Temperature profile experimental

Nonisothermal Gas Absorption. The computation of nonisothermal gas absorption processes is difficult because of all the interactions involved as described for packed columns. A computer is normally required for the enormous number of plate calculations necessary to estabUsh the correct concentration and temperature profiles through the tower. Suitable algorithms have been developed (46,105) and nonisothermal gas absorption in plate columns has been studied experimentally and the measured profiles compared to the calculated results (47,106). Figure 27 shows a typical Hquid temperature profile observed in an adiabatic bubble plate absorber (107). The close agreement between the calculated and observed profiles was obtained without adjusting parameters. The plate efficiencies required for the calculations were measured independendy on a single exact copy of the bubble cap plates installed in the five-tray absorber. 

Fig. 27. Computed and experimental Hquid temperature profiles in an ammonia absorber with 5 bubble cap trays (107). Water was used as a solvent.

One potential problem with this approach is that heat loss from a small scale column is much greater than from a larger diameter column. As a result, small columns tend to operate almost isotherm ally whereas in a large column the system is almost adiabatic. Since the temperature profile in general affects the concentration profile, the LUB may be underestimated unless great care is taken to ensure adiabatic operation of the experimental column. 

When used for superresolution, the laser beam is incident on b, which hides the domains in s. During read-out, b is heated and the domains in s are copied to b. The optical system sees only the overlap area between the laser spot and the temperature profile which is lagging behind, so that the effective resolution is increased. Experimentally it is possible to double the linear read-out resolution, so that a four times higher area density of the domains can be achieved when the higher resolution is also exploited across the tracks. At a domain distance of 0.6 pm, corresponding to twice the optical cutoff frequency, a SNR of 42 dB has been reached (82). 

Schematic DRD shown in Fig. 13-59 are particularly useful in determining the imphcations of possibly unknown ternary saddle azeotropes by postulating position 7 at interior positions in the temperature profile. It should also be noted that some combinations of binary azeotropes require the existence of a ternaiy saddle azeotrope. As an example, consider the system acetone (56.4°C), chloroform (61.2°C), and methanol (64.7°C). Methanol forms minimum-boiling azeotropes with both acetone (54.6°C) and chloroform (53.5°C), and acetone-chloroform forms a maximum-boiling azeotrope (64.5°C). Experimentally there are no data for maximum or minimum-boiling ternaiy azeotropes. The temperature profile for this system is 461325, which from Table 13-16 is consistent with DRD 040 and DRD 042. However, Table 13-16 also indicates that the pure component and binary azeotrope data are consistent with three temperature profiles involving a ternaiy saddle azeotrope, namely 4671325, 4617325, and 4613725. All three of these temperature profiles correspond to DRD 107. Experimental residue cui ve trajectories for the acetone-...

Another instance in which the constant-temperature method is used involves the direc t application of experimental KcO values obtained at the desired conditions of inlet temperatures, operating pressure, flow rates, and feed-stream compositions. The assumption here is that, regardless of any temperature profiles that may exist within the actu tower, the procedure of working the problem in reverse will yield a correct result. One should be cautious about extrapolating such data veiy far from the original basis and be carebil to use compatible equilibrium data. 

From Tolmin s theory and experimental data (e.g., Reichardtthe relationship between velocity profile and temperature profile in the jet cross-section can be expressed using an overall turbulent Prandtl number Pr = v /a, where Vf is a turbulent momentum exchange coefficient and a, is a turbulent heat exchange coefficient ... 

Near room temperature, Ea is roughly 2.5 kJ mol-1 (or 0.6 kcal mol-1) larger than AW. These treatments assume that both AW and Ea are temperature-independent. That is, the temperature profiles according to Eqs. (7-3) and (7-5) are both linear. Most data (but not all, see Sections 7.2 and 7.3) conform to that model, although Eq. (7-8) says that, literally, AW and Ea cannot both be temperature-independent constants. In fact, the RT term in Eq. (7-8) is usually much smaller than the others. Thus, the temperature independence of both Ea and AW can in practice be sustained with good accuracy. The choice of the T to use in Eq. (7-8) is not crucial—a midpoint value of the experimental range does the job. In the same vein, the value of AS and A are related by (see Problem 7.4)... 

An implicit assumption of the foregoing treatment is that A// remains independent of temperature over the range investigated. This is very nearly correct in general, and is particularly the case given that studies of reactions in solution are usually conducted over a temperature interval of only some 30-50°. In certain circumstances the temperature profiles show curvature outside the experimental error. Such cases have, or appear to have, temperature-dependent activation enthalpies. Here we explore one of the reasons for that another is given in Section 7.3. 

Continuous Polymerizations As previously mentioned, fifteen continuous polymerizations in the tubular reactor were performed at different flow rates (i.e. (Nj g) ) with twelve runs using identical formulations and three runs having different emulsifier and initiator concentrations. A summary of the experimental runs is presented in Table IV and the styrene conversion vs reaction time data are presented graphically in Figures 7 to 9. It is important to note that the measurements of pressure and temperature profiles, flow rate and the latex properties indicated that steady state operation was reached after a period corresponding to twice the residence time in the tubular reactor. This agrees with Ghosh s results ). 

Figure 7. Theoretical and experimental temperature profiles for steel mold...

Figure 11.7 Experimental temperature profile for an early SMR reactor with an unoptimized design.

Boundary layer similarity solution treatments have been used extensively to develop analytical models for CVD processes (2fl.). These have been useful In correlating experimental observations (e.g. fi.). However, because of the oversimplified fiow description they cannot be used to extrapolate to new process conditions or for reactor design. Moreover, they cannot predict transverse variations In film thickness which may occur even In the absence of secondary fiows because of the presence of side walls. Two-dimensional fully parabolized transport equations have been used to predict velocity, concentration and temperature profiles along the length of horizontal reactors for SI CVD (17,30- 32). Although these models are detailed, they can neither capture the effect of buoyancy driven secondary fiows or transverse thickness variations caused by the side walls. Thus, large scale simulation of 3D models are needed to obtain a realistic picture of horizontal reactor performance. 

A simplified version of the model in Table IX, neglecting accumulation of mass and heat as well as dispersion and conduction in the gas phase, predicts dynamic performance of a laboratory S02 converter operating under periodic reversal of flow direction quite well. This is shown by Fig. 13 taken from Wu et al. (1996). Data show the temperature profiles in a 2-m bed of the Chinese S101 catalyst once a stationary cycling state is attained. One set of curves shows the temperature distribution just after switching direction and the second shows the distribution after a further 60 min. Simulated and experimental profiles are close. The surprising result is that the experimental maximum temperatures equal or exceed the simu-... 

Fig. 13. Comparison of simulated and experimental temperature profiles in a 2-m, near-adiabatic, packed-bed S02 reactor using a Chinese S101 catalyst and operating under periodic reversal of flow direction with r = 180 min, SV = 477 h"1, and inlet S02 = 3.89 vol% and T = 25°C. (Figure adapted from Wu et at., 1996, with permission of the authors.)...

Barriers to heat transfer produce corresponding temperature differences in a freeze-drying system, the actual temperature profile depending upon the rate of sublimation, the chamber pressure, and the container system as well as the characteristics of the freeze dryer employed. An experimental temperature profile is shown in Figure 5 for a system where vials were placed in an aluminum tray with a flat 5 mm thick bottom and a tray lid containing open channels for escape of water vapor. Here, heat transfer is determined by four barriers ... 

Under common experimental conditions the convective term does not contribute significantly to the temperature profile, and thus Eq. (27) can be simplified to... 

We now repeat the derivation of the steady-state heat transport limited moisture uptake model for the system described by VanCampen et al. [17], The experimental geometry is shown in Figure 9, and the coordinate system of choice is spherical. It will be assumed that only conduction and radiation contribute significantly to heat transport (convective heat transport is negligible), and since radiative flux is assumed to be independent of position, the steady-state solution for the temperature profile is derived as if it were a pure conductive heat transport problem. We have already solved this problem in Section m.B, and the derivation is summarized below. At steady state we have already shown (in spherical coordinates) that... 

Ehase Inversion Temperatures It was possible to determine the Phase Inversion Temperature (PIT) for the system under study by reference to the conductivity/temperature profile obtained (Figure 2). Rapid declines were indicative of phase preference changes and mid-points were conveniently identified as the inversion point. The alkane series tended to yield PIT values within several degrees of each other but the estimation of the PIT for toluene occasionally proved difficult. Mole fraction mixing rules were employed to assist in the prediction of such PIT values. Toluene/decane blends were evaluated routinely for convenience, as shown in Figure 3. The construction of PIT/EACN profiles has yielded linear relationships, as did the mole fraction oil blends (Figures 4 and 5). The compilation and assessment of all experimental data enabled the significant parameters, attributable to such surfactant formulations, to be tabulated as in Table II. 

The agreement between the experimental and calculated values of Ce, is excellent. The data shown in Figure 2 are for a constant bake time of 17 minutes. The upper and lower limits on define a cure window. The cure window for the low solids coating is 50 C. The model was further tested by measuring extents of reaction and temperature profiles for samples attached to different parts of a car body which passed through a pilot plant oven. This simulation tested the model under conditions where the substrate temperatures were far from constant. As shown in Table II, the agreement between the experimental and calculated values of Ce is again excellent. 

No rate enhancement was observed when the reaction was performed under microwave irradiation at the same temperature as in conventional heating [47]. Similar reaction kinetics were found in both experiments, presumably because mass and heat effects were eliminated by intense stirring [47]. The model developed enabled accurate description of microwave heating in the continuous-flow reactor equipped with specific regulation of microwave power [47, 48]. Calculated conversions and yields of sucrose based on predicted temperature profiles agreed with experimental data. 

Fig. 15 Thennogravimetric weight loss-temperature profiles for the two dihydrate polymorphs of the experimental compound L-706000-001T. Upper plot type D lower plot type A.

From this illustration we can see that the added detail of the radial temperature profile near the wall that could be provided by CFD simulations does not help in obtaining better estimates for the standard heat transfer parameters. It also implies that experimental efforts to measure temperatures closer to the wall are, in fact, counter-productive. Finally, it is clear that the standard model with plug flow and constant effective transport parameters does not fit satisfactorily to temperature profiles in low-Abeds. These considerations have led us to look for improved approaches to near-wall heat transfer. 

Assay of Homogenate for Aldrin Epoxidation. The following experimental sequence was designed to determine the optimum in vitro conditions for aldrin epoxidation in larval whole body homogenates 1) the effect of component chemicals generally included in an incubation mixture, 2) a pH profile, 3) a temperature profile, 4) a molarity profile, 5) a reaction time profile, 6) a larval concentration (enzyme concentration) profile, 7) a substrate concentration profile, and 8) a restudy of the effects of component chemicals in the initial incubation mixture (Step 1) upon aldrin epoxidation under optimum conditions as defined by steps 2-7 above. The effect of PBO, FMN, and FAD upon enzyme activity was also tested. 

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