-------
120
no
in
(X.
1670 PSIA,
RE ~
£ Forces on Inlet Valve = 0 and
Inlet Valve Will be Forced Open
by Cylinder Steam Pressure
PD_ = Recompression Pressure
KE
For Inlet Valve Opening
Pc = 49 PSIA Acceptable
Even With No
Recompression Valve
40
I
I
I
500
1000 1500 2000
Rotative Speed - RPM
2500
3000
Figure 4.4-12.
Maximum Allowable Condenser Pressure for Inlet Pressure
of 1000 PSIA and 78 Ib. Recompression Valve Spring
Preload.
81
-------
y = - fer (3 a2£-a3) = -1.924 x 10~4 W in.
obJ.
where a = 5.75 in.
H = 6.83 in. (fixed end of rod to top of piston)
W = applied load, Ib.
E and I same as previous values
The top of the piston will be deflected 0.010 in. radially by a
side load of 52 Ib. applied at the free end of the rod (approximately
the bottom of the piston).
82
-------
5.0 EXPANDER MATERIALS
5.1 Materials Technology Base
In recently designed steam reciprocating engines for automotive
(12}
applications ', maximum steam temperatures of about 700°F have been utilized.
Temperatures have been restricted to this general range by materials
limitations, in one form or another. For example, the GM SE-101 engine
had a maximum steam temperature of 700°F in recognition of the possible
thermal degradation of the organic lubricant used to prevent excessive
wear at the steam piston cylinder interface. A design pressure of 800 psia
arose from considerations of expander bearing loads. These operating
conditions (700°F, 800 psia) might be considered to be current practice
in the design of steam engines for automotive applications. Engine thermal
efficiency is quite limited, under these conditions, as compared with
modern central steam power stations operating at 1000° to 1050°F and
2400-3500 psia. These advanced conditions have been reached by increasing
technology over the last 40 years, the same period in which there has been
essentially no increase in the cycle conditions for the steam car engine.
The higher temperatures associated with the more efficient automotive
steam engine place constraints on materials due to mechanical strength
and corrosion considerations. To insure the long life (*»• 3000 hours) re-
quired for satisfactory use in the reciprocating steam expanders, the
materials must be resistant to oxidation, corrosion by steam and have
adequate mechanical properties (yield strength, resistance to creep and
cyclic loading, hardness, etc.) at the anticipated operating conditions.
Surface properties can be favorably altered by the application of a sur-
face coating; use of surface coatings are considered where appropriate,
" * -v
i.e., wear surfaces. In addition to the constraints imposed on materials
by the engine operating conditions, economic constraints can limit the
utilization of existing technology in the solution to materials problems.
83
-------
Cost is a major factor in large quantity production of automotive com-
ponents. Economy dictates that the lowest cost alloy or material, with
properties suitable to the service conditions, be used.
Probably, the most limiting factor in the successful operation of
highly efficient steam engines will be lubrication. The dry conditions
which superheated steam imposes on a system at the operating condi-
tions can cause severe galling and excessive wear problems in those com-
ponents whose surfaces are in relative motion, i.e., the cylinder liner/
piston ring interface and the inlet valve face/seat and inlet valve stem/
guide interfaces. Lubrication of these components are discussed in
Section 6.0.
A literature search and industrial survey was made to identify and
evaluate the current materials technology in support of the expander de-
signs described in Section 4.0. Personal visits were made to the Ford
Motor Company Automotive Research and Engineering Center and the General
Electric Company Diesel Engine Department, Large Steam Turbine and
Generator Department and Medium Steam Turbine and Generator Department
to identify materials used in current reciprocating and steam turbine
components. To supplement the information obtained from these visits,
library searches were made covering automotive components and oxidation
and wear resistant materials for use in dry steam environment through
the General Electric Technical Information Center Library and Automatic
Information Retrieval System, the SAE and ATME-ASM Transactions and the
Engineering Index.
A summary of the results obtained from the technology review follows:
5.1.1 Cylinder Block, Cylinder Head, Intake Manifold. Exhaust Manifold
Experience with materials in large central station steam power station
has shown that the useful life of steel is essentially unlimited if the proper
steel is selected and the service conditions and water chemistry are properly
controlled. Some steels have been in service for forty to fifty years.
Invariably when a failure occurs the cause can be attributed to impurities in
the steam and the problem usually is solved by correction of the service
conditions. Plain carbon and low alloy steels are commonly used for the
84
-------
fabrication of pressure-containing components such as feedwater piping,
steam piping, valve bodies, flanges, discs, etc.,that are regularly found
in steam generating systems. The most commonly used steels are listed in
Table 5.1-1 together with the appropriate ASTM designations and maximum
service temperatures.
The use of these steels up to the maximum temperature listed in
Table 5.1-1 will vary with the anticipated service conditions and from user
to user. For example, one manufacturer limits the use of 1.25% Cr-0.5% Mo
steel to 800°F. In actual practice carbon steels have been used success-
fully in steam service at temperatures to 750°F for 25 to 30 years. Simi-
larly, the 0.5% Mo, the 1.25% Cr-0.5% Mo and the 2.25% Cr - 1.0% Mo alloy
steels have proven to be satisfactory for service at temperatures up to
850°, 1000°, and 1050°F, respectively. However, there is a major restriction
in the use of 0.5% Mo steel in that it is not recommended for use in welded
structures. Welded 0.5% Mo steel has a strong tendency to graphitize in
high temperature service. All of these steels can be considered for use
in the major static components for automotive steam engines.
In contrast to the high pressure-containing components in steam generating
systems, the materials currently being utilized in the production of the
cylinder blocks, cylinder heads and intake manifolds of gasoline and diesel
fueled internal combustion engines for passenger cars, trucks and locomo-
tives are gray cast iron and cast steel. The most common gray cast iron
being used for the production of these components in passenger cars is
SAE G4000 (ultimate strength - 40,000 psi). SAE G4000 contains 3 - 3.3% C,
1.8 - 2.1% Si, 0.6 - 0.9% Mn and consists of a lamellar pearlite matrix
with Type A flake graphite. For parts subjected to higher pressure or
heavy duty engines, SAE G6000 (ultimate strength 60,000 psi) gray cast
iron is used. In one large diesel engine the cylinder head is produced
from a cast, weldable Mn-Mo steel (0.20% C - 1.35% Mn - 0.25% Mo). The
steel casting is normalized at 1650°F (during the hardening of the valve
seats) to a hardness of 180 - 225 BHN. This treatment produces an ultimate
strength of 90,000 psi, a yield strength of 60,000 psi, tensile elongation
of 10% and reduction-in-area of 25%. Special heat resistant parts are
frequently made from Ni-Mo alloy steel. Intake manifolds in some internal
combustion engines are also being produced from cast aluminum alloy 355
85
-------
Table 5.1-1
1.
2.
3.
4.
5.
6.
7.
8.
9.
10.
11.
12.
13.
14.
15.
16.
17.
18.
Plain Carbon
Pressure Containing
Alloy
Carbon Steel
(0.35% C max)
Carbon Steel
(0. 25% C max)
Carbon Steel
(0.35% C max)
Carbon Steel
(0. 25-35% C max)
Carbon Steel
(0.06-0.18% C -
0.35% Si max)
Carbon Steel
(0. 27% C max)
0.25% C - 0.5% Mo
0.25% C - 0.5% Mo
0.15% C - 0.5% Mo
0. 20% C - 0. 5% Mo
0.15% - 1.0% Cr -
0. 5% Mo
0.15% C - 1.0% Cr -
0. 5% Mo
0.20% C - 1.25% Cr -
0.5% Mo
0.15% C - 1.25% Cr -
0.5% Mo
0.15% C - 1.25% Cr -
0. 5% Mo
0.15% C - 1.25% Cr -
0.5% Mo
0.18% C - 2.5% Cr -
1.0% Mo
0.15% C - 2.25% Cr -
1.0% Mo
and Low Alloy Steels Used for
Components in Steam Generating Plants
ASTM
Designation
A216 (Grade WCB)
A216 (Grade WCA)
A105 (Grades 1,
ID
A106 (Grades A,
B,C)
A192
A210
A217 (Grade WC1)
A182 (Grade Fl)
A335 (Grade PI)
A209 (Grade Tia)
A182 (Grade F12)
A225 (Grade P12)
A217 (Grade WC6)
A182 (Grade Fll)
A225 (Grade Pll)
A213 (Grade Til)
A217 (Grade WC9)
A182 (Grade F22)
86
(13)
Max. Service
Form Temp., °F
Casting
Casting
Forging
Pipe
Boiler Tube -
High Pressure
Superheater
Tube
Casting
Forging
Pipe
Boiler and
Superheater
Tubes
Forging
Pipe
Casting
Forging
Pipe
Boiler & Super-
heater Tubes
Casting
Forging
850
850
850
850
850
850
850
850
850
950
1050
1000
1050
1050
1050
1100
1050
1050
-------
Table 5.1-1 (Cont'd.)
Plain Carbon and Low Alloy Steels Used for
Pressure Containing Components in Steam Generating Plants
Alloy
19.
20.
21.
22.
23.
24.
25.
26.
0.15% C -
1.0% Mo
0.15% C -
0.5% Mo
0.15% C -
1.0% Mo
0.15% C -
1.0% Mo
0.15% C -
0.5% Mo
0.15% C -
1.0% Mo
Type
(18%
Type
(18%
304
Cr -
321
Cr -
2.
2.
2.
3.
5.
9.
25% Cr -
0%
25%
0%
0%
0%
Cr -
Cr -
Cr -
Cr -
Cr -
SS
10% Hi)
SS
10% Ni-Ti)
ASTM
Designation
A225
A213
A213
A213
A213
A213
A213
A213
(Grade
(Grade
(Grade
(Grade
(Grade
(Grade
(Grade
(Grade
P22)
T3b)
T22)
T21)
T5)
T9)
TP304)
TP321)
Form
Pipe
Boiler
heater
Boiler
heater
Boiler
heater
Boiler
heater
Boiler
heater
Boiler
heater
Boiler
heater
\
Max. Service
Temt).. 8F
& Super-
Tubes
& Super-
Tubes
& Super-
Tubes
& Super-
Tubes
& Super-
Tubes
& Super-
Tubes
& Super-
Tubes
1050
1200
1200
1200
1200
1200
1500
1500
87
-------
(Al-5% Si - 1.25% Cu - 0.5% Mg - 0.1% Zn - 0.2% Ti - 0.1% Mn - 0.2% Fe).
Materials used for the exhaust manifold in internal combustion
engines are subjected to the highest temperatures in the engine and would
correspond to the intake manifold of the steam expander. Alloyed cast
iron is most commonly used for the exhaust manifold in the internal com-
bustion engine. The SAE G4000d, a Cr-Mo alloyed cast iron and SAE G4000e
a Cr-Mo-Ni alloyed cast iron are typical. The range of operating tempera-
tures that have been reported for the higher temperature components of
the internal combustion engine that require the use of alloyed gray cast
iron are:
Component Operating Temperature Range. °F
Cylinder Heads 450 - 1000 (locally)
Exhaust Manifolds 300 - 1200
Exhaust Valve Seat 800 - 1300
The alloyed cast irons can be considered for components of the steam
expander that operate at lower temperatures and pressures. Where strength
is critical and greater reliability (safety) and oxidation resistance is
required, as is encountered in high pressure steam systems, nodular cast
iron is utilized. Typical nodular cast irons are SAE D5506, D4512 and
D7003. SAE D5506 has a yield strength (0.2% offset) of 55,000 psi and
a tensile elongation (^ 2 inches) of 6%; the structure is ferritic-pearlitic.
SAE D4512 has a yield strength of 45,000 psi, a tensile elongation of 12%
and has a ferritic structure. SAE D7003 has a yield strength of 70,000
psi, a tensile elongation of 3% and a pearlitic structure.
5.1.2 Cylinder Liners/Piston Kings
As stated previously, the high-speed relative motion between the
cylinder liner and the power piston rings in combination with the high
temperatures and pressures that exist within the cylinder chamber make
these components two of the most critical components in the engine. Al-
though the materials currently in use as cylinder liners and piston rings
in gasoline and diesel fueled internal combustion engines were reviewed,
the technology is not directly applicable to the reciprocating steam
expanders being designed to operate with dry solid lubricants. However,
88
-------
cylinder liner/piston ring material combinations used in oil free air
compressors are applicable and will be discussed in Section 6.0.
The materials most commonly used for cylinder liners under oil
lubricated conditions are alloyed gray cast iron, the Meehanite cast
irons, chromium plated gray cast iron and nitrided gray cast iron. Typical
of the alloyed gray cast iron is a composition of 3.35% C - 2.15% Si -
0.65% Mn - 0.4% Cr; the liner is hardened by quenching from 1600°F into
salt at 475°F followed by cooling in air. The resulting hardness is
Rockwell C 45 minimum. The Meehanite cast irons are high strength, fine
grained castings combining the properties of cast iron and steel. A
commonly used Meehanite, Grade GA, has an ultimate strength of 50,000
psi. Liners that are chromium plated or nitrided are produced from centri-
fugally cast gray cast iron. The liners are centrifugal cast in order to
achieve the bore surface integrity (minimum porosity) that is required
for the hard chromium plating and nitriding processes. A typical centri-
fugally cast gray cast iron used in liners for large diesel engines has
an ultimate strength of 35,000 psi and a hardness of ^ 250 BHN.
Gray cast iron also is the most common material used for oil lubricated
compression and oil piston rings. Occasionally AISI 1070 and 52100 bearing
steels are used. However, the steel segment or steel rail types of oil
rings are almost always made of AISI 1070 to 1095 steels. The chromium
plating of piston rings has been found to reduce wear of the rings and
cylinders by •*• 75%. However, chromium cannot be run against itself so
that the rings cannot be chromium plated if the liner is chromium plated.
Chromium plating is used primarily on compression rings in heavy duty
engines and on oil rings of the segmented type. Since chromium plate
reduces the fatigue endurance limit of cast iron, higher strength cast
irons must be specified for chromium plating applications. Two examples
of cylinder liner/piston ring combinations are shown in Table 5.1-II.
The K-28 nodular cast iron (3.3% C - 2.2% Si - 0.5% Mo - 0.5% Cu -
0.05% Mg) is a centrifugally cast martensitic ductile iron and has a
hardness of 372 - 437 BHN. The K-27 ring also is a centrifugally cast
martensitic nodular cast iron with a hardness of 250 - 283 BHN. The K-6E
ring is a statically cast gray iron (3.7% C - 2.85% Si - 0.45% Mo - 0.3% Cr)
consisting of free graphite flake in a fine pearlite matrix and has a
H'J
-------
Table 5.1-11
Cylinder Liner/Piston Ring Material
Combinations for Heavy Duty Diesel Engines
Piston Rings
Cylinder Liner Type Material
1. Hard Chromium Plated Top Compression K-28 Nodular Cast Iron
Centrifugally Cast Gray _ _- _„_„ _ T
Cast Iro Bottom Seal K-6E Gray Cast Iron
Oil K-Iron
2. Nitrided Centrifugally Top Compression Hard Chromium Plated
Cast Gray Iron _
Bottom Seal K-Iron
Oil K-Iron
90
-------
hardness of 230 - 297 BHN. The K-iron (3.7% C - 2.65% Si) is a statically
cast unalloyed gray cast iron consisting of free graphite flake in a
matrix of pearlite and with an ultimate strength of 30,000 psi.
5.1.3 Piston
The pistons for automotive internal combustion engines are produced
primarily from aluminum die castings. The light weight and excellent
thermal conductivity of aluminum alloys made them very attractive as the
trend in engines went toward higher speeds and higher compression ratios.
Most pistons are made from SAE 328 (Al, 11 - 12.5% Si, 0.9% Fe, 1-2% Cu,
0.5 - 0.9% Mn, 0.4 - 1.0% Mg, 0.05% Ni, 1.0% Zn, 0.25% Ti) and SAE 332
(Al, 8.5 - 10.5% Si, 1.2% Fe, 2-4% Cu, 0.5% Mn, 0.5 - 1.5% Mg, 0.5% Ni,
1.0% Zn, 0.25% Ti). These alloys are usually heat treated (T5) to a
hardness of approximately 100 BHN and are tin plated for scuff resistance.
For heavy duty engines, the piston head usually is produced from
alloyed (Cr-Mo) cast iron. However, in order to achieve greater reliability,
at least one manufacturer has changed to a cast Mn-Mo steel (0.20% C -
1.35% Mn - 0.5% Si - 0.25% Mo) in the production of power piston heads.
The Mn-Mo steel has a minimum yield strength of 60,000 psi and a hardness
of 180 BHN. In order to reduce weight of the cast steel power piston,
the lower portion is constructed from a forged aluminum alloy - usually
4032 alloy (11 - 13.5% Si, 0.8 - 1.3% Mg, 0.5 - 1.3% Cu, 0.5 - 1.3% Ni)
in the T-6 condition. The 4032-T6 alloy has a minimum yield strength of
42,000 psi and a hardness of 115 BHN.
One of the major reasons that the high silicon content aluminum
alloys are used in the fabrication of aluminum pistons is to minimize
the difference in thermal expansion between the cylinder wall material
and the piston material. The coefficient of thermal expansion for 4032
alloy between RT and 572°F is reduced to 11.7 x 10 in./in./°F from
13.1 x 10~ in./in./°F for unalloyed aluminum.
5.1.4 Piston Pin
•* " ; ~ - - - ^
Failure of piston pins rarely occurs. However, 0.001 - 0.002 inch
wear of the piston pin results in a noisy engine so that selection of
materials to minimize wear is very important. Piston pins art: normally
91
-------
produced from AISI 1117 (1.15% Mn), 1016, 5115 or 5120 (0.8% Cr) steel
for automotive engines and AISI 8620, 8640 (0.55% Ni - 0.5% Cr - 0.2% Mo)
and 5046 (0.4% Cr) alloy steel for heavy duty engines. The low carbon
steels are carburized and the higher carbon steels are case hardened by
induction to a surface hardness of Rockwell C 60. Usually the surface
is polished to a high finish, i.e., on the order of 2 RMS. It has been
determined experimentally that the load that the piston pin can withstand
without scoring is directly proportional to the surface finish - the
better the finish the higher the permissible load. Under severe condi-
tions of loading, the surface of the pin can be roughened by shot peening
(followed by lapping to remove the displaced metal) to form depressions
in the surface in 'order to better hold the lubricant.
5.1.5 Piston Pin Bearing
There are a relatively large number of metals and metal alloys that
are used in sleeve bearings. The bearing alloys can be grouped in the
following classes: tin-base alloys, lead-base alloys, copper-lead alloys,
tin-bronze, silver, aluminum alloys, zinc-base alloys, gray cast irons,
non-metallic materials (PTFE) and overlays. The proper selection of the
bearing material will depend upon the type of load, i.e., steady state or
cyclic, amount of load, speed, temperature, corrosive conditions, oil
supply, dirt contamination, and shaft hardness.
Type of load - the load carrying capacity of cyclically loaded
bearings, as in piston pin bearings, is usually determined by the bearing
fatigue strength of the material. This property is of minor Importance
for bearings under constant load.
Amount of load - cyclic loads up to 1500 psi are considered low and
can be supported by ordinary babbitt bearings (tin or lead base alloys).
Thin babbitt overlays can operate up to 2500 psi. The limit of alternating
loads for the copper-lead bearing material is approximately 4000 psi. A
load of 5000 psi can be supported by a trilayer bearing consisting of a
mild steel (AISI 1010) back, a 25% Pb-Cu alloy intermediate layer and a
plated babbitt overlay. The latter bearing is used in some heavy duty
engines.
92
-------
Speed - babbitt alloys are best suited for high speed operation.
Temperature - strength and corrosion rates are affected by the
service temperature. Certain excellant bearing materials, such as babbitts,
are restricted by their low melting points. The tin-bronzes such as
SAE 62 (86.7% Cu - 10% Sn - 10.3% Pb - 2.0% Zn - 1.0% Ni) and phosphor
bronzes such as SAE 64 (78% Cu - 10% Sn - 10% Pb - 0.75% Zn - 0.5% Ni -
0.5% Sb) can be used as bearings at temperatures approaching 600°F. Maxi-
mum load for these materials is *• 4000 psi. Both materials have excellant
fatigue properties but poor anti-seizure, conformability and embeddability
properties; anti-seizure can be improved by silver plating.
Corrosion Resistance - the tin-base babbitt alloys are the most
resistant to the corrosive action of acids that are formed in lubricating
oils.
Oil Supply - a common cause of failure in bearings is the loss of
the oil supply or breakdown in the oil film. This results in direct
contact between the pin and the bearing surfaces and a large increase in
friction. When this happens, the anti-seizure characteristics of the
bearing are most important. The rating of materials for this property
in order of decreasing anti-seizure qualities are: tin babbitt, lead
babbitt, aluminum alloys, copper-lead alloys, leaded bronzes, silver and
bronze.
Dirt Contamination - in bearing applications where there is a high
level of contamination by particulate matter, the bearing material must
possess good embeddability. Dirt can be embedded in soft babbitt ma-
terials without doing harm.
Shaft Hardness - the harder the bearing, the harder the shaft must
be to prevent damage to the shaft. In general applications, soft babbitt
bearings can be used satisfactorily with steels as soft as 130 - 165 BHN;
for copper lead (to 25% Pb) the minimum hardness of the steel shaft must
be 165 - ZOO BHN; for aluminum alloys the minimum hardness of the steel
shaft must be 200 - 300 BHN; and for leaded bronze the minimum hardness
of the steel shaft must be 300 - 400 BHN.
93
-------
A listing of recommended bearing materials for various bearing loads
in internal combustion engines is given in Table 5.1-III. The materials that
are currently being used for piston pin bearings in the two leading large
diesel engines for locomotives are (1) a leaded tin bronze (80% Cu -
10% Sn - 10% Pb), both as a solid bushing and with a AISI 1010 steel
backing (SAE 792) and (2) a tri-metal bearing consisting of a AISI 1010
steel backing, a 0.015 in. thick silver intermediate layer and a 0.0003
in. thick lead overlay. These bearings require a good finish and many
axial grooves for pressurized oil lubrication. A crankshaft bearing
material that is currently in use in automotive engines is a 0.030 in.
thick SAE 780 (8280) aluminum alloy sheet (6% Sn - 1% Cu - 0.5% Ni -
1.5% Si) bonded to a 0.001 in. thick Ni plated AISI 1010 steel backing.
A recommendation received from one of the leading aluminum producers
also called for the use of a leaded tin bronze bearing for the piston
pin bearing and a AISI 1010 steel backed aluminum alloy bearing for the
crankshaft.
5.1.6 Connecting Rod
The high alternating loads imposed on the connecting rod in re-
ciprocating engines require the use of relatively high strength steels
of high quality. For this reason connecting rods are generally produced
from contour forgings made from vacuum degassed ingots. The most common
steels used for the production of connecting rods are AISI 1041 for use
in automotive engines and AISI 8640 (0.55% Ni - 0.5% Cr - 0.2% Mo), 4140
(0.95% Cr - 0.2% Mo) and 4340 (1.8% Ni - 0.8% Cr - 0.25% Mo) for heavy
duty and diesel engines. One automotive engine manufacturer utilizes
SAE 80002 malleable iron castings for the production of connecting rods.
5.1.7 Camshaft/Cam/Cam Follower (Tappet)
The cyclic loading and high localized surface stresses that occur
as a result of the action between the cam surface and cam follower (tappet)
makes the materials selection for these components very important. In
some engines, the localized Hertzian stress on the cam surface approaches
200,000 psi. Numerous combinations of camshaft/tappet materials are in
common use. Hardenable gray cast iron, such as SAE C4000 is most widely
used for automotive camshafts. A typical composition IH 3.3% C - 2.25% SI -
0.6% Mn - 0.95% Cr - 0.5% Mo. As cast, the camshaft has a hardness of
94
-------
Table 5.1-III
Bearing Materials Used in Internal Combustion Engines
Bearing Load, psi
Lowest Cost Material
Alternate Materials
Advantages of
Alternate Materials
1200 max
1200 - 2200
vo
2200 - 2800
1200 max
(Small and Medium Size Engines)
Pb Babbitt on steel
(0.015-0.030 in.)
Pb Microbabbitt on steel
(0.002-0.005 in.)
Babbitt-impregnated sin-
tered Cu-M on steel
Babbitt-impregnated sin-
tered bronze on steel
Cu-25% Pb on steel
Al alloy on steel
Pb Babbitt on steel or
bronze (0.015-0.030 in.)
Zr or cast Al alloy
Sn Babbitt on steel
(0.015-0.030 in.)
Babbi tt-impregnated
sintered Cu-Ni on steel
Babbitt-impregnated
sintered bronze on steel
Sn Microbabbitt on steel
Al alloy on steel or solid
Cu-35% Pb on steel
Solid Al alloy or Al alloy
on steel
Above materials + 0.001 in.
Pb alloy overlay
Cu-25% Pb on steel + 0.001 in.
Pb alloy overlay
Al alloy on steel + 0.001 in.
Pb alloy overlay
(Large Engines)
Sn Babbitt on steel or bronze
Solid Al alloy
(a) (b) (c)
(d)
(a) (b) (d)
(a) (b) (d)
(d)
(a) (b) (d)
(a) (b)
(a) (b) (d)
(b) (e)
(d) (e) (f) (g)
(d) (e) (f) (g)
(d)
(c) (d)
-------
Table 5.1-III (Cont'd.)
Bearing Materials Used in Internal Combustion Engines
Bearing Load, psi
Lowest Cost Material
Alternate Materials
Advantages of
Alternate Materials
\o
1200 - 1500
1500 - 2000
2000 min.
Pb Babbitt on steel or
bronze (0.020 in. max)
(Large Engines)
Pb Microbabbitt on steel or
bronze
Pb Microbabbitt on steel
or bronze (0.002-0.005 in.)
Solid Al alloy
Aolid Al alloy
Cu-Pb (30-40% Pb) on steel
Solid Al alloy or Al alloy on
steel
Above materials + 0.001 in.
Pb alloy overlay
Cu-25% Pb. on steel + 0.0005-
0.002 in. Pb alloy overlay
Al alloy on steel + 0.0005-
0.002 in. Pb alloy overlay
(a) (b)
(c) (d)
(a) (b)
(c) (d)
(b) (e)
(a) (b) (c)
(d) (g)
(a) Yield strength
(b) Fatigue strength
(c) Possible cost savings
(d) Corrosion resistance
(e) Conformability
(f) Embeddability
(g) Inherent lubricity
-------
248 - 311 BHN; the cam surfaces are flame or induction hardened to a Rockwell
C hardness of 54 min. For heavy duty engines the camshafts are forged and
are usually produced from water quenched carbon steels and low alloy steels
of 0.5 - 0.7% C, AISI 4340, or a curburizing grade such as AISI 8620 steel.
Heat treatment of the high carbon steels to achieve the necessary hardness
of the cam surfaces is accomplished by either conventional through-hardening
or case hardening processes using flame or induction. One diesel engine
manufacturer uses a forged AISI 1080 shaft which is quenched and tempered
to a hardness of 235 BHN. The cam and bearing surfaces are induction case
hardened to a hardness of Rockwell C 60 - 65 to a depth of 0.060 in. Care
must be taken to insure sufficient depth of case as thin, brittle surface
hardened cases tend to spall in service.
Tappets usually fail by scuffing or rapid loss of surface. The most
common automotive tappet material is a gray cast iron of a composition similar
to 3.2% C - 2.25% Si - 0.8% Mh - 1.1% Cr - 0.6% Mo - 0.55% Ni. The cast iron
tappet is hardened by quenching in oil from 1550°F and tempered to a hardness
of Rockwell C 55 - 60. Steel tappets also are used and are produced from
hardenable steel such as 52100 bearing steel or high-carbon molybdenum steels
of the AISI 4000 series or from carburizing grades of steel such as AISI 51200
or 8620.
In some heavy duty engines, where tappet wear has been a severe problem,
it was found that the utilization of material containing varying amounts of
carbides in the microstructure improved the service life. In one case the
successful use of hardened and tempered D-2 tool steel in solving the wear
problem was attributed to free carbides in the microstructure. The D-2 alloy
has a high carbon content in conjunction with stable carbide formers. The
composition is 1.5% C - 12% Cr - 0.8% V - 1.0% Mo. A higher volume carbide
content in the microstructure can be obtained with alloys such as Stellite
Star J (2.5% C - 32.5% Cr - 17.5% W - Bal Co) and this material at Rockwell
C 60 hardness is being used for tappets in some heavy duty engines. Experience
has shown that the Star J alloy shows superior performance against a nitrogen
rich surface in comparison to a carburized surface. It should be noted that
-*
if a nitrided surface is to be used for the cam, a carbonitride process should
be utilized in order to obtain the desired case depth to carry the high
loads. The core depth of a straight nitrided steel surface is too thin for
application on cam or tappet surfaces. A solid tungsten carbide cermet also
97
-------
is being considered for use as a tappet material. In some very heavily loaded
cam/tappet designs a roller cam follower is employed. An AISI 8620 steel
carburized to a surface hardness of Rockwell C 55 min. and a case depth of
0.060 inch has performed satisfactorily in roller type cam followers. Again
a surface finish of 2 - 4 RMS is recommended.
Tappet faces are usually finished to about 6 RMS for the same reasons
a good finish is required for the piston pin, i.e., improved load carrying
ability. Shot peening of the tappet surface is sometimes done to improve
the retention of the lubricant. Coating of the highly finished tappet surface
to aid in the wearing-in process and improve frictional characteristics is
common practice. An oxide coating (Fe,~0,) is generally applied to chilled
cast iron and a phosphate coating is generally applied to hardened steels or
gray cast irons. The phosphates can be manganese phosphate, zinc phosphate or
iron-manganese phosphate. Other coatings that have been used are oxidates,
manganates, and sulfides.
The following material combinations were considered as candidates for
the cam and tappet in the single cylinder engines:
Cam
1. AISI 8620/cairburized
(Re 58 min.)
2. AISI 8620/carburized
(Re 58 min.)
3. AISI 8620/carbonltrided
(Re 60-65)
4. Hardenable Cast Iron
(Re 54 min.)
5. Hardenable Cast Iron
(Re 54 min.)
Tappet
Chilled Cast Iron (Re 54)
Ferrox Coating (Fe-O,)
Carboloy 883 (WC + 6% Co)
(RA 92)
Star J (Re 61)
Hardenable Cast Iron
(Re 55-60)/Mh-Fe Phosphate
Coating
D-2 Tool Steel (Re 58-60)/
Mh-Fe Phosphate Coating
98
-------
Bearing materials used for camshafts in heavy duty diesel engines are
generally cast aluminum alloys. Aluminum alloy 750 (6.25% Sn - 1.0% Ni -
1.0% Cu) is commonly used.
5.1.8 Inlet Valve
Operating conditions for valves in large central steam power stations
vary considerably from the operating conditions of valves for automotive
and diesel internal combustion engines. The operating conditions for steam
valves for central power stations are much less severe than for the valves
in internal combustion engines. Actuation of many steam valves is on an
infrequent basis so that fatigue and wear of the valve face and seat is not
a severe problem. Service temperatures of the steam valve are significantly
lower than those for the exhaust valve of an internal combustion engine and
they do not fluctuate as they do in the internal combustion engine; thus
creep and thermal fatigue problems are less severe. In addition, corrosion
as a result of high temperature combustion products is not a problem in steam
valves.
In sumnary, the years of experience in the operation of steam valves
or exhaust valves in internal combustion engines are not directly applicable
to the selection of materials for inlet steam valves in reciprocating
steam expanders. Selection of materials for the various components of the
steam inlet will have to be based on the knowledge of the basic properties
of the materials with respect to corrosion in steam, mechanical properties
and wear behavior under conditions of no lubrication and as supplemented
with valve experience in reciprocating internal combustion engines.
Inlet steam valves in large central power steam systems are generally
fabricated from a martensitic 12% Cr stainless steel (AISI 410) or Crucible
422 alloy (12% Cr - 1% Mo - 1% W - 0.8% Ni - 0.25% V); the stem and face
are nitrided. Both the valve seat and valve guide are made from Stellite
6B. The Stellite 6B on the valve seat usually is applied as a weld over-
lay on a low alloy steel (2.25% Cr - 1.0% Mo alloy).
99
-------
The materials in use in valve components for internal combustion
engines will vary considerably with the operating temperature of the valve.
The lower strength and increased corrosion and wear rates that accompany
increased service temperatures necessitates the use of more highly alloyed
materials as the temperature of the valve increases. Since the inlet
valves in internal combustion engines operate at temperature under 600°F
and exhaust valves operate at temperatures as high as 1550°F, it is ob-
vious that the most severe materials problems are associated with the
exhaust valve. The development of improved valve materials over the years
has been primarily by empirical methods. This stems from the fact that
it is not possible•to simultaneously reproduce all of the service condi-
tions of the valve by any other means than in an actual engine.
Materials that are recommended for use in inlet and exhaust valves
in various types of engines are listed in Table 5.1-IV. The steels in Group
A of Table 5.1-IV are used for light duty inlet valves that operate at low
temperature or for short times. They also are used for stem materials
in two-piece valve construction where they can be welded to higher alloy
heads and used for heavier duty inlet and exhaust valves.
Steels in Group B through E are especially made for valve applica-
tions with the Group B steels being the least expensive and the Group E
steels being the most expensive. The sigma phase forming steels in Group
D have good hot hardness and superior resistance to wear than the austenitic
steels of Group E. However, they are more brittle and have less resistance
to creep than the Group E steels. The nickel-rbase super alloys listed
in Group F are only used where valve temperatures are very high because
of their high cost.
Short operating lives of valves can usually be traced to permanent
dimensional changes that take place in the valve components during service
as a result of inadequate creep strength or poor wear and/or hot corrosion
characteristics. The terms used to describe the dimensional changes are:
a. elongation,
b. projection of the stem,
c. face runout, and
d. tip recession.
100
-------
Table 5.1-IV
Materials Used for Inlet and Exhaust Valves in Internal Combustion Engines
Application
Valve Materials
Approx. Max.
Exhaust Valve
Temp. °F
A. Intake Valve-Light Duty
B. Intake Valve-Heavy Duty
C. Intake Valve-Heavy Duty
Exhaust Valve-Light Duty
D. Exhaust Valve-Heavy Duty
Carbon Steel
SAE NV-1 (AISI 1041)
(a) SAE NV-2 (AISI 1047)
(AISI 1050)
Low Alloy Steels
SAE NV-4 (AISI 3140)
(AISI 4150)
SAE NV-6 (AISI 5150)
(AISI 6145)
SAE NV-5 (AISI 8645)
Martensitic Steels
SAE HNV-2 (Sil F) 4 Si-2.25 Cr (0.4C)
4 Si-2.25 Cr 1.5 Ni-0.85 Mo (0.4C)
Martensitic Steels
(b) SAE HNV-3 (Sil 1) 3.25 Si-8.5 Cr (0.45C)
2.75 Si-7.5 Cr-1.5 Ni-0.85 Mo (0.3C)
(b) SAE HNV-6 (XB) 2.25 Si-20.0 Cr-1.5 Ni (0.8C)
Austenitic-Sigma Phase Alloys
SAE EV-1 (SCR) 23.5 Cr-4.75 Ni-2.75 Mo (0.45C)
SAE EV-2 (TXCR) 24.0 Cr-3.75 Ni-1.35 Mo-3.75 Mn (0.4C)
1350
1550
-------
Table 5.1-IV (Cont'd.)
Materials Used for Inlet and Exhaust Valves in Internal Combustion Engines
Application
Valve Materials
Approx. Max.
Exhaust Valve
Temp. °F
E. Exhaust Valve-Heavy Duty
o
NJ
F. Exhaust Valve-Heavy Duty
Austenitic Steel Alloys 1550
SAE EV-7 (2155N) 21 Cr-5 Ni-5.5 Mn (0.2C, 0.25N)
21 Cr-4 Ni-7 Mn (0.4C, 0.1N, 0.22P)
(a) SAE EV-8 (21-4N) 21 Cr-4 Ni-9 Mn (0.4C, 0.4N)
SAE EV-5 (Sil 10)19 Cr-8 Ni-3 Si (0.4C)
SAE EV-4 (21-12N)21 Cr-12 Ni (0.2C, 0.2N)
(Cast) 25 Cr-12 Ni (0.2C)
SAE EV-9 (TPA) 14 Cr-14 Ni-2.4 W-0.35 Mo (0.45C)
(Cast) 15 Cr-15 Ni-3.5 Si-0.4 Mo (l.OC, 0.25 Cu)
Austenitic Ni Alloys 1650
SAE HEV-2 (Inconel M)
SAE HEV-3 (Inconel X-750)
(c) SAE HEV-5 (Nimonic 80A)
SAE HEV-6 (Nimonic 90)
16 Cr-Bal Ni-3 Ti-0.5 Al (0.03C)
15 Cr-Bal Ni-1 Cb-2.5 Ti-0.9 Al
(0.04C)
20 Cr-Bal Ni-2.5 Ti-1.2 Al
(0.05C)
20 Cr-Bal Ni-18 Co-2.5 Ti-2.2 Al
(0.05C)
(a)
Used in current automotive engines
Used in current high performance automotive engines - intake valves
(c)
Used in current heavy duty engines
-------
Elongation ("wire drawing") is the permanent increase in length as measured
from the valve face to the tip. This causes a decrease in the "lash"
(clearance between valve tip and tappet) so that the valve face does not
seat properly and results in "blowby" of the exhaust gases. This can be
corrected by using a material with greater resistance to creep. Face
runout is caused by hot corrosion which also allows leakage of the ex-
haust gases. Stem projection is similar to elongation in that it reduces
"lash"; it is the total elongation as measured through the valve guide
and includes any wear of the seat and valve face and elongation of the
stem. Tip recession results in an increase "lash" and leads to fracture
of the valve. Face runout and stem projection, due to corrosion and wear
of the seat and valve face and tip recession, can usually be corrected by
appropriate use of valve seat inserts and/or corrosion resistance and
hard facing materials. Commonly used insert and facing materials are
listed in Table 5.1-V. The Group D and E materials are used most often
with the cobalt base alloys, Group E, are preferred for improved re-
sistance to corrosion.
Examples of material combinations used in exhaust valve components
for current automotive and large diesel internal combustion engines are
given in Table 5.1-VI.
5.2 Expander Materials Selection Study
The structural materials used in the containment of high tempera-
ture (1000° - 1050°F) and high pressure (2400-3500 pai) steam as well as ma-
terials used for specialized components such as steam valves have been
studied in depth for years and their properties are well documented.
Similarily, materials in use in heavy duty and high performance recipro-
cating engine components are well established and most of these materials
are adaptable to use in corresponding components of the reciprocating
steam expander. For this reason,..there was no need to conduct experi-
mental investigations of candidate expander materials for this program.
103
-------
Table 5.1-V
Materials Used for Valve Inserts and Hard Facing Applications
Material
Nominal Composition, %
SAE Alloy
Designation
C
A.
B.
C.
D.
E.
Cr-Mo and Cr-Mo-W Steels 0.
(Inserts) (a) 1.
1.
0.
1.
Cr-Mo Cast Iron 2.
(Inserts) 2.
W Steels 0.
(Inserts)
Ni Alloys
(Facings)
Co-Cr-W-Ni Alloys
(Facings)
0.
2.
0.
1.
(b) 1.
1.
2.
(c) 2.
65
00
00
65
35
50
25
50
55
00
20
00
25
60
50
40
Mn Si Cr
0.6 1.00 3.
0.6 2.50 4.
0.6 0.25 5.
0.35 0.25 5.
0.35 0.45 3.
0.60 2.00 3.
1.00 1.00 2.
4.00 -
0.25 3.
0.40 0.30 25.
1.00 1.00 19.
28.
- 28.
- 24.
- 30.
- 29.
0
0
0
0
5
0
5
5
0
5
0
0
0
0
0
Ni Mo
W
- 5.00 -
8.50 -
- 1.25 -
3.00 3.5
1.0 6.50 5.5
5.00 -
7.50 -
9.5
- -
60.0 -
78.0 -
_ _
3.0 -
24.0 -
- -
39.0 -
13.0
8.5
—
4.0
4.5
12.5
11.0
15.0
C.o
0.
1.
0.
25 (V)
00 (V)
20 (Cu)
HNV-7
_
—
67.
55.
37.
52.
10.
0
0
0
0
0
VF-4
VF-1
VF-2
VF-5
VF-3
(X-782)
(80-20 NiCr)
(Stellite 6)
(Stellite F)
(Eatonite)
(a)
(b)
(c)
Light duty on large gasoline and diesel engines.
Strongest and toughest of the Co base alloys.
High hot hardness and resistance to shock and pitting.
-------
Table 5.1-VI
Exhaust
Valve Component
Material Combinations Used in Exhaust Valve Components
for Internal Combustion Engines
Material
(a)
Light Duty
Spark Ignition
Heavy Duty
Spark Ignition
Heavy Duty
Diesel
Head
Face (overlay)
Stem
Seat (Insert)
Guide
SAE EV-8 (21-4N SAE HEV-5 (Nimonic)
SAE VF-5 (Stellite F) SAE VF-3 (Eatonite)
SAE EV-8 (21-*4N)(b^ SAE HEV-5 (Nimonic) 80A
(Eatonite or L.E. Jones Alloy)
(b)
-(Gray Cast Iron>
Inconel 751
(c)
SAE VF-2 (Stellite 6)
AISI 3140(c)
AISI 410 SS
(Re 40-49)
Weld overlay
Gray Cast Iron
(Type A Flake in
Pearlite)
(a)
(b)
(c)
Material compositions are listed in Tables 5.1-IV and 5.1-V
One piece construction.
Two piece construction - flash welded.
-------
However, prior to the selection of the structural materials for
the two single cylinder steam expanders, their strength properties and
compatibility in a steam-air environment were reviewed. In addition,
material and process specifications were established for each component
in the steam expanders.
5.2.1 Material Properties
5.2.1.1 Low Carbon and Low Alloy Steels
The use of any one steel for structural applications and contain-
ment of high pressure steam primarily depends on the sustained metal
temperatures. Low carbon steel is used for steam generation tubes and
the low temperature regions of the superheater where the metal tempera-
tures do not exceed 750° - 800°F. As the temperatures in the steam
generating system increases, steels of higher alloy content are employed.
The C-0.5% Mo steel can be used up to about 850°F; the low chromium-
molybdenum steels (1.0 - 2.25% Cr) up to 1000° - 1050°F; the intermediate
chromium-molybdenum steels (3 - 9% Cr) up to 1100°F; and the austenitic
stainless steels above 1100°?.
The temperature limits imposed on the various steels are the result
of loss in elevated temperature strength (yield and creep), structural
changes in the microstructure of the steel that may detrimentally affect
the properties, i.e., ductility, and oxidation-steam corrosion considera-
tions. The maximum allowable design stresses for the low carbon, low
alloy and austenitic stainless steels used in boiler construction as a
function of temperature are shown in Figure 5.2-1. The maximum allowable
stress is based on the ASME Boiler and Pressure Vessel Code which limits
the allowable stress to the level at which the creep rate is 0.01% in
1000 hours at the design metal temperature. Additional restrictions on
the allowable stresses are imposed by the code and they are based on the
stress-rupture strength of the material as modified by suitable allowances
for loss in metal cross-section due to oxidation and corrosion. From
Figure 5.2-1, the allowable stress for low carbon steels begins to drop
above 700°F and at 850°F is only half the value at room temperature. The
maximum allowable stresses for the low alloy Cr-Mo steels are sustained
up to about 800°F whereupon they start to drop and reach 50% of the room
106
-------
Carbon steel
Tube SA-192-A
Carbon steel
Tube SA-210
Pipe SA-106-B
Plate SA-201-B
C-0.5Mo
Tube SA-209-T1
1.25Cr-0.5Mo
Tube SA-213-T11
Pipe SA-335-P11
Plate SA-387-C
2.25Cr-lMo
Tube SA-213-T22
Pipe SA-335-P22
Plate SA-387-D
18Cr-10Ni-Ti
Tube SA-213-TP321H
SA-312-TP321H
SA-376-TP321H
SA-240-321
16Cr-13Ni-3Mo
Tube SA-213-TP316H-
Pipe SA-312-TP316H
SA-376-TP316H
SA-240-316
Pipe
Pipe
Plate
Pipe
Plate
I
200 300 600 800 1000 1200 1400 1600
Metal Temperature, °F
Figure 5.2-1.
Maximum Allowable Design Stresses for
Low Carbon, Low Alloy and Austenitic
Stainless Steels in Large Steam
Generating Systems. Reference 14
107
-------
temperature values at 1000°F. The 50% loss in the maximum allowable de-
sign stresses for the austenitic stainless steels occurs at about 1200°F.
Of the low alloy, Cr-Mo steels, the 1.25% Cr-0.5% Mo alloy has superior
creep rupture properties at temperatures on the order of 1000°F, Table 5.2-1
As chromium is added to achieve superior oxidation resistance, the creep-
rupture strengths are reduced. Care must be exercised in the fabrication
of components from these steels, i.e., amount of cold work, heat treat-
ment, etc., such that grain size and phase morphology are not adversely
affected with respect to their influence on strength and ductility.
The thermal stability of the low alloy steels also restricts the
temperature at which the steels can be used. Phase changes, such as the
graphitization of carbon steels above 750°F and C-Mo steels above 850°F,
can result in drastic reduction in strength and ductility of the steel.
Chromium additions of more than 0.5% in the steel eliminate this problem.
Compatibility with the surrounding environment is another important
consideration in the selection of the structural material. In the steam
expander, as is the case for the boiler, the exterior surfaces of the
containment materials are subject to oxidation by oxygen in the air and
the internal surfaces are oxidized by the oxygen in the steam. Each
steel has a threshold temperature above which rapid oxidation or corrosion
takes place generally due to the formation of a thick porous scale. As
a result the load carrying ability is reduced and, in the case of boiler
tubes, restricted flow and reduced heat transfer efficiency occurs. Suf^
ficient oxidation and corrosion resistance can be achieved in the steels
by the addition of small amounts of chromium usually in excess of 1%. The
chromium in the steel improves the oxidation resistance by promoting the
formation of a tightly adhering scale which inhibits further oxidation.
The degree to which chromium additions improve the oxidation resistance
of low alloy steels in steam can be seen in Figures 5.2-2 and 5.2-3^ .
In Figure 5.2-2, the data are plotted as penetration in mils/yr. vs. tem-
perature and in Figure 5.2-3, the data are plotted as penetration in mils
as function of time at 1100°F. It is important to point out that these
data (Figure 5.2-2) were obtained under actual service conditions from
samples periodically cut from a large steam plant (Detroit Edison) and
show higher corrosion rates than in steady state laboratory tests by a
factor of *v/ 2.
108
-------
Table 5.2-1
Creep Rupture Properties of Cr-Mo Alloy Steels at
Stress (ksi) to Produce „ - _ . , .
7 Stress for Rupture, ksi
Alloy
1
1.
2.
3
5
7
9
Cr
25
25
Cr
Cr
Cr
Cr
- 0.
5
Cr -
Cr -
- 1.
- 0.
- 0.
- 1.
0
5
5
0
(a)
Mo
0.5 Mo
1.0 Mo
Mo
Mo
Mo
Mo
a iiireep nani uj. j./0 in
10,000 hours
11
12
12
10
9
.8
.0
.0
.5
.0
8.0
12
.0
L
1000 hrs.
28.
27.
20.
18.
19.
18.
22.
4
0
4
0
0
8
8
10,
18.
21.
15.
14.
14.
14.
19.
000 hrs
8
0
4
7
6
0
1
Annealed condition.
109
-------
70
60
50
40
30
t-l
CO
Q>
10
01
U
CO
OS
C
O
•H
to
o
o
o
e
3 20
en
u 10
>
950
lCr-0.5Mo
u2.25Cr-0.5Mo
C-O.SMo
Carbon Steel
1000
5Cr-0.5Mo_
—"
\ 9Cr-lMo
1200
1050 1100 1150
Temperature, °F
Figure 5.2-2. Effect of Temperature on Long Time Steam Corrosion Rates.
1250
0.20
0.15 —
§
0)
c
•i
0.05
0
7.5
Figure 5.2-3.
Time, hours x 1
Effect of hong Time Exposure to Steam at 1100°F.
110
-------
Similar long time oxidation-corrosion tests in full size commercial
tubing (2 in. OD x 0.5 in. wall) were conducted at the Philip Sporn plant
of the American Electric Power Company. In these tests, scaling data
were obtained on the internal and external surfaces after varying periods
of exposure at 1100°, 1200°, 1350°, 1500°F and 2000 psi steam pressure.
Data for the 1100°F exposures are given in Table 5.2-II. There is little
difference in scale" thickness between the two surfaces ' .
5.2.1.2 Medium Carbon Low Alloy Steels
A number of dynamic components in the steam expander require the
use of high quality relatively high strength materials. For these com-
ponents, i.e., power piston head, piston rod, connecting rod and bolts,
the following medium-carbon, low alloy steels were reviewed:
Nominal Composition, %
Alloy C Ni Cr Mo V
AISI 4140 0.4 - 1.0 0.2
AISI 4340 0.4 1.8 0.3 0.25
17-22-A 0.45 - 1.0 0.55 0.3
H-ll 0.35 - 5.0 1.5 0.4
The 0.2% yield strength of these alloys are compared in Figure 5.2-4.
H-ll steel has an exceptional combination of high strength and toughness
and is used for critical and highly stressed components. When the alter-
nating stress endurance strength is the limiting criteria, vacuum melted
grades of the steels are specified. For example, the endurance limits
for vacuum melted and air melted AISI 4340 alloy steel heat treated to
a strength level of 200,000 psi are 105,000 psi and 90,000 psi^respectively,
at room temperature.
The oxidation-steam corrosion characteristics of the medium-carbon,
low alloy Cr-Mo-V steels are similar to the low-carbon, Cr-Mo alloy steels.
Data for a Cr-Mo-V steel are given in Table 5.2-II.
5.2.1.3* Cast Iron
For reasons of cost and in certain wear applications, cast iron
should be specified wherever possible. Cast iron grades usually are
111
-------
Table 5.2-II
fa")
Oxidation-Corrosion of Cr-Mo Steels at 1100°FV '
Alloy
-------
280
S
240
200
co
P.
CO
CO
0)
CO
160
120
80
40
Note: All Steels in Hardened and
Tempered Condition
H-ll
17-22-A
4340
4140
510-560 BHN
311-363 BHN
340 BHN
360 BHN
.H-ll
4340
I
I
0
200
400 600
Temperature, °F
800
1000
Figure 5.2-4. 0.2% Yield Strength Medium Carbon Low Alloy Steels,
113
-------
specified to a specific strength level rather than to a chemical composi-
tion, i.e., SAE G4000, and numerous grades of cast iron are available.
The ultimate strengths and endurance limits of gray cast iron generally
are not affected by temperature until above 800°F. For example, a 2.84C -
1.5 Si gray cast iron with an ultimate strength of 48,400 psi has a fatigue
endurance limit of 20,000 psi. With the exception of a slight dip in
strength levels at the intermediate temperatures, to 42,000 psi ultimate
strength and 18,000 psi endurance limit, the initial strength levels are
retained to a temperature of approximately 800°F after which they decrease to
corresponding levels of 35,000 psi and 14,000 psi. Notched fatigue strength
usually are within 10% - 20% of the unnotched strength level.
Although improvements in the high temperature strength of gray cast
irons are possible by alloying with molybednum, where reliability is re-
quired in critical components of high pressure systems, the use of nodular
cast iron is preferred over gray cast iron. Nodular cast iron can be
made to significantly higher strength levels than gray cast iron in addi-
tion to being able to deform plastically. Measurable tensile elongation
values of 3 - 30% are possible with the nodular irons. They also have
good resistance to mechanical shock. Typical room temperature properties
for a 80-60-03 Pearlitic Grade of nodular iron are as follows:
(a)
Endurnace Limit, ksi
Ultimate, psi Yield, psi Elong., % Unnotched Notched
95-130 55-80 3-9 40 24
^a'For casting with 100 psi ultimate strength.
One of the major concerns in the use of cast irons in steam (or air)
at elevated temperatures is their dimensional instability due to oxida-
tion. This growth in dimensions can be significantly large in gray cast
iron because of the continuous network of graphite. However, additions
of chromium and the use of nodular cast iron or, for more severe condi-
tions of corrosion, the use of the Ni-Resist cast irons can eliminate or
minimize the dimensional changes due to oxidation. Data on dimensional
growth for various cast irons as a result of exposure to steam at 900°F
are given in Table 5.2-III.
114
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Table 5.2-III
Dimensional Growth of Cast Irons
in High Temperature SteamQ-8,19)
Cast Iron
Growth at 900°F. in./in.
500 Hrs. 1000 Hrs. 2500 Hrs.
Gray Cast Iron
Gray Cast Iron
+ 0.46% Cr
Gray Cast Iron
+ 0.6% Cr
Ni-Resist Type 2
Ni-Resist Type 3
(b)
(b)
Ni-Resist Type 2D
(Nodular)
Ni-Resist Type 3D
(Nodular)
(b)
(b)
0.0023
0.0003
0.0052
0.0000
0.014
0.001
(a)
0.0005
(a)
0.0005
0.0003
0.0003
0.0010
0.00045
0.0005
0.0015
0.00048
0. 0005
0.0000
(a)
(b)
1000°F/2000 hours.
Ni-Resist Type 2, 2D:
Ni-Resist Type 3, 3D:
O.OC max-2.25 Si-20 Ni-2.1 Cr.
2.6C max-2.1 Si-30 Ni-3.0 Cr.
115
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5.2.1.4 Aluminum Alloys
The use of aluminum alloys for dynamic components and cylinder walls
in internal combustion engines is highly desirable because of the low
density and high thermal conductivity of aluminum alloys. However, the
calculated temperatures in the crown of the power piston head, particularly
for the crosshead piston expander, and the calculated stresses in the
connecting rods and the piston rod of the crosshead piston expander pre-
cludes the use of aluminum alloys for these applications. At this stage
in the expander designs, only the skirt portion of the trunk piston and
the crosshead piston itself warrant consideration. For these applications,
it is desirable to utilize alloys with as low a coefficient of thermal
expansion as possible that is consistent with suitable mechanical prop-
erties.
As previously stated, the following alloys (high silicon contents)
are used in piston heads and skirts in automotive spark ignition and large
diesel engines: SAE 328, SAE 332 and 4032. The nominal compositions and
the coefficients of thermal expansion are given in Table 5.2-IV. Two other
high strength forging alloys (2014 and 2219) that have been recommended
for high temperature engine components also are listed. For comparative
purposes, the room temperature minimum tensile properties of these alloys
are given in Table 5.2-V.
Typical properties of the forged 4032 alloy that was selected for
the piston components in this program together with typical properties
for 2014 and 2219 alloys are listed in Table 5.2-VI. From these data it
is apparent that the 4032 and 2014 alloys have poor thermal stability
at the higher temperatures in contrast to the excellant thermal stability
of 2219 alloy. Significant degradation in the tensile properties is not
observed in the 2219 alloy until a temperature in excess of 600°F is
reached or after very long time exposures at 600°F (10,000 hrs.).
Careful consideration must be given to the use of aluminum alloys
in steam. Although aluminum alloys should perform satisfactorily in dry
steam at elevated temperatures, i.e., 600° - 700°F, erosion may be a
problem in wet steam. Further, it is known that conventional commercial
aluminum alloys will corrode very rapidly in high purity water at tem-
peratures over 400°F. However, aluminum alloys have been used as cladding
116
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Table 5.2-IV
Nominal Compositions of Aluminum Alloys
Nominal Composition, %
Alloy Si Fe Cu Mn Mg Ni Zn Ti Other
SAE 328(a^ 11.75 0.9 1.5 0.7 0.7 0.0005 1.0 0.25
SAE 332^a) 9.5 1.2 3.0 0.5 1.0 0.5 1.0 0.25
4032^ 12.25 1.0 (c) 0.9 - 1.1 0.9 0.25(c) - 0.1 Cr(c)
2014(b) < 0.8 - 4.4 0.8 0.4 -
2219(b) 0.2(c) 0.3(c) 6.3 0.3 0.02(c) - 0.1(c) 0.06 °*^0V,
U. -Lo Zr
VjU
in.
11.
12.
11.
13.
13.
ej. . rjtpciiis j-uti
/in./°F x 10~6
5 (68-392°F)
0 (68-392°F)
7 (68-572°F)
6 (68-572°F)
6 (68-572°F)
(a)
(b)
(c)
Casting - automotive engines.
Forging - large diesel engines,
Maximum.
-------
Table 5.2-V
Room Temperature Tensile Properties
of High Silicon Aluminum Alloys
Minimum Tensile Properties
Ultimate, 0.2% Yield, Elong.,
Alloy Condition Temper psi psi %
SAE 328 Cast T5 32 26 low
T65 42 37 low
SAE 332 Cast T5 31 - low
4032 Forged T6 52 42 5
2014 Forged T6 65 55 10
118
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Table 5.2-VI
Typical Mechanical Properties for Aluminum Alloy Forgings
(4032, 2014 and 2219 in J6 Condition)
vo
Test
Temp . ,
Alloy °F
4032 75
300
500
*
2014 75
300
500
2219 75
300
500
600
700
Exposure
Time Prior
to Test,
hours
1/2
4.000
1/2
1000
-
1/2
1000
1/2
1000
w
1/2
1000
1/2
1000
1/2
1000
1/2
1000
Tensile Properties
Ult. , ksi
55
46
44
23
10
70
51
34
25
11
64
51
49
30
30
22
18
12
5
Yield, ksi Elong. , %(4D)
46
41
40
22
7
63
49
30
24
9.5
45
39
37
22
22
16
14
10
4.5
9
9
9
12
45
15
14
20
18
43
10
20
20
25
25
26
28
30
80
oung s Fatigue Properties
ksi x 10^ Cycles Stress, ksi
11.3 1 x 10® 18.0
5 x 10 16.5
10.5 1 x 10® 13.0
5 x 10 11.5
9.3 1 x 10® 5.5
5 x 108 5.0
10.5
9.6
9.6
8.5
8.5
10.6 1 x 10® 16.5
X
5 x 10° 15.0
9.9
9.0
8.6
7.6
6.4
-------
materials in pressurized water reactors operating in excess of 300°F.
In fact, an Al-Ni-Fe alloy (X8001) has been developed specifically for
this purpose and has shown outstanding resistance to corrosion in high
purity water at elevated temperatures. Corrosion tests on the X8001 alloy
containing 0.5% Fe and 1.0% Ni in high purity water exhibited the following
corrosion rates:
Exposure
Temperature, °F Corrosion Sate, mils/yr
550 1.7
600 3,1
680 8.0
(a)
Test duration - approximately 1500 hours.
5.2.1.5 Case Hardened Low Alloy Steels
Dynamic components that are relatively low stressed but are subject
to wear in service require that their surfaces be hardened. An excellent
example is the camshaft where cyclic Hertzian stresses on the surface
approach 200,000 psi. In these applications, induction hardening of high
carbon steels (AISI 1080), carburizing or nitriding of low-medium carbon
low alloy steels (AISI 8620, 4340) are employed. Since fatigue fractures
generally initiate at the surface, it is extremely important that the
surfaces of case hardened components be free of defects or irregularities
that will provide stress-risers. Assuming defect free surfaces, the
surface hardening treatments will substantially increase the fatigue en-
durance limit, partly because of the induced compressive stresses in the
surface. As an example, where the fatigue endurance limit of heat treated
AISI 4340 is approximately 75,000 - 80,000 psi, shot peening the surface
will increase the endurance limit to approximately 90,000 - 100,000 psi
and nitriding the surface will increase the endurance limit to 120,000
to 135,000 psi. In a crankshaft application, designed for a minimum
/
115,000 psi tensile strength, a AISI 4140 steel heat treated to a 85,000
psi tensile strength and nitrided on the wear surfaces, provided the
necessary wear and fatigue resistance.
120
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5.3 Materials Recommendations
The materials recommendations for the single cylinder steam expanders
were based on the materials technology survey (Section 5.1) and the
materials selection study (Section 5.2). In general, considerable conservatism
was exercised in the selection of materials for the engines in this program.
The conservatism in materials selection was based on the fact that the engines
were to be used as test vehicles. The selection of materials was further
compromised by the fact that only one or at most two components of any design
were required. For this reason no castings were utilized for major components
of the expander other than the valve guide and tappet. A listing of the
materials selected for each component in the single cylinder steam expanders
together with the recommended materials specification and heat treatment
condition are given in Table 5.2-VII. These materials were selected on the
basis of a high probability of success rather than cost. Future studies
would be useful, directed toward materials optimization with respect to
cost. For example, the addition of approximately 1% Cr + 0.20% Mo (AISI 4140)
to AISI 1045 steel represents an increase in price of the finished mill
product to about $0.05/lb for quantities in excess of 10,000 pounds. In
the final selection of materials, mechanical properties and environmental
compatibility must be balanced against cost. Materials must be used up to
the limit of their acceptable life before specifying a more costly material.
121
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Table 5.2-VII
Materials Recommendations for Single Cylinder Steam Expanders
Expander
Component
1. Inlet Manifold
Cylinder Head
Valve Body
2. Exhaust Manifold
3. Cylinder Block
4. Cylinder Block-
Lower (Crosshead
Piston Expander
Only)
5. Cylinder Liner (A)
(B)
6. Piston Compression
Rings (A)
(B)
(Trunk Only)
Rider Ring
Lower Seal Ring
Scraper Seal Ring
Oil Cutter Ring
Recommended
Material
2.25% Cr-1% Mo Steel
1.0% Cr-0.5% Mo Steel
1.25% Cr-0.5% Mo Steel
AISI 304 SS/Gas Nitride
Bore
17-22-A/Hard Cr Plate
Bore
AISI 440C
Sb Impregnated
Carbon-Graphite
Koppers K-1051/over
Inconel X-750
Carbon-Graphite
Koppers K-35 Ni-Resist
Koppers K-Iron
Koppers K-6E
Material
Specification
Condition
ASTM A387 Grade D, Plate
ASTM A335 Grade P22, Pipe
ASTM A387 Grade B, Plate
ASTM A335 Grade P12, Pipe
ASTM A387 Grade C, Plate
ASTMA335 Grade Pll, Pipe
ASTM A240, Plate
ASTM A376, Pipe
GE/P11BYA11, Nitride
AMS 6304C/GE B5F5, Forging
GE/P16DYA10, Chromium Plate
AMS 5630C, Forging
See Section 6.0 of Report
Normalize 1700°F, Temper 1250°F
Normalize 1700°F, Temper 1250°F
Normalize 1700°F, Temper 1250°F
Anneal 1900°F/Rapid Cool
Normalize 1750°F/l-l/2 Hrs.
Mar Quench 650°F/10-15 min.
Temper 318 - 363 BHN
Austenitize 1900°F/Oil Quench
Double Temper 675°F/4 Hrs.
Hardness Re 52-56
-------
Table 5.2-VII(Cont'd)
Materials Recommendations for Single Cylinder Steam Expanders
U)
Expander
Component
7. Power Piston
8. Piston Rod
(Crosshead Only)
9. Rod Seal
(Crosshead Only)
10. Crosshead Piston
11. Connecting Rod
12. Piston Pin
13. Piston Pin Bushing
Bushing
14. Camshaft/Cam,
Bearing Surfaces
15. Tappet
16. Tappet Housing
Recommended
Material
17-22-A (1% Cr-0.55% Mo-
0.3% V)
H-ll Steel
PTFE-Bronze, MoS_
Filled
Al 4032-T6 Alloy
AISI 4140
AISI 8620/Carburized
80 Cu-10 Sn-10 Pb/Mild
Steel Backing
AISI 8620 Carburized
Chilled Cast
Antichafing Coating
Nitralloy 135 Mod/
Gas Nitride Bore
Material
Specification
AMS 6304C/GE B5F5, Forging
AMS 6487C, Bars & Forgings
GE/P16DYA10,Chromium Plate
ASTM B247, Forging
AMS 6382G, Forging
AMS 6276C,Bars & Forging
GE/P11BYA11,Carburize
SAE 792
AMS 6276C, Bars & Forgings
GE/P11BYA11, Carburize
Eaton EMS 91 or EMS 80
ASTM A355 Class A, Bar
Nitride - Drawing Control
Condition
Normalize 1750°F/l-l/2 Hrs.
Mar Quench 650°F/10-15 min.
Temper 311-363 BHN
Austenitize 1850°F/Air Cool
Double Temper 1000°F/2 Hrs.
Hardness Re 52-55
Solution Treat
955 ± 15°F/2 Hrs.
Water Quench 150-212°F
Age 340°F/10 Hrs.
Austenitize 1550°F/Oil Quench
Double Temper 1000°F/2 Hrs.
Hardness 320-360 BHN
Core Hardness Re 58 min.
Core Depth 0.040-0.070 in.
2 RMS Finish
Stabilize 1000°F/4 Hrs.
Core Hardness Re 58 min.
Core Depth 0.040-0.070 in.
Austenitize 1750°F/Oil Quench
Temper 1100°F/2 Hrs.
Core Hardness 330-370 BHN
Case Hardness RC 65 (converted)
Case Depth 0.015-0.020 in.
-------
Table 5.2-VII (Cont'd)
Materials Recommendations for Single Cylinder Steam Expanders
Expander
Component
Recommended
Material
17. Push Rod
18. Valve Head.Facing
Valve Stem
NJ 19. Valve Seat
** (Facing)
20. Valve Guide
21. Valve Rings
22. High Temp.
Static Seals
AISI 8620/Carburize
H-ll Steel/Stellite 6
H-ll Steel/Gas Nitride
Stellite 1
Ni-Resist D-3
(Nodular)
H.S.25/LPA-101 Coat
H.S.25
Material
Specification
Condition
AMS 6276C, Bars & Forgings
GE/P11BYA11, Carburize
AMS 6487C, Bars & Forgings
ASTM.A399 Type R Co-Cr-A
Hard Surface Rod
Nitride - Drawing Control
ASTM A399 Type R Co-Cr-C
Hard Facing Rod
ASTM A439, Ductile Iron
Casting
AMS 5759
Coating - Drawing Control
AMS 5759
Core Hardness Re 58 min.
Core Depth 0.040-0.070 in.
Austenitize 1850°F/Air Cool
Double Temper 1000°F/2 Hrs.
Hardness Re 52-55
Coating Thickness - 0.125 in.
Case Hardness Re 65 (Converted)
Core Depth 0.010-0.015 in.
Coating Depth 0.125 in.
Coating Hardness Re 48 min.
Stabilize 1600°F/2 Hrs.
Furnace Cool to 1000°F
@ 100°F/Hr.
Hold 1000°F/1 Hr.
Slow Cool
Cold Worked & Aged @ 1100°F/5 Hr.
Hardness Re 45-50
Coating Thickness 0.003 in.
Material Cold Reduced to
Re 45-50
-------
6.0 LUBRICATION
6.1 Lubrication Technology Base - Solid Lubricants
One of the most critical problems associated with the design of an
efficient steam engine is that of lubrication within the cylinders and
the inlet steam valve. In order to fully utilize the potential thermal
efficiency of the steam cycle, it is necessary to operate the expander
with high inlet steam temperature. With increased steam temperature,
there is a corresponding increase in temperature of the materials within
the cylinders and valves. The vast majority of steam expander designs,
from the early Stanleys to the most modern engines, obtain upper cylinder
lubrication by means of hydrocarbon lubricants pumped in with the steam.
Due to thermal degradation, oxidation, and sludging of these lubricants,
their use generally limits steam temperatures to about 750°F. In recog-
nition of this limitation, the General Motors SE-101 engine was limited
to a maximum steam temperature of 700°F.
The Pritchard engine is lubricated in a similar way. A slow
speed mechanical pump forces small quantities of oil into the steam to
act as an upper cylinder lubricant. Approximately one pint per 500 miles
is used in this fashion. After 10,000 miles of operation with a closed
steam cycle, 20 pints (2.5 gallons) of oil will have been pumped into
the steam system. A number of other more recent steam engine designs
also have expanders with the lubricant added to the working fluid side
of the engine.
Even though lubrication technology for steam engines has been es-
tablished for large, low-speed engines operating at moderate temperatures
and pressures, an adequate lubrication system has not been identified
for highly efficient steam engines that operate at high inlet steam tem-
peratures and at high inlet pressures. There are two approaches which
125
-------
appear most feasible for lubrication within the cylinders of high tem-
perature, high pressure steam engines: (a) to inject water-miscible
fluids, which have been extensively developed for cutting and grinding
applications, into the steam; or (b) to use solid lubricants which have
been developed for extreme condition applications. With the injection
of a water-miscible fluid, it would be necessary to add a separating system
to separate the lubricant from the water to avoid fouling of the steam
generator tubes; also, this approach would result in a high consumption
of water-miscible fluid. For this reason, use of solid lubricants appears
to be the most feasible method of lubricating the cylinder walls in order
to permit the utilization of higher steam temperatures.
A literature search and industrial survey was made to identify and
evaluate the current technology in solid lubricants and wear resistant
materials. Searches were made through the following resources: General
Electric Company Automatic Retrieval System, SAE Transactions, Engineering
Index, NASA Scientific and Technical Information, AGAKD and the Defense
Documentation Center. The industrial survey was made to determine those
solid lubricants that either are in commercial production or in the stage
of advanced development.
6.1.1 Criteria for Solid Lubrication
The most important function of any lubricant is to keep the rubbing
surfaces separated, both to prevent wear of the moving parts and to
minimize the high coefficients of friction associated with metal-to-metal
contact. A solid lubricant keeps the moving surfaces separated solely
through its own strength. However, to do so, the solid lubricant must
remain in the clearance between the moving surfaces for the entire de-
sign life of the engine. It follows then, that in order for a material
to be an effective solid lubricant, it must possess most or all of the
following properties and characteristics depending upon the form of the
lubricant:
1. Good bonding properties or high tendency to adhere to rubbing
surfaces
2. Low shear strength
3. High compressive strength
126
-------
4. Laminar crystal structure
5. High bulk fracture strength
6. Good thermal stability
7. Chemical compatibility with surrounding environment (including
structural materials)
8. Low wear
9. Good resistance to thermal shock
10. Good fabricability.
Although all of the above listed properties are important with
respect to achieving satisfactory performance by solid lubrication, the
ability of the lubricant to adhere to the rubbing surfaces is absolutely
essential in order to protect the surfaces and prevent scoring, galling,
gross welding and resultant metal transfer. This corresponds to wetting
in liquid lubricated systems. For self-lubricating composites, it is
necessary that the lubricant be transferred to the mating surface to form
an adherent thin film. For example, the ability to form a transfer film
is what makes graphite an excellant lubricant.
Low shear strength is Important from the standpoint of achieving
a low coefficient of friction. The coefficient of friction for solid
lubricants in general is high in comparison to oils, as shown in the
following tabulation:
Typical
Lubricant Coefficient of Friction
Oil 0.001
Solid Lubricant-Bonded 0.01
Solid Lubricant-Composite 0.1
Metal-Metal (Steel) Approaching 1.0
From these data it can be seen that it is desirable to utilize materials
with as low a shear strength as possible.
A high compressive strength is desirable from the standpoint that
the higher the compressive strength of the lubricant, the higher is the
load that can be applied before metal-to-metal contact is made. In
effect, the lubricant should have a low shear strength/compressive strength
ratio for best performance.
127
-------
The laminar crystal structure of the lubricant is associated with
low shear strengths and low coefficients of friction. Although effective
lubrication can be achieved with solids that do not have a laminar struc-
ture, the most effective solid lubricants are those that do have the
laminar structure and have good surface adherence properties. The laminar
structure consists of alternating layers of atoms in which there are
strong bonds between atoms within the layer (covalant or ionic forces)
and weak bonds between atoms in adjacent layers (Van der Waals' forces).
The Van der Waals'forces are easily broken in the "layer-lattice" struc-
ture resulting in the successive atomic layers readily sliding over each
other such as basal plane slip in the hexagonal structure of graphite.
Solid lubricants that have a laminar crystal structure are highly
anisotropic with respect to mechanical properties. In cases where the
lubricant is fabricated into a self-lubricating composite, the lubricant
matrix must have sufficient strength to withstand high pressure differ-
entials and frictional drag loads at high temperatures without changing
shape or fracturing. To endure the start-up cycle of a steam engine,
the self lubricating composite also must have adequate resistance to thermal
shock.
Thermal stability of the lubricants is required to assure acceptable
performance over the entire temperature range and throughout the design
life of the engine. Since the total surface temperature of the lubricant
includes the heat of the environment and the heat of friction, the in-
fluence of both must be considered. Thermal stability includes resistance
to thermal decomposition and crystallographic changes as well as being
resistant to oxidation in air and steam. Crystallographic changes will
result in changes in frictional characteristics. However, the effects
of surface heat are not always detrimental since some solid lubricants
are designed to work as a liquid phase.
Another important consideration is compatibility with all engine
materials and small quantities of organic lubricants. Consideration of
free energy exchange between steam and the candidate materials indicates
that a large number of candidate solid lubricants are available for
evaluation. However, free energy exchange is somewhat complicated by
the addition of cylinder and piston materials.
128
-------
Finally, an optimum solid lubricant is one that has low wear and
the ability to replenish the transfer film such that long life is ob-
tained. Also, for self-lubricating applications, the solid lubricant
must be readily fabricable into usable shapes.
6.1.2 Forms and Types of Solid Lubricants
The lubricating solids can be used in several different forms:
(a) as loose powder dusted or rubbed on the mating surfaces or supplied
to the parts in a controlled stream of a carrier gas; (b) as bonded ma-
terials attached to the surfaces with an adhesive binder; (c) as self-
lubricating bearing materials where the lubricating solid is dispersed
throughout a metal or plastic bearing or where the lubricating solid
forms the matrix of a fabricated bearings, and (d) as additives to oils
and greases to replace the chemical extreme pressure (EP) additives
normally used. The latter application of solid lubricants is not appli-
cable to the lubrication of cylinder walls and will not be discussed
further. Of the remaining three forms in which a solid lubricants can
be applied, each form has its own advantages and disadvantages.
Loose powders, which have the advantage of minimizing fabrication
or assembly problems5have the disadvantage of providing a very thin film
of limited life; the film either has to be replenished by external means
in order to achieve the desired life or be used only to assist in the
wear-in period.
Bonded solid film lubricants generally provide longer life than
powders with the performance of the bonded solid film varying with the
specific lubricant, the bonding agent and the method of application.
Some dry films depend solely on small particle size for bonding. The
lubricant particles are held together only by the attraction arising
from the extremely fine particle size and there is little bonding to the
substrate. A limited amount of dispersing agent is sufficient to bring
about bonding. Particle*-bonded dry films can be dispensed from aerosol
containers dr dispersed in water, organic solvent or other volatile
carrier. Most of the particle-bonded lubricants are air-drying and re-
quire no baking. However, the particle-bonded lubricants have the same
problem as the loose powders in that they have very limited life unless
129
-------
they can be replenished. An organic resin (phenolic, epoxy, polyimide)
may be added to increase binding between particles and bonding to the
substrate. The resin produces an adherent paint-like film that offers
the best wear life and widest range of use. The lubricant particles also
can be bonded together by a silicone, a suitable ceramic or metal salt
(sodium silicate, aluminum phosphate) that is hardened upon removal of
the solvent. This type of film can be very hard and is more temperature-
resistant than the organic resin bonded films. The type of binder that
is used depends upon the adhesion required, the anticipated service
temperature and environment, and the desired wear life. This type of
bonded film can be applied by spraying, dipping or brushing. Newer
application techniques include electrophoretic deposition, plasma spraying
and sputtering.
One of the major problems associated in the bonded films is that
the lubricant is usually limited by the binder itself. For example, the
resin bonded solid lubricants are limited to use temperatures of up to
about 400°F; silicones and silicates have permitted service temperatures
to increase to about 600°F at the expense of higher friction and lower
wear life. The organic resin and metal salt bonded solid lubricants also
have a limited wear life that can be attributed to the very small volume
of lubricant available. This becomes evident when one considers that the
film is only 0.0002 - 0.0005 in. thick.
Because of the relatively short life of the bonded solid lubricants,
self-lubricating components containing solid lubricants have been de-
veloped for longer life bearing and seal applications in order to take
full advantage of the good lubricating qualities of solid lubricants.
The solid lubricants are incorporated into the self-lubricating composites
by impregnation, sintering, and various forms of high-temperature, high-
pressure compaction. Self-lubricating bearings generally fall into two
major classes: (a) solid lubricant matrix utilizing graphite or the
lubricant plastics such as nylon, polytetrafluoroethylene (PTFE), etc.,
with reinforcing materials as fiberglass, graphite, molybdenum disulfide,
and soft metals; and (b) metal matrix in which lubricating solids such
as graphite, PTFE, metal disulfides, selenides, or tellurides and low
melting point metals are mixed with metal/alloy powders of copper, iron,
130
-------
nickel, molybdenum, etc., and sintered under pressure or impregnated
into a porous metal matrix skeleton.
In the application of solid lubrication for cylinder walls in high
performance reciprocating steam expanders where steam temperatures approach
1000°F, the peak steam pressure is 1000 psi and peak piston ring side
loads are on the order of 200 - 300 psi at an average linear velocity of
1200 -1600 feet/minute, it was assumed that self-lubricating composites
would be required to achieve the necessary life. Therefore, in the
survey of solid lubricants, emphasis was placed on the self-lubricating
solids or composites. Bonded solid film lubricants were considered
primarily as an auxiliary lubricant during the wear-in process or as an
aid in establishing a transfer film. The following materials were re-
viewed as lubricants to minimize wear in the cylinder wall/power piston
interface and valve stem/guide interface:
Self-Lubricating Solids
Carbon-graphites
Polytetrafluoroethylene (PTFE)
Sulfides/Selenides
Metals
Porous metal composites
Hard surfacing materials
Bonded Solid Films
Sulfides/Selenides
PTFE
Graphite
Soft Oxides
6.1.2.1 Self-Lubricating Solids
6.1.2.1.1 Carbon-Graphites
Graphite was found to show promise of being a very effective solid
lubricant for steam engine applications. Graphite has been used as a lubricant
for many years and is used extensively for brushes in electric motors.
In fact, it was excessive wear of motor brushes in high altitude aircraft
s.
131
-------
that led to the detailed study of the mechanism of graphite lubrication.
(21 22 23)
In a detailed study of this problem by the General Electric Company ' '
it was discovered that the low friction and wear normally exhibited by
graphite is due not to any lubricant quality inherent in the graphite,
but to adsorption on the surface of substances from the atmospheric en-
vironment. The laminar structure of graphite undoubtedly is essential,
but alone it is not sufficient. It was found that water vapor, adsorbed
on the surface, caused reduction in friction and wear rate. The minimum
pressure of water vapor that is necessary to achieve minimum wear rates
is on the order of 3 - 5 torr. Oxygen showed a similar effect but at
much higher pressures, i.e., 300 - 500 torr. It has been proposed that
only the edges of the graphite crystals need be covered with the adsorbed
film to achieve the low friction and wear rates by permitting the laminar
layers of the hexagonal lattice to slip. In the absence of water vapor
or oxygen the edges of the laminar layers of atoms are locked by free
radicals formed by the evaporation of volatile oxides of carbon. Graphite
then is expected to maintain good lubricity in the cylinder wall because
of the presence of high pressure steam.
In extremely high temperature applications, graphite tends to oxidize
and its uses are restricted to neutral or reducing atmospheres. However,
graphite has been successfully used as seals for steam turbine driven
electrical generating equipment which tends to verify its stability in
high temperature steam. The threshold oxidation temperature of graphite
(24)
in steam has been listed as about 1300°F and good resistance to
oxidation by steam is expected at the maximum proposed cycle temperature
of 900°F. This may be further verified by examination of the free energy
changes for the two oxidation reactions:
C + H20 = CO +
and
The free energies of these reactions are plotted against temperature in
Figure 6.1-1. As may be seen on this plot, the driving force for the
reactions (negative free energy) does not become appreciable until tem-
peratures greater than about 1200°F. This is in essential agreement
with the temperature limit cited above. In air, graphite cannot be used
132
-------
20
15
8
0)
P.
10
4-1
U
cfl
H-l
O
00
S-i
0)
0)
1-1
0
500 - 1000
Temperature, °F
1500
FiRiire 6.1-1.
Free Energy of Oxidation Reactions Involving Graphite
and Steam.
133
-------
much above 750°F without severe oxidation reactions.
I
Even though graphite displays excellent properties from the stand-
point of thermal stability, thermal conductivity, low elastic modulus,
low shear strength and thermal shock resistance, it has poor strength
properties in tension. Higher strengths are possible with the carbon-
graphite grades of which there are an infinite number to choose; lower
wear rates also are possible with the carbon-graphite materials. These
materials are available commercially and, although it is an oversimplifica-
tion, are produced in four general classes: (a) straight carbon-graphites
with varying carbon/graphite ratios, (b) resin impregnated, (c) metal
impregnated (Cu, Cu-Pb, Cu-Sn, Pb, Ag, Babbitt) and (d) high temperature
oxidation inhibited grades. The maximum service temperatures for these
materials in air are reported to be up to 500°F for the resin Impregnated
grades, 700"F for most of the straight carbon-graphites and metal im-
pregnated grades, and 1000° - 1200°F for the high temperature grades.
Properties of candidate carbon-graphite grades that are believed to be
suitable for application as piston compression rings in a steam environ-
ment are listed in Table 6.1-1. Care must be exercised in the use of the
oxidation-inhibited, high temperature grades and some of the resin im-
pregnated grades in steam and water because of the tendency for some of
the resins and oxidation inhibitors to exude to varying degrees in these
environments. Consideration also must be given to the compatibility of
graphite with metals used for the impregnation process as well as mating
metals/alloys that are in contact with the graphite in service. Studies
/25^
by McGeev ' at General Electric have shown that metal oxides that can
be reduced to lower oxides or to the metallic state by graphite in the
presence of oxygen appear to function as catalysts in the oxidation of
the graphite. The relative activity of various metal oxides with respect
to their effect on the oxidation of graphite is shown in Table 6.1-II
and is based on the change in ignition temperature of pure polycrystalline
graphite. The data show that lead oxide is particularily effective in
lowering the ignition temperature of graphite.
Although there are numerous influencing factors, the coefficient
of friction for the various grades of carbon-graphite will generally
range between 0.1 to 0.25. Wear rates also will vary significantly with
134
-------
Table 6.1-1
u>
in
Properties of Carbon-Graphite Grades
'
Vendor Grade
VS Graphite 86
Div' 110
2980
102
103
107
Pure Carbon P5NR
Co. pQ3
P658RC
X3310
Graphite- Bronze
Metallizing Graphalloy
Corp.
UCC CDJN
CJP
Carbone 5890
Corp. 589Q
JP-500
, ..
TypeW
CG-R
CG
CG-HT
CG-Cu
CG-Ag
CG-Sb
CG
G
CG
CG-HT
CG-Cu- Sn
CG-HT
CG-HT
G
G-HT
CG
(a) G Graphite
CG Carbon-Graphite
CG-R Resin Impregnated
Hardness
Shore
Compr
Transv.
Fracture
Tensile.
(5cleto*cope) Str.,ksi Str.,ksi Str.,ksi
100
92
67
88
89
92
80
75
90
85
-
105
65
CG-HT
CG-Sb,
Cu,Ag,
Cu-Sn
36
32
15
38
35
38
30
20
38
29
-25
36
26
12.5 '
10.0
4.5
9.5
10.5
10.0
8.5
8.0
11.0
9.4
_
8.8
8.0
High temperature/oxidation
Metal
Impregnated
9.5
8.5
4.0
7.5
7.5
7.5
7.0
5.5
8.0
8.5
~8.0
7.0
7.0
inhibited
E
106 psi
3.0
3.3
1.8
2.6
2.8
3.5
2.7
1.7
3.1
2.6
-
3.2
2.1
Coef .Th.
Exp.,
in/in/°F
£
x 10~6
(RT-500°F)
2.3
2.3
2.3
3.1
3.1
2.4
2.2
1.9(b)
2.2
2.3(c)
2.3
2.3
2.3
Th.Cond.
Btu/F.t/Hr
"F
7.6
8.0
30.0
8.7
9.3
7.0
_
-
_
-
_
4.4
13.3
Apparent
Density
gm/cc
1.90
1.90
1.85
2.35
2.70
2.20
1.70
1.82
1.80
2.05
_
1.76
1.77
Borosity
Vol. %
0.5
1.0
4.0
4.0
3.0
1.0
20.0
10.0
2.0
0.4
_
-
-
.
Max.
Service
Temp. ,Air
500
700
1000
700
700
700
500
900
500
1000
750
875
1100
(b) RT to 900CF
(c) RT to 1000'F
-------
Table 6.1-II
Catalytic Activity of
Catalyst
(acetate or oxide)
Pb
V
Mn
Co
Cr
Cu
Mo
Ag
Cd
Fe
Pt
Ni
Ir
Rh
Ru
Pd
Ce
Zn
W
Hg
Sn
Uncatalyzed
Oxides in Graphite
w/o as Metal
0.15
0.20
0.45
0.33
0.95
0.20
0.15
0.16
0.21
0.13
0.03
0.45
0.40
0.20
0.30
0.30
0.72
50.00
0.02
0.10
0.10
Oxidation
Ignition
Temp., °C
384 (738°F)
490
523
525
540
570
572
585
590
593
602
613
638
622
640
659
692
700
718
720
738
740 (1364°F)
136
-------
the grade of carbons-graphite. Overall, the metal Impregnated carbon-
graphite appears to be the most suitable for use as compression piston
rings in the steam expander. Data indicate that after an initial high
wear rate during the wear-in period, the metal impregnated carbon-
graphites have superior wear rates for long time service; however, the
coefficients of friction for the metal impregnated carbon-graphites
generally are higher than the other grades. In addition, the metal im-
pregnated carbon-graphites have low porosities and permeabilities, are
strong and are compatible with steam. With the exception of their tendency
to exude in the presence of water, some of the oxidation inhibited, high
temperature grades also are attractive. They have good friction and wear
characteristics and low values of porosity and permeability.
Additional possibilities involve composites of graphite with rela-
tively soft oxides. A cadmium oxide-graphite mixture has friction coef-
ficients in the order of 0.1 or less over most of the temperature range
( 26^
from 100 to 1000°F v '. The cadmium oxide is presumed to improve the
adherence of graphite to the surface and, hence, to improve the lubricity
effectiveness.
Materials that would be satisfactory as mating materials in contact
with the carbon-graphites under dynamic conditions are AISI 440C or hard
chromium plate with the latter being limited to service temperatures of
less than 700°F. A high hardness is desirable, i.e., greater than
Rockwell C 45, and the surface finish should be smooth but not too smooth
so as not to be able to hold the transfer film. A finish on the order
of 8 - 12 RMS appears satisfactory.
6.1.2.1.2 Polytetrafluoroethylene (PTFE)
Plastic bearings made from polytetrafluoroethylene (PTFE) are
available for nearly every type of application. They have been used
successfully as inserts to plain bearings, as reinforced thin sheets on
plain spherical bearings, and as the retainer material for ball bearings.
PTFE is one of those materials that does not have a laminar layer-lattice
structure and yet has an exceptionally low coefficient of friction. The
low friction has been attributed to its low surface energy. This results
in very weak adhesion and shearing takes place primarily at the inter-
(27)
face rather than in the bulk material . It also has the necessary
.137
-------
characteristic of establishing an adherent transfer film on most rubbing
metal surfaces. The major disadvantage of PTFE is its low compressive
strength which results in cold flowing under relatively low loads and
limits its use to temperatures on the order of 400 - 450°F (it also begins
to thermally decompose at about 540°F). Other disadvantages of PTFE are
its poor thermal conductivity and high coefficient of thermal expansion.
To improve the flow characteristics of PTFE, PTFE composites have
been developed. The three most common materials that have been used to
reinforce the PTFE are bronze, glass and carbon-graphite. In one recent
(28)
symposium of PTFE seals for reciprocating air compressors , it was
generally considered that no one filler material has demonstrated superior
performance over the others and that lack of consistency in the composites
was a major factor in the erratic performance. Three typical compositions
that are in use for seals in reciprocating compressors are:
1. PTFE + (15% Bronze + 5% MoS,2)
2. PTFE + 20% Glass
3. PTFE + 10% Carbon
(% by volume)
(29)
G. L. Griffin has suggested that non-uniformities in the struc-
ture as a result of molding practice of filled polymers could account for
differences in performance. In particular, the surface of molded com-
posites is usually deficient in filler material and this could result in
varying friction and wear behavior.
Other developments to improve the poor creep properties of PTFE
are the use of fibers of PTFE. Compacts can be made from PTFE fibers
which have 25 times the' tensile strength of conventional PTFE compacts
made from powders.
Halliwell of the U.S. Navy Marine Engineering Laboratory has con-
ducted tests with PTFE reinforced with metallic filament windings
in an attempt to develop a piston seal for high-pressure air compressors.
The filament winding approach has resulted in a superior reinforced ma-
trix as compared to randomly dispersed particles and fibers previously
reported. This new approach to the problem has resulted in extended
compressor life with low rates of wear and leakage. Halliwell found that
138
-------
reliable compressor operation could be obtained at 5000 psig for periods
beyond 1000 hours using such seals in lieu of conventional split rings.
The operational life of PTFE rings in a reciprocating steam expander will
be affected by the quality (dryness) of the steam. Wear rates of PTFE
decrease with dryness until a dewpoint of ^ -40°F is reached whereupon
the wear rate increases drastically. However, small amounts of condensed
water will cause an increase in wear rates and should be avoided.
Polyimides are the next generation of lubricating plastics pushing
the use temperature from the 400° - 450°F for PTFE to 500° to 600°F.
The polyimides also have superior friction and wear properties. Unfor-
tunately the polyimides cannot be used in high temperature (> 212°F)
steam or water because they will hydrolize.
6.1.2.1.3 Metals
In some applications, metals are used on bearing surfaces to achieve
improved friction and wear characteristics. The metals (and alloys) used
in these applications can be categorized in three classes: (a) classical
layer-lattice crystal, (b) low shear strength and (c) soft oxide formers.
(31)
Buckley and Johnson^ ' investigated the influence of crystal structure
of metals on their friction and wear behavior and found that relatively
low coefficients of friction could be obtained for cobalt in which the
hexagonal crystal structure is stabilized. A coefficient of friction of
0.3 was measured for an oxide-free 25% Mo-Co alloy at speeds of 2000 ft/
—9
min (750°F; pressure of 10 torr). The coefficient of friction of cubic
cobalt under similar conditions is on the order of 0.7.
A number of metals that do not have the layer lattice structure
are still widely used in bearing applications. These metals are listed
(32)
in Table 6.1-III; they are soft and shear easily . For example, the inetals
lead, tin, indium and silver are used in many common plain bearings, i.e.,
piston pin bearing. However, the coefficients of friction of these
metals below their melting points are relatively high (> 0.3). When
used above their melting points (liquid film) friction coefficients are
considerably lower, on the order of 0.17 for lead.
139
-------
Table 6.1-III
Soft Metals Used in Bearings
Metal Moh hardness Melting point,
°C
Indium
Thallium
Lead
Tin
Cadmium
Gold
Silver
Platinum
Rhodium
1
1.2
1.5
1.8
2
2.5
2.5-3
4.3
4.5-5
155
304
328
232
321
1063
961
1755
1955
A commercial material incorporating a soft, low shear strength metal
(33)
for lubrication is the Bishiralloysv . These alloys are produced by
pressing and sintering alloy powders, followed by impregnation with lead
and finally rolling to size. The compositions and hardnesses of these
materials are given in Table 6.1-IV. Alloy C is reported to have the best
wear resistance; the coefficient of friction is on the order of 0.2. A
continuous film of lead is maintained on the surface by flow of the lead
in the pores of the alloy. However, because of the low melting point
of lead (621°F) the maximum ambient service temperature for the Bishiralloys
is about 575°F.
The metal gallium is a liquid at 86°F and effective boundry lubrica-
tion is obtained with gallium rich films on AISI 440C stainless steel.
The coefficient of friction of gallium coated AISI 440C vs. AISI 440C at
500°F in air was measured to be 0.104 (sliding speed 390 ft/min, load
1000 grams)(34) .
Oxidation of metals will result in significant reductions in the
(35)
coefficients of friction and protection of the metal surfaces . The
transition temperatures for improved friction and wear behavior were
determined for iron, copper, nickel, molybdenum and chromium to be 100° -
200°F, 400° - 500°F, 1200° - 1400°F, 800° - 900°F, 800° - 1100°F, re-
spectively. Coefficient of friction values of 0.2 - 0.3 were obtained
140
-------
Table 6.1-IV
Chemical Composition and Hardness of BISHISALLOY
Basic Composition of BISHIEALLOY Matrix _ „. „ ,.
wt « wt % Range of
'
BISHIRALLOY-A
BISHIRALLOY-B
/
Fe Cu Ni Mo Cr C
Bal 5.0 — — — less than
0.2
Bal 2.5 1.5 2.5 — less than
0.2
.impregnated x.eaa naraness
in BISHIRALLOY Brinell
15 ~ 30 75 - 95
15 ^ 30 105 - 135
BISHIRALLOY-C Bal 2.5 1.0 1.0 5.0 less than
0.2 15 ^ 30 65 - 95
-------
upon the formation of the oxides in comparison to values of 0.5 - 1.0
for clean surfaces at lower temperatures. The formation of so-called
soft oxides MoO , WO , Cu 0, ZnO, Co 0 , PbO, CdO were the most effective
in preventing surface damage. The lowest friction coefficient was ob-
tained with PbO.
A commercial product is available which reportedly provides low
friction and wear characteristics as the result of the properties of its
oxide surface film. The material is Clevite 300, an iron base alloy con-
taining cobalt and molybdenum. The alloy is produced by powder metallurgy
techniques and its properties are listed in Table 6.1-V.
6.1.2.1.4 Sulfides/Selenides
The sulfides and selenides of molybdenum, tungsten, tantalum and
columbium have the layer-lattice structure similar to graphite and have
extremely low shear strengths. Further, they do not require water vapor
or oxygen for their lubricating ability. Another difference between the
sulfides or selenides and graphite is that the weak easily sheared bonds
occur only between every third layer of atoms instead of every layer as
is the case for graphite. In MoS_, for example, each layer of molybdenum
atoms is strongly bonded to the adjacent layers of sulfur atoms on each
side but the layer of sulfur atoms is weakly bonded to the sulfur atoms
in the next adjacent layer so that shear occurs through the weak sulfur-
sulfur bonds. Recent studies on the lubricating mechanism of MoS^ were
(29)
reported at the solid lubrication conference in Denver
Since MoS» occurs in nature, it is considerably less expensive than
the other sulfides and the selenides or tellirides which are synthetically
produced. For this reason most of the common solid lubricant composites
based on this group are made from MoS_. The effectiveness of MoS? in air
is limited to about 750°F where oxidation begins to effect its lubricity.
The oxidation products of MoS. are MoO and a sulfur compound which can
result in abrasion and corrosion and an increase in the coefficient of
friction at the lower temperatures. (It should be noted that at tempera-
tures above 1300°F - 1400°F, Mo03 can be an effective lubricant; the melting
point of Mo(>3 is 1463°F).
142
-------
MECHANICAL PROPERTIES
Table 6.1-V
Properties of Clevite 300
•e-
co
Temp. (F)
75
600
900
1200
FRICTION
Yield
Strength, psi
70,500
-
62,250
48,000
AND WEAR PROPERTIES
Ultimate
Strength, psi
106,000
107,000
95,800
71,000
Elongation, %
-
2.8
2.7
6.6
Elastic
Modulus
psi x 106
28.4
25.8
24.4
22.5
Compression
Strength, psi
210,000
-
-
-
Hardness
RA
66 (31
57
-
35
Re)
(Based on face-seal tests of Clevite 300 against itself at 18 psi load and 150 fps surface speed,
in air.)
Temp.. °F Coefficient .of Friction, "f" Wear Rate, in/hr
500
750
1000
1200
0.072
0.063
0.052
0.047
0.00073
0. 00062
0. 00035
0. 00030
-------
Composites based on the solid MoS7 lubricant have been developed by
(37)
Boeing and described by Devine . The composites developed by Boeing
were quite brittle and even the most successful composite, 90% MoS2~8% Fe-2%-Pt,
would have strength problems because of the brittle iron sulfides that form at
the sintering temperatures. Piston rings have been produced from iron bonded
MoS_ and evaluated by Midwest Research(38) . The solid composite called "Navy
(37)
Lube"v is a MoS2-Graphite-Sodium Silicate (NaOSi02) and is not only weak
but the NaOSiOp is hygroscopic and would not be suitable for use in the steam
environment.
Another series of composites based on MoS_ and WS0 are described
(39)
by Hopkins ' . -Three of these composites have shown low wear rates and
low friction in air. The compositions of these materials are:
1. 53% WS2 - 12% Co - 35% Ag
2. 80% MoS2 - 20% Ta
3. 50% MoS2 - 38% Ta - 12% Fe
Overall, the use of solid lubricant composites with MoS_, or other
disulfides or diselenides, as the base do not appear attractive for use
as piston rings in steam expanders primarily because of their low strengths
and fragile nature.
6.1.2.1.5 Porous Metal Composites
In order to utilize the good lubricating qualities of some of the
better solid lubricants that are too weak to be used as the base of
self- lubricating composites, porous metal composites have been developed.
These composites are generally produced by powder metallurgy techniques
in which the metal or metal alloy and the lubricant are blended, pressed
and sintered to form a metal matrix throughout which are discrete pockets
of solid lubricant. In one investigation, it was found that a minimum
concentration of 5% of the solid lubricant was required to achieve a
transfer film and thus low wear rates. Friction coefficients of 0. 2 were
measured*- '. On the other hand, too high a concentration of the lubricant
results in a weak structure. Examples of this type of porous metal com-
posite are: (a) 5-15% MoS2 in a 95% Ag - 5% Cu matrix, (b) 5-40% MoS2
in a nickel or 80% Ni - 20% Cr matrix, (c) the commercial Molalloy - MoS2
in a refractory metal (Ta/Mo) matrix.
144
-------
Another method of producing the porous metal composites is to prepare
a metal skeleton with controlled porosity with respect to volume and
distribution and impregnate the pores with the lubricant by application
of pressure and heat. Metal skeletons with up to 65% void volume have
been produced with Inconel 600, Inconel X-750, Hastelloy X, Nichrome V
alloys in which PTFE, graphite, MoS have been impregnated. A Inconel
600 (60 - 65%voids) impregnated with PTFE has performed well in seal
applications at temperatures up to 500°F.
The metal fluorides have shown promise for use as lubricants for
high temperature application in air. Sliney at NASA Lewis Research
Center has investigated the use of fluorides for a number of years ' ' .
Coefficients of friction of self-lubricating porous metal composites (40%
void volume in Inconel X-750), vacuum-impregnated with fused fluoride
eutectic of 62% BaF - 38% CaF« were measured to be less than 0.1 at
1000°F and at a sliding velocity of 2000 ft/min while under load. The
coefficient of friction increases as the temperature and speed decrease;
at 500°F and a speed of 1000 ft/min the coefficient of friction is approxi-
mately 0.3. More recent studies have shown that coefficients of friction
of 0.2 + 0.05 could be obtained at speeds as low as 500 ft/min over the
temperature range of RT to 1700°F. These data were obtained with a porous
nickel alloy impregnated with a high temperature enamel (NBS-418) and
an overlay of the BaF« - CaF eutectic.
Finally, this approach is compatible with classical piston ring
materials technology where gray cast irons are used because of the ex-
cellent lubricity imparted by flakes of graphite. Similarly, the nodular
Ni-Resist alloys of 20 - 30% nickel content with its spheroids of graphite
has been shown to have good friction and wear characteristics in conjunc-
tion with good corrosion resistance in superheated steam. It is believed
that the use of strong porous metal composites impregnated with a solid
lubricant of low shear strength and ability to establish a transfer film
is one of the most attractive materials for application as piston rings
in reciprocating steam expanders.
*
6.1.2.1.6 Hard Surfacing Materials
In applying the adhesion theory in friction and wear, relatively
low coefficients of friction are achieved between two sliding hard ma-
145
-------
terials that have no affinity for each other, i.e., materials with little
or no mutual solubility or tendency to form intermetallic compounds. The
hard, wear resistant surfacing materials fall into this class of materials
and warrant investigation.
Plasma sprayed coatings are of interest because they can be readily
applied. Koppers has plasma sprayed rings with their K-1051 coating
(metal bonded Cr«C2) for use in lubricated, reciprocating
steam expanders running against nitrided steel or AISI 440C stainless steel
which was heat treated to a hardness of Rockwell C 50-55. Since the
hard metal carbides are metallic in nature with respect to their atomic
bonding, a more suitable materials combination with respect to "alloying
tendency" might be an oxide coated cylinder liner running against a
carbide coated ring. Possible combinations are Union Carbide Corp. LC-19
(Cr20 + A1o°3^ sPrayed on the cylinder wall and either LW-1 (Cobalt
bonded WC) or LC-1 (Nichrome bonded Cr C ) sprayed on the ring surface.
In dry rubbing tests between LC-1 (Cr3C_-15% Ni-Cr) and LA-2(A12C>3),
coefficients of friction of 0.15 - 0.27 were measured in the temperature
range of RT to 1400°F.
E*I. du Pont de Nemours and Co. have recently made available a
series of anti-friction and anti-wear LP alloys offering outstanding ad-
vantages. Metallurgically, the LP Alloys consist of hard grains of an
intermetallic compound with a Laves Phase structure dispersed in a softer
matrix, that provides good embeddability characteristics in contrast to
other hard facing materials. Currently preferred compositions contain
Laves-phase hexagonal close-packed intermetallic compounds of cobalt,
molybdenum and silicon (Co Mo~Si and CoMoSi) in a cobalt-rich matrix.
The Laves phase, depending upon its composition has a micro-hardness of
1000 to 1500 DPH and the matrix has a hardness of 200 to 800 DPH. Over-
all, the bulk hardness ranges from R 30 to R 60. Tests have shown
c c
that, under extreme loading and boundary lubrication conditions, LP Alloys
exhibit very little wear.
• LP Alloys are produced as medium and fine powders for plasma spraying
and powder metallurgy applications; they also can be cast as semi-finished
parts. The compositions (in wt. %) of LP Alloys are shown in Table 6.1-VI.
146
-------
Table 6.1-VI
Chemical Composition of IP Alloys
Co Mo Si Cr Vol. % Laves
LPA 100
LPA 200
LPA 300
LPA 400
35
70
45
62
35
28
48
23
10
2
7
2 8
65
20
98
50
In one application, the sleeve and roller bearings and the piston skirt
and rings of a 2-cycle engine were coated with LPA-100 and the engine
operated without oil in the fuel for 50 hours with satisfactory results.
Another new material of interest that is available from E.I. du Pont
de Nemours is the compound Ni B. The material is applied to the surface
by electro-chemical methods and has a microhardness of 900 - 1500 knoop
(R 60-70). Friction data of Ni_B against hardened steel show a coef-
ficient of friction of 0.1 at 600°F in air, unlubricated.
Other materials in this class that could be considered are: various
grades of SiC, Si.N, and TiC in the form of pressed and sintered compacts
or applied as coatings.
6.1.2.1.7 Mixed Composites
Multicomponent self-lubricating composites have been developed that
contain a metal matrix, a film former, and a load-carrying component ^^'.
Examples of these types of composites are Ag-PTFE-WSe2,Cu-PTFE-WSe2,
Ag-Bronze-PTFE-WSe2,Ag*Hg-PTFE-MoSe2. The latter material has shown ex-
cellant friction and wear characteristics at 600°F ^ \ The modified
PTFE melts at 590°F and is thought to provide the excellent friction and
wear characteristics.
Other mixed component self-lubricating composites that have been
investigated are: WSe2-Ga/In ^45\and. Graphite-WS -NaF ^6\
6.1.3 Bonded jSolid Film Lubricants
The primary disadvantage of the bonded solid film lubricants is
147
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that there is little stored lubricant in the film and as a result, service
lives are limited. The bonding agents are either organic thermoplastic
or thermosetting resins (phenolic, epoxy, polyimide) or inorganic metal
salt (sodium silicate) and fused ceramic (aluminum phosphate). Because
the bonded films have so little strength, the films are kept as thin as
possible, usually on the order of 0.0005 inch or less. For these reasons,
the use of bonded films is generally limited to high load/low sliding
velocities or low load/moderate velocities. Most commercially available
bonded solid film lubricants are limited in temperature by the bonding
agent. For example, at temperatures above 400°F, the organic bonding
agents tend to thermally decompose causing failure of the film because
of lack of adhesion. Commercial products of this type incorporating
various solid lubricants are:
Solid
Lubricant Bonding Agent Service Temp.. °F
MoS2 Epoxy 275 - 300
90% MoS2-10% Graphite Phenolic 300
PTFE Polyimide 450 - 500
A newer bonded film based on the polyimide for use at higher tem-
(47)
peratures was reported by Campbell and Hopkins . It consists of MoS
and Sb-O- dispersed in the polyimide binder (MLR-2) and it shows excellent
wear properties. However, it should be pointed out again that the
polyimide resins cannot be used in high temperature steam and water.
The metal salt and fused ceramic bonded films were developed for
use at temperatures generally above 600°F. Examples of these types of
bonded films were reported by Hopkins ™8^: MLF-5, is a sodium silicate
bonded MoS--graphite - Au film and MLF-9 is a aluminum phosphate bonded
MoS -graphite - Bi film. Ceramic bonded MoS -graphite films have been
used in excess of 700°F in air0 As mentioned previously sodium silicate
is hygroscopic and cannot be used on steam or water.
148
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The solid lubricants of interest that are being developed at Midwest
Research are the FSAL 28 (AlPO^ bonded Ba2F + Ca2F) and FSAL 29 (AlPO^
bonded Ba2F + Ca-F + Mg-F). These lubricants are applied as bonded films
and can be sintered at temperatures below the fluoride eutectic temperatures.
About 6 vol.% A1PO, has been found to be the optimum binder content.
Coefficients of friction of the FSAL 28 versus Inconel X in air are on the
order of 0.1-0.15 at 500°F and 0.10-0.12 at 1000°F. Loads have varied
between 6000 - 25,000 psi at speeds of "180 ft/min.
At the higher temperatures, >1000°F, PbO has been shown to exhibit
good friction and wear properties. Because PbO oxidizes to PbJ3, at
temperatures below 1000°F resulting in loss in lubricity, additions of SiO~
(49)
are made to PbO to inhibit the oxidation reactions . Very low friction
characteristics (0.1 coefficient) have been achieved with 10% Si02-PbO bonded
film at temperatures of 500° to 1200°F.
Another very promising new material that is being studied is graphite
fluoride, (CF ) . Coefficients of friction equal or superior to MoS9
3c n &-
and graphite were reported at temperatures up to 750°F in air. Superior
wear properties are attributed to greater adhesive qualities. Although
moisture is beneficial, it is not a prerequisite for low shear strength.
The use of bonded solid film lubricants does not appear to be as
promising for the lubrication of the cylinder walls in a steam reciprocating
steam expander as do the self-lubricating composites. It would appear
that if the bonded solid films were to be evaluated for this application,
they should be applied to the cylinder wall rather than the piston rings
in order to provide a greater amount of available lubricant for increased
life.
149
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6.2 Lubrication Technology Base - Liquid Lubricants
Lubrication of the large, slow speed reciprocating steam expander
operating under steam at relatively low temperatures and pressures has
been well established for years. On the other hand, lubrication of the
small, high speed and high performance steam expander operating at high
steam temperatures and pressures is considerably more complex and may be
more complicated than the lubrication of internal combustion engines.
Using conventional techniques of lubricating steam cylinders and valves
by injecting lubricants directly on the parts and/or by injecting the
lubricants into the inlet steam, two separate lubricating systems have
been used, one for lubricating the components in the crankcase and one
for lubricating the cylinder walls and valves that are operating in the
high temperature steam. This is accomplished through the use of a cross-
head piston design; both the General Motors SE-101 and SE-124 steam
engines employed this concept. A further complication is the fact that
the cylinder walls cannot be cooled without decreasing the efficiency of
the engine cycle so that the fluid lubricating the hot cylinder walls and
valves will come in contact with metal surfaces that are at temperatures
of 600° - 800°F. Thermal decomposition products of the lubricant resulting
from being in contact with metals at these temperatures will tend to foul
up the steam generator tubes unless the lubricant is separated from the
condensate water. Frequent lubricant changes would be expected.
150
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Successful use of solid lubricants to lubricate the cylinder wall/
piston ring and valve stem/valve guide interfaces will eliminate or greatly
minimize many of the problems encountered with liquid lubricants. In fact,
it may be possible to utilize a standard trunk piston design. However,
due to the fact that the engine cylinders cannot be cooled, as stated
previously, even the bulk temperature of the crankcase lubricant may be
exceptionally high in comparison to the crankcase temperature in con-
ventional internal combustion engines. Crankcase lubricant bulk tempera-
ture may be as high as 300°F unless a special cooling system is employed
to cool the lubricant. Even so, the limiting temperature may be the hot
spot temperature that the lubricant will see on the bottom surface of
the power piston head of the trunk piston design or the internal surface
of the piston rod in the crosshead piston design. The average bulk
lubricant temperature and maximum hot spot temperatures for the two engine
designs in this program are:
Max. Lubricant Temp.. °F
Engine Design Bulk Hot Spot
Trunk Piston 250 540
Crosshead Piston 250 425
Although these temperatures are not unusually high for some
conventional petroleum oils, it may be desirable to select a lubricant
capable of operating at higher temperature in order to achieve a longer
service life or some specific property. It may be possible to provide
a lubricant that only requires changing every two years or even the life
of the engine as is the case for oils in large steam turbines. A trade-
off can be made between the cost of the lubricant and the service life.
Hydrocarbon oils that are extracted from petroleum sources and
synthetic lubricants both were considered for use as the crankcase lubricant
in the steam expanders. In the selection of a suitable lubricant, it is
necessary to: first, select a suitable base and second to select the
proper chemical additives that are compatible with the base and the en-
vironment.. To be considered a good lubricant for the intended application,
the lubricant must have the following important characteristics:
151
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1. Good oxidation resistance
2. High thermal stability
3. Low volatility
4. Suitable viscosity
5. Low pour point
6. Adequate lubricity
7. High ignition temperature
8. Good hydrolytic stability
The last characteristic is particularly important with respect to
steam expander applications in that some contamination of the crankcase
with water is highly probable, particularly during start-up and shutdown.
For this reason, the lubricant needs better rust protection and much
better demulsibility than do current motor oils. Additives similar to
those used in hydraulic oils, i.e., calcium sulfonates could be used
effectively for rust protection. The partial organic acid esters and
phosphorus containing acid esters also are used for this purpose. Rust
inhibitors usually have a high polar attraction toward metal surfaces
and form a continuous protective film over the surfaces. Care must be
exercised in the selection of rust inhibitors to avoid corrosion reactions
with nonferrous metals or emulsions with water. Chemical additives are
rarely effective in improving the hydrolytic stability of lubricants and
are seldom used to improve this property; hydrolytic stability is an
inherent quality of the base lubricant. Hydrocarbons have excellent
hydrolytic stability and major concern is with the additives and with
synthetic lubricants. The straight alkyl-chain compounds generally are
more easily attacked by water than the highly branched structures. For
example, the esters have long chains and generally have poor hydrolytic
stability at temperatures of 400°F and for this reason the esters will
not be considered as a base for this application. Hydrolysis and water
contamination can result in the following reactions: (a) change in physical
properties, (b) generation of sludges and other insoluble compounds,
(c) decrease in solubility of essential additives, (d) corrosive attack,
(e) release of volatile compounds due to chemical breakdown. Obviously
all of these changes are detrimental and must be kept to a minimum. It
is expected that a complete reformulation of the crankcase lubricant will
be required because most of the extreme pressure agents and inhibitors
currently in use have little tolerance to water.
152
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Because the bulk temperature of the crankcase lubricant is expected
to be higher than the temperatures in the current internal combustion
engines, oxidation stability of the lubricant must be superior to the
current premium-grade motor oils. Oxidation of the hydrocarbons results
in the formation of oil-soluble acid compounds which increases the vis-
cosity of the oil and can be corrosive; oxidation of the synthetics are
more complex depending upon the type of structure, i.e., some form volatile
gases and do not change the viscosity. Additives to inhibit oxidation
include sulfur and phosphorus compounds and the amine and phenolic com-
pounds.
Other additives which are believed to be required in the lubricant
are defoamants and anti-wear additives. The most common defoamant is the
organo silicon oxide polymer and only small concentrations are required
(1 - 20 ppm) to inhibit foam formation caused by the action of the engine
crankshaft. Anti-wear additives are required to minimize friction and
wear in highly loaded components such as the cam/tappet interface where
Hertzian stresses can be as high as 200,000 psi. Extreme-pressure (EP)
additives are used and they react with the rubbing metal surfaces to form
a lubricating film which protects the metal surface when the lubricating
oil film is lost. The most common EP additives are zinc, phosphorus,
sulfur and chlorine compounds.
In the survey for a suitable crankcase lubricant for the steam ex-
panders to be tested in this program, one of the more attractive solutions
appeared to be tne use °f a highly refined turbine oil. Many of these oils
are the result of extensive development work at the Pennsylvania State
University, where processes have been developed for "super-refining"
mineral oils to achieve lubricants which have lower pour points, better
viscosity-temperature characteristics and higher thermal and oxidative
stability. Within the family of super-refined mineral oils, it is possible
to use either paraffinic or aromatic-base oils. There are distinct dif-
ferences in the physical properties of these two types of hydrocarbons.
Because of their intermolecular spacing, the paraffinic oils are more com-
pressible,' show less change in vfscosity with temperature, have lower
densities and lower pour points.
153
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The production of super-refined mineral oils is essentially a clean-
up operation where the oils are carefully dewaxed, many of the polar im-
purities are removed, and selective catalytic hydrogenation is used to
minimize unsaturation. This process removes not only undesirable im-
purities, but also naturally-occurring oxidation inhibitors and polar
compounds, which are essential for boundary lubrication. Thus, it is
necessary to compound these oils with oxidation inhibitors and anti-wear
additives. Since thermal decomposition will also proceed by the forma-
tion of free radicals, a small percentage of an aromatic disulphide was
found to be effective in protecting the oil. The use of the aromatic
disulphide gives an added bonus since it will also act as an anti-wear
agent by surface reaction between the metal and the sulphide.
The use of blending techniques opens up many possibilities as far
as the control of properties such as viscosity and pour point. A wide
range of properties can be obtained by blending suitable base stocks to
achieve the bulk viscosity and pour point required for the application.
A common turbine oil that has proven to have satisfactory service
(*>» 20 years) at temperatures < 250°F in steam turbines and generators is
produced by a number of refining companies and is known as Teresso 65,
Industrial Oil 61, Tellus 69 or DTE Extra Heavy. The oil has oxidation
and rust inhibitors but no EP anti-wear additive which would have to be
formulated. Characteristics of the oil are as follows:
Flash point, deg F, min 420
Viscosity, Saybolt Universal, sees 100 deg F, min 540
max 700
Viscosity index, min • 85
Pour point, deg F, max +25
Neutralization value (total acid number)
mg KOH/g, max 0.80
Oxidation stability test, hours, min (a) 1000
Sludge, %, max (b) 0.10
Rust prevention test Shall Pass
(at)
Test carried out to total acid number of 2.0 mg KOH/g.
In oil taken from oxidation test.
154
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The major disadvantages to the use of this oil are the relatively
high pour point and the low hot spot temperature capability. Although a
modified turbine oil would probably be satisfactory for the crosshead
piston design, it may not be satisfactory for the trunk piston design be-
cause of the anticipated 540°F hot spot temperature.
The synthetic hydrocarbon lubricants have about a 100°F advantage
over the mineral oil based turbine oils and appear to have the necessary
properties for the application. For relatively long time service, the
following generalized maximum service temperatures can be stated:
Base Max. Bulk Oil Temp.. °F Max. Hot Spot Temp.. °F
Mineral Oil 250° - 275° (possibly * 500
300°F)
Synthetic Hydro- 350° - 375° (possibly *> 600
carbon 400°F
A special synthetic hydrocarbon formulation (XKN-1301-C) was pre-
pared by Mobil Research and Development Corp. for evaluation in the
program. The lubricant has suitable oxidation, rust and foam inhibitors
and an anti-wear additive. The properties are given in Table 6.2-1. The
lubricant has a low pour point (< -65°F) and a viscosity in the SAE 30
range. However, its high viscosity index (149) puts it in the SAE 40
range at elevated temperature. The lubricant appears to have good anti-
foam performance and good rust protection and although it forms a small
amount of emulsion, no emulsion problem would be expected in the engine.
A general comparison of various other synthetic lubricants is made
in Table 6.2-II . Known disadvantages of some of these synthetic lubri-
cants include poor hydrolytic stability of the esters and poor viscosities
and high pour points of the polyphenyl ethers (in addition to their high
cost). Projected costs of the synthetic hydrocarbons are not expected
to exceed current premium grade motor oils.
The ESSO Research and Engineering Co. (Govt. Research Lab) is under
contract to Steam Engine Systems Corp. to develop a suitable lubricant
for the lubrication of cylinder walls in a reciprocating steam expander
(52)
operating at 1000°F v '. In their work they found the inhibited blends
of hydrorefined paraffinic distillates and residua to be the most suitable
155
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Table 6.2-1
Properties of Experimental
Synthetic Hydrocarbon Lubricant
XRN 1301 C
Gravity, °API (ASTM D-287) 37.4
Specific Gravity 0.837
Pour Point, °F (ASTM D-97) <-65
Flash Point, °F (ASTM D-92-1) 460
Viscosity: cs @ 210°F 11.05
SUS @ 210°F 63
cs @ 100°F (ASTM D-445) 74.5
StS @ 100°F 345
cs @ -40°F 34,000
Viscosity Index 149
Acid Number (ASTM D-664-1) 0.08
Base Number (ASTM D-664-3) 0.11
Foam (ASTM D-892)
Sequence I Tendency, ml 330
Stability, ml 0
Sequence II Tendencey, ml 40
Stability, 0
Sequence III Tendency, ml 380
Stability, ml 0
Rust Test, 48 Hrs, Distilled Water (ASTM D-665-2) Pass
48 Hrs, Syn.Sea Water (ASTM D-665-4) Pass
Emulsion Test (ASTM D-1401)
Total Emulsion, ml 5-10
Time for 3 ml, Minutes 31-33
Time for Complete Break, Minutes 34-37
Water Trace, ppm 40
Surface Tension, Dynes/cm 30.7
Panel Coker, 24 Hrs @ 600°F, Deposit, mgs 27
Mobil B-10 Catalytic Oxidation, 40 Hrs @ 260°F
% Viscosity Increase @ 210°F 1
Lead Loss, mg 1.4
Sludge Nil
Neutralization Number Increase 0.2
156
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Table 6.2-1 (Cont'd)
Properties of Experimental
SynthetIc Hydrocarbon Lubricant
(a)
Mobil B-10 Catalytic Oxidation, 40 Hrs @ 325°F
% Viscosity Increase @ 210°F 17
Lead Loss, mg 17.3
Sludge Nil
Neutralization Number Increase 2.8
Mobil B-10 Catalytic Oxidation, 72 Hrs @ 325°F
% Viscosity Increase @ 210°F 230
Lead Loss, mg 225
Sludge Trace
Neutralization Number Increase 12.4
SAE Wear Test, Steel-on-Steel, 30 Min @ 150 Ib Load,
250°F Oil Temperature
Total Weight Loss 0.018 mg
Almen Load, psi 8000
Mobil Thin Film Oxidation Test -600°F
(100 = Clean)
Mobil Thin Film Oxidation Test -625°F
(100 = Clean)
93
69
(a) Mobil Research and Development Corporation
157
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Table 6.2-II
Performance of Selected^ Types
of Synthetic
Probable
Maximum Thermal Resistance Extreme
Stability Limit, to Water Pressure
Class of Compound °F Degradation Lubricity
Polyglycols 600 Excellent Good
Phosphate esters 800 Good Excellent
Dibasic acid esters 600 Good Good
Chlorofluorohydrocarbons 600 Excellent Excellent
Silicones 900 Poor Poor (Steel)
Silicate esters 800 Poor Fair
Fluoroesters 600 Good Poor
Neopentyl polyol esters 600 Good Good
Polyphenyl ethers 600 — Good
Silanes 700 — Poor
158
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of the petroleum base oils. A combination tricresyl phosphate (TCP) and
ethyl-702 appears to offer the best anti-wear and anti-oxidant properties
in the selected bases. The candidate lubricant (designated 2415-50-1) has
a higher viscosity (SUS) rating at the standard temperatures of 100°F
and 210°F, 663 and 74 respectively, but a lower viscosity index (102) than
the Mobil XRN-1301-C lubricant. However, as discussed previously, there is
no requirement to lubricate the cylinder walls or valve parts in this program.
6.3 Lubricant Recommendations
6.3.1 Solid Lubrication
In the steam expander design, there are three areas requiring solid
lubrication; these include the inlet valve, the piston/cylinder and the shaft
connecting the power piston with the crosshead piston. The inlet valve stem
will slide in a guide that is exposed to steam and that is sealed by means of
compression rings. The shaft connecting the power piston to the crosshead
piston will slide in a seal that is exposed to steam on one side and oil on
the other. The most severe lubrication problem is the piston/cylinder wall,
and either the power piston rings or the liner can serve as the lubricant.
Although the qualitative correlation of the coefficient of friction and
wear rate is generally recognized, the exact correlating function between the
two is difficult to obtain. Theoretical relationships may be developed which
are useful in ranking material and lubricant combinations. However, the
reduction of the theoretical relationships to useful design parameters by
testing at the expected service conditions is necessary.
Friction effects arise from the tangential forces transmitted across
the interface of contact between two bodies. The wear phenomena consists of
the removal of material from the contacting surfaces. Adhesion, the ability
of contacting bodies to withstand tensile forces after being pressed together,
is the primary interacting phenomena. The concept of surface energy makes
it possible to develop parameters for making rudimentary predictions of the
performance of specific materials under sliding conditions.
159
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While 90% or more of the resistance to sliding arises from the need
to shear strongly adherent surface atoms of the contacting materials, other
factors have to be considered. A roughness component arises from the necessity
of lifting one surface over the other's asperities, which usually contribute
about 0.05 to the friction coefficient. Hydrodynamic lubrication can eliminate
this component (liquid phase lubrication). A plowing component is present
when a hard, sharp surface digs into a softer surface and produces a groove.
For really rough surfaces this term may be large, but otherwise it is usually
negligible.
Adhesive wear is best combatted with hard materials with low interaction
tendencies (or one member of the pair a non-metal), or by use of two metals
with a high tendency for interaction reduced by a good boundary lubricant.
The use of hard materials does not in itself produce significantly lower wear
rates, but the probable low alloying tendency may produce a factor-of-10
change in the wear rate.
Although adhesive wear (sometimes called galling and scuffing) resulting
in welding of the surface asperities is usually associated with metallic
materials, the adhesion theory is also applied to solid polymers. The
difference between the metals and plastics is that the deformation in plastics
may be partly elastic over a wide range of load. Also, the flowing term
and surface roughness are of greater importance in plastics. The mechanism
of wear with carbon-graphites is primarily abrasion which occurs when asperities
of two moving surfaces touch and wear fragments are formed from one or both
surfaces. Grooves are generally plowed in the "softer" carbon graphite.
Based on the current technology of solid lubrication and wear resistant
materials that was discussed in Section 6.1 and in the preceding paragraphs,
a list of candidate material combinations for use in sliding contact was
compiled for the piston ring/cylinder application and is tabulated in Table
6.3-1. The primary selections were a metal impregnated carbon-graphite
material for the piston rings rubbing against an AISI 440C SS or hard
chromium plated steel cylinder liner. The same material combination was
selected for the back-up design for the piston rod seal in the crosshead piston
design. The primary selection for the rod seal was a 15% bronze + 5% MoS,,
filled PTFE material rubbing against a hard chromium plated H-ll alloy rod.
160
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Table 6.3-1
CANDIDATE PISTON RING/CYLINDER LINER COMBINATIONS
Ring
Liner
Hard Cr Plate or AISI 440C SS
Carbon-Graphite, Sb Metal Impregnated
1. Carbon-Graphite, Sb Metal
Impregnated ^a'
2. LC-1C (Cr3C2 + 15% Ni Cr) or
K-1051 (Cr3C2 Cermet)
3. Graphite 5890 Graphite 5890
4. Hast X Porous Structure (35 v/o)Hard Cr Plate(b)
+ Graphite Filled
5. Hast X Porous Structure(35 v/o) Hard Cr Plate^
+ NBS418 + (BaF2 + CaF2)
Eutectic Filled
6. LPA-100 (Laves Phase - 65 v/o) LC-19 (Cr203 + A1203) or Nitrided
(a)
+ 20% Ni
7. Ni3 B
8. LW-1 (WC + 9% Co)
9. LC-1C (Cr3C2 + 15% Ni Cr) or
K-1051 (Cr3C2 Cermet)
10. LSR-1 (TiC) over AISI 440C
11. K-35 (Ni-Resist) or H.S.31
12. K-35 (Ni-Resist)
13. K-35 (Ni-Resist)
14. LW-1 (WC + 9% Co)
Steel or Ni-Resist D3
LC-19 (Cr203 + A1203) or
AISI 440C SS
LC-19 (Cr203 + A1203) or Nitrided
Steel or Meehanite
LC-19 (Cr203 + A1203)
LC-19 (Cr203 + A1203) or Nitrided
Steel
LC-19 (Cr203 + A1203)
Hard Cr Plate + MoS2 Bonded Film
Hast X Porous Structure (35 v/o)
+ Graphite Filled
Hast X "Porous Structure (35 v/o)
+ (BaF2 + CaF2) Eutectic Filled
(b)
(a) Primary recommendation for first engine.
(b) Limited to temperature below 700°; for temperatures in excess of
700°F,, -one of the following liner coatings can be used:
• LC-19 (Cr203 + A1203) or LC-4 (Cr203)
• LC-K (Cr3C2 + 15 Ni Cr)
• Chromize
161
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A hard surfacing material (LPA-100 + 20% Ni) and Ni-Resist (D-3 type)
cast iron materials combination was selected for the valve stem ring/valve
guide components with cold worked Haynes alloy 25 as a back-up material
for the valve stem ring.
6.3.2 Liquid Lubrication
The Mobil synthetic hydrocarbon oil designated XRN 1301-C was
selected for use in the first single cylinder steam expander test. This
oil was specially formulated by Mobil for use in this program. The synthetic
hydrocarbon base was selected because of its excellent thermal stability,
temperature-viscosity characteristics and low pour point. Appropriate
non-metallic inhibitors have been added to the base to provide the necessary
oxidation resistance, rust protection, anti-wear properties and defoatnant
characteristics. Properties of the oil are shown in Table 6.2-1. In
general, the XRN 1301-C oil should have approximately a 100°F advantage in
use temperature (bulk oil temperature and hot spot temperature) over the
oils based on the petroleum base stocks. Probably the greatest disadvantage
of the XRN 1301-C oil over the natural hydrocarbon oils is its poorer
demulsibility characteristics. However, it is believed the demulslbility
of the XRN 1301-C is satisfactory for use in the steam expander.
One of the major reasons for not selecting the inhibited hydrofined
paraffinitic oil (2415-50-1) being developed for SES by ESSO Research and
Engineering Company for use in the program was the fact that it was not
fully developed. Further changes were anticipated to be made in the anti-
rust inhibitor; also, a pour point depressant was to be added later to lower
the relatively high pour point of + 20°F. Upon completion and evaluation
of the final formulation, the 2415-50-1 type oil should be reconsidered
for evaluation as the crankcase lubricant. No existing commercial liquid
lubricant was found to satisfy the crankcase lubrication segments of the
reciprocating steam expander.
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7.0 TEST FACILITY
7.1 Facility Description
A schematic representation of the steam expander test facility is
shown in Figure 7.1-1.
The steam generator was an electrically heated once-through
boiler-superheater which was designed to provide 1000 psi, 1000°F,
superheated steam at a rate of 400 Ib/hr. Three hundred and eight (308)
feet of 3/8" OD x .031" wall, Type 316 stainless steel tubing was
coiled on a two (2) ft. diameter with a pitch of I'1 to form the once-
through boiler and superheater section. Calculated pressure drop
through the steam generator at rated output was 80 psi. The coil
terminates in a small vapor drum which discharged steam through a 1"
schedule 40 pipe to a throttle valve and then to the expander. The
steam generator was equipped with a pressure relief valve, and was code
stamped in accordance with the ASME Power Boiler Code. Design pressure
was 1100 psi, and design temperature was 1200°F. Two separate three
phase, saturable core reactors (180 KWe) controlled power to the steam
generator coil. Automatic and manual modes of power control were
provided.
A by-pass control valve and desuperheater permitted checkout of
the test facility prior to installation of the test expander. The by-
pass valve and expander throttle valve also permitted inlet steam flow
control to the expander with a constant boiler heating rate. This form
of control was used during startup and shutdown of the expander.
A General Electric Company Model TLC-65/50 HP, 2500/750 RPM D.C.
type dynamometer was used to load (up to 65 HP) the steam expander. It
was capable of motoring (operating as a DC motor) up to 50 HP. Speed
regulation of ± 1% of rated speed was maintained during both modes of
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STEAM EXPANDER
TEST STAND
^.atil-a
TEMP
BECMW»TEST ENGINE
I -""• 'i — i«
i.4 ^ i n-yxyyQ
I l : i11! «1-»|-:^:
• MINI i
?] _*""i|KK>
V r TT'U«Bff*»=
.Ai-TT--'
Tin* mr m.',.*!,—.
,o.ors,. V ^P^OPS, TyC-lT''"^ ,, PUMP* ,f
RESIN OUTLET
DEMINERALIZEO
H2O STORAGE TANK
^-PACKAGED STEAM SYSTEM
Figure 7.1-1. Steam Expander Test Facility Schematic (GE Dwg. 7O7E676).
-------
operation. The dynamometer was capable of delivering up to 500 ft.lbs
of torque, and was used to rotate the expander during checkout of the
expander and during startup on steam operation.
The steam expander was coupled to the dynamometer through a flexible
Spicer shaft with universal joints at each end and a splined telescoping
center section. A shear pin adapter at the dynamometer shaft connection
limited inertial torque loading of the expander by the massive dyna-
mometer in the event of "freeze-up" or other failure which might
suddenly stop the expander. The dynamometer was mounted on a separate
foundation from the expander, and was isolated from the floor of the
test cell for minimum vibration. An electronic (strain gage type load
cell) system was used for torque measurement.
The expander exhaust was piped to a water cooled condenser which
was capable of 400 Ib/hr throughput of steam, and which could maintain
20 psia discharge pressure. Steam discharge from the desuperheater
also entered the condenser during by-pass operation.
A water storage tank equipped with a liquid level indicator, fill-
ing system and oil removal system received the condensate from the con-
denser. Demineralized makeup water was supplied to the tank directly
from the demineralizer as required during operation. Open loop opera-
tion could be conducted by discarding the condensate and by supplying
demineralized water for boiler feed. Most of the expander testing was
done in that manner.
A plunger type motor driven pump with automatic stroke adjustment
supplied the boiler with water. A small heat exchanger ahead of the
pump provided cooling to prevent pump cavitation. The pump was capable
of providing up to 1.0 gpm at 1200 psi to the boiler. A built-in pres-
sure relief valve prevented an overpressure of the pump. Pump discharge
pressure and flowrate were controlled automatically by sensing boiler
discharge pressure. Steam flowrate was determined by measuring boiler
feed water flow. A turbine type meter was provided for water flow
measurement.
165
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A control console located in the control room contained all the
operating controls and instrumentation for both the facility and steam
expander. All critical parameters were recorded. Protective circuits
for overspeed, overtemperature, and overpressure were provided. Tem-
perature, pressure or expander speed above set limits would result in
an automatic shutdown of the facility.
Water and steam samples were taken from the test loop for analysis
at several locations. An oil/steam sampling valve was provided at the
outlet of the expander. An oil/water removal valve was provided near
the top of the pump head tank, and a water sampling valve was provided
near the inlet to the boiler. All of these valves provided different
methods of removing fluid samples to be analyzed for oil or other
contaminants.
2
Figure 7.1-2 shows the I R boiler with a portion of its outer
insulated casing removed. Figure 7.1-3 shows the high pressure plunger
type water pump in the foreground with condenser and water supply tank
toward the rear.
Figures 7.1-4 and 7.1-5 show the crosshead expander installed and
instrumented ready for testing.
Considerable difficulty was experienced with the Spicer shaft and
couplings (reference Figure 7.1-4) between the expander and dynamometer.
Several shaft failures were encountered during initial testing of the
crosshead expander.
An analytical study of the expander-dynamometer rotating system
indicated that the peak torque (for non-resonant condition) on the drive
shaft was 219 ft/lb. When assuming that the stiffness or torsional
spring constant of the "Spicer" shaft was approximately that of a rigid
tube, the calculated natural frequency was above 2800 RPM. However,
the exact stiffness of the shaft which included two universal joints
was not known, and apparently the stiffness was much lower than the
value assumed.
166
-------
I R ELECTRIC BOILER
AND SUPERHEATER
PUMP WATER ,
SUPPLY TANK !
Figure 7.1-2. Single Cylinder Expander Test Facility (P72-2-5F)
167
-------
Figure 7.1-3. Single Cylinder Expander Test Facility (72-2-5G)
168
-------
o
-
Figure 7.1-4. Installation of Crosshead Expander (P72-4-4M)
-------
Cooler
Figure 7.1-5. Installation of Crosshead Expander (P72-4-4N).
170
-------
Numerous discussions were held with engineering personnel at the
Dana Corporation, Waukesha Motor Company, T.B.Woods Company, and others
in an attempt to resolve the shaft breakage problem. It was concluded
that the rotating mass system had a critical torsional speed within the
0 - 2500 RPM operating range. As a result, a new shaft was designed
which included provisions for torsional damping and reduction of shock
loading. Also, instrumentation was installed to detect torsional
critical speeds. Calculations indicated the torsional natural frequency
of the "resilient" shaft to be 610 RPM. Newly installed instrumentation
indicated the natural frequency to be in the vicinity of 700 RPM, and
therefore operation between 600 and 800 RPM was avoided. Operation at
500 RPM with high inlet steam pressure resulted in excessive temperature
rise of the shaft. This was due to oscillations caused by cyclic torque
of the expander as it went from power to recompression during each
revolution. Therefore, only test points of low cylinder pressure were
obtained at 500 RPM.
7.2 Expander Instrumentation
Each expander was well instrumented with 33 thermocouples, 4 pres-
sure transducers, 2 vibration pickups, 2 speed pickups, 2 steam flow
sensors, a shaft torque meter, cam push rod force sensor, plus other
miscellaneous sensors.
Measurement of steam pressure in the expander cylinder as a func-
tion of crank angle was a challenging task. It was generally concluded
that the difficulties which were encountered with the cylinder pressure
transducer were related to overtemperature or poor temperature distribu-
tion throughout the transducer.
Three different types of pressure transducers were tried for
cylinder pressure measurement. All three transducers required water
cooling. Two of the transducers were strain gage types, and one utilized
a piezoelectric crystal. All of the transducers had a temperature limit
in the range of 300 - 400°F. If the diaphragm of the transducer
exceeded'its temperature limit, transducer output response became
erratic and sensitivity decreased rapidly. Temperature of the diaphragm
171
-------
for the strain gage transducers was determined from gage resistance
measurements during testing. Calibration curves; were then used to
determine gage sensitivity. Zero shift with temperature change was
significant. Therefore, only peak pressure measurements were con-
sidered to be reliable. To help reduce the diaphragm temperature, a
small amount of argon gas was continuously injected into the vicinity
of the gage diaphragm. The argon, a noncondensable gas, reduced the
condensation rate of steam on the water cooled diaphragm - thus reducing
the temperature of the diaphragm. All cylinder pressure data were
obtained by using this technique. The piezoelectric transducer was
very temperature sensitive and no valid pressure measurements were
obtained with this unit. The best cylinder pressure data was obtained
with a Dynisco Model PT49A transducer. This unit contained water
cooling passages in the diaphragm.
172
-------
8.0 TEST RESULTS
CROSSHEAD PISTON EXPANDER
8.1 Component Performance(s)
The crosshead piston expander was first rotated by motoring with the
dynamometer on March 30, 1972. The expander was first operated on steam
on April 26. Problems arose due to failure of the expander-dynamometer
coupling shaft because of shaft torsional vibratory fatigue. A temporary
fix allowed the first test point data to be taken on April 27-28, May 2-3,
and May 10-11 while awaiting the delivery of a torsionally-damped shaft.
Performance testing of the crosshead expander with graphite rings and
with the inlet valve set to open 10° BTDC was completed on June 8. The
total operating time of the expander on that date was 87.2 hours. An addi-
tional 152.3-hour endurance test at 1000 psia, 1000°F, 1500 RPM was com-
pleted on June 16. The total operating time with the original graphite
rings was 241.9 hours.
Visual examination and dimensional measurements of the expander com-
ponents were made at various times during the 87.2 hours of performance and
152.3 hours of endurance testing of the first engine build-up with carbon
graphite piston rings. Major inspection of component parts occurred after
motoring checkout and after 7.5, 31.3, 52.4, 152.6 and 241.9 hours of test
under steam. Following is a summary of the findings of the post test in-
spections performed on crosshead expanders.
8.1.1 Camshaft/Valve Lifter (Tappet)
The AISI 8620 camshaft exhibited light pitting on the carburized lobe
adjacent to the contact area. The noncritical areas of the chill-cast
Cr-Mo cast iron tappet also showed evidence of pitting which appears to be
caused by a form of pitting corrosion. However, the cam lobe/tappet con-
tact surfaces were highly polished and no evidence of pitting or other
forms of damage were observed.
173
-------
Early'in the test program during the checkout of the expander,
severe damage to the tappet/cam lobe surfaces was observed after only 2
hours of motoring without steam. In this case the materials combination
was a cobalt base cast Stellite tappet against a carburized AISI 8620 cam
lobe. However, inadvertently a cast Stellite 6B alloy tappet with a hard-
ness of Re 38-39 rather than the intended Stellite Star J alloy with a
hardness of Re 60-61 was used. The failed Stellite 6B tappet had the
classical appearance of spalling (pitting), i.e., irregular shaped holes
apparently caused by rupture of the metal below the surface, and scuffing
due to a galling action between the tappet and cam surfaces. Visual
examination of the carburized AISI 8620 steel cam surface revealed a build-
up of metal on the nose and on each side of the lobe of the cam surface.
The metal transfer to the cam surface is believed to be caused by a
galling action between the tappet and cam surface as a result of a com-
bination of a high tendency for adhesion between the two surfaces and a low
bulk fracture strength of the Stellite 6B.
The chill-cast Cr-Mo cast iron tappet with a steam tempered black
oxide (Fe-0,) antichafing coating was selected for the second engine build-
up in preparation for performance testing. The chill-cast tappet was*
selected on the basis that experience has shown that the chill-cast iron
tappet generally is superior to hardenable cast iron against hardened steel
(53)
cam surfaces, The major disadvantage of chill-cast cast iron tappets
is its greater tendency to failure by spalling (pitting) especially in the
presence of certain EP additives in the oil^ » » ' . However, since
the tappet material was changed to the chill-cast Cr-Mo cast iron and an
auxiliary oil pump was installed to insure pre-startup lubrication no
further cam lobe/tappet damage has occurred other than the slight pitting
corrosion of non-critical areas.
8.1.2 Inlet Steam Valve/Housing
The H-ll alloy inlet steam valve stem assembly required a light force
in order to remove it from the cast Ni-Resist Type D-3 alloy housing. The
cold worked (Re 49) Haynes alloy No. 25 upper stem compression seal rings
had worn a shallow groove (^0.002") in the ID surface of the housing; no
significant wear of the housing was observed at the location of the bottom
174
-------
seal rings. Possibly the application of a wear-in coating on the Ni-
Resist housing (or rings) would reduce the observed wear by providing
lubrication until the Ni-Resist surface has a chance to work harden and
achieve a harder and more wear resistant surface. Evidence of some
pitting also was observed on all components of the housing. However, in
spite of the slight pitting and wear in the housing, the steam leakage
\
past the inlet valve stem seal was on the order of only 0.5% of the total
steam flow at steam conditions of 1000°F/1000 psi, Figure 8.1-1.
Other components of the inlet steam valve assembly were in excellent
condition. The nitrided spherical end of the H-ll alloy push rod and the
spherical Stellite 6B seat were unpitted and highly polished. Both the
Stellite 1 and Stellite 6 alloy valve seat and valve facing respectively
were in good condition and were reuseable. Although the low alloy steel
inlet valve spring relaxed approximately 1/8" in the first 52.4 hours of
testing, no further problems were encountered with spring relaxation
after the addition of oil jets to the housing to improve cooling of the
spring.
LPA 101 plasma spray-coated 17-7 PH alloy compression seal rings were
incorporated in the inlet steam valve assembly for the first 52.4 hours of
engine testing. Difficulties were encountered with the LPA 101 coating
during fabrication of the rings and during the period of testing due to
frequent chipping of the coating. An increasing rate of steam leakage
through the inlet valve stem (Figure 8.1-1) was observed over the initial
52.4 hours of testing as a result of continued chipping of the coating.
Upon replacing the LPA 101 coated 17-7 PH rings with the cold worked
Haynes Alloy No. 25 rings, no further increase in the leakage rate occurred
and after a period of time the leakage rate decreased to the ^0.5% value.
The LPA type material (laves phase in cobalt matrix) may have potential
as a ring material but not as a coating. Possibly a cast ring of solid
LPA alloy may warrant evaluation.
8.1.3 Recompression Valve ,
All the materials of construction of the recompression valve assembly
are identical to the materials used for the fabrication of the inlet steam
valve with the exception of the spring. After (100 hours) of operation
175
-------
2400 -
2000 -
1600 -
ti
X^
- §§§
(400-700 psi
,\^j < 700-900 °F
v^v^ (500-2000 RPM
\(400 psi
WOO°F
(500-20QO RPM ,
0 25 50
1
1
A
\
v '
Haynes-25
1000 psi
1000 °F
1000-2000 RPM
^§$
^"X-^^V. ^X^^^V^
1
75
I i i 1 1 '
Ring
/
l\s\ss^s
{400-700 psi
700- 1000 °F
500-2000 RPM
i I l i i i
100 125 150 175 200 225
.
-
_
— -
25
1.2
1.0
0.8
0.6
0.4
0.2
0
0)
•U
Expander* Operating Time, Hours
Figure 8.1-1 Inlet Valve Stem Steam Leakage Rate
-------
the low alloy steel recompression valve spring relaxed nearly 0.100" and
was replaced with a spring fabricated from Inconel X-750. Although even
the Inconel X-750 recompression valve spring had relaxed 0.040" at the
end of the final 141.9 hours of test, it still was reuseable.
All other components of the recompression valve were in excellent
condition.
8.1.4 Piston Rings/Cylinder Liner
At the end of the 241.9 hours of testing, the Sb impregnated carbon-
graphite power piston rings (CC-5A material) exhibited severe wear and
damage. Each ring segment was minus at least one lap joint and approxi-
mately one half of each ring segment located in the bottom ring groove
was missing. The pressure balancing circumferential grooves had all but
disappeared from the rubbing surfaces of those segments in the upper ring
groove. From probe measurements of the ring groove depth with the rings
installed in the expander and actual micrometer measurements at the time the
expander was disassembled after 52.4 hours and at the completion of the test-
ing of 241.9 hours, it appears that the wear of the carbon-graphite rings
is linear out to about 153 hours. Beyond 153 hours, the wear rate of the
top ring increased and the wear rate of the middle and lower rings appeared
to decrease. The wear data are shown in Table 8.1-1 and Figure 8.1-2.
Photographs of the rings after 52.4 and 241.9 hours on test are presented
in Figures 8.1-3 and 8.1-4.
The Inconel X-750 ring spring of the power piston bottom ring groove
exhibited substantial wear at the end joint. This ring spring has shifted
radially outward due to the missing graphite segments and had scored the
I.D. of the Type 440C SS cylinder liner, particularly near the exhaust
ports. Visual examination of the cylinder liner after 152.6 hours of test-
ing revealed no scratches or damage of any kind.
8.1.5 Power Piston Head
The Cr-Mo-V alloy power piston exhibited little, if any, wear except
for light scratches in the lower ring groove surfaces caused by the Inconel
X-750 ring spring.
177
-------
Table 8.1-1
00
PISTON RING WEAR
(Carbon-Graphite Grade CC5A)
Engine Hours
On Steam
0
7.5
31.3
52.4
152.6
241.9
• Ring Groove »
Depth, In.w
0.071
-------
0.100
0.080
0.060
0.040
0.020
87.2 Hours
Performance
Tests
400-1000 psi
700°-1000°F
500-2000 RPM
152.3 Hours
Endurance Tests
1000°F, 1000 psi, 1500 RPM
6 Top
X
__ () Middle
— — Q Lower
Code
Micrometer
Probe
O'
100
200
243
TIME, HOURS
Figure 8.1-2 Piston Ring Wear at Middle of Segments
(Carbon Graphite Grade CC5A).
300
-------
SEGMENTED PRESSURE BALANCED
GRAPHITE RING
Figure 8.1-3. Crosshead Expander Carbon-Graphite Rings (CC-5A)
after 52.4 hours on test (P72-5-4A)
180
-------
. .,..-.,
1
Figure 8.1-4.
Crosshead Expander Carbon-Graphite Rings (CC-5A)
after 241.9 hours on Test (P72-6-3L)
181
-------
8.1.6 Power Piston Rod Seal
The power piston push rod chevron seal was in excellent condition.
The seal was fabricated from 15% bronze + 5% MoS- filled PTFE and in-
spection of the seal after completion of the 241.9 hours of test revealed
no measurable wear. The mating surface was a hard chromium plated H-ll
steel.
During the initial stages of checkout testing without steam, traces of
oil entered the expander steam chamber through piston rod seals. It was
concluded that the oil was bypassing the seal by a hydrodynamic pumping
action. By modifying the mounting of the chevron seal to take advantage
of the hydrodynamic action of the seal, the problem was solved. The chevron
seal was installed with a floating housing. An additional wave spring was
added to the seal assembly and all wave springs were placed on the oil side
of the seal. With this modification to the seal, oil leakage into the steam
expander condensate was maintained at less than 4 parts per million as shown
by Figure 8.1-5.
Leakage of water past the piston rod seal into the crankcase also was
monitored at intervals during the performance and endurance testing. Total
water leakage was determined by 1) analyzing the % water content in the oil
(emulsion), 2) measuring the quantity of water (demulsified, condensed
steam) in ml that accumulated in the lower section of the crankcase, and 3)
measuring the quantity of water that' boiled off through the crankcase vent.
These data are given in Table 8.1-II. Examination of the data show that
relatively little water has leaked into the crankcase. The water content
of the oil after the final 89.3 hours endurance test analyzed only 0.16%;
however, even the VL.0% water content in the oil after the first 63.5 hours
of endurance testing would not be expected to be detrimental to the lubri-<
;
eating properties of the oil.
8.1.7 Crosshead Piston
Both the aluminum alloy crosshead piston and the mating nitrided
Type 304SS lower cylinder wall exhibited little, if any, wear. Only very
shallow pitting was observed on the lower position of the nitrided ID sur-
face of the cylinder.
182
-------
12
oo
C
5
4J
g
O
I
•H
O
10
0
Seal Failure
700-1000°F
400-1000 psi
500-2000 RPM
1000°F, 1000 psi
1500 RPM
25
Modified
Seal
,Instailed
50
75 100 125 150
Expander Running Time, Hours
175
200
225
250
Figure 8.1-5 Oil Concentration in Steam Condensate .
-------
Table 8.1-II
Water Leakage into Crankcase
Total Water Drained from Vapor
Engine Hours . Water Content Crankcase before Boil-Off
on Steam Crankcase Oil, Analytical Sample Crankcase Vent
% ml ml/hr
0 0.0008 (8ppm)
12.9 0.028 - nil
31.3(a) 0.60 5 4.5
152.6(b) 1.00 (1 ml) 113 6.3
241.9(c) 0.16 50 1.1
Piston rod seal failure; crankcase oil drained 2050g oil, 12.3g water.
63.5 hours on endurance test, all water drained prior to start of
63.5-hr run.
(c)
89.3 hours on endurance run, all water drained prior to start of 89.3-hr run.
184
-------
8.1.8 Wrist Pin Bushing
The diametral clearance of the wrist pin/bushing assembly increased
approximately 0.004". The high load areas of the wrist pin exhibited
burnished markings but to no discernable depth.
8.1.9 Other Components
All other components of the expander were in excellent condition and
reuseable for further testing.
8.1.10 Crankcase Lubricant
A sample of the XRN 1301C oil which had accumulated 187.7 hours of
operation was drained from the crankcase after the completion of the 152.3-
hour endurance test and sent to Mobil Research and Development Corporation
for analysis. The results of the analyses are presented in Table 8.1-III.
The data show relatively little change of the used oil from new oil and no
serious oxidation or degradation. The slight viscosity increase is possibly
due to the significantly higher water content of the used oil. The increase
in neutralization number reflects some slight oil oxidation that was con-
firmed by differential infrared analysis. The ash content is very low and
shows expected traces of iron wear metal and possibly dirt as indicated by
the silicon content. Overall, the XRN-1301C oil performed well and pro-
vided satisfactory lubrication with little evidence of degradation over
187.7 hours of operation.
Buildup of the crosshead expander with Cr_C9 coated Inconel-X piston
J *•
rings was completed on June 23. The expander was run for about one hour at
400 psig and 700° Finlet steam conditions on June 23 for preliminary check-
out. No difficulty was encountered during the checkout test, and visual
inspection of the piston rings by observation through the exhaust port re-
vealed no ring or cylinder liner damage.
Performance testing of the crosshead expander with the Inconel-X rings
began ort June 26, with initial^performance being satisfactory. However,
after approximately six hours, expander performance began to decay, as in-
dicated by poor cylinder recompression and increased steam consumption.
The expander was shutdown and the piston rings were inspected by viewing
185
-------
Table 8.1-III
Change in Properties of XRN-1301C Oil
After 187.7 Hours of Engine Operation
Property New Used
Viscosity @ 100 cs 75.45 80.20
@ 210 cs 11.07 11.60
Viscosity Index 147 147
Neutralization Number 0.06 0.5
Water Content, ppm 38 431
Ash, wt % 0.001 0.006
Metal Contents, ppm by Emission Spectrograph
Al -(a) --"- 0
Cr 0
Cu 0
Fe - 78(b)
Mg 0
Ba 0
Si 4
Differential infrared analysis of new oil vs. used oil shows loss of about
10 - 20% of antioxidant and slight oil oxidation.
Insufficient ash for determination
Computer extrapolation value might be slightly high
186
-------
through the exhaust port. Extensive ring wear and cylinder wall rough-
ness were observed. Upon disassembly, the Cr_C2 coated Inconel-X rings
were found to be worn approximately 60 mils which included the 3 to 6 mil
Cr-Cj outer coating. The hardened Type 440-C liner contained a number of
axial grooves in the order of 1 to 3 mils deep, and there were also axial
streaks of metal buildup on the cylinder wall 2 to 3 mils thick. The
stepcut tabs on all rings were broken off, and 20 to 30% of each ring was
missing as shown by Figure 8.1-6. All other expander components were
virtually undamaged.
Oil leakage into the steam condensate was high initially, i.e., 55 ppm
after 3.2 hours; however the oil concentration rapidly decayed to 25.2 ppm
after 4.2 hours and 4.0 ppm after 5.4 hours. It is assumed that some oil
had been trapped in the system prior to engine start-up which resulted in
the high initial oil content in the condensate. The piston rod seal was
functioning as expected and further testing would have resulted in oil
concentration of <4.0 ppm.
8.2 Thermodynamic Performance
8.2.1 Thermodynamic Performance - Graphite CC-5A Piston Rings
Test results for the crosshead expander are presented in Table 8.2-1.
Figure 8.2-1 shows the increase in brake horsepower with speed and
with increasing capacity of the steam to perform work as pressure increases.
More efficient breathing and.lower friction power losses at low speeds
cause the deviations from straight lines. The effects of greater blowby
at lower speeds and higher pressures are also present. A reference line
of theoretical indicated power at 1000 psia and 1000°F is shown (assumes no
blowby). The brake data compare favorably with the prediction, since the
brake power would be expected to be less than the indicated power due to
blowby and friction losses.
The brake mean effective pressure is shown in Figure 8.2-2 as a
function- of speed and inlet steam conditions. It reflects the same effects
which the power curves show. The theoretical mean effective pressure at
1000 psia and 1000°F is shown as a dashed line. The difference between the
187
-------
Figure 8.1-6.
Crosshead Expander Cr-$C2 Coated Inconel-X Rings
After Approximately 10 Hours of Testing (P72-6-3N)
188
-------
Table 8,2-1
Crosshead Expander Test Data
May 10 - June 7. 1972
H
00
Test
Point
101E
101G
10 2E
102G
102H
103H
104H
105E'
105H
106E
106H
10 7H
108H
11 7H
119H
122H
123H
124H
127F
12 7H
128H
129G
129H
129HA
Brake
Torque
Ib-in
636
509
530
551
509
403
288
636
594
1356
1272
1060
933
594
254
1102
933
721
1590
1823
1442
1102
1208
1187
BMEP
psi
115
92
95
99
92
73
52
115
107
244
229 -
191
168
107
46
199
168
130
286
328
260
199
218
214
Speed
rpm
475
500
850
1000
1000
1500
1995
546
500
590
500
1008
1495
1001
1992
999
1500
1995
1000
1001
1500
2000
2000
2000
Brake
H.P.
4.79
4.04
7.15
9.15
8.08
9.59
9.13
5.51
4.71
12.69
10.09
16.95
22.12
9.43
8.04
17.47
22.20
22.82
25.23
28.95
34.31
34.97
38.35
37.67
Cond.
Press.
psia
21.0
19.8
22.0
20.4
19.5
19.5
19.5
22.0
19.5
21.6
19.8
19.8
20.0
19.7
19.6
19.6
19.9
20.0
23.0
19.5
19.5
23.7
20.0
19.8
Inlet
Press.
psig
380
403
400
435
408
405
412
400
405
700
700
700
698
440
400
705
700
690
912
1020
990
950
990
1000
Inlet
Temp.
°F
604
700
640
700
694
700
706
634
694
682
712
688
688
1012
988
988
988
1000
1020
1010
1008
940
1035
1040
Steam
Rate
Ib/hr
92
110
102
-
150
170
167
74
91
176
175
280
340
105
127
187
235
282
415
340
402
415
430
432
Specific Operating
Rate Time
Ib/hp-hr hr
19.3
27.2
14.3
-
18.6
17.7
18.3
13.3
19.4
13.9
17.3
16.5
15.4
11.1
15.9
10.7
10.6
12.4
16.4
11.7
11.7
11.9
11.2
11.5 87.2
-------
60
50
M
V
g
(X
QJ
to
0)
20
10
70° valve event
10° valve lead
Graphite CC-5A Piston Rings
in
in
0
500
1000 psia, 1000°F
Theoretical IHP
1015, 1025°F~
BHP 945, 980°FA
715, 990°F O
715, 690°F D
BHP
435, 1000 °FO
420, 675°F O
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 8.2-1. Crosshead Expander Performance
190
-------
400
co
0.
0)
t-1
en
co
350
300
250
200
o
tu
14-1
w 150
100
50
70° valve event
10° valve lead
Graphite CC-5A Piston Rings
500
in
T.
in
1000 psia, 1000°F
Calculated IMEP
Measured IMEP
1015, 1025°F " '
1015, 1025°F
100
BMEP
945, 980°F
A
BMEP
715, 990°F
715, 690°FQ_
BMEP
435, 1000°FO-
420, 675°F O
175
150
125
75
50
25
1000 1500 2000
Rotational Speed - RPM
2500
3000
S
cr
Figure 8.2-2. Crosshead Expander Performance
191
-------
measured and theoretical IMEP is chargeable to the theoretical assumption
of no blowby. The difference amounts to about 6 hp or 13 percent.
The difference between the IMEP and BMEP curves at 1015 psia and
1025°F represent frictional power losses of 3.9 hp at 1500 rpm and 4.7 hp
at 2000 rpm. These losses are almost exactly what is expected from two-
stroke diesel engines operating at the sam
mechanical efficiencies are 0.90 and 0.89.
stroke diesel engines operating at the same IMEP . The corresponding
The specific steam consumption is shown in Figure 8.2-3 as a function
of speed and inlet conditions. The influence of speed is strongest at
500 rpm,but the SSC is practically invariant at the higher speeds. A
strong influence of the steam enthalpy is also seen. The influence of
other parameters upon the brake SSC curves is not apparent. Friction and
blowby would contribute to the difference between the 1005 psia, 1005°F
curve and the theoretical indicated SSC.
The engine efficiency as a function of speed and inlet conditions is
shown in Figure 8.2-4. The definition of engine efficiency as shown by
the figure is:
o Q°Ut = BHP
ne Q±n wAh
The theoretical indicated efficiency is shown for reference.
One cylinder pressure-volume plot is shown by Figure 8.2-5. Since
this curve is drawn from a small oscilloscope photograph, such as shown
by Figure 8.2-6, absolute accuracy is not expected, but the correlation
between predicted and actual pressures appear good. The recompression
curves match very closely until shortly before TDC. The match is also
3
quite good from 7 in (approximate point of inlet valve closure) through
expansion. The variation through this interval is probably primarily due
to blowby.
Following completion of approximately 87 hours of performance testing,
the crosshead expander was held at 1500 rpm with 1000°F and 1000 psia
steam inlet conditions for an additional 150 hours. Figure 8.2-7
shows the decay of shaft horsepower, and Figure 8.2-8 shows a steady
192
-------
t
o
CO
I
u
•H
14-,
•H
O
0)
o.
en
30
25.
20
15
10
0
0
70° valve event
10° valve lead
Graphite CC-5A Piston Rings
p.
in
T.
in -
D
u
BSSC 420 psia, 674°]
-BSSC 715, 690°F
•BSSC 1005, 1005 °F
1000, 1000°F
Theoretical ISSC
500
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 8.2-3. Crosshead Expander Performance
193
-------
1.0
0.9
0.8
u
I 0-7
•H
(4-1
VU
W
01
c
0.6
0.5
0.4
0.3
0
70° valve event
10° valve lead
Graphite CC-5A Piston Rings
500
1000 1500 2000
Rotational Speed - RPM
Theoretical
Indicated
Efficiency
P. = 1000 psia
,
in
= 1000° F
Brake
Efficiency
.,
In
,
in
1015 psi«
= 1025°F "
2500
3000
Figure 8.2-4. Crosshead Expander Performance
194
-------
VO
in
1200
1000 -
2000 BPM
70° valve event
10° valve lead
Graphite CC-5A Piston Rings
CO
p.
I
a
(0
CO
5
iH
£
800
600 -
1000 psia, 1000°F
Measured
1015 psia, 1040°F
400 -
200 -
'Dynisco Model PT 49A
pressure transducer
15 20
Cylinder Volume
Figure 8.2-5. Crosshead Expander Typical P-V Diagram
-------
c
-
INL! ' VALVE OPENING
EXHAUST PORT UNCOVERED
INLET VALVE PUSH ROD FORCE
EXPANDER SPEED - 2000 RPM
INLET PRESS. - 1000 PSIG
INLET TEMP. - 1000°F
VALVE EVENT - 70°
VALVE OPEN5 - 10° BTDC
FORCE - 215 LB/CM
CYLINDER PRESSURE
EXPANDER SPEED - 2000 RPM
INLET PRESS. - 1000 PSIG
INLET TEMP. - 1000°F
PRESSURE - 155 PSI/CM
CRANK ANGLE 360C
Figure 8.2-6. Oscilloscope Photos
-------
M
CO
M
O
EC
CO
60
50
40
30
20
10
0
Grade CC-5A Graphite Rings
50
100
150
200
250
Accumulated Test Hours
150 Hour Endurance Test
Figure 8.2-7. Single Cylinder Crosshead Expander - 1500 RPM
197
-------
,n
iH
I
§
•H
CO
c
o
cj
a)
-p
CO
u
•H
u
(U
o.
co
30
25
20
15
0
0
Grade CC-5A Graphite Rings
50
100
150
200
250
Accumulated Test Hours
150 Hour Endurance Test
Figure 8.2-8. Single Cylinder Crosshead Expander - 1500 RPM
198
-------
increase in SSC with time. Both figures reflect an increase in steam
blowby as ring wear progressed. Figure 8.2-9 indicates a decline of
engine efficiency, but there is a strong indication that the decline is
asymptotic. This would be expected since the expander would continue to
produce power with no rings.
All test data shown by Figures 8.2-1 through 8.2-9 were taken from
the single cylinder crosshead expander which contained antimony impregnated
graphite (grade CC-5A) piston rings and a Type 440C stainless steel
cylinder liner. Figure 8.2-10 shows the measured temperature distribution
for the crosshead expander.
8.2.2 Thermodynamic Performance - Cr-C» Coated Inconel X-750
Piston Rings f
The most meaningful comparison of the performance of the crosshead
expander with graphite and with Cr,C~ coated Inconel X-750 power piston
rings is the specific steam consumption (SSC) of each assembly. Figure 8.2-11
shows that the Cr.jC2 coated rings operating at an average inlet condition
of 396 psia, 687°F had a SSC which was 17 percent higher at 500 rpm to
4 percent higher at 2000 rpm than that of the graphite rings operating at
an average inlet condition of 420 psia, 675°F. The Cr~C2 coated rings
operating at 695 psia, 702°F had a SSC which was approximately 16 percent
higher than that of the graphite rings operating at 715 psia, 690°F
from 1000 - 2000 rpm.
The greater SSC of the Cr-C2 coated rings could be caused by increased
steam blowby past the rings and/or by greater friction power losses. The
brake horsepower (BHP) and brake mean effective pressure (BMEP) were
greater for the Cr,^ coated rings than they were for the graphite rings,
as shown by Figures 8.2-12 and 8.2-13.
The greater SSC of the C^oCo coated rings despite higher BHP and
BMEP appears to be due to higher ring friction as indicated by Figures
8.2-14 and 8.2-15 which compare pressure-volume for similar test conditions
for the two ring types. These .figures fail to show any strong evidence of
greater blowby for the Cr-C^ coated rings since the expansion and recompression
portions of the diagrams are very similar to those for the graphite rings.
199
-------
1.0
0.9
Grade CC-5A Graphite Rings
0.8
§ 0.7
W
0)
•H 0.6
oo
0.5
0.4
0.3
0
50 100 150
Accumulated Test Hours
150 Hour Endurance Test
200
250
Figure 8.2-9. Single Cylinder Crosshead Expander - 1500 RPM
200
-------
-Ii50
STEAM INLET TEMP - 10**0°F
STEAM INLET PRESS - 1000 PSI
EXPANDER SPEED - 2000 RPM
STEAM EXHAUST-
RECQMPHESSION
VALVE VENT LINE-1
STEAM INLET
RECOMPRESSION
RELIEF VALVE
STEAM INLET VALVE
CYLINDER LINER WITH
EXHAUST PORTS
OIL JET
ROD SEAL-OIL
CROSSHEA
WRIST PIN AND
BEARING INSERT ,
L-—CRANKCASE
OIL TEMP - 238°F
CAM-TAPPET
OIL JET
CRANKSHAFT
1625STROKE
Figure 8.2-10. Measured Temperature Distribution of Crosshead Expander
201
-------
M
JS
PQ
.0
i-t
4J
(X
(0
g
u
0)
4J
w
o
0)
30
25
20
15
10
0
0
Cr3C2/Inconel X Rings
396 psia, 687°F =•_
695 psia, 702°F =f
Graphite Rings
420, 675°F
715, 690°F
500
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 8.2-ll.< Comparison of Crosshead Expander Specific Steam Consumption
with Graphite and with Cr.C- Coated Inconel-X Piston Rings
202
-------
60
0)
I
fr
(0
n
o
PC
I
pq
50
40
30
20
10
Cr3C2/Inconel X R
— — 695 psia, 702
Graphite Rings
715, 690°F
Cr_C9/Inconel X
.^^v -^^"" J *•
405, 1004°F =0
396, 687°F =<
Graphite Rings
435, 1000°F
420, 675°F
ngs
F
lings
0
500
1000
1500
2000
2500
3000
Rotational Speed - RPM
Figure 8.2-12. Comparison of Crosshead Expander Power Production with
Graphite and with Cr3C_ Coated Inconel-X Piston Rings
203
-------
400
•H
CO
O,
350
300
0)
S 25°
CO
2
£ 200
Q)
MH
H
0)
100
50
0
0
Cr C /In cone 1 X Rings
695 psia, 702°F
Graphite Rings
715, 690°F
Cr-C2/Inconel X Rings
=~ __ 405, 1004°F =0
0 396, 687°F =•
Graphite Rings
435, 1000 °F
420, 675°F
500
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 8.2-13. Comparison of Crosshead Expander BMEP with Graphite and
with Cr_C9 Coated Inconel-X Piston Rings
J £
204
-------
600
0
•Cr-CL Rings, P, - 690 psia, T. = 704°F
o £. in in
\ (TP 108, 6/26/72, SSC = 18.0 Ib/BHP-hr, BMEP
\
191 psi, nen = 0.466)
-Graphite Rings, P = 698 psia, T = 688°F
(TP 108, 6/7/72, SSC = 15.A Ib/BHP-hr,
BMEP = 168 psi, n
eng
0.544)
Cylinder Volume - In'
Figure 8.2-14. Comparison of Crosshe-ad Expander P-V Diagrams with Graphite and with
Cr C_ Coated Piston Rings at 1500 RPM and Similar Steam Inlet Conditions
-------
600
500
3
D.
I
400
p
in
•Cr3C2 Rings, P±n =400 psia, Tin - 704°F
(TP 101A, 6/26/72, SSC = 21.8 Ib/BHP-hr, BMEP = 115 psi, n
0.431)
-Graphite Rings, P = 405 psia, T = 712°F
(TP 105, 6/7/72, SSC = 19.4 Ib/BHP-hr,
BMEP = 107 psi, n
eng
= 0.480)
0
Cylinder Volume - IJT
Figure 8.2-15. Comparison of Crosshead Expander P-V Diagrams with Graphite and with
Cr_C0 Coated Piston Rings at 500 RPM and Similar Inlet Steam Conditions
3 f.
-------
The areas within the P-V diagrams are very nearly the same for each type
of ring - thus the expander indicated horsepower would be very similar.
Also measured engine efficiencies (BHP/wAh. . ) are lower for the
0 isentropic
^r3^2 coated rings. It is therefore concluded that the principal cause
of the increased SSC of the crosshead expander with Cr,C_ coated rings
was due to higher ring friction.
207
-------
9.0 TEST RESULTS
TRUNK PISTON EXPANDER
9.1 Component Performance '
The trunk piston expander was assembled with Cr_C. coated Inconel
X-750 piston rings and with a carbon-graphite (grade P5NR) piston rider
ring. Oil exclusion rings in the aluminum piston skirt were made of
Type K6E and Type K (cast iron).
A 1.5 hour motoring checkout test was conducted on July 6. The
speed was taken to 2000 RPM for approximately 15 minutes. The expander
performance during checkout was quite satisfactory with the exception of
the presence of more oil than anticipated in the cylinder. However, a
condenser pressure of 20 psia during steam testing was expected to cor-
rect this.
V
Performance testing began July 7. Seven test points (6.75 hrs.
testing time) were taken before the expander was shut down for inspection.
Borescope inspection of the piston and cylinder liner revealed minor scoring
of the liner but the compression rings appeared to be in good condition.
The oil exclusion rings and aluminum piston skirt were not visible for
inspection. A small amount of water (< 50 ml) was drained from the crank-
case during this shutdown. Valve stem leakage during the 6.75 hour test
period was approximately 940 ml/hr or approximately 0.5% of total steam
flow. An analysis of the steam condensate from the condenser for oil
after 3.4 hours of testing indicated an oil concentration of 192 ppm.
The high oil concentration in the condensate probably due to residual oil
being in the system or a result of the high speed motoring checkout test.
Because of the short duration of the test it was not possible to compare
the relative oil leakage between the crosshead piston and the trunk piston.
The expander was restarted but 2.5 hours later the recorded speed
208
-------
trace indicated a drop in speed from 2000 RPM to 0 KPM in approximately
4 seconds. An attempt to rotate the expander was unsuccessful. The ex-
pander was removed from the test facility and dismantled.
Examination of the expander during and after disassembly revealed
the following:
1. The cylinder liner was scored in the area contacted by com-
pression rings as shown by Figure 9.1-1. Oil was present in
the piston ring grooves and on the cylinder wall at shutdown.
2. The aluminum piston skirt was badly worn and scored as shown
in Figure 9.1-2. It appeared that the piston skirt to liner
clearance was insufficient at sometime during testing. Review
of temperature recorder charts revealed that the crankcase oil
temperature inadvertently exceeded the maximum allowable by
about 50°F. The high oil temperature resulted in excessive
thermal expansion of the aluminum skirt - thus causing inter-
ference with the cylinder liner.
3. The rear main crankshaft bearing was seized to the crankshaft.
Particles of aluminum were present in the main bearing oil
ports.
4. The oil filter appeared completely plugged with a dark slurry
of what appeared to be carburized oil. . Filter plugging results
in the opening a filter bypass if the filter pressure drop
exceeds 12 psi. The bypass could allow oil with aluminum chips
from the piston skirt to enter the main crankcase oil gallies
and reach the rear main bearing - thus restricting oil flow to
the bearing. No damage to other bearings, gears, etc. was
found. Emission spectrographic chemical analysis of the dark
deposit that had plugged the filter shows high concentrations
of iron, copper lead, zinc and chromium, plus trace amounts of
tin, aluminum and nickel that would be expected to come from
wear of engine components. A high concentration of sodium
indicates the possibility of contamination with the treated
water. This combination of oil, water, metal fines, plus the
209
-------
Figure 9.1-1. Trunk Piston Liner Following 9.2 Hours Test (P72-6-3Z)
210
-------
Figure 9.1-2. Trunk Piston Skirt Following 9.2 Hours Test (P72-6-3W)
211
-------
Figure 9.1-3. Cam Lobe From Trunk Piston Expander (P72-6-3U)
212
-------
presence of sodium ions at high temperatures, could react to
form the grease-like deposit. Analysis of oil from the crank-
case for water content indicated a concentration of 1.3% in the
first 500 ml sample and 0.13% in the second 500 ml sample.
5. The camshaft lobe, Figure 9.1-3, and cam tappet appeared worn
or burnished over the contact areas.
Just prior to bearing seizure, a large amount of condensate and
oil mixture was observed coming from the crankcase vent. Bearing failure
and excessive cam wear may have resulted from poor lubrication: (1) ex-
cessive oil temperature, (2) too much water in the oil, (3) restricted
oil flow, and (4) a combination of the above three.
9.2 Thermodynamic Performance
The brake mean effective pressure (BMEP) of the trunk expander is
compared on Figure 9.2-1 to those of the crosshead expander with both
carbon-graphite (Grade CC5A) rings and a Type 440C stainless steel liner,
and Koppers K1051 (Cr_C coated Inconel X-750) rings and a Type 440C
stainless steel liner. Three sets of inlet steam conditions are shown.
The first test series (400 psia, 712°F) of the trunk expander shows
BMEP which are initially considerably lower than those of the crosshead
expander, but which later in time are only slightly lower. A strong in-
fluence of "wear-in" time is suggested. Both ring friction and leakage
losses are involved, but insufficient data is available to determine
which is the more influential.
The second test series (417 psia, 998°F) shows BMEP which are some-
what lower than those of the crosshead expander but which decrease with
increasing speed as do the crosshead data. This suggests that "wear-in"
has been completed, but that a higher friction coefficient exists.
The third test series (690 psia, 703°F) shows its first point having
a BMEP very close to that of the crosshead expander with graphite rings,
with its final point very close to that of the crosshead expander with
Cr^C^ coated Inconel rings. Similar conclusions relating to "wear-in"
as a function of time may be drawn from the brake horsepower versus speed
213
-------
curves of Figure 9.2-2 and from the brake specific steam consumption
curves of Figures 9.2-3 through 9.2-5.
The change of brake specific steam consumption (BSSC) with time is
more clearly shown in Figure 9.2-6 for the trunk and crosshead expanders
with Cf_C_ coated rings versus Type 440C liners. In any pressure-tem-
perature sequence of testing, speed was varied from 500 to 2000 BFM.
Therefore, speed effect could distort the performance versus time trend.
However, it appears from Figure 9.2-6 that testing time had a stronger
influence on improving BSSC (up to the time of ring or expander failure)
than operating speed, pressure, or temperature.
214
-------
300
U>
p.
0)
w
(0
CO
0)
1-1
u
0)
w
0)
•a
V)
pq
250
200
150
100
50
0
End Test
Trunk
690 psia, 703°F
I
695 psia, 702°F
Crosshead
715 psia, 690°F Graphite
405 psia, 1004°F
417 psia, 998°F
-Trunk
!C"396_£sia, 687°F
Crosshead
t
psia, 675 F
0
500
1000 1500 2000
Rotational Speed - EPM
2500
3000
Figure 9.2-1 Trunk and Crosshead Expander Brake Mean Effective Pressure
As A Function of Rotational Speed, Inlet Steam Conditions,
and Piston Ring/Cylinder Liner Materials
Expander
Trunk
Crosshead
Crosshead
Piston Ring Material
Cr3c2~Coate<*
Inconel X-750
Carbon-Graphite Grade CC5A
Cylinder Liner Material
Type 440C stainless steel
Type 440C stainless steel
215
-------
J690 psia, 702°F, Trunk I
695 psia, 702 °F
-- Crosshead
psia, 690°F Graphite
LAOS psia, 100A°F_!Cr C
- 396 psia, 687°F_j 3 2
A35 psia,_10000FlCrosshead
A20 psia, 675°F [Graphite
;
I 417 psia,
i
500
1000 1500 2000
Rotational Speed - EPM
2500
3000
Figure 9.2-2
Trunk and Crosshead Expander Brake Horsepower As A Function
of Rotational Speed, Inlet Steam Conditions, and Piston
Ring Material
216
-------
g
o>
g
o
0)
4J
CO
U
0)
P.
CO
30
25
20
10
Trunk
400 psia, 712°F
Crosshead
coated
j
__, 396 psia, 687°F Cr^ coated
— 420 psia, 675°F Graphite
0
500
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 9.2-3
Trunk and Crosshead Expander Brake Specific Steam Consumption
As A Function of Rotational Speed and Piston Ring Material
for Steam Inlet Conditions of Approximately 400 psia, 700°F
217
-------
30
f 25
i
g
20 -
.
§
CO
s 15
0)
4J
CO
o 10
*rt
«M
•d
0)
£
o) 5
1
PQ
i
Trunk J
417 psia,998°F Cr coated
Crosshead
405 psia, 1004°F Cr coated
I
-I
420 psia, 1000°F Graphite
I
500
1000 1500 2000
Rotational Speed - BPM
2500
3000
Figure 9.2-4
Trunk and Crosshead Expander Brake Specific Steam Consumption
As A Function of Rotational Speed and Piston Ring Material
for Steam Inlet Conditions of Approximately 400 psia, 1000°F
218
-------
PQ
I
g
•rl
•U
I
1
o
0)
4J
CO
0
0)
(X
CO
0)
•3
U
pa
30
25
20
15
10
T
End Test
Trunk
690 psia, 703°F Cr3C2 coated
Crgsshead
~ 695 psia, /u2cr CrC coated
715 psia, 690°F Graphite
I
I
500
1000 1500 2000
Rotational Speed - RPM
2500
3000
Figure 9.2-5
Trunk and Crosshead Expander Brake Specific Steam Consumption
As A Function of Rotational Speed and Piston Ring Material
for Steam Inlet Conditions of Approximately 700 psia, 700°F
219
-------
t-l
43
§3
C
o
CO
§
o
0)
4J
CO
O
•H
O
<0
p.
CO
rt
70
72
Test Hours (Crosshead + Graphite)
74 76 78 80 82
60
50
40
30
20
10
0
T
T
MOO psia
700°F
500-2000 RPM
500-2000
700 psia
700°F
- 1000-2000 RPM
^400 psia, 1000°F
1000-2000 RPM
A
Crosshead
•x-700 psia, 700°F
1000-2000 RPM
Trunk Piston
^^*
400 psia, 1000°F
1000-2000 RPM
6 8
Test Hours
10
12
14
Figure 9.2-6
Crosshead and Trunk Expander Brake Specific Steam
Consumption As A Function of Testing Time. Piston
Ring Material was CroC2 Coated Inconel X-750 and
Cylinder Material was Type 440C Stainless Steel
220
-------
REFERENCES
1. Rabinowcz, E., Friction and Wear of Materials, John Wiley & Sons,
Inc., New York, (1965).
2. MacGregor, C.W., et al., Handbook of Analytical Design for Wear,
Plenum Press, New York, (1964).
3. Summers-Smith, D., "Operating Experience with Filled Piston Rings",
(Personal Communication).
4. Brooks, R.D., "Design of Reciprocating Single Cylinder Expanders
for Rankine Cycle Engines", Quarterly Report, Jan. 4, 1972 to
April 4, 1972, EPA Contract No. 68-01-0408.
5. Liston, J.E., Aircraft Engine Design, McGraw-Hill, 1942, p. 199.
6. Marks, L.S., Mechanical Engineers Handbook, 7th edition, McGraw-
Hill Book Co.
7. Roark, R.J., Formulas for Stress and Strain, 3rd ed., McGraw-Hill
Book Co., New York, (1954).
8. Rathbart, H.A., et al., Mechanical Design and Systems Handbook,
McGraw-Hill Co., Inc., New York, (1964), pp. 6-54 and 6-67.
9. Brooks, R.D., "Design of Reciprocating Single Cylinder Expanders
for Rankine Cycle Engines", Quarterly Report, Oct. 4, 1971 to
Jan. 4, 1972, EPA Contract No. 68-01-0408.
10. Taylor, C.F., The Internal Combustion Engine in Theory and Practice,
M.I.T. Press, Cambridge, Mass., (1966), Vol. I, p. 354.
11. Brooks, R.D., "Design of Reciprocating Single Cylinder Expanders
for Rankine Cycle Engines", Monthly Progress Report No. 7, July 4
to Aug. 4, 1972, EPA Contract No. 6^-01-0408.
12. Vickers, P.T., et al., "The Design Features of the GM SE-101-A
Vapor Cycle Powerplant", SAR Paper 700163 (Jan. 1970).
13. Lyman, Taylor, (Ed.) Metals Handbook, Vol. I, Properties and Selec-
tion of Metals, ASM, (1961).
14. Lorentz, R.E. and Harding, W.L., "Selection of Materials for Boilers
and Nuclear Reactors", Metals Progress, (April 1967).
15. Steels for Elevated Temperatures, U.S. Steel Corp., (1965).
16. Uhlig, H.H., (Ed.) Corrosion Handbook, Wiley, (1958), p. 578.
17. Eberle, F., and Kitterman, J.H., "Scale Formation on Superheater
Alloys Exposed to High Temperature Steam", Behavior of Superheater
Alloys in High Temperature, High Pressure Steam, ASME, (1968).
221
-------
18. Engineering Alloys Digest, Cl-28, (April 1960).
19. "Properties of Cast Iron at Elevated Temperatures", ASTM STP 248.
20. Testimony of E. Pritchard, "Automobile Steam Engine and Other External
Combustion Engines", Joint Hearings of Committee on Commerce, U.S.
Senate, May 28, 1968 (USGPO Ser. No. 90-82).
21. Van Brunt, C. and Savage, R.H., "Carbon Brush Contact Films", General
Electric Review 47, p. 16 (July 1944).
22. Savage, R.H., "Physically and Chemically Adsorbed Films in the
Lubrication of Graphite Sliding Contacts", Annals of the New York
Academy of Sciences, Vol. 53, Article 4, (June 1951).
23. Savage, R.H. and Schaefer, D.L., "Vapor Lubrication of Graphite
Sliding Contacts", Journal of Applied Physics, Vol. 27, No. 2,
(Feb. 1956).
24. Campbell, I.E., ed., High Temperature Technology, p. 110, John Wiley
& sons, Inc., New York (1956).
25. McKee, D.W., "Metals Oxides as Catalyst for the Oxidation of Graphite",
Carbon, Vol. 8, Pergamon, (1970).
26. Wilson, D.S., et al., "The Development of Lubricants for High-Speed
Rolling Contact Bearings Operating at 1200°F", TR 60-732, WADD
(Jan. 1961).
27. Bowers, R.C. and Murphy, C.M., "Status of Research on Lubricants,
Friction and Wear", NLR Report 6466 (Jan, 19, 1967), p. 10.
28. Summers-Smith, D., "A Review of the Symposium on TFE Seals in
Reciprocating Compressors", ASME Annual Meeting, New York, (Dec. 1970).
29. Lancaster, J.K., "Solid Lubrication", ASME Conference, Denver, Aug.
1971, Tribology, Vol. 4, No. 4., (Nov. 1971), p. 234.
30. Halliwell, H., et al., "An Application of Self-Lubricated Composite
Materials", presented at the Annual ASLE Meeting, Toronto, Canada,
May 1-4, 1967, preprint 67 AM 8A-5.
31. Buckley, D.H. and Johnson, R.L., "Marked Influence of Crystal Structure
on Friction and Wear Characteristics of Cobalt and Cobalt Base Alloys
in Vacuum to 10~9 mm of Mercury", NASA TN D-2524 (Dec. 1964).
32. Peterson, M.B., Florek, J.J., and Murray, S.F., "Consideration of
Lubricants for Temperatures Above 1000°F", ASLE Trans., Vol. 2, No. 2,
(May 1960), p. 225-234.
33. Mitsubishi Metal Mining Co., Ltd., (Nov. 23, 1971).
34. Buckley, D.H. and Johnson, R.L., "Gallium-Rich Rilms as Boundary
Lubricants in Air and in Vacuum to 10~8 mm Hg", ASLE/ASME Lubrication
Conference, Pittsburgh (Oct. 12-18, 1962).
222
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35. Peterson, M.B., Florek, J.J. and Lee, R.E., "Sliding Characteristics
of Metals at High Temperatures", ASLE Conference, New York (Ovt. 1959).
36. Campbell, M.E. and Van Wyck, J.W., "Development of Design Criteria
for a Dry Film Lubricated Bearing System", ASD-TDR-62-1057, (March
1963).
37. Devine, M.J., Lawson, E.R., and Bower, J.R., Jr., "Anti-Friction
Bearing Design Considerations for Solid Lubrication", ASME Preprint
63-MD-43 (May 1963).
38. McDonnell, R.D., (Ed.), "Proceedings AFML-MRI Conference on Solid
Lubricants", AFML-TR-70-127, (July 1970).
39. Hopkins, Vern, et al., "Development of New and Improved High Tempera-
ture Solid Film Lubricants", ML-TDR-64-37, Part II (April 1965);
Part III (August 1966).
40. Bisson, E.E., "Non-conventional Lubricants", Advance Bearing Technology,
NASA SP-38 (1964), p. 217.
41. Sliney, H.E., Strom, T.N., and Allen, G.P., "Fused Fluoride Coatings
as Solid Lubricants in Liquid Sodium, Hydrogen, Vacuum and Air",
NASA TND-2348 (Aug. 1964).
42. Sliney, H.E., "Self-Lubricating Composites of Porous Nickel and
Nickel-Chromium Alloy Impregnated with Barium Fluoride-Calcium
Fluoride Eutectic", NASA TN-D-3484, (July 1966).
43. Sliney, H.E., "An Investigation of Oxidation-Resistant Solid Lubricant
Materials", NASA TM-X-6785, (Aug. 1971).
44. Bowers, R.C., and Murphy, C.M., "Status of Research and Lubricants,
Friction and Wear", NLR Report 6466, (Jan. 1967), p. 29.
45. Chaseman, M.R., "Solid Lubrication for Aero Propulsion Systems",
AGARD LP-84-71.
>>_
46. Abe, W., et al., "Friction and Wear Characteristics of Solid-Lubricants
Embedded Plain Bearing at High Temperature", Giles Industries Co.,
Ltd., Japan.
47. Campbell, M.E. and Hopkins, V., "Development of Polyimide Bonded
Solid Lubricants", ASLE Conference, Toronto, Preprint No. 67-A-7A-1
(May 1967).
48. Hopkins, Vern, et al., "MLF-5, An Inorganic Solid Lubricant Film",
USAF-SWRI Aerospace Bearing Conference, (May 1964).
49. Sliney, H.E. and Johnson, R.L., "Bonded Lead Monoxide Films as Solid
Lubricants for Temperatures up to 1250°F, NASA RM E57B15 (1957).
50. Fusaro, R.L. and Sliney, H.E., "Graphite Fluoride (CFX) A New Solid
Lubricant", ASLE/ASME Lubrication Conference, Houston, (Oct. 1969).
223
-------
51. Wagner, T.O., "Fuels and Lubricants for Steam Propulsion Systems",
SAE 700662, National Conference, (Aug. 1970).
52. Nixon, J. and Gagabrant, A.R., "Steam Expander Lubrication Program",
Quarterly Report No. 3, CTR No. 68-04-0004, (Sept.-Nov. 1971).
53. Personal Communication, Eaton Manufacturing Co., (April 4, 1972).
54. Jarret, M,P., "Material Considerations for Automotive Camshafts",
SAE PP 710545, (June 7-11, 1971).
55. Ambrose, H.A. and Taylor, J.E., "Wear, Scuffing and Spalling in
Passenger Car Engines", SAE Trans., Vol. 63, (1955).
56. Havely, T.W., Phalen, C.A. and Bunnell, D.6., "Influence of Lubricant
and Material Variables on Cam and Tappet Surface Distress", SAE
Trans., Vol. -63, (1955).
57. Ayres, V., Bidwell, J.B., Pilger, A.C., Jr., and Williams, R.K.,
"Valve Train Wear as Affected by Metallurgy, Driving Conditions
and Lubricants", SAE Trans., Vol. 66, (1958).
58. Brooks, R.D., "Design of Reciprocating Single Cylinder Expanders
for Rankine Cycle Engines", Quarterly Report, April 4, 1972 to
July 4, 1972, EPA Contract No. 68-01-0408.
59. Engineer's Handbook of Piston Rings, Seal Rings, Mechanical Shaft
Seals, p. 10, Koppers Company, Inc., Baltimore, (1967).
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