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Unformatted text preview: Heat Transfer Properly Employ
O verhead Condensers
for Vacuum Columns
In a vacuum distillation system, the
condenser is an important determinant
of overall performance. Follow these
guidelines for specifying, selecting and
designing overhead vacuum condensers. David Greene
George J. Vago
Consultant V acuum distillation may be used to enhance
the separation of hydrocarbons because of
the improvement in relative volatility at the
lower pressure. Vacuum may also be employed to
break azeotropes by using two columns in series
that operate at different pressures. And, it may be
required to reduce the boiling point of the bottom
composition so conventional heat-transfer fluids
can be used for heating.
The improved performance of vacuum column
operation must be weighed against the additional
capital and operating costs of the vacuum system.
Because of these high costs, vacuum operation is
normally reserved for situations where the components are either very difficult to separate at higher
pressures or where the products would be degraded
at elevated temperatures.
Once a decision is made to operate under vacuum,
it is generally advantageous to provide a condenser upstream of the vacuum source to reduce the size and
cost of the vacuum system and recover process material. A vent condenser after the main condenser, possibly
using a colder cooling medium, may be used to further
reduce the load on the vacuum system.
The main process parameters involved in specifying
vacuum systems are the absolute pressure and the
amount of noncondensables. Other considerations in- 38 www.cepmagazine.org February 2004 CEP Typical overhead condenser installed on a vacuum distillation column. clude the cooling medium and its temperature, the desirability of recovering process vapors, and the need to treat
either the gaseous or liquid effluent stream, or both. In a
batch system, the time necessary to pull the system down
to the desired vacuum level is also relevant, and in fact
might control the sizing of the vacuum system. Pressure
Vacuum can be classiﬁed as high, medium or low,
where “high vacuum” corresponds to the lowest absolute pressure: low vacuum ≈ 760–100 torr; medium Heat Transfer Noncondensables
Noncondensables can be introduced through in-leakage,
they may be entrained or dissolved in the process ﬂuid, or
they may be produced in the distillation column by cracking of heavier components. During start-up, additional
noncondensables in the form of contaminants, such as oil
or grease, may be present in the system.
Two methods are available to estimate in-leakage. A
crude technique developed by the Heat Exchange Institute
(HEI) (1) provides an estimate of in-leakage as a function of
absolute pressure and system volume for tight systems. Although quick and simple, there seems to be a general consensus that the results are overly conservative (2).
A somewhat more scientiﬁc basis is to count the ﬁttings, valves, ﬂanges and any other connections between
the process and atmosphere, and estimate in-leakage
based on either in-house or published data. In this approach, dynamic connections are assigned more weight
than static connections.
The amount of noncondensables dissolved in the process stream or produced in the process must also be estimated and added to the noncondensables in-leakage estimate to determine a total noncondensable load. If a safety
factor is desired, it is added to the total noncondensable
load. Remember that the control system should be designed to accommodate the difference between the actual
or expected noncondensables and the design allowance
without introducing additional inerts.
The next step is to saturate the noncondensables with the
process components and determine the total load to the condenser and the physical properties of the process stream. It is obviously desirable to minimize in-leakage by
eliminating unnecessary ﬁttings and avoiding situations
that permit the ingress of noncondensables. To the extent
possible, the system should be welded and the distance
from the process source through the condenser to the vacuum system minimized. Temperature
A vent condenser will further reduce the load on the
vacuum system, recover process material and reduce effluent-treatment requirements. It may be desirable to add refrigeration to maximize recovery.
At the same time, the design should be based on a reasonable coolant temperature and a realistic temperature
approach. The cost of designing for the worst-case cooling
water temperature and a close approach will be very high
for a vacuum condenser.
Also, use caution when dealing with materials with
high freezing points to avoid the possibility of freezing. A
recirculating tempered-cooling loop may be provided to
obtain a close approach while maintaining the process
temperature above the freezing point.
TEMA data sheet
After the heat duty has been calculated, a condensing
curve, as illustrated in Figure 1, is prepared. This curve
will be used to calculate a weighted mean temperature difference to most accurately reﬂect the temperature difference throughout the heat exchanger.
A data sheet, such as provided in the Tubular Exchanger
Manufacturers Association (TEMA) standards (3), is used to
transmit pertinent ﬂuid properties and mechanical preferences
from the process engineer to the heat transfer specialist. 240.0 Temperature, ˚F vacuum ≈ 100–10 torr; and high vacuum ≈ 10–1 torr. The
costs of producing vacuum increase exponentially with the
level of vacuum.
Once the level of vacuum is established by evaluating the process requirements against cost and performance, the system can be specified. The first decision is
how to allocate the pressure drop between the process
user and the vacuum system and how to split this pressure drop between the piping and condenser. A rule-ofthumb is to allocate about 10% of the absolute pressure
to the pressure drop in the piping and condenser. Because condensation increases with pressure, the pressure loss between the process equipment and condenser
should be minimized. In the case of medium- and highvacuum systems, the preferred arrangement is to integrally connect the process equipment and condenser.
Eliminating the piping and mounting the condenser directly on the column minimizes the pressure drop between the column and condenser.
If the condenser is mounted away from the column, the
piping should be generously sized, ﬁttings minimized and
valves avoided. Valves are expensive in large-bore piping
and are sources of both in-leakage and pressure drop. 220.0 200.0 180.0 160.0 0.0 20.0 40.0 60.0 80.0 100.0 Duty, % s Figure 1. Typical heat release curve. Relationship between condenser and column
Piping. The condensate piping must be properly designed to ensure good drainage at all times. This includes a correctly designed barometric seal with an ap- CEP February 2004 www.cepmagazine.org 39 Information Technology propriate allowance for the density of the condensate
and any tendency to foam. For a fluid density of 50
lbs/ft3 and a tendency to foam, 45 ft is an appropriate
height. The seal piping should be run vertically. If this
is not possible, one or two 45-deg sections may be permitted, but this should be considered as a last-resort
approach. Horizontal runs should never be used, as
they might allow gas bubbles to form and agglomerate
into vapor pockets, which would impair draining and
might cause backup to the condenser. Remember also
that condensate at sub-ambient temperature may warm
up in the seal leg or hotwell and flash. To avoid this, it
may be necessary to insulate the pipe.
In addition, the hotwell seal needs to have sufficient
volume to ﬁll the pipe to avoid breaking vacuum if the
system is shut down. A 12-in. seal between the pipe outlet
and hotwell overﬂow is suggested to accommodate minor
Condensate collection and reflux distribution. Distribution of reflux is a critical factor in the operation of a
vacuum column because of the liquid loading. It is also
important to avoid entrainment of the returning reflux
into the exiting vapor stream. The most common reflux
arrangement is to remove part of the condensate as
product (distillate) and return the remainder as reflux.
Some columns return all of the liquid to the column and
remove the product somewhat below the condenser
from a specially designed drawoff stage. In either case,
the condensate must first be collected in a central location for further distribution.
The most common way to collect condensate for a directly mounted condenser is to provide a welded annular
trough around the inlet nozzle. Liquid can be withdrawn
from this trough via either an overflow weir or a pipe to
direct the flow to the column below or to a product line.
Using an internal splitter is simple, but the reflux
flowrate cannot be directly measured. At one time, it was
common to withdraw the entire condensate and measure
the reflux and product streams from the pumped discharge. With today’s control systems, the reflux can be
determined by performing a heat balance on the condenser using coolant temperatures and flows to calculate
the total condensate and subtracting the distillate flow
measurement, thereby permitting the use of the simpler
internal splitter for reflux control.
Special consideration must be given to the reflux distributor in the column. As vacuum level increases, the
column cross-section gets larger, but the reflux volume
remains the same and may even decrease. In order to
provide good distribution for wetting packing or ensuring uniform flow across trays, a specially designed distributor is normally used. These distributors become
quite sophisticated if turndown is required, and it is
most important that the distributor be level to ensure
uniform flow across the column. 40 www.cepmagazine.org February 2004 CEP Selecting the design
The selection of the vacuum condenser type is mainly a
function of the operating pressure and the allowable pressure drop. The size of the unit is then determined by the
speciﬁed vapor mass ﬂowrate and operating pressure.
Condensers operating at low vacuum (100–760 mm Hg)
are the easiest to design; difficulty increases exponentially as
the vacuum level increases. Experience shows that the most
critical design factor for vacuum condensers is the proper allocation of pressure drop throughout the unit. Since the
largest losses occur at the inlet nozzle before condensation
starts, this region must be carefully analyzed. As the size of
the condenser increases, the supporting structure becomes
another important cost factor in optimization of the design.
The following criteria must be factored into the selection of a shell-and-tube condenser:
Condensation inside vs. outside of tubes. Condensation
inside tubes is not appropriate for vacuum overhead condensers because of the high tubeside pressure drop and the
difficulty in piping and supporting a vertically mounted
unit. Therefore, condensation on the outside of the tubes is
the best choice.
Horizontal vs. vertical shell orientation. Vertical orientation for outside-tube condensation is justiﬁed for tube
bundles mounted inside the column or possibly inside a receiving tank. However, it is more difficult to meet the low
pressure-drop objectives when compared to horizontal designs because it is not possible to split the ﬂow in a vertical arrangement. In addition, the use of a U-tube bundle
may not be allowed, since the coolant is not drainable and
a more costly ﬂoating head would have to be provided.
The optimal design can be achieved in a horizontal shell.
Shellside ﬂow arrangements. Dividing the ﬂow crossing
the tube bank can minimize pressure drop in a horizontal
unit. Using divided shell ﬂow and double segmental bafﬂes can achieve this very effectively. Further decrease in
pressure drop can be achieved by increasing the tube layout pitch, which will open up the cross-ﬂow area. Typically, the vapor enters at the bottom in the center of the shell,
the ﬂow is divided inside the tubeless shell bottom, and
noncondensables exit at the top in the center. The inlet
nozzle is specially designed for collecting the condensate.
The applied LMTD correction factor is the same as in a
conventional divided-ﬂow unit.
Mechanical design and materials of construction. The
minimum piping pressure drop is achieved by eliminating the piping between the column and condenser.
Mounting the condenser at the top of the column does
present some mechanical and maintenance difficulties
and may require a booster pump to supply coolant to tall
columns, but there is a significant process advantage in
minimizing pressure drop.
The unit must be self-supporting from the vapor inlet
nozzle and designed to prevent in-leakage of the ambient
air. This is important in order to avoid overloading the vac- Heat Transfer Nomenclature
A = total required heat transfer surface, ft2 D = tube outer diameter, in. g = gravitational constant, 4.17 × 108 ft/h2 hc = condensing heat-transfer coefficient, Btu/h-ft2-°F hv = convective heat-transfer coefficient of vapors, Btu/h-ft2-°F hw = tubeside heat-transfer coefficient of coolant, Btu/h-ft2-°F k = thermal conductivity, Btu/h-ft-°F L = effective tube length, ft N = number of vertical tube rows ∆P = pressure drop, psi
∆P2 = two-phase ﬂow pressure drop, psi
∆PL = liquid-phase ﬂow pressure drop, psi
Q = heat duty, Btu/h Qv = vapor cooling duty, Btu/h
R = thermal resistance, h-ft2-°F/Btu Re = Reynolds number
T = temperature, °F ∆T = temperature difference, °F
U = overall heat-transfer coefficient, Btu/h-ft2-°F Wc = total mass ﬂowrate of condensate, lb/h
x = weight fraction of vapor (vapor quality) Xtt = Lockhart–Martinelli parameter for liquid and
vapor in turbulent ﬂow
λ = latent heat, Btu/lb µ = viscosity, lb/ft-h ρ = density, lb/ft3 ΦL2 = Lockhart-Martinelli two-phase pressure drop multiplier Subscripts
f = fouling i = vapor/condensate interface L = liquid phase m = metal n = local value at n-th section along the condensing curve o = outer tube surface v = vapor phase uum system and reducing the condenser performance. At
the same time, maintenance requirements may dictate that
the unit be removable. The best solution is to use a ﬁeldwelded connection. This also assures minimum leakage.
If a welded design cannot be used, a mounting ﬂange
rated for 300 lb with spiral-wound gaskets would be the
next-best choice. This arrangement can be quite expensive
for large high-alloy-steel applications. Leakage of coolant into the shell must be minimized.
Using strength-welded tube-to-tubesheet joints per UW20 of the ASME Pressure Vessel Code, section VIII, is
The choice between a U-tube or fixed tubesheet tube
bundle depends mainly on the thermal expansion stresses between the tubes and the shell. For a fixed tubesheet
design, when the temperature difference between the
tube wall and the shell at operating or any upset conditions results in excessively high stresses, a shell expansion joint is required. To avoid an expansion joint, a Utube bundle may be selected. By using a non-removable
channel with cover (TEMA N-U type), the shellside
leakage would be minimized.
Trace impurities such as chlorides may concentrate at
the phase change. If this is anticipated, corrosion-resistant
materials of construction are required.
Entrainment of liquid. Entrainment should be minimized through proper baffling and deﬂector plates at the
shell exit section of the heat exchanger. Excessive entrainment will reduce condenser capacity and prevent the optimum recovery of product.
Freezing of condensing liquid on the tube wall. Some
process liquids may have a freezing point close to the
coolant temperature. The heat-transfer specialist will take
this into consideration when designing the heat-transfer
surface by checking the outer tube-wall-temperature,
which must be kept at a safe minimum. This temperature
can most effectively be controlled by increasing the
coolant inlet temperature. The most critical part of the
condenser is the bundle exit section, where the heat ﬂux is
at minimum and the heat-transfer coefficient is dominated
by convective heat transfer.
Heat transfer calculations
The heat transfer mechanism in the condenser is partial condensation. Due to the presence of noncondensable gases in the multicomponent vapor mixture, the
condensing vapors must be cooled along the condensing
curve. The cooling process is a combination of heat and
mass transfer that can be calculated by the rigorous
method of Colburn and Hougen, which is described
with practical examples by Hewitt, et al. (4). The complexity of mass transfer correlations and a lack of data
on diffusivities for multicomponent mixtures of vapors
led to the use of a simplified calculation method developed by Bell, et al. (5), which is the basis for the calculations presented here.
In this method, the incremental heat ﬂux into the
vapor/liquid interface must be equal to the heat ﬂux from
the condensing liquid ﬁlm to the coolant:
dQ/dA = U(Ti – Tw) (1) dQv/dA = hv(Tv – Ti) (2) CEP February 2004 www.cepmagazine.org 41 Information Technology Solving for Ti and introducing the simplifying parameter Z = dQv/dQ yields:
dQ U (Tv − Tw )
dA (1 + UZ hv ) (3) The overall heat-transfer resistance is:
+ R fo + Rm + R fi + o U hc Di hw ( 4) (1 + UZ hv ) dQ
U (Tv − Tw )
0 Q (5) A=∫ The shellside vapor-phase convective heat-transfer coefficient, hv, is calculated using the Tinker type streamanalysis method, as described by Hewitt, et al. (4).
The condensing liquid-ﬁlm heat-transfer coefficient is
established by the correlations for single vapor condensation as given by McAdams (6) for two regions of ﬂow.
Gravity-controlled condensation occurs at ReL < 2,100,
and the heat-transfer coefficient in this region is:
0.25 (6 ) For ReL > 2,100, the heat-transfer coefficient becomes
shear-controlled and can be established from the graphical
interpretation given by McAdams (6), where the heattransfer parameter is plotted as a function of ReL:
hc 3 2 kL ρ L g ReL = 0.33 = f ( ReL ) (7) 4Wc
NLµ L (8) Pressure drop calculations
The pressure drop through the condenser is typically
calculated from the inlet to the outlet nozzle of the shell,
and it consists of following components:
∆P = ∆Pinlet + ∆P90-deg-turn + ∆Pdome
+ ∆Pentrance + ∆P2Φ + ∆Poutlet 42 www.cepmagazine.org February 2004 ΦL2 = f(Xtt) (10) The local Lockhart-Martinelli parameter is deﬁned as: Integrating Eq. 3 reveals that the required heat-transfer
surface area will be: k 3 ρ 2 λg hc = 0.725 L L Nµ L D∆T The single-phase shellside pressure drops are calculated
by correlations as deﬁned by Kern (7). The two-phase
pressure drop across the bundle is established by applying
the Lockhart-Martinelli method (8), where the pressure
drop is calculated for a single phase (all liquid ﬂow) and
then corrected by a two-phase ﬂow multiplier, shown
graphically as a function of the Lockhart-Martinelli/Nelson parameter Xtt in the form: (9) CEP 1− x
Xtt = x 0.9 ρL ρv 0.5 µL µV 0.1 (11) The total two-phase pressure drop is then: ( ∆P2 Φ = ∑ ∆PL × Φ 2
n n ) (12) Example: Application of computer programs
Vacuum condensers for one installation were sized by
applying a combination of manual and computer calculations with full recognition of cost optimization.
The TEMA (3) designation of the condenser was type
BJM, meaning divided-ﬂow, ﬁxed tubesheets, and bonnet
heads with the condensing-vapor inlet nozzles located on
the top of the shell. This case was a special design in which
the ﬂow arrangement was reversed and one vapor inlet nozzle was located at the bottom to also catch returning condensate. The full ﬂow was returned to the column as reﬂux
via drain holes at the shell’s bottom and a half-pipe drain.
The detailed design is shown in the setting plan in Figure 2.
For the thermal design, the HTRI condenser computer
program was used. This program follows, in principle, the
calculation methods outlined here, but with more exact and
improved correlations. Since the program is not completely
adaptable to the special features of this design, some modiﬁcations were made to satisfy the standard input:
1. The HTRI designation of BJ21M was speciﬁed, denoting two inlet nozzles and one outlet nozzle. Therefore,
the single inlet had to be split into two nozzles with an
equivalent mass-ﬂow velocity. The effect on heat transfer
of the upward ﬂow at the inlet section with the liquid ﬂowing down was neglected and checked for ﬂooding only.
2. The shell size was speciﬁed as the equivalent diameter of the real shellside ﬂow area, which corresponds to the
full shell area minus the segmental-inlet-ﬂow area.
The pressure drop determined by the computer program
was modiﬁed by manual calculations to account for all
contributing components, in accordance with Eq. 9. The
shell size was also established by manual calculations so
that the diameter would agree with the bundle tube count
and more accurately simulate the shellside ﬂow areas. Heat Transfer Vapor Vapor Outlet Outlet Cooling Cooling Outlet Outlet Cooling Splash Baffle Inlet
Cooling PI Inlet and TI
Pipe Inlet and Column Mount Inlet and
Column Mount s Figure 2. Setting plan for condenser.
60 Pressure drop distribution
and condenser size
The frictional losses in the dome
area and the outlet nozzle losses of the
noncondensable vapors are negligible.
The distribution of the total pressure
drop among the components presented
in Eq. 9 can be expressed as a percentage of the total pressure losses for a
typical design lumped into three
groups, as shown in the table.
It can be seen that the inlet losses
represent an appreciable portion of the
pressure loss. Because these pressure
losses are mainly a function of shell diameter, they directly affect the size and
cost of the condenser. Figure 3 illustrates the dependence of the shell diameter on the operating pressure using the duty and ﬂow
from the design example and allocating 10% of the operatCEP
ing pressure to the condenser pressure drop.
Cut Baffles Shell ID, in. 50 40 Table. Pressure drop distribution in a typical condenser. 30 Inlet nozzle + impingement on
the dome plate . . . . . . . . . . . . . . . . . . . . . . . .35%
10.0 100.0 1000.0 Pressure, torr 90-deg turn of the ﬂow into the bundle . . . . .25%
Tube bundle in two-phase ﬂow . . . . . . . . . . .40% s Figure 3. Condenser shell diameter dependence on operating pressure. Literature Cited DAVID GREENE (35 Sycamore Way, Warren NJ 07059; Phone: (908)
7578183; E-mail: firstname.lastname@example.org) was process director for Aker 1. Heat Exchange Institute, “Standards for Steam Jet Vacuum
Systems,” 4th ed., HEI, Cleveland, OH (1988). 2. Peress, J., “Estimate Emissions from Vacuum Operations,”
Chem. Eng. Progress, 98 (5), pp. 40–42 (May 2002). 3. Tubular Exchanger Manufacturers Association, “Standards
of the Tubular Exchanger Manufacturers Association,” 8th ed.,
TEMA, Tarrytown, NY (1999). 4. Hewitt, G. F., et. al., “Process Heat Transfer,” Begell House
Publishers, New York, NY (1994). 5. Bell, K. G., and M. A. Ghaly, “An Approximate Generalized
Method for Multicomponent/Partial Condensers,” AIChE
Symposium Series, No. 131, Vol. 69, pp. 72–79 (1973). 6. McAdams, W. H., “Heat Transmission,” McGraw Hill, New
York, NY (1954). 7. Kern, D. Q., “Process Heat Transfer,” McGraw Hill, New
York, NY (1950). 8. Lockhart, R. W., and R. C. Martinelli, “Proposed Correlation of Data for Isothermal Two-phase, Two-component Flow
in Pipes,” Chem. Eng. Progress, 45 (1), pp. 39–48 (Jan. 1949). Kvaerner Pharmaceuticals when this article was written. He had been
with Kvaerner for 25 years and was responsible for the process design
of biotechnology facilities for the past 20 years, including a wide
variety of industrial microbial and mammalian-cell-culture facilities for
clients in the U.S. and Europe. He holds a BSChE from the Univ. of
Rhode Island and an MBA from Northeastern Univ. He is a registered
Professional Engineer in Massachusetts and New Jersey and is a
member of AIChE.
GEORGE J. VAGO (45 Fairway Green, Fairfield, CT 06432; Phone: (203) 3673933; E-mail: email@example.com) is a consultant in heat-transfer
equipment design for the chemical process industries and has over 35
years of experience in the field of heat transfer. He held the position of
manager of heat transfer with Kvaerner E&C for 18 years. His experience
stretches into the design of various special types of heat exchangers for
chemical and power plant applications. He holds an MS in mechanical
engineering from the Technical Univ. in Prague and is a life member of
ASME, where he was active on the Heat Transfer Equipment Committee. CEP February 2004 www.cepmagazine.org 43 ...
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This note was uploaded on 12/29/2011 for the course CHE 128 taught by Professor Scott,s during the Fall '08 term at UCSB.
- Fall '08