Cooling tower approach

Theoretical possible heat removal per pound of air circulated in a cooling tower depends on the temperature and moisture content of air. An indication of the moisture content of the air is its wet-bulb temperature. Ideally, then, the wet-bulb temperature is the lowest theoretical temperature to which the water can be cooled. Practically, the cold-water temperature approaches but does not equal the air wet-bulb temperature in a coohng tower this is so because it is impossible to contact all the water with fresh air as the water drops through the wetted fill surface to the basin. The magnitude of approach to the wet-bulb temperature is dependent on tower design. Important factors are air-to-water contact time, amount of fill surface, and breakup of water into droplets. In actual practice, cooling towers are seldom designed for approaches closer than 2.8°C (5°F).  [c.1162]

Time of contact between water and air is governed largely by the time required for the water to discharge from the nozzles and fall through the tower to the basin. The time of contact is therefore obtained in a given type of unit by vaiying the height of the tower. Should the time of contact be insufficient, no amount of increase in the ratio of air to water will produce the desired coohng. It is therefore necessary to maintain a certain minimum height of cooling tower. When a wide approach of 8 to II°C (15 to 20°F) to the wet-bulb temperature and a 13.9 to I9.4°C (25 to 35°F) cooling range are required, a relatively low cooling tower will suffice. A tower in which the water travels 4.6 to 6.1 m (15 to 20 ft) from the distributing system to the basin is sufficient. When a moderate approach and a cooling range of 13.9 to I9.4°C (25 to 35°F) are required, a tower in which the water travels 7.6 to 9.1 m (25 to 30 ft) is adequate. Where a close approach of 4.4°C (8°F) with a 13.9 to I9.4°C (25 to 35°F) cooling range is required, a tower in which the water travels from 10.7 to 12.2 m (35 to 40 ft) is required. It is usually not economical to design a coohng tower with an approach of less than 2.8°C (5°F), but it can be accomphshed satisfactorily with a tower in which the water travels 10.7 to 12.2 m (35 to 40 ft).  [c.1164]

Figure 12-14 shows the relationship of the hot-water, cold-water, and wet-bulb temperatures to the water concentration. From this, the minimum area required for a given performance of a well-designed counterflow induced-draft cooling tower can be obtained. Figure 12-15 gives the horsepower per square foot of tower area required for a given performance. These curves do not apply to parallel or cross-flow coohng, since these processes are not so efficient as the counterflow process. Also, they do not apply when the approach to the cold-water temperature is less than 2.8°C (5°F). These charts should be considered approximate and for preliminary estimates only. Since many factors not own in the graphs must be included in the  [c.1164]

Figure 30-lC is distinctly different from the first two in the type of SO2 control processes used and the sequence of the particulate matter and SOj controls. It is a promising approach for up to 90% SO2 control of western United States coal, and there is a single waste product. Other features include the collection of particulate matter at temperatures below 90°C and the possibility for spray dryer cooling tower water integration. This. system may or may not include a catalytic NO unit.  [c.492]

This equation is good if the air temperature is 50°F or above, the cooling tower s approach to the wet bulb temperature is 5°F or above, and Hog is within a range of about 0.1 to 8.  [c.158]

Cooling tower 10°F approach to design wet bulb temperature.  [c.408]

These values coupled with the flow rate and wet bulb temperature allow the selection of a cooling tower. Those new to cooling towers should make several selections at different wet bulb temperatures to test how wet bulb relates to cooling tower size. It becomes clear that the tower size increases as the wet bulb rises and that the size increase becomes dramatic as the approach is in the less than ten degrees area. This exercise demonstrates how to oversize a cooling tower... just use an inflated design wet bulb temperature. This is better than artificially inflating the flow rate and possibly over sizing the spray nozzles. Increasingly, manufacturers offer software to make selections easier.  [c.67]

The tendency for any system to grow biological material depends on several factors. Cooling tower design is one. Crossflow towers and counterflow towers without louvers, for example, tend to grow more algae due to the increased amounts of sunlight in contact with the system water. Water quality also comes into play. Make-up water that is reclaimed from a sewage treatment plant, for example, can be rich in nutrients. Also, some food processing operations where beer, tomato paste, milk, sugar, etc. enter the cooling system can have severe corrosion and biological problems. Another potential lies with air quality. Cooling towers located near bakeries, for example, show an increased tendency to grow biological material due to the molds and yeast. Biological concerns run the gamut from nearly zero to very substantial. Whatever the case, an appropriate solution must be developed. The traditional approach is for the operator to alternate between two liquid biocides adding them at a predetermined frequency. Two different formulations are often used to avoid an immunity being developed to just one. Unlike scale and corrosion chemicals that are metered into the system frequently, biocides are typically administered every few days to shock the system. Other chemicals for biological control include chlorine, iodine, bromine and ozone. These are not rotated with other chemicals They are fed by themselves. Ozone is also used to prevent scale. Be certain to specify Viton pump seals when using ozone Standard seals will fail quickly.  [c.87]

Design Conditions Thermal parameters for which the cooling tower is purchased. They consist of a given gpm flow of water entering the tower at a specific temperature, cooling through a given range, leaving the tower at the required temperature, and having a designated approach to a stated wet-bulb temperature.  [c.91]

FIGURE 4.19 Mathematical approach for a cooling tower.  [c.97]

Curve showing effect of wet bulb temperature, approach to wet bulb, and cooling range on cooling tower size. The normal tower is assumed to be designed for 15 degree cooling range and a 15 degree approach to a 70 degree wet bulb. If all other factors remain the same, reducing the approach to the wet bulb to 6.3 degrees will double the size of the tower, or decreasing the cooling range to 6.1 degrees will permit the use of a tower only half as large or designing for a 53.7 degree wet bulb instead of a 70 degree wet bulb will require a tower - A times as large because of the lower water absorbing capacity of colder air.  [c.393]

The Cooling Tower Institute (CTl) publishes sets of graphs that give demand in terms of three design temperatures and L/G. The CTl graph first published in 1967 gives KaV/L demand plotted against L/G with the approach temperature as a parameter. Each curve applies to a specified combination of wet bulb temperature and cooling range. Each cooling tower pack has its own KaV/L and responsible suppliers will supply performance graphs similar to those of the CTl.  [c.526]

The factors which should be applied at the design stage cover the water flow rate, the design wet bulb figure, the required return temperature at the design point, the cost of power and land, and the water analysis. Water flow is normally determined by the equipment that the cooling tower is serving (for example, heat exchangers). Process designers historically leave the cooling tower until last (it is, after all, the final heat sink). When water costs were negligible, this was acceptable, but with the increase in costs and, in certain cases, the restrictions on availability of water, this approach has had to be modified. Greater consideration must be given to the overall system. Experience in the last ten years has shown that economic optimization can lead to a more efficient cooling tower, with a corresponding drop in the cost of heat exchanger. This is particularly true in power generation and industrial processes.  [c.527]

Molecular beams are most often coupled to time-of-flight instruments for reasons that are discussed in section (Bl.7.5). The important advantage that molecular beams have over the other methods discussed in this section is their ability to cool the internal degrees of freedom of the sample. Collisions between a carrier gas (such as helium or argon) and the sample molecule in the rapidly expanding gas mixture results in rotational and vibrational cooling. Using this approach, the effective internal temperature of the sample can be significantly less than ambient. One example of the benefits of using molecular beams is in photoionization. The photon energy can be more readily equated to the ion internal energy if the initial internal energy distribution of the neutral molecule is close to 0 K. This cooling also allows weakly bound species such as neutral clusters to be generated and their resulting PI or El mass spectrum obtained.  [c.1331]

In practice, algiaate gels are obtaiaed usiag three principal methods, aamely, diffusioa settiag, internal settiag, or settiag by cooling (21). Diffusioa settiag is the simplest technique and, as the term implies, the gel is set by aHowiag calcium ions to diffuse iato an algiaate solutioa. Siace the diffusioa process is slow, this approach can only be effectively utilized to set thin strips of material (eg, pimiento strips), or to provide a thin gelled coating on the surface of a food product such as an onion ring.  [c.432]

The chilled surface drying process is similar to atmospheric drying, with the cooling process being driven by a chilled surface as opposed to air flow through a tall tower. The wet soap is superheated in high pressure, nonboiling heat exchangers. Drying is achieved by the release of steam when this superheated soap is introduced into a chamber with a slight negative pressure, which is commonly referred to as a flash chamber. The resulting hot, dry soap melt is cooled through the formation of a thin film on a chilled surfaced, commonly in the form of a roU (rotating cylinder). The hot dry soap falls into the small gap ( 10 50 fiTo) formed at the interface between a large chilled roU and a smaller, temperature-controlled (may be heated or cooled) apphcator roU that aids in uniform film formation. As the chilled roU rotates, the dry, cold soap is removed via scraping with a doctor blade and emerges in the form of flat flakes. The amount of soap drying is governed by the temperature at which the soap is introduced and the air flow in the flash chamber. This process is exceptionally good for modem synthetic surfactant containing formulations because it is amenable to more sticky in-process materials. This drying approach can also be achieved using a chilled belt in place of the chilled roU.  [c.156]

Chlorine is a toxicant with a typical dose-response pattern for biota (Fig. 7). There is a time-dependent mortaUty at high concentrations and a low concentration above which long-term chronic effects are shown. Early methods for using the time-dependent effects of high temperature for predicting safe temperatures and exposure times during cooling-water exposures led to a similar approach for chlorine (3). Chlorine toxicity data were scant and different aggregates had to be developed for freshwater and marine assemblages because of the markedly different chlorine chemistry in fresh and salt water.  [c.477]

Mechanical thermocompression may employ reciprocating, rotary positive-displacement, centrifugal, or axial-flow compressors. Positive-displacement compressors are impractical for all but the smallest capacities, such as portable seawater evaporators. Axial-flow compressors can be built for capacities of more than 472 mVs (1 X 10 ftVmin). Centrifugal compressors are usually cheapest for the intermediate-capacity ranges that are normally encountered. In all cases, great care must be taken to keep entrainment at a minimum, since the vapor becomes superheated on compression and any liquid present will evaporate, leaving the dissolved solids behind. In some cases a vapor-scrubbing tower may be installed to protect the compressor. A mechanical recompression evaporator usually requires more heat than is available from the compressed vapor. Some of this extra heat can be obtained by preheating the feed with the condensate and, if possible, with the product. Rather extensive heat-exchange systems with close approach temperatures are usually justified, especially if the evaporator is operated at high temperature to reduce the volume of vapor to be compressed. When the product is a sohd, an elutriation leg such as that shown in Fig. 11-1225 is advantageous, since it cools the product almost to feed temperature. The remaining heat needed to maintain the evaporator in operation must be obtained from outside sources.  [c.1143]

Another approach is to divide the tower arbitrarily into a lean section (near the top), where approximate methods are valid, and to deal with the rich section separately. If the heat effects in the rich seclion are appreciable, consideration could be given to installing cooling units near the bottom of the tower. In any event a design diagram showing the operating and equilibrium curves should be prepared to check on the applicability of any simplified procedure. Figure 14-8, presented in Example 6 is one such diagram for an adiabatic absorption tower.  [c.1355]

Because alterations to equipment design can be cumbersome and expensive, a more economical approach may be to change the metallurgy of affected components. Metals used in typical cooling water environments vary in their resistance to erosion-corrosion. Listed in approximate order of increasing resistance to erosion-corrosion, these are copper, brass, aluminum brass, cupronickel, steel, low-chromium steel, stainless steel, and titanium.  [c.249]

When optimizing the natural cooling of the enclosure its diameter may have to be increased but this would add to the cost of the IPB, on the one hand and make it too large to terminate at the generator end, on the other. The total design is thus a matter of many permutations and combinations. Most manufacturers have optimized their designs, based on experience and test results, through complex computer programming. In this example we will suggest only a basic approach to establish a design. Its optimization would require an elaborate cost analysis as noted earlier although its authenticity can be verified through laboratory tests. Assume the following parameters of a turbo-alternator (Table  [c.945]

In this arrangement, in contrast to the previous approach, the coolant is kept from evaporating by maintaining it under an inert gas pressure higher than its vapor pressure. A centrifugal pump is used to achieve high circulation rates. Besides the previously mentioned three coolants (water, tetralin, and Dowtherm A), other nonvolatile heat transfer oils as well as molten salts or molten metals can be used. These coolants are advantageous primarily at higher temperatures where, even if exothermic reactions are conducted in laboratory reactors, the heat losses are usually more than the reaction heat generated. A simple temperature controller, therefore, can be used to keep the electric heater adjusted by sensing with a thermocouple immersed in the return line. At a high recirculation rate of the liquid coolant, the constant wall temperature can again be approximated, but not as well as with boiling-type cooling.  [c.41]

In the basic continuous adsorption cycle illustrated previously in Fig. 5b two adsorbent beds are heated and cooled out of phase in order to provide continuous heating or cooling. It is possible to recover some of the adsorption heat rejected by each bed and use it to provide some of the heat required by the other bed. This might be done by the use of a circulating heat transfer fluid or a heat pipe. Meunier [5] first systematically looked at the potential gain in COP that might be obtained by such heat recovery, both as a function of the approach temperature of the beds donating and receiving heat and of the number of beds. The number of beds is not limited to two and the COP increases with the number of beds that it is possible to transfer heat between. There is of course a practical limitation, but it is possible to calculate the theoretical benefit of employing heat exchange between any number of beds.  [c.323]

There are several papers in the literature which give details of cycle calculations, and include details of how the cooling flow quantity may be estimated and used. Here we describe one such approach used by the author and his colleagues. Initially, we summarise how i/rcan be obtained (fuller details are given in Appendix A). We then illustrate how this information is used in calculations, once again using a computer code in which real gas effects are included.  [c.71]

Figure 6-3 shows a condensate stabilizer system. The well stream flows to a high pressure, three-phase separator. Liquids containing a high fraction of light ends are cooled and enter the stabilizer tower at approxi-  [c.132]

For high temperatures one would naively expect that r/7 is near to 1, and then for lower temperatures slowly starts to deviate from equihbrium, meaning that this ratio should decay. However, as Fig. 12 shows, for high temperatures this ratio is significantly smaller than 1, increasing with decreasing temperature and then breaking down to approach 0. The deviation at high temperatures comes from the fact that the simple definition of Eq. 4 takes into account only uncorrelated intra-chain effects. However, when one simulates a melt of chains the degeneracy of the states is significantly different, due to the constraints on a given monomer induced by all the other monomers. If one corrects for this, one actually finds that for high temperatures this ratio is very near to 1 and slowly decays as a function of temperature until it starts to break down around the glass transition temperature. However, the general scaling picture as shown in Fig. 12 does not change significantly up to a shift. Since all the data for the different cooling rates more or less coincide and define a characteristic temperature where the curve changes its behavior, this figure suggests that there might be a possi-bihty of defining a universal glass transition temperature for the cooling rate going to 0. This procedure actually allows one to compare the cooling rate dependent glass transition temperature with the asymptotic freezing temperature of a system. Certainly, there has to be much more research in this  [c.503]

A 500-ml, three-necked, round-bottom flask is fitted with a mechanical stirrer, a pressure-equalizing addition funnel and a reflux condenser (drying tube). The flask is charged with 100 ml of dry dimethylformamide, 9.2 ml (0.1 mole) of dry -butyl alcohol, and 28 g (0.107 mole) of dry triphenylphosphine. The flask is cooled in a water bath, stirring is begun, and bromine (approx. 16 g, 0.1 mole) is added over about 15 minutes.  [c.46]

Much work has been done on the modeling of wet cooling-tower plumes the ultimate aim was to determine their effect on the environment (42—44). The basic approach involves writing the equations of continuity, conservation of energy and momentum, and the equations of motion for the conditions of the plume. These are then solved simultaneously using iterative and numerical methods on large computers. The accuracy of the results depends on whether the boundary conditions and simplifying assumptions are reaUstic. These are difficult to accomplish because conditions change rapidly, ground configurations are seldom simple, and plume behavior is influenced by a host of casual, nonrepeated situations such as the passing of an airplane or the presence of cloud cover. Modeling owes much to meteorology and especially to the theory of cumulus clouds (see Atmospheric modeling).  [c.106]

It is never appropriate to add any type of anti-freeze solution to an open cooling tower. Closed (fluid cooler) systems, however, can be protected from freeze-up by the addition of ethylene glycol or other fluids. Fluid cooler casing sections can also be insulated to reduce heat loss thereby protecting the coil from freeze-up. Counterflow, blowthrough towers tend to be more popular as the freeze potential increases. Crossflow towers tend to freeze water on their air inlet louvers under extreme conditions. Fans (propeller type) can be arranged to reverse direction on such towers to melt ice. This process should never be automated. Instead, the operator should weigh the situation and reverse the fan only as long as required. The designer must select components suitable for reverse rotation. Fan discharge dampers are a capacity control accessory item for centrifugal fan cooling towers. They fit in the fan scroll. In the open position, they are much like a thin piece of sheet metal in a moving airstream oriented parallel to airflow. The airstream doesn t know its there. As the dampers close- the sheet metal becomes less parallel to airflow- turbulenee disrupts the air stream. Airfoil dampers essentially ruin fan housing effieiency to achieve a reduction in airflow. Dampers can set and locked when a manual locking quadrant is specified but it is more common to use electric or pneumatic actuators that close the dampers as the exiting water temperature becomes too low. While reducing airflow is the correet method of reducing capacity, dampers are not the best approach. They offer the poorest energy savings and the actuating mechanisms tend to fail long before the average eooling tower life span.  [c.79]

In the automobile industry the road trials on test tracks are validation tests as are the customer trials conducted over several weeks or months under actual operating conditions on pre-production models. Sometimes the trials are not successful as was the case of the Copper Cooled Engine in General Motors in the early 1920s. Even though the engine seemed to work in the laboratory, it failed in service. Production was commenced before the design had been validated. The engine had pre-ignition problems and showed a loss of compression and power when hot. As a result, many cars with the engine were scrapped. Apart from the technical problems GM experienced with its development, it did prove to be a turning point in GM s development strategy, probably resulting in what is now their approach to product quality planning.  [c.265]

The apparatus described in the preceding experiment is employed. To 415 g of 98 % sulfuric acid is added 200 g of 30% fuming sulfuric acid and the solution cooled to approx. 10°. About 3 ml of 88% formic acid is added and the rapidly stirred solution cooled to 5°. After 5 minutes, 46 g of 88 % formic acid and 50 g of decahydro-2-naphthol are added from two dropping funnels over hour. The solution should foam greatly during the additions. After stirring at approx. 0° for hour longer, the solution is poured on ice. The oil, which soon crystallizes, is dissolved in ether and extracted into 10% sodium carbonate solution. Acidification of the aqueous layer gives about 80% of 9-decalincarboxylic acid which is largely cis. After three recrystallizations from acetone, the pure product is obtained, mp 122-123° (yield about 7 g).  [c.135]

The effects of wet bulb, approach and range on mecbanical draft cooling tower size is indicated in Figure 9-118.  [c.389]

The temperatures or enthalpy change for the streams (and hence their slope) cannot he changed, but the relative position of the two streams can be changed by moving them horizontally relative to each other. This is possible because the reference enthalpy for the hot stream can be changed independently from the reference enthalpy for the cold stream. Figure 6.16 shows the same two streams moved to a different relative position such that AT ,in is now 20°C. The amount of overlap between the streams is reduced (and hence heat recovery is reduced) to 10 MW. More of the cold stream extends beyond the start of the hot stream, and hence the amount of steam is increased to 4 MW. Also, more of the hot stream extends beyond the start of the cold stream, increasing the cooling water demand to 2 MW. Thus this approach of plotting a hot and a cold stream on the same temperature-enthalpy axis can determine hot and cold utility for a given value of Let us now extend this approach to many hot  [c.161]

At microwave frequencies, direct absorption teclmiques become less sensitive than those in the THz region due to tire steep dependence of the transition intensities with frequency. A variant of heterodyne spectroscopy, pioneered by Flygare and Balle [28], has proven to be much more sensitive. In this approach, molecules are seeded into or generated by a pulsed molecular beam which expands into a high-g microwave cavity. The adiabatic expansion cools the rotational and translation degrees of freedom to temperatures near 1-10 K, and thus greatly simplifies the rotational spectra of large molecules. In addition, the low-energy collisional enviromnent of the jet can lead to tlie growdi of clusters held together by weak intennolecular forces.  [c.1244]

Batch processiag of nylon-6 is generally used only for the production of specialty polymers such as very high molecular weight polymer or master batch polymers for special additives. In a typical modem batch process (147—150), the caprolactam is mixed ia a hoi ding tank with the desired additives and then charged to an autoclave with a small amount (2—4%) of water. During the two-stage polymeriza tion cycle, the temperature is raised from 80 to 260°C. In the first stage, water is held ia the reactor, the pressure rises, and the hydrolysis and addition steps occur. After a predetermined time the pressure is releasedand the final condensation reaction step occurs. The molecular weight of the polymer can be iacreased by means of a vacuum finishing step, if desired. The entire process can take three to five hours. The final polymer is then drained, often with a forcing pressure of iaert gas, through a die to form ribbons of polymer, which are then cooled ia water and cut iato pellets. Because nylon-6 has such a high monomer and oligomer content, 10—12% by weight, ia the cast pellets, which would significantly reduce the quaUty of the final fiber or resia products, it must be extracted. This is usually done ia hot water under pressure at 105—120°C for 8—20 h. Most of the caprolactam and higher oligomers that are released with the steam from the autoclave or extracted from the pellets ia hot water are then recycled. The pellets must be carefully dried because excess water decreases the molecular weight of the polymer duting subsequent melt processiag. The fiaal polymer processed through water extractioa and drying can have an oligomer level of <0.2% and a moisture level of <0.05%. Alow level of total oligomers is necessary because on remelting and further processiag, the oligomers coateat will iacrease owiag to the reestabUshmeat of the equiUbrium distributioa of molecular species that occurs for all coadeasatioa polymers (151). Because the approach to equihbrium progresses at a moderate rate, it is possible to utilize extracted ayloa-6 ia a remelt process without increasing the oligomer coaceatratioa above 2—3% and thus avoiding any significant drop ia fiaal properties.  [c.234]

From Fig. 12-14, the water concentration required to perform the cooling is 1.75 gal/(min-fF), giving a tower area of 1145 fF versus 1000 ft for a 70°F wet-hiilh temperature. This shows that the lower the wet-hiilh temperature for the same cooling range and approach, the larger is the area of the tower required and therefore the more difficult is the cooling joh.  [c.1165]

The performance of the natural-draft tower differs from that of the mechanical-draft tower in that the cooling is dependent upon the relative humidity as well as on the wet-bulb temperature. The draft will increase through the tower at high-humidity conditions because of the increase in available static pressure difference to promote air flow against internal resistances. Thus the higher the humidity at a given wet bulb, the colder the outlet water will be for a given set of conditions. This fundamental relationship has been used to advantage in Great Britain, where relative humimties are commonly 75 to 80 percent. Therefore, it is important in the design stages to determine correctly and specify the density of the entering and effluent air in addition to the usual tower-design conditions of range, approach, and water quantity. The performance relationship to humidity conditions makes exact control of outlet-water temperature difficult to achieve with the natural-draft tower.  [c.1169]

The supersaturation in condensers arises for two reasons. First, the condensable vapor is generally of higher molecular weight than the noncondensable gas. This means that the molecular (iffusivity of the vapor will be much less than the thermal diffusivity of the gas. Restated, the ratio of NsJNp is greater than 1. The result is that a condenser yields more heat-transfer units dT / T — T,) than mass-transfer units dYg/(Yg — Y ). Second, both transfer processes derive their driving force from the temperature difference oetween the gas Tg and the interface T,. Each incremental decrease in interface temperature yields the same relative increase in temperature driving force. However, the interface vapor pressure can only approach the hmit of zero. Because of this, for equal molecular and thermal diffusivities a saturated mixture will supersaturate when cooled. The tendency to supersaturate generally increases with increased molecular weight of the condensable, increased temperature differences, and reduced initial superheating. To evaluate whether a given condensing step yields fog requires rigorous treatment of the coupled heat-transfer and mass-transfer processes through the entire condensation. Steinmeyer [Chem. Eng. Frog., 68(7), 64 (1972)] illustrates this, showing the impacl of foreign-nuclei concentration on calculated fog formation. See Table 14-14. Note the relatively large particles generated for cases 1 and 2 for 10,000 foreign nuclei per cm. These are large enough to be fairly easily collected. There have been veiy few documented problems with industrial condensers despite the fad that most calculate to generate supersaturation along the condensing path. The explanation appears to be a limited supply of foreign nuclei.  [c.1414]

The strucmres of the first group of glasses are consistent with the suggestion that tire smaller metalloid atoms fill holes in tire metal strucmre, and enable a closer approach of tire metal atoms and an increased density. The atTangement of the metal atoms can either be in a random network or in the dense random packing model of Bernal, in which the co-ordination number of the metalloid can be 4, 6, 8, 9 and 10, some of tire latter tlrree involving 5 co-ordination of the non-metal. The formation of glass requires tlrat the rate of cooling from the melt must be greater than the rate at which nucleation and growth of a crystalline phase can occur. The minimum rate of cooling to attain the glass structure cair be obtained for airy system by dre observation of dre rate of crystallization as a function of supercooling below the liquidus. The extremum of dre TTN (time, teirrperamre, nucleation rate) curve shows dre maximum rate of nucleus formation as a function of undercooling temperature, and hence dre minimum in dre rate of cooling required to achieve dre formation of a glass.  [c.298]

Activation of the porous composite was carried out in moisture saturated He, or CO2, in the temperature range 800-950°C. Attaining uniform activation in large composite pieces is challenging. The use of a special gas distribution manifold partially alleviated this problem, yet attaining uniform activation in sections > 25 mm thick has proven to be very difficult. Jagtoyen and Derbyshire [4] applied the O2 chemisorption technique of Nandi and Walker [5] and Quinn and Holland [6] and achieved improved uniformity of activation in monoliths up to 12 x 7 x 6 cm in size. At ORNL, we have also applied this approach and successfully activated porous composite monoliths of 10 cm in diameter and 25 cm in length [7]. The chemisorption/activation procedure adopted was performed in a three zone Lindberg furnace fitted with a 20-cm diameter Inconel retort. The porous composite samples were dried in a vacuum at 300°C, cooled to 200°C for the chemisorption treatment in flowing O, and then heated again to 850°C in He. The chemisorption and activation steps were repeated until the desired bum-off was attained.  [c.172]

An alternative approach is shown in Fig. 4.6c. The cooled efficiency (77)101 may be presented as a unique function of the rotor inlet temperature (T5), for a given x and component efficiencies. But from the El-Masri expression, Eq. (4.25), it can be deduced that the cooling air quantity tfj is a function of the combustion temperature T, for a given x (and Tj) and a selected blade temperature Tbi, so that from the. steady flow energy equation for the mixing process, there is a value of 7 3 conesponding to the rotor inlet temperature Ts. Analytically, we may therefore state that (i7)ici =f(T[c.57]

An empirical development of the approach described above uses experimental cascade data, obtained with and without coolant discharge, to obtain an overall relationship between the total cooling flow through the blade row (i/() and the extra stagnation pressure loss arising from injection of the cooling air. In film cooling, the air flow leaves the blade surface at various points round the blade profile causing variable loss (noting that injection near the trailing edge causes little total pressure loss—it may even reduce the basic loss in the wake). If there is an elementary amount of air di/ at a particular location where the injection angle is , then an overall figure for the extra total pressure loss due to coolant injection in a typical blade row can be obtained by integrating the Hartsel equation (4.47) round the blade profile f,1]. An overall exchange factor for the extra blade row stagnation pressure mixing loss in the row can thus be obtained in the form  [c.64]

N,N-Dimethylcyclohexylmethylamine A 500-ml three-necked flask fitted with a reflux condenser (drying tube), a pressure-equalizing dropping funnel, and a magnetic stirrer is charged with a mixture of 5 g of lithium aluminum hydride in 60 ml of anhydrous ether. A solution of 20 g of A, A-dimethylcyclohexanecarboxamide in 50 ml of anhydrous ether is added to the stirred mixture at a rate so as to maintain a gentle reflux (about 30 minutes). The mixture is then stirred and heated (mantle) at reflux for 15 hours. The flask is fitted with a mechanical stirrer and cooled in an ice bath. Water (12 ml) is added slowly with vigorous stirring, and the stirring is continued for an additional 30 minutes. A cold solution of 30 g of sodium hydroxide in 75 ml of water is added and the mixture is steam-distilled until the distillate is neutral (about 225 ml of distillate). The distillate is acidified by addition of concentrated hydrochloric acid (approx. 15 ml) with cooling. The layers are separated, and the ether layer is washed with 10 ml of 3 A hydrochloric acid. The combined acidic solutions are concentrated at 20 mm pressure until no more distillate comes over at steam-bath temperature. The residue is dissolved in water (approx. 30 ml), the solution is cooled, and sodium hydroxide pellets (17 g) are added slowly, with stirring and cooling in an ice-water bath. The layers are separated, and the aqueous phase is extracted with three 15-ml portions of ether. The combined organic layer and ether extracts are dried over potassium hydroxide pellets for 3 hours. The solvent is removed by fractional distillation at atmospheric pressure, and the product is collected by distillation under reduced pressure. The yield is about 16 g (88%) of A,A-dimethylcyclohexylmethylamine, bp 76729 mm, 1.4462-1.4463.  [c.20]

In a 250-ml three-necked flask fitted with a magnetic stirrer, a pressure-equalizing dropping funnel, and a thermometer is placed a solution of l,4-cyclohexanediol(l 1.4g, 0.10 mole), 35 ml of chloroform, and 27 ml of dry pyridine. The solution is cooled in an ice bath to 0-5 and is maintained below 5 throughout the addition. A solution of benzoyl chloride (14 g, 0.10 mole) in 30 ml of dry chloroform is added with stirring at a rate so as to keep the temperature below 5° (approx. 40 minutes). After completion of the addition, the mixture is allowed to come to room temperature and stand overnight. The chloroform solution is washed four times with 50-mI portions of water, once with 50 ml of 5 % sulfuric acid solution, and finally with saturated sodium chloride solution. The chloroform solution is then dried (sodium sulfate), and the solvent is removed. Fractionation of the residue gives a cis and trans mixture of 4-benzoyloxycyclohexanol, bp 175-17870.2 mm, as a very viscous oil, yield about 55%.  [c.64]

A 500-ml, three-necked, round-bottom flask is fitted with a mechanical stirrer, a condenser (drying tube), and a dropping funnel. The flask is charged with 150 ml of absolute ethanol, and sodium (11.5 g, 0.5 g-atom) is added cautiously (stirring unnecessary) to prepare a solution of sodium ethoxide. The flask is cooled to 10° with stirring, whereupon a previously cooled (ice bath) mixture of cyclohexanone (49 g, 0.5 mole) and diethyl oxalate (73 g, 0.5 mole) is added over about 15 minutes. (Rapid stirring is advisable to prevent solidification of the mixture.) After the addition, stirring in the cold is continued for 1 hour followed by stirring at room temperature for 6 hours. The stirred flask is again cooled to 5-10° and a mixture of 14 ml of concentrated sulfuric acid and 110 g of ice is cautiously added to decompose the reaction mixture, the temperature being maintained at 5-10°. The solution is then diluted with water (approx. 1 liter), and the oily ethyl 2-ketocyclohexylglyoxalate separated. The aqueous phase is extracted four times with 100-ml portions of benzene and the extracts are combined with the original organic phase. The benzene solution is washed twice with water and is transferred in portions to a 250-ml round-bottom flask set up for distillation on a steam bath. Distillation is continued until benzene no longer comes over. The distillation setup is now heated with an oil bath under vacuum, and the forerun is collected below 105°/10-12 mm. The product is collected at 105-165710-15 mm (final bath temperature of about 200°), yield about 65 g (65 %). The distillate is transferred to a 125-ml flask set up for distillation, and approx. 1 mg of iron powder and 0.5 g of finely ground soft glass are added. The mixture is distilled at 40 mm, the decarbonylated product being collected over the range 125-140° (bath temperature 165-175°). The crude 2-carbethoxycyclohexanone, about 50 g (59%), may be purified by distillation, bp 93-94°/l mm.  [c.88]

See pages that mention the term Cooling tower approach : [c.66]    [c.70]    [c.1216]    [c.33]    [c.161]    [c.217]   
Applied Process Design for Chemical and Petrochemical Plants, Volume 2 (1997) -- [ c.382 ]