefficients based on the solutes, plastic and contacting liquid phases involved. Table 4-9 shows approximate upper and lower limits for partition coefficients one may normally encounter in plastic/food systems based on the polarities of the solutes, plastics and foods. The table also gives approximate ranges of partition coefficient values for various solutes between typical food contact plastics and liquid phases. 

With chromatographic systems involving polypeptides and protein purification, the physical properties of the system and operating conditions are often predetermined. For polypeptides and proteins of known molecular weight, the liquid-phase mass transfer coefficients can be evaluated as discussed earlier. However, the forward surface interaction rate constant (k j) must be determined from experimental data via a parameter-fitting program... 

Since dissolved gas concentrations in the liquid phase are more difficult to measure experimentally than the liquid reactant concentration, Equation 8 evaluated at the reactor exit 5=1 represents the key equation for practical applications involving this model. Nevertheless, the resulting expression still contains a significant number of parameters, most of which cannot be calculated from first principles. This gives the model a complex form and makes it difficult to use with certainty for predictive purposes. Reaction rate parameters can be determined in a slurry and basket-type reactor with completely wetted catalyst particles of the same kind that are used in trickle flow operation so that DaQ, r 9 and A2 can be calculated for trickle-bed operation. This leaves four parameters (riCE> St gj, Biw, Bid) to be determined from the available contacting efficiency and mass transfer correlations. It was shown that the model in this form does not have good predictive ability (29,34). 

As discussed in Chap. 3, there are a large number of models proposed to evaluate macromixing in a trickle-bed reactor. A brief summary of the reported experimental studies on the measurements of RTD in a cocurrent-downflow trickle-bed reactor is given in Table 6-7. Some of these experimental studies are described in more detail in a review by Ostergaard.94 Here we briefly review some of the correlations for the axial dispersion in gas and liquid phases based on these experimental studies. 

The adopted heat capacity values for Rb(t) are those recommended by Fink and Lelbowitz ( ). These authors reviewed the liquid phase heat capacity and enthalpy studies for all alkali metals and recommended a Cp" equation for Rb(t) based on their evaluation of four enthalpy (2, 3, 4, 5) and three heat capacity (6, 7, 8) experimental studies. These studies provided information in the liquid region up to 1334 K. 

All processes are modeled as series of countercurrent equilibrium cells. Parameters were determined experimentally (section 3). A liquid-phase reaction is accounted for by Da = (rate constant)x(cell volume)/(solid flow rate). Adsorption is described by the bi-Langmuir model. All equations were implemented in the simulation environment DIVA [3] details on the implementation of a largely analogous model can be found in [1,4]. The following set of performance parameters were used to evaluate each process ... 

The surface layer at a vapor-liquid interface is in tension and will contract to minimize the surface area. Qualitatively, the surface tension is due to the larger attractive forces that molecules at the interface experience from molecules in the dense liquid phase than from those in the low-density gas phase. Quantitatively,. surface tension is defined as the force in the surface plane per unit length. Jasper [Jasper, J. J., J. Phys. Chem. Ref. Data, 1 (1972) 841] has made a critical evaluation of experimental surface tension data for approximately 2200 pure chem-icms and correlated surface tension a (mN/m = dyn/cm) with temperature as... 

The standard entropy of a-SnSe was evaluated to be 86.93 J-K" -mol , corresponding to Af5° (SnSe, a, 298.15 K) = - 6.3 J-K -mor, in the thermodynamic optimisation and assessment of the Sn-Se system in [96FEU/MAJ]. The value originates mainly from the modelling and assumptions made about the liquid phase in the system and the recalculation to 298.15 K by the use of enthalpies of phase transformations and heat capacities. The only experimental determination of the entropy at low temperatures was made by Melekh, Stepanova, Fomina, and Semenkovich [71MEL/STE] who performed emf measurements on the galvanic cells... 

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