U.S. patent application number 13/242801 was filed with the patent office on 2012-01-19 for compact, high-effectiveness, gas-to-gas compound recuperator with liquid intermediary.
This patent application is currently assigned to DOTY SCIENTIFIC, INC.. Invention is credited to F. David Doty.
Application Number | 20120012293 13/242801 |
Document ID | / |
Family ID | 40801524 |
Filed Date | 2012-01-19 |
United States Patent
Application |
20120012293 |
Kind Code |
A1 |
Doty; F. David |
January 19, 2012 |
Compact, high-effectiveness, gas-to-gas compound recuperator with
liquid intermediary
Abstract
A liquid-loop compound recuperator is disclosed for
high-.epsilon. heat exchange between a first shell-side fluid
stream and a second shell-side fluid stream of similar thermal
capacity rates (W/K). The compound recuperator is comprised of at
least two fluid-to-liquid (FL) recuperator modules for transfer of
heat from a shell-side fluid, usually a gas, to an intermediary
tube-side heat transfer liquid (HTL). Each FL module includes a
plurality of thermally isolated, serially connected, adjacent
exchanger cores inside a pressure vessel. The cores are rows of
finned tubes for cross-flow transfer of heat, and they are arranged
in series to effectively achieve counterflow exchange between the
HTL and the shell-side stream. The HTL may be water, an organic
liquid, a molten alloy, or a molten salt.
Alumina-dispersion-strengthened-metal fins, superalloy tubes, and a
lead-bismuth-tin alloy HTL may be used for high temperatures.
Cumene may be used as the HTL in cryogenic applications.
Inventors: |
Doty; F. David; (Columbia,
SC) |
Assignee: |
DOTY SCIENTIFIC, INC.
Columbia
SC
|
Family ID: |
40801524 |
Appl. No.: |
13/242801 |
Filed: |
September 23, 2011 |
Related U.S. Patent Documents
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Application
Number |
Filing Date |
Patent Number |
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12673974 |
Feb 19, 2010 |
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PCT/US08/67008 |
Jun 13, 2008 |
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13242801 |
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61034148 |
Mar 5, 2008 |
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61016247 |
Dec 21, 2007 |
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Current U.S.
Class: |
165/182 |
Current CPC
Class: |
F28D 15/00 20130101;
F28D 7/08 20130101; F28D 1/05316 20130101; F28D 1/0477 20130101;
F28D 7/163 20130101 |
Class at
Publication: |
165/182 |
International
Class: |
F28F 1/12 20060101
F28F001/12 |
Claims
1. A method for heat exchange between a first shell-side fluid
stream at mean pressure p.sub.1 and a second shell-side fluid
stream at mean pressure p.sub.2, said method using a first set of
serially connected thermally isolated cross-flow exchanger cores
for transfer of heat between an intermediary tube-side heat
transfer fluid (HTF) and the first shell-side stream, a second set
of serially connected thermally isolated cross-flow exchanger cores
for transfer of heat between the HTF and the second shell-side
stream, said HTF characterized as being substantially liquid phase
throughout all cores and having critical temperature not less than
370 K, wherein a core is characterized as comprising at least one
row of finned tubes, said finned tubes are further characterized in
that the length of the tube fins per row in the shell-side flow
direction is typically less than 80 mm and the fin pitch is
typically less than 8 mm.
2. The method of 1 further characterized as having more than 4
thermally isolated cores exchanging with each shell-side stream and
having effectiveness .epsilon. greater than 60% at design operating
conditions.
3. The method of 1 in which said shell-side fluids are further
characterized as selected from the set comprised of organic liquids
having viscosity greater than 1 cP at 310 K and gases at pressure
greater than 0.05 MPa.
4. The method of 1 where said HTF is further characterized as
having flow rate G.sub.L kg/s, specific heat C.sub.PL J/kg-K, and
W.sub.L=G.sub.LC.sub.PL, said first shell-side fluid has flow rate
G.sub.1, specific heat C.sub.P1, and W.sub.1=G.sub.1C.sub.P1, said
second shell-side fluid has flow rate G.sub.2, specific heat
C.sub.P2, and W.sub.2=G.sub.2C.sub.P2, said geometric mean
shell-side conditions defined by W.sub.S=(W.sub.1W.sub.2).sup.0.5,
said tube-side conditions further characterized in that
W.sub.L>0.7 W.sub.S and W.sub.L<1.4 W.sub.S.
5. The method of 1 further characterized in that said HTF is
selected from the set comprised of water, organics, molten alloys,
and molten salts and is further characterized as having F.sub.D
greater than 2E5 J.sup.2/(s-m.sup.4-K.sup.2-cP) at the mean
operating temperature, where F.sub.D=k.sub.t.rho.C.sub.P/.mu.,
where k.sub.t is in W/m-K, .rho.C.sub.P is in J/m.sup.3-K, and .mu.
is in cP.
6. The method of 1 further characterized in that said tube-side HTF
has F.sub.D greater than the lesser of the F.sub.D of either of the
shell-side streams by more than a factor of 10 at mean operating
conditions.
7. The method of 5 further characterized in that each of said
shell-side streams has F.sub.D less than 2E5
J.sup.2/(s-m.sup.4-K.sup.2-cP) at the operating conditions.
8. The method of 1 further characterized as including a plurality
of liquid pumps and liquid reservoirs for circulation of a
plurality of HTFs.
9. The method of 1 in which said HTF is further characterized as
substantially selected from the set comprised of polyphenyl ethers,
polyol esters, polyalphaolefins, phosphate esters, phthalates,
silicones, fluorocarbons, polymer esters, organic liquid mixtures
that include alkylated polynuclear aromatics, and engine oils.
10. The method of 9 further characterized as including a liquid
reservoir with overhead gas space, said gas having H.sub.2 partial
pressure greater than 0.01 MPa, O.sub.2 partial pressure less than
1 kPa, H.sub.2O partial pressure less than 10 kPa, and total
pressure greater than 0.15 MPa.
11. The method of 1 in which said HTF is further characterized as
substantially comprised of a lead-bismuth-tin alloy.
12. The method of 1 further characterized in that the mean pressure
in said HTF is between 50% and 200% of the mean of p.sub.1 and
p.sub.2.
13. The method of 1 wherein one of said shell-side fluids is
further characterized as an organic solvent containing a dissolved
gas that effervesces when the fluid is heated, and means are
included between cores for separating the effervesced gas from the
liquid.
14. The method of 1 wherein one of said shell-side fluids is
further characterized as a gas containing a vapor that condenses
when the fluid is cooled, with means for draining the condensed
liquid from a core.
15. The method of 1 in which said HTF is further characterized as
an organic liquid, and means are included for separation of
reaction products from said HTF.
16. The method of 1 further characterized in that p.sub.2 is
greater than 3p.sub.1 and the typical fin pitch in said second set
of cores is less than 70% of the typical fin pitch in said first
set of cores.
17. The method of 1 further characterized as including transverse
passages between thermally isolated cores to equilibrate shell-side
pressures across the faces of said cores, wherein cores are
considered thermally isolated if fewer than 30% of the fins are
continuous between adjacent cores in the shell-side flow direction
and the tube pattern is not interleaved between adjacent cores.
Description
CROSS-REFERENCE TO RELATED APPLICATIONS
[0001] This application is a divisional application of U.S.
application Ser. No. 12/673,974 filed Jun. 13, 2008. U.S.
application Ser. No. 12/673,974 claims the benefit of U.S.
application No. 61/016,247 filed Dec. 21, 2007, and the benefit of
U.S. application No. 61/034,148 filed Mar. 5, 2008. Each of the
above-mentioned patent applications is incorporated herein by
reference for all purposes.
FIELD OF THE INVENTION
[0002] The field of this invention is heat exchangers, and more
particularly, compact, gas-to-gas recuperation at high
effectiveness for clean gases of similar heat capacity rates using
compound recuperators with liquid intermediary.
BACKGROUND OF THE INVENTION
[0003] Gas-to-gas recuperation with both high thermal effectiveness
and order-of-magnitude improvement in cost effectiveness is
critical to addressing global energy needs, as shown in at least
two co-pending patent applications. From a manufacturing
perspective, the challenges arise from the fact that it is not
practical to produce heat exchangers with closely-spaced fins on
both the inside and outside of tubes, and alternative approaches
thus far have had limited success.
[0004] An enormous number of heat exchangers have been well
optimized for numerous purposes over the past four decades.
However, most have not been directed at high thermal effectiveness
.epsilon. for cases where the heat capacity rates in the two
streams are similar. The heat capacity rate W of a stream is given
by GC.sub.P (its SI units are W/K), where G is the mass flow rate
(kg/s) and C.sub.P is the specific heat (J/kg-K). By the standard
definition of e (the ratio of heat transferred to the theoretical
limit), high .epsilon. is most easily achieved when W.sub.min (that
of the weaker stream) is much less than W.sub.max (that of the
stronger stream). However, exergy destruction can be minimized only
if W.sub.min is close to W.sub.max. The terms "recuperator" and
"regenerator" have usually implied the streams have similar W's,
and that will be the usage and regime of primary focus in this
invention. However, the streams need not be in the same state--one
may be liquid while one is a gas.
[0005] Common examples of cost-effective heat exchangers with high
exergy loss include automobile radiators and air-conditioning
condensers. In the automobile radiator, for example, the warmed air
leaves at a temperature much below that at which the hot water
enters. Thus, most of the water's exergy (energy availability) has
been destroyed, irrespective of precisely how one chooses to define
it. Other examples of cost-effective compact exchangers for
unrelated purposes include micro-channel, compact, fluid cooling
systems, as seen for example in U.S. Pat. No. 6,907,921.
[0006] The subset of fluid heat exchangers directed at high
.epsilon. have mostly addressed one of the following cases:
condensing-vapor-to-liquid, condensing-vapor-to-gas,
boiling-liquid-to-liquid, boiling-liquid-to-gas, liquid-to-gas, or
liquid-to-liquid. In all of these cases, the fluid thermal
conductivities, k.sub.t, W/m-K, are fairly large on at least one
side (generally over 0.2 W/m-K), or phase change is present to
drive small-scale turbulence on one side. A common gas-to-gas
exchange application is in steam power-plant superheaters. However,
the steam here has high thermal conductivity and rather high
density (for example, 0.067 W/m-K and 40 kg/m.sup.3 at 10 MPa, 650
K). Moreover, high .epsilon. there is not an objective, as the flue
gas will be used subsequently for boiling. Gas-to-gas exchange is
also sometimes seen in air preheaters in steam power plants. Here,
moderately high .epsilon. may be seen, though usually the minimum
flue-gas exhaust temperature is .about.400 K to limit corrosion
from acid condensation, and this limits .epsilon. of these
recuperators.
[0007] Achieving high .epsilon. in gas-to-gas exchange with low
pumping power has been challenging because volumetric specific
heats are much lower than seen in liquids and thermal
conductivities are usually low. Challenges are also seen in
achieving high .epsilon. in recuperators for organic liquids just
above their pour point where viscosity is quite high.
[0008] Doty, in U.S. Pat. No. 4,676,305, disclosed a compact method
of achieving highly effective recuperation with low pressure drop
for gases of similar W's. However, this microtube recuperator has
not yet been shown to be commercially competitive with the brazed
plate-fin type, in wide usage in recuperated open Brayton cycles in
the 301250 kW range and occasionally up to 25 MW. See, for example,
the microturbines available from Capstone Turbines Corporation, of
Chatsworth, Calif. These too have limited cost effectiveness and
limitations in accommodating applications where there are large
pressure differences (greater than .about.0.7 MPa) between the two
streams at high temperatures (above .about.750 K).
[0009] Optimized, compact high-.epsilon. gas-to-gas recuperators
require low flow velocities (several percent of the sonic
velocity), total flow-path exchange lengths in the range of 0.1 to
2 m, and passage hydraulic diameters of 0.5 to 8 mm, with the
larger diameters corresponding to pressures near 0.1 MPa and the
smaller sizes corresponding to pressures above 0.5 MPa. They have
also required the use of construction materials having fairly low
thermal conductivity, though that is not required in this
invention.
[0010] An alternative to paralleling tens of thousands of
microtubes that has seen rather little usage but appears to be the
most competitive for some compact recuperation applications is the
rotating honeycomb regenerator, as used in some turbine engines
where system mass is critical. Oda et al in U.S. Pat. No. 4,304,585
disclose an early ceramic design. Regenerators have seen very
little usage largely because of the difficulties in obtaining
adequate isolation between the high-pressure and low-pressure
streams and because of the shedding of ceramic particles, leading
to turbine abrasion.
[0011] Ceramic is usually selected for honeycomb regenerators in
recuperated aero-turbine applications because of the need for
oxidation resistance at high temperatures and the advantage of low
thermal conductivity in the flow direction. Rotating ceramic
honeycomb regenerators have demonstrated .epsilon. above 98% while
the brazed plate-fin recuperators seldom achieve more than 87%
.epsilon., primarily because of cost and mass optimization
constraints. The honeycomb regenerators can be an order of
magnitude more compact and an order of magnitude less costly for a
given exchange power and .epsilon. than plate-fin microturbine
recuperators--which are an order of magnitude more compact than the
gas-to-gas exchangers seen in most current chemical engineering and
power generation applications.
[0012] Oxidation resistance is irrelevant in some applications, and
there honeycomb regenerators can be made at lower cost and with
much higher reliability from a low-conductivity alloy honeycomb,
such as silicon bronze, stainless steel, or some magnesium or
aluminum alloys. The thermal conductivity of silicon-nickel-bronze
can be below 40 W/m-K, and 120 W/m-K is sufficiently low except
perhaps for the most compact applications. For example, a magnesium
alloy with thermal conductivity .about.90 W/m-K has been used
experimentally in a helicopter turboshaft engine. Titanium alloys
would be better, and their relative cost should decrease over the
next decade. The much higher thermal stress tolerance of metals
compared to ceramics is extremely beneficial with respect to
durability, as thermal stress is a primary factor limiting ceramic
regenerator design and contributing to shedding of particles from
ceramic regenerators.
[0013] Regenerator cost for a given performance is typically near
minimum when pore diameters are about 0.7 mm for many mobile
gas-gas exchange applications. The relevant design theory, well
understood for more than three decades, has recently been reviewed
and updated by David G Wilson in "Design and Performance of a
High-Temperature Regenerator Having Very High Effectiveness, Low
Leakage and Negligible Seal Wear", paper GT 2006-90096, Turbo-Expo
2006. The use of a metal for the honeycomb, possibly with the
innovations in Wilson's U.S. Pat. No. 5,259,444, may permit a
satisfactory solution of the sealing and wear problems in larger
recuperators where the pressure difference between the two streams
is small.
[0014] However, the rotating honeycomb regenerator still has
substantial limitations, either where there are substantial
pressure differences between the two streams, or where the size is
small (below .about.100 kW), or where the lower-pressure stream is
above .about.0.4 MPa. This last condition leads to greater
difficulties in limiting leakage and carry over, and it leads to
unreasonably low porosity requirements (or high solidity) in the
honeycomb (for sufficient thermal storage). High solidity
exacerbates axial thermal conduction losses and makes the
regenerator more massive and perhaps more prone to stress-related
failure. When two or more of the above conditions are present
simultaneously, the honeycomb suffers markedly.
[0015] High-.epsilon. recuperators are essential in many cryogenic
processes. A common and extremely effective design in cryocoolers
uses micro-multi-port (MMP) tubing with one of the gases flowing in
one direction through some of the "ports" (passages) and the other
flowing in the opposite direction through the other ports. The
viscous losses in very long lengths (4-20 m) of microtubes (under 1
mm ID, inside diameter) are often fully acceptable for gases at
very high pressures (over 1 MPa) and low temperatures (below 140
K). Many cryogenic recuperators operate at such conditions, where
outstanding counterflow recuperators can be made from MMP tubing or
a similar construction. For many cryogenic applications outside the
above conditions, the novel compound recuperator presented herein
will be superior.
[0016] The basis for the innovation presented herein begins by
learning from the highly developed liquid-to-gas exchangers best
exemplified in air-conditioning (AC) condensers and automobile
radiators. To achieve the high .epsilon. sometimes needed in
gas-to-liquid recuperation, it is simply necessary to arrange 5 to
30 of such exchangers in series, with the liquid flowing serially
from the first to the last and the gas also flowing serially, but
from the last to the first. Such a counterflow exchanger can be an
order of magnitude less massive and less costly than conventional
shell-and-tube gas-to-liquid counterflow exchangers of similar flow
rates, pressure drops, and .epsilon..
[0017] The dry-air condensers ubiquitous in AC condensers have been
extremely well optimized by numerous air-conditioning companies
over the past four decades. For example, "80-ton" (280 kW of
cooling) air conditioners are widely produced. The air-flow passage
lengths in these condensers are often under 3 cm per row of tubes;
and air-passages, though perhaps several centimeters wide,
typically have thicknesses of .about.1.5 mm. This corresponds to a
hydraulic diameter of .about.3 mm for the air flows, which
interestingly is that predicted to be optimum at 0.1 MPa by the
alternative analysis presented by Doty in U.S. Pat. No. 4,676,305.
The condenser in such a unit typically rejects about 350 kW.sub.T
at a .delta.T (dry air) of about 10.degree. C. Some large
commercial freezer systems utilize refrigerant R744, CO.sub.2,
where condenser pressures can exceed 6 MPa, so clearly high
tube-side pressures can be accommodated by cross-finned tubes
produced by automated manufacturing processes as used in AC
condenser cores. These exchanger cores are usually intended for use
with two-phase flow tube-side over a significant portion of their
length. However, predominately single-phase tube-side liquid flow,
as seen for example in U.S. Pat. No. 3,922,880 in a design for use
in an AC unit, can also be very cost effective.
[0018] In U.S. Pat. No. 4,831,844 Kadle discloses that for
condensing two-phase tube-side flow, substantial improvement is
obtained by a step-down approach in which the tube-side vapor flow
begins in two parallel tubes and then combines to a single tube
about two-thirds of the way through the condensing process. Several
advantages are noted for many AC applications, but the drawings
therein also appear to show interleaved tube-side flow between
parallel rows of finned tubes. Both step-down and interleaved
tube-side flow would generally be disadvantageous for single-phase
tube-side liquid flow in high-.epsilon. exchange, as addressed
herein; but with such patterns avoided, common AC condenser cores
may be utilized for high-.epsilon. recuperation.
[0019] Another common approach to improving tube-side heat transfer
with a low-velocity liquid of low k.sub.t is to use MMP tubing for
the liquid-phase flow, as discussed by Guzowski et al (IMechE 1999,
C543/083) and Guntly et al (U.S. Pat. No. 4,998,580). A simpler
method is to insert turbulators, such as open-pitch coil springs
inside the tubes. This can be quite beneficial with single-phase
tube-side flow of liquids under certain conditions.
[0020] The solution for order-of-magnitude improvement (compared to
shell-and-tube exchangers) in cost-effective high-.epsilon.
recuperation between clean gases shell-side at moderate
temperatures and liquids tube-side is to simply use a series
arrangement of several AC condenser cores (of proper design),
serially connected inside a pressure vessel. As seemingly obvious
and advantageous as the above approach is for high-.epsilon.
gas-liquid exchange, it does not appear to have been practiced as
such--liquid-only tube-side flow through a series of thermally
isolated cores. Related exchangers, in which the shell-side gas
goes cross-wise back and forth several times over the length of
cross-flow tubes, are commonplace; and often the tubes have fins
(though usually spaced 3 to 15 mm). However, the above differences
are of enormous importance with respect to manufacturing,
compactness, and cost effectiveness.
[0021] Single- and multi-row cores similar to what are suitable for
a component in the instant invention are produced by Armstrong
under the product name Duralite.TM. Plate Fin Coils. But apparently
the value of thermally isolated serially connected cores inside a
pressure vessel has not previously been appreciated as optimum to
achieve high-.epsilon..
[0022] Perhaps series arrangements of thermally isolated cores
similar to those used in AC condensers have not been considered for
high-.epsilon. gas-liquid exchange because most large applications
also require dealing with moisture, acids, and particulates in the
gas stream. For many such cases, available shell-and-tube
exchangers, developed primarily for condensing shell-side steam,
with typical tube diameters of 12 to 50 mm and shell-side fins
usually spaced .about.6 mm, may be the best option, especially when
the gas pressure is below 0.12 MPa and high .epsilon. is not
desired.
[0023] The heat pipe is in some sense related to the compound
exchanger disclosed herein, as it too uses an intermediary fluid.
However, the heat pipe uses a self-pumped two-phase fluid
tube-side, and it is poorly suited to gas-gas recuperation. So the
relationship to the heat pipe is tenuous at best. A complex,
finned, device cooler shown in U.S. Pat. No. 7,296,619 may
incorporate heat pipes, though that document tries to distort and
confuse the standard meaning of "heat pipe". Regenerators are also
somewhat related, as they utilize an intermediary, but there it is
a solid.
[0024] The standard air conditioner is most closely related to the
inventive compound recuperator, as it too provides heat transfer
between two gases using a fluid intermediary. There, however, the
large majority of the heat transfer in each exchanger includes
phase change, and a very energy-intensive vapor pump is required.
It is possible that some air-to-air recuperators for heat recovery
in buildings have utilized proprietary concepts somewhat related to
those presented herein, but apparently all such have relied upon
phase change in the fluid intermediary for most of the heat
transfer, and there is no evidence that they have achieved high
.epsilon..
[0025] Tube-side phase change has previously been desired because
it greatly increases tube-side heat transfer coefficient, h.sub.t,
W/m.sup.2-K, and thus generally allows significant reduction in
exchanger size. However, phase change is not desired in the instant
invention, as it makes minimization of irreversibilities
impractical (because it requires a very large number of
intermediary loops). The instant invention allows for enormous
reduction in exchanger size without phase change, and thus it also
readily permits high .epsilon.. Not surprisingly, commonly used
"refrigerants" are the worst type of fluids that can be imagined
for the applications envisaged by the instant invention.
[0026] It is noteworthy that the chemical engineering process
simulation software we have evaluated is not capable of handling
the case where a tube-side liquid stream is being heated by gas in
a cross-flow finned-tube exchanger, as seen in the instant
invention.
[0027] Two co-pending patent applications disclose enormous,
emerging applications for high-.epsilon. low-cost recuperation
between clean gases where good solutions are not currently
available: (A) where the hot gas stream enters above 550 K and at
more than 0.2 MPa, especially if the pressure difference between
the streams exceeds 1 MPa, (B) where some liquid condensation or
frosting can be expected in one or both of the gas streams, and (C)
where both gases are at pressures below 1 MPa, the pressure
difference exceeds 0.1 MPa, the temperatures are above 90 K, and
cross-contamination must be avoided. There also appears to be an
enormous, emerging application for high-.epsilon. low-cost
recuperation between viscous organic liquids. The invention
presented herein addresses these and many other situations most
optimally.
[0028] The instant invention is, in practice, usually implemented
as a minimum of two separate modules with one or more liquid
intermediary loops between them. Naturally, each independent module
is usable as a fluid-to-liquid recuperator, where the shell-side
fluid is usually a gas but may be a viscous liquid of low thermal
conductivity.
Relevant Art
[0029] 1. M M Guzowski, F F Kraft, H R McCarbery, J C Noveskey,
"Alloy and Process Effects on Brazed Automotive Condenser Tubing",
http://www.ent.ohiou.edu/kraft/VTMS4paper.pdf, presented at IMechE
1999, C543/083. [0030] 2. F D Doty, G Hosford, J B Spitzmesser, and
J D Jones, "The Micro-Tube Strip Heat Exchanger", Heat Transfer
Engr., 12, 3, 31-41, 1991. [0031] 3. DG Wilson and J Ballou,
"Design and Performance of a High-Temperature Regenerator Having
Very High Effectiveness, Low Leakage and Negligible Seal Wear",
paper GT 2006-90096, Turbo-Expo 2006, Barcelona. [0032] 4. Trane
Product Literature, "Installation, Operation, Maintenance: Series
R",
http:www.trane.com/webcache/rf/rotary%20liquid%20chillers%20(rlc)/service-
/rtaa-svx01a-en 09012005 pdf, RTAA-SVX01A-EN, 2005. [0033] 5. F P
Incropera and D P Dewitt, "Introduction to Heat Transfer", Wiley,
NY, 2002. [0034] 6. R K Shah, A D Kraus, D Metzger, "Compact Heat
Exchangers", Hemisphere Pub., NY, 1990. [0035] 7. L R Rudnick,
"Synthetics, Mineral Oils, and Bio-based Lubricants: Chemistry and
Technology", CRC, Boca Raton, 2006. [0036] 8. K Weissermel, H J
Arpe, Industrial Organic Chemistry, 4th ed., Wiley, 2003. [0037] 9.
C H Bartholomew and R J Farrauto, Industrial Catalytic Processes,
Wiley, 2006. [0038] 10. E Prabhu, "Solar Trough Organic Rankine
Electricity System (STORES)", NREL/SR-550-39433,
http://www.nrel.gov/docs/fy06osti/39433.pdf, 2006. [0039] 11.
DESIGN II for Windows, Version 9.4, 2007, by WinSim Inc.,
documentation available from
http://www.lulu.com/includes/download.php?fCID=390777&fMID=810115.
[0040] 12. Armstrong Duralite.TM. Plate Fin Coils, product
information, Granby, Quebec, 2008,
http://www.armstronginternational.com/files/common/allproductscatalog/pla-
tefincoils.pdf
TABLE-US-00001 [0040] U.S. PATENT DOCUMENTS 3,922,880 December 1975
Morris 62/498 3,994,337 November 1976 Asselman et al 165/119
4,304,585 December 1981 Oda et al 65/43 4,645,700 February 1987
Matsuhisa et al 428/116 4,676,305 June 1987 Doty 165/158 4,831,844
May 1989 Kadle 62/507 5,259,444 September 1993 Wilson 165/8
5,435,154 July 1995 Nishiguchi 62/476 6,907,921 June 2005 Insley
165/170 6,957,689 October 2005 Ambros et al 165/41 7,225,621 June
2006 Zimron et al 60/651 7,296,619 November 2007 Hegde
165/104.33
U.S. Patent Application Publication
[0041] US 2006/0211777 9/2006 Severinsky
SUMMARY OF THE INVENTION
[0042] A liquid-loop compound recuperator is disclosed for
high-.epsilon. heat exchange between a first shell-side fluid
stream and a second shell-side fluid stream of similar thermal
capacity rates (W/K). The compound recuperator is comprised of at
least two fluid-to-liquid (FL) recuperator modules for transfer of
heat from a shell-side fluid, usually a gas, to an intermediary
tube-side heat transfer liquid (HTL). Each FL module includes a
plurality of thermally isolated, serially connected, adjacent
exchanger cores inside a pressure vessel. The cores are rows of
finned tubes for cross-flow transfer of heat, and they are arranged
in series to effectively achieve counterflow exchange between the
HTL and the shell-side stream. The HTL may be water, an organic
liquid, a molten alloy, or a molten salt.
Alumina-dispersion-strengthened-metal fins, superalloy tubes, and a
lead-bismuth-tin alloy HTL may be used for high temperatures.
Cumene may be used as the HTL in cryogenic applications.
BRIEF DESCRIPTION OF THE DRAWINGS
[0043] FIG. 1 illustrates schematically a multi-stage, liquid-loop,
compound recuperator.
[0044] FIG. 2 illustrates the preferred liquid routing for a
portion of a compound exchanger.
[0045] FIG. 3 is a perspective, cut-away view of a typical
fluid-liquid exchanger module.
[0046] FIG. 4 illustrates a typical, single-row, finned-tube
core.
[0047] FIG. 5 illustrates a serpentine pattern in a finned-tube
core.
[0048] FIG. 6 illustrates five thermally isolated series tubes.
[0049] FIG. 7 illustrates a radial-flow version of an FL
module.
DETAILED DESCRIPTION OF THE PREFERRED EMBODIMENT
[0050] FIG. 1 illustrates a 4.times.3 array of 12 liquid-gas
cross-flow exchanger cores with 2 liquid pumps and two different
heat transfer liquids as an example of a method of achieving
high-.epsilon. recuperation between two isolated fluids of low
thermal conductivity, gas-1 and gas-2, identified in the figure
using hollow lines. These fluids have mean thermal conductivity
less than 0.4 W/m-K (that of H.sub.2 at .about.720 K) and will
usually be gases with k.sub.t less than 0.06 W/m-K. Thus, for
improved clarity, they are generally referred to as gases herein,
though applications where these fluid streams would be viscous
organic liquids are seen in a co-pending patent application. Both
gas-1 and gas-2 are shell-side, sometimes also called "fin-side".
In this example, gas-1 is the hot source stream, and gas-2 is the
cold stream being heated to nearly the entry temperature of gas-1.
Often, the hotter gas will be at lower pressure than the cooler
gas, but the reverse relationship is also possible.
[0051] In the example of FIG. 1, there are four sets of exchangers
(A, B, C, D). The heat transfer liquids (HTLs) are identified in
the figure with heavy solid lines. Here, each is directed serially
through three cross-flow exchanger cores for each gas stream. The
HTLs are all tube-side.
[0052] In this example, gas-1 enters 1 fin-side into exchanger
labeled D1 at 760 K and exits 2 fin-side from exchanger B3 at 400
K. Gas-2 enters 3 fin-side into exchanger labeled A1 at 320 K and
exits 4 fin-side from exchanger C3 at perhaps 680 K. For such
temperatures with similar W's, .epsilon. would be about 78% by the
standard definition.
[0053] Here, each gas stream passes through 6 cross-flow exchanger
cores, three on each side of each liquid loop. In practice, this
will often be a minimum number, though it also depends on how one
defines a cross-flow exchanger core. For example, a typical AC
condenser "core" contains 2, 3, or 4 rows of finned tubes, often
connected serially. Hence, a typical, 3-row, serial "AC core" could
perform the three serial exchanges as shown in FIG. 1 for each side
of each loop. Herein, 3 rows of thermally isolated finned tubes,
serially connected, is considered to be three cross-flow exchanger
cores in series. For the rows to be considered thermally isolated,
it is required that the fin metal not be continuous from one row to
the next (at least on most of the fins) and that the tube flow
pattern not be interleaved--that is, that the tubes not return back
to a first row after leaving that row and going to a second row.
From a functional perspective, the rows may be considered to be
thermally isolated if the thermal conduction of the metal between
adjacent rows is less than twice the thermal conduction of the
fluids (the sum of the shell-side and tube-side) between the
rows.
[0054] For variety in presentation, a thermally isolated, serially
connected, cross-flow exchanger core may be referred to as a
"finned tube row". The complete serial group of tube rows in a
single HTL loop for one of the gases will be referred to as a "core
set". The core sets will be inside a pressure vessel to contain the
shell-side pressure, and often all the sets associated with the
first gas stream would be inside one pressure vessel, and those
associated with the second gas stream would be inside a second
pressure vessel. For example, sets B and D of FIG. 1 would normally
be inside one pressure vessel and sets A and C would normally be
inside a second pressure vessel. The pressure vessel with the cores
it contains may be referred to as a fluid-to-liquid (FL)
recuperator module or a gas-to-liquid (GL) recuperator module, as
the shell-side fluid will usually be a gas.
[0055] The combination of two FL recuperator modules coupled with
an intermediary HTL may be referred to as a liquid-loop recuperator
or a compound recuperator. At least one liquid pump 5 and surge
tank 6 are also required for each compound recuperator. FIG. 1
illustrates a dual-loop compound recuperator.
[0056] For minimization of .delta.T-related irreversibilities, the
thermal capacity rate W.sub.L=G.sub.LC.sub.PL of the HTL through a
core set in a compound recuperator should be close to the geometric
mean W.sub.S of the thermal capacity rates W.sub.1 and W.sub.2 for
the two shell-side gas streams, G.sub.1C.sub.P1, and
G.sub.2C.sub.P2,
W.sub.L.about.W.sub.S=(W.sub.1W.sub.2).sup.0.5. [1]
Moreover, the ratio W.sub.1/W.sub.2 should be fairly close to 1,
though the compound recuperator will also be advantageous for other
conditions. Normally, W.sub.L would be between 0.7 W.sub.S and 1.4
W.sub.S. Of course, G.sub.L is proportional to n.rho.vd.sup.2,
where n is the number of parallel tubes in a core, .rho. is the
fluid density, v is the flow velocity, and d is the tube inside
diameter (mm).
[0057] The practical .epsilon. limit (for similar W's) is
essentially determined by the total number of rows, n.sub.r, of
isolated, serially connected, finned tubes (or cores) and the
"number of heat transfer units", NTU, where
NTU=h.sub.tSA.sub.X/W.sub.S, [2]
where A.sub.X is the heat transfer surface area. The .epsilon.
suggested for FIG. 1 is probably above a cost-effective limit for
just 12 total cores with liquid intermediaries, though it is
certainly possible. On the other hand, with 16 cores per set, four
sets, and two liquid loops, a practical limit of about 94% would be
expected. The same practical limit would be expected with a single
liquid loop and 32 cores per set. Such a design would be preferred
when the temperature difference between the hot source gas and the
cold source gas is rather small, as this requires only one liquid
pump. Having multiple loops, as shown in the multi-stage compound
recuperator of FIG. 1, allows for the use of different HTLs in
different temperature ranges, which allows for improved performance
in large recuperators operating over a large temperature range.
[0058] FIG. 1, though drawn acceptably by diagramming conventions
and chosen for its clarity, does not convey flow details that
improve maximum practical effectiveness per tube row. The fluid
routing shown in FIG. 2 better conveys liquid routing details that
significantly improve effectiveness per stage. There, the liquid
enters each row from the same side relative to the shell-side flow,
which is always distributed across the face of the tube rows, as
indicated by using parallel gas-flow arrows. The object is to make
the direction of the thermal gradient along each row the same and
maintain a fairly uniform change in the gas temperature per row
across the face of each core. Such a tube-side flow pattern is
uncommon in AC condenser cores, as .epsilon. there is not so
important.
[0059] Suitable finned-tube AC condenser or evaporator cores,
though generally without the most optimum tube-side flow routing,
are readily available for efficient heat transfers at power levels
from about a hundred watts to tens of kilowatts, and heat transfers
of hundreds of megawatts can be handled just as cost effectively by
paralleling tens of thousands of suitable AC cores. AC condenser
cores are available at low cost because efficient production
methods have been so highly optimized from the competitive
pressures of high-volume manufacturing over the past four decades.
Accommodating high shell-side pressures is straightforward--one
simply places the assembly in a large pressure vessel with suitable
baffles, as seen in U.S. Pat. No. 4,676,305, for example, and
addressed later in more detail. Although AC condenser cores are
usually intended for operation near 310 K, they are sometimes
constructed using copper tubing with aluminum or copper fins brazed
on using a filler material having liquidus near 870 K. It is also
not too uncommon to use 90Cu-10Ni alloy C706 for the tubing with
copper fins. In larger sizes these cores typically use tubing of 9
to 13 mm diameter, and the fin pitch (center-to-center spacing) is
often under 2 mm. Fin length in the direction of air flow is
typically .about.25 mm per row, though sometimes up to 80 mm per
row. Fin pitch in the FL recuperator up to 8 mm may be desired if
the shell-side fluid is a very viscous liquid, such as an oil just
above its pour point.
[0060] Only minor modifications of available cores are needed to
permit operation to about 700 K at limited pressures with
non-oxidizing clean gases. Moreover, operation to 900 K in many
non-oxidizing conditions is possible by simply changing to an
alumina-dispersion-strengthened copper such as C15720 (0.4%
Al.sub.2O.sub.3, bal Cu) for the fins and to the common 70Cu-30Ni
alloy C715 for the tubing (.about.70 MPa yield strength at
.about.900 K for C715, compared to .about.750 K for alloy
C706).
[0061] The benefits of this approach may not be immediately clear
to those accustomed to evaluating heat transfer primarily on the
basis of surface area, as (A) changing from a conventional
shell-and-tube exchanger to a typical AC condenser core may
increase the shell-side surface area per volume by only a factor of
5 to 10 (from .about.200 m.sup.2/m.sup.3 to 1000 or perhaps even
2000 m.sup.2/m.sup.3), (B) the tube-side "compactness ratio" may
decrease by a factor of 2 or more, and (C) the heat has to be
transferred twice. What may be overlooked is that the shell-side
heat transfer coefficient, h.sub.t, W/m.sup.2-K, will also
typically increase by a factor of 5 to 10 because the passage
thicknesses are decreased, so the shell-side total benefit can be a
factor of 25 to 100. By selection of an optimum HTL and flow
velocity, the tube-side h.sub.t can easily be made over 30 times
(possibly even more than 200 times) that of most gases--usually
without adding tube-side turbulators. Hence, the novel
compound-exchanger can permit an order of magnitude improvement in
compactness compared to shell-and-tube exchangers for gas-gas
exchange (for comparable powers, flow rates, .epsilon., and pumping
losses) even though the heat has to be transferred twice.
[0062] The increased complexity associated with compound
recuperators using a liquid intermediary would not be justified
below some size threshold. That cutoff is dependent on many
variables, including desired .epsilon., temperatures, gas
compositions, the importance of mass reduction, cleanliness of the
gas streams, and gas pressure differences. It will also depend on
availability of appropriate finned-tube cores for the relevant
conditions, a factor that is likely to change markedly over time.
Even today, it seems that compound recuperators would be preferred
for many cases with non-oxidizing gases if .epsilon. above 70% is
desired at (A) temperatures below 700 K, (B) exchange power levels
above 20 kW, and (C) mean gas pressures above 0.05 MPa. Suitable
cores for competitive compound exchangers for a much wider range of
conditions should become available.
[0063] A few more comments are useful to help elucidate the value
and hence inventiveness of the instant invention. The shell-side
thermal specific conductance, W/kgK, under the typical shell-side
flow conditions (largely laminar) will be inverse with the square
of the pitch; but the mass will be nearly independent of the pitch
for a given core volume. Clearly, as materials become steadily more
expensive, there will be a strong incentive to minimize the pitch
to permit ever higher heat transfer per exchanger mass. Of course,
the shell-side pressure drop will increase inversely with the pitch
for a given flow velocity. However, most applications will be with
shell-side pressure well above 0.2 MPa, and the shell-side pumping
power losses will often be inverse with the square of the gas
density. Hence, smaller fin pitch than is generally seen in AC
condenser cores will often be optimum, as long as the shell-side
flow-section area A.sub.S (the frontal section area, not A.sub.X)
is kept large and the flow path length is kept short, as discussed
later in more detail.
[0064] Minimum channel thickness in most prior-art, compact,
high-effectiveness exchangers is ultimately limited by the need to
establish highly uniform flow. Hence, manufacturing tolerances
limit minimum spacing. Channel thickness tolerance is not as
critical in the instant invention because flow mixing can readily
occur between successive thermally isolated cores.
[0065] Current AC condenser practice (.about.2 mm fin pitch) is
probably close to optimum in the compound exchanger for mean gas
pressures of .about.0.3 MPa, mean k.sub.t of .about.0.04 W/m-K, and
.epsilon. of 75-90%. The fin pitch can often be reduced (for
further reductions in exchanger mass and cost) at higher gas
pressures or low temperatures. However, there are limits, as the
fin thickness must be sufficient to provide the needed thermal
conduction and stiffness, and corrosion lifetime may be an issue in
some cases. A strong advantage of the liquid-loop compound
exchanger compared to regenerators is that the gas-passage
thickness of the high-pressure gas stream may readily be made much
less than that for the low-pressure gas stream, as desired for
maximum performance.
[0066] It is not uncommon in AC condensers for the tubes to have
internal features such as ribs, fins, or undulations to increase
h.sub.tL, though this adds to tubing cost and increases stress
concentrations. Such surface enhancement is mostly beneficial in
the initial portion of a condenser where no condensation is
occurring (the tube-side vapor is still superheated) and in the
final portion (subcooling) where the liquid velocities are very
low. Surface enhancements are much less beneficial in the compound
recuperator because the tube-side flow is liquid-only, of
essentially constant velocity, which may be better optimized.
[0067] Heat-Transfer Liquids (HTLs). The primary requirements in
the HTL are chemical stability at the relevant conditions, low
viscosity, low vapor pressure, high thermal conductivity, fairly
low cost, low health hazard, and high autoignition temperature
(AIT). It is also beneficial to have freezing point above the
minimum start-up temperature, though thaw-out measures can be
implemented. The AIT is also of minor importance, as inert or
reducing-gas pressurization over the HTL would normally be
incorporated; but it is still of some concern, should liquid leaks
develop. Water, organic fluids, molten alloys, or molten salts will
normally be selected, based mostly on the temperature range. Table
1 presents some pertinent data, some of which are estimates, on
some HTLs at 500 K. The column labeled "Risks" gives a single,
overall indication of the three hazards normally
considered--health, flammability, and reactivity.
[0068] The (turbulent-flow) tube-side heat transfer coefficient may
be calculated by,
h.sub.tL=B.sub.1G.sup.0.8k.sub.t.sup.0.6(C.sub.P/.mu.).sup.0.4
d.sup.-1.8
where .mu. is the dynamic viscosity (cP, centipoise, which is
identical to 1 mPa-s, or 0.001 kg/m-s) and B.sub.1 is a dimensioned
factor that is nearly constant over a wide range of conditions but
varies with surface features and other exchanger design details.
(Note: a fluid with .mu.=1 cP and .rho.=1000 kg/m.sup.3 has
kinematic viscosity, .mu./.rho., of 1 cSt, centistokes.) A few
simple manipulations and calculations are useful:
G.sup.0.8=B.sub.2(.rho.v).sup.0.8d.sup.1.6 [4]
h.sub.tL=B(.rho.v).sup.0.8k.sub.t.sup.0.6(C.sub.P/.mu.).sup.0.4d.sup.-0.-
2 [5]
F.sub.H=.rho..sup.0.8k.sub.t.sup.0.6(C.sub.P/.mu.).sup.0.4 [6]
-h.sub.tL=Bv.sup.0.8F.sub.Hd.sup.-0.2 [7]
where the B's are dimensioned constants, and F.sub.H is a
convenient, composite fluid property. A typical magnitude of B at a
Reynolds number of 10,000 to 20,000 inside smooth tubes is
.about.5.6, assuming the parameters are in the units shown above.
For 40 wt engine oil at 500 K, for v=10 m/s in tubes of 0.0077 m ID
(Re.apprxeq.15,000), this gives h.sub.tL.apprxeq.9000 W/m.sup.2-K.
For comparison, FP Incropera gives a representative value of
overall h.sub.t for air in cross-flow with water inside finned
tubes as .about.35 W/m.sup.2-K, and maximum overall h.sub.t of 6000
W/m.sup.2-K for steam condensers.
[0069] From eq. 7 and the above example, it might appear that one
simply needs to increase the HTL flow velocity to make h.sub.tL
very large compared to mean shell-side heat-transfer coefficient
h.sub.S (as desired for economic optimization), but of course that
consumes power--which increases almost as the cube of v. The
pumping power also increases with increasing .rho., .mu., and flow
length. Considering this, a better HTL figure of merit (composite
fluid property) for its selection than the above F.sub.H is the
following F.sub.M,
TABLE-US-00002 TABLE 1 HTL Properties at 500 K pour .rho., C.sub.P,
.mu., k.sub.t, F.sub.D Name point, K n.b.p. K AIT, K kg/m.sup.3
kJ/kg K cP W/m-K Risks F.sub.H F.sub.M kDt acetone 185 329 738 411
3.44 0.05 0.08 2 146 1430 2200 ethanol 200 352 636 475 3.43 0.06
0.09 1 164 1460 2400 butanol 210 380 699 581 3.78 0.094 0.10 1 179
1260 2300 water 274 373 -- 835 4.57 0.11 0.646 0 743 5130 22000
toluene 190 384 808 640 2.51 0.12 0.077 2 127 657 1030 cumene 130
425 697 661 4.57 0.15 0.108 2 188 1090 2220 ethylene glycol 260 470
673 935 3.16 0.34 0.2 1 221 669 1740 1-butylnaphthalene 260 561 800
824 2.1 0.35 0.093 2 106 269 450 Delo 100 30 wt 243 570 550 670 2.5
0.3 0.09 0 100 300 500 PAO, Delo 400 5W40 230 580 620 670 2.5 0.3
0.09 0 100 300 500 Delo 6170 40 wt 255 620 640 680 2.5 0.35 0.09 0
96 260 440 POE, Mobil 254 212 640 672 700 2.5 0.4 0.1 1 99 250 440
dioctyl phthalate 250 657 780 798 2.1 0.5 0.11 1 98 200 365
1-dodecyl-naphthalene 305 676 800 795 2.5 0.41 0.092 1 103 250 445
tri-o-cresyl phosphate 260 693 680 950 2.2 0.4 0.11 2 136 340 680
TBPP-100 phosphate 270 708 795 900 2.2 0.5 0.13 1 123 255 515
polyphenyl ether 5P4E 280 749 860 970 1.9 0.6 0.13 1 114 200 400
60NaNO.sub.3--40KNO.sub.3 480 870 870 1950 1.4 4.5 0.45 2 166 77
270 55Bi--45Pb 400 1800 -- 10000 0.15 2.7 4 2 1150 300 2220
38Pb--37Bi--25Sn 400 1900 -- 9000 0.18 2.5 8 1 1770 520 5200
F.sub.M=k.sub.t.sup.0.6(.rho.C.sub.P).sup.0.8/.mu. [8]
[0070] The combination of the need to achieve
h.sub.tL>>h.sub.S at moderate v and flow length, along with
good W matching, imposes constraints on the tube diameters and the
tube paralleling scheme. The HTL would usually have nearly constant
velocity throughout most of the cores, so tube diameters would be
nearly constant throughout. However, in cores containing several
parallel tubes, it may be beneficial for them to combine at their
core entrances and exits to simplify tubing inter-connections
between cores. Clearly, the HTL velocities in the interconnections
could be very different from the typical value in the cores.
[0071] As seen in Table 1, F.sub.M is low for organics compared to
that for water or molten alloys, but it is usually higher than that
of molten salts--a concept that has previously been misunderstood.
Other advantages of organics may include no freezing problems, no
metal erosion, lower corrosion, lower density, lower toxicity,
lower cost, lower viscosity, and simpler disposal problems.
Pressurized water can be used well beyond 500 K, but exchanger
costs are increased because of the very high stresses. An organic
of low vapor pressure is often better, though in some applications
lower-boiling fluids, such as ethanol or even acetone, could meet
the specific requirements and be preferred. Note that the relative
merits of the HTLs are temperature dependent.
[0072] Silicone fluids (such as Dow Corning 550, AIT of 755 K, but
not suitable for long term usage above 550 K) and low-grade
hydrocarbon (HC) mixtures, such as Exxon Calorie HT-43 (AIT of 627
K) have been used. Some more attractive organic fluids with n.b.p.
and AIT both above 660 K, pour point below 320 K, and acceptable
chemical stability and safety are: (A) polyphenyl ethers (PPEs,
aerospace lubricants and diffusion-pump oils, 5-ring type 5P4E has
AIT-880 K, n.b.p.=749 K, 290 K pour point, .DELTA.G.sub.t.about.2
kJ/g, non toxic, has been used in short-term vapor-phase
lubrication up to 870 K), (B) polyol esters (POEs, most type-2
aviation turbine oils, AIT usually .about.670 K, but AIT and n.b.p.
can be over 740 K), (C) polyalphaolefins (PAOs, a major component
in type 5W50 synthetic engine oil, 16 cSt at 100.degree. C., AIT
often .about.650 K, but AIT can be .about.700 K in heavy PAOs), (D)
phosphate esters (used in aviation hydraulic fluids), (E) phenyl
silicones, (F) fluorocarbons, (G) polymer esters (PEs), (H)
phthalates, and (I) mixtures of the above and high-boiling
alkylated polynuclear aromatics. See Table 1 for data on two
alkylated polynuclear aromatics.
[0073] Highly branched alkanes are preferred to n-alkanes in engine
lubrication applications, as they have much better oxidation
resistance, much lower viscosity for a given boiling point, and are
more resistant to dehydrogenation and cracking. The relative price
of such synthetic oils, similar to PAOs, should drop substantially
over the coming decade.
[0074] Inexpensive tin-lead alloys may be acceptable as an HTL at
high temperatures with stainless or superalloy tubing. The
solubility of iron in tin is about 0.1% at 650 K, and this may lead
to excessive exchanger erosion with low-alloy steel tubing (even
after the molten alloy becomes saturated with iron, as there will
be some thermal gradients in the liquid). The solubility of iron in
both bismuth and lead is at least an order of magnitude lower than
in tin. However, alloys of more than 50% bismuth expand upon
freezing (if not immediately, then after several days), and this
could produce unacceptably high stresses within the exchangers.
Lead-bismuth-tin alloys of relatively low tin content should be
fine with some low-cost steel alloys for the tubing. The
38Pb-37Bi-25Sn alloy shown in Table 1 has an excellent balance of
low iron solubility, low vapor pressure, low toxicity, high
F.sub.M, low cost, and low liquidus temperature, though perhaps
lower Bi and Sn contents with increased Pb and minor additions of
antimony (Sb) would give an even better balance.
[0075] Molten salts, especially mixtures of NaNO.sub.3, KNO.sub.3,
NaNO.sub.2, and Ca(NO.sub.3).sub.2, have often been used for HTLs.
Some have freezing points lower than those of some lead alloys, but
their upper temperature limits are lower. For example, KNO.sub.3
decomposes at 670 K and NaNO.sub.3 decomposes at 650 K, though some
mixtures, such as the eutectic listed in Table 1, have higher
stable temperature limits. There are some security risks, as all
can easily be used to make powerful explosives of limited
stability. Moreover, their NEPA health ratings are usually "2,
highly hazardous", and their F.sub.M is quite inferior to other
options. When hot, they react vigorously with most pump lubricants
and elastomeric seals, and they slowly attack many alloys at grain
boundaries. Another complication with salts, alloys, and heavy
polynuclear aromatics is that they are solid at room
temperature.
[0076] The thermal conductivities of the gases expected in some
emerging applications typically range from 0.04 to 0.06 W/m-K at
500 K (for CO, C.sub.4H.sub.10, air, and some H.sub.2/CO.sub.2
mixtures of interest), gas densities are often .about.5 kg/m.sup.3,
C.sub.P is often 1 to 3 kJ/kg-K, and .mu. is typically 0.01-0.03
cP. For gases under the relevant shell-side (substantially laminar)
conditions, a composite fluid property more useful than F.sub.M for
estimating the ease with which heat transfer may be achieved is
F.sub.G=k.sub.t(.rho.C.sub.P).sup.2/.mu. [9]
A useful expression for comparing liquid and gas heat transfer
fluids for similar flow geometries (same hydraulic diameters, flow
lengths, etc.) is (F.sub.MF.sub.G).sup.0.5, and this suggests the
heat-transfer challenges for the gases could be two to three orders
of magnitude greater than for liquids for similar geometries. A
simpler parameter, F.sub.D, for comparing diverse fluids is
included in Table 1 and discussed in the last section with
reference to applications for shell-side liquids. The important
point here is that there is usually little need to worry about
tube-side heat-transfer enhancement. This allows enormous
manufacturing simplifications. The focus needs to be primarily on
reducing passage thickness and increasing surface area on the
shell-side.
[0077] FL Module Implementations. As noted previously, the core
sets will be inside a pressure vessel sufficient for the shell-side
pressure. Often, all the sets associated with the first shell-side
stream (usually a gas) would be inside one pressure vessel, and
those associated with the second shell-side stream would be inside
a second pressure vessel. The pressure vessel with the cores it
contains is referred to as a fluid-to-liquid (FL) recuperator
module, and a typical embodiment is shown crudely in FIG. 3.
[0078] A typical FL recuperator module might contain 30
series-connected, thermally isolated finned-tube cores 31 (though
the figure shows just 8 cores for better clarity), each having
typical external dimensions of about 1 m.times.1 m.times.0.03 m.
The shell-side entrance and exit ports 32, 33 are normally at the
opposite ends from the tube-side entrance and exit ports, 34, 35. A
typical core is better illustrated in FIG. 4, though again not
likely to scale. Each 1 m.times.1 m core might typically have 40
parallel finned tubes 41, each of 8 mm ID and 10 mm OD, each
traversing the full width, with typical center-to-center spacing of
25 mm, and tube-side entrance and exit manifolds 42, 43. FIG. 4, on
the other hand, shows 20 tubes and 64 fins, which is closer to
being typical for a 30 cm.times.30 cm core, though even there the
number of fins would likely be greater than shown by a factor of 2
to 4.
[0079] If the fin thickness is 0.5 mm and the fin pitch 1.6 mm,
then the example shell-side flow-section area A.sub.S is about 0.7
m.sup.2 and the tube-side flow-section area A.sub.T is .about.0.002
m.sup.2. Hence, A.sub.S is .about.350 times A.sub.T. For the
typical core dimensions noted earlier (1 m.times.1 m.times.0.03 m),
the mean shell-side flow length L.sub.S is about 0.03 m per core,
and the mean tube-side flow-length L.sub.T is 1 m per core. Hence,
L.sub.T is about 30 times L.sub.S. Note that this ratio is
independent of the number of cores when they are serially
connected, as both flow lengths increase by the same factor. The
flow-section area ratio is also independent of the number of
serially connected cores. There may be a substantial gap between
adjacent cores, as depicted in FIG. 3, in the shell-side flow
direction for pressure equalization across the face of the cores
and some transverse mixing, but the shell-side flow is
substantially axial with respect to the pressure vessel.
[0080] The tube-side HTL flow is shown in FIG. 3 as being ducted 36
from an exit manifold on one side of a core diagonally across to
the entrance manifold on the next core. Note that the HTL enters
all the cores on the same side and exits all the cores on the
opposite side. The diagonal HTL ducting pattern is one way to
improve tube-side flow homogeneity. Other measures may also be
taken, and often a primary measure will be judicious selection of
the diameter of the tubes 41 such that the flow velocity within
them will achieve pressure drops that are large compared to the
pressure drop in the manifold while simultaneously meeting the
other previously noted requirements with respect to pumping power,
h.sub.t, and W.sub.L. Support structure for the cores is not shown,
though clearly some is needed. The flow-cage 37 for constraining
the shell-side flow within the cores is only partially shown. For
the conditions normally addressed, the shell-side volumetric flow
rate will generally be relatively high (especially when compared to
tube-side), so shell-side pressure drops from inlet 32 to outlet 33
must necessarily be low (to achieve low pumping power) and
differential stresses on the cage can easily be handled. The
pressure vessel would preferably have a burst pressure greater than
twice the mean shell-side relative pressure and generally much
greater than 0.3 MPa.
[0081] One may define an HTL conductance Y.sub.F [W] between
adjacent cores as
Y.sub.F=T.sub.dW.sub.L [10]
where T.sub.d is the mean temperature difference between adjacent
cores. Adjacent cores are herein considered to be effectively
thermally isolated if the heat conducted between cores through
solid materials is less than one-third of Y.sub.F. Such a condition
is not easily met if more than 20% of the fins are continuous from
one core to the adjacent core in the shell-side flow direction, but
such a condition is easily met if none of the fins are continuous
between adjacent cores and the tube pattern is not interleaved
between adjacent cores. However, adequate thermal isolation will
sometimes be possible if up to 30% of the fins are largely
continuous, except for holes for transverse pressure equalization,
between adjacent rows.
[0082] It is not necessary for all tube rows to be thermally
isolated. A "compound core" may have several rows of finned tubes
thermally coupled by continuous fins between them for improved core
robustness. However, the practical effectiveness limit is strongly
dependent on the total number of thermally isolated cores in
series. Hence, it will often be desirable for this number to be
more than 20, though there will be some cases when as few as two
thermally isolated cores per FL module are sufficient. It is
unlikely that a compound core would contain more than four rows of
thermally coupled rows of finned tubes. In most cases, each
thermally isolated core would be a single row of finned tubes, as
shown in FIGS. 3 and 4.
[0083] Shell-side flow homogeneity is also essential for high
.epsilon., at least when the W's are similar. In most cases,
allowing for pressure equalization across the faces, as is readily
achieved when none of the fins are continuous between adjacent
cores, will be sufficient, as shown in FIG. 3. In the prior art,
all the fins are usually continuous between adjacent rows, such as
seen in the Armstrong Duralite.TM. Plate Fin Coils products. A
minor fraction could still be continuous. In such a case, pressure
equalization across the faces of thermally isolated cores can
readily be achieved if holes or cut-outs are included in fins
joining adjacent cores.
[0084] With series-connected cross-flow exchangers, the flow
homogeneity may be further improved by inserting turbulent mixers
in the gas flow streams between cores. (This obviates the benefit
from the flow routing illustrated in FIG. 2, but is better than the
alternative of channeling--where, because of the viscosity
dependence on temperature, the shell-side velocity may become
higher than mean on one side of all the cores when the shell-side
gas is being heated a large amount in each core.) The use of
separate, series-connected FL modules further simplifies the
insertion of turbulent mixers into the shell-side stream.
[0085] FIG. 5 illustrates a portion of a core with a serpentine
pattern that may be desired to better meet the HTL velocity and
pressure-drop objectives in some cases. If the tube-side flow for
the core of FIG. 4 were instead handled by 10 parallel tubes of 8
mm ID, each traversing the full width 5 times in a serpentine
pattern with center-to-center spacing of 20 mm, the tube-side
flow-section area then would be 5E-4 m.sup.2. With a reasonable
allowance for the bends at each end, each tube may then need to be
.about.6 m long. In this case, L.sub.T would be about 200 times
L.sub.S, and A.sub.S would be 1400 times A.sub.T. As in FIG. 4, the
shell-side flow in FIG. 5 is normal to the plane of these
figures.
[0086] A method of arranging thermally isolated, serially connected
finned tubes without manifolds between them is shown in FIG. 6.
Such an arrangement may have manufacturing advantages in some
cases. An option is to stack a large number of serpentine finned
tubes normal to the plane of FIG. 6, with the shell-side flow as
indicated. For the shell-side flow direction assumed in FIGS. 4 and
5, the fins would normally be continuous between the tubes as shown
within a core. However, the fins could not be continuous between
the tubes for the flow direction shown in FIG. 6 and achieve
thermal isolation between these serially connected tubes.
Conventional usage may not refer to such an arrangement of a
single, serpentine finned tube as a "core". When a large number of
the serpentine finned tubes as shown in FIG. 6 are stacked normal
to the plane of the Figure, five thermally isolated planar cores
are effectively formed.
[0087] As discussed earlier, it will sometimes be preferable to
utilize more than one liquid loop. Hence, in some applications,
there may be two or even three liquid loops servicing the cores in
a single FL module. In some applications, it may be preferable to
utilize separate pressure vessels for the high-temperature cores,
mid-temperature cores, and low-temperature cores, and in such cases
in particular a small number of thermally isolated cores per FL
module may be sufficient. In larger applications, it will often be
desirable to arrange modules in parallel, as it may not be optimum
to manufacture modules larger than can easily be transported by
truck. Note that paralleling also does not affect the ratios
A.sub.S/A.sub.T or L.sub.T/L.sub.S, but the ratio A.sub.S/L.sub.S
steadily increases with capacity in an optimum design.
[0088] For very large modules, a hexagonal arrangement of the cores
as shown in FIG. 7 (depicted without the containment vessel) may be
preferred, as it permits a larger ratio of A.sub.S/L.sub.S within
practical trucking constraints. Here, the shell-side fluid flow is
generally radial, and the core arrangement of FIG. 6 is assumed,
though the arrangement of FIG. 4 could also be used. Similar
arrangements of finned-tube cores, except square rather than
hexagonal, are commonplace in the AC industry, where shell-side air
flow through a condenser exhausts to atmosphere. However, the
prior-art condensers (A) utilize tube-side phase change for most of
the enthalpy change, (B) are not enclosed in a pressure vessel, and
(C) may not include serially connected thermally isolated
cores.
[0089] For the hexagonal arrangement shown in FIG. 7, the
shell-side face flows are all functionally in parallel. Hence, the
tube-side flows must also be functionally in parallel. In other
words, all of the innermost cores would connect to the same HTL
port, and all of the outmost cores would connect to the same HTL
port. Obviously, pentagonal, octagonal, or other circumferential
arrangements of cores could also work well. The pressure vessel
would normally have its axis aligned perpendicular to the
shell-side flows through the cores.
[0090] Cores of significantly different characteristics may also be
combined, either in series or in parallel, with predictable
results, though the analysis is more complex. Clearly, many
variations in dimensions and patterns are possible, but generally
A.sub.S would be more than 100 times A.sub.T and L.sub.T would be
more than 10 times L.sub.S. Such ratios appear to be well outside
the prior art in multi-pass, finned-tube, shell-and-tube heat
exchangers.
[0091] Core Modifications for Severe Conditions. For high
performance at demanding conditions (high temperatures, oxidizing
atmospheres, or large differences in pressures between the two
gases), appropriate changes in choice of materials for the tubing,
fins, and braze are required. The tubing material is selected
primarily for yield strength at the required temperature,
formability, brazability, and corrosion resistance. The fin
material is selected primarily for thermal conductivity, cost,
corrosion resistance, melting point, and brazability. In some
cases, the fins have been pressed on rather than brazed on, but
this approach is less desirable for extremes of temperature, for
closely spaced thin fins, or if much vibration is likely.
[0092] Alumina-dispersion-strengthened copper, aluminum, or nickel
are particularly good choices for the fins, though cobalt and
alloys are also possible for high temperature fins. While most
superalloys have poor thermal conductivity compared to pure metals
near room temperature, some with superior oxidation and corrosion
resistance, such as Haynes 214 (16Cr, 4.5AL, 3Fe, 0.2Y, bal-Ni),
have fairly good thermal conductivity at high temperatures (32.4
W/m-K at 1255 K).
[0093] Some superalloys, such as Haynes 188 (38Co, 22Cr, 22Ni,
14.5W, 2Fe, 1 Mn, 0.3Si, 0.1C, 0.07La), have good brazability and
formability in the annealed state as well as outstanding oxidation
resistance and high-temperature strength (1400 K for 70 MPa yield
strength in alloy 188). An alloy similar to Haynes 188 would be
well suited for high-temperature exchanger tubing, though
modifications to reduce cost and improve formability and
brazability, particularly by reducing Co, W, and Cr, may be
preferred. The tubing material can have poor thermal conductivity
with little consequence on performance. If the hydrostatic pressure
on the HTL(s) is maintained near the mean of the pressures of the
gas streams, preferably within a factor of two of this mean, the
stresses on the tubing are reduced.
[0094] Brazes compatible with the higher temperatures and materials
are required. Nickel-plated dispersion-strengthened-copper fins
could be brazed to such using filler BNi-7 (890.degree. C.
liquidus, 85Ni, 14Cr, 10P). Superalloy or
alumina-dispersion-strengthened nickel fins could be brazed to
Haynes 188 or similar tubing using BNi-5 (1135.degree. C. liquidus,
70Ni, 19Cr, 10Si) for operation at still higher temperatures.
Methods for applying chromium platings to the finned tube rows can
be developed, based on the prior art.
[0095] Organic HTLs at High Temperatures. An organic HTL may be
used quite satisfactorily at a much higher temperature than that
for which it has normally been recommended if suitable measures are
taken. First of all, it is most important that the surfaces in
contact with the hot oil (tubing interiors, etc.) be catalytically
deactivated with a thin layer (0.1 micron is sufficient) of
coke--carbon and very heavy condensed polynuclear aromatics.
Thermal (non-catalytic) reactions require much higher temperatures
than catalytic, and most metallic or oxide surfaces have some
catalytic activity. Secondly, since water catalyzes reactions on
many metal surfaces, it is important to maintain the liquid
pressure well above the maximum external gas pressure (the greater
of the pressures in shell-side gas-1, gas-2, and ambient) at all
times to prevent ingress of air and moisture through minute leaks.
Of course, it is important to insure that any organic HTL is
initially de-gassed of dissolved O.sub.2 and H.sub.2O.
[0096] In general, there are four primary types of thermal
reactions that will dominate for most of the heavy HCs likely to be
used for a high-temperature HTL: cracking, dehydrogenation,
de-isomerization, and aromatic polymerization or condensation. All
but de-isomerization (conversion from a highly branched to a less
branched structure) are somewhat inhibited by moderate H.sub.2 and
CH.sub.4 concentrations within the HTL--or perhaps it is more
proper to say that high H.sub.2 and CH.sub.4 concentrations
increase the rate of the reverse of many undesired reactions.
[0097] As previously noted, a small surge tank is required for the
HTL to accommodate expansion and contraction. To extend the
lifetime and useable temperature limit, the gas overhead 7 in this
reservoir should have an H.sub.2 partial pressures of at least 0.01
MPa and possibly as much as 5 MPa, though excessive H.sub.2 partial
pressures will increase cracking (especially of n-alkanes) and
hydrogenation of aromatics into lower-boiling cyclics. Hence, it
may be desirable to also have significant methane partial pressure,
possibly as much as 15 MPa, as it is less reactive. For some HTLs,
such as water, glycols, phthalates, silicones, polyol esters, and
polyphenyl ethers, partial pressurization with argon and perhaps
N.sub.2 may be preferred. Maintaining an excessive total pressure
on the HTL increases the cost of the high-temperature cores and
exacerbates problems with dynamic seals, but an HTL static pressure
about 0.1 to 1 MPa above the higher of the shell-side gas pressures
would normally add little to the system cost.
[0098] The concentrations of CH.sub.4 and H.sub.2 dissolved within
the HTL are determined largely by their partial pressures and the
liquid temperature in the HTL reservoir. Solubilities of H.sub.2 in
HCs (A) are generally higher for alkanes than for aromatics, (B)
they increase with increasing temperature, (C) they approximate a
Henry's law behavior, and (D) they decrease slowly with increasing
molecular mass of the HTL. The solubility of H.sub.2, in moles
H.sub.2 per kg liquid per MPa, at 460 K are 0.068 and 0.044 for
hexadecane (C.sub.16H.sub.34) and tetralin (C.sub.10H.sub.12)
respectively, for example. Solubilities at 520 K are about 30%
higher. Solubilities in very heavy oils are about half that for
hexadecane. Methane solubility is much higher (by perhaps a factor
of 20 at 460 K) and much less dependent on temperature. When the
HTL cools during power-down, it may effervesce H.sub.2.
[0099] It will normally be preferable to have the liquid pumps at
the low-temperature points in the loops, as shown in FIG. 1, as
this simplifies problems associated with dynamic seals. It may also
be preferable to have the reservoir near the low-temperature point
in the loop to avoid H.sub.2 super-saturation within the HTL at its
cooler points in the loop, as super-saturation could lead to
hydrogen effervescence in the cooler exchangers and reduced heat
transfer. However, some level of H.sub.2 super-saturation is
usually quite stable in HCs, and this may further inhibit
production of coke precursors with some HTLs. Hence, it may be
preferable to have the surge tank at a higher temperature point in
the loop, even though this increases its cost a little.
[0100] Even with the above measures, operating at temperatures near
the upper practical limits will result in the production of
reaction products, both light and heavy, that are undesirable
beyond a certain level but are quite tolerable at low levels. In
most cases, this will simply mean that periodic HTL changes will be
required. For large installations there are other options. Cracking
produces light alkenes, some of which will be hydrogenated to light
gases such as C.sub.2H.sub.6, C.sub.3H.sub.8, and C.sub.4H.sub.10,
which are less than optimum for reservoir pressurization. An easy
way to deal with such is to continually, slowly vent some
pressurization gas and maintain the desired pressure with fresh gas
of optimum mixture. Of course, membranes and other separations
methods could be used to separate the vented gas into useful
product streams if desired. Some of the alkenes will alkylate with
other alkanes or aromatics to heavy HCs and coke precursors in the
HTL. One way to maintain the HTL at an acceptable composition is to
steadily bleed HTL from the reservoir and maintain the desired
level with fresh supply. Various separations methods could be
applied to the used fluid for reclamation. More examples of
reaction-product separation processes are disclosed in a co-pending
patent application on Dual-source Organic Rankine Cycles.
[0101] In summary, the following are required to operate with
organics at high temperatures:
1. Deactivate all surfaces in contact with the hot HTL. 2. Maintain
sufficient HTL pressure to preventingress of air and moisture
through minute leaks. 3. Maintain an optimum gas mixture
pressurizing the HTL. 4. Remove primary HTL reaction products
before they lead to excessive coking. 5. Select a fluid with high
chemical stability with optimum gas pressurization.
[0102] For the temperatures indicated in the example of FIG. 1 with
appropriate gas pressurization over the HTLs, the HTL for sets A
and B could be dioctyl phthalate, a PAO oil, or a POE oil. For sets
C and D, a molten alloy, a molten salt, PPE-5P4E, or possibly an
alkylated polynuclear aromatic could be used.
[0103] Cryogenic Applications. While some of the largest
applications may be at elevated temperatures in chemical processes
and power plants, there will also likely be enormous applications
at cryogenic temperatures, as very high .epsilon. in exchange
between gases is often required there. Moreover, gas viscosities
(and hence pressure drops) there are often so low that it is very
difficult to establish the uniform flow conditions that are
essential for high .epsilon.. As previously noted, the use of
separate, series-connected cores or FL modules allows simple
insertion of turbulent mixers into the shell-side stream between
modules.
[0104] For cryogenic applications, the fin pitch can be further
reduced--because (A) viscous losses are much smaller, (B) the
fin-metal thermal conductivity can be an order of magnitude higher,
(C) the gas thermal conductivity is often much lower, (D) the HTL
may have a higher F.sub.M, and (E) corrosion is more readily
controlled.
[0105] One HTL, cumene (isoproplybenzene, C.sub.9H.sub.12), is
listed in Table 1 that is particularly advantageous down to 130 K,
and others are suitable for lower temperatures. Propane, for
example is usable down to 90 K and is easily liquefied at room
temperature, as its critical temperature T.sub.C is 370 K. For
lower temperatures, gases with T.sub.C well below 300 K are
required, and this complicates start-up somewhat, as a rather large
compressed-gas reservoir is needed. Oxygen (T.sub.C=155 K) is an
excellent HTL for the range of 60-130 K. Fluorine oxide, F.sub.2O
(T.sub.C=215 K), is suitable for the 55-170 K range, and other
gases can be used over other, narrow ranges. For example, H.sub.2
(T.sub.C=33 K) can be used over the 15-30 K range. However, very
high pressures are required to condense these gases near the upper
ends of their maximum liquid ranges, and this increases exchanger
cost.
[0106] In principle, a gas could be used as the heat transfer
intermediary, where the obvious choice for the 35-60 K range would
be hydrogen. However, high (tube-side) h.sub.tL with low pumping
power cannot be achieved with a gas as the intermediary above its
T.sub.C, as its density is much too low at practical pressures. The
best way to increase h.sub.tL with gases is to use MMP tubing,
which indeed works beautifully at very high pressures.
[0107] The minimum competitive size of the compound exchanger for
cryogenic applications will be smaller than for most
high-temperature applications--because mass is often much more
critical and mean temperature difference between the gas streams
may need to be an order of magnitude smaller. The compound
exchanger seems likely to be preferred in many cryogenic
recuperators down to 90 K at exchange powers above 1 kW for gas
pressures below 0.5 MPa.
[0108] Compact Recuperator Variations. An advantage not yet noted
for the compound exchanger is that it can greatly reduce the
ducting costs in large plants where the heat generated in one
process is needed in another process hundreds of meters or even
tens of kilometers away. In such a case, it will sometimes be
easier to achieve optimum thermal balancing--matching the gas Wand
temperature to those of the HTL within each module--by splitting
and re-combining the HTL streams at numerous points. When streams
are combined, their temperatures should be similar for minimal
exergy destruction. A portion of an HTL may be split out from an
intermediate point in an exchanger module to exchange energy with
another process and then be re-combined at an appropriate point
where the temperature is similar.
[0109] Of course, it is not uncommon to transfer heat long
distances using either phase change (usually water) or liquids
(including many of those mentioned earlier as good HTLs for
compound recuperators). For example, Severinsky in US publication
2006/0211777 notes that it can be advantageous to transfer heat
throughout a large plant using a number of different phase-change
heat-transfer fluids (HTFs).
[0110] While it is important to emphasize that exergy destruction
is more readily minimized by avoiding substantial phase change when
there is a large temperature difference between the hot source gas
and the cold source gas (so the number of HTL loops can be
reduced), a minor amount of boiling and condensing may take place
within the HTLs without departing from the spirit of this
invention. Hence, the HTL may be referred to as an HTF, as is
customary in the prior art, though in a high-.epsilon. recuperator
the enthalpy associated with phase change would be small compared
to that associated with temperature change.
[0111] The description of the shell-side fluids as "clean gases" in
earlier discussions requires further clarification. It is
anticipated that in many cases the amount of condensation, acid
formation, ice formation, corrosion, and particulates will be
minor, though such are not precluded. When fouling mechanisms are
negligible, the fin pitch may be reduced for improved compactness.
However, the FL recuperator will still be advantageous in many
applications where these mechanisms are substantial--though perhaps
not where they are strongly dominant.
[0112] Fouling will often be significant in only one of the gas
streams, and often only at either the hot or cold end of that
stream. A strong advantage of the compound recuperator is that it
may readily permit individual modules to be switched off-line for
rejuvenation (defrosting, cleaning, re-plating, etc.) while a fresh
module is put into service. In some case, the fouled module may
need to be shipped back to the factory for service, but often it
will simply need to be drained, defrosted, burned out, or solvent
washed. In many cases, it will simply be necessary to orient those
modules in which significant condensation occurs so that the
condensate readily drains while in use--as for example the draining
of moisture from the common AC evaporator on a humid day. The
compound exchanger will often permit a dramatic reduction in the
number of replacement exchanger modules that will need to be kept
on hand in a large process plant.
[0113] Large applications are also anticipated where the shell-side
fluids are viscous organic liquids, since such exchanges also
benefit from the very short flow passages that can more easily be
obtained in the inventive module. While the benefits may be
greatest with oils of high viscosity, even moderate-viscosity oils,
such as 1,3-diphenylpropane at 310 K, where .mu.=4.4 cP,
k.sub.t=0.12 W/m-K, .rho.=968 kg/m.sup.3, and C.sub.P=2 kJ/kg-K,
would benefit when high effectiveness is needed, especially if
effervescence is also present in one of the streams. In such a
case, a phase separator or flash drum can be inserted between
modules or even between cores to separate the evolved shell-side
gas so the fluid's volumetric flow rate (and hence velocity)
remains low--to limit viscous losses.
[0114] A composite fluid property that is dimensionally much
simpler than (F.sub.MF.sub.G).sup.0.5 and nearly as valid for
comparing diverse fluids for similar flow geometries is
F.sub.D=k.sub.t.rho.C.sub.P/.mu. [11]
The HTL in Table 1 having the lowest F.sub.D (i.e., least
desirability) at 500 K is (again) the salt, where F.sub.D=2.7E5
J.sup.2/(s-m.sup.4-K.sup.2-cP) in these mixed reduced units, which
herein will be abbreviated Dt (for Doty). (In SI units, 1 Dt=1000
J.sup.2/(kg-m.sup.3-K.sup.2).) For comparison (again at 500 K),
F.sub.D is seen to be .about.440 kDt for 40 wt engine oil and 22
MDt for water.
[0115] In contrast, the shell-side fluids have lower F.sub.D. A
typical value for the gas conditions indicated earlier (500 K, 5
kg/m.sup.3, 0.05 W/m-K, etc.) would be .about.25 kDt. Some liquids
for which high-performance heat recovery will be needed have
F.sub.D well below those of a preferred HTL, and in such cases, a
compound recuperator can be advantageous, particularly if the
temperature permits the use of a tube-side HTL of very high
F.sub.D, such as water or a molten alloy.
[0116] For diphenylmethane at 310 K for example, F.sub.D is
.about.100 kDt, and for 1,3-diphenylpropane F.sub.D is 52 kDt. For
heavy oils, F.sub.D can be an order of magnitude smaller yet, even
at temperatures where substantial heat recuperation may be needed
in some situations.
[0117] The FL recuperator will be useful for heat recovery in many
fluids where F.sub.D is less than 200 kDt at the operating
conditions, which generally implies .mu.>1 cP for organic
liquids. When the shell-side fluid has rather high F.sub.D (as for
some low-viscosity liquids and gases at very high pressures), a
tube-side HTF would be needed with very high F.sub.D, such as water
or a molten alloy. However, a tube-side HTF with F.sub.D as low as
200 kDt would be satisfactory when operating with shell-side fluids
of rather low F.sub.D. Preferably, the tube-side fluid would have
F.sub.D more than 10 times that of the shell-side fluids (which of
course can be very different, and at very different
conditions).
[0118] Although this invention has been described herein with
reference to specific embodiments, it will be recognized that
changes and modifications may be made without departing from the
spirit of the present invention. All such modifications and changes
are intended to be included within the scope of the following
claims.
* * * * *
References